Reliability Problems of Reactor Pressure C om po ne nts
Proceedings of a Sym posium , Vienna, 10-13 October 1977
RELIABILITY PROBLEMS
OF
REACTOR PRESSURE COMPONENTS
V O L . I I The following States are Members of the International Atomic Energy Agency:
AFGHANISTAN HOLY SEE PHILIPPINES ALBANIA HUNGARY POLAND ALGERIA ICELAND PORTUGAL ARGENTINA INDIA QATAR AUSTRALIA INDONESIA ROMANIA AUSTRIA IRAN SAUDI ARABIA BANGLADESH IRAQ SENEGAL BELGIUM IRELAND SIERRA LEONE BOLIVIA ISRAEL SINGAPORE BRAZIL ITALY SOUTH AFRICA BULGARIA IVORY COAST SPAIN BURMA JAMAICA SRI LANKA BYELORUSSIAN SOVIET JAPAN SUDAN SOCIALIST REPUBLIC JORDAN SWEDEN CANADA KENYA SWITZERLAND CHILE KOREA, REPUBLIC OF SYRIAN ARAB REPUBLIC COLOMBIA KUWAIT THAILAND COSTA RICA LEBANON TUNISIA CUBA LIBERIA TURKEY CYPRUS LIBYAN ARAB JAMAHIRIYA UGANDA CZECHOSLOVAKIA LIECHTENSTEIN UKRAINIAN SOVIET SOCIALIST DEMOCRATIC KAMPUCHEA LUXEMBOURG REPUBLIC DEMOCRATIC PEOPLE’S MADAGASCAR UNION OF SOVIET SOCIALIST REPUBLIC OF KOREA MALAYSIA REPUBLICS DENMARK MALI UNITED ARAB EMIRATES DOMINICAN REPUBLIC MAURITIUS UNITED KINGDOM OF GREAT ECUADOR MEXICO BRITAIN AND NORTHERN EGYPT MONACO IRELAND EL SALVADOR MONGOLIA UNITED REPUBLIC OF ETHIOPIA MOROCCO CAMEROON FINLAND NETHERLANDS UNITED REPUBLIC OF FRANCE NEW ZEALAND TANZANIA GABON NICARAGUA UNITED STATES OF AMERICA GERMAN DEMOCRATIC REPUBLIC NIGER URUGUAY GERMANY, FEDERAL REPUBLIC OF NIGERIA VENEZUELA GHANA NORWAY VIET NAM GREECE PAKISTAN YUGOSLAVIA GUATEMALA PANAMA ZAIRE HAITI PARAGUAY ZAMBIA PERU
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Printed by the IAEA in Austria June 1978 PROCEEDINGS SERIES
RELIABILITY PROBLEMS OF REACTOR PRESSURE COMPONENTS
PROCEEDINGS OF A SYMPOSIUM ON
APPLICATION OF RELIABILITY TECHNOLOGY TO NUCLEAR POWER PLANTS
ORGANIZED BY THE INTERNATIONAL ATOMIC ENERGY AGENCY AND HELD IN VIENNA, 10-13 OCTOBER 1977
In two volumes V O L. II
INTERNATIONAL ATOMIC ENERGY AGENCY VIENNA, 1978 RELIABILITY PROBLEMS OF REACTOR PRESSURE COMPONENTS, VOL. II IAEA, VIENNA, 1978 STI/PUB/467 ISBN 92-0-050178-9 FOREWORD
This Symposium was a follow-up to the Symposium on Reliability of Nuclear Power Plants, held in Innsbruck, Austria, in April 1975; the pro ceedings of that Symposium, organized by the IA E A in co-operation with the Secretariat of the U N Economic Commission for Europe, were published by the IA E A in the same year. The scope of the previous Sym posium ranged from nuclear power plant reliability data systems to the practical application of reli ability analysis in plant design, and questions of testing, operation and maintenance. The present Symposium, held by the IA EA in Vienna on 10— 13 October 1977, paid special attention to the reliability problems of reactor pressure com ponents, because of their importance for the safety of nuclear power plants. Reliability of reactor pressure vessels and other pressure com ponents is of major concern to reactor owners, designers, material suppliers, manufacturers, and regulatory and licensing bodies. The Symposium considered the general approaches for evaluating the reliability of components, operating experience, failure data and their analysis, and touched on some economic aspects. Papers discussed reliability criteria for design, materials selection, fabrication, operation and inspection of components, and studies and models evaluating potential behaviour of materials in reactor service. One session dealt primarily with national and international codes and practices related to pressure components. In a concluding Round Table discussion, subjects covered included the question as to what were the m ost profitable areas of w ork to improve reliability, the scope offered by im provements in design and materials, and reliability through inspection. A total of 154 participants from 29 member states and four international organizations took part. In publishing the proceedings, which include 52 papers presented and a record of the discussions, the Agency wishes to thank all those who contributed to the success of the Sym posium , particularly the Chairmen of the Sessions and the Chairman and Members of the Round Table, and hopes to have made a worthwhile contribution to the subject of reliability technology and its applications. EDITORIAL NOTE
The papers and discussions have been edited by the editorial staff of the International Atomic Energy Agency to the extent considered necessary for the reader’s assistance. The views expressed and the general style adopted remain, however, the responsibility o f the named authors or participants. In addition, the views are not necessarily those of the governments of the nominating Member States or o f the nominating organizations. Where papers have been incorporated into these Proceedings without resetting by the Agency, this has been done with the knowledge o f the authors and their government authorities, and their cooperation is gratefully acknowledged. The Proceedings have been printed by composition typing and photo-offset lithography. Within the limitations imposed by this method, every effort has been made to maintain a high editorial standard, in particular to achieve, wherever practicable, consistency of units and symbols and conformity to the standards recommended by competent international bodies. The use in these Proceedings o f particular designations of countries or territories does not imply any judgement by the publisher, the IAEA, as to the legal status of such countries or territories, o f their authorities and institutions or o f the delimitation o f their boundaries. The mention of specific companies or of their products or brand names does not imply any endorsement or recommendation on the part o f the IAEA. Authors are themselves responsible for obtaining the necessary permission to reproduce copyright material from other sources. CONTENTS OF VOLUM E II
FRACTURE MECHANICS AND DESIGN CONSIDERATIONS (Session V)
Subcritical crack growth in the ligament between the instrumentation r o d s o f th e B B R p re ssu re v e sse l b o t t o m ( I A E A - S M - 2 1 8 / 1 5 ) ...... 3 G. Marci, E. Bazant, H.P. Kautz The growth of underclad flaws in fatigue loading (IA EA -SM -218/16) ...... 17 K. Rahka, J. Forsten Analyse tridimensionnelle de la propagation en fatigue d’un defaut de f o r m e s e m i-e llip tiq u e ( I A E A - S M - 2 1 8 / 4 3 ) ...... 33 J.M. Boissenot Discussion on papers IAEA -SM -218/15, 16, 43 ...... 4 9 Som e aspects relating to the reliability of steel reactor pressure vessels ( I A E A - S M - 2 1 8 / 3 ) ...... 53 G. P ran tl, T. Varga, D .H . N jo D i s c u s s i o n ...... 61 Som e aspects of the assessment of pipework integrity ( I A E A - S M - 2 1 8 / 1 2 ) ...... 63 B.J.L. Darlaston D i s c u s s i o n ...... 81 Som e design considerations for improved reliability of nuclear reactor c o m p o n e n t s ( I A E A - S M - 2 1 8 / 2 8 ) ...... 83 A. Kakodkar D i s c u s s i o n ...... 9 2 Prediction of failure risk in a pressure vessel due to a manufacturing d e fe c t ( I A E A - S M - 2 1 8 / 3 3 ) ...... 9 5 D.L. Marriott, J.M. Hudson D i s c u s s i o n ...... 113
CODES, STANDARDS AND PRACTICES (Session Via)
Codes, standards and practices and their influence on the reliability of nuclear plants (IAEA-SM -218/29) ...... 1 1 9 L. J. Chockie D i s c u s s i o n ...... 132 World-wide standardization of power reactor technology by ISO (IAEA-SM-218/30) ...... 133 K. B e c k e r D i s c u s s i o n ...... 142
The role of the American Society for Testing and Materials (A STM ) in providing standards to support reliability technology for nuclear power plants (IAEA-SM -218/31) ...... •...... 145 L. E. Steele Nuclear reactor pressure vessel surveillance capsule examinations: application of American Society for Testing and Materials standards (IAEA-SM-218/24) ...... 161 J.S. Perrin D is c u s s io n o n p a p e r s I A E A - S M - 2 18 / 3 1 a n d 2 4 ...... 173
PRACTICAL EXPERIENCE WITH COMPONENTS RELIABILITY (Session VIb)
S t e a m g e n e r a t o r r e lia b ilit y — C a n a d ia n p ra c tic e ( I A E A - S M - 2 1 8 / 2 5 ) ...... 177 R. I. Hodge, J.E. LeSurf, J.W. Hilborn D is c u s s io n ...... 197 Reliability considerations for L M F B R steam generators ( I A E A - S M - 2 1 8 / 2 7 ) ...... 199 G.A. de Boer, M. de Hes, P. W.P.H. Ludwig D i s c u s s i o n ...... 2 1 6 PW R steam generator tubing: corrosion problems (IA EA -SM -218/26) ...... 2 1 7 D. van Rooyen
QUALITY ASSURANCE AND QUALITY CONTROL (Session VII)
The contribution of quality assurance to safety and reliability in n u c le a r p o w e r p la n t s ( I A E A - S M - 2 1 8 / 5 1 ) ...... 2 3 7 N. R a isic Comparison of quality systems standards (IAEA-SM -218/49) ...... 2 4 9 J.L. Fowler D is c u s s io n o n p a p e r s I A E A - S M - 2 1 8 / 4 9 a n d 5 1 ...... 2 5 5 Quality assurance requirements for the reliability of nuclear power plants in developing countries (IA EA -SM -218/35) ...... 2 5 9 S. M. Bhutta D is c u s s io n ...... 2 6 7 Some aspects of quality control problems experienced in nuclear component manufacture (IAEA-SM -218/36) ...... 2 6 9 V .S.G . R a o D i s c u s s i o n ...... 2 8 2
INSPECTION AND TESTING (Session VIII)
A tentative approach to a more rational preparation of in-service inspection programmes (IAEA -SM -218/39) ...... 2 8 5 C. B. Buchalet, G. Martin, M. Vauterin D i s c u s s i o n ...... 2 9 2 Contröles non destructifs et methode devaluation des defauts en tant que moyen pour ameliorer la fiabilite des composants de reacteurs (IAEA-SM-218/37) ...... 2 9 3 A.C. Prot, R. Saglio, M. Asty, M. Pigeon Inspection en service du circuit primaire des reacteurs ä eau sous pression (IAEA-SM-218/38) ...... 311 M . T. Destribats, A.M. Touffait, M. Route, M. Pigeon, M. Asty, A. Samoel, R. Saglio Discussion on papers IA EA -SM -218/37 and 38 ...... 3 2 9 A method for estimating flaw detection probability from inspection data ( I A E A - S M - 2 1 8 / 4 0 ) ...... 331 D. H. Shaffer D is c u s s io n ...... 3 3 9 The reliability of ultrasonic inspection (IA EA -SM -2 18/41) ...... 341 N . F. H ain es D i s c u s s i o n ...... 3 5 6
ROUND TABLE DISCUSSION ...... 3 5 9
C h a ir m e n o f S e s s i o n s ...... 3 7 5 S e c re ta ria t o f th e S y m p o s i u m ...... 3 7 6 L i s t o f P a r t i c i p a n t s ...... 3 7 7 A u t h o r I n d e x ...... 391
S e s s i o n V
FRACTURE MECHANICS AND DESIGN CONSIDERATIONS C h a ir m e n
T . V A R G A Switzerland
P.M. PETREQUIN F r a n c e IAEA-SM-218/15
SUBCRITICAL CRACK GROW TH IN THE
LIGAMENT BETW EEN THE INSTRUMENTATION
RODS OF THE BBR PRESSURE VESSEL BOTTOM
G. M ARCI, E. BAZANT, H.R. KAUTZ Babcock Brown Boveri Reaktor GmbH, M a n n h e im , Federal Republic of Germany
A b stra c t
SUBCRITICAL CRACK GROWTH IN THE LIGAMENT BETWEEN THE INSTRUMENTATION RODS OF THE BBR PRESSURE VESSEL BOTTOM. A fracture mechanics fatigue analysis is made for an assumed crack emanating from the bore of an instrumentation rod. This assumed crack has partially penetrated the Inconel buttering of the 22 Ni Mo Cr 37 on which the structural Inconel welds are laid. Our analysis shows that the assumed crack could only penetrate 26% of the remaining ligament of the Inconel structural weld as a result of the fatigue crack growth during the entire operating life of the pressure vessel. Therefore a leak caused by a flaw missed during pre-service and in-service non-destructive testing can be excluded.
1. I N T R O D U C T I O N
The primary system of nuclear power plants is designed according to “safe life” - and “fail safe” concepts. The analysis to be presented is in three parts, to document the fulfilm ent of one of the “safe life” requirements, namely that a leak m ay not occur in a nuclear pressure vessel below its inlet and outlet nozzles. A n analysis has been made in three parts. The first part investigated the fatigue crack growth of an assumed crack in the beltline region and its effect on structural integrity [ 1 ] while the second part investigated the fatigue crack growth potential of stress relief cracks assumed to be present in the H A Z of structural welds of the pressure vessel [2 ]. This paper represents the third part : a crack is assumed present in an area where the geometric configuration leads to local stress concentration and where non-destructive testing is more complicated than in other regions of the lower half of the pressure vessel. The B B R pressure vessel bottom contains instrumentation rods which are welded into it. It is assumed that a crack is present underneath the structural weld between the vessel bottom and the instrumentation rod. The fatigue crack growth of this crack through the structural weld is investigated.
3 4 MARCI et al.
nozzle for control assembly drive
pressure vessel closure head
upper support plate
plenum assembly
control assembly guide tube
outlet nozzle inlet nozzle upper grid
core support assembly
core basket
fuel assembly
lower grid assembly
lower support plate
flow distributor head
incore instrument guide tubes
The assum ption is made that the crack is already present at the beginning of operating life and if the crack cannot penetrate half of the structural weld during the 40 years of operation, a leak can be excluded. It should be noted that the position selected and the crack size assumed in no way reflect a weakness of structural design, non-destructive testing or design calculations, but merely demonstrate the damage tolerance of the structure. IAEA-SM-218/15 5
Nozzle No. 2 Nozzle No. 1
FIG.2. Instrumentation rods in the bottom of the BBR pressure vessel (dimensions in mm).
2. DEFINITION OF THE PROBLEM
The B B R pressure vessel bottom contains instrumentation rods which are welded into it (Fig. 1). The pressure vessel, and therefore the bottom too, is made of 22 N i M o Cr 37 forgings. A circular area of approximetely 1 inch (25 m m) width and 3/4 inch (19 m m) depth is grooved out around the instrumen tation bore and buttered with Inconel. The structural weld for the attachment of the Inconel instrumentation rod is made with Inconel, too. For this sample problem the instrumentation nozzle No. 1 at the centre of the vessel bottom was selected. Figure 2 shows the geometry of the vessel bottom between the No. 1 nozzle and the nozzle closest to it. The Inconel buttering and Inconel structural welds are indicated. In Fig. 3a details of the No. 1 instrum entation nozzle are illustrated, with the assumed crack included. The crack is emanating from the generator of the cylindrical bore hole and its surface is parallel to it. This is shown in Fig. 3c. The assumed crack has a length “J2” equal to 0.59 in (15 m m) and a width “2a” equal to 0.394 in (10 m m ) and is of semi-elliptical shape. The centre of the crack lie s 1 .1 0 0 in (28 m m ) underneath the vessel’s inside surface; the crack periphery penetrates the Inconel buttering 0.120 in (3 mm). Therefore, the Inconel material remaining between the crack’s periphery and the inside surface of the pressure vessel has a width of 0.900 in (23 mm). 6 MARCI et al.
FIG.3. Assumed crack configuration in the pressure vessel bottom (dimensions in mm).
TABLE I. TRANSIENTS AND STRESS EXCURSIONS FOR THE BBR PRESSURE VESSEL BOTTOM
S t r e s s -E r c u rs io n s T ra n sie n t N u m b e r ot C y c le s M e m b ra n e S t r e s s e s B e n d in g S t r e s s e s N /m m Z k s i N / m m 2 k s i 1 255 143 20.732 106 15.368 II 80 23 3.335 115 16.673 III 510 37 5.364 106 15.368 IV 40 61 8.848 30 4.350 V 36000 11 1.595 37 5.364 VI 610 38.8 5.625 0 0 VII 40000 7.75 1.123 0 0
ksi = 6895 X 103 Pa IAEA-SM-218/15 7
TABLE II. DESCRIPTION OF TRANSIENTS AND RESPECTIVE NUMBER OF CYCLES FOR THE BBR PRESSURE VESSEL
T ra n Transient Description E x p e c te d D e s ig n sien ts C y c le s C y c le s
1 Heat-up and Cooldown 160 240
II Rapid Depressurization 50 80
III Power change between 325 510 0 % and 15 %
IV Rod Withdrawal Accident 25 40
V Power change between 24000 36000 8 % and 100 %
VI Reactor Trip 320 480 Loss of Feedwater to One Steam G e n e rato r 15 20 Loss ot Station Power 25 40 Control Rod Drop 25 40 Change of Flow 20 30 I'g T f f
VII 10 % Power Change 25000 40000
Table I gives the excursions in membrane and bending stresses for seven operating transients and their respective number of cycles for the centre region of the pressure vessel bottom. The description of operating conditions included . in each transient, the m axim um expected cycles and the cycles on which the follow ing analysis is based are given in Table II. In analysing the crack configuration as shown in Fig. 3 with respect to fatigue, it has to be realized that the stress intensity factor under a given loading situation varies around the periphery of the crack. In addition, with the problem described above, the crack fronts intersecting the nozzle bore are located in different materials. Therefore the lateral expansion of crack will be different on the two sides, even when the acting stress intensity factors are approxim ately equal.
3. THEORETICAL AND EXPERIMENTAL BASIS OF THE ANALYSIS
W ith fracture mechanics fatigue analysis the crack growth per cycle is put in relation to the excursions of the stress intensity factors,
— = function of ДК (1) d N 8 MARCI et al.
where Д К = K imax - K im in • The excursion of the stress intensity factors characterizes the variation of stresses and strains in close proxim ity of the crack front. There are m any equations correlating the incremental crack growth to ДК, but they generally have the disadvantage of being derived for particular loading spectra or a particular material. The Paris — Erdogan equation
da — =С(ДК)П (2)
is the most general equation available for which the input parameters “C ” (fitting parameter) and “n” (slope of the crack growth data) have been extensively investigated. The parameters “C ” and “n” are chosen such that Eq. (2) predicts fatigue crack growth rates which are upper bound curves for the material in question. For the material 22 Ni M o Cr 37 the A SM E Code, Section XI, Appendix A (Figure 4300-1) gives the values: for 22 Ni Mo Cr 37:
n = 3 .7 2 6 (3 a )
c = 2.67 X 1СГ 11 ( 3 b )
for an internal crack. From data available to us on fatigue crack growth of Inconel 600 (Fig. 4), values for the upper bound curve have been calculated.
for Inconel 600:
n = 3 .2 7 (4 a )
C= 6.87 X 1(T 10 ( 4 b )
Currently no closed form solution exists for the stress intensity factor for an elliptical crack emanating from a cylindrical hole. Shah [3] derived an approximate solution for this category of cracks emanating from a cylindrical hole and this solution is in very good agreement with the respective closed form solutions available for some of the crack configurations (Fig. 5). Shah’s solution gives the F IG .5. Stress in ten sity fa c to r com parison fo r cracks em anating from cylindrical holes. cylindrical from anating em cracks r fo parison com r to c fa sity ten in Stress .5. IG F
Fatigue Crack Growth Rate, da/dN, inch/cycle FIG .4. Fatigue crack grow th behaviour o f In con el 600. el con In f o behaviour th grow crack Fatigue .4. FIG tes nest fco rne AK, ( )^/2 /(m N M , K A range, factor intensity Stress -t-LJ > 101 K>’ ______15 /1 8 1 2 - M S - A E A I ■ I » ■ * » I 1- ■ ■ ■ 9 10 MARCI et al.
FIG. 6. Shah on cracks at fastener holes.
stress intensity factors along the crack periphery as a function of the ratio between the crack length “2” and the radius of the cylindrical hole. He then normalizes the stress intensity factors by dividing them through the respective stress intensity factors for an elliptical crack in an infinite solid with length “ 2 2 ” and the same width “2a”. The result is non-dimensional parameter which characterizes the increase in stress intensity factor due to the presence of the cylindrical hole. The graphic representation of these non-dimensional parameters is shown in Fig. 6 . IAEA-SM-218/15 11
Know ing the dimensions of the subject crack, namely “2a” and “2”, the stress intensity factor for an elliptical crack in an infinite solid (here assumed to be the pressure vessel bottom ) with dimension “ 2 a ” a n d “ 2 2 ” is calculated in accordance with A SM E Code, Section XI, App. A. The stress intensity factors calculated from the A SM E Code are then multiplied with the non-dimensional parameter obtained from Fig. 6 for the crack front intersecting the cylindrical hole and for the crack front on the longitudinal axis. In this way the stress intensity range for the leading edges of the crack can be determined and, using Eq. (2), the fatigue crack growth in these directions. These directions are identified in Fig. 3b, with “u” being the longitudinal direction, “v” being the direction of lateral expansion of the crack into the 22 N i M o Cr 37 and “w ” the direction of lateral expansion of the crack into the Inconel weld. The aspect ratio of the subject crack is fixed by assuming the crack configuration, but might change during crack growth. The non-dimensional parameters taken from Fig. 6 have therefore to be checked during the calculation. However, this does not cause a problem, since the change generally is toward the conservative side. The conclusion to be derived from this analysis is that the co-ordinate system, as shown in Fig. 3b, remains fixed for the calculation of the fatigue crack growth. A Simpson-type numerical integration is made with step sizes up to 500 cycles. After each step the half-width of the crack is revised. The crack growth in the directions “u”, “v” and “w ” are added to the original dimensions and the stress intensity factor is calculated, based on the instantaneous half-width of the crack.
4. DETERMINATION OF INPUT VALUES
Since the flaw size is revised after each integration step (max. 500 cycles), the follow ing equations are used for the determination of the stress intensity range for an internal elliptical crack.
D i A a m (5 a )
D 2 = Д а ъ М ь ( 5 b )
AKeiiipt. = (D i ■*" D 2) y / a ( 5 c )
A K j — AKeiiipt. Foi ( 5 d ) •12 MARCI et al.
The flaw shape parameter “Q ”, membrane correction factor “M m ” and the bending correction factor “M b ” are not revised but merely checked for conservatism at the end of the calculation. The same applies to the non- dimensional parameter F 0j of the Shah-solution. The flaw parameters are defined:
a = 0.197 in (5 mm)
1 *- 1.181 in (1* = 2 -1 = 30 m m)
a / * = 0 .1 7
e = 0 .3 2 t
t = 6 .2 2 in (158 mm)
2a/t = 0.064
2e/t = 0.65
R = 0.876 in (22.25 mm)
1/R = 0.674
The membrane correction factor “M m ” is obtained from A SM E Code, Section XI, Appendix A, using Fig. A - 3300-2. For the 2a/t value of 0.064 the membrane correction factor is determined to be approximately 1.05 for both crack fronts. The value for M m was chosen as 1.1. This value is conservative because it corresponds to a higher 2 a/t ratio than the actual one. The bending correction factor “M b ” is obtained from Fig. A — 3300-4 of the same Code. The value chosen for both crack fronts is 0.66. The flaw shape parameter Q was obtained from Fig. A — 3300-1 of the 0щ + Ob above-mentioned code, based on the assumed ratio ------= 0.8. Its value is 1.17. a y s The non-dimensional parameter obtained from Shah’s calculation (Fig. 6 ) b a s e d o n the ratio 1/R = 0.674 is:
For the “v” and “w ” directions:
F ov = F q w - 2.5 IAEA-SM-218/15 1 3
TABLE III. RESULTS OF THE F.M. FATIGUE ANALYSIS: ELLIPTICAL CRACK EMANATING FROM THE INSTRUMENTATION BORE
T ra n s. C y c le s ДК. Д К и lu 8v 8w I u ДК« R a d iu s
N /m m N /m m [m m ] [m m ] [m m ] S t a r t o( O p e ra t in g Life 15.0 5.0 5.0 0.674 1 255 2054 1356 15.146 5.690 7.765 0.681 II 80 1062 700 15.151 5.710 7.866 0.681 III 510 1166 770 15.186 5.878 8.735 0.683 IV 40 950 626 15.187 5.893 8.770 0.683 V 36000 413 272 15.240 6.134 10.851 0.685 VI 610 503 332 15.242 6.141 10.917 0.685 VII 40000 100 66 15.243 6.142 10.940 0.685
For the “u ” direction:
F ou = 1.65
The fitting parameter “C ” and the slope parameter “n” for the Paris - Erdogan equation for 22 Ni M o Cr 37 and Inconel 600 have been determined in the fo r e g o in g .
5. RESULTS AND DISCUSSION
The results of the calculation are shown in Table III. There the stress intensity ranges are shown for the transverse and longitudinal directions together with the crack dimensions in terms of the co-ordinate system “u”, “v” and “w ”, i.e. in relation to the original crack centre. To get a criterion for the safety against the development of a leak, a comparison is made between the size of the remaining 1 4 MARCI et al.
FIG. 7. Original and final crack configuration.
ligament between the crack front and the inside surface of the pressure vessel bottom crack initially and after 40 operating years. One can see from Table III that approximately 26% of the Inconel ligament has been penetrated by the crack as a result of fatigue crack growth. This is graphically illustrated in Fig. 7. The analysis contains some simplifications which are discussed in the following. Shah’s solution, used in the analysis, is for two cracks emanating from a hole, lying in one plane, with symmetry to the hole axis. For the subject problem with one crack emanating from a hole the following correction [3] should be applied.
2 R + 2 К one crack ------: X К two cracks (6) 2 R + 2 £
W hich gives for the subject crack-and-hole problem:
К one crack “ 0-9 К two cracks (7) IAEA-SM-218/15 1 5
Since the non-dimensional parameters of Shah’s calculations were used as shown in F ig . 6 for our analysis, i.e. the correction factor for only one crack as given by E q . ( 6 ) was neglected, the results of our calculation are conservative. The crack shape factor was obtained based on the ratio a/2ß= 0.17 which corresponds to Ch = 1.17 for am + 0 b =0.8 ays. The final crack configuration has a ratio a/2ß = 0.27 which gives a crack shape factor of Q 2 = 1.40. The change in shape factor has also been neglected in the present calculation, resulting in an additional conservatism. Shah’s solutions for cracks emanating from a hole were derived for plate structures and correlated to cracks in plate structures. The pressure vessel bottom has spherical configuration. Yet, comparing the pertinent dimensions (crack size and hole radius) to the radius of pressure vessel bottom , it is felt that the results would need m inor corrections only if the curvature of the bottom was taken into account. In addition, the calculation of the stress intensity factor according to ASM E-Code procedure incorporates to some degree the correction for curvature. One can assume that the non-dimensional parameters given by Shah are still valid if, for a crack emanating from a cylindrical hole penetrating a spherical shell, the stress intensity factor is calculated based on the reference crack in a spherical shell.
6 . CONCLUSION
The results presented in Table III and Fig. 7 show that the assumed crack can only realize half of the fatigue crack growth considered tolerable for sufficient safety against the development of a leak. Therefore, a leak in the pertinent areas cannot develop as a result of a flaw missed by present non-destructive testing m e t h o d s . In view of the non-destructive testing requirements, the flaw size and location assumed for the present analysis and the results of the analyses performed with respect to cracks in the beltline region [ 1 ] and the fatigue crack growth potential of stress relief cracks [2 ], it is concluded that a leak cannot develop in the lower half o f the pressure vessel.
REFERENCES
[1] PALME, H.S., “Evaluation of sample problem, CASE 1 — embedded flaw located in the beltline region wall of a PWR reactor vessel”, Babcock & Wilcox - Contributed Paper to Nondestructive Examination Conference, Washington, D.C., Nov. 18, 1976, sponsored by: Babcock & Wilcox and Technischer Überwachungsverein Rheinland. [2] MARCI, G., “Beurteilung der durch das Schweißverfahren eventuell verursachten Risse in der Wärmeeinflußzone”, Reaktortagung 1977, Mannheim, March 29 — April 1, 1977, (1977) 756. [3] SHAH, R.C., “Stress intensity factors for through and part-through cracks originating at fastener holes”, (Proc. 8th Nat. Symp. on Fracture Mechanics) ASTM, STP 590, p. 429. 16 MARC! et al.
APPENDIX
NOMENCLATURE
d a / d N Crack growth per cycle KImin M i n i m u m stress intensity factor during a cycle Kim ax M a x i m u m ДК Excursion of the stress intensity factor during a cycle c F it t in g - parameter of the Paris-Erdogan equation n S lo p e - u, v, w , Co-ordinates used to identify crack growth Q Crack shape parameter °m> M e m b r a n e - stress and its excursion during a cycle ОЪ, А о ь B e n d in g - Oys Yield stress Foi Non-dimensional parameter correlating K i for a crack emanating from a hole to K i for the same crack in an undisturbed structure. IAEA-SM-218/16
THE GROWTH OF UNDERCLAD FLAWS IN FATIGUE LOADING
K. RAHKA, J. FORSTEN Technical Research Centre of Finland, E s p o o , F in la n d
A b stra c t
THE GROWTH OF UNDERCLAD FLAWS IN FATIGUE LOADING. In this report attention is focused on the influence of possible underclad flaws on the fatigue performance of a cladded component. A comparison is made between the well-known ASME III Sa-N fatigue design curve for essentially flawless parts with fatigue design curves for flaw-containing welds proposed for inclusion in an acceptance standard by Harrison, Burdekin and Young. The performance of a cladding under which a simulated flaw has been placed has been evaluated. The results are in conformity with quality class performances given by Harrison, Burdekin and Young. The incremental polynomial technique has been applied in the evaluation of fatigue crack growth rate, which has been plotted against the stress intensity factor range for the very final stages of cladding rupture also.
1. I N T R O D U C T I O N
According to A SM E III (ASM E, 1974) [1 ] the number of fatigue load cycles to failure in a region of a com ponent is determined by the alternating shear stress, which is evaluated principally according to the m axim um shear stress hypothesis. For some cases the value of the “stress amplitude” Sa used in A SM E III is equal to a norm al stress amplitude, as for instance in a closed-end pressure vessel, where the only fluctuating loads are induced by a fluctuating pressure, and the stress observed is the hoop stress. For such cases a comparison between postu lated fatigue performances of unflawed (A SM E III) and flaw-containing vessel walls [4] can be made (see Fig. 1). In the comparison it should be noted that the position of the A S M E III curve is influenced by safety factors which have been applied on test result mean values, whereas the curves given by Harrison et al. are lower bound estimates for test results without any safety factors. It is noted that the occurrence of “worst class flaws (W — Z )” would probably render the A SM E III fatigue performance estimate optimistic for such an area. Figure 2 shows a definition of the weld quality classes V - Z according to Harrison, 1974 [4]. As underclad flaws m ay exist and possibly lead to premature failure of the cladding, thus exposing more water-sensitive base metal to the coolant, this investigation was performed in order to clarify details of such, possibly hypothetical, cladding performance.
17 18 RAHKA and FORSTEN
FIG.l. Fatigue design curves for welds with internal flaws according to a British Standard proposal [5]. A modification (see text) o f a part o f the ASME III design curve for carbon and low alloy steel (UTS < 560 Nmm~2) is given for comparison.
FIG.2. The maximum acceptable flaw size, 2a, for different section thicknesses in the quality classes, W, X, Y, Z according to Harrison [4\ No flaws are accepted in the quality class V. IAEA-SM-218/16 19
3 - LAYER CLADDING L
L / s r T c 60 FLAW, 4
______thickness 10
300______
FIG.3. Specimens used in the tests (dimensions in mm).
FIG. 4. Experimental set-up. The specimen (on the left) is fixed into the fatigue-testing machine and illuminated by a stroboscope. The crack growth is measured with a travelling microscope.
2. EXPERIMENTAL DETAILS
Four panel specimens of the configuration shown in Fig. 3 were fatigued until the cladding failed in a servo-hydraulic testing machine at room temperature at a sine wave loading frequency of 10 Hz. The initial flaw falls into the Z class, according to Fig. 2. The test set-up, which also shows the crack length measuring device, i.e. a travelling microscope, is shown in Fig. 4. The ends of the specimens 2 0 RAHKA and FORSTEN
FIG. 5. Flaw size in the 10-160 kN fatigue loading as a function of the number of loading cycles. The position of the mid-point (m) of the flaw is indicated by the dashed line.
FIG. 6. Flaw size in the 50-160 kN fatigue loading as a function of the number of the loading. The position of the mid-point of the flaw is indicated by the dashed line.
were rigidly fastened to the grips of the testing machine, resulting in an even gross-stress distribution over the specimen cross-section. The material of the test specimen was a Cr-M o-V steel cladded with three layers o f austenitic stainless steel [2 ].
3. TEST RESULTS
The prim ary test results for the four tested specimens are shown in Figs 5, 6 , 7 a n d 8 . The figures show the locations of the two crack fronts as a function of the number of fatigue cycles. The m idpoint (m ) of the instantaneous flaw is indicated by the dashed line. In Fig. 9 a plot is shown where the number of cycles to rupture of the cladding is shown as a function of applied stress range. IAEA-SM-218/16 21
FIG. 7. Flaw size in the 90-160 kN fatigue loading as a function o f the number o f loading cycles. The position of the mid-point of the flaw is indicated by the dashed line.
The numbers given refer to the instantaneous (initial) flaw width (2a) from which the number of cycles is counted. It is seen that the results fall well on the safe side of the design curves given by Harrison et al [3]. The locations of the points which give the number of cycles to failure for smaller initial flaw sizes have been estimated using a simple fracture mechanics calculation based on a (da/dN) — Д К straight line approximation and Д К esti mation rules according to A SM E X I (ASM E, 1974) [1 ]. In order to evaluate the material’s resistance to crack extension in fatigue, the instantaneous fatigue crack growth rate was calculated for the part of the results which belongs to cladding rupture, using the incremental polynom ial technique developed under the supervision of the A ST M E-24.04.01 Task Group in the U SA [8 ]. This part of the a-N results is shown with an enlarged N-axis in Fig. 10. The (da/dN) values thus obtained are 'shown in Figs 11,12 and 13 as a function of the stress intensity factor range ДК, which was calculated according to the formula, (see Ref. [7])
(1) В • 6 0 m m 22 RAHKA and FORSTEN
FIG. 8. Flaw size in the 50-120 kN fatigue loading as a function of the number of loading cycles. The position of the mid-point of the flaw is indicated by the dashed line.
