FIFTH CANADIAN CONFERENCE ON NONDESTRUCTIVE ... - IAEA

q 11 

Atomic Energy 

of Canada Limited 

L'Énergie Atomique 

du Canada, Limitée 

AECL-8707 

FIFTH CANADIAN CONFERENCE ON 

NONDESTRUCTIVE TESTING 

CINQUIÈME CONFÉRENCE CANADIENNE 

SUR DES ESSAIS NON DESTRUCTIFS 

i October / Octobre 28-31,1984 

TORONTO, ONTARIO 

"NDT in Canada - The Next 20 Years' 

20th Anniversary Celebration 

Les essais non destructifs au Canada 

Les 20 années prochaines" 

20ieme Anniversaire 

f-dito' Éditeur . 

C.A Kittmer 

Chalk River Nuclear Laboratories 

Laboratoires nucléaires de Chalk River 

Chalk River

ATOMIC ENERGY OF CANADA LIMITED AECL-8707 

FIFTH CANADIAN CONFERENCE ON 

NONDESTRUCTIVE TESTING 

CINQUIÈME CONFÉRENCE CANADIENNE 

SUR DES ESSAIS NON DESTRUCTIFS 

October 1984 Octobre 

TORONTO, ONTARIO 

PROCEEDINGS / 

NOTES DE LA CONFÉRENCE 

Editor / Editeur 

C.A. KITTMER 

Atomic Energy of Canada Limited 


L'Énergie Atomique du Canada, Limitée 

Laboratoires nucléaires de Chalk River 

Chalk River 

January 1985 Janvier

ATOMIC ENERGY OF CANADA LIMITED 

FIFTH CANADIAN CONFERENCE ON NONDESTRUCTIVE TESTING 

EDITOR 

C.A. KITTMER 

ABSTRACT 

The theme for the Fifth Canadian Conference on Nondestructive Testing was "NDT in Canada - The 

Next 20 years". The three day conference with 42 presentations provided a short overview of NDT 

in Canada, a look at NDT in pipeline, materials, offshore, nuclear and training applications, and a 

glimpse into the next 20 years with recent advances in research and development as related to this 

"hi-tech" field of work. 

The three keynote speakers, Andrew B. Mitchell (Canada), Harold Berger (U.S.A.) and Roy S. Sharpe 

(England), were united in their concern for increased dialogue between research and field 

application. This conference was one small step in the right direction. 


Chalk River, Ontario KOJ 1J0 

1985 January 

AECL-8707

L'ENERGIE ATOMIQUE DU CANADA LIMITÉE 

CINQUIÈME CONFÉRENCE CANADIENNE DES ESSAIS NON DESTRUCTIFS 

ÉDITEUR 

C.A. KITTMER 

RÉSUMÉ 

La cinquième conférence Canadienne des essais non destructifs (END) avait pour thème: "END au 

Canada - Les 20 années prochaines". La conférence, d'une durée de trois jours, a permise une vue 

d'ensemble rapide des END au Canada grâce à 42 présentations. L'utilisation des END dans 

plusieurs domaines; construction de pipelines, examens de matériaux, exploration au large des 

côtes, domaine nucléaire et d'entraînement de même qu'un aperçu des progrès récents en 

recherche et développement de cette haute-technologie furent examinés. 

Les trois orateurs invités; Andrew B. Mitchell (Canada), Harold Berger (U.S.A.) et Roy S. Sharpe 

(Royaume-Uni) furent unanimes dans leurs demandes pour un dialogue accru entre les chercheurs 

et les usagers. Cette conférence représente un premier pas dans cette direction. 

Laboratoires Nucléaires de Chalk River 

Chalk River, Ontario KOJ 1J0 

Janvier 1985 

AECL-8707

CINQUIEME CONFERENCE CANADIENNE SUR DES ESSAIS NON DESTRUCTIFS 

PRÉFACE 

II a été décidé de tenir la cinquième conférence Canadienne sur des essais non destructifs pour 

deux raisons principales. La première raison est qu'il s'est écoulé huit ans depuis la dernière 

conférence Canadiennne. De grands progrès ont été accomplis à la technologie Canadienne depuis 

la dernière conférence mais peu de gens sont au courant. Il y a certes des problèmes de 

communication dans l'industrie Canadienne puisque l'industrie Canadienne est éparpillée sur une 

étendue de 150 km de large par 7500 km de long. Les investissements principaux quitte le pays 

étant donné que les Canadiens n'ont pas connaissance que la technologie des essais non 

destructifs est disponible au Canada. Conséquemment, un des objectifs de la conférence était de 

promouvoir la communication et de rassembler les gens de l'est à l'ouest et également d'établir les 

progrès établis à l'intérieur du Canada d'un océan à l'autre. 

Une seconde raison pour tenir la conférence était qu'elle coincide avec le 20ième anniversaire de la 

Société Canadienne des Essais Non Destructifs et corresponds à 20 ans de progrès et de succès 

dans le domain des essais non destructifs. Depuis la fondation de la société Canadienne en 1964 

(qui était initiallement appelée le Centre Canadien pour la Technologie Non Destructif jusqu'en 

1970), la croissance de la société a été tout à fait remarquable atteignant 1500 membres et même 

plus. Durant cette époque, la société a été active en publiant son propre journal (maintenant 

intitulé le journal CSNDT), en présentant des cours et des séminaires et en participant à de 

multiples conférences et expositions. Cependant, c'était la première fois que la société décidait de 

promouvoir une activité majeure par elle-même. 

Le succès de la conférence indique le travail acharné et la persévérance du committé d'organisation. 

Président: 

John H. Zirnhelt 

Président technique: 

Charles A. Kittmer 

Président de l'exposition: 

Marvin Horenfeldt 

Présidents des sessions techniques: 

Jean F. Bussiere 

Norman G. Harding 

Mirek Macecek 

Lynda B. Manzer 

Secrétaire de la conférence: 

Yvonne A. Carroll 

Bureau Canadien de Soudre 

L'Énergie Atomique du Canada, Limitée 

Du Pont Canada Limited 

Conseil National de Recherches Canada 

la Société Canadienne des Essais Non 

Destructifs 

Techno-Scientific Inc. 

la Fondation de la Société Canadienne 

des Essais Non Destructifs 

Conference and Seminar Management 

La présence des notes de la conférence est un hommage aux personnes impliquées dans les essais 

non destructifs au Canada. Nous tenons encore à exprimer notre gratitude aux auteurs et 

également à ceux qui ont présenté les publications pour leur compréhension et leur collaboration 

tout en respectant les horaires. Nous espérons qu'ils maintenant reconnaissent que l'effort était 

valable et que la conférence a apporté une contribution à la technologie Canadienne des essais 

non destructifs pour maintenant et pour le futur. 

La technologie Canadienne, un chef de têtel Tel était un fait reconnu tout au long de la 

conférence. Voici maintenant les notes de la conférence... à vous d'en juger. 

Charles Kittmer

FIFTH CANADIAN CONFERENCE ON NONDESTRUCTIVE TESTING 

FOREWORD 

It was decided to hold the Fifth Canadian Conference on Nondestructive Testing (NDT) for two 

main seasons. First, it had been eight years since the last Canadian conference. A great deal had 

happened to Canadian NDT technology in that time, but few knew about it. With Canadian industry 

located in a strip of land 150 km wide by 7500 km long, communications is a problem. Major 

funding is leaving this country because Canadians are not aware of what NDT technology is 

available in Canada. Therefore, one of the objectives of the Conference was to provide a forum for 

communication.... to bring east and west together and give people an opportunity to find out what 

is happening in Canada from coast to coast. 

A second reason for holding the Conference was to mark the 20th Anniversary of the Canadian 

Society for Nondestructive Testing (CSNDT). What more fitting way to celebrate twenty years of 

achievement in NDT? Since formal establishment in 1964 (it was initially called the Canadian 

Council for Nondestructive Technology until 1970) the society has grown remarkably to a 

membership of over 1500. In that time, it has been an active society publishing its own newsletter 

Inow the CSNDT Journal), sponsoring education courses and seminars, and participating in 

numerous conferences and exhibitions. However, this was the first time for the society to sponsor 

a major activity on its own. 

The success of the Conference reflects the hard work and perseverance of the working committee: 

Chairman: 

John H. Zirnhelt Canadian Welding Bureau 

Technical Chairman: 

Charles A. Kittmer Atomic Energy of Canada Limited 

Exhibitor Chairman: 

Marvin Horenfeldt Du Pont Canada Limited 

Session Chairmen: 

Jean F. Bussiere National Research Council of Canada 

Norman G. Harding Executive Director, CSNDT 

Mirek Macecek Techno-Scientific Inc. 

Lynda B. Manzer CSNDT Foundation 

Conference Secretariat: 

Yvonne A. Carroll Conference and Seminar Management 

The presence of these proceedings reflects, and is a tribute to the hard work of the many people 

involved in NDT in Canada. We again express our appreciation to our keynote speakers and 

authors who responded with understanding and cooperation as they were exposed to our very 

tight time schedule. We hope in retrospect they feel the effort worthwhile, and that the Conference 

has made a contribution to awareness of NDT in Canada both now and for the future. 

Canadian technology, second to none! This was a stated theme throughout the Conference. Here 

are the proceedings... you be the judge. 

Charles Kittmer

iii 

ACKNOWLEDGEMENT 

A special note of thanks to Donna McCarthy of the Nondestructive Testing Development Branch, Chalk 

River Nuclear Laboratories, for her long hours spent in bringing these proceedings to print.

FOREWORD 

TABLE OF CONTENTS 

DAY 1 

KEYNOTE ADDRESS: Five Years Experience with Eddy Current Testing of NBEPC 1 

Generating Stations 

- A.B. Mitchell, H.M. O'Connor, R. Pickles 

Canadian Forces Non-Destructive Program Origins 11 

- J.S. Cavers 

Reducing Unwanted Effects -- Recent Developments in Eddy Current Equipment 16 

t-eatures 

- A.G. Julier 

Nondestructive Examination Plays Key Role in Reactor Rehabilitation Program 27 

- CA. Wallis 

Evaluation of Ultrasonic Methods for the Detection and Sizing of Real Defects in the 40 

Base of Rails 

- L. Piche, J.F. Bussiere, J.P. Monchalin 

Optical Detection of Ultrasound at Distance 51 

- J.P. Monchalin 

Automated Ultrasonic Testing System for the Characterization of Defects in Weldments 66 

- M. Macecek, K. Luscott, J. Wells, D.K. Mak 

The Influence of Stress on the Inspection of Steel with Particular Reference to Gas 79 

Pipelines 

- D.L. Atherton, D.C. Jiles, C. Welbourn 

Defect Characterization and Sizing in Pipeline Weldments by Analysis of Ultrasonic 97 

Echo Features 

- J.P. Monchalin, J.F. Bussiere, R.W.Y. Chan, D. Sharp, D.R. Hay 

Capabilities/Limitations of Focused Ultrasonic Beams for Sizing Pipeline Girth 104 

Weld Defects 

- CA. Kittmer, J.B. Halle«, V.N. Sycko 

Flaw Characterization Using the Time-of-Flight Method and Ultrasonic Frequency 115 

Analysis 

- D.K. Mak 

The Introduction of Real-Time Radiography for the Inspection of Butt Welds in 130 

Offshore Pipelines 

- F.E. Reynolds Jr. (presented by L. Rathwell) 

Continuous Acoustic Emission Monitoring 145 

- J.S. Mitchell, M.N. Bassim

DAY 2 

KEYNOTE ADDRESS: Advanced Radiography for Transportation and Energy Systems 152 

- H. Berger 

NDT - Radiographie Processing Quality Control, Making the Transition to 163 

Automatic Processing in Radiographie NDT, Sensitometric Values for Industrial 

Radiography 

- W.E.J. McKinney 

Recent Microfocus X-Ray Imaging Applications (abstract only) 176 

- R.S. Peugeot 

Wet Channel Inspection Sysiems for CANDU Nuclear Reactors....CIGAR and 178 

CIGARette 

- M.D.C. Moles, M.P. Dolbey, K.S. Mahil 

An Advanced Heat Exchanger Eddy Current Inspection System 194 

- M. DeVerno, H. Ghent, H. Licht 

Wet Channel Measurement of Pressure Tube to Calandria Tube Spacing in CANDU 207 

Reactors 

- J.H. Sedo 

Checking for Cracks in Gas Turbine Rotor Discs 222 

- J. van den Andel, A.B. Nieberg 

Eddy Current Inspection of Mildly Ferromagnetic Tubing 235 

- W.R. Mayo, J.R. Carter 

On the Relation Between Ultrasonic Attenuation and Fracture Toughness in Type 250 

403 Stainless Steel 

- F. Nadeau, J.F. Bussiere, G. Van Drunen 

Acoustic Emission Testing of Man-Lift Devices 262 

- J.A. Baron 

Acoustic Sorting of Grinder Balls 276 

- F. Nadeau, J.F. Bussiere 

The Benefits of NDT Training for Canadians 289 

- L.B. Manzer 

Fear Detection and Removal - The Psychological Implications of the Technological 297 

Age (abstract only) 

- T. Helliwell 

Certification of NDT Operators in Canada - An Update 298 

- V. Caron

DAY 3 

KEYNOTE ADDRESS: Looking into the Future 303 

- R.S. Sharpe 

NDT of Structural Ceramics by High Frequency Ultrasonics 307 

- A. Fahr 

Hot Pressed Piezo-Electnc Ceramic Elements for Ultrasonic Transducers 316 

- N.D. Patel, J. van den Andel, P.S. Nicholson 

Ultrasonic Analysis of Voids in Glass: Theory and Practice 331 

- A..I. Stockman, J. van den Ande!, P.S. Nicholson 

Computer Simulation of Ultrasonic Testing 345 

- D.B. Duncan 

A Novel Approach to Eddy Current Imaging of Defects 356 

- D. Leemsns, M. Macecek 

Developments in X-Ray Stress Measurement, The CANMET Portable Stress 372 

Diffractometer 

- R.A. Holt 

Applications of Neutron Diffraction to Engineering Problems 387 

- T.M. Holden, G. Dolling, S.R. MacEwan, J. Winegar, B.M. Powell, R.A. Holt 

NDE in Polymers, an Example: Ultrasonics for the Determination of Density in 398 

Polyethylene 

- L Piche, A. Hamel 

An Industrial Application of Computer Assisted Tomography: Detection, Location 408 

and Sizing of Shrink Cavities in Valve Castings 

- P.D. Tonner, G. Tosello 

Computer Operated Composite Panel Testing (abstract only) 424 

- S. DeWalle 

Some Unconventional Techniques for the Inspection of Layered Materials 426 

- P. Cielo 

Materials Effects on Acoustic Emission During Deformation and Fracture 445 

- M.N. Bassim 

List of Delegates 454

DAY1 

KEYNOTE ADDRESS: Five Years Experience with Eddy Current Testing of NBEPC 1 

Generating Stations 

- A.B. Mitchell, H.M. O'Connor, R. Pickles 

Canadian Forces Non-Destructive Program Origins 11 

- J.S. Cavers 

Reducing Unwanted Effects -- Recent Developments in Eddy Current Equipment 16 

Features 

- A.G. Julier 

Nondestructive Examination Plays Key Role in Reactor Rehabilitation Program 27 

- CA. Wallis 

Evaluation of Ultrasonic Methods for the Detection and Sizing of Real Defects in the 40 

Base of Rails 

- L. Piche, J.F. Bussiere, J.P. Monchalin 

Optical Detection of Ultrasound at Distance 51 

- J.P. Monchalin 

Automated Ultrasonic Testing System for the Characterization of Defects in Weldments 66 

- M. Macecek, K. Luscott, J. Wells, D.K. Mak 

The Influence of Stress on the Inspection of Steel with Particular Reference to Gas 79 

Pipelines 

- D.L. Atherton, D.C. Jiles, C. Welbourn 

Defect Characterization and Sizing in Pipeline Weldments by Analysis of Ultrasonic 97 

Echo Features 

- J.P. Monchalin, J.F. Bussiere, R.W.Y. Chan, D. Sharp, D.R. Hay 

Capabilities/Limitations of Focused Ultrasonic Beams for Sizing Pipeline Girth 104 

Weld Defects 

- CA. Kittmer, J.B. Halle«, V.N. Sycko 

Flaw Characterization Using the Time-of-Flight Method and Ultrasonic Frequency 115 

Analysis 

- D.K. Mak 

The Introduction of Real-Time Radiography for the Inspection of Butt Welds in 130 

Offshore Pipelines 

- F.E. Reynolds Jr. (presented by L. Rathwell) 

Continuous Acoustic Emission Monitoring 145 

- J.S. Mitchell, M.N. Bassim

FIVE YEARS EXPERIENCE WITH EDDY CURRENT TESTING AT NBEPC GENERATING 

STATIONS 

A.ß. Mi tali ell 

Wen 1 Bïunnvick Raea-tcli and P r o d u c t i v i t y Council, Fxede x i c t o n , N.B. 

H.M. 0'Co 11 not 

{ büxme tiij with New Brunswick Research and Productivity Council) 

R. Pickles 

" a . G \ itii M.'(' d: f (' e c t '; • c Powe.fi C o m m i s s i o n . F x e d e ï i c t o n . N . B . 

SUMMARY 

Five years ago NBEPC established an eddy current test program at its 

generating stations in which all condensers were to be periodically inspected to 

determine future maintenance requirements. This program has proved to be 

extremely successful. 

By reference to specific situations encountered, the present paper 

describes how the aims of the program, its logistics, the techniques used and 

the assimilation of the results have gradually been upgraded to provide a 

powerful tool to assist in preventive maintenance. 

Faced with the operation of several older plants having a high incidence of 

condenser tube failure, New Brunswick Electric Power Commission (NBEPC) took a 

major step forward when five years ago it established an eddy current test 

program at fossil fuelled generating stations. All condensers were to be 

periodically inspected to determine future maintenance requirements. This 

program proved to be extremely successful in improving condenser reliability 

ana since that time has expanded to include numerous auxiliary heat exchangers 

and also heat exchangers at hydraulic and nuclear generating stations. The 

New Brunswick Research and Productivity Council (RPC) has worked with NBEPC 

throughout the program, providing technical advice and carrying out all the 

inspections. 

The three co-authors of this paper (and the program) represent what should be 

an iaeal balance of interests in any preventive maintenance program. There was 

an operations engineer whose interests are plant efficiency, preventive 

maintenance ana cost effectiveness, an NDE specialist who developed many of the 

special inspection techniques used and a metallurgist interested in the 

performance of materials and the implications of the inspection results on tube 

life expectancy. However, even with all this combined expertise, it has still 

taken several years to overcome all the problems inherent with such a program 

and only in the last year or so have we been able to consider most inspections 

as "routine". Experience has taught us that each large heat exchanger must be 

considered on an individual basis and will probably have its own unique problem 

requiring a "tailor-made" inspection program. 

In the short time available in this presentation, we would like to concentrate 

on some of the lessons learned and problems encountered in establishing the 

program. However, before we get to that, we should say something about the 

program in general.

MB Power has a total generating capacity of about 2500 MW of which 

approximately b0% is fossil fuel, 35« hydro and 15% nuclear. The Eddy Current 

inspection program has been concentrated on the larger fossil and nuclear units 

in which aluminum brass, cupro-nickel and titanium tubes are used in a wide 

range of different condenser sizes and designs. Station ages range from 

2 years to 20 years and several condensers have been partially retubed in the 

last five years as a result of the program. Station details and corresponding 

condenser tube materials are shown in Table 1. 

Because of the type of corrosion detected, it has been found beneficial to 

inspect most Al-brass and Cu-Ni tubed condensers on an annual basis. Results 

obtained from such inspections have been used in: 

- establishing accurate corrosion rates; which in turn have been used tc 

monitor chemistry control effectiveness and to calculate tube life 

expectancy: 

- defining susceptible areas within the condenser; such as, air removal zone, 

regions affected by flow channelling, regions effected by flow induced 

vibration, poor performance due to variable tube quality, etc., and: 

- guiding preventive maintenance by identifying individual tubes for plugging 

and in defining the extent of zones for retubing. 

Titanium tubed condensers have received an initial inspection and then, since 

corrosion damage is minimal, have only been reinspected to investigate specific 

technical problems occurring during service. As described later, tube damage 

due to vibration required a major inspection program. At another station 

damage due to weld spatter which occurred during construction was investigated 

when discovered at a later stage. 

Although several techniques for mechanizing inspection have been tried, RPC 

have always reverted to manual inspection methods as being the most versatile 

for condenser inspection. An experienced team of an inspector and helper can 

achieve an inspection rate of 100-150 tubes per shift, the slower rate of 

inspection being required for tubes exhibiting numerous significant 

indications. An immediate diagnosis of each tube is made at the time of 

inspection and all signals are recorded for later review. Tubes with defects 

deeper than 25?; of wall thickness are noted for future reference and those with 

deeper than 7b% recommended for immediate plugging. 

The first lesson learned in the program was how to be cost effective. 

At the outset of the program it was thought that condenser tube leakage could 

be brought down to extremely low frequencies by inspecting very large numbers 

of tubes. This illusion was soon dispelled by the realization that a 

significant number of leaks developed quickly by random events (impact damage, 

erosion at mussel shells, etc.) for which inspection provided no forewarning. 

It also became clear after the first few inspections that a large sample 

approaching a complete inspection was uneconomical both in cost and outage 

time. Experience has now shown that a sample of 300 to 600 tubes (3 to 7'= of 

the condenser) if judiciously chosen, is both adequate for monitoring generic 

degredation and of a short enough duration to be acceptable to plant 

operations. With this sample size, the program objective became clarified as 

monitoriny the overall condition of the condenser rather than preventing 

individual tube leakage.

- 3 - 

The second lesson learned was the value of "credibility" and how this is only 

established over a period of time. 

The commonest form of corrosion found in copper based condenser tubes is 

condensate grooving (Figure 1). This type of corrosion is worst in air 

removal regions and is attributed to ammonia and oxygen dissolved in condensate 

running down the baffle plates. It forms a narrow band of wastage around the 

tube which steadily progresses with years of service. Experience in station 

operation worldwide suggests a "typical" condenser tube life is 15 years and 

our own experience with condensate grooving fits this projection. NBEPC has 

found condensate grooving to be troublesome in most aluminum brass tubed 

condensers. 

Evidence of severe condensate grooving was found during an inspection made 

after only one year's operation at at one oil-fired generating station 

(Figure 2). If remedial action were to be taken to save the tubes it had to be 

immediate since grooving in some tubes had already apparently exceeded 25% of 

wall thickness. However, this was an unexpected problem for a brand new 

station, and a problem which had been identified by a new team still developing 

the required technique. No remedial action could be taken until these findings 

were verified and our task over the next three years became that of carefully 

monitoring the rate of progress of the grooving. This proved possible by 

comparing the phase rotation of the defect signal with calibration grooves of 

similar width machined in a piece of new tubing. The accuracy of this cross 

correlation was verified by the removal of typical corroded tubes from the 

condenser. After three years, an accurate corrosion rate had been established 

that left no doubt that the air removal zone of the condenser would need 

complete retubing after five years service. It also identified the need for 

mechanical changes to improve venting of the condenser. This knowledge allowed 

ample time for retubing preparations and also the selection of a more resistsnt 

material (Cu-Ni) for the air removal zone. Inspection was also able to define 

the exact extent of the zone requiring retubing. 

The third lesson learned was that an inspector must be knowledgeable enough to 

handle the unexpected. 

The inspector has an ongoing responsibility to correctly diagnose all signals 

picked up by his °ddy current instrument. Unexpected signals which cannot be 

related to previous interpretations are always worrying and can significantly 

reduce the rate of inspection by dominating the inspector's attention. Many 

such signals are generated by various magnetic effects and these have often 

proved troublesome to interpret. Weld spatter, magnetic inclusions, material 

of variable permeability and the effect of high stress can all produce 

anomolous signals needing very careful analysis. Our ability to recognize such 

signals has slowly improved with familiarity but we are still sometimes misled. 

Another annoying magnetic effect has been encountered with cupro-nickel tubes 

which show variable permeability. This causes signal drift and slows down 

inspection because of the need for frequent balance point adjustments. 

Metallurgical studies have demonstrated that the permeability of a piece of 

cupro-nickel tube varies widely depending on its residual cold work and heat 

treatment.

The fourth lesson learned was that there is a practical limit to an inspector's 

ability to interpret indications and that there should be no "loss of face" 

involved in requesting a tube to be removed to diagnose an unusual indication. 

As a demonstration that an inspector can be tempted to extrapolate theoretical 

experience too fer, we give the following example: 

At another plant in the examination of a conaenser tube which had been in 

service for many years, our inspector attributed a particular type of signal to 

a "magnetic" corrosion film which he felt must be on the inner surface of the 

tube (Figure 3). He was even able to demonstrate how this signal could be 

generated by magnetic tape wrapped around the inside of the tube (Figure 4). 

the signal was of completely the wrong phase angle to be interpreted as either 

an internal or an external defect. Metallographie examination of a tube pulled 

from the condenser a year later, failed to reveal this "magnetic film". It 

did, however reveal numerous corrosion pits on the internal tube surface 

(Figure 5). It was then demonstrated that the combined effect of closely 

spaced pits could rotate the angle of the resultant signal to such a degree 

that it would be unrecognizable as a partial through-wall defect. Five years 

ago when this incident occurred, the effect of multiple defects was not 

reported in the literature. Some research is now being carried out to 

elucidate these effects. 

The fifth and final lesson learned was that certain problems were beyond the 

capability of conventional eddy current techniques and that research laboratory 

back-up was essential. 

As a last illustration, I would like to talk about the detection of 

circumferential cracks in titanium condenser tubes at Pt. Lepreau G.S. All 

eddy current inspectors know that conventional bobbin probes used for tube 

inspections are very insensitive to circumferential cracks. When it was 

discovered that flow induced vibration had caused fatigue failure of some outer 

condenser tubes at Pt. Lepreau, we were faced with the problem of identifying 

within a few hours, a sufficiently sensitive method to detect other cracked 

tubes before they became major leakers. The solution was a new technique 

called "3D"*. 

This novel technique is so called because of its ability to add another 

dimension to eddy current testing — in addition to defect depth and volume, 

this equipment can also be used to determine its anqular position in the tube 

Wäll. 

A "3D" eddy current instrument is similar to a conventional eddy current 

instrument except that the 3D generates a three phase sine wave source. The 

special 3D probe is tx three coil probe connected in a three phase star 

configuration. 

* Developed and manufactured by Eddy Current Technology.

_ c _ 

If the 3 phase exitation is a constant voltage source of equal magnitude in 

each coil and the test object is symmetrical with respect to all 3 coils, then 

the offset voltage will be zero. However, any non-symmetrical defect, change 

in dimension, electrical or magnetic property will cause an impedence in 

balance and a defect indication. 

In practice this means that symmetrical features such as; support plates, tube 

fins, rolled joints, etc., will be invisible to "3D", whilst a non-symmetrical 

defect such as; an axial crack, corrosion or uneven fretting will give strong 

indications. In the present instance, the equipment was proved sensitive to an 

incomplete circumferential crack. 

In Figure 6 we can see an example of a circumferential crack in the titanium 

tubing. To the left we can see a stainless steel stake which in actual 

practice was installed exactly over the cracks. In Figure 7 we have compared 

the signals of conventional differential eddy current with "3D" for a support 

plate and a circumferential crack. The advantages of 3D for the latter are 

clearly apparent. The stainless steel "staking" proved to be an added 

complication since this was also unsymmetrical to the probe. Fortunately, 

however, the application of a very high frequency could be used to almost 

eliminate the staking signal whilst still retaining sufficient sensitivity in 

the 3D probe for the radial cracks. 

In this inspection program, several cracked tubes were found by "3D" in a water 

box where sensitive chemical detection equipment was unable to detect leakage. 

To conclude, we would like to summarize our experiences with regard to 

condenser inspection as follows: 

(a) Retubing a condenser costs 50 to 500 K$. We consider the cost of a 

typical inspection small, for the benefits accrued. 

(b) Each condenser inspection program must be "tailor-made" to its own 

particular operating problems. 

(c) Whereas we do not recommend annual inspections throughout the life of the 

condenser, we do recommend a couple of inspections during early in-service 

life to identify corrosion degradation at a stage where counter-measures 

are possible and worthwhile. 

(d) A first inspection should be considered both as baseline and an 

opportunity to identify any special techniques that need to be developed 

to measure a particular deterioration. 

(e) We are convinced that a well planned condenser inspection program will 

provide a great amount of information on material performance, degradation 

rate, susceptible zones and remaining tube life which will be of great 

benefit to plant operations.

STATION UNIT 

Coleson Cove 

Coleson Cove 

Coleson Cove 

Courtenay Bay 

Dalhousie 

Grand Lake 

Pt. Lepreau 

1 

2 

3 

3 

4 

1 

2 

8 

1 

TABLE 1 - DETAILS OF NEW BRUNSWICK ELECTRIC POWER COMMISSION 

STATIONS AND CONDENSERS 

CAPACITY 

(MW) TYPE CONDENSER MATERIALS 

202 

195 

202 

100 

100 

108 

180 

60 

600 

Oil 

Oil 

Oil 

Oil 

Oil 

Oil 

Oil/Coal 

Coal 

Nuclear 

Aluminum Brass/Partly 

Retubed Cu-Ni 


Retubed Cu-Ni 


Retubed Cu-Ni 

Aluminum Brass 



Titanium 


Retubed Cu-Ni 

Titanium 

NUMBER OF 

TUBES/CONDENSER 

6,800 

6,800 

6,800 

6,600 

6,600 

8,200 

8,950 

6,900 

12,000 

NUMBERTOF 

CONDENSERS

- 7 - 

Figure 1. Condensate grooving in aluminum brass condenser tubing. 

80 

70 

60 

50 

40 

30 

20 

10 

1 

1 

1 

1 y 

Is' 

*' 

1 

/I 

y 

0 2 4 6 8 10 12 14 16 

YEWS OF SERVICE 

Figure 2. Progression of condensate grooving in an oil fired 

generating station. 

y 

y

0*'? 

- 8 - 

Ä | 

'••• '~-!.if~ z '--Y- 

'••

- 9 - 

Figure 5. Corrosion pitting on inside surface of aluminum brass 

condenser tube. 

Figure 6. Cracked titanium tube. Note "staking" bar in practice 

covers crack.

- 10 - 

7/8 inch Titanium Tube 

3D 

i ' 

t ; , 

-r SuDOort Plate 1 

Figure 7. Comparison of signals obtained with conventional and "3D" techniques.

- 11 - 

CANADIAN FORCES NON-DESTRUCTIVE PROGRAM ORIGINS 

J.S. Ca.ve.fLA 

Ae.tioipa.ee. Maintenance Development Unit 

Canadian ¥oicei> Base T.'iznton, k&tn.a, On.tan.io 

ABSTRACT 

This paper presents an overview of how non-destructive testing in the 

Canadian Forces has evolved since its inception in 1959, with review of 

present training programs, development projects, condition monitoring, 

and a few related problems. 

CANADIAN FORCES NON-DESTRUCTIVE PROGRAM ORIGINS 

Traditional non-destructive testing in the Canadian Forces began in 

1959 with two Senior Non-Comissioned Officers being sent on course to the 

Department of Energy, Mines and Resources in Ottawa. On-job training for 

the next year and a half was carried out by an individual known to many of 

us; Mr. Bill Havercroft. In 1962 these two Senior Non-Commissioned Officers 

were transferred to what was then called the No. 6 Repair Depot (now the 

Aerospace Maintenace Development Unit) to establish an Non-Destructive 

Testing Shop. 

Between 1962 and 1964 approximately ten personnel had been absorbed 

into the NDT and were trained at various locations throughout North America 

and the United Kingdom. These ten personnel provided NDT services 

to all RCAF Bases in Canada, flying Bases, that is,as the CF NDT program was 

then, and still is today, an Air Element organization. Services to the land 

and sea elements were provided on a "when required" basis. These ten personnel 

personnel were constantly on the road carrying out inspections, in 

fact so much so that expansion became a must. As a result, the initial buildup 

began in 1965 with the first formal course being held at AMDU Trenton. 

Plans were now being formulated to estalish facilities at most aircraft 

operation locations. The facilities would not only be responsible for 

their own flying base, but for their geographical area as well- 

By 1967 there were approximately 40 trained technicians at Trenton, 

with support work to other Bases still dominating their workload. From then 

to present day, there has been a drastic expansion with facilities being 

established at 11 locations across Canada and Europe.

The facility at CFB Greenwood was closed in 1982, due mainly to the 

lark of N'DT inspectioi-.a required on the new Aurora aircraft, which replaced 

'he Ar.i;us. At that sant? tine, considerable support work was being carried 

• ut .'.t CF:i Sumraerside by CFB Greenwood personnel. With an increasing ND1 

requirement anticipated in the future, a division was made to establish a 

i.ev facility at CFB Summeraide. 

The total number of personnel now employed in NDT is approximately 

'•••ht with the largest number (40) being employed at AMDU Trenton. Table 1 

• : i splays total Service strength versus the number of personnel employed in 

."'I ". from 1959 to tha present. Our N'on-Dcstructive Testing Branch is currently 

'-: -ffpa with one Senior Engineering Officer with a post-graduate degree in 

r.,:,-destructive testing, and 3 Junior Officers, one of whom also has a postradu.i'-.e 

degree in nc --destructive testir.g. In addition, 27 Junior and 9 

Seniur NI "'• technicians with varying qualifications and experience arc- on 

s t n i f . 

1967 Edmonton - 6 

1968 Baden - 10 

Ottawa - 5 

1970 Moose Jaw - 5 

SKi.rXTlON PROCESS 

TABLE 1: NDT Service Strength 

1971 Chatham - 4 

1973 Cold Lake - 12 

1974 Shearwater - 6 

1975 Bagotville - 8 

1982 Summerside - 6 

Personnel selected for NDT training nu.st have at least 

live ve.ir' in their b.isic aircraft trade, and tl.ey must meet specific 

. iitii-.it. i ••ii requirener I;;. They also must have at l'.'ast seven years remaining 

i :, the Service. Volunteers are subjected to a fairly extensive interview by 

ilie Ideal NDT facilitv Supervisor, who will then forward ):iü rocomme dations 

to tiic Career M'.napel at National Defence Headquarter iii Ottawa, who is 

re ,jM.iTi;> ihle for NDT personnel selection. 

It sf.i.-il.i bo pointed out thnt NDT in the Canadian i-j;-ces is a Branch 

•il the basic aircraft trades. As such, many individuals will be re-employed 

in their original trade at a future date. Normally personnel will be employed 

in *:IJT for 7-10 years, allowing the Canadian Forces to utilize a fully- 

'i lined technician for a period of 2 - 5 years. 

Since J9'>2 approximately 250 aircraft technicians have been 

M)T-t r.i Lned, with, a larp(; number of those who have retired irom the Service, 

, rfi'Mtlv b' 1 ing enplcr.-ed In industry in the NDT field.

AMVU NON-DESTRUCTIVE TH.STING BRANCH 

- Li - 

Two of the responsibilities ot the Non-Destructive Testing Branch are 

a. to monitor development in industrial NDT, which includes 

evaluating equipment, and new or improved inspection 

b. develop and approve NDT inspection technique for Canadian 

Forces aircraft. 

As an example of how a technique may be developed, one of our NDT 

facilities, in co-operation with the local Aircraft Maintenance 

and Engineering Branch, identified the need to develop an inspection 

technique for the forward rotor shaft of the CH113/113A 

helicopter. The situation was addressed to Air Command who, in 

turn, recommended NDHQ that an NDT technique be developed for 

carrying out this inspection. NDHQ then created a project to 

carry out this tasking and forwarded it to AMDU for action. One 

of the areas which required inspection was the threaded portion 

of the rotor shaft at the carrier bearing retaining nut, which is 

approximately 27 inches from the top of the shaft opening. 

AMDU received the tasking on a Friday afternoon, and by the 

following Tuesday, two technicians had designed and manufactured 

a test apparatus and were carrying out successful inspections. 

It goes without saying that this is an 

exception rather than the norm for the amount of time personnel 

would require, or be given, to develop a viable inspection. However, 

it is typical of the type of work that is carried out at 

AMDU. There are between 30 and 40 inspections developed each 

year utilizing all NDT disciplines. 

Other responsibilites of the Non-Destructive Testing Branch include: 

a. carry out all NDT training in the Canadian Forces, provide 

training material, and screen all examination applications 

from Area Facilities; 

b. provide technial assistance to all ten NDT Area Facilities. 

This may be an aid in the interpretation of inspection results 

to guidance on equipment repair, and the operation of new 

equipment; 

c. provide supplementary personnel support to Area Facilities when 

they are short staffed, thus ensuring priority inspection deadlines 

are met; 

d. provide NDT support on aircraft periodic and sampling inspections 

at AMDU Trenton; and 

e. evaluate condition monitoring programs for possible inclusion in 

routine aircraft maintenance.

CONDITION MONITORING 

- 14 - 

The Condition Monitoring Program in the Canadian Forces is based on 

the following four analysis methods: 

1. Spectrometric Oil Analysis was first introduced to the air 

Element of the Canadian Forces in 1968, with the first 

laboratory being located at AMDU. This was instituted in 

response to a requirement to monitor the wear state of the 

J-79 engine of the CF 10A Starfighter. As additional 

components were monitored, a shorter response time was 

considered desirable, thus additional labs were opened at 

CFB Edmonton in 1971 and at CfB Baden (Germany) in 1972. 

As maintainers of various other aircraft realized the 

validity of the SOA program, labs were installed at 

Comox, Moose Jaw, Shearwater, Bagotville, Chatham and 

Cold Lake. These labs all operated with atomic absorption 

(AA) instruments. In the late 1970's the Canadian Forces 

began a transition to Atomic Emission (AE) spectrometers, 

and are presently in the final stages of conversion. This 

conversion from AA to AE spectrometers was deemed necessary 

for the following reasons: 

(1) to standardize guidelines with the U.S. Forces, 

since a large number of our oil wetted systems 

are the same as our American counterparts; 

(ii) operation of the AE spectrometers is simpler and 

faster; and 

(iii) accessories such as nitrous oxide, acetylene and 

keytone are eliminated. 

Approximately 35,000 samples per year are analyzed with a considable 

number providing enough information on various components to 

allow for early detection of problems which could have led to 

catastrophic failure and has proven to be a valuable maintenance 

tool. 

2. Flash Point Testing is utilized todetect fuel contamination of 

engine oils. It is carried out as a regular part of maintenance 

on aircraft engines equipped with fuel/oil heat exchangers. 

3. Ferrography was found to be an excellent complementary tool to 

the spectrometric oil analysis method. For this reason we use 

it to further analyze those oil samples which are found 

suspicious by spectrometry. In addition, we have started an 

evaluation on filter debris analysis-

- 15 - 

4. Vibration Analysis is also used extensively in the 

Canadian Forces, and has been adopted as a maintenance tool for 

Sea King helicopters on the east coast. There have been several 

successful maintenance actions resulting from this program, and 

in time, other aircraft will be included in the Vibration Analysis 

program. 

The Condition Monitoring Section is presently preparing a course for 

technicians involved in this field to meet the requirements which will ensue 

from the expansion of this program. 

Several long-range NDT-related projects are being initiated, amongst 

them is one to define an ultrasonic transducer performance, with the 

objective being acceptance/rejection specifications. Another project is on 

reliability of NDT inspections; the aim of which is to update our present 

data to allow us to more accurately determine inspection frequencies, and as 

a double check on our training and equipment. 

A Real Time Microfocus Radiography System has been purchased for our 

Branch, and will be installed shortly. It will be used for training, 

technique development and for the inspection of composite materials. 

CONCLUSION 

In conclusion, the Canadian Forces has been supporting the technology 

of non-destructive testing for the past 25 years. Our expertise has grown 

from a nucleus of two technicians to an organization of 106 personnel, and 

our Area Facilities have grown to ten. Our main purpose is to provide the 

Air Element with non-destructive testing support for both operational and 

maintenance environments. To this end, we have created our own training 

system, developed an in-depth technique development program, and just 

recently, a Condition Monitoring program. We also monitor new equipment 

entering the market, and liaise with other Serices in order to keep abreast 

of the latest developments in this ever advancing field.

- 16 - 

REDUCING UNWANTED EFFECTS ~ RECENT DEVELOPMENTS IN EDDY CURRENT 

EQUIPMENT FEATURES 

A Cati G. Jut let 

Hocking Et cet ion les Inc. 

Haio Iiiit'iumeufi Div-Lii.cn 

Temp te H(££.i, MP 

1. ABSTRACT 

It has long been established that the distribution of eddy currents is 

affected by many variables, some of which are desirable and some are not. 

This paper will seek to illustrate by reference to recent applications how 

eddy current equipment features have been developed to overcome some of these 

problem effects. 

2. PRESENTATION 

It has long been established that the distribution of eddy currents is 

affected by many variables, some of which are desirable and some of which are 

not. Within the time available today, it would be impossible to try to 

enumerate all of the problems and all of the solutions. So I will 

concentrate on four particular applications—two involving probes and two 

involving instrumentation. I will illustrate how their designs have been 

developed in order to overcome the associated problem effects. 

A recent task was the requirement to detect a fatigue crack at the root of 

small titanium helicopter engine rotor blades (Figure 1). In this example, 

the distance of the crack from the fir tree root was unknown, but expected to 

be less than 2 mm and could occur anywhere between the leading and trailing 

blade edges. 

Using conventional high frequency pencil probes the unwanted effects in this 

application were: 

1. Zero changes due to the proximity of edges and corners 

2. Variable surface geometry 

3. Difficulty in identification of indications 

The design, manufacture and use of shielded probes has been described at 

conferences and in technical publications previously, and so I do not propose 

to dwell on that aspect. Nevertheless, it should be remembered that a 

shielded probe has a smaller magnetic field than a normal unshielded pencil 

probe, and that the field is further reduced by incrasing the frequency. 

For this application, we chose to use an instrument that had three switch 

selectable test frequencies of 500 kHz, 2 MHz and 6 MHz. Using this unit we 

were able to construct some curves which allowed us to examine the effect of 

shielding at each of the three frequencies.

- 17 - 

Figures 2 and 3 show the response to edge effect of the unshielded probes of 

each frequency compared with the same response from the shielded type. 

A sinusoidal scan pattern was described within the radius of the blade 

between the fin and fir tree root. The output from the eddy current 

instrument was connected to a simple strip chart recorder. 

Figure 4 shows the results obtained for shielded and unshielded 2 MHz probes, 

followed by those for the shielded and unshielded 6 MHz probes, Figure 5. 

It can be clearly seen that the shielded 6 MHz probes provide a distinct 

advantage to this application. The final step was to design it for in situ 

usage. 

My next example involves the application of multi-frequency eddy current 

equipment to the examination of nonferrous heat exchanger tubing such as that 

found in air conditioner, condensor and nuclear steam generator systems. 

It is nothing new to be using eddy current techniques to examine heat 

exchanger tube bundles. We have been doing it for the last 10 to 15 years. 

Throughout all this time we have been troubled by one major problem. 

Tubes can vary from a few feet in length in a simple air condition unit up to 

nearly 100 feet in a nuclear steam generator, but in general the tube is 

supported along its length by what are known as support or baffle plates. 

These are generally made of steel and, as such, generate a large eddy current 

signal when the probe passes underneath one. 

The eddy current method can discriminate between any defects that are on the 

inside, the outside and also through wall of the tube. However, when a 

defect exists, as they often do, under the support plate area, the eddy 

current signal produced is a combination of that from the defect with that 

from the support plate. If the defect is large, then the resulting modified 

signal would indicate to an operator that there was a defect within that 

area. However what he is not able to determine is whether the defect is on 

the inside, outside or through wall. 

Even worse, if the defect is very small, such as a pin hole, then the 

modification of the support plate signal may be so slight as to prevent its 

detection. Thus the opertor fails to detect it. 

By driving the probe coils at two frequencies simultaneously, we can 

electronically mix the two frequencies and use this technique to effectively 

eliminate the support plate signal. 

In effect, we choose two different frequencies and display the support plate 

signal for each frequency on the screen of the eddy current instrument. By 

using what are known as "mixing" contrôle, we attempt to make the size, shape 

and phase angle of each signal identical to each other.

- 18 - 

When this happens, the instrument subtracts the second frequency from the 

first and, of course, if they are identical, the resulting signal is a zero 

(Figure 6). So by making the support plate signal identical in both 

channels, we can effectively eliminate it. 

For those of you who are interested in the mathematics, they are given in 

Figure 7. 

Taking the example of a titanium tube with a mild steel support, we can show 

how this occurs. 

If a defect such as a 1 mm diameter hole is introduced into the tube, we can 

show the resultant signals before and after mixing, with and without the 

support plate (Figure 8). 

We can extend this application into the aerospace industry for the 

examination of cracks extending from bolt or fastener holes which have steel 

bolts or fasteners inserted (Figure 9), and again show the resultant signal 

before and after mixing (Figure 10). 

My next example still, keeps us in the heat exchanger tubing industry, but 

this time deals with a probe design. 

In the development of heat exchangr tubing, the major criterion is thermal 

performance. Improvements in the heat transfer characteristics can 

frequently be gained by increasing the surface area of the tube relative to 

its length and also by altering the flow characteristics of the fluid inside 

the tube. This has led to tubes which are thinned, roped or studded on both 

the inside and the outside as illustrated in Figure 11. 

A further complication has been the presence of skip-finned areas which occur 

at the support plate regions (Figure 12). 

Even with dual frequency techniques, saturation of the instrument occurs at 

fairly low gain levels, therefore the sensitivity to defects in a 

conventional test is poor. A standard differential type probe will also not 

satisfactorily detect the long splitting type of defect, commonly called 

"zipper" cracks found in these tubes. 

One attempted solution to the problem of this type of defect developed here 

.in the USA is the cross axis probe which has absolute type detection 

characteristics with a differential response. 

The latest design is a special dual differential probe in which signals from 

opposing sides of the tube are cancelled. This enables the cylindrically 

symmetrical signals to be balanced out and gives good sensitivity to metal 

loss type defects. 

A comparison of the relative sensitivities to a hole defect and defects in 

the skip-fin area for the differential probes is illustrated in Figure 13. 

Finally, my last example involves the problem of eddy current instrument 

calibration. There is a requirement by both user and manufacturer to have an

- 19 - 

absolute method of ensuring that eddy current instruments are able to respond 

in the correct manner. Normally we use test blocks having machined or 

electrical discharge machined (EDM) slots for this purpose. However, a 

number of unwanted effects can occur and these are: 

1. Variations in slot depth from test block to test block 

2. Variations in slot width from test block to test block, but also 

possible on the same block 

3. Variations in test block material conductivity 

4. Variations in probe impedance due to age, wear and probe type 

There are normally three positions of the probe which are critical to the 

calibration procedure: 

1. Probe on a test block 

2. Probe with a small lift-off 

3. Probe over the calibration slot 

Using absolute values of inductance and resistance, we have devised an 

electronic reference unit which simulates these probe actions and test block 

variables, but without the need for either. 

Consider a probe on an aluminum test block. It will have an impedance Zl 

which can be represented by values of inductance LI and resistance Rl. There 

is also a dampening effect which is a function of resistance and is really an 

energy loss. 

When the probe is lifted off the material slightly, there will be an 

impedance change. In this case, the inductance will rise, the resistance 

will rise and the energy loss, or dampening, will be reduced. 

Finally, when the probe is over a machined slot, the inductance and dampening 

will increase, but the resistance will decrease. 

The electronic reference unit connects to the eddy current instrument through 

the normal probe connection and allows for simulation of the setting up 

procedure as well as for different materials and different test frequencies. 

It can be set up to simulate the response from a certain size slot. 

At a certain sensitivity level, we should obtain a certain meter deflection. 

This provides an opportunity for users to check and calibrate their 

Instruments and be confident the Instruments are performing and responding 

correctly. 

After doing this, then probes can be connected in the normal way and 

calibrated against standard test blocks in order to achieve the necessary 

sensitivity for the particular task in hand.

- 20 - 

(CJUIMTTCJW 

ROTOR BLADE SHOWING 

nrncu. pomion or ocncT 

TYPICAL POSITION OF DEFECT 

AND SCAN PATTERN 

Figure 1: Rotor blade showing typical position of defect and scan 

pattern. 

Voter 

Rc.iding 

F SO 

00 

50 

\ 

TOR UH tVITH STANDARD PENCIL PROBE ON STAINLESS STEEL 

2MHJ 

\ 

\6MHi \ 

\ 

\ 

\ 

\ 

\ 

\ 

\ 

A 

\\500KHz 

\ 

\ 

\ 

\ 

V \ 

\ \ 

\ \ 

\ \ 

\ \ 

1 2 3 4 8 8 

Distance of centre of probe from edge (mm) 

Figure 2: Edge effect of a standard probe on stainless steel at 

frequencies of 500 kHz, 2 MHz and 6 MHz.

Figure 3: 

Meier 

• -.1 ! i I 

.0 

00 

50 

OMHt 

FLAT 

/ . 

\ \ 

• > V M M » 

\ \ 

1 \ 

- 21 - 

EDGE EFFECT CHARACTERISTICS - ZERO CHANl.l 

1 [X.ATOR UM WITH SHIM Of 0 PHOtit S ON S1AINI | S'. silli 

\ 

\ 

\ 

t 

\900KHz 

\ROUND 

\ENO 

\ 

\ 

\ 

\ 

\ \ 

\ \ 

1 

N 

I s I 1 I 

1 1 3 4 5 • 

Distance of centre of probe from edge (mm) 

Edge effect of a shielded probe on stainless steel at 

frequencies of 500 kHz, 2 MHz and 6 MHz. 

Unihlaldad prob« - good oiada 

UniW«td*d pfoba - dafaellv* blade 

f Maftfad proba • good blada 

Shlaldad proba - dafacllva blada 

RESULTS WITH SCAN PATTERN FROM FIN TO ROOT - 6MH/ pmh.-. 

Figure 4: Results of inspection of good and defective blades with 6 MHz 

shielded and unshielded probes.

- 22 - 

Un»hl«ld«d prob* - good biad« 

8hl*ld*d prob* - good bisd« 

Shlvldsd prob« - d*t*cllva blad« 

RESULTS WITH SCAN PATTERN FROM FIN TO ROOT - ?MH; proh.- 

Figure 5: Results of inspection of good and defective blades with 2 MHz 

shielded and unshielded probes. 

TITAMUM TUBE WITH MILD STEEL SUPPORT 

CHANNEL 1 

15OKH1 

PHASEC D6 OUTPUT: 

MIX 

CANCELLATION OF SUPPORT SIGNAL 

Figure 6: Illustration of cancellation of support plate signal by using 

dual frequency "mixing" of signals.

- 23 - 

BASIC THEORY OF DUAL 

FREQUENCY INSPECTION 

Ai fraquancy F1 d«f«ct signal D1, IM«rf«rlnf 

F2 • • D2 " • a 

it In» rasullanla al tha alngls fraquanclaa ara RI and R2 

R1 = D1 • 11 . 

R2 = D2 • 12 ...(2) 

NO. ,i MI2 = 11

STEEL FASTENERS 

EXAMINATION OF FASTENER HOLES WITH IN-SITU STEEL FASTENER 

Figure 9: Schematic of examination of fastener holes with in situ steel 

fastener. 

INSPECTION FOR CRACKS NEAR 

STEEL FASTENERS 

PHASEC 06 OUTPUT FOR NORMAL FASTENER 

l'HASEC D6 OUTPUT FOR FASTENER AND SLOT 

Figure 10: Illustration of the effects of dual frequency mixing to inspect 

for cracks in fastener holes near steel fasteners.

Figure 11: Examples of developments in heat exchanger tubing to increase 

surface area. 

Figure 12: Examples of skip-finned areas in heat exchanger tubing at 

support plate locations.

- 26 - 

SKIP FINNED TUHt 

•< IG 5 INSPECTION OF SKIP FINNED TUBE 

WITH DIFFERENT PROBE TYPES 

Figure 13: A comparison of inspection results for a skip-finned area using 

differential, cross axis and dual differential probes.

NONDESTRUCTIVE EXAMINATION PLAYS KEY ROLE IN REACTOR REHABILITATION 

PROGRAM 

C liante.-!, A. Watt is 

Ontario Hijdio 

Bruce Nuclear Poivcn Vnvc ïcpmeu t, Tiverton, Ontario 

ABSTRACT 

In October 1983 a program of major proportions was initiated to 

reposition spacers supporting reactor pressure tubes in several Candu 

Reactors. Nondestructive examination played a major role in the 

successful rehabilitation program. Novel radiographie and eddy current 

techniques were developed to locate and monitor the movement of spacers 

during the program. 

REACTOR REHABILITATION PROGRAM 

On August 1, 1983, a pressure tube failed in Unit 2 at Pickering Nuclear 

Generating Station. During the investigation of the failed tube it was 

discovered that the central spacers, more commonly referred to as garter 

springs, were out of their design location. It was felt at the time that 

this was a contributing factor to the failure. Although the failure has 

since been attributed to metallurgical causes, the location of garter 

springs remains a concern because of heat economy and the uncertain 

effects their position may have on pressure tube life. 

The spacers have two main functions. One is to maintain concentricity 

between the calandria and pressure tubes so that when the reactor is in 

operation the hot pressure tubes will not come in contact with the 

relatively cool calandria tubes. Contact would result in poor heat 

economy and also a high temperature gradient in the material at the point 

of contact. The springs also have the ability to roll and thus act as 

bearings between the two tubes preventing fretting and allowing for 

unequal thermal expansion (see Figure 1). 

In the Bruce reactors there are 480 zirc-niobium pressure tubes which are 

approximately 6.5 meters long and have a nominal wall thickness of 4.1 

mm. Each of the tubes is supported by four centrally located garter 

springs each being spaced a meter apart. The 130 mm. diameter springs 

are made from square sectioned zirc-niobium wire wound 7 1/2 coils per 

cm. with a cross section of 4 1/2 mm. Running through the center of the 

spring is a girdle wire which has its ends resistance welded to each 

other. This maintains the circular shape of the spring assembly (see 

Figure 2). Pressure tubes and springs are installed at site as part of 

the construction program. Once springs are installed there is no 

possible access to them by any other means than cutting into the pressure 

or calandria tubes.

To gain primary statistical information/ tests were initiated soon after 

the failure to determine spring locations in uncommissioned units at 

Bruce and Pickering. It was soon evident that numerous springs had moved 

from their design locations in a seemingly random manner. 

Tests were conducted in a pressure tube mock-up to simulate actual 

reactor conditions. It was found that the springs/ which are compressed 

and fitted in the calandria tubes in the vertical position, had relieved 

the spring compression by assuming a slanted orientation (see Figure 2). 

The spring would then "walk" away from the design location in small 

incremental steps each time the pressure tube moved in relation to the 

calandria tube. 

By experimentation it was determined that mechanical work around the 

reactor accounted for a large portion of spring movement. It was also 

discovered that devices such as soil compactors working up to 500 meters 

away could also cause motion. Inspections conducted before and after 

commissioning indicated that circulation of water through pressure tubes 

and around calandria tubes would also contribute a significant amount of 

movement. 

Investigations in uncommissioned reactors indicated that by the time a 

reactor was ready for fuelling, the cumulative out of tolerance positions 

of springs varied from 200 to 500 meters and that springs would be 

randomly spaced. In some cases the four springs were clumped together at 

one end of a pressure tube which provided no effective support for that 

tube. 

There were three difficult choices that could be made: leave as is and 

accept lower efficiency and the risk of shortened pressure tube life; cut 

and replace using new or shortened tubes with redesigned spacers; or 

reposition existing spacers and somehow immobilize them in their design 

location. 

After an in-depth study, Ontario Hydro decided to correct spring 

locations in all uncommissioned units before going critical. 

Repositioning appeared to be the most attractive of the remaining two 

options as replacement would be costly in terms of time and materials. 

The repositioning option boiled down to; "How is it possible to 

selectively move springs in a 6 meter annulus without being able to 

either touch or see them?" 

A number of methods of moving springs were tabled during the numerous 

brainstorming sessions that ensued. Interest spread and ideas flowed 

in. The policy was to fully review any suggestion no matter how 

unconventional it sounded. Psychokinesis, exploding hydrogen and filling 

the annulus with urea formaldehyde were just a few that left the option 

list quietly and quickly. After sifting, the short list of options 

centered around electrical, vibratory and simple mechanical levering 

methods. Development programs were started in all three areas.

- 29 - 

In the electrical field, the discharge of large scale capacitors through 

a coil looked promising. This method was vigorously pursued. Mechanical 

levering was the simplest to perform but would only work successfully on 

springs slanted in the direction of desired travel. This was, in real 

life, a rare occurance as the springs were normally tilted away from 

their starting point. A system using a large sonic vibrator was also 

experimented with and eventually proved successful. 

The capacitor discharge method was furthest along in development when it 

came time to start repositioning on the first unit. It was to the state 

where it would consistently move springs in the desired direction. The 

system was made up of an extremely large capacitor bank, low resistance 

cables and an insulated copper coil that was placed in the pressure tube 

adjacent to the garter spring. The capacitor bank discharged through the 

coil causing a high intensity magnetic field that propelled the spring 

away from the coil. The maximum potential of the bank is 200,000 amps. 

Typical settings used on site were in the range of 140-160 ka. Each 

discharge of the bank would move the spring an average of 5 mm. 

Moving the springs solved half the problem. The other half was finding 

where the springs wWere and monitoring their movement during the 

repositioning program. 

Information on spring location was needed on a large sample very early in 

the program. The only workable examination method with immediate 

availability was radiography. Fifteen film cassettes were attached end 

to end on eight long aluminum rods. The rods were then inserted into 

reactor pressure tubes in a square pattern surrounding an unoccupied 

tube. An iridium source was then advanced in half meter increments down 

the pressure tube located in the center of the square. When the source 

had traversed the full length of the tube it was retracted, and the films 

were removed for developing. It took approximately one full week working 

around the clock to perform the initial inspection on the first reactor. 

It was obvious that radiography was a too costly and a too time consuming 

method of looking for springs over the long term. Eddy current, real 

time radiography, acoustic emission, ultrasonics and infrared detection 

were all methods that had the capability to find springs, but of these 

eddy current stood out as the method most suited to our needs. 

At first glance it appeared to be a simple task. Using basic theory, it 

seemed quite likely that a large absolute pancake type probe working at a 

low frequency would have sufficient penetrating power and sensitivity to 

locate springs. However it turned out that the geometry of the springs 

was such that insufficient eddy currents were generated by this method to 

produce a suitable signal. The eddy current signal is in fact generated 

by the girdle wire. Springs by themselves, or springs with broken girdle 

wires produce relatively no signal at all.

- 30 - 

The next step in the development program was the designing and building 

of a large inside differential coil close to the inside diameter of the 

pressure tube. The first probe consisted of two, two hundred turn coils 

separated by 15 mm. This arrangement provided a strong signal from 

vertically oriented spring assemblies. However, springs that had flopped 

over and lay slanted produced almost no signal at all (see Figure 3). A 

probe was quickly produced that had the coils on a 20° slant. This set 

up worked fine on slanted springs providing the spring was tilted in the 

same direction as the coils. To pick up springs inclined in the opposite 

direction or slued somewhere in between meant rotating the probe until 

the maximum signal was achieved. This was far too time consuming a 

procedure when it is considered that 1720 springs per reactor had to be 

located and mapped. 

AECL's Chalk River Laboratories had also tackled the problem of designing 

a single probe that could locate springs irrespective of orientation. 

Their solution was a one 25 mm. wide exitation coil with 6 mm. receive 

coils on either side. This configurate gave consistently strong signals 

from all springs with the exception of those within the 150 mm. long 

belled portion at each end of a calandria tube. Unfortunately, many of 

the springs had migrated to the bells and these had to be located as 

well. As an interim measure to solve this problem, a shielded isotope 

camera holder was fabricated that allowed a camera to be directly 

attached to an end fitting. It was then possible to radiograph the 

adjacent bell without evacuating the area. Paper radiography with 

instant development speeded the process to the point where results were 

available in less than five minutes. Later, Chalk River was able to 

provide a probe with segmented coils that was capable of detecting 

springs in the belled ends. 

The next hurdle was that of accuracy. In order to have any effect on the 

garter spring, the capacitor discharge coxl had to be placed with an 

accuracy of + 1 mm. and this up to eight meters down a tube. 

The initial scans had been performed by a probe mounted on an automatic 

stem unit. The heart of the stem unit is two foil elements wrapped on 

drums that form a strong lightweight tube as they are unrolled. They 

extend in much the same manner as a tape measure. A wheel encoder riding 

on top of the elements provides a digital readout equivalent to the axial 

distance (see Figure 4). 

To perform an inspection, the stem unit is placed on an end fitting and 

by means of a controller the probe is driven to the far end of the tube. 

The probe is retracted and as it does so, the axial distance and the eddy 

current signals are recorded on a strip chart. The finished chart is 

about a meter long and is scaled in proportion to the tube length. The 

accuracy of the chart is approximately + 10 mm. The encoder that 

provides the axial scan distance is also subject to error due to skipping 

on the element, deformation of the rubber rim of the wheel and slight 

variations in the wheel diameter. With a fine tuned system, accuracies 

of + 3 mm. can be achieved. However, during general use, a tolerance of

- 31 - 

jf 25 mm. was achieved 85% of the time. So although the automatic system 

was fast (5 minutes per tube) and gave a permanent record of acceptable 

accuracy, it was not accurate enough to place the capacitor discharge 

coils. 

To achieve the 1 mm. accuracy required for setting coils, a rigid system 

with an attached measuring tape was developed. A 30 mm. diameter 

aluminum push rod 7 meters long with a 3 meter extension was used. 

Cables ran to the probe through the center of the rod, which was 

centrally supported in the calandria tube by nylon disks. The probe was 

mounted on a wheeled carriage which helped reduce the lift-off effect. 

The carriage assembly was attached to the push rod by means of a 

universal joint (see Figure 5). 

The probe assembly is manually inserted into a pressure tube to a 

selected spring while watching the eddy current instrument. As the probe 

passes a spring, a distorted Figure 8 appears on the cathode ray tube. 

The probe is then slowly moved back and forth over the spring until the 

signal dot is at dead center of the Figure 8. At this point the center 

of the probe would be aligned with the center of the mass of the spring. 

The tape measure is then consulted and the distance recorded. To place 

the capacitor discharge coil, the probe is backed off an amount 

equivalent to the predetermined distance from the center of the coil to 

the bumper at the end of the probe. The capacitor discharge coil is then 

slowly fed in from the opposite end of the tube until it comes in contact 

with the bumper. It is then accurately positioned for firing. 

It took approximately three months of around the clock work to reposition 

the springs in the first reactor. Nine complete sets of eddy current 

equipment, each manned by a two-man crew, were available at all times. 

Seven units were committed to production work at the reactor face. The 

remaining units acted as standbys and also served the ongoing development 

programs. As with any program of this magnitude, there were a multitude 

of problems. 

The supplier of the eddy current equipment claimed the instrumentation 

was capable of operating without breakdown during normal usage. We were 

also advised that portions of the internal circuitry were so fine that 

static electricity from a person's hand could cause destruction. Our 

needs and expectations were for equipment that would operate on a 100% 

duty cycle for months at a time within meters of equipment producing 

currents in the neighbourhood of 150,000 amps. The external 

electromagnetic fields are of such magnitude that as the capacitor banks 

fire the cathode ray tubes are completely cleared. Because of this 

numerous breakdowns were experienced at the beginning. The situation has 

now largely been overcome by the use of isolating transformers and minor 

modifications to the electronic circuits. The equipment now operates 

almost without fault.

Each firing of the capacitor banks causes localized heating in the tube 

adjacent to the coil. The rise in temperature changes the material 

resistivity to the point where meaningful eddy current readings are 

unobtainable. Natural cooling through conduction and convection takes up 

to thirty minutes. This is an intolerable delay so cooling manifolds 

were fabricated that provided a cooling air blast through several jets on 

the effected area. This, in conjuction with working on several adjacent 

tubes in parallel, reduced down time to an insignificant amount. 

Unfortunately there are always mysterious signals that must be explained 

away. Reactivity mechanisms run vertically through the reactor adjacent 

to the calandria tubes. These produce a signal easily misinterpreted by 

the inexperienced as they are the mirror image of a garter spring 

indication. A spring adjacent to a mechanism produces signals that 

cancel each other out and results in undetectable springs. For this 

situation and where it was expected that two springs were nestled 

together, radiographie equipment had to be always ready. 

A general sequence of events established for repositioning in any tube is 

as follows: 

(1) Perform eddy current inspection and locate all four springs. 

(2) Feed spring locations into a computer which would suggest 

various options with respect to which springs to move and by how 

much. 

(3) Select a spring to move, place capacitor discharge coil and fire. 

(4) Cool tube and eddy current inspect to determine motion. 

The above sequence or slight variations on the above would be 

repeated until the tube was accepted by the computer. 

One of the alternate systems for moving a spring involved flexing the 

pressure tube repeatedly with a hydraulic jack located in the center of a 

2 1/2 meter long rigid beam. An eddy current coil was built into the 

beam adjacent to the jack and it was possible to monitor spring movement 

in real time. It turned out to be self defeating as the coil diameter 

interfered with the flexing of the tube and only modest movement was 

achieved. The probe was eventually mounted on the end of the beam which 

allowed progress readings to be taken by partially withdrawing the jack 

assembly. 

Another system that was developed by Ontario Hydro research utilized a 

large sonic vibrator that generated high intensity sound over a wide 

range of frequencies. The vibrator was located near the end of the 

selected pressure tube. An expanding head was fed into the prassure tube 

and clamped near the spring to be moved. An aluminum tube coupled the 

vibrator to the head. Various combinations of frequency and intensity 

were selected until the real time eddy current system indicated spring

movement in a positive direction. This system depended on having the 

tube free to vibrate and for that reason an air floatation system was 

used to support a lightweight eddy current probe. The sonic vibration 

method worked well but because of large sound deadening insulation 

cabinets it took up great quantities of valuable space in a very 

congested area. This coupled with vibrator breakdowns and improvements 

to the other available systems has led to its present limited use. 

As the first reactor program got underway, we were prepared for defeat. 

Opinion varied as to whether the repositioning was a cost effective 

solution as compared to retubing. It was also impossible to guarantee 

the equipment would perform for sustained periods. It was quite possible 

that coils would fail and cause arcing or that the large discharge of 

current through the coils would cause deformation to the tubes. Knowing 

of all the possible problems, it was easy to be pessimistic. The very 

starting of this program took a large measure of faith. 

It was well known the program could only function with teamwork and the 

constant exchange of ideas and the positive attitude that we could always 

do better. The first few days on the reactor were disheartening as only 

minimal travel was achieved and some days negative secondary motion on 

adjacent springs would result in a net loss for the day. Slowly daily 

movement totals increased as techniques, materials and teamwork 

improved. During the early stages four meters of travel a day was 

considered acceptable progress, whereas twenty meters of travel a day was 

not unusual by the time the second reactor had been completed. The first 

reactor required 72,000 firings of the capacitor banks and as many eddy 

current inspections to bring the springs to acceptable positions. 

Experimentation during commissioning has shown that when a tube is loaded 

with fuel the resulting sag pinches the springs between the calandria and 

pressure tubes and thus in most cases immobilizes them. It has therefore 

been our practice to load each tube with fuel as soon as the computer 

indicates acceptance based on maximum allowable spans between springs. 

Because of this pinch effect only minimal movement is expected in 

service. Periodic inspection during the life of the reactor will be used 

as confirmation. 

Garter springs have now been redesigned in such a way that they will stay 

fixed in their design location in any future units. The major change is 

having the spring in tension around the pressure tube rather than in 

compression inside the calandria tube. Bruce Unit 8 now has a complete 

set of the new springs installed and their positions are being monitored 

periodically during the construction program. So far, none have moved. 

Methods are now being developed to find and reposition springs on 

reactors that went into service before this problem was identified. The 

present plan calls for inspection and repositioning in tubes still filled 

with water.

- 34 - 

At the Bruce we have completed the rehabilitation of two reactors and are 

presently working on the third. The second unit was completed at about a 

quarter of the cost of the first. The garter spring repositioning 

program has been a high profile work assignment for quality control. In 

most cases, quality control inspections are the last task to take place 

in a long chain of production operations. By the time it gets to us, 

it's usually behind schedule and everyone is quite willing to point the 

finger at quality control for holding up production. This, however, has 

been one of the gratifying experiences where quality control was an 

integral part of the production process and as such played a positive 

role in achieving a rather remarkable goal. 

I would like to thank the following for the immeasurable help they have 

provided : 

- Ontario Hydro Central Nuclear Services for their help in getting us 

started. 

- AECL's Chalk River Nuclear Laboratories who did miracles in designing 

probes and automatic scanning equipment. 

- Hy staff of forty-four inspectors and supervisors who made the job look 

easy.

END FITTING BELLOWS 

1 

END FITTING (TYR) v • IOOO mm. 

(TYP. )mmJ—«- 

'•> I 

2750 mm. — *- 6500 mm. 

—TUBE SHEET 

o> I 2000 mm. 

GARTER SPRING 

GARTER SPRINGS (TYP.) 

SIMPLIFIED VIEWS OF PRESSURE TUBE 

( NTS. ) NOTE= All Measurements Approximate 

•CALANDRIA TUBE 

PRESSURE TUBE 

, CALANDRIA TUBE 

r . 

-*— M 2750mm. 

FIGURE 1

1 

GARTER SPRING 

IN VERTICAL POSITION 

GIRDLE WIRE 

ENDS OF GIRDLE WIRE 

RESISTANCE WELDED 

GARTER SPRING 

Zirc - Niobium 

of 

Travel 

z-Direction 

TILTED SPRING 

GARTER SPRINGS 

(NTS.) 

CALANDRIA TUBE 

GARTER SPRING 

PRESSURE TUBE 

FIGURE

- 37 - 

VERTICAL SPRING TILTED SPRING 

ISJ KSI 

1 

BET PU 

fr 

 

1 : Differential Probe 

2- Send/Receive Probe 

Trace on Cathode 

Ray Tube 

Strip Chart 

VARIATIONS IN RESPONSE FROM 

VERTICAL and TILTED SPRINGS with 

DIFFERENTIAL and SEND/RECEIVE PROBES 

FIGURE 3

DIGITAL READ OUT 

3.595 

TOIL ELEMENTS 

•TELESCOPING TUBE 

STRIP CHART RECORDER 

PRESSURE TUBE, 

MKI" STEM DRIVE 

(N.T.S.) 

FIGURE 4 

00' 

I

MEASURING TAPE 

HAND HELD PROBE 

(N.T.S.) 

FIGURE 5 

I

- 40 - 

EVALUATION OF ULTRASONIC METHODS FOR THE DETECTION AND SIZING OF 

REAL DEFECTS IN THE BASE OF RAILS 

l. Picke, and J.F. 

National Resea-tc/i Council Canada, Bsucii^ iv iHe , Qunhcc 

J.-P. Monchalin 

Encïgij, M-ôiei and Reiou-ïcciJ, 

Ottawa, Canada 

Abstract 

The presence of a seam or a crack in the base of rails is not uncommon and 

can, in some cases, lead to catastrophic failure if their depth should exceed 

= 0.5 mm (0.020 in). Such flaws should be detected during the production 

stage and it is required that the evaluation be done automatically in a 

continuous way and, of course, be nondestructive. We have carried out a 

survey of the different ultrasonic techniques which were applicable to this 

specific problem. In particular, the possibility of using surface (Rayleigh) 

waves was investigated, and compared to the more classical bulk wave 

pulse-echo techniques. Both artificial (machined) flaws and real flaws were 

studied. In the case of real flaws, results were found to be sensitive to 

whether the crack was closed and/or filled, especially when Rayleigh waves are 

used. 

1 INTRODUCTION 

An important task of NDE is that of detecting and sizing surface breaking 

cracks. Such defects act as stress concentrators and depending on their 

origin, nature, and shape may lead to catastrophic failure by fast fracture. 

An example is giv&n by railways in Canada where the danger is enhanced since 

the rails are submitted to particularly harsh weather conditions. For this 

reason Canadian railroads impose strict specifications on newly manufactured 

rails; the head and the web [1] are inspected but also the base, where 

experience shows that the presence of surface flaws can be highly hazardous. 

Defects in the base occur usually in the form of seams and cracks, along the 

principal axis of the rail and located near the center. The main criterion 

for acceptability is that their depth be less than 0.020 in (0.5 mm). 

Different techniques are used and surface flaws are detected by visual, 

dye-penetrant, magnetic particles and ultrasonic techniques. Ultrasonics has 

often proven to be a very powerful tool for NDE, however, for surface flaws, a 

reliable [2] inspection procedure is not available as yet. Various methods 

have been proposed which may give accurate results for artificial defects but 

their relevance [3] in characterizing real defects can be debated. Here, we 

describe the work that we have done to appraise different techniques which use 

standard NUT equipment applicable to plant conditions and purposely leave 

aside more sophisticated but less easily applicable techniques. In

- 41 - 

particular, we investigated surface (Rayleigh) wave propagation and compared 

the results to shear wave pulse echo measurements performed both on artificial 

idealized cracks and on real cracks, which were then cut open for micrographie 

evaluation of the depths. 

II METHODOLOGY 

The basic physical properties which are used to describe the propagation of 

ultrasound are rather well established [4]. Unfortunately, the application is 

not simple and the results are often qualitative rather than quantitative. 

The techniques and possible approaches are numerous and have been given full 

reviews [5, 6], and we shall refer to them as we try to find and evaluate a 

practical method of characterizing our surface defects. 

A Experimental arrangement 

We have used "off the shelf" commercially available contact longitudinal 

transducers. By fitting these with plastic wedges of different angles both 

shear waves and Rayleigh surface waves could be launched. After some 

experience, we settled for 0.5 inch diameter Panametrics probes of the high 

resolution type (well-damped). We used the smaller probes instead of the 

larger ones because a) the resolution was better (typically 3 or 4 

oscillations); b) the signal to noise ratio was greater (circa 6dB); c) the 

companion wedge being smaller, there was less attenuation and dead time due to 

the plastic; d) the coupling to the irregular surface of the rail was easier. 

The transducers were excited with a Metrotek (Mod. MP 203) puiser, the 

receiver also Metrotek (Mod. MR 101A) had a 60 dB gain and a 20 MHz 

bandwidth. The output was fed to a gated peak detector and to an 

oscilloscope. In sum, our arrangement is classical and corresponds to what is 

usually found where ultrasonic NDT is routine. 

B Description of the samples 

We shall report on two different kinds of defects that can be found in rails. 

These were detected in the production plant with magnetic particles and their 

depth estimated by shear wave pulse-echo measurements. The first (SI) is an 

obvious crack 18 cm (7 in.) in length with a depth (h) estimated to exceed 

0.7 mm (0.030 in.) near the center; the other (S2) is a small hairline, hardly 

detectable crack 8 cm (3 in.) long. 

Before we attempted measurements on real defects, we evaluated the different 

techniques on artificial flaws. These artificial flaws were narrow slits 0.25 

mm; 0.010 in.), 5 cm (2 in.) in length which were EDM machined near the middle 

of the base of the rail. Twenty such slits with depths ranging from 0.15 mm 

(0.006 in.) to 1.3 mm (0.050 in.) were fabricated.

III MEASUREMENTS AND RESULTS 

- 42 - 

In the midst of available techniques we can distinguish those which involve 

time of flight considerations and those which rely on the measurement of the 

scattered amplitude. Methods of the first category are based on the measurement 

of the delay introduced in the time of flight of a sound pulse by the 

presence of an obstacle. In principle these methods are very accurate and 

reliable because a) time measurements can be made with a high degree of 

precision; b) only the arrival time of the signal is considered and not its 

amplitude so that the influence of variations in transducer coupling and 

effects of attenuation in the material are minimized. The possible approaches 

are numerous [7] using either bulk waves or surface waves, but as a general 

rule the delay corresponds to the time for a wave to travel up and down the 

lateral faces of the flaw. 

We attempted several variations of the time of flight technique but to no 

avail. The cause can be traced to the probes, for which the damping time is 

still not strong enough. Immersion type transducers may allow to produce the 

very short pulses but even then the sensitivity to such small cracks would 

still be small. So for this first approach to our problem, time of flight 

techniques will be considered as advanced NDT, which is outside the scope of 

this paper. We shall thus turn to the more usual technique of scattered 

amplitude measurements, first with bulk waves then with surface waves. 

A Bulk wave techniques 

The amplitude of the signal, which is scattered when an acoustic wave encounters 

an obstacle, is used as a signature of the defect. Simple theory [8] 

assumes that the defect produces a mirror-like reflexion with an amplitude 

that increases regularity with the size of the mirror. Actually the interaction 

is much more complex [9], In even the simplest case, one has to consider 

that part of the energy will be reflected, part will be transmitted and 

that mode conversion will occur along with diffraction. For surface breaking 

cracks, there will be multiple scattering from the surface and this will further 

complicate [10] the picture. For real situations, the number of factors 

which will influence the results increases dramatically: transducer coupling, 

sonic frequency bandwidth and mode [11], crack shape and orientation [8], the 

internal roughness [12] of the crack, the state of stress [13] etc. However, 

some of these difficulties might be overcome if the defects are more or less 

of the same type and this appears to be the case for the problem at hand. 

We needed first to establish the ability of the technique to detect a flaw. 

The investigation procedure is illustrated in Fig. 1 where the inspection is 

shown to be performed from the base. The frequency is 1 MHz, which was found 

to be an adequate compromise between resolution and signal to noise ratio. 

An angle of 45° was found to be the most suitable. In Fig. 1, the flaw is a 

1.2 mm (0.048 in) EDM slit. The transducer is located near the edge of the 

rail and the ultrasonic beam is first reflected by the face opposite the base 

where it spreads before it reaches the flaw. Because of this wide angle

- 43 - 

search beam, the defect is more easily found then by direct insonification and 

results are not so sensitive to the position of the probe. However the 

amplitude of the signal is reduced. The calibration curve for the amplitude 

(A) of the echo versus flaw depth (h) was found to be linear. 

B Surface (Rayleigh wave) 

One way to circumvent the problem of geometry and sensitivity to exact probe 

location is to propagate surface (or Rayleigh) waves [14]. This technique has 

been used mainly for the detection of surface breaking fatigue cracks [15] and 

the published results indicate that it should be a valuable tool here also. 

The simplest way to launch a surface wave is to use a 90° angle probe, as 

illustrated in Fig. 2. We noted that in the absence of a crack, the 

propagation is not very sensitive to the surface finish: the attenuation does 

not increase noticeably when going from a polisheo. surface to the rough 

surface of the rail, however, the surface must be clean of water or traces of 

oil. In Fig. 2, the photograph shows the signal from a 0.5 mm (0.020 in) EDM 

defect. The frequency is 1 MHz and the gain is half that of Fig. 1: (1) 

corresponds to the initial pulse, (2) is the wedge/rail interface, (3) the 

signal from the flaw, (4) a reflexion from the far end of the rail. The flaw 

signal is large and very easily found, so that the technique is as a very 

sensitive means of detecting the presence of even the smallest flaws (0.2 mm, 

0.006 in). We have also experimented with leaky Rayleigh waves [16]. In this 

approach, where the surface is immersed, the sensitivity is even better. 

However, the surface wave is highly damped by the liquid and the inspected 

zone is reduced; the alignment of the transducers is critical and great care 

must be taken to eliminate spurious echos. 

Using the contact method, we went on to establish a correlation between the 

amplitude of the echo (R) and the depth of the flaw (h). The results for R 

versus the ratio crack depth/wavelength (h/X) are shown in Fig. 3, for 20 EDM 

defects at different frequencies (.5,1, 2.5 and 4 MHz). In the domain of long 

wavelengths or small defects, the relation between the amplitude of the echo 

signal and the crack depth can be reasonably approximated by a straight line 

(at 1 MHz for h = 1.3 mm; 0.050 in). The oscillatory behaviour which is 

observed in Fig. 3 is well-known [14, 17] and it is the basis for the spectral 

analysis [18, 19] approach. One may see the similarity with the problem of a 

transmission line, the crack acting as a cavity: the non-linearity corresponds 

to a multimode, multifrequency operation. We tried this approach where the 

crack is identified by the spectral density of the reflected pulse. The 

effects are rather small and the technique, if it holds promise, belongs to 

what we classified as advanced NDE. 

The defect-cavity analogy brings out the importance of the geometry factor: 

the propagation will be sensitive to the shape [19] of the defect (cavity) and 

more so at higher frequencies. If the EDM defects differ by factors other 

than their depth (h), this will manifest itself by scatter of the data points, 

as observed in Fig. 3 where the scatter exceeds our experimental uncertainty.

C Real defects 

- 44 - 

Here we report our results using both techniques in evaluating real defects in 

rails. The probe (1 MHz) is displaced along the length (x) of the rail; the 

depth (h) is estimated (.) using the linear calibration curve for bulk, waves 

and Fig. 3 for Rayleigh waves. The results obtained on rail sample SI using 

the bulk wave technique are shown in Fig. 4 and those using surface waves in 

Fig. 5. For the case of sample S2, the bulk wave signal was too weak to allow 

sizing of the depth; with surface waves we estimated a uniform dîpth h =0.13 

mm (0.005 in) along the length. Measurements were also performed at .5 and 2 

MHz, which were overall identical. 