FIG. 9. The number of cycles to perforation of the cladding as affected by the initial flaw size (2a) and applied tensile stress range. The symbols are:
о load range 10-160 k.N (Fig. 10) □ load range 90-160 k.N (Fig. 12) л load range 50-160 kN (Fig. 11) A load range 50-120 kN (Fig. 13). IAEA-SM-218/16 2 3
FIG. 10. Crack front positions as a function o f cycles for the 50-120 kN, 90-160 kN, 50-160 kN fatigue-loaded specimens for the cladding rupture part.
where ДР = load range В = specimen thickness a = half width of flaw b = distance between m idpoint of flaw and outer cladding surface
This form ula gives the stress intensity factor with a mathematical accuracy of 0.5% for any a/b value, specifically for a centre cracked sym m etric panel. It is assumed that consideration of the drifting of flaw m idpoint also gives the same degree of accuracy for the present more irregular case. For comparison, the Д К value was evaluated for the crack penetrating into the base metal of the specimen loaded 90 kN-160 kN, (Table II), using for b the distance between the flaw m idpoint and the specimen edge on the base metal side of the crack. This difference seems obvious, although a more accurate analysis has not been made. The (da/dN) — Д К values fall into a general area for steels, which justifies the statement that the base material in itself shows adequate resistance against fatigue crack growth. 2 4 RAHKA and FORSTEN
FIG.11. da/dN versus AK for the 50-160 kN fatigue-loaded specimen. The dm /dN value on the top o f the figure shows the rate o f flaw m iddlepoint drift. IAEA-SM-218/16 25
FIG.12. da/dN versus AК for the 90-160 kN fatigue-loaded specimen. The dm /dN value on the top o f the figure shows the rate o f flaw m iddlepoint drift. 6 2 he top of he gure he rat flaw middlepoint dri t. if r d t n i o p e l d d i m w a l f f o te a r e th s w o h s e r u ig f e th f o p o t e th n o G.13. for t fat oaded speci The dm/dN ue lu a v N d / m d e h T . n e im c e p s d e d a lo - e u g ti a f N k 0 2 1 - 0 5 e th r o f K A s u s r e v N d / a d . 3 1 . IG F t o -0.01 Q О s £ - £-0.001 2 e ш *D О _l TJ_l II e Ü
CRACK GROWTH RATE da/dN (цт/cycle) 0.001 0.01 0.021 01 - -0.1 0.1 0.1 7 7 1 - 0 1000 100 TESITNIYFCO RNE AKlkNmm-^ INTENSITY RANGE STRESS FACTOR ------AK n FORSTÜN and RAHKA 1 ______t А А t T ä / I ______/ А А А LOAD 50-120 LOAD kN I /2) 10000 7 7 IAEA-SM-218/16 27
Figures 11,12 and 13 also show the velocity dm /dN with which the flaw m idpoint drifts, giving the positive direction towards the cladding and vice versa. Addition of this dm /dN value to the da/dN value gives the velocity with which the crack front approaches the cladding surface. Subtraction will give the velocity by which the opposite crack edge penetrates into the base metal. Tables I, II and III show the computed results.
TABLE I. CRACK GROWTH ANALYSIS RESULTS FOR THE 50-160 kN FATIGUE-LOADED SPECIMEN
HESSU2
N Z*A DA/PN DM /ON DK
9500 14.3318 1 • 5480 1E-04 - 4 .1B177E-05 1003.93
11200. 14.8586 1•689918-04 -3.98079E-05 1033.29
12100 15.1589 1 •779548-04 -3.72928E-05 1050.46
1Л200 15.9563 2 .07489E-04 -3.20963E-05 1098.35
15000 16.2711 2 .15867E-04 -2.79132E-05 1118.42
16B00 17.(1532 2.39596E-04 -1 .65192E-05 1170.45
18700 18.0261 2.83548E-04 - 1 .12786E-05 1239.49
19900 18.7304 3 .19692E-04 4. 16156E-07 1293.73
20800 19.3164 3.49873E-04 1.2Z601E-05 1342.45
21800 20.0461 4 ■08182E-04 5 .01952E-05 1409.11
22200. 20.375 4.45663E-04 7.95270E-05 1442.14
22500 20.6365 4 .7B242E-04 1.05993E-04 1469.89
23100 21.231 5 •387B1E-04 1.54307E-04 1539.ei
23500 21.6859 5.60713E-04 1.68090E-04 1600.61
23700 21.9192 5.B8431E-04 1•91734E-04 1634. 12
24000 22.2679 6.24126E-04 2.21316E—04 1687.16
24300 22.5983 7.01912E-04 2.92989E-04 1740. 1
24800 23.3792 9•34453E-04 5. 15343E-04 1899.43 28 RAHKA and FORSTEN
TABLE II. CRACK GROWTH ANALYSES RESULTS FOR THE 90-160 kN FATIGUE-LOADED SPECIMEN
HESSU2
N 2*A d a / on DM/ON DK
11700 1 1.8594 2.69139E-05 2 .182B6E-05 555.74
15500 12.0921 3.26547E-05 1.973B6E-05 564.627
16800 12. IBB 3.27955E-05 1.72005E-05 568.305
18400 12.295 3.69604E-05 1.80681E-05 572.376
22000 12.574 4. 14663E-05 1.51554E-05 5 8 3 .06Б
28000 13.0956 5 .00949E-05 1 .14195E-05 603.203
32000 13.5268 5 .54743E-05 В. 55605E-06 620.37
33500. 13.6956 5.51797E-05 5 .17034E-06 627.174
35500. 13.9407 5.62416E-05 2.11076E-06 637.242
37600. 14.1643 5.73539E-05 -1.1 0449E-06 646.203
39500. 14.3778 5.90484E-05 -3.32539E-06 654.852
42000. 14.6732 6 • 27298F.-0 5 -4.795Я2Е-06 666.968 (565.1
43700. 14.8997 6 .63994E-0 5 —4 .62950E-06 676.575
47200. 15.3956 7.45323E-05 -3.70922E-06 6 9 B .16
48900. 15.6495 7•9 1790E-0 5 -2 .5657BE-06 709.498
52100. 16.1752 8.6B072E-05 -1 .53192E-06 734.046
52700. 16.2B7 B.65672E-05 -3 .00835E-06 739.561
53700. 16.4638 8.81923E-05 -3 .44400E-06 748.219
56200. 16.925 9 .6 3 7 9 1E-05 -4 •09076E-07 771.573
57400. 17.1553 9 .75927E-05 -1 .66838E-06 783.66
58100. 17.2947 9 .79603E-05 -2.74330E-06 791.173
50400. 17.3702 1.05338E-04 4.01593E-06 795.585
61200. 17.9253 1.02705E-04 -4.3B693E-06 B26.452
63700. 18.4316 1.05318E-04 -6.925B7E-06 8 5 5 .9Э
65200. 18.7431 1 .10061E-04 -5.27421E-06 874.911
6660 0• 19.0523 1 .17903E-04 -3 . 17130E-07 894.763
68000. 19.382 1 1.27B90E-04 6.7B507E-06 917.27
69300. 19.7403 1.37341E-04 1•35569E-05 944.508
70500. 20.071 1 .37239E-04 1.09819E-05 970.955
71100. 20.247 1.44089E-04 1 • 65957E-05 985.968
72200. 20.5523 1.51433E-04 2 .16732E-05 1012.39 IAEA-SM-218/16 29
t a b l e i i ( c o n t . )
HESSU2
N 2*A DA/DN DM /ON DK
73800. 21.0433 1.64577E-04 3 .15202E-05 1059.95
74600• 21 .3099 1.71006E-04 3•62999E-0 5 1088.73
75700. 21.6997 1.86555E-04 4 .95827E-05 1135.16
76700. 22.1016 2.004 15E-04 6. 13B20E-05 1190.73 (770.)
77200. 22.293 2 .23446E-04 8.33822E-05 1219.01
77600. 22.4669 2.23938E-04 8 .30496E-05 1247.47
78400. 22.8656 2.48262E—04 1.05725E-04 1322.95
79000. 23.1726 2.63694E-04 1 .19921E-04 1390.39
79800. 23.5992 2 .85766E-04 1 .40344E-04 1502.5
80200. 23.8032 3.39855E-04 1.93609E-04 1564.4
80900. 24.297 3.95362E-04 2.47674E-04 1768.66
8 1 2 0 0 . 24.5567 4 .18896E-04 2.70589E-04 1917.77
4. DISCUSSION
It was pointed out in the introduction that a fatigue analysis based on crack initiation and performed according to A SM E III could potentially lead to opti mistic results if flaws exist in the structure. The results of this investigation show that data exist for the analysis of the influence of underclad flaws using the approach of Harrison et al. and taking the perforation of the cladding as the fatigue failure criterion. When assessing an A S M E -III evaluation it is important to remember, however, that fatigue failure according to A S M E III is dictated by the number of shear stress fluctuations whereas the fatigue failure according to Harrison et al. is dictated by the number of normal stress fluctuations. The comparison between the A S M E -III analysis and the analysis of Harrison et al. can thus be made only when these stress values are technically the same. Estim ation of the applicability of the present results in the assessment of the performance of a pressure vessel or a structure in a service environment must take into account the possible combined effects of crack propagation at high temperature and of cyclic loading at low frequencies, which at present seem to be largely unknown. This question has been treated recently by Pense and Stout [5], and it appears that an enhancement factor of 1.5 might be applied to the room temperature crack growth rates in order to get representative values 3 0 RAHKA and FORSTEN
TABLE III. CRACK GROWTH ANALYSIS RESULTS FOR THE 50-120 kN FATIGUE-LOADED SPECIMEN
HESSU2
N 2*A DA/DN DM/ON DK
46(100. 15.7388 4.32712E-05 -B.77245E-06 684.515
74000. 18.6157 6 .18880E-05 -1.43548E-05 600. 14
76000. 18.8449 6.86785E-05 -1 .21764E-05 810.282
79000. 19.2616 6.95761E-05 -1.46417E-05 829.332
82000. 19.6629 7.33094E-05 -1 .42713E-05 847.565
87000. 20.4421 8.90866E-05 -9 .73816E-06 885.925
90000• 21.0184 1•09161E-04 -9.93e33E-06 916.513
94000. 21.9809 1•29423E-04 - 5 .0 335 1E-06 9 7 4 .70Б
95300. 22.2731 1.44950E-04 5 .1037BE-06 992.008
97000. 22.8165 1.59131E-04 1.22349E-05 1031.45
98000. 23.1439 1.68629E-04 1•75862E-05 1056.66
100000. 23.8322 1.91130E-04 3 .17946E-05 1114.38
100500. 24.0236 1.94572E-04 3 .31635E-05 1132.23
101200. 24.304 2.05987E-04 4 .16758E-05 1159.94
101700. 24.5178 2 .15764E-04 4.93787E-05 1182.59
102900. 25.0383 2•29894E-04 5 .B5332E-05 1242.23
103800. 25.4588 2 .59087E-04 8 .39940E-05 1297.15
104400. 25.7882 2.61535E-04 8. 39543E-05 1347.25
105000. 26.1185 2•73 086E-04 9.30174E-05 1402.6
105400. 26.3162 2.86791E-04 1.05063E-04 1436.96
106400. 26.8986 3.45134E-04 1.59260E-04 1559.92
107000. 27.308 4.1 1719E-04 2 .23357E-04 1670.55
107500. 27.7327 4•6968 IE-04 2.79246E-04 1820.6
for the estimation of structural performances at 300°C. According to Pense and Stout, however, only a m inor reduction o f fatigue life is due to the high service temperature. It is evident that several other factors besides those described above can affect the fatigue performance of structures or pressure vessels with underclad cracks, but these are outside the scope of this report. Such factors include vacuum environment, variation in flaw location with respect to cladding weld passes, variation in flaw shape, m ulti-axial and variable loadings, etc. As the IAEA-SM-218/16 31 da/dN values obtained with relatively thin specimens have been found to be in conform ity with data obtained with thick sections [6 ], a judgement of thick section behaviour is justified based on the present da/dN values. It could be argued, however, that the effects of any o f the factors listed above are m inor provided that the stress intensity factor and flaw shape at the start of fatigue loading are comparable.
5. CONCLUSIONS
A qualitative assessment of the significance o f underclad cracks has been made in this paper. The fatigue crack growth rates of simulated underclad cracks in a Cr-M o-V steel cladded with three layers o f austenitic stainless steel have been investigated, and the follow ing conclusions are reached:
Fatigue behaviour of underclad cracks prior to access of primary coolant water to the crack tip in the base metal can be reasonably well predicted by the design curves of Harrison et al. [3]. Higher fatigue mean-stress levels lead to more rapid damage in the cladding
than lower levels. Due attention should be focused on possible optimism in an A SM E -III fatigue evaluation as affected by possible flaws in the area analysed.
ACKNOWLEDGEMENTS
The authors thank Mrs. Hilkka Vainikainen for drawing the figures, Mrs. Maija Virtaniemi and Miss Ulla Vartiainen for typing the manuscript and Mr. Heikki Saarelma for doing the crack growth analysis computations.
REFERENCES
[1] ASME, Boiler and Pressure Vessel Code, 1974, (a) Section III, Nuclear Power Plant Components, (b) Section XI, Rules for In-service Inspection. [2] HAKALA, J., HAKKARAINEN, T., JUVA, A., On the effect of microstructure on the corrosion behaviour of an overlay welded austenitic stainless steel cladding. Presented at the IIW-meeting in Tel Aviv (1975). [3] HARRISON, J.D., BURDEKIN, F.M., YOUNG, J.G., A proposed acceptance standard for welded defects based upon suitability for service, 2nd Conf. on Significance of Defects in Welds, London, 29-30 May, 1968. [4] HARRISON, J.D., British Work on the Significance of Weld Defects, Svetsen 33 (2) 23. 32 RAHKA and FORSTEN
[5] PENSE, A.W., STOUT, R.D., Elevated temperature fatigue of pressure vessel steels, Welding Journal 54 (8) Research Supplement (1975) 247. [6] SULLIVAN, A.M., CROOKER, T.W., The effect of specimen thickness upon the fatigue crack growth rate of A516-60 pressure vessel steel, ASME Paper No.76-PVP-57. (1976). [7] TADA, H., PARIS, P., IRWIN, G., The Stress Analysis of Cracks Handbook, Del Research Corporation, Pa. (1973). [8] CLARK, Jr., W.G., HUDAK, Jr., S.J., Variability in Fatigue Crack Growth Rate Testing (ASTM E-24.04.01) Task Group Report, Scientific Paper 74-1E7-MSLRA-P2, Westing- house Research Laboratories, Pittsburgh, Pa. (1974). IAEA-SM-218/43
ANALYSE TRIDIMENSIONNELLE DE LA PROPAGATION EN FATIGUE D’UN DEFAUT DE FORME SEMI-ELLIPTIQUE
J.M. BOISSENOT Centre technique des industries mecaniques, S e n lis, F ra n c e
Abstract-Risu m£
A THREE-DIMENSIONAL ANALYSIS OF FATIGUE PROPAGATION OF A SEMI ELLIPTICAL FAULT. The study of the propagation of a semi-elliptical fault in a three-dimensional elastic environment is based on the concept of a maximum rate of variation in potential energy associated with a fatigue law. With this technique one can follow the development of the fault as far as the critical size, under various types of stress. The technique shows that the shape of the crack front, the critical size of the flaw and the number of cycles to fracture are closely linked with the type of stress to which the cracked piece is subjected.
ANALYSE TRIDIMENSIONNELLE DE LA PROPAGATION EN FATIGUE D’UN DEFAUT DE FORME SEMI-ELLIPTIQUE. L’etude de la propagation d’un defaut de forme semi-elliptique dans un milieu elastique tridimensionnel est basee sur un concept de taux maximum de variation d’energie potentielle associe ä une loi de fatigue. Cette methode permet de suivre revolution du defaut, jusqu’ä la taille critique, sous Faction de divers types de soilicitations. Elle met en evidence le fait que la forme du front de fissure, la taille critique du defaut ainsi que le nombre de cycles a rupture sont etroitement lies au type de sollicitation auquel est soumise la piece fissuree.
INTRODUCTION
L ’etude de l’admissibilite et de la propagation de defauts tridimensionnels dans les pieces mecaniques en service ont fait l’objet de travaux recents dans divers dom aines de l’industrie et notam m ent dans le cadre de la sürete nucleaire [1— 5]. Ces travaux necessitent souvent la mise en oeuvre de programmes de calculs puissants capables de determiner avec precision des parametres de rupture caracteristiques d’une fissure sous Taction de soilicitations statiques ou cycliques. Dans la presente etude, on se propose de suivre par le calcul la propagation d’un defaut debouchant, de forme semi-elliptique, soumis ä un chargement cyclique. Une serie de calculs tridimensionnels dans le domaine elastique bases
33 3 4 BOISSENOT sur la methode des equations integrales permet dans un premier temps de determiner l’energie potentielle du Systeme en fonction des parametres geometriques du defaut et d’en deduire un taux de variation d’energie potentielle G. Afin de suivre la progression du front de fissure en fonction du nombre de cycles, on utilise une loi de fatigue reliant la surface balayee par cycles au parametre global G caracteristique du defaut. Cette loi, determinee experimentalement, est introduite dans le calcul sous une forme analytique approchee. On definit ensuite, pour chaque accroissement dS de la surface du defaut, un ensemble de parametres geometriques tendant ä minimiser l’energie potentielle du Systeme, de sorte que Гоп est amend ä resoudre simultanement un Probleme d’integration d’equation differentielle lie ä la loi de fatigue utilisee et un Probleme d’optim isation lie ä l’dquilibre du Systeme.
1. DESCRIPTION CINEMATIQUE
On considere un domaine tridimensionnel dlastique arbitraire V representant une piece mecanique ou une partie de cette piece, limite par une surface 2 et contenant une fissure plane schematisee par une surface singuliere S. On suppose que les sollicitations auxquelles est soumis le domaine V sont telles que le defaut se propage dans son plan et que la surface S est determinee ä chaque instant par un nombre fini de parametres.
S = ^ ( a b a2, . . . an) (1)
S o it^ (V ) l’energie potentielle relative au domaine V. Dans le cas de l’elasticite classique et en l’absence de forces de volume, ^ ( V ) s’ecrit:
(2)
oü W represente la densite d’energie de deformation, u; les deplacements et 2 j la partie de 2 sur laquelle sont imposees les tensions Tj. Lorsqu’au cours de la propagation le front de fissure balaye une surface dS, il existe en general, pour un m im e accroissement dS, une infinite de solutions aj verifiant les relations:
S = S 0 + d S
(3) S = .^(aj, a2, . . . , an)
ai = a° + daj, avec da, > 0 IAEA-SM-218/43 35
De ce fait il est necessaire d’introduire un concept supplem ental permettant de connaftre la variation reelle d’energie potentielle associee, fonction des aj, afin d’en deduire revolution de la forme du front de fissure.
1.1. Critere devolution
Ce concept, propose par Lemaitre [6 ] est base sur le principe du travail m axim um de Hill. Lemaitre a etabli ce concept dans toute sa generality mais, si on se limite au cas de l’elasticite, le critere s’enonce de la facon suivante: «Parmi tous les accroissements possibles de fissures de т ё т е aire dS, l’acroisse- ment reel est celui qui m aximise le taux de variation d’energie potentielle». Le taux de variation d’energie potentielle etant defini comme la borne superieure de la variation dP d’energie potentielle par rapport ä l’accroissement dS de la surface du defaut,
G = Sup [—(dP/dS) (ab a2, . . . , an)] (4) *4 relation dans laquelle les aj verifient les relations (3). II reviendrait au т ё т е de dire que parmi tous les accroissements possibles verifiant les relations (3), on choisit celui qui tend ä minimiser l’energie potentielle.
1.2. Loi de fatigue
L’evolution du defaut sous faction du chargement (ou de l’environnement) est regie par une loi de fatigue reliant la variation de surface S par rapport au nombre de cycles (ou au temps) ä plusieurs parametres, dont le taux de variation d’energie potentielle.
d S / d N —& [ ( G ,...), o u b ie n (5)
S = J ^ (G , . . . )
D ’une fa?on generale, ces lois ne sont pas tres connues et on est souvent conduit ä utiliser des formules plus ou m oins empiriques determinees ä partir de courbes experimentales, mais il est toujours possible d’associer le critere devolution du front de fissure ä la loi de propagation en fatigue du defaut et de resoudre simultanement les equations relatives ä la m inim isation de l’energie potentielle d’une part et l’equation differentielle regissant la fatigue d’autre part. 3 6 BOISSENOT
1.3. Critere de rupture
Si de plus on admet que le defaut atteint une taille critique lorsque le parametre G atteint une valeur critique G c, le Probleme de la propagation dans le cas d’un chargement cyclique est completement defini par les relations suivantes
dS/dN= ^ (G , . . . )
G = Sup [—(dP/dS)] ( 6 )
G < G C
avec c^”et G c donnes, P, S et aj definis par les relations (2) et (3).
2. CAS D’UN DEFAUT DE FORME SEM1-ELLIPTIQUE
On considere un defaut de forme semi-elliptique dans une plaque epaisse soumise ä un chargement cyclique (fig. 1). Si l’on suppose que le defaut reste de forme semi-elliptique au cours de sa propagation, le nombre de parametres definissant le front de fissure est limite ä deux, par exemple le petit axe b et le grand axe a de l’ellipse consideree. Dans ce cas on obtient pour le critere devolution les conditions suivantes:
dS = (П/2) (ab - a 0b 0)
a = a 0 + da da > 0
b = b0 + d b db > 0 (7)
G = Sup [—(dP/dS)] a,b
G < G C
critere auquel on associe une loi de Paris du type:
dS/dN =С(ДК)т ( 8 )
ой C et m sont des constantes du materiau, К = [EG/(1 — v2)]1/2 dans le cas d’un chargement cyclique simple (charge-decharge) et avec l’hypothese qu’il у a globalement au voisinage du defaut un etat de deform ations planes. IAEA-SM-218/43 37
FIG.l. Defaut de forme semi-elliptique dans une plaque epaisse soumise ä un chargement cyclique.
2.1. Calcul par la methode des equations intögrales
La resolution des equations (7) et ( 8 ) passe par la connaissance de l’energie potentielle en fonction de a et b. On est done amend ä faire une Serie de calculs pour differentes formes de defaut dans le but de construire un polynom e de deux variables P(a, b). Ces calculs ont ete menes par la methode des equations integrales qui consiste, dans son principe, ä transformer les equations locales qui decrivent le comportement des fonctions ä l’interieur du domaine (deplacement u ou traction t imposes) en une equation integrale liant les fonctions inconnues et certaines de leurs derivees aux fonctions connues en surface [7—9]. Autrement dit, pour tout point x de la surface lim itant la structure consideree, l’equation integrale relie les deplacements et les tensions en tous les points de cette surface; certains de ces termes sont connus par les conditions aux limites, les autres sont les inconnues du probleme. Apres resolution, on obtient les deplacements, les tensions et les contraintes en tous les points de la surface et une relation donne les memes inform ations en 38 BOISSENOT
tous les points du domaine. L ’avantage de cette technique est evidente puisqu’elle necessite uniquement la «discretisation» de la surface. On remplace ainsi un Probleme ä trois dimensions en un Probleme ä deux dimensions, ce qui simplifie la mise en donnees et l’exploitation des resultats. La discretisation de la surface de la structure est realisee ä l’aide d ’elements quadrilateraux ä 24 degres de liberte, avec variation parabolique de la geometrie et des fonctions inconnues. En plus des deplacements et des contraintes en chaque noeud du maillage de la piece (fig. 2 ), le programme fournit l’energie potentielle pour chaque geometrie de defaut. Dix ä quinze geometries sont necessaires pour etablir le polynom e P(a, b) avec une bonne precision. IAEA-SM-218/43 3 9
FIG.3. Front de fissure.
2.2. Analyse des resultats de calcul
Com m e on peut le voir sur la figure 2, qui represente un quart de la plaque, le front de fissure est borde de deux couches d’elements de surface isopara- m e triq u e s. Conform em ent ä une idee developpee en elements finis [10] puis en equations integrales, il est possible d’introduire la singularite en y^r du champ de deplacements au voisinage du front de fissure en mettant le point «milieu» au quart du cöte perpendiculaire au front de fissure (fig. 3). On introduit ainsi dans les fonctions de forme une singularite du т ё т е type qu’en mecanique de la rupture. La figure 4 montre differentes solutions obtenues pour le facteur d’intensite de contraintes le long du front de fissure. La courbe du dessus correspond ä la solution obtenue par extrapolation des deplacements lorsque les points «milieu» sont au quart du cöte sur les elements situes de part et d’autre du front de fissure. La courbe mediane correspond ä la solution obtenue avec des elements standard (points au milieu). La solution qui consiste ä utiliser des elements singuliers (points au quart) du cöte de l’ouverture et des elements standard du cöte de la matiere donne des resultats intermediaries. La troisieme courbe, qui donne les resultats les plus faibles, correspond ä la solution semi-elliptique [11]. La difference qui existe entre cette solution et les solutions numeriques est due en grande partie ä des effets de bords car les dim ensions du defaut sont grandes devant les dimensions de la plaque. II est interessant de calculer G par l’intermediaire du facteur d’intensite de contrainte et de le comparer ä celui obtenu directement par derivation de l’energie potentielle. 4 0 BOISSENOT
FIG.4. Differentes solutions obtenues pour le facteur d ’intensite de contraintes le long du front de fissure.
------en deformations planes 1 - p2 G K = K 2/E' E' = E en contraintes planes + S
G k = Y s J GKdS (9) -S
On obtient pour un defaut de rapport profondeur sur epaisseur egal a 0,3 et de rapport b/a egal a 0 ,6 :
5 ,6 0 X 1 0 -2 Мра X m (elements singuliers)
G K = 5 ,1 9 X 1 0 ~2 Мра X m (elements standards)
3 ,7 3 X 1 0 ~2 Мра X m (solution analytique) IAEA-SM-218/43 41
La solution analytique n’est donnee ici qu’ä titre indicatif puisque les effets de bords sont preponderants. Par ailleurs, la solution obtenue directement de la derivation de l’energie potentielle donne G = 5,69 X 1СГ 2 Мра X m. Cette valeur est pratiquement independante du choix des elements au voisinage de la fissure; par contre les solutions obtenues localement en dependent fortement, et leur comparaison montre que l’utilisation d’elements singuliers fournit une meilleure p re c isio n .
2.3. Etude de la propagation du defaut
La resolution des equations (7) et ( 8 ) necessite la mise en oeuvre d’un programme sur ordinateur qui calcule le polynom e P(a, b) ä partir des valeurs obtenues par la methode des equations integrales, calcule G par une methode iterative et integre 1’equation differentielle relative ä la fatigue par une methode de Runge-Kutta d’ordre 4.
Effet du chargement
Partant d’une forme initiale semi-circulaire, on etudie revolution du defaut sous Paction de trois chargements cycliques simples (charge-decharge): traction, traction + flexion, pression.
La figure 5 montre revolution du defaut sous Paction de ces trois types de sollicitations jusqu’ä rupture (ligne R), les lignes 1,2, 3, . . . indiquent la position du defaut au bout de 100 000, 200 000, 300 000, . . . cycles. Toutes ces valeurs dependant evidemment des constantes du materiau choisi pour le c a lc u l.
Dans le cas de la traction (fig. 5A) la forme du front tend vers une demi ellipse de rapport b/a = 0,6 ä la rupture qui intervient au-delä de 500 000 cycles. Dans le deuxieme cas de charge, ou l’on superpose ä la traction un moment ' de flexion, on constate que le nombre de cycles ä la rupture diminue de facon sensible puisqu’il est inferieur ä 400 000; on observe egalement que le front progresse moins vite en profondeur que sur les bords. Le rapport b/a au moment de la rupture est de l’ordre de 0,5. Pour le troisieme cas de charge, on observe que l’effet d’une pression exterieure est tres important car le nombre de cycles ä rupture tombe ä 2 2 0 0 0 0 ; par ailleurs, le defaut se propage essentiellement sur les bords.
Effet de la loi de fatigue
Partant d’une taille de defaut beaucoup plus petite et de forme semi- elliptique, on se propose maintenant d’etudier l’effet de la loi de fatigue sur Involution du defaut. 1 г BOISSENOT
FIG.5. Evolution du defaut sous Faction: A) dune traction; B) d ’une traction et d ’une flexion; C) d ’une pression exterieure. IAEA-SM-218/43 4 3
ь
FIG. 6. Nombre de cycles ä rupture en fonction de la loi utilisee.
t ^
FIG. 7. Piquage representant la jonction d ’un tube avec uh cylindre de grand diametre. 4^ -P- BOISSENOT
FIG.8. Defaut en form e de quart d'ellipse se propageant en coin. IAEA-SM-218/43
FIG .9. Action sur le piquage: A) de la pression seule; B) et C) d ’un effort tranchant superpose a la pression. 4 6 BOISSENOT
On considere pour cela un enlargement cyclique simple en traction, la loi de Paris notee ^ et une loi logarithmique notee ^ 2.
3F X = dS/dN = C,(A K )m
L o g [ K / K 0 ]P =dS/dN = C2 [K/K0]n Log [Kc/K]q
ой C b m, C 2, n, K 0, K c, p et q dependent du materiau choisi et sont donnes. La taille du defaut etant petite comparee aux dimensions de la plaque, de sorte qu’on peut admettre que le milieu est infini, on constate contrairement au cas precedent que le front de fissure a tendance ä prendre une forme semi- circulaire. La figure 6 montre egalement que le nombre de cycles ä r u p tu r e depend fortement de la loi utilisee.
2.4. Etude d’un piquage comportant un defaut
Un piquage representant la jonction d’un tube avec un cylindre de grand diametre (fig. 7) est soum is ä un cyclage en pression couple ä un effort tranchant agissant sur l’extremite de la tubulure. On suppose que ce piquage comporte un defaut en forme de quart d’ellipse se propageant en coin, comme il est indique sur la figure 8. On etudie revolution de ce defaut sous Paction conjuguee des deux types de sollicitations. L ’energie potentielle du Systeme en fonction des parametres geometriques a et b de l’ellipse est determinee ä partir d’une serie de calcul par la methode des equations integrales. Pour chaque calcul, on a pris soin de traiter les types de chargement separement afin d’obtenir l’energie potentielle en fonction de la pression et de Peffort tranchant. Dans un premier cas de chargement cyclique, on suppose que la pression seule agit sur le piquage et qu’elle oscille entre 135 et 165 bar; la figure 9A donne revolution du front de fissure jusqu’ä rupture. Dans ce type de propagation, obeissant ä la loi de Paris, le rapport b/a est legerement superieur a 1; il atteint 1,14 au moment de la rupture. Dans un deuxieme cas de chargement, on etudie l’effet d’un effort tranchant superpose ä la pression du premier cas de charge. On considere que Peffort tranchant varie en phase avec la pression entre 270 et 330 bar. L ’effort tranchant agit peu sur la geometrie du front de fissure, le rapport b/a reste constant et egal a 1; par contre, la rupture apparaft bien plus töt (fig. 9B) que dans le premier cas. Dans le troisieme cas, on reprend les conditions initiales de chargement du deuxieme cas, mais on suppose qu’au bout de 180 000 cycles la pression chute et n’oscille plus qu’entre 108 et 132 bar. On enregistre alors un retard dans la IAEA-SM-218/43 47 progression du front de fissure. La encore le rapport b/a reste voisin de 1 (fig. 9C). II faudrait un effort tranchant cinq fois plus grand pour obtenir un rapport de l’ordre de 0,5.
CONCLUSION
L ’approche globale qui consiste ä determiner le taux de variation d’energie potentielle pour l’ensemble de la fissure pourrait etre mise en defaut dans la mesure oil la definition des parametres geometriques du front de fissure ne serait pas assez fine. Nous avons presente des exemples de propagation en supposant que le defaut restait de forme semi-elliptique; cette hypothese est restrictive mais interessante parce qu’elle simplifie les calculs; on aurait pu choisir d’autres parametres, comme par exemple les coordonnees des noeuds du maillage situes sur le front de fissure; dans ce cas, la determination des inconnues geometriques serait plus difficile, mais cette technique permettrait sans doute de suivre n’importe quelle forme de fissure. D ’u n e f a 9 on generale, les resultats obtenus semblent conformes ä ceux que Гоп observe experimentalement. Le choix de la loi de fatigue est important si Гоп veut prevoir le nombre de cycles a rupture de facon precise. Nous envisageons une Serie d’experiences qui permettront peut-etre de dire s’il est possible de relier un essai de laboratoire sur une eprouvette de fatigue ä u n cas reel tridimensionnel et de connaitre egalement l’importance du champ de contraintes residuelles genere par la zone plastique en fond de fissure. Par ailleurs d’autres developpements sont encore possibles dans l’analyse tridimensionnelle par la methode des equations integrales. La methode proposee par Weaver [12], qui consiste ä schematiser la fissure par une seule surface sur laquelle on impose des conditions aux limites appropriees, faciliterait le calcul et s’adapterait particulierement bien ä la methode proposee.
REFERENCES
[1] BERGAN, P.C., AAMODT, B., Finite element analysis of crack propagation in three- dimensional solids under cyclic loading, Nucl. Eng. Des. 29 2 (1974) 180—88. [2] KOBAYASHI, A.S., et al., «Corner crack at a nozzle», Proc. 3rd Int. Conf. on Pressure Vessels, ASME, Tokyo, 2 (1977) 507-16. [3] BROEKHOVEN, M.J.G., «Fatigue and fracture behaviour of cracks at nozzle corners», Proc. 3rd Int. Conf. on Pressure Vessels, ASME, Tokyo, 2 (1977) 839—62. [4] HELLEN, T.K., DOWLING, A.R., Three-dimensional crack analysis applied to an LWR nozzle-cylinder intersection, Int. J. Pressure Vessels Piping 3 (1975) 57—74. [5] LABBENS, R., PELLISSIER-TANON, A., HELIOT, J., Application de la theorie lineaire de la mecanique de la rupture aux structures metalliques epaisses, Rev. Phys. Appl. 9 (1974) 587-97. 48 BOISSENOT
[6] LEMAITRE, J., Extension de la notion de taux d’energie de fissuration aux problemes tridimensionnels et non-lineaires, C.R. Hebd. Acad. Sei. T 282B (1976) 157-60. [7] CRUSE, T.A., «Elastic singularity analysis», Innovative Numerical Analysis in Applied Engineering Sciences (C.R. Coll. Int. Versailles, 1977). [8] LACHAT, J.C., WATSON, J., Application de la methode des equations integrales au calcul des structures, Memoire technique du CETIM n° 25, mars 1976. [9] LACHAT, J.C., WATSON, J., Effective numerical treatment of boundary integral equations: a formulation for 3-D elastostatics, Int. J. Numer. Methods Eng. 10 (1976) 991-1005. [10] HENSHELL, R.D., SHAW, K.G., Crack tip finite elements are unnecessary, Int. J. Numer. Methods Eng. 9 (1975) 495—507. [11] ROOKE, D.P., CARTWRIGHT, D.J., Compendium of Stress Intensity Factors, H.M.S.O., London (1976). [12] WEAVER, J., Three dimensional crack analysis, Int. J. Solids Structures 13 (1977) 321-30. DISCUSSION
on papers IA EA -SM -218/15, 16,43
G.I. SC H U ELLER: Mr. Marci, you say near the end of your paper that, where the pressure vessel is concerned, “a leak in the pertinent areas cannot develop as a result of a flaw missed by present non-destructive testing m ethods”. But is there really no probability of failure at all in this area? A s far as reliability analysis is concerned, 1 think your statement might not be valid for all cases. G. M A R C I: For the area of the pressure vessel and the crack problem assumed for our analysis, the probability of the occurrence of a leak is equal to or smaller than the probability of the assumed crack size not being detected by present m ethods o f non-destructive testing. W. SCHIM M EL: May I ask any or all of the three authors how large a crack needs to be for a leak to develop during the operating time of a reactor vessel (i.e. 20— 30 years)? K.A. RA H K A : The point we wanted to make in our paper only concerns cladding integrity and no true leak was dealt with. The British design rules would apply more or less when one is estimating the risk of coolant attacking the base metal. G. M A R CI: The size of crack which will cause a leak depends on the structure under investigation. For the problem posed in m y presentation a leak would occur if a crack could grow through the Inconel structural weld. W. SC H IM M EL: How fast is the growth of a crack under leak conditions? G. M A R C I: Crack growth rates depend on the state of stress at the particular position of the crack front and on the material in which the crack is growing. W ith respect to the material one has to consider temperature, R = K imin/K jmax, environment and frequency when determining the input data for the crack growth equation; for example, for the Paris-Erdogan equation this would be the input data C and n. C.B. BU CH ALET: Mr. Boissenot, I would like to know at what place on the crack you calculate the value of K, since К varies all along the crack front. J.M. BO ISSEN O T: It is not necessary to calculate К all along the crack front for this type of approach which is a global approach. Figure 4 shows the variation in К along the crack front for a fault of arbitrary shape. It will be observed that at the beginning of the process the fault tends to propagate more quickly in the area where К is at its m aximum , where an arbitrary initial shape is concerned. But during propagation, where the type of load is the same, the fault tends to take on a shape which causes the crack front to become an iso-K line (K has the same value on the whole of the front). C.B. BU CH ALET: Could you comment further on the calculations described in your presentation?
49 50 DISCUSSION
J. M. BOISSENOT: About ten calculations are required to determine G as a function of the geometrical parameters of the fault. Calculations for a nozzle require 350 seconds TT (250 seconds CPU) on C D C 7600.
K. W.L. BRANDSTÄTTER: The technique of integral equations operates in the arbitrary elastic region. Could one introduce into the integral equation technique a criterion for exceeding the elastic lim it? J.M. BO ISSEN O T: At present the integral equation technique cannot be used for non-linear phenomena such as plasticity. Feasibility studies have shown that this could be done but as far as I know there are no operational programmes. The three-dimensional analysis was therefore made in elasticity and effects due to plasticity at the bottom of the crack were ignored. D.J. BU RN S: At Waterloo we have also given some thought to the use of energy form ulations for predicting crack growth direction and shapes. However, all calculations o f subsequent growth are dependent on the generation of basic crack growth data and equations similar to Mr. Boissenot’s Eq. (8). The constants C and m in Eq. (8) are usually generated for pressure vessel applications with specimens since the crack tip is predom inantly of the plane strain type. When Eq. (8) is used to predict su rfa ce (plane stress) growth of a semi-elliptical crack it usually gives predicted lives (growth) very different from those observed in experiments. Are you assuming that Eq. (8) at К = [EG/(1 - v 2 )'\in allows you to obtain from published data values of C and m that can be applied in your energy approach? Are you ignoring the likely differences in C and m when considering growth along the minor (a) and major (b) axes of the semi-elliptical crack? Do you agree that an energy formulation does not obviate this problem? If not, please explain how you generate your basic crack growth data and give details of your test specimen. J.M. BO ISSEN O T : I agree that the plane strain hypothesis is an approximate one but it becomes more or less obligatory if one wants to use a fatigue law which involves use of the stress intensity factor K, since К depends on the state of stresses. If experimental data involving the use o f an energy balance, and there fore a A G rather than а ДК, were available, one could approach the problem more rigorously, since one would im plicitly be taking the particular state into account. However, it should be remembered that the fatigue law applies to the increase dS swept by the whole of the crack front and not to the lengths a and b (these axes are only for defining the geometry of the front), so that overall one is nearer a state of plane strain than of plane stress. We intend to carry out tests on CT-type specimens in order to determine experimental laws using G: (dS/dN) = F(G , . . .) and we propose to use these laws to study the propagation of a semi-elliptical fault in an interior surface under pressure. This study will be both theoretical and experimental. R.L. RO CH E: If calculations of fatigue crack growth are necessary for estimating the reliability of pressure components, one should be careful when DISCUSSION 51 evaluating the evolution of the shape of the crack front. Grow th rate depends on the state of stress and strain. The rate is very different for states of plane stress as opposed to plane strain. The empirical relationships used for analysis, such as the Paris law, are designed for states of strain. They are suitable at considerable distances from free surfaces but near these surfaces the state is rather one of plane stress and the growth laws are different. M ay I ask Mr. Boissenot and Mr. Marci to say what they think about the corrections that need to be made because of this effect in the growth case under study, especially near the edges of specimens?
J.M. BO ISSEN O T : Our analysis is a three-dimensional one. The question of whether one has plane stress rather than plane strain arises only when one wants to use а Д К rather than AG. In the example of the bar, I used the Paris law and therefore а ДК, on the assumption that overall there is a state of plane strain (a pessimistic hypothesis), although at the edge there is a plane stress state. G. M A R C I: It is realized that the Paris-Erdogan equation represents an empirical relationship between the incremental crack growth and the excursion of the stress intensity factor per cycle. But when one is making a fracture mechanics fatigue analysis in order to show a certain margin of safety, the analysis has to be based on the most unfavourable assumptions. The assumed crack size has to be such that it is found with certainty by the applicable non-destructive testing method. For the sample problem described in my presentation of paper SM-218/15 the assumed crack size is more than 10 times larger than the reference reflector for ultrasonic inspection. The location of the assumed flaw is unfavourable with respect to detectability apart from damage caused to the structural integrity. It is an experimentally documented fact that fatigue crack growth rates are lower in the plane stress region than in the plane strain region of a structure. Furthermore, the existence of a threshold stress intensity range is amply verified. Both phenomena have been neglected, since their incorporation into a fracture mechanics analysis would lead to less crack growth under the particular conditions found and therefore they are covered by the worst case assumption. The curve-fitting parameter C and the exponent n of the Paris- Erdogan equation are chosen so that the predicted results are upper bound values for the material in question. If all the above features are considered and the appropriate values are chosen, the analysis will give a crack size which is an upper bound value based on the assumed initial crack size. The safety analysis is then based on the upper bound crack size.
D.G.H. LATZKO : The influence of crack growth retardation due to crack tip closure under overload conditions is not treated in any of the three papers, not even in that of Mr. Marci (SM-218/15), in which the loading spectrum considered specifically includes overload conditions. 52 DISCUSSION
I would also like to point out that, apart from Mr. Antalovsky’s proposal (paper SM-218/46) that Priddle’s law should be used, Mr. Boissenot’s use of a logarithmic crack growth law described in section 2.3 of paper SM -218/43 constitutes the only evidence presented at this Sym posium of an effort to use an alternative to the Paris-Erdogan equation. S.M. BUSH: Since the subject of this Symposium is reliability, it might be worth while to examine the fatigue behaviour of reactor systems. I would estimate that about 75% of the observed cases of fatigue in reactor pressure vessels have been due to thermal loads, specifically the mixing-tee problem where injected cold water mixes with hot water but may impinge on the hot wall prior to complete mixing. In such a case the different coefficients of expansion of stainless steel as opposed to carbon steel serve as a driving force for initiation and propagation to a depth where the pressure loads are sufficient to continue fatigue. M y comment applies to all three papers, but particularly to that of Mr. Rahka (SM -218/16). K.A. RA H K A : I very strongly agree with you about the possible influence of the factors affecting cladding integrity which you mentioned. I think, though, that the British failure curves, as compared with the A SM E III design curve, give a good idea of the influence of an underclad flaw on cladding integrity. The case referred to should thus be added to a complete cladding integrity assessment, with due attention being paid to the thermal stresses and the “crack opening pressure effects”. Naturally, thermal stresses of the kind you referred to have to be estimated, and especially the number of them, so that some disagreement on this point has to be expected! The best thing to do here would be to aim at modifying the local designs so that the risk of thermal turbulence is m in im iz e d . IAEA-SM-218/3
SOME ASPECTS RELATING TO THE RELIABILITY OF STEEL REACTOR PRESSURE VESSELS
G . P R A N T L Swiss Federal Institute for Reactor Research (EIR), Würenlingen
T . V A R G A Sulzer Brothers Limited, W in t e r t h u r
D . H . N J O Swiss Nuclear Safety Division (A SK ) Würenlingen, Switzerland
A b stra c t
SOME ASPECTS RELATING TO THE RELIABILITY OF STEEL REACTOR PRESSURE VESSELS. Fracture mechanics methods are used to assure the reliability of steel reactor pressure vessels. In principle the two-stage concept for protection against non-ductile failure as described in IAEA Technical Document IAEA-189 is applied. In the first stage a preliminary verification based on standard values and data from codes is required. The design and analysis at this stage uses mainly the method described in the ASME Code, Section III. For protection against non-ductile failure a preliminary analysis in accordance with Appendix G of Section III has to be performed. In contrast to the ASME Code the residual stresses have to be taken into account at this stage of the safety verification. This is followed in the second stage by a detailed analysis to assess more realistically the fracture-safety of the pressure vessel, using the actual component data. With these, the actual conditions relating to stresses, material characteristics and existing defects should be considered. If the above mentioned factors are considered, a more realistic and comprehensive verification of protection against non-ductile failure can be achieved, not only with respect to defect sizes but also in connection with operating temperatures.
1. I N T R O D U C T I O N
The demonstration of safety from non-ductile failure usually follows a procedure divided into two successive steps [1 ]. These two steps differ with respect to the input data used for the evaluation of the safety margin, but both apply the same quantitative fracture mechanics approach.
53 54 PRANTL et al.
In the early stages of a project, no data for the individual materials proposed are available and the exact stresses in the components are still unknown. Never theless a first, though not very elaborate, knowledge of the behaviour of the component is required at that time. It comprises an important part of design procedure and allows the licensing authority to perform a first assessment of the safety of the project. For the reasons mentioned, this stage-one evaluation is usually based on methods and data as given, for instance, by the ASM E-Code, Sections 111 and XI. Later in the course of the project the properties of the actual materials become know n and the results of extensive stress analysis become available. Applying these data, the preliminary safety evaluation has to be reviewed and a final demonstration of safety has to be made. Usually the preliminary calculations already indicate which load and temperature combinations are going to present the lim iting case for the different parts of the pressure vessel. M ost attention has to be paid to these cases when perform ing the final, so-called stage-two evaluation. The purpose of this paper is to describe Swiss practice, to discuss some points relating to testing techniques and to mention a few problems raised by the inter connection between the three fields involved, namely:
(1) Testing of the materials for properties relevant to the mode of operation of the structure. (2) Completeness of the stress analysis. (3) Consideration of the non-destructive testing technique and its capability for detecting crack-like defects.
2. MATERIAL DATA
In order to obtain from laboratory specimens test results that can validly be applied to the real structure, several points should be considered. Most important am ong these are:
(1) The mode of operation and loading of the structure. (2) The expected behaviour of the material in the structure. (3) The limitations inherent in the calculational techniques used to apply the behaviour of the specimen to that of the structure.
In addition it is reasonable to stay within the framework of well-established testing techniques and to make use of the broad basis for comparison offered by numerous results derived elsewhere. A good point to start with is a brief review of the particular properties of the structure we have in mind. The pressure vessel of a light water reactor is a thick- walled welded structure, subjected to fairly well controlled operating conditions, IAEA-SM-218/3 55 as compared to many non-nuclear components. Under normal operating conditions the loading is basically static with a cyclic component, due to reactor power fluctuations being superimposed. During the life of a vessel, there are relatively few cycles between zero and full power, with a carefully controlled loading path chosen to exclude situations dangerous with respect to non-ductile failures. Hydrotests and transient events due to faulted operational conditions leading to local thermal stresses must be considered in the fracture analysis. According to this, the temperature range for testing the materials must be fixed. It is important to obtain, at temperatures well above room temperature, a quantitative measure of the margin against fracture. Since even large compact tension specimens behave in a ductile manner at these temperatures and the same is expected for a vessel with small defects, linear elastic fracture mechanics is no longer a suitable method. In order to extend the testing range to higher temperatures, the crack opening displacement is measured by means of clip-in gauges installed in the slot of the test specimens. Fracture behaviour is also determined in terms of ‘J-Integral’ or ‘Equivalent Energy’ values. Instrumented tests with fatigue pre-cracked Charpy specimens are very valuable for supplementary screening tests and particu larly for irradiated materials testing [2]. For each of the methods chosen the onset of failure of the specimen has to be defined. It is in any case safe to choose the initiation of first stable crack growth and to develop a technique which detects this event reliably. From the initiation point on a load-displacement line to the point of m axim um load wheee the load carrying capacity of the specimen is exhausted, there is a certain safety margin. This safety margin depends on the shape of the load-displacement curve. If the com m on pressure vessel steels are tested at temperatures corresponding to a hydrotest or at even higher temperatures, where ductile behaviour is usually found, it is im portant to distinguish between the margin relative to load and the one relative to displacement. Whereas the reserve expressed in load increase m ay be rather small, the reserve in displacement increase is large. This means that the tolerance of the structure, assuming it shows a behaviour similar to the test specimen, to overloads is small but the possibility for further deformation in a gross yielding process is large. Due to geometrical effects the upper portion of a load displacement curve of a cylindrical or spherical shell is much steeper than the one of a plane test specimen made of the same material. Consequently, the margin relative to the load-carrying capacity is higher than indicated by a laboratory test using a plane specimen (Fig.l). It may be useful to refer briefly to the evaluation of a test and the application of the result to a real structure by looking at the failure criterion. Irrespective of the method, this criterion is always in principle the same. A relevant property of the material is compared to the corresponding quantity, which is derived, for instance, from a mechanical analysis of the specimen or the structure, assuming a continuum. If the property of the material is an elastic-plastic one, the mechanical 56 PRANTL et al.
FIG.l. Schematic presentation o f the safety margins of a specimen and a vessel from the same material.
General Failure Criterion
Mechanical <; Material Quantity Property
Example S = COD ° elast. plast. elast. plast. correct
c = Л Л Г ) ° elast. plast. elast.
safe
c _ p n n ° small sc. yield, ^^^elast. plast. unsafe
FIG.2. Failure criterion as comparison of a material property with the corresponding mechanical quantity. IAEA-SM-218/3 57
0 tolerable > 0 detectable nom. safety factor uncertainties in analysis
FJG.3. Relation between fracture mechanics analysis and capability of non-destructive testing technique. quantity should also be based on an elastic-plastic analysis. If, due to the lack of elastic-plastic solutions, a mechanical quantity derived from a purely elastic analysis is used in connection with elastic-plastic material data, the failure criterion is unsafe. If on the contrary an elastic-plastic solution for the structure is used together with an elastically derived property of the material, the analysis is on the safe side. In certain cases, where the same calculational procedure is used to evaluate the test results and to apply these results to the structure, the errors m ay nearly cancel each other out. W ith respect to this, the crack opening displacement is in a somewhat peculiar position, since it is measured directly on the test specimen and applied to the structure via a quantity derived from either an approximate model or a numerical analysis. (Fig.2).
3. DEFECT SIZE
In the case of a reactor pressure vessel the given material and the stresses, determined by the chosen design, form the starting point for the fracture analysis. This analysis aims at deriving critical and tolerable defect sizes. Sometimes, a tolerable stress or a permissible operating temperature is sought, starting from postulated or real crack-like defects in certain regions of the component. In this way, non-destructive testing techniques enter the analysis, as is illustrated in Fig.3. This diagram shows the following:
(1) The capability for crack detection by the chosen testing technique as applied to the particular structure treated. This curve is not generally valid, because 58 PRANTL et al.
it depends on the testing technique, the material and the geometry of the com ponent as well as the skill and experience of the personnel. (2) The critical crack size, which is a property of the component operating in given conditions.
It is immediately clear from this figure that the comparison between the testing technique and the tolerance of the structure to defects yields the m aximum real safety m argin possible. Very often, a tolerable defect size is derived from the critical defect size by application of a safety factor (a better name would be ‘factor of ignorance’) large enough to cover all the uncertainties involved in the input data and the analysis. This presents a means to specify the m inim um capability of the testing techniques, because the tolerable defect size must always be equal to or larger than the detectable defect size. If it were possible realistically to estimate all those uncertainties, a com parison of the tolerated defect with the detectable defect would furnish the real safety margin of the com ponent against reaching the critical defect size. In the case o f easily inspectable areas on a reactor pressure vessel and crack orientations favourable for detection, the ultrasonic test method could, for instance, show flaw indications corresponding to artificial reflectors of about 3 m m diameter with almost 100% certainty. According to experience, these indications correspond to real cracks o f about twice this size. So far, fatigue growth of existing cracks has not been mentioned. This aspect is now discussed with special reference to ‘non-inspectable areas’. There are regions on a pressure vessel which are, due to geometrical lim itations in the design of the component and its environment, inaccessible to in-service inspection by volumetric testing techniques1. One example is the bottom head of a BW R vessel. This deficiency has to be balanced by a particularly careful and complete analysis and a rigorous control of the initial condition of this part of the pressure vessel. Fig.4 shows schematically two possible situations. At the beginning of service the m axim um possible crack present in the structure is determined by the final non-destructive test. This original crack may grow to some final size undetected, as long as it remains smaller than the dimensions of the non-inspectable area. Consequently, if the tolerable crack can be shown to be significantly larger than this area, the real crack approaching it could be detected, provided the in-service inspection is performed with sufficient frequency. If, on the other hand, the end-of-life tolerable crack is smaller than the non- inspectable area, a real crack approaching it could not be detected. In this case, which is also illustrated in Fig.4, a rigorous analysis of the defects has to be
1 Definition of inspectable area: An area is considered to be inspectable, if this area permits adequate probability of detection of a tolerable flaw using an appropriate method of established NDT-technique such as the ultrasonic testing method. crack length FIG.5. Fatigue growth of surface cracks towards critical critical size. towards cracks surface of growth Fatigue FIG.5. FIG.4. Non-inspectable areas and actual crack growth. crack actual areasand Non-inspectable FIG.4. IAEA-SM-218/3 59 60 PRANTL et al. performed. This analysis has to make use of data measured on the actual material in this area and of a precise and complete stress analysis. Especially the shape of growing cracks has to be determined in a realistic way. Figure 5 shows schematically that, depending on the crack depth-to-length ratio, the A S M E postulation2 [3] is sometimes very incorrect and consequently the distance in time and size between a real crack and the critical crack it would eventually approach is overestimated.
4. STRESS DATA
Experience shows that even in stress-relieved welded structures there are residual stresses which must be accounted for3 [4]. Unpublished measurements made by different manufacturers of pressure vessels show that residual stresses as high as about 150 M Pa have to be expected in parts of a complex geometry. In another case the measurement yielded 60 MPa in a spherical pressure vessel head. These values are not negligible in comparison to load or thermal stress. Unfortunately the residual stresses are often highest in areas not very accessible to inspection by non-destructive testing techniques.
5. CONCLUSIONS
Earlier on, the safety factor in terms of crack size was discussed. It is interesting to analyse this factor with the help of an example: Assum e that a fracture analysis were carried out using nom inal values of the fracture toughness K Ic and the stress a for the derivation of the critical crack size. Then a nom inal safety factor of 5 would be applied to arrive at the tolerated crack size. Now, assume that, due to scatter in materials data and incomplete estimation of the stresses, the local K )c is 30% less than the nom inal one and the local stress exceeds the nominal stress used in the analysis by 50%. The con sequence would be that the real safety factor in this area of the structure would be only about 1.2 instead of the nom inal one, which was taken as 5. If on the other hand, one insisted in a real safety m argin of 5, the nominal analysis would have to assume a nom inal safety factor of about 22. For this reason nominal factors а ^ /а ^ of at least about 5 must be used if the input data are reliable. W ith less reliable nom inal data the required nom inal safety factor m ight be up to 20, in order to assure a sufficient real margin.
2 A-5 300 [4] indicates that eventual crack growth shall be assumed to take place maintaining the a/c ratio of the original flaw. 3 ASME XI A-3200 also requires consideration of residual stresses. IAEA-SM-218/3 61
Our present approach is deterministic. It has been shown that many uncertainties are involved in the analysis of the behaviour of the pressure vessel. It is necessary to estimate these uncertainties in order to derive bounds for the real safety margin of the structure as opposed to the nom inal safety factor, which is used in the analysis. This is made possible by introducing probabilistic elements into the procedure. However, the proper use of the well-developed mathematical tools must be based o n re a listic statistical data for the various items, such as materials toughness, distribution of flaws, possible variation in stress, etc. It appears that these data are not yet available to the extent desirable.
REFERENCES
[1] NJO, D.H., VARGA, T., PRANTL, G., “The concept for the protection of the RPV against nonductile failure as it is applied in Switzerland” , Fracture Mechanics Applications: Implications of Detected Flaws, Techn. Doc. IAEA-189, IAEA, Vienna (1975). [2] PRANTL, G., VARGA, T., NJO, D.H., “Reactor vessel surveillance, Present practice and future trends in Switzerland”, Reactor Pressure Vessel Surveillance (Proc. Techn. Committee Meeting Plzen, 1976) Techn. Doc. IAEA-202, IAEA, Vienna (1977). [3] ASME Boiler and Pressure Vessel Code, Section XI, 1974 Edition, A -5 300 (4). [4] ASME Boiler and Pressure Vessel Code, Section XI, 1974 Edition, A—3200.
DISCUSSION
G. SC H U ELLER : You stated at the end of your presentation that a more realistic analysis of the problem would be one which involves the application of probabilistic m ethods in order to take into account the uncertainties involved. I certainly agree that, as Freudenthal showed more than 25 years ago, the safety factor can be related quantitatively to the reliability of a structure. This approach, which has since been taken further and refined by a number of researchers, should be applied to the problem discussed in your paper. G. P R A N T L: I agree. In our paper we tried to illustrate that the considera tion of the various uncertainties involved in a deterministic analysis quite naturally calls for the introduction of probability techniques. On the other hand, it appears that at present the main difficulty is not the handling of the mathematical tool but the obtaining of realistic input data on, for instance, the distribution of flaw sizes and orientations in the structure, or on the variability of the material properties. This kind of information has to be collected prior to a successful application of reliability techniques, which is more than just an academic exercise. A.B. L ID IA R D : Your Fig.5 shows crack growth different from that which would follow from the A SM E recommendation. Can you say whether this is the 62 PRANTL et al. result of new measurements and, if so, could you describe the conditions under which they were obtained and say when and where they will be published? G. PR A N TL: The data shown come from preliminary results of a current research programme at the Swiss Federal Institute for Reactor Research (EIR). The aim is to investigate the change in size and shape of cracks growing under fatigue loading in order to supplement the numerous data on one-dimensional crack growth available elsewhere. The intention is to test specimens of various different geometries. The data shown here have been obtained from measure ments of plate-shaped tensile specimens with semi-elliptical surface flaws with a view to establishing a leak-before-break criterion. The first results will be published in the open literature in about a year. IAEA-SM-218/12
SOME ASPECTS OF THE ASSESSMENT OF PIPEWORK INTEGRITY
B.J.L. DARLASTON Berkeley Nuclear Laboratories, CEGB, Berkeley, Gloucestershire,
United Kingdom
A b stra c t
SOME ASPECTS OF THE ASSESSMENT OF PIPEWORK INTEGRITY. The paper represents a collation of the work undertaken and so far reported within the CEGB on the assessment of pipework integrity. The three main aspects discussed are the failure assessment route, the effect of combined bending and internal pressure on failures and leak before break. The CEGB has been promoting a method of assessment of defects in structures based on a two criteria approach. The bounds of the method are where the crack tip events are controlling (LEFM) and where the ligament behaviour is controlling (plastic limit theory). Using this as a basis for analysis, further work has been carried out which is directly applicable to pipework. The interaction of internal pressure and bending has been examined experimentally and a design rule postulated. Work has been undertaken on leak before break and a wide range of pipe and defect geometries have been tested to provide validation of the assessment method. Work continues on all these subjects and this paper represents, in part, an interim statement of the position.
\ 1. INTRODUCTION
This paper represents a collation of documents produced within the Central Electricity Generating Board on various aspects relating to pipe work integrity.
The three main subject areas are:
1. The failure assessment procedure 2. The effect of combined bending and internal pressure on pipe failure 3. Leak before break
Work continues on all these topics and this paper is, in part, an interim statement of the position.
63 64 DARLASTON
FIG.l. Failure assessment - two-criteria approach.
2. THE FAILURE ASSESSMENT PROCEDURE (1)
Most pressure vessels are not inherently brittle. They often operate at temperatures where toughness is high and the possibility of large scale yielding prior to failure has to be recognised. Under these conditions LEFM is unconservative and codes such as Section XI of the ASME Boiler and Pressure Vessel Code resort to the use of the factors to guarantee safety. This situation has led to much activity in developing failure assessment procedures based on elastic plastic fracture mechanics. One such procedure has been developed in the UK in the Central Electricity Generating Board (1).
The proposed procedure was built upon existing codes and practices so that as far as possible tried and tested techniques can be used which w ill be familar to the user. It is accepted that certain of the steps may not be fully established particularly for all foreseeable geometries and to offset this, notes are included which discuss the source and status of the various steps taken. Insofar as the final document was to be a working document such notes have been kept to a minimum and separated from the user instructions. The procedure is therefore presented in a code format. The basis of the assessment for failure is the Dowling-Townley approach (2), which states that failure can be described by two criteria, one where failure is controlled by crack tip properties, that is, LEFM, and that where
failure is controlled by the properties of the ligament, that is,plastic IAEA-SM-218/12 65
FIG.2. Failure assessment diagram.
lim it theory. Figure 1 shows the approach forumulated by Dowling and Townley. The two criteria define the two lim its by which failure of a structure is bounded. Of the many ways of interpolating between the two lim its the Dugdale (3) Bilby-Cottrell-Swinden (4) model for strip yielding ahead of the crack provides the most appropriate forumula since it can be used, suitably modified in accord with the above criteria, in a simple and versatile way. The resultant expression has been shown to be a lower bound to the analytical failure curves which can be reasonably derived, it also fits the experimental data. To sim plify the procedure the equation was re-written as
- i К = S ^2 I n sec (-| S r ) r r TT l a o , a f where К = —------S = — r K,lc r 1 a,
This has been used to produce a failure assessment diagram as shown in F ig u r e 2.
К is a measure of how close the structure is to linear elastic failure and can be written where L is the applied load or pressure and is the load or pressure to produce the linear elastic failure. Similarly is a measure of how close the structure is to plastic collapse and can be written where L is as above and is the load or pressure to cause plastic collapse. The calculations of and for any loading 66 DARLASTON
о, OUCOL W30A1 WORTHINGTON, PJ. Д DUCOL W30BJ C.E.GB. NOTE RO/L/N2A8/74 a A533B BEGLEY. J.A., AND LANDES, J O. 1972 ASTM STP5H,1
FIG.3. Validation of application of failure assessment diagram.
condition yields a point on the failure assessment diagram. Failure is predicted if this point is on or outside the assessment line. The validation of the assessment curve is shown in Figure 3. In all cases lower bound values on data have been taken. These examples w ill form a supplement to the manual to be issued shortly.
Let us now turn to the procedure for evaluating the degree of criticality of a given flaw. First of all it is necessary to define the shape and size of the flaw and define the stresses acting in all relevant loading conditions. The next step is to determine the lower bound material properties for a ll the relevant temperatures which the component w ill experience. The collapse lim it of the flawed structure is then calculated as is the K^. value for the flawed structure. The ratios and are then evaluated and the point plotted on the failure assessment diagram for each loading condition. If the flaw is acceptable it is then necessary to proceed to estimate the amount of flaw growth over the remaining life of the component and the failure analysis on the end of life flaw size is
reassessed. The steps in the procedure are listed in the Appendix. IAEA-SM-218/12 67
In the failure assessment route a separate section is devoted to each of the steps describing what is required and suggesting a suitable basis for performing the calculation. A worked example is considered in the document.
Having proceded through the steps of the assessment let us consider some of the points emerging.
In general the ASME XI methods of flaw characterisation are suggested as the methods to be used. A flaw recharacterisation procedure is included which is designed to assess the possibility of a stable through-thickness flaw arising from a surface flaw which snaps through in an elastic-plastic failure. A factor of 2 is built into the recharacterisation procedure in order to cater for dynamic effects at the crack tip. This is an area of uncertainty which is being actively pursued within our Laboratories.
The lower bound fracture toughness may not always be a controlling parameter, for example where is less than 0.4,that is, where failure is entirely collapse dominated. In such instances it is suggested that use may be made of Charpy correlations to assess toughness. It is also noted that in some instances the possibility of slow ductile crack growth has to be recognised and procedures are recommended which can test and allow for this possibility by using initiation test data obtained from specimens of thickness equal to that of the cracked component. This possibility may or may not be related to the fracture toughness and it is suggested that where possible KI(, should also be obtained from valid tests.
The collapse lim it of the crack structure is obtainable from a variety of sources, for example, slip line field theory, (5) empirical solutions of the type obtained by Kiefner el al. (6), finite element calculations (7). Scale models of components suitably cracked can be tested if it can be ensured that failure w ill be by plastic collapse and this also provides a very useful tool (8).
To form a satisfactory failure analysis the ratios and have to be calculated in a way that includes safety factors and uncertainty factors on the input data. These factors w ill depend on the confidence in the data and the confidence that is required in the safety assessment. Let us now turn to an example to illustrate the flexibility of the method particulary with respect to the sensitivity of the predicted failure condition to the data input and assumptions. 68 DARLASTON
In order to do this vessel 5 of the HSST intermediate vessel test program has been taken, as this test is well known and well reported (9).
An elastic-plastic finite element computation gave a collapse pressure -2 of 224 MN.m for the uncracked intersection. It was assumed conser vatively that the presence of a flaw reduced this pressure linearly with increasingly flaw depth. This led to a value of of 0.94.
From Rashid and Gilman (10) the stress intensity К was obtained. -3 / 2 1 Taking the fracture toughness of 270 MN.m , the value of was shown to be 0.95. By plotting and on the failure assessment diagram the flaw was shown to be unacceptable and that the vessel would fail.
For any failure assessment it is essential to be able to define the way in which assumptions made in the assessment influence the final result, or which parameters have the greatest effect on safety margins. A great advantage of the present procedure is that this is easily done and the results are graphically presented.
Let us consider three aspects, the effect of applied pressure, fracture toughness and crack depth.
Both and are linearly dependent on pressure and reduced to zero at zero applied pressure. Hence failure is predicted at the intercept of the line constructed through the origin and the assessment point К with the assessment line. This is illustrated in Figure 4, which shows that the pressure has to be reduced by a factor of 0.86 of the measured -2 . . failure pressure, that is to 157.5 MN.m . This factor arises simply from the usual fracture mechanics convention of pessim istic analysis and reflects the fact that fracture mechanics is here used to avoid rather than to predict failure.
For many assessments the relevant values of fracture toughness are not available and assessments had to be made. Potential effects of
error in the estimates are easily displayed as shown in Figure 5. On the failure assessment diagram is clearly not at all dependent on fracture toughness, whereas К is inversely proportional. It is shown that if all
the other assumptions and calculations are exact then a toughness in -3 / 2 excess of 377 MN.m would have prevented failure of the vessel at a -2 pressure of 183.5 MN.m
Figure 6 shows the effect of varying the flaw depth only and is of particular interest in that it can be used to show an important point IAEA-SM-218/12 69
Pf к 1-6gm J n ä Sr Sr = ---^ ---- Kr- l'6g" ' ^ ° Pucd-^f) Г K,c Puc«1-^ ) KIC
FIG.4. Effect of applied pressure. FIG.5. Effect o f fracture toughness.
Sr _ P l ___ K 1-6gm J n ä Puc(1-a/t) K.c
FIG. 6. Effect of crack depth.
with regard to factors of safety. The flaw depth to avoid failure is seen to be 20 mm,compared with the LEFM prediction of 34 mm. For certain conditions ASME XI Appendix A requires a factor of 10 on the LEFM calculated critical flaw size to define the maximum allowable flaw size. Under elastic conditions this provides a safety factor of 3 on applied pressure. Figure 6 shows that where there is large scale plasticity this 70 DARLASTON
factor of safety is much reduced. The safety factor is defined in this
instance by drawing a straight line between the assessment point and the origin and examining where the intercept lies on the assessment line. The required safety factor is then the ratio of the distance of the assessment point from the origin to the distance at the intercept from the origin. It is seen from the Fig.6 that the maximum value for this factor is obtained in the absence of the flaw, that is when К = 0 and is -2 r for an applied pressure of 183.5 MN.m only 1.22.
In the proposed assessment procedure factors of safety are not defined nor are they hidden in design curves or formulae. Instead the user is encouraged to consider each area of his input and the influence
of the significance of failure of the component and to select factors accordingly. In other words the user is encouraged to decide the factors.
It is worthwhile to re-emphasize some of the points that the text has attempted to convey. First of all the proposed procedure provides a single unified route for the assessment of structures over the range from LEFM to plastic collapse. Secondly the procedure enables the user to first of all determine the maximum tolerable flaw size, secondly to determine the acceptability of a discovered or postulated flaw, thirdly to examine the sensitivity of the results to variations in input data and all methods of calculation and finally to define appropriate margins of safety.
3. PIPE FAILURE UNDER BENDING AND INTERNAL PRESSURE (8)
A programme of work is being pursued to examine the plastic collapse of structures and in particular the combined effects of internal pressure and bending. This is the first stage of the programme which will sub sequently cover pipe bends and pipe tee junctions.
The overall objective of the work has been to identify the limitations in combining bending and pressure effects as independent failure modes and applying them to situation in which the bending and pressure are combined. The work reported to date on the failure of pipes with defects has been concerned mainly with internal pressure. Although there has been con siderable interest over the years in collapse criteria for structures it is only comparatively recently that attention has been focused on structures with defects.
The experimental results obtained so far are for through thickness defects in bending and part-thickness defects in bending only, under IAEA-SM-218/12 71 internal pressure only and combinations of bending and pressure. Three defect lengths have been examined and two part-through defect depths. Although further work is still required for a complete validation,a proposed design criteria is put forward,based on the pattern of behaviour which has emerged from the experimental results.
The main body of the programme was carried out on 3J" diameter pipes1 with 0.064" wall thickness. Various lengths of defects were assessed but only two depths, those with ^ = 0,7 and those with ^ = 0.94. Some full penetration defect tests were carried out under bending. All the defects were 0.012" wide and nominally flat-bottomed. Results from another experimental programme (11) with a range of pipe diameters from 6" to 24"- were used in the analysis of the pressure-only effects. Thus experimental data on pressure-only loading was available on pipe diameters from 3.5" to 24" with a range of defect geometries. The tests were carried out with defects on the compressive and tensile side as well as on the neutral axis. The load displacement characteristics of the pipe were monitored throughout the test. A specially designed rig shown in Figure 7 was used also for the combined bending and pressure tests. For the majority of tests a specific bending moment was applied and then the pressure increased until failure . In other words the pressure at which the ligament collapse was observed was determined in the presence of an applied bending moment.
The Kiefner et al. (6) analytical approach together with the Folias (12) bulging-correction factor has been used as a basis for comparison of the pressure-only test results. These are shown in Figure 8. For the larger diameter vessels the results agree well with the prediction method over a range of defect geometries. However, in the case of the 6" and 3j" diameter pipes the deviation appears to be due to lack of uniformity in the depth of the artificially induced defects. It is possible, however, that with the 3J" diameter pipes there is a general inapplicability of the Kiefner-Folias approach as all the results predict much lower failure pressures than found in practice.