The samples were then cut open and micrographs such as those of Fig. 7 (sample 

SI) and Fig. 6 (sample S2) were made. The actual depth of the defect from 

mouth to tip was measured. The values (V) which have been reproduced in Fig. 

4 and 5 show that the total depth is underestimated by both methods. This is 

even more so for the case of rail S2 for which the micrographs indicate that 

depth is of the order of 1 mm (.040 in). However, it can be seen in Fig. 6 

that the cracks are very often closed or filled, at least partly. Those 

sections of the cracks which are partly filled or closed will scatter 

ultrasound non coherently [12]. Since we have measured the amplitude of the 

reflected signal, it is to be expected that the main contribution was from 

that part of the flaw which is open. Even though the exercise is not 

foolproof, we have attempted to evaluate the open length of the crack. The 

values obtained are shown ([]) in Figs. 4 and 5. For the bulk wave 

experiments (Fig. 4), the agreement between the ultrasonically determined and 

the measured open depths appears to be better. 

In the case of surface wave reflexion, the analysis is more delicate, given 

that the scattering mechanism is more complex [20]. Various [18] studies have 

shown that reflected signals can be seen as the added contribution of a 

reflexion from the lip of the crack and one from the bottom. Even if the 

crack is closed, the lip will cause a signal to be reflected with an amplitude 

that is not related to the actual depth. This is what occurs in the case of 

sample S2. Concerning the contribution from the bottom of the crack, it is by 

definition correlated to the open depth. Then the correlation will also 

depend on the details of the geometry as was mentioned above. The calibration 

we have used was for 0.2 mm wide EDM slits, whereas real defects have widths 

which are two orders of magnitude smaller. This difference in geometry might 

in part explain why the agreement between the estimated depth (.) and the 

measured open flaw depth ([]) is not as good as for the bulk wave technique 

(Fig. 4). The description for the contribution of the closed or filled 

section of the crack becomes very involved. In a recent theoretical paper, 

W.M. Visscher established [21] that the response function of such a defect is 

multivalued and that the inverse scattering problem is undetermined.


Classical bulk and surface wave pulse-echo techniques were evaluated in their 

ability to detect and size surface breaking cracks in the base of rails. 

Although both of these techniques give good results for EDM slots, echo 

amplitudes from real cracks depend on crack closure and the presence of 

filling material. Rayleigh waves offer considerable advantages in terms of 

singal to noise and ease of detection [22] but their high sensitivity to 

factors other than crack depth make them unusable for sizing. Classical shear 

wave pulse-echo techniques were found to be in good agreement with the open 

flaw depth when calibrated on EDM slots. For proper sizing of the full depth, 

the techniques used were found inadequate. Since the closed or filled part of 

cracks are partially transparent to ultrasound, it is likely that more 

advanced NDT techniques using feature extraction in the time or frequency 

domain would also be inadequate without adaptation. This work points out the 

need for theoretical modelling and experimentation on the interaction of 

ultrasound with closed cracks and in particular of the scattering phenomena 

occuring at the tip [23], 

REFERENCES: 

[1] K.G. Hall; Non-Destructive Testing, p 121, June (1976) 

[2] H.J. Mayer; Mater. Eval., j42, 793 (1984) 

[3] J. Drury; Brit. J. NDT, p 200, July (1979) 

[4] E.A. Kraut; IEEE Trans. Sonics Ultrasonics, Su-^3_, 162 (1976) 

[5] P.A. Doyle, CM. Scala; Ultrasonics, 16, 164 (1978) 

[6] R.B. Thompson, D.O. Thompson; J. Metals, p. 29, July (1981) 

[7] K. Date, H. Shimada, N. Ikenaga; NDT Intl, JJ5» 315 (1982) 

[8] S. Serabian; Mater. Eval., J39_, 1243 (1981) 

[9] K. Harumi, H. Okada, T. Saito, T. Fujimori; IEEE Ultrasonics 

Symposium, p. 904 (1982) 

[10] T. Kundu, A.K. Mai; Trans. ASME, 48, 570 (1981) 

[11] V.U. Schmitz, F.L. Becker; Mater. Eval., _40, 191 (1982) 

[12] L. Adler, K. Lewis, M. de Billy, G. Quentin; "New Procedures in 

Nondestructive Testing" P. Holler, Ed. (Springer-Verlag 1983) p. 163 

[13] R.B. Thompson, C.J. Fiedler; "Review of Progress in Quantitative 

Nondestructive Evaluation 3"; D.O. Thompson and D.E. Chimenti, Eds. 

(Plenum Press, NY, 1984) 

[14] I.A. Viktorov; "Rayleigh and Lamb Waves", (Plenum Press, NY 1967) 

[15] J.M. Carson, J.L. Rose; Mater. Eval. p. 27, April (1980) 

[16] L. Adler, M. de Billy, G. Quentin; J. Appl. Phys. j)3, 8756 (1982) 

[17] M. Hirao, H. Fukuoda, Y. Mirura; J. Acoust. Soc. Ara. 7^, 602 (1980) 

[18] G.P. Singh, A. Singh; Mater. Eval. J39_, 1232 (1981) 

[19] Y.C. Angel, J.D. Achenbach; J. Acoust. Soc. Am., 7j>, 313 (1984) 

[20] B.W. Reinhardt, J.W. Dally; Mater. Eval., ^8, 213 (1970) 

[21] W.M. Visscher; ibid [15] 

[22] H. Seiger; NDT Intl., p. 131, June (1982) 

[23] R.B. Thompson; Applied Mechanics Division of ASME, Symposium on Wave 

Propagation in Inhomogeneous Media, San Antonio, Texas, June 1984

Flaw 

Figure 1: Schematic of experimental set-up used for bulk waves and 

oscilloscope trace for an EDM defect 0.040" deep. 

Flaw t 

Figure 2: Schematic of Kayleigh wave set up and oscilloscope trace obtained 

with a 0.020" deep EDM slot using half the gain used in Figure 1.

R 

.8 

.7 

.6 

.5 

A 

.3 

2 

.1 

— 

— 

A / 

/ ° 

D A 

D /A 

— /° 

/ ° 

j 

f 

STo • 

i 

B 

I 

\ \ 

o \ 

AA I 

t 

IvV - 47 - 

/ 

I 

O / 

/ 

/A 

/ A 

A 

A • 0.5 

° 1.0 

o 2.25 

* 4.0 

i 

0.5 1.0 1.5 

(h/X) 

i 

— 

— 

— 

— 

MHz " 

MHz 

MHz " 

MHz — 

Figure 3: Reflection coefficient for Rayleigh waves, versus flaw depth 

normalized to ultrasonic wavelength, h/X, for EDM slots. The 

solid curve was obtained from the literature. 

i

0.5 

E10 

Q. 

8 

1.5 

2.0 

- 48 - 

Position along defect 

01 2 3 4 5 6 7 8 9 10 X(cm) 

\ I 

• Bulk wave estimate 

v True full depth 

o Open flaw depth 

S1 

10 

20 

30 

40 

D 

CD 

O 

50 _ 

Figure 4: Depth of defect SI along its length obtained using shear wave 

pulse-echo (•) compared to the full depth (A) and open flaw depth 

([]) measured metallographically. 

60 

70 

80

CD 

Û 

0.5 

1.0 

1.5 

2.0 

- 49 - 

Position along defect 

0 "i : 2 : 3 4 ! 5 ( 

H. 

— 

> ; 

— 

i 

i 

i 

i 

| 

• 

i i1 

i 

i 

i I 

i 

i 

I 

i 

! 

\ . 

i 

\ 

\ 

\ 

) 

L 

i 

i 

• 

i 

ir 

< 

•j 

7 8 ino 

X(cm) 

I"' 

< 

i 

1 ' 

î *~ i 

""* 1 

- 50 - 

Figure 6: Optical micrograph of seam SI used for the comparison between 

Rayleigh wave and shear pulse-echo mearurements. (64X) 

mmm 

Figure 7: Optical micrograph of seam S2 showing large closed regions. (40X)

- 51 - 

OPTICAL DETECTION OF ULTRASOUND AT DISTANCE 

Jean-Piz um Monchalin 

Enzngu, Mines and Re souicc s, Ottawa, Ontario 

+ Work performed in the laboratories of the Industrial Materials Research 

Institute, Boucherville. 

ABSTRACT 

Various probes which are based on optical interferometry and are used to 

detect ultrasound at distance (distances from less than 1 millimeter to 

several meters are considered) are reviewed. Their properties and sensitivity 

(detection limit) are critically discussed. We analyze a Michelson 

interferometer probe which senses the surface displacement and another one, 

working as an optical frequency discriminator, which senses the surface 

velocity. A differential arrangement is also discussed. We present also 

preliminary experimental results obtained by laser generation and optical 

detection on steel at distances ranging from 1 to 2 meters. 

1. INTRODUCTION 

In recent years, considerable attention has been focused on the generation of 

ultrasound using pulsed lasers and numerous results have been published (1, 2, 

3, 4). This technology has the advantage to enable truly generation at 

distance, so it enables to produce ultrasound in hot pieces (e.g. in the steel 

industry) and at locations with difficult access. It is also worth noting 

that very short ultrasonic pulses can be produced, which enables an excellent 

resolution close to the surface. This technology, in order to be useful, has 

however to be coupled with a detection means which can also operate at 

distance. Optical interferometry has been recognized to be such a means. 

Numerous interferometric systems for the detection of ultrasound have been 

previously described, but if we except the one developped by a private company 

(5), none can really be applied to an industrial environment. Industrial 

applications require a system with a sufficiently high sensitivity to be able 

to be used on pratical rough surfaces. These surfaces have not been specially 

polished, as it can be done in the laboratory for more fundamental 

experimentation, and they scatter incident light in various directions, 

leaving a small fraction to be collected by the detection apparatus. This 

system has also to be immune from ambient vibrations. Also, except for very

- 52 - 

specialized applications such as microscopy, a probing area of the order of 

that produced by a conventional transducer is desirable. In this paper, the 

techniques for the detection of ultrasound based on optical interferometry are 

reviewed and discussed, especially their potential for application in the 

industry. We will also present some experimental results obtained by 

generating and detecting ultrasound at distances over 1 meter (both). 

First, it is important to understand how the ultrasonic vibration of a surface 

affects the light which impinges on it. It is well known that the surface 

motion shifts the frequency of the reflected or scattered light by the Doppler 

effect, but it is also useful to see the phenomenon in a different way. 

Let us assume that the surface moves by S(t) at a location where a light field 

Eo cos ut is incident,

- 53 - 

performing the beating right on the surface by creating a light grating on 

it. This technique is similar to differential Doppler anemometry (9) and is 

described in section III. 

II THE DISPLACEMENT MICHELSON INTERFEROMETER 

Such a system is sketched in fig. 2. The beam from the laser is focused onto 

the surface which acts as one of the mirror of the interferometer. Without a 

frequency shifter in the reference arm (homodyne interferometer), the detected 

signal can be written as follows (excluding constant terms): 

ID = 2 IL fk Js cos [4 IT 6 (t)/X -*(t)] (1) 

where Ij^ is the laser power, R is the effective transmission coefficient in 

intensity for the beam in the reference arm and S is that for the beam 

reflected off the surface (S is much less than unity for practical unpolished 

surfaces), $(t) is a phase factor which depends upon the interferometer path 

difference and is affected by vibrations. In practice, the reference mirror 

is mounted on a piezoelectric pusher which enables to ajust the path 

difference in such a way that $=+_ IT/2 + 2mr (n is an integer). It then 

follows that the signal is proportional to the displacement. In practice, 

electronic stabilization circuitry should be used to keep this adjustment in 

spite of ambient vibrations (10, 11, 12). It is also possible to have a 

system insensitive to vibrations by generating inside the Michelson 

interferometer two optical beams in quadrature (13, 14). This last design 

needs critical alignment and both have the drawback to be not easily 

calibrated (14). 

If the detection is performed about a higher frequency by shifting the optical 

frequency in the reference arm (heterodyne interferometer), calibration can be 

easily performed in real time: the signal from the detector (-2Tifgt is added 

to the phase in eq. 1) includes a carrier at the shifting frequency fg and 

two sidebands (as shown in fig. 1, except the spectrum is now centered on 

fg) and the calibration can readily be performed by comparing the sidebands 

amplitude to that of the carrier. An interferometric probe based on such a 

design has been realized (the frequency shifting is given by an opto-acoustic 

cell driven at 40 MHz). This probe has already been presented (15), its 

performances for the detection of pulsed ultrasound will be described 

elsewhere (16), as well its electronic circuitry for continuous (17) and 

pulsed ultrasound (18). It can be shown that the signal-to-noise ratio, 

assuming quantum noise limited detection and detection of both sidebands, is 

expressed by: 

is/iN = (4 n 6/A) ^2 11 S n/(B h v) (2) 

where n is the quantum efficiency of the detector, e is the electron charge, 

hv is the energy of a quantum of light of frequency v and B is the detection 

bandwidth.

- 54 - 

As the probed surface is not polished to a mirror like finish, the scattered 

light shows the well known speckle phenomenon. It can also be shown that the 

best signal-to-noise ratio is obtained when one speckle is detected (18) and 

when the speckle size is of the order of the incident beam size, which can be 

obtained by focusing on the surface with diffraction limited optics (19). 

Then, if the beam diameter on the lens is noted a and if the focusing distance 

between the lens and the surface is noted D, the solid angle of one speckle is 

" (a/D) . Assuming isotropic scattering and neglecting the losses produced by 

various optics, we then find that S » (a/2D) 2 . Taking IL = 5 mW, a = 4 mm, 

D = 15 cm a detection limit of = 1.5 Â is found for a bandwidth of 10 MHz. A 

detection limit of the order mentionned above has been observed with our 

system on nearly isotropic scattering surfaces (18). This limit can be 

lowered by using a higher laser power (the detection limit scale as l/^/T^), 

or a closer working distance D (the detection limit scales as 1/D) (20, 21), 

but closer focusing produces a smaller spot size and a shorter depth of 

focus. These results are consistant, as shown below, with the antenna theorem 

for heterodyne detection (22) which states that the étendue of the receiving 

system (i.e. its light gathering efficiency, exactly the product of the viewed 

area by the solid angle of the system aperture from the surface) is X at 

most. The maximum received power is obtained for the minimum spot size on the 

surface and is then equal, to II X /TT WQ , where WQ is the focal spot 

radius. Since WQ - XD/a we retrieve equation 5. 

Since the displacement Michelson interferometer probe produces a useful 

detection limit only for sharp focusing and close working range, its 

usefulness for industrial NDT applications is rather limited. However, it 

should be noted that it does not require a single frequency laser and as far 

as the path lengths in the interferometer are compensated, it can use a rather 

broad band laser. It can be very useful for ultrasonic field mapping and 

ultrasonic transducer characterization, since, in this case the signal being 

repetitive, the signal-to-noise ratio can be greatly increased by simple 

averaging. This probe measures most easily displacements normal to the 

surface, but it can also be used when inclined to measure in-plane motion 

(17). 

In conclusion, this type of optical receiver which is rather simple to realize 

has however a small étendue (its figure of merit) of the order of X . We will 

see below that velocity probes can be made with a much larger étendue. 

Ill THE DIFFERENTIAL DISPLACEMENT PROBE 

(OR OPTICAL GRATING DISPLACEMENT PROBE): 

As seen in fig. 3, this probe is made by having two beams issued from the same 

laser and separated by an angle 26 focused at the same location on the surface 

(18, 23). A grating is then produced on the surface and a part of the speckle 

field scattered from the surface is viewed by a detector. The grating is 

fixed when the frequency of the 2 beams is the same (homodyne probe) and 

moving when it is heterodyne. If N speckles of mean intensity Isp are seen

- 55 - 

by the detector, the detected signal in the case of the heterodyne probe can 

be written (the speckles add up incoherently) (18): 

l D = sß Isp cos f 2irf ßt + (?i -t2).t(t) + * (t)] (3) 

where fjj is the shifting frequency kj and k2 are the wavevectors of the 2 

beams, (ki - k2>.ô = 4n sin 6 (sin a 6x + cos a 6z)/A, a being the angle 

between k^ - k2 and the normal to the surface, ôx and Sz being respectively 

the in-plane and out-of-plane displacements. As seen above, this expression 

gives a carrier signal at fß and two sidebands, which enables easy 

calibration. In-plane and out-of-plane displacements can be detected, but 

this probe is particularly useful to detect displacements parallel to the 

surface (in this case ki-&2 is perpendicular to the surface and a = TT/2). 

When both sidebands are detected, it can be shown that the signal-to-noise 

ratio is given by (it is indépendant of the number of speckles): 

is/iN = (?i -k2).t \/n Isp/(2 B h v) (4) 

Eq. 4 also shows that, in order to improve the signal-to-noise ratio, 

ISpShould be maximized, which means that the speckle size should be of the 

order of the detection aperture and consequently that this aperture should be 

large enough to collect a large part of the scattered light. This means, in 

turn, very sharp focusing. When typical numerical values are used, a 

detection limit of the same order as the one found before is obtained. 

Therefore, this probe has the same drawback as the one described in the 

previous section. Furthermore, it tends to be more bulky since 0 cannot be 

too small for a reasonable detection limit. 

IV THE MICHELSON VELOCITY INTERFEROMETER 

A typical setup is sketched in fig. 4. We will note the main difference 

between this setup and the one of fig. 2: here the interferometer does not 

make use of the surface as a mirror, but views the light scattered by it. 

This type of system has been used in shock wave research (25) and has been 

also considered for the detection of ultrasound (26). The detected intensity 

for a given direction 6 of an incident ray is: 

ID = Ai + A2 cos (2TT V/AV + $) (5) 

where A, B and are constant, the free spectral range Av = 1/T =c/2Ad(8), 

(T is the delay time, Ad(8) is the difference of arms lengths for the incident 

ray inclined by 6. As sketched in the boxed diagram of fig. 4, which plots 

the spectral response given by eq. 5, frequency discrimination is obtained 

when the laser frequency (a single frequency laser source is necessary) is 

tuned to a zero crossing 2ir v/Av + = +_ IT/2 + 2mir, m being an integer). The 

surface being rough and irregular acts as an incoherent source, so the fringes 

are only observed at infinity (24) or at the focus of a lens. It can be 

readily shown that Ad(6) = Ad cos 6, where Ad is the path difference for the 

ray perpendicular to the mirrors. The fringes are then concentric rings, and

- 56 - 

in practice the central fringe is selected by using a circular hole at the 

lens focus (at the expense of a much higher complexity a mask pattern matching 

the fringes can be used). Taking a maximum path difference of A/4, it follows 

a limit aperture 654 « >J A/4Ad. In order to be able to detect ultrasound in 

the range 1 to 10 MHz, the free spectral range Av (the bandwidth can be 

defined as Av/2) should not be too large: taking Av = 25 MHz yields Ad = 6 in 

(even if folding mirrors are used, the system will be large and bulky) and 

8j4 = 2 10~ rd for A = 1.06 pm. Using an entrance aperture of 10 cm in 

diameter (and mirrors of this size), the system has an étendue of = 10" mm . 

This is much larger than the optical heterodyne systems described in section 

II and III, but not sufficient to use all the light scattered by an area of 

1/4 inch in diameter, 2 m away from a 10 cm aperture: the required étendue is 

- 0.06 mm . We present below a way to increase the étendue. 

The time response of this interferometer can be understood by considering that 

the path length from the laser source is changed by 6(t) for one arm and by 

6(t-T) for the other (27). Then, when the laser frequency is properly 

adjusted to the zero crossing, the detected signal is given by (6

used, a front lens too large to he feasible. Therefore, it can be concluded 

that the étendue will be limited by the viewed area and the front optics size 

(= 0.06 mm for example chosen) and that there is no need to increase further 

that of the interferometer (29). 

We evaluate below the sensitivity of such a system. We assume a beam splitter 

with 50% reflexion and transmission and neglect any other reflexion losses in 

the interferometer. We also assume isotropic scattering, a surface 

displacement U cos(2irfut + ) with fu ^> Av and quantum noise limited 

detection. Then it can be shown that the signal-to-noise ratio is given by: 

is / iN = 4(fu/Av) (U/X) JIT iL (1-A) Ü n/(hv a S)' (6) 

where II is the laser power, E is the étendue, A is the absorption 

coefficient Oi. the surface and S is the viewed area. Taking II = 10 W, A = 

0, X = 1.06 pm, fu = 2.5 MHz, Av = 25 MHz, B =10 MHz, n = 0.5, E = 0.06 mm 2 , 

S = area corresponding to a 1/4 inch in diameter, we find a detection limit of 

0.2Â. This limit is generally appropriate for the detection of laser 

generated ultrasound on a metal surface. For more absorbing surfaces, such as 

carbon fibers composites, a higher power may be required. One should note 

some improvement with a narrower bandwidth Au/2 and a higher étendue E, 

although there are limitations on the physical size permissible for such a 

system. The detection limit calculated above assumed a single frequency laser 

source which is very stable, both in amplitude and frequency. In practice, 

the laser should be stabilized with respect to the interferometer and the 

fluctuations of amplitude and frequency should be compensated by dividing or 

substracting electronically the signal coming from the sample with a signal 

representative of these fluctuations (28). 

V EXPERIMENTAL RESULTS OBTAINED WITH A VELOCITY INTERFEROMETER 

CANMET and IMRI are engaged in the development of an ultrasound laser 

generating and receiving system. We present below some results obtained with 

a preliminary version of this system. 

Fig. 7a shows the present experimental setup, which uses as sample a half inch 

steel plate. The generating laser is a Q-switch Nd-YAG laser (made by 

Lasermetrics, typically half a Joule multimodcs, 10 to 20 ns pulses), which is 

focused with a 2 m focal length lens on thfî plate. Ultrasonic displacements 

are detected in the present setup from the opposite side of the plate. Fig. 

7b shows the echos obtained from multiple reflexions in the plate. The noise 

observed on the picture is quantum noise, except the one at the beginning of 

the trace which originates from electromagnetic interference from the laser. 

The ultrasonic displacement for the generating conditions used is unipolar, 

but since the interferometer detects (as seen above in first approximation) 

the optical frequency change and the surface velocity, each echo appears as a

- 53 - 

bipolar pulse, as it is seen, using an expanded scale, in Fig. 7c. Although 

the laser pulse lasts 10 to 20 ns, the ultrasonic pulse, because of the very 

high attenuation of high frequencies, lasts about 10 times longer, as it is 

also observed. It should be mentionned that the surface which was illuminated 

by the laser pulse was covered by a water film, which had the effect to 

increase the ultrasonic source strength, but also that only a low power 

helium-neon laser (2 mW) was used. Developments are scheduled which should 

boost the signal by several orders of magnitude and bring the whole setup to a 

practically useful stage for defect detection and also thickness gauging. 

VI CONCLUSION 

We have presented a review of the various interferometric means to detect 

ultrasound at distance. Displacement probes based on a Michelson 

interferometer and optical heterodyning are relatively simple, have a high 

spatial resolution, a wide bandwidth, but a limited sensitivity- Their field 

of application is the mapping of ultrasonic fields, and they can be very 

useful for the evaluation of various ultrasonic NDT procedures and equipment. 

In particular, these probes could be used to check the operation ultrasonic 

transducers (in time and space over the transducer surface) and also could 

avantageously replace such ultrasonic field evaluation procedures using 

artificial reflectors (e.g. drilled holes). When shear displacements parallel 

to the surface have to be measured, a differential arrangement should be 

prefered, but both types of probes ("absolute" or differential) can be used to 

detect shear or compression displacements and have similar detection limits, 

although the differential probe tends to be more bulky. Velocity probes which 

use an interferometer such as the Michelson interferometer as an optical 

filter or an optical frequency discriminator are more sensitive and can be 

used for ultrasonic NDT at distance in an industrial environment. When 

coupled with ultrasonic generation by a pulsed laser, a NDT ultrasonic system 

with unmatched performances can be realized, enabling in particular to probe 

very near a surface, to work with geometries with difficult access, to probe 

pieces at elevated temperature and an easy scanning. 


This work has benefited from the skillful technical assistance of Mr. R. Hgon. 

REFERENCES 

1. C.A. Calder and W.W. Wilcox, "Noncontact material testing using laser 

energy deposition and interferometry", Materials Evaluation, 38, 86 

(1980). 

2. C.B. Scruby, R.J. Dewhurst, D.A. Hutchins, and S.B. Palmer, "Laser 

generation of ultrasound in metals", in Research techniques in NDT, 

R.S. Sharpe Ed., Vol. 5, Academic Press, 1983, pp. 411-413.

- 59 - 

3. D.A. Hutchins, "Non-contact ultrasonics using lasers", in Advanced NDE 

Technology, Proceedings of the Conference held in Montreal May 31 - 

June 1, 1982, J.F. Bussiêre ed., p. 125. 

4. P. Cielo, F. Nadeau and M. Lamontagne, "Laser generation of 

convergent acoustic waves for materials inspection", to be published 

in Ultrasonics, 1984. 

5. Krautkramer-Branson Inc., Lewistown, PA, Commercial note on the Laser 

Ultrasound system. 

6. A recent review is found in M.D. Levenson "Introduction to Nonlinear 

Laser Spectroscopy", Academic Press, New York, 1982. 

7. Wolf Bickel, "Method and apparatus for receiving ultrasonic waves by 

optical means", US patent // 4,345,475, Aug. 1982, assigned to 

Krautkramer-Branson Inc. 

8. Wolf Bickel, "Method and apparatus for receiving ultrasonic waves by 

optical means", US patent it 4,129,041, Dec. 1978, assigned to 


9. L.E. Drain, "The Laser Doppler Technique", J. Wiley, New York, 1980, 

see chap. 5. 

10. C.H. Palmer and R. Green Jr., "Optical detection of acoustic emission 

waves", Appl. Opt. 16, 2333 (1977). 

11. M. Kroll and B.B. Djordjevic, "A laser stress-wave probe with 

sub-angstrom sensitivity and large bandwidth", IEEE Ultrasonics 

Symposium Proceedings, p. 864 (1982). 

12. J.-P. Monchalin, Progress report IGM 83-231-02, Appendix B 

(unpublished). 

13. D. Vllkomerson, "Measuring pulsed picometer-displacement vibrations b 

optical interferometry", Appl. Phys. Lett. ^9_, 183 (1976). 

14. J.-P. Monchalin and R. Hgon, Progress Report IGM 84-402-01, section I 

(unpublished). 

15. Presented at the Canadian Steel Industry Research Association Sensor 

Research Workshop, Hamilton, Ontario, May 8-9, 1984. 

16. J.-P. Monchalin, "Optical detection of ultrasound at distance", to be 

published in the Canadian Society for Nondestructive Testing Journal. 

17. J.-P. Monchalin, Progress Report IGM 84-402-01, Appendix A 

(unpublished).

- 60 - 

18. J.-P. Monchalln and R. Héon, Progress Report IGM 84-402-02, in 

preparation. 

19. A. E. Ennos, in Progress in Optics, vol XVI, edited by E. Wolf, North 

Holland, (1978), see section IV, Speckle Interfereomety. 

20. J.E. Bowers, R.L. Jungerman, B.T. Khuri-Yakub and G.S. Kino, "An all 

fiber-optic sensor for surface acoustic wave measurements", J. 

Lightwave Tech. LT-1, 429 (1983). 

21. R.L. Jungerman, B.T. Khuri-Yakub and G.S. Kino, "Characterization of 

surface defects using a pulsed acoustic laser probe", Appl. Phys. 

Lett. 4£, 392 (1984). 

22. A.E. Siegman, "The antenna properties of optical heterodyne 

receivers", Appl. Opt. 5, 1588 (1966). 

23. R. Dändliker and J.-F. Willemin, "Measuring microvibrations by 

neterodyne speckle interferometry", Optics Lett., 6, 165 (1981). 

24. M. Born and E. Wolf, "Principles of Optics", Pergamon Press, 4th ed. 

1970, chap. 7.5.4, p. 301. 

25. L.M. Barker, "Laser Interferoraetry in Shock-wave Research", Exp. 

Mech., May 1972, p. 209. 

26. W. Kaule, "Interferometrlc method and apparatus for sensing surface 

deformation of a workpiece subjected to acoustic energy", US patent // 

4,046,477, (Sept. 1977), assigned to Krautkramer-Branson Inc. 

27 R.J. Clifton, "Analysis of the Laser Velocity Interferometer", J. 

Appl. Phys., hl_, 5335 (1970). 

28. W. Kaule, "Method and apparatus for receiving ultrasonic waves by 

optical means", US patent // 4,388,832, (June 1983), assigned to 


29. C.W. Gillard, "Optical differential interferometer discriminator for 

FM to AM conversion", US patent it 3,503,012 (March 1970), assigned to 

Lockeed Aircraft Corp. Calif.

JMU V J*fu Optical 

frequency 

a. Continuous ulrasonic 

excitation 

Sample 

- 61 - 

b. Pulsed ultrasonic 

excitation 

V Optical 

frequency 

Fig. 1 Optical spectrum of the light scattered from the surface. 

Ultrasound 

Lens 

Mirror 

S Frequency shifter 

-.--J 

Detector 

Beam splitter 

Fig. 2 Schematic of the displacement Michelson interferometer: without the 

frequency shifter, the probe is called homodyne and with it in the 

reference arm, it is called heterodyne.

Sample 

Lens 

Ultrasound 

- 62 - 

Detector 

Frequency shifter 

Beam splitter 

Mirror 

Fig. 3 Schematic of the differential displacement probe (or optical grating 

displacement probe): the probe in homodyne without the frequency 

shifter and heterodyne with it.

- 63 - 

Sample •D 

Av 

Optical 

P^ frequency 

Fig. 4 Schematic of the Michelson velocity interferometer. The boxed diagram 

indicates the response of the interferometer to optical frequency.

CD 

CO 

c 

O 

Q. 

W 

0 

- 64 - 

Sin(7TfuT) 

Ultrasonic frequency 

Fig. 5 Response of the Michelson velocity interferometer to ultrasonic 

frequency (the laser frequency is set at a zero crossing, as shown in 

the boxed diagram of fig. 4). 

Mirror M2 

f u 

Image of M2 

Mirror M1 

Fig. 6 Schematic of the Michelson interferometer with increased étendue.

Nd-YAG 

generating laser 

Interferometer 

& detector 

Illuminating laser 

Steel plate (Vz inch thick) Scope 

Fig. 7 Generation and detection of ultrasound at distance. a) Present 

experimental setup for the system under development, b) Ultrasonic 

echos sequence observed from a i inch steel plate, c) One echo from 

the sequence.

- 66 - 

AUTOMATED ULTRASONIC TESTING SYSTEMS FOR THE CHARACTERIZATION 

OF DEFECTS IN WELDMENTS 

M. Macecefe, K. Lu&cott, 3. Wells, 

Techno Scientific Inc., Voioniview, Ontario 

V.K. Mafc 

CAWMET, Ottama, Ontatio 

ABSTRACT 

An ultrasonic immersion inspection system, intended for research 

work on large, heavy section we laments, has been built. During 

its design special attention was given to three factors. First 

was the capacity and rigidity of the mechanical scanner and 

immersion tank. Second was the importance of flexibility in the 

data acquisition subsystem, to support a number of different data 

collection modes. Finally, there was a desire to provide the 

operator with a variety of display methods for test data. 

The resulting system provides for simultaneous collection of up 

to 8 channels of ultrasonic data, and allows digitization and 

storage of ultrasonic waveforms. Processing of the collected 

data is performed by a general purpose microcomputer, with test 

results displayed using interactive 3-d'' n ension colour graphics. 

It is expected that future work will involve the addition to the 

system of more advanced algorithms for the location and sizing of 

weld defects.


- 67 - 

Because of increased activity in the offshore exploration and 

petro-chemical fields, a need has arisen within the research 

community for a means of performing ultrasonic immersion 

inspection on large, heavy section weldments. 

Techno Scientific Inc. was contacted by the Physical Metallurgy 

Research Laboratories (PMRL), at the Canadian Centre for Mineral 

and Energy Technology (CANMET) to provide an inspection system 

which would be suitable for this type of work. It was specified 

that the system must: 

i) include an immersion scanner having a capacity of 

roughly 6 feet by k feet by 3 feet deep; 

ii) support simultaneous collection and processing of data 

from up to 8 transducers; 

iii) allow long-term storage of test data with later 

retrieval for processing and display; 

iv) provide a level of flexibility in data collection and 

processing methods appropriate to a research "tool". 

The last point should be emphasized. While many automated 

ultrasonic inspection systems have been described in the 

literature (1-7), all have been designed for a particular type of 

inspection, such as that of pipelines or composite panels, with 

the display of test results limited to 1 or 2 standard modes. 

2. SYSTEM DESCRIPTION 

The system can be divided into k major modules or subsystems: 

î) a scanning bridge and immersion tank, 

ii) a scanner controller, 

iii) an ultrasonic signal processing subsystem, 

iv) a data acquisition, storage, and display computer. 

Two separate computer systems are included. The first, referred 

to as the "host", provides the basic processing power and data 

storage for the system, as well as coordinating the operations of 

the the various subsystems. The second, or "slave" system, 

includes a programmable scanner contr . 1er (which drives the 

scanning bridge motors) and the ultrasonic signal processing 

hardware. Figure 1 is an overall block diagram of the system, 

showing the division between the host and slave units.

- 68 - 

2.1 Scanning Bridge and Immersion Tank 

The 3-a*es scanning bridge provides control over the movement of 

the transducer(s) within a volume of 72" (183 cm) by 48" (122 cm) 

by 40" (1O2 cm), corresponding to the size of the stainless—steel 

immersion tank. Each axis is controlled by a leadscrew driven by 

a 200-step-per-revolution stepper motor, and has a linear 

resolution of 0.0005" (0.01 mm). The scanning bridge and 

immersion tank are shown in Figure 2. 

Also included as part of this subsystem is a stainless—steel, 

stepper-motor-driven, drop-in turntable unit for the inspection 

of cylindrical testpieces (see Figure 3)« 

2.2 Scanner Controller 

The microprocessor-based scanner controller converts high-level 

commands from the host, such as "move" and "scar,", to the lowlevel 

control functions required by the stepper motors driving 

the scanning bridge and turntable. This frees the host computer 

to perform data acquisition and processing. 

Any two of the four possible axes X, Y, Z or T (for turntable) 

may be selected for a scan, with the step size and velocity 

specified independently for each. While the scan pattern used is 

fixed, the size of the scan and the sampling density i.e., the 

distance between ultrasonic samples along the scan axis, are 

programmable. 

Although normally controlled by the host computer, the scanner 

controller may also be used as a stand-alone unit by connecting a 

computer terminal to one of its RS-232 serial ports and entering 

commands directly to initialize parameters and begin a scan. 

2.3 Ultrasonic Subsystem 

Figure 4 shows a block diagram of the ultrasonics subsystem. It 

can utilize up to 8 transducers in transceive mode, or 8 pairs in 

transmit-receive mode. Under the control of the slave processor, 

each transmit transducer is pulsed in sequence while an eightchannel 

time-division multiplexer directs the received echo 

through the signal processing circuitry. Either a linear or a 

logarithmic amplifier is used to increase the amplitude of the 

ultrasonic signal, while an adjustable attenuator is available to 

provide control of the signal level in discrete logarithmic 

steps. Separate gating is supplied for each of the 8 ultrasonic 

channels, permitting independent control of the gate delay and 

width for each transducer. Additional signal processing can be

performed by a video detector stage. 

Peak detector and x-y display modules art also included to allow 

the ultrasonics subsystem to be used, along with the scan 

controller, in a stand-alone mode. During scanning, the x-y 

display interface combines the output from the peak detector with 

position information from the scanner controller to produce realtime 

modified c-scan images on the built-in crt display or on an 

external x-y chart recorder. 

Outputs are provided from each stage to allow viewing on an 

oscilloscope or digitization by the host computer. 

2.h Host Computer 

The host computer is a Multibus-based system utilizing an Intel 

8086 microprocessor and 8087 math co-processor, 768 kilobytes of 

program memory, and both fixed and floppy-based mass storage. A 

digital transient recorder (DTR) installed on the Multibus acts 

as a high-speed analog-to-digital (a/d) converter, permitting the 

host to digitize and store ultrasonic signal data. The DTR can 

digitize arbitrary signals at rates up to 20HHz, while repetitive 

waveforms may be effectively sampled at up to 160 MHz. 

A colour graphics module, consisting of 3 additional Multibus 

boards, provides a resolution of 1024 by 769 points. Up to 16 

colours can be displayed simultaneously, chosen from a pallette 

of 4096. Support for 3~dimensional graphics is provided by an 

ACM-SIGGRAPH Core library of subroutines (8). 

3. SYSTEM OPERATION 

Initially, the host prompts the user for all the necessary 

scanning parameters: the axes to be used, the scan dimensions and 

velocity, the sample interval, and the number of transducers. 

These are then sent to the slave, followed by a signal for 

scanning to begin. 

During an inspection, the slave signals the host computer 

whenever the scanning bridge reaches a position at which data is 

to be acquired. Then, for each of the transducers in use, the 

slave enables the appropriate ultrasonic channel while the host 

triggers the puiser and the DTR. Once triggered, the DTR 

collects 4096 amplitude data points, each with a resolution of 1 

part in 256. While these may be stored directly for later 

analysis, additional peak detection processing is usually 

performed to reduce the stored data to a manageable amount. 

If time-of-f1ight data is being collected, the host extracts from

- 70 - 

each waveform digitization the time position of the two largest 

peaks, assumed to represent a front reference surface, and a flaw 

or back reference surface. If the user has requested that peak 

amplitude data be collected, it is assumed that he has gated out 

the reference surfaces and so a single value corresponding to the 

amplitude of the highest peak in the data is stored instead. 

In some situations it may be useful to combine the 2 methods of 

data acquistion, by collecting both amplitude and time-of-f1ight 

values for each peak. This is done, however, at the cost of 

increased data storage space requirements. 

Data is stored on disk, from which it can be retrieved for 

further processing and display. There are k types of displays 

that the system can currently produce: 

i) rf or video waveforms (A-scans), 

ii) 2-dimensional sections through 3~dimensional arrays of 

data (including B-scans and C-scans), 

iii) modified C-scans 

iv) true 3-dimensional displays created by converting timeof-flight 

data into linear measurements. 

3.1 Waveform Display 

Figure 5 shows a display of the 4096 data points collected from 

an rf signal which had been sampled at 20 Mhz. This type of 

display is most useful for retaining a representative waveform 

from a scan for the evaluation of noise levels, signal strength, 

etc. at some later time. 

3.2 2-Dimensional Sections 

Amplitude or time-of-f1ight data collected from a 2-dimensional 

rectilinear scan is generally stored, either implicitly or 

explicitly, in a 3~ d imensional array as a function of scanning 

bridge position. The common B- and C-scans can be thought of as 

displaying subsets of this array in the form of sections or 

slices normal to one of the 3 dimensions. For example, the Cscan 

of Figure 6 is simply a slice made normal to the "amplitude 

axis" in a peak amplitude data array. Similarly, a B-scan 

results from making a slice through a time-of-f1ight data array, 

normal to the scan increment axis. 

However the display subsystem is not limited to standard B- and 

C-scan modes. For any given data array, 3 different types of 

sections can be displayed, with the thickness and location of 

each under user control.

3.3 Modified C-Scans 

- 71 - 

If a series of amplitude profiles are drawn on the same page or 

screen, each offset from the previous one by a small horizontal 

and vertical displacement, the effect is much like that of a 3dimensional 

projection. This method of display is known as 

modified C-scan, and can be thought of as all the possible Cscans 

of the testpiece, stacked one on top of another in order of 

increasing amplitude threshold. 

Beam profiles, used to aid in the characterization of ultrasonic 

transducers, are a common form of modified C—scan display. 

Figure 7 is an example, showing the echo response of a 3.5 MHz 

transducer to a 2 mm diameter target as a function of their 

relative positions. It shows graphically the ultrasonic beam 

strength of a transducer at various points within it's field. 

3-Dimensional Time-of-F1ight Displays 

Since the tirne-of-f1ight from a transducer to a reflector is 

proportional to their separation, 3-dimensional reflector 

coordinates can be calculated for a scan knowing only the 

inspection geometry and the measured time-of-f1ight values. 

These can then be displayed in 3-dimensional form using standard 

perspective or parallel projections (9) as supplied by the Core 

library (10). Figure 8 shows time-of-f1ight data for the 

testpiece of Section 3.2 plotted in this way. By combining the 

information from a complete set of B-scans into one image, it 

shows not only the extent of defects, but also allows some 

estimation of their depth. 

Because of the general nature of the 3~dimensional projections 

used, it is possible to view the object from any angle or 

distance. 

h. DISCUSSION 

The most important criterion for the system, as well as the most 

difficult one to satisfy, was that of flexibility. It has been 

achieved, we feel, through the use of a modular ultrasonic 

subsystem and by the incorporation of the DTR into the system. 

In particular, the ability of the DTR to digitize a received 

waveform allows almost any desired numerical algorithm to be 

applied to the data, while the storage of results as numeric data 

files simplifies the data processing task. The softwarecontrolled 

display subsystem also aids flexibility, 

by simplifying modification and improvement ot' the display routines).

- 72 - 

Further development of the system will proceed along 2 divergent 

paths. On one hand, there is a desire to build a production 

version of the current research system. In this case flexibility 

can be sacrificed, to some extent, for the sake of reducing 

costs. Two changes likely to be made will be the use of a 

somewhat lower resolution display, and the replacement of the DTR 

with a peak detector and a hardware circuit for time-of-f1ight 

measurements. 

The second development path will see the implementation of more 

sophisticated defect sizing and location techniques. Since the 

system already offers the capability to digitize rf echo signals, 

no specialized hardware should be necessary in order to 

experiment with the use of methods such as amplitude-time locus 

curves (ALOK) and the synthetic aperature focussing technique 

(SAFT), among others (11). 

ACKNOWLEDGEMENTS 

The authors wish to thank the staff of Physical Metallurgy 

Research Laboratories, CANMET and in particular Mr. V. Caron and 

Dr. J. T. McGrath for their numerous helpful suggestions, 

continued interest in the project, and a number of special 

references.

REFERENCES 

- 73 - 

1. Udagawa, T., et al. "Automatic Ultrasonic Testing and 

Recording System for Welds Utilizing a Microcomputer," 

Materials Evaluation (March 1982), pp. 305-311. 

2. Woodmansee, W.E. "An Interactive Graphics System Developed 

for NDT," Materials Evaluation (March 1980), pp. 33-36. 

3. Nielsen, N. "P-Scan System for Ultrasonic Weld 

Inspection," Proc. 3rd Int. Conf. NDE in Nuclear Industry, 

(Salt Lake City, 1980), pp. 171-195. 

4. Coote, R.I., et al. "Ultrasonic Inspection of Girth 

Welds," Proc. Int. Conf. Pipeline Inspection, CANMET 

ERP/PMRL 83-68 (OP-J), pp. 263-282. 

5. de Raad, J.A., et al. "Mechanical Ultrasonic Test 

Systems for Pipeline Welds," ibid., pp. 283-303. 

6. "Improved Ultrasonic System for Nondestructive 

Evaluation of Composite Materials," NTIAC Newsletter, 

vol.10, no.2, pp. 1-2. 

7. A. Singh, et al. "Automated Inspection of Corroded Steel 

Structures," Materials Evaluation (April 1983), pp. 568-570. 

8. Newman, W.M., Van Dam, A. "Recent Efforts Towards Graphics 

Standardization," Computing Surveys, vol.10, no.A, pp. 365- 

380. 

9. Foley, J., Van Dam, A. "Viewing in Three Dimensions," 

Fundamentals of Interactive Computer Graphics. Addison- 

Wesley, 1982, pp 267-316. 

10. Bergeron, R., Bono, P., Foley, J. "Graphics Programming 

Using the Core System," Computing Surveys, vol.10, no.k, pp. 

389-W3. 

11. Schmitz, V., et al. "Improved Methods for ultrasonic Defect 

Classification Reconstruction and Reliability," Proc. Int. 

Conf. Quantitative NDE in the Nuclear Industry (San Diego, 

1982), pp. 258-264.

TRANSDUCER SELECT/ 

PULSE CONTROL 

- 74 - 

Figure 1 — Block diagram of the inspection system, showing 

division between host and slave units. 

TRANSDUCER 1 

TRANSDUCER 8 

PULSES 

FROM SCANNER 

. CONTROLLER

Figure 2 — Overall view of the immersion tank 

Figure 3 — Drop-in turntable unit, with immersion tank and slave 

system in background

SCAIIIIING 

BRIDGE 

COLOUR 

GRAPHICS 

DISPLAY 

HA S S 

STORAGE 

SCANNER 

CONTROLLER 

HOST 

COMPUTER 

OPERATOR 

TERMINAL 

- 76 - 

OPTIONAL 

X-Y 

PLOTTER 

X-Y CRT 

DISPLAY 

ULTRASONIC 

SUBSYSTEM 

DIGITIZER 

PRINTER 

Figure U — Block diagram of the ultrasonic subsystem 

SLAVE 

SYSTEM 

fr 

uHOST 

SYSTEM

- 77 - 

Figure 5 — A096 point rf waveform as displayed on colour monitor 

Figure 6 — C-scan image of test block

Figure J — Beam profile of 3.5 MHz transducer 

Figure 8 — 3-dimensional perspective image of test block using 

t iiTie-of-f 1 ight

- 79 - 

THE INFLUENCE OF STRESS ON THE INSPECTION OF STEEL WITH PARTICULAR 

REFERENCE TO GAS PIPELINES 

V.L. Athenian, U.C. Jil&i and C. Wzlbouin 

Q.ueew'4 Un.lve.mIty, Kingiton, Ontanio 

*Research supported in part by the Department of Energy, Mines and Resources 

and the National Research Council (IMRI). 

ABSTRACT 

It is commonly recognized that stress induced permeability changes are a factor 

capable of giving signals or spuria in magnetic or electromagnetic non destructive 

inspection of steel. In fact the situation is more complex because ferromagnetic 

materials exhibit hysteresis with the result that their magnetic 

behaviour is dependent not only on the ambient stress and magnetic fields but 

also on both the magnetic and stress histories. 

Although stress effects usually complicate N.D.I, they can also be exploited, 

for example, far side corrosion signals can be enhanced by the application of 

stress which is concentrated in the regions of wall thinning. Magnetometer 

surveys of pipelines also show stress anomalies and changes in stress can be 

detected. The technique appears promising for monitoring stresses in other 

high performance steel structures such as offshore drilling rigs. 

Introduction 

It is well known that stress alters the magnetic behaviour of ferromagnetic 

materials and that changes in permeability give signals or spuria in magnetic 

or electromagnetic non destructive testing of steel. Generally permeability 

variations are an unwanted complication, particularly in the case of eddy 

current inspection, but they can be exploited when properly understood. The 

effect of stress on the magnetic behaviour of ferromagnetic materials is more 

complex than had been generally recognized until recently (1) because ferromagnetic 

materials are hysteretic with respect to both magnetic field and 

stress. Their magnetic behaviour therefore depends not only on the ambient 

stress and magnetic fields but also on both the magnetic and stress histories. 

It is possible that magnetic monitoring may therefore t' used to indicate 

ambient stress and stress history. 

Figure 1 shows as an example a log of a magnetometer survey above a gas pipeline. 

There are large fluctuations in the magnetic fields and some of these 

are attributed to the stresses introduced in the line at cold bends made in 

the field in order to follow undulating topography. Principally, however 

the fluctuations are due to the individual joints from which the line is 

constructed behaving as randomly orientated individual bar magnets as shown in 

Figure 2. The effect of stress introduced at cold bend deformations is to add

- 80 - 

an additional reverse magnet to the pipe bar magnet to give the magnetic 

profiles shown in Figure 3. 

These figures show that stress in ferromagnetic structures can produce significant 

but complex magnetic effects. : now review the basic results of the 

effects of uniform stress on simple sarii£-\es of ferromagnetic material such as 

line pipe steel. 

Effects of Constant Stress 

Figure 4 shows hysteresis loops for simple samples of line pipe steel under 

constant tension or compression. The effect of stress is to give a slight 

tilt to the hysteresis loops. At low field tension tends to increase the 

magnetisation, compression to decrease it although both tend to decrease it 

at high field. At first this seems to be explicable by applying Le Chatelier"s 

principle (which states that the application of a constraint to a system in 

equilibrium tends to shift the position of equilibrium to reduce the applied 

constraint) to the magnetostrictive effect (the change in length on magnetisation) 

. Figure 5 shows typical magnetostrictive behaviour for iron. The 

magnetostrictive coefficient (strain) is dependent on both field and applied 

stress. A positive magnetostrictive coefficient i.e. expansion on magnetisation, 

would suggest that tensile strain would cause an increase in 

magnetisation in order to reduce the applied stress and this seems to be in 

good accord with Figure 4. unfortunately the results of changing the stress 

in a constant field are more complex. 

Effects of Changing Stress in a Constant Field 

Figure 6 shows the changes in magnetisation occur ing from the same initial 

conditic.is during a tension cycle and a compression cycle under constant 

applied field. Both of these cycles result in increased magnetisation and 

are therefore contrary to predictions based on Le Chateliei's principle. 

Figure 7 shows the changes in magnetisation occuring as a result of similar 

stress cycles applied at different fields around a hysteresis loop. The 

results depend on the initial conditions and are due to a reduction in the 

magnatic hysteresis. 

Magnetic hysteresis occurs because magnetic domains are not free to move to 

their equilibrium magnetisated but are pinned (2). The effect of stress is 

to reduce this domain pinning and to let the magnetisation approach the true 

equilibrium (anhysteretic) value. The dotted line in Figure 7 shows the 

anhysteretic magnetisation curve, explained in Figure 8. The stress induced 

shift towards the anhysteretic magnetisation is further exemplified in 

Figure 9 which shows the changes in magnetisation induced by similar stress 

cycles. For cycles originating at points on the initial magnetisation curve 

the changes appear to be related to .he differences between the initial and 

anhysteretic magnetisation curves and this is confirmed by the very small 

changes resulting from cycles originating on the anhysteretic curve itself.

Irreversible and Reversible Changes 

- 81 - 

Le Chatelier's principle cannot be directly applied to ferromagnetic mate-ials 

since they exhibit hysteresis and are therefore not in true reversible 

equilibrium. The first effect of stress applied to a ferromagnet is to cause 

a reduction in the hysteresis, i.e. a shift toward the anhysterstic magnetisation. 

This effect is irreversible and is generally the dominant effect during 

the initial application of stress. The second effect is that the anhysteretic 

curve is itself stress dependent (3) as shown in Figure 9. This effect is 

essentially reversible. It could therefore be correlated with the magnetostrictive 

effect except that magnetostrictive coefficients are generally 

measured along the initial magnetisation curve which is irreversible and are 

therefore not directly applicable. 

The two effects appear to be approximately addative. In steel, the reduction 

in hysteresis is the dominant effect during the initial stressing but during 

subsequent stress cycles the reversible shift of the anhysteretic becomes 

relatively more significant. 

In stress monitoring applications both these effects may need to be considered. 

Stress Enhancement of Far Side Corrosion Signals 

A simple example showing the exploitation of stress effects is illustrated by 

Figure 10. The signal obtained from magnetic flux leakage detectors are 

greater for near side defects such as corrosion pits than for far side defects 

because the anomalous leakage fluxes are concentrated near the defect. If 

however a gas pipeline with external corrosion is stressed by the gas pressure 

external corrosion pitting will cause anomalous highly stressed regions on 

the inside of the pipewall. These may be detected using active or residual 

leakage flux techniques and used to enhance the far side corrosion signals 

as shown. The optimum procedure needs rather careful consideration as 

discussed below. 

Bending Stress Monitoring 

We have built (4,5) a laboratory scale demonstration of pipe bending stress 

monitoring using two 3.4 m long 11 cm diameter pipes sprung together. The 

apparatus is shown in Figure 11. Changes in magnetisation as indicated by 

the peak to peak amplitude of the external radial field profile logged with 

a fluxgate magnetometer one pipe diameter from the wall, are shown during 

two stress cycles in Figure 12. There are large irreversible changes occuring 

during the initial stressing after magnetisation and smaller reversible 

changes occuring thereafter. Figure 13 shows similar changes in magnetisation 

occuring during five bending stress cycles followed by remagnetisation and 

further stress cycles. The processes are shown conceptually on an M-H 

diagram in Figure 14. The effect of remagnetisation is basically to

- 82 - 

reinitialise the stress dependence so that the next application of stress 

will again cause relatively large irreversible changes in magnetisation. 

It should be noted that the magnetisation changes occuring in a complex case 

such as a non uniformly magnetised, non uniformly stressed pipe, indicated by 

external leakage field measurements as shown in Figure 13 are well understood 

in terms of the M-H diagram of Figure 14 even though it derived from the behaviour 

of a small sample with uniform magnetisation and stress. 

The technique shows promise for monitoring the maximum stress that a pipeline 

has been subjected to since last magnetisation (usually by magnetic inspection 

pigging). It might therefore be used to detect dangerous anomalous bending 

stresses due to foundation subsidence in pipelines constructed over permafrost 

or river crossings etc. Other applications to high performance steel structures 

not subject to periodic remagnetisation such as offshore drilling platforms 

may be simpler, particularly for statically loaded structures where stresses 

increase steadily without cyclic variations. 

Conclusions 

Stress has long been known to have major effects on ferromagnetic properties. 

For many N.D.T. techniques this is an undesireable complication although it 

can also be exploited both to enhance defect signals and as a potential 

technique for monitoring stress. This requires some knowledge of the basic 

effects of stress on a ferromaghet and until recently the situation has been 

confusing. It is now clear that stress causes both reversible and irreversible 

changes in magnetisation and that these are not directly related to magnetostriction 

as had generally been supposed. The irreversible changes, which are 

relatively large and occur mainly during the initial application of stress, 

are due to overcoming magnetic domain pinning and thereby reducing the 

hysteresis. The reversible effects are related to the stress dependence of 

the reversible (anhysteretic) magnetisation. We hope that with this recent 

improvement in understanding it will be possible to devise and develop 

further techniques for exploiting the effects of stress on ferromagnetic 

behaviour to augment current N.D.T. methods. 

References 

(1) D.L. Atherton and D.C. Jiles, 'Effects of stress on the magnetisation of 

steel', IEEE Trans, on Magnetics, Vol MAG-19, pp. 2021-2023 Sept. '83. 

(2) D.C. Jiles and D.L. Atherton, "Theory of ferromagnetic hysteresis', 

J. Appl. Phys., Vol. 55 pp. 2115-2120 March '84. 

(3) L. Dobranski, "The effects of uniaxial isostress on the anhysteretic 

and initial magnetisations of various pipeline steels", M.Sc. (Eng.) 

Thesis, Queen's University at Kingston, April '84. 

(4) D.L. Atherton, L.W. Coathup, D.C. Jiles, L. Longo, C. Welbourn and 

A. Teitsma, 'Stress induced magnetisation changes of steel pipes - 

Laboratory tests', IEEE Trans, on Magnetics Vol MAG-19, pp. 1564-1568 

July '83. 

(5) D.L. Atherton, C. Welbourn, D.C. Jiles, L. Reynolds and J. Scott-Thomas, 

'Stress induced magnetisation changes of steel pipes - Laboratory tests. 

Part II', IEEE Trans, on Magnetics Vol MAF-20, November "84.

FIELD 

VARIATION IN 

KILOGAMMA 

WELD WEcD 

FACTORY 

GIRTH WELD 

1° 

SAG 

BEND 

7Y 

LEVEL OF EARTHS FIELD COMPONENT 

(APPROX. 60 KILOGAMMA) 

Figure 1: A fluxgate magnetometer traverse made 1.2 m above a 200 m long string 

of 1.2 m diameter 13 mm wall pipeline showing fluctuations in the 

vertical component of magnetic field. These are due to the constituent 

20 m jointe acting as randomly aligned bar magnets superimposed 

on which are the effects of cold bends. 

I 

oo

BAR 

MAGNET 

IDEAL 

FIELDS 

MEASURED 

FIELDS 

N 

Figure 2: Pipelines are constructed typically from 7 m joints welded together. 

Each individual joint behaves rather like a bar magnet. There are 

north and south magnetic poles located a little way in from the 

ends. The external magnetic field lines run from north to south. 

If a magnetometer scan is made parallel to the pipe axis and at a 

constant distance outside it this gives radial and axial field logs 

as shown ideally and as measured above a real pipe.

IDEAL 

BEND 

FIELDS 

MEASURED 

BEND 

FIELDS 

5 ky 

Sky 

- 85 - 

Figure 3: Magnetometer radial and axial magnetic field component profiles above 

a real pipeline joint before and after cold bending show almost ideal 

field patterns after. The more complex patterns before arise 

because this joint consists of two joints factory welded together 

and there are small relict poles near the center of the double 

joint. The stress induced changes are a significant percentage of 

the earth's field itself and much greater than the natural variations 

in the earth's field which are normally less than 100 gamma in 

a day.

- 86 - 

Figure 4: Hysteresis loops for a small sample of pipeline steel measured 

under constant applied stresses.

EXTENSION 

CONTRACTION 

- 87 - 

COMPRESSED 

FIELD (H) 

UNSTRESSED 

TENSIONED 

Figure 5: Magnetostriction is the fractional change in length, the strain of 

a ferromagnetic material when magnetised. Initially iron expands 

on magnetisation but at higher fields it begins to contract, as 

shown. The effect is also stress dependent. Compression tends to 

enhance the region of expansion, tension to reduce it. The inverse 

effect, the change of magnetisation with stress variation, is the 

magnetomechanical effect which was expected to have a reciprocal 

behaviour.

-150 

H= 1.6 kA/m 

COMPRESSION 

-100 -50 

0 03 

à B 

TEH. 

0 02 - • 

0 01-- 

H= 1.6 kA/rn 

TENSION 

50 100 

Figure 6: The changes in magnetisation, AB, during a compression cycle and a 

tension cycle both starting from the same initial conditions. 

a MPQ 

150 

00 

CO

- 89 - 

Figure 7: The effects of identical stress cycles at different points around 

a magnetic hysteresis loop. The results appear to depend on the 

initial conditions. The dotted line is the anhysteretic magnetisation 

curve described in Figure 8.

- 90 - 

MAGNETISATION 

ANHYSTERETÎC 

INITIAL 

FIELD 

Figure 8: Magnetisation curves. The initial or normal magnetisation curve is 

obtained with an unmagnetised or demagnetised sample as the field 

is increased. If the field is then decreased, reversed and cycled 

a repeatable but irreversible path is followed. The area inside 

this hysteresis loop gives the irreversible energy requirement or 

loss. The anhysteretic curve is reversible and is the magnetisation 

curve which would be obtained if hysteresis effects were 

suppressed. Points on it are obtained, rather similarly to demagnetisation, 

by using an alternating field of steadily diminishing 

amplitude but superimposed on a steady bias field.

M 

Stress, 

MPa 

•*-200(cr) 

0(H) 

H, kA/m 

Figure 9: The anhysteretic surface as a function of applied field at constant 

stress for a sample of line pipe steel.

(b) 19 mm 

diameter 

o 

Hard 

spot 

19mm 

diameter 

o 

Interior 

- 92 - 

19mm 50 mm long 

diameter % 12 mm wide 

12 mm long 

x 50 mm wide 

19 mm 

diameter 

o 

V7A///7- LJ -(/ r/rV// / WT////77 

-500 mm- 

Exterior 

Figure 10: (a) Hall probe scan of the radial residual field leakage flux 

above a short previously magnetised piece of pipe with the simulated 

defects shown in (b). (c) The same, but with reduced scale, after 

a stress simulating the effects of line pressure has been applied. 

There are now clearer superimposed signals from the anomalously 

stressed regions due to pipewall thinning. These 'stress shadows' 

tend to preferentially enhance the signals from far side corrosion 

since their anomalous stresses are on the near side to the detector.

Figure 11: Laboratory scale monitoring of stress near pipes stressed by bending 

or by internal pressure. The pipes are supported by a central 

saddle and tied together at one end. They are bent by jacking 

the other ends together. Bending stresses are determined by strain 

gauges bonded to the peak stress point of the pipe and are verified 

by stress calculations from both the measured deflection and the 

required force. Internal pressure is provided hydraulically. A 

fluxgate magnetometer is being used to monitor the stress induced 

changes in the external field. This is a sensitive device used 

for geomagnetic prospecting, mine detection, etc. The pipes can 

be arranged at any horizontal and vertical angle to the earth's 

magnetic field in order to investigate its effect.

Q. 

I 

CO 

zL 2 

(Ky) 

TYPICAL 

UNCERTAINTY 

0 20 40 60 80 

MAXIMUM BENDING STRESS (MPa) 

Figure 12: Peak to peak values of changes in radial fields measured at one pipe 

diameter as a function of maximum bending stress during two bending 

cycles of a demagnetised pipe. 

100 120 

i

(K/) 

-fr— BASELINE; MAGNETISATION DUE TO 

FIRST 3A CURRENT PULSE 

O — BEFORE SECOND 3A PULSE! ^O*'lNITIAL STRESSING 

"^^tf^NEXT 5 STRESS CYCLES 

— AFTER SECOND 3A PULSE: *D*'lNITIAL STRESSING 

20 40 60 80 

MAXIMUM BENDING STRESS (MPa) 

Figure 13: Peak to peak radial field changes as a function of maximum bending 

stress for a pipe magnetised by a 3A current pulse in a coaxially 

wound solenoidal coil then subjected to 5 bending stress cycles 

followed by a similar magnetising current pulse and a further 

stress cycle. 

STRESS CYCLE 

TYPICAL 

UNCERTAINTY 

100 110

- 96 - 

Figure 14: Conceptual M-H behaviour during the processes of Figure 12. Initial 

magnetisation 0-a-b. 1st stressing: b-c, remainder of 1st 5 cycles: 

c *•—^ d, remagnetisation: d-e-f, 6th stress cycle: f-g-h. The 

application of stress tends to cause an initial large irreversible 

shift in magnetisation towards the anhysteretic magnetisation, zero 

in this case because there is no applied field. During subsequent 

stress cycles there are smaller nearly reversible changes. 

Remagnetisation produces a slightly greater magnetisation than from 

the initial unmagnetised state and tends to reinitialise the stress 

situation.

- 97 - 

DEFECT CHARACTERIZATION AND SIZING IN PIPELINE WELDMENTS BY 

ANALYSIS OF ULTRASONIC ECHO FEATURES 

J. P. MonckaZin 

Encigij, Wines and Resources Canada 

Ottawa, Ontario 

J. F. Bus sie-in 

Nat ional Re seaicli Council o {

- 98 - 

characterization program based on these lines. Classification uses the 

algorithms based on a linear discriminant function, an empirical Bayesian 

classifier and a K-nearest neighbor method previously developed for acoustic 

emission by Tektrend Inc. In this paper, we report preliminary work made on 

this project, the application being considered at the beginning being the 

characterization and sizing of pipeline weld defects. 

EXPERIMENTAL 

Since the application considered first in this program is pipeline welds, 

samples with defects were prepared from 1/2 inch (12.5 mm) thick steel 

plates. Planar flaws were obtained by butt welding two plates laid side by 

side: one weld pass was made with insufficient penetration, then the plates 

were turned upside down and a second pass was made, leaving in the middle an 

unfused area, running the whole plate length. In this way, planar defects, 

which perfectly simulate internal lack of side wall fusion, were obtained. 

Long defects (10 inches or more) having uniform through-wall dimension of = 1, 

2 and 3 mm were obtained. A flaw of 1 mm or less is generally considered as 

benign whereas larger flaws are considered critical (1). 

Four plates (labelled #3, #4, #5 and #6) were made in this way and then cut in 

half. One half of each plate was machined to remove the weld crown, whereas 

the other half was left unaltered. Defect size was measured visually from 

both ends of the plates. The following results were found: 

Plate #3: 2.75 mm - 3 mm 

Plate #4: 2 mm - size not well defined from one sick? 

Plate #5: 3.5 mm - 3 mm 

Plate #6: 1 mm - 1 mm 

Experimental data were taken from these samples at several locations along the 

defect using the pulse-echo immersion technique and broad-band commercial 

transducers at 5 MHz. Both planar and focused transducers were used with 

incidence angles of 45°, 60° and 75° in steel. Transducer height was adjusted 

in order that the defect was in the far field region of the planar transducer 

or in the focal zone of the focused transducer. The position of the 

transducer along the direction perpendicular to the defect was determined by 

maximizing the echo height for the planar transducer and somehow by trying to 

observe the maximum number of peaks with the focused one. The signal echo was 

digitized using a digital oscilloscope and the data was then stored for 

further computer processing. So far data has only been taken from the plates 

with the crown removed. 

DISCUSSION OF OBSERVED ECHO FEATURES 

The data taken with the planar transducer at 45° and 75° do not present any 

readily obvious distinguishing features, except some higher frequency 

oscillation at = 7 MHz for the 3 mm defects at 45°. At t>0°, trailing pulses 

are observed. The transducer being well damped poor signal-to-noise ratio is 

encountered and diffracted edge waves as predicted by elementary theory could 

not be clearly observed: theory predicts for a vertical defect that two pulses

- 99 - 

separated in time by 2d cosr/vs should be observed where d is the defect 

size, r is the refraction angle in steel and vs is the shear velocity. This 

time interval amounts to about 1.4 \is, 1 ys and 0.5 ys for a 3 mm defect at 

45°, 60° and 75" refraction angle, respectively. 

With the focused transducer properly adjusted, a much higher signal-to-noise 

ratio was obtained and several pulses could be observed as shown in figs. 1, 2 

and 3 obtained at 60° with plate //3, plate ilk and plate //6, respectively. The 

first two pulses can be linked to shear waves diffracted by the edges and the 

observed time delay corresponds approximatively to that given by elementary 

theory (these pulses can be hardly separated for the 1 mm flaw sample). The 

other peaks which follow (one or two are observed) are linked to conversion 

into a compressional wave, to reflections of this wave by the surfaces of the 

plate and reconversion into a shear wave. This phenomenon was demonstrated by 

using a second compression transducer (planar) oriented normally to the plate 

and located right above the defect (see fig. 4). Using this transducer in a 

pitch-catch mode, the multiple reflections of the converted compressional wave 

can be seea as shown in fig. 5. The first pulse seen with this transducer is 

produced by a parasitic wave directly scattered from the first surface. 

Further evidence was obtained by milling a layer of constant thickness from 

the bottom of the plate. The periodic spacing observed with the compression 

transducer was observed to diminish by a time interval corresponding to the 

propagation of a compressional wave through the removed thickness. In the 

same way, the time interval between the two edge pulses and the third echo 

(also the fourth when observable) was decreased by the same amount. The 

converted wave was observed to be very strong at the 60° incidence angle and 

was generally very weak at 45° and 75°. This observation is consistent with 

the fact that 60° corresponds to 30° incidence on the defect surface which is 

close to the limit angle for longitudinal wave generation. It should be noted 

that in the experiments conducted so far, the weldment crowns were machined 

off. In the presence of the crowns, the use of the second transducer is not 

obviously possible and the crowns may diminish or enhance the conversion into 

shear waves observed in pulse-echo according to their shapes. Experiments are 

planned to observe the effect of the weld crowns. 

CLASSIFICATION BY PATTERN RECOGNITION 

Using the classifier and the features extraction programs developed by 

Tektrend Inc., we have analyzed about 400 ultrasonic signals taken as 

mentioned above from the 4 samples, at 45°, 60° and 75 C , with a planar and 

with a focused transducer. The collected signals for each experimental setup 

(angle and transducer) were grouped according to their defect size and 

geometry, and they were further divided into two sets; one set of signals was 

used to train the classifier to recognize the signals based on their defect 

sizes and the other set was used to verify the performance of this 

classifier. The tests performed are indicated below with the recognition 

rates (expressed as percentages):

- 100 - 

1. Planar transducer: 

l.a. Signals from plate #3 and plate 

representative of a 3 mm flaw. 

lmm vs 2mm 

lmm vs 3mm 

2mm vs 3mm 

45 C 

91.67 

100.00 

75.00 

#5 are taken together as signals 

60 c 

83.33 

91.67 

91.67 

75° 

91.67 

91.67 

83.33 

l.b. The classifier was trained on signals from plate #3 and then tested 

with signals from plate #5. 

lmm vs 3mm 

2mm vs 3mm 

45° 60 c 

100.00 

100.00 

2. Focused transducer: (same tests as above) 

2.a Signals from plate #3 and plate #5 

representative of a 3mm flaw. 

lmm vs 2mm 

lmm vs 3mm 

2mm vs 3mm 

45° 

90.00 

93.75 

81.25 

80.00 

80.00 

75° 

80.00 

80.00 

are taken together as signals 

60 c 

80.00 

93.75 

75.00 

75 C 

85.00 

93.75 

81.25 

2.b The classifier was trained on signals from plate #3 and then tested 

with signals from plate #5. 

lmm vs 3mm 

2mm vs 3mm 

45° 60° 75° 

100.00 

78.57 

83.33 

83.33 

87.50 

71.43 

These classification results show that 100% recognition rates can be 

obtained in some cases. Also, they show that a resolution of lram in size 

which is required by fracture mechanics analysis is possible. Worse 

recognition rates are likely to be caused by a variation of the flaw size at 

the location where the data is taken. This problem will be resolved later 

when the samples will be sectioned. Also work is required to elucidate what 

are the most significant features for the planar transducer data as well as 

the focused one.

CONCLUSION AND PERSPECTIVES 

- 101 - 

We have reported preliminary work, on sizing of planar weld defects by 

ultrasonic echo feature analysis. With proper ultrasonic excitation 

diffracted edge waves were observed and found to be easily interprétable for 

sizing. However, in many cases, the echo pattern is complicated by other 

phenomena. In particular, this work has shown that the flaw cannot be 

considered as isolated since interactions with the sample boundaries produce 

extraneous features which have to be taken into account. We have also shown 

that classification of planar flaws according to size can be accomplished with 

high recognition rates and resolutions of = 1 mm. The approach presented 

therefore has a good potential for practical applications such as 

distinguishing between benign and critical flaws in pipeline weldments. 


This work has benefited from the technical assistance of Jean-Guy Allard and 

Claude Michel for making the samples and of René Héon for taking the data. 

Software development and classification work performed at Tektrend Inc. is 

supported by government contract it ISQ83-00025, jointly funded by CANMET, 

Energy, Mines and Resources Canada and IMRI, National Research Council of 

Canada. 

REFERENCES 

1. J.L. Rose, "A feature based ultrasonic system for reflector 

classification", in New Procedures in Nondestructive Testing 

(Proceedings), ed. by P. Höller, Springer-Verlag, Heidelberg, 1983, 

p. 239-250. 

2. M.F. Whalen and A.N. Mucciardl, "Application of Adaptive Learning 

Networks for the Characterization of two-dimensional and 

three-dimensional flaws in solids", Review of Progress in Quantitative 

NDE, 5th Annual report, p. 482-508, July 1980. 

3. M.F. Whalen and A.N. Mucciardi, "Inversion of physically recorded 

ultrasonic waveforms using adaptive learning network models trained on 

theoretical data", Review of Progress in Quantitative NDE, 4th Annual 

Report (July 77 to June 78), p. 341-367, January 1979. 

4. D.R. Hay and R.W.Y. Chan, "Identification of deformation mechanism by 

classification of acoustic emission signals", 14th Symposium on 

Nondestructive Evaluation, San Antonio, Texas, April 1983. 

5. R.P. Reed, H.I. MeHenry and M.B. Kasen, A Fracture-Mechanics Evaluation 

of Flaws in Pipeline Girth Welds, Welding Research council Bulletin # 

245.

Fig. 1 

1/JS 

- 102 - 

Echos observed from plate #3 (vertical planar defect, = 3 mm) 

refraction angle in the plate = 60°, 5 MHz focused transducer. 

1/iS 

Fig. 2 Echos observed from plate #4 (vertical planar defect = 2 mm), 

refraction angle in the plate = 60°, 5 MHz focused transducer. 

1//S 

Fig. 3 Echos observed from plate #6 (vertical planar detect = 1 

refraction angle in the plate - 60°, 5 MHz focused transducer. 

mm} t

Transducer 

(Pulse-echo mode) 

- 103 - 

Transducer 

Edge shear waves / Converted 

Back converted compression wave 

shear wave 

Fig. 4 Experimental set-up to observe diffracted and converted waves. 

Fig. 5 Oscilloscope traces observed with the transducers shown in figure 

4. Top: pulse-echo signal from angled transducer - (the first pulse 

is parasitic). Bottom: echos from transducer directly above flaw - 

the first pulse is a direct reflection from the top surface, 

following pulses are reflections of mode-converted wave inside the 

plate.

- 104 - 

CAPABILITIES/LIMITATIONS OF FOCUSED ULTRASONIC BEAMS FOR SIZING 

PIPELINE GIRTH WELD DEFECTS 

C.A. Kittrwi, J.8. Hallztt, V.N. Sycko 

Atomic Energy o {

- 105 - 

orientation) it is possible to compare it against a "tolerable" defect as 

determined by fracture mechanics which could safely be left in the pipeline 

with minimum risk of failure. This permits a realistic decision to be made 

as to repairs and avoids unnecessary rework of welds. 

Although ultrasonic inspection is capable of detecting significant defects 

(many of which are missed by radiography), defect sizing is still a major 

concern. Using standard transducers it is difficult to determine the size 

of girth weld defects by standard mapping techniques(3). This arises 

because the defect (or characteristic) being sized is comparable to or 

smaller than the beam used for mapping. The tendency therefore is to 

profile the beam rather than size the defect. Size predictions for a 

defect would consequently vary depending on the size of the beam used. 

One solution is to use a focused beam. This can be achieved simply and 

economically by attaching an acoustic lens to the face of a standard 

transducer. A test program was sponsored by the Pipeline Research 

Committee of the American Gas Association to investigate the capability of 

focused ultrasonic beams in sizing girth weld defects. 

APPROACH 

The main objective of the program was to investigate the defect sizing 

accuracy of an inspection procedure using focused ultrasonic beams with 

standard mapping techniques presently in use throughout industry. A 

secondary objective was to evaluate the technique as applied to field 

usage. Accordingly, the following approach was chosen: 

i. Use standard ultrasonic transducers which are readily available 

commercially "off-the-shelf". 

ii. Use contact inspection (as opposed to immersion) requiring various 

lens/wedge/probe combinations. 

iii. Use a simple scanning device (e.g. linear potentiometers to monitor 

probe position) so as to de-emphasize the equipment and concentrate 

on the ultrasonic technique itself. (See Figure 1 for the inspection 

equipment used). 

Lens Design 

Theory has been developed to design acoustic lenses for focused beams in 

defect sizing (4,5). The fundamental equations of interest are: 

1 - _ak 

N 

!|k _ 0.635 1-^1+0.2128 l-^l I [1] 

(Deviation + 9%)* 

* Deviation of analytical approximation from the true relationship

- 106 - 

V = f* - 0.9 

ak 

N [2] 

h. /^k)7__i 

N " l N / \1+ I 

It ^[ 1 + (1 -^' 

(Deviation + 5%) 



Ultrasonic theory also shows that the beam diameter within the acoustic 

focal range can be given approximately by: 

where fa^ = acoustic focus (note f ,/N = focusing factor, Z) 

fopt = optical focus 

Lg = the 6 dB working range 

AL = the part of L distributed before f ak 

V = gain (increase in sensitivity) 

N = transducer nearfield length 

The theory is relatively easy to apply, and focused beams can be readily 

reproduced (6,7) to meet the various inspection requirements. 

Weld Specimens 

The pipe specimens were in a range of sizes (610 mm diameter x 6.7 mm wall, 

to 1067 mm diameter x 15 mm wall thickness) typical of what is being used 

in the field. They were fabricated using gas-metal-arc (one sample) and 

shielded-metal-arc (seven samples) welding processes with representative 

beveled joint specifications. Defects were manually introduced by the 

operator during the welding process and included: inadequate penetrations, 

incomplete root fusions, lack of sidewall fusions, internal undercuts, slag 

inclusions (wagon tracks) and root cracks. They ranged in size from over 

50% wall thickness (root cracks) to a few millimetres (slag inclusions). 

Radiographic inspection carried out prior to ultrasonic inspection detected 

all but one of the manufactured defects (incomplete root fusion), but could 

give little information concerning defect sizes. 

[3]

ACOUSTIC LENSES 

- 107 - 

Ten acoustic lenses were designed to cover the range of wall thicknesses 

and weld preparations suggested for inspection. Selection of defects for 

inclusion in the program finalized the inspection requirements (inspection 

angle, defect location), and four lens/wedge/transducer combinations were 

then chosen to be used. Each assembly was calibrated by signal amplitude 

and time—of-flight using a calibration block containing 1.5 ram diameter 

side drilled holes spaced 2.5 and 4.0 mm apart (Figure 2a). The high 

resolution capability of focused beams is readily capable of 

differentiating between such small targets (Figure 2c) in contrast to the 

broad beam results for a standard transducer (Figure 2b). The calibration 

data was used to determine the beam inspection angle and time-of-flight of 

sound in the lens/wedge assembly. 

DEFECT SIZING 

Each weld specimen was examined ultrasonically from both sides of the outer 

surface weld bead using a simple scanning mechanism (Figure 1). The 

pulse-echo technique, which is commonly used in the field, was used 

throughout in contact mode with a light oil as couplant. The inspection 

data was recorded as multiple A-scans. In this technique, the probe is 

incrementally moved axlally away from the weld bead. The gate of the 

ultrasonic instrument is swept across the CRT screen, and the signal 

amplitude versus time display (shown on the CRT) is plotted for each probe 

position (see Figure 3). As the path length to the defect increases, the 

time-of-flight increases and the defect response on the A-scan moves to the 

right on the time axis. Conversely, any extraneous signal (e.g. from the 

lens/wedge assembly) remains at constant time and is readily 

distinguishable on the multiple A-scan record. The through-wall depth of 

the defect can then be calculated based on the extent of the defect 

response along the time axis. 

Metallurgical Examination 

After ultrasonic examination, each specimen was broken open by cooling it 

in liquid nitrogen and using three point impact loading to fracture along 

the centre of the weld. Where samples did not break through the defect 

(primarily for the smaller buried defects) the sample was sectioned and 

metallographically prepared to reveal the defect size. 

RESULTS 

A typical inspection record (Figure 4a) and specimen cross-section (Figure 

4b) are shown for a planar root crack giving a comparison of predicted 

versus actual defect size. These results are representative, as in all 

cases size predictions were in good agreement with the actual defect 

depths. This excellent correlation is further illustrated in Figure 5 

where the majority of data points fall within a band +1.5 mm of the line as 

determined by regression analysis. The 95% confidence limits for the 

predicted mean flaw depths are shown. There was little distinction in 

sizing accuracy between volumetric and planar defects. The technique is 

conservative in that defects tend to be slightly oversized (average error 

was 0.37 mm).

DISCUSSION 

- 108 - 

The results of this project illustrate that focused beams can be used 

successfully in conjunction with a simple mapping technique for defect 

sizing. Other, more sophisticated, techniques have shown comparable or 

better accuracy (3,8). The added complexity, however, does tend to isolate 

them to the laboratory with several years of development before being 

ready for field use, and require a higher qualification of inspectors than 

is presently employed for pipeline inspection. 

The project also restated standard problems encountered during contact 

inspection due to weld bead geometry. Although the leading edges of the 

wedge assemblies were ground off, the outer weld bead was sometimes still 

wide enough to prevent full coverage of a known defect. Depending on the 

nature of the defect and the working range limits of the lens, this could 

result in loss of sizing accuracy. However, by lens design for longer 

steel paths (i.e., in the region of 1 to 2 skips of the reflected beam, 

instead of 0 to 1 skip as used here) this can be readily overcome. 

Weld bead geometry can also confuse interpretation by introducing 

extraneous signals, or significantly altering the sound path length. When 

the ID weld bead is offset from the centre of the weld, sound may reflect 

off the bottom of the weld bead, instead of off the pipe ID, before 

reaching the target defect. In such instances an ID defect may appear to 

be alternately located at midwall or at the weld root, depending on the 

relative position of the transducer. Are there two defects located here, 

or one? Focused beam inspection is as vulnerable to operator 

interpretation in this area as any other technique which relies on 

time-of-flight information to perform its function. 

Transducer selection is also an area of vulnerability for this technique. 

The use of transducers with inappropriate beam characteristics would 

significantly degrade the outcome of any sizing attempt. Similarly, 

variations due to instrumentation effects (instrument/probe interaction) 

must be considered for optimum accuracy(6). At the extreme, the unfocused 

beam characteristics must be determined using the actual instrument/ 

transducer combination intended for the inspection for input to the lens 

design process. 

Although the basic ultrasonic theory is well developed, it does not readily 

lend itself to overcoming the specific limitations outlined above. In 

response to such needs, computer modelling is presently being used as an 

aid to understanding the physical processes involved. One such example is 

based on solution of the first order equations of elasticity using finite 

difference methods to model sound wave propagation through solids(9). The 

results of such modelling (e.g. see Figure 6) will be used to develop 

guidelines for optimization of transducer selection and inspection 

procedure specification. 

CONCLUSIONS 

- Focused ultrasonic beams can be used in conjunction with simple mapping 

techniques used in standard practice to give accurate sizing of 

pipeline girth weld defects.

- 109 - 

- The technique can be applied without the use of sophisticated support 

equipment and is readily field implementable. Operator interpretation 

is minimized through the use of a multiple A-scan format for recording 

inspection data. 