For the bending only tests the collapse load was defined from the load deflection curves in the conventional manner. In a number of tests, particularly those in which the through thickness defect was subjected to compressive bending moment, local collapse occurred in the vicinity of
1 1" = 2.54 cm. 72
BENDING MOMENT (TONS-IN.) defects. axial with pipes to failure pressure measured against Predicted FIG.8. IPAEET INTERNAL PRESSURE DISPLACEMENT F IG .9. B en d tests on p ipes w ith fu ll thickness defects. thickness ll fu ith w ipes p on tests d en B .9. IG F FIG. 7. Schematic of test facility. test of Schematic DARLASTON PIPE 18A IN. 12л IN. 3 • 5 IN. 20« IN. 6 IN. о DIAM
IAEA-SM-218/12 73 the defect at a load which was lower than the collapse load of the structure. When the defect was on the tensile side the mode of failure was by collapse of the pipe wall on the compressive side. Results obtained in the series which relate to through thickness defects are shown in Figure 9 as bending moment against defect length with the bending moment taken conventionally as that for structural collapse, ignoring the effects local to the defect. The defect-free pipe gave a collapse load of approximately220 tons, in and it is reasonable to assume that the collapse load of the defect-free and of the full-thickness-defect pipes represents the bounds of behaviour. Tests were also carried out on part-through thickness defects at varying lengths. Local plastic deformation effects were not noticed as in the through-thickness-defect case.
It should be noted that all the results of the tests on combined bending and pressure were on 3.5" diameter pipes. The results have been presented in Figure 10 as internal pressure to failure against applied bending moments. The sign of the bending moment is taken as relative to the orientation of the defect. With the 0.040" deep defect results present a reasonably systematic picture for each defect length. It is reasonable to conclude from these results that for ductile failure,providing the bending moment does not exceed half that for collapse, the failure pressure is unaffected by the applied bending moment. This represents a simple design rule which is similar to that at present in the codes for defect-free pipes which allows the bending stress to be equal to half the hoop stress due to pressure. Taking the moment to first yield at 13.8 tons the design limit would be 9.1 tons, in . From the experimentally derived rule it is observed that the applied moment should not exceed half that to collapse that is 9.27 tons, in which illustrates that there is a reasonable correlation in the sense that the experimental behaviour for pipes containing defects follows a similar pattern to that for defect- free pipes.
The observations from this experimental work lead to the postulation of a design rule relating to the effect of defects in pipes under combined internal pressure and bending. It only applies to ductile situations in which the mode of failure is collapse mechanism. The postulated design rule is: if the failure of a pipe containing an axial defect occurs by
2 1 ton (short) = 907.18 kg. 74 DARLASTON
FAILURE DEFECT PRESSURE l e n g t h
FIG.10. Pipe failure for 0.044 in. deep defect. Combined bending and pressure, experimental results.
INITIAL DEFECT I
GROWTH I FAILURE OF LIGAMENT I LENGTH OF CRACK
CRITICAL SUB-CRITICAL 1 CRACK GROWTH 1 1 1 UNDETECTABLE DETECTABLE LEAK LEAK
i — 1 TOTAL FAILURE I FAIL SAFE I
FIG.ll. Leak-before-break concept. IAEA-SM-218/12 75 plastic collapse,then, provided the bending moment does not exceed half that for collapse due to bending alone, it will have negligible effect on the failure pressure.
4. LEAK BEFORE BREAK (13)
The leakage mode of failure usually relates to a very localised area of the vessel wall. An initial defect grows such that the vessel wall is penetrated at a defect length below the critical value for total failure.
The other mode of a failure, that is total failure, occurs when the critical crack size is reached before the vessel wall is penetrated. The first mode of failure presupposes that a stable full-penetration defect can exist. Describing the extremes of behaviour is relatively easy but this gives little guidance on where to draw the line in the assessment of acceptable behaviour of a vessel. In order to draw this line, assumptions must be made not only in the assessment of critical crack lengths but also
in the assessment of leakage rates. In this respect the important issue is what size of crack is required before the leak can be detected, and this may depend on the loading conditions, crack surface roughness, what fluid is in the vessel, and the volume into which the fluid is being discharged.
In the development of the failure, the initial defect grows to a point just before penetration, at which stage the ligament snaps through. The subsequent stable growth of the defect, and finally the relationship between the full-thickness critical defect length and the length of the defect at a particular time, are the important features. A flow diagram is
shown in Figure 11 which illustrates the alternative paths which may be followed in the leak-before-break argument. The starting point in any integrity assessment is the assumption that the structure has been correctly designed and manufactured to eliminate the possibility of failure in the absence of localised defects. It is noted that these defects do not necessarily exist initially in the structure and the assessment should take into account the possibility of defects arising during operation by fatigue. However, this dependence of failure on the presence of a defect provides a useful starting point. Concerning the pressure vessel, a further assumption may be made, that no defects form a leakage path at the start of operation. One route is for the flaw to grow in a manner such that, when the ligament snaps through, it does so at a length which is greater than the critical crack length, leading to total failure. The other route is for the ligament to snap through at a subcritical length leading to a stable situation. It is worth noting here that snap-through 76 DARLASTON in the remaining ligaments may be either brittle, ductile or of mixed mode. However, the mode of snap-through is of considerable importance and interest, as it may well control whether the crack runs or not. Returning to the flow diagram, several possibilities exist at this point. One alternative is for cyclic crack growth to occur, maintaining the shape of the initial defect, giving a small initial penetration which grows in a controlled manner. This may be regarded as a completely stable penetration process, no instability occurring before some leakage of fluid. On the other hand, instability might occur before full penetration and once again there are two possibilities. Firstly if the material is very brittle, the stress intensity factor for the defect may exceed its critical value and the defect will then propagate through the wall thick ness. Secondly if the material is tough, the stress in the uncracked ligament may reach a critical value and failure in the ligament will occur. In both cases subsequent behaviour depends on the stability or otherwise of the full penetration defect, taking into account any dynamic effects which may result from the nature of the snap-through failure. Once the defect has penetrated the wall and has become stable, two factors are of importance. First of all, can the leak be detected? And secondly, how much crack growth needs to occur before the critical crack length is reached? If the leak is undetectable for one of several reasons, the situation is no different from that of total failure - that is, there is no early warning. On the other hand, if the leak is detectable, it is essential to know the safety margin in terms of the operational cycles necessary to grow the crack to a critical length. The magnitude of this safety margin represents the adequacy of the fail-safe concept as applied to the pressure boundary.
In order to carry out a rigorous leak-before-break assessment three important aspects need to be understood. The first point is that we require the ability to define the crack profile as it develops. The second point is that a criterion must be defined which identifies the point of instability of a part-penetration defect. The third point is that there must be a criterion which defines whether the defect is stable following snap-through.
The series of internally pressurised pipes discussed in section 3 are used to illustrate the method of analysis and the principles of the concept. The tests relate not to the fatigue crack growth process but to the snap-through of the ligament and subsequent behaviour of the pipe. IAEA-SM-218/12 77
In this context a leak is said to occur when the pipe wall fractures over a length which is resticted to the original length of the surface flow. The break is said to occur when the fracture of the pipe wall extends beyond the length of the original flaw.
Failure of pipes containing external axial flaws can be understood as taking place in potentially two stages. First of all, the ligament beneath the flaw fails, causing a full-thickness defect to be formed. The full-thickness defect then may or may not extend beyond the original length of the flaw to cause either a leak or a break. A full-thickness defect will extend axially under a given pressure if the defect length exceeds a critical length. Thus, in order to predict whether a particular defect will leak or break, it is necessary to know the pressure at which the ligament fails and whether the length exceeds the critical length for full thickness defect.
For ligament failure the plastic failure pressure was taken to be
p = £t / 1 ~ a/t ( u R ( 1 - a/mt ( in accordance with the expression given by Kiefner et al. (6) m is a stress concentration factor taking into account the pipe bulging around the flaw, and is given by Folias, (12)
21 1.05c m Rt > c being the flaw semi-length, a the flaw depth, R the mean radius of the pipe, and t the pipe wall thickness.
Similarly the pressure to cause brittle fracture was taken as
t КIc P к R/тга F(a/t) where F(a/t) is a geometrical correction factor. Values of F(a/t) have been obtained by extrapolating the stress intensity factor, K^, calibrations of Rooke and Cartwright (14) to high values of a/t.
Let us now consider the criticality of the full-thickness defects formed as a result of the ligament failures. A full-thickness defect which is supercritical at the pressure causing ligament failure would extend axially, and the overall failure would be a leak Conversely, a subcritical 78 DARLASTON
FIG. 12. Failure assessment diagram o f the leak/break results.
full-thickness defect would not extend, and would be a leak. The basis for the analysis was the Failure Assessment Route given in section 2
of this paper.
The expressions given for the plastic collapse and brittle fracture pressures, Рц and P^, can be used to investigate the criticality of the full-thickness defects according to the failure assessment method (1), by rearranging the equation in terms of the normalised variables = P^/P^, and = Р^/Рц . The expression on which the analyses were based were taken from Kiefner et al (6) and Rooke and Cartwright (14). The equation
of the assessment diagram
8_ К S In sec (i Sr) r r 2 7Г
should therefore delineate a region which contains only leaks. Figure 12 shows this curve, together with the {K^.S^} points derived from the experimental pipe failure pressures. A logarithmic scale has been employed to accommodate the large range of the data. The value /О 0 K_ = 250 ksi/in (272.5 MN.m ' ) has been used for the fracture toughness, Ic _ and the flow stress, o, for a test was taken to be the appropriate batch
mean flow stresses o, .
ksi = 1000 lbf/in2 = 6.895 X 106 Pa. IAEA-SM-218/12 79
Figure 12 effectively separates leaks and breaks. No breaks occur in the region predicted to be occupied by leaks, while only two leaks occur in the break region. The values of the latter are both greater than unity, while the S values are less than unity. This suggests that, for г _ “3/2 these tests, a somewhat greater toughness than 250/in (272.5 MN.m ) was appropriate.
5. CONCLUDING REMARKS
There are a number of topics within the CEGB programme framework in addition to those discussed in this paper.
The analytical work uses the BERSAFE suite of computer programs and those programs are under continuous development particularly with respect to fracture mechanics and plasticity aspects.
An important feature of leak before break is to develop an understanding of the way fatigue cracks grow in various structural geometries. The basic work on simple specimens has been completed and attention is now focused on cylinders and cylinder/cylinder interactions. These latter tests also provide data on 'snap through' of ligaments and subsequent crack stability. The evaluation of plastic collapse present a problem with the more complex structures. Modelling techniques using polycarbonate and epoxy resins are being employed in this context as well as analytical techniques which require further development.
ACKNOWLEDGEMENT S
The author wishes to acknowledge those contributions made by Dr. R.P. Harrison, Dr. I. Milne, Dr. D.C. Connors, Mr. R.A.J. Hellen and Mr. K. Loosemore, all of the CEGB.
The paper is published by permission of the Central Electricity Generating Board.
REFERENCES
[1] H A R R I S O N , R.P., LO O S E M O R E , K., MILNE, L, CE G B Rep. R/H/R6 (1976). [2] D O W L I N G , A.R., TO W N L E Y , C.H.A., Int. J. Pressure Vessels Piping 3 (1975). [3] D U G D A L E , D.S., J. Mech. Phys. Solids 8 (1960) 100. [4] BILBY, B.A., CO T T R E L L , A.H., SWINDEN, K.H., Proc. R. Soc. (London) A272 (1963) 304. 80 DARLASTON
[5] GREEN, A.P., HUNDY, B.B., J. Mech. Phys. Solids 4 (1956) 128. [6] KIEFNER, J.F., MAXEY, W.A., EIBER, R.I., DUFFEY, A.R., ASTM Rep. ASTM-STP-536 (1973). [7] HELLEN, T.K., HARPER, P.G., CEGB Rep. RD/B/N3597 (1976). [8] DARLASTON, B.J.L., HARRISON, R.P., in Structural Mechanics in Reactor Technology (Proc. 4th Int. Conf. San Francisco, 1977). [9] MERKLE, J.G., Pretest Analysis Information for HSST Program Intermediate Test Vessel V5, Oak Ridge Natl. Lab. Rep. ORNL-TM-5090 (1975). [10] RASHID, Y.R., GILMAN, J.D., in Structural Mechanics in Reactor Technology (Proc. 1st Int. Conf. Berlin, 1971) North-Holland, Amsterdam (1972). [11] BAUM, M„ BUTTERFIELD, J.M., CEGB Rep. RD/B/N3742 (1977). [12] FOLIAS, E.S., “On the fracture of nuclear reactor tubes” , Structural Mechanics in Reactor Technology (Proc. 3rd Int. Conf. London, 1975) North-Holland, Amsterdam 1 (1975) Paper C4/5. [13] CONNORS, D.C., HELLEN, R.A.J., CEGB Rep. RD/B/N4039. [14] ROOKE, D.P., CARTWRIGHT, D.J., Compendium of Stress Intensity Factors, HMSO, London (1976).
APPENDIX
FAILURE ASSESSMENT PROCEDURE
1. Determine actual flaw configuration from measured indication.
2. Resolve actual flaw into a simple shape.
3. Determine stresses at location of flaw for all normal, emergency
and faulted conditions.
4. Calculate stress intensity factors for each condition.
5. Determine collapse load for structure containing flaw.
6. Determine necessary material properties.
7. Evaluate a) ratio K^/K^ = Kr b) ratio of applied load/collapse load = S r 8 . Factorise, if necessary, to define position of assessment
point, К , S , on Failure Assessment Diagram. r r a ’ r a 9. For acceptance the assessment point must lie within the failure
assessment line.
10. Determine sensitivity of position of assessment point to flaw size. IAEA-SM-218/12 81
11. Determine - maximum size to which detected flaw is calculated to grow during life of component a^ - minimum critical flaw side under operating conditions
a^ - minimum critical flaw size under emergency and faulted conditions.
12. Determine acceptability of detected flaw.
DISCUSSION
W. RED PA TH : In your paper I have seen no mention of reliability. Is it the intention of the Central Electricity Generating Board (CEG B) to follow a probabilistic route in the future in order to quantify the reliability of pressurized components? B.J.L. D A R LA ST O N : A t present we are following a deterministic approach using statistical data input where sufficient inform ation is available. The fact is that the United Kingdom licensing authorities currently require deterministic approaches. Consideration is being given to probabilistic m odels but it is accepted that insufficient inform ation is available to make it possible to use such models with an acceptable level of confidence.
IAEA-SM-218/28
SOME DESIGN CONSIDERATIONS
FOR IMPROVED RELIABILITY OF NUCLEAR REACTOR COMPONENTS
A. KAKODKAR Bhabha Atom ic Research Centre, Department of Atom ic Energy, B o m b a y , In d ia
A b stra c t
SOME DESIGN CONSIDERATIONS FOR IMPROVED RELIABILITY OF NUCLEAR REACTOR COMPONENTS. High system reliability has to be aimed at in the design of nuclear reactor components to ensure safety of the public and workers as well as to ensure minimum energy cost. This is of particular importance in nuclear reactors because of their high capital cost and high hazard potential. A number of general principles for achieving high reliability can be very effectively put to use in the design of reactor components. Some of the important ones are fail-safe design, use of assessed quality components, fault indicators, routine diagnostic inspection, good repairability and redundancy. While it is true that tills type of provisions along with a detailed systematic design leads to high costs, the reliability gained in this manner pays off much larger dividends. In this paper various such considerations have been highlighted in the context of reactors being designed in India.
1. I N T R O D U C T I O N
The current Indian nuclear power programme is based on Pressurized Heavy Water Reactors (PHW R). Fast Breeder Reactors are planned to be built in the second stage of the Indian nuclear power programme. Reactors capable of utilizing thorium would be built in the third stage. The discussion in this paper is restricted to pressure tube type PHW Rs. The scope of the paper has also been restricted to pressure components. Other hardware is not discussed. The PH W R type of system has proved to be adaptable to indigenous manufacture and the raw materials available within the country. Having established design and manufacturing capability [ 1 ], a sta g e h a s n o w b e e n reached when more efforts could be directed towards im proving the component reliability through design and better quality manufacture, so that overall improvement in the system reliability can be realized.
83 SAFETY AND SEPARATOR /REHEATER STEAM RELEASE VALVES oo г 56Q000 lb/h______580 “ I r G i f e a ,
COOLING WATER FROM LAKE. R A K D O K A K
HEATERS
7 HELIUM h RECOMBINATION !> ° A c o n t r o l DUMP | J | VALVE VALVE -±=_Г1 020 JET
— ----- ' 02<3 f PURIFICATION- У7П COOLANT ЕЩЗSTEAM HEAVY ORDINARY W MODERATOR WATER CONDENSATE WATER СИ HELIUM GAS □ LAKE WATER, и MODERATOR COOLER.
FIG.l. Simplified flow diagram. 1 psia = 1 lbf/in2 (absj = 68 9 5 Pa 1 lb = 0.4536 kg. GR OUT OIITPR
SOLID SHUT-OFF ROD ASS'Y REGULATING ROD (SHIM) REGULATING ROD (FINE) BOOSTER ROD __ LIQUID SHUT-OFF ROD ASS'Y f==i=" -ф- FLUX MONITOR --- OF OVER FLOW PIPE ----- OPR OVER PRESSURE RELIEF LINE MODERATOR INLET ОТ MODERATOR OUTLET IAEA-SM-218/28
PRESSURE RELIEF LINE —/^MAX P20 LEVE END SHIELD CALANDRIA ION CHAMBER Z — ~ MODERATOR INLET • FUEL CHANNEL 1306) VAULT LINER • - END SHIELD CARBON STEEL
00 FIG.2. Narora atomic power project reactor components general arrangement. on 8 6 KAKODKAR
The reactor core of a typical PH W R is housed in a cylindrical vessel containing a heavy water moderator at near atmospheric pressure. Fuel is placed inside the coolant channels, which penetrate the reactor core horizontally. The hot pressurized heavy water coolant is circulated through these channels by the primary heat transport system of the reactor. Figure 1 shows a typical flow sheet of a PHW R. A general arrangement of the reactor components is shown in Fig. 2. Capability for on-load refuelling, use of natural uranium and m aximum neutron economy are some of the major advantages of such a system. Since the reactor is basically of tube construction, most of the components that form the pressure boundary of prim ary heat transport system are not as heavy as the components that one com m only comes across in a pressure vessel type construction. These components are therefore well within the manufacturing capability available in a developing country like India. It has therefore been possible to direct a sizeable proportion of manufacturing effort towards achieving the high quality standards that are necessary to ensure the high reliability expected of these components. It is to be expected that should there be a large gap between the technology available and the technology required for m aking a particular component or equipment, the achievement of high quality standards would be correspondingly difficult and would tend to become uneconomic. The choice of tube type PH W Rs is thus relevant in the context of a developing country such as India.
2. RELIABILITY OBJECTIVES
Failures can be categorized depending on their effects on system performance, costs and safety. While safety-related components and systems must be highly reliable, the reliability level of other components must be commensurate with the requirement of efficient and economic system performance. Some of the objectives in addition to safety might therefore be: minimum m anrem expenditure, m inim um cost penalties due to downtime, part load operations, active coolant losses (heavy water in the case of PHW Rs) and m aximum efficiency for operation and system performance. Depending on the number of components involved and the fault tree configuration, the achievable reliability level could be worked out if reliability parameters for individual components were known. The com plexity of present-day systems calls for considerations broader than mere hardware failure probabilities. The time at which failure occurs m ay determine whether the failure is catastrophic or a mere nuisance [2, 3]. By proper selection of components and arrangement of the system it is possible to achieve optimum reliability. Various design considerations that are useful in this direction are highlighted in the follow ing section. IAEA-SM-218/28 87
3. STRUCTURAL DESIGN
One major safety function assumed by most components in the Primary Heat Transport (PH T) system is that they form the second barrier for fission product radioactivity, the first barrier being the fuel sheath itself. In the unlikely event of a failure of a PH T system component, a loss of coolant accident situation occurs with possibilities of fuel failure. High reliability of the PH T system as a whole is therefore of prime importance. Reduction in variability of material properties, fabrication variables, loading of com ponents etc. with an appropriate difference between load and strength distributions can result in improved reliability [4]. While a high reliability can be built into the individual components by adopting high standards of design analysis, inspection and quality control, achieving the same value for the system as a whole is impractical in view of the large number of components involved. In the case of pressure tube construction the existence of a large number of independent high-pressure channels running through the core needs special attention in this respect. A ll these tubes are subjected to hostile conditions of irradiation, pressure, temperature and dynam ic disturbances due to flow. Failure of any one tube could lead to outage of the entire system for a long time. One cannot resort to high safety factors to overcome this situation as this would mean more structural material in the core and consequent losses in neutron economy. Adaptation of the best possible quality standards, and routine in-service inspection thus seem to be essential. Depending on the reliability that one can achieve for an individual channel and the overall expected reliability level, the m axim um number of channels that can be allowed for a reactor is limited. W ith the continuing trend towards increasing the size of reactors the number o f channels per reactor is also on the increase. This aspect therefore needs careful consideration. Use of materials that have critical crack length at the w orking stress level far in excess of their thickness is also of importance in this connection as this could result in “leak before burst”. If one provides by design suitable indicators for such leaks, reliability against sudden bursts is greatly improved. In the case of coolant channels this can be conveniently done by m onitoring leaks in the gas annulus surrounding the coolant c h a n n e l.
Sim ilar gains can be achieved by arranging high reliability construction where access for repair is not possible. Alternatively, in areas where high reliability construction is not possible, access for repair by a proven method and feasibility of repair must be provided by design, so that any flaw detected during any one of the routine in-service inspections can be prom ptly eliminated. In the opinion of the author, rolled joints, which are com m on in tube type construction, and partial penetration welds, which occur on tube sheets and other places, are examples where provision for access for purpose of repair should be provided by design. In tube-type reactors such access is norm ally not difficult 8 8 KAKODKAR to provide. With further improvements in system designs, it m ay become possible to come closer to this goal. Rolled joints between Zircaloy calandria tubes and stainless steel calandria tube sheets and the calandria tubes themselves are some areas which are not accessible either for repair or for maintenance. Shop-assembled coolant channels which can be easily removed and re-installed could provide a way of solving the access problem to calandria tubes and their rolled joints. Furthermore, some improvement in the reliability of the coolant channels themselves could be expected because of better conditions during shop assembly. It must however be recognized that significant effort will be required to convert such a proposition into reality. There are other interconnected areas which must not be disadvantaged because of such a change. Tube-to-tube sheet joints in heat exchangers also warrant similar consideration, particularly because of the large number of joints that are norm ally involved. The proper selection of materials and techniques employed for m aking tube-to-tube sheet joints, inspection m ethods adopted, etc., are some of the factors that determine the level of reliability that can be achieved. Provision of access at each of the tube-to-tube sheet joints to enable examination on a routine basis is e sse n tia l. The use of new techniques for design, fabrication, inspection etc., must be regarded with caution. Their premature use may bring in a number of unknown factors detrimental to system reliability. Such new techniques should therefore be subject to an extensive assessment programme, enabling evaluation of all aspects that are involved. Till sufficient confidence is generated, it is worth while using a higher safety factor if such construction has to be adopted. The most powerful volumetric non-destructive methods available today for use during manufacture are radiography and ultrasonic testing. In the case of stainless steel welds, which are very com m on in PH W R components, a number of problems a rise with ultrasonic examination. Because of low pressure, most of the im portant components, such as calandria, shields, etc., tend to have corner joints which are difficult to radiograph and hence need to be covered extensively by ultrasonic examination. A number of theories have been advanced to explain the cause of problem s in ultrasonic examination of such welds in stainless steel. Special teclmiques to take care of such situations have also recently been reported to be available [5]. However, more careful thought needs to be given at the design stage to arrive at a geometrical configuration which could be more extensively examined by radiography.
4. IN-SERVICE INSPECTION
From the above discussion it is clear that sufficient reliability can be built in by modern design, manufacturing and inspection techniques. Further IAEA-SM-218/28 89 improvements are possible by resorting to well organized in-service inspection programmes at regular intervals. Various volumetric, surface and dimensional examination methods are available for this. During design, care has to be taken to ensure the feasibility of adopting these methods. Access to various regions, suitability of geometry, space for in-service equipment, etc., are some of the im portant design parameters in this regard. Provisions of surveillance coupons also give valuable inform ation on the state of the structural materials at a particular point of time. Apart from conventional methods, a number of additional techniques have now become available which could be powerful diagnostic tools to judge the state of health of a system or component. Acoustic emission technique, noise analysis, and vibration analysis using pseudo-random binary sequences are some of the techniques which could be developed into potential early warning systems. W ith the large scale application of these techniques, a significant decrease in the failure rate can be expected. Successful application of these techniques will however require a good understanding of the design of the component and the technique itself. The PH W R system, being of tube construction, has the special advantage of a large number of access points. The system can therefore be instrumented more easily and more extensively. Availability of such diagnostic instrumentation also assists the system assessment programme during early stages of commissioning. Vibrations, leakage, corrosion product transport, radioactive product transport system dynam ic behaviour, system structural behaviour, temperatures at various points in the system,etc., are some of the im portant areas of interest in such assessment programmes. Such assessment brings out very effectively the corrective steps that are necessary to m inim ize the loss of reliability during the initial ‘burn-in’ period of the system. Baseline data on the system and components necessary for com parison at any time during the later life of the system can also be generated more easily and more m eaningfully when such facilities are available.
5. ENGINEERED SAFEGUARDS
Engineered safeguards of a reactor system come into play when the normal system controls cannot take care of an accident situation. These safeguards ensure that the public is not exposed to any danger. Containment systems, emergency core cooling systems, etc., are typical examples of engineered safeguards. These systems are not required to act when normal system controls are functioning, however, in case of need they must act with high reliability. In view of the importance of these systems, they have been subjected to extensive reliability a n a ly s is [6 , 7, 8 ]. 9 0 KAKODKAR
Judicious system arrangement and layout can provide a number of advantages. Systems can be designed solely depending on, and actuated by the effects created by, an accident. Thus even though these systems are not required to act in norm al situations, high reliability is ensured when the system is indeed required. In the absence of features like this, one has to depend on relatively low reliability com ponents and in such cases, to achieve the desired system reliability, one has to resort to redundant systems. In such cases also a m inim um number of active components with facilities for routine checks are desirable. A vapour suppression pool is one example of an engineered safeguard system which is independent of any mechanical component and depends solely on the pressure differential created by the accident itself. In the case of PH W Rs built in India, vapour suppression pool systems have been adopted instead of dousing tanks for this reason, though the geometrical and layout configuration of a PH W R system is not very favourable for effective performance of a vapour suppression pool. The system configurations have in fact been progressively improved to obtain m axim um absorption of released energy in the suppression pool. Thus system adaptability and system effectiveness have to be considered to g e th e r. The fraction of released energy that can be absorbed in the suppression pool depends on the relative volumes of dry well and wet well spaces. To increase the fraction of energy absorbed the volume in the dry well should be reduced and volume in the wet well should be increased. In the case of tube type PHW Rs the dry well space is located at either end of the reactor. If these two areas are isolated to form two separate dry well spaces connected independently to the suppression pool, the optim um ratio for dry well space to wet well space can be achieved. In fact it has been calculated that more than 90% of the released energy can be absorbed this way without a very significant increase in pressure loading on the primary building containment wall. Sim ilar advantages can be achieved by m aking proper use of gravity in the case of safety systems. If possible, it is preferable to arrange a gravitationally stable configuration for the system in shut-down state. Water storage at high elevation can also be used very reliably for a variety of im portant functions such as emergency cooling, emergency motive power, etc.
6 . EASE OF MAINTENANCE
Ease of maintenance has to be built into a system to ensure high reliability against failures during operation and also to ensure high system availability. Apart from conventional considerations to ensure ease of maintenance, nuclear system design has also to take care of problems created by radiation and needs IAEA-SM-218/28 91 to ensure uninterrupted availability of some of the systems. This means provision of adequate stand-by arrangements for bringing in and taking out of components or equipment for maintenance. The layout should ensure proper access to all areas requiring maintenance, with radiation fields brought down to manageable level at least during maintenance. The selection of materials, system cleanliness prior to start of operation, and chemistry control are some other parameters which have to be properly monitored to prevent future maintenance from being jeopardized due to man-rem problems.
7. VALVES AND SEALS
Heavy water losses have to be kept to a low level in PHW Rs. All components having m oving parts at the pressure boundary of heavy water systems have to have high reliability against leakage. Valves and seals are com m on components of this type. The best way of im proving the reliability of the overall system against leakage is to eliminate as m any valves as possible and to use high quality minimum-leakage valves where valves are essential. Special valves for this kind of service have been developed. In m any system designs a number of lines are in the form of very-small- diameter tubing. Because of their limited strength and high flexibility, there is always some chance of rupture of these lines due to over-vibration or some accidental condition. This chance can be minimized by providing stronger tubes and proper supports along the tube run.
8 . S U M M A R Y
While there are a number of ways of im proving overall system reliability, one has to depend on high quality design, manufacture and inspection to achieve the desired reliability for reactor components. Proper material selection and sim plicity in operation are also desirable. In-service inspection at regular intervals helps reduce the probability of failure. A number of advantages with regard to reliability, ease of maintenance and efficiency of performance can be gained by careful design.
REFERENCES
[1 ] MEHTA, S.K., et al., “Innovations in PHWR design, integration of nuclear power stations into power systems and the role of small nuclear power plants in a developing country”, Nuclear Power and its Fuel Cycle (Proc. Int. Conf. Vienna, 1977) 6, IAEA, Vienna (1977) 403. 92 KAKODKAR
[2] WEISS, D.W., “Prediction of mechanical reliability — a review” , Proc. Annual Symposium on Reliability, Washington, D.C. (1967) 30—39. [3] HINELY, Jr., SHELIEY, B.F., “Reliability optimisation in the conceptual phase”, Proc. Annual Symposium on Reliability, Washington, D.C. (1967) 209-217. [4] CABIE, C.W., VIRENE, E.P., “Structural reliability with normally distributed static and dynamic loads and strengths”, Proc. Annual Symposium on Reliability, Washington, D.C. (1967) 329-336. [5] HERBERG, G., et al., “Preliminary results for practical ultrasonic testing of austenitic steel welds”, NDT International (October 1976) 239—241. [6] BALFANZ, H.P., et al., “A reliability approach to inspection stragety applied to an emergency core cooling system” , Structural Mechanics in Reactor Technology (2nd Int. Conf. Berlin, 1973) V, Part M. [7] BALFANZ, H.P., et al., “Principles of reliability analysis methods applied to an emergency core cooling system” , Nucl. Eng. & Des. 29 (1974) 384—394. [8] BALFANZ, H.P., et al., “Prediction of failure rates and failure modes of mechanical and electrical equipment components by the fault free method and failure effect analysis” , Kerntechnik 13 9 (1971) 392-399.
DISCUSSION
S.M. BH U TTA: Have the suggestions for design modifications which you mentioned in your paper, such as removal and re-installation of coolant channels and provision of suitable indicators in the gas annulus surrounding the coolant channel, been incorporated in the designs for your future reactors? A. K A K O D K A R : We have been progressively moving towards self- sufficiency. Initially, attention was paid m ainly to fabrication and installation. Later, achievement of design capability became the target. N ow we have reached the stage where reliability consciousness plays a part. Certain aspects, such as the m onitoring of the coolant channel annulus for leaks, have already been considered for designs on which work is currently being done. S.M. BH U TTA : What steps have been taken to minimize or prevent heavy water leakage? A. K A K O D K A R : Heavy water leakage has been a problem and much attention is being paid to it in the design for the Narora Atom ic Power Project. Steps have been taken to isolate possible points or areas of leakage in order to minimize the leakage and degradation of heavy water. S.M. BH U TTA: Do you carry out formal quality assurance programmes in
In d i a ? A. K A K O D K A R : Yes, the importance of quality assurance is fully recognized
in our country. IAEA-SM-218/28 93
J.S. M A C LE O D : In the abstract of your paper it is stated that in order to achieve high reliability, use is made of redundancy. Is diversity of systems or components also inherent in your philosophy or only redundancy? A. K A K O D K A R : I agree that diversity of systems and components helps towards the improvement of system reliability. However, in m y opinion, as far as pressure-retaining com ponents are concerned, it is more im portant to adopt better quality assurance procedures in design, fabrication and installation along with improvements in design philosophy.
IAEA-SM-218/33
PREDICTION OF FAILURE RISK
IN A PRESSURE VESSEL DUE TO A MANUFACTURING DEFECT
D.L. M ARRIOTT, J.M. HUDSON Licensing Branch,
South African Atom ic Energy Board, P e lin d a b a , South Africa
A b stra c t
PREDICTION OF FAILURE RISK IN A PRESSURE VESSEL DUE TO A MANUFACTURING DEFECT. A significant contribution to the probability of failure of a reactor pressure vessel is from fracture initiated by undetected manufacturing defects. Defect distributions are not necessarily similar in all vessels, therefore the risk to an individual vessel may differ markedly from the average of a vessel population. A procedure has been developed to derive a defect population from the manufacturing process, and use this in calculating the probability of failure of a specific vessel. The analysis is in two parts: (a) a study of the manufacturing process using techniques of systems analysis such as fault trees; (b) an approximate stress analysis to assess the risk of failure for a given defect distribution. Findings are limited because the methods depend on specific information obtained during manufacture but, by using typical data, where available, it has been possible to draw some general conclusions about the characteristics of risk analysis. These are: (a) access to data is a major problem. The paper indicates what information is required for assessment of an individual vessel; (b) weld defects and small zones of high stress dominate other contributions to failure probability. This finding is not novel but the degree of dominance is surprisingly large; (c) it is feasible, on the basis of simple analysis, to establish an allowable defect size and distribution for a given failure probability which can be used to set manufacturing standards; (d) the systems analysis can be used to identify phases of manufacture with the greatest corrective influence, if the defect population predicted is likely to be critical.
1. INTRODUCTION
No direct inform ation on failure rates in reactor pressure vessels (R PV s) is available, although several attempts have been made to infer such information from non-nuclear vessels [1,2]. It has been claimed, on the basis of these studies, that the expected risk of failure of a typical R P V is of the order of 10 " 6 p e r vessel year. This figure would be acceptable if it could be assigned to a specific vessel, but unfortunately it is an average value and applies only to a hypothetical
95 96 MARRIOTT and HUDSON population of similar vessels. In safety assessment it is the specific vessel which is of concern. Before average failure rate data can be used it is necessary to ensure that the construction and operation of the pressure vessel under examination is representative of at least the average vessel of the population. This entails a more detailed examination of the potential for variation in quality between vessels, and the effect of this variation on the probability of failure. A major contribution to the failure probability of RPV s is due to fast fracture from a defect. Besides the potentially catastrophic nature of this mode of failure, it is also one for which the risk can vary greatly from one vessel to another. This is because the risk of failure due to the presence of an undetected defect depends, firstly, on the defect population in the specific vessel concerned, and secondly, on the success of any subsequent non-destructive examinations (NDE), both of which are dependent upon the standards of workmanship and quality control used in manufacture. Reports of Westinghouse pressure vessels [3] mention variations of between zero and 6 6 defects per vessel, which suggests that, for the fracture m ode of failure at least, an average figure is unlikely to be applicable to individual vessels. From the above discussion it is apparent that there is a need for methods to predict the incidence of defects in a particular vessel. The problem would be less severe if N D E were 100% effective, but this is not so. In fact the risk of not detecting defects of critical size is too high to be accepted without some prior estimate of the probability of such a defect existing in the first place. The object of this paper is to describe one method of assessing the likelihood of failure due to the presence of undetected defects. The method has two parts: (a) a synthesis of defect populations from detailed considerations of manufacture; (b) an approximate analysis of risk from a given defect distribution.
It has not been possible to make a complete quantitative assessment of
fracture because of the lack of data. The analysis outlined below should be con sidered as a proposal for dealing with the problem, with the initial objectives of identifying what information needs to be collected during and after manufacture, and how this information can be used. In some cases limited data do exist and these have been used to draw some conclusions.
2. PREDICTION MODEL FOR PROBABILITY OF AN UNDETECTED DEFECT IN AN RPV
2.1. Types of manufacturing faults
During the manufacture of an RPV, it is inevitable that defects will occur to some degree, regardless of the manufacturing procedures, techniques and IAEA-SM-218/33 97 controls employed. A knowledge of the origin of significant defects, their detection and subsequent identification (important for determining the acceptability of the product) enables an assessment to be made of the integrity of the manufactured parts stage by stage throughout the m anufacturing process. The majority of defects form during casting and since only a cursory visual examination of the ingot surface is possible at the time, m ost of the defects escape detection until a later stage of manufacture. Sophisticated N D E techniques are impractical owing to the large section thickness of the ingot, coarse grain structure, etc., prior to forging. Defects due to sh rin k a g e, a contraction of the ingot during the solidification process, can be classified as either primary pipe o r secondary pipe. The former occurs as a large cavity at the extreme top end of the ingot and is caused by an insufficient volume of molten metal in the ingot head. It is oxidized and there fore can never be completely closed by forging. However, this type of defect is rare, easily prevented and can be detected visually when removing the top end discards after forging. Secondary pipe is an internal shrinkage caused by insufficient feeding of molten metal to the body of the ingot when a “bridge” of solidified metal forms at the head of the ingot. Since it is not oxidized, it can usually be healed by forging, otherwise the only means of detection is by ultrasonic testing (unless it outcrops on a machined end-face, when visual and surface N D E can be e m p lo y e d ). The preferential solidification of elemental com ponents of the steel, a d verse segregation, will always occur to some extent, the effect being more pronounced in large ingots. As such, segregation is not regarded as a defect unless it is associated with flaws due to shrinkage, non-metallic inclusions or hydrogen cracking. Segregated areas contain the constituent melts in a different proportion to the overall analysis. The most effective means of detection is by sulphur printing, the sulphur acting as a tracer for the other alloying elements that segregate simultaneously with the sulphur. Non-metallic inclusions m ay originate from the slag and refractories of the steel-making process (exogenous) or from the products of de-oxidation (indigenous). The severity and extent of inclusions, which generally favour the region towards the bottom of the ingot, are determined by the casting conditions, in particular the casting temperature. Such inclusions can affect the integrity of the manufactured R P V in a variety of ways, but the probability of failing to detect their presence can be reduced to a low level provided full use is made of the available detection techniques, i.e. ultrasonic testing (UT), visual examination, mechanical testing and surface N DE, including sulphur printing. Similar comments apply to the defect known as sk u ll, which is an area of non-metallic inclusions trapped beneath a prematurely solidified raft o f steel. The accidental entrapment of metal other than the parent cast, foreign metal, is a rare defect which ought to be prevented altogether by observing proper 98 MARRIOTT and HUDSON control and reporting procedures. However, this defect is important from the point of view of difficulty of detection should it occur. The forging process can give rise to forging bursts, caused by an insufficiently high forging temperature and/or insufficient soak time, and to forging laps, a mechanical defect arising during the forging operation when there is an incomplete fusion between two surfaces squeezed over each other. Forging bursts will not be a problem for hollow parts since the axial cavities so formed will be removed with the centre of the ingot. Forging laps are heavily oxidized and are readily detected visually on the machined surface. A potentially harmful type of defect, which could affect the weldability of the steel and which form s during the heat treatment process, is hairline (or hydrogen) cracking. The cracks are very fine, form ing in the body of the forging, and are associated with hydrogen in the steel. Vacuum degassing has helped to reduce significantly their probability of occurrence. Detection of the defect is difficult — visual and sometimes dye-penetrant testing fail if the defect does not outcrop on the machined surface — and U T and magnetic-particle testing must b e u se d . The machining operation is a potential source for grinding cracks w h ic h form by localized heating of the metal surface during grinding. Since this is a surface defect, detection is not unduly difficult. The major source of harmful defects stems from the welding together of the R P V components. It is imperative that the m ost stringent control procedures are observed and adhered to during welding in order to minimize the probability of significant and/or extensive flaws in the welds. Likewise, all available techniques
should be fully utilized to ensure that cracks which have occurred do not escape d e te c tio n .
2.2. Manufacturing defect model
Figures 1 and 2 set out in the form of simple fault tree logic the variables
associated with the propensity for defects to form in a vessel casting or weld as manufactured. Figure 3 shows how, using the results generated from Figs 1 and 2 ,the probability of having an undetected defect in the vessel Рди m ay be synthesized. The reactor vessel of a three-loop PW R is considered. Although Fig.3 depicts weld defects as only one box of the tree, the problem of generating a probability figure for this input is, of course, not so deceptively simple. Weld defects can be split into defects in the initial weld and defects in welds subsequently repaired. For weld repair defects, Fig.l is likely to require some modification. Modification to Fig.2 is likely, depending on which forged part is considered. As an example, secondary pipe is likely to be an irrelevant consideration for parts
where forging reductions are high. IAEA-SM-218/33 99
CONTROLS
Pouring Temperature 8 Supervision and Control Method of Manufacture 9 Orying of Electrodes Observation and Vacuum Oegassing Records Mould Drying Surface NOE before Heat Treatment Welding Pouring Temperature Control of Material De-oxidation Control defects in Weld Area Melt and Refining Sequence 13 Pre-Heat System Bottom End Oiscard and Controls Control of Dimensions Content of Electrode and Fit-up Analysis of Electrode Machine Shop Controls Weld Procedure and Orawing and Design Procedural Tests Controls Welder Quaiificotion Machining Allowances Design for inspection In Process NOE Heat Treatment Controls
UT = ultrasonic testing NDE r non-Destruction Examination
FIG.l. Fault tree of undetected surface weld defects.
Further questions arise as to the controlling variables associated with data on weld defects. Should the data be collected per metre length of weld, or by volume? Is there a significant spread in results attributable to different welders? How does the degree of weld difficulty affect weld quality? The answer to these and other significant related questions will influence the com position of the m odel as outlined in Fig.3. 100 MARRIOTT and HUDSON
Dimensional ч inspection _( j as forged 4^/
Dimensional inspection f j
CONTROLS ( continued from FIG.1)
E. 1. Control and Observation Ouring Forging 2. Efficient Scarfing
1. Ingot Head Size 2 Top End Discard 3. Visual Observation of Stripped Ingot L Visual Observation After Oiscard Removal 5. Method of Manufacture ( i e. remove core and hollow forge)
1. Ingot Design 2. Pouring Temperature 3- Method of Manufacture i. Forging Reduction
1. Casting Observation 2. Pit Controls
1. Method and Control of Scarfing and Grinding to Prevent Damage Oue to Local High Temperatures 2. Scarfing / Grinding Records
xujru Mechanical test results Surface onc-oc-moreFailure of overheat u - n or burn — caused by scar ling or grinding Fail to detect bysurface NDE of 4 4—1' )
FIG. 2. Fault tree o f undetected defects in vessel casting. IAEA-SM-218/33 101
FIG.3. Fault tree o f undetected defects in RPV as manufactured. 102 MARRIOTT and HUDSON
FIG.4. Size distributions for different defect populations.
FIG. 5. Schematic representation of NDE Probability of failure to detect.
2.3. Use of the model
The obvious missing ingredient necessary as input in the calculation of Pdu is data. Those parts of the tree which read Failure of one-or-more controls ... could be replaced by defect distributions of perhaps the forms shown in Fig.4, where j is the probability of a defect of type 1 occurring; P d2 is the probability o f a t y p e 2 defect occurring, and so on. Associated with each defect type is the probability of its detection using all the methods available. For each defect type, a curve of the form shown in Fig.5 m ight be appropriate, where Pdt is the probability of detecting a defect of a given ty p e . It m ust be pointed out that Fig.5 relates only to cracks that are present in the vessel. Indications are that occasionally cracks are detected which on subsequent repair action are found to be non-existent. This will necessitate a refinement to the m odel since subsequent repair action m ay reduce the integrity o f the vessel. IAEA-SM-218/33 103
Once suitable data are available, the problem is one of establishing the value of Pdu. For m utually exclusive events 11
P d u O c ) p dip dtdc О) - 1 7 i = 1 c
where n is the number of defect types considered to be relevant and c is any suitable defect size o f interest. Equation (1) expresses the probability of a defect exceeding a certain size. In order to determine whether this leads to fracture it is necessary to introduce a fracture criterion and perform a stress analysis. In practice both the fracture criterion and stresses existing at any point in the structure are subject to uncertainty, so that a full analysis is extremely complex. The following section describes a simplified approximate analysis which enables a number of results to be obtained without extensive computation.
3. APPROXIMATE ANALYSIS OF FAILURE PROBABILITY
There is no fundamental difficulty in predicting failure probability due to random defects. Given variabilities of stress, material properties and defect population, it is possible to obtain a numerical solution by simulation. This approach has been used by W ilson [4] to predict pipe failure and by Becher and Pederson [5] for reactor pressure vessels (RPVs). Unfortunately it is difficult to obtain general conclusions from this type of analysis without considerable computation. The object of the analysis given here is to extract some general results about probabilistic behaviour of vessels by deliberately retaining as simple a treatment as possible. M any of the assumptions are approximate, but it is considered that the resulting m athematical model retains the essential qualitative characteristics of the physical problem. Although no attempt at quantitative prediction was originally intended it was found to be possible in some cases.
3.1. Assumptions
(a) Only the initial defect problem is considered. Defect growth can be dealt with, but is omitted here due to lack of space. It is claimed by Becher and Pederson [5] and Rasmussen [6 ] that defect growth in RPVs will be small compared with initial defect size. The problem considered here is not therefore unrealistic. 104 MARRIOTT and HUDSON
(b) Defects are characterized by a single size parameter c, which is related to applied stress through the criterion,
К > KIc (2) w h e re К = 0 \/W c , the stress intensity (3)
a is the applied stress, and K jc is the fracture toughness. This simple linear elastic fracture mechanics criterion (L E F M ) is inaccurate, but not seriously so for the order-of-magnitude calculations considered here. A more serious problem is the question of high-stress fracture. It will be seen later that the high-stress areas are im portant contributors to failure probability, but the theory of high stress (or post-yield) fracture is insufficiently developed to be incorporated in a probabilistic study at this time. (c) Fracture toughness Kjc is assumed normally distributed,
f ( K i c ) = ( 2 ir a l Г 1/2 e x p (- K ^ 'j ( 4 ) ' l o t /
K i c is a v e rage
ctk is st. deviation
Results obtained from the Heavy Steels Section Technology Programme (H SST) [5] show fracture toughness variations to be very approximately normally distributed, and for the purpose of this analysis the assumption is adequate. (d) Defects are assumed to be exponentially distributed in size and the number uniform ly distributed in volume.
c f ( c ) = ! e x p ( 5 ) c c
where f(c) is the frequency of defects of size c p is the volum e density and c is the mean value of c. Inform ation on defect distributions is difficult to obtain. The very limited data sources, e.g. [1 ], support an exponential form and, in the absence of any better information, this has been accepted by other workers [5]. The work described in the first part of this paper clearly shows that defects are neither uniform ly distributed through the vessel, nor are they necessarily all of one type. However, it is possible to deal with this problem by subdividing the IAEA-SM-218/33 105 vessel into subsections, e.g. welds, heat-affected zone (H A Z), parent plate, etc., which defects of certain types are assumed to be uniform ly distributed. Different types of defects can be dealt with as a series of separate problems. The total failure probability is given by sum m ing over all defect populations as specified in Eq.(l). All subsequent analysis herein will assume a single volume and a single defect population. The extension to the more complex situation is straightforward.
3.2. Analysis
Consider an element of volume dV under stress a. By integrating Eq.(5) the probability of finding a defect size greater than some value c is
dp(defect > c) = pdV'exp (6)
Changing the variable to stress intensity K, the probability of a given value of stress intensity exceeding K, given stress a is , 2
P(Stress Intensity > K ) = pdV-exp ( 7 ) w h e re
К = O a/W
From Eqs (4) and (7) the probability of the stress intensity exceeding a particular value of K jc can be established, and integrating over all values of K, the probability of fracture in element d V under stress a is
1/2 K 2 _ ~ K Ic P ( F ) d V = PdV| ) e x p (8) -K 2 + 2 a 2K K 2 + 2 o \
In a vessel under load W, the total probability of failure is
i l l K 2 ~ KIc P ( F ) v e x p d V (9 ) = /" ( p 2ak K 2 + 2fffr
К = ff(W )
and a(W ) is the varying stress distribution through the volume due to the load W. 106 MARRIOTT and HUDSON
Generally speaking, Eq.(9) can only be integrated numerically. Calculation of P(F) for “exact” stresses will not be pursued any further. The problem of concern here is the situation where the exact stresses are unknown, either because the stress analysis is inaccurate, or because there is some uncertainty due to residual stresses. Retaining only the requirement that stresses should be in equilibrium it has been possible to obtain bounds on. failure probability under certain circumstances which are independent of the details of the stress distribution. The integral in Eq.(9) can be shown to have a stationary value given by
/2 Keff -Kfc p V - e x p (10) K eff + 2 a K K eff + 2 4 w h e re
Keff = (Jeff
and aeff is an effective stress given by
V
The stationary value of P(F)V is a maximum (upper bound) or minimum (lower bound) depending on the relative magnitudes of K eff and K lc. The relation between P(F)V and (K eff/Kjc) is given in Fig. 6 . Where the curve is convex (AB) Eq.(l 0) gives a lower bound. Where it is concave (BC) the bound is an upper bound. These two zones also correspond approximately to two basically different types of crack population whose characteristics can be summarized as follows:
Type (i): (K eff/Klc ) 1. Equation (10) gives lower bound. p V can be large. Average crack size c" is small. P (F )y is very sensitive to stress changes since the exponential term in Eq.(10) dominates.
Type (ii): (K eff/KIc) > 1. Equation (10) gives upper bound. p V must be small because the exponential term tends to unity. Large average defect size P(F)V depends predominantly on the defect frequency, pV, and is insensitive to stress analysis. IAEA-SM-218/33 107
e w ^ d
FIG. 6. Dependence of failure probability per unit volume on (Keff[K\c)-
The limited evidence to date suggests that defect populations in RPV s tend to be of type (i). This means that the bound obtained from Eq.(10) is limited to establishing a component to be definitely unsafe. An upper bound would be a more useful tool. Nevertheless the fact that some bounding is possible suggests that further work could produce more practical bounding techniques. The fact that the exponential term in Eq.( 8 ) is very sensitive to stress changes causes considerable difficulty. Taking some typical values from Westing- house data [1], 12 defects greater than 12 m m were found in 44 pressure vessels examined. It can be taken that the defect frequency is of the order of 1. For the purpose of illustration it will be assumed that
K ic = 2 2 0 M N - m ‘ 3/2 ° K = 17MN-m-3/2 Yield stress, a Y = 420 M N-nT 2 Belt line stress, = 175 M N-nT 2
Residual stress, °R = 55 M N - n T 2
Under steady operational conditions in a Pressurized Water Reactor, significant am ounts of material at nozzles and other stress concentrations experience stress of the order of the yield stress ay • As will be shown later, the exact percentage of highly stressed material is not critical. Taking a figure of 1% of highly stressed material, the contribution to failure from this material is, from Eq.(9), 108 MARRIOTT and HUDSON
р(р Ч 10 2 ( p V ) e x p (11) чоутгс + 2
р(р Ч will be of the order of 1 0 6 or less so that
) .,0- a \ 7ГС + 2 /
This is achieved if the average crack size is
c = 8 .7 5 m m
Evidence quoted by Hellan [7] indicates that, even with refined finite element analysis, errors in stress levels can be ±5%. This variation causes P (F )0y to change by a factor of 20 in the presence of residual stresses up to 55 M N -л г 2 in amplitude but of unknown position. For the same crack population, at lower stresses, the contribution of an element of material d V to total failure probability is m any times smaller than the contribution of high stress, e.g. in the belt line re g io n ,
P(F) IO " 12 P ( F )uy a
The only sensible interpretation of figures of this magnitude is that low stress areas provide a negligible contribution fo failure probability. Areas of low
stress are considerably more sensitive to P(F). At ctb = 175 M N-m -2 v a r ia t io n of ±5% in stress causes P(F) to vary by nearly 4 orders of magnitude. It seems unlikely, in the circumstances, that any stress analysis, however carefully performed, can be relied on to give a reasonable estimate of failure
probability. Prediction of failure probability for a given crack population is analogous to calculating deform ations in an elastic, perfectly plastic structure as it approaches collapse. Sm all variations in the input lead to large variation in the result. However, it is more useful to determine an allowable size of defect for some given probability of failure. Then Eq.(l 1) m ay be generalized to represent a volume Vj in which the dom inant stress level is a j, and inverted to give the permissible crack size Cj
K Ic (lnCPiV^-lnCPifFVJr1^ (12) Traf
where Pj(F)Vi = contribution to P(F)V from volume V; at stress a;. IAEA-SM-218/33 109
For representative values of Pj(F)V i = 1СГ6, and (PjVj) = 1СГ 1 the average
defect size is virtually independent of the term (PjVj). cj varies by ±50% for a variation of (PjVj) from 10 ~2 to 102. Even the crudest of calculations should enable the value of material at a given stress level to be estimated with far greater accuracy than ±2 orders of magnitude. Even fairly gross overestimates of Vj to include material at lower stress levels will not change the value of cj significantly.
3.3. Strategy for ensuring a given level of reliability
A simple strategy emerges for establishing a desired level of reliability. (a) Divide the component into sub-volumes Vj based on significant stress le v e ls 0 j in each sub-volume (e.g. weld, parent material, stress-concentration r e g io n s ). (b) Calculate the values of c] for each section using conservative estimates ofCOjVj). (c) The values cj represent the admissible crack sizes in each of the regions in order to achieve the target reliability.
N B (i) This calculation is not only simple, it is more consistent with design or safety assessment than attempting to predict failure probability for a given defect distribution, in that the performance is specified and the acceptability lim it is the derived quantity. (ii) Residual stresses are automatically accounted for, as long as their magnitude is known. This is done by increasing Oj to (oj + aB ). Uncertainty over the position and extent of the peak residual stresses is not critical due to the insensibility of Cj to the assumed value of (PiVi)-
4. IMPLICATIONS OF ANALYSIS
4.1. Some preliminary results of predicting the incidence of defects
A relatively simple exercise was undertaken by the authors where the model was used to generate the point probability estimate that undetected defects greater than 25 mm would be present in an RPV. It must be pointed out that the data used were of a m ost rudimentary nature in that they were purely the collective judgement of a few qualified people. The results are interesting and in some respects predictable.
Pdu (>25 mm) = 7 X К Г 5 п о MARRIOTT and HUDSON
FIG. 7. Critical crack sizes based on linear elastic and post-yield fracture theories.
Weld defects were estimated to account for something in excess of 85% of this figure for Pdu (the probability of having an undetected defect). The principal com ponent in this estimate for weld defects is the internal one, contributing almost 100%. This reflects the difficulties associated with U T and radiography, e.g. inaccessibility, extent of inspection, defect orientation, m asking of defects by inclusions, and operator limitations.
4.2. Further consideration of zones of high stress
L E F M is not accurate for fracture stresses above about 0.6ay. Post-Yield
Fracture Mechanics (P Y F M ) predicts a relation between failure stress oF and defect size c as shown in Fig.7. (See Heald et al. [8].) Taken at face value, this curve leads to even higher fracture probability in zones of high stress than predicted by the L E F M model, when intuitively one m ight expect less likelihood of fracture where yielding has occurred. This is an obvious misinterpretation of the meaning of the fracture curve in Fig.7, but it has been done deliberately to illustrate a shortcom ing in the present state of PYFM . The problem is that the m o d e of failure changes at high stresses and the consequence of reaching the failure stress is ductile yielding. W hile it is possible to identify the failure m odes for extremely large cracks and extremely small cracks as brittle fracture and plastic yielding respectively, the behaviour at inter mediate stress levels is not defined, and the im plications of this behaviour in a redundant structure are by no means clear. Until this question has been answered there is little purpose in extending probabilistic studies to take post yield fracture into account. IAEA-SM-218/33 11
4.3. Effects of material variability
Some conclusions can be drawn from Eq.(12) regarding material variability. It can be seen that for q to be positive
( 1 3 )
For a wide range of (p^V;) and Pj(F)Vi this means that
< a p p r o x 0 .2
Values for aK quoted by Marshall suggest that R P V steels experience variability (aK /Kjc) яв 0.08. More confirmation of this degree of control would be helpful. Under conditions of deterioration, such as irradiation damage, K Ic can be expected to decrease and aK to increase. The lim it expressed by inequality (13) cannot be dismissed as unattainable. Even if this limit is not reached an increase of (a^/Kj,.) can seriously lim it the admissible crack size.
4.4. Effect of N D E
The probability of undetected defects entering service is obtained by com bining knowledge of the pre-existing defect distribution with a measure of the probability of “failure-to-detect” of the N D T procedure used. If the defect population in a single vessel is reasonably large the probability
distribution of undetected defects can be calculated relatively easily. The probability of failure-to-detect p(d) varies with defect size. It is very high for large defects and is thought to decrease to a constant value at large defect sizes as shown in Fig.5. The form of p(d) is even less understood than other defect parameters. However, to examine trends p(d) can be modelled as fo llo w s ,
P ( d ) = a + (1 - a ) e x p (-/3 c ) (14)
I f c' represents the size of defect remaining undetected after inspection,
f(c') = f(c)p(d) 112 MARRIOTT and HUDSON
From Eqs (5) and (14), and integrating, the probability of undetected defect exceeding some value c is
P(defect > c ') = pV ( 1 5 )
Calculation of failure probability after inspection is identical to the procedure for failure probability due to pre-existing defects with different constants in the exponential expression for crack distribution. When the defect population is sparse there are considerable practical difficulties in inferring the likelihood of undetected large defects. Unless a grossly conservative estimate is to be accepted it is necessary to obtain further inform ation by indirect means. A t least three possibilities exist: (a) Inform ation from manufacturers on the variability to be expected between individual pressure vessels. (b) Synthesis of critical defect distributions from information obtained earlier in the construction process. (c) Inference from statistics of frequent but less critical defects to obtain the incidence of critical defects. The form at for investigating the last two possibilities has been developed and has been presented in this paper. However, all three possibilities depend on much freer exchange of N D E information than is the case at present.
5. CONCLUSIONS
(1) There is a need for more information on defect statistics. In m any cases information is non-existent or presented in a form unsuitable for analysis. This paper m ay be helpful in illustrating how the data m ight be used if presented in the right form. (2) In the context of probabilistic fracture analysis it is possible to answer a number of im portant questions without recourse to complex stress analysis. The m ost im portant of these questions is the size of defect considered acceptable for a given failure probability. It emerges that the average crack size is an im portant parameter. (3) The problems of high-stress and post-yield fracture emerge as dominant in component failure probability calculations. Identifying the significance of P Y F M is not original but the degree of importance, and the need to understand the probabilistic consequences of high stress fracture, need to be recognized. IAEA-SM-218/33 113
REFERENCES
[1] UKAEA Report of Study Group (Chairman: Dr. W. Marshall) “An Assessment of the Integrity of PWR Pressure Vessels”, HMSO (1st Oct. 1976). [2] PHILLIPS, C.A.G., WARWICK, R.G., UKAEA Report AHSB(S) R 162 (1968) 35. [3] NICHOLS, R.W., “The assessment and assurance of pressure vessel reliability”, Proc. Inst. Mech. Engrs. 189 47 (1975) 13. [4] WILSON, S.A., AEC Research and Development Report GEAP-10452 (1972). [5] BECHER, P.E., PEDERSON, A., Application of statistical linear elastic fracture mechanics to pressure vessel reliability analysis, Nucl. Eng. Des. 27 (1974) 12. [6] Reactor Safety Study, USNRC Report WASH-1400 (1975). [7] HELLAN, T.K., DOWLING, A.R., 3-D crack analysis applied to an LWR nozzle-cylinder intersection, Int. J. Pressure Vessels Piping 3 (1975) 17. [8] HEALD, P.T., et al., Post yield fracture mechanics, Met. Sei. and Eng., 10 (1972) 9.
DISCUSSION
T. VARG A: The distribution curves you have shown are, in the light of our experience as pressure vessel manufacturers, true for rather small defect sizes. If one goes above the 3 m m reference reflector, very few indications are found for a full pressure vessel. O f course this depends very much on the manufacturing and testing techniques used. There have been vessels manufactured a few years ago which were tested only with X-rays; on one of these a weld seam which had already been accepted showed under ultrasonic examination a defect one metre long and twenty-five millimetres deep. But if the manufacture and testing procedures are up to date, the distributions shift towards rather small defect sizes. D.L. M A R R IO T T : I agree with everything you have said; in fact the object of our visit here is more to obtain information rather than supply it: almost everything we have done has been based on speculation because our data are very sparse. This is an area in which, generally speaking, much extrapolation is done: from very little information, and that is why we have produced no absolute value for failure probability. However, it would be useful if people with information on this topic would publish it. N.F. H A IN ES: The probability of detecting a defect by ultrasonic techniques is not purely a function of size. In fact, orientation of the defect relative to the incident wave direction is the single m ost im portant factor in the detection of large defects. D.L. M A R R IO T T : We have found that the concept of the probability of the failure to detect a large defect is a rather too simple one. For example, if you ask ultrasonic inspectors what the probability of finding a defect is, they will say that it is 100%. But there is a large difference between their certainty that they will find something and the extent to which they are prepared to inter pret their findings. 114 MARRIOTT and HUDSON
A.B. L ID IA R D : I agree that the value of calculations of this kind lies at the present time not so much in terms of absolute predictions, as in what they show to be the failure probability sensitivity to such things as the incidence of cracks, which we do not know very well. However, in this connection I think it is im portant that one calculates the right failure probability function. I believe you should be looking at the function which gives the failure probability per unit time and that this function should allow for the fact that the vessel has passed the pre-service pressure test.
D.L. M A R R IO T T : We are aware of the time-dependence factor, but since our organization is a small one we are starting with the modest objective of seeing a pressure vessel getting into service and only then will we be looking at how it can remain in service. Where the hydro test is concerned, it is very often argued that the test eliminates defects above a certain size. Although I have not yet formed an opinion on this, it has also been argued that the types of stress applied in the test are not necessarily those which will reveal the flaws which might occur under various upset and emergency conditions. Perhaps the failure of the Cockenzie boiler drum at the third test is of relevance here.
A.B. L ID IA R D : From the formal point of view, there is no doubt that the cold-hydro test should apply stresses more severe than those which the vessel is subjected to in norm al service. However, it is said by some that it is hard to believe that the cold-hydro test has the validity ascribed to it by linear-elastic fracture mechanics. This is an im portant question which needs discussion. J. FO RSTEN : What information do you need apart from the data obtained in ordinary quality control tests? Do you propose any modifications to existing destructive or non-destructive examinations or new and additional tests? D.L. M A R R IO T T : Ideally we would like complete information on fracture toughness variations, typical defect distributions for different kinds of defect, etc.
However, what we actually can have is decided at a higher level. I do not in fact believe that we shall be able to have all this inform ation; in any case we are not interested in vessel populations. This is because, however effectively surveys may show, for example, that the probability of PW R vessels failing is generally less th a n 1 0 ~6, they do not help us because we need to be sure that our vessel is not the one out of 10 6 which does fail. So we need information specific to our vessel. How this information is acquired is rather a complicated question. We are not going to get information on fracture toughness, for example, which will show variations in a particular vessel. That means that we will have to do some kind of correlation between, on the one hand, surveillance tests on Charpy, with equivalent Charpy tests on generic data and, on the other hand, what we expect our material to be relative to the generic material. For instance, put in simple terms, if generic fracture toughness data on A508 are satisfactory, if equivalent Charpy data on that particular material are available, if we have results of Charpy tests on the material of which our vessel is made and if our Charpy IAEA-SM-218/33 115
tests are superior to the generic tests we will have some reason to believe that our fracture properties will also be superior. As far as defect data are concerned, quite useful information can be gained from simple numbers, such as that of the weld repairs which have been made during the construction of the vessel. I do not think anyone is going to make a complete ultrasonic or radiographic inspection of a vessel and evaluate every single flaw picked up. There will have to be a cut-off somewhere. But this is the sort of information which would be used to estimate defect distributions. However, perhaps we can find a middle way by comparing
the number of defects noted in a particular run of weld with the number of other vessels. If we find fewer defects on average in our vessel, we can expect it to be somewhat better than the average vessel. So our probabilistic analysis will have to be carried out at a level one step below the ideal level at which one would have all the distributions and combine them in a complete Monte Carlo simulation, for example, in order to find the actual failure probability. In any case, if one takes the stress distributions from, say, two separately performed finite-element analyses for a fairly complicated structure and one works out the failure probability for each of those analyses, although they m ay both have been done in good faith, a difference of orders of magnitude can be found, sim ply because of extreme sensitivity to stress. S.H. BU SH : There are programmes being carried out which should provide Dr. Marriott with the sort of information he requires. (1) A SM E X I has a programme for studying the effects of flaws in the beltline region, nozzles and
flanges when subjected to all design loads. This work will be published by the Electric Power Research Institute (EPRI). (2) The Pressure Vessel Research Committee (PV RC ) non-destructive examination test blocks 203, 155 and 204 have been evaluated by Buchanan using a two-point coincidence. A report on this should be available soon. (3) Acoustic emission during the hydro test may help to locate the flaws with which Dr. Marriott is concerned. (4) A SM E III is considering a requirement for both ultrasonic and radiographic testing in the construction code.
Session V ia
CODES, STANDARDS AND PRACTICES C h a ir m e n
R . L . R O C H E F r a n c e
K . B E C K E R Federal Republic of Germany IAEA-SM-218/29
CODES, STANDARDS AND PRACTICES AND THEIR INFLUENCE ON THE RELIABILITY OF NUCLEAR PLANTS
L.J. CHOCKIE* General Electric Company, Nuclear Energy Systems Division, S a n Jo se , California, United States of America
Presented by S.H. BUSH
A b stra c t
CODES, STANDARDS AND PRACTICES AND THEIR INFLUENCE ON THE RELIABILITY OF NUCLEAR PLANTS. Codes and Standards applicable to the construction and operation of pressurized equipment have been in use throughout the world for many years, yet only recently have studies and statistical analysis directed attention to the effect that Codes, Standards and current practices have on the reliability of the pressurized systems. The major effort in assessing the effect of reliability has been in the area of nuclear power plants, where the results of the studies and assessments show an enhancement in the reliability of the system, and also have led to revisions of the nuclear Codes and Standards to effect even greater improvement of the reliability of the nuclear power plant. Application of Codes and Standards provides a documented record stipulating the minimum set of requirements and provides data which can be used as input for statistical studies to compare the reliability of one system or plant to that of another system or plant. When the entire assessment is completed, it is readily concluded that the Codes provide an enhancement in the reliability of non-nuclear pressure-retaining components. The nuclear Codes provide for an order of magnitude increase in the reliability of nuclear components over that assessed for non nuclear uses. As an example, just one aspect of the Code, the timing of examinations and inspections during the lifetime of the nuclear plant is evaluated as providing a reliability enhancement of over 300 times when the programme is divided into uniform increments of time, compared to no programme at all. But an enhancement in reliability of over 8000 is credited when the programme is optimized as to the time increments during the life of the plant. The ASME Code, as well as the proposed IAEA Safety Guide on Plant Inspections, takes advantage of the optimized timing, thus providing for improved reliability.
* Chairman, Subcommittee on Nuclear Inservice Inspection (Section XI) of the ASME Boiler and Pressure Vessel Committee.
119 120 CHOCKIE
Introduction
The reliability of pressurized systems in a nuclear plant is a func tion of the quality of the as-built condition plus a program of continuing inspection during operation. Quantitative reliability figures cannot be based upon failure data for nuclear plants since such failures are so lim ited in numbers. Instead it is necessary to utilize information from comparable non-nuclear systems such as power boilers.
Much of the basis which provides the assurance of integrity of the nuclear reactor primary systems stems from the application of, and enforcement of, the Codes (Ref. 1). Codes, by definition, are a group of administrative and technical rules and standards covering combina tions of materials, designs, construction, installation, inspection and operation of equipment which is prepared for ready adoption into law or the legal jurisdiction. The Codes also provide a documented minimum set of requirements, providing input for statistical studies, and fault- tree analysis so that comparisons of the reliability of one system may be compared to the reliability of another system.
One of the most complete Codes for the construction and operation of power plants is the ASME Boiler and Pressure Vessel Code. This Code was established in the United States of America in response to appeals for the preparation of technically sound rules which would be acceptable on a national basis. Presently the ASME Code is referenced in the regulations of over seventy jurisdictional authorities. In North America, the States, the major cities in the States, and all of the Prov inces in Canada reference the Code in their statutes. In addition, a number of U. S. Federal Agencies, one of which is the U. S. Nuclear Reg ulatory Commission, have referenced the ASME Code in their respective regulations. Internationally, the Commission of European Communities study shows that the ASME Nuclear Code is used to some extent by all countries who have or are building nuclear power installations; either entirely or as a basis of judging other standards and practices.
Every aspect of plant construction and operation is considered and included in the Code affecting the primary pressure retaining ability of the nuclear system. Requirements are imposed on the administrative practices of the Operating Organization, the Manufacturer, the Fabrica tor, and the Materials Supplier in the form of a mandatory Quality Assurance program. Loadings must be identified which result in de tailed considerations for the design, with the Code imposing limits on the stresses that may be used for the materials selected. Controls are exercised by the Code on the fabrication, installation, testing, inspec tion and operation of the components and systems in the plant.
Detailed assessments have been made of the materials variabili ties, the results of fabricating, welding, and heat treating practices, and have demonstrated the capability of nondestructive examination meth ods to locate and measure the size of flaws. IAEA-SM-218/29 121
Failure Statistics and Analysis
Three recent studies (Ref. 2, 3 and 4) in the United States conclude that an improvement in failure rates for non-nuclear systems can be attributed to those pressurized components manufactured to recognized non-nuclear Codes and Standards. Quantifying the improvement is dif ficult, and inferred to be only one to two orders of magnitude. Use of the nuclear Codes (Ref. 1) for construction improves the reliability another order of magnitude, while the probabilistic studies infer that periodic inservice inspection by the Code for inservice inspection (Ref. 5) improves reliability one to two orders of magnitude, depending upon the type and completeness of the inspections.
Bush reports (Ref. 4) that the statistics of defects found during periodic inspection of boilers and pressure vessels have been compiled in United Kingdom and Federal Republic of Germany and failure rates and probabilities have been derived for those defects that are pertinent to pressure vessels. This portion of the studies brings into sharp focus the confusion resulting from imprecise definition and inconsistent use of terms such as "failure statistics" and "catastrophic failure". Many of the "failures" cited, because of their nature and locations, were probably generated during the fabrication stage and remained virtually unchanged until detected during a periodic inspection; they did not direct ly interfere with functional use. The significance of such defects and the possibility of their growth even to detectable but noncritical size is difficult to assess.