Contact inspection of girth welds using focused beams is subject to the 

same problems with weld bead geometry encountered during standard 

ultrasonic inspections. 

REFERENCES 

1 COOTE, R.I., FINGERHUT, M.P., 'Ultrasonic Inspection of Girth Welds', 

Proceedings of the International Conference on Pipeline Inspection, 

Edmonton, June 1983, pp. 263-279. 

2 GLOVER, A.G., COOTE, R.I., PICK, R.J., 'Engineering Critical Assessment 

of Pipeline Girth Welds', Fitness for Purpose Validation of Welded 

Connections, Paper 30, Welding Institute U.K., London, 1981. 

3 GRUBER, G.J., HENDRIX, G.J., SCHICK, W.R., 'Characterization of Flaws 

in Piping Welds Using Satelite Pulses', Materials Evaluation, Volume 

42, April 1984, pp. 426-432. 

4 WUSTENBERG, H., KUTZNER, J., MOHRLE, W., 'Fokussierende Prufkopfe zur 

Verbesserung der Fehlergro enabschatzung bei der Ultraschallprüfung von 

dickwandigen Reaktorkomponenten', Materialpruf, Volume 18, number 5, 

May 1976, pp. 152-161. 

5 SCHLENGERMANN, U., 'Schallfeldausbildung bei ebenen Ultraschallquellen 

mit fokussierenden Linsen', Acustica, Volume 30, Issue 6, 1974, pp. 

291-300. 

6 KITTMER, CA., 'Acoustic Lenses - Focusing in on Defects', Proceedings 

of the International Conference on Pipeline Inspection, Edmonton, June 

1983, 161-178. 

7 LARSEN, R.E., 'Zone-Focused Search Units for Bore Ultrasonic 

Examination Series', Proceedings of the Twelfth Symposium on 

Nondestructive Evaluation, San Antonio, April 1979, pp. 207-210. 

8. MACECEK, M., 'Precise Ultrasonic Depth Measurements of Internal 

Undercut in Pipeline Girth Welds', Proceedings of the Conference on 

Pipeline and Energy Plant Piping, Calgary Canada, November 1980, pp. 

291-299. 

9. DUNCAN, D., 'Computer Simulation of Ultrasonic Testing', To be 

published in Proceedings of the 5th Canadian Conference on 

Nondestructive Testing, October 1984.

- 110 - 

FIGURE 1: Equipment Used for Sizing Defects showing Ultrasonic 

Instrument, X-Y Recorder and Simple Scanning Mechanism.

TIME 

TRANSDUCER 

LENS 

WEDGE 

* 1.5 mm DIA. 

' SDH 

- Ill - 

NORMAL 

POSITION 

FOCUSSED 

FIGURE 2: (a) Test Calibration Configuration and Ultrasonic A-Scans 

Comparing Signal Response for (b) Standard Transducer Beam and 

(c) Focused Beam. 

POSITION 123456789 

I I II I I I I I 

DEFECT THROUGHWALL 

OEPTH, h = 3.23 lU-Uj COS a 

FIGURE 3: Schematic of Test Setup Illustrating the Muliple A-Scan Format 

for Recording Inspection Data.

FIGURE 4: 

- 112 - 

5 10 IS 10 

Depth From Inspection Surface (mm) 

(a) Typical Results of a Multiple A-scan Corresponding to the 

Right Hand Most Position as shown in Figure 4(b). 

FIGURE 4: (b) Cross-Sectional View of Planar Root Crack in 15 mm Thick 

Specimen showing Actual Versus Predicted Defect Sizes.

3 

(UIUI)1 DEPTH 

CTED Fl 

PREDI 

IUM 

MAXIM 

12 

10 

8 

6 

2 

- 113 - 

95% CONFIDENCE y 

INTERVAL ' ' 

Y = 0.735 • 0.879X 

O VOLUMETRIC DEFECTS 

• PLANAR DEFECTS 

2 l* 6 8 10 

MEASURED FLAW DEPTH (mm) 

FIGURE 5: Plot of Measured Versus Maximum Predicted Flaw Depth Using 

Focused Ultrasonic Beams for Sizing Purposes. Also shown is 

the 95% Confidence Interval for the Predicted Mean Values. 

The Defects were Generally Slightly Oversized with an Average 

Error of 0.37 mm. 

12

FIGURE 6: Output of Computer Modelling of Longitudinal Wave Entering a 

Test Sample and Reflecting off the Backwall. Note the 

Presence of a Surface Wave Travelling Across the Top of the 

Test Sample at Half the Longitudinal Wave Velocity.

- 115 - 

FLAW CHARACTERIZATION USING THE TIME-OF-FLIGHT METHOD AND 

ULTRASONIC FREQUENCY ANALYSIS 

V.K. Mak 

EnzKgij, Mitiei and Re.ioun.ce.i, Ottawa, Ontanio 

ABSTRACT 

The time-of-flight method and ultrasonic frequency analysis are valuable tools 

for measuring the sizes and shapes of defects. Experiment work was carried 

out using both methods simultaneously to measure the diameter and orientation 

of circular rods with diameters ranging from 2 mm to 7 mm. A 5 MHz broadband 

transducer was used as a transmitter/receiver. Each rod was immersed in a 

water bath and located in the far field region of the transducer. The difference 

in travel time (At) from opposite edges of the circular rod was measured. 

Sound waves diffracted from the circular scatterer were analyzed at 

the same time to yield a frequency spectrum and the spacing between consecutive 

frequency maxima (Af) was determined. The measurements were repeated by 

changing the orientation of the transducer, Both methods have the capability 

to determine independently the diameter and orientation of the circular rods. 

It was also shown that the product AtAf is approximately equal to 1. Time and 

frequency are mutually complementary variables. The two parameters can reinforce 

each other in terms of the information each one produces. A merger 

of the time-of-flight method and ultrasonic frequency analysis will provide a 

powerful nondestructive testing method in the future.

- 116 - 

"Events on earth consist in the interplay of two opposed forces, the yin and 

the yang, the female and male principles respectively . The interaction 

of these two forces together produce most of the universal phenomena. One 

yin and one yang together equal the Tao (the Great Ultimate)". 

- James K. Feibleman [1]. 

INTRODUCTION 

The ultrasonic examination of materials has played a major role in nondestructive 

testing. One technique for sizing defects is called the time-of-flight 

method [2]. It is based on measuring the travel time of sound waves diffracted 

or reflected from a defect, and has usually been used to measure crack 

depths in one dimension [33. Recently the method has been extended to twodimensional 

[t,5] and three-dimensional [6,7] cases. Another technique which 

has flourished is the ultrasonic frequency analysis method [8], This method 

makes use of the analysis of ehe frequency spectrum of the signal scattered 

from the defect and attempts to extract information about its size and physical 

properties. Both techniques have provided valuable information and have 

been used in research and development for the last ten years. An experiment 

is described below in which both techniques are used simultaneously to measure 

the size and orientation of simulated flaws. 

EXPERIMENTAL PROCEDURE 

Figure 1 is a block diagram of the time-of-flight method and ultrasonic frequency 

analysis system. The immersion technique is illustrated schematically 

with a single transducer acting as both puiser and receiver (pulse-echo). A 

Metrotek high energy pulser MP215 was used to excite the 5 MHz 1.27 cm diam 

transducer. The highly damped transducer emits a relatively broadband multifrequency 

pulse with strong frequency components from 1 to 9 MHz. A brass rod 

placed in water acted as an idealized flaw whose size and orientation could be 

easily controlled. The flattened smooth ends of brass rods with various diameters 

(Fig. 2) were used to scatter the broadband ultrasonic pulse. The signal 

was detected by a Metrotek wideband receiver MR106 and fed into a Metrotek 

stepless RF gate MG701 where the relevant signals were gated. Both the signal 

and the gated signal were displayed on an oscilloscope. The time-of-flight 

was measured by using the delayed sweep magnification feature of the oscilloscope. 

The gated signal was also applied to an HP8557A Spectrum analyzer 

connected to an HP853A Spectrum analyzer display which produced a display of 

the relative amplitude as a function of frequency on a cathode ray tube (CPT). 

The instrument had a measurement range of 10 kHz to 350 MHz. Adjustable 

center-frequency and bandwidth enabled any portion of a displayed spectrum to 

be examined in detail. 

A mechanical system provided accurate control of the orientation of the transducer 

and allowed its orientation to be measured to *0.1°. The distance 

between the transducer and the specimen (the water path ) was ~18 cm. The 

specimen was thus located in the far field region of the transducer. 

Each circular brass rod was positioned vertically at the bottom of the water 

tank. The transducer was set at a specific orientation. It was then X- and 

Y-scanned to provide an optimum signal from sound waves diffracted from the

- 117 - 

rod. After time and frequency measurements were taken, the transducer was 

rotated about the x-axis to a different orientation and was Y-scanned to 

optimize the signal. 

THEORY 

T .e geometrical theory of diffraction assumes that in addition to reflected, 

refracted and incident rays, diffracted rays are produced when a ray hits an 

edge [91. The incident ray produces an infinite number of diffracted rays, 

traveling in directions determined by the law of diffraction. This law stateo 

that each diffracted ray which lies in the same medium as the incident ray 

makes the same angle with the edge as the incident ray. Furthermore, the incident 

and diffracted rays lie on opposite sides of the plane normal to the 

edge at the point of diffraction. However, the diffracted ray need not lie 

in the same plane as the incident ray and the edge. Therefore the diffracted 

rays form the surface of a cone with its vertex at the point of diffraction. 

When a plane wave generated by a transducer is incident on a circular scatterer 

at an angle, the only sound wavelets diffracted back into the transducer 

originate from two opposite edges of the scatterer (Fig. 3). The wavelets are 

coherent sources that interfere at the face of transducer. Because these 

wavelets contain a broad range of frequencies, the condition of constructive 

interference will always be satisfied by some frequencies. From geometrical 

consideration, the frequencies of the wavelets that interfere constructively 

at the center of the transducer may be expressed as [10] 

f = 

n 

nv 

dsine + [D + Dd sin 9 + ^ à r' - [D - Dd sin 6 + d ] 1 /? (1) 

where n is an integer 

v = velocity of sound in the medium 

d = diameter of the circular scatterer 

D = distance of the circular scatterer from the center of the 

transducer 

9 = orientation of the axis of the transducer with respect to the 

axis of the circular rod. 

In the far field, d«D, Eq. (1) reduces to 

nv 

f n = 2dsine 

Let e = 8 + y 

where ß is the orientation of the axis of the transducer with respect to 

the vertical and can be measured and y is the orientation of the axis of 

the circular scatterer with respect to the vertical and is an unknown. 

Substituting Eq. (3) into Eq. (2) and rearranging terms 

2 _n cos S 

n 

= d cos y tan 6 + à sin 

(2) 

(3)

- 118 - 

now _n = Af (5) 

n 

where Af is the spacing between consecutive frequency maxima which is the 

slope of the plot of f vs n. 

Plotting 2hfa—ß" vs tan ß wil1 y ield a strai Sht line with 

gradient = d COSY (6) 

and intercept = d sinY (7) 

Thus, the diameter and orientation of the circular scatterer can be found 

from 

d = \ (gradient 2 + intercept 2 ) (8) 

-1 /intercept \ 

' I gradient / 

For d«D, it can be shown from simple trigonometry that 

VAt = 2dsine (10) 

where At is the difference in round trip travel time between the two opposite 

edges of the circular scatterer. 

Substituting Eq. (3) into Eq. (10) and rearranging terms 

VAt 

2 cos 

= d cos y tan s + d sin Y 

Plotting „ —— vs tan B will yield a straight line with gradient 

c. COS p 

and intercept again given by Eq. (6) and (7). Thus d and y can be found 

from Eq. (8) and Eq. (9). 

The methods described above thus show that ultrasonic frequency analysis and 

time-of-flight method can independently yield information about d and Y of 

a circular scatterer. 

Using Eq. (2), (5) and (10), it can be shown that 

AtAf ~ 1 (12)

Results 

- 119 - 

Figure *l shows the signals diffracted from opposite edges of a circular seatterer. 

Constructive interference of the sound wavelets can be seen lying 

between the two signals. Figures 5-10 show the frequency spectrum of sound 

waves diffracted from the 7 mm brass rod at different orientations. Plotting 

fn vs n, the slope yields if as illustrated in Figs. 11 and 12. A frequency 

analysis plot and a time analysis plot yielding d and y are shown in 

Figs. 13 and It. The actual diameters of the brass rod were measured with a 

pair of vernier calipers to ± .02 mm. The values are compared in Table 1. 

The axes of the brass rods were vertical (i.e., y = 0) in all cases because 

no mechanical system providing accurate control of y was available at the 

time of experimentation, but there was no loss of generality in the method 

presented here as experimentally, the transducer sees only the relative orientation 

of the circular rod with respect to itself. 

Figure 15 compares the diameters deduced from frequency analysis and time analysis 

with those measured from the vernier calipers. Figure 16 compares y 

deduced from frequency analysis and time analysis with the known y. The 

average of the error is ~1°. Table 2 illustrates that Ataf~l. The 5.02 mm 

diam brass rod was taken as an example. 

CONCLUSIONS AND DISCUSSIONS 

The time-of-flight method and ultrasonic frequency analysis have been used to 

measure the diameters and orientation of circular rods. Commercially available 

equipment has been used. In the present experiment, both methods yield 

the same information. However it is possible that in other cases, the two 

methods can yield information which complement each other. It should be noted 

that time measurement can be deduced from the rectified rf signal (Fig. 17). 

When the signal is rectified, a lot of information is lost. For unrectified 

rf signals frequency analysis can be used to find the bulk and surface properties 

of the flaws. 

Time and frequency are mutually complementary variables. The two parameters 

can reinforce and supplement each other in terms of information each one produces. 

The properties of a physical system can be described in terms of pairs 

of mutually complementary variables or properties. Other examples of complementary 

aspects are the position and linear momentum of a particle, the angular 

orientation of a system and its angular momentum, and so on. 

The time-of-flight method and ultrasonic frequency analysis have become important 

research tools in characterizing flaws for the last ten years. The merger 

of the two methods will become a powerful tool for nondestructive testing 

in the next twenty years. And hopefully, Canada will play an important role 

in their marriage. 


The author thanks Mr. M. Hafer for help with plotting the graphs.

- 120 - 

REFERENCES 

1. J.K. Feibleman, Understanding Oriental Philosophy, p. 86 (Mentor; New 

York 1976). 

2. M.G. Silk, Research techniques in nondestructive testing; Vol. 3; ed. 

R.S. Sharpe; p. 51 (Academic Press; London 1977). 

3. K. Date, H. Shiraada, and N. Ikenaga, NDT International V15, p. 315 

(1982). 

4. D.K. Mak, Physical Metallurgy Research Laboratories Report ERP/PMRL 

8*1-1(J), CANMET, Energy, Mines and Resources Canada (1984). 


84-3(J), CANMET, Energy, Mines and Resources Canada (1984). 


84-2KINT), CANMET, Energy, Mines and Resources Canada (1984). 

7. M. Macecek, K. Luscott, J. Wells and D.K. Mak, presented at the Fifth 

Canadian Conference on Nondestructive Testing, (1984). 

8. D.W. Fitting and L. Adler, ultrasonic spectral analysis for nondestructive 

evaluation; (Plenum Press; New York 1981). 

9. J.B. Keller, J. Appl. Physics Vol. 28, p. 426 (1957). 

10. L. Adler and H.L. Whaley, J. of Acoustical Society of America; V. 51, 

p. 881 (1972).

d (mm) 

Y (deg) 

d (mm) 

Y (deg) 

d (mm) 

Y (deg) 

d (mm) 

Y (deg) 

d (mm) 

Y (deg) 

d (mm) 

Y (deg) 

- 121 - 

Table 1 - Comparison of diameters and orientations of the brass rods 

measured from frequency and time analysis with the caliper 

values. All brass rods were held vertically 

Caliper 

values 

2.02 

0.0 

2.96 

0.0 

4.02 

0.0 

5.02 

0.0 

6.01 

0.0 

7.00 

0.0 

Frequency 

analysis 

1.78 

2.2 

2.99 

-0.3 

3.81 

0.8 

4.76 

0.4 

4.93 

2.7 

5.99 

2.1 

% error 

-12.1 

1.1 

-5.3 

-5.2 

-18.0 

-14.5 

Time 

analysis 

2.20 

-1.5 

2.92 

0.3 

3.94 

0.8 

4.27 

3.2 

5.93 

0.1 

6.77 

0.9 

Table 2 - The table shows that AtAf~l for all orientations 

The 5.02 mm diam brass rod was taken as an example 

8 (deg) 

6.4 

12.7 

18.9 

25.1 

31.1 

at (us) 

0.82 

1.60 

2.20 

2.80 

3.20 

if (MHz) 

1.29 

0.69 

0.47 

0.36 

0.30 

AtAf 

1.06 

1.11 

1.04 

1.00 

0.96 

% error 

8.7 

-1.4 

-2.1 

-15.0 

-1.4 

-3.2

PULSER 

IMMERSION 

TRANSDUCER 

WATER 

^CIRCULAR 

//ROD 

SYNC 

RECEIVER 

- 122 - 

INPUT 

OUTPUT 

TRIG/EXT 

, GATE 

GATE 

GATED 

OUTPUT 

DUAL TRACE 

EXT. TRIGGER OSCILLOSCOPE 

SPECTRUM 

ANALYZER 

1. Block diagram of the time of flight method and ultrasonic frequency analysis 

system. 

4 cm. 

2 cm. 

dem 

2 cm 

d~0.2.0.3,0.4,0.5,0.6.0.7 

2 Brass rods used in experiment. 

3. Sound waves diffracted from a circular soatterer. The incident plane 

waves are denoted by P.

- 123 - 

4. Signals gated for spectrum analysis. 

5. Ultrasonic frequency spectrum of signals reflected from the 7 mm diam 

brass rod, e = 0.0°.

- 124 - 

6. Ultrasonic frequency spectrum of signals diffracted from the 7 mm diam 

brass rod, e = 6.0°. 


brass rod, e = 12.5°.

- 125 - 




brass rod, 9 = 25.0°.

i 

I 

1 

i l 

- 126 - 

—j 



11. Plot of frequency vs n for 2.02 mm diam rod, S = 25.0°.

PLOT OF FREQ .vs. n CflSE 1- 3 

O , KNOHN - 7.000 MTI. 

ß - ia.800 DEGREES 

0 2 4 6 » 10 12 14 16 IS 20 

13. Ultrasonic frequency analysis 

plot for the 4.02 mm diam rod. 

- 127 - 

12. Plot of frequency vs n for 7.00 ram 

diam rod, B = 18.8°. 

FREQUENCY flNRLYSIS PLOT CflSE 1 

0.0 0.1 0.2 0.3 0.4 0.S 0.6 0.7

TIME RNFOSIS PLOT CRSF: 

- 128 - 

1H. Time-of-flight analysis plot for 

the 4.02 mm diam rod. 

0-0 1.0 2.0 3.0 *.O 5.0 6 0 7.0 8 0 

KNOWN DIFIMETER , MM. 

°- FREQUENCY RNHLÏSIS 

+ - TIME RNHLYSIS 

15. Experimental diameters measured from ultrasonic frequency analysis and 

time analysis plotted against the diameter measured from a pair of vernier 

calipers. Points in exact agreement would fall on the straight line with 

gradient = 1.

- 129 - 

°- FREOUENCï flNHLïSIS ! 

•- TIME HNHLÏSIS I 

16. Orientation of the circular rod measured from ultrasonic frequency analysis 

and time analysis plotted against the diameter measured from a pair of 

vernier calipers. As the circular rods are all held vertically, points in 

good agreement would fall on the x-axis. 

-4- 

A 

17. rf signals and rectified rf signals.

- 130 - 

THE INTRODUCTION OF REAL-TIME RADIOGRAPHY FOR THE INSPECTION OF 

BUTT WELDS IN OFFSHORE PIPELINES 

T.E. Re.ynold& In. 

Bfiown S Root Inc. 

Houston, Texas, U.S.A. 

This paper was presented at the 16th Annual OTC in Houston, Texas, May 7-9, 

1984. The material is subject to correction by the author. Printed with 

permission of the Offshore Technology Conference. 

1. ABSTRACT 

Real time radiography may significantly reduce the costs and efforts, improve 

the reliability, and eliminate some subjectivity arising from the conventional 

film radiographie inspection of butt welds in offshore pipeline 

production. This paper describes this state of the art technology and discusses 

its introduction to the production line of traditional pipe laying 

barges. 


Traditional offshore pipeline installation operations involve an assembly 

line process. Welding, inspecting, and coating the pipe occur simultaneously 

along the assembly line. New weld methods and pipe handling systems are 

being developed to speed this fabrication operation. In order to avoid 

having the inspection station become the slowest, and thereby the controlling 

station in the timing of this process, quicker inspection techniques must be 

developed. Because conventional film radiography is currently being pushed 

to its limits, other inspection methods were investigated. Furthermore, the 

current mood among pipeline contractors' clients is to remain with radiography 

and not stray into other non-destructive examination techniques such 

as ultrasonics. 

Real time radiography provides the required speed of inspection and helps 

solve the problem of subjectivity in interpretation. This paper is the 

result of efforts to bring a real time radiographie system to field testing 

on a pipe laying vessel. Potential manufacturers of these systems were 

contacted and encouraged to pursue development of this technology. Field 

testing is to be conducted once successful prototype demonstrations have been 

completed. This paper attempts to define the considerations to be made in 

real time radiography systems by describing the function of the systems and 

mentioning current applications of them. The current status of various codes 

governing pipeline inspection is reviewed with respect to their acceptance of 

real time radiography as a viable inspection method. 

Interpretation of radiographs can be a subjective art. Disputes arise among 

the three parties whose concern is weld quality: the client, the contractor, 

and the inspector. Limiting the range of interpretability of radiographs is

- 131 - 

a useful aspect of real time radiography. Completely automated inspection 

for the pipeline industry is possible in the future with real time radiography. 

3. BASIS FOR THE WORK 

The effort to product x-ray or gamma ray pictures in a real time fashion 

produced the fluoroscope, a device which converts radiation directly into a 

visible image. These are currently employed in some pipe mills around the 

country. Real time radiographie imaging is the instantaneous collection and 

conversion of x- or gamma radiation into an interprétable image. However, 

even with the advancement of technologies in image collectors, cameras, and 

displays, the fluoroscope alone has its limitations. In this paper, real 

time radiography will refer to computer enhancement of radiographie 

"television images" to perform weld inspection. 

When coupled with a computer, a fluoroscope type of system takes on new 

dimensions. Computer enhanced images expand the range of human perception 

and decrease the subjectivity of interpretation, while sacrificing only 

seconds in inspection time. Theoretically, computer enhanced radiography is 

just short of being real time; it takes a small amount of time for the 

computer to execute the functions that produce the enhanced image. The 

computer assigns values to each small picture element (pixels) which comprise 

the whole raw image. These values are assigned in terms of the location and 

intensity of the pixels. By performing algorithms on these values, image 

analyses can be performed so that scrutiny of the radiography can be 

facilitated. With smaller computers now having the capability to quickly 

process huge amounts of data, the gap between theoretical real time and 

actual image creation time is becoming negligible for inspection purposes. 

Advances in computer technology and simplification of computer programs has 

allowed for relatively inexpensive systems to be developed which meet the 

requirements of a laybarge operation. 

4. STANDARD FOR COMPARISON (Conventional Radiography) 

Real time radiography is applicable to the inspection of butt welds in 

pipelines only if it exceeds the following performance parameters of conventional 

radiography: 

1. Qualification of the image. 

2. Cycle time. 

3. Total overall cost. 

4. Operating reliability. 

5. Record keeping ability. 

6. Safety to personnel. 

A word on these parameters as they relate to conventional film radiography is 

necessary before continuing. Conventional film radiography refers to the 

inspection process in which a film media is exposed to x- or gamma radiation, 

chemically developed to form an image, and then interpreted to help ascertain 

weld acceptability.

5. QUALIFYING THE IMAGE 

- 132 - 

In comparing two methods of producing radiographie images, the basic consideration 

is the qualification of the image against a specified standard. 

Using a given exposure geometry or technique, an image must be qualified 

before it is interpreted. Qualified images meet sensitivity and density 

requirements; they are also properly identified for purposes of traceability. 

Sensitivity of the film is assured with the use of properly placed image 

quality indicators (IQI's). Placed next to the weld at specified increments, 

image quality indicators assure the inspector that image contrast sensitivity 

and resolution are adequate to visibly detect discontinuities in accordance 

with the governing standard. In the United States, the standard API 1104 1 

or ASME^ hole penetrameter is the image quality indicator. "2%-2T" sensitivity 

Is a common requirement of the film Images. The penetrameters are 

chosen with regard to pipewall thickness and weld reinforcement. "2%" refers 

to the thickness of the penetrameter as related to the thickness of the steel 

that the radiation must penetrate. When "2T" is specified, it is required 

that the "2T" hole in the penetrameter appears in the radiograpic image. 

In Europe, wire "DIN" image quality indicators are the standard. The ability 

to resolve wires of decreasing diameter on a given specimen relates directly 

to increasing image sensitivity. A lesser known image quality indicator is 

called the CERL IQI. It is composed of three parts; the first part is a 

flexible step wedge which is used to measure thickness sensitivity. The 

second and third parts are metal wires which are closely spaced and dimensionally 

graded in geometric progression. The unsharpness of the image is 

measured by the least discernable spacing between two wires. CERL IQI's 

might be best suited for measuring the resolution and contrast sensitivity of 

television type radiographie images, where the resolution and sensitivity are 

partially a function of the pixel density that comprises the image. 

The density or relative darkness of an image is a second consideration in 

qualifying a radiograph. Density is measured on a scale called Hurter- 

Dreffield (H&D) units. Lighter film obtains lower valued H&D units while 

darker film obtains higher values. For film radiography, the range of 

acceptable density falls approximately between 1.8 and 4.0. 

Most specifications require that identification of weld images appear on the 

radiograph so that the image'can be traced back to the weld. Improper identification 

can be cause for rejection of the radiograph. Currently, lead 

numbers and flashing are the primary means for identifying radiographs. 

6. INSPECTION TIME 

The inspection cycle time is the time commencing with the arrival of the pipe 

at the inspection station and ending with the final interpretation decision 

on weld quality. In conventional film radiography, the following sequence of 

events occur during the cycle time:

- 133 - 

1. Pipe stops at inspection station. 

2. Radiation source is positioned at the weld location. 

3. Film is wrapped around the joint. 

4. Film exposure is made. 

5. Film is transported to the dark room. 

6. Film is developed. 

7. Film is dried. 

8. Film is interpreted. 

The minimum cycle time for conventional radiography on pipeline operations is 

four minutes to read film wet and four and one-half minutes to read film dry. 

These numbers represent the physical minimum times using hot chemicals and 

high speed dryers while not sacrificing image quality. 

7. COSTS OF CONVENTIONAL RADIOGRAPHY 

The obvious cost items for provision of offshore film radiographie inspection 

services are labor, materials, and transportation. Based on estimates 

obtained from proposals submitted to Brown & Root during the past year, an 

average cost breakdown was developed for an example job. The chart is based 

on a single joint 30" 0 pipeline contract where 180 joints are radiographed 

on the firing line every twelve hour shift. This assumes a four minute cycle 

time. Total cost of the job, neglecting mobilization and demobilization, is 

figured at $53,800 when a 10 day duration of work is assumed. One major cost 

item in the job is the cost per joint which is 50% of the total cost of the 

work. This single item accounts for the expendable items involved in the 

work, film and processing chemicals. These figures were derived from formal 

telexed quotes for work from inspection companies for jobs in 1983. As pipe 

becomes smaller in size, the expendable costs become a smaller fraction of 

total operation costs. For any diameter pipe, the total day rate for labor 

and equipment remains constant. 

Some costs of film radiography are not so obvious but are substantial. In 

bidding work, the difference in cost between the requirement of reading wet 

film and reading dry film might be disparate. In bidding, the difference has 

translated to time and money lost. Another subtle cost factor is attributed 

to barge downtime caused by the inspection station. The time and cost of 

this factor have not yet been quantified. Also, radiographers do not know 

until at least four minutes after an exposure is made whether or not the 

radiograph is qualifiable. The costs of reshoots and crawler breakdowns 

comprise other unseen costs of film radiography. 

8. OPERATING RELIABILITY 

Mechanical and human operations are the main parameters which govern the 

operating reliability of conventional radiographie systems. Operating 

personnel control such functions as exposure time, chemical temperatures, 

placement of film in cassettes, lead shielding, record keeping, and 

interpretation. Exposure time is currently regulated manually and controlled 

with a timer. Film development has become almost fully automated and less

- 134 - 

CONVENTIONAL FILM RADIOGRAPHIC INSPECTION 

(Typical Costs in 1983 Dollars) 

Single joint operation. 30" 0 pipeline, 180 joints per 12 hour shift, 10 day 

duration of job, internal radiography. 

6 man radiographie crew 

(3 men per shift) 

2 Level II inspectors 

($400 each/day) 

4 Level I inspectors 

(§250 each/day) 

1 crawler technician on 

24 hour call ($400) 

Cost per joint (expendables) 

$1.00/llnear foot of weld 

($8.00 per joint) 

Cost of equipment (crawler 

facility, dark room, 

dryer, maintenance room) 

10 day job cost 

800 

1,000 

400 

$ 2,200 Labor 

$ 2,880 Expendables 

300 

$ 5,380 

$ 53,800 

Equipment 

Day Rate 

Note: This table does not account for transportation costs. 

subject to human variability. Shielding must sometimes be replaced to avoid 

scratches appearing on the radiographie images. Film is interpreted on a 

high intensity light viewer screen. The viewing area must be kept clean and 

dry. Film interpreters must stay alert throughout their shift. 

A major source of mechanical downtime is due to malfunction of the Internal 

crawler. Crawlers provide the radiation source when internal shots are made. 

They have been known to run down the line, lose power, and generally break 

down. 

9. TRACEABILITY OF RECORDS 

Accurate record keeping is required when radiographie weld examination is 

performed. The ability to trace the radiographie record of a weld back to 

the original weld and specific location on the weld is necessary. Currently 

lead number belts and the flashing of identification Information onto the 

film are common means of keeping identification information. Discontinuities 

and sections deemed unacceptable are noted on paper separate from the film. 

For retrieval of records, the paper record must be matched to the film 

radiograph. Also, with film processes running at such high speeds, the

- 135 - 

archivability of film or its ability to retain the original film record, is 

very questionable. 

10. SAFETY OF PERSONNEL 

Radiation safety must also be considered when comparing different systems. 

Currently, the total time of radiation exposure is calculated and safety 

areas are delimited so that personnel will not be overexposed. The total 

time of exposure at a given radiographie station on the laybarge during one 

day usually will not exceed two hours if 180 joints per day (4 minute cycle 

time) is assumed. This also assumes a 20 second exposure. Radiation safety 

zones are determined by marking off an area outside of which exposure will be 

limited to two milliroentgen per hour (2mR). 

11. REAL TIME RADIOGRAPHY 

11.1 System Description 

Real time radiography is the non-destructive examination procedure which 

collects radiation that penetrates a specimen, converts the radiation to an 

electronic image, and then visibly displays the image. The technology of 

real time radiography poses three problems: 

1• the physics of image formation, 

2. the computer management of information, and 

3. the application of the complete system to its environment. 

11.2 Physics of Image Formation 

The formation of real time radiographs follows the same principle as the 

formation of conventional film radiographs. Sources utilized for real time 

radiography are identical to those utilized in conventional film radiography. 

The quantity of radiation energies absorbed by the specimen results in the 

contrast density differences that form the image. Radiation sources have the 

same influence in the quality of real time image produced as they do in the 

quality of film image produced. Focal spot size and energy output directly 

influence geometric unsharpness and contrast sensitivity, respectively. 

The physics of producing a real time image is more complicated than producing 

a film radiographie image. X- or gamma radiation must be converted into a 

useable electronic image. Conversion of the radiation to an electronic image 

begins when x- or gamma radiation interacts with special phosphors, metals, 

or crystalline structures sometimes called scintillators. Scintillators 

convert x- or gamma rays into luminescence or electron charges that can be 

readily made into a visible image. Common equipment which convert x- or 

gamma ray energies into analog information are intensifiers, linear arrays, 

and charged coupling devices (CCD). New technologies are aiding in the 

development of miniaturized and solid state detecting components. Regardless 

of the converter, this analog information must be digitally addressed so that 

the computer can create and retain an image. The computer image is called a 

digital image because the computer has assigned numerical values to the 

densities (brightness) and locations of each picture element (pixel).

- 136 - 

11.3 Computer Management of Image Information-* »^ 

Radiographie images provide information beyond the human range of perception. 

The digital values assigned to each pixel in terms of location and intensity 

are combined and stored in matrix array as one complete image. 

Enhancement is never a substitute for good x-ray technique. Maximizing the 

signal (primary radiation) and minimizing noise (scattered radiation) are 

necessary; a high signal to noise ratio (S/N ratio) is desirable. Therefore, 

an unvarying technique that keeps scatter and geometrical unsharpness to a 

minimum is critical to obtaining a qualified radiographie image. In butt 

weld radiography of pipelines, the technique is held constant throughout the 

job. 

Once the raw digitized image is stored in the computer memory, the computer 

can begin to enhance it. The S/N ratio of the raw image depends not only or, 

the technique, but also depends on the efficiency of the detector. 

Brightness transfer functions can be applied to images so that different 

results occur. For example, if the intensity variation (film density 

variation) across an image falls within some limited narrow range, then that 

range can be expanded to encompass the human range of perception. The 

result is an image with broader intensity (density) range. Even though real 

time radiography provides positive images, the field can be reversed to give 

negative images which the film interpreter is accustomed to seeing. By 

relating intensity values to quantity of penetrating energies, the computer 

can calculate the thickness properties of the inspected piece. 

Sometimes the density variation across an image is variable. In the case of 

radiographing a pipe section using a film set flat against the pipe rather 

than rounded about the pipe, the edges of the developed film appear much 

lighter than the center. This is referred to as blowout. And, though image 

detectors are analogous to flat film, the computer can assign a function to 

the variation of this intensity across the image and, with enhancement, 

provide a consistent intensity across the image. This process is called 

gradient removal. A shortcoming of gradient removal might be that more of 

the image appears to be qualified because of the consistent density. 

Histograms relating the thickness of the weld to the distance along a line 

drawn across the weld can be drawn by the computer. Any discontinuities can 

be measured for length and width. By changing slightly the orientation of 

the source on the weld, it is also possible to locate and determine 

discontinuities in the third dimension. With such capabilities, three 

dimensional topographical radiographie images of a weldment can result. 

Computers excel at organizing volumes of information, so records management 

is greatly facilitated with real time radiography. The computer can record 

information such as operator identification and operating functions 

performed. The traceability of welds is made easy by the ability of these 

systems to record joint identification, location, and discontinuity data 

directly on the imaging media. A variety of media can be used to store 

digitized data. These include videotape, magnetic discs, and laser discs.

11.4 Fully Automated Inspection 

- 137 - 

Three major phases have been identified in moving toward fully automated real 

time radiography: 

1. The first phase would establish the feasibility of replacing 

conventional radiographie film with real time radiography while 

using the weld inspector to make all decisions regarding 

acceptability. 

2. The second phase would use the computer to scan over acceptable 

regions of the weld image and stop at questionable areas. The 

interpreter would be shown the areas and would be asked to make 

a decision regarding acceptance. Image enhancement features 

would be available to aid in the decision. 

3. Finally, fully automated inspection would involve having the 

computer make the radiographie acceptability decisions. The 

computer will compare the image against a specification criteria 

and make the "go" or "no go" decision. 

Laboratory real time radiographie systems are currently entering the second 

phase of development toward full automation. Enhancement of questionable 

areas within images is possible, and software being developed will enable 

computers to locate those questionable areas. 

11.5 Application of the System to the Production Environment 

The transition from the laboratory to the production environment presents the 

single development step that must be bridged before a real time radiographie 

system is placed on a laybarge. Although many systems are now capable of 

producing qualified images of thick, steel sections, the functional reliability 

and efficiency of the systems on a laybarge must be proven. The 

following is a list of some considerations to be made in judging the functional 

applicability of a system on a laybarge: 

Front end of the system. Any pipe mounting band or ring must be 

easily detachable to allow for passage of the pipe and miscellaneous 

joints, such as valves and tees, through the system. The band or 

ring must also be easily attachable and must allow for the image 

collector and, if necessary, the source of radiation to travel over 

or on it. Consideration must be made for concrete coating on the 

pipe. 

The image collector should be rugged enough to function in extreme 

climatic conditions and should not be affected by the heat of cooling 

evolved from the weld. Magnetic fields should not harm its function 

either. If double wall exposure, single wall viewing is employed, 

the radiation source and detector should track each other at 180° 

separation and should remain within strict tolerances.

- 138 - 

Back end of system. The computer facility must also be housed to 

endure the rigors of transportation and environmental extremes. A 

climate controlled, insulated housing facility will protect the 

computer system from these extremes. The computer facility must be 

free from failure after power outages and should not be affected by 

the mechanical vibration of operating laybarges. 

12. REAL TIME IMAGE QUALIFICATION 

Real time radiographie images should be qualified in the same ways as film 

images. The ultimate factor should be the interpreter's ability to discern 

the required portion of the image quality indicator. In contrast to film 

radiography, consideration should be made to the "graininess" of the digital 

image. Because these images are comprised of pixels, the spatial resolution 

is higher when there are more pixels comprising a unit area of the image. If 

conventional image quality indicators prove insufficient to determine image 

sensitivity and contrast resolution, it may be necessary to use the CERL 

image quality indicator or one of tailored design. 

13. REAL TIME RADIOGRAPHIC INSPECTION CYCLE TIME 

In real time radiography, the sequence of events involved in an inspection 

cycle is: 

1. Pipe stops at inspection station. 

2. Radiation source is positioned at the weld location. 

3. Radiation detector is positioned at the weld. 

4. Radiographic image is collected and interpreted simultaneously. 

Based on demonstrations provided by manufacturers of these real time radiography 

systems, it is estimated that image collection will require, at most, 

five seconds per diameter inch. Total cycle time will be a little more than 

one minute for 8" 0 pipe, and a little more than three minutes for 40" 0 

pipe. This allows 30 seconds for positioning of the source and mechanical 

equipment. There is also the possibility of mounting more than one source 

and detector on the pipe. In turn, the cycle time would decrease by a factor 

equal to the reciprocal of the number of detectors. 

14. REAL TIME RADIOGRAPHY COSTS 

Pipeline contractors and owner companies traditionally contract inspection 

work to third party inspection companies. If these companies will be 

required to invest in the equipment that comprises a real time radiographie 

system, then they will have to recover the cost of their investment through 

charges to their clients. 

One such scenario might be the following. Given that real time radiographie 

systems range in price from $60,000 to $200,000 for a laybarge type system; a 

price of $150,000 will be assumed. The inspection contractor would like to 

recover the cost of the system, plus $30,000 each for interest and profit. 

The inspector estimates that the useable life of the system will be three 

years, and it will be used 60 days per year. The total cost which must be 

recovered is $210,000. Based on the assumptions stated, the inspection

- 139 - 

contractor will have to charge a day rate of $1,170. If another twenty percent 

is added for equipment spares, the total day rate for equipment totals 

$1,400. The cost of video and computer recording media in relation to the 

cost of the expendables has been determined to be about $0.25 per linear foot 

of weld. Labor costs are assumed to remain constant. The trend will be to 

use less, but more technically trained personnel; therefore, the cost tradeoff 

will balance. The following chart is included for comparison to the 30" 

0 job described for conventional radiography. 

REAL TIME RADIOGRAPHIC INSPECTION 

(Estimated Cost in 1983 Dollars) 

Single joint operation. 30" 0 pipeline, 180 joints per 12 hour shift, 10 day 

duration of job, internal radiography. 

Labour (same as conventional film $ 2,200 Labor 

radiography) 

Cost per joint (expendables) 

$0.25/linear foot of weld 

($2.00/joint) 

720 

Equipment and spares 1,400 

10 day job cost 

$ 4,320 

$43 ,200 

Expendables 

Equipment 

Day Rate 

Although the above is a hypothetical presentation of how an inspection company 

may charge for work undertaken with a real time radiographie system, it 

points out two factors: 

1. The cost of real time radiography might be essentially the same, 

or might be greater than conventional film radiography, depending 

on the cost of expendables and size of the pipe. 

2. Where the daily cost of operation of a laybarge may range from 

$75,000 ($3,125/hour) for a small spread in the Gulf of Mexico 

to $250,000 ($10,420/hour) or more for a large operation in the 

North Sea, total inspection cost for a whole job can be returned 

in a few hours of increased reliability. Pipeline contractors 

may want to consider the purchase of a real time radiographie 

system and the subcontracting of inspection work performed with 

that system. The return on investment could be substantial. 

15. OPERATING RELIABILITY 

The reliability of real time radiographie systems is more dependent on 

mechanical devices than on personnel. Film and proper screens do not have

- IAO - 

to be loaded properly into cassettes; no wrapping of the cassette around the 

weld joint is necessary. Only the front end mechanical system has to be 

latched to the joint. In real time radiography, the image is collected 

automatically, scanning rate and radiation output are preset. Once the 

scanning begins, the operator has only to begin interpreting radiographie 

images, a qualified image is assured by presetting and establishing a 

standard radiographie technique. Only minor adjustments may be required 

during a shift; any adjustments or enhancement features used are documented 

alongside the image. Once scanning of the weld is completed, the image 

collecting system is removed from the weld to await the arrival of the next 

joint. 

Reliability becomes a function of the ability of the equipment to endure the 

nonstrop operation on the laybarge. This reliability will be ascertained in 

field testing. 

To further increase the reliability at the inspection station, investigation 

into eliminating the internal crawler is being conducted. One possibility 

that exists with real time radiography is the double wall exposure, single 

wall viewing (DWE/SWV) of large diameter pipes. With a fractional sacrifice 

in image collecting time, internal crawlers might be eliminated from use by 

placing the source diametrically opposite from the image collector. Though 

size limits of this technique have not been established, demonstrations have 

shown qualified real time radiography images using the DWE/SWV technique on 

28" 0 x 0.68" W.T. pipe. 

16. TRACEABILITY 

Information about joint identification, discontinuities observed, enhancement 

features used, operator identity, and time and date can be stored alongside 

the radiographie image. Rather than have a dual system of paper and film 

records, all records for real time radiographie systems are stored on one 

medium. To locate repairs, the real time radiography system can refer to any 

location on the weld in terms of degrees or linear distances. Hard copies of 

images can be made on polaroid film and paper, although they lose some 

clarity. If desired, radiographie images can be transmitted, with minimal 

lag time, to land based operations via phone or satellite link. The 

archivability of the images is a function of only the imaging medium, not of 

the development technique as it is in conventional film radiography. 

17. SAFETY 

Although the time of radiation exposure is substantially increased with the 

use of real time radiographie systems, demonstrations have shown that the 

radiation intensity required to develop an image will be less than that 

required by conventional film radiography. Coupled with the fact that the 

DWE/SWV technique will use a collimated radiation beam, rather than the 

panoramic beam of conventional film radiography, radiation will pose no new 

health hazards to personnel. Safety zones within the 2roR range must still be 

delimited.

- 141 - 

18. EXISTING APPLICATIONS OF REAL TIME RADIOGRAPHY 

Fluoroscopic inspection of pipelines has been ongoing for years in many pipe 

mills around the United States. Kaiser Steel Corporation has achieved 

qualifiable radiographie weld images on pipelines up to 36" 0 in their Napa 

Valley Plant using an unenhanced real time radiography system. They claim 

that the savings generated by using this sytem for one and one-half years was 

more than $750,000, or more than twice the cost of the system. Kaiser also 

sees a great reduction in record storage space with use of video tape.5 

Arco Alaska, Inc. desired to assess and quantify corrosion damage in some of 

their insulated crude oil flow lines at Prudhoe Bay. They measured corrosion 

using a consistent radiographie technique and relating film density (or 

radiation penetration) directly back to the quantity of corrosion pitting. 

When they applied real time radiography to their problem, they increased 

their inspection rate from 27 feet of line per day to 480 feet. Total 

inspection cost per foot of pipe inspected was reduced 90%.6 

The primary application of real time radiography today is in ordinance 

inspection. Defense related manufacturing requires high quality assurance. 

To meet these demands, defense contractors have manufactured sophisticated 

radiographie systems which can inspect such products as solid fuel in 

missiles for porosities or honeycomb sections of jet wings for minor 

deformities. With cost not being a major consideration in utilizing these 

systems, the resolution and sensitivities obtained are unsurpassable. 

19. STATUS OF GOVERNING CODES 

API 1104. At its annual meeting in September, 1983, the API 1104 governing 

committee considered amending its code to allow for imaging systems other 

than film systems to obtain radiographs. Subsquently, a subcommittee met and 

reworded Section 8.0 of the code to allow for other imaging systems. The 

rewording will be put to a vote before the full committee at the 1984 annual 

meeting. Any revised standard will probably not appear until 1985. For now, 

Section 5.2 of the code states, "Nondestructive testing may consist of 

radiographie inspection or another company specified method". This is the 

sentence in the code which implies that real time radiography inspection 

systems might be used if they are specified by the owner company. 

BS 4515. The BS 4515 committee is currently considering a rewritten draft of 

the complete specification with inspection and radiographie testing being 

subsections under it. With regard to draft section 25, "Inspection and 

Testing", and draft section 26, "Radiographie Examination", both are written 

with a generalized wording so that the use of real time radiography is not 

prevented. BS 4515 refers one to BS 2910 for radiographie 

techniques.''»® 

DNV. DNV has no current intentions to rewrite its 1981 code. In its present 

form, the DNV code is slanted more toward conventional film radiography and 

gives little leeway for real time radiography."

20. OPERATIONAL UTILIZATION 

- 142 - 

Film interpreters do not, at present, have tools available with which to 

enhance conventional film radiographs. It is envisioned that the main 

functions of real time radiography will be to inspect more quickly with 

higher quality and reliability. If real time radiographie images qualify 

according to specifications, no enhancement schemes need be used. The 

inspector will continue to interpret images the same as he did with 

conventional radiography. If questions arise, enhancement might help answer 

them. It is not foressen that special enhancement functions will be 

performed during normal operations• The keys to the real time radiography 

system will be to avoid a bottleneck in production and to provide a better 

system with which to perform radiography. 

21. CONCLUSION 

This paper has outlined the basis for comparison of conventional film 

radiography to real time radiography. Technical evaluations of existing real 

time radiographie systems have shown that it is feasible to bring a 

laboratory system onto a laybarge and achieve results which will outshine 

conventional radiography. Quantified comparative data will result when field 

testing is complete. Field testing will also help the offshore industry warm 

to the idea of real time radiography as a viable non-destructive weld 

examination technique for offshore pipelines. 

21.1 Acknowledgements ; 

The author wishes to thank the management of Brown & Root for continued 

support of this project, and Mr. James H. Walker for his contagious 

enthusiasm and guidance from the inception of this work. 

22. REFERENCES 

1. "Standard for Welding Pipelines and Related Facilities", American 

Petroleum Institute, Fifteenth Edition, 1980. 

2. "Standard Method for Controlling Quality of Radiographie Testing 

SE-142", American Society of Mechanical Engineers, Section V, Article 

22, "ASME Boiler and Pressure Vessel Code", 1983. 

3. Jacoby, M.H., "Image Data Analysis", Section 15, The Nondestructive 

Testing Handbook on Radiography and Radiation Testing, 1982. 

4. Pratt, W.K., Digital Image Processing John Wiley and Sons (1978). 

5. Cutting, C.C., "X-Raying Pipe on the Line", Welding Design & 

Fabrication, April 1983. 

6. Hill, D.E. and Galbraith, J.M. "Radiographie Surveys for the 

Evaluation of Corrosion Damage in Crude Oil Lines in the Eastern 

Operating Area of Prudhoe Bay", paper 52, presented at the 

International Corrosion Forum, Anaheim, California, April 18-22, 

1983.

- 143 - 

7. "Specification for the Process of Field Welding of Steel Pipelines", 

British Standards Institute, draft standard of BS 4515, July 1983. 

8. "Methods for Radiographie Examination of Fusion Welded 

Circumferential Butt Joints in Steel Pipes", British Standards 

Institute, BS 2910, 1973. 

9. "Rules for Submarine Pipelines", Section 10, Det Norske Veritas, 

1981.

% 

-•'v/ FILM AND SCREENS 

' \ INSIDE CASSETTE 

INTERNAL MECHANICAL CRAWLER 

PROVIDES X- OR GAMMA RAY 

SOURCE 

IMAGE COLLECTING 

A J 

f ^\ Uj I / DEVICE 

EXTERNAL RADIATION SOURCE 

FOR DOUBLE WALL EXPOSURE 

DARKROOM FACILITY 

. FOR RAW AND 

ENHANCED 

IMAGES 

MINICOMPUTER 

(HARDWARE AND 

SOFTWARE IMAGE 

MANIPULATIONS)

- 145 - 

CONTINUOUS ACOUSTIC EMISSION MONITORING 

J.S. Mitchell 

Viatec Ra-ioufice Sij.bte.mt> Inc., Calgaiij, Albe-ita 

M.SI. Baiiim 

Unive-ï-Sity o f 6 Manitoba. 

Winnipeg, Manitoba 

1. Abstract: 

A novel concept that essentially reconfigures the architecture of current 

acoustic emissions (A/E) monitoring instruments has been developed and is 

now the focus of prototype construction. This system will enable the 

continuous A/E monitoring of structures with enhanced reliability. 

Applications include integrity monitoring of large structures (pressure 

vessels, pipelines, etc.) and process control monitoring (valve operation, 

catalyst performance, etc.). Each application requires a thorough 

understanding if the metallurgical and acoustic regimes operating in the 

structures of interest. With this information the A/E monitor is 

programmed to observe and to some extent identify significant acoustic 

emissions, and present this information to personnel otherwise not skilled 

in interpretation of A/E data. 

2. Review: 

The present method of acoustic emission monitoring of large structures is 

primarily based upon equipment developed within the past 15 years. This 

development was enabled by the availability of rugged peizo-electric 

transducers, and the advent of "portable" mini computers. There have been 

some improvements with this equipment (primarily in the software areas) but 

the overall configuration has essentially remained unchanged, see Figure 1.

- 146 - 

By design, these A/E monitoring equipments have been targeted toward 

inspection type applications. These are typically short term monitoring 

tasks that attempt to identify (and in some cases locate) the presence of 

active defects in a structure undergoing some form of stress. This 

equipment is versatile in that they may be used to monitor several 

structure types (pressure vessels, piping, bridges, cranes, etc). However, 

with few exceptions, the operation of the equipment and interpretation of 

the results requires the constant attention of a skilled operator. These 

skills range from an understanding of electronic instrumentation and their 

software operating systems, to metallurgical damage mechanism that are 

active in the structure being monitored. Hence, the A/E monitor operator 

is an incredible individual. Their experience on a given structure is 

acquired over a relatively short term test. Also, they are asked to 

interpret A/E data without the benefit of learning what affect the service 

history of a structure may have had on its acoustic emission behavior. 

This is further complicated if the structure is monitored while being 

stressed under conditions that do not replicate the service environment 

(eg. hydro-static pressure testing of pulp digesters). 

Where this equipment has been used for continuous monitoring of large 

structures (eg. offshore platforms in the North Sea) the long range 

transmission of analog signals appeared as one problem that affects the 

overall A/E monitoring reliability. This hardware problem has several 

remedies, one of which is to improve the packaging, connectors and to 

transmit the A/E signal on a frequency modulated carrier. The overall 

system configuration is not changed by this, but it is an improvement in 

reliability. 

3. Present Concept: 

The present concept for the continuous A/E monitoring of large structures 

borrows heavily from seismological techniques. As shown in Figure 2 the 

concept can be described in terms similar to the operation of several 

seismic observatories co-ordinated by a central seismic analysis centre. 

In addition, the type of data processing performed by seismologists may 

also prove beneficial for the interpretation of A/E signals (ie. assessing 

the relavence of A/E signals amongst background noise). Aside from this 

concept, the methods of A/E signal analysis presently being investigated 

is not possible on conventional A/E monitoring equipment; nor was it 

intended to be.

- 147 - 

Figure 3 shows a block diagram of the prototype system currently under 

development. This system may be considered a redeployment of the 

components that make up an advanced conventional multi-channel A/E monitor; 

but the advantages that accrue to continuous monitoring of large 

structures, to advanced signal analysis, and to reduced system hardware 

cost are substantial. 

Similar to seismic laboratories located about the earth, this system uses 

acoustic emission surveillance units located about a subject large 

structure. The surveillance unit is capable of a variety of A/E signal 

processing tasks. The current prototype is designed to measure six A/E 

parameters that are dealt with during conventional A/E analysis. With the 

advent of digital signal processing techniques, subsequent units will have 

a variety of enhanced capabilities enabled by this technology. 

A surveillance unit is placed in close proximity to each A/E transducer. 

In some applications the transducer may be installed within the 

surveillance unit enclosure, thus simplifying installation and minimizing 

the risks to low level analog signal transmission. The surveillance unit 

consists of three electronic subsystems. These are: 

a) Analog section for A/E signal processing; the 

b) Digital section for measuring A/E transient parameters; and the 

c) Microprocessor section which controls the function of the surveillance 

unit, may perform some A/E data processing and storage, and transmits 

this data to a central control point. 

Data transmission is possible by several mechanisms, the most common using 

wires to conduct electronic signals. This concept is not restricted to 

hard wired systems, but the prototype under development will use a four 

wire Modem type data link. 

The control point for a network of surveillance units will vary between a 

simple status indicator display, to mini computer based data collection and 

analysis facilities. Generally, as the complexity of the structure 

increases, or as the number of variables that affect the acoustic emission 

regime increases, the required computing facility will increase. To 

illustrate this, the required central computing facility to monitor the 

operation of several valves would be less complex than that to monitor the 

acoustic emissions from an offshore platform (the number of A/E transducers 

being the same in both cases). 

This equipment deployment scheme essentially supports the two levels of 

software necessary when establishing the interpretive functions of a 

continuous A/E monitoring system. It is the interpretive function that:

- 148 - 

a) enables an A/E monitor system to be a part of, or stand alone as, a 

structural integrity monitoring system; and/or 

b) permits A/E monitoring for process control operations. 

Hence, unlike conventional inspection type A/E monitoring, the source and 

characteristics of the acoustic emission within a subject structure must be 

learned before a continuous A/E monitoring is commissioned. Further, once 

continuous monitoring begins, this interpretive function is further refined 

in recognition of unique service or operational environments. This can 

also be expressed as allowing experience while monitoring a subject 

structure (learning curve) to improve the reliability of results. These 

results (or output) of the system are presented to otherwise unskilled A/E 

operators, or as control signals to other equipment. In addition, the 

results of this form of continuous monitoring may be interfaced with other 

monitoring functions. The most promising of which has been vibration 

analysis; but temperature, pressure, strain, etc. can also be incorporated. 

4. Application: 

From a technical point of view, the number of applications for continuous 

A/E monitoring are considerable. Three applications are currently under 

investigation, and at least one other will be undertaken when resources 

permit. The structures investigated to date have been a cross country 

transmission pipeline, an inlet manifold to a thermal power boiler, and a 

semi-submersible drilling platform. In each case, a metallurgical study is 

conducted under laboratory conditions using the materials (steel types and 

damage mechanisms) that make up the subject structure. Also, acoustic 

emission data is collected from the structure, and this information is 

combined to program the microprocessor supported surveillance units and the 

computer based controller. The pipeline application has involved field 

testing of the system hardware, and preliminary results have been very 

positive. 

Other structures that are readily applicable to continuous A/E monitoring 

are: 

a) Plant wide piping systems (steam lines); 

b) Pressure vessels (pulp digestors, mixing vessels, packed reaction 

towers); 

c) Boilers and heat exchangers (deairators, manifolds);

d) Offshore structures (subsea pipelines, production platforms). 

This is not a complete list, but these applications satisfy both the 

criteria for large structures and continuous monitoring for either 

integrity or process control. Applications such as ice breaker monitoring 

have been considered, but will require a significant effort in acoustic 

emission signal processing and analysis before its feasibility can be 

established. 

5. Summary: 

A continuous acoustic emission monitoring concept has been described. The 

theme of continuous monitoring is to provide information regarding the 

integrity of a structure, or provide feedback for process control, or 

both. The acoustic emission behavior of each structure would be determined 

through a study of materials, damage mechanisms, and service conditions. 

With this information the continuous A/E monitoring system may be 

configured to present results to personnel otherwise unskilled in A/E data 

interpretation. The system hardware has been developed to prototype form, 

and essentially represents a redeployment of conventional, advanced multichannel 

A/E instrumentation. However, the seismological resemblance 

discussed above is a better model of the system and its operational 

concept. 

6. Acknowledgements: 

We wish to acknowledge the support of Petro-Canada Resou-ces with many 

aspects of this concept; and the encouragement and logistics support of 

Ontario Hydro Research Division, and Petro-Canada Exploration.

- 150 - 

OTHER . 

CHANNELS 

CONVENTIONAL Mil LTI- CHANNEL 

A/E MONITOR 

Figure 1: Hardware Deploynent of a 

Conventional A/E Monitor 

CONTINUOUS SEISMIC MONITORING 

• SEISMIC OBSERVATORY 

(77| SEISMIC ANALYSIS CENTER 

Figure 2: Seismic Model of A/E Monitoring 

/£ SIGNAL 

OUTPUT 

MINI 

coup.

R*NSoucE ft 

SURVEILLANCE 

UNIT 

- 151 - 

MULTI-DROP 

OATA LINK TO 

OTHER S.U.'« 

CONTINUOUS A/E MONITOR 

Figure 3 Hardware deployment of 

Continous A/E Monitor 

COMPUTER 

COM TROLLER

DAY 2 

KEYNOTE ADDRESS: Advanced Radiography for Transportation and Energy Systems 152 

- H. Berger 

NDT - Radiographie Processing Quality Control, Making the Transition to 163 

Automatic Processing in Radiographie NDT, Sensitometric Values for Industrial 

Radiography 

- W.E.J. McKinney 

Recent Microfocus X-Ray Imaging Applications (abstract only) 176 

- R.S. Peugeot 

Wet Channel Inspection Systems for CANDU Nuclear Reactors....CIGAR and 178 

CIGARette 

- M.D.C. Moles, M.P. Dolbey, K.S. Mahil 

An Advanced Heat Exchanger Eddy Current Inspection System 194 

- M. DeVerno, H. Ghent, H. Licht 

Wet Channel Measurement of Pressure Tube to Calandria Tube Spacing in CANDU 207 

Reactors 

- J.H. Sedo 

Checking for Cracks in Gas Turbine Rotor Discs 222 

- J. van den Andel, A.B. Nieberg 

Eddy Current Inspection of Mildly Ferromagnetic Tubing 235 

- W.R. Mayo, J.R. Carter 

On the Relation Between Ultrasonic Attenuation and Fracture Toughness in Type 250 

403 Stainless Steel 

- F. Nadeau, J.F. Bussiere, G. Van Drunen 

Acoustic Emission Testing of Man-Lift Devices 262 

- J.A. Baron 

Acoustic Sorting of Grinder Balls 276 

- F. Nadeau, J.F. Bussiere 

The Benefits of NDT Training for Canadians 289 

- L.B. Manzer 

Fear Detection and Removal - The Psychological Implications of the Technological 297 

Age (abstract only) 

- T. Helliwell 

Certification of NDT Operators in Canada - An Update 298 

- V. Caron

VAV TOO

- 152 - 

ADVANCED RADIOGRAPHY FOR TRANSPORTATION AND ENERGY SYSTEMS 

Haiold BzKge.1 

lndui,t?i.Lal Quality, Inc. 

Gaithe.iibu.iQ, UV, U.S.A. 

ABSTRACT 

Radiography is a code-approved NDT method for inspections involving pipelines, 

pressure vessels and many other products. Radiography provides several advantages 

for inspection including image quality indicator (IQI) sensitivity of 

2% or better, and an image for ease of interpretation and long-term inspection 

record retention. New advances in radiographie testing are expected to broaden 

the use of radiography even more. One example involves improved real-time 

radiographie methods; these are now demonstrating an IQI sensitivity of 2% 

or better. Real-time radiography is being applied in areas that include cast 

aluminum automobile wheels, aircraft maintenance and pipeline welds, as examples. 

Microfocus radiographie equipment is available in energies up to 160kV 

in both conventional and rod anode configurations. The geometric magnification 

one can achieve improves the radiographie results by displaying extremely 

small discontinuities and by the improved contrast that results from the distance 

between object and detector. Applications in inspection of composites 

and ceramics have been particularly useful. The rod anode equipment has been 

applied to welds in small tubing such as used in heat exchangers. Relatively 

low energy radioisotope sources such as ytterbium-169 have also been used in 

similar applications. A notable advance in the energy field is the development 

of a portable linear accelerator for in-service inspection. This new 

Minac, a 3.5MeV machine whose radiation head weighs about 100 kg, has been 

used in several electric utility inspections for components such as pumps and 

piping. A new smaller radiation head, the Shrinkac, has been particularly useful 

for piping inspections where access is often limited. Neutron radiography 

is now available for on-site industrial applications because of developments 

in transportable sources. Corrosion detection in aircraft and other maintenance 

inspections are attractive applications for this novel radiographie method. 

These examples illustrate that there are many new concepts available in 

radiographie testing and that many of the applications for this new technology 

are in the transportation and energy fields. 


Inspection requirements for the transportation and energy industries are often 

critical. Lives and safe use of property depend on the performance of systems 

in fields such as aerospace, rail, highway, ship and pipeline transportation 

and the production and distribution of energy from nuclear and fossil fuel 

plants. It is often these industries that lead the way in the use of advanced

- 153 - 

nondestructive testing (NDT) equipment and procedures. This report includes 

discussions of some advanced radiographie methods that have been applied or 

show promise in these industries [1]. 

Current methods for the use of radiographie inspection [2,3] mainly involve 

X-ray film. These methods yield radiographs that typically display penetrameter 

or image quality indicator (IQI) sensitivity [A] values of 0.5 to 2 percent. 

The films themselves provide an inspection record that can be retained; 

in some cases such as nuclear power, the inspection record retention requirement 

may be 40 years or more. 

Some new radiation methods that are coming into broader use, particularly in 

the transportation and energy industries are electronic real-time detection 

methods, improved radiation sources and the use of other radiations such as 

neutrons. The following sections expand on these ideas. 

REAL-TIME METHODS 

Fluoroscopy to present a prompt picture of an X-ray image has been used for 

many years for both medical and industrial applications. Television methods 

to bring these radiation images to a safe, remote location and provide electronic 

control of brightness and contrast prompted widespread medical use beginning 

in the 1950's. Industrial interest in these systems clearly existed 

but broad use of real-time methods is becoming a strong factor in industrial 

inspection only in this decade [5]. 

Improvements in radiation sensitive screens and image intensifier tubes have 

been factors in this increased interest in real-time radiography [5-10], Better 

electronic methods for handling and retrieving data have also contributed 

to increased use of electronic-based radiation inspection systems [5,11]. 

Although capital costs for real-time systems can be relatively high, the promise 

of efficient, relatively low-cost operation has contributed significantly 

to increasing industrial use of real-time methods [7]. 

Modern real-time X-ray inspection systems are capable of presenting images that 

display contrast sensitivity in the 1-3% range and resolution in the 2-5 line 

pairs (lp)/mm range. Applications have included inspections of assemblies, 

castings, welds, electronic components and a variety of parts and materials 

[12], Some specific examples pertinent to this report include the real-time 

X-ray inspection of cast aluminum wheels at the rate of 3/min. [ll and the 

possibility of inspection of pipeline welds in a quantitative manner [11]. 

Practical approaches for pipeline weld inspection by mechanized equipment have 

been described; diagrams of equipment for both double-wall and single-wall 

inspection are shown in Fig. 1 [11]. 

RADIATION SOURCES 

Radiation sources obviously are a critical part of any radiation inspection 

system. A number of advances have been made in recent years including field 

emission X-ray tubes for flash radiography [13] and metal-ceramic X-ray tubes

- 154 - 

for light weight and resistance to shock [14]. Advances of particular interest 

in transportation and energy include a portable linear accelerator [15,16], 

a newly available, medium energy radioisotope source [17,18] and new microfocus 

X-ray tubes that permit inspections with high geometric magnification 

[19,20]. 

The portable linear accelerator, the Minac, is a lightweight (about 100 kg) 

X-ray machine that provides an output of 100 R/min. at 1 meter at an X-ray 

energy of 3.5MeV. The unit is portable and has been used for a variety of 

field applications including inspections of pumps, valves and piping in nuclear 

power plants. Wall thicknesses up to about 20 cm of steel have been inspected. 

A new modification, called the Shrinkac, is extremely small (about 

25 cm on a side). It is coupled to the main radiation head by a flexible waveguide 

up to 6 m in length. This provides excellent versatility for placing 

the radiation head in the tight quarters often found in field situations. 

At the low energy end of the source scale is the radioactive gamma source 

ytterbium-169. This source yields gamma rays in the energy range 53 to 310kV, 

with much of the radiation intensity in the 180 to 240kV energy range [17,18]. 

Therefore, this source is useful for many medium thickness inspections in 

which the higher energy of conventional sources such as iridium-192 or cobalt- 

60 makes them unsuitable. The first half-value-layer for 1 °^Yb is 4.3 mm of 

steel (about 180kV effective). The radiation output is 0.125 R/hour at 1 

meter/Ci. Sources up to 10 Ci in size are available. The short half-life, 

31 days, is a disadvantage but the portability and medium energy make it useful 

for jobs such as panoramic inspections of welds in tubes, a common inspection 

problem in the transportation and energy fields. 

X-ray units with very small focal spots, 100 jam or less, permit a geometric 

magnification of the X-ray image of 5, 10, 20X or more. In addition to providing 

magnified views so that small detail, such as porosity in cast turbine 

blades, can be seen more readily, these microfocus X-ray images also display 

improved contrast because less scattered radiation from the inspection object 

reaches the detector. Several commercially available units claim focal spots 

in the 1 to 10 um range and yield X-rays up to 160kV in energy. Rod anode 

units are available in rod diameters from 8 to 15 mm and lengths from 300 to 

1000 mm. Such units are especially useful for panoramic radiography of tube 

to tubesheet welds in heat exchangers and other similar applications. 

The relatively low energy of these microfocus systems and the excellent detail 

and contrast that can be obtained make them particularly useful for inspections 

of nonmetallic materials and components such as composites and ceramics 

[21], Ar. example of a microfocus X-ray inspection of a ceramic sample showing 

small inclusions is shown in Fig. 2. In this case the microfocus X-ray was 

combined with a real-time detection system so that the part could be oriented 

to give the best inspection view; microfocus X-ray and real-time detection 

can provide excellent inspection results.

NEUTRON RADIOGRAPHY 

- 155 - 

Neutron radiography has been used for many years for inspections of nuclear 

fuel and control material, explosive components in aerospace, and high performance 

turbine blades from jet engines to cite a few examples from the transportation 

and energy fields [2,22,23]. It has also been recognized for some 

time that neutron radiography would provide particular sensitivity to hidden 

corrosion because the hydrogen-containing components in the corrosion products 

would attenuate slow neutrons while most metals that might be in the radiation 

path would be relatively transparent [24,25], Recent developments in 

transportable neutron sources, coupled with improved neutron detectors (including 

real-time methods [26]) have increased interest in neutron radiography. 

Much of this new interest involves detection of corrosion in aircraft and space 

structures. Neutron radiographie equipment now available includes a mobile, 

accelerator-based system developed for the Army [27], An example of corrosion 

detection in aluminum structures with this unit is shown for both film and 

real-time detectors in Fig. 3. 

Other systems soon to be available include higher neutron output acceleratorbased 

systems for the Air Force and the Navy [28]. In addition, a large neutron-inspection 

facility for aircraft maintenance is being planned at McClellan 

Air Force Base; aircraft will be positioned in a special inspection hangar 

for neutron radiography [29]. The inspection of aluminum honeycomb structure 

will be particularly useful for detecting water and corrosion. 

A novel application of real-time neutron radiography is the observation of 

fluid movements inside metallic structures such as engines, pumps, etc. Studies 

of lubrication and fuel movements in engines by a Rolls-Royce, Harwell and 

Burmah-Castrol team [30-32] have excited interest in this technique for better 

understanding of the movements of these fluids in operating systems. An example 

of results from these studies is shown in Fig. 4. 

DISCUSSION AND CONCLUSIONS 

Several novel radiation test methods have been described along with inspection 

applications in the transportation and energy industries. The techniques described 

represent some highlights concerning novel radiation inspection advances 

but the review was not meant to all-inclusive. Other developments such as 

semiconductor array detectors, image enhancement, scattering methods and automated 

inspection [1,2,5,12,33-35] have not been covered here; these also represent 

significant advances in radiography. 

Also not covered are advances in computerized axial tomography [36], There is 

now much industrial interest in tomography as an inspection method for missiles, 

engines, turbine blades, space hardware and other components and assemblies. 

Tomography provides a cross-sectional view of the inspection object to 

yield information about dimensions and discontinuities. An example of a tomographic 

scan of a turbine blade to show wall thinning is given in Fig. 5. 

This limited review has shown that many new radiation test techniques are

- 156 - 

available. Application examples in the transportation and energy fields have 

been presented. 

REFERENCES 

1. H. Berger and V.J. Orphan, "New Developments in Radiography - An Overview," 

Quantitative NDE in the Nuclear Industry, R.B. Clough, ed., Metals Park, 

OH, American Society for Metals, 1983, p. 267. 

2. L.E. Bryant, ed., "Radiography and Radiation Testing," Vol. 3, Nondestructive 

Testing Handbook, Columbus, OH, American Society for Nondestructive 

Testing, 1984. 

3. R.C. McMaster, ed., "Nondestructive Testing Handbook," 2 volumes, New 

York, Ronald Press, 1959. 

4. Anon., "Standard Method for Controlling Quality of Radiographic Testing," 

ASTM E142, Philadelphia, PA, American Society for Testing and Materials, 

1983. 

5. D.A. Garrett and D.A. Bracher, eds., "Real-Time Radiologie Imaging: 

Medical and Industrial Applications," ASTM STP 716, Philadelphia, PA, American 

Society for Testing and Materials, 1980. 

6. S. Nudelman, H.D. Fisher, M.M. Frost, M.P. Capp and T.W. Ovitt, "A 

Study of Photoelectronic-Digital Radiology - Part I: The Photoelectronic- 

Digital Radiology Department," Proc. IEEE, V. 70, No. 7, 700 (1982). 

7. S. Nudelman, J. Healy, M.P. Capp, "Part II: Cost Analysis of a Photoelectronic-Digital 

Versus Film-Based System for Radiology," Proc. IEEE, V. 70, 

No. 7, 708 (1982). 

8. S. Nudelman, H. Roehrig and M.P. Capp, "Part III: Image Acquisition 

Components and System Design," Proc. IEEE, V. 70, No. 7, 715 (1982). 

9. P. Schagen, "X-Ray Imaging Tubes," NDT International, V. 14, No. 1, 9 

(1981). 

10. W. Kuhl, "Detection of X-Rays for Real-Time Imaging," ref. 5, 33-44 

(1980). 

11. S.A. Wenk, R.A. Brokaw and R.G. Schonberg, "The Potential of Real-Time 

Radiography for API 1104," Materials Evaluation, V. 41, No. 9, 1069 (1983). 

12. Anon., "Real-Time Imagery Topical Meeting," Atlanta, August 14-16, 

1984, Materials Evaluation, V. 42, No. 8, 957 (1984). 

13. L.E. Bryant, ed., "Flash Radiography Symposium," Columbus, OH, American 

Society for NDT, 1977.

- 157 - 

14. H. Putzbach, "Metal-Ceramic X-Ray Tubes," Proc. 9th Uorld Conf. NBT, 

Parkville, Australia, Australian Inst. for NDT, Paper 4C-4, 1979. 

15. R. Schonberg, "Portable Linear Accelerator Developments," Report 

EPRI NP-2831, Palo Alto, Electric Power Research Institute, 1983. 

16. M.E. Lapides, "Radiographie Detection of Crack-Like Defects in Thick 

Sections," Materials Evaluation, V. 42, No. 6, 788 (1984). 

17. D. Pullen and P. Hayward, "Gamma Radiography of Welds in Small Diameter 

Steel Pipes Using Enriched Ytterbium-169 Sources," Brit. J. NDT, V. 21, 

No. 4, 179 (1979). 

18. J.D. Hislop, "Radiography Using Ytterbium-169 - Conference Report," 

NDT International, V. 14, No. 5, 295 (1981). 

19. R.W. Parish and D.W.J. Cason, "High Definition Radiography of Cast 

Turbine Blades as a Method of Detecting and Evaluating the Incidence of Microporosity," 

NDT International, V. 10, No. 4, 181 U977). 

20. B.E. Foster and R.W, McClung, "A Study of X-Ray and Isotopic Techniques 

for Boreside Radiography of Tube-to-Tubesheet Welds," Materials Evaluation, 

V. 35, No. 7, 43 (1977). 

21. D.S. Kupperman, H.B. Karplus, R.B. Poeppel, W.A. Ellingson, H. 

Berger, C. Robbins and E. Fuller, "Application of NDE Methods to Green Ceramics: 

Initial Results," Report ANL/FE-83-25, Argonne, IL, Argonne National 

Laboratory, March, 1984. 

22. H'. Berger, ed., "Practical Applications of Neutron Radiography and 

Gaging," ASTM STP 586, Philadelphia, PA, American Society for Testing and Materials, 

1.976, 

23. J.P. Barton and P. von der Hardt, eds., "Neutron Radiography - Proc. 

First World Conference," Dordrecht, Holland, D. Reidel Publishing Co. (1982). 

24. J. John, "Californium-Based Neutron Radiography for Corrosion Detection 

in Aircraft," ref. 22, 168 (1976). 

25. H. Berger, "Neutron Radiographie Detection of Corrosion," Proc. Symposium 

on NDT and Electrochemical Methods of Monitoring Corrosion in Industrial 

Plants, Philadelphia, PA, American Society for Testing and Materials, in press. 

26. H. Berger, "Real-Time Neutron Radiographie Observations of Fluid 

Motion," 1982 Paper Summaries, Columbus, OH, American Society for Nondestructive 

Testing, 487 (1982). 

27. W.E. Dance, LTV Aerospace and Defense Co., Dallas, TX, private communication, 

1983.

- 158 - 

28. V.J. Orphan, Science Applications, Inc., San Diego, CA, private communication, 

1983. 

29. D. Froom, McClellan Air Force Base, CA, private communication, 1983. 

30. P.A.E. Stewart, "Cold Neutron Imaging for Gas Turbine Inspection," 

réf. 5, 180 (1980). 

31. D.A.W. Pullen, "Radiographic Photogrammetry Yields Valuable Data in 

Studies of Running Gas Turbines," Mechanical Engineering Technology, 27 (July, 

1981). 

32. P.A.E. Stewart and J. Heritage, "Cold Neutron Fluoroscopy of Operating 

Automotive Engines," in ref. 23, 635 (1982). 

33. J. McGinnis, "X-Ray Sensitive Photodiode Array," Industrial Res. & 

Dev., 143 (March, 1980). 

34. Anon., "Digital Radiography," Physics News in 1981, P.F. Schewe, ed., 

New York, American Institute of Physics, 75, 1981. 

35. CG. Gardner, "Automated Radiography - A State-of-the-Art Review," 

Report NTIAC-78-1, San Antonio, TX, Nondestructive Testing Information Analysis 

Center, 1978. 

36. H. Berger, ed., "Topical Issue on Tomography," Materials Evaluation, 

V. 40, No. 12, 1982.

FIGURE CAPTIONS 

- 159 - 

Fig. 1. Illustrations of drive mechanisms for real-time radiographie inspection 

of pipeline welds for the source outside the pipe (upper view, double 

wall technique) and for the source inside the pipe (lower view, single wall 

technique). Photo from ref. 11; courtesy Southwest Research Institute. 

Fig. 2. A microfocus, real-time X-ray image of a ceramic spinel disc 

across the diameter showing images of dark inclusions in the size range 50 

to 100 um. Geometric magnification was 14X. 

Fig. 3. Examples of film and real-time neutron images taken with a mobile, 

accelerator-based neutron system. The dark areas show neutron attenuation 

due to corrosion in the upper and lower views; test pieces, including a modified 

ASTM E545 Type A IQI (middle, left) are visualized in the middle region. 

Reference 27; courtesy LTV Aerospace and Defense Co. 

Fig. 4. Two neutron radiographs of an engine, a static view at the left and 

a dynamic view at the right. Differences can be seen such as the pipe full 

of oil (in the middle of the lower part) in the static view while the dynamic 

view at the left shows much less neutron attenuation showing aeration of the 

oil in the scavenger pipe. Reference 31, courtesy NDT Centre, Harwell. 

Fig. 5. Tomographie image of the tip of a turbine blade showing thinning of 

the wall at the left lower region. Wall thinning in the order of 0.05 mm can 

be detected. Courtesy, Scientific Measurement Systems, Inc.

CROSS SLIDE 

- 160 - 

AXIAL DRIVE : 3" 

AZIMUTH DRIVE 

UKE CAMERA 

PLATFORM IMAGE 

INTENSIFIES 

it TV CAMERA 

ASJUITAIll 

TIM« 

COUNTERWEIGHT 

AXIAL DfflVf 

IMME '»» "««* » • 

INTENSIFIE* 

» TV CAMERA 

Fig. 1. Illustrations of drive mechanisms for real-time radiographie inspection 

of pipeline welds for the source outside the pipe (upper view, double 

wall technique) and for the source inside the pipe (lower view, single wall 

technique). Photo from ref. 11; courtesy, Southwest Research Institute. 

Fig. 2. A microfocus, real-time X-ray image of a ceramic spinel disc across 

the diameter showing images of dark inclusions in the size range 50 to 100 

urn. Geometric magnification was 14X.

- 161 - 

MOBILE ACCELERATOR N-RAY SYSTEM 

DETECTION OF CORROSION IN ALUMINUM STRUCTURES 

• 1 I« 

a 

o 

ni» ma Ni A« RIAL :i"i I 1 . :r«.-i 

VOUGHT 

Fig. 3. Examples of fiIra and real-time neutron images taken with a mobile, 

accelerator-based neutron system. The dark areas show neutron attenuation 

due to corrosion in the upper and lower views; test pieces, including a modified 

ASTM E545 Type A IQI (middle, left) are visualized in the middle 

reeion. Reference 27; courtesv. LTV Aerospace and Defense Co. 

Fig. 5. Tomographie image of the tip of a turbine blade showing thinning of 

the wall at the left lower region. Wall thinning in the order of 0.05 mm 

can be detected. Courtesy, Scientific Measurement Systems, Inc.

- 162 - 

Fig. 4. Two neutron radiographs of an engine, a static view at the left and 

a dynamic view at the right. Differences can be seen such as the pipe full 

of oil (in the middle of the lower part) in the static view while the dynamic 

view at the left shows much less neutron attenuation showing aeration of the 

oil in the scavenger pipe. Reference 31, courtesy, NDT Centre, Harwell.

- 163 - 

NDT - RADIOGRAPHIC PROCESSING QUALITY CONTROL 

W.E.J. McKinncy 

E.I. Vu Pont Vz Nzmoun.* S Co., Inc. 

iflilm-ington, 

U.S.A. 

NDT - RADIOGRAPHIC 

PROCESSING QUALITY CONTROL 

ABSTRACT 

In radiography much skill and equipment go into the latent image formation: 

source, film, holder, technique, set-up time, exposure time, and safety. 

After the exposure the film is processed in a chemical to form the useful 

visible image. Actually many films, hundreds of films, are processed by one 

central processor. The best technique and the best film, subjected to 

inferior processing will produce an inferior visible image. This paper 

discusses ways to control processing quality. 


The fundamental tenant of Statistical Quality Control is to define the 

process, monitor the process, and respond when the process fails. The 

process in processing is the oxidation/reduction chemical reaction that 

converts the latent image into the visible image. It is a process that 

operates close to optimal. When and how it deviates from optimal directly 

affects film quality equally.

INTRODUCTION (Continued) 

- 164 - 

Processing, being a chemical process, is controlled by elements of time, 

temperature, and activity. "Time" is the immersion time the film is in the 

chemical and is controlled by a human in manual processing or transport speed 

in the transport system of an automatic processor. Temperature control is 

very specific and is related to the time factor and the activity factor. The 

shorter the time for a given developer activity level or quality, the higher 

the temperature required to produce a certain density. Thus, there is a 

balance between time and temperature based on a given chemical activity. 

This balance allows longer times at lower temperatures or shorter times and 

higher temperatures. However, these relationships only work for a givtn 

activity. This means the developer must be of a certain quality as a result 

of correct mixing, starter solution, and an accurate replenishment/regeneration 

rate. Thus, to control processing quality there are three test approaches: 

TESTING 

o Electromechanical 

o Chemical 

o Sensitometric 

Testing is performed periodically to determine if the process is in control. 