Bush's summary of 99 percent confidence upper bound values for probability of disruptive failure of non-nuclear vessels derived from each of the sources of data reviewed is shown in Table I.
The data sets for the various sources listed are arranged in order of increasing size of data base; that is, in order of increasing number of vessel-years of operation. In each of the first five cases the number of disruptive failures is zero. For zero failures, the statistical 99 percent confidence upper bound is determined by the number of vessel-years cov ered in the set; i. e ., it equals 4. 6 divided by the number of vessel-years. Accordingly, the upper bound value is seen to decrease in the table from 46 x 10_® for the EEI-TVA data (10 000 vessel-years) to 0. 63 x 10” ® for the ABMA data (725 000 vessel-years). The most significant and useful of these values is the ABMA number, that derived from the largest data base. It is important to recognize that, even if the true disruptive fail ure probability were much lower than the 0. 63 x 10"^ ABMA value, it would be impossible to establish that fact with a data base limited to 725 000 vessel-years. For example, if the next 725 000 vessel-years of operation accrued still involved no disruptive failure, the upper bound failure probability which would then have been demonstrated is not sig nificantly lower; viz, 0. 32 x 10“5 per vessel-year. In fact, even if one disruptive failure were to occur in the additional period, the upper bound value then substantiated (0.46 x lO- -*) would be lower than the value demonstrated by the data available to date with no failures. 122 CHOCKIE
T A B L E I
SUMMARY OF 99% CONFIDENCE UPPER BOUND VALUES FOR PROBABILITY OF DISRUPTIVE FAILURE OF PRESSURE VESSELS
Num ber of 99% Confidence Upper Data Source D isruptive Ves sei Years Bound, Failures Per F ailu res of Operation Vessel Year
US-EEI-TVA 0 10 000 46 x IO '5
US-EEI 0 22 000 21 x IO '5
FRG-Kellermann et al. 0 67 000 6. 9 x 10'5
UK-Phillips & W arw ick 0 100 300 4. 6 x 10'5
US-ABM A 0 725 000 0.63 x 10'5
FRG -Kelle rmann and Seipel 1 700 000 0.27 x IO'5 to 4 .0 x IO '5
Periodic Inservice Inspection
The value of any assessment of reliability of the as-built or as- fabricated quality will be quickly lost unless some program of continued inspection is pursued,with each inspection followed by corrective action for any deficiencies noted. The most comprehensive set of rules for periodic inspections has been written for nuclear plants, Section XI of the ASME Code. This Code also provides the basis for the IAEA Code and Guides for the operational phase of nuclear power plants. The fol lowing discussion presents the development of Section XI of the ASME Code and recent revisions which improve, even further, the assessed reliability for nuclear plants.
ASME Section XI
ASME Section XI, "Rules for Inservice Inspection of Nuclear Power Plant Components" has been developed by a committee of the American Society of Mechanical Engineers to provide a set of mandatory requirements for the examination, testing and inspection, analysis, repairs and replacements for components and systems in operating nuclear power plants. Responsibility for developing a program which IAEA-SM-218/29 123 meets the requirements is assigned to the Owner of the operating nuclear power plant. An Inspector, who represents the Regulatory or Jurisdic tional Authorities, and who may not be an employee of the Owner, is assigned the duty to verify that the responsibilities of the Owner and that the requirements of the Code are met. Results of examinations are compared to criteria to determine acceptability, and in the event that repairs or replacements are necessary to correct a condition, the Code includes the requirements for repair by welding as well as simple mechanical repairs and replacements.
The rules are separated into three divisions covering the inservice inspection and testing of components of light-water-cooled plants, gas- cooled plants and liquid-metal-cooled plants. The format and philosophy developed for the water-cooled systems, outlined below, is modified only as necessary to adapt to the other systems. These rules detail:
• Owner Responsibilities The Owner is assigned the responsibility to determine the appropriate classification for each component or system in the plant and to design an arrangement and clearance adequate to conduct the required examinations.
• Inspector's Duties Section XI rules for inservice inspection of nuclear power plants are mandatory and the Authorized Inspector is assigned the duty to verify, to assure, or to witness that the responsibilities of the Owner and the requirements of Section XI are met.
• Examination and Inspection Examination denotes the performance of all visual observa tions and nondestructive tests and is detailed to the type of examination, the area to be examined and the frequency for each component and specific portions thereto. Examination personnel may be employed by the Owner.
Inspection denotes verifying the act of performing the exam inations by an Inspector representing a state or municipality of the United States, Canadian Province, Authorized Inspec tion Agency or other enforcement authority having jurisdic tion over the nuclear power plant.
• Standard for Acceptance Indications of flaws detected during any inservice examina tion must be evaluated to determine acceptability by com parison to tabulated limits on flaw sizes or to the results of a detailed analysis. 124 CHOCKIE
о Pumps and Valve Testing The rules for testing pumps and valves represent a major addition to the original philosophy of inspecting the pressure boundary in that emphasis is placed upon operating char acteristics to improve assurance that the pumps and valves will perform as required when needed.
о Repair Procedures
A direct consequence of exceeding the permissible accep tance standards is the need for in situ repair procedures. To provide the maximum flexibility in making in situ repairs, five repair procedures were developed for Section XI.
о Hydrostatic Test Requirements
Hydrostatic tests or pressure tests are required at the end of each inspection interval and following any repair or replace m en t.
о Data Report A Data Report is included to provide the official documentation that the requirements in Section XI have been met.
Referring to the Owner's Responsibility, those components and sys tems which are classified by the Owner as being Class 1 are then to be given the inservice inspection program written for Class 1 components and systems in Article IWB "Requirements for Class 1 Components of Light-Water-Cooled Power Plants. " Class 2 components are covered in Article IWC, and Class 3 in Article IWD.
Requirements for Class 1 Systems
The requirements for the inspection of Class 1 components (Refs 6 and 7) and systems are the most comprehensive in terms of the extent of examinations to be performed, the amount of equipment to be exam ined, and the number of times during the lifetime of the plant that the examinations are to be performed. All Class 1 pressure-retaining com ponents and their supports are subject to inservice inspection unless specifically excluded from the requirements. Components which are specifically excluded from the volumetric and surface inspections in clude some components which are capable of being isolated from the Class 1 systems, or some components by virtue of their size (for example piping of diameter one inch or smaller is excluded), some pipe joints are excluded when the joints meet certain criteria for postu lated pipe breaks, and some reactor vessel head connections are ex cluded due to the inconsequential leak which would be present should the connection fail. In any case, however, all Class 1 components and their supports must be subjected to a visual examination; there are no exemp tions from the visual examination requirements. IAEA-SM-218/29 125
An important aspect of the inservice inspection requirements is the rule for a preservice inspection. One hundred percent of all pres sure retaining welds in Class 1 systems must have a preservice inspec tion meeting the rules of Section XI before the component or plant may be entered into service. The methods of the preservice examination are to be as close as possible to the methods expected or anticipated to be used for the later inspections. Shop inspections may be used when the methods and records meet the requirements of Section XI.
FIG.l. Typical weld seams in a BWR reactor pressure vessel. 126 CHOCKIE
VOLUME REQUIRED TO BE EXAMINED AREA TO BE SCANNED BY ANGIE BEAM 4 ^ BEAM NORMAL TO WELD FROM RIGHT SIDE
AREA TO BE SCANNED 4 Г BEAM FROM LEFT SIDE
VOLUME REQUIRED TO BE EXAMINED BY FOR LAMINATIONS. BOTH SIDES STRAIGHT BEAMFOR WELD AND HAZ ^ /
FIG.2. Extent o f ultrasonic scanning o f weld seam.
A sa result of the development of improved automated inspection devices to conduct inservice ultrasonic examination of reactor vessel components, increased attention has been focused upon the extent of vessel examination that could practically be accomplished during each inspection interval. The typical weld joints of a Boiling Water Reactor (BWR) reactor vessel are shown in Figure 1.
Experience gained from conducting preservice examinations of reactor vessels has demonstrated that the assembly and installation of the examination equipment rather than the conduct of the scanning process for ultrasonic examination consumes most of the time required to perform an inspection. The application of remotely operable inspec tion equipment, under properly planned procedures, has also reduced the personnel irradiation exposure below the level estimated during the earlier period when the inservice inspection rules were developed. These factors prompted a reassessment of the extent of examination required for reactor vessels (i. e. , Examination Category В-A in Section XI), and the Code increased the amount of examination.
A review of the ultrasonic examination procedures of the Code also revealed that the multiple scans required by angle beam and IAEA-SM-218/29 127
WELD VOLUME EXAMINED A-B-C-D SEAM DURING ULTRASONIC EXAMINATION D It cP 5EHZD
SURFACE SCANNED A-B u DURING ULTRASONIC EXAMINATION
FIG.3. An example of extent of volumetric examination in a typical PWR reactor vessel.
straight beam techniques to detect reflectors in the metal involved cov erage of a substantial proportion of the surface area of the vessel proper. Figure 2 illustrates the typical ultrasonic scans to which a weld of a reactor vessel is subjected as required by the rules of Article 1-5000 in the Code. Noteworthy is the wide band of vessel surface parallel to a weld seam that must be scanned for reflectors in order to volumetrically examine the metal within the boundary of the required volumes on both sides of the weld seam (e.g. , Figure IW B-3511. 1 (a) ).
In a typical reactor vessel intended for a Pressurized Water Reactor (PWR) system, shown in Figure 3, the preservice examination requires scanning of approximately 50 percent of the entire vessel sur face in order to examine the pressure retaining welds in the shell and 128 CHOCKIE heads. Similarly, the examination of the four inlet and four outlet recirculating reactor coolant nozzles entails scanning 75 percent of the surfaces within the nozzle-shell junction boundaries.
Alternative Inspection Programs
The initial inspection program specified in the earlier editions of the ASME, Section XI, was developed on the basis of repetitive inspec tions to be performed during each 10 years of plant service, (i. e. , the inspection interval). Specified percentages of the total examinations of components required during the inspection interval were to be completed in, at least, three inspection periods (approximately 3-1/3 years) cor responding to one-third of the inspection interval. Although inspections could be performed during any plant outages, current practices favor at least three scheduled examinations during an inspection period coincid ing with plant refueling outage . The total number of scheduled exam inations that may be expected over a plant service lifetime of 40 years is 36. This inspection program of equal-spaced periods of examination is now identified as Inspection Program "B". The basis upon which this program was developed included recognition of the factors that could contribute to structural degradation. These factors were considered to be directly related to service periods and as such periodic examinations would be meaningful if conducted at equal-spaced intervals.
The increased attention (Ref. 2, 3 and 4) on the part of the United States regulatory authority in assessing the failure probabilities of reac tor vessels built in accordance with the rules of the ASME Code, Section HI (Ref. 1), and inspected periodically in accordance with the rules of ASME Code, Section XI (Ref. 5), prompted a reassessment of inspection p ro g ra m s.
An analysis of the influence of periodic inspection of nuclear reac tor vessels is contained in Appendix В of the WASH-1318 (Ref. 4) report issued by the regulatory staff of the U.S. Atomic Energy Commission. A statistical methodology was adopted to develop a model for vessel failure in order to evaluate the influence of various inspection programs in terms of enhanced reliability factor. This factor, R, is defined as:
R P /Р о i w here P failure probability of a component о without the benefit of inservice examinations
P. failure probability of the component l with the benefit of inservice exam inations
The inservice examinations are assumed to be followed by corrective actions, if necessary, to restore the component to an acceptable level of structural integrity for continued service. IAEA-SM-218/29 129
T A B L E II
ENHANCED RELIABILITY FACTORS
Inspection Program Inspection Program Inspection S e rv ice A S erv ice В r In terval Y e a rs Reliability Factors Y e a rs Reliability Factors
1 st 3 20 10 9 2nd 10 400 20 42 3rd 23 8100 30 134
This statistical methodology has been extended in a more general study (Ref. 8) to assess the merits of inspection programs with uniform and non-uniform inspection intervals. The method which is used to optimize the benefits of a recurrent inspection, demonstrates that the enhanced reliability factor is strongly influenced by the times at which the inservice examinations are performed. This statistical approach formed the basis for the development of Inspection Program "A" as an alternative program to Inspection Program "B", of ASME Section XI (Ref. 9). The optimization study revealed the marked advantages of performing inservice inspections early in the service lifetime rather than at equal-spaced inspection intervals. To illustrate the advantages, the example shown in Table II provides a measure of the benefit that may be gained by the application of Inspection Program "A" over that of Inspection Program "B", when 100 percent inspections are optimized.
Figures 4 and 5 illustrate graphically the spacing of the examina tions in time, as well as the percentage of required examination per formed during each plant outage for Inspection Program nA" and "B", respectively. Although the number of inspections to be performed dur ing the service lifetime are equal under both programs, 50 percent of the total number of inspections under Inspection Program "A" would be completed by the end of the 10th year of plant service, whereas, the same percentage would not be completed before the 20th year of plant service for Inspection Program "B".
A distinct advantage of the Inspection Program "A" would accrue from the earlier examinations of components subject to irradiation at times when the irradiation levels are low and when corrective repairs, if needed, could be more readily accomplished. Analyses of inservice inspections of nuclear power plant components performed to date, al though limited in number, appear to indicate that structural deteriora tion is more likely to develop in the earlier years of plant service. This characteristic is also confirmed by operating experience in non-nuclear sy stem s (Ref. 10). Such se rv ice -in d u ce d c h a r a c te r is tic s obviously favor 130 CHOCKIE
INSPECTION INTERVALS
FIG.4. Inspection program-B examinations per plant outages.
FIG.5. Inspection program-A examinations per plant outages. the application of Inspection Program "A", for those components whose early examinations reveal indications which, upon evaluation, qualify the component as conditionally acceptable for continued service.
SUMMARY
Application of Codes and Standards and Practices provides a record which can be evaluated to determine the improvement in relia bility of nuclear power plants for both the constructional phase as well IAEA-SM-218/29 131 as the operational phase. The results of studies and evaluation indi cate, that the nuclear codes provide for an improvement of one or two, or more, orders of magnitude over that assessed for non-nuclear uses.
The studies and assessments, resulting from data provided by the application of the Codes, also provided the basis and foundation for effecting revisions to the Codes themselves; revisions which, when implemented, provide for even greater improvement in the reliability of nuclear plants.
REFERENCES
(1) ASME Boiler and Pressure Vessel Code - Section III - "Nuclear Power Plant Components". Published by American Society of Mechanical Engineers, New York, New York 10017.
(2) WASH-1285, Report by the Advisory Committee on Reactor Safeguards on "Integrity of Reactor Vessels for Light-Water Power Reactors, " January 1974.
(3) WASH-1400, Reactor Safety Study - An Assessment of Accident Risks in U. S. Commercial Nuclear Power Plants, U.S. Atomic Energy Commission, October 1975.
(4) WASH-1318, Technical Report on "Analysis of Pressure Vessel Statistics from Fossil-Fueled Power Plant Service, and Assess ment of Reactor Vessel Reliability in Nuclear Power Plant Serv ice", May 1974, U.S. Atomic Energy Commission.
(5) ASME Boiler and Pressure Vessel Code - Section XI "Rules for Inservice Inspection of Nuclear Power Plant Components" - 1974 Edition, inclusive of all published Addenda to Winter 1975.
(6) BUSH S. H. and MACCARY R. R., "Development of Inservice Inspection Safety Philosophy for U.S.A . Nuclear Power Plants". I. Mech. E ., Periodic Inspection of Pressure Vessels, 1972.
(7) JOHNSON W. P. , BUSH S. H. and MACCARY R. R., "Augmented Scope of the 1974 ASME Section XI Code Inservice Inspection of Nuclear Power Plant Components". I. Mech. E ., Periodic Inspection of P ressu re V essels, 1974. 8910
(8) ARNETT L. M., "Optimization of Inservice Inspection of Pressure Vessels", presented at Conference on Nondestructive Testing in Nuclear Industry, ASM, Denver, CO., December 1975.
(9) CHOCKIE L. J. , BUSH S. H. , and MACCARY R. R. , "Extended Rules of the 1974 ASME Section XI Code 'Inservice Inspection of Nuclear Power Plant Components' ", I. Mech. E ., Third Con ference on Periodic Inspection of Pressurized Components, London, 1976. (10) BUSH S. H. , "Pressure Vessel Reliability, " Journal of Pressure Vessel Technology, Vol. 97, Series J, No. 1, February, 1975. 132 CHOCKIE
DISCUSSION
T.A. SM ITH: Fig. 3 shows the extent of volumetric examination for a typical PW R vessel. I notice that it does not include the nozzle and ligament region of the top dome. Is this because of problems of access or is it because you do not consider it necessary to examine this section of the vessel? S.H. BUSH: Examination of the ligament region is not excluded. It must be done on a best effort basis, in recognition of the fact that such a region poses substantial difficulties for ultrasonic testing (UT). Surface examination is used to complement the UT. There has been some success with UT but not necessarily with a 60° shear wave. G. PACKM AN: In the paper the emphasis appears to be placed on the improvement to be obtained by applying a particular code. Does the committee responsible for that (or any other corresponding) code have a specific safety target in mind? If so, is the target defined in two parts: (i) a minimum, which must be achieved; and (ii) a higher safety standard, to be considered on a cost- benefit basis? S.H. BUSH : Historically, an acceptable risk of failure in a reactor pressure vessel (RPV ) was considered to be less than one major failure in ten million years. A report of the Advisory Committee on Reactor Safeguards concluded that RPVs were sufficiently reliable, subject to the proviso that appropriate testing and examination would be continued so as to assure this level of reliability. To provide such assurance was the initial purpose of the A SM E X I Code. Although the charter of A SM E has been extended, its main purpose is still R P V reliability. Cost-benefit aspects continue to be examined but they do not play a decisive role. IAEA-SM-218/30
WORLD-WIDE STANDARDIZATION OF POWER REACTOR SAFETY AND TECHNOLOGY BY ISO
K . B E C K E R D IN /N Ke German Nuclear Standards Committee, B e r lin , Federal Republic of Germany
A b stra c t
WORLD-WIDE STANDARDIZATION OF POWER REACTOR SAFETY AND TECHNOLOGY BY ISO. The Nuclear Standards Committee (NKe) of the DIN German Standards Institute serves as the national counterpart of the International Organization for Standardization (ISO) Technical Committee (TC) 85 on Nuclear Energy. In particular, it acts as ISO secretariat in several subcommittees (SC) of TC 85, including SC 5,“Nuclear Fuel Technology”, and SC 3/WG 6,“Pressure Boundaries of Primary Circuits” . After an introduction describing briefly the scope of ISO in general, past, current, and future activities in the reactor safety and technology field are discussed. Working Group 6 consists currently of the four sub groups Materials, Design, Manufacture, and Inspection and Testing. It developed a working document, N 130, “Rules for steel reactor pressure vessels”, which is by now somewhat out dated, and an US proposal based on the ASME-III Code. In recent meetings of WG 6 and SC 3, it was decided to develop a revised version, taking into consideration a new German Code which is currently being developed by consensus of the five parties involved in the licensing process (licensing authorities; manufacturers and suppliers; utilities; the inspection organization, TUV; and other interested institutions including DIN). A five-year-plan calls for completion of the draft proposals before the end of 1978. They are then to be circulated, revised and to be submitted to the ISO Central Secretariat in Geneva in 1979. After approval by the ISO member bodies and Council they are expected to become final ISO standards in 1981. Some more general considerations on the role of international standards, in particular in technology transfer to third-world countries, are also discussed briefly.
1. STRUCTURE AND SCOPE OF ISO
It m ay be appropriate to start with a brief description of the International Organization for Standardization (ISO ) and its activities in the important field of power reactor safety. After a meeting in London in 1946, delegates from 25 countries pursued an initiative by the newly formed United Nations to create a new international
133 134 BECKER organization “whose object shall be to facilitate the international co-ordination and unification of industrial standards”. The new organization, ISO, began to function officially thirty years ago. A t the same time, the already existing International Electrotechnical Com m ission (IEC), which covers electrotechnical questions, was affiliated to ISO and, while preserving its autonom y, has served since then as the Electrical Division of ISO . ISO is now universally recognized as the specialized international agency for standardization, comprising currently the national standards bodies of about 85 countries. The work of ISO is aimed at world-wide agreement on international standards with a view to the expansion of trade, the improvement of quality, the increase of productivity and the lowering of costs. A member body of ISO is the national body “most representative of standardization in its country”. Mem ber bodies are entitled to participate and exercise full voting rights, are eligible for Council membership and have a seat in the General A s s e m b ly . M ost member bodies are either (as in m any planned-economy and developing countries) governmental institutions, or organizations incorporated by public law, as is the case with m any of the older standards organizations in free-economy, highly industrialized countries such as the U S A and the Federal Republic of Germany, which both play a major role in international nuclear standardization. The remainder have in all cases close links with the public administration in their respective countries. There are also correspondent members, which are norm ally organizations in developing countries which do not yet have their own national standards body. IS O ’s work covers virtually every area of technology. It brings together the interests of producers, users (including consumers), governments, and the scientific com m unity in the preparation of international standards. The technical work of ISO is carried out by more than 1600 technical bodies (for details on ISO work, see Ref.[ 1 ]). It is decentralized, and consequently the Technical Secretariats (TC ) are located in 29 countries and are served by some 400 people, while the ISO Central Secretariat in Geneva has a staff of about 100 persons drawn from 16 countries. At present about 3000 International Standards are available, and more than 1000 Draft International Standards are registered at the ISO Central Secretariat, representing some 35 000 pages of concise reference data and technical know-how. Several regional organizations are active within ISO, for example A SM O for the Arabic countries and CO PA N T for Latin America. The decision to set up a technical committee is taken by the ISO Council, which also determines the scope of the committee. A technical committee usually creates subcommittees (SC) and working groups (W G) to cover different aspects of the work. A working group is composed of individuals and not of national delegations. More than 100 000 voluntary experts from all over the world are currently engaged in ISO activities, and both the number and importance of the IAEA-SM-218/30 135 resulting drafts and standards are expected to grow rapidly for several reasons, in c lu d in g
the increasingly supranational nature of m any modern technologies such as nucleonics, aerospace, electronics and transportation
the desirable technology transfer to developing countries which is possible by means of international standards.
The latter aspect was stressed, for example, during the recent 4th Meeting of the United Nations Conference on Trade and Development (U N CT A D ) in Resolution 87 IV and at a special conference on the role of standardization in the developing world in Algiers in 1976 which was sponsored by ISO and the U N Industrial Development Organization (U N IDO ). The following resolution
was adopted:
“The role of standardization in the transfer of technology should be brought to the attention of governments at the highest level by national standards bodies. ISO should take steps to ensure that this role is fully taken into account by the various organizations of the United Nations wherever technology is discussed. In particular:
standardization should be included in the topics for consideration at the U N Conference on Science and Technology for Development, scheduled fo r 1 9 7 9 .
U N ID O , through its Industrial Operations Division, should become more actively involved in the work of ISO and IE C at the technical level, in order to ensure that this work is fully taken into account in technical assistance programmes, and also to obtain some feedback on needs of developing countries.
ISO, in collaboration with the IEC, should produce guidance material on the widespread influence of standardization on technical assistance pro grammes, with a view to ensuring that standardization considerations are fully present when such programmes are being planned” [2].
The subject of the role of nuclear standards in nuclear technology transfer has recently been discussed in more detail elsewhere [3].
2. ISO/TC 85, “NUCLEAR ENERGY”
Within ISO, Technical Committee TC 85, “Nuclear Energy”, is responsible for world-wide standardization in the nuclear field. The decision to form TC 85 was made by ISO in 1956, and a first plenary session took place almost exactly twenty years ago with 64 delegates representing 17 countries. The secretariat 136 BECKER has during the last few years been with the American National Standards Institute AN SI, but will shortly be transferred, probably to the D IN German Nuclear Standards Committee. TC 85 is divided into the following five sub-committees:
(1) ISO /ТС 85/SC 1: Terminology, Definitions, Units and Symbols (Secretariat: AN SI, USA).
(2) ISO /ТС 85/SC 2: Radiation Protection (Secretariat: AFN O R, France); Scope: standardization in the field of protection against ionizing radiation and radioactive substances, m ainly in three areas: criteria for the safe design, testing, and use of devices and facilities incorporating sources of ionizing radiation; procedures and equipment for protection against external irradia tion and radioactive contamination; instruments and methods for evaluation of external irradiation and radioactive contamination.
(3) ISO /ТС 85/SC 4: Radioactive Sources (Secretariat: Institute of Nuclear Research, Poland); Scope: standardization in the field of performance, quality assurance, containment, and packaging of radioactive sources.
(4) ISO /ТС 85/SC 5: Nuclear Fuel Technology (Secretariat: DIN, Federal Republic of Germany); Scope: standardization in the field of nuclear fuel manufacture, transportation, storage, fuel reprocessing, and radio active waste management, incorporating such aspects as criticality preven tion, fissile material control, physical protection and quality assurance.
In the context of this symposium the most important one is undoubtedly
(5) ISO /ТС 85/SC 3: Power Reactor Technology (Secretariat: SIS, Sweden); Scope: standardization in the field of power reactor technology covering all functional, mechanical, and structural aspects of a nuclear-power generating facility during siting, design and construction, pre-operative and start-up testing, operation, and in-service inspection, testing and maintenance. SC 3 has currently 13 participating and 15 observing member countries. In Table I, the currently active W orking Groups of SC 3 are listed.
3. ISO/ТС 85/SC 3/WG 6, “PRESSURE BOUNDARIES OF PRIMARY CIRCUITS”
Especially relevant for this sym posium is W G 6. Its secretariat is with the Nuclear Standards Committee (Normenausschuss Kerntechnik, NKe) of the Federal German National Standards Institute, DIN . The organizational structure IAEA-SM-218/30 137
TABLE I. ISO/ТС 85/SC 3 “POWER REACTOR TECHNOLOGY”
WG 2 “Metereological Aspects of Reactor Safety” (Secretariat: UNICEN, Italy)
WG 3 “Containment Structures” (Secretariat: ANSI, USA)
WG 6 “Pressure Boundaries of Primary Circuits” (Secretariat: DIN, Federal Republic of Germany)
WG 7 “Seismic and other Hazards Consideration” (Secretariat: AFNOR, France)
WG 8 “Quality Assurance in the Design, Manufacture of Equipment, Construction, Operation and Maintenance of Nuclear Power Plants” (Secretariat: BSI, UK)
WG 9 “Reliability Data” (Secretariat: UNICEN, Italy)
WG 10 “Concrete Reactor Vessels” (Secretariat: BSI, UK)
FIG.l. Organizational Structure of the DIN/NKe German Nuclear Standards Committee. 138 BECKER of this German nuclear standards body (Fig.l) does, incidentally, closely reflect ISO /T C 85 and this body is, therefore, particularly suitable as a national counter part for ISO. It has presently about 400 voluntary experts in forty working groups and steering committees. Alm ost one hundred nuclear D IN standards and draft standards have been developed so far — more than thirty in 1976 alone (for details, see Ref. [4] or contact the N K e Secretariat). W G 6 is divided into four sub-groups: Materials (Secretariat: Sweden), Design (France), Manufacture (Germany), and Inspection and Testing (Italy). A new co-ordination subgroup to consider the relevant more general aspects such as quality assurance and the role of “third-party” inspection authorities is being formed, with Switzerland likely to serve as secretariat. W G 6 has met seven times, most recently in Berlin in March 1977, to review progress and prepare a status report for the subsequent eighth meeting of SC 3 in Stockholm . The most important result of its work so far has been the pre paration of two rather substantial documents. The first one, Draft N 130, entitled “Rules for Steel Reactor Pressure Vessels”, has 145 pages. As it was submitted in 1972 and is based on earlier drafts, it is by now somewhat outdated. Moreover, being a “common-denominator” type of document, it is not specific enough to be used directly. In an attempt to solve this problem, SC 3 requested the preparation of another draft working document. This document, entitled “Safety Standard for Construction of Metal Primary Coolant Circuit Pressure Retaining Com ponents”, has 347 pages, consists of the U S A SM E -III Code, and was submitted in 1975. During the subsequent discussions of experts and secretariats, the feeling emerged that this draft was not sufficiently “denationalized” to serve as an adequate new basis for an ISO standard, but that another serious attempt to harmonize the US and the European proposals is required. This harmonization is, however, somewhat difficult because of delays in the development of generally approved reactor safety codes in the Federal Republic of Germany, which is one of the world’s largest suppliers of nuclear facilities. This is m ostly due to the fact that a special body closely associated with the Federal Ministry of Interior was set up in Germany to develop and approve all basic reactor safety standards. This body, called Kerntechnischer Ausschuss,KTA (freely translated as “Nuclear Technology Board”), consists of the five parties involved in the licensing process, namely the licensing authorities; manufacturers and suppliers; utilities; the independent “third- party inspection” organization T Ü V (Technischer tfberwachungsverein), and other interested institutions and authorities, including D IN . A high degree of consensus between all parties is required for the approval of a KTA-standard. So far, a total of 12 KTA-standards, among them 5 prepared by D IN /N Ke, have had final approval. The federal structure, and the unique role of independent experts in the German licensing process, have been complicating IAE A-SM-218/30 139 factors in developing basic nuclear safety standards. But there are also other problems: members of the W G 6 subgroups are currently proceeding to prepare standards which are specific enough to be used directly, for example by a develop ing country im porting nuclear technology from different supplier countries, as opposed to general guidelines to be used only as a framework for detailed national standards. This creates some problems with regard to the time and effort required from the relatively few highly qualified experts available, and — more im portantly concerning the updating and interpretation of the resulting standards. ISO has not yet established mechanisms for doing this which are as sm ooth and fast as desirable for such highly com plex standards. The present schedule can be summarized as follows: All subgroups have prepared draft documents, which are currently being circulated for comments and will lead to revised drafts in the near future. A “synthesis” report from the available up-dated U S and emerging European standards will also have to be developed, but manpower shortage m ay delay this process. A five-year-plan was adopted at the latest plenary meeting of SC 3 in Stockholm in March 1977, calling for completion of all draft proposals before the end of 1978. These are then to be circulated, revised and to be submitted to the ISO Central Secretariat in Geneva late in 1979. After approval by the ISO member bodies, the standards have to be submitted to the ISO Council and are expected to become final ISO standards late in 1981 or (in the case of the Inspection and Testing part) before the middle of 1982. The original scope of W G 6, namely light water power reactors, has also recently been expanded to encompass at least parts of the high-temperature reactor field. It has been requested by SC 3 “to include pressure circuits with materials in the creep range as soon as this becomes possible”.
4. CONCLUSIONS AND OUTLOOK
Only fifteen ISO nuclear standards have already been published in final form or exist as a draft, and m any others are in different stages of preparation (Table II). However, compared with the actual need for international nuclear standards, the progress which has been made is not overly impressive. To give only a few numbers: According to a recent catalogue and classification of the world’s nuclear standards, rules, codes and regulations, prepared originally in 1975 for the European Com m unity [5] and which has recently been carefully updated and revised [6], there were at the end of 1976 about 2700 nuclear standards in the world, prepared by 145 organizations and institutions in 31 countries.
Clearly, this implies a tremendous redundancy in efforts, which could be much reduced by carefully co-ordinated international activities. Moreover, 140 BECKER
TABLE II. NUCLEAR STANDARDS ESTABLISHED BY ISO/TC 85 “NUCLEAR ENERGY”
Note: All items are available from National Standards Institutes
ISO/361 —1975 Basic ionizing radiation symbol
ISO/921 —1972 Nuclear energy glossary
ISO/1677—1977 Sealed radioactive sources — general
ISO/1709—1975 Principles of criticality safety in handling and processing fissile materials
ISO/R 1710-1970 Fundamental principles for protection in the design and construction of installations for work on unsealed radioactive materials
ISO/R 1757-1971 Personal photographic dosimeters
ISO/1758—1976 Direct reading pocket type electroscope exposure meters
ISO/1759—1976 Indirect reading capacitor type pocket exposure meters and accessory electrometers
ISO/2855 —1976 Radioactive materials — Packagings — Tests for contents leakage and radiation leakage
ISO/2889—1975 General principles for sampling airborne radioactive materials
Draft Standards:
DAD 921.1 Add.l to ISO/921.1972
DAD 1757.1 Add.l to ISO/R 1757
DIS 2919 Sealed radioactive sources — classification
DIS 3925 Unsealed radioactive sources — identification
DIS 3999 Specifications for apparatus for gamma radiography
substantial technology transfer to developing countries could be provided — not so much with the more general common-denominator type of standards, but with those which include all necessary details of nuclear safety and technology. But why has progress in ISO work been so slow? One of the key problems is the limited availability of the experts required for this type of work. W ith the IAEA-SM-218/30 141
ISSUED BY
GOVERN MANDA- MENT *■ TORY
NATIONAL STANDARDS VOLUN INSTITUTE TARY
FIG.2. Hierarchy of nuclear standards.
costs o f translations, travel, secretariat personnel, etc., also increasing rapidly, higher priority is frequently given to pressing problems of national nuclear standardization. Another difficulty lies in the potential overlap of ISO activities with those of other international organizations. Some people are afraid that very similar standardization work is done, for example by ISO and the IA EA , which latter has been very active recently in the development of Safety Guides [7]. However, both on the national and the international level a hierarchy of nuclear standards exists which can be illustrated by a schematic pyram id (Fig.2). It is representative for m any industrialized countries such as the U S A and the Federal Republic of Germany. An atomic energy act and radiation protection regulations apply the generally recognized basic standards such as IC R P recommendations to the specific conditions of the country. They are issued by the government, are mandatory, and may be supplemented by some kind of regulatory guides also issued by the government or its associated or advisory organizations. The next level consists of the basic safety standards to be applied prim arily in the licensing process, which are set up by all interested parties. These standards also require some kind of recognition or approval by the governmental authorities, are mandatory and may be issued by the government or by the National Standards Institute after approval by the authorities. The base of the pyram id consists of nuclear and conventional standards which need no official approval and are not mandatory. They include, for example, standards on terminology, dimensions and systems of components, on codes of good practice in the handling of radio active materials, etc. 142 BECKER
Sim ilarly, one can distinguish on the international level between voluntary standards developed and promoted by ISO and, to a lesser degree, IE C and the European Standards Organization CEN , and on the other hand standardization by international organizations directly sponsored by the governments of the participating countries. Activities of the IA E A and regional organizations such as the European Com m unity (EC) are typical examples of the latter. Clearly, the main goal of IA E A and EC efforts is to provide a basis for national atomic energy acts and regulations. (Good examples are the IA E A “Regulations for the Safe Transport of Radioactive Materials” and the EU R A T O M “Basic Safety Standards for the Health Protection of General Public and Workers against Dangers of Ionizing Radiations”). On the other hand, a typical ISO goal would be to provide detailed instruction on, for example, optimized techniques for ultrasonic testing of reactor components. In conclusion, there is obviously a need and room for both basic and applied international nuclear standards. It should be the goal of those associated with this w ork to strengthen their efforts, optimize procedures, minimize delays and to motivate more experts to create a universally recognized system of nuclear safety standards as an im portant contribution to a safe and sound future for nuclear energy.