Walter A. Shewhart in the early 1940's created the "Trend Chart" which had two 

goals: indicate a problem requiring a decision or action, and help indicate 

the right decision. He stated that if we control the major variables we do 

not have to worry about the minor variables. In radiographie processing there 

is time, temperature, and activity or three major variables. Shewhart states 

that a trend chart has upper and lower control limits to either side of a goal 

or aim value. If there is a large spike or fall-off from the average track of 

the chart this is due solely to one of the major variables being out of 

mni-rol. Two other interesting facts about trend charts: 

o The uncontrolled variables will prevent a straight, 

line track. There should be some random fluctuation. 

o Although the trend chart demonstrates a problem, the 

goal of a quality control program is to never have a 

problem.

TESTING (Continued) 

- 165 - 

Anything can be monitored and placed on a chart. However, not everything 

need be tested all of the tine. For instance, unless there is some indication 

of a problem then line voltage might only be checked annually. Transport 

rates in inches or centimeter per minute might be measured monthly. 

Replenishment/regeneration rates should be measured weekly. Temperature is 

often monitored daily. Chemical tests are difficult to perform, but could 

include bromide, pH, hydroquinone, and archival quality testing performed by 

a film manufacturer. Clearing time tests, silver content tests, and specific 

gravity tests could be performed periodically in-house. However, the sum 

total of the process is the effect of the chemicals, controlled by the 

electromechanics, on a controlled exposure in a controlled film emulsion. 

This then is sensitometric quality control and is the best test and the most 

common test. 

SENSITOMETRIC QUALITY CONTROL 

Ferdinand Huerter and Vero Driffield in the late 1890's were experimenting as 

amateur photographers with different photographic emulsions for scientific 

recording of events. They found different emulsion formulas produced different 

results so they set out to characterize how a certain film emulsion 

responds to a controlled exposure and controlled development. By giving the 

candidate emulsion a controlled exposure and controlled processing they could 

see how it compared to other emulsions. Thus, there are three components to 

this test: film, exposure, and processing (development). If any two components 

or variables are held constant, the third may be tested. Sensitoroetric 

quality control is a measure of the sum total of the image forming process. 

We borrow from the past to see us into the future. But, do not confuse 

sensitometry with densitometry which is the measurement of a density in a 

pentrameter to meet a work specification. 

In sensitometry there are four measurements: 

o D-Max: Maximum density for maximum exposure 

o Contrast: Difference between two density levels 

o Speed: The amount of exposure to produce a density of 2.00 

o D-Min: Minimum density for minimum exposure 

Base plus fog (B+F) is a measurement of the inherent fog, base, and tint 

values combined and occur prior to an exposure. B+F is not a sensitometric 

value.

- 166 - 

SENSITOMETRIC QUALITY CONTROL (Continued) 

The sensitometric values may be detennined directly from a stepped wedge or 

from calculations derived from an H&D curve. The indicated value for speed 

is measured by reading the density of the step on a wedge closest to D 2.00, 

but above D 2.00. This is adeguate for quality control work because it 

states that for a given film emulsion and a given exposure we wish to achieve 

a certain density. The actual density level is unimportant as long as it is 

above D 2.00. The step chosen is called the speed point. 

Indicated contrast is measured by selecting two steps, one to either side of 

the speed point, on the stepped wedge. The resulting densities are substracted 

to produce the difference or contrast value. This is also called the 

delta D (AD) meaning difference of density or contrast. 

Calculated speed is derived frcm the H&D curve. The coordinate is D 2.00 

plus B+F. At this point on the H&D curve a vertical line is dropped to the 

abscissa where the relative mAs is read. An alternative is to use Du Pont 

Graph Paper which has a built-in Bit Speed Scale which relates to relative or 

arithmetic speed. Step 11 or Bit Speed of 16 is assigned a relative value of 

100. 

Calculated contrast may be calculated one of two ways: 

o Gross or Work which has coordinates of 1.00 and 2.00 

o ANSI, with coordinates of 1.50 plus B+F and 3.50 plus B+F 

In both cases a line is drawn between the coordinate points to form the 

average gradient slope line. There are three common ways to measure the 

slope or rate of inclination: 

o Tangent of the angle 

o Parallel rule method 

o Rise divided by run (Y/X) 

Obviously, indicated values are faster and they can be just as accurate for 

quality control work. The values derived should be tabulated for a month and 

then an average found. This forms the aim point or index or standard which 

is placed on a chart. Upper and lower control limits are established at +/- 

10%. 

WHEN TO TEST 

Test before a work film shows a problem because then it is often too late and 

the work film, plus all the other films in processing, must be repeated which 

is very expensive. Test prior to the processing of work film such as at the 

beginning of each shift. Test during a shift to insure the process is 

holding consistent; that the process is not changing. Record the data on the 

trend chart. Testing is cheaper than repeats and downtime.

HOW TO TEST (Continued) 

- 167 - 

Set aside a box of film, sufficient in quantity to last a month or two, in 

a good environment such as 20 C (70 F). Label the box for test purposes 

only and record the film type, brand, class, emulsion number and expiration 

date. Choose a film that is the same as the film most often used. If both 

Class I and II films are regularly used choose a higher speed Class I film. 

When the first batch of film is about to be used up choose another batch and 

check for similarity by proocessing both emulsions together. 

Choose an easy, fast exposure technique. Expose a stepped wedge and, if 

possible and practical, a defective part that has known flaws. Label this 

part "for test purposes only". Attach appropriate penetrameters to the wedge 

and part. The film holder or cassette, screens, FFD, etc. must be constant. 

The essence of sensitcmetrie quality control is to give a controlled piece 

of film a controlled exposure and then process it to see if the processing 

is in control. Hold everything reasonably constant in order to monitor one 

of the variables. 

To help you get started, and everyone must have a sensitometric quality control 

program, the Du Pont Company offers a Sensitometric Quality Control Kit consisting 

of 50 pre-exposed NDT 55 Class I film, clearing time strips, hypo retention 

solution and reference strips for sensitometry, and archival quality. There is 

an instruction booklet and a copy of the Du Pont Radiographer's Reference with 

technique and processing specifications. Du Pont produces this kit as an 

industry service to help get everyone on QC, it is a no-profit item, and it is 

intended only to get you started. You should be able to start-up your own 

program within the first month. 


Sensitometric quality control measures the sum total effect of processing 

chemicals, time, temperature on a controlled exposure to a controlled piece 

of filai. An image is measured to see if processing is in control. This is 

the best method to effect quality control on Quality Control. Radiography 

is a quality control tool of production, but, is it, itself, under control? 

Has the quality been defined and is it consistent? Quality Control is an 

economic as well as a technical tool. It is absolutely necessary today... 

not twenty years frcm now. 

REFERENCE READING 

"Radiographic Latent Image Processing" by William E. J. McKinney, Section 

Seven, ANSI Nondustructive Testing Handbook, Columbus, Ohio, 1982. 

WEJM/daw

- 168 - 

MAKING THE TRANSITION TO AUTOMATIC PROCESSING IN RADIOGRAPHIC NDT 

W.E.J. UcKinnzij 

E.I. Va Vont Ve. Ue.moan.is ê Co., Inc. 

Wilmington, Ve.la.waie. 

U.S.A. 

MAKING THE TRANSITION TO AUTOMATIC PROCESSING IN RADIOGRAPHIC NDT 

ABSTRACT 

Automatic processing offers the potential for high volume, consistent quality 

radiographie visible image formation. But, it is only partially automatic 

and careful planning and execution are required to achieve the inherent 

value. This paper discusses the advantages, and problems of automatic 

processing and the procedures necessary in making the transition fran manual 

to automatic processing. 


Radiographie latent image processing, whether manual or automatic, by hand or 

by a machine, is a chemical reaction wherein the latent image is amplified 

into the useful visible image. This is an oxidation/reduction reaction. The 

developer is oxidized, releasing electrons which selectively attack exposed 

silver halide crystals and reduces them to black metallic silver. This 

reaction must be controlled by elements of time, temperature, and activity. 

A radiographer chooses appropriate time, temperature, and activity levels. 

In manual processing a person controls these values. In an automatic processor 

the processor controls these functions. To many, this appears to be a 

loss of control. Actually, by eliminating the human variable there is 

increased consisfency and productivity. But, what if the technique is 

incorrect? In manual it can be salvagedi This observation is a classic 

dicotomy because yes the film can be saved, but should it be since obviously 

the wrong technique was used. Allowing that processing, whether manual or 

automatic, must neither overdevelop or underdevelop, why would it be manipulated 

to save a poor technique? The answer is, in order to save an exposure 

and this is referred to as "sight development". Its major failing is that, 

it provides a way to mask the need for technique reform or control.

INTRODOCTIOM (Continued) 

- 169 - 

Quality control means consistent, measurable quality. An automatic processor 

provides this to a degree superior to controlled manual processing. 

It is important to understand the prejudices and history before trying a new 

method, if the new method is to succeed. 

SELECTING A PROCESSOR 

Many people gravitate toward the smallest, cheapest unit available. Unfortunately 

this means a medical unit slowed down. The result is low capacity 

and perhaps increased mechanical load. The ideal processor for NOT applications 

is one with 5 to 10 gallon tanks, preferably plastic for greatest 

temperature stability, bellows replenishment pimps which are the most accurate 

and a transport rate sufficient to produce 50 films per hour on an eight 

minute cycle. Of course price must be considered, but it should be considered 

in relationship to film capacity and darkroom personnel costs. This 

translates into a cost per film. If the processor has built-in variable 

drive, cold water washing, water saver, and automatic standby then other 

savings are applied to reduce the cost per film. 

INSTALLING A PROCESSOR 

Processors may be located through darkroom walls, or be placed totally in 

the dark. Or it may be totally in the light as when a Du Pont Daylight 

Module is used. Some processors can also be placed in trucks, trailers, 

airplanes, and on barges and ships. The only criteria is to provide service 

room to all sides and to consider the ambient conditions. In the first 

case, about two feet is needed on all sides of a processor. In the second 

case, ambient conditions include the room air, moisture, and quantity of 

metal filings and/or general dust present. The processor has a powerful 

blower that draws in 100-200 cubic feet per minute of air to dry the film. 

The air should be relatively clean, cool, and dry. Metal filings affect all 

developers, but would tend to drawn toward an automatic processor more than 

a manual one. Physical installation is described in the installation manual 

of the processor. 

ADVANTAGES/DISADVANTAGES 

The automatic processor offers high volume, consistent quality in a small 

package. An untrained darkroom person can produce good films immediately, 

or in eight minutes. If the processor is monitored periodically, which is 

referred to as sensitometric quality control testing, to insure that it is 

indeed consistent, then it is a simple device important to production.

(Continued) 

- 170 - 

The primary mechanical disadvantage is the presence of 50-100 rollers which 

touch the film. But, in manual we learned how to avoid or control air bells, 

reticulation, streaks, and hanger scratches. In an automatic processor we can 

learn how to keep the rollers clean and because it is automatic it does not have 

any of the problems listed above. 

The primary advantage is consistency. This is achieved with automatic replenishment 

that is only +_ 3% inaccurate for Gorman Rupp Bellows Pumps. It is 

achieved with solid state electronic temperature control with inaccuracies of 

only +_ 1/4°F in a Du Pont NOT 100 processor or similar unit. It is achieved 

with motors, gears, pumps, switches, and thermostats that should last 5 to 10 

years. 


More and more radiographers are buying automatic processors and more important 

is the fact that more radiographers are expressing confidence in automatic 

processing. Automatic processing is more consistent, it handles higher 

film volumes with less mess in a smaller space, it is cost effective. But most 

of all, it is more consistent than manual processing. Once this is realized 

the next step is check out the costs. But what about the occasinal jam or 

roller nark or guide shoe nark? Artifacts such as these will occur occasionally 

but are not generic to automatic processing. They are controllable. In 

fact, more films will be produced artifact-free than are now possible with the 

human manually processing. And improved quality, consistent quality with 

improved capacity and reduced costs is attainable only through the modern NOT 

radiographie processor. 

And, finally, how does one make the transition to automatic processing? By 

matching the automatic processing quality to the manual, or if automatic is 

superior then adjust techniques. In either case density and contrast are 

maintained. 

READING 

"Radiographic Latent Image Processing" by William E. J. McKinney Section 

Seven, ASNT Nondustructive Testing Handbook, Columbus, Ohio, 1982. 

WEJM/daw

- 171 - 

SENSITOMETRIC VALUES FOR INDUSTRIAL RADIOGRAPHY 

W.E.3. UcK.Lnne.ij 

E.I. Vu Font Vz Ne.mou.ii, ä Co., Inc. 

Wilmington, Vzta.wa.ne. 

U.S.A. 

SENSITOMETRIC VALUES FOR INDUSTRIAL RADIOGRAPHY 

ABSTRACT 

In measuring sensitometric values of different film products, or in the use 

of sensitometry as a quality control tool, several methods may be used. This 

paper discusses indicated versus calculated values. Examples are given of 

the various calculated values. 


In NOT Quality Control, it is often necessary to produce film that meets 

specific sensitometric values. Perhaps more important is to log the values 

of test films and to monitor the trends so as to insure consistency. Trend 

charts may be constructed frcm indicated or calculated values as outlined 

below. 

SPEED 

o Indicated: Read the density of the step closest to Density 2.00 If 

one step reads D 1.85 and the next higher step reads D 2.20, 

choose the higher step as standard. Remember this "value" 

is influenced by base plus fog. Indicated speed may also be 

derived by "eye-balling" (visual comparative analysis). 

o Calculated: Locate the coordinate of D 2.00+B+F on the H&D curve. Draw 

a line down from this point to intercept the Bit speed calculation 

bar*. If your graph paper lacks the calculator set 

one up beginning at Step 11 = 16.0 Bits, Step 9 = 17.0 Bits, 

etc. A Bit Speed of 16.0 = Arithmetic Speed of 100; 17 = 200; 

15 = 50. Alternately, if Step 11 represents a 100 speed, then 

Step 12 is a 41% increase in energy or about 20% slower speed. 

Step 13 represents half the speed of Step 11 or 50. 

*Du Pont H&D Graph Paper.

UJNXKAST 

o Indicated: 

o Calculated: 

- 172 - 

Read the density of the steps one above and one below the 

speed point. Subtract these densities. The answer is the 

"difference" or "delta D" (4 D). 

Gamma (

CONTRAST (Continued) 

NO. 1 PARALLEL LINES: 

- 173 - 

Once the slope or line is drawn, any one of the following 

three methods may be used to produce the same answer 

{*_ 5%). The following graphs demonstrate each method. 

In the lower right hand corner of the graph paper, there is a slash line or 

a full line at a 45 angle that intercepts the density scale at D 1.00. 

Using parallel rulers, opposing triangles or by measuring with a ruler, move 

the H&D average gradient slope line over so that it intersects with the slash 

reference line at the bottom of the graph paper. Project the line upward; where 

it meets the vertical rise of numbers, read the number as the contrast value. 

As you can easily see the rise in numbers stops too soon. By measurement, the 

vertical scale can easily be extended. Another method would be to cut out a 

portion of graph paper and make a scale extender. Another approach is to change 

the density scale readings. To do this move the parallel slope line to the 

right, just past Step 19. Draw a 45 angle and line so that it exactly meets 

D 0.50 on the vertical rise. Relabel this as 1.00; 1.00 becomes 2.00, etc. The 

scale has been doubled. 

NO. 2 TANGENT OP THE ANGLE: 

Using a protractor read the angle of the incline or slope (H&D average gradient 

line) to the half or quarter of a degree, if possible. Look up the 

tangent of that angle in the trignometric functions of a mathematics book or 

use a scientific calculator. The answer is the contrast value. 

NO. 3 Y OVER X: 

This is also called X into Y or "Rise (Y) over Run (X)". From the upper 

coordinate (Y) of the average gradient slope line, draw a vertical line down, 

very straight and parallel. From the lower coordinate (X) draw a line horizontal 

to the right very straight and parallel. Where they intersect should 

be a 90 angle. Measure the vertical Y axis in centimeters. Measure the 

horizontal X axis in centimeters and divide into the Y value. The answer is 

the contrast value. 

It is only necessary to perform one of these methods; the one method you 

prefer. There are, also, other methods such as the difference in relative 

log exposure divided into the density difference range of 2.00. 

WEJM/daw

- 174 - 

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- 175 - 

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- 176 - 

RECENT MICROFUCUS X~RAY IMAGING APPLICATIONS 

R-icha/id S. Pzagtot 

Rldgz Incoipoiatzd 

Tu.ck.zK, Go.on.Qia, U.S.A. 

The advent of reliable, commercially-available high output microfocus 

x-ray generating systems has enabled real time x-ray imaging to successfully 

compete with film radiography. The advantages of a near point-source 

of radiation in reducing x-ray image penumbral unsharpness are well known. 

While tnicrofocus x-ray equipment has existed for more than four decades, 

available microfocus x-ray equipment has been characterized by low x-ray 

output, rendering the technique unsuitable for real time x-ray imaging. 

Recently, advances in microfocus x-ray tubehead design have made available 

microfocus x-ray equipment operating at energies up to 160kV with focal 

spots on the order of 10 microns while at the same time producing x-ray 

output intensities comparable to conventional x-ray generating systems. 

The result is that microfocus x-ray imaging can be successfully applied 

to a wide range of test projects ranging from thin, lightweight materials 

to materials of substantial cross section and density up to the equivalent 

of approximately 5/8" of steel. 

The advantages of using microfocus imaging techniques are twofold. First, 

image resolution is improved by orders of magnitude according to the 

amount of geometric enlargement used. Resolutions on the order of 25 lint: 

pair per millimeter are routinely achieved at 10X geometric enlargement 

using a commercially-available imaging system having a resolving power of 

only 2.5 line pair per millimeter. Second, image contrast is improved. 

Thii» comes about through a reduction in the amount of test object scatter 

radiation reaching the imaging system due to two factors: 1) the x-ray 

illuminated area of the test object is reduced, thereby reducing the 

amount of scatter generated, and 2) the test object is positioned away from 

the image receptor reducing the scatter intensity and intercepted solid 

angle. 

The result is that high output microfocus x-ray imaging systems can 

routinely produce equivalent penetrameter sensitivities on the order of 

2-1T or even 1 — IT in some instances. The resulting high quality real time 

video x-ray image may be processed using analog and/or digital image

- 177 - 

processing techniques. Images may be archived by electronic means, thereby 

eliminating costly x-ray film. 

This paper will highlight several practical applications of microfocus 

x-ray imaging in the area of micro-electronics, investment castings and 

composite aerospace materials. Test part sizes range from small microcircuits 

to large sections of aerospace structures. Geometric magnifications 

ranging from 10X to more than 200X are illustrated. 

One interesting application which will be discussed involves the automated 

in-place inspection of bonded components on a complete airframe. Robotic 

positioning of the microfocus x-ray imaging system over an 85' x 90' 

working envelope is provided. All inspection functions except the interpretation 

of the actual microfocus x-ray image are under computer control. 

Planned future system enhancements include a digital optical disk archival 

image storage and retrieval system allowing past inspection images to be 

compared with present inspection images for tost part performance tracking. 

It is also anticipated that much of the operator-dependent real time x-ray 

image interpretation can be placed under computer control providing a truly 

automated x-ray inspection system m-ide possible by the high quality real 

time x-ray image resulting from the use of high output microfocus x-ray 

imaging techniques.

ABSTRACT 

- 178 - 

WET CHANNEL INSPECTION SYSTEMS FOR CANDU NUCLEAR REACTORS,.. 

CIGAR AND CIGARette 

M.V.C. l\olzi>, M.P. Volbzy and K.S. Uakil 

Ontaftlo Hydto, To Konto, Ontario, Canada 

This paper describes and compares the operation of the CIGAR and CIGARette wet 

channel inspection systems that have been developed by Ontario Hydro for CANDU 

nuclear reactor pressure tubes. Initially, the highly sophisticated, 

automated CIGAR (£hannel inspection and Gauging Apparatus for Reactors) is 

described. When the hydride blistering problem was discovered in Pickering 

NGS Units 1 and 2 in late 1983, CIGAR was not operational, so a less 

automated, slower system called CIGARette was rapidly designed and built. The 

operation and some problems experienced with CIGARette are also described. 

1. INTRODDCTIOH 

Pressure tubes in CANDU nuclear reactors contain the bundles of natural 

uranium fuel and act as a pressure boundary (see figure 1). There are several 

hundred pressure tubes per reactor, mostly made of Zr-2.5%Nb, and these are 

surrounded by a calandria tube containing dry gas between the two tubes. 

Pressure tubes are 103 mm inside diameter, 4.0-4.5 mm wall thickness and 

about 6m long. The pressure tube and calandria tube are separated by garter 

spring spacers, from two to four per pressure tube, depending on the unit. 

There are two general reasons for inspecting pressure tubes - regulatory and 

economic. So far the mandatory regulatory requirements have been less 

demanding than special requirements, only requiring gauging of a few pressure 

tubes per year/1/. The economic reasons vary from wanting information on 

pressure tube behaviour, to predicting and correcting pressure tube problems, 

to major, well-publicized outage problems. The latter include the rolled joint 

delayed hydride cracking in Pickering units 3 and 4 in 1973-74/2/, and the 

more recent hydride blistering problem in Pickering units 1 and 2, which was 

the impetus for this paper. 

There are two general approaches to inspecting pressure tubes, namely 'dry' 

and 'wet' channel systems, and Table 1 gives a sample list of both types. Dry 

channel systems require pressure tubes to be defuelled, isolated, drained and 

opened, so they require more time and man-rem than wet systems. Also 

ultrasonic volumetric inspection is difficult to perform in a dry channel. Wet 

channel inspections are performed in a defuelled channel still connected to 

the primary coolant circuit. Channels can he defuelled on-line beforehand to 

save time and costs with either approach.

- 179 - 

This paper describes the concept and evolution of the sophisticated CIGAR 

pressure tube inspection system, and the very rapid development and 

operational experience with the less automated CIGARette system,both 'wet' 

systems. 

2. CIGAR 

2.1 Concept 

Following the problems experienced with the rolled joint cracking in Pickering 

units 3 and 4 in 1973-74, it was decided that a sophisticated system for 

inspecting and gauging pressure tubes (PT) was required/1/. The system was 

called CIGAR (Channel Inspection and Gauging Apparatus for Reactors), and the 

specifications have evolved somewhat with time/2-4/. The inspection objectives 

are listed in Table 2 and the operational objectives in Table 3. 

The CIGAR system can be sub-divided into three areas 1) the delivery system 

(drive mechanism and drive rods); 2) the nondestructive evaluation systems 

(eddy current,sag and ultrasonics); and 3) control and data handling (storage 

and display). These are described briefly below. 

2.2 The Delivery System 

The drive mechanism, shown on the Ontario Hydro Research mock-up (Figure 2), 

was developed by CGE in Peterborough. It is designed to be moved into the 

reactor vault and mounted on the fuelling machine bridge, which moves it to 

the appropriate channel for inspection. The CIGAR drive is then connected to 

the remote control console which is outside containment (Figure 3) by a cable 

approximately 100 m long. 

The CIGAR inspection head (described in more detail later) has been loaded 

beforehand into a defuelled channel by the fuelling machine. The drive 

mechanism pushes a drive rod containing the instrumentation cables through a 

special closure plug/carrier after connecting to the enclosed CIGAR inspection 

head. This connects the nondestructive evaluation (NDE) systems transducers 

directly to their instruments in the remote control location. Later the CIGAR 

drive automatically inserts a second section into the drive rod to make it 

long enough to push the inspection head to the end of a Darlington channel 

with 150 mm of elongation. 

2.3 The NDE Systems 

The CIGAR NDE system utilizes the following transducers: 

1) an eddy current probe to measure the gap between the PT and calandria 

tube (CT); 

2) an inclinometer to measure the sag deflection of the pressure tube; 

3) four ultrasonic probes mounted to produce 50° shear waves for volumetric 

inspection; 

4) a three transducer, four channel ultrasonic gauging system for measuring 

PT diameter and wall thickness.

- 180 - 

5) an eddy current probe to locate the position of garter spring spacers. 

These transducers are operated from instruments in the remote control location 

using up to 100 metres of cable. 

The NDE transducers are mounted in a CIGAR inspection head (Figure 4), which 

consists of two centering modules, a universal joint connector, plus eddy 

current, sag measurement and ultrasonic modules. The heads are modular so 

inspection systems can he readily interchanged. The limiting factor at the 

present time is the number of miniature coaxial cables and connectors that 

can be inserted down the drive rods, namely fifteen. A radiation resistant 

in-channel multiplexer is under development that will allow more transducers 

to share the 15 available cables. 

2.4 Control and Data Handling 

CIGAR inspections are controlled by a DEC PDP11 minicomputer at the remote 

console, using a series of preprogrammed scanning patterns. For example, the 

operator might perform 1) a high speed helical ultrasonic inspection using a 

3mm axial step (100 minutes for the whole channel); 2) sag measurements, 

performed axially (15 minutes), 3) eddy current and ultrasonic wall thickness 

measurements for gap (15 minutes), 4) garter spring location if required (15 

minutes), 5) detailed ultrasonic inspection of a selected area (say 20-50 

minutes), for a total inspection time of about 3 hours per channel. The order 

is flexible, depending on the operator. 

CIGAR generates a vast amount of data, more than an operator can absorb while 

watching the various screens and displays, so some method of recording and 

displaying the data is necessary. Much of the data is recorded by a 14 channel 

magnetic tape recorder in analog form (Figure 5). This device records the 

axial and circumferential positions from the console, 2 channels (X and Y) 

from the eddy current instrument, eight channels of data from the flaw 

detection ultrasonic instrument (two channels each from four probes), plus a 

voice channel. These can he played back through a second computer to give, for 

example, isometric displays of the ultrasonic flaw detection data. The 

ultrasonic gauging data and sag data are recorded on digital cassette. Also, 

the ultrasonic flaw detection system triggers alarms to notify the operator 

when signals above threshold are observed, and these are printed out on the 

alarm print module, along with the axial position. 

3. CIGARette 

When Pickering unit 2 pressure tube G16 burst on August 1, 1983, due to 

'hydride blister' related cracking, CIGAR was not ready for operation. A mini- 

CIGAR, quickly nicknamed CIGARette, was improvised using many CIGAR components 

and basic inspection system concepts, and was under test in about a month. A 

small friction drive mechanism was designed and built by AECL Sheridan Park 

using standard components for speedy delivery/5/.

- 181 - 

Though CIGARette was based on CIGAR, it differed in a number of ways (see 

Table A). First, if CIGAR is automated (ie) computer controlled, CIGARette is 

mechanized (ie) operator controlled. The operators run the CIGARette console 

directly and watch and control the NDE instrumentation. Consequently CIGARette 

scans a lot slower than CIGAR. The CIGAR drive is much larger, more powerful 

and automatically aligns and connects to the channel being inspected whereas 

the CIGARette drive is smaller and is clamped to the end-fitting by two 

operators (Figure 6). The two drive rods are the same for both systems, but 

are loaded by hand for CIGARette. Both CIGAR and CIGARette use CIGAR heads, 

closure plugs and drive rods, with only minor modifications. 

CIGAR records ultrasonic data on 1A channel tape and plays back directly 

through a computer for quasi-on line print outs. In contrast, CIGARette 

records data on a similar recorder for later playback, and operators are 

required to make assessments of the data in situ. CIGARette play back is 

supplementary. Also, the CRT screen of the ultrasonic flaw detector was 

videotaped during CIGARette inspections to assist in post test assessment. 

CIGAR has a slip-ring unit that allows the inspection head to rotate 

continuously. CIGARette does not, so it has to scan back and forth with a 

maximum rotation of A00°. CIGARette is operated from an AECL-designed 

console/5/,(Figure 7) which can be operated manually or be made to perform a 

continuous scan pattern with rotational and axial steps of preset size and 

velocity. 

3.1 CIGARette NDE Inspections 

When CIGARette was developed, the CIGAR inspection head did not contain a 

garter spring spacer location system. The head was modified by installing an 

eddy current coil for this purpose in place of the CIGAR Sag Module. 

Subsequently, the CIGAR head was changed to contain both these systems. The 

CIGARette head includes all but the Sag Module. Once the head was in channel 

with both drive rods and end connector in place, the CIGARette inspection 

typically took the following sequence: 

1) An axial pass (with ultrasonic probes at 6 o'clock or bottom dead centre) 

to locate the garter spring spacers using the eddy current probe; 

2) If one or both of the two garter springs was found to be out of position, 

ultrasonic inspections were performed on the longest span. If the two 

garter springs were in their design locations (or could not be found), 

ultrasonic inspections were performed on the middle of the outlet span 

(Figure 8). Typically a number of locations along the PT were inspected 

if reflectors were found. 

3) Eddy current gap measurements were performed along the length of the PT. 

A) Ultrasonic gauging of the wall thickness was performed, along nominally 

the same path as the gap measurements, to provide correction information 

for the gap measurement. 

This sequence is illustrated in Table 5. Ultrasonic volumetric inspections 

were performed over an arc of only 45 degrees around bottom dead centre, using 

a 2mm axial step. The scans were performed over a limited length of pressure

- 182 - 

tube, say 0.5-1.0 metre, so in practice only a few percent of the total PT 

volume was inspected for hydride blisters. In contrast, CIGAR inspects 

nominally 100% of the PT in a similar period of time. The eddy current gap 

operation will be described elsewhere in this conference/6/. 

3.3 Problems with CIGARette 

After the expected teething problems, CIGARette worked pretty well. A total of 

66 PTs were inspected in Pickering 1 and 2, of which about one third showed 

reflectors. A further six PTs were inspected in Pickering unit 3. Some of the 

problem areas that had to be resolved are discussed below. 

Drive Rods/Connectors 

The drive rods were loaded by operators working in plastics on the FM 

bridge. First, they had to be clean so no slippage occurred in CIGARette's 

friction drive. Second, they had to be connected slowly so that the connecting 

pins would engage without damage. The connecting pins between the drive rods 

and end/head connectors were occasionally damaged during loading. Since about 

100 pins are used in each drive rod/head assembly, a critical shortage of 

these components arose. Occasionally an inspection had to be performed with a 

probe not functioning. 

End Connectors 

This was a major problem in the early stages with individual wires in the 

catinery cable becoming tangled in the pulleys and being cut. Also, the 

oscillating rotational scan pattern of the ultrasonic inspection fatigued the 

cables at a sharp bend where they came out of the end connector. These 

problems were solved by 1) binding the loose wires with tape, and 2) by using 

a strain relief spring on the end of the connector. 

Rotational Slippage 

After only a few inspections it was recognized that progressive rotational 

slippage between the drive rod and friction drive (containing the position 

readout) could occur. This had serious implications for the ultrasonic 

inspection. Blisters were expected only about the 6 o'clock position and a 

rotation scan of only 45° was performed about this position. Unrecognized 

slippage of only 30° would put the inspection out of the range of the 

potential defects. The problem was solved by putting mercury switches in the 

end connectors to trigger at the 6 o'clock position to indicate if slippage 

occurred. 

Operator Comfort 

Since CIGARette inspections centre round the operator, fatigue was a potential 

problem. The first inspection was done inside the vault in plastic suits, at 

great effort. Later, containment penetrations were made, and cabins 

constructed, which greatly improved operator conditions.

Problems with Eddy Current Systems 

- 183 - 

These are described in more detail in/6/. The garter spring location system 

coil went through several development stages before a satisfactory reliability 

and signal-to-noise ratio was achieved. The gap measurement system, which had 

been under development before the hydride blistering problem/4/, saw its first 

active service during these inspections and required considerable procedural 

modifications before satisfactory results were obtained. 

Problems with Ultrasonics 

The CIGAR heads and flaw detection (volumetric inspection) worked well, 

largely because they had been tested extensively in the laboratory. The probes 

lasted longer than expected in the reactor radiation fields. Most of the 

problems involved failed connectors and radiation decay of cables. Problems 

with standing signals were solved by modifying the ultrasonic modules. 

We had some problems selecting a recording threshold since hydride blisters 

were a new problem. This was solved in practice by using a 'grass-level plus 1 

approach. Probe power checks were performed in each PT using a transmit— 

receive method between opposite probes, then the gain increased till grass was 

observed on the inside surface gate. All signals a certain level above grass 

were noted. 

Initially there were problems playing hack the data from the 14 channel analog 

tape. Videotaping the flaw detector CRT was useful, but a composite print-out 

was also required. Eventually the CIGAR computer play back system was modified 

to provide an off line isometric playback facility. Figure 9 shows a typical 

computer generated plot of a blistered area, produced from the composite data 

of four shearwave probes. 

The ultrasonic gauging system in CIGARette was not a great success. The 

digital thickness gauges selected for this joh, could not be made to operate 

in a stable manner with the long cables used. Consequently discrete 

measurements were made on the CRT of the flaw detectors. This was very time 

consuming and did not provide as high a precision as desired for gap 

measurement correction. 

4. CLOSING COMMENTS 

The CIGAR system will be used for active channel pressure tube inspection in 

1985. It will provide the capability to perform a wide range of inspections 

on a number of tubes quickly, with minimal manrem expenditure and with 

comprehensive data analysis and presentation either on-line or immediately 

after the inspection. The CIGARette system that was improvised for the 

Pickering 1 and 2 inspections filled the gap adequately and is now operating 

well. CIGARette may he used for quick inspections and for special purposes 

such as carrying video cameras. Many of the problems experienced with 

CIGARette, such as those due to instrumentation cables and connectors, have 

not been experienced with CIGAR due to the different delivery methods. From an 

NDE viewpoint, the modular CIGAR head worked well, and in the future, one can 

predict that new inspection modules will be developed and used for a variety 

of different applications.


- 184 - 

Many people in Ontario Hydro's Research, Design and Development, Central 

Nuclear Services and Pickering, plus AECL-Sheridan Park and Canadian General 

Electric at Peterborough have been involved in the design, development, 

manufacture and operation of CIGAR and CIGARette. Unfortunately there are too 

many to list by name, as we would like. 

REFERENCES 

1. M.P. Dolhey and O.A. Kupcis, I. Hech. E. Conf. publication no C32/79. 

2. O.A. Kupcis, Proceedings of the Third Conf on Periodic Inspection of 

Pressurized Components, I. Hech. E. Conf. Publication 1976-10. 

3. J.A. Baron, M.P. Dolbey, J.H. Erven, D. Booth and D.W. Murray, I. Mech. 

E. Conf. Publication no C139/82. 

4. J.A. Baron, M.P. Dolbey, J.H. Erven, D. Booth and D.W. Murray, Nuclear 

Engineering International, December 1981, p.45. 

5. P. Braun, Paper presented at Conf on Robotics and Remote Handling in the 

Nuclear Industry, Toronto, Sept 23-27, 1984. 

6. J. Sedo - Paper presented at this conference.

- 185 - 

TABLE 1 

Sample List Of Dry And Wet Channel Inspection Systems 

Dry Wet 

Closed Circuit TV Rolled Joint Inspection System 

Orenda Sag CIGAR 

Diameter and Profilometry CIGARette 

Inspection 

STEM Eddy Current Surface Flaw 

Inspection 

Laser Sag Measurement 

TABLE 2 

Inspection Objective Of CIGAR/3/ 

Conduct mandatory in-service inspections, as required. 

Conduct in-service inspections to monitor pressure tube conditions 

to predict pressure tube life. 

Obtain information that could indicate required action to extend the 

inservice life of pressure tubes. 

Obtain design-related information for future reactors.

- 186 - 

TABLE 3 

Operational Objectives Of CIGAR/3/ 

Minimize outage time required for inspections. 

Minimize radiation exposure to personnel. 

Operate on a shutdown reactor with the heat transport system cold 

and depressurized. 

Inspect defuelled channel without draining or isolating, 

To be operated remotely from outside containment. 

Acquire inspection information at the remote console and process the 

data while the equipment is in channel. 

Provide archive data for comparison with future inspections. 

To operate on Bruce, Pickering and Darlington reactors. 

TABLE 4 

Differences Between CIGAR and CIGARette 

CIGAR CIGARette 

Control Automated, computer Mechanized, 

controlled operator controlled 

Inspection Speed Fast Fairly Slow 

Drive Mechanism Large, mounted on F.M. Small, mounted by 

Bridge operator 

Drive Rod Loading Automatic Manual 

Data Playback Essentially on-line Off-Line, later 

Scan Pattern Helical or back and Back and Forth 

Forth only

Step 1 

Step 

2 

Step 3 

Step 4 

No 

- 187 - 

TABLE 5 

Sequence for CIGARette Inspection 

Action Objective 

Axial Scan Using Eddy 

Current Garter Spring 

Probe 

Garter Springs Located? 

\ f Yes 

Garter Springs in design location? 

No 

Ultrasonic Inspection 

of longest span 

Record any indications 

Yes 

Ultrasonic Inspection 

of span between outlet 

rolled joint& nearest 

garter springs 

Axial/rotational scans to measure 

PT/CT gap using eddy current 

Ultrasonic Gauging of pressure 

tube wall thickness 

Report Results 

Locate Garter 

Springs 

Detect 

Hydride 

Blisters, 

if present 

Eddy Current 

and ultrasonics 

to measure PT/ 

CT gap

Pressure 

Tube 

Fuel 

-Carter Spring Spacer 

- 188 - 

FIGURE 1 

— End Fitting 

CANDU REACTOR FUEL CHANNEL SHOWING 

PRESSURE TUBE AND CALANDRIA TUBE 

End Fitting 

FIGURE 2 

Instrument Contaminant Enclosure 

CIGAR DRIVE MECHANISM IN MOCK-UP 

Drive Shafts

Equipment 

air loch 

FIGURE 3 

- 189 - 

RamoU control 

SCHEMATIC SHOWING OPERATION OF CIGAR 

ON REACTOR FACE FROM REMOTE CONTROL CONSOLE 

Centering Modules 

FIGURE H 

CIGAR/CIGARETTE INSPECTION HEAD SHOWING 

CENTERING AND INSPECTION MODULES 

to 

• Ultrasonic Moiule 

•Eddy Current Carter 

Spring Coils 

Eddy Current Cap Coil

«AC UUSU«O«HII 

INCUMKTM 

- 190 - 

sur««««— FIGURE 5 

CIGAR DATA ACQUISITION AND PLAYBACK 

• End Fitting Sleeve 

Clamps onto Pressure 

Tube 

FIGURE 6 

Rotational 

Drive Motor 

CIGARETTE DRIVE MECHANISM IN MOCK-UP 

Axial Drive Motor 

Hydraulic 

Levelling 

System

- 191 - 

FIGURE 7 

: (Axial Drive Index 

( Rotational Drive Index 

\ Digital Position Readouts 

Function Controls 

CIGARETTE CONTROL CONSOLE IN CONTROL ROOM

Inlet or 

Outlet £nd 

Outlet 

2m 

- 192 - 

6 m 

2m 

Design Locations 

of Carter Springs 

~Area Scanned by 

Volumetric Ultrasonics 

(a) Out of Location Carter Spring 

7 

Area Scanned by 

Ultrasonics 

Design Locations 

Actual Locations 

(b) Correctly Located Carter Sprints 

FIGURE 8 

2m 

Actual Locations 

of Carter 

Pressure Tube 

J 

Calandria Tube 

•• Rolled Joint 

SCHEMATIC ILLUSTRATING INSPECTION AREAS FOR PRESSURE TUBES WITH 

(a) CARTER SPRINGS OUT OF LOCATION AND (b) SPRINGS IN THE DESIGN LOCATIONS

TYP|CAL CIGARETTE 

.TAOUTPUT. 

ULTRASONIC DATA OUTPÜ. 

. • * • • - . * 

One Blister 

^C: \One Blister 

[Detected by 

)Four Probes 

OUTERSIDE 

9 

OUTER 

SURFACE 

SHOW.NC BUSTERS ON

- 194 - 

AN ADVANCED HEAT EXCHANGER EDDY CURRENT INSPECTION SYSTEM 

M. Ve.Ve.Mno, H. Ghznt, H. Licht 

Atomic Energy oß Canada Limi.te.cl 

Chalk Rivz/i, Ontario 

ABSTRACT 

Heat exchangers are important components in many process industries, 

especially nuclear power plants. Nondestructive testing techniques and 

equipment are required for fast, reliable testing of heat exchanger tubes 

during both unanticipated and scheduled inspections. An advanced eddy 

current inspection system has been developed at Chalk River Nuclear 

Laboratories to meet this need. This system was designed to meet the needs 

of the operator, analyst, station manager and regulatory agencies. The 

inspection system provides real-time data acquisition, display, recording and 

hardcopying to allow on-line analysis by a qualified eddy current analyst. 

In real time, a one page printout containing impedance images and strip chart 

traces of eddy current signals is produced for each tube. Tube coordinates, 

inspection data, print number and data acquisition validity messages are also 

included for record keeping purposes. This record enables the signal analyst 

to readily identify, analyze and comment on anomalous signals. Three pairs 

of eddy current test data (X and Y components) can be fed into the system. 

All three data sets plus probe position and tube information are stored on FM 

tape. The operator can select one data set for real-time display. A 

sophisticated tape playback feature permits a convenient means of further 

analyzing data. The real-time hardcopying allows the inspection team to 

provide the utility manager with a copy of the data and inspection results 

immediately following the inspection. 

This inspection system, known as R'Eddy Record, requires two people to 

operate, can realistically inspect 400 tubes/8 h shift, and is compatible 

with all commercially available single and multifrequency eddy current 

instruments using either absolute or differential type probes. 

A particularly impressive feature of the format used to present the data is 

the inherent ability of comparing previous inspection data with current 

inspection data. This provides a convenient means of monitoring a heat 

exchanger's tube performance, and permits long range planning for tube 

plugging and/or heat exchanger replacement. 

This paper describes the complete inspection system, how it is operated and 

field experience to date.


- 195 - 

The development of an advanced eddy current testing (ET) nondestructive 

Inspection system for heat exchanger tubing in nuclear facilities and other 

industries presents a challenging set of problems. For the operator the 

inspection system must be reliable, easy to operate and maintain. For the 

analyst the system must present the data to facilitate on—line analysis. 

Stored data required for signal confirmation must be comprehensive and easily 

retrievable. The system must enable the complete inspection to be done 

quickly and reliably to reduce the overall downtime of the heat exchanger, 

and yet provide the utility manager and control agencies a convenient means 

by which the conclusions of the inspection can be verified. To provide a 

method of monitoring a heat exchanger's tube performance and permit long 

range planning of maintenance and replacement, a convenient method of 

comparing collected data must be available. The R'Eddy Record system, 

designed to meet these needs of the operator, analyst, utility manager and 

control agencies, represents a significant advance in eddy current inspection 

technology. 

Eddy current inspection is focused on the collection and analysis of two 

voltage signals, Vx and Vv, which together represent the impedance of the 

eddy current coil and make up one channel of eddy current data. The Vx and 

Vy signals respectively represent the XY cartesian mapping of coil 

impedance and the Lissajous figure, Vx versus V„, represents the eddy 

current impedance image or 'feature'. The measured coil impedance is a 

function of frequency and the material environment surrounding the coil. 

Conventional single and multifrequency eddy current instrumentation can 

provide X and Y outputs at both single and multiple test frequencies, 

frequency mixes, and provide a storage oscilloscope impedance image display. 

Signal analysis of the eddy current data involves interpretation of the 2—D 

image with the separate X and Y versus distance traces providing information 

on signal direction within the image. 

A rudimentary inspection would consist of manual probe insertion down a heat 

exchanger tube. An operator would analyze the tube condition by visually 

monitoring the oscilloscope impedance image. A single storage scope display 

superimposes images, thus signals can be missed and/or misinterpreted. If 

long terra data collection or hardcopy results were needed the data could be 

put on FM analog tape and/or strip chart. On-line signal analysis is very 

difficult and prone to errors in this type of inspection. Post inspection 

analysis must then be accomplished. On large inspections, literally miles of 

strip chart can be generated making data retrieval and record identification 

cumbersome and difficult. Data playback from analog tape also presents the 

same problems of retrieval, identification and feature display 

superpositioning. 

A logical extension to manual inspection and the CRNL forerunner to the 

R'Eddy Record system was a motorized probe pusher/puller (probe drive) to 

speed up data collection, and a separated feature playback display that 

allowed an operator to display non-superimposed ET features on an analog XY

- 196 - 

storage terminal. However, the major problems associated with off-line 

analysis, data retrieval and record keeping still remained. The R'Eddy 

Record system solves these problems. 

2. SYSTEM DESIGN OVERVIEW 

2.1 General 

The R'Eddy Record heat exchanger inspection system provides computer 

controlled data acquisition, probe positioning, data storage and real-time 

graphical data display with hardcopy printout. The data display and hardcopy 

were specifically designed to facilitate on-line analysis and provide a long 

term quality assurance document. This one page printout provides a permanent 

record of the X and Y signal traces, separated features, tube Identification 

and pertinent scan information of a single channel of eddy current test data 

for immediate analysis and future review. Three channels of eddy current 

data can be collected and stored along with record identifiers and position 

information on magnetic tape and strip chart. These data records can quickly 

be retrieved and displayed through the playback feature. The system will 

collect and display any twin set of analog signals making it a general data 

acquisition system independent of the type of probe or probe configuration 

and the particular signal generating instrumentation. 

The overall system configuration can be divided into a probe drive control 

subsystem (Figure 1) and a data processing subsystem (Figure 2). The modular 

design allows for flexibility in configuration, and ease of maintenance and 

updating. The overall weight and physical dimension of any system module 

(weight

- 197 - 

This subsystem was designed for repetitive scanning of heat exchanger tubes 

with automatic stop and retract upon reaching a preprogrammed position or 

under control from the data processing subsystem. The probe drive control 

subsystem communicates status, position information, tube coordinates and 

three twin analog channels to the data processing subsystem through a single 

multi-channel link. 

2.2.2 Probe Drive 

The probe drive comprises four segments designed for easy access to congested 

areas and assembly in minutes. The probe drive provides 400N (90 lbf) drive 

force using four drive wheels for both plastic and steel wound cables. A 

compact motor driven take-up reel with mercury-wetted slip rings is utilized. 

An incremental shaft encoder allows for high resolution data collection and 

accurate positioning. A flexible probe guide tube directs the probe cable 

and probe to the heat exchanger tube. The probe guide normally connects to a 

calibration tube and zero limit switch. This switch allows the system to 

correct for positioning hysteresis effects. A second mechanical switch 

placed within the probe guide tube just ahead of the drive wheels prevents 

the probe from being withdrawn through the drive wheels. 

2.2.3 Probe Drive Interface 

The probe drive interface module contains the probe drive power supply, speed 

regulation controller and manual probe drive controls. Manual probe drive 

operation without position readout is possible with this module alone. 

2.2.4 Probe Drive Controller 

The probe drive controller is a microprocessor controlled module responsible 

for the execution of preprogrammed scan patterns and manual scanning with 

full distance display plus slip and jam detection. The probe drive 

controller accepts input via two remote keyboards from either the operator or 

the analyst. Shaft encoder and limit switch inputs provide positioning 

information and three twin differential channels transmit eddy current data 

to the data acquisition subsystem. 

2.3 Stand-Alone Probe Controller 

The stand-alone probe controller module is a scaled down version of the probe 

drive interface and controller. This single module is fully programmable by 

the operator via a hand held keyboard. The module contains the probe drive 

power supply, speed regulator, and microprocessor control system. This 

module allows repetitive scanning with automatic retraction, time delayed 

repeat and jam detection. Programmed speeds and distance regions allow scan 

speed variation in the U-bend region in both the retraction and insertion 

phases of a scan. No eddy current signal transmitting capabilities are 

included in this module.

2.4 Data Processing Subsystem 

2.4.1 General Design 

- 198 - 

The data processing subsystem is the heart of the R'Eddy Record system. It 

provides user friendly operation of the overall system by means of an 

interactive menu which leads the operator through the entire inspection. 

The major functions of the data processing subsystem are: 

i) Scanning and display parameter set generation through the interactive 

nenu. Up to 16 distinct parameter sets can be created and stored in 

memory. 

ii) Storage and retrieval of scanning and display parameters on digital 

cassette for quick recall. 

iii) Transfer of scan pattern to probe drive controller, 

iv) Data acquisition, 

v) On-line graphic display of data, 

vi) Automatic hardcopy at the end of each scan, 

vii) Tubesheet signal pattern recognition to allow the scan to end at the 

far tubesheet. This allows tubes of varying length to be scanned 

without the probe exiting the far end of the tube and getting caught 

on retraction, 

viii) The generation of record identifiers and distance encoding for both 

strip chart and FM analog recordings, 

ix) The automatic control of FM analog recorders, 

x) Voice synthesis of the tube number (as keyed in by the operator) and 

scan status on scan completion. This voice synthesis can be broadcast 

over a speaker and on the voice channel of the analog recorder, 

xi) Automatic playback of tape records incorporating high speed search for 

tube or print numbers, 

xii) Overall system coordination. 

The data processing subsystem is modular and structured; operator interaction 

is centered around the menu. A main menu page gives the operator a number of 

tasks which the system will execute and which the operator invokes by 

number. Tasks that do not lead directly to scanning automatically return to 

the main menu page with the task listing. Tasks that lead directly to tube 

scanning allow the analyst the option to return to the menu before scanning 

commences. During scanning, allowed operator inputs are displayed on a two 

line display on the operator terminal. The operator can exit the scanning 

sequence back to the menu by single key operation at selected points. 

Data acquisition is based on an interrupt driven data acquisition format that 

ensures a minimum of one point (X,Y) of data is collected for every 0.4 mm of 

probe travel. As data is collected, it is displayed as separated features on 

a monochrome monitor.

2.4.2 Repetitive Scanning 

- 199 - 

The design of a repetitive scan function revolves around maintaining the 

shortest overall inspection turnover time. The system reacts to slip and jam 

conditions of the probe drive in order to prevent probe and probe cable 

damage. Maximum insertion speed of 50 cm/sec and retraction speed of 75 

cm/sec and the use of selected speed/location changes allow the fastest 

possible speed without signal distortion and cable jamming. The need to 

re-scan due to false tube identification is minimized since the system 

recognizes when the tube number has not been updated, alerts the operator and 

suspends any subsequent scan until updating has taken place. 

The operator is alerted to signal deterioration by the calibration signals 

that are presented on the first feature of every display. Due to the nature 

of the display, signal anomalies can be readily picked out by an analyst, 

thus allowing concurrent analysis. An average scan time of roughly 1 

minute/tube, provides enough time for such analysis. Also, to reduce 

inspection downtime, the system can be temporarily controlled by a single 

operator. 

The analyst selects a scan parameter set for use, then enters the main 

inspection sequence through the menu. The operator then positions the probe 

guide at the tube to be inspected and either the operator or analyst 

initiates the scan by depressing the scan 'Ready 1 key on the remote keyboard. 

As the probe enters the tube, data is collected, displayed and stored on 

tape. The insertion sequence terminates with recognition of the tubesheet or 

upon reaching a preprogrammed position. During probe insertion the tube 

coordinates are entered via the remote keyboard. At the end of the probe 

insertion phase the probe automatically retracts, the display hardcopy is 

produced and the voice synthesis announces the scan status and tube 

coordinates. If the tube number is not entered by the end of the insertion 

phase then a 'label' light is lit on the keyboard; the probe still retracts 

but the hardcopy and voice synthesis are delayed until the tube coordinates 

are entered. The end of printing, voice synthesis and probe retraction to 

zero, mark the end of the retraction phase. The operator then repositions 

the probe guide and the sequence repeats. During a scan the operator can 

stop and start probe motion by using the keyboard. The scan can be 

terminated (scan stopped and probe retracted) or aborted (probe drive motion 

stopped pending reinitialization) by the analyst through the terminal. The 

operator and remote keyboard could be replaced by an automatic tubesheet 

walker for scanning tubes in hostile environments such as nuclear steam 

generators. 

3. SYSTEM OPERATION AND DATA DISPLAYS 

3.1 General Operation 

System operation begins with equipment setup. A series of calibration scans 

would be done, which provides a quick mechanism to calibrate the ET 

instrumentation and check overall system operation. The inspection and

- 200 - 

display parameter set could then be loaded from cassette or entered via the 

menu. Feature stepping locations are normally based on the positions of tube 

supports and baffle plates. If these positions are known from drawings of 

the heat exchanger, the parameter sets could be prepared prior to the 

inspection. Otherwise, the information can be acquired by scanning to the 

end of the tube with evenly spaced feature stepping and the locations of 

supports and baffle plates determined from the printout. After the parameter 

set is generated, it would be stored on cassette and repetitive inspection 

scanning would then take place. 

3.2 Generation of a Parameter Set 

The generation of a parameter set is accomplished through the menu. The 

operator responds to a series of questions to determine the following scan 

parameters: 

a) Scan length and speed/location profile. 

b) The location of blind regions, which are regions where the data is not 

displayed on the impedance images and can be used to blank out the 

signals generated as the probe enters the calibration tube. 

c) The number of features displayed and stepping positions. Stepping can 

occur at operator entered positions, evenly spaced positions, on speed 

changes, after blind regions, or automatic stepping upon threshold 

detection. 

d) Options for system control (what to do on slip or jam conditions, and 

whether the operator or analyst or both control the start of scan). 

e) Tubesheet information for tubesheet recognition if required. 

After each parameter set is generated, it is automatically printed for 

permanent reference. 

3.3 Calibration Scanning 

Calibration of the eddy current instrument for the probe in use is aided by 

the R'Eddy Record 'Calibrate and Scan' program. The system requires minimum 

operator input and produces a hardcopy printout, as seen in Figure 3. In 

this case, the operator requested a 9 cm scan with 3 separated features, 

whlci reflects the calibration tube used. This calibration tube contained a 

through-wall hole, an 0D groove and an ID groove. The upper row of Impedance 

features was obtained by plotting X vs Y with stepping occurring as the probe 

passed the 3 and 6 cm positions. Below the impedance features are the 

individual X and Y component traces. A row of 'tic' marks below the Y 

component trace indicates the stepping location of each feature with the last 

'tic' mark denoting end of scan. Below the 'tic' marks lies bookkeeping 

information. The scan parameters used in the calibration run are 

automatically printed after being entered; thus both permanent parameter set 

information and the calibration scan hardcopy are obtained. No tube number 

Is associated with a calibration scan.

3.4 Inspection Scanning 

3.4.1 General Tube Scanning 

- 201 - 

Scanning can take place as soon as a parameter set has been generated and 

calibration has taken place. The repetitive scan task is invoked via the 

menu. Entry back to the menu is available to the analyst at the end of each 

scan. 

A model of a heat exchanger tube is seen in Figure 4. This tube contained 

two defects, a dent located under a broach plate and fretting wear (20% 

eccentric) under a baffle plate. The calibration tube and retract limit 

switch used to zero the distance count are shown. 

The printout collected from the mockup heat exchanger tube scan is shown in 

Figure 5. The first feature contains the calibration signals. The two 

defect signals are easily picked out of the printout. Twelve features were 

requested along with tubesheet recognition. The label 'Full Length Scan 1 

indicates an inspection scan. The label L:555 denotes that the maximum 

length of scan was 555 cm and the label T:555 indicates the tubesheet 

recognition stopped the probe at 555 cm. If desired, up to 50 impedance 

images could be displayed. If data collection and display could not meet the 

0.4 mm data resolution, the message 'Invalid Scan' would be displayed. The 

printout can be written upon (analyst notes) and photocopied. A complete, 

easy to access inspection record is thus available to reassure the station 

manager and regulatory agencies about inspection credibility. 

3.4.2 Scan of Selected Tube Regions 

The system gives the operator a method of scanning up to 10 tube regions with 

expanded signal traces on the X and Y traces of the display. This option, 

called 'Expanded Scan', suppresses the signals from tube regions not 

requested for display. As in the full length scan, the calibration tube 

signals appear in the first feature. An expanded scan display of the two 

regions of the mockup tube containing defects appears in Figure 6. 

4. VERIFICATION OF ANALYSIS AND INSPECTION PROCEDURE 

Verification of both inspection procedure and data analysis requires 

recording the original signals and their interpretation as permanent records. 

Analog FM recording of data including encoding of tube number, print number 

and distance markers, acts as a primary record and retrieval of selected 

records is automated. By using the voice synthesized tube number on the 

voice channel of the FM recorder, a taped record can be retrieved without the 

R'Eddy Record system. The hardcopy printouts of inspection parameters and 

data display records with analyst remarks can be bound and act as secondary 

records. A third record set is available, comprised of standard strip chart 

recordings of data with an ana.log bar code of tube number, print number and 

distance pulses on a separate channel. All these records are automatically 

generated and make verification a straight forward task.

5. PLAYBACK 

- 202 - 

A tape playback program allows easy access to R'Eddy Record data on magnetic 

tape. High speed search for up to 10 records identified by tube or print 

number, with record replay and sequential record display after search 

completion allows great flexibility In playback. Three sets of single or 

multi-frequency data can be retrieved and regular R'Eddy Record displays 

produced. This allows detailed post inspection analysis if required. The 

tape playback does not require the use of the probe drive control subsystem. 

6. SYSTEM APPLICATIONS AND FIELD EXPERIENCE 

The system can be used for a wide range of tube inspections. System set up 

time and calibration time are minimal. The set-up of the Inspection system 

components typically takes 2 experienced personnel approximately 30 minutes. 

Additional setup time is required for instrumentation calibration. An 

experienced inspector can enter a particular scanning parameter set in 

approximately 2 minutes, provided feature stepping distances are known 

beforehand. The highly regular scans and displays are a valuable Inspection 

asset. The capability of providing on-line analysis coupled to a high tube 

scan repetition rate has demonstrated an inspection time saving of up to 40%. 

This is based on three field Inspections, the smallest of which was approximately 

450 tubes. Experience has shown that 400 tubes per 8 hour shift Is a 

reasonable expectation. A scan repetition rate of 1 tube/minute suggests 

that a higher number of inspected tubes per shift can be achieved under 

optimum conditions. Because the system was designed and constructed as a 

prototype, minor problems have occurred, but the overall system performance 

has been very good. 

The probe drive modular design has proven to be very important. Transportation 

and setup are now done more quickly, and with less physical effort. The 

new technology used in the friction drive components has enabled slippery, 

wet push tubes to be pushed and pulled through even partially clogged heat 

exchanger tubes. The jam detect feature on the probe drive controller has 

prevented the drive from damaging push tubes. No mechanical malfunction of 

this probe drive has yet occured. 

An example of inspection data from a condenser is shown in Figure 7. This is 

a direct photocopy of the silver-oxide coated paper used by the system 

printer. An analysis of the eddy current signals indicates there Is substantial 

steam erosion between the second and third tube supports (between 25% 

and 50% of wall loss), and lesser amounts before the first tube support, and 

after the third tube support. A through-wall defect is suspected just prior 

to the second tube support. 

7. CURRENT SYSTEM STATUS 

The R'Eddy Record inspection system is currently in the process of being 

licensed to a manufacturer of eddy current instrumentation. It is anticipated 

this sophisticated inspection system will be commercially available in 

the near future.

s 

KEYBOARD 

HEAT 

EXCHANGER 

„ET PROBE 

PROBE DRIVE PROBE DRIVE COMPUTER OR 

INTERFACE COHIHQLLER CONSOLE ~ 

Figure 1: Configuration of Probe Drive Control Subsystem 

with Probe Manipulation by Operator 

PROBE DRIVE 


|- [g] CASSETTE DECK 

•JÏ I TV MONITOR 

US VIOEO 

PRINTER 

TERMINAL 

KEYBOARD 

CONSOLE 

ZJ (ZJ\ TAPE RECOROER 

STRIP CHART RECORDER 

))) AUDIO 

Figure 2: Configuration of Data Processing 

Subsystem During Inspection 

o

" * 

0 0 

MOkE* 

i 

SUN! SOW 

A 

1 

3* 

~V 

i 

1 

CALIBRATE 

1 

p»oeî 

00 GROOVE * 

J 

A1UH flUN 

* 

10 W00VE 

s«r 

Figure 3: Typical Printout when a Calibration Tube with 

Three Defects is Scanned (Note: *Denotes Comments 

added for Clarity, not Generated by the System) 

3«5 P77Jp77i BAFFLE PLATE 

Si I 

5E0 

S35£ 

555 

Zl BAFFLE PLATE 

1 FRETTING WEAR. 

30% ECCENTRIC 

445 V/>\r7-7\ BAFFLE PLATE 

223 BAFFLE PLATE 

ENVIRONMETAL 

SHICLO 

ID 

0? 

HOLE 

«TRACT -"— 

LIUIT SWITCH 

DISTANCE (cm) 

BROACH PLATE 

|4j 3R0ACH PLATE 

BROACH PLATE 

TUBE 

SHEET 

CALIBRATION 

TUBE 

Figure 4: Mockup of Heat Exchanger Tube 

O 

-P-

x.y FEATURES 

IIMPtOANCE PLANE) 

X-CGMPONENT 

VERSUS OISTANCE 

IT-COMPONENT 

VERSUS OISTANCE 

STEPPING MARKERS 

CALIBRATION' 

SIGNALS 

r 

lr 

SLANT HOW 

123/45 

- 205 - 

IUBESHEH* OENT A! BROACH* • NORMAL BROACH* 

r 

PLAIE PLATE 

U-8EN0 SUPPORT* NORMAL BAFFLE* 20V. FRETTINO WEAR 

PLAH UNOER 8AFFLE PLATE 

r 

C 

PRINT IS 

21-JUH-l) 

LA8tL HOTl.tG-1 

FREU 210 KHZ 

L VA T VA 

ENVIRONMENTAL 

SHIELD 

FULL iCNGln SCAH 

TUBESHEET 

c 

afOO» 

RECORO 

(BNt 

Figure 5: Printout Generated when Testing the Mockup Tube. 

(Note': *Denotes Comments Added for Clarity, not Generated by the 

System). 

123/45 

\i—• 

Figure 6: Printout Generated by an Expanded Scan of Two Selected Regions 

of the Mockup Tube Containing Defects. Only the Calibration 

Signals (0-9 cm region) and Signals in Regions 70 to 90 cm and 

385 to 405 cm are Displayed. 

(Note: ^Denotes Comments Added for Clarity, not Generated by the 

System). 

V

SLHHT POU 

422/14 

- 206 - 

x •* 

F-RIHT 33 PRO8t:yC015 ) 

15 FIJLL Lt:!GTH :t 

•d>) 

R-EDDY 

RECORD 

CRML 

Figure 7: Photocopy of Printout Generated During Actual Field Inspection 

using the R'Eddy Record System. Comments were Written Directly 

onto the Printout by the Analyst.

- 207 - 

WET CHANNEL MEASUREMENT OF PRESSURE TUBE TO CALANDRIA TUBE 

SPACING IN CANDU REACTORS 

J.tf. Sado 

Ontario Hydn.0 

ToA.on.to, Ontario 

ABSTRACT 

The pressure tube (PT) to calandria tube (CT) spacing in CANDU reactors is an 

important parameter that relates to the general condition of the fuel 

channels. The measurement system that was developed to measure this parameter 

during the wet channel inspections of Pickering Units 1 and 2 is described in 

this paper. A send-receive eddy current probe was designed which is 

primarily sensitive to variations in PT/CT spacing but is also affected by 

pressure tube wall thickness. A computer simulation showed that the phase 

angles of the response to these variables are similar for all usable 

frequencies, thus eliminating the possibility of mul tifrequency compensation. 

A marriage of technologies was proposed involving the ultrasonic measurement 

of wall thickness values which are then used to extract the spacing 

information from the eddy current signal. The accuracy of the system is 

aproximately ±(30% +.lmm) which has been sufficient to determine if and where 

any of the pressure tubes have come in contact with their calandria tube. 

Field experience with the new system is discussed and areas for development 

are also outlined. 


Each fuel channel in commercial size CANDU Reactors is installed horizontally 

and contains an unsupported center span approximately 6 meters long 

consisting of a pressure tube surrounded by a dry gas and enclosed by 

calandria tube. These tubes are firmly centered at the rolled joint positions, 

that is at the ends, but spacers (garter springs) are required to ensure 

proper separation in the center span. After the reactor has been in operation 

for some time, the pressure tube will exhibit sagging between the garter 

springs as shown in Figure 1. However, if a garter spring is out of position 

then accelerated sagging may occur resulting in premature pressure tube (PT) 

to calandria tube (CT) contact. 

Two techniques of measuring the PT/CT spacing (gap) have been investigated 

during the development of a system for the remote inspection of pressure tubes 

(CIGAR) /I/. One method processes the output of an eddy current probe during a 

full 360° rotational scan in order to arrive at an estimate of the PT/CT 

eccentricity. The second method utilizes both an eddy current probe and an 

ultrasonic probe to produce a direct measurement of the change in gap from a 

reference point.

- 208 - 

The recent failure of a Zircaloy-II pressure tube in Pickering Unit 2 

resulted in a decision to construct a simplified version of CIGAR 

(CIGARette)/2/ in order to inspect the condition of the remaining fuel 

channels in both Units 1 and 2 . As part of the program it was decided to 

develop the direct gap method since it showed the potential to be more 

accurate over a wider range of values and because the measured data 

corresponds in a more straightforward manner to the unknown gap. This is an 

important advantage during the interpretation of data and also when trouble 

shooting in the field. 

2. THE EDDY CURRENT PROBE 

The eddy current gap probe contains two differentially connected receive coils 

encircled by a send coil as shown in Figure 2. All of the materials used in 

the construction of the probe are radiation resistant in order to withstand 

the 10 rad/hour fields in the shutdown reactors. Although the probe was 

originally designed for use in fuel channels with Zr-2.5% Nb pressure tubes, 

the offsetting wall thickness and resistivity changes in Zircaloy-II tubing 

result in similar eddy current behaviour in fuel channels of Pickering 1 and 2 

and allow the existing design to be used. 

2.1 Computer Predicted Response 

The complex plane response of the probe (Figure 3) was simulated by a 

computer program based on the equations and coding developed by Dodd for a 

similar probe configuration over multilayered planar conductors/3/. The 

diameter of the gap probe is small in comparison with that of the pressure 

tube and so it was assumed that the planar case would closely approximate the 

actual situation. Another departure from the configuration used by Dodd is 

that instead of placing the receive coils symetrically within the send coil, 

their position is adjusted during manufacture so as to produce a minimum 

signal when the probe is in contact with a piece of Zr-2.5% Nb pressure tube. 

This situation more closely approximates the probe's normal operating 

environment thus minimizing the offset signal and making available a larger 

overall gain from the eddy current instrument. These approximations plus the 

neglect of the computer program to take into account the effect of the 

stainless steel wear cap that surrounds the probe may result in some departure 

of the actual response from the predicted response. 

The frequency and phase settings of the eddy current instrument are chosen so 

as to place the signals due to lift off changes on the X axis and those due to 

gap changes predominantly on the Y axis. Therefore, monitoring only the Y 

component of the eddy current signal practically eliminates any lift off 

influence and substantially reduces that of pressure tube resistivity. Signals 

due to pressure tube wall thickness changes are also primarily in the Y 

direction, but the exact phase angle is influenced somewhat by the magnitude 

of the PT/CT spacing beneath the probe. This condition persists over a wide 

range of operating frequencies, illustrating why multi frequency techniques 

were originally rejected for wall thickness compensation. Instead, an 

ultrasonic system will be used to measure the pressure tube wall thickness 

allowing the gap information to be recovered from the eddy current signal.

2.2 Primary Factors 

- 209 - 

Computer simulation results verified that vectorial addition was valid for 

signals due to gap and wall thickness changes, thus allowing these effects to 

be studied independently. A number of sets of data were collected, each at a 

different fixed wall thickness and containing the response of the eddy current 

probe to a range of gap values. One such set is shown in Figure 4 where the 

non-linear response of the probe is evident. In order to relate these 

independent data sets and determine the effect of wall thickness changes, it 

is only necessary to determine the response of the probe to wall thickness at 

one fixed gap. This was done by performing a rotational scan on a pressure 

tube sample with a large wall thickness variation around the circumference and 

no calandria tube (infinite gap). The result, shown in Figure 5, was a good 

linear fit with a regression coefficient of 0.9887. The slope of this line 

was used to shift the sets of previously collected data so that they all 

referenced a common balance point at infinite gap and 5.0 mm wall thickness. 

Straight lines were fitted to the points of common gap with a great deal of 

success, the smallest regression coefficient being 0.9960. This overview, 

shown in Figure 6, formed the data base of the probe's response which was to 

be used during field inspections. 

2.3 Pressure Tube Resistivity 

The calculated complex plane response of the probe indicated that pressure 

tube resistivity could significantly influence the eddy current signal if 

large enough variations were encountered during the inspections. This 

possibility was explored by measuring the resistivity and wall thickness 

around the circumference of a pressure tube sample and then scanning the same 

section with the gap probe in order to compare both influences. 

The resistivity scan in Figure 7c shows a well defined change of about 1 

just after the 180° mark. Using Figure 3, it was calculated that a 

resistivity change of this magnitude should cause approximately the same 

effect in the Y component of the probe response as a 0.17 mm wall thickness 

change. Figure 7a shows that this is about half of the total wall thickness 

variation measured in the pressure tube sample. However, the response of the 

gap probe shown in Figure 7b does not agree with this prediction as the wall 

thickness influence was followed very closely with almost no sign of the 

resistivity influence. 

The apparent insensitivity of the probe to resistivity variations was 

encouraging for the immediate goal of the Pickering 1 and 2 inspections, 

however, further investigations are required for the long term development of 

the gap measurement system. 

2.4 Temperature 

The temperature sensitivity of the probe in its operating environment and in 

isolation are shown in Figures 8 and 9 respectively. The response in Figure 9 

is somewhat unexpected since send-receive probes are generally expected to be 

insensitive to temperature variations due to the small effect that variations 

in coil resistance have on the output signal. In this case the response is 

thought to be caused by a physical deformation of the probe and is currently 

under investigation. The linear response of -0.036V/°C obtained from the

- 210 - 

probe in its operating environment is the combined result of this effect plus 

the temperature related resistivity changes in the pressure tube and in the 

stainless steel wear cap. 

The temperature variations in each fuel channel are expected to be minimal due 

to the flow of coolant that exists during shutdown. As a result the 

temperature sensitivity of the probe should not cause finy major problems 

during inspections and may be covered by the stated accuracy of the results. 

2.5 Probe Variations 

The data presented thus far were obtained with one specific probe. 

Experiments comparing subsequent probes to the original showed a constant 

difference in signal outputs due to identical gap variations. The wall 

thickness and temperature sensitivities of the alternate probe were also 

verified to differ from the reference probe by the same factor. This allowed 

the assignment of a compensation factor to each probe generally ranging in 

value from 0.7 to 1.1, so that for an identical set of circumstances the 

change in output of the alternate probe could be referred to a standard. 

3. GAP ESTIMATION 

The procedure for estimating gap assumes that lift off, resistivity, 

temperature or any other factors do not influence the eddy current signal 

obtained in channel. Before the inspection data can be used, these signal 

values must be related to those recorded in the data base, and so it is 

necessary to select a reference point in each channel where gap, wall 

thickness and the eddy current signal value are known. This generalized 

requirement does not necessarily involve the point at which the instrument is 

balanced, thus reducing the function of balancing to one of locating the 

output trace somewhere on the strip chart record. The wall thickness and eddy 

current values may be measured at the reference point but the gap must be 

inferred from the structure of the fuel channel. Two good reference points 

are located near the ends of the pressure tube, as close as possible to the 

rolled joints. Here the pressure tube is centered firmly in the calandria 

tube and the gap may be assumed to be 8 mm at all rotational positions. 

Once a reference point has been chosen in the channel, the corresponding point 

in the data base may be located using the wall thickness and gap information. 

From this point, the changes in eddy current signal and wall thickness data, 

as measured when the probes travel from the reference point to the measurement 

point in-channel, can be followed individually to arrive at the mesurement 

point in the data base. It is quite likely that the measurement point in the 

data base will end up between two lines of constant gap in which case simple 

linear interpolation is used to determine the estimated gap value. The number 

of lines of constant gap is large enough that this approximation to the true 

non-linear nature of the gap response will not introduce any significant 

error. This procedure may easily be implemented on a computer using only the 

stored slopes and intercepts of the lines in the data base.

4. FIELD EXPERIENCE 

- 211 - 

The gap measurement subsystem was incorporated into the CIGARette inspection 

equipment along with garter spring and flaw detection capabilities for use in 

Pickering Units 1 and 2/2/. Unfortunately the gap measurement function was 

not one that worked right away. 

4.1 Wall Thickness Measurements 

Wall thickness measurements were originally to be obtained using a modified 

digital thickness gauge and a highly focussed radiation resistant ultrasound 

probe. This method is similar to that planned for the CIGAR inspection system 

with the exception that a technician would be responsible for logging the 

data from the digital display instead of a computer. Unfortunately, stable 

readings could not be obtained and eventually the instrument used for the flaw 

detectors, a Branson KB-6000 was utilized. Thickness measurements using this 

modified system required that the probe be stationary so that the time beween 

the reflection peaks of the inner and outer pressure tube walls could be 

measured on a CRT display. The result was then converted to a thickness value 

by multiplying by 2.389 mm/p s (the velocity of sound in Zircaloy-II divided by 

2). This slow procedure limited the number of points that could be measured 

in each channel and introduced a random observation error that was estimated 

to be about±l% or a little more than the thickness of the screen trace. 

4.2 Eddy Current-Signal Quality 

Although the procedures for calculating gap outlined in section 3 indicate 

that data need only be collected at the reference and measurement points, it 

was soon realized that the eddy current signal would have to be recorded 

during a continuous scan in order to better ascertain the signal quality. 

Once this procedure was adopted, it became evident that in many cases the eddy 

current signal contained excessive noise and drift which was traced to 

various problems with the 80 meters of signal cabling and 8 connection points 

in the CIGARette system. Finally, diligent trouble shooting and the 

availability of tested spare components reduced this problem significantly, 

resulting in smooth and stable signal traces for many of the channel 

inspections. 

4.3 Types of Scans 

Two types of scans were used during the inspections, an axial scan along the 

bottom of the channel to establish a gap profile and a rotational scan, 

usually used at points of interest in an attempt to confirm the axial scan 

results. The rotational scans were evaluated assuming an infinite gap at the 

12 o'clock position but were generally not as successful as *he axial scans. 

The reasons for this difficulty were never firmly established, however the 

quite rapid wall thickness changes around the circumference of the pressure 

tube and the relatively few data points collected required exacting 

performance from the system for proper results. Axial scans were much more 

successful with the added benefit that the eddy current signal trace was 

usually indicative of the gap profile since very few significant axial wall 

thickness variations were encountered.

- 212 - 

The eddy current and wall thickness probes are located at the same rotational 

position on the inspection head, however they could not be used simultaneously 

during an axial scan due to the conflicting requirements of the two systems 

(the eddy current scan must be continuous whereas the ultrasonic probe must be 

stopped for each measurement). This separation of the data collection scans 

caused some problems on numerous occasions as the inspection drive mechanism 

would tend to slip rotationally and corkscrew without affecting the rotational 

display counter. This had serious consequences for the gap system since a 

small rotation in a typical pressure tube during an axial scan would result 

in a sharp rotational wall thickness change being sensed as a gradual axial 

change when in fact there was none. Such a situation, especially if only 

experienced during the eddy current or wall thickness scan, would result in 

misleading gap results. 

4.4 Sample Results 

The axial gap profile of a fuel channel exhibiting accelerated pressure tube 

sag due to the shift of a garter spring from its design location is shown in 

Figure 10. The gap system was originally conceived to discover situations 

such as this where the PT/CT spacing is very small in some areas. 

4.5 Sources of Errors 

Errors can generally be grouped into two categories, those that affect the 

accuracy of the measurements and those that affect the reliability of the 

system. The latter group tend to severely distort the data and would include 

such conditions as; cabling problems, misalignment of the wall thickness and 

eddy current probes, operator or analyst error and excessive temperature or 

resistivity changes. Many times these problems will cause the data or results 

to be unusual or even unreasonable when compared to historical data and known 

characteristics. Some checks that can be made to give credibility to the data 

are listed here: 

1. The eddy current signal trace should be smooth with no 

abrupt discontinuities or excessive noise or wandering. 

2. The wall thickness data is usually fairly constant in an 

axial direction. 

3. Some sign of both gap and wall thickness changes should 

be seen in the eddy current signal trace since it is 

highly unlikely that gap and wall thickness would vary in 

exactly opposing manners so as to cause the eddy current 

signal to remain constant. 

4. The gap profile should be fairly smooth. 

5. The gap near the rolled joints at both ends of an axial 

scan should be 8 mm.

- 213 - 

6. The axial gap profile should be similar when computed 

using either end of the pressure tube as a reference. 

Note that if the gap at both ends of the pressure tube 

are not equal to 8 mm, repeating the calculations using 

the other end as a reference will not just shift the gap 

profile by the difference. The shape of the profile will 

also change due to the non-linearity in the data base. 

The accuracy of the gap measurements is primarily determined by the effect of 

the ±0.05 mm uncertainty in the wall thickness measurements at both the 

reference and measurement points. Referring to Figure 3, the uncertainty in 

the location of the reference point, which must be located along the line of 

constant 8 mm gap for axial scans, can also be expressed in terms of the eddy 

current signal voltage as approximately ± 0.08V. Doing so allows the 

uncertainty to be transferred to the measurement point and to consider the 

reference point fixed. The additional uncertainty in the measurement point 

due to its associated wall thickness measurement can also be expressed in 

terms of the eddy current signal, but the exact conversion factor will depend 

on the size of the gap. The wall thickness measurements can be assumed to be 

indépendant and normally distributed and so both of the uncertainties in the 

measurement point may be combined as the square root of the sum of the 

squares. The result will vary from a high of +0.11V for a gap of 8 mm to a 

low of +0.083V for a gap of 0 mm. An error bound of ±(30% + 0.1 mm) on 

computed gap values approximates this overall uncertainty and is the magnitude 

of the error bars that have been placed around the sample results in Figure 

10. 

S. AKEAS ft» DEVELOPMENT 

The field experience in Pickering has served to solidify the gap measurement 

procedure and also to uncover various weak points in the equipment which 

effect the results. Further development is being done using the CIGAR 

inspection system with the emphasis on Zr-2.5%Nb pressure tubes now that the 

Zircaloy-II tubes in Pickering Units 1 and 2 are to be replaced. 

The use of the CIGAR inspection system addresses a number of equipment 

reliability problems encountered in Pickering which made it difficult to 

isolate and solve problems originating with the measurement technique. The 

most important advancement is that both the eddy current and wall thickness 

data are collected automatically during the same axial scan. This ensures the 

proper alignment of both probes and enables a large amount of data to be 

collected which can then be smoothed, reducing any unbiased error in the 

individual measurements. Another improvement is the gentler and largely 

automatic handling of the signal cables and their connections which has 

resulted thus far in stable and reliable eddy current signals. Very promising 

and repeatable results have been obtained to date but more testing is still 

required. Some development work still remains to be done in the area of 

pressure tube resistivity variations and in the quality of the eddy current 

probe, both in terms of its temperature sensitivity and its consistency in 

manufacture.

6. SUMMARY 

- 214 - 

The failure of a pressure tube in Pickering Unit 2 created a need to measure 

the pressure tube to calandria tube spacing in the remaining Zircaloy-II fuel 

channels in order to help ascertain their condition. A technique, previously 

investigated for CIGAR and which utilizes ultrasonic wall thickness 

measurements to compensate the response of an eddy current probe was 

integrated into the CIGARette inspection system. Some procedural and 

reliability problems were initially experienced but eventually the system was 

made to work with an accuracy of about ±(30% + .1 mm). This was sufficient to 

be able to discern regions where the pressure tube had come in contact with 

the calandria tube. Work is continuing to integrate this capability into the 

CIGAR inspection system for use with the rest of the commercial size CANDU 

reactors. It is hoped that the use of CIGAR will increase both the 

reliability and the accuracy of the measurements although final results are 

still pending. 

ACKHOWLEDGEMEHTS 

Appreciation is extended to Dr. D. Leemans of Cross Currents Research & Policy 

Consulting Ltd for discussions on the eccentricity method and the design of 

the gap probe and to H. Ghent of CRNL for his assistance with the computer 

simulation of the eddy current response. 

REFERENCES 

1. J.A. Baron, M.P. Dolbey, J.H. Erven, D. Booth, and D. Murray, I. Mech. E. 

Conf. Publication No. C139/82. 

2. M.D.C. Moles, M.P. Dolbey and K.S. Mahil. Paper presented at this 

conference. 

3. C.V. Dodd et al. 0RNL Report No. TM-4107, Oak Ridge, Tennessee, 1973.

ROLLED JOINT 

CALANDRIA TUBE 

rs.5 

- 215 - 

PRESSURE TUBE 

6.03 H 

.5 

MM, 

ROLLED JOINT 

FIGURE 1 FUEL CHANNEL EXHIBITINC PRESSURE TUBE SAC 

Pressure Tube 

Recei ve 

Coils Send Coil 

FIGURE 2 

Send 

Coil 

THE EDDY CURRENT GAPPROBE 

ä 

3 

Receive 

Coils 

(Wound in 

Opposition ) 

> =J Pressure 

Calandria Tube Tube

+2.5%p 

Lift Off •* 

- 216 - 

+5% Wall Thickness 

-5% Wall Thickness 

-5% Wall Thickness 

FIGURE 3 

Cap = 8 mm 

p = 74 \iti-cm (Zirc II) 

Wall Thickness = 5.0 mm 

COMPUTER PREDICTED COMPLEX PLANE 

RESPONSE OF THE CAP PROBE 

-2.5%p

in 

in Vol 

lent i 

Compor 

10 

c 

o 

o 

0 

-0 .5 

-1 .0 

-1 .5 

-2.0 

0.3 

-0.3 - 

- 217 - 

Gap in mm 

FIGURE 4 

Zirc II 

Wall Thickness =5.Omm 

RELATIONSHIP BETWEEN EDDY CURRENT SIGNAL 

AND GAP (NOT ALL DATA SHOWN) 

Zirc II 

Cap = S mm 

Slope = -1.71 V/mm 

R = 0. 9887 

5.1 5.2 5.3 5.4 

Wall Thickness in mm 

FIGURE 5 

RELATIONSHIP BETWEEN EDDY CURRENT 

SIGNAL AND WALL THICKNESS 

5.5

0 

-0.2 

-0.4 

-0.6 

-0.8 

-1.0 

c .1 2 

o 

o -1.4 

-1.6 

-1.8 

-2.0 

-2.2 - 

-2.4 

5.0 

5.1 

- 218 - 

Wall 

_L 

5.2 

Thickness 

5.3 

in mm 

Cap in mm 

5.4 5.5 

FIGURE 6 

EXPERIMENTALLY DERIVED DATA BASE USED FOR THE 

COMPUTATION OF CAP IN ZIRC II CHANNELS 

(NOT ALL LINES SHOWN)

u 

- 219 - 

180° 

Rotational Position 

FIGURE 7a WALL THICKNESS OF PRESSURE TUBE SAMPLE 841-207(ZIRC II) 

-375 

180° 


360' 

FIGURE 7b SCAN OF SAMPLE 841-207 USING THE EDDY CURRENT CAP PROBE 

76.8 

0° 180° 360 e 


FIGURE 7c RESISTIVITY OF PRESSURE TUBE SAMPLE 841-207 

360--

0 

-0.25 

-0.50 

-0.75 

-1.00 

-1.25 

i 

1/1 

• 

1 

c 

iponeni 

Con 

> ~ 

- 220 - 

Rate of Temperature Rise 

0.5°C/min 

Slope = 

-0.036V/°C 

10 30 50 

Temperature in °C 

FIGURE 8 

TEMPERATURE SENSITIVITY OF THE 

CAP PROBE MOUNTED IN ITS HOLDER 

AND AGAINST ZIRC II TUBING 

.00 

0 .75 

0 .50 

0 .25 

0 

Slope = 0. 035 V/°C A 

/ 

\f i I i 

10 30 50 

Temperature in °C 

FIGURE 9 

TEMPERATURE SENSITIVITY OF THE 

GAP PROBE ALONE

»6 

5 

^ 4 

13 

2 

IL 

0 

- 221 - 

•Rolled Joints 

£t 11Uiiiiitii 

Detected Positions 

Of Carter Springs 

X 

1,X 

Rotational Position = 180° 

Error Bars are t (30% + 0. I mm) 

t K 

3 4 5 6 7 

Axial Position As Read By Cigarette Drive (m) 

FIGURE 10 

GAP RESULTS FROM CHANNEL P1-G14 

I 

f A •. A 

X X X 

X

- 222 - 

CHECKING FOR CRACKS IN GAS TURBINE ROTOR DISCS 

J. ran den Aude I and 4.8. Niebciq 

d'eifiufflicuic Canada Inc., Hamilton, Cnta'iic 


A gas turbine consists of a compressor and a turbine. The turbine drives the 

compressor and a mechanical output shaft. The compressor becomes warm (from 

compressing air), the turbine becomes hot from the flame which burns in the 

compressed air in the combustion chamber. The higher the flame temperature, 

the higher the turbine's efficiency, but the material is also used closer to 

its limitations. The turbine discs (Figure 1) appear to have a finite life 

and although most may last seemingly forever, there is still a fair percentage 

that does not. This means that discs are considered to fail eventually 

and that a close watch is to be kept on the disc quality at overhauls, usually 

in the field. The type of failure is most often a crack in the tooth-serrations 

(Figure 2) and so far, only two nondestructive tests appear useful 

for the detection of root cracks, i.e., Pénétrants and Eddy currents. The 

former is tedious, messy and difficult to evaluate in tight locations, the 

latter is quicker, more definite and very sensitive, but is affected by material 

edges and can therefore not test the entire length of the serration. 

2. PENETRANT TESTING 

The third party is now faced with a possible residue of two previous penetrant 

applications in a crack and has to go through an extensive cleaning process 

which might or might not remove all the old penetrant residue. 

The cleaning process involves glass beading to remove scale, vapour degreasing 

to dissolve trapped penetrant and perhaps a wash with acetone to enhance 

a difficult cleaning process. At this point, it should be realized that we 

are dealing with an entire turbine rotor from which only the turbine blades 

were removed. Turbines come in different sizes, but a length of 7 m and a 

mass of 12 tonnes, is not uncommon. 

The pénétrants currently used are biodegradable and water washable. It was 

found that the post emulsifiable pénétrants although superior when properly 

used, lose all of their advantages when the emulsif1er is left on too long. 

Apart from having no control over an operator in the field, the operator has 

a huge component on his hands that has to be treated in its entirety, left to 

soak a precise time and then wash. Such a procedure is impossible in the time 

slots available. Water washable pénétrants are more tolerant and appeared to 

be much better suited for the inspection of large components.

- 223 - 

After washing, drying and developing, the examination is another problem. 

Often a number of rows of teeth have to be examined and accessibility is limited; 

an operator has to work with tiny pieces of mirror attached to popsickle 

sticks which he tries to position in the bottom groove, shine the black light 

onto them and look at the same time for defects. 

Even if defects were found, the familiar relationship between the crack depth 

and the amount of penetrant-bleeding is often not there because of the very 

tight nature of this type of defect. The cracks may be intergranular and 

tightly filled with combustion or corrosion products which did not wash away 

in the cleaning operation. The small indication of Figure 3 proved to be 

from a substantial flaw (Figure 4). 

Operators with less sophisticated cleaning facilities would find even less 

cracks. This becomes a problem where no cracks are found in a disc which 

does contain them. Fortunately, small cracks are not instantly causing a catastrophic 

failure; most cracks grow slowly and are found during the next overhaul 

at which time the disc can be replaced. Nevertheless, a flawed disc 

which was given a clean bill of health because the testing method was inadequate, 

may fail before the next overhaul and cause expensive damage. The test 

therefore is inadequate and better testing means ought to be found. 

3. EDDY CURRENTS 

An obvious solution to the testing of disc teeth seems to be an Eddy Current 

test. However, these tests are not so obvious when one actually tries them. 

Rotors become slightly scaled and the rough surface so far has caused problems 

with the Eddy Current test. Descaling by means of glass-beading has been 

practised before an Eddy Current test was attempted and leaves a sort of rough 

surface. 

Figure 5 shows an Eddy Current signal of a defect-free tooth. There is some 

horizontal wandering and an exit signal veering off to the upper left. Figure 

6 shows the signal when the probe passes over a defect close to the exit. The 

display shows loops which betray defects. The sensitivity setting of the 

instrument was relatively low and when increased to find small defects, the 

displays would grow much larger than the screen and could no longer be interpreted 

(Figure 7). 

The Eddy Current tester was adjusted such that the lift-off signal as well as 

the wandering of the dot while the probe travels were horizontal on the screen. 

The vertical component is then always related with defects and end-of-groove 

effects. The vertical component has an output available on the instrument 

which was connected to the Y direction of an ordinary storage osciloscope. A 

relatively simple fixture was made to sense the position of the probe and fed 

to the x-direction of the scope. Figure 8 shows how a differential probe senses 

a crack halfway along the tooth. The tooth is 91 mm long, the crack 13 mm 

but it shows 20 mm on the base line which means that the probe sees 4 mm more

- 224 - 

on the fringes. (The sensitivity setting on the instrument was very low, hence 

the defect would show even larger with the sensitivity increased to normal 

testing levels.) 

4. VERIFICATION 

ïhe position of the cracks depends on the design of the turbine. Depending on 

the model, the experienced operator could tell ahead where the cracks will be 

found. That means that a probe can be designed with the highest sensitivity 

at the location of the anticipated cracks. Nevertheless, one needs a probe 

which covers a reasonably wide swath, in case defects with another orientation 

are present which ought to be detected. 

Test blocks were manufactured from homogeneous teeth by Electro Discharge 

Machining small artificial defects in the grooves (Figure 9)• One groove contained 

notches of different depths for calibration purposes, another groove 

contained identical notches which were offset to determine field-of-view of 

the ET probes. 

It was found before, that absolute probes have advantages when only one phenomenon 

is sought. 1 An absolute probe was made with both a wide and a limited-length 

field of view. A once-through test with this probe on the reference 

notches with respect to their position is shown in Figure 10. 

By changing the frequency, the optimum response was found, and it was not surprising 

that it was close to the design frequency. It was expected that a 

change in ratio of signals from small to large reference notches would also be 

found with a change of frequency but it was not (Figure 11). 

Since notch 2 is supposed to be exactly 0.2 mm deep and notch 3, 0.4 mm deep, 

the ideal signal ratio would be two. From 800 to 900 kHz the ratio was 2.3, 

considered the best obtainable. It was also decided that it was probably better 

to use a firm signal as a set-up rather than a smaller but noisier one and 

the 0.4 mm deep notch was a tentative choice. 

At 800 kHz a series of measurements with different instrument sensitivities 

was made (Figure 12). This showed that without increasing the noise level, 

signals could be boosted to show the 0.1 mm notch (No. 4, Figure 9) clearly. 

With this level of sensitivity, any dangerous crack would cause the indication 

at the lower sensitivity. The latter appears better for testing since too 

many indications would go off-screen if the sensitivity were too high. With a 

reduced sensitivity test may be more realistic and could be used to as close 

as 5 mm from the end. (Notches are at 20, 40, 60 and 78 mm from the left 

edge, Figure 13.) With this setting, the real crack in Figure 13 showed a 

start in steps but is fully established at 50 mm and carries on to the end. 

(The microscope showed it to fade at 4 mm from the end or 87 mm from the left 

which is revealed in the picture although not clearly.) 

1. G. Van Drunen and V.S. Cecco, CSNDT Journal, Vol. 5 No. 1, 1983 October

5. CONCLUSIONS 

- 225 - 

A practical Eddy Current test for disc root cracks is possible with a sensitivity 

high enough to detect 0.1 mm deep imperfections. The position of these 

defects is shown on the screen of the scope, using relatively simple position 

fixtures which have to be made for every type of turbine disc. The defect 

size is displayed as distinct measurable spikes, while testing is quick and 

precise. Pictures made of questionable teeth are self explanatory and can be 

compared with those made of the reference. 

Limitations of the test are the lack of definition within 5 mm from the end 

and a field-of-view, limited to only approximately A mm wide. 

Overall, a greatly simplified more precise and much quicker test is now available 

for testing blade roots in discs for cracks. 

6. ACKNOWLEDGEMENTS 

The encouragement given by Messrs. M. Metala and E. Fuselier of Westinghouse 

Electric Corporation have been greatly appreciated.

Figure 1 18 MW gas turbine rotor, partially de-bladed. 

S3

- 227 - 

Figure 2 Multiple power stage

- 228 - 

Figure 3 Weak penetrant indication on carefully prepared sample. 

II 

Figure 4 Previous small 

indication was a large defect.

Figure 5 

- 229 - 

Eddv Current sienal of defect-free groove. 

Figure 6 Large multiturn loop runs into exit signal.

- 230 - 

Figure 7 All signals superimposed. 

Figure 8 Vertical Eddy Current signal displayed on probe position.

REF NOTCHES ON 

CRACK LINE 

CENTRIFUGAL 

FOHCE 

DISC 

Figure 9 

- 211 - 

Location of reference notches. 

LINE WHERE CRACKS OCCUR 

NOTCHES OFF CRACK LINE TO 

DETERMINE FIELDOF VIEW 

• STANDARD REF. NOTCH 

2 mm long 

0.1 mm wide 

0.2 mm deep 

Figure 10 Test on groove with reference notches (absolute probe).

- 232 - 

Figure 11 Frequency response on reference notches.

- 233 - 

Sens. 

35 

Figure 12 Influence of sensitivity setting on output. 

45 

55

- 234 - 

row 3 sent to 0 4 n 

1Vor?d.w 

v 3 wi«. to 0 à mm notch 

ObVorl div. 

(not»; testing prwible ui 5mm from 

ind. blade width 91 mini 

rtrf notch sensitivity ai -iljon- 

MMI cr.jck from 50 mm to tn>l 

w:n4 20 

Figure 13 Two possible sensitivity settings and a real crack.

- 235 - 

EDDY CURRENT INSPECTION OF MILDLY FERROMAGNETIC TUBING 

(U.R. Mayo, J.R. Cantan. 

Atomic Enzigy o {, Canada 

Chalk Rivzn, Ontatii.o 

ABSTRACT 

The past decade has seen the development of eddy current probes for 

inspection of the mildly ferromagnetic alloy Monel 400. Due to the rapid 

advances in permanent magnet technology similar probes have been upgraded to 

magnetically saturate, and hence inspect, the duplex stainless steel Sandvik 

3RE6O, which has saturation induction more than twice that of Monel 400. 

Prototypes of these probes have been tested in three ways: saturation 

capability, quality of typical eddy current data, and ability to eliminate 

permeability induced signals. Successful laboratory testing, potential 

applications, and limitations of these type probes are discussed.


- 236 - 

In-service eddy current testing (ET) of heat exchanger tubing is often 

complicated by the use of magnetic materials such as nickel-copper alloys, 

austenitlc-ferritic stainless steels, and mild carbon steel. Since many 

factors must be considered when a choice of tubing material is made for a 

particular unit, including initial cost and resistance to deterioration, 

inspectability is often of secondary importance. For this reason the alloys 

Monel 400* and Sandvik 3RE60** are encountered as examples in the nuclear 

industry, and mild carbon steel is used extensively throughout industry in 

general. 

It is commonly known that magnetic materials can only be inspected by ET once 

they have been magnetically saturated[l]. This condition can easily be met 

during manufacturing testing, but places a severe restriction upon in-service 

inspection, where usually only the tube ID is accessible. 

We may non-rigorously define materials as being mildly ferromagnetic if they 

have saturation induction less than 1 tesla. This puts the alloys Monel 400 

and 3RE60 in the mild class. Carbon steel, on the other hand, is highly 

ferromagnetic, having saturation induction over 2 tesla. A considerable 

amount of ET development work in the Canadian nuclear industry has been 

directed toward saturation probes for inspection of mild magnetic materials. 

This was largely due to the use of Monel 400 in the Pickering Generating 

Station heat exchangers. Probes to inspect Monel 400 tubing have seen 

successful use in the field, and recently we have succeeded in developing 

probes for 19 mm 3RE60 tubing, which is more difficult to saturate. The 

ability to inspect 3RE60 tubing is very likely the highest level to which 

probes of this type, using high energy-product permanent magnets, can be 

elevated. Highly ferromagnetic mild carbon steel will require some new 

approach and is currently the subject of laboratory investigation. 

This paper presents data obtained from a prototype probe suitable for 

inspection of 19 mm 3RE60 heat exchanger tubing. Impedance monitoring shows 

the magnetic tubing is fully saturated by the probe. Comparison of 3RE60 

signals (eddy current impedance plane display) to those from a nonmagnetic 

316 stainless steel tube of the same dimensions also demonstrates complete 

saturation. Finally, stress induced permeability variation signals 

effectively "dissappear", as would be expected at full saturation. The 

results indicate 3RE60 tubing of this size is now inspectable (with 

limitations), representing a further advancement in the in-service inspection 

of ferromagnetic materials. 

* Reg. trademark of International Nickel Co. 

** Reg. trademark of Sandvik Tubular Prod.

2. BACKGROUND 

2.1 Theory 

- 237 - 

The thickness of conducting material inspectable by ET Is limited by the 

exponential decay of eddy current density with depth. This is characterized 

by a decay constant known as the "standard depth of penetration" (6), given 

by [2]: 

where: 

6 = J— (1) 

u ~ electromagnetic field angular frequency 

a » electrical conductivity of medium 

y - magnetic permeability of conducting medium 

As expressed in equation (1), 6 is the depth at which the alternating field 

intensity, and hence eddy current density, falls to 1/e (e = 2.71828...) of 

its surface value. (This is satisfactory as an approximation, but is only 

exact for a semi-infinite conducting medium with uniform external magnetic 

field intensity). Since 6 is inversely proportional to the square root of 

permeability, it must be much less in magnetic materials than in nonmagnetic 

cases. * The former have permeabilities that are typically orders of 

magnitude higher than the latter. This is indicative of one of two problems 

associated with ET in magnetic materials; eddy currents cannot penetrate the 

specimen sufficiently. Since the eddy current technique relies upon the 

presence of defects to perturb the flow of currents in the metal, there can 

be little sensitivity to sub-surface defects if eddy current density is 

vanishingly small. The application of a biasing field will increase the 

depth of eddy current penetration, by effectively reducing the permeability 

of the sample, but here the second problem arises. Large amplitude signals 

due to permeability variations will be detected by the ET probe unless 

saturation is complete. Apart from these signals overwhelming the 

conventional ones, we must cope with the fact that permeability variations 

are not restricted to defect locations. Internal stresses, non-uniform heat 

*Actually the situation is far more complicated for magnetic media, which 

have nonlinear induction vs field intensity relationships. The value of 

permeability that we use in equation (1) must also be carefully chosen in the 

magnetic case, but detailed consideration of electromagnetic fields in 

conductors is beyond the scope of this discussion.

- 238 - 

treatment, and bending are examples of non-defective sites where signals will 

occur in unsaturated samples. These are often indistinguishable from real 

defects, resulting in unnecessary tube removal or plugging. It is therefore 

essential, for reliable inspection, that the medium be fully saturated. 

To understand the phenomenon of magnetic saturation we must consider the 

magnetic dipole moments of the constituent atoms in a magnetic material. 

These are arranged in a random pattern of domains when the sample is in the 

demagnetized state. Application of an external field causes large scale 

alignment of dipoles, which contributes to the total magnetic induction. 

Eventually, as applied field intensity is increased, the material 

contribution reaches a saturating value; all dipoles are aligned parallel to 

each other. Any further increase in induction is then equal only to the 

increase in applied field. The slope of the B vs H curve becomes equal to a 

constant (no), where B and H are the moduli of induction and field 

intensity respectively, and conventional eddy current testing is possible. 

2.2 Saturation Probe 

Because of the geometrical configuration of tubing in a heat exchanger, the 

only source of applied saturating field must be the eddy current probe 

itself. For mild magnetic materials it has been found that permanent magnets 

housed in the probe are the most convenient method of field application. 

This eliminates the need to supply any high magnetizing currents, and the 

subsequent cooling requirements that these imply. In general, for ET 

saturation probe development, it is more desirable to choose permanent 

magnets, whenever possible, over electromagnetic coils. 

We have been fortunate in developing probes to saturate mild materials, due 

largely to the advances over the past two decades in permanent magnet 

manufacturing technology. Of particular use has been the fabrication of Rare 

Earth - Cobalt alloys (e.g., S111C05) with maximum energy-products of order 

200 kJ/m 3 (25 MG-Oe). Only with hard magnetic materials of this quality 

can alloys such as 3RE60, with saturation induction approximately 0.6T 

(6000G), be saturated. This is achieved by placing three magnets in the 

configuration shown in Figure 1, where like poles are arranged facing each 

other across a space filled with soft magnetic material (mild steel or 

Permendur "keepers"). This configuration was initially developed by Cecco 

and co-workers[3]; its effect is to focus magnetic flux lines into the tube 

wall. Higher flux levels, and hence inductions, can be obtained in the tube 

using this method than from a single magnet. The eddy current sensing coil 

normally rides with the assembly, surrounding the central magnet. As 

expected, the magnets experience considerable demagnetizing fields from each 

other, and must therefore have a very "long and flat" magnetization vs 

coercive field curve. This is a property of Rare Earth-Cobalt alloys, making 

them Ideal for this application.

- 239 - 

The choice of optimum magnet sizes depends largely upon the material type and 

tube dimensions. Assembling a probe for a given material and tube 

configuration can be aided by computer analysis, such as finite element 

solutions to Maxwell's equations for magnetostatics, but the magnetic 

properties for the materials involved (tube, magnets, keepers) must be known 

accurately. Laboratory verification is still required and constitutes the 

major portion of the work involved. Once a magnet configuration is 

"optimized", usually by impedance monitoring of magnetization levels, 

conventional eddy current probe design considerations, such as impedance 

matching and coil winding, can be determined. This is described in a paper 

by Cecco [3]. A cut-away view of a probe is shown in Figure 2. 

The probe may be described as three magnets and two keepers, arranged in the 

manner of Figure 1, with the sensing coil surrounding the central magnet and 

adjacent to the region of tube wall saturation. Some shielding is necessary 

to prevent coupling into the keepers by the alternating field. A reference 

coil is located at the rear and is shielded to prevent it from sensing the 

tube material. Spacer material is usually some nonconducting medium such as 

Dupont Delrin. All components are housed in an Inconel* (nonmagnetic) casing 

with electrical connections through a four pin connector at the rear. 

A final comment, in relation to probe dimensions, should be inserted here. 

Dimensions shown in Figure 1 are typical of the magnet assembly in a 19 mm OD 

tube, and they effectively determine the minimum probe length and diameter. 

In many instances it may not be possible for the probe to traverse U-bend 

tubes completely, restricting usage to straight tubing, or the straight 

sections of U-bend tubing. 

3. PROBE TESTING 

The first step in developing a saturation probe is to optimize the magnet 

configuration to the tube size. This cannot be done without some trial and 

error measurement, and a fast method of verifying saturation in a tube has 

been developed to accommodate it. If an AC encircling coil is placed around 

the outside of a nonmagnetic tube, it will have an electrical impedance that 

is dependent upon the tube material resistivity and geometry. Replacing the 

tube with a magnetic one of identical dimensions and resistivity will change 

the coil Impedance, but it can be brought back to its original value by 

saturating the magnetic tube. Eddy current impedance monitoring of 

saturation is based upon this phenomenon. 

international Nickel Co.

- 240 - 

With an AC coil surrounding a magnetic tube, magnets are slid down the inside 

and saturation is verified if the coll impedance equals the value it would 

have for an identical, but nonmagnetic, tube. In practice neither electrical 

resistivity nor dimensions can be matched identically between magnetic and 

nonmagnetic tubes. The former problem, that of electrical resistivity, can 

be tolerated if the match is within 15 to 20 percent. The latter problem, 

different tube dimensions, is overcome by closely matching only the outer 

diameter and sampling a shallow surface layer using high frequency. Figure 3 

shows the change in impedance of a coil encircling a 19.1 mm 00 (1.84 mm wall 

thickness) 3RE60 tube as a saturating magnet assembly is moved through the 

tube. 

After saturation is attained and the complete prototype probe is assembled, a 

comparison of ET signals from artificial defects in a nonmagnetic tube and a 

3RE60 tube can be made. This is shown in Figure 4, where impedance plane 

signals from machined calibration defects in 316 stainless steel and 3RE6O 

are presented. These signals were recorded at a test frequency of f9o, the 

frequency at which shallow ID and 0D defect indications have 90° phase 

separation[4]. 

As a final check on the ability of the probe to prevent unwanted 

"permeability signals", a strip chart recording is taken in a 3RE6O tube with 

machined ID, 0D and through hole indications, and a shot-peened 

circumferential band. Shot-peening induces a localized permeability 

variation due to internal stresses, but is not a real defect site. On the 

strip chart recording, which is shown in Figure 5, we would expect the real 

defects to appear, but not the shot-peened area. 

The key procedures and components for probe development are given in the 

following summary. 

SUMMARY OF PROCEDURES AND COMPONENTS 

Test 1 Impedance Monitoring to Verify Saturation 

Detector: AC Encircling Coil 

Frequency 1 MHz 

Nonmagnetic Reference Tube: 304 stainless steel 

resistivity p = 72 x 10~ B fi-m 

dimensions 19.13 mm 0D x 1.38 mm wall 

Method: Monitor encircling coil impedance as magnets pass through 

tube. 

Results: Figure 3 - Impedance plane display of encircling coil.

- 241 - 

Test 2 Comparison of ET Signals in Nonmagnetic and Magnetic Tubing 

Detector: ET Saturation Probe 

Frequencies: £,„ (316 SS) = 70 kHz 

£,0 (3RE60) = 75 kHz 

Nonmagnetic Tube: 316 stainless steel 

p = 75 x 10~ 8 J2-m 

dimensions 19.05 mm OD x 1.85 mm Wall 

Method: Record ET signals from OD, ID, through hole machined 

calibration defects in nonmagnetic tube and 3RE60 tube 

Results: Figure 4 - Impedance plane display of ET probe signals 

Test 3 Verify Saturation Through Absence of Permeability Induced Signal 

Detector: ET Saturation Probe 

Frequency f90 (3RE60) = 75 kHz 

Method: Record ET signals from 3RE60 tube having OD, ID, through 

hole machined calibration defects and shot-peened OD band. 

Results: Figure 5 - Strip chart recording to show absence of signal 

from shot-peened area. 

For all tests: 3RE60 tube - resistivity p = 85 x 10~ e Q-m 

dimensions 19.1 mm OD x 1.84 mm wall 

4. DISCUSSION OF RESULTS 

Eddy Current Instrument - Automation Industries EM33OO 

Because there is a great deal of procedural background involved in 

understanding Figure 3, some discussion is warranted. The results shown may 

be briefly explained as follows. 

An AC encircling coil operating at 1 MHz In air has an impedance vector locus 

at point 0. Phase is set up such that an inductive increase gives a 

vertically upward signal. Insertion of a 304 stainless steel tube into the 

coil moves the Impedance locus to point 0', whereas if a 3RE60 tube is used, 

impedance is shifted to point A. From point A three traces can be generated. 

These are labelled as trace A, trace B and trace C. Trace A is generated by 

magnetizing the tube to saturation by means of an external solenoidal coil, 

bringing the impedance locus to point A 1 . If the 304 stainless steel and 

3RE60 tubes had identical resistivity and geometry points 0' and A' would 

coincide. If the vertical component of trace A Is plotted against solenoid 

magnetizing current, trace B is generated, where B 1 represents the point of 

saturation. If now, instead of a solenoid, magnets are used inside the tube 

and the vertical component of impedance change (from point A) is plotted

- 242 - 

against magnet position, trace C is generated. By definition the vertical 

distances from A to A' and B to B 1 , which are equal, represent the vertical 

impedance change necessary to saturate the tube. For the magnets to also 

saturate the tube, the vertical distance from C to C must equal the former 

two. 

It is evident from Figure 3 that the magnet configuration tested has 

saturated the 3RE60 tube at point C. This occurs when the magnets are 

centred under the AC sampling coil. We see three lobes in trace C, each one 

due to one of the three magnets in the configuration. With saturation 

verified, a probe can be constructed by winding an AC coil around the central 

magnet. 

The eddy current impedance plane signals in Figure 4 compare signals from 316 

stainless steel and 3RE60. The saturation probe is used for both, since in 

the case of 316 stainless steel the magnets do not influence the eddy current 

signals. Signals from 3RE6O and 316 stainless are essentially identical. 

For this type of signal analysis, eddy current instrument phase is set to 

give a horizontal deflection to the left for a concentric ID groove. Working 

at fgo, OD groove signals will be rotated 90° clockwise from the ID signals. 

This 90° phase separation is in itself often used as a verification of 

saturation, and it is amply demonstrated in the 3RE6O signals. The OD signal 

from the 3RE60 tube is slightly distorted near the bottom. This is 

considered to be due to demagnetizing effects near the corners of the groove, 

preventing full saturation in these areas. Some samples of 3RE60 tubing are 

more difficult to saturate than others. Those that are more difficult 

generally show this distortion; it does not occur in all tube samples. The 

central portion of the groove is fully saturated, as can be seen by the 

disappearance of the distortion as the signal grows vertically. Overall the 

eddy current signals from the 3RE6O tube closely natch those from the 316 

stainless, indicating the probe is capable of providing data that is readily 

interprétable in the light of present eddy current knowledge. 

The strip chart recording of Figure 5 shows signals of the type in Figure 4 

resolved into X and Y components. Each one is displayed on a separate 

channel. Although they are labelled for easy identification, there is no 

difficulty determining that signals are present from the machined calibration 

defects. We note however that between the two event markers there Is no 

discernable signal, although this is the region of an external shot-peened 

band. Without full saturation, a large signal would be seen in this region. 

The absence of such a signal further proves the probe can eliminate 

"permeability noise", and saturates the tube under examination.


- 243 - 

It would be premature to conclude that full field inspections could now be 

performed in all mildly ferromagnetic materials up to those of saturation 

induction approximately 0.6 tesla. The laboratory testing of saturation 

probes for 3RE60 is to a great degree complete, and the probes must now prove 

themselves in the field. Laboratory results on calibration defects are 

encouraging. We know of at least one prototype probe that has seen field 

use, but at this time its performance has not yet been assessed in depth. 

It is believed that 3RE60 tubing of "average" magnetic characteristics (for 

that material) can be inspected by eddy current, using probes of the type 

discussed here. Although the results presented in this paper dealt 

exclusively with 19 mm tubing, previous work using 16 mm tubing has also been 

successfully carried out. With a minimal amount of laboratory work the 

application of these probes could likely be extended to other 

austenitic-ferritic stainless steels. 

To avoid presenting an overly optimistic picture of the situation, it is 

necessary to point out several important considerations. These are as 

follows. 

i) Because of the sizes of the magnets, U-bend tube sections are largely 

uninspectable. 

ii) Tube regions under magnetic support plates are uninspectable since 

these areas cannot be fully saturated. 

iii) The variation of magnetic characteristics between tube samples is not 

yet well known. Only a field test can determine whether all tubes in 

a given heat exchanger are saturable. (This is a problem also 

encountered in Monel 400 Inspection). 

iv) It is possible that the existence of heavy magnetic deposits might 

sufficiently perturb the saturating field to render tubing 

uninspectable. 

v) The probe cannot be used effectively in the hands of an operator 

unaccustomed to ferromagnetic inspection and its associated problems. 

The above points serve to underline that although the probe makes 3RE60 

tubing Inspectable, it cannot be expected to perform satisfactorily outside 

of its limitations.

6. SUMMARY 

- 244 - 

Conventional eddy current inspection of ferromagnetic materials requires full 

magnetic saturation of the test item. This has been realized In 19 mm 3RE60 

tubing by using a focussed arrangement of permanent magnets. The results 

shown here are proof that the laboratory problems associated with 3RE60 

inspection have been overcome, and field confirmation of probe ability is the 

logical next step. There is no reason to believe that the use of these 

probes cannot be extended to similar mildly ferromagnetic alloys. Properly 

utilized, this generation of saturation probes is of great potential value to 

ferromagnetic tube inspection, so long as they are used within the bounds of 

their applicability. 

7. REFERENCES 

[1] Van Drunen, G. and Cecco, V.S., Recognizing Limitations in Eddy 

Current Testing, Atomic Energy of Canada Report AECL-7508, (1981). 

[2] Lorrain, P. and Corson, D.R., Electromagnetic Fields and Waves - 2 ed 

(Freeman and Co., 1970), 476. 

[3] Cecco, V.S., Materials Evaluation, _3_7, (1979), 51. 

[4] Cecco, V.S., Van Drunen, G., Sharp, F.L., Eddy Current Testing Manual 

on Eddy Current Method Vol. 1, Atomic Energy of Canada Report 

AECL-7523, (1981), 124.

End Magnet 

- 245 - 

Soft Magnetic 

Keepers 

30 mm 

FIGURE 1: Magnetic Focusing Configuration in 

Ferromagnetic Tube. 

End Magnet

10 

1 Centre Magnet 

2 Keepers 

3 End Magnets 

4 Brass Shield Rings 

5 Delrin Spacers 

- 246 - 

6 Delrin Reference Coil Mount 

7 Brass Reference Coil Shield 

8 Eddy Current Coil 

9 Reference Coil 

10 Inconel Casing 

11 Casing Closure Plug (Inconel) 

FIGURE 2: Cruss Sectional View of Probe. 

11

Induct ive 

Increase 

Inserti on 

of 304 

Stainless Steel 

Tube 

0' 

Trace A 

B r 

Solenoid Currrnt 

(amperes) 

External Magnetization 

50 0 

Trace C 

Magnet 1'usition (cm) 

FIGURE 3: Impedance Monitoring of Saturation. 

50 100 150 200 

Permanent Magnet Saturation 

•p-

A - II' Concentric Groove 

0.1 mm deep x 2.5 mm 

long 

- 248 - 

316 Stainless Steel Signals: fy0 = 70 kHz 

B - OD Concentric Groove 

O.-i ram deep x 2.5 mm long 

3RE60 Signals: f 0 = 75 kHz 

A - ID Concentric Groove 

0.5 mm deep x 2.5 mm long 

C - OD Hole 1.5 ma Dia. 

0.5 mm deep 

D - Through Hole 

1.5 mm Dia. 

B - Through Hole 

1.5 mm Dia. 

E - OD Eccentric Groove 

0.8 mm deep x 

3.5 mm long 

C - 01) Fxi'i-ntric Gro 

0.7 S mm deep x 

3.5 mm long 

I'lGURE M : Kddv Current Impedance Plane Signals.

Vertical Signal Component (Y-Channel) 

A_A 

ID Through OD 

Groove Hole Groove 

Horizontal Signal Component (X-Channel) 

- 249 - 

Both Channels - 10 volts full scale deflection 

ET Frequency - 78 kHz 

ID 

Groove 

FIGURE 5: Strip Chart Recording. 

Shot-Peen 

(Permeability) 

event 

marker 

event 

marker

- 250 

ON THE RELATION BETWEEN ULTRASONIC ATTENUATION AND FRACTURE 

TOUGHNESS IN TYPE 403 STAINLESS STEEL 

F. Nadeau and J.-F. Bui.iie.-xe. 

National Reieaicli Council o & Canada, Bouckexv itie, Quebec 

G. Van Dianen 

Atomic Eneigij oj Canada Limited 

Chalk Rive-%] ' Ontaiio 

ABSTRACT 

Previous studies have indicated the possibility of using ultrasonic 

attenuation as a nondestructive tool for predicting the fracture toughness of 

metals. In all these studies, however, the fracture toughness was varied by 

changing microstructure or composition. In the present paper, ultrasonic 

attenuation is measured in a single sample of constant microstructure (type 

403 stainless steel) in which fracture toughness is changed by nearly a factor 

of three by varying the temperature. Results show that the frequency and 

temperature dependence of attenuation are in agreement with conventional grain 

scattering theory and do not correlate with the large changes in Kic. It is 

concluded that one of the main difficulties of nondestructively characterizing 

fracture toughness is that a small variation in one of the material's 

properties, such as yield stress, can alter the fracture mechanism, leading to 

very rapid changes in Kjc. In most cases, the modest change in yield stress 

will be accompanied by similarly modest changes in attenuation, which do not 

reflect the rapid changes in Kjc. 


As the ability to detect and characterize discrete defects in structures is 

improved through the use of better nondestructive methods, more attention is 

being focused on novel techniques for the determination of microstructure 

'1' and mechanical properties (2). of particular interest is the 

possibility of using ultrasonic measurements to determine engineering 

mechanical properties such as tensile strength, yield strength, fracture 

toughness and the ductile to brittle transition temperature (DBTT). Earlier 

attempts in this area have been based on empirical relations between 

ultrasonic velocity and yield (or tensile) strength in various steels and 

iron-carbon alloys (3-5)t More recently, the bahavior of third-order 

elastic constants was shown to correlate with the heat treatment history of 

aluminum alloys ("'. 

However, most attempts at characterizing microstructure and mechanical 

properties ultrasonically have been based on measurements of attenuation. 

Attenuation results from the combined effects of scattering, the reorientation 

and mode conversion of acoustic energy at microstructural interfaces, and

- 251 - 

absorption, the conversion of acoustic energy into heat. Scattering and 

absorption are both caused by the presence of inhoraogeneities such as 

precipitates, inclusions, voids, grain and interphase boundaries, 

dislocations, interstitial impurities, etc. Since these same inhomogeneities 

also play a primary role in determining mechanical properties, it is expected 

that measurements of attenuation will reflect mechanical properties. The 

sensitivity of attenuation to small changes in microstructure which affect 

mechanical properties has been demonstrated for aluminum alloys »'' and 

other materials (8)# in cases where grain boundary scattering is dominant, 

various models (l>°,10) offering good agreement with experiment can be used 

to calculate the grain size from the frequency dependence of attenuation. The 

ultrasonically measured grain size, 6, can then be correlated with mechanical 

properties via the use of the Hall-Petch relationship. 

U = uo + kS-i (1) 

where M is the mechanical property of interest, n, is a constant dependent 

on composition, and k is a proportionality constant. This latter approach was 

used by Klinman et al (11) to successfully predict yield strength, tensile 

strength and DBTT's in carbon steels of known compositions. Another approach 

was taken by A. Vary ("), who suggested an empirical relation relating 

fracture toughness, K.ic, yield strength oy, and ultrasonic attenuation: 

Klc 2 /Oy = 8.12 x 10 6 (Va&j) 0 ' 31 " 1 (2) 

where V£ is the velocity of longitudinal waves in the material and ßg is 

the slope of the attenuation versus frequency curve at a frequency, f = 

Vg/6, where the wavelength is equal to the grain size. Since f is 

generally outside the range of most normal attenuation measurements (eg. f = 

0.5 GHz in steel with 6 = 10 Mm), the value of fJ$ is extrapolated from 

measurements in the 10 to 50 MHz range. Two ma rag ing steels and a titanium 

alloy, aged at various temperatures, were found to obey relation (2). 

However, studies on Fe-C alloys (13) have shown that increasing the grain 

size or increasing the cabon content, both cause an increase on the DBTT but 

have opposite effects on ultrasonic attenuation. This is attributed either to 

diminished absorption due to an increase in dislocation damping with 

increasing C content, or to diminished scattering due to a decrease of elastic 

anisotropy of the grains related to the presence of pearlite or cementite 

around the ferritic grain boundaries. Thus it is difficult to isolate one 

component of the ultrasonic attenuation, and in particular, to differentiate 

between absorption and scattering. It is also difficult to vary one material 

property such as fracture toughness while others remain constant. 

In the present study, this is accomplished by correlating the ultrasonic 

behavior and mechanical properties of a 403 stainless steel between -60 and 

+40°C. In this temperature range, the plain strain fracture toughness, Kjc, 

changes by =250% because of a smooth brittle to ductile transition, while 

composition and microstructure remain unchanged.

- 252 - 

Following a brief description of the sample, its mechanical properties, and 

the experimental techniques, attenuation measurements are presented as a 

function of temperature and frequency. They are interpreted in terms of 

scattering mechanisms and discussed in terms of possible correlations with 

mechanical properties. 

EXPERIMENTAL 

Sample 

The material chosen is an AISI type 403 stainless steel which has well 

characterized properties because of its use as end fittings to hold zirconium 

pressure tubes in Canada-Deuterium Uranium (CANDU) nuclear reactors. The 

sample was part of a compact tension (CT) specimen used in a previous study on 

the effects of impurity concentration and temperature on the fracture 

toughness of type 403 stainless steel (14). The particular sample used 

here, identified as LR in reference 14, has chromium as main alloying element 

(11.8 wt %) and a total impurity concentration of 1620 ppm. (See reference 14 

for exact composition and heat treatment). The microstructure consists of 

tempered martenslte. Carbides outline the original austenite grains (6= 25um) 

and also the martensite lath boundaries. Fig. 1 is a scanning electron 

micrograph of the sample. 

The Kic values were measured using compact tension (CT) specimens according 

to ASTM Test for Plane-Strain Fracture Toughness of Metallic Materials 

E 399-74. The yield stress, ov, and the plane strain fracture toughness, 

Kic, taken from reference 14 are given in Fig. 2 for the temperature range 

-60°C to +40°C. Note that the yield stress decreases by = 7% and Klc 

increases by approximately a factor three over this temperature range. 

Ultrasonic attenuation 

Ultrasonic attenuation was measured using the pulse-echo technique with two 

undamped 12 mm diameter X cut quartz piezoelectric transducers having center 

frequencies f0 of 5 to 10 MHz used successively. Using overtones, 

measurements were made at frequencies of 10 to 50 MHz in 5 MHz intervals. The 

ultrasonic pulses were generated and detected with a Matec heterodyne 

pulser-receiver and the attenuation was measured by electronically fitting an 

exponential curve over the decaying echos. The velocity of longitudinal waves 

was also measured (6075 m/s at 20°C varying by less than 1% in the temperature 

range studied) and used to convert the decay time into neper/m attenuation 

units. Temperature was varied by mounting the 5 mm thick sample on a Peltier 

element which allowed both cooling and heating of the sample depending on the 

polarity of the current supplied to the element. The temperature was measured 

with a thermocouple mounted directly on the sample. The experimental 

apparatus, along with a typical oscillogram showing a series of echos fitted 

with an exponential decay curve, is illustrated in Fig. 3.

Results and Discussion 

- 253 - 

Ultrasonic attenuation results are given in Fig. 4 as a function of temperature 

for frequencies ranging from 10 to 50 MHz. The attenuation exhibits a 

strong dependence on frequency but the variations with temperature are small 

(within experimental error) and somewhat random. The frequency dependence of 

the attenuation is illustrated in Fig. 5 using values averaged over the 

temperature range studied and corrected for diffraction losses by the method 

described in (15). The boundary that distinguishes Rayleigh scattering from 

stochastic scattering, X = 2it6, corresponds for the case of a grain size 6 = 

25 ym to a value of f = 38.7 MHz. We see, on Fig. 5, the initial f 1 * dependence 

of a typical of Rayleigh scattering, gradually transforming into an f 

dependence in the X = 2IT6 region where stochastic scattering is expected. 

Thus, the frequency dependence of a is consistent with grain boundary 

scattering theory. 

In order to examine the temperature dependence of a in the perspective of the 

temperature related changes in plastic properties and fracture toughness, the 

values of 0£ were computed, as described in ref. 2, and compared with the 

empirical relation given in the introduction, equation (2). Fig. 6 is a plot 

of V£0£ values against Kic /cv. The data points lie more or less 

along a horizontal line indicating no net effect on ultrasonic attenuation of 

the almost 3 fold change in fracture toughness. 

It is well-known O0) that ultrasonic attenuation in steel at typical 

N.D.E. frequencies (1-50 MHz) is mainly due to scattering by grain boundaries 

and the present results appear to corroborate this fact. Although scattering 

in a multiphase material such as type 403 stainless steel is not easily 

calculated theoretically because of the complexity of the structure and 

insufficient knowledge of elastic constants, the results of Fig. 5 are 

consistent with measurements by Papadakis and others ( 10 ), which indicate 

that the main scattering features are the original austenitic grains, and that 

the effect of the intragranular structure is to modify the effective anisotropy 

of the grains. In general, the presence of multiple phases is to reduce 

the anisotropy compared to single phase materials. This explains the 

relatively low attenuation of the present sample. Based on the anisotropy, A, 

of pure iron (**•) where A = Cilit-(Cn-C12)/2 and the Cy 's are the usual 

elastic constants for a cubic system, any increase of =6% in attenuation with 

temperature is expected between -60°C and +40°C. This is within experimental 

error of the results of Fig. 4. 

The temperature dependence of yield stress (and therefore indirectly that of 

Kic) may be expected to correlate better with absorption, which can be 

related to dislocation density and mobility ( 17 \ rather than scattering 

which is associated with a constant grain size and varying anisotropy. An 

effort was therefore made to separate any absorption contribution to a by 

analyzing the data of Fig. 5 in the form of plots of ct/f versus f 2 as in 

reference (18). Such an analysis, however, did not give a clear separation of 

scattering and absorption, largely because of the lack of data lying clearly 

^ ideated that the absorption contribution is


- 254 - 

In order to verify tô what degree ultrasonic attenuation can be used to nondestructively 

predict the fracture toughness of a material, the ultrasonic 

attenuation behavior of an AISI type 403 stainless steel exhibiting a smooth 

ductile to brittle transition was studied as a function of temperature. Over 

the temperature range studied (-60°C to +40°C) the yield strength and fracture 

toughness changed respectively by 7 and 250%. The ultrasonic attenuation, 

however, varied by less than the 6% expected from grain boundary scattering 

theory, giving no indication of the large change in Kic. An analysis of the 

results in terms of Vary's empirical model (equation 2) which appears to give 

good results for cases where Kic is varied by changes in microstructure 

(grain size, heat treatment) gave no correlation between attenuation and 

K ic / a y 

One of the main difficulties of nondestructively characterizing a material's 

fracture toughness is that a small variation in one of the material's 

properties, such as yield stress, can cause the fracture mechanism to change 

(ductile to brittle transition) leading to very large variations in Klc 

values. If that property is indirectly measured by a method such as 

ultrasonic attenuation which is mostly sensitive to other properties of the 

material, the chances of obtaining an accurate K\c value are reduced even 

more. 

Nondestructive evaluation of Kic can be achieved when the method directly 

measures the material properties controlling its fracture toughness, for 

example ultrasonic attenuation for samples where variations in the grain size 

changes their toughness (**). Similarily, a method allowing for the direct 

measurement of ultrasonic absorption may provide better correlation with K^c 

variations caused by differences in the plastic properties of a material for a 

given microstructure. Such a method is currently under development (19) at 

the Industrial Materials Research Institute and it is hoped that it will prove 

to be a powerful tool for the nondestructive characterization of materials. 

Acknowledgement 

Supply of sample material for this study by R.R. Hosbons of the Metallurgical 

Engineering Branch at Chalk River Nuclear Laboratories is gratefully 

recognized.

REFERENCES 

- 255 - 

1. K. Goebbels, "Structure Analysis by Scattered Ultrasonic Radiation", in 

Research Techniques in Nondestructive Testing, Vol. IV, ch. 4, edited 

by R.S. Sharpe, Academic Press New York (1980). 

2. A. Vary, "Ultrasonic Measurement of Material Properties", ibid, ch. 5. 

3. J. and H. Krautkrämer, "Ultrasonic Testing of Materials", second 

edition, Springer-Verlag, New York, N.Y. (1977). 

4. L.Y. Levitan, A.N. Fedorchenko and A.V. Sharko, "Ultrasonic Testing of 

the Strength of Steel 45", Defektoskopiya, Vol 3, pp. 129-130, May-June 

1976. 

5. E.R. Leheup and J.R. Moon, "Yield and Fracture Phenomena in 

Powder-Forged Fe -0.2C and their Prediction by NDT Methods", Powder 

Metallurgy, April 1980, p. 177. 

6. J.S. Heyman and E.J. Chern, "Characterization of Heat Treatment from 

Ultrasonic Determinarion of the Second and Third Order Elastic 

Constant", 1981 IEEE Ultrasonics Symposium p. 936, IEEE New York 

(1981). 

7. M. Rosen, "NDE Characterization of Precipitation Hardening Phenomena in 

Aluminum Alloys" in Advanced NDE Technology 1982, National Research 

Council of Canada, Boucherville, Québec, Research Report IGM82-0-59 

J.F. Bussière ed., p. 54. 

8. R.E. Green, "Effect of Metallic Microstructure on Ultrasonic 

Attenuation", p. 117 in Nondestructive Evaluation: Microstructural 

Characterization and Reliability Strategies, edited by 0. Buck and 

S.M. Wolf, Conference Proceeding, The Metallurgical Society of AIME, 

New York, N.Y. (1981). 

9. Mason, W.P., Physical Acoustics and the Properties of Solids, New York, 

Van Nostrand, New York, (1958). 

10. E. Papadakis, "Scattering in Polycrystalline Media", ch. 5, pp. 2237 in 

Methods of Experimental Physics, Vol. 19 - Ultrasonics, P.D. Edmonds 

ed., Academic Press, New York, N.Y. (1981). 

11. R. Klinman and E.T. Stephenson, "Ultrasonic Prediction of Grain Size 

and Mechanical Properties in Plain Carbon Steel", Mat. Eval. Vol. 39 

(1981), No. 12 pp. 1116 - 1120 Nov. 1981. 

12. A. Vary, "Correlations among Ultrasonic Propagation Factor and Fracture 

Toughness Properties of Metallic Materials", Mat. Eval. Vol. 36 N.7 

p. 55-64, June 1978.

- 256 - 

13. R.L. Smith, F.L. Rusbridge, W.N. Reynolds and B. Hudson, "Ultrasonic 

Attenuation, Microstructure and Ductile to Brittle Transition 

Temperature in Fe-C Alloys", Mat. Eval. Vol. 41, pp. 219-222, (Feb.) 

1983. 

14. R..R. Hosbons, A.J. Pacey, B.L. Wotton, "Effect of Impurity 

Concentration on the Change in Fracture Toughness of AISI Type 403 

Stainless Steel with Fast Neutron Radiation", Properties of Reactor 

Structural Alloys After Neutron or Particle Irradiation, ASTM STP 570, 

pp. 103-116, (1975). 

15. H. Sepi, A. Granato, R. Truell, "Diffraction Effects in the Ultrasonic 

Field of a Piston Source and Their Effects upon Accurate Ultrasonic 

Attenuation Measurements", Journal of the Acoustical Society of 

America, Vol 28, p. 230, (1958). 

16. G. Simmons and H. Wang, "Single Crystal Elastic Constants and 

Calculated Aggregate Properties: A Handbook", The M.I.T. Press, 

Cambridge, Mass., 1971, p.34. 

17. A.B. Bathia, "Ultrasonic Absorption", Clarendon Press, Oxford, England 

(1967), pp. 338-370. 

18. R.L. Smith, W.N. Reynolds and S. Perring, presented at the Fourth 

European Conference on Internal Friction and Ultrasonic Attenuation, 

Lyon, France 1983 (to be published). 

19. J.P. Monchalin, J.F. Bussiêre, "Measurement of Near-Surface Ultrasonic 

Absorption by Thermo-Emissivity" presented at the Symposium on 

Nondestructive Methods for Material Property Determination, Hershey, 

Pennsylvania, April (1983), (Plenum Press, New York, to be published).

- 257 - 

Fig« I' Scanning electron micrograph of the type 403 stainless steel 

sample used in the present study. 

i 

820 

810 - 

800 

790 

780 

770 

760 

1 

I | 

750 

I 

i i 

-60 -40 -20 0 20 40 

T(°C) 

i 

- 

100 

80 

o 

i 

60 I 

- 40 

Fig. 2: Yield strength, ay, and plane strain fracture toughness, 

Kic, of the 403 stainless steel sample between -60°C and 

+40°C.

Ultrasonic 

Instrumentation 

- 258 - 

«ÏOTÏYB 

- Styrofoam 

Transducer 

Sample 

Pelletier Element 

— Heat Sink 

Fig. 3: Schematic of the experimental setup used to measure ultrasonic 

attenuation as a function of temperature.

-Ê 

70 

60 

50 

40 

6 30 

20 

— 

- 

"I 

I 

• 

a 

1 

* 

a 

O 

A 

X 

50 MH^. 

30 MH^.. 

15 MH^ 

10 MH^. 

- 259 - 

10 I 

I 

1 

r 

-60 

0 

-40 -20 

T(°C) 

0 

* 

I 

• 

O 

A 

* 

D 

0 

* 

• 

O 

* 

• 

A 

* 

a 

A 

I 

* * 

* 

D 

1 

a 

1 

* * 

D 

D 

- 

I. 

- 

20 40 

Fig. A: Ultrasonic attenuation data as a function of temperature for 

frequencies ranging from 10 to 50 MHz.

100 

a (m" 1 ) 

10 

f Slope = 4 

10 

- 260 - 

f (MHZ) 

Slope = 2 

100 

Fig. 5: Frequency dependence of ultrasonic attenuation.

- 261 - 

V ßß =8.07x10" 2l (k 2 icÖy) 

K 2 1c/ay(MJ/m 2 ) 

Flg. 6: Relation between ultrasonic parameters V^ßg and mechanical 

properties Kic/Oy. The solid line is equation (2), which 

Vary used to fit his data on Titanium alloy and maraging 

steels. In this case Kic and oy were changed by heat 

treatment. The crosses correspond to the experimental data 

obtained for 403 stainless steel in the temperature range -60°C 

to +40°C.

- 262 - 

ACOUSTIC EMISSION TESTING OF MAN-LIFT DEVICES 

J.A. Baton 

Ontario Hydto Reie.an.ah 

Toronto, Ontario 


Safety is a very controversial issue. Whereas we may find it an acceptable risk 

to cross a busy street or to ski down a steep slope, we may have some reservation 

about flying in a commerical airliner or to be raised 20 m above the ground 

in a bucket of an aerial lift device unless we have some positive perception 

regarding the structural Integrity of the equipment. Much of this assurance is 

developed through nondestructive test procedures but, until recently there has 

been no method which could provide reliable information as to the mechanical 

worthiness of the fibre reinforced plastic (frp) boom of man-lift devices 

("bucket trucks" or "cherry pickers"). 

Dielectric testing of booms is mandatory but this test, usually performed semiannually, 

is a poor indicator of mechanical integrity. Under dry conditions a 

boom on the point of mechanical failure could pass this electrical test. 

Conversely, seepage of moisture Into surface layers would result in dielectric 

failure but may not have implications in terms of mechanical fitness. 

Tradition NDT methods do not work well on composite materials such as frp. 

Radiography will show fibre direction and gross voids but little else. Ultrasound 

suffers high attenuation plus the fibres give rise to a myriad of reflections. 

Liquid pénétrants will show surface imperfections but may also cause 

chemical degradation of the plastic resin. However, acoustic emission monitoring 

can potentially detect and locate "active" defects in both frp and metallic 

components and can be used in a straightforward, all encompassing test. 

FAILURE MODES 

As a generality, nondestructive testing is performed to detect potential future 

failure of a component or structure. If such a test is to be worthwhile, it 

must be sensitive to the precusors of possible failure modes. Further, the test 

must be performed at a frequency such that the progression from the earliest 

reliable detection of incipient failure to actual failure cannot occur. This 

point will not be dealt with here but does raise some questions as to the 

efficacy of periodic testing of frp booms.

- 263 - 

The two components of the frp composite are very different in mechanical 

behaviour. The glass fibres are brittle, exhibiting little strain prior to 

tensile fracture, whilst the matrix is a plastic which exhibits visco-elastic 

flow [1] rather than sudden failure. To some degree, the composite tends 

towards the characteristics of the matrix when in compression and the fibres 

when under tension. Under normal load regimes, a boom will experience tenoion 

on its upper surface and compression on its lower surface. The tensile properties 

are superior to those under compression thus, if the load is such as to 

cause failure, the normal mode is compressional collapse. 

Corapressional failure is, from an acoustic emission standpoint, very active. It 

is initiated by failure of the matrix and results in the large scale expulsion 

of fibres, as in Figure 1. All the processes involved are good generators of 

acoustic emission and thus defects that would lead to this failure are readily 

detectable. 

In rare but no unknown circumstances [2], tensile failure can occur. An hypothesis 

describing this behaviour has been presented [3] but has not been 

confirmed. Briefly, the scenario is as follows. Below a certain threshold, the 

fibres will maintain a tensile load Indefinitely. At loads above this threshold, 

small scale fibre breakage will occur. The load previously supported by 

the broken fibres will be instantaneously transferred to the surrounding 

matrix. The matrix then undergoes visco-elastic flow and, in turn, transfers 

its load to the remaining fibres, thus adding to their overload. Again, this 

results in more fibre breakage and more load shedding through the matrix to the 

remaining fibres. It is reasonable to suggest that the time taken for the 

matrix to transfer its load due to fibre breakage, ie, through the flow process, 

decreases as the load to be transferred increases. Thus, once the process 

starts it accelerates to an avalanche type failure, assuming the load is 

maintained. This can result in a very sudden failure in contrast to the 

relatively slow compressional failure mode. 

Figure 2 illustrates the relationship between fibre breakage and time and also 

indicates the effect of changing load. Figure 3 illustrates how the time to 

failure is dependent on load and also the effect of increasing the fibre count. 

Both of these figures are generated based on the above hypothesis. However, 

McElroy [2] 1 s demonstrated the relationship in the laboratory, his results 

being illustrated in Figure 4. Figure 4, reproduced from reference 2, also 

illustrates the effect of removing or changing the load during the process. 

Initially, acoustic emission monitoring was applied to bucket trucks solely to 

enhance the inspectability of the boom. It soon became apparent that the metal 

components also generated acoustic emission under the proof test conditions. At 

first this was considered a nuisance and "noise" sources were eliminated eg, a 

poorly lubricated pivot pin, but advantage was taken of this and the test was 

soon expanded to cover all major structural items in a single, comprehensive 

test. In fact, current indications are that defects in metallic components are 

discovered far more frequently than defects In the frp components [4],

- 264 - 

Metal components with classical acoustic emission sources (eg, propagating 

crack) exhibit the Kaiser effect. This effect describes the phenomenon whereby 

acoustic emission will only be generated when stress exceeds the previous highest 

value. A new component will tend to yield copious acoustic emission when 

first loaded. If the load were removed and reapplied, no acoustic emission 

would result until the load was further increased above its previous value. 

This can be beneficial in that if a structure is loaded to the onset of acoustic 

emission, an indication of load history is gained. If a periodic proof load is 

imposed, such as 150% working load, acoustic emission will only be generated if 

active degradation has occurred since the last proof load. 

The Kaiser effect is less well exhibited by frp composites. This is because the 

resin material flows and thus relaxation occurs on load removal. This was 

demonstrated in the laboratory, Figure 5. Here an frp specimen was subjected to 

a three point bend test such that acoustic emission occurred. The load was then 

removed. When it was reapplied with little delay, almost no acoustic emission 

was generated, in accordance with the Kaiser effect. However, if a delay of 

24 h ensued prior to reloading, much of the acoustic emission was recovered. 

This phenomenon has implications where a test is performed following an accident 

to a boom, such as a known overload. 

Some reference has been made in the literature to the ultimate strength of frp 

components with respect to acoustic emission monitoring. The ultimate tensile 

strength is truly the load that just fails to initiate the avalanche failure 

mechanism. Whilst loads much higher than this can be sustained for finite 

periods of time, they do impart damage and do reduce the load level required to 

initiate avalanche failure. Also, as failure is a time dependent phenomenon, 

above the threshold load acoustic emission will continue to be generated up to 

ultimate failure. Initially this may be very sparse but becoming more prolific 

as fracture approaches. 

For loads below the avalanche threshold, the Kaiser effect (short-term) tends to 

hold. For loads above the threshold, acoustic emission will always occur. If a 

boom is loaded, unloaded and then reloaded and acoustic emissions recur, then 

the threshold load has been reached. In this case, damage will have been done 

and acoustic emission will be re-established at a slighly lower load dependent 

on the degree of degradation. The ratio of these two loads is known as the 

"Felicity Ratio" and is used as a measure of damage. 

PERFORMANCE OF THE ACOUSTIC EMISSION TEST 

With the boom in the lowered state, but with outriggers operating, sensors are 

attached. It is important that the sensors are well coupled to the component. 

This is as simple as having a clean surface, using a small amount of viscous 

fluid such as vacuum grease between sensor and the surface and holding the 

sensor firmly in position. Sensor hold-downs are extremely varied and may range 

;rora a magnetic clamp for steel components to a simple band of adhesive tape 

for the boom Figure 6. Some care must be taken to ensure that the hold-down is 

not going to move, and hence produce its own acoustic emissions, during the 

test.

- 265 - 

Sensors used on the major metallic components will have a maximum sensitivity in 

the range 100-400 kHz. However, attenuation in frp at these frequencies would 

demand an unacceptable number of sensors necessary for coverage of the complete 

boom. For this reason, it is more usual to employ sensors for the boom with 

peak response in the 20-50 kHz range. This is strictly a compromise as the 

lower frequency increases the susceptibility to extraneous mechanical noise. 

It is rare to use "time of flight" source location on either the boom or metal 

components. Sources of acoustic emission that may occur are simply identified 

with the particular component. In this way, two sensors may be fed into a 

single amplifier chain. This reduces the capital cost of equipment at the 

expense of discrimination. 

Sensor spacing may be quite important. An artificial acoustic emission source, 

such as a puiser or pencil lead break [5], must be used to ensure that the 

required degree of coverage is achieved. However, this rigorous process need 

only be performed once for a given model of aerial lift truck. Sensitivity, 

which includes sensor performance, couplant and amplifier gain, must be checked 

and equalized for each test. This is a simple process of stimulating each channel 

with the artificial source and adjusting amplifier gains for equal 

response. Figure 7 illustrates commonly used sensor locations. 

The boom is then deployed in a manner such that it is representative of working 

configurations and such that a controlled load may be applied to the bucket 

end. in Figure 8 the load is applied through a cable winch and load cell. 

A predetermined load regime is followed consisting of the gradual imposition of 

load which is then held for a few minutes and then repeated [6], The applied 

load must be consistent with the total vehicle design [7], but this may be as 

high as three times, but more typically twice, rated load [8]. 

TEST RESULTS 

For each stage of the load profile acoustic emission activity is recorded for 

each channel. The data are plotted in a number of ways, but would normally 

include count versus load, Figure 9, amplitude distributions, Figure 10 and 

perhaps an indication of individual channel activity through the test. 

Experience plays a large part in both conducting the test itself and in interpreting 

the results. At the present, there are no solid, universal accept/ 

reject criteria. However, if acoustic emission continues into the hold period 

or recurs on the second load, there is strong indication that a problem exists. 

If the general activity, that is the overall count, is unusually high, concern 

would depend on which channels were active. For example, this type of indication 

could be due to a worn or poorly lubricated pin. The shape of the count 

versus load plot is also a significant factor used in interpretation. 

Fibre breakage represents high energy acoustic emission generation. By observation 

of the amplitude distribution, Figure 10, some assessment of fibre breakage 

during the test is available.

- 266 - 

Any recognizable fibre damage may be interpreted as cause for rejection, or at 

least further investigation. 

THE VALUE OF THE TEST 

Many utitlities in both Canada and USA either perform the test for themselves, 

contract out to testing agencies or are on the point of implementating a testing 

program. Frequency of testing varies from semi-annual, consistent with dielectric 

testing, to only at rebuild time, every ~8 years. There is no doubt 

that defects are discovered, but primarily in metal components, and there is 

some debate as to whether improvements to maintenance and inspection procedures 

would be of greater benefit than an acoustic emission test. However, acoustic 

emission seems to be the only reliable monitor of the structural integrity of 

the boom. 

Industry experience is that structural problems with the boom are rara, and 

personal injury related to such problems is rarer still. However, it may be 

argued that safety is not an issue that can be weighted against the cost of 

downtime, but the question must be asked as to whether the acoustic emission 

test enhances safety. Is the frequency of testing such that the test will 

identify incipient failure before it can progress to actual failure? The answer 

is probably yes if the vehicles are used within the design limits specified by 

the manufacturer and enforced through the codes [7], but no if the vehicle is 

liable to be subjected to abuse« 

One USA utility has attempted to address this problem by installing on-board 

monitors. These are single channel, battery powered acoustic emission systems 

which are attached to the critical area of the upper boom. In an overload 

situation which results in acoustic emission the monitor produces an audible and 

visual alarm. Unfortunately, this is not a panacea as false alarms do occur 

which may require the vehicle being taken out of service for verification, only 

a very small portion of the critical areas are covered and it is possible that 

the device can generate a false sense of security in the work crew. On the 

other hand, the most common failure site is at the lower insert, Figure 11, 

which is well covered by the onboard monitor, and there is no question that 

enormous amounts of acoustic emission will be generated even prior to a sudden 

avalanche failure in tension. 

A recent exhaustive inspection of forty-three manlifts and forty-nine digger 

derricks [4] discovered no mechanical defects classed as critical or needing 

immediate attention associated with frp boom components. However, very many 

other defects were discovered but, again, whether acoustic emission testing 

would have uncovered all of these (eg, missing bolts holding pedestal to truck) 

or is even the appropriate procedure is debatable. 


There is no question that the frp material used for manlift booms generates 

acoustic emission at the early stages of structural degradation. Both compressive 

and tensile modes of failure are acoustically "noisy".

- 267 - 

The utilities using an acoustic emission proof test discover defects in metallic 

and frp components. Most of these are trivial and could be addressed in 

alternate approaches* but the bottom line is that critical defects are being 

discovered by the test, thus increasing the level of safety in the workplace. 

The proof test cannot protect against critical failure brought about through 

abuse. The on-board monitor can potentially provide some assurance during 

operation, exactly where it is needed, but may generate unwanted downtime due to 

false alarms. 

Catastrophic failure of bucket trucks is rare. This makes evaluation of an 

acoustic emission program based on statistical evidence alone very difficult if 

not impossible. 

REFERENCES 

1. J.R. Mitchell and D.G. Taggart, Acoustic Emission Testing Electric Utility 

Fleet Management, V. 2, No. 5, October/November 1983. 

2. J.W. McElroy, On-Board Acoustic Emission Monitoring of Fiberglass Boom 

Aerial Lift Trucks Society of Automotive Engineers, Congress on Exhibition, 

Detroit, February 1980. 

3. J.A. Baron, Status Report on Assuming the Structural Integrity of Aerial 

Personnel Devices Canadian Electrical Association, Study SD-124. 

4. K. Moore, Aerial Equipment Requires Thorough, Regular Inspection 

Transmission and Distribution, January 1984. 

5. A. Nielsen, Acoustic Emission Source Based on Pencil Lead Breaking, 

Svejsecentralen, Copenhagen, Publication 80.14. 

6. Acoustic Emission Test Method for Insulated Aerial Personnel Devices 

American Society for Testing and Materials proposed standard, Committee F18. 

7. CSA Standard C225, Canadian Standard for Vehicle-Mounted Aerial Devices, 

Canadian Standards Association. 

8. Procedure for Acoustic Emission Testing of Aerial Lift Devices, PSE&G 

Research Corporation, 1980.

- 268 - 

FIGURE 1 

EXPULSION OF FIBRES 

IN COMPRESSIONAL FAILURE

a 

< 

o 

TIME 

- 269 - 

LOAD 2L 

FIGURE 2 

LOAD L 

THEORETICAL RELATIONSHIP BETWEEN NUMBER OF FIBRES 

TO BREAK AND TIME (CONSTANT LOAD) 

TÎME TO FAILURE 

^ — — - _ 

FIGURE 3 

THEORETICAL RELATIONSHIP BETWEEN LOAD AND TIME 

TO FAILURE FOR GIVEN NUMBER OF FIBRES IN BOOM 

(c.f. FIGURE 4)

100 

95 

LU LU 

5 S 90 

^"-85 

K 

H H 80 

Utu 75 

70 

- 270 - 

VARIABLE 

TIME CONSTANT 

PRE-LOAD STRENGT 

DEGRADATION 

TIME TO 

SPECIMEN FAILURE 

65 

100 200 400 1000 2000 4000 10000 20000 40000 

TIME (SECONDS) 

FIGURE 4 

THE RELATIONSHIP BETWEEN LOAD, TIME. AND 

RESIDUAL STRENGTH FOR FIBERGLASS SAMPLES 

REPRODUCED FROM REFERENCE 

50% 

100000

1000,- 

- 271 - 

(C) RELOAD AFTER 21 HOURS 

(B) RELOAD AFTER 15 MINUTES 

TIME 

FIGURE 5 

ACOUSTIC EMISSION AT FIXED LOAD 

SHOWING TIME DEPENDENT RELAXATION 

10 mins

- 272 - 

FIGURE 6 

SENSORS AND PREAMPLIFIERS ATTACHED 

TO BOOM USING ADHESIVE TAPE

OUTRIGGER 

UPPER BOOM 

- 273 - 

FIGURE 7 

TYPICAL SENSOR DISTRIBUTION 

FIBREGLASS INSERT

900 

800 

700 

i/> 600 

H 

Z 

O 500 

U 

400 

300 

200 - 

100 - 

0 

- 274 - 

FIGURE 8 

BOOM CONFIGURATION DURING TEST 

LOAD APPLIED BY CABLE WINCH 

ALL FIBERGLASS CHANNELS 

r 

.J" 

ALL METAL CHANNELS 

250 500 

LOAD(kg) 

FIGURE 9 

SUM OF COUNT VERSUS LOAD BY 

BOTH FIBRECLASS AND METAL CHANNELS

ENTS 

LU 

LL 

o 

/IBER 

ID 

z 

400 

300 

200 

100 

1 

- 275 - 

_ 1 4 

MATRIX FIBRE 

t 

r •) 

I 

/ 

BEHAVIOR 

yr i 

BREAKAGE •- 

20 tO 60 80 

100 

AMPLITUDE (dB) 

FIGURE 10 

NON-ACCUMULATED AMPLITUDE DISTRIBUTION 

(LIN-LOG SCALE) 

FIGURE 11 

BOOM FAILURE AT LOWER INSERT 

IN A DESTRUCTIVE TEST

- 276 - 

ACOUSTIC SORTING OF GRINDER BALLS 

F. Na.de.au. and J.F. Zixalnfin 

national Re.itax.ch Council o i Canada, Boucke.n.v±lle., Qntan-io 

An automated non destructive testing device for ball mill grinder balls 

was developed using acoustic flaw detection. An initial study of the sound 

generated in air by the impact of grinder balls has shown that the changes 

in ball velocity upon impact produce an initial low frequency pressure 

pulse and that various resonances exited in the impacted spheres produce the 

higher frequency oscillations that follow. Resonances of uncracked balls 

were related to a particular spheroidal mode of vibration while resonances 

of cracked balls are attributed to "tuning fork" modes of vibrations. 

Both a laboratory and an industrial prototype that perform the sorting 

of cracked and uncracked grinder balls using the differences in their 

acoustic properties were built and tested. A 97% success rate was achieved 

for 3" diameter balls. The device is being fitted to the production line. 


Many of the more sophisticated Non Destructive Testing (N.D.T.) 

techniques are expensive. This often restricts their use either to the 

aerospace or the nuclear industries. When an N.D.T. problem involves a low 

mark-up non critical product, emphasis is placed on the per unit cost of the 

test. Automation becomes important while 100% reliability is no longer 

required. This paper presents such an example where a cheap, fast, 

automated N.D.T. method was developed using audio range acoustics to detect 

flaws in grinder balls. Grinder balls are hardened steel balls used in 

stone grinding "ball mills". Those used in this study are forged from high 

carbon steel rods to roughly spherical shapes and waterquenched to a 

hardness of about 60 Re. A non négligeable proportion of the production 

exhibits deep running cracks and therefore tend to rupture under operating 

conditions inside the ball mill. This problem is familiar to ball mill 

operators who often have to manually remove the broken parts from inside the 

mill to ensure smooth operating conditions and prevent exercise wear of the 

mill's inner sleeve. Testing of the production is done on a statistical 

basis (samples of each batch) by manual inspection. Cracks can be detected 

visually and also acoustically: two balls are knocked together and a 

characteristic high pitch sound is heard if one of them is cracked. 

Acoustic, or "sonic" inspection is probably one of the oldest N.D.T. 

method. The sound of a body's natural resonances has been used to evaluate 

material properties (1, 2, 3), size and shape (4), and also for flaw 

detection (5). In the case of grinder balls, a study of the acoustical 

properties of good and cracked balls confirmed that repeatabledifferences in 

their natural sounds could be electronically detected. These results are 

presented in the first part of this article. 

An automatic grinder ball tester was designed and both a laboratory and 

an industrial prototype were built and tested. The second part of this 

article contains description of the laboratory machine along with some test 

results.

- 277 - 

1. THE ACOUSTIC HAVE GENERATED IN AIR BY THE IMPACT OF GRINDER BALLS 

When two grinder balls are knocked together in ambient air, an impact 

sound is generated which has several distinct features. Fig. 1 shows two 

typical waveforms picked up by a Bruel & Kjaer 1/4" microphone (# 4135) flat 

to 70 kHz and captured on a digital waveform analyzer. The waveform on the 

left was produced by the impact of two good balls while the one on the 

right, by the impact of a good and a cracked ball. The balls, 3" in 

diameter, are hand held with the target ball near the microphone and the 

"anvil", always a good ball, on the opposite side. A close examination of 

these waveforms reveals several distinct features. First of all, an initial 

pulse is always present at the beginning of each waveform, whether the balls 

are cracked or not. Also common to both waveforms is the low frequency 

rumble that follows. This rumble is due to reflections on surrounding 

structures and low frequency room resonances. Tests conducted in an 

echo-free room did not show such low frequencies. Following the initial 

pulse, high frequency resonances are present in both waveforms. In the case 

of the "good" impact waveform, the amplitude of the resonance is small, its 

frequency is around 40 kHz and its damping is low. In the case of the 

"cracked" impact, the resonance's amplitude is much greater, while its 

frequency is lower (around 5-15 kHz) and it is more damped. These 

features were individually studied and the results are presented in the 

following discussion. 

The initial pulse 

An initial pulse of width around 200 ys such as the one of Fig. 1 

appears at the beginning of every waveform recorded, whether good or cracked 

balls are used for the impact. This relatively low frequency pulse 

( 3 kHz) can be explained by considering the overall motion of the balls as 

a source of acoustic waves in air. 

Let us first recall the relation between particle displacement and 

acoustic pressure p for plane waves in air (6): 

p = - poc 2 

where: po is the density of air 

c is the velocity of sound in air 

x is the equilibrium position. 

We also have for plane waves propagating in one direction in air: 

Replacing in (1), we obtain: 

Se 

ôx 

e = f(ct - x) 

6e 1 6e 

i.e. — = 

6x c 5t

- 278 - 

6e 

P = Po c - St 

or p = p0 eu where u is the particle velocity 

Thus an infinite flat surface moving in air at a velocity u(t) will 

generate a pressure wave p(t) as given by Eq. (4), considering that the 

particles in contact with the surface have the same displacement. For 

spherical surfaces, one can no longer assume plane waves which means that a 

directivity factor along with a radial dependence on amplitude would 

appear. Still the time dependence of p(t) and p(t) would essentially be the 

same for points very near the surface. Hence, when an object is, for 

example, accelerated, its front face generates a compressive wave while its 

back face generates a rarefaction wave. This is easily verified 

experimentally as illustrated in Fig. 2: a grinder ball is accelerated as 

it is struck by an other one. When the microphone is placed near the front 

face of the target ball, it registers a positive pressure pulse. When 

placed at the rear of the target ball, it registers a negative pressure 

pulse. The ball thus acts as a dipolar source of length approximately equal 

to d, the diameter of the ball. 

This was also verified by measuring the width At of pulses obtained 

using grinder balls of various diameter. The microphone was located on the 

collision axis as illustrated in the upper part of Fig. 2. Therefore we 

expect the fall in pressure to be delayed with regards to the rise in 

pressure by the propagation time over a distance equal to the diameter "d" 

of the ball, thus producing a positive pulse of width At equal To d/c, where 

c is the velocity of sound in air. Fig. 3 is a plot of t vs. d. We see a 

good agreement with this model as the experimental data points cluster 

around the straight line of slope equal to 1/c. 

We can also look at the shape of the pulse and relate it to p(t), the 

time evolution of the velocity of the ball. In particular, the rise time of 

the pulse is expected to be equal to the rise time of the velocity of the 

ball i.e. the frequency content of the pulse would be limited by the 

interaction time T of the collision. We therefore calculate this 

interaction time T for the simple case of two spheres of diameter d and mass 

m colliding at a velocity v0 the relative velocity of one sphere with 

regards to the second one is 2 . v0. 

The symmetry of the problem allows the second sphere to be replaced 

with a fixed and infinitely rigid wall. With position x = 0 and time t = 0 

chosen when the sphere touches the wall, we have (7): 

F = - 5.182 fi (d.x 3 H 

where F is the reaction force of the wall on the sphere. When t = T/2, 

the work of that force on the center of mass of the sphere has brought it to 

a halt. Therefore we have: 

jxm F(x)dx = 1/2 m vo 2

- 279 - 

where x8 is the maximum value of x. Solving for xm yields: 

/ -VA 

xm=l.A22l _ I 

\ E 2 d / 

We can approximate the equation of motion of the sphere by a sine: 

x xm sin uit 

and knowing that when t = 0, dx/dt = vo, we have: (o = vo/xo 

The interaction time, T, corresponds to a half cycle of the sine wave: 

Thus, finally: 

T = TT/ÜJ 

= IT Xm/Vo 

In the case of 3" diameter balls, the numerical parameters are: 

m = 1.8 kg 

d = 76 ram 

E = 2 x 10 5 MPa 

vo= 1 m/s 

width yields: 

T = 280 us 

This value is in good agreement with the experimental data and offers 

and explanation to the absence of frequency components higher than 2-3 kHz 

in the initial pulse. 

Resonances of uncracked grinder balls 

Sound measurements of uncracked grinder ball impacts were made using a 

high pass filter to attenuate the low frequency noise associated with the 

initial pulse (up to 3 kHz as mentionned above). The filtered signals show 

the enhanced presence of a high frequency resonance in the acoustic waveform 

ranging from 35 kHz for 3" dia balls to 65 kHz for H" dia balls. This 

oscillation can be related to one of the lower modes of resonance of a solid 

steel sphere. Fig. 4 is a plot of the inverse of the resonance frequency as 

a function of ball diameter. The straight lines represent theoretical 

curves for the first 3 modes of vibration of solid steel spheres as 

calculated from relations given in (8), and with elastic constants given in 

1/5

- 280 - 

(9) for hardened steel. All other modes fall below the 2S1 mode and so were 

not considered. The data points indicate that either the 1S2 mode or the 

1T2 mode is the one that generates the high frequency acoustic waves. 

However, for torsional or "T" modes in spheres there is no radial 

displacement and therefore it is assumed that these modes do not radiate any 

sound in air. Only spheroidal or "S" modes can therefore couple with air 

and generate an acoustic wave so that the 1S2 mode is believed to be the one 

responsible for the high frequency tone. A representation of the 

displacement associated with the 1S2 mode is included in Fig. 4. It shows 

that this mode is very likely to be strongly exited by an impact because 

such a uniaxal force will ovally deform the sphere. Other modes are 

probably exited as well, including higher S modes. Some inharmonic 

distortion was indeed observed on several high frequency waveforms. This is 

expected if other modes are exited since their frequencies are not exact 

multiples of each other. 

Resonances of cracked grinder balls 

When similar acoustic measurements are made using cracked grinder 

balls, the appearance of medium range frequencies (5 - 15 khz) in the 

waveforms is observed. These oscillations look like damped resonances. 

Disappearance of the high frequency 1S2 vibration is also observed, 

especially for severely cracked balls for which the amplitude of the 

mid-range resonances is often the highest. This amplitude also depends on 

the point of impact i.e. the orientation of the cracked ball relative to the 

anvil. This effect however varies from one cracked ball to another and no 

repeatable pattern emerged. 

The mid-range resonances, characteristic of cracked grinder balls, are 

attributed to some sort of "tuning fork" mode of vibration as the ball is 

partially divided by deep running cracks. Acoustic measurements made on 

artificially flawed grinder balls (saw-cut) reveal the presence of at least 

two modes of vibration at frequencies lower than the 1S2 mode of the solid 

sphere. The data is presented in Fig. 5 where the frequencies of the 

various resonances observed are plotted against the depth, a, of the 

saw-cut. We see that mid-range frequencies similar to those produced by 

naturally cracked balls appear for values of a in the range of 30% to 70% of 

the diameter d of the ball. The 1S2 amplitude becomes unmeasurable when a 

reaches about half the diameter of the ball. The assumed displacements 

corresponding to the two "tuning fork" modes are also illustrated in 

Fig. 5. This assumption is in agreement with the observed variations in the 

relative amplitude of the two frequencies with the orientation of the 

impact: when directed perpendicular to the plane of the saw-cut, the lower 

frequency mode was preferably exited whereas when directed parallel to the 

bottom of the saw-cut, the higher frequency "shear" mode was preferably 

exited. The frequency of both modes understandably tends towards zero when 

the value of a approaches the value d.

281 - 

Finally, resonances of saw-cut balls were practically undamped compared 

to those of naturally cracked balls. Evidently the latter are damped by 

friction occuring along the crack interface. Consequently, one can also 

assume that the shear motion depicted in Fig.5 for the "shear" mode is 

unlikely to occur in naturally cracked balls. 

2. DEVELOPMENT OF AN ACODSTIC CRACK DETECTOR FOR SORTING GRINDER BALLS 

Principle of operation 

We have seen that noticeable differences exist between the sound 

produced by the impact of uncracked grinder balls and the sound produced by 

the impact of cracked ones. From this, it was concluded that a device that 

would use these differences to acoustically detect and sort out the cracked 

grinder balls from the good ones was technically feasible and a laboratory 

scale prototype was developed. Figure 6 is a photograph of the device. The 

mechanical assembly allows the balls to drop from a constant height on a 

fixed anvil. It is composed mainly of 4" PVC tubing supported by a bolted 

"Dexion" frame. The anvil is a 2" steel cube mounted in a foam filled box. 

A 6" long aluminium rod mounted on a hinge inside the horizontal sorting 

tube and activated by a solenoid deflects the cracked balls toward the first 

opening, while the others exit by the end opening. 