REFERENCES
[1] ISO Memento (1976). Available from ISO, C.P. 56, CH-1211 Geneve 20. [2] ISO Bulletin January 1977, p.4. Available from ISO, C.P.56, CH-1211 Geneve 20. [3] NEIDER, R., BECKER, K., in Proc. Iran Conf. on Transfer of Nuclear Technology, Persepolis/Shiraz (1977), in press. [4] BECKER, K., in: “Regeln und Richtlinien in der Kerntechnik”, Deutsches Atomforum 1977, p.128 (English version to be published in Nuclear Safety). [5] BLOSER, M., FICHTNER, N., NEIDER, R., EUR-5362e (1975). Available from EG, B.P.1003, Luxembourg. [6] FICHTNER, N., BECKER, K., BASHIR, M., Catalogue and Classification of Technical Safety Standards, Rules and Regulations for Nuclear Power Plants and Fuel Technology, in press. Available from EG, B.P. 1003, Luxembourg. [7] IANSITI, E., KONSTANTINOV, L., Nuclear Power and its Fuel Cycle (Proc. Int. Conf. Salzburg, 1977) 5, IAEA, Vienna (1977) 51.
DISCUSSION
R.W. N IC H O LS: What is the overall position and the timetable as regards an ISO standard for pressure vessels and what would be the relationship between any ISO standard for the nuclear reactor pressure boundary and an ISO pressure vessel standard? IAE A-SM-218/30 143
К . B E C K E R (Chairman)'. Although the work of ISO /ТС 85/SC 3/WG 6 has not yet led to an ISO standard or draft standard on the nuclear reactor pressure boundary being drawn up, two drafts have been prepared which need to be updated or harmonized in a new document within the next year or so. Nuclear standards for this field are based on the existing ISO standards for conventional boilers (ISO 831) and on the Draft International Standard on Pressure Vessels (ISO 2694) which has been developed by ISO /TC 11.
IAEA-SM-218/31
THE ROLE OF THE AMERICAN SOCIETY FOR TESTING AND MATERIALS (ASTM) IN PROVIDING STANDARDS TO SUPPORT RELIABILITY TECHNOLOGY FOR NUCLEAR POWER PLANTS
L.E. STEELE* Naval Research Laboratory, Washington, D.C., United States of America
A b stra c t
THE ROLE OF THE AMERICAN SOCIETY FOR TESTING AND MATERIALS (ASTM) IN PROVIDING STANDARDS TO SUPPORT RELIABILITY TECHNOLOGY FOR NUCLEAR POWER PLANTS. ASTM is an international society for managing the development of standards on characteristics and performance of materials, products, systems and services and the promotion of related knowledge. This paper provides on overview of ASTM, emphasizing its contribution to nuclear systems reliability. In so doing, the author, from his perspective as chairman of ASTM committee E 10 on “Nuclear Applications and the Measurement of Radiation Effects and the Committee on Standards”, illustrates ASTM contributions to the understanding and control of radiation embrittlement of light-water reactor pressure vessels. Four major related tasks are summarized and pertinent standards identified. These include: (1) surveillance practice (5 standards), (2) neutron dosimetry (8 standards), (3) specification for steels for nuclear service (7 standards) and (4) basic guidelines for thermal annealing to correct radiation embrittlement (1 standard). This illustration, a specific accomplishment using ASTM standards, is cited within the context of the broader nuclear-related activities of ASTM.
1. INTRODUCTION
The American Society for Testing and Materials (ASTM), as the full name implies, grew up as a standards organization con cerned with testing methods and products or materials. In recent years that scope has been broadened to include as well definitions, classifications, and procedures or prac tices. Within that framework, there are specialties which ad dress specific technologies. One of those is nuclear technology. In such new areas, another long-term activity of ASTM has pro vided the vehicle for initial activities and standards prepara tion, that is, the research information gathering and dissemina tion role which is considered complimentary to standards pre paration role. As the industry grew, an evolutionary development
*Chairman, ASTM Committee E10 on Nuclear Applications and the Measurement of Radiation Effects and the Committee on Standards.
145 1 4 6 STEELE in nuclear standards occurred. With several committees perform ing actively the task of information exchange, standards pre paration for the nuclear industry became a real, though initially secondary, activity within ASTM.
This evolution has paralleled the projected rather than actual nuclear systems, being founded basically by research- oriented people. That is, the efforts have looked ahead of actual application and consequently have made a significant contribution to the nuclear industry through research informa tion exchange as well as by the preparation of standards.
This paper is intended to review the role of ASTM in this process through citation of specific examples of standards developed to support the nuclear industry, particularly those having a bearing on the realiability of reactor power plants, and specifically the primary pressure boundary. In addition, the nature of the ASTM organization and the activities of a key nuclear committee, E10, are outlined.
2. ASTM
In words provided by ASTM Ll], "ASTM is a management system for the development of voluntary full consensus standards. It provides a legal, administrative, and publications forum within which producers, users, and those representing the general inter est can meet on a common ground to write standards which will best meet the needs of all concerned."
The ASTM scope is defined Cl] as follows: The American Society for Testing and Materials is a non-profit corporation formed for the development of standards on characteristics and performance of materials, products, systems, and services, and the promotion of related knowledge.
By ASTM definition [1] standardization is, "the process of formulating and applying rules for an orderly approach to a specific activity for the benefit and with the cooperation of all concerned." Within this framework, ASTM promulgates the following type of standards [1].
a. Standard definitions which create a common language for a given area of knowledge.
b. Standard recommended practices which suggest accepted procedures for performing a given task.
c. Standard methods of test which prescribe ways of making a given measurement.
d. Standard classifications which set up categories in which objects or concepts may be grouped.
e. Standard specifications which define boundaries or limits on the characteristics of a material, product, system, or service.
The writing of standards is done by 26,000 knowledgeable volunteer members of 130 different committees. For committees IAEA-SM-218/31 147 writing standards applying to commercial products a balance of producers and users members is required.
Financial support comes mainly from the sale of the book of standards (about 75%) and membership fees (about 25%).
As noted in the scope, "promotion of related knowledge," is considered a high goal along with developing standards. To this end, ASTM provides several publications in addition to the book of standards which now comprises 48 volumes totalling 39,000 pages. Among the knowledge promoting or information dis seminating publications are: 1) periodicals, "ASTM Standardiza tion News," and, "ASTM Journal of Testing and Evaluation," 2) special technical publications (many comprising proceedings of technical conferences), and 3) data series or miscellaneous publications. In addition, there are two key operational publica tions, "Rules Governing the Operation of Technical Committees," and, "Form and Style for ASTM Standards."
ASTM Standards and the Standards writing committees are organized as follows: () is number of committees.
A. Ferrous Metals (5)
B. Nonferrous Metals (9)
C. Cementitious, Ceramic, Concrete and Masonry Materials (2 0 )
D. Miscellaneous Materials (Petroleum based-wood, paint rubber, textiles, etc.) (31)
E. Miscellaneous Subjects (testing and analysis methods, measurements, criteria, nuclear, etc.) (39)
F. Materials for Specific Applications (largely end pro ducts for industry or personal use, safety related, etc.) (20)
G. Corrosion, Deterioration and Degradation of Materials (4.)
The ASTM is thus basically a collection of volunteers with the goal of preparing consensus standards or procedures to help systematize specific activities. In many ways it resembles a technical or professional society but differs in the diversity of specialists and the uniqueness of its primary product - s t a n d a r d s .
3. DEVELOPMENT OF STANDARDS TO SUPPORT NUCLEAR POWER
The development of nuclear power to its current status in the United States has depended largely upon two things, a grow ing understanding of the physical phenomenon of nuclear fission and the amplification of various existing or new technologies to permit the capture of the energy, at least a part of it, from the fission process. This has involved a large number of exist ing and newly developing technologies, those relating especially to heat transfer, fluid flow and electricity generation which were in existence but had to be tailored to the peculiar condi tions of the light water reactor, the primary nuclear system in 148 STEELE the United States. Besides that, a number of technologies which are new or relatively new including environmental science and technology, health physics, nuclear fuel preparation and radio active waste handling as well as the solution to some peculiar heat transfer problems within the nuclear power plant have resulted in the product we have today, namely sixty boiling water and pressurized water reactors.
In every facet of ASTM activities, there are standards which impact the safe and reliable application of nuclear power. Nevertheless, there are those which impact more directly. Those committees which have the greatest role or the most direct role in nuclear power development are, Committee E-10 on nuclear applications, and Committee C-26 on the fuels, moderator and control materials, as well as secondary contribution from A-l on steels, B-10 on reactive and refractive metals, C-5 on graph ite, D-19 on water, E-23 on sampling the environment, G-l on corrosion, and a large number which are important contributors whose efforts are less focused on problems affecting peculiarly the nuclear industry. Among these would be those concerned with methods such as for testing metals for fatigue, E-9, mea suring or detecting flaws in metals, E-7, developing methods for testing the mechanical properties of metals, E-24 and E-28, as well as a large number that have to do with standards which support all advanced technology and hence nuclear power as well.
For the purposes of this paper, it is impossible to identify all the pertinent activities of ASTM which contribute directly to the nuclear power safety and reliability, but it is important to identify Part 45 of the Book of ASTM Standards [2] which in corporates approximately 121 standards considered to be directly nuclear related. The specific subject matter of Part 45 stand ards are divided by subject essentially as follows: [() is number of standards] .
"A" Committees - steels for nuclear application (7).
"B" Committees - other metals and metal products for nuclear applications (hafnium, nickel, tantalum, titanium, zircon ium) (19).
"C" Committees - nuclear grade materials and analytical methods (graphite, concrete; uranium and plutonium; aluminum, boron, silver-indium-cadmium, [ and compounds of these metals]) (34).
"D" Committees - primarily tests for water purity, radio activity, specific nuclear related impurities (30).
"E" Committees - fuel burnup and radioactive component analysis, neutron dosimetry, operating reactor sur veillance for radiation effects (40).
"G" Committees - corrosion of reactor materials (1).
A list of about 700 nuclear related ASTM standards generated by committees Al, B2, BIO, C5, D9, Dll, D14, D19, E2, and G1 are also listed for reference purposes. This one volume, though im perfect, in citing all nuclear related standards, comes close to identifying the contribution of ASTM toward the nuclear field and it is important to note that most of these largely relate IAEA-SM-218/31 149 to product quality or reliability and to public safety. Further, it should be noted that these are constantly being revised and updated for better applicability utilizing the latest data from an advancing field of science and technology.
Further subdividing for purposes of illustration, the activ ities of Committee E-10 are highlighted. In this way it is hoped that ASTM's contributions to standards to support nuclear system reliability can be illustrated. A similar illustrative story might be written about several other ASTM committees.
3.1 ASTM Committee E10 on Nuclear Applications
The scope of committee E10 is defined as follows: Scope: To promote the advancement of nuclear science and technology and the safe application of nuclear energy in all forms by:
(1) Standardizing measurement techniques and specifications for radiation effects and dosimetry including materials response, instrument response, and fuel burnup.
(2) Standardizing the nomenclature and definitions used in or relating to testing methods or instruments in support of nuclear industry.
(3) Maintaining a broad expertise in application of nuclear science and technology, especially the measurement of radiation effects from environments of nuclear reactor, particle accelerators, indigenous space, spacecraft, and radioisotopes.
(4) Maintaining a broad expertise in the applications of radioisotopes.
(5) Sponsoring scientific and technical symposia and pub lication in our fields of specialization.
(6) Performing liaison with related ASTM committees and other technical societies and organizations, both na tional and international.
(7) Advising other technical committees of the Society in our field of expertise.
The emphasis on maintaining expertise, on sponsoring symposia and publications and on liaison is to optimize information gathering and e x c h a n g e , our goal being to stay on the advancing edge of current nuclear technology.
The organization of E10 is outlined in Table I. In addi tion to these formal subdivisions of E10 which are roughly descriptive of the activities of the committee, there are a num ber of task groups in several of the committees, especially E10.02.
Subcommittee E10.01 is concerned with standards for analyz ing nuclear fuels for degree of burnup. These were developed years ago when their need was largely for laboratory assay but recently have been modified for larger scale application in nuclear power plants. 150 STEELE
TABLE I. SUBCOMMITTEES OF COMMITTEE E 10
E10.01 Nuclear Fuel Burnup E10.02 Behavior and Use of Metallic Materials in Nuclear Systems E10.03 Tracer Applications E10.04 Measurements Using External Radiation Sources E10.05 Dosimetry E10.06 Description of Radiation Damage in Metals Using Electron Microscopy E10.07 Radiation Effects on Electronic Materials and Devices and Pulsed Radiation Effects E10.07.01 Radiation Effects on Organic Materials E10.08 Procedures for Neutron Radiation Damage Simulation E10.09 Radiation Effects on Sensing and Controlling Instruments for Nuclear Power Reactors E10.10 Matrix Approaches to Standards for Nuclear Systems Tech nology E10.10.01 ASTM Generic Matrix Standard E 1 0 .10.02 Standards Development Matrix to Implement the National Energy Plan E 1 0 .10.06 Nomenclature and Definitions E 1 0 . 10.10 Matrix Approaches for Performance of Reinforcements for Nuclear Applications of Concrete E 1 0 .10.20 Matrix Standards for LMFBR E 1 0 .10.30 Standards Development Matrix to Establish the Long Term Availability of Fuels and Other Basic Resources E10.10.34 Overall Matrix for All Fuels E10.10.38 Matrix Development Plan to Establish the Avail ability of Fuels and Other Basic Resources E10.40 Matrix for the Development of Voluntary Consensus Standards to Help in Safeguarding Nuclear Fuels and Materials E10.il Matrix for the Identification of Standards for Nuclear Fuel Cycles E10.90 Advisory (Executive) Committee E10.91 Coordinating Committee for Nuclear Standards E10.92 Executive Coordinating Committee for Nuclear Standards
Subcommittee E10.02, whose name was recently changed to, "Behavior and Use of Metallic Materials in Nuclear Systems," was the focal point for much early activity in radiation effects in reactor alloys. This activity will be cited further as source of examples for reliability related ASTM standards. Subcommittees E10.03 and .04, now inactive, were founded to assist in specific standardization tasks relating to internal and external applications of radioisotopes.
The neutron dosimetry aspects of nuclear applications have been centered in Subcommittee E10.05. This activity has gener ated a systematic series of standards involving the determina tion of neutron dosage especially as related to experiments or in-service studies of neutron damage of metals.
Advanced laboratory techniques for measuring or simulating nuclear radiation damage are covered by two relatively new sub committees, E10.06, concentrating on electron microscopy tech- IAEA-SM-218/31 15 niques, and E10.08, concentrating on neutron simulation tech niques. These subcommittees have goals aimed toward more ad vanced systems which are now only in the research and develop ment stages. Nevertheless, they contribute significantly by standardizing procedures for laboratory study of radiation damage phenomena in alloys , possibly useful in providing the structural components of advanced nuclear systems.
Subcommittee E10.07 has the function of developing stand ards for measuring radiation effects of a pulsed nature and for defining such effects in electronic components. This group began as an outgrowth of space related activities but now has broadened its scope to include dosimetry for short-term or pulsed radiation sources. E10.09 has a similar function but is charged with concern for the effects of radiation on the sensing and controlling instruments for the current generation of nuclear power plants.
The newest activity of committee E10 is in the development of matrices for identifying possible standards needs in several areas. It's scope is as follows:
(1) To identify and systematize critical matrix components for nuclear standards.
(2) To outline these within the National Energy Plan (a moving target).
(3) To provide nomenclature and definitions.
(4) To assess performance of reinforcements for nuclear applications of concrete.
(5) To provide for the LMFBR.
(6) To outline available nuclear fuels resources.
(7) To provide for safeguarding nuclear fuels and materials.
(8) To support the full nuclear fuel cycle.
The last item has matured so that a new subcommittee E10.il, Matrix for Identification of Standards for Nuclear Fuel Cycles, was established to develop matrices to support critically impor tant aspects of the full nuclear fuel cycle.
Closely related to E10.10 and E10.il is the Coordinating Committee on Nuclear Standards, E10.91, whose function it is to coordinate the development of all nuclear standards required of ASTM. This group reviews all requests for new nuclear standards (except those generated by ASTM committees) identifies those needed and pertinent to ASTM activity and follows up with an agreement to have needed standards developed by specialists as these are found within ASTM. The Coordinating Committee seeks close liaison with other standards writing organizations as well as those organizations such as the Nuclear Regulatory Com mission and the Energy Research and Development Administration which often generate requests for new nuclear standards. Thus, a very useful function is performed by a technical committee for the whole of ASTM in a critical new technology. 152 STEELE
Administrative and general technical matters are managed by an Advisory or Executive Committee, E10.90, which includes all E10 officers and subcommittee chairmen.
4. ILLUSTRATION OF ASTM CONTRIBUTIONS TO RELIABILITY THROUGH STANDARDS
One important problem, the discovery of significant neutron radiation induced embrittlement of ferritic steels, provides a basis for describing how ASTM standards contribute to the integ rity and reliability of the primary reactor pressure vessel. During the 1950's decade, the extent of the radiation embrittle ment problem was initially realized and, since there was very limited experience with full scale power plants, the potential gravity of this problem was only lightly acknowledged. However, a limited number of investigators using research or test reactors learned enough to suggest a potentially serious problem.
Also in the 1950's the ASTM organized Committee E10 on Radioisotopes and Radiation Effects, primarily to attend to ques tions raised by the new nuclear technology especially through the project known as aircraft nuclear propulsion. Scientific and technical conferences were sponsored and some initial data on radiation damage of steels was presented. One E10 subcommittee addressed the question of standards in the area of radiation damage in materials (metals, ceramics and other inorganic mate rials). Primary emphasis gravitated to metals and the question was raised in 1958 as to how radiation embrittlement of reactor pressure vessels might be determined during service. From this question grew a task group and from the task group came the first skeleton of a standard to provide for surveillance of pressure vessel radiation damage. The nature of embrittlement is illus trated in Figs. 1, 2 and 3 [3]. Fig. 1 is a core cross-sectional diagram of surveillance capsule locations in a pressurized water reactor, Turkey Point 3. Fig. 2 shows surveillance results for a radiation sensitive weld of the Turkey Point plant. The in crease in brittle-ductile transition noted is extremely high for the fluence indicated. The cause is high copper (0.31%). The related Type A508 forging contained only 0.06% copper and showed no measurable radiation embrittlement for the exposure of ^5x10*8 (Fig. 3). The value of ASTM standards in providing a procedure for learning these critically important facts is apparent.
4.1, Surveillance Tests on Structural Materials in Nuclear Reactors
With this early beginning relative to the pace of installa tion of nuclear power plants, it was necessary to construct a projected matrix for critical components of the standard. The factors identified were: (a) types and numbers of specimens and their selection from vessel components, (b) controls for the in reactor exposure (capsule design and location, temperature mon itors, etc.), (c) neutron dosimetry procedures, (d) postirradia tion evaluation (removal schedule, testing procedures, etc.), and (e) analyses and reporting requirements. Where possible reference to standard specimens, tests and dosimetry techniques were cited. The earliest effort, which was standardized in 1962, necessarily lacked detail and was weak in references to other (supporting) standards in the nuclear field. IAEA-SM-218/31 153
о0
FIG.l. Diagram of a pressurized water reactor showing cross-section of core and location of all surveillance capsules designated by letters, S, T, U, V, W, X, Y and Z. Numbers in parentheses after capsule designation are ratio o f peak inner vessel wall flux to capsule flux [3].
The next version was published in 1966 following a burst of nuclear power development in the USA and remained essentially un changed until 1973. During this period significant advances in the knowledge of the potential vessel problem was attained as an IAEA review publication indicates [4]. Accordingly, the 1973 version of E185 addressed only the nuclear reactor vessel while earlier versions addressed both vessel and internal structural components. For purposes of this report, it is adequate to com pare these two versions. Key attributes and major changes from 1966 to 1973 are outlined in Table II [5,6]. Furthermore, a new revision is being completed now [7]. The new version contains major changes in materials selection requirements emphasizing 1 5 4 STEELE
TEMPERATURE
FIG.2. Curves showing Charpy V-notch fracture transition for irradiated weld metal (high copper, 0,31% — sensitive to radiation). Transition increases are shown at 30 ft-Ibf (40.7J) and 50 ft- Ibf (67.8J) [3].
the RT n d t , composition control and selection of low (shelf) toughness materials as well as others defined by 1973 version. However, it seems that the detailed annex on materials selection may be removed. More specific requirements for specimen orienta tion, location and total numbers of specimens were defined along with a plan for neutron dosimetry capsules which exceed prior requirements. A detailed plan for capsule withdrawal is being added but agreement has not been reached. Added guidance on neutron environmental determination is provided largely through reference to new standard E482 (Neutron Dosimetry for Surveil lance). New specific mechanical test requirements include a change to add 50 ft-lb (67.8 J) and upper shelf test data as well as an optional requirement for use of 50% fracture appear ance transition temperature at mid-range (middle of the transi tion energy). Another significant change being suggested is to require extrapolation of surveillance capsule results to both pressure vessel wall (inner) surface and to i thickness (inner) location in the vessel.
4.2 Neutron Dosimetry
Paralleling the development of E185 for radiation damage surveillance were the critically important series of standards for defining the neutron flux in the surveillance position or IAEA-SM-218/31 155
(-73) (-18) (38) (93) °C TEMPERATURE FIG.3. Charpy V-notch fracture data for irradiated ASTM Type A508 forging of low copper (0.06% Си) and low radiation sensitivity. No shift in transition due to radiation is . m easurable [3]. for accelerated studies in test or research reactors. This too was an evolutionary process and the story is told very well by a numerically ordered list of Committee E10 produced dosimetry s t a n d a r d s : E170 Definition of Terms Relating to Dosimetry E181 Analysis of Radioisotopes E182 Analysis of Phosphorus 32 E261 Measuring Neutron Flux by Radioactivation Techniques E262 Measuring Thermal Neutron Flux by Radioactivation Tech niques E263 Measuring Fast-Neutron Flux by Radioactivation of Iron E264 Measuring Fast-Neutron Flux by Radioactivation of Nickel E265 Measuring Fast-Neutron Flux by Radioactivation of Sulfur E266 Measuring Fast-Neutron Flux by Radioactivation of Aluminum E343 Test for Fast-Neutron Flux by Analysis of Molybdenum-99 Activity from Uranium-238 Fission E393 Measuring Fast-Neutron Flux for Analysis of Barium-140 Produced by Uranium-238 Fission E418 Fast-Neutron Flux Measurements by Track-Etch Technique E419 Guide for Selection of Neutron Activation Detector Mate rials E481 Measuring Neutron Flux Density by Radioactivation of Cobalt and Silver E482 Neutron Dosimetry for Reactor Pressure Vessel Surveillance E523 Measuring Fast-Neutron Density by Radioactivation of Copper E526 Measuring Fast-Neutron Flux by Radioactivation of Titanium E560 Extrapolating Reactor Vessel Surveillance Dosimetry Results TABLE II. KEY REQUIREMENTS AND CHANGES TO ASTM STANDARDS FOR SURVEILLANCE OF
REACTOR PRESSURE VESSELS [3, 4]
Volume Measurement ( Y e a r ) ______T o s t S p e c im e n s ______I r r a d i a t i o n C o n d i t i o n s ______of Neutron Exposure
BASIC STANDARD REQUIREMENTS
1966 M aterials. One heat of base metal, one Location. To match highest exposure, S ignificance. Stresses im port weld, one HAZ. neutron spectrum, tem perature as ance, need for accurate Fabrication H istory. Representative of w ell as is possible. records of flux, energy v e s s e l . Accelerated or Reduced Exposure. s p e c t r u m . Test Coupons Location. Base, highest Evaluate variation by dosimeters Neutron Dosimeters. Cites transition temperature; weld and if possible, if not, by calcula ASTM E261, "M easuring HAZ from same coupon. tion using conservative practice. Neutron Flux by Radioacti Type o f Specimens. Tension by ASTM E8; Thermal Control Specimens. Place to vation," which cites speci Charpy-V notch (Cy) by ASTM E23 evaluate therm al effects alone. fic techniques, E262, E263, (notched perpendicular to surface, Test Capsules. Protect specimen sur E 2 6 4 , E 2 6 5 , E 2 6 6 , a n d E 2 6 7 . taken in major working direction); face in corrosion resistant capsule HAZ, care that notch root is in HAZ. unless corrosion effects are sought; Number of Specimens. 2 tension, 8 or include low m elting alloys to assess more Cy per irradiation location highest tem perature, provide for but enough for TT curve from b rittle rig id capsule support. STEELE to 100% shear; unirradiated, 15 Cy, Specimen W ithdrawal. At least three 3 t e n s i o n . over design life , one to match C orrelation M onitors. W ell documented end-of-life. steel not part of reactor (see E184).
CHANGES FROM BASIC (emphasis sh ifte d to pressure vessel only)
1 9 7 3 * M aterial. Choose m aterial which may Accelerated Exposure. Not to exceed Neutron Dosim eters. Cites lim it life . (Annex covers procedures 3X vessel w all flux. new specific methods using in detail). Introduces lim it Flux Measurements. For accelerated fission detectors, E343 and of ASME Code Sect. Ill and effects of capsule, add flu x capsule at w all E393; removes reference to copper and phosphorus. or in a vessel w all capsule in E267. Recommends iron and •»Chemical Requirements. Section added; addition to those in accelerated unshielded cobalt in every calls for analysis to include: P, c a p s u l e . c a p s u l e . о, , auu * . •»Specimen O rientation and Location. Defines new orientation, normal to major working direction. Spells out details of weld and HAZ location. Number of Specimens. (Two cases; А “ ATT of <100UF, ф <5x1018, 12 Cv req. В “ ATT of >100°F, ф>5х1018, 12 Cv req. 2 tension for B; none for A.
1970 revision developed but 1966 version not sign ifica n tly d iffe re nt. New requirem ent. IAEA-SM-218/31 157
The first key group of these is the series E261-266. These were developed largely to support research reactor irradiation experiments though the general approach described in E261 . The thermal neutron technique, E262, as well as the technique for iron activation for fast neutron dosage, E263, have become central to surveillance needs. Added to these is the more recently devel oped procedures for fast neutron damage determinations by fission product analysis of 99Mo and 1Ц0Ва. Later, as surveillance needs grew with the growing number of nuclear power plants, it was found useful to add the two "how to" standards in addition to the earlier one, E419 for guiding the selection of dosimeter materials. These two are, E482 for specific neutron dosimetry needs to support pressure vessel surveillance and E560 for extrapolating dosimetry results from surveillance capsule to vessel wall. Standard E482, Neutron Dosimetry for Reactor Vessel Surveillance [6], provides specific guidance for determination of neutron spectrum including advice on modeling the reactor computer codes and use of cross-section libraries as well as specific advice on computer iterative analysis procedures. Other valuable specific guidance is provided on choice of neutron dosimeters (thermal and fast), on procedures for analysis and possible sources or error.
Standard E560, Extrapolating Reactor Vessel Surveillance Dosimetry Results [2], is, as the title implies, a source of analytical and analytical-experimental procedures for determin ing the variation in neutron flux density and energy spectrum between surveillance capsule location and pressure vessel wall. Procedures for reporting results of these analyses are also p r o vided.
4.3. Controlling Embrittlement Through Steel Specifications
As the understanding of the causes of radiation embrittle ment has grown, the opportunity for its control has grown also. The role of copper, phosphorus, and sulfur in affecting steel toughness after irradiation was defined from experiments in research reactors and verified by surveillance results. Within ASTM specifications most USA steels are standardized. Thus, it was appropriate to petition those committees with responsibility for steel standards to modify those for pressure vessel steels, especially the plate steel A533 and the forging grade A508, for limitations on copper, phosphorus and sulfur. This has been done with the result that anyone who seeks nuclear reactor pres sure vessel steels from among those accepted by ASME for such service will be alerted to the potential problem and specifica tions which minimize the problem. The immediate and easy inter action of ASTM committee members probably enhanced the modifica tion of standard steel specifications to control irradiation embrittlement in future nuclear power plants. For older plants, the work of E10 members in providing research results offered a different but potentially vital solution as well. This was the plan for correction of radiation embrittlement through thermal annealing to remove damaging defects.
4.4. Annealing to Correct Radiation Embrittlement
In order to provide some general guidance on the advantages of and possible procedures for annealing, a standard, E509-74, "Standard Recommended Guide for In-Service Annealing of Water Cooled Nuclear Reactor Vessels," was developed and published. 158 STEELE
The standard outlines critical considerations to the assurance of an effective annealing heat treatment covering especially: (a) factors to be analyzed in order to determine the feasibility of annealing, (b) the availability of (or need for) information on certain critical vessel properties, especially fracture tough ness before annealing, (c) possible approaches to vessel anneal ing and guides for required information to permit the choice of a best approach, and (d) surveillance procedures to verify an nealing results. The latter, it should be noted, will be most effective if planned with the initial vessel surveillance program.
While this standard provides only general guidance, it also illustrates the closing of the cycle on one nuclear power plant reliability question, that of vessel embrittlement during opera tion. Thus we illustrate an important role for ASTM standards in assuring or enhancing nuclear pressure boundary reliability, a critical reactor safety requirement for light water reactors.
5. C ONCLUSIONS
The ASTM, a collection of volunteer standards specialists, uses the due process-consensus approach to develop standard definitions, practices, methods of test, classifications, and specifications. The ASTM product (standards) extends to areas of high technology such as the nuclear, where in combination with other USA standards and code writing bodies, it contributes much to the assurance of reliability of nuclear reactor pressure components.
ASTM's contribution is illustrated by showing a cycle of activities whereby knowledge is advanced and procedures pro vided for minimizing the effects of nuclear radiation embrittle ment of reactor pressure vessels. The latter is illustrated through a sequence of related standards covering: (a) surveil lance tests for radiation effects in nuclear pressure vessels, (b) neutron dosimetry for quantitative analysis of such effects, (c) modified specifications for vessel steels to optimize radia tion resistance, and (d) annealing heat treatment for correcting radiation damage in existing reactor vessels. Collectively, ASTM standards provide a dynamic tool for guiding nuclear reactor operators away from this potential danger zone through a full cycle of pertinent standards. Other examples can be cited which illustrate this ASTM contribution to the reliability of sophis ticated engineering systems.
The other major contributing arm of ASTM in this critical area of reliability is that of information exchange through con ferences, symposia, and technical periodicals and other special publications.
REFERENCES
[1] ASTM Brochure, "ASTM and Voluntary Consensus Standards," ASTM, Philadelphia, PA, USA, undated.
[2] Annual Book of ASTM Standards, Part 45, Nuclear Standards, ASTM, Philadelphia, PA, USA (1977). IAEA-SM-218/31 159
YANICHKO, S.E., PHILLIPS, J.H., ANDERSON, S.L., "Analysis of Capsule T from the Florida Power and Light Company, Turkey Point Unit No. 3, Reactor Vessel Radiation Surveil lance Program," WCAP-8631, Westinghouse Electric Company, Pittsburgh, PA, USA (December 1975).
STEELE, L.E., "Neutron Irradiation Embrittlement of Reactor Pressure Vessel Steels," IAEA Technical Report Series 163, IAEA, Vienna, Austria (1975).
Annual Book of ASTM Standards, Part 30, General Test Methods, ASTM, Philadelphia, PA, USA (May 1967).
Annual Book of ASTM Standards, Part 45, Nuclear Standards, ASTM, Philadelphia, PA, USA (November 1975).
KOZIOL, J.J., private communication, Task Group Draft of Revised ASTM Standard E185 (August 1977).
IAEA-SM-218/24
NUCLEAR REACTOR PRESSURE VESSEL SURVEILLANCE CAPSULE EXAMINATIONS: APPLICATION OF AMERICAN SOCIETY FOR TESTING AND MATERIALS STANDARDS
J.S. PER RIN Colum bus Division, Battelle Mem orial Institute, Columbus, Ohio, United States of America
A b stra c t
NUCLEAR REACTOR PRESSURE VESSEL SURVEILLANCE CAPSULE EXAMINATIONS: APPLICATION OF AMERICAN SOCIETY FOR TESTING AND MATERIALS STANDARDS. A series of pressure vessel surveillance capsules is installed in each commercial nuclear power plant in the United States. A capsule typically contains neutron dose meters, thermal monitors, tensile specimens, and Charpy V-notch impact specimens. In order to determine property changes of the pressure vessel resulting from irradiation, surveillance capsules are periodically removed during the life of a reactor and examined. There are numerous standards, regulations, and codes governing US pressure vessel surveillance capsule programmes. These are put out by the US Nuclear Regulatory Commission, the Boiler and Pressure Vessel Committee of the American Society of Mechanical Engineers, and the American Society for Testing and Materials (ASTM). A majority of the pertinent ASTM standards are under the jurisdiction of ASTM Committee E-10 on Nuclear AppUcations and Measurements of Radiation Effects. The standards, regulations, and codes pertaining to pressure vessel surveillance play an important role in ensuring reliability of the nuclear pressure vessels. ASTM E 185-73 is the Standard Recommended Practice for Surveillance Tests for Nuclear Reactors. This standard recommends procedures for both the irradiation and subsequent testing of surveillance capsules. ASTM E 185-73 references many additional specialized ASTM standards to be followed in specific areas of a surveillance capsule examination. A key element of surveillance capsule programmes is the Charpy V-notch impact test, used to define curves of fracture behaviour over a range of temperatures. The data from these tests are used to define the adjusted reference temperature used in determining pressure-temperature operating curves for a nuclear power plant.
Introduction
Irradiation of the steels used in fabricating nuclear pressure vessels affects the mechanical properties. In particular, the tendency for failure by brittle fracture is increased [1], The levels of certain residual elements, particularly copper and phosphorus, have been shown to increase the radiation- induced embrittlement of pressure vessel steels [2]. Newer commercial nuclear plants are being constructed with steels which have a higher resistance to
161 162 PERRIN
radiation-induced embrittlement, but many plants now in service have steels with a lower resistance to radiation-induced embrittlement. In both cases, it is important to be able to monitor the changes in mechanical properties of the pressure vessel.
Commercial nuclear power plants in the United States each contain a series of pressure vessel surveillance capsules. These capsules are installed in a reactor at the beginning of the plant lifetime. A capsule typically contains neutron dosimeters, thermal monitors, tensile specimens, and Charpy V-notch impact specimens. Some capsules also contain fracture mechanics specimens such as wedge-open-loading or compact tension specimens. The mechanical property specimens are machined from the actual plates and weldments used for each pressure vessel. Surveillance capsules are periodically removed from a reactor and examined to determine changes in mechanical properties resulting from irradiation.
There are various standards and regulations which pertain to the surveillance capsule programs of U.S. reactors. Appendix G, "Fracture Toughness Requirements", and Appendix H, "Reactor Vessel Material Surveillance Program Requirements", to 10 CFR part 50, "Licensing of Production and Utilization Facilities" describe U.S. Nuclear Regulatory Commission require ments [3] . In addition, the U.S. NRC has recently issued Regulatory Guide 1.99, "Effects of Residual Elements on Predicted Radiation Damage to Reactor Vessel Materials" [4]. The U.S. NRC documents reference other documents which must be followed. These include Sections III and XI of the Boiler and Pressure Vessel Code of the American Society of Mechanical Engineers[5,6] and numerous standards prepared by the American Society for Testing and Materials (ASTM) , including ASTM E 185-73, "Surveillance Tests for Nuclear Reactor Vessels" [7].
Standards play an important role in increasing the reliability of reactor components such as pressure vessels. In the case of surveillance capsule examinations, data are obtained which are used to adjust reactor pressure-temperature heatup and cooldown curves for normal and hydrotest operation. In addition, the data obtained from Charpy V-notch impact specimens are used to determine the drop in the upper shelf energy level; this energy level in some cases may lead to a shortened life of a pressure v e s s e l.