The impact sound is picked up by a nearby microphone and electronically 

analyzed. The ball finally exits by one of two ports depending on whether 

or not a crack has been detected. The electronic circuit "listens" for 

those mid-range resonances typical of cracked grinder ball impacts that 

immediately follow the initial pulse. It measures the "amount" of such 

resonances and uses it as a criterion to decide whether the ball is good or 

bad. Figure 7 is a simplified diagram of the circuit. The microphone is a 

Bruel & Kjaer i" microphone and power supply flat to 30 kHz. It is located 

in the target area, as near as possible to the trajectory of the grinder 

balls, positioned downwards to prevent any buildup of dust or other 

pollutants on its membrane. The detector consists of a band pass filter 

(10-20 kHz), a rectifier and a gated integrator. The output voltage thus 

represents the amount of acoustic energy between 10 and 20 kHz integrated 

over the time window defined by a gate signal. It is compared to a 

reference voltage to determine whether the ball is cracked or not. The 

logic block carries out the measurement sequence. It uses the "initial 

pulse" as a trigger event, resets the integrator and holds it on reset for 

about 1 ms. This adjustable delay determines the start of the time window 

and allows the blanking out of the initial pulse. It then frees the 

integrator for about 10 ms and then interrupts the integration. This 

defines the width of the time window, the end of which is positionned just 

before the occurence of a second collision due to the recoil of the anvil 

bouncing back on the grinder ball. While the time window is open, the 

output of the integrator is compared to an adjustable reference and the 

output of the comparator commands the operation of the sorting device. As 

the time window closes, the trigger of the logic block is disabled for about 

100 ms to prevent retriggering on the second collision mentionned above or 

any other spurious noises. After that delay, it is ready to be retriggered 

when an other ball hits the anvil. Until then, the integrator is on hold,

- 282 - 

but not reseted, so the comparator, and therefore the sorting device, stays 

in the same state. 

One major constraint is the signal-to-noise ratio. Sources of acoustic 

noise, especially those occurring at the moment of impact and in the 

frequency range of cracked ball resonances, have to minimized. Of 

particular importance in this matter is the anvil. It has to be designed so 

that the frequency of its lowest mode of resonance is much higher than 

cracked ball resonant frequencies. The resonant frequencies of the 

prototype's anvil, a 2" steel cube, were measured and found to be all higher 

than 25 kHz. The anvil has to be mounted on some sort of damped suspension 

(foam rubber in our case) so that when it is struck by a grinder ball, the 

high frequency part of the impact force is not transmitted to the supporting 

structure and therefore does not ring high frequency resonances in it. For 

this reason and also to maximize the portion of the impact energy 

transmitted to the grinder ball, one would tend to make it as massive as 

possible. The limiting factor is that the resonant frequencies diminish 

with increasing size. It is possible to increase the mass of the anvil 

without decreasing its resonant frequency by optimizing its shape. In this 

respect, the sphere has the highest mass . frequency product followed by the 

square cylinder (length equal to diameter), followed by the cube etc... The 

cylinder is probably the most convenient compromise for automated systems 

but, in theory, the best anvil to test a particular size of grinder ball is 

another uncracked one of the same size. 

Another design constraint is that the rate at which grinder balls can 

be tested is limited by the travel time of a ball from the moment of impact 

to the moment it clears the mechanical sorting device; the electronics are 

not the limiting factor, as the good or cracked diagnosis is done before the 

ball actually leaves the anvil. This travel time was more than half a 

second for the rudimentary sorting tube of the laboratory prototype but 

could easily be improved for an industrial machine. 

Field Test Results 

Sorting tests were conducted mainly on 3" diameter balls- A sample of 

approximately 100 3" diameter balls was carefully hand sorted so as to form 

a "good" pile and a "cracked" pile. Each time a ball was put through the 

machine, the integrator output level was recorded. The data is presented in 

the form of histograms (Figure 8). A normal curve has been drawn through 

the good ball distribution and the 2.5 V. threshold is indicated. 

Ideally, the distributions of good balls and of cracked balls would be 

two delta functions and no matter how close to one another they would be, 

one would always be able to position the threshold between them and achieve 

100% success rate. However, in reality, a number of factors widen the 

distributions causing overlapping and, thus, diminishing the success rate. 

In the case of cracked balls, a number of those factors are uncontrollable. 

They are: severity of the crack, geometry of the crack, direction of impact 

with regards to crack and microphone orientation with regards to crack. The

- 283 - 

effect of the last two could be averaged out by having each ball run through 

the test several times but this is not necessary to achieve satisfactory 

success rates. Other factors that widen all the distributions are 

controllable to a degree. These are: velocity of the incoming ball, 

direction of impact and location of impact on the anvil. All these 

parameters are rerated to the trajectory of the incoming ball. The more 

repeatable it is, the finer the distributions will be. This allows the gain 

to be increased, moving the good ball distribution closer to the threshold 

and thus pushing the cracked ball distribution further away from it. The 

distributions obtained with this prototype were narrow enough to obtain very 

good success rates (97X of cracked balls recognized with no good balls 

misdiagnosed). However, one parameter in particular, ball velocity, varied 

significantly and is believed to cause most of the spreading of the 

distributions. A significant improvement is expected when the feeding 

system is redesigned to fit a tester on the production line. 


The author wishes to thank Mr. Ghislain Vaudreuil of the Industrial 

Materials Research Institute and also Mr. George Chapman of the Steel 

Company of Canada for their precious contributions to this work. 

REFERENCES 

1. Grime, G., "Determination of Young's Modulus for Building Material by a 

Vibration Method", Philosophical Magazine, Vol. 7, No. 20, 1935, 

pp. 304-310. 

2. Prigge, R.E., "Correlation of Modulus of Rupture and Modulus of 

Elasticity", ß.S. Thesis, New York State College of Ceramics, 

(May 1951). 

3. Hornibrook, F.B., "Discussion on Sonic Method for Modulus of 

Elasticity", A.S.T.M. Proceedings, Vol. 39, 1939, pp. 996-998. 

4. Rubin, G.A., Leach, M.F., "Particle Size and Shape Characterization from 

Acoustic Emissions", Proc. 82nd Annual Conference of the ACS/CEC/NCE, 

Chicago, April 1980. 

5. Rowe, K..G., "Vibration Apparatus for Testing Articles", U.S. Patent 

#2 486 984 (Nov. 1949). 

6. Kinsley, L.E., Frey, A.R., Fundamentals of Acoustics, 2nd edition, 1962, 

p. 112. John Wiley & Sour, Inc., New York. 

7. Roark, K.J. & Young, W.C.. Formulas for Stress and Strain, 5th edition, 

1975, p. 516. Me Graw-Hill, Inc., New York. 

8. Schreiber, E., Anderson, O.L. & Soga, N. Elastic Constants and their 

Measurement, 1st edition, 1973, pp. 126-142. Me Graw-Hill, Inc., 

New York. 

9. Krautkramer, J. & H., Ultrasonic Testing of Materials, 2nd edition, 1977, 

p. 620. Springer-Verlog, Berlin.

- 284 - 

GOOD CRACKED 

0 2 4 6 8 0 2 4 6 8 

t(ms) 

Figure 1: Acoustic waveforms produced by the impact of good balls (left) 

and cracked balls (right). 

0 2 4 6 8 0 2 4 6 8 

t(ms) 

Figure 2: Influence of the nicrophone location on the polarity of the 

initial pulse.

At (/xs) 

300 r 

200 

100 

- 285 - 

y 

/i-.\ Slope = c 

_ i i i i i 

0 20 40 60 80 100 

d (mm) 

Figure 3: Width At of the initial pulse, as a function of d, the ball 

diaaeter. 

(X10 5 S) 1T2 

j i I i i i 

20 40 

d(mm) 

60 80 

Figure 4: Besonance period as a function of ball diaaeter.

0 

- 286 - 

Figure 5: Resonance frequencies of saw-cut balls as a function of the depth 

"a" of the cut.

- 287 - 

Figure 6: Grinder ball crack detector and sorter. 

DETECTOR 

Figure 7: Schematic diagran of the crack detector.

- 288 - 

GOOD 

0 .8 1.6 2.4 3.2 

Level (V) 

4 5.6 

Figure 8: Histograms of integrator output level for good and cracked balls.

- 289 - 

THE BENEFITS OF NDT TRAINING FOR CANADIANS 

Lynda B. Uanzzn. 

Canadian Soc.ie.ty ^on. Monde.* tractive. Tz&ting Foundation 

HAMILTON, Ontario 

Although industry is becoming increasingly aware of the 

benefits to be derived from nondestructive testing, many of these 

same firms have yet to realize the importance and impact of effectively 

educating NDT personnel. 

The author will outline the history of NDT training in 

Canada, the types of programs presently available, the typical 

student profile, and who the target market could (and should) be. 

The author will then discuss the effects of NDT training and education 

in Canadian industry.

- 290 - 

THE BENEFITS OF NDT TRAINING FOR CANADIANS 

It is generally ac'cepted that the role of nondestructive testing is 

to prevent human injuries and to save lives, to increase productivity, 

and to make a profit for the user. For those of us who have heard, 

or used, this phrase over and over through the years, let us not allow 

repetition to desensitize us to the enormous responsibility that this 

places on those involved in nondestructive testing (NDT). 

This inherent factor of responsibility demands that personnel responsible 

for, or involved with, a company's NPT section receive some 

form of post-secondary school education in this technology. A technology 

which is essential in an industrial nation such as Canada. 

Yet the history of NDT training in Canada, and the numbers and types 

of students presently in training, would seem to indicate that the 

educational component has not kept pace with the phenomenal growth 

in the use and applications of nondestructive testing. That may 

seem a shocking statement from someone such as myself, a representative 

of the Canadian Society for Nondestructive Testing Foundation 

which is the leading educator in NDT in Canada. But who better to 

relay to you our concern that not enough people in business and industry 

realize the importance of training their personnel, or hiring 

well-trained, competent individuals to carry out NDT functions, or 

to supervise these functions. 

In the early days of NDT, much of the training was conducted on-thejob; 

personnel became proficient (or, did not become proficient) 

through trial and error. In 1940, the Society for Nondestructive 

Testing (SNT) was established in the United States, and in 1953, the 

first Canadian Section of SNT was formed. In 1954, a second section, 

the Eastern Canada Section of SNT, was formed in Montreal. A small, 

select group of Canadians realized the importance of nondestructive 

testing and the importance of disseminating information on NDT technology 

. 

These Sections conducted a few continuing education courses and, in 

effect, the pattern was set for NDT training in Canada to be the 

short, intensive-type of course. In the 1950's, and even today, 

equipment suppliers offered training courses which followed the same 

pattern - courses of three to five days duration. 

In 1964, the Canadian Council for Nondestructive Technology (C.C.N.D.T.) 

was formed with the following objectives: 

-to advance scientific, engineering, and technology knowledge 

in the field of nondestructive testing. 

-to gather and disseminate information relating to nondestructive 

testing useful to the individual and beneficial to the 

general public. 

-to promote nondestructive testing through courses of instruction, 

lecture, meetings, publications or other means.

- 291 - 

The Canadian Society for Nondestructive Testing (C.S.N.D.T.) was 

formed from C.C.N.D.T., and with the similar objectives, in 1967 

and the NDT community had a truly Canadian Society. The increased 

interest in NDT an,d support of a National Society in the late 50's 

and early 60's paralleled a new emphasis, in Canadian education, on 

mathematics and sciences. It may be that the Vocational Training 

Act, passed in 1961, emphasized the need for training in new technologies, 

although the federal dollars did not filter through to a 

Canadian institution specifically for NDT training. 

It was also in 1960 that the Canadian Government Specifications 

Board (C.G.S.B. and now the "Canadian General Standards Board") introduced 

C.G.S.B. Standard 48-GP-4 "The Certification of Industrial 

Radiographie Personnel". This Standard established a Junior and 

Senior level of certification, outlining the requisite amount of 

work experience and the responsibilities of each level. Standards 

for certification in Ultrasonics, Liquid Penetrant, Magnetic Particle 

and Eddy Current followed. 

C.S.N.D.T. Chapters across Canada continued to offer continuing education 

programs and increased the educational input of the Society by 

conducting seminars. By the mid 70's, it was evident that assistance 

was needed to provide industry with trained NDT personnel from the 

Canadian labour force. It should be noted that some companies were 

still finding it necessary to hire Europeans who had good NDT backgrounds 

and to send personnel to the United States for training 

courses. Also at this time, the Department of Energy, Mines and 

Resources (the Examining Authority for C.G.S.B.) indicated their need 

for assistance in conducting the certification program. 

In 1976, the educational arm of C.S.N.D.T. was established as the 

Canadian Society for Nondestructive Testing Foundation. This nonprofit 

organization, headquartered in Hamilton, Ontario, has three 

main objectives: 

- to improve the quality of education in nondestructive 

testing throughout Canada. 

- to assist Canadian industry in its use of nondestructive 

testing 

- to assist the Department of Energy, Mines and Resources 

in conducting the written and practical examinations for 

certification of nondestructive testing personnel, in 

accordance with the Standards of the Canadian General 

Standards Board. 

These objectives outlined a major undertaking; previously, there had 

not been an organization charged with the responsibility of providing 

intensive, daytime NDT programs. The Nondestructive Testing Section 

of the Department of National Defense, Aircraft Maintenance Development 

Unit, based in Trenton, Ontario had been (and still is) a major 

contributor to education and certification in Canada. The Foundation's 

responsibility, however, is to the entire spectrum of NDT in Canadian 

i ndus try.

- 292 - 

C.S.N.D.T. Chapters continue to conduct evening educational programs, 

some of which are in co-operation with local community colleges. 

This automatically limits the geographical area served. The Foundation, 

in order to meet its objectives, must provide programs wherever 

and whenever they are needed. 

In its first full year of operation, 1977, the C.S.N.D.T. Foundation 

conducted eight (8) courses with a total student enrolment of one hundred 

and fifteen (115). These courses were the short, intensive-type 

and classed as "Regular Courses"; in other words, open to industry in 

general. 

In 1978, the Foundation added Practical Workshops and In-Plant Trainin] 

to its list of available programs. The following chart gives a statistical 

account of student enrolment from 1977 to 1983: 

FIGURE 1 

C.SMJDX FOUNDATION 

STUOGMT ENROLMENT 

I97Y ~1963 

1977 1976 1979 1980 )9Q\ (982 1983 

Regular courses -Practical Workshops In-Plant Training

- 293 - 

These figures are much more than an exciting and encouraging indication 

of the Foundation's growth and its established reputation. Student 

enrolment in these types of programs, conducted by the Foundation, 

grew from a total of 115 in 1977 to 758 in 1983 - an increase of 559%! 

Furthermore, the C.S.N.D.T. Foundation is not the only NDT educator in 

Canada. Companies run their own in-house educational programs, equipment 

suppliers offer NDT courses, and C.S.N.D.T. Chapters have continued 

with their evening courses. There is, very definitely, an increase 

in the recognition of the importance of educating NDT personnel. 

In Figure 1, note the changes in the enrolment figures for Regular 

Courses and In-Plant Training Programs between 1981 and 1983; a decrease 

in attendance at Regular Courses and a dramatic increase in 

In-Plant Training Programs. 

There has also been a change in the students' training objectives 

since 1981. Prior to this, I estimate that 80% of the students in any 

of these programs were preparing for C.G.S.B. certification. After 

1981, there has been little change in the training objectives of students 

in Regular Courses, but a decrease to approximately 50% of the 

students in In-Plant Programs who are preparing for certification. 

Many of the companies now contracting with the Foundation to teach "onsite" 

are providing their employees with a general knowledge of nondestructive 

testing. These are not necessarily NDT technicians; they 

may be engineers, marine surveyors, procurement inspectors, quality 

assurance managers, quality control managers, production superintendents 

and foremen - all of whom are representative of the disciplines which 

must inter-relate with the actual NDT personnel in order to ensure an 

economical, quality product. The increase of programs of the "need to 

know" type, as opposed to the "how to do", underlines the growing 

recognition that knowledge of nondestructive testing is essential to 

the decision makers. 

I am not advocating, nor am I suggesting, that there has been a swing 

away from the certification of NDT personnel. To the contrary, Canada 

has one of the best certification systems in the world, and if more 

companies would require C.G.S.B. certification in their inspection 

procedures, at least, we would be assured of the minimum level of 

competence of their NDT technicians. We are well aware of the ramifications 

of a nondestructive test improperly performed or improperly 

interpreted. 

Then why aren't more of us, who know nondestructive testing and understand 

our responsibility for the welfare and safety of the Canadian 

public, demanding that all NDT technicians be educated and trained in 

specific applications, and their knowledge verified? In order to ensure 

top quality, professional and ethical technicians, we need: 

- a responsible employer 

- a receptive trainee 

- a responsible educator 

- a responsible regulatory body 

AND 

- a responsible, professional National Society

- 294 - 

The groundwork is in place. The increased emphasis on education, 

which has already been pointed out, has to reflect on the employers. 

In my own contact with hundreds of students, I have noticed a change 

in attitude; more ßtudents have a real desire to learn, and to learn 

more than just the basics. We certainly have responsible educators, 

the Foundation being one of the leaders in this field. 

Through the Canadian General Standards Board, and its Examining 

Authority (E.M.R.), we have a responsible regulatory body. The certification 

standards have been changed to a three level system (rather 

than Junior and Senior) and classroom training is now mandatory, in 

addition to the required work experience. There are approximately 

3,000 certified NDT technicians in Canada and, until the recent economic 

recession, there had been a steady increase in examination candidates. 

The C.S.N.D.T. Foundation, fulfilling its objectives, conducts 

about 60% of the required practical and written examinations; Figure 

2 shows the numbers of candidates who have made use of the Test Centre 

facilities at the Foundation from 1977 to 1983: 

FIGURE 2 

300 

CERTIFICATION EXAMINATIONS CC.G-S.ß.) 

CONOUCTBO BY THE C.S.N.P.T. FoUMDATIOIV 

\977 ~ \9S3 

1977 1978 1979 1980 

- V«nrrer* EXAMINATION* 

EXAMINATIONS 

1981 1982 1983

- 295 - 

In the Canadian Society for Nondestructive Testing, we have a responsible 

National Society. The.Society provides its members with an excellent 

technical Journal, communicates at the national and international 

level with other technical Societies, provides representation 

on the C.G.S.B. subcommittees for certification of NDT personnel, 

and conducts seminars and conferences. These services are vital in 

keeping the NDT community in Canada up-to-date on recent developments 

and providing a forum for discussion. C.S.N.D.T. has taken a major 

step in ensuring high quality, uniform education for nondestructive 

testing personnel in Canada. I refer to "C.S.N.D.T. Procedure - ED-1" 

which states : 

This document identifies the conditions and procedures 

under which the Canadian Society for Nondestructive 

Testing Inc., and the Canadian Society for Nondestructive 

Testing Foundation (hereafter referred to as the 

Society and the Foundation) will issue certificates of 

education to students who have attended in any of the 

NDT methods identified in this document. 

The use of this procedure is mandatory and is to be 

adopted by the Foundation and all Chapters of the 

Society. Other organizations may adopt this procedure 

in order to conform with standardization of NDT 

training in Canada; such organizations must comply 

with ED-1 in its entirety for approval by the National 

Commi 11ee. 

ED-1 is meant to supplement the C.G.S.B. certification standards, not 

to replace them. The Society has simply gone a step further and provided 

recommended guidelines regarding the educational material, 

instructors' credentials, and examination format for the educational 

programs. Again, the objective is to ensure high quality, uniform 

education for NDT personnel. 

N.D.T. training in Canada has come a long way in a relatively short 

period of time. At this conference, we are celebrating the major 

strides that the Society has taken in just twenty years and I congratulate 

the Society for its emphasis on education. 

What strides will be taken in the next twenty years? The strides that 

are taken are going to have to be in step with a rapidly changing technology. 

Many of the registrants at this conference will listen to presented 

papers with NDT applications which they had not thought of, or 

perhaps, thought possible. But how many people, and how many of the 

disciplines 'hat inter-relate with NDT, are we reaching? 

In any technology, there is a frustrating (and potentially dangerous) 

gap between the engineer and the technician. Within the next five years 

I believe those of us in the NDT sector of Canadian industry are going 

to be the first to realize the impact of a fairly recent addition to 

some community college semester programs. In a course entitled "Materials 

Testing", students are taking Mathematics, Chemistry, Physics, Foundry 

Practice and Metallurgy, Electrical Systems and Instrumentation, Welding 

Computer Concepts, Statistics, Mechanics of Solids, Nondestructive and 

Destructive testing, and Language and Report Writing.

- 296 - 

The graduates of these programs, along with the growing numbers of 

Level III NDT technicians, will help to fill the gap. However, 

engineers must also be educated in NDT technology. I recently had 

the pleasure of hearing Dr. David W. Hoeppner, of the University 

of Toronto speak on "Failure Analysis and NDT Interface". Dr. 

Hoeppner, an educator at the university level, is appalled at the 

lack of a solid component of nondestructive testing techniques in 

the education of our engineers. 

What will be the outcome of an even greater emphasis on nondestructive 

testing education in Canada? The saving of human lives, an 

increased prevention of human injury, higher productivity and the 

saving of billions of dollars. 

In the United States, the National Bureau of Standards and Battelle 

Columbus Laboratories conducted a study to assess the costs of 

material fracture to the United States for the year 1978. In their 

report'-, it states 

"The costs were large. In 1982 dollars the total cost was 

estimated to be $119 billion per year, about 4 percent of 

the gross national product. An estimated $35 billion per 

year could be saved through the use of currently available 

technology. Costs could be further reduced by as much as 

$28 billion per year through fracture-related research... 

The study concluded that substantial material, transportation, 

and capital investment costs could be saved if technology 

transfer, combined with research and development, 

succeeded in reducing the factors of uncertainty related to 

structural design. Equally safe or safer structures could 

be produced with substantial cost savings in material, 

transportation, and capital investment. This could be done 

by reducing the uncertainty currently related to structural 

design through better predictions of structural performance 

from materials properties, better process and quality control 

and in-service flaw monitoring, and less materials 

variability." 

The same solutions and cost savings, albeit with a lower dollar 

figure, apply to Canadian industry and, what dollar figure do we 

apply to the value of human life. 

The Economic Effects of Fracture in the United States 

Special Publication 647-1 

U.S. Department of Commerce - National Bureau of Standards 

R.P. Reed, J.H. Smith, B.W. Christ 

Issued March 1983

- 297 - 

FEAR DETECTION AND REMOVAL 

THE PSYCHOLOGICAL IMPLICATIONS OF THE TECHNOLOGICAL APE 

Tanii HdlliiMe.il 

With the advent of the technological age many age old fears of 

mankind have resurfaced. What is Fear? How does fear impact on all of our 

lives today? How can we learn to detect and remove fear from our lives and' 

help others to do the same? 

Fear for most of us indicates a loss of control in a situation and it 

leads to feelings of helplessness and powerlessness in our life. With the 

ever increasing speed and frequency of change surrounding us in our world 

many of us are thrown off balance, out of control and into the ocean of 

fear. 

Fear of failure, fear of rejection, fear of change and fear of the 

unknown are particularly relevant for people in industries which are becoming 

increasingly high tech such as the petrochemical, nuclear, plastic, 

steelraaking, and refineries. Older workers worry "Have I got what it takes 

to keep up?" Younger workers question "Will I be the first to go if 

employees are laid off and computers take over?" Managers wonder "How can I 

motivate my employees and allay their fears when I'm not sure what is going 

to happen myself?" 

These fears can lead us into depression, burnout, passivity, 

aggression, and work aholism. Through recognizing our fears and our reaction 

to them we gain insight into our roles and behaviours, our relationships and 

our communication with others. We must learn to free ourselves from our 

fears if we are to motivate ourselves and others to adapt to the changing 

world around us so that we can be active participators and not victims of 

these changes.

- 298 - 

CERTIFICATION OF NONDESTRUCTIVE TESTING PERSONNEL IN CANADA 

AN UPDATE 

I'. Care n 

CAWUET, Ottawa, Ontario 

ABSTRACT 

This paper surveys both the national and international scenes regarding the certification 

of nondestructive testing (NDT) personnel. In Canada, the increasing 

demand for NDT personnel and the implementation of the M-Standards have led to the 

need for automating the administration of the program. Concurrently, the increased 

interest towards harmomizing national certification schemes has steered activities 

in both the ISO and the ICNDT forums. 

A) THE NATIONAL SCENE 

As is well-known, certification of nondestructive testing (ndt) personnel in 

Canada is a fully bilingual operation which is carried out through a central agency 

which is the federal Department of Energy, Mines and Resources. The certification 

standards are issued by the Canadian Government Standards Board (CGSB) 

and they contain all the requirements and procedure details underlying the certification 

process. Some 35 to 40 persons representing a broad range of interests 

are members of the national CGSB Committee for the Certification of NDT Personnel 

which meets every six months to review overall activities and to resolve any major 

issue arising from the certification programme. A Steering Committee supports the 

main committee. 

Certification initially started with Industrial Radiography in 1960, followed by 

ultrasonics in 1970, magnetic particle and liquid penetrant in 1971. Certification 

in eddy-currents did not start until 1982. The graph of Fig 1 illustrates 

the growth of the certification programme since 1970 by showing the number of 

certificates issued annually for the first four methods. Obviously, level I 

radiographers had the strongest growth during the past thirteen ( L3) years but 

also dropped considerably in 1983. 

The total number of certificates issued as of January 1, 1984, in each method, 

since their respective implementation date, is as follows: 

LEVELS RT (I960) UT (1970) MT (1971) FT (1971) ET (1982) 

I & II 3708 1878 1385 1490 175 

for a grand total of about 8636 excluding Level III. Well over 60% of the certified 

personnel holds certification in at least two (2) methods or more.

- 299 - 

A major event took place in 1979 when the new CGSB M- standards started being 

implemented. The advent of a new approach to ndt personnel certification followed 

a long period of monitoring the world trend towards a three—level system, particularly 

under the infLuence of SNT-TC-1A which dates from 1965. The resulting Mstandards 

retained a general approach and a central examination system but were 

extended to include a third level of competence. Provisions for employer's specific 

certification were added to the standards. In the process, training courses 

became compulsory and industrial requirements became less stringent. 

A major difference between the two A- and M- standards involved the amount of 

industrial experience needed for eligibility to examination and the new requirement 

of formal training courses. The A- standards required considerably more 

experience but training was not mandatory. As an example, to appreciate the 

trade-off between training and experience for purpose of eligibility, one calculates 

an equivalency or a equivalent trade-off of one week of training for one year 

of industrial experience for the particular case of radiography Level II. 

Impact of m-standards on administration 

This heading describes the status of a few problems associated with the introduction 

of the M-Standards. 

(a) The change from A- to M- Standards meant that over 30 papers totalling over 

60 hours of examination time would have to be prepared instead of the wellestablished 

11 papers lasting 19 hours. Accordingly, it was necessary to replace 

descriptive by multichoice questions. Considerable time savings can be achieved 

mainly from marking papers. 

(b) In order to mimimize personnel needs preparing written examinations, a bank 

of coded examination questions stored on diskettes is approaching completion and 

will contain both French and English versions. Print-outs and answer sheets allow 

to save much time in respectively preparing and marking examination papers. 

(c) The mandatory annual renewal of certification has become over the years a 

cumbersome operation; for instance, more than 6000 pocket certificates were issued 

to some 2800 persons in 1984. Software is being developed to automate those applications 

for renewal. Progress was impeded in 1984 when the Division undertook 

a global analysis of in-house computer requirements for both research and administrative 

operations. However it is hoped to achieve progress and run dry tests in 

the coming months once the computer mainframe become available. 

(d) The problem of space has become critical as it must compete with R&D laboratory 

needs. It was necessary to radically change filing customs from previously 

holding all examination material to nearly discarding all such material, except of 

course for resuLts. Even at that, space needs are increasing but resources are 

not presently available to put in a computer all the desirable information material, 

statistics, background information etc. However, the basic information which 

is presently recorded in the Cardex systems is being transferred onto diskettes: 

name, address, age, examination results, ndt method, date of certification.

- 300 - 

(e) A differing but persisting problem since the M- Standard on Radiography was 

introduced in 1979 concerns the Aircraft Structures category. The General category 

includes Welding, Casting and Forgings. The Aircraft Structures is a 

category which now pre-requires the General certification whereas the absolete Astandards 

did not have such a requirement. That extra requirement has always encountered 

strong opposition from both the civilian and military users who are only 

interested in maintenance work. That issue has re-surfaced in recent times and 

will be given further study during the coming months by the national CGSB 

Committee. 

( f ) The certification network now includes nine (9) practical test centres but 

it would need to be enlarged in Eastern Canada but progress is slow, one reason 

being the difficulty to obtain defective metal parts from users and producers and 

the considerable time needed to analyse them and prepare documentation. 

A new certifying agency (?) 

Following an offer made by the Canadian Welding Bureau (CSA) to assume the responsibility 

of Certifying Agency, the possible withdrawal of DEMR is currently being 

given consideration within the interested forum. Considerable resources are being 

devoted to both building and operating the new certification programme and those 

resources are indeed conflicting with the basic mission of PMRL/CANMET which is 

R&D. A feasibility study will be conducted for submission Lo the next meeting of 

the CGSB Committee in April 1985. 

B) THE INTERNATIONAL SCENE 

This section briefly surveys the international developments within the optics of 

possible harmonization of national standards or practices. There is no doubt that 

the SNT-TC-1A practice for certification of ndt personnel which was introduced in 

1966 has exerted an important influence on the development and contents of such 

national documents. Their growing number reflects the necessity for measuring the 

qualification of ndt personnel in order to cope with the increasing demand and 

complexity of ndt requirements during manufacturing or service. 

Efforts at harmoning certification have been carried out for several years by Lhe 

ICNDT whereas ISO has become active only during the past year although its 

official commitment dates back to 1975. The secretariat for ISO-TC 135/SC7 on 

certification is now held by Canada which gives DEMR the responsibility for 

drafting an international document. 

(a) Perhaps the most striking feature which differentiates the various national 

schemes is the "general" and "sector" components of certification as being conducted 

by an independent central body, leaving the employer responsible for any 

additional features. From a survey of several national documents, it would appear 

that SNT-TC-1A is the only procedure giving the employers the sole responsibility 

for the whole certification. 

(b) The number of levels of certification may vary but are usually conceived to 

roughly correspond. Thus, a country with two levels will correspond to levels I 

and II. The range of qualifications ascribed to each level shows extensive

- 301 - 

sirailiarities but there may exist differences. For instance, a few standards 

allow Level I personnel to interpret/evaluate some results, whereas others do not 

allow such responsibility. 

(c) The categories or sectors of certification are often identical with 

castings, forgings, welds and aircraft structures/aerospace being commonly found 

but others may also be found. Specific areas may also be covered such as wrought 

steel plates and bars, types of welds (plate, pipe, nozzle), etc. 

(d) Examination schemes show some common features. Written and practical tests 

are applied by all schemes with oral tests being encountered in but a few schemes, 

either as a requirement or as a complement. Most schemes include a general 

written test, a sector written and practical tests. 

(e) Training requirements are common to most schemes but not all. Minimum 

hours may be specified and even related to a given degree of schooling. In order 

to avoid the implications of assessing and comparing national education systems, 

the ICNDT has endeavoured to determine a syllabus without a corresponding number 

of hours and schooling level. Industrial experience is usually mandatory. 

Several other headings are not examined herein but must also be considered, such 

as validity period, renewal conditions, health requirements, certificate ownership, 

etc. 

The above description points to some of the difficulties that might be expected 

towards harmonizing present and future national schemes for the certification of 

nondestructive testing personnel but there also exist definite commond grounds. 

The next ISO TC135/SC7 meeting will be held in February 1985 and the Board of 

Directors of the American Society for Nondestructive Testing recently expressed 

their objective to have an international standard finalized by 1988. Time will 

tell whether an international document or standard will be acceptable to the concerned 

countries and if so, what the time frame will be to achieve that objective.

250 

200 

150 

100 

50 

Magnetic Particle 

Level II O 

Liquid Penetrant 

Level II 

1970 72 74 76 78 80 82 1970 72 7A 76 78 80 82 1970 72 74 76 78 80 82 

FIGURE 1: Number of Certificates Issued Annually Between 1970 and 1983 Inclusive. 

I 

o 

I

DAY 3 

KEYNOTE ADDRESS: Looking into the Future 303 

- R.S. Sharpe 

NDT of Structural Ceramics by High Frequency Ultrasonics 307 

- A. Fahr 

Hot Pressed Piezo-Electric Ceramic Elements for Ultrasonic Transducers 316 

- N.D. Patel, J. van den Andel, P.S. Nicholson 

Ultrasonic Analysis of Voids in Glass: Theory and Practice 331 

- A.J. Stockman, J. van den Andel, P.S. Nicholson 

Computer Simulation of Ultrasonic Testing 345 

- D.B. Duncan 

A Novel Approach to Eddy Current Imaging of Defects 356 

- D. Leemans, M. Macecek 

Developments in X-Ray Stress Measurement, The CANMET Portable Stress 372 

Diffractometer 

- B.A. Holt 

Applications of Neutron Diffraction to Engineering Problems 387 

- T.M. Holden, G. Dolling, S.R. MacEwan, J. Winegar, B.M. Powell, R.A. Holt 

NDE in Polymers, an Example: Ultrasonics for the Determination of Density in 398 

Polyethylene 

- L. Piche, A. Hamel 

An Industrial Application of Computer Assisted Tomography: Detection, Location 408 

and Sizing of Shrink Cavities in Valve Castings 

- P.D. Tonner, G. Tosello 

Computer Operated Composite Panel Testing (abstract only) 424 

- S. DeWalle 

Some Unconventional Techniques for the Inspection of Layered Materials 426 

- P. Cielo 

Materials Effects on Acoustic Emission During Deformation and Fracture 445 

- M.N. Bassim 

List of Delegates 454

LOOKING INTO THE FUTURE 

R.S. Sh.an.pe. 

Mat-tonal NVT Ce.ntn.0. 

AERE Hawo.ll, England 

ABSTRACT 

- 303 - 

The recent rapid growth in scientific maturity of NDT is resulting in greater 

confidence in finding defects, higher inspection reliability, better 

understanding of the principles and limitations of the techniques 

conventionally used and a more quantitative approach to the assessment of the 

nature and sizes of defects through improved signal and data processing 

routines. 

In addition, the boundaries of nondestructive testing are reaching out to 

encompass new objectives and also provide a wider horizon than just 

inspection after manufacture, for which the majority of the conventionally 

applied techniques and application philosophies were primarily designed. 

The paper will show how the National NDT Centre at Harwell is operated to 

provide a focus of activity in the UK for spearheading this progress towards 

greater technological maturity. Examples taken from the current programmes 

of work will be used to illustrate moves towards a quantitative science, 

moves towards greater sensitivity and higher reliability and accuracy, moves 

towards linking the technology more overtly to a better understanding of 

fracture processes and moves towards better process control which, 

paradoxically, should make NDT, as an inspection process at 'the end of the 

line', edge towards redundancy!

- 304 - 

Nondestructive testing Is a subject that one might say was born of necessity, 

nurtured and kept alive by dedicated and experienced practitioners and has 

developed empirically, often through panic and expediency, to satisfy 

specific needs as they arise in industry. 

In recent years the subject has developed in a much more structured way, by 

encouraging formal training and operator certification, by the setting up of 

societies and conferences both nationally and internationally and, perhaps 

most significantly, by attracting the interest and participation of 

scientific research groups in many countries. 

One great advantage of this input of 'fresh blood' into the subject has been 

that NDT has developed much firmer foundations; the conventional techniques 

are, themselves, much more clearly understood and the basis for forward 

development is rapidly reverting to one of 'fact' and 'scientific reasoning' 

rather than 'hunch' and 'lets have a go'. The disadvantage, which should 

certainly not be underestimated, is that we now have a very dichotomous 

situation with, on the one hand, those practised and experienced in 

conventional practices and in the needs of industry, who have not amassed 

detailed knowledge or understanding of recent scientific developments or of 

research practices associated with signal and data processing; and, on the 

other hand, the 'scientific recruits' to the subject who have virtually no 

feel for the applications side of practical field NDT, scant allegiance to 

the technological society framework and little concern for the empirical 

'footslogging' of those who built the 'string and sealing wax' launch pad 

which they, the scientists, have used to get themselves into orbit. 

So if we look first into the near future one of the major needs is to bridge 

this "gulf before it becomes an unbrldgable chasm. One solution is to build 

up combined teams in which practical experience and scientific erudition can 

be brought together, hopefully to the mutual benefit of both disciplines. We 

have tried to tackle it this way in the UK with the national NDT Centre, 

which was set up at Harwell as long ago as 1967 with this as one of its main 

objectives. Both NDT applications work and research programmes are combined 

in the Centre's work programme and there is the added stimulus that need (as 

measured fairly quantitatively by contract income) can be used as the arbiter 

to decide the balance and range of activity. 

Our experience is that the type of organisation we have set up in the UK can 

be made to cater for both the short-term 'fire—bridgade ' type of NDT problem 

that tends to arise unexpectedly but, seemingly, with amazing regularity in 

manufacturing industry; and for the longer term R&D needs of larger 

organisations (such as the power generation, nuclear construction, oil and 

gas distribution and aerospace industries) who are perhaps more likely to 

anticipate and plan for the NDT needs associated with new manufacturing 

designs or new constructional materials and hence support underlying 

research.

- 305 - 

Another 'near future 1 need is undoubtedly going to be one of technology 

transfer. The rapid increase in research is resulting in a rapid fall-out of 

innovative Ideas and laboratory demonstration systems. The scientist often 

tends to lose interest when the 'physics has been worked out 1 and is then 

keen to transfer his innovative skills to the solution of the next 

challenging problem. This translation from a laboratory demonstration 

'happening' to an engineered, marketable and user-friendly package that can 

be taken over (and championed !) by the NDT practitioner is indeed a very 

significant problem- Too often laboratory solutions relate to a very 

specific problem, and the further development costs to produce an engineered 

system are then difficult to justify in terms of potential marketable units. 

This situation can sometimes lead to 'prototype systems' perpetuating the 

laboratory concepts, and hence requiring a 'PhD driver', getting out into 

practical shop-floor application and quickly leading to disillusionment and 

engendering loss of credibility - for the technique, the research laboratory 

that produced it and ultimately for NDT itself. 

Another 'near-future' need relates to the importance of getting the newer NDT 

techniques and systems through the barriers of standards, codes of practice, 

training and certification syllabi that collectively tend to set the norm of 

currently accepted NDT practices. Changes in this area are generally slow 

and effected by those often unfamiliar with what is happening at the 'sharp 

end' of research. 

So much for the near future, what of the longer term future for NDT and of 

the 'fresh fields' into which NDT practices might be expected to develop. 

Firstly, I detect a significant 'fresh fields' outlook away from simple 

defect detection of manufactured products back into the monitoring of the 

materials processing techniques and into the closer control of the properties 

of the materials themselves from which products are manufactured. 

The overall objective in this new approach is to provide better monitoring of 

the materials and the materials processing treatments so that the need for 

conventional NDT inspection 'at the end of the line' hopefully may become 

both less necessary and less critical. 

Then there is definitely a 'fresh fields' approach in studying and monitoring 

the early stages of crack formation. The objectives here are to develop a 

better understanding of the mechanisms of the process of crack development 

and to monitor the factors that contribute to failure occurring. Work on 

monitoring crack—tip stress and fracture toughness are very much in their 

infancy but could soon become a major area of study. 

Coming back to the more conventional defect detection role of NDT, there is a 

'fresh fields' approach now building up in which maximum use will be made of 

all test data so as to obtain more informative detail about the defect being

- 306 - 

interrogated, from which an interpretation of 'significance' then becomes 

easier to make, in support of fracture mechanics predictions. 

A whole range of advanced data assimilation and signal processing systems 

have recently been developed. We have recently developed and evaluated one 

such system which rapidly digitises and stores every A-scan and allows recall 

at will in a variety of presentational modes. 

We are also helping to contribute to a new generation of non-amplitude 

dependent ultrasonic sizing techniques with what we call the time-of-flight 

diffraction (TOFD) technique. We and others are also working on the more 

difficult 'classification' problem of identifying and characterising defects 

from the fine-scale information contained in the defect signal. 

Another fresh-fields approach in NDT that I see developing is to move away 

from 'eye-ball' defects in order to detect significant imperfections that are 

of much smaller dimension. We are contributing to this in the further 

development and evaluation of high definition radiography and fluoroscopy and 

in studying the problems associated with inspecting engineering ceramics, 

where the critical defects for initiating brittle failure are measured in 

microns rather that millimetres. 

Yet another trend, that should appeal to many, is aimed at hastening the 

obsolescence of NDT on the product line by modelling the effects of variables 

in the manufacturing process so as to try to reduce the incidence of 

defective products, by providing greater control of the process itself. As 

an example of this approach we are at the present time theoretically 

modelling the paint spray process in order to understand the effects of 

process variables on the final product, so that it should hopefully be 

possible to produce a more consistent 'quality', in this case in a robotised 

paint spraying process. 

Perhaps a final fresh fields approach as far as NDT is concerned should be 

that of being less introverted; by spending more time looking over one's 

shoulder to see which way other technologies are developing and making 

maximum use of such parallel experience wherever it is relevant to the 

furtherance of NDT practices. 

Bearing in mind that NDT only a few years ago was a 'bottom of the heap' 

technology kept alive only by a dedicated band of enthusiasts, it is 

particularly encouraging that it has recently attracted so much academic and 

management interest and as a consequence is opening up its frontiers in so 

many new and challenging directions. Looking into the future of NDT is, in 

fact, quite a stimulating and heartening experience at the present time and 

the possiblities certainly open up a whole new horizon of opportunity whilst, 

at the same time, exposing many problems and pitfalls associated with communication 

(up, down and sideways) that will need to be tackled and overcome.

- 307 - 

NDE OF STRUCTURAL CERAMICS BY HIGH FREQUENCY ULTRASONICS 

A. fakn. 

National Re.izan.ck. Council o I Canada 

ABSTRACT 

A high frequency ultrasonic system with computer controlled data acquisition and 

analysis has been described. The system has been used to detect surface flaws in 

the size range of 20-300 um in hot pressed silicon nitride. The detection 

technique is based on Rayleigh waves which are generated as a result of the mode 

conversion of compressional waves incident from water onto the surface of the 

test material at near the Rayleigh angle. The scattering of the Rayleigh waves 

by surface cracks has been detected, analysed in the time and frequency domains, 

and correlated with the size of the cracks. 

This work was supported by the Defence Research Establishment, Pacific, 

Victoria, B.C. and was carried out at the Ontario Research Foundation, 

Mississauga, Ontario.


- 308 - 

Ceramics such as silicon nitride, silicon carbide, and zirconia have excellent 

corrosion resistance, low thermal conductivity, high wear resistance, and good 

strength at elevated temperatures. These properties have generated considerable 

interest in the use of these ceramics in energy conversion systems, e.g., diesel 

engines, turbines, and heat exchangers. 

The hard and brittle nature of ceramics, however, inhibits the release of 

stresses at flaws under load by plastic deformation. This implies that a very 

small flaw can cause a catastrophic failure in ceramics. Therefore, flaws in the 

size range of 20-100 vim are considered as "critical" in ceramics for such 

applications as turbine blades. 

The reliable use of ceramics as structural parts requires nondestructive 

evaluation techniques which are capable of detecting and characterizing such 

small defects. Microfocus x-radiography and high frequency ultrasonics have been 

found^ ' to be the most suitable NDE methods for ceramics. 

This paper describes a high frequency ultrasonic system with computer controlled 

data acquisition and analysis capability and its application in the detection and 

characterization of microscopic surface cracks in silicon nitride. 

2. EXPERIMENTAL STUDIES 

2.1 Material and standard cracks 

Hot pressed silicon nitride is one of the most promising materials for high 

temperature structural applications. The majority of failures in this material 

originate from surface cracks introduced during grinding and machining operations. 

These produce arrays of semi-elliptical cracks with random inclination to the 

surface, but a preferred alignment parallel to the direction of motion of the 

abrading particles ,. A Knoop indentation technique was used to simulate 

surface machining cracks. In silicon nitride the Knoop indentation produces 

sharp semi-circular cracks with a mouth opening of the indentation less than one 

fifth of its depth, a, and its surface length about 2a. 

Knoop indentations in the depth range of 10 to 300 ym were introduced on a 

polished surface of a hot pressed silicon nitride (NCI32) plate. These cracks 

were carefully measured under an optical microscope prior to ultrasonic 

examination. 

2.2 High frequency ultrasonic system 

Since the critical flaw size in ceramics is very small and the high stiffness to 

density ratio generally gives comparatively high values of acoustic velocity and 

hence long wavelength at conventional ultrasonic frequencies, it is necessary to 

use high frequency ultrasonic waves.

- 309 - 

The system described in this paper (see Fig. 1) can be operated at frequencies up 

to 100 MHz. The puiser of the system consists of a 1 Hz to 100 MHz pulse 

generator (Philips PM5771) and a broadband (0.15 to 300 MHz) power amplifier 

(EIN403LA). The pulse generator is capable of providing pulses as narrow as 5 ns 

which are required for 100 MHz transducers. The high frequency transducers* 

employ a quartz buffer rod with a piezoelectric lithium niobate crystal mounted 

on one end and a lens machined on the opposite end. Focused transducers (12 mm 

focal length in water) with a narrow beam diameter (500 ym at focal point) were 

used. 

The detecting system uses a broadband (0.1-1300 MHz) amplifier with 48 dB gain 

(HP8447F) and a digital scope (HP1980). The latter has a bandwidth of 5 Hz to 

100 MHz and consists of two 100 MHz analogue measurement channels and an analogue 

to digital converter. When used together with a digital waveform storage unit 

(HP19860), repetitive ultrasonic signals of up to 100 MHz can be averaged (up to 

64 times), digitized (up to 501 points per waveform in main or delay modes), and 

stored in the memory. The stored waveforms are transferred to a microcomputer 

(HP9845) for analysis in the time and/or frequency domains. 

2.3 Detection Technique 

Surface or Rayleigh waves were used to detect surface cracks in silicon nitride. 

These waves are generated by the mode conversion of compressional waves incident 

at the Rayleigh angle, 9R, onto the surface of the test material. The 

conventional method of generating surface waves is to use a wedge transducer with 

a wedge angle of 9 g in contact with the test specimen. Preliminary work*- ' ' 

however, indicated that the contact method is not suitable for testing ceramics 

at high frequencies due to coupling problems. Consequently, the technique of 

non-contact mode conversion in an immersion tank was used in the present work. 

Figure 2 shows the basic configuration for the generation and detection of the 

Rayleigh waves. In this configuration the Rayleigh angle is given by: 

6R = arc sin -y— (1 ) 

R 

where V is the acoustic velocity in water and VR is the Rayleigh wave velocity in 

the test material. For silicon nitride 9R=15.7 . The Rayleigh waves reflected 

by sharp surface discontinuities radiate energy back into water at an angle of 

9Rand this energy is received by the transducer. These are referred to as "leaky 

Rayleigh waves". Thus, as the transducer is scanned parallel to the surface of 

the test material, strong signals from surface flaws are detected. Figure 3 

shows a typical signal from a 100 pm Knoop indentation detected by leaky Rayleigh 

waves at a frequency of 50 MHz. 

The detectability of surface flaws by leaky Raleigh waves is dependent on the 

transducer frequency. This is demonstrated in the experimental results of Fig. 4 

which show the smallest Knoop indentation detectable at different frequencies. 

This figure indicates that for the experimental conditions of the present work, a 

* Precision Acoustic Devices, Inc.

- 310 - 

frequency of about 50 MHz is optimum for the detection of surface cracks as small 

as 20 um in silicon nitride. This frequency is equivalent to an acoustic 

wavelength of 112 um in silicon nitride. Beyond 50 MHz, the detectability tends 

to level off due to the increase of attenuation losses. Thus, the crack 

measurements in this investigation were carried out at 50 MHz. 

The leaky Rayleigh waves propogating along the water-solid interface, lose their 

energy rapidly and particularily at high frequencies. Thus, only those flaws 

which are located near the point of incidence of the impinging beam are detected. 

As a result, the resolution of this method is very good since reflections from 

neighbouring flaws or specimen edges do not overlap with the signal from the flaw 

of interest. The resolution is particularily good when a narrow Rayleigh beam is 

employed by focusing the incident beam on the surface of the test piece using a 

focused transducer. 

Overall, the scanning simplicity, the high resolution and high sensitivity of 

leaky Rayleigh waves make this technique a practical and reliable method of 

detecting small surface flaws in ceramic materials. 

2.4 Measurement techniques 

The normal procedure for flaw size measurement by any ultrasonic technique 

utilizes the acoustic reflection coefficient, S, of the flaw for size estimation. 

S is defined as the ratio of the peak amplitudes of the flaw signal to input 

signal at the transducer terminal. In the case of cracks, however, the acoustic 

reflection is very much dependent on orientation of the crack with respect to the 

incident beam. This is demonstrated in the experimental results of Fig. 5 

obtained for a 100 ym Knoop indentation. Analytical solutions are given by 

Auld^ ' for calculation of the reflection coefficient at oblique angles. Both 

the analytical and experimental results indicate that as the direction of the 

crack relative to the Rayleigh beam deviates from normal, a smaller portion of 

the crack face reflections are received by the transducer. Thus, for complete 

characterization of cracks from the reflection coefficient, it is necessary to 

scan the test material in more than one direction. 

Another important factor affecting the acoustic signal amplitude from a flaw is 

the wavelength compared to the flaw size. Since at 50 MHz the Rayleigh 

wavelength in silicon nitride is about 112 pm and this is in the flaw size range 

of our interest (i.e. 10-200 um), therefore both the long wavelength (X^>a) and 

the short wavelength (A^

and 

- 311 - 

S = a 2j for XR>a (2) 

S = a'.a for XR

References 

- 312 - 

1. W.N. Reynolds, and R.L. Smith, British Journal of NDT, 145, 1982. 

2. J.J. Mecholsky, S.W. Freiman, and R.W. Rice, J. Amer. Ceram. Soc, 6^, 116, 

1977. 

3. S. Johar, A. Fahr, and M.K. Murthy, Proc. Conf. Advanced NDE Tech. sponsored 

by IMRI, NRC, 1982, Montreal. 

4. A. Fahr, S. Johar, M.K. Murthy, and W.R. Sturroek, Proc. Review Prog. 

Quantitation NDE, Vol. 3A, Plenum Press, 1984. 

5. B.A. Auld, Wavemotion, J_, 3, 1979. 

6. G.S. Kino, J. Appl. Physics, h% (6), 1390, 1978. 

7. B.T. Khuri-Yakub, G.S. Kino, and A.G. Evans, J. Amer. Ceram. Soc, 63, 

(1-2), 65, 1980. 

8. S. Ayter, and B.A. Auld, Proc. DARPA/AFWAL Rev. Prog. Quant. NDE, AFWAL-TR- 

80-4078, 344, 1974. 

9. V. Domarkas, B.T. Khuri-Yakub, and G.S. Kino, Appl. Phys. Lett., 3_3, (7), 

557, 1978. 

10. N. Saffari, and L.J. Bond, IEEE Ultrasonics Symposium, 1983.

TRANSCEIVER 

digital 

signal 

display 

1 

A to D 

converter 

waveform 

storage 

- 313 - 

focused 

transducer 

Specimen 

A A 

Plotter 

7- 

3 

Mag netiC 

tap 

Pf_J 

Fig. 1. A schematic of the set-up. 

LEAKY RAYLEIGH 

WAVES 

Fig. 2. The basic configuration for 

the generation and detection of leaky 

Rayleigh waves. 

5,us 

„'User's 

"i Graphics 

terminât 

CRACK 

REFLECTION 

EDGE 

REFLECTION 

Fig. 3. An oscilloscope view of a 

signal from a Knoop indentation (100 pm) 

detected by 50 MHz leaky Rayleigh waves. 

The indentation was located 2 mm from 

the specimen edge.

PLANE OF 

CRACK 

100 

20 40 60 60 

SAW Frequency (MHz) 

1OO 

Fig. 4. The detectability of surface cracks in silicon 

nitride by leaky Rayleigh waves as the function of frequency. 

EXPERIMENTAL 

(KNOOPCRACK) 

40 

60 

80 

I00 

I40 

THEORETICAL • 

20 

160 ISO' ANGLE OF 

INCIDENCE 

CNORMAL 

/ ÄCOUSTI) 

REFLATION 

.COEFFICIENT 

Fig. 5. The experimental data obtained for the acoustic reflection of a Knoop 

indentation at different orientations compared to the theoretical curve obtained

m 

"O -30 

z 

o 

Ul -40 

u. 

Ul 

ce 

o 

o 

Ü 

-50 

-60 

- 315 - 

CRACK DEPTH (Aim) 

50MHz Ä * 112 um 

100 200 300 

10 20 30 

Ko 

Fig. 6. The acoustic reflection of various size 

Knoop indentations obtained by 50 MHz leaky Rayleigh waves. 

O 

0. 

< 

Ul 

Ul 

50 100 

FREQUENCY 

CRACK SIZE 

(/im) 

60 ^ 

150 

Fig. 7. Spectra of different size Knoop 

indentations deconvolved by a reference signal.

- 316 - 

HOT PRESSED PIEZOELECTRIC CERAMIC ELEMENTS FOR ULTRASONIC 

TRANSDUCERS 

U.V. Patdt, J. van den Ande.1 and P.S. Nicholson 

HamiCtcr., Ontanic, Canada 

ABSTRACT 

Piezoelectric ceramics (PZT4, PZT5, PZT7 and SPN) were investigated for 

development of high frequency (30-100 MHz) normal and focussed transducers. Hot 

pressed piezoelectric materials with a controlled grain size of ~2jum and 

approaching theoretical density can be lapped down to 30 jum thickness without 

structural damage whereas conventional sintered piezoelectric material with grain 

sizes ranging frcm 2-15jum and at least 3% porosity can only be lapped down to 

100>um. Hot pressed piezoelectric ceramics when compared with conventional 

piezoelectric ceramics, show higher values of coupling coefficient (10-20%), 

elastic compliance ^12% and moderate increases in mechanical quality factor 

(Q ) and dielectric constant (€ ). The ccmpressional sound velocity measured 

along the poling direction is very sensitive to the switching of dipoles other 

than 180 ones. The velocity increases as the polarization increases and this 

constitutes a useful method for quality control of the piezoelectric ceramic as 

it gives the correct thickness resonant frequency. By accurately measuring the 

change by a comparison method, the dipole behaviour can be better understood and 

the degree of polarization and depolarization can be established. Temperature 

dependent polarization and depolarization is also discussed in the light of 

dipole switching. 

Hot pressed SPN and PZT5 are sufficiently transparent to be considered for 

use in opto-ultrasonic applications such as in medicine. 

Because of the high value of the thickness-frequency constant and low O , 

SPN seems a better choice as a high frequency transducer material. The signal 

spectra of a high frequency transducer constructed frcm these materials is 

presented. 

I INTRODUCTION 

High frequency transducers in the range of 30 to 100 MHz are of great 

importance for detection of minute flaws (~-'5/*m) in high performance ceramics. 

The detection of these flaws require powerful high-frequency transducers which 

are now being developed. As the frequency increases, the absorption in the 

medium being investigated also increases. Therefore, to compensate this loss, 

the transducer piezoelectric element should have a high transmitting coefficient 

d,. and receiving constant g„. Also, the sound beam should be capable of

- 317 - 

deep penetration in same focussed fashion. 

Most ccmmercially available piezoelectric discs are sintered. These 

sintered materials have grain sizes between 2 and 10 /um with a few percent 

porosity and occasionally have large flaws. In order to achieve high frequencies 

in the range 30-100 MHz, these elements have to be lapped down to as thin as 20 

jm without any structural danage, however, sintered piezoelectric ceramics, when 

lapped below 100 /un thickness, receive structural damage, eg. grain pull-out and 

have insufficient strength. Therefore, piezoelectric ceramics used for high 

frequency transducers should have at least 7-12 grain layers devoid of porosity. 

Large flaws also contribute to the electrical failure of piezoelectric ceramics 

on polarisation. 

It is veil known that grain size and bulk density markedly influence the 

electrical properties of sintered piezoelectric ceramics. By choosing another 

fabrication method, i.e., hot pressing, one can control the grain growth and 

densification process. This method can yield 100% dense ceramics with 

controlled grain sizes suitable for the extreme thinning required. The ccmpressional 

wave sound velocity in the direction of polarisation for the hot pressed 

material is higher than their sintered counterparts. This gives an extra layer 

of thickness when compared with the sintered material for the same thickness 

resonant frequency. The piezoelectric properties of hot pressed elements are 

also superior to sintered elements. 

Another important factor governing these ceramics is polarisation. One has 

to optimise the poling condition for each composition when such ceramics are to 

be lapped down to «OO/fm after polarisation. If one exceeds these cptimim 

poling conditions (Ref.l) (e.g. temperature, poling field or poling time), 

micrccracks will develop causing element failure when excited by the voltages 

necessary to generate sufficiently powerful bursts of high frequency ultrasound. 

Whereas the switching of normal dipoles (180 dipoles) does not involve 

any structural changes, the alignment of "mechanical dipoles" (other than 180°) 

during polarisation requires such changes. Since sound waves are associated with 

interatomic forces, one can detect these mechanical dipoles by measuring sound 

velocities in the material. Such measurements facilitate an understanding of 

dipole behaviour during polarisation and depolarisation. It is the mechanical 

dipole alignment associated with the thickness resonant frequency of the 

piezoelectric ceramic which causes the change in sound velocity. Therefore, the 

thickness resonant frequency will depend upon the number of mechanical dipoles 

aligned during polarisation. 

II SAMPLE PREPARATION 

PZT 4, PZT 5 and PZT 7 calcined powders and SPN were pressed at 20 MPa 

(3000 psi) into slugs ^15 nm long, isopressed at 340 MPa (50,000 psi) and fired 

in a closed alunina crucible at 1270 -1350 c in a controlled atmosphere for 

an hour. Heating and cooling rates of 450 c/h were maintained. 

Samples of the same powders were pre-pressed to obtain slugs and these were 

placed in a molybdenum mold and packed with zirconia/alunina (200 mesh) powder. 

This composite was pressed in a vacuum induction furnace for ~10-20 minutes at 

900-1200 C at pressures between 3 and 30 MPa (500-5000 psi) employing the same

- 318 - 

heating and cooling rates of 450 C/h. Bulk densities were measured in butanol 

and distilled water. 

Both hot presses and sintered ceramics were sliced into « 6 mm thick discs. 

Tfi^ hot-gressed slices were oxidized in a controlled atmosphere for 4-24 hours at 

350-1200 C, depending on composition. Both sides of these discs were lapped to 

a thickness of 400/urn with optimum flatness and parallelism. Fired silver electrodes 

and sputtered gold electrodes were deposited on them. The fired silver 

electrodes were polished to M 2/tm on both sides of the piezoelectric discs. 

III ELECTRICAL AND ACOUSTIC CHARACTERIZATION 

All materials were poled in oil at 80-180°C. The poling temperature, time 

and voltage were varied to identify the optimum poling conditions for each 

cemposition. All poled specimens were aged for 24 hours, prior to measurement. 

The resonant (f ) and antiresonant frequencies (f ) were measured according 

to IEEE standards. Following measurement of the ccmpressional and shear 

wave velocities, the Poisson's ratio for each cemposition was obtained by calculation. 

The piezoelectric constants, the planar coupling factor, k ; the piezoelectric 

strain constant, d.,, and the piezoelectric 'voltage' constan? g-., ware thus 

obtained and are tabulated in Table I. Capacitance and loss factors were 

measured at 1 kHz using an impedance bridge. 

The ccmpressional sound wave velocities at 50 MHz ware measured in both 

polarized and unpolarized piezoelectric discs. In the polarized materials, 

velocity was measured in the polarized direction. The pulse-echo overlap nethod 

was employed and velocities were measured to one part in 10 . A low-viscosity 

oil was used as a couplant between the 50 MHz transducer and the piezoelectric 

discs. The thickness variation of the discs was monitored before and after 

polarization and also during the velocity measurements. 

The microstructural grain size was determined from photomicrographs (optical 

and electron optical) of the test specimens. These were mechanically'polished to 

a surface roughness of 0.3A

- 319 - 

the density by »w3 to 7%. 

These increases are due to the increases in density, the lack of texturing 

and the uniform grain size of the hot pressed materials. Such materials are 

clearly advantageous for use in high frequency transducers. The higher sound 

wave velocities allow extra thickness for the same frequency, e.g. for 100 MHz 

longitudinal sound wave velocity, the required thickness for hot pressed PZT 7 is 

•* 25 /«m whereas for the sintered material it is 23 yum. The hot pressed 

piezoelectric ceramics drew less current during poling than the sintered ones 

(Figure 1). For all materials the anount of current drawn increases with the 

applied voltage. Current fluctuation or considerable increases in 

voltage during poling, indicate the development of microcracks in the ceramic. 

This causes a decrease in the coupling coefficient (9) as indicated in Figure 2 

beyond 3,600 V/mm. If either time, temperature or field are increased, the 

specimens fail. The dipole alignment on application of an electric field, leads 

to the development of internal stresses which eventually leads to microcracking. 

In the absence of microcracking, the dipole alignment increases k ; with microcracking, 

k drops. p 

In assembling high frequency transducers, either indium bonding or a 

conducting epoxy is used to bond the piezoelectric discs to the backing (damper) 

medium. For both cases, a temperature of 80-150°C is required to construct the 

transducer. Therefore, the elements should be able to withstand these temperatures 

without losing their piezoelectric properties. Figure 3 illustrates the effects 

of overheating PZT 5 discs. The curves show the value of k as a function 

of increasing temperature up to 175 C on the same test specBnen. It is clear 

that k is reduced by «'Si when the piezoelectric element is heated to ~200 o c 

No sucn property change was observed when specimens were heated to 140°C; 

k peaks at «/HO C and this can be attributed to the transition frcm one 

ferroelectric phase to another. 

To construct a high frequency transducer it is necessary to know the correct 

thickness resonant frequency of the element and this is difficult to measure on 

thin discs. By measuring the sound velocity in the polarization direction, 

a high degree of accuracy can be realized. Fran such data, the thickness 

resonant frequency; (f = c/2t, c = velocity, f = frequency and t = thickness) 

can be obtained and dipole behaviour during poling, elucidated. 

During poling the 180 -demain reversal is complete and other discrete 

angle dipoles are incompletely switched (the percentage of such switching depends 

upon the phases involved and the impurities added to the ceramic). Stress free 

piezoelectric ceramics would be those whose domains switch only by 180°, i.e., 

experience no structural changes. Therefore, the sound velocity measured along 

the polarization direction for such a material, should be unchanged when compared 

with the virgin ceramic. This was true for the initial polarization of PZT 4, 

PZT 5 and PZT 7 in a low field, i.e., where only the 180 dipoles are involved 

(Figures 4,5,6 and 7). The k values obtained frcm alignment of these normal 

dipoles for PZT 4 (rhcmbohedral/tetragonal phase), sintered and hot pressed are 

~23% and 29%, respectively; for PZT 5 (rhanbohedral), sintered and hot 

pressed are 27% and 38%; for PZT 7 (tetragonal phase), sintered and hot 

pressed are 18% and 23%, and for SPN, sintered and hot pressed are '"19% and 

«/26%, respectively.

- 320 - 

The alignment of the mechanical dipoles involves structural changes and 

this causes internal stress; this will change the sound velocity in the polarization 

direction. The velocity change is directly proportional to the extent of 

mechanical dipole alignment during polarization. For hot pressed, PZT 4 the 

change was ~9%, for PZT 5 «'10%, for PZT 7 "7% and for SPN ~8%. Their sintered 

counterparts showed a few percent less sound velocity. The initial temperaturedependent 

depolarization occurs via the mechanical dipoles as stresses are 

relieved. For the PZT compositions and the SPN, the mechanical dipoles alone 

switch back up to 250°C and 325°C respectively. As temperatures are increased 

beyond 250 C for PZT canpositions and 325°C for SPN, the normal 

dipoles also switch back. The time and frequency spectra of hot pressed elements 

and transducers constructed fran them are shown in Figures 8 and 9. 

SPN, with a high sound velocity, low dielectric constant and low O is 

potentially useful for the construction of high frequency transducers. Also the 

acoustic impedance for SPN (31.3 x 10 _ g/an sec) is low compared with that of 

the PZT canpositions ( ~39 x 10 g/an sec). This property facilitates 

matching with transducer backing materials. 

Hot pressed PZT 5 and SPN can be sufficiently transparent for consideration 

in opto-ultrasonic applications. 

V CONCIDSIONS 

As a result of the reported investigation, the following conclusions can be 

made; 

(1) hot pressed piezoelectric ceramics ( /v 100% dense) can be used to construct 

high frequency transducers in the range 50-100 MHz. 

(2) the piezoelectric properties of the hot pressed materials are considerably 

better than their sintered counterparts. 

(3) the high values of sound velocity along the polarisation direction (i.e., 

the high frequency constant) for hot pressed material allows an extra layer 

of grain thickness for the same frequency as compared to similar sintered 

materials. 

(4) the grain size of the final ceramic can be controlled by hot pressing 

techniques. 

(5) the ccmpressional wave velocity measurements in the polarisation direction 

facilitate an understanding of the behaviour of dipole switching during 

polarisation and depolarisation. 

REFERENCES 

(1) R.W. Rice and P.C. Pohanka, J. An. Ceram. Soc., 56 (8) 420-23 (1973). 


SPN calcined powder was supplied by BM Hi Tech (Collingwood, Ontario).

density 10 3 kg/m 3 

dissipation factor % 

relative dielectric 

kp % 

k 31 % 

d31 

pC/N 

g31 10- 3 Vm/N 

Vp m/s 

Vu m/s 

om 

required thickness In 

Vim for 100 MHz 

TABLE 1: Physical Properties of Four Piezo-Electric Ceramics 

sintered 

7-05 

0-4 

1100 

54 

32 

.115 

-11-8 

4420 

4100 

4SO 

22 

PZT 4 

hot 

pressed 

7*09 

0-3 

1250 

02 

37 

.138 

-12-3 

4820 

4420 

490 

24 

sintered 

7-73 

1-8 

15SO 

«3 

38 

-170 

-12-4 

4510 

4180 

70 

23 

PZT 5 

hot 

pressed 

8-01 

1-8 

2100 

72 

43 

• 230 

-12-4 

4850 

4420 

80 

24 

sintered 

7-45. 

1-5 

400 

32 

19 

-30 

-8*5 

4820 

4400 

1025 

23 

PZT 7 

hot 

pressed 

7 »97 

0.6 

450 

38 

23 

-40 

-10-0 

4940 

4640 

1180 

25 

sintered 

4-20 

2.2 

260 

37 

22 

-31 

-12.5 

6520 

8170 

140 

28 

SPN 

hot 

pressed 

4>49 

1-e 

380 

45.. 

27 

-46 

-13-1 

6980 

6430 

210 

35

o PZT 7 

o PZT 4 

• HP 7 

• HP4 

- 322 - 

1-5 3-0 

Poling field (kV/mm) 

development of 

microcracks, 

piezo-eiectric 

properties 

drop 

Fig. 1: Dissipation Current as a Function of Poling Field at 120°C 

for sintered and Hot-Pressed PZT 4 and PZT 7 Ceramics.

(0 

60 

g40 

«MM 

"5. 

3 

O ü 

Cö 

§20 

A-HP4 

A-PZT4 

- 323 - 

120 °C 

DEVELOPMENT 

OF 

MICROCRACKS 

O-PZT7 

• -HP7 

i 

00 1500 3000 

Poling field (V/mm) 

i 

4500 

Fig. 2: The Dependence of Planar Coupling Factor kp on Poling 

Field at 120°C for Sintered and Hot-Pressed PZT 4 and PZT 7 

Ceramics.

70 

-65 

b. 

o 

•4-» 

O CO 

= 60 

a 

3 O 

£55 

PZT5 

- 324 - 

I I 

50 150 

Temperature (°C) 

•DECREASE IN PIEZO-ELECTRIC 

PROPERTIES 

Fig. 3: Loss of Planar Coupling Factor by Repeated Overheating of 

a Piezo-Electric Element. 

250

60 

PZT7 

Vp " 

- 325 - 

o A —polarization 

• * —depolarization 

/Vu 

8 10 

Fig. 4: The Planar Coupling Coefficient as a Function of Change 

in the Compressional Wave Velocity, with Depolarization 

caused by Heating (PZT 7.).

o 

CO 

O) 

60 

- 326 - 

2 4 6 

V P » v u /Vu 

A o—polarization 

*• —depolarization 

Fig. 5: As Fig. 4, but for PZT 4. 

8 10

- 327 -

- 328 - 

2 4 6 

o —polarization 

• —depolarization 

vP - vu /Vu ) 

Fig. 7: As Fig. 4, but for SPN. 

8 10

- 329 - 

Fig. 8: Typical Outputs of Three Hot-Pressed, Unmounted SPN Elements 

of Different Thickness and Determination of Their Time Period.

- 330 - 

60 MHz 

Fig. 9: Output of Two Transducers Made with Hot-Pressed PZT 4 Elements: 

1. Highly Damped 

2. Lightly Damped 

3. Frequency Spectrum of 2

- 331 - 

ULTRASONIC ANALYSIS OF VOIDS IN GLASS; THEORY AND PRACTICE 

Aitkux J. Stockman, Jan van den Ande.1 and Patrick S. Hlcholion 

McMaiie-t linivdtisitij 

Hamilton, Ontaiio, Canada 

ABSTRACT 

The characterization of voids in glass enables the study of reflected ultrasonic signals off . known but 

undisturbed defects which can also be characterized by optical means. The reflected ultrasonic 

signals can be compared with theoretical signals transformed to a frequency spectrum by a computer 

and a frequency spectrum can be transformed to a time signal (A-scanl. 

Flaw information can now be read from such signals, and such will eventually help to characterize 

flaws in opaque ceramics. 

I INTRODUCTION 

The characterization of defects in high performance ceramics is important when the components are 

to be used in situations of high stress. Size, shape and composition of the target defects will determine 

whether or not the ceramic will fail. Interest in ultrasonic testing to determine these factors has led 

to many advances in the field of defect characterization by scattered signal analysis.' 1 ' 

On the link between theory and practice, it was found by the authors that a beam entry of exactly 90° 

with the entrance surface of the component was important.' 2 ' The essence ofthat article was that the 

frequency components of the ultrasonic beam pass the interface undistorted if a perfect 90° angle-is 

maintained. A slight misalignment will cause enough distortion to make the frequency spectrum of 

the reflected beam lop-sided and random misalignments from sample to sample means that the 

analysis will no longer yield accurate results. 

The importance of proper alignment is brought out even more in a study of reflections of previouslyfound 

defects. 13 ' Each defect responds differently to a frequency component so their reflected signals 

contain information about their type and shape. The change in frequency component amplitude can 

also be calculated, measured and the resulting spectrum compared to the perfectly aligned spectrum, 

i.e. the only distortion is caused by the defect rather than by off-normal set-up conditions 

The technique was applied to ceramics at frequencies above 10 MHz in order to test for flaws 

< 100 urn. A precision immersion scanner with a resolution of 20 urn was built especially for locating 

these small defects. Beams were focussed to obtain higher power densities and even though these 

beams do not provide a true plane wave, the wave theory as reported by Ying and Truell' 4 ' was 

applied along with the "high frequency decomposition" part, reported by Tittman et al.' 5 '

- 332 - 

As the scattering theory requires spherical defects, this work examines the scattering of a focussed 

sound beam from spherical and near spherical bubbles in glass. The use of glass rather than opaque 

ceramics facilitates optical characterization, the only way a small defect can ever be diagnosed. 

II. THEORY 

The pressure profile for the sound beam from a focussed transducer operating at a single frequency is 

a complicated function of the density p and speed Î sound cL of the medium, the diameter D and 

radius of curvature A of the transducer, and the frequency f. At the focal distance Zf from the 

transducer the pressure is: 

where 

i2nflt - (ß + z.)] 

2c, ' 

c, S fe 

L zf 

8nS 

S = » 

V 1 - A/Zf 

(Zf - h)" + (D/2)" 

sin {nf (p 1 - Zf )} 

1/2 

, h = A - A" - (D/2)- 

and So is the displacement of the front face of the transducer. N'ear the focus the pressure function 

can be approximated in the form of a plane wave as: 

i2nf(t - AZ/Cjj 

where AZ is the distance along the z-axis from the focal point. L'sing the theory of Ying and Truell 

(4), the relative pressure amplitude for the longitudinal wave backscattered to the transducer from a 

spherical void of radius, a, is: 

p (f,a) -2n-e 

f L 

(2m 

m = 0 

where the scattering coefficients Am are determined from the stress and strain relationships at the 

surface of the void. Now if the sound pressure at the site of the void were known it would be possible 

to determine the pressure of a backscattered monochromatic longitudinal wave. 

Reflections from the sample surfaces make single frequency studies impractical, fnstead, short 

pulses of ultrasound are generated by shock excitation of the transducer so that the defect can be 

located. A pulse of ultrasound is composed of many waves of different frequencies The distribution of 

wave amplitudes and phases as a function of frequency is called the response function of the 

transmitter-receiver system, R(f). Since the sound pressure is maximum at the focus the waves will 

be in phase so that only the amplitude component of the response function needs to be considered 

Experimentally the function is determined by routing to a spectrum analyzer the signal reflected 

from a flat surface in the focal plane of the transducer. For the calculations the system response 

1/2

- 333 - 

function is approximated by a Gaussian distribution then it is multiplied by the relative pressure 

amplitude function to produce the expected frequency spectrum of a signal reflected from a void. 

The expected frequency spectra so calculated have both amplitude and phase components making it 

possible to obtain the real time signals and the magnitude frequency spectra expected. However, only 

relative amplitudes of the time signals and magnitude spectra can be determined since the exact 

pressure at the focus is not known. Shape comparisons between measured and calculated signals and 

spectra are possible and are shown below. 

III. EXPERIMENTS 

Figure 1 shows the experimental set-up used. The commercial focussed immersion transducer has a 

piezo-electric ceramic element and is labelled as 25 MHz resonant frequency. It has a 6.2 mm element 

and a focal length of 25 mm in water making it a F/4 class. Signals reflected from the sample are 

picked-up by the same transducer and routed to the amplifier and gating electronics; they are then 

passed to an oscilloscope and a spectrum analyzer. First, the samples were scanned ultrasonically to 

pinpoint the defects. Then the transducer was positioned and refocussed to yield the maximum signal 

from the void on the oscilloscope and the trace was recorded photographically. A frequency spectrum 

was made of the returned signals by means of a spectrum analyzer and this was also photographed; 

however, noise from the gating electronics obscured the spectrum of signals from small bubbles 

pointing to the need of improved gating circuitry, to improve the signal-to-noise ratio. Oscilloscope 

trace photographs for select voids were enlarged and digitized on a graphics tablet so that they could 

be Fourier transformed by computer to obtain frequency spectra without interfering noise. From 

these spectra, frequency measurements could be made accurately. Figure 2 shows the typical signal 

reflected from the flat surface of the samples and its corresponding frequency spectrum. The shape of 

this spectrum was theoretically approximated by a Gaussian distribution with a centre frequency of 

30 MHz and a full width at half maximum of 16 MHz (Figure 3) for use in the calculations. 

IV. RESULTS 

Figure 4 a-d is representative of the voids examined. Shown in (a) is a photograph of the void giving 

the major pnd minor cross-sectional dimensions; in (b) the signal trace of the void from the 

oscilloscope; in (c) a photograph of the spectrum analyzer trace, and in (d) a plot of the frequency 

spectrum as calculated from the Fourier transform of the digitization of the oscilloscope trace. (Note 

the absence of noise on this trace). The shape of the signals and spectra, changes with void 

dimensions. In order to quantify these changes, the frequency at maximum amplitude and frequency 

full width at half maximum (FWHM) were measured. 

Figures 5 and 6, a and b are measured and calculated frequency spectra for void diameters of 49 urn 

and 132 urn. The calculated spectra have been determined for a shock pulse and glass matrix 

parameters ct = 5660 ms 1 and C|/cr = 1.67, which are average values for crown glass. Rather than 

a one to one comparison of measured to calculated frequency spectra, repeated calculations have been 

performed to generate plots of frequency at maximum amplitude versus void diameter (Figure 7a) 

and frequency FWHM versus void diameter (Figure 7bl

- 334 - 

It is also possible to do the reverse and transform the complex frequency spectra into time spectra; 

Figure 8 is an example o[ this for a 90 pm void. Transformations for void sizes from 28 urn to 132 urn 

diameter show that the fargest deflection from zero was negative-going. In Figures 9a-d, peak 

deflections versus time have been plotted with the largest negative-going deflection normalized to - 

100%. From this figure, it is evident that as the void diameter increases, the largest positive-going 

deflection changes from the one following the largest negative-going deflection to the positive-going 

deflection preceding it. This behaviour has been noted previously in the experimental results. 

Plotting ratios of these two positive going peaks against the void dimensions (fig. 10) shows that there 

exists a reasonable correlation with the minor dimension of the void. Although a similar relation 

exists with the major void dimensions it is less apparent, especially for the larger dimensions which 

approach or exceed the wavelength, \c, of the main frequency component.. 

V. CONCLUSIONS 

Calculations have shown that trends apparent in-signals reflected from voids in the focal zone of 

pulsed focused transducer can be reproduced by a model based on scattering of monochromatic plane 

waves convoluted with the frequency distribution of the incoming wavepacket. Although direct 

signal amplitude comparison of theory and experiment is not possible due to lack of information about 

the focal zone pressure, comparison of normalized signals show strong correspondences in signal 

shape. The calculated changes of the frequency distribution as a function of void diameter as 

measured by the frequency at maximum amplitude and frequency full width at half maximum have 

been followed closely by the measured values. 

The ratio of the positive going signal peaks in the time spectra have proven to be related to the size of 

the reflector and the wavelengths used. It can be used as a practical tool for absolute sizing without 

the use of artificial reference. 

REFERENCES 

1. Proceedings of the DARPA/AFWAL Review of Progress in Quantitative NDE (1981). 

2. Van den Andel, J Stockman, A., and Nicholson, P.S., "The Importance of Correct Alignment 

in Ultrasonic Testing", Advanced NDE Technology, Bussiere, J.F. ed., NRC Canada, 1982, p. 

21. 

3. Nicholson, Stockman and Van den Andel, "Ultrasonic Detection of Defects in High 

Performance Ceramics", CSNDT Journal, Vol. 4, No. 2, Nov. 1982. 

4. Ying, CF. and Truell, R., "Scattering of a Plane Longitudinal Wave by a Spherical Obstacle 

in an Isotropically Elastic Solid", J. Appl Phys.36, 1086(1956). 

5 Tittmann, B.R., Cohen, R.E. and Richardson. J.M., "Scattering of Longitudinal Waves 

Incident on a Spherical Cavity in a Solid", J. Acoust. Soc. Am, 63, 68(1978).

PULSER/RECEIVER 

OSCILLOSCOPE 

SPECTRUM 

ANALYZER 

- 335 - 

GRAPHICS 

TABLET 

FOCUSED 

TRANS- 

DUCER 

Figure 1: Experimental Set-Up 

MINI- 

COMPUTER 

SCANNER 

CONTROL 

FOURIER 

TRANSFORMS 

MODEL 

CALCULATIONS •

- 336 - 

-2.0 -1.0 0.0 1.0 2.0 

TIME (MS) 

10 20 30 40 50 

FREQUENCY (MHz) 

Figure 2: Signal reflected from flat surface of a sample and its corresponding 

frequency spectrum

to 

•a 

I a. a 

Ê £ 

< IS 

i-o 

00 

- 

- 337 - 

Gaussian Frequency Spectrum 

Centre Frequency = 30 MHz 

% 

4» 

4» 

/ 

/ 

•*—F W V 

HM—- 

(5 30 45 

Fre quency ( MHz) 

Full Width at Half Maximum (FWHM) = 16 MHz 

Figure 3: Gaussian frequency distribution showing measured parameters of frequency 

at maximum amplitude and full width at half maximum 

\ 

60

- 338 - 

130ymx 163 y m 

12 22 31 42 52 

FREOUENCY\MHz) 

130 x 163 pm Void 

Figure 4: Representative of characterized voids: 

(a) optical character 

(b) ultrasonic signal 

(c) frequency spectrum obtained from spectrum analyzer 

(d) frequency spectrum from Fourier transform of digitized signal 

* .

Figure 5: Frequency spectra from: 

(a) measured signal off a 50 ym x 53 ym void and 

(b) theory for 49 ym diameter void

Figure 6: Frequency spectra from: 

(a) measured signal off a 130 yra x 163 pm void and 

(b) theory for 132 ym diameter void

- 341 - 

100 150 

Void Diameter (jjm) 

50 100 150 

Void Diameter (pm) 

LEGEND 

+ Minor Diameter 

x Major Diameter 

Model Prediction 

• Input Parameter 

( fc =30 MHz, FWHM =1SMHz ) 

Figure 7: Variation of frequency spectrum parameters with void diameter: 

(a) frequency at maximum amplitude and 

(b) full width at half maximum (FWHM)

TD 

QJ 

o 

o 

c/) 

CL -£ 

E 5 

-1 

-1 

- 342 - 

'0 15 30 45 60 

Frequency (MHz) 

A * 

/ l\ 

l\ l\ 

^ /M 

V !' ^ 

\| u 

0-10 0-20 

Tj me (JÜS) 

Figure 8: Theory prediction for a 90 ym void: 

(a) magnitude frequency spectrum 

(b) time si signal obtained from inverse Fourier transform of frequency 

spectrum 

75 

0-30

- 343 - 

TIME (ns) 

Figure 9: Peak time signal deflections predicted from theory. Signal amplitudes 

are normalized to the largest negative-going signal. Shown are signals 

for void sizes of (a) 28 pm (b)66 ym (c) 94 ym and (d) 132 pm 

100

if) 

a .1.0 

> 0-9 

o ",T; 

to 

O 08 

ÛL 

07 

Model Trend 

Data Trend 

- 344 - 

Minor Diameter + 

Major Diameter x 

Model Prediction • 

fc = 3 0 MHz, Ac =189jjm 

' h-1 

50 100 150 ^200 

Void Dimension (juin) 

Figure 10: Ratio for positive-going peak amplitudes for the peaks preceding and 

succeeding the largest negative-going peak. Lines are used to indicate 

trends rather than represent a fit to the data.

- 345 - 

COMPUTER SIMULATION OF ULTRASONIC TESTING 

V.B. Duncan 

KtomLc. Enz/igy o


- 346 - 

We are concerned here with the computer simulation of ultrasonic testing techniques 

using finite difference approximations of the equations of infinitesimal 

elasticity. These equations are used to describe ultrasonic waves in metals and 

in models of seismic waves in the Earth. Note that we are interested in the 

dynamic (time dependent) form of the equations and that solution of these equations 

is only superficially related to the solution of problems in elastostatics. 

Another spurious connection is with the simpler Acoustic Wave 

Equation; it does not describe the interplay between elastic waves of different 

types. 

Computer simulation is used in nondestructive testing (NDT) in a variety of 

ways. One application is as an aid to the understanding of the physical 

processes involved in, for example, wave interactions with reflecting surfaces, 

cracks, inhomogeneities and other features. Computer models can be much more 

informative than experiments in this work because the displacements and 

stresses are predicted for all points in space and time. A more commercial 

application is the modelling of specific transducers and testing configurations 

where the model is already used to add detar'l to experimental work and will be 

used to select transducers before experimental work is started. 

We avoid the use of detailed mathematics in what follows and instead give a 

general description of the techniques used with references for more detailed 

study. The equations of elasticity are stated in two-dimensional form, but 

approximation techniques are the same in any number of space dimensions. The 

construction of finite difference schemes and the modelling of features such as 

cracks and transducers are discussed. The use of the first order form of the 

Elastic Wave Equation is highlighted, and its various advantages are explained. 

The numerical approximation of the first order form that is used in the 

computer model developed by the author is briefly described. 

2. EQUATIONS OF INFINITESIMAL ELASTICITY 

The equations of infinitesimal elasticity with plane strain can be written in a 

number of different ways. They are most often expressed in Cartesian coordinates 

in the second order form relating displacements and stresses as follows

pu = T XX + T Xy 

tt x y 

pV = T Xy + T yy 

tt x y 

- 347 - 

where U and V are the displacements in the x and y directions respectively, the 

T's are the stresses, and p is the density of the material. The stresses are 

defined in the following way 

T XX = \(U + V ) + 2uU 

x y x 

T yy = \(U + V ) + 2MV 

x y J *" y 

T Xy « u(U + V ) 

y x' 

where X and p. are the lame elasticities. These two sets of equations can be 

combined to give a pair of linear, second order, hyperbolic equations involving 

only the displacements U and V. 

Reformulating the equations in first order form was proposed by Clifton [1] and 

later used by Smith [2] in the form given below. The equations are written in 

terms of the 5-vector 

We have 

where 

W T = (pUt,pVt,T xx ,T yy ,T xy ). 

-t -x ° -y ' 

A =/0 0 1 0 0\ , B =/0 0 0 0 

0 0 0 0 l \ /OOOIO 

cOOOO) lOaOOO 

aOOOO/ \ 0 c 0 0 0 

b 0 0 0/ \b 0 0 O 0, 

and a » A./p, b » u/p, c • (\+2(x)/p. 

Again we have a linear, hyperbolic system of equations. The displacement 

velocities and the stresses are calculated simultaneously in this first order 

form of the equations. The displacement at any point in the space domain can 

be found by integrating the velocity of displacement with respect to time.

- 3A8 - 

3. PROPERTIES OF THE ELASTIC WAVE EQUATION 

We have already noted that the Elastic Wave Equation is hyperbolic (in both 

second and first order form) and therefore has travelling wave solutions. In 

the interior of the space domain disturbances are propagated with speeds /p/p 

and /(\+2|i)/p which are often denoted by c2 and el, respectively. 

Compressional waves are disturbances with displacements parallel to the 

propagation direction and travel with speed cl. Shear waves are disturbances 

perpendicular to the direction of travel and have speed c2. In ultrasonic 

testing, we are particularly interested in waves which are localized in space, 

have short wavelengths and are fast-moving relative to the length and time 

scales of the problem. 

The energy in a perfectly elastic material is the sum of its Kinetic and 

Elastic Strain Energies. The Kinetic Energy Density Is given by 

and the Elastic Strain Energy Density by 

1/2 (\+2(i)(U + V ) 2 + |i/2 /(U +V ) 2 - 4U V ) 

x y V y x x y/ 

The total energy is found by integrating these densities over the space region 

of interest. We note that the energy density can be expressed as an algebraic 

combination of the components of the solution vector In the first order form, 

but that it involves partial derivatives of the second order form solution. 

Energy is conserved in a perfectly elastic material, that Is, the rate of 

change of energy in a space region with respect to time is equal to the total 

of the energy flux through the boundary of the region. This conservation 

property is fundamental and should be modelled accurately. 

4. FINITE DIFFERENCE APPROXIMATION 

Finite difference approximations of the Elastic Wave Equation were first 

developed In the 1960's when their main application was seismology. A major 

contributor to this field was Alterman (see [3]). Most of the work with finite 

difference approximations in NDT simulation has concentrated on approximating 

the second order form of the equations, but two exceptions are mentioned 

above. General surveys of these methods in NDT simulation are given by Aboudi 

[4] and Bond [5]. A general description of finite difference approximations of 

hyperbolic PDE's is given in chapter 4 of Mitchell and Griffiths [6J. 

Schemes of implicit type, that is, schemes in which the approximation formula 

involves coupling between the points at the latest time level, can be rejected 

almost immediately. The reason for this is the prohibitive cost of solving the

- 349 - 

linear system of equations coupling the latest time level of the solution. 

Marfurt agrees with this view in a recent review paper [7]. We, therefore, 

restrict our attention to schemes which calculate the solution at the latest 

time level explicitly. 

The usual finite difference approximation to the second order form of the 

equations is constructed by replacing the partial derivatives with second order 

accurate central difference approximations. (The order of accuracy of a finite 

difference approximation is the order of the leading term in its truncation 

error. See [6], p 21). For example, the second partial time derivative of U is 

approximated by 

U(x,t) = -^ (U(x,t+At) - 2U(x,t) + U(x,t-At)) 

At 

and the second partial space derivative of U by 

U(x,t) = -i- (U(x+Ax,t) - 2U(x,t) + U(x-Ax,t)) 

xx Ax 2 

Bond [5] gives a detailed description of this scheme. 

There are a large number of approximation schemes for first order hyperbolic 

systems and one must consider their accuracy, efficiency and reliability. The 

numerical simulation developed by the author uses a "leap-frog" scheme proposed 

by Kreiss and Öliger [8] which is accurate to fourth order in space and to 

second order in time. These schemes are particularly useful because they can 

be implemented efficiently, give good resolution of the solution and do not 

dissipate energy. In addition, they have been used successfully in weather 

prediction codes and have been the subject of extensive theoretical 

investigation. 

Detailed descriptions of the Kreiss and Öliger leap-frog scheme can be found in 

[8] and [9], and so we give only a brief description here. The scheme is 

constructed by approximating the time derivative of U by 

U(x,t)t = Y^r (U(x,t+At) - U(x,t-At)) 

and the space derivatives by 

U(x,t) = yi-^- |U(x-2Ax,t) - 8U(x-Ax,t) + 8U(x+Ax,t) - U(x+2Ax,t)J 

The difference approximation of the space derivatives must be modified at 

points adjacent to and on the boundaries of the space region to avoid the use 

of points outside the region.

- 350 - 

5. ADVANTAGES OF USING THE FIRST ORDER FORM 

An important advantage of the first order form is that the most common boundary 

conditions can be expressed without involving space derivatives of the solution 

components. This makes numerical approximation at the boundaries much easier. 

For example, the free surface (stress free) boundary condition involves only 

two linear combinations of the three stresses. More general conditions on the 

normal and tangential stress and displacement are equally easy to apply. 

In general, the space domains of interest include a wide variety of shapes that 

do not lend themselves to modelling on a Cartesian grid. Many cases are suited 

to a polar coordinate system or require the use of a non-uniform grid. It is 

generally much more difficult to modify the second order form of the equations 

to use non-rectangular grids than it is to modify the first order form. The 

modelling of inhomogeneous materials is also easier when the first order form 

is used, and the modified set of equations which results is similar to the 

modified equation set for a non-uniform grid. In both cases, the advantage of 

using the first order form is due simply to the ease of manipulation of lower 

order space derivatives. 

Little is known about the properties of numerical approximations to the second 

order form of the equations, but approximations of first order hyperbolic PDE's 

have been investigated in great detail. The stability of difference 

approximations to initial-boundary value problems, such as we are discussing, 

is not easy to guarantee; results are given in [9] and [10] for the first order 

equations. The stability of approximations to the second order form of the 

Elastic Wave Equation has only been investigated experimentally and should not 

be assumed to hold in all cases. 

6. LIMITING THE COMPUTATIONAL AREA 

In most cases, it is not necessary to model the whole object being tested. One 

can safely ignore the parts that cannot contribute to the response signal at 

the transducer during the time period of interest. When modelling the testing 

of weld joints in pipelines, for instance, there is no need to consider more 

than the small region that includes the transducer and weld. Any signal 

spilling out of this region is effectively lost down the pipeline. 

To limit the extent of the computational region, an artificial boundary is used 

to truncate the object and absorb the energy of any disturbance reaching it. 

Ideally, there should be no reflected disturbance at this boundary, and waves 

should appear to pass through it undistorted. It is possible to impose 

perfectly absorbing boundary conditions for hyperbolic systems in one 

dimension, but it is usually impossible to do so for multi-dimensional 

problems. A study of absorbing boundary conditions for wave equations can be 

found in Clayton and Enquist [11]. 

The simplest absorbing boundary conditions have been used in our numerical 

model. They are designed to absorb waves travelling perpendicularly to the 

boundary perfectly, but have proved to be most satisfactory for waves of all 

orientations tried so far. In fact, the energy reflected by this artificial 

boundary usually amounts to only a few percent of the energy reaching it.

- 351 - 

These boundary conditions are constructed by assuming that the solution is 

uniform in the direction parallel to the boundary, then applying the perfectly 

absorbing conditions for the resulting one space dimension equation 

perpendicular to the* boundary. For example, the equation is approximated at 

the boundary x=0 (i.e. a vertical surface) by the one-dimensional equation 

W_ = AW . 

—t —x 

The perfectly absorbing boundary conditions for this equation are 

) - T XX = 0 

/ET(pV ) - T xy = 0 . 

7. CRACKS AND CAVITIES 

We are primarily concerned with the detection of cracks and cavities in 

metals. In an unbounded elastic material it is possible to model a wide 

variety of cracks and cavities by fitting the coordinate system of the 

simulation to the geometry of the flaw. If boundaries and flaws must all be 

modelled, then it can be much more difficult to fit a coordinate system to the 

geometry of the problem. 

Planar cracks with stress-free surfaces have been successfully modelled. 

Cracks with various orientations can be included in the approximation by using 

a coordinate system which is skewed so that one axis is parallel to the crack. 

The crack tips have so far been assumed to be stress free, but further investigation 

of this is required. Some cavities with a simple parallelogram shape 

have also been modelled, but more cojplicated geometries have not yet been 

attempted. 

8. TRANSDUCERS 

The simulation of transducer generated ultrasonic pulses has two closelyrelated 

purposes. It can simply be regarded as a way to send a pulse with 

reasonable physical properties into the computational region to gain insight 

into the physics of interactions with cracks and other features. Alternatively, 

accurate modelling of the input can be used to study the properties of 

specific transducers. We note that accurate models of transducers can be used 

to predict both transmission and reception of signals.

- 352 - 

At Chalk River, we have concentrated on modelling the physics of an ultrasonic 

pulse inside the object being tested and not on the detailed modelling of 

specific transducers. Pulses of both shear and compressional types can be 

input at arbitrary angles to free surfaces and absorbing boundaries. This is 

achieved at free surfaces by applying stresses and at absorbing boundaries by 

specifying a combination of stress and displacement. 

9. COMPUTATIONAL COSTS 

Computer simulation of the Elastic Wave Equation is limited on any computing 

device by two main factors: the availability of fast access storage; and the 

cost of executing the program. The execution cost is simply the cost of 

performing the arithmetic required to calculate successive time levels of the 

solution. Past storage and retrieval of the values of the approximate solution 

at the two preceding time levels is important because of the large amount of 

data _nvolved and the fact that it must be done every timestep. Information 

storage and retrieval becomes very expensive if it cannot be done in the 

machine's central memory (or equivalent) and alternative storage, such as disk, 

is used. 

Execution and storage costs are related to the number of grid points used in 

the construction of the finite difference approximation. To illustrate this we 

consider an approximation in which the number of grid points for each of the 

space dimensions is directly proportional to N (a positive integer). The total 

number of space grid points is then proportional to N to the power D, where D 

is the number of space dimensions. The amount of storage required for the 

approximation and the number of arithmetic operations per timestep are 

multiples of the number of space points. The number of timesteps is proportional 

to N and so the total number of arithmetic operations is a multiple of N 

to the power (D+l). 

The number of space points required to approximate the solution to a given 

accuracy depends on the physics of the problem. Probably the most significant 

factor is the ratio of the wavelength of the ultrasonic pulse to the length 

scale of the object it is traversing. The number of subdivisions in each space 

dimension is inversely proportional to this ratio, making it more expensive to 

model shorter wavelength pulses in the same object. 

To achieve reasonable accuracy in typical problems encountered at Chalk River 

it was necessary to use values of N of at least 100. To model all the problems 

we have considered so far, values of N of 400 or more would be required. The 

storage and execution costs of two-dimensional modelling of the easier problems 

severely stretched the capabilities of our CYBER 175, and difficult 

two-dimensional problems would stretch most Supercomputers. General threedimensional 

modelling of most problems is certainly not possible with current 

technology; the execution and storage costs are prohibitive on any existing 

machine.

10. EXAMPLE 

- 353 - 

The following example demonstrates the application of our numerical model to a 

simple test problem. It represents the generation, propagation and reflection 

by a crack of a shear wave in steel. The computational region is a 10 mm by 10 

mm square and all its edges are energy-absorbing except for a 1 mm crack along 

the centre part of the right edge. The figures show the geometry of the region 

and plot the energy density by contour levels which are graded to pick out the 

low amplitude features. 

The first two figures show the generation of the wave at the left edge uf the 

region. It is generated by specifying a combination of stress and 

displacement, and it has the form of an exponentially damped wave packet wir.h 

frequency 3.5 MHz. Small amplitude compressional waves are generated at tho 

edges of the input region due to the rapid variation in stress and displacement 

there. There are no signs of spurious reflection by the absorbing boundaries. 

Figures 3 and 4 show the shear wave after reflection by the crack. A large 

amount of energy is lost through the absorbing boundary, but a significant 

portion is reflected by the crack. Another important feature is the scattering 

by the crack tips which shows itself mainly in the two compressional waves 

emanating from the tips. 

11. SUMMARY AND CONCLUSIONS 

We have described the construction of numerical approximations of the Elastic 

Wave Equation and outlined the particular advantages of using the first order 

form of the equations. We have discussed the current and potential uses of 

computer simulation in ultrasonic testing and also suggested that powerful 

computers are required if codes are to be run reasonably quickly and 

efficiently. The example given shows that the physical phenomena of reflection 

and crack-tip scattering can be successfully modelled and we have been equally 

successful in modelling other physical phenomena. 

12. REFERENCES 

1) R.J. CLIFTON, Quart. Appl. Math., V. 25, p.97 (1967). 

2) P.D. SMITH, Int. J. Num. Meth. Eng. , V. 8, p.9x11 (1974). 

3) Z. ALTERMAN and D. LOEWENTHAL, Methods in Computational Physics Vol. 12, 

Ed. B.A. Bolt, pp.35-164, Academic Press (1972). 

4) J. ABOUDI, Modern Problems in Elastic Wave Propagation, Eds. J. Miklowitz 

and J.D. Achenbach, pp. 45-65, Wiley (1978). 

5) L.J. BOND, Research Techniques in NDT Vol. 6, Ed. R.S. Sharpe, 

pp. 106-150, Academic Press (1983).

- 354 - 

6) A.R. MITCHELL and D.F. GRIFFITHS, The Finite Difference Method in Partial 

Differential Equations, Wiley (1980). 

7) K.J. MARFURT, Geophysics, V. 49, pp.533-549 (1984). 

8) H. KREISS and J. ÖLIGER, Methods for the Approximate Solution of Time 

Dependent Problems, Global Atmospheric Research Programme Publications 

Series No. 10 (1973). 

9) J. ÖLIGER, Math. Comp., V. 28, pp.15-25 (1974). 

10) D.M. SLOAN, Math. Comp., V. 41, pp.1-11 (1983). 

11) R. CLAYTON and B. ENQUIST, Bull. Seis. Soc. Amer., V. 67, pp. 1529-1540 

(1977).

3 

"/A ! 

- 355 - 

1 I 2 I 

Figs. 1-4. Energy density at times 0.43, 0.86, 3.85 and 4.27 ys, respectively, 

in a 10 mm by 10 mm region. The dashed lines show the energy 

absorbing boundaries and the thick line on the right edge 

represents a 1 mm crack.

- 356 - 

A NOVAL APPROACH TO EDDY CURRENT IMAGING OF DEFECTS IN 

ALUMINUM SHEETS 

V. Le.zma.ni, M. Macecefe 

Tzchno Sc-cewi-cjj.tc Inc. 

Ontcifi-lo 

ABSTRACT 

Eddy current testing relies on the disturbance of eddy currents 

induced in a conducting material by an alternating 

electromagnetic field. While the data acquisition is very 

simple, its interpretation is complex and a lifetime of 

experience of the operator is sometimes required. 

This paper describes a new approach where the flaws in aluminum 

sheets will be imaged in a fashion similar to an X-ray or 

ultrasonic C-scan image. An array of coils is used to generate 

the image of the flaw. Both wire wound and printed circuit coil 

arrays can be used. The inversion of the data is discussed and 

future work outlined.


- 357 - 

The inspection for defects by Eddy Currents is based on an indepth 

analysis of the impedance changes experienced by a single 

coil (1,2). This can, after analysis, yield the location of the 

defect and an estimate of its depth. However, the evaluation of 

the defect type and depth from the impedance plane image or chart 

recording is tedious (Figure 1). Special calibration is required 

to obtain semi-quantitative data on flaw size. Our work proposes 

to take a different approach: developing an eddy current system 

capable of imaging flaws directly, with the goal of designing a 

real time eddy current imaging system to detect and image flaws. 

EDDY CURRENT IMAGING 

A breakdown of eddy current methods for imaging is based on two 

characteristics: 

1) the movement of the testing system in relation with 

the test object. 

2) the number of coils used. 

One method of eddy current imaging is eddy current holography, 

described in detail by B.P. Hildebrand (3). This technique is 

based on the observation that phase is dependent on the distance 

from the receiving coil and uses the backward wave reconstruction 

algorithm of Van Rooy (A). The results reported are encouraging 

but there are some distinct problems with focus and resolution. 

The experimental set-up is of the SD (single coil dynamic) type, 

scanning a single coil, preprocessing the signal in a multifrequency 

eddy current instrument to minimize the noise due to 

variations in lift-off, and using a digital reconstruction 

algorithm developed originally for ultrasonics. 

More sophisticated inversion procedures are discussed in the 

1 itérât - e (5) based on the convolution of some parameters of the 

test object, and a response function. C-scan type pictures have 

been made for eddy currents (6). This practice was also used 

extensively to distinguish between types of defects (pits, 

cracks). 

The AS (array-static) type of test is quite different in nature 

from the previous one. It consists of N coils arranged in an 

array which are then placed on top of the test object to produce 

an image. 

There are several approaches to array imaging. One can separate 

the driver and receiver function of the coils; the array of 

receiving coils is placed within a large driver coil. This has 

the advantage that phase lags can be measured with respect to one 

driver coil. The signal information can be processed in a binary 

fashion. This method is being used by DREP (7)»

- 358 - 

The impedance (or the phase lag) of each coil is measured and if 

there is a sufficiently large deviation from the measurement 

above a good test object a 1 is assigned to the location of the 

coil in question. Another approach has the coils acting 

simultaneously- as driver and receiver. The induced eddy currents 

would be similar around each probe. If the eddy current signal 

could be analysed at the coil location for depth (by measuring 

amplitude and phase at frequencies), an immediate profile of 

the defect could be shown. A holographic method of imaging is 

also possible in the AS mode. It is assumed that the response of 

a coil is not altered by the presence of neighbours. If no 

reasonable pictures can be developed from the static case (AS), 

it will be difficult to develop these from the dynamic case (AD). 

There are similarities with early radiographie techniques 

indicating the presence of something but no indication of the 

depth. Therefore a binary picture will not give the necessary 

information to fully evaluate the importance of a defect. 

However, it suffices for the detection of anomalies. 

In our approach, the response of each individual coil is 

processed, resulting in a depth profile. The experiments 

addressed the questions below. 

- Are all the coils identical or can their response be made 

identical? 

- Are neighbours altering the coil response? 

- Can a calibration curve be drawn for all probes? 

EXPERIMENTAL 

Design of Coils and Building of the Array 

We used the Hocking Phasec D6A Eddy Current instrument, which is 

a multi-frequency device (ranging from 1 kHz to 1 MHz), highly 

flexible and easy to operate. The Hocking Phasec D6A prefers 

driver coils with impedances larger than 50 ohms. This puts 

certain limitations on the choice of coils and the working 

frequency. 

Two tools assisted in coil design. The first is a computer 

program AIRCO, developed by Dodd et al.,(9) which allows for the 

calculation of the impedance of coils with air cores, taking 

into account geometry and wire gage. This program was adapted 

for the microcomputer in C BASIC. Ferrite cores were also 

introduced to increase the inductance of the coils. 

Probes were manufactured in-house or were commissioned from a 

reputable manufacturer for the final array. An HP 4192A 

impedance meter was used to measure the impedance and inductance 

of the coil in air. The resonance frequency of the coils was over 

600 kHz.

- 359 - 

A 3 by k array was constructed with twelve nominally identical 

coils. It was made by potting the ferrite ends in holes which 

were drilled carefully in a fiberglass plate. This assembly was 

then potted and the leads connected to a mechanical switch. See 

F igure 2. 

Mechanical Scanner ^R j g 

In order to assess the effect of coil location on the defect 

response, a number of scans were carried out. A computer 

controlled ultrasonic immersion X-Y scanner provided accurate 

encoding of the X,Y position of the coil, as well as an isometric 

display. A number of repeatable scans with variable scanning 

patterns could then be made. A block diagram of the experimental 

set-up is shown in Figure 3. 

INFLUENCE OF NEIGHBOURS ON THE PERFORMANCE jDF CO ILS 

A series of experiments was developed to assess the influence of 

neighbours. The first test consisted of moving the coils with 

respect to each other while recording the differences in signal 

output of the eddy current instrument. Impedance of a coil 

was also measured before and after it was placed into an array. 

The principal reason for the change in impedance when neighbours 

are present < Lnought to be the presence of ferrite cores. Due 

to this higl- permeability material there is an increase in the 

impedance uf the coil. Figure U shows the change in impedance 

as a function of distance between two coils. It shows the 

exponential drop off with distance, with the effect disappearing 

at about *i mm. In our array,coils are placed within a distance of 

less than two millimeters and are therefore strongly affected 

by neighbours. The impedances and inductances were also measured 

on twelve coils before and after they were arranged in an array 

with distance between nearest neighbours equal to 1.7 mm. The 

change in inductance increases with increasing Neighbouring 

Factor. While the standard deviation was 3-6 JJH for individual 

coils, it became 8.9 fiH when the coils were in an array. The 

changes between individual coils entail that: 

1. rebalancing of each coil in the array is 

important for traditional eddy current 

system approach; 

2. the lift-off direction will be different 

for each coi1 ; 

3. calibration curves will be coil dependent. 

These coils were wound within one turn of each other on 

Identical ferrite slugs. It is unlikely that the 

reproducibility of the coils can be improved with existing 

techniques. Photo lithography was an alternative which was 

exploited.

Printed Circuit Coils 

- 360 - 

To assure repeatability of the coil parameters, a printed circuit 

coil method was applied. Printed circuit coils (PCC) were made 

with planar technology. A master was first drawn on a high 

contrast medium, then photographically reduced (100 X) to actual 

size. The spiral coils were drawn by a microcomputer scanner. 

The masters were reduced in a photolithographic laboratory on 

Kodak high resolution plates. 

Figure 5 shows the layout of a multi turn planar inductor in a 

spiral configuration. A simplified design equation giving the 

inductance L was reported by Brown (8) as 

L = 0.8(r - nd/2) n 

(6r + 7nd) 

L t Inductance (nH) 

r : Outer radius of coil (mil) 

d : turn width and spacing between turns (mil) 

n : number of turns 

The agreement with the equation was within 10%; wound coils 1 

inductances were higher. The coils were first made on regular 

printed circuit material but a serious underetching problem 

prevented the use of the thick (0.002") Cu plate. Pre-thinning of 

the copper plate by etching did not produce an even surface. 

Therefore, vacuum deposition of aluminum was attempted giving an 

Al layer in excess of 1000 A. Shipley positive photoresist was 

applied and coil patterns were exposed to ultraviolet light. 

After developing, the aluminum was etched in an etchant based on 

phosphoric acid, HNO , and acetic acid. The strip width is about 

12 /urn. A number of substrates were experimented with including 

glass, plexiglass, machinable ceramics and nylon. Best results 

were obtained with machinable ceramics, but the plexiglass and 

other flexible materials were also mastered. The coils produced 

in this way had a desired geometry and repeatability but due to 

tne thin layer of aluminum, very high resistance. To increase 

the cross section of the conductor, a copper or gold layer was 

grown on the aluminum pattern by electrochemical means. Finally, 

contact holes 0.006" (0.15 "*") were drilled and contacts secured. 

Using this technology, repeatable and identical coils could be 

produced in any array configuration. The particular coil was 

2.18 mm in diameter and had 29 turns 

We have also experimented with larger coils (21.8 mm in 

diameter.) Kodak negative photoresist and standard copper clad 

plates were utilized. This type of coil has 29 turns, a 

resistance of A ohms and an inductance of 10.3 pH at 1 kHz. 

Resonance did not occur below 13 MHz. The requirement of the 

instrument for 50 ohms impedance indicated that higher 

frequencies (above 700 kHz in this case) should be used. The 

coils picked up all the notches including the .1 mm deep by 2 mm 

long flaw.

INVERSION PROBLEM 

- 361 - 

The inversion is a mathematical process which yields the true 

dimensions of the flaw based on the signals obtained with sensing 

coils. A number of works address purely theoretical inversion 

protocols and very little work is done on experimental data. 

The modelling of lift-off as well as the response of practical 

probes is well understood. A flaw response in both twodimensional 

and three-dimensional geometries has also been 

developed. Although great progress has been made in improving 

mathematical modelling of eddy current responses, the theory has 

not been adequately tested against experimental practice. A 

flowchart of the inversion is shown in Figure 6. 

We developed a novel image producing method based specifically on 

array concepts. A hexagonal array (AB type) is suggested as 

shown in Figure 7« Calculations by W.G. Simpson et.al. (9) have 

shown that the change in impedance due to a point defect can be 

approximated by a complicated relationship represented by the 

equation below, where F (x)=signal response, x=distance from 

centre of coil and r=effective coil radius. Simply put, the 

coil is not sensitive under its centre. 

F (x) = A M - cos 

F (x) = 2A xlexp - [x - l\ | L< x < r 

2A xlexp - /x - i\ 1 

F (x) = A exp |-2fx-r 

x ] 

r) 

~ I V 2 JJ 

The parameter of interest (eg. crack depth) in matrix form is: 

T = 

t(i,l) t(i,m) 

t(n,i) t(n,m) 

where (i,k) is the location of the coil with (i,k) coordinates. 

The response of the n X m coils is also described by the matrix 

S = 

s(n,i) s(n,m) 

The measured S matrix should be transposed into a T matrix. 

Let us assume that the response of the coil at location n,m is a 

weighted sum of the crack depth values (t) at the nearest 

neighbour positions. 

s(m,m) = t(n-1,m-1) + t(n, m-1) + t(n-1,m) + t(n, m+1) 

+ t(n+1, m-1) + t(n+1,m)

- 362 - 

This is a reasonable approximation if t(n,m) is the average value 

of the depth over a circular area. The sum of the responses over 

the six nearest neighbours of the location that is to be 

investigated (n,,m) is calculated overt 

F(n,m) = s(n-1,m) + s(n, m-1) + s(n, m+1) + s(n-1, m-1) 

This can be expressed as 

+ s(n+1, m-1) + s(n+1,m) 

F(n,m) = 6t(n,m) + 2 t(n-1,m) + 2t(n,m-l) 

+ 2t(n+l,m) + 2t(n,m+l) + 2t(n-1,m-1) 

+ 2t(n+T,m-1) + 2t(n+2,m) + 2t(n-2,m) 

+ 2t(n+1,m-2) + 2t{n+1,m+i) + 2t(n-1,m-2) 

+ 2t(n-1,m+1) + t(n-2,m-1) + t(n-2,m+1) 

+ t(n+2,m-1) + t(n+2,m+1) + t(n,m-2) + t(n,m+2) 

The first sufficient approximation is t(n,m) = F(n,m) 

6 

This solution can be used for an iterative procedure for the 

solution of n X m linear equations with n X m unknowns, thus 

completing the transformation of the S matrix into a T matrix. 

The values of the T matrix can then be color coded to produce an 

image with depth indications. The above methodology applies both 

to wire wound and printed circuit coils. This procedure, which 

plots depth estimates at the center of each individual coil 

core, can also be used in the dynamic case. 

TESTS PERFORMED WITH THE TSI ARRAY 

This array was prepared in house. In short, twelve coils were 

arranged in a rectangular pattern of three by four coils, with 

the distance between the nearest neighbours being 1.75 nm - The 

individual coils serve both as driver and receiver. 

The performance of individual coils was assessed, namely the 

sensitivity to detect the shallow (.1mm deep by 2mm long) 

slot (Figure 8). The coils picked up the defect with no 

difficulties. 

This array was also mechanically scanned over an aluminum plate 

containing 3 3" hole. Figure 9 shows two isometric plots 

obtained by scanning the array over 33 m hole. 

We experienced some asymmetries: coils 1 to 6 had a larger left 

peak than right peak with the opposite being the case for ^ to 

12. This was due to the higher concentration of flux lines to

- 363 - 

the right of coils 1 to 6, with the opposite being the case for 

coils 7 to 12. These flux lines were constrained by the presence 

of the neighbours. 

Tests were performed with single coils over the source plate and 

symmetrical curves were observed as shown in Figure 10. The 

asymmetries observed in the array studied were due solely to the 

effect of neighbours and / or asymmetries in the induced field. 

The coils had to be individually balanced, and lift-off 

directions and multipliers had to be adjusted for each coil. 

The eddy current signal was influenced by individual coil 

parameters and to a small degree by the location of the coil. It 

was acknowledged that coil parameters such as impedance were 

affected by the location of the coil in the array. Also, for our 

array, a simple signal processing method was not directly 

apparent. However, a novel inversion protocol was devised which 

would lead to advanced true flaw imaging. 


Eddy Current imaging is feasible with arrays of coils. Coils, 

when arranged in an array, influence each other. Interactions 

between neighbours must be considered. The changes in impedance 

due to the presence of defects are very small (Az = 10" 3 - 10" 2 ). 

With present technology there is a need for individual treatment 

of each coil in the areas of balancing, lift-off orientation and 

calibration. We believe the balancing problem can be solved by 

means of software. 

The test notch (,1mm X 2 mm) is detected with no difficulties in 

the coil design used in the array. The coils should be made 

small, identical and powerful. Printed circuit coils can result 

in identical elements in the array. However these coils operate 

at higher frequencies (250 KHz and above) and would be useful for 

the inspection of open shallow cracks. 

We have developed a novel inversion procedure to size flaws. A 

hexagonal stacking of identical coils is advocated and the 

summing of coil responses in groups of six is shown to have 

significant signal to noise improvement. Utilizing arrays to 

produce images of defects yields improved S/N and prompt imaging, 

speeds the inspection and provides superior flaw 

characterization. 

The signal processing and three dimensional colour graphics 

display techniques developed elsewhere can be easily applied to 

this problem (10).


- 36A - 

The work was supported by the Defence Research Establishment 

Pacific. We acknowledge the input of Mr. W. Sturrock, his 

critical comments as well as permission to publish the work. 

References 

1. H.L. Libby, Introduction to Electromagnetic Nondestructive 

Test Methods, (New York; Wiley- Interseience, 1971). 

2. R. L. Stoll, The Analysis of Eddy Currents, (Oxford : 

Clarendon Press, 

3. B. P. Hildebrand,Eddy Current Imaging, Spectrum Development 

Laboratories, Inc., 1982. 

k. D. L. Van Rooy, Digital Ultrasonic Wavefront Reconstruction 

in the Near Field, IBM Report # 320.2402. 

5« B.A. Auld, et.ai., "Quantitative Modelling of Flaw 

Responses in Eddy Current Testing", Nondestructive 

Testing, Vol. 7, (London: Academic 'Press, 198A), 

PP 37-76- 

6. D.C. Copley, "Eddy Current Imaging for Defect Characterization 

", Review of Progress in Quantitative NDE, (1982.) 

7. W. R. Sturrock, Defence Research Establishment Pacific, 

private communication to M. Macecek, 1984. 

8. R. B. Brown, "Experimental Study of Q Dependence on Thin 

Film Inductor Geometry", International Microelectronics 

Symposium, 1973., paper 5A, pp 1-13. 

9. W. A. Simpson, C. V. Dodd, J. W. Luquire and W. G. Spoeri, 

Computer Programs for _sqnie Eddy - Current Problems -1970, 

ORNL-TM-3295 

10. M. Macecek, K.L. Luscott, J.D. Wells and D.K. Mak, 

"Automated Ultrasonic Testing System for the 

Characterization of Defects in Weldments", Paper to the 5th 

Canadian Conference on NDT, Toronto, Oct. 29-31, 1984.

—- 

••;•••• 

'IM 

i lüiiaiiüiiii! au. 

BR JSH ACCUCHART 

1 

t— 

If* »fff 

—I—\ 

M 2 

éi 

| 

- 365 - 

•"1 

H 1 1- 

I?» 

rt: 

4 5 

M, 

1 - .005" notch 

2 - .015" notch 

3 - .030" notch 

U - lift-off pad 

5 - ferromagnetic 

inclusion 

Figure 1. Eddy Current impedance plane images of notches 

and corresponding chart recording

\ 

X-Y SCANNER 

OBJECT 

- 366 - 

Figure 2. Potted coil array with switch arrangement 

IE.C . PROBE 

TWO-FREQUENCY 

EDDY CURRENT 

TESTER 

COMPUTER 


~" 

-- 

Figure 1. Block diagram of experimental set-up 

MIXER

- 3hl 

2 3 4 

OISTANCE (mm) 

Figure h. Influence of neighbours on coil impedance 

and the physical arrangement of the coils 

Figure 5. Layout of the printed circuit coil and 

finished probe (X 50)

- 368 - 

Multiposiiion/mullifrequency measuremenl 

Gross crack size 

Position Length Width Opening 

Figure 6. Flow diagram illustrating an inversion protocol 

(after Auld [5]) 

Figure 7. Hexagonal array (AB type)

- 369 - 

a) 0.080 X 0.010 X 0.004" 

55 dB gain 

f, = 900 KHz 

f2= 700 KHz 

b) 1/4 X 0.010 X 0.004" 

55 dB gain 

f, = 900 KHz 

f.,= 700 KHz 

Figure 8. Impedance plane images of EDM notches

DISTANCE (in cm) 

Figure 9a. Isometric display of the response of individual 

TSI array coils to a 3 mm hole; probe TSI #11

- 371 - 

DISTANCE ;,n(m) 

Figure 9b• Isometric display of the response of individual 

TSI array coils to a 3mm hole; probe TSI #9 

Figure 10. Isometric plot of the response of a single coil 

to a 3mm hole

- 372 - 

DEVELOPMENTS IN X"RAY STRESS MEASUREMENT 

THE CANMET PORTABLE STRESS DIFFRACTOMETER 

R.A. Holt 

Enzfigy, H-Lne.i> and Rz&ou.fic.Q.i, Canada 

Ottawa, Ontaftlo 

ABSTRACT 

The measurement of stress in a surface with X-rays makes use of 

the Bragg law for the diffraction of monochromatic X-rays from a crystal 

to measure precisely the small changes in lattice parameter corresponding 

to elastic distortion, until recently, accurate stress measurements 

using this technique have been confined to the laboratory. 

The recent development of compact position sensitive proportional 

counters, solid state high voltage supplies and compact X-ray 

tubes and the availability of powerful, inexpensive computers make it 

possible to design a diffractometer to measure stress in engineering 

structures and components in the field under clement conditions. 

The CANMET Portable Stress Diffractometer has been designed to 

make accurate stress measurements in the field while retaining adequate 

versatility for dedicated laboratory use. It has several unique features 

including the precisely controlled angular alignment of the incident 

X-ray beam relative to the specimen surface which allows the measurement 

of stress in coarse grained materials and those with pronounced gradients 

in crystallographic texture. A commercial prototype of the instrument 

is currently being built.

- 373 - 


X-ray diffraction is one of a few nondestructive techniques for 

measuring stress in engineering components and structures. The basic 

principle is simple and techniques using X-ray sensitive film for measuring 

stress in crystalline materials were developed more than 40 years 

ago (1). Recent developments in electronics and X-ray detector technology 

have allowed a reduction in the size of the equipment required and 

an increase in the speed with which accurate measurements can be made. 

Furthermore, an improved understanding of the relationship between the 

stress state in a polycrystal and the distortions measured by X-ray diffraction 

allows a more reliable interpretation of data than was previously 

possible (2,3). 

THE PHYSICS OF X-RAY DIFFRACTION 

The measurements of stress by X-rays is based on the Bragg law 

for the diffraction of monochromatic X-rays by the planes of a crystal, 

i.e.: 

nX = 2d sin 6 Eq. 1 

where: n is the order of diffraction 

X is the wavelength 

d is the interplanar spacing of the diffracting planes 

e is the diffraction angle 

Unique values of X are obtained by making the anode of the 

X-ray tube from a pure metal ; the X-ray wavelengths normally used correspond 

to the transition of electrons from the L to the K orbitals of the 

atoms in the anode. The nature of diffraction from a three dimensional 

crystal lattice is such that the X-rays appear to reflect at a unique 

angle (S) corresponding to a unique interplanar spacing (d). As ö 

approaches 90° - the back reflection range - changes in d corresponding 

to elastic strains of less than 10"^ can be measured with acceptable 

accuracy. 

The X-ray wavelengths useful for diffraction in metals are in 

the range 0.15-0.25 nm and these have a penetration depth in most metals 

of a few hundredths of a mm. Hence X-ray diffraction is suitable for 

measuring elastic distortions at the surface. 

PRINCIPLES OF MEASUREMENT 

We define a set of axes x, y and z in which z is normal to the 

specimen and x and y are aligned with some axes of symmetry in the specimen 

(Fig. 1). In practical terms we can measure the lattice spacing in 

any direction, m, within a cone 5O~60° from the specimen normal. We

- 374 - 

define the axes u, v and w where w is the projection of m into the specimen 

surface. The angles * and 0 define the orientation of m, u, v 

relative to x, y and z. 

The elastic strain along m is 

d d-d d-d 

n m _ in o „, m o 

G s j_n 

m d d d 

o o 

where: dm is the interplanar spacing along m, 

d0 is the interplanar spacing with no stress and 

the change from the exact value d0 to the approximate value 

d, causes only a small error in era. 

Then: 

em = ew(cos 2 ip) YUW cosipsini|/ 

and, for material exhibiting isotroplc elasticity: 

uw 

uw 

G 

where: ew, eu, yuw are normal and shear strains in the u, 

v, w coordinate system, 

a u> a v> a w anc ' T uw are n 01 "" 13 ! an( l shear stresses 

in the u, v, w coordinate system 

E, v, G are Young's modulus, Poisson's ratio and the shear 

modulus. 

The X-rays normally sample only a thin surface layer in which 

stress and structure gradients are small. With no surface tractions 

= Tuw = 0 

Eq. 

Eq. 3 

Eq. 3a 

Eq. 3b 

Eq. 3c

and 

- 375 - 

e a -^(a + a ) + —-— .a .sin ty Eq. 

m E u v E u 

Referring to equation 2, au may be found from the slope of 

a plot of the measured lattice spacing dm versus sin"^. If the assumptions 

of isotropic elasticity and absence of large stress gradients in 

the measured volume are valid, such plots are linear (Fig. 2) and only 

two measurements are necessary to obtain ou. If two detectors are 

used, the two measurements can be made simultaneously, intercepting diffracted 

beams on each side of the incident beam within the u-w plane 

(Fig. 3(a)). If only one detector is used, the X-ray source and detector 

must be rotated by several 10 f s of degrees relative to the specimen 

surface to obtain lattice parameter measurements at two values of t|i 

(Fig. 3(b)). The two methods are called single exposure and double exposure 

techniques respectively. With either technique, measurements in 

three planes normal to the specimen surface (usually 0 = 0,^5 and 90°, 

Fig. 1) are sufficient to define fully the surface stress tensor. 

If the assumption of isotropic elasticity is invalid, or if 

stress gradients are sufficiently steep that the stress varies significantly 

within the layer penetrated by the X-ray beam, or if a triaxial 

stress distribution can be sustained within the surface layer, then plots 

of d vs. sin 2 i|i may no longer be linear. In this event, both the single 

and double exposure techniques may give incorrect results (3), and measure 

ments must be made for a number of values of ty at each value of 0 (multiple 

exposure technique). From such measurements, a triaxial stress 

tensor can be extracted, or corrections made for effects of anisotropic 

elasticity or depth gradients in stress (see references 2 and 3 for detailed 

treatments). 

For instruments in which both the X-ray source and specimen remain 

stationary during a single lattice parameter measurement and the 

X-ray line profile is sampled by a moving slit or fixed position sensitive 

detector, errors occur if the grain size is coarse or if steep gradients 

in crystallographic texture exist. Because of the reflecting 

geometry, the diffracted profile arises from a range of orientations of 

diffracting planes corresponding to half the range of 2e over which the 

profile is sampled (Fig. 4). If the number density of diffracting planes 

varies with orientation, the diffraction profile is affected. The effects 

of coarse grain size can be partially compensated by averaging the 

diffracted profile for a small range of incident beam orientations (3) 

but precise control of the incident beam direction combined with data 

manipulation are required to make a rigorous correction (4). 

The single exposure technique has an important advantage over 

double or multiple exposure techniques. Because the effective radius of

- 376 - 

the instrument is the same for both detectors, the error arising from 

misalignment of the specimen surface relative to the center of the 

instrument when the instrument is rotated is eliminated (5,6). This is 

particularly important for a portable instrument designed to examine 

large specimens, because the instrument rotates about the specimen and 

providing an accurate center of rotation can be difficult. 

ADVANCES IN INSTRUMENTATION FOR X-RAY STRESS MEASUREMENT 

The explosion in the development of solid state electronics in 

the past 15 years has resulted in many advances which make it possible 

for accurate measurements of stress to be made with a portable instrument. 

The new developments include inexpensive, compact but very powerful, 

computers, compact solid state high voltage supplies capable of 

delivering 100-200 W and stable, accurate electronic circuits for signal 

analysis. Very small X-ray tubes have been developed to complement the 

solid state power supplies (Fig. 5(a)). 

However, the most important advance has been the development of 

small position sensitive proportional counters (7). These measure the 

distribution of X-ray intensity in one dimension with a resolution of 

50-60 um. Detectors as small as 105 x 50 x 50 mm, with an active length 

of 50 mm (Fig. 5(b)) and weighing only = 0.6 kg are now commercially 

available. A position sensitive proportional counter consists of a fine 

wire centered in a chamber filled with Xe-lOJCHij or Ar-10?CHij (Fig. 

6(a)). For the counters shown in Fig. 5 the wire is a 25 ym diameter 

quartz filament, coated with carbon, which has a resistance of 8 kjj/mm. 

The wire is at a high positive voltage relative to the walls of the chamber. 

X-rays passing through the chamber ionize the gas causing electron 

cascades onto the wire. The voltage pulses observed at each end of the 

wire, arising from each cascade, have rise times proportional to RC, 

where R is the resistance from the end of the wire to the point of the 

electron cascade, and C is the capacitance between the end of the wire 

and ground (Fig. 6(b)). The two voltage pulses are shaped and amplified 

to produce bipolar pulses whose cross-over times are proportional to the 

rise times (Fig. 6(c)). Each cross-over time is marked by a very sharp 

voltage pulse, one of which is delayed a constant amount 6 (Fig. 6(d)) 

so that it always follows the other. Thus, their time separation T is 

proportional to the position of the X-ray photon, x, in the counter. 

Electronic circuitry to perform this analysis is available commercially. 

DESIGN OF THE CANMET PORTABLE STRESS DIFFRACTOMETER 

These new technologies have been applied to design a versatile 

diffractometer capable of accurate stress measurement by either single 

or multiple exposure techniques. The CANMET diffractometer will be suitable 

for field use under clement conditions but with sufficient versatility 

and accuracy for dedicated laboratory use.

- 377 - 

The commercial prototype is expected to be ready for testing in 

December 1984. It consists of the head (Fig. 7), a stand, a 10 m umbilical 

cord and a services package containing analysis electronics, computer, 

X-ray power supply, closed circuit cooler, detector gas supply and 

motor controllers. The head will weigh ~15 kg, the simplest stand 

~10 kg, the umbilical cord ~12 kg and the services package ~90 kg. 

The instrument may be used with either one or two position sensitive 

proportional counters of 50 mm active length. A unique feature 

(H) is the precisely controlled angular rotation of the X-ray tube about 

the focal spot (a-rotation, Fig. 8(a)) combined with data manipulation 

to allow accurate determination of the Bragg angles for specimens with 

large grain size or pronounced gradients in crystallographic texture. 

For single exposure measurements, only the a-rotation is used. With 

a-rotation the diffraction peaks shift in the detectors, and hence the 

a-rotation can be used to calibrate the position sensitivity of the 

detectors. 

For multiple exposure measurements, the X-ray tube and detector 

(s) can be rotated simultaneously about the focal point (w-rotation, 

Fig. 8(b)). The uniformity and concentricity of this rotation must be 

precise, but a high degree of angular accuracy is not required. The a 

and ai rotations are coplanar and the stress is determined in the direction 

of the line of intersection between the plane of rotation and the 

specimen surface. 

Because the diffractometer is designed to measure stresses in 

structures as well as small specimens, the design incorporates stages to 

position the diffractometer relative to the specimen. The adjustment 

along the direction of the specimen normal is most critical and is incorporated 

into the diffractometer head (Fig. 7). A position sensor is 

provided to sense the position of the diffractometer in this direction, 

relative to the specimen surface, and allow accurate automatic positioning. 

The diffractometer head also rotates about the specimen normal 

(^-rotation) to allow stress measurement in any direction in the specimen 

surface. 

Translations in the plane of the specimen surface are provided 

to allow one or two dimensional mapping of stress. The stages providing 

these motions will be part of a laboratory stand. 

Table 1 summarizes the important features of the CANMET portable 

stress diffractometer. 

PERFORMANCE OF THE CANMET DIFFRACTOMETER 

To date, some aspects of the design have been evaluated with an 

engineering version of the diffractometer.

- 378 - 

It incorporates many of the features of the commercial prototype 

although it lacks the w-rotation. Relative accuracies of ±18-20 MPa 

(standard deviation) have been obtained on annealed ferritic steel and 

cold worked Inconel 600 calibration beams, for example Fig. 9. The 

counting times used to obtain this accuracy are ~150s with another 90s 

in executing the o-motion of the X-ray tube. With streamlining of the 

software and improvements in data reduction measurement, times for fine 

grained ferritic steels will be further reduced (possibly by a factor of 

2). 

In actual applications, the engineering version of the diffractometer 

gives results comparable with those obtained using a standard 

Phillips X-ray diffractometer modified for stress measurement (8), for 

example Fig. 10. 

SUMMARY 

The CANMET Protable Stress Diffractometer is designed for field 

and laboratory use. The design takes advantage of the error compensating 

feature of the single exposure technique, but still allows the investigation 

of stress distributions requiring multiple exposures. New detector 

and solid state electronics technologies have been applied to make 

it transportable for field use under clement conditions. It will be as 

accurate as standard laboratory equipment and have sufficient versatility 

for dedicated laboratory applications. A unique feature (the a-drive) 

allows the elimination of texture gradient and coarse grain size effects. 

A commercial prototype incorporating these features is currently under 

construction. 


The CANMET Portable X-Ray Stress Diffractometer was conceived, 

and until his retirement January 198 1 !, developed, by CM. Mitchell of the 

Physical Metallurgy Research Laboratories (PMRL), CANMET, Department of 

Energy, Mines and Resources Canada. The engineering model of the instrument 

was designed and made by Proto Manufacturing Ltd., Oldcastle, 

Ontario, in 1983 with funding from CANMET. The commercial prototype of 

the diffractometer will be made by the same company within the framework 

of the Program for Industry Laboratory Projects (PILP). We wish to thank 

J.A. Hunter of NRC for his enthusiastic support in the role of PILP 

Project Manager and D.W.G. White of PMRL for useful discussions and help 

in other ways.

- 379 - 

REFERENCES 

1) Klug, H.P. and Alexander, L.E. X-ray Diffraction Procedures, John 

Wiley and Sons, New York (1954) p. 539 

2) Dolle, H. Journal of Applied Crystallography, V.12, 489 (1979). 

3) Maeder, G., Lebrun, J.L. and Spravel, J.M. N.D.T. International, 

235 (October 1978). 

4) Mitchell, CM. Proc. International Conference on Pipeline Inspection, 

Canadian Government Publishing Centre, Ottawa (1984). 

5) Mitchell, CM. Advances in X-ray Analysis, V. 20, 379 (1977). 

6) Ruud, CO. and Snoha, D.J. Journal of Metals, 33 (February 1984). 

7) Schelten, J. and Hendricks, R.W. Journal of Applied Crystallography, 

V. 11, 297 (1978). 

8) Winegar, J.E. Report No. AECL 6961, Atomic Energy of Canada Ltd., 

Chalk River, Ontario, June 1980.

- 380 - 

Table 1 - Important Features of CANMET Portable Stress Diffractometer 

HARDWARE - uses 2 position sensitive proportional counters with analog 

timing electronics 

- small 200 W X-ray tube and compact solid state power supply 

- IBM-PC compatible microcomputer 

- for large components or structures and field use under clement 

conditions 

- transportable but with adequate versatility and accuracy for 

dedicated laboratory use* 

- capable of single or multiple exposure methods» 

- a-drive controls angle of incident X-ray beam* 

- (o-drive controls tilt for multiple exposure technique 

- positioning stages for manipulation of diffractometer head* 

SOFTWARE - complete software package for calibration, operation of diffractometer 

and data manipulation to eliminate texture gradient 

and grain size effects* 

•unique features of CANMET Portable Stress Diffractometer

- 381 - 

incident x-ray beam 

diffracted x-ray beam 

Fig. 1. Coordinate system for X-ray stress analysis. z is 

normal to the specimen surface and m is the direction 

of measurement of the lattice spacing. An angle of 

180°-29 separates the incident and diffracted beam. 

The measurement direction, m, falls on the bisector of 

this angle. 

0.804 

z 0.803 

a! 

0.802 - 

O.I 0.2 0.3 

SIN 2

- 382 - 

incident x-ray 

beam 

detector I incident 

'x-ray beam 

letector 2 

Fig. 3. Representations of the u-w plane of Fig. 1 showing the 

geometry of: A - the single exposure technique and B - 

the double exposure technique for the measurement of 

stress along the "u" direction using lattice parameter 

measurements m, and m 2' 

incident x-ray 

beam 

Fig. 4. Representation of the u~w plane of Fig. 1 showing three 

diffracted beams, d.. , d and d~, for a given incident 

beam direction. That corresponding to the centre of the 

diffraction profile, d„, arises from lattice planes 

normal to m„ while those corresponding to the high and 

low angle tails of the profile, d. and d,, arise from 

lattice planes with normal m. and m., respectively. 

Because of the reflecting geometry, the angular separation 

of the diffracted beams is twice that of the plane 

normals. 

m,

- 383 - 

Fig. 5(a). 200 W X-ray tube manufactured by KEVEX 

Fig. 5(b). Small, commercially available position sensitive 

proportional counters manufactured by Technology for 

Energy Corporation (TEC) and M. Braun GmbH.

- 38A - 

VOLTAGE L S T S 

Y Y Y 

s 

TIME 

Fig. 6(a). Schematic diagram of a position sensitive proportional 

counter. 

Fig. 6(b). Voltage pulses produced at the ends of the counter 

wires by an X-ray absorbed by the counter gas at X. 

Fig. 6(c). Shaped voltage pulses with cross-over times proportional 

to rise times of the initial pulses. 

Fig. 6(d). Sharp, pulses marking the cross-over times. The pulse 

at S is delayed by an amount S to S' so that the 

separation T is proportional to X.

- 385 - 

Fig. 7. Head of the commercial prototype of the CANMET Portable 

Stress diffractometer. It includes tube, T, collimator, 

C, two detectors, D.. , and D~, a and u) rotations about 

the axis F-F^ in the specimen surface, translation 

along the specimen normal, z, and rotation about the 

specimen normal, . 

Fig. 8 a-rotation. The X-ray tube, T, rotates about the focal 

point F, while the angular positions of the detectors 

Dl and D2 remain fixed. w-rotation. The X-ray tube and 

detectors rotate simultaneously about the focal point.

- 386 - 

APPLIED STRAIN 

Fig. 9. Elastic strain (measured using X-ray diffraction) as a 

function of applied strain (measured using strain 

gauges) in a flat beam of Inconel 600 stressed in four 

point bending. Note the large compressive residual 

stress in the beam (indicated by the negative intercept 

with an applied strain of zero). 

MEASUREMENT POSITIONS 

Fig. 10. Longitudinal residual stress as a function of circumference 

(relative to the tube axis) on the outside 

surface of "u" bend in an Inconel 600 nuclear steam 

generator tube. Measurement position 1 is at the 

inside diameter of the bend, measurement, 15 at the 

outside diameter.

- 387 - 

APPLICATIONS OF NEUTRON DIFFRACTION TO ENGINEERING PROBLEMS 

T.M. Holdan, G. Veiling, S.U. MacEwan, J. Wlne.gai and B.M. Pouie.ZZ 

Atomic. Ena-igy o () Canada Limited 

Chalk Rl\mK, Ontaulo 

R.A. Holt 

Physical Me.taZlu.igy Rziaafich Laboiatonle.!, 

Ottauia, Ontailo 

ABSTRACT 

Neutron diffraction may be applied with advantage to the study of engineering 

problems in areas currently served by X-ray diffraction. At Chalk River, neutron 

diffraction has been applied to the measurement of lattice strains in an 

over rolled Zr 2.5 wt% Nb pressure tube and in a bent Incoloy-800 steam generator 

tube. Basically, the interplanar lattice spacing, deduced from the diffraction 

peak, acts as an internal strain gauge. Our experiments on the pressure 

tube showed that a major contribution to the lattice strain comes from 

grain to grain interactions and allowed us to measure the strain tensor associated 

with the overrolling procedure. The experiments on the generator tubing 

showed that, even in a cubic material, the strain in the longitudinal direction 

is higher by a factor of two for grains aligned with an [002] crystallographic 

axis longitudinal than for grains aligned with a [111] axis longitudinal. The 

differences originate in the elastic anisotropy of Incoloy-800. A number of 

other potential applications of neutron diffraction are discussed. 


Neutron diffraction can be applied to the study of engineering problems in many 

areas currently served by X-ray techniques. The advantages of neutron 

diffraction over X-ray diffraction originate in the low neutron absorption and 

scattering cross sections of most elements. For example a 1 Â thermal neutron 

beam is only attenuated by a factor of 3, 1.3 and 1.1 by a 1 cm thickness 

respectively of steel, zirconium and aluminum. Hence one is usually able to 

measure peaks diffracted by atomic planes whose normals lie along any chosen 

direction in the component. With this flexibility one can examine, for 

example, the hoop strain in a tube by passing the beam through one wall and 

scattering off the second wall oriented so that the tangent to the tube bisects 

the incident and scattered beams. With a thin-walled tube, the property (be it 

texture, or strain) is averaged through the wall with neutrons whereas X-rays 

measure the property in the first few microns from the surface. With a carefully 

collimated beam of neutrons one can, by moving the component through the 

point of intersection of the incident and scattered beam, map out the properties 

through the bulk [1].

- 388 - 

The disadvantages of neutron diffraction are few but significant. Sufficiently 

intense neutron beams are available only from nuclear fission reactors or from 

spallation reactions in the target of a charged particle accelerator; neutron 

beams and neutron spectrometers are both costly and few in number throughout 

the world. With the NRU reactor at Chalk River and its associated neutron 

spectrometers, we have equipment in Canada that is as good as any in the world 

for these kinds of experiments. 

The paper is arranged as follows. The basis of the neutron technique is 

described in section II, while section III contains two examples of recent 

neutron experiments at Chalk River, viz. the measurement of residual strain in 

overrolled Zr 2.5 wt% Nb pressure tubes and in bent Incoloy-800 steam generator 

tube. A number of other applications are mentioned. 

2. NEUTRON DIFFRACTION TECHNIQUES 

A typical neutron spectrometer mounted at a fission reactor can select a beam 

of neutrons with any desired wavelength in the range » 1-5 Â, that can be 

directed onto the samples. Figure 1 shows such a spectrometer mounted at the 

L3 beam facility of the NRU reactor, Chalk River. Typically about 10 neutrons 

cm" s~ impinge on the sample, and the intensity of powder diffraction is of 

the order 10-100 counts/sec when the diffracting volume is of order one cm . 

The spectrometer, because of stringent shielding requirements, weighs several 

tons and can accommodate samples of up to about a ton on the sample table. The 

maximum cross-sectional area of the neutron beam is 50 mm * 50 mm; smaller 

sizes are obtained by suitable collimators that cut down the beam with absorbing 

masks. The neutron beam size in any experiment is chosen to match the 

spatial scale of the feature under investigation. For example in the experiment 

on Incoloy-800 steam generator tubing there were marked changes in strain 

in one dimension over a distance of order 1 mm, so the masks were chosen to be 

1 mm wide and 10 mm long. The angular resolution is easily altered to suit the 

demands of the measurement, by appropriate choice of wavelength and collimation 

of the incident and diffracted beams. For crystallographic texture measurements, 

low resolution is required to integrate over the appropriate Bragg 

peaks; on the other hand, the highest available resolution, of order 0.01% is 

often desirable for residual strain measurements. Typically the time taken to 

characterize the texture for a particular set of crystallographic planes is 

about a day, whereas the determination of the angle of diffraction 29j,kji in a - 

strain measurement takes a few hours. The interplanar spacings of the crystal 

lattice constitute miniature internal strain gauges. The angle of diffraction 

is related to the corresponding spacing, dhkJi, by Bragg's law. 

X = 2

peaks which are easily measurable. 

3. APPLICATIONS 

- 389 - 

3.1 Lattice strain measurements in overrolled cold-worked Zr 2.5 wt% Nb 

pressure tubes 

Several measurements of lattice strains in model engineering components by neutron 

diffraction have recently been reported. [1,2]. Our experiments were 

initiated in response to a pressing engineering problem: the understanding of 

the strains that developed in overrolled présure tubes in CANDU reactors in 

Ontario. In a CANDU reactor, the seal between the zirconium alloy tube and its 

steel end-fitting is a rolled joint, made by inserting a rolling tool and 

expanding the tube into the end-fitting. The mouth of the fitting was slightly 

tapered for easy location and insertion of the tube in the fitting. In many 

cases the rolling tool was inserted too far, so that the tube was expanded into 

the taper, giving rise to a large circumferential residual tensile stress [3]. 

Under operating conditions, hydrogen diffused to the regions of high tensile 

stress. Platelets of zirconium hydride were formed perpendicular to the 

residual tension, and cracking eventually occurred. We set out to map the 

lattice strain in the overrolled region with the neutron diffraction technique. 

Figure 2 shows measurements of lattice strain in the hoop direction in a complete 

overrolled tube as a function of distance from a point in the tube within 

the end fitting (0 mm) through the point of maximum Insertion of the rolling 

tool (17.5 mm) to a region characteristic of the production tube. The lattice 

strain is determined with respect to the interplanar spacing at the free 

"unstrained" end and corresponds to the incremental elastic strain induced by 

rolling. Well within the end-fitting, where the tool presses the tube against 

the fitting the lattice is compressed, as one would expect; however, in the 

unsupported region there is a sharp increase In the (0002) interplanar spacing 

corresponding to the hoop direction. This tensile strain extends beyond the 

insertion point of the tool. A coupon was subsequently cut from the complete 

overrolled tube for X-ray line broadening measurements and Fig. 2 also shows 

the lattice strain In the hoop direction of this coupon. Although the macroscopic 

hoop stresses no longer exist, as much as 50% of the original lattice 

strain Is still present In the coupon. We believe that this phenomenon arises 

from grain-to-grain interactions. One conventional method of measuring 

residual stress involves observing the change in strain gauge readings when 

such a coupon Is cut from a tube, and assuming that no stresses remain in the 

coupon. This approach would clearly fail to reveal the effects we have 

observed with neutron diffraction. Also shown In Fig. 2 is the lattice strain 

profile for a complete tube that was given a stress-relief heat treatment after 

the rolled joint was made. 

Measurements of the anlsotropy of the residual strain were made with this 

coupon and some of these are shown in Fig. 3 for the plane in the coupon

- 390 - 

containing the tube radius (X^) and the hoop direction (X3). At the place in 

the coupon where the (0002) plane spacing in the hoop direction shows residual 

tensile strain, the *(2ffO) plane spacing in the radial direction shows a 

compressive strain. The third major direction X2 (axial, corresponding to 

(1010)) shows a small residual tensile strain- Analysis of 19 interplanar 

spacings in three orthogonal planes X1X2, X2X3 and X3X1 gives the complete 

strain tensor including sizeable shear terms. 

To make this approximate one-to-one correspondence between plane normals and 

direction in the sample, we have made use of the marked texture of a production 

pressure tube. The vast majority of grains have their [01Î02 axes along the 

tube axis, the [0002] axes along a hoop direction and the [2110] axes along a 

radial direction. This is the first time such a detailed mapping of the strain 

has been carried out and it should lead to a realistic model of the distortion 

caused by overrolling. 

3.2 Residual stresses in Incoloy-800 steam generator tubes 

The steam generator tubes in CANDU nuclear generating stations are formed into 

hairpins. Residual stresses in a given tube after it has been bent to its 

design shape are undesirable because they make the tube prone to stresscorrosion 

cracking when a concentration of salts in areas of localized dryout 

coincides with high tensile stress. 

The spacings of the {ill} and {200} planes in the longitudinal direction of the 

tube were measured as a function of position around the circumference. The 

"relaxed" value of plane spacing was measured on a piece of straight tubing. 

The strains derived from {220}, {200} and {ill} planes in the radial and hoop 

directions were also measured as a function of circumferential position. 

Figure 4 shows a schematic view of the experimental arrangement in which the 

neutron beam beam passes first through the near wall of the tube but the 

incident and scattered beams only intersect in the far wall which in this case 

is oriented so that its axial direction lies along the scattering vector. The 

neutron counter is shielded so that it cannot see the scattering from the near 

wall. This kind of arrangement is very useful for studying tubes nondestructively. 

The results are shown in Fig. 5. The longitudinal (or axial) 

lattice strain shows a maximum just above the neutral plane of the bent tube 

and a minimum just below it. The maximum longitudinal strain derived from 

{200} planes is almost twice that derived from {ill} planes. The magnitude of 

the strains within the wall came as a surprise, since conventional X-ray 

methods [4] had indicated lower strains. The fact that the strain in the 

longitudinal direction could be different in differently aligned grains of a 

cubic material was also a surprise. Model calculations for a thin walled tube 

with elastic/plastic bending [5], following by elastic unloading, confirm the 

shape of the residual strain curve. They suggest, moreover, that the anisotropy 

of Young's modulus in Incoloy accounts for the difference between the {ill} 

and {200} strains. The Incoloy tubing is textured such that there are 5 times 

more {ill} planes oriented along the tube axis than {002} planes. The {ill} 

behaviour, showing a balanced bending moment, is more representative of the 

tube. However, the texture is not so strong that we can determine the complete 

strain tensor. The variation of lattice strain in both the hoop and radial

- 391 - 

directions around the circumference is the reverse of the longitudinal strain 

and about 25% of its value. It appears that the relationship between the 

strains, 

given by 

EL, ER, £H in the longitudinal, radial and hoop directions is 

£ R 

V V 

where v is Poisson's ratio, which for Incoloy-800 lies between 0.28 and 0.34 

[6] . If we assume that the strains are uniaxial the residual stress may be 

obtained by making use of Young's moduli derived from X--ray measurements. The 

residual stresses measured with neutrons are roughly 1.5 times those derived 

from X-ray diffraction on the same tube. This may indicate major gradients 

through the wall or surface relaxation effects. It is clear from these 

measurements that neutrons are an excellent diagnostic tool for measuring the 

residual elastic strain in engineering components, capable of giving Important 

insights Into the mechanical behaviour of deformed metals. 

3.3 Texture measurements 

Measurements of crystallographic texture by neutron diffraction are not particularly 

new. For a recent review, see Bunge [7] . However, they have major 

advantages since the number of grains sampled by a neutron beam Is much larger 

than with typical X-ray beams. There are few geometrical constraints for most 

materials and it is often possible to place an entire component in the neutron 

beam. 

3.4 Measurement of second phases in materials 

The existence of second phases often controls the strength and fracture 

behaviour of materials. The sensitivity of neutrons to second phase material 

is often higher than X-rays partly because the signal is an average over many 

grains, and partly because the light element constituents of many precipitates, 

such as carbon, nitrogen and oxygen, have scattering cross sections that are 

just as large as most metals or other materials of interest to the engineer. 

Detection of volume fractions less than 1% is fairly routine. As an example of 

a recent application of neutron diffraction, Windsor et al. [8] have studied 

the development of fee second phase material in bec maraging steels. 

3.5 Small angle scattering of neutrons 

Small angle scattering relies on an effective contrast in scattering between 

the matrix and the embedded "small particles" that are the objects of interest. 

The sizes of small particles in the range 10-1000 Â may be readily determined; 

they may be lattice distortions around impurity atoms, nucleating phases barely 

detectable by other means, small precipitates etc. There is already a large 

literature on the technological applications of small angle scattering (see for 

example ref. [9]).

4. THE FUTURE 

- 392 - 

The neutron techniques are straightforward and well understood. The 

penetrating nature of neutron beams allows examination of real components from 

all angles, and the scale of neutron equipment is commensurate with the size 

and weight of large components such as pipeline and rail steel. Our first 

experiments have been on topics of interest to the nuclear industry, but the 

extremely detailed information that we have obtained suggest a much wider 

domain of usefulness in the future. 

5. REFERENCES 

[1] L. Pintschovius, V. Jung, E. Macherauch and 0. Vöhringer, Materials 

Science and Engineering, jjl_, 43 (1983). 

[2] M.J. Schmank and A.B. Kravitz, Met. Trans. A., 134, 1069 (1982). 

[3] E.C.W. Perryman, Nucl. Energy, J7, 95 (1978). 

[4] J. Winegar, AECL Report 6961 (1980). 

[5] S.P. Timoshenko and J.M. Gere, Mechanics of Materials, New York, Van 

Nostrand-Reinhold, p. 113 (1970). 

[6] R.A. Holt, G. Dolling, B.M. Powell, T.M. Holden and J.E. Winegar. Proceedings 

of the 5th International Symposium on Metallurgy and Materials 

Science, Riso, Denmark, 3-7 Sept. 1984 (in press). 

[7] H.-J. Bunge, Texture Analysis in Materials Science, London, 

Butterworths, 1982. 

[8] CG. Windsor, R.N. Sinclair, S. Faulkner, V.S. Rainey and G.F. Slattery. 

Proceedings of the 5th International Symposium on Metallargy and 

Materials Science, Riso, Denmark, 3-7 Sept. 1984 (in press). 

[9] W. Schmatz, in The Neutron and Its Applications, ed. P. Schofield, The 

Institute of Physics (London, 1982) , p. 301; these conference proceedings 

contain several other papers dealing with special applications of 

small angle neutron scattering.

I.-3 TRIPLE-AXIS NEUTRON SPECTROMETER 

LIQUID NITROGEN COOLED 

• CALANDRIA 

VNEUTRON FILTER .-AUXILIARY GATE 

RREACTOR HOLE GATE R QUARTZ 

1 

r LIQUID NITROGEN 

•• 

-COLLIMATOR K 

RMONOCHR0MAT0R 

\ 

CRYSTAL 

J COLLIMATOR ; 

r-SHIELDING DRUM 

!! r MONITOR 

SHIELDING YOKE 

NEUTRON BEAM 

Fig. 1 Neutron spectrometer at the NRU reactor, Chalk River. 

NEUTRON DFTECTOR 

' SPECIMEN 

I

I 2 

- 394 - 

AXIAL DISTANCE (CM) 

OVER-ROLLED PRESSURE TUBE 

STRESS RELIEVED TUBE 

COUPON FROM PRESSURE TUBE 

' "**.. 

O •- 

Fig. 2 Measured lattice strain in the hoop direction derived from the 

interplanar spacing of (0002) planes in Zr 2.5 wt% Nb pressure tubes. 

The dashed line indicates the result for the over-rolled pressure 

tube, the solid line indicates the result for the coupon cut from the 

tube, and the dotted curve indicates the result for the stress 

relieved tube.

Ad hki£ 

d hkil 

xlO 3 

- 395 - 

INTERNAL STRAIN IN Zr"2.5wt%Nb PRESSURE TUBE 

X( X3 plane 

3 4 

AXIAL DISTANCE (CM) 

0002 

SECOND GRAIN 

Fig. 3 Measured anisotropy of lattice strain in the coupon cut from the 

overrolled Zr 2.5 wt% Nb pressure tube. A tensile strain is observed 

in the hoop direction X3, (0002) planes, and a compressive strain in 

the radial direction Xj, (2ÏÏ0) planes. A very small compressive 

strain is observed for intermediate directions corresponding to (2ÏÏ2) 

and other related planes.

IRRADIATED 

VOLUME 

- 396 - 

BEAM APERTURE 

ANGLE OF 

SCATTERING 

20 

Fig. 4 Arrangement of steam generator tube in the neutron beam for diffraction 

from (200) planes in one wall of the tube, directed along the 

bisector of the incident and scattered beams. 

Cd

- 397 - 

Fig. 5 Longitudinal, tangential (hoop) and radial strains measured around the 

circumference (to - 0 corresponds to the neutral plane) of an 

Incoloy-800 steam generator tube at the apex of the bend. The strains 

in differently oriented grains but directed along the same axis in the 

tube are markedly different, showing the effects of the elastic 

anisotropy. 

120

- 398 - 

NDE IN POLYMERS, AN EXAMPLE: 

ULTRASONICS FOR THE DETERMINATION OF DENSITY IN POLYETHYLENE 

L. Vichu and A. Hame.1 

NatZonai. Re.i zafich* Co an.c-lt o & Canada 

Bouche.ivA.Zle., Quebec 

Abstract 

An ultrasonic immersion method is presented for the rapid and accurate 

measurement of longitudinal sound velocity (v) in polyethylene (PE). The 

method is based on the sole measurement of time delays and can be automated. 

The samples are the same press molded plaques that are prepared for the 

standard density (p) determination using a density gradient column. Our 

measurements on a representative group of various types of PE with densities 

in the range .92 to .96 g/cm show without ambiguity that velocity is 

proportional to density. The degree of correlation is such that we come to 

suggest that the acoustic method be used for the characterization of density. 

We show that the accuracy is good and compares with that of the density 

gradient column and then we conclude on the practical advantages of the 

ultrasonic method.

I - INTRODUCTION 

- 399 - 

Polyethylene, be it low density (LDPE) or high density (HDPE) is one of the 

most important, if not the most important product of the plastics industry. 

As it is the case for all industrial materials, it is of primary importance 

for the manufacturer and the processor to- have means of characterizing the 

material. 

From a somewhat simplistic point of view, polyethylene can be considered as a 

binary mixture of an amorphous and a crystal phase and the degree of 

crystallinity will affect the elastic properties (elastic modulus, impact 

strength and melting point). On the other hand, it is known (1) that density 

is directly related to the degree of crystallinity: so that in the 

relationship between the basic parameters of PE and its end properties, the 

value of density is an important link. This explains why the value of density 

has always been the major criterion for the characterization of polyethylene. 

The most commonly used and accepted method for the determination of density is 

that of the "Density Gradient Column" defined by ASTM D1505 standard (2). The 

sample is placed in a column of liquid which exhibits a known density gradient 

and after it has sunk to its equilibrium, the density is obtained from the 

reading of its position in the column. 

As such, the procedure appears to be simple but it is in fact quite involved. 

First, the column must be prepared from a mixture of different solutions of 

known densities. This is time consuming, requires much care at all stages and 

is much an art. The column which is placed in a constant temperature 

environment is then calibrated, using standard glass floats (the calibration 

has to be verified periodically). To perform the measurement, the samples are 

cut to size, cleaned, wetted down and gently introduced in the column where 

they settle in roughly 20 minutes. When, after a number of measurements, the 

number of samples in the column is too important, these are retrieved 

carefully with a basket, without upsetting the gradient. 

Here we propose a method and technique based on the propagation of ultrasound 

which allows for an accurate and reliable determination of density. The 

technique is simple and can be used in an uncontrolled environment; the 

measurement is quick and can be made automatic. 

II - EXPERIMENTAL METHOD 

For a number of polymers, the ultrasonic data are available from the 

literature (3); sparse results are found for the sound velocity and density 

(4,5). The special case of polyethylene has, however, been given closer 

attention (6,7,8,9) and the various results tend to show that there exists a 

correlation between sound velocity (v) and density (p). 

The most complete and conclusive study is that of Davidse et al (7). They 

have measured, using a resonance method, the propagation velocity of low 

frequency oscillations (5 to 10 kHz) in bars of PE and their results on a

- 400 - 

great number of samples show that the velocity is proportional to the 

density. The technique they have used is involved: it requires preparation of 

samples and the determination of velocity takes into account geometrical 

effects and a priori knowledge of the Poisson ratio for the material. 

It is our objective here to verify the existence of a possible correlation 

between the velocity and the density, to describe the correlation and to 

propose a technique by which the velocity can be obtained in a rapid yet 

precise way so as to allow for the determination of density. 

In our study, we use an immersion technique (4,10) because of the ease with 

which the measurements can be made and repeated. The principle of the method 

where a single transducer is used as both the emitter and the receiver is 

illustrated in Fig. 1. A high frequency (MHz) sound pulse travels with a 

normal incidence onto the sample (a plate of thickness d). There it sees part 

of its energy reflected back (E^) and reaching the transducer at time t\, and 

part of it going through the sample where it later meets the second interface: 

again part of the energy is reflected, part is transmitted and so on. This 

gives rise to the formation of échos (Ej, E2 ...) which are seen at times tj + 

T, ti + 2T...(T is the time for the sound to travel back and forth in the 

sample). If v is the velocity of sound in the sample, then 

T = 2d/v 

In order to obtain v, one needs to know the value of d with sufficient 

precision so as to obtain the required accuracy. This measurement could be 

made using a micrometer but this would be both annoying and not very 

accurate. The samples being rough from press show surface and thickness 

irregularities so that the value of d must be some average taken at the 

location of insonification. We have chosen to take the thickness into 

account through acoustics. Part of the energy after having traversed the 

sample is transmitted forward in the liquid. This signal can be reflected 

back to the transducer by use of an acoustic reflector M. If t'i is the 

time it takes the sound pulse to go from the sample to the mirror and back, 

a signal will be detectecd at time 

t'=ti+T+t'i=ti+ 2d/v + t'i 

The sample being removed, the signal coming from the mirror appears at time 

t" = t! + 2d/c + t 1 ! 

where c is the velocity of sound in the liquid. 

Considering the difference in the times of flight yields 

t' - t" = At = 2d (1/c - 1/v)

- 401 - 

This together with the expression of T gives 

v = c (At/T + 1) 

Thus, from the sole measurement of time delays, and the knowledge of c, one 

obtains the value of the velocity of sound in the sample. 

Ill - EXPERIMENTAL PROCEDURE AND RESULTS 

The samples (35) corresponding to a broad range of densities were furnished 

by DuPont Canada and Union Carbide Canada Ltd. They were in the form of press 

molded plaques (80 X 50 X 1.9 mm) prepared according to ASTM standards from a 

variety of commercial grades of P.E. In the measurements, they were used as 

is with no preparation. A number of pieces (2 to 4) were cut from the plaques 

and their density measured in a density gradient column following the standard 

ASTM procedure. The values are given with an accuracy of +_ .0007 g/cm at a 

constant temperature t = 23°C. 

The ultrasonic set up was designed to accomodate such samples and placed in 

a bath of demineralized water at 23°C. The transducer is of the 

commercially available immersion type. In order for the echos to be well 

separated in time, the pulses must be short: this calls for a highly 

damped transducer operating at high frequency. However, the attenuation of 

sound in P.E. rapidly increases with frequency and this sets an upper limit. 

We found that 3 MHz constitutes a good compromise between time resolution and 

signal amplitude. The reflector, located roughly 15 cm from the transducer, 

is a flat piece of glass with an absorbant backing so that no signals (echos) 

are issued other that the reflexion from the front face. The set up is rather 

simple and the only requirement is that the samples and reflector be 

perpendicular to the sound beam. 

Concerning the electronics, we used a Metrotec puiser (Mod. MP 203) capable of 

delivering short (100 n sec) high voltage (300 volts) pulses with a repetition 

rate or 50 to 100 Hz. The receiving amplifier was also Metrotec (Mod. MR 101) 

wiht a 60 dB gain factor and a 20 MHz bandwidth. The RF signal is monitored 

on an oscilloscope (H.P. Mod. 1743A) which is equipped with a dual time base 

that allows time delay measurements with an adequate resolution (3 n sec or 

better). 

The sound velocity in the water has to measured: this is accomplished by 

displacing the reflector by a known distance and noting the arrival times. We 

obtain a value of c = (1.483 +_ 0.001) X 10 5 cm/sec at 23°C, in good agreement 

with what is found in the literature from which we quote the temperature 

dependence: 200 cm/sec°C. 

The results for our measurements of sound velocity (at 3 MHz) in 

polyethylene samples of different densities are illustrated in Fig. 2. To 

a good approximation, the velocity is proportional to density.

- 402 - 

Since the velocity measured is actually a value taken over the whole region of 

insonification, we looked at the effect of changing the acoustic beam 

diameter. Our observations using transducers of increasingly larger diameters 

(from 0.2 cm for à focused beam up to 4 cm) show that the scatter diminishes. 

Increasing the diameter amounts to measuring a value that is averaged over a 

wider area. In order to appraise the scatter, we have measured the velocity 

at different locations on the same plaques. The results are illustrated in 

Fig. 3 for a 1.25 cm (i in) diameter transducer. The standard deviation is 

found to be (Sv/v)expe. = _+ 0.0018. However, if we repeat the measurement 

several times at the same location on the plaque, the standard deviation is 

(6v/v)repeat = _+ 0.0010. This tends to show that the scatter partly has its 

origin in non homogeneities found across the plaque. 

In turn, we can estimate the theoretical experimental uncertainty by 

considering the relation from which the velocity is calculated and the various 

instrumental errors. We obtain a value of (ôv/v)theo. = i 0.0012, which is 

to be compared with (

- 403 - 

(5v/v)fit must be considered as more realistic since it integrates errors 

from all sources. 

From a statistical point of view, the fit can be considered as very good and 

this leads us to the conclusion that for PE having a density in the range of 

0.92 to 0.97 g/cm , the room temperature (23°C), sound velocity is 

proportional to density. We may now examine the possibility of using the 

measurement of SDund velocity for the determination of density in PE with a 

model such as p = Av + B. Working the regression analysis with p as the 

dependent variable we obtain A = (0.0965 +_ 0.0012)10~ ? g sec/cm 4 and B = 

(0.07218 +_ 0.0026) g/cm . The correlation factor is of course the same as 

above (r = 0.998) and the standard error of estimate becomes 0.00098 g/cm 

yielding the relative average value (ôp/p) = _+ 0.0010. As mentioned above 

this integrates errors from all sources, and yet it compares quite well with 

what can be expected when using a density gradient column for which 

(6p/p)co^# = _+ 0.0007. Given this result and bearing in mind that the small 

difference most probably originates in the additive content, we believe that 

the ultrasonic approach will allow the determination of density in 

polyethylene with the required accuracy for quality control purposes. 

IV - CONCLUSION 

We have demonstrated that in commercial grades of polyethylene, there exists a 

linear relationship between the sound velocity at ultrasonic frequencies and 

the density of the material. We have shown experimentally that the velocity 

measurements can be performed with sufficient accuracy to allow the 

computation of density. The method we describe is simple and since it is 

based on the sole measurement of time delays, it can be made automatic through 

digitalization. The advantages of the method are numerous when compared to 

the classical density gradient column: a) no preliminary preparation, the 

setup is completely portable, suffers little influence from the environment 

and requires no special care to upkeep, b) the sample is not destructed and 

can be kept for files, c) the measurement is fast, reliable and repeatable, 

the accuracy is good (ôp = +^ 0.001 g/cm ) and the operation being simple 

requires no special attention, care nor skill. 

Further work is needed to check whether such a technique could be applicable 

to other polymers and also to examine the influence of fillers (silica, carbon 

black...). Further work is also needed in order to establish the physical 

foundation for the observed behaviour. 

ACKNOWLEDGMENT 

The author wishes to acknowledge the gracious participation of Dr. P. Kelly 

(DuPont Canada Inc, Kingston, Ontario) and Mr. I.R. Lambie (Union Carbide, 

Sarnia, Ontario), who provided both the incentives and the samples for this 

work.

REFERENCES 

- 404 - 

1. S. Kavesh, J.M. Shultz; Polymer eng. Sei. _9> 331 (1969) 

2. Annual book of ASTM Standards 

3. For a revue, see: Hartmann B., Methods of Experimental Physics, 

vol. 16c; R.A. Fava Ed.; Academic Press Inc., 1980. 

4. B. Hartmann, J. Jarzianski, J. Acous. Soc. Am _56_, 1469 (1974). 

5. Biing-Nan Hung, Goldstein A., IEEE Trans-sonics, Ultrasonics Su3£, 4 

(1983). 

6. J. Schuyer; Jour. Poly. Seien., J36_, 475 (1959). 

7. P.D. Davidse, H.I. Watermann, J.B. Westerdijk; Jour. Poiy. Seien. _59, 

389 (1962). 

8 A. Levene, W.J. Pullen, J. Roberts; Journ. Poly. Seien., A3, 687 (1965). 

9. Adachi K., Harrison G., Lamb J., North A.M., Pethrick R.A,: Polymer _22, 

1032 (1981). 

10 G.W. Paddison, Ultrasonics Symposium, IEEE, p. 502, (1979). 

Patent Pending

Transducer 

Signal 

- 405 - 

P.E.(p.v) 

M 

Immersion Liquid: Water, (c) 

T 

Reflector 

t" time 

Figure 1: a) Schematic description of the path followed by the sound beam as 

it travels through the water and interacts with the sample; b) 

representation of the different signals which are to be observed on 

an oscilloscope.

sec.) 

2.5 

F 2.4 

b 

n O 

Î2-3 

S CO 

f 2.2 

Longi 

— 

— 

I I 

2.1 - •/ 

- 406 - 

I I I 

2 

.91 

/ 

i 

.92 

i 

.93 .94 .95 

Density (g/cm 3 ) 

Vv 

/ ; 

I I I 

— 

.96 .97 

Figure 2: Results obtained for the measurement of ultrasonic (3 MHz) sound 

velocity in plaque samples of commercial grade polyethylene, the 

density of which was measured in a density gradient column at 23°C.

ü 

.CO 

ü 

m O 

o 

CÖ 

g 

•4—> 

Q 

2.1 

2.09 

2.08 

2.07 

±.18% ±.12% 

- t 

1 

0 5 10 

Measurement N°. 

Figure 3: Standard deviation as estimated (6v/v)theo. = .12% and observed 

(6v/v)expe = .18% from sampling at 16 different locations on the 

same plaque. 

1 

o 

5 o 

o 

o 

0 

0 

0 

0 

15 

o 

o 

o 

o 

4 o 

0 

o" 

16 o 

o 

I

- 408 - 

AN INDUSTRIAL APPLICATION OF COMPUTER ASSISTED TOMOGRAPHY: 

DETECTION, LOCATION AND SIZING OF SHRINK CAVITIES IN VALVE CASTINGS 

P.V. Tonne* and G. Toitllo 

Atomic Enz-igy o & Canada 

Chalk R-ti/e*, OntaKlo 

ABSTRACT 

Computer assisted tomography (CAT) scanning is a nondestructive testing technique 

used to obtain quantitatively accurate mappings of the distribution of 

linear attenuation coefficients inside an object. To demonstrate the potential 

of the technique for accurately locating defects in three dimensions a sectioned 

5 cm gate valve, with a shrink cavity made visible by the sectioning, was 

tomographically imaged using a Co-60 source. The tomographic images revealed a 

larger cavity below the sectioned surface. The position of this cavity was 

located with an in-plane and axial precision of approximately ± 1 mm. The 

volume of the cavity was estimated to be approximately 40 mm 3 . 


Computerized Tomography (CT) is a nondestructive testing technique used to 

obtain quantitatively accurate mappings of the distribution of linear attenuation 

coefficients inside an object. One possible application of CT is to reinspect 

items that have failed to conform to standard QA acceptance tests in 

order to characterize the defect in greater detail. Since CT images contain a 

wealth of distortion free spatial information, CT re-inspection may be especially 

useful if the non-conforming item is to be repaired or upgraded. The 

present study demonstrates the use of CT re-inspection to detect, locate and 

size shrink cavities in valve castings earmarked for upgrading. 

Radiographs of valve castings are routinely used for QA, with the choice of 

acceptance standards depending on the intended end use of the valve. If an 

unacceptable shrink cavity is indicated on the film the casting may be upgraded 

by gouging out the cavity and backfilling the entire area with weld material. 

Because of the complex shapes involved, it is sometimes difficult to accurately 

locate the cavity in three dimensions by referring to radiographs alone. In 

some cases this had led to the removal and backfilling of considerably more 

material than necessary. In other cases the cavity has not been adequately 

repaired on the first attempt and, hence, it has been necessary to upgrade the 

casting a second time.

- 409 - 

The present study was carried out to demonstrate how CT images can be used to 

eliminate such costly upgrading errors. Because tomography is not presently in 

widespread use as an industrial NDT technique, the theory of tomography is 

briefly reviewed with special emphasis on the differences between conventional 

radiography and CT imaging. This is followed by a description of the general 

purpose laboratory scanner used to perform the study. Finally, results obtained 

with a test valve casting are presented and analyzed. 

2. THEORY OF TOMOGRAPHIC IMAGING 

As indicated in Figure l(a), the attenuation of a monoenergetic collimated beam 

of gamma rays as it passes through a uniform slab of material is given by 1 

where 

-uL 

I e 

o 

Io is the initial intensity of the gamma ray beam, 

I is the intensity after passing through the material, 

u is the linear attenuation coefficient of the material (cm" 1 ), and 

L is the path length in the material (cm). 

Equation (1) can be solved for the product (J.L: 

(1) 

In Io/I = nL (2) 

If the ray passes through a non-homogeneous material the path can be considered 

to consist of a number of elements of width w with attenuation coefficients (j-i » 

H2» *• *|J-n» as indicated in Figure l(b). 

Equation (2) then becomes 

In Io/I = + (3) 

That is, the logarithm of the measured transmission along a particular ray 

represents the sum of the attenuation coefficients of all the elements that the 

ray traverses. In tomography, the quantity In Io/I is normally called a ray 

sum. 

A single ray sum cannot, by itself, give any indication of the distribution of 

attenuation coefficients in the material. For example, the homogeneous sample 

illustrated in Figure l(c) has precisely the same ray sum as the nonhomogeneous 

sample shown below it. 

While it is not possible to use a single ray sum to learn the distribution of 

attenuation coefficients in an object it is intuitively obvious that ray sums 

obtained from other directions will provide useful information about that distribution. 

Multiple radiographs of a single object obtained from many viewing 

angles around the object are sometimes obtained for that very reason: i.e. to

- 410 - 

allow the radiographer to visualize the relative location of different materials 

inside the object. In tomography then, one obtains sets of ray sums at 

many different angles about the object being tested. A mathematical algorithm 

is then used to reconstruct the unique distribution of attenuation coefficients 

within the object that gave rise to the experimentally measured ray sums. 

In its simplest form, a tomographic scanner is designed to obtain sets of 

parallel ray sums in a single plane from many directions around the object. A 

translate-rotate tomographic scanner is schematically illustrated in Figure 2. 

It consists of a collimated source of radiation, either an isotopic source or, 

as shown here, an x-ray tube, and a collimated detector. Each parallel line in 

Figure 2(a) represents one ray sum. One set of parallel ray sums, obtained by 

translating the source and detector relative to the object (or vice versa) is 

called a projection. Projections from many directions around the object are 

obtained by rotating the source and detector relative to the object (or vice 

versa) and repeating the translation procedure at each new orientation. Figure 

2(b) illustrates a typical tomographic scan pattern consisting of translations 

at 1 degree increments over 180 degrees. For clarity only the 1st, 60th and 

120th translations are shown. More sophisticated scanners employ arrays of 

detectors to speed up data acquisition and make better use of the available 

photon flux. 

The rules of thumb normally applied in collecting tomographic data on a translate-rotate 

scanner (the type used in this study) are: 

1. The total number of projections, m, should be greater than or equal 

to nTc/4, where n is the number of ray sums per projection. 

2. When formed with identical source and detector collimator apertures, 

the effective ray width (i.e. the full-width at half-maximum) of the 

beam midway between source and detector is approximately half the 

width of the collimator aperture. 

3. The ray sura sample spacing should be less than or equal to half the 

effective ray width (Nyquist criterion). 

Because of the large amounts of data involved, all practical tomographic scanners 

are computer controlled. Hence, the acronyms CT for computed tomography 

or CAT for computer assisted tomography, are generally used to describe the 

technology. A computer is also required to reconstruct the tomographic image 

from the measured ray sums. The reconstruction technique most often used, and 

that used in this study, is termed filtered back projection. The details of 

this technique are described in reference 1. 

Tomographic images are, by their nature, digital images. They are constructed 

on a grid of picture elements or pixels typicaily 64 x 64 to 1024 x 1024 in 

dimension. The width of each pixel is normally chosen equal to the ray spacing. 

Each pixel is assigned a value which is proportional to the attenuation 

coefficient, JJ., at the corresponding location in the object.

- 411 - 

Since tomographlc Images are reconstructed from gamma-ray intensity measurements 

which are subject to statistical variation, the reconstructed |j.'s will 

also have an associated statistical variance a^ 2 . The standard deviation, 

ff„, is the precision with which a given |a can be determined from a single 

pixel in a tomographic image. 

It can be shown [2,3,4] that an average a^ 2 can be estimated using 

where 

o 2 = î_ï 

11 3 

12At N 

tot (4) 

Ntot is the total number of photons passing through the object, 

D is the average diameter of the object (cm), and 

At is the width of a pixel (cm). 

Equation (4) is only exactly true if the number of counts in all ray sums passing 

through the object are equal. The equation is, however, approximately 

correct even when this situation does not exist. As can be seen, UM, At and 

Ntot are related such that, with Ntot constant, an increase in At by 

a factor of 2 results in an eight-fold improvement (decrease) in o^ 2 . 

3. TOMOGRAPHIC SCANNING APPARATUS 

The tomographic scanner used in this study is illustrated schematically in 

Figure 3. It comprises a control and data acquisition computer, scintillation 

detector with associated electronics, a radioactive source, source and detector 

collimators and shielding and a computer controlled motorized assembly for 

translating and rotating the test object. 

The computer system consists of a DEC LSI-11/23 microcomputer with 256 k bytes 

of on board memory, a 20.8 M byte (Winchester) disk and a ^ M byte floppy 

disk. 

In the present study an Nal scintillator, hermetically sealed to a photo multiplier 

tube, was used as the detector. The detector preamp and single channel 

analyzer have been customized to enable measurement of pulse rates of up to 

50 kHz without software correction for pile up. With software correction this 

limit can be extended by more than an order of magnitude. Detector output 

pulses are counted in a standard counter/timer and interfaced to the DEC Q-bus 

through the Canberra 1788 and DEC DLV11-J serial interfaces. 

The test object is placed on a turntable mounted on a translatable carriage. 

The turntable and carriage are driven by individual stepper motors and their 

absolute positions are continuously monitored by shaft angle encoders. The 

stepper motors and encoders are interfaced to the bus through ADAC 1600 series 

interface cards.

- 412 - 

The photon source may be Cs-137, Ir-192 or Co-60, depending on the requirements. 

A Co-60 source with a useful penetration of up to 24 cm of steel was 

used in the present study. 

The source and detector collimators are Identical tungsten-alloy cylinders, 

165 mm in length, mounted in V blocks. The apertures are continuously variable 

in width from 0 to 6.4 mm and are available in three discrete heights; 1.59 mm, 

3.175 mm and 6.35 mm. 

Tomographie reconstruction is executed on the LSI-11/23 with the assistance of 

the Computer.Design and Applications (CDA) MSP-3 array processor. Typical 

reconstruction time is about 90 seconds for a 128 x 128 reconstruction grid. 

After reconstruction, the images are displayed on a colour graphics screen 

using the CDA MDP-3B graphics processor. Images can be photographed off the 

screen for further study and can be archived in digital form on disk or magnetic 

tape. 

4. CT DEMONSTRATION 

4.1 Description of the Sample Valve Casting and Defect 

A class 600 flanged-end gate valve casting with a 5.08 cm (2") bore was examined 

in the present study. The casting had been bisected along a plane passing 

through the bore and the stem port axes (see Figure 4). As indicated, a shrink 

cavity in one of the crotch areas is exposed at the sectioned face. Shrink 

cavities generally represent a tear rather than a smooth internal void and 

therefore they may propagate, like a crack, under vibration. For this reason 

they are considered a serious defect. They are caused by the uneven solidification 

of liquid metal in the mold and often occur in regions where external 

and internal surface contours change rapidly. The crotch area of a gate valve 

is a typical example of this. 

The fact that shrink cavities are likely to occur near contoured surfaces presents 

special difficulties in the interpretation of radiographs. In particular, 

determining the coordinates of the defect in three dimensions from radiographs, 

as required for upgrading, is subject to error because of the changing 

film exposure due to variation in the path lengths through the irregularly 

shaped object. This difficulty is compounded by the ever present magnification 

distortion due to the divergence of the x-ray beam on its path from source to 

film. 

4.2 CT Scanning Procedure 

To demonstrate the effectiveness of CT in circumventing these problems inherent 

in conventional radiography, the valve casting was toraographed in the region of 

the shrink cavity. The casting was mounted in the scanner as shown in Figure 

4. In order to delineate the defect in three dimensions, six adjacent slices, 

each perpendicular to the axis of the bore and separated by 1.59 mm along the

- 413 - 

bore axis, were scanned. The general location of these slices is indicated by 

the scan plane illustrated in Figure 4. For ease of reference in the discussion 

that follows, the height of the scan plane above the bottommost surface of 

t_>e casting will be referred to as distance Z in the axial direction. The 

distance Z to the midpoint of the topmost slice was 109.5 mm and this was reduced 

for each of the six slices by slipping a succession of 1.59 mm (1/16") 

shims underneath the casting. 

The details of the scanner set-up are summarized in Table 1. The 4.4 x 

(12 curie) Co-60 source delivered an unattenuated gamma-ray exposure of approximately 

29.7 R/hr at the face of the Nal detector. The corresponding count 

rate, as measured using the test assembly counting electronics and corrected 

for pile-up, was approximately 7 x 10 counts/second. 

A modulation transfer function (MTF) test pattern was also included in each 

scan. As described in Appendix A, it was used to determine the system MTF. 

The 50% point on the system MTF corresponds to 2.1 line pairs/cm (see Figure 

A-2). 

4.3 Results 

Tomographie images of the six parallel slices are shown in Figure 5. 

A useful method of understanding these images is to imagine that the casting 

and MTF test piece have been cut through the scanning plane indicated in Figure 

4, and that the material above the cut has been removed. The tomographic 

images represent what an observer, standing over the source and facing in the 

direction of the detector, would see when looking down on the surface from 

above, except that the tomographic image is a map of linear attenuation 

coefficients, not a true photographic image. The series of images (a) through 

(f) in Figure 5 represents a succession of these surfaces revealed by a 

succession of cuts from top (Z = 109.5 mm) to bottom (Z = 101.6 mm). 

The images shown in Figure 5 are photographs taken from a CRT display. The CRT 

display consists of a 256 x 256 array of square picture elements or pixels too 

small for the human eye to discern. Each pixel represents a volume (voxel), 

0.85 mm on a side, and 1.59 mm high, in the object itself. Each pixel is 

assigned an integer number between 0 and 255 that is related to the linear 

attenuation coefficient in the corresponding voxel. Rather than attempt to 

display these integer numbers themselves, each number is assigned a unique 

shade of grey. In this particular assignment the largest integers (i.e. largest 

linear attenuation coefficients) are assigned the lightest shades of grey 

and the smallest integers are assigned the darkest. The operator is free to 

use any assignment of grey scale, or colour, he chooses. Certain choices may 

tend to enhance some features and suppress others. However, the data, i.e. 

the actual reconstructed attenuation coefficients, remain unaltered. 

It is immediately apparent from Figure 5 that the shrink cavity is not confined 

to the surface of the sectioned face of the valve but that a larger cavity is 

located somewhat below the surface. This subsurface cavity is particularly

- 414 - 

evident in Figures 5 (c) and (d) but is detectable in all the images. Also 

noted is a seccnd smaller subsurface cavity in Figure 5(a) that is not evident 

in the other images. 

To demonstrate the dimensional accuracy of the tomographic images, known dimensions, 

the lengths a,b,c and d defined in Figure 5(b), were compared with those 

determined from the tomographic reconstruction. The reconstructed dimensions 

were measured by counting pixels and converting this to mm using the conversion 

factor 1 pixel = 0.86 mm. Reconstructed and measured dimensions agreed to 

within 1 mm (see Table 2). 

The in plane coordinates of the centre of the main body of the subsurface 

shrink cavity can thus be determined in relation to a reference coordinate 

system. The location and orientation of one possible reference coordinate 

system is shown in Figure 5(c). The axes are oriented parallel and perpendicular 

to the straight edges of the flange, thereby providing useful directional 

references both before and during upgrading. With reference to this coordinate 

system, the centre of the main body of the subsurface shrink cavity lies 28 

±1 mm below the surface in the x direction and 72 ±1 mm from the edge of the 

flange in the y direction. 

The axial (Z) coordinate of the centre of the main subsurface shrink cavity is 

apparently somewhere between Z = 106.4 mm and Z = 104.8 mm, corresponding to 

Figure 5(c) and (d), respectively. The intermediate value Z = 106 ±1 mm is a 

reasonable compromise. 

The volume of the main body of the subsurface cavity may be estimated by assuming 

the cavity is approximately spherical in shape with a diameter of approximately 

6 mm, as estimated from Figure 5(c). On this basis the volume is estimated 

to be approximately 40 mm 3 . 

Tomography not only provides quantitatively accurate dimensional information 

but it can also be used to determine, again quantitatively, the linear attenuation 

coefficients of the materials in the image. The theoretical linear attenuation 

coefficents, (i^h' °^ brass, iron and air for 1.25 MeV gammas are compared, 

in Table 3, to the corresponding average coefficients determined from 

the reconstructed images, {Trec. The excellent agreement between nth, which 

has an associated uncertainty of about 5%, and |Trec is typical of tomographic 

imaging. 

The ]Irec of each material was obtained by averaging the [i's assigned to a 

number of pixels representing that material. Thus the uncertainty in the 

ÏÏreo a iï> is considerably smaller than the pixel-to-pixel standard deviation 

in M-reo represented by o^. In fact, op is a„//n where n is the number 

of f pixels il used d to determine d i Mrec* ^ i i Fi 5 i 

or tne i raa S es i- n Figure 5 cy is approxi- 

mately 0.011 cm" 1 . Since 120 pixels were use to determine free» a \î *- s 

0.011//Î2Ô or 0.001 cm" 1 , as quoted in Table 3. 

Expressed as a percentage of the n of iron, a^ is approximately 3%. It is 

important to note that this 3% is not equivalent to radiographie contrast

- 415 - 

normally measured with a shim type penetrameter. Radiographie penetrameter 

contrast refers to an integrated contrast along the line between source and 

film, taking into account all the material along this line, whereas o^ is the 

precision with which the n of a specific pixel anywhere within the interior of 

the object can be determined. If the difference between the reconstructed ^'s 

of two pixels is greater than 2a„ then there is a 95% chance that these two 

pixels represent different M-'s in the object. Since in tomographic images any 

two pixels or groups of pixels can be compared in this way, a^/y, is equivalent 

to a much smaller radiographie contrast, typically two orders of magnitude 

smaller. 

5. CONCLUSIONS AND DISCUSSION 

The tomographic scanning and reconstruction method has been used to detect and 

locate a subsurface shrink cavity in the crotch area of a sectioned 5.08 cm 

(2") valve casting. 

The main body of the cavity has been located in three dimensions with a precision 

in each dimension of approximately ±1 mm. The volume of the main body of 

the cavity has been estimated to be about 40 mm 3 . The precision with which the 

cavity can be located would be of considerable assistance in upgrading the 

valve. 

The ability to accurately reconstruct linear attenuation coefficients has been 

demonstrated. The standard deviation in a given reconstructed linear attenuation 

coefficient was determined to be approximately 0.011 cm"* or 3% of iron. 

This is typical of CT images and represents an improvement of one to two orders 

of magnitude over the contrast obtained with film radiography. 

Although the demonstration scans in this study each required approximately 14 

hours for data acquisition, the scan time can be greatly reduced by employing 

arrays of detectors rather than a single detector. Scan times of 4 seconds or 

less are typical of current generation medical CAT scanners. There is no fundamental 

reason why CAT scanners intended for industrial purposes cannot achieve 

similar scan times. 

6. REFERENCES 

1. L.M. Zaty, "Basic Principles of Computed Tomography Scanning", 

Radiology of the Skull and Brain, Vol. V: Technical Aspects of 

Computed Tomography, 1981. 

2. D.A. Chesler, R.J. Stephen and N.J. Pele, "Noise Due to Photon Counting 

Statistics in computed X-Ray Tomography", Journal of Computer 

Assisted Tomography, 1(1), 1977. 

3. R.A. Brooks and G.D. Di Chiro, "Statistical Limitation in X-ray 

Reconstructive Tomography", Medical Physics, Vol. 3, No. 4, 1976. 

4. P. Reimers and J. Goebbels, "New Possibilities of Nondestructive 

Evaluation by X-Ray Computed Tomography", Materials Evaluation, 41, 

1983.

Source 

- 416 - 

Table 1: Details of Scanner Set-up 

Source type Co-60 

Source strength 4.4 x 10 11 Bq (12 curies) 

Detector 

Detector type Nal (Tl) 

Detector size 5 cm x 5 cm 

Geometry 

Collimator aperture opening (w x h x 1) 3.175 mm x 3.175 mm x 16.5 mm 

Source-to-detector spacing 711 mm 

Centre of rotation to source 394 mm 

Scanning Pattern 

Total number of images 6 

Vertical spacing between images 1.59 mm 

Projections per image 201 

Angular spacing between projections 0.9° 

Rays per projection 256 

Spacing bewteen rays (At) 0.86 mm 

Count time per ray 1 s 

Table 2: Comparison of Reconstructed and Actual Dimensions 

Dimension (see Figure 5(b)) Reconstructed (mm) Actual (mm) 

a 

b 

c 

d 

174.5 

34.4 

85.1 

49.0 

Table 3: Comparison of Reconstructed and Theoretical 

Linear Attenuation Coefficients 

175.0 

35.0 

85.5 

50.0 

Material "£,.»„ (cm" 1 ) art(cm~ i ) li^.cm" 1 

^th c 

brass 0.450 ± .001 0.45 

iron 0.422 ± .001 0.42 

air -0.001 ± .001 0.00

(a) L 

(b) 

ic) 

1 

- 417 - 

•HwK- 

|3AJ|3JU|3/JL^J 

^ S >^ s^\ 

1JU|6JU|2AJLJ 

J 

« J 

,e iM- 

e-9/jw 

•=>l=lo 

Figure 1 : Attenuation of gamma-rays of intensity Io incident on 

(a) a homogeneous slab, 

(b) a non-homogeneous slab, and 

(c) homogeneous and non-homogeneous slabs having identical ray 

sums. 

X-RAY / 

TUBE \f 

1st 

TRANSLATION 

(a) 

DETECTOR 

1st 

TRANSLATE 

1° INCREMENTS 

60th ""^ 

TRANSLATION \ 

(b) 

P 120th 

\ TRANSLATION 

Figure 2 : A schematic illustration of a translate-rotate CT scanner. 

(a) The source and detector are first translated relative to the 

object. Each parallel line represents one ray sura. 

(b) Multiple projections are obtained by rotating the source and 

detector around the object.

DEC 

LSI 11/23 

CPU 

SOURCE 

SOURCE' 

FLASK 

ROM/FLOPPY/ 

WINCHESTER 

MEMORY 

ADAC 1604 

PULSE OUTPUT CONTROLLER 

STEPPER 

MOTOR (SM) 

DRIVERS 

O 

SOURCE 

COLLIMATOR 

PHOTON 

BEAM 

Figure 3 

PRINTER/CRT 

li-TRANSLATION 

IjSAE 

^TRANSLATION SM 

LEAD 

SCREW 

CDA MSP-3 

ARRAY 

PROCESSOR 

CDA MDP-3B 

GRAPHICS 

PROCESSOR 

ADAC 1664 

TTL I/O INTERFACE 

SHAFT ANGLE 

ENCODER (SAE) 

DECODERS 

H.V. 

Nal DETECTOR 

DETECTOR 

COLLIMATOR 

CARRIAGE 

TURNTABLE 

PREAMP 

POWER 

PREAMP 

Q.-BUS 

GRAPHICS 

DISPLAY 

A block diagram of the CRNL toraographic scanner apparatus. 

DLV11-J 

SERIAL 

INTERFACE 

CANBERRA 

1788 

INTERFACE 

•Pi- 

1 

oo

Figure A : Orientation of the casting and MTF test piece on the scanner turntable. 

The scanning plane was changed by slipping 1.59 mm (1/16") 

shims underneath the casting. 

I

(e) 

- 420 - 

Figure 5 : Tomographie images of the six parallel slices. The height of the 

scanning plane, Z, is 109.5 mm for image (a) and decreases in 

steps of 1.59 mm in each of the subsequent images (b) through (f). 

The large subsurface cavity is particularly evident in images (c) 

and (d) but is detectable in all images. A second smaller subsurface 

cavity is evident in (a) only. The modulation transfer function 

(MTF) test pattern is analyzed in Appendix A.

- 421 - 

APPENDIX A 

Determination of the System Modulation Transfer Function (MTF) 

To determine the modulation transfer function (MTF) of the imaging system an 

MTF test pattern was included in each scan. It consisted of a cylindrical 

block of brass containing 21 holes 4, 3, 2, 1.5, 1.0, 0.75 and 0.5 mm in diameter 

arranged in seven parallel rows of three. The diameters of holes in a 

given row were equal and the centre to centre spacing between holes in each row 

was twice the hole diameter. 

Each of the seven sections in Figure A-l (a) through (g) show, at the left, the 

portion of Figure 5(a) containing the image of the MTF test pattern. The linear 

attenuation coefficients of the pixels along the horizontal line superimposed 

on each reproduction of the test pattern are plotted at the right of the 

image. Plots through all seven sets of holes, from largest to smallest, are 

shown in Figure A-l (a) through (g), respectively. 

The three smallest sets of holes (Figures A-l (e), (f) and (g)) show no evidence 

of separation. In all the others (Figures A-l (a), (b), (c) and (d)) 

separation is evident. However, the n's associated with the holes do not extend 

all the way down to the air level, nor do the n's representing the brass 

between the holes fully recover to the n of brass. The peak to peak variation 

between the reconstructed \i of the holes and that of the brass between the 

holes, expressed as a percentage of the full range from air to brass and 

plotted as a function of inverse hole diameter, is a measure of the system 

MTF. 

The system MTF, based on the data shown in Figure A-l, is plotted in Figure 

A-2. The conversion between line pairs/mm (lp/mm) and hole diameter is 

1 lp/mm = l/2d. The 50% point on the MTF corresponds to 2.1 lp/cm.

- LOO - 

Figure A-l: Plots of the linear attenuation coefficients along lines passing 

through the seven rows of equal size holes in the modulation 

transfer function (MTF) test pattern of Figure 5(a). Hole diameters 

are: (a) 4 mm, (b) 3 mm, (c) 2 mm, (d) 1.5 mm, (e) 1 mm, 

(f) 0.75 mm and (g) 0.5 mm.

ÜJ 

I/) 

1 

ce 

TO PE/ 

a. 

100 

90 - 

80 - 

70 

60 

50 

40 

30 

20 

10 

- 423 - 

2 3 

Up/cm (= 1/2d) 

Figure A-2: The system modulation transfer function (MTF) determined, as 

explained in the text, from the data in Figure A-l•

- 424 - 

COMPUTER CONTROLLED CARBON COMPOSITE PANEL TESTERS 

5. VaWalle. 

Canadian HVL Technology Limited, Rzxdate, Ontario 

When we speak of composites these days we generally refer to panels composed 

of a number of layers of cloth made of carbon, graphite or kevlar fibers held 

together and impregnated with a binder or adhesive which after the 

application of heat and pressure within a vacuum becomes a strong light 

panel. 

Since the price of the materials from which these panels are made and the 

construction as well as the testing has been very costly, the use of these 

panels has been restricted primarily to military airplanes, frames and skins. 

However as the cost of the material and construction is coming down and more 

applications are being found more aluminum and other metal components are 

being replaced with this superior material. 

The superiority of these panels for their function in airplanes or other 

structures are in their weight-to-strength ratio as well as in the rigidity 

in one direction versus flexibility in other directions. 

However, the graphite or carbon composites can exhibit these properties only 

if the layers are bonded together perfectly. Even small areas of unbond 

between adjacent layers can significantly weaken the panels. Particularly if 

these multilayer composites are glued against both sides of a metal or fibre 

honeycomb, this so called composite sandwich can only be inspected by means 

of ultrasonic through-transmission. 

The heart of all present ultrasonic panel testing systems consists of two 

transducers, one transmitter and one receiver, connected to an appropriate 

ultrasonic instrument. 

Since the honeycomb centered panel is extremely light and often very large, 

up to 50" x 8" wide or larger, it is difficult to submerge it and perform an 

immersion test on it. 

Therefore, the two transducers are housed in so called water jets or 

squirters. These water coupling devices produce a small diameter water 

column in front of the transducers which allows the sound beam to be conveyed 

from the transmitter transducer to the receiver. 

When the panel is placed between the transducers, the sound beam passes 

through the panel except where it is interrupted by an area of unbond. This 

loss of signal amplitude is related to an amount or area of unbond.

- 425 - 

This paper will concentrate on testing systems which incorporate a number of 

novel and some unique patented devices and features. Probably the most 

important of these are the recently patented so-called "Laminar Link" water 

column coupling devices, which produce truly laminar water columns at any 

orientation, horizontal or vertical at virtually any water pressure. Since 

this laminar flow is achieved without any object such as vanes etc. in the 

water stream these units can be used in pulse echo or through-transmission to 

allow not only the detection but also the location of unbond areas. 

In order to be capable of testing large, thin and fragile panels, besides 

thick rigid ones, the discussed systems were designed to hold these panels 

vertically so the water will not settle on the one side as is the case with 

systems holding the panels horizontally. 

Three types of systems will be presented: 

1) With opposed water coupling devised mounted on a U-shaped arm to test 

flat and slightly curved panels of reasonable width. 

2) Wif.h opposed coupling devices locked together to test very long flat 

panels up to 10 ft. wide 

3) Independently mounted coupling devices to test very long flat panels as 

well as complexly curved panels having widths up to at least 10 ft. 

Also discussed will be a newly developed T)ot-Matrix C-scan recorder which 

prints scan results in GO-NO-GO, lined grey scale shading or numbered defect 

severity indications. This unit requires no special paper but uses plain 

high quality bond paper. 

All systems are under complete computer control and are operating by means of 

simple menu prompted direction which will be discussed in some detail.

- 426 - 

SOME UNCONVENTIONAL TECHNIQUES FOR THE INSPECTION OF LAYERED 

MATERIALS 

P. Cietc 

National Reiea-tc/i Council o & Canada 

Beuche\vMi, Quebec 

ABSTRACT 

Optothermal methods using a pulsed laser have been developed for the detection 

of delaminated areas in layered materials such as Al-epoxy or graphite-epoxy 

laminates. The pulsed laser be?m heats the laminate surface to some degrees C 

inducing a slight thermoelastic bend of the unbonded layer. 

Two techniques have been tested for the detection of delaminatlons. In the 

first technique, an infrared detector is used to monitor the thermal decay 

rate of the surface after the absorption of the pulse. The slope of the 

signal shows a strong variation when the thermal wavefront reaches an 

interface such as a delamination or a layer of different thermal properties. 

This technique allows to detect subsurface defects and to locate their depth, 

or alternatively to evaluate the thermal properties of the layer. 

Graphite-epoxy composites were evaluated by this method. The second technique 

uses a laser interferometer to measure the thermoelastic bending of the layer, 

which is much larger If the layer is unbonded. The displacement of optically 

rough surfaces can be measured with great accuracy by using a focused 

interferometer. Delaminations on an Al-epoxy honeycomb laminate were detected 

by this technique. 

These techniques are attractive for the inspection of large and complex 

structures because they require no contact with the inspected surface, are 

easy to scan and require access from one side only. Moreover, both techniques 

exert a lifting thermoelastic moment on the surface which tends to widen the 

delamination gap. This facilitates the detection of lack-of-adhesion defects 

when the two layers are initially in physical contact. 


Laminate structures such as metal or composite lap joints and honeycomb 

sandwich structures are increasingly used in the transportation and 

construction industries. Their combination of high stiffness and strength 

with low weight is particularly appreciated in the aerospace industry. Such 

structures are formed by adhesively bonding several layers of metallic or 

fiber-reinforced-plastics materials in order to obtain a more uniform stress 

transfer and an increased fatigue life. However, these materials require a 

careful inspection after being assembled to detect any extended bonding 

defect, such as an internal delamination or an unbonded area, which would 

severely affect the structural resistance to transverse or shear loading.

- 427 - 

Ultrasonic and radiographie methods are extensively used for the 

nondestructive inspection of laminate structures » . Commercially available 

instruments based on such methods have reached an advanced stage of 

development, and are now capable to detect most types of bonding defects. 

Research is nevertheless quite active for the development of alternate 

techniques which would lead to faster and less cumbersome devices, requiring 

no coupling media or physical contact with the inspected structure, and which 

can be applied where access is difficult and possible from one side only. 

Moreover, none of the presently used techniques is satisfactory for the 

detection of lack-of-adhesion defects between the layer and the adhesive, 

i.e. areas where the adherents are in physical contact with each other, but 

with a zero bond strength > . 

A research program has thus been established at the Industrial Materials 

Research Institute to explore unconventional inspection techniques for 

stratified materials. Laser-holographic and therr.ographic techniques are 

attractive because they are non-contact and provide in a relatively short time 

a full image of the inspected surface. However, after nearly 20 years of 

development efforts such techniques have found few applications out of the 

research laboratory. Part of the reasons for such an inertia are technical: 

in particular, thermography is surface-emissivity dependent and provides 

little quantitative Information, while holography is very sensitive to ambient 

vibrations and to background light, usually requiring that the inspected 

object be mounted on a cumbersome vibration-isolated table in a ' dark 

environment. Such requirements make it difficult to apply such techniques to 

bulky structures on-the-field. A research effort is thus under way at the 

Industrial Materials Research Institute to develop alternative optical 

techniques which, while maintaining the non-contact and scanning rapidity 

advantages, may provide more quantitative information and show a potential for 

on-the-field applicability. 

Two optothermal approaches under evaluation at our Institute are described in 

this paper. The first is a pulsed thermographie technique which has been 

applied to a quantitative thermal analysis of graphite-epoxy laminates. The 

second Is a novel thermoeleatic technique using a localized heating source and 

focused Interferometer to inspect the transient thermomechanical behaviour of 

the unbonded layer. A description of such methods and of the experimental 

results as well as a discussion of the relative merits and fields of 

application of each of these approaches Is presented. 

PULSED PHOTOTHERMAL INSPECTION OF LAYERED MATERIALS 

The effectiveness and convenience of thermographie techniques for the 

detection of delamlnattons and other subsurface defects has been demonstrated 

by a number of authors " . Usually, the surface of the layered structure is 

heated by a heat lamp or a hot-air gun and the temperature distribution during 

or after heating Is observed with a thermographie camera. The presence of a 

delamination is revealed by the appearance of a hot spot because of the 

thermal-barrier effect of the unbonded interface. More detailed analyses have 

been performed recently » » in order to better quantify the retrieved 

information. Part of the research effort at our Institute is directed toward 

the assessment of the thermographie method and the expansion of the amount of 

Information that can he obtained from the detected signal. In particular, the 

possibility of a thermal analysis of the Inspected materials by properly

- 428 - 

processing the time resolved signal is pointed out in this section. 

One of the problems encountered in thermographie NDT is related to the spatial 

fluctuations of the recorded thermal image caused by changes in the surface 

emlssivity or by non-uniformities of the heating-source distribution«. A 

differential approach using a short heating period and monitoring the 

temperature decay rate of the heated area is an effective way to avoid such 

problems . The temperature history within a homogeneous sample after the 

surface absorption of a short heat pulse can be represented in a first 

approximation by the one-dimensional model for an instantaneous pulse > : 

T(z,t) - — exp (1) 

(ïïKpct) 1 / 2 

where z and t are respectively the depth and time variables, Q is the injected 

energy density, K is the thermal conductivity, p is the density, c the 

specific heat, and a = K/pc is the thermal diffusivity. We can see from ea. 

(1) that the surface temperature of a homogeneous sample decays as (t) ' 

after the absorption of the pulse. 

The effect of the finite pulse duration and non-uniform heat distribution are 

best analyzed by a numerical model. A two-dimensional (axially symmetric) 

finite-difference model was thus used to study the heat propagation en the 

stratified material. under transient surface heating . The surface 

temperature history obtained from such a model in a configuration typical of 

our experimental tests is shown in fig. 1. In this figure, curve (a) 

represents the thermal history after the absorption of a 50 ps pulse by a 

homogeneous steel sample (K = 0.7 W/cm °C, p = 7.8 g/cm and c = 0.5 j/g °C). 

We can see that, apart from an initial perturbation caused by the finite pulse 

duration, the thermal decay slope is equal to - 1/2 on the log-log scale, in 

agreement with eq. (1). Curve (b) corresponds to a lower-conductivity layer 

(K * 0.14 W/cm °C), 100 pm thick, perfectly bonded to a steel (K. = 0.7 W/cm 

°C) substrate. The curve slope presents a discontinuity when the thermal 

front reaches the interface after a period which is of the order of the 

thermal propagation time t = I /4a through the thickness I of the coating. 

The position of this discontinuity can thus be used to obtain the coating 

thickness I if its thermal diffusivity is known. Moreover, the vertical 

displacement Tjj between the coated-sample curve (curve b) and its asymptote 

for large values of t (curve a) can be used to obtain the ratio between the 

effusivity e = Kpc of the coating and the substrate: 

T,i = log (2) 

2 

as can be inferred from eq. (1). Curve (c) corresponds to an unbonded layer 

where the interface defect has been represented by a I um-thick air layer. 

The presence of the defect is apparent in this curve. 

\ graphite-epoxy laminate comprising four layers of Hercules AS 3501-6 prepreg 

bonded together with equal orientation was inspected by such method. Each 

layer has a thickness of 125 \m and is composed of a large number of 

equally-oriented, 8 )im-diameter graphite fibers embedded in an epoxy matrix.

- 429 - 

A micrograph of the laminate is shown in fig. 2. As can be seen from fig. 2a, 

Che upper layer was partially unbonded because of interface stress during the 

mechanical manipulation. The laminate was inspected by the experimental 

apparatus shown in ftg. 3. A CO2-TEA laser with volumetric ratio equal to 

0.05 provided 60mJ, 50 ps pulses with pulse shape optimized to' obtain 

efficient surface heating without overheating . 

The single-transversal-mode beam had a gaussian distribution with 1/e 2 radius 

of 4 mm. The surface temperature was monitored by an InSb infrared detector 

pointed in the center of the irradiated area. Such a detector has a spectral 

sensitivity range from 2 to 5.5 pm, so that it is insensitive to the 10,6 pm 

reflected intensity of the CO2 laser. 

The experimental results obtained with such a system are shown in fig. 4. 

Curve (a) corresponds to a well-bonded region, while curves (b) and (c) were 

obtained on regions where the upper layer was partially unbonded. The 

discontinuity in the slope of curves (b) and (c) after approximately 10 ms 

corresponds to the arrival of the thermal front on the thermally resistive 

air—filled unbonded interface. The thermal propagation time of 10 ms over a 

125 vim thickness corresponds to the expected thermal properties of the 

graphite-epoxy material. 

While the slope of curve (a) is nearly equal to -1/2, as for a uniform sample, 

curve (c) has a significantly different slope, even before 10 ms. This may be 

interpreted by taking into account the heterogeneous nature of this material. 

As can be seen in fig. 2b, the fibers are randomly distributed in the matrix, 

some fibers being in contact with each other while other fibers are separated 

by a significant thickness of the resin. As the thermal conductivity of 

epoxy, 4 . 10~ W/cm °C, is much smaller than the conductivity of graphite 

fibers 17 ' 18 , 0.5 w/cm °C, the fiber distribution may significantly affect 

the transverse thermal properties of the composite. 

Figure 5 shows the results of the fiber distribution in the composite matrix. 

In the three-dimensional model used to obtain such curves, 25 ym x 25 urn x 1 

cm graphite elements are embedded in an epoxy matrix with volume ratio of 

0.5. A contact thermal resistance equivalent to 2.5 urn of epoxy is assumed 

between adjacent graphite elements. Curves (a) and (b) in Figure 5 correspond 

to a well-bonded and delaminated layer, respectively, with a uniformly 

distributed graphite-epoxy mesh, as shown in Figure 6a. Curve (c) in Figure 5 

corresponds to a delaminated layer with a random matrix obtained by a 

random-number generator, Figure 6b. We can see that curve (c) has a more 

gradual slope, similar to the experimental curve (c) of Figure 4. These 

results suggest that photothermal techniques may provide information not only 

on the presence of sub-surface delaminations, but also on the fiber content 

and distribution in fiber-resin composite materials. 

As a conclusion to this first section, it can be said that pulsed photothermal 

techniques are an attractive unconventional approach for the inspection of 

coated or layered materials. Their rapidity and non-contact nature makes such 

techniques convenient to use for scanning the surface of large structures. 

Moreover, the possibility to use such techniques for an on-the-field thermal 

evaluation of layered materials may open new opportunities in the NDE field.

- 430 - 

PULSED PILATOMETRIC INSPECTION OF UNBONDED LAYERS 

Thermographie techniques are particularly convenient for the inspection of 

high emissivity surfaces such as graphite-epoxy laminates or anodized 

coatings. However, the thermographie signal is normally lower by an order of 

magnitude when inspecting metallic surfaces such as Al-epoxy honeycomb panels 

such as the one shown in fig. 7. This is a consequence of the low emissivity 

of bare aluminium surfaces, which is typically of the order of 0.1. On the 

other hand, thermography measures an indirect property of the material i.e. 

thermal resistance of the layer-to-layer interface, rather than the property 

which is of real interest i.e. the adhesive bond at the interface. A lack of 

adhesion between the aluminium skin and the honeycomb core in a laminate such 

as the one shown in fig. 7 would be quite difficult to detect thermographically, 

because the thermal conductivity of the air-filled honeycomb core 

is lower by several orders of magnitude than the thermal conductivity of the 

aluminium skin. 

i 2 u 

The holographic interferometry approach • » , on the other hand, provides 

a direct bond evaluation by mechanically lifting the bonded layer between the 

two hologram exposures. Although either vibration, pressure or heat can be 

applied to lift the unbonded layer, thermal stressing is the most convenient 

technique for a fast and non-contact inspection of laminates ~ . With this 

technique, the structure is holographically inspected before and after surface 

heating by a lamp or a hot-air gun. unbonded areas are revealed by the 

appearance of localized fringe patterns if the therraoelastic bending forces 

produce a- normal displacement of the unbonded layer of more than half a 

wavelength. 

In order to take advantage of the mechanical inspection capability of 

holography while avoiding its sensitivity to ambient vibration and stray 

light, a thermoelastic interferometric technique has been developed at our 

Institute 21 . 

The basic principle is shown in fig. 8. The layer to be inspected is heated 

by a pulsed and focused laser beam. The pulse duration, typically 1 ms or 

less, is comparable to the thermal propagation time across the layer 

thickness. Consequently, a transient therraoelastic moment causes the layer to 

lift, if unbonded, while its position is monitored by a sensitive focused 

interferometer. Because of the short duration of each pulse, large surfaces 

can be scanned rapidly. Apart from eliminating fringe-counting and heat 

uniformity problems, this approach has the following advantages with respect 

to holography: 

1. Every point being inspected during a very short period of time, 

low-frequency ambient vibration noise can be avoided. Operation on the 

industrial floor Is thus possible with our approach. High-power pulsed 

lasers have been used in the past to avoid hologram blurring during the 

exposure time, but ambient vibrations during the heating period between 

the two exposures cannot be avoided with such a method. Moreover, the 

colllmated interferometer beam can be spatially filtered to avoid stray 

light, so that our system can operate under conditions of normal ambient 

illumination.

- 431 - 

2. Holography requires displacements of some wavelengths, while laser 

interferometers are sensitive to much smaller displacements > . 

Heterodyne holographic techniques can be used to detect small 

deformations , but this requires the use of complex double-wavelength 

exposure techniques demanding an accurate control of the alignment'and of 

the ambient conditions while still requiring point-by-point scanning. 

3. In the holographic approach, a broad heat source is used to heat a large 

area of the bonded structure. The generated thermoelastic stress 

distribution inside the unbonded layer is of the type shown in fig. 9a. 

A large longitudinal stress is produced in a direction parallel to the 

surface, but the bending component which makes the.-unbonded area to lift 

outward is quite small, unless the layer is initially slightly convex. 

Bending of the unbonded layer appears to be more related to differential 

thermal expansion between the layer and the substrate and to longitudinal 

thermal gradients produced by the thermal barrier effect of the unbonded 

interface » . Such lifting mechanisms are weak and unreliable. On 

the other hand, the transient strain configuration under localized pulsed 

heating shown in fig. 9b is much more effective in producing a bending 

moment, as was verified using both analytical and finite-element 

thermoelastic modelling. 

The advantages of conventional holographic techniques over interferometry are 

that a full picture is obtained without scanning and that scanning surfaces 

can be inspected. However, focused interferometers are now commercially 

available which can probe perfectly rough surfaces at any angle of incidence 

and at operating distances of several meters . As to the scanning speed, a 

rate of the order of 100 cm /s could be obtained with 1 ms pulses and a 3 mm 

resolution; such a speed compares favorably to the exposure, heating and 

development time required with the conventional holographic approach. 

The interferometric technique was experimentally tested on different 

adhesively-bonded samples. Some preliminary results were obtained using a 

Cu-Be layer, 125 um thick, epoxy-bonded to a massive plexiglass plate. Some 

unbonded areas at the metal-plexiglass interface could be easily localized 

visually through the plexiglass. A nearly 1 J, 1 ms pulse from a Nd:YAG laser 

was used to heat the layer surface to some degrees C above ambient. Fig. 1Ü 

shows two curves obtained by scanning the two concentric laser beams over two 

delaminated areas of different dimensions. Curve (a) was obtained over a 1 

cm-diameter unhond with a 6 mm-diameter heating YAG beam, while curve (b) was 

obtained over a 3 mm-diameter delamination with a 0.7 mm-diameter heating 

beam. The unbonds are clearly visible, and the signal level of some \m is 

much higher than the ultimate sensitivity of the interferometer. 

The same interferometric system was also used to inspect an Al-epoxy laminate 

of the kind shown in fig. 7. Some artificial bonding defects were produced by 

inserting some Teflon foils between the aliminium skin and the epoxy or 

between the epoxy and ths honeycomb core. The skin thickness was of 300 ym, 

and the honeycomb cells had a size of 3 mm. Fig. 11 shows some curves 

obtained by scanning the skin surface over a 2 cm-diameter unbonded area. The 

heating-beam diameter was of 1 mm, 3 mm and 6 mm, respectively for curves (a), 

(b) and (c). All of the three curves show a clear evidence of the unbonded 

area, which is situated between the 5 mm and the 25 mm positions on the 

abscissa. A nearly constant background signal is also obtained on the

- 432 - 

well-bonded region, because of the linear thermal expansion as well as the 

elastic deformation of the relatively compliant epoxy interface. It is 

interesting to note that the 3 mm underlying honeycomb structure can be seen 

only when the heating laser beam diameter is smaller than the cell size, in 

agreement with the thermoelastic considerations illustrated in Fig. 9 . ' 

CONCLaSION 

The optothermal approach to the nondestructive inspection of layered materials 

appears to be an attractive alternative to conventional materials inspection 

techniques. Its main advantages are that it is fast, non-contact, does not 

require sample preparation nor access from both sides of the laminate, and 

provides information about the mechanical bonding properties of the layer. 

Two inspection techniques are presented in this paper, and the results of our 

experimental investigation are discussed. The infrared technique appears to 

be a relatively inexpensive method for the inspection of high emissivity 

materials such as graphite-epoxy laminates, while the interferometric approach 

is more appropriate for the inspection of metal-honeycomb structures and for 

quantitative analyses of bonded laminates. 

REFERENCES 

1. Scott, I.G. and Scala, CM.- A review of NDT of Composite Materials, NDT 

Int. Vol. 15, p. 75, 1982. 

2. Segal, E. and Rose, J.L., NDT Techniques for Adhesive Bond Joints, in 

"Research Techniques in NDT", R.S. Sharpe ed., Academic Press, London, 

1980. 

3. Chaskelis, H.H., and Clark, A.V., - Ultrasonic Nondestructive Bond 

Evaluation: an analysis of the problem, Mater. Eval. Vol. 38, p. 20, 

1980. 

A. Birnbaum, G. and Vest, CM. - Holographic NDE: status and future, in 

International Advances in NDT, vol. 9, p. 257, 1983. 

5. Vavilov, V.P. and Taylor, R. - Theoretical and practical aspects of the 

thermal NOT of bonded structures, in "Research Techniques in NDT", R.S. 

Sharpe ed., Academic Press, London, 1982. 

6. Vogel, P.E.J., - Evaluation of bonds in armorplate and other materials 

using infrared NDT techniques, Appl. Opt. Vol. 7, p. 1739, 1968. 

7. Lawson, W.D. and Sabey, J.W. - Infrared Techniques, in "Research 

Techniques in NOT", R.S. Sharpe ed., Academic Press, London, 1970. 

8. Trezek, G.J. and Balk, S. - Provocative Techniques in Thermal NDT 

Imaging, Mater. Eval., Vol. 34, p. 176, 1976. 

9. Engelhardt, R.E. and Hewgley, W.A. - Thermal and Infrared Testing, in 

"Nondestructive Testing, a Survey", NASA-SP-5113, 1973.

- 433 - 

10. Green, D.R., Schneller, M.D. and Sulit, R.A. - "Thermal NDE Method for 

Thermal Spray Coatings", 10th Int. Thermal Spraying Conf., Essen, West 

Germany, May 1983. 

11. Williams, J.H., Mansouri, S.H. and Lee, S.S., - One-dimensional Analysis 

of Thermal Nondestructive Detection of Delamination and Inclusion Flows, 

British Journal of NDT, Vol. 22, p. 113, 1980. 

12. Griffiths, D.H. - Toward Quantitative Industrial Thermography, SPIE 246, 

152, 1980. 

13. Carslaw, H.W. and Jaeger, T.C. - Conduction of Heat in Solids, Chapt. 10, 

Oxford University Press, Oxford, 1959, 

14. Ready, J.F. - Effects of High-Power Laser Radiation, Academic Press, New 

York, 1971. 

15. Cielo, P. - Analysis of Pulsed Thermal Inspection, 14tli Symposium on 

Nondestructive Evaluation, San Antonio, Apr. 19-21, 1983. 

16. Dallaire, S. and Cielo, P. - Pulsed Laser Treatment of Plasma-sprayed 

Coatings, Metall. Trans. Vol. 13B, p. 479, 1982. 

17. Fritz, W. Huttner, W., Maler, K. and Brandt R. - Thermophysical 

Properties of Carbon-fibre Carbon-matrix Composites at High Temperatures, 

Rev. Int. Hautes Temp. Refract., Vol. 16, p. 350, 1979. 

18. Balageas, O.L. and Luc, A.M. - Non Stationary Thermal Behaviour of 

Directional Reinforced Composites. Limit of Application of Thermal 

Property Horaogenization, 18th AIAA Thennophysics Conf., Montreal, June 

1-3, 1983, 

19. Kersch, L.A. - Laminate Structure Inspection, in "Holographic NDT, 

R.K.ERF ed., Academic Press, N.Y., 1974. 

20. Harris, W.J. and Woods, D.C. - Thermal Stress Studies Using Optical 

Holographic Interferometry, Mater. Eval. Vol. 32, p. 50, 1974. 

21. Cielo, P., Rousset, G. and Bertrand, L. - Nondestructive Interferometric 

Detection of Unbonded Layers, Optics and Lasers in Engineering, 1984 (to 

be published). 

22. Moss, G.E., Miller, L.R. and Forward, R.L. - Photon-noise Limited Laser 

Transducer for Gravitational Antenna, Appl. Opt. Vol. 10, p. 2495, 1971. 

23. Wickramasinghe, H.K., Martin, Y., Ball, S. and Ash, E.A. - Thermodisplaceraent 

Imaging of Current in Thin-Film Circuits, Electron Lett. 

Vol. 18, p. 700, 1982. 

24. Damm, V., DISA, Copenhagen, and Wedgwood, F.A., AERE Harwell,' England, 

Private communications.

0.1 

- 434 - 

0.01 0.1 1 

TIME (m s) 

Fig. 1 Calculated thermal decay curves for a 

plasma-sprayed coating heated by a 50 |xs pulse. 

Curve (a): bare steel sample; 

Curve (b): well-bonded Ni-Cr coating on steel substrate; 

Curve (c): delaminated coating

- 4 35 - 

Fig. 2 (a): cross-section of the 4-ply graphite-epoxy laminate 

used in the experiments. The upper layer was partially 

disbonded under stress; 

(b): close-up view of the fiber-epoxy matrix 

(a) 

1OO|o.m 

(b)

IR Detectors- 

Fig. 3 Experimental apparatus used to thermally inspect 

the four-ply laminate shown in Fig. 2.

10 _ 

0.1 

1_ 

0.01 0.1 1 

- 437 - 

TIME (m s) 

(c) 

10 100 

Fig. 4 Experimental curves obtained with a graphite-epoxy 

laminated sheet. 

(a): well-bonded area; 

(b) and (c): delaminated regions.

o ü 

LU 

LU 

OC 

Ü 

LU 

CC 

ICC 

LU 

Q. 

LU 

LU 

Ü 

Lf 

CC 

œ 

1O_ 

1_ 

0.1. 

0.01 

I 

0.1 

- 438 - 

1 

TIME (m s) 

(a) 

10 100 

Fig. 5 Theoretical curves obtained for a graphite-epoxy 

laminated sample. 

Curve (a): well-bonded sample; 

Curve (b): delaminated layer, uniform fiber distribution; 

Curve (c): delaminated layer, random fiber distribution.

• •' '

i M H 

I ! ! : • 

niy 

ill 

Fig. 9 Thermoelastic siress distributiow on a partially 

unbonded layer under: 

(a): continuous broad healing ami 

(b): pulsed focused heatin. j 

(a)

HEATING 

PULSES 

DETECTOR 

" -—FILTER 

LASER 

BONDED 

LAYER 

SUBSTRATE 

Fig. 8 Basic principle of the thermoelastic interferometric 

method for detecting an unbond. A surface-heating 

pulse causes the layer to thermoelastically bend, 

while its vertical displacement is detected by a 

focused interferometer. 

.c-

Fig. 7 Al-honeycomb sandwich structure to be Inspected 

for unbond defects.

- 443 - 

5 10 

POSITION (mm) 

Fig. 10 Interferometrically measured vertical 

displacement of the heated surface of an 

epoxy-bonded Cu-Be layer over: 

(a): a 1 cm-diameter unbond and 

(b): a 3 mm-diameter unbond.

10 15 20 25 

POSITION (mm) 

30 35 

Fig. 11 Interferometric scan of a 2 cm-diameter 

deiamination on an Al-epoxy honeycomb laminate, 

with a heating beam diameter of: 

(a): 1 mm, (b): 3 mm and (c): 6 mm. 

I

- 445 - 

MATERIALS EFFECTS ON ACOUSTIC EMISSION DURING DEFORMATION AND 

FRACTURE 

M. Nabit Ba&6im 

Univzuity o (j Wa.ni.toba. 

Winnipeg, Manitoba, Canada 

ABSTRACT 

The technique of acoustic emission is increasingly used for nondestructive 

testing of large structures. Because of its passive nature in detection of 

the stress wave emissions which occur during plastic deformation and/or crack 

propagation, it is possible to monitor large areas by using a relatively low 

number of transducers. While the application of the technique is being 

perfected, understanding of the nature and origin of acoustic emission is 

less clear due to the complexity of quantifying the nature of the acoustic 

sources and the effect of the medium where the stress waves are propagating. 

In this investigation, models describing the origin of acoustic emission in 

terms of dislocations are reviewed. The effect of attenuation on the stress 

waves is examined and the extent of acoustic emission activity from a 

specific material is related to its microstructure. A comprehensive analysis 

of the nature of acoustic emission as observed in a typical application is 

thus presented. 


Considerable effort has taken place in recent years to use acoustic emission 

as a nondestructive testing method for evaluation of the integrity of 

structures and components. The bulk of this effort has been directed to 

applications involving large structures, namely pressure vessels, nuclear 

reactors...etc where flaw localization is the main concern and evaluation of 

the severity of these flaws is the primary goal of the application of the 

technique. Commercially available systems, which are equipped with very 

sophisticated digital processing techniques including micro-processors are 

available for performance of such tests. The application of acoustic 

emission is usually conducted during proof tests of these structures and 

record.-, of performance of the acoustic emission activity, in terms of 

acoustic emission parameters such as total ringdown count, count rate, peak 

amplitude and energy of the signals are obtained. As well, the position of 

the flaw is determined from the differences in time of arrival to several 

transducers positioned at various locations on the structures. Bassim(l) and 

and Houssny-Eman(2) have summarized the different techniques used for

- 446 - 

analysis of the acoustic signals in the time and frequency domains and have 

provided a theoretical basis for the flaw localization techniques most used 

in industrial applications. 

While examples abound in the literature on successful application of acoustic 

emission to large structures by use of general purpose large systems which 

are computer-controlled to give, in real time, the location and severity of 

flows, the operation of such systems depends, to a large extent on the 

presence of highly skilled operators which can provide almost immediate 

interpretation of the raw data obtained during the tests. Furthermore, these 

sophisticated pieces of equipment have only a limited capability to perform 

over a long period of time beside the fact that they are costly, bulky and 

difficult to transport from one place to another. These drawbacks limit the 

use of existing first generation acoustic emission equipment to applications 

which have a finite length of time, such as proof tests, where they can be 

assembled to perform tha test, proceed with the actual test, analyse the data 

and disassemble to move on to another application. 

An alternate process to provide continuous monitoring of large structures 

with acoustic emission is to use surveillance units which continuously 

analyse acoustic emission signals, determine the relevant signals vhlch are 

indicative of impending failure and use these signals to trigger a warning 

system. These surveillance units would be compact, inexpensive, and a number 

of these units can be applied to a very large structure such as a pipeline, 

an offshore platform or a nuclear reactor for continuous monitoring of these 

structures. The units will each include a micro-processor which is 

software-controlled to perform the different functions assigned to each of 

these units. This new approach to use of acoustic emission for continuous 

monitoring of large structures has been patented by Bassim and Tangri(3). 

The description of the system is given by Mitchell and Bassim(4). 

An important feature of these surveillance units is that they are 

preprogrammed to identify signals corresponding to impending failure. This 

requires an intensive research program to identify the sources of acoustic 

emission in the structure under consideration and to determine the effect of 

attenuation on the waveform of the signal detected by the transducer. In 

this paper, this research program is outlined with regard to an industrial 

application, involving the use of the acoustic emission system to monitor a 

long pipeline. 

2. OUTLINE OF MATERIAL TESTING PROGRAM 

The material testing program for a continuous monitoring system used in 

conjunction with pipelines can be summarized in the following steps: 

2.1 Laboratory tests of pipeline materials 

The tensile and fracture properties of a pipeline material were determined 

using tensile and three-point bending specimens respectively. The effect of

- 447 - 

orientation of these specimens with respect to the rolling plate was 

investigated. In the fracture tests, parameters such as the stress intensity 

factor K-£C and the J-integral, Jjc which is used to characterize ductile 

fracture, were determined. 

Together with the mechanical testing, acoustic emission was obtained using a 

system similar to that in Fig. 1. Acoustic emission parameters, namely the 

total count, count rate and root mean square of the amplitude were determined 

as a function of the applied load and elongation. In the fracture tests, 

acoustic emission was obtained by positioning a transducer on the side of the 

specimen and following the emissions as the plastic zone ahead of the crack 

is developed followed by stable crack growth. 

2.2 Model tests on pipeline materials 

In these tests, two aspects were considered: Firstly, the continuous 

acoustic signals due to leaks were detected with transducers of different 

frequencies located at different distances from leaks induced by drilling 

holes in the pipe as described by Bassim and Tangri(5). Air flow at 

different pressures was maintained through the pipe. The experimental set-up 

is shown in Fig. 2. Secondly, the attenuation coefficient a was determined 

using ultrasonic methods described in (6). The attenuation characteristics 

of the pipe were thus established and calculations of the optimum spacing 

corresponding to a transducer with a given resonant frequency was 

calculated. 

2.3 Theoretical consideration of acoustic emission sources 

Theoretical modelling of the sources of emissions in terms of dislocation 

motion and interaction was carried out to arrive at the optimal frequencies 

corresponding to an acoustic emission signal. Also, once an emission has 

occurred, modelling of the type of waves and its propagation characteristics 

was carried out to determine the proper position of transducers on the pipe 

and their optimum spacing. The effect of pipe (or specimen geometry) on the 

destructive and constructive interference of acoustic signals, following the 

analysis of Hamel, Ballon and Bassim(7) was also investigated. 

3. RESULTS AND DISCUSSION 

Following are representative data obtained in the course of the research 

program as well as an interpretation of their physical meaning and the extent 

that they can be used in the design of a dedicated acoustic emission system. 

3.1 Material characterization 

Flat tensile samples, from two orientations with respect to the rolling 

direction, were obtained from A53B pipeline steel. Fig. 3 shows a typical 

stress-strain and RMS data from a longitudinal specimen. It is observed that 

the A53B deforms initially by a complex Lüders band propagation. Several

- 448 - 

bands are initiated right after the first yield drop. High acoustic emission 

activity occurs in the Lüders strain region of the stress-strain curve. Each 

peak on the RMS curve corresponds to a stress drop on the stress-strain 

curve. This indicates that the major source of acoustic emission in this 

steel is due to inhomogeneous deformation. No systematic difference in 

bahaviour was observed between longitudinal and transverse samples. Thus 

acoustic emission activity is detected mostly within ehe first 3% plastic 

strain during tensile testing. 

In fracture testing, three-point bend specimens, 12.7 mm thick with a notch 

in the TL and LT directions were obtained from the as-received A53B pipeline 

steel. Fig. 4 shows a typical RMS voltage of the acoustic signal during a 

test as a function of load point displacement. The acoustic emission 

activity remains high throughout the displacement range investigated which 

represents the development of the plastic zone ahead of ehe crack tip without 

extending the crack length. Most of the emission were found to occur during 

the plastification of the ligament ahead crack tip and no significant 

emission accompany the stable crack growth process itseLî. This is in 

agreement with Blanchette, Bassim and Dickson(8) for A516 grade 70 steel. 

3.2 Leak and attenuation characterization 

The leak tests were performed on a segment of pipe, 400 cm long with inside 

and outside diameters of 160 and 170 mm respectively. The pressure and flow 

rate of the air through the pipe were measured accurately. Two series of 

tests, one using an accelometer with a range of up to 25 KHz and the other 

with an acoustic transducer with a resonant frequency of 1140 KHz were 

conducted. An artificial leak was made by drilling a hole directly in the 

middle of the pipe and the signal was analysed for its frequency content. 

Fig. 5 represents a plot of the RMS from different leak sizes as a function 

of pressure while Fig. 6 shows the effect of hole diameter on the leak 

frequency obtained at different air pressures through the pipe. It is 

noticed that for a given pressure, as the hole diameter decreases, the leak 

frequency also decreases. On the other hand, the value of the R.M.S. which 

represents the relative amplitude of the signal increases with the increase 

in hole diameter as well as increase of the pressure. Finally, measurement 

of the signal amplitude at different points on the pipe shows no significant 

attenuation for this relatively short segment of pipe. Turbulence due to end 

effects appears to increase the signal amplitude to some extent. 

In the attenuation tests, the aim is to determine the attenuation coefficient 

aj which is defined for a plane wave as 

P = Po e~ ad (1) 

where PQ is the initial pressure and P is the sound pressure at distance d. 

In this study the attenuation coefficient a, expressed as dB/m, was

- 449 - 

determined experimentally for pipeline steels through the thickness of plate 

as well as in the pipe segment described earlier. The procedure for 

estimation of a through plate thickness using the pulse reflection method was 

described by Krautkramer(9), and will not be repeated here. For X70 pipeline 

steel, the attenuation coefficient was found to be equal to 1.09 dB/m while 

for X42 steel, the attenuation was much higher with a = 2.125 dB/m. 

In the tests on the pipe segment, a wave generator was used to produce 

frequencies swept between 01 to 1 MHz. An exciter transducer was connected 

with the wave generator. A receiving transducer was placed at specific 

distances from the exciter and the signal, after amplification, pass-as 

through an R.M.S. voltmeter and is plotted as a function of frequency. The 

value of a determined in this test corresponded to 0.89dB/m. This value is 

affected by the reflection of the elastic waves in the free ends of the 

pipe. 

3.3 Theoretical modelling 

Significant effort in the theoretical modelling of the acoustic emission 

system and its functioning was made. This ranges from modelling the sources 

of burst emission in terms of lattice perturbation and dislocation 

interaction (Bassim and Wassef(lO)), to the prediction of the stresses and 

displacement in a plate due to a propagating crack (Bassim, Wassef and 

Tangri(ll)) and finally the prediction of the type and frequency spectre! of 

the noise due to the flow of a gas (or fluid) in a circular hole or 

rectangular slit (Wassef et al(12)). 

The above studies have for general conclusions that acoustic emissiori is 

caused by the perturbation of an otherwise perfect lattice to cause phonon 

emission. Involved quantum mechanics approach shows that the interaction of 

two or more dislocations is capable to producing a phonon emission 

corresponding to about 200 KHz which is the range where acoustic emission is 

most observed. 

Once acoustic emission is produced, its propagation will depend on the 

position of the source with respect to the structure and on the relative 

dimension of the emitting crack and the rest of the structure. Modelling, 

involving spectral representation, Green's functions as well as solutions to 

Fredholm integral equations are used to predict the stresses and displacement 

fields in a body containing a crack and subjected to a loading. These 

predictions give an accurate indication of the type of transducer to be used 

in conjunction with a given geometry as well as the most appropriate 

transducer configuration to use in a structure. 


A summary of a research program which was carried out in conjunction with 

development of a new acoustic emission system is presented. The specific 

application is the continuous monitoring of pipelines. The program involves 

materials studies on pipeline steels, model tests to measure signals

- 450 - 

characteristics of leaks and theoretical modelling of the sources of acoustic 

emission and wave propagation in a solid medium. 

5. ACKNOWLEDGEMENTS 

The author acknowledges the efforts of his associates, Drs. W. Wassef, M. 

Houssny-Emam and D. Tseng and Mssrs. Gauthier and Hartle, in producing some 

of the data presented. The financial support of Petro-Canada and the Natural 

Sciences and Engineering Research Council is acknowledged and appreciated. 

6. REFERENCES 

(1) M.N. Bassim, in Microstructural Characterization of Materials by 

Non-Microscopical Techniques, 5th Riso International Symposium on Metallurgy 

and Materials Science, 1984, pp. 193-198. 

(2) M.N. Bassim and M. Houssny-Emam, in Acoustic Emission, J.R. Matthews 

ed., Gordon and Breach, 1983, pp. 139-164. 

(3) M.N. Bassim and K. Tangri, U.S. Patent 570866 (pending). 

(4) J. Mitchell and M.N. Bassim, Fifth Canadian Conference on Nondestructive 

Testing, 1984. 

(5) M.N. Bassim and K. Tangri, Proceedings of the International Conference 

on Pipeline Inspection, Published by CANMET, 1984, pp. 529-544. 

(6) J. Krautkramer and H. Krautkramer, ultrasonic Testing of Materials, 

Spring-Gerlag, 1969, pp. 89-95. 

(7) F. Hamel, J.P. Bailon and M.N. Bassim, Ultrasonics, 1977, pp. 125-127. 

(8) Y. Blanchette, M.N. Bassim and J.I. Dickson, in Proceedings of the 5th 

Canadian Fracture Conference, Pergamon Press, 1981, pp. 191-199. 

(9) .T. Krautkratner, Ultrasonic Nondestructive Testing of Materials, 1961. 

(10) M.N. Bassim and W. Wassef, Physical Review, Submitted for publication, 

1984. 

(11) M.N. Bassim, W. Wassef and K. Tangri, Int. J. Fracture, in press. 

(12) W.A. Wassef, M.N. Bassim, M. Houssny-Emam and K. Tangri, submitted for 

publication in J. Acoustical Soc. of America. 1984.

STRUCTURE 

•^ TRANSDUCER (S) 

- 451 - 

TIME DOMAIN 

ANALYSIS 

FREQUENCY 

DOMAIN ANALYSIS 

RINGOOWN COUNT 

ENERGY 

ROOT-MEAN SQUARE OF AMPLITUDE 

AMPLITUDE DISTRIBUTION 

COMPUTER 

FLOW LOCATION 

Figure 1: Typical Experimental Set-up 

-CONSTANT PRESSURE ^PIEZOELECTRIC TRANSDUCER 

VALVE-2 / OR 

fi ^ 7 /—PRESSURE GAUGE-2 /— ACCELEROMETER 

-FLOWMETER TRANSDUCER SfTES PRESSURE REGULATOR- 

AMPLFIER FILTER 

--- 

t i_ 

TERMINAL 

PRESSURE GAUGE- I 

CONSTANT PRESSURE VALVE-1 - 

- 

R. M. S. 

METER 

SPECTRUM 

ANALYSER 

BE!» 

Figure 2: Leak Testing on Pipeline

72Or 

- 452 - 

STRESS 

2.30 SX» 7.50 10.0 12 5 ISO 17.5 20.0 

TRUE STRAIN (%) 

Figure 3: Acoustic emission from tensile test of 

pipeline steel. 

7S0 - 

*• 500 - 

250 

0.5 1.0 1.5 2.0 

LOAD - POINT DISPLACEMENT ( mm ) 

Figure 4: Load and RMS vs displacement during 

fracture test.

- 453 - 

500 

PRESSURE (KPO) 

LEAK DIAMETERS • 0.5 mm 

* 1.0 mm 

A 1.5 mm 

• 2.0 mm 

X 2.5 mm 

Figure 5: R.M.S. vs pressure in leak tests 

335 900 

PRESSURE (Kpo) 

Figure 6: Frequency vs pressure in leak tests 

670 

670

Ahmad, A. 

Babcock & Wilcox Canada 

Coronation Blvd. 

CAMBRIDGE, Ontario 

N1R 5V3 

Allan, Carolyn 

CSNDT Foundation 

Mohawk College 

135 Fennell Avenue W. 


L8N 3T2 

Allinson, J.O. 

Westinghouse Canada Inc. 

Nuclear Products Dept. 

PORT HOPE, Ontario 

L1A 3V4 

Andersen, H. 

Gulf Canada Products Co. 

TORONTO, Ontario 

Andress, B.L. 

Stelco Inc. 

Box 490 

WELLAND, Ontario 

L3B 5R2 

Atherton, D. 

Queen's University 

KINGSTON, Ontario 

K7L 3N6 

Barker, P. 

Atomic Energy of Canada Ltd. 

Sheridan Park Research Community 

MISSISSAUGA, Ontario 

L5K 1B2 

Baron, J.A. 

Ontario Hydro Research 

800 Kipling Avenue 


M8Z 5S4 

Bassim, Nabil 

University of Manitoba 

Dept. of Mechanical Engineering 

WINNIPEG, Manitoba 

R3T 2N2 

- 454 - 

LIST OF DELEGATES 

Beards, David 

Ontario Hydro 

Toronto, Ontario 

Behal, V.G. 

Dofasco Inc. 


Berger, H. 

Industrial Quality Inc. 

9832 Canal Road 

P.O. Box 2397 

GAITHERSBURG, MD 20879 

Bradbury, T. 

Sonco Steel Tube Ltd. 

14 Holtby Avenue 

BRAMPTON, Ontario 

L6X 2M1 

Buby, J. 

ASCO Ltd. 


Burnay S. 

Viatec Resource Systems Inc. 

Suite 425, 20 Richmond St. East 


M5C 2R9 

Burton, J. 

Babcock & Wilcox 

Coronation Blvd. 


N1R 5V3 

Bussiere, J.F. 

National Research Council 

IMRI 

75 DeMortagne Blvd. 

BOUCHERVILLE, Quebec 

J4B 6Y4 

Butler, M. 

McMaster University 

1280 Main St. West 


L8S 4K1

Caron, V. 

Energy Mines & Resources 

CANMET 

568 Booth Street 

OTTAWA, Ontario 

K1A0G1 

Carroll, Y.A. 


Cavers, J. 

Canadian Armed Forces 

Aircraft Maintenance Dev. Unit 

TRENTON, Ontario 

Chittim, K. 

CWB 

254 Merton Street 


Churchill, Paul 

Suncor Oil Sands Group 

P.O. Box 4001 

FORT MCMURRAY, Alberta 

T9H 3E3 

Cielo, P. 


IMRI 


BOUCHERVILLE, Quebec J4B 6Y4 

Cobill, P. 


Radiochemical Company 

413 March Road 

P.O. Box 13500 

KANATA, Ontario 

K2K 1X8 

Dalrymple, D. 



CHALK RIVER, Ontario 

Derkx, S.J. 

ASCO Ltd. 


Deverno, M. 




- 455 - 

DeWalle, S. 

CANDET 

18 Canso Road 

REXDALE, Ontario 

Dumoulin, G. 

Sonco Steel Tube Ltd. 

14 Holtby Avenue 


L6X 2M1 

Duncan, D.B. 




Dunn, A.L. 

Intertechnology Limited 


Fahr, A. 

Ontario Research Foundation 

Sheridan Research Community 


Foster, M.L. 


Radiochemical Company 

413 March Road 

P.O. Box 13500 

KANATA, Ontario 

K2K 1X8 

Gerrior, David 

Eastern Technical Services 

686 Rothesay Avenue 

SAINT JOHN, N.B. 

E2L 4E6 

Gilbert, R.E. 


Whiteshell Nuclear Research Est. 

PINAWA, Manitoba 

ROE 1L0 

Gooch, K. 

CSNDT 


Grabarczyk, J.A. 


Sheridan Park Research Community 


L5K 1B2

Hanrath, D. 

Ministry of Consumer & Commercial Rel. 


Harding, N. 

CSNDT 


Hartling, M. 

McMaster University 


Havercroft, W.E. 

Consultant NDT 

2373 Ridgecrest Place 


K1H 7V4 

Hearn, M. 

Stelco Inc. 

General Delivery 

NANTICOKE, Ontario 

NOA 1LO 

Helliwell, T. 

191 Waverley Road 


M4L 3T4 

Hirning, C. Ross 

Ministry of Labour 

Special Studies & Services Branch 

400 University Avenue 

8th Floor 


M7A 1T7 

Holden, T. 




Holt, R. 

CANMET Dept. E.M.&R. 



K1A 0G1 

Horenfeldt, M. 

Du Pont Canada Ltd. 


- 456 - 

Huether, Doug 

Ontario Hydro 

Darlington G.S. 

Box 1000 

BOWMANVILLE, Ontario 

L1C 3W2 

Julier, Alan 

Hocking Electronics Inc. 

6631 Wakefield Drive 

ALEXANDRIA, VA 22307 

Kennedy, W. 

Canadian Welding Bureau 



M4S 1A9 

Kittmer, C. 




Levesque, Yvon Robert 

Alcan International Kingston Labs. 

12 Michael Grass Cres. 

KINGSTON, Ontario 

K7M 2W3 

Macecek, Mirek 

Techno Scientific Inc. 

205 Champ jne Drive 

DOWNSVIEW, Ontario 

Mak, D. 

Dept. Energy Mines & Resources 

CANMET 



K1A 0G1 

Manzer, L. 


135 Fennell Avenue West 


Marr, David 

Saint John Shipbuilding & Drydock Ltd. 

SAINT JOHN, N.B. 

Mayo, W. 



CHALK RIVER, Ontario

Mclntyre, Dent 

Babcock & Wilcox Canada 

Coronaton Blvd. 


N1R 5V3 

McKinney, W. 

E.I. duPont de Nemours & Co. Inc. 

Photo Products Building No. 1 

Chestnut Run 

WILMINGTON, DE 19898 

McNeill, Rob 

Ontario Hydro 

Darlington G.S. 

Box 1000 

BOWMANVILLE, Ontario 

L1C 3W2 

Mitchell, A. 

NB Research & Productivity Council 

FREDERICTON, N.B. 

Mitchell, J. 

Viatec Resource Systems Inc. 

Digital Building 

6815-8th Street N.E. 

CALGARY, Alberta 

T2E 7H7 

Moles, M.D.C. 




M8Z 5S4 

Monchalin, J.P. 


75 DelVortagne Blvd. 


J4B 6Y4 

Mullins, K. 

Stelco Inc. 

P.O. Box 2030 


L8N 3T1 

Nadeau, F. 


IMRI 



J4B 6Y4 

- 457 - 

Nieberg, A.B. 

Westinghouse Canada 

Box 510 


L8N 3K2 

Peugeot, R. 

Ridge Inc. 

4432 Bibb Blvd. 

TUCKER, GA 30084 

Piche, L. 


75 DeMortagne Blvd 


J4B 6Y4 

Rathwell, L. 

Monenco Offshore Limited 

1200 Monenco Place 

801-6th Avenue Southwest 

CALGARY, Alberta 

T2P JW6 

Reykdal, L. 


50 Paramount Road 

WINNIPEG, Manitoba 

R2X 2W3 

Rotter, R. 

McMaster Engineering Society 

1, Leith CT. 

ANCASTER, Ontario 

L9G 3V8 

Roy, M. 

Guardian inspection Atlantic Ltd. 

552A Windmill Road, Bay A 

DARTMOUTH, Nova Scotia 

B3B 1B3 

SEDO, J. 




M8Z 5S4 

Sharpe, R.S. 

Atomic Energy Research Est. 

Harwell, United Kingdom

Sinclair, A. (Tony) 

University of Toronto 

5 King's College Road 


M5S 1A4 

Staples, R. 

Stelco Inc. 

Box 2030 


L8N 3T1 

Tim leck, D. 




M4S 1A9 

Norscan NDE Ltd. 

94 Ambleside Drive 


Tonner, Paul 




Tremblay, S. 

Laboratoire Ferex Inc. 

2323 Versant-Nord, Suite 110 

SAINTE-FOY, Quebec 

G1N 4P4 

Truss, K.J. 


Whiteshell Nuclear Research Est. 

PINAWA, Manitoba 

ROE 1L0 

van den Andel, J. 

Westinghouse Canada Inc. 

P.O. Box 510 


L8N 3K2 

Veray, Mark 

Pratt & Whitney 


- 458 - 

Wallar, Jon 


Mohawk College 

135 Fennell Avenue West 


L8N 3T2 

Wallis, C. 

Ontario Hydro 

R.R. #4 

KINCARDINE, Ontario 

NOG 2G0 

Wallis/Assistant 

Ontario Hydro 

R.R. #4 

KINCARDINE, Ontario 

NOG 2G0 

Wells, John 

Techno Scientific Inc. 

205 Champagne Drive 

DOWNSVIEW, Ontario 

Yih, G. 

Atomic Energy Control Board 

Supplies and Services 


Zirnhelt, J.H. 




M4S 1A9

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