The data used for making decisions relating to plant operation conditions and pressure vessel life obviously must be as accurate as possible; decisions based on pressure vessel surveillance capsule data are no better than the data on which they are based. In this paper, the role of current ASTM standards in obtaining high quality data in U.S. pressure vessel surveillance capsule examinations w ill be discussed. The discussion w ill include an overview of the pertinent ASTM standards, with an example of how one particular ASTM standard is used to obtain experimental results in a surveillance capsule examination.
ASTM Standards Used in Surveillance Capsule Examinations
The overall ASTM standard relating to surveillance capsule exami nations is E 185-73, Surveillance Tests for Nuclear Reactor Vessels. As discussed in the presentation by L. Steele[8], it has been developed over a period of many years. It is continuing to be modified under the jurisdiction of ASTM Committee E-10 on Nuclear Applications and Measurement of Radiation Effects, and the latest version is expected to be published in 19781.
1 The author is secretary of ASTM Committee E-10 and a member of the Task Group on the Rewriting of ASTM E 185. IAEA-SM-218/24 163
As stated in ASTM E 185-73, its scope includes "procedures for irradiating and testing mechanical test specimens for the purpose of monitoring and evaluating at periodic intervals, the radiation-induced changes occurring in the mechanical properties of the reactor vessel steels of light water-cooled nuclear power reactors". The standard includes recommendations in two major areas. The first is in regard to the overall design of a surveillance program for a specific reactor and the resultant capsule design and irradiation. This part of the standard is of use to the reactor manufacturer who sets up the program for each reactor being built. The second major area of ASTM E 185-73 covers the postirradiation examination and evaluation of the specimens contained in each capsule of a surveillance program. This second part of the standard is used by the utility owning the reactor and the hot laboratory chosen by the utility to examine and evaluate the capsules.
In the area of overall design of a surveillance program, ASTM E 185-73 covers selection of the material used in the program (including specific vessel plates and base/weld/heat-affected zone areas), characterization of materials (chemical composition and fabrication history), choice of mechanical property specimens (type, number, and location/orientation of specimens with respect to parent plate), choice of neutron dosimeters, thermal monitors, and capsule locations within a reactor.
In the area of capsule examination, ASTM E 185-73 covers the recommended time of removal of the surveillance capsules from a reactor, the examination and evaluation of tensile specimens, the examination and evaluation of Charpy V-notch impact specimens, the documentation of neutron exposure of a capsule, and the reporting of results.
In addition toASTM E 185-73, there are many specialized ASTM standards which cover various elements of a surveillance capsule examination program. The way in which a number of these standards relate to a typical capsule examination are shown in Figure 1.
The initial step in a program is the removal of a capsule from a reactor and the shipment of the capsule to a Hot Laboratory. This is followed by visual examination of the external capsule body for damage, disassembly of the capsule, and inspection of the specimens in the capsule [9].
Many capsules contain thermal monitor specimens. These are typically wires of an eutectic alloy which melt if they are heated above a specific temperature. By using two or more alloys wijth different melting points, subsequent examination w ill indicate whether or not the thermal monitors have been heated above the pertinent melting points. ASTM E 185 states the desirability of knowing both the reactor vessel and the capsule temperature, and suggests that thermal monitors be incorporated into surveillance capsules.
Knowledge of the neutron exposure of the surveillance capsule, and the subsequent relating of the capsule fluence to the vessel wall fluence are of paramount importance to a meaningful surveillance capsule examination. As indicated in Figure 1, there are a number of ASTM standards in this area. In addition to ASTM E 185-73, these include the following standards:
-ASTM E 184-62 (1968) "Effect of High-Energy Radiation on the Mechanical Properties of Metallic Materials"[10] -ASTM E 482-76 "Neutron Dosimetry for Reactor Pressure Vessel Surveillance"[11] -ASTM E 523-76 "Measuring Fast-Neutron Flux Density by Radioactivation of Copper"[12] -ASTM E 261-70 "Measuring Neutron Flux by Radioactivation Techniques"[13] -ASTM E 263-70 "Measuring Fast-Neutron Flux by Radioactivation of Iron"[14] -ASTM E 264-70 "Measuring Fast-Neutron Flux by Radioactivation of Nickel"[15], 6 PERRIN 164
CAPSULE SHIPMENT FROM REACTOR TO HOT LAB
CAPSULE D.[SASSEMBLY AIJD INSPE CTION
i ------\______NEUTRON DOSIMETRY MECHANICAL PROPERTIES THERMAL MONITOR ASTM E 1 8 4 -6 2 ASTM E 2 6 3 -7 0 ASTM E 1 8 4 - 62 ASTM E 1 8 5 -7 3 ASTM E 1 8 5 -7 3 ASTM E 2 6 4 -7 0 ASTM E 185- 73 ASTM E 2 6 1 -7 0 ASTM E 4 8 2 -7 6 ASTM E 5 2 3 -7 6
* 1______TENSILE CHARPY IMPACT COMPACT TENSION ASTM E 8-69 ASTM E 2 3 -7 2 ASTM E 3 9 9 -7 4 ASTM E 2 1 -7 0 ASTM A 3 7 0 -7 5
FIG.l. ASTM standards associated with a surveillance capsule examination programme. IAEA-SM-218/24 165
FIG.2. Charpy V-notch impact specimen located against anvils, about to be impacted by Charpy machine tup.
Figure 1 also indicates the three types of mechanical property specimens that are evaluated in a surveillance capsule program. These are the Charpy V-notch impact and tensile specimens in all surveillance capsules, and the compact tension or wedge-open-loading specimens in some capsules. The main standard for Charpy impact testing is E 23-72, "Notched Bar Impact Testing of Metallic Materials" [16]. Additional recommendations are listed in A 370-75 "Methods and Definitions for Mechanical Testing of Steel Products”[17], The primary standards for tensile testing are E 8-69, "Tension Testing of Metallic Materials" [18], and E 21-70 "Elevated Temperature Tension Tests of Metallic Materials" [19]. Compact tension and wedge-open-loading specimen evaluation is covered in E 399-74, "Plain-Strain Fracture Toughness of Metallic Materials" [20]. A new ASTM standard for these latter two types of specimens when tested under conditions not covered by E 399-74, namely, substantial elastic-plastic deformation preceeding fracture, is now under consideration.
Application of Charpy Impact Testing Standard ASTM E 23-72 to Surveillance Capsule Examinations
In this section an example of how one particular ASTM standard, ASTM E 23-72,"Notched Bar Impact Testing of Metallic Materials”, is used in pressure vessel surveillance capsule examinations w ill be presented. It is a key standard in such examinations. The use of it ensures that valid and consistent Charpy impact data are obtained. 166 PERRIN
2500
2000
E 3. 1500
1000
500 Lateral Expansion
0
в O' л;
>i O' и cШ и V о a(О E
- 5 0 0 50 100 150 200 Temperature,°C
FIG.3. Charpy impact properties for unirradiated and irradiated material. IAEA-SM-218/24 167
1 0 m m 8m m
- L / 2 T
- L 10 m m 5 5 m m
r a d i u s
Note - Permissible variations shall be as follows:
Adjacent sides shall be at 90° ±10 min Cross-section dimensions + 0 . 0 2 5 m m Length of specimen (L) +0, -2.5 mm Centering of notch (L/2) + 1 mm
Centering of notch ±1 m m Angle of notch ±1° Radius of notch ± 0 . 0 2 5 m m Dimensions to bottom of notch 8 ± 0 . 0 2 5 m m Finish requirements 2 pm on notched surface and opposite face; 4 pm on other two surfaces.
FIG.4. Charpy V-notch impact specimen as defined by ASTM E 23-72.
The Charpy machine has a pair of anvils against which the Charpy specimen rests. The specimen contains a notch, and the specimen is positioned in the anvils so that the center of the notch is located midway between the two anvils. The Charpy machine has a pendulum which is used to impact and break a specimen in a single blow. The part of the pendulum that strikes the specimen is the tup. Figure 2 shows a Charpy specimen positioned against anvils about to be struck by the tup [9]. The amount of energy absorbed by the specimen during impact is measured. A series of specimens tested over a range of temperatures is used to construct a curve of impact energy versus temperature.
In addition to the impact energy, two other types of data are obtained from a tested Charpy specimen. One is the fracture appearance, which is a visual estimate of the amount of shear or ductile type of fracture appearing on the specimen fracture surface. The other type of data obtained is the amount of lateral expansion. The lateral expansion is a measure of the deformation produced during fracture of the Charpy specimen. It is the change in specimen thickness of the section directly adjacent to the notch location. 168 PERRIN
TABLE I. CALIBRATION DATA FOR THE BATTELLE HOT LABORATORY CHARPY IMPACT MACHINE
Average Nom inal Battelle Energy Energy a V a r ia t io n Specimen Group (kg ■ m) (kg • m) A c tu a l Allow ed
Low Energy 1.70 1.72 -0 .0 2 kg-m +0.139 kg-m Medium Energy 7.07 7.32 -3.4% +5.0% High Energy 9.52 9.90 -3 .8 % +5.0%
Established by U.S. Army Materials and Mechanics Research Center.
Figure 3 is a typical curve for unirradiated and irradiated pressure vessel steel showing the energy-temperature and lateral expansion-temperature curves for weld metal [21]. The steel is SA 533 Grade В Class 1 and was irradiated to a fluence of 3.02 x 10^® nvt (E>1 MeV).
The Charpy data for the unirradiated material is used to establish a reference temperature, RTjjdt[5 ]. To d e fin e RT^d t * the nil ductility transi tion (NDT) temperature using drop weight specimens is first determined [22]. At a temperature not greater than NDT + 33 °C, each of three Charpy V-notch specimens tested must exhibit at least 890 /яn lateral expansion and at least 6.9 kg-m of absorbed energy. When a series of three unirradiated Charpy specimens meet these requirements, the NDT is the reference temperature. If these requirements are not met, additional Charpy tests are conducted in groups of three at higher temperatures until the requirements are met. The ЯТщ)Т is then the Charpy test temperature minus 33 °C. Using irradiated specimens from a surveillance capsule, an adjusted reference temperature can be determined. This is done by adding to RT^^ the amount of the temperature shift in the Charpy test curves between tne unirradiated material and the irradiated material, measured at the 6.9 kg-m level or the 890 дш lateral expansion level, whichever temperature shift is greater [3].
ASTM E 23-72 has extensive recommendations pertaining to Charpy testing,including machine installation, specimen design, machine inspection including dimensional tolerances or important parts, machine calibration, and machine operation.
Figure 4 is an example of how ASTM E 23-72 precisely specifies dimensional control. The figure shows the Charpy V-notch specimen used in pressure vessel surveillance capsule programs. Variations in specimen dimensions, including the notch, have been shown to significantly affect the Charpy impact data obtained in impact testing [23]. For example, a ±0.13 mm change in the radius at the root of the notch from the standard 0.25 mm radius was observed in reference [23] to cause a variation of up to 46% in impact energy. Therefore, the use of ASTM E23-72 guidelines in specimen machining can substantially improve the data generated.
ASTM E 23-72 provides recommendations for simple daily and weekly machine checks to determine if the machine appears to be in calibration. The standard provides detailed specifications which are to be checked annually. This includes an annual proof test using standardized specimens. The standard ized specimens are provided by the U.S. Army Material and Mechanics Research Center. A set is composed of 15 Charpy V-notch impact specimens, five at each IAEA-SM-218/24 169 of three known energy levels. The average value obtained during testing of the five specimens at each energy level is to correspond to the nominal value of the standardized specimens. The three values obtained during proof testing are to correspond to the nominal values within 0.139 kg’m or 5.0%, whichever is greater. Table I compares the results obtained during a typical proof test of the Charpy impact machine at the Battelle’s Columbus Laboratories. Each of the three values obtained experimentally are seen to be within the nominal values. This proof testing is an example of the high degree of effort to produce the best possible data incorporated into ASTM E23-72.
Future ASTM Standards
An important aspect of standards such as the surveillance capsule standards formulated by ASTM is that present standards be modified and new standards be introduced when the need arises. The need itself may be due to new plant designs or materials, new operating conditions, new knowledge obtained from present surveillance capsule programs, or changing government regulations. There are presently a number of task groups active within Subcommittee E-10.02 of ASTM Committee E-10 working in areas which may lead to modification of present standards or introduction of new standards relating to surveillance capsule programs. Examples of these current task group activities.are as follows:
Revision of ASTM E 185-73. As discussed earlier, ASTM E 185-73, "Standard Recommended Practice for Surveillance Tests for Nuclear Reactors", is being modified. The present version was approved and published in 1973. The task group which has rewritten E 185-73 expects to have the revision approved and published in 1978. The modifications incorporated into the new version result both from new information obtained in recent surveillance capsule examinations and changes in government regulations.
Interpretation of Charpy Data. A Task Group on Interpretation of Charpy Data has been active in discussing current problems in this area and the possibility of a new standard. Surveillance capsules typically contain only eight to twelve specimens of a particular material. Figure 3 shows an example of a curve generated on irradiated weld metal with only eight specimens. In addition to generating the general curve, it is desirable to be able to specifically determine the tempera tu re co rre sp o n d in g to 4 .1 kgm-rn, the energy co rre sp o n d in g to 6.9 kgrn-m, the tem perature corresponding to 890 jum lateral expansion, and the upper shelf energy level. Task group considerations include how to pick test temperatures, how to draw the most meaningful curve through the data and how to handle data scatter, with a special consideration to the problem of the lim ited number of specimens typically available from a surveillance capsule program.
Supplemental Test Methods. A Task Group on Supplemental Test Methods for Reactor Vessel Surveillance has been active in discussing new test methods. These methods could possibly be applied to future surveillance capsule programs, either to supplement present methods or to possibly replace current test methods at some point in the future. Supplemental methods which have been discussed include the use of equivalent energy or J-integral methods for evaluating compact tension specimens, a miniature dynamic tear specimen, and instrumented testing of both conventional Charpy specimens and precracked Charpy specimens having a small crack at the root of the V-notch.
Sum m ary
In order to ensure the continued safe operation of U.S. nuclear power plants, the effects of irradiation on the pressure vessel steel in each reactor are monitored through the use of a series of pressure vessel surveillance capsules. These capsules are periodically removed and the mechanical property and neutron dosimetry specimens contained in the capsules evaluated. A substantial number of ASTM standards have been written for surveillance capsule programs. These 1 7 0 PERRIN
standards cover both the irradiation and the subsequent evaluation of capsules. The use of these standards increases the accuracy of the data obtained, and this contributes to increasing the reliability of the pressure vessels of nuclear power plants.
REFERENCES
[1] BUSH, S.H., "Structural materials for nuclear reactors", J. Test, Eval. Testing and Evaluation 2_, 6 (1974) 435.
[23 STEELE, L.E., "Neutron Irradiation Embrittlement of Reactor Pressure Vessel Steels", Technical Report Series No. 163, IAEA, Vienna (1975) 185.
[3] Title 10, Code of Federal Regulations, Part 50, "Licensing of production and utilization facilities, Appendices G and H, U.S. Government.
[4] "Effects of residual elements on predicted radiation damage to reactor vessel materials", Regulatory Guide 1.99, Revision 1, Office of Standards Development, U.S. Nuclear Regulatory Commission (April, 1977).
[5] ASME Boiler and Pressure Vessel Code, Section III, "Rules for construction of nuclear power plant components", American Society of Mechanical Engineers (1974 Edition).
[6] ASME Boiler and Pressure Vessel Code, Section XI, "Rules for inservice inspection of nuclear power plant components", American Society of Mechanical Engineers (1974 Edition).
[7] ASTM E 185-73, "Sta n d a rd Recommended P ra c t ic e fo r S u r v e illa n c e T e sts fo r Nuclear Reactor Vessels", 1976 Annual Book of ASTM Standards, Part 45, American Society for Testing and Materials, Philadelphia, Pennsylvania, U.S.A. (1976), 737.
[8] STEELE, L. E., "Role of American Society for Testing and Materials in providing standards to support reliability technology for nuclear power p la n t s ", these Proceedings, paper IAEA-SM-218/31.
[9] PERRIN, J. S ., FROMM, E. 0 ., and LOWRY, L. M ., "Remote d isa sse m b ly and examination of nuclear pressure vessel surveillance capsules", (?roc.25th Conference on Remote Systems Technology, 1977) American Nuclear Society, Hinsdale, Illinois, U.S.A. (1977).
[10] ASTM E 184-62 "Sta n d a rd Recommended P ra c t ic e f o r E f fe c t s of H igh -En e rgy Radiation on the Mechanical Properties of Metallic Materials", 1976 Book of ASTM Standards, Part 45, American Society for Testing and Materials, Philadelphia, Pennsylvania, U.S.A. (1976) 733.
[11] ASTM E 482-76, "S ta n d a rd Recommended P ra c t ic e f o r Neutron D osim etry fo r Reactor Pressure Vessel Surveillance", 1976 Book of ASTM Standards, Part 45, American Society for Testing and Materials, Philadelphia, Pennsylvania, U.S.A., (1976) 921.
[12] ASTM E 523-76, "Standard Method for Measuring Fast-Neutron Flux Density by Radioactivation of Copper", 1976 Bo o k of ASTM Standards, Part 45, American Society for Testing and Materials, Philadelphia, Pennsylvania, U.S.A. (1976) 947. IAEA-SM-218/24 171
[13] ASTM E 261-70, "Standard Method for Measuring Neutron Flux By Radioactivation Techniques", 1976 Book of ASTM Standards, Part 45, American Society for Testing and Materials, Philadelphia, Pennsylvania, U.S.A. (1976) 776.
[14] ASTM E 263-70, "Standard Method for Measuring Fast-Neutron Flux by Radioactivation of Iron", 1976 Book of ASTM Standards, Part 45, American Society for Testing and Materials, Philadelphia, Pennsylvania, U.S.A. (1976) 795.
[15] ASTM E 264-70, "Standard Method for Measuring Fast-Neutron Flux by Radioactivation of Nickel", 1976 Book of ASTM Standards, Part 45, American Society for Testing and Materials, Philadelphia, Pennsylvania, U.S.A. (1976) 801.
[16] ASTM E 23-72, Standard Methods for Notched Bar Impact Testing of Metallic Materials", 1976 Book of ASTM Standards, Part 10, American Society for Testing and Materials, Philadelphia, Pennsylvania, U.S.A. (1976) 197.
[17] ASTM A 370-75, "Standard Methods and Definition for Mechanical Testing of Steel Products",•1976 Book of ASTM Standards, Part 10, American Society for Testing and Materials, Philadelphia, Pennsylvania, U.S.A. (1976) 28.
[18] ASTM E 8-69, "Standard Methods of Tension Testing of Metallic Materials", 1976 Book of ASTM Standards, Part 10, American Society for Testing and Materials, Philadelphia, Pennsylvania, U.S.A. (1976) 120.
[19] ASTM E 21-70, "Standard Recommended Practice for Elevated Temperature Tension Tests of Metallic Materials", 1976 Book of ASTM Standards, Part 10, American Society for Testing and Materials, Philadelphia, Pennsylvania, U.S.A., (1976) 187.
[20] ASTM E 399-74, "Standard Test Method for Plane-Strain Fracture Toughness of Metallic Materials", 1976 Book of ASTM Standards, Part 10, American Society for Testiig and Materials, Philadelphia, Pennsylvania, U.S.A. (1976) 471.
[21] "Surry Unit No. 2 Pressure Vessel Irradiation Capsule Program: Examination and Evaluation of Capsule X", Virginia Electric and Power Company (1975) U.S. Nuclear Regulatory Commission Public Documents Room, Washington, D C, U.S.A.
[22] ASTM E 208-69, "Conducting Drop-Weight Test to Determine Nil-Ductility Transition Temperature of Ferritic Steels", 1976 Book of ASTM Standards, Part 10, American Society for Testing and Materials, Philadelphia, Pennsylvania, U.S.A. (1976) 330.
[23] FAHEY, N. H., "Effects of variables in Charpy impact testing", Mater. Res. Stand. 1, 11 (1961) 872.
DISCUSSION
on papers SM-218/31 and SM-218/24
M. BRU M O VSKY: One of the most important parameters of irradiation embrittlement is the transition temperature shift. It depends, to a large extent, on the shape of the transition curve (dependence of notch toughness on Charpy-V specimens and on temperature). Is it planned to look for a more objective method for the evaluation of these results, for example by m aking a com putation on the basis of the least squares method and a form ula (e.g. hyper bolic geometric function)? It would be very useful if some such method could be standardized in order to ensure that all experimental tests were evaluated in the same way, so that they could be compared with each other. J.S. PER R IN : The A STM task group on Charpy data interpretation has considered various possible standardized techniques for the plotting of Charpy data. One method considered is that presented by W. Oldfield in the article “Curve Fitting Im pact Test Data: A Statistical Procedure”, published in the November 1975 issue of “A ST M Standardization News”. However, statistical curve fitting will not solve the problem of how to produce a definitive Charpy transition curve when there are insufficient data points. D.G.H. LA T ZK O : In view of the well-known uncertainties mentioned in Mr. Perrin’s answer to Mr. Brum ovsky, is replacement of standard Charpy-V test pieces by standardized SE N or C T specimens for J-integral testing being considered for the future? L.E. ST EELE: Yes, it is. However, the specimen and test method should not take on the status of standard through the A ST M Committee E-24 on Fracture Testing before its acceptance by A ST M Committee E-10 on Nuclear Applications, the A SM E Code and the Nuclear Regulatory Commission or before its use is fully validated. Nevertheless, a number of activities are leading us towards an approach to surveillance based on elastic-plastic fracture mechanics, though I believe the next step will be an acceptable surveillance specimen of the LEF M type. This is likely to be the £T compact tension specimen. A.B. L ID IA R D : With respect to your description of the A STM procedure for agreeing on a standard, what do you mean when you say that it is voluntary and based on a full consensus? L.E. ST EELE: It is voluntary in the sense that all contributors to standards writing are volunteers; no voting member of any A ST M Committee is paid. There are paid consultants for special purposes but they cannot vote. Full consensus means that any technical committee member m ay vote against any standard and this negative vote, if supported by written or oral justification, must be considered by the standards writing committee. It m ay be reversed by an affirmative vote of 60% (soon to be 90%). If reversed, the negative voter m ay
173 174 DISCUSSION appeal to a final authority, the Com m ittee on Standards, which gives a final ruling. Furthermore, product-oriented committees must have a voting member ship with producer members balanced against user and general interest members. A. N IELSEN : Is A ST M considering instrumented impact testing (e.g. to evaluate the J-integral)? L.E. STEELE: Within ASTM Committee E-10 on Nuclear Technology and Applications, Sub-committee E l 0.02 has Task Groups on the use of instrumented Charpy-V specimens, on alternative fracture test specimens for surveillance and on possible standard techniques for Charpy curve fitting. Committee E-24 on Fracture Testing is working on methods for elastic-plastic fracture. The Chairman of E-10 has encouraged and has him self recently developed an approach for co-ordinating activities in which the needs of E-10 for pressure vessel surveillance are considered by E-24 when developing new fracture test m e t h o d s . A. N IE L S E N : If only 8 Charpy test pieces are available from a surveillance capsule it m ay be useful to forget about the transition curve and to concentrate on upper shelf energy (2 pieces) and the transition shift (so that there are 6 pieces left to try to hit the centre of the transition range). J.S. P E R R IN : It is difficult to define a complete Charpy transition curve with only 8 specimens, particularly if the material is one which exhibits substantial data scatter. In some United States surveillance capsule examinations, in which there are only 8 Charpy specimens for each material, 3 specimens are used to define the upper shelf energy level and the remaining 5 are used to define the transition region, including the 35 and 50 ft Tbf levels. A.L. LOW E: One fact that Dr. Perrin did not mention was that the A STM specifically describes m inim um requirements for surveillance programmes and that suppliers of surveillance programmes for reactor vessels have a duty to anticipate the future needs of a specific reactor vessel by providing any additional specimens to ensure the future licensability of the plant. This m ay require putting specimens in capsules that have not been accepted by A ST M or the use of analytical techniques which have not been fully developed. It is difficult to anticipate the licensing needs of a reactor 10 years hence when regulations m ay be changed while the capsules are being irradiated. However, if A ST M standards are available they should be used to the extent applicable.
METHODS OF ASSESSING THE PROPERTIES OF HEAT-AFFECTED ZONES
Following this discussion a presentation was made by H.A. Maurer of work done by J. Ewald and K. Kussmaul of the Staatliche Materialprüfungsanstalt (MPA) of the University of Stuttgart. As this is to be published under the above title in Atomic Energy Review early in 1978, the text is not reproduced here. Session VIb
PRACTICAL EXPERIENCE WITH COMPONENTS
RELIABILITY C h a ir m e n
G . Ö S T B E R G S w e d e n
D.G. DALRYMPLE C a n a d a IAEA-SM-218/25
STEAM GENERATOR RELIABILITY - CANADIAN PRACTICE
R.I. HODGE, J.E. LeSURF, J.W. H ILBO RN Atom ic Energy of Canada Limited, Chalk River, Ontario, C a n a d a
A b stra c t
STEAM GENERATOR RELIABILITY - CANADIAN PRACTICE. The paper updates “Steam Generator Reliability — The Canadian Approach” prepared by the authors for the XIX Nuclear Congress of Rome in March 1974. At that time we reported three steam generator tube leaks, one each in the 20 MW(e) Nuclear Power Demonstra tion (NPD), the 200 MW(e) Douglas Point and the 540 MW(e) Pickering 2 units. We have experienced no further failures. Since the repair of those leaks the accumulated operating time has doubled on NPD and tripled on Douglas Point and Pickering Unit 2 (increased more than nine times if we include the three other identical Pickering Units). The steam generators in the Bruce stations recently brought on line, and in the 600 MW(e) stations under construction, differ from those mentioned above, having a greater size and surface heat transfer rate; furthermore some of the 600 MW(e) stations will be sea-water cooled, which increases the vulnerability of the steam generators to condenser cooling water in-leakage. This paper reports on developments to secure the reliability of these new steam generators; experiments to optimize chemical control of corrosion, improvement of techniques to analyse internal flow patterns for design calculations on dry-out, vibration, etc., and development of special equipment for in-service examination of tubes.
1. INTRODUCTION
This paper updates “Steam Generator Reliability — The Canadian Approach” [ 1 ] prepared in 1974; for a complete statement of our present practice the earlier paper should be read together with this one. We do not, for example, include in this paper any description of the boiler repair technique involving explosive plugging of tubes: that is described in some detail in the earlier paper and the repair service is now available commercially from Canadian boiler manufacturers. Also, an outline of our tube vibration analysis technique was given in Ref.[ 1 ] and is here updated principally by references to other recent p a p e rs. Steam generator tube failures were reported in 22 of the 62 water-cooled nuclear power plants we surveyed in 1975 [2], and to judge by current magazine articles the situation in 1976 and 1977 might be worse. However, the boilers in the operating stations in Canada have been trouble-free since the last tube
177 178 nta ikrn ecos2 hafe huto . n tdow u sh fter a h 24 reactors Pickering d an t in o P F IG .l. Average boiler room radiation field s a t (a) N PD 24 h after sh u tdow n , (b) Douglas (b) , n tdow u sh after h 24 PD N (a) t a s field radiation room boiler Average .l. IG F 801LER ROOM RADIATION FIELDS ( R/h) HODGE et al. et HODGE
IAEA-SM-218/25 179 failure in Pickering 2 in January 1974. On December 31, 1976 we had accumulated 122 500 effective full power hours in the four Pickering units: m ultiplying by the 2.23 X 10 6 m of boiler tubing in these units yields ~ 3.1 X 10 7 m years in which we have experienced 1 failure. Thus so far the Pickering A station (i.e. units 1 — 4) is giving better service than the Ufetime reliability objective o f less than 1 leak per year per 3 X 10 6 m of tubing. It should be pointed out that the condenser cooling water at our operating stations is not aggressive in the steam generator environment, so we are relatively immune to small leaks in the condensers. This will not be the case in two of the 600 M W (e) reactors that we are now building, which will have sea water cooling. These 600 MW (e) units (identified as Post-Bruce in Ref.[l]) are also larger, which increases the problem of keeping the tube-sheet clear of deposit, and have higher surface heat transfer rate (20% more than Bruce, 100% more than Pickering), which increases the propensity to concentrate solubles in near-stagnant zones adjacent to heated surfaces. Our awareness o f the increased potential for damage in these new stations has led us to further studies o f material corrosion control by tubing selection and water chemistry, of geometric and hydraulic arrangements to minimize stagnancy, of vibration and fretting which limit these hydraulic arrangements, and of techniques to periodically examine the tubes in situ. These topics form the subject matter o f this paper.
2. PRIMARY SIDE CORROSION
2.1. Radiation Fields as a Consequence of Corrosion
Radiation fields in the boiler rooms at CAN DU -PH W 1 plants operating in Canada are shown in Fig.l. The radiation fields in the boiler rooms at N PD and Douglas Point continue to decrease as a consequence of actions which are described elsewhere [3]. The radiation fields in Pickering units 1 to 4 are rising slowly,as predicted by the mathematical models of activity transport [4]. The radiation fields in future C A N D U units should be even lower than those in Pickering A because of actions which are outside the scope of the present paper, but are described fully elsewhere [ 6 ]. From present experience and predictions based on proven principles, it is expected that boiler room radiation fields will not contribute significantly to radiation doses in C A N D U units being built or planned.
1 Canada Deuterium Uranium — Pressurized Heavy Water. 1 8 0 HODGE et al.
2.2. Intergranular Corrosion
In the previous review [1 ] the possibility of intergranular corrosion occurring from the primary side was cited as one of the reasons for preferring Incoloy-800 to Inconel-600 for boiler tubing. The occurrence of this phenomenon at non- C A N D U reactors with Inconel-600 boiler tubes [ 6 ] endorses the previous assessment. Primary side corrosion has not been reported from stations with Incoloy-800 boiler tubing, nor has it been produced in the many laboratory tests under conditions which have produced intergranular corrosion in Inconel-600 [7, 8 ]. A restricted specification of Incoloy-800 continues to be the recommended boiler tube material for all CAN DU -PH W reactors.
3. SECONDARY SIDE CORROSION
3.1. Caustic Cracking
Further laboratory studies [9] have extended the knowledge of caustic cracking of high nickel alloys beyond that published previously [ 1 ]. A summary of the current status is: highly stressed Inconel-600 will crack in caustic concentrations > 1 wt% NaOH; highly stressed Incoloy-800 and Monel-400 will crack in concentrations > 4 wt% NaOH; cracks are predom inantly intergranular in Inconel-600 and trans- granular in Incoloy-800; both inter- and transgranular features are present on cracks in Monel-400; the threshold stress to produce cracking in caustic concentrations of between 4 and 20 wt% NaOH, is between 48 and 110 MPa (7 to 16 kpsi) 2 for Inconel-600 and Incoloy-800, and between 110 and 170 MPa (16 to 25 kpsi) for Monel-400. The broad conclusion is that Incoloy-800 is marginally superior to Inconel-600 in resisting caustic cracking. However, it is better to avoid form ing N aO H than to rely on superior materials.
3.2. Acid Wastage
Inadequate control of the chemistry within crevices in the boilers at many P W R s 3 has resulted in corrosive thinning (“acid wastage”) of the Inconel-600 boiler tubes, leading to some tube leaks and requiring m any tubes to be plugged because of loss of wall thickness [Ю ].
2 kpsi = 1000 lbf/in2 = 6895 X 103 Pa. 3 Pressurized Water Reactor IAEA-SM-218/25 181
Current knowledge on wastage may be summarized as follows: the corrosion rate of Inconel-600 in phosphate solutions increases rapidly as the N a:P0 4 molar ratio falls below 2; the rate of attack increases rapidly with temperature in the range 275°C to 325°C; attack is initially in the form of pitting, but with time the pits coalesce to give the appearance of general attack; from its chemical com position Incoloy-800 should be more resistant than Inconel-600 but any practical benefit is insufficient to confer a satisfactory lifetime under adverse chemistry conditions.
3.3. Denting
Cracking and wastage in PW R steam generators led to m odification of their boiler water chemistry. In particular, the permissible range of N a:P0 4 r a tio s was restricted to avoid exceeding 3 (caustic cracking) or dropping below 2 (acid wastage). Because the restricted range could not be maintained with the manual sampling procedures used, phosphate addition was replaced by an A ll Volatile Treatment (AVT). Hydrazine and volatile amines had previously been incorporated with the phosphate treatment but now stood alone. Shortly after this change in chemistry, the denting phenomenon appeared. Initially, it was observed only in boilers which had been exposed to phosphate before switching to AVT. The degree and frequency of denting was particularly severe at sea-water-cooled stations [11]. Residual phosphate deposits were blamed. Experiments by A E C L [12] and others predicted that denting would be observed in sea-water-cooled units with boilers that had operated with A V T only. This prediction has been fulfilled with the recent detection of denting at Ringhals [13] and at other stations which have seen little or no phosphate. A general summary of the current understanding of denting is: deposits form within the annulus between the tube and tube support plate (especially with phosphate treatment, but corrosion products and im purity salts will also deposit there with the A ll Volatile Treatment); sea-water salts (notably chlorides) trapped within the deposits trans form to very corrosive, acidic mixtures; the carbon steel support plate is corroded; the corrosion products distort the tube and the support plate; very high stresses generated within the dented tube cause intergranular corrosion to occur on the primary side of the tube, leading to cracking and leakage in some cases; distortion of the support plate causes primary side cracking to occur at tight U-bends; 182 HODGE et al.
Incoloy-800, Monel-400 or any other boiler tube material would not be significantly more resistant to denting than is Inconel-600, but Incoloy-800 is much more resistant to primary side attack than is Inconel-600.
4. BOILER WATER CHEMISTRY
4.1. The Nature of Boiler Water
Boiler water may contain: volatile impurities condensed with the steam (ammonia, amines); gas ( 0 2, C 0 2) by air in-leakage caused by vacuum in the turbine and c o n d e n s e r ; salts from cooling water leaking in at the condenser; traces of salts and gases introduced with the make-up water; corrosion products generated in the condenser, feed train and boiler; chemical additives to suppress oxygen and corrosion. The study o f boiler water chemistry is difficult experimentally because samples of liquids and deposits need to be taken at the operating conditions, since changes occur (e.g. in pH, solubility of salts) on cooling down an experi mental rig for examination. Despite these difficulties, considerable advances have been made both theoretically and experimentally towards an understanding of the chemical reactions in a boiler, both in the bulk water and locally within heated crevices. Some types of cooling water are m uch more aggressive than others when they leak into an operating boiler. The degree of aggressiveness may be estimated from theoretical prediction of changes in com position [14].
4.2. Phosphates in Boiler Water
The chemistry of sodium phosphate in pure water and the interaction between phosphate and some other ions in solution (e.g. corrosion products) has been reported [15]. O f particular significance is a series of experiments in a m odel boiler [ 12] which demonstrated that: sea water ingress was extremely corrosive towards a carbon steel appendage on the tube; A V T did not suppress this attack by sea water; controlled addition of phosphate did suppress the corrosion by sea w ater. IAEA-SM-218/25 183
5. BOILER FEED WATER TREATMENT
5.1. Condensate Polishing
A possible way to prevent impurities from the condenser entering the boiler is to pass all of the condensate through ion exchange resins (“condensate polishing”). This strategy has been adopted at m any fossil-fuelled power stations and at some water-cooled nuclear plants. The positive features of condensate polishing are simply stated; when operated correctly it prevents cooling water salts entering the boiler. Som e of the negative features are listed here - opinions vary as to the weight to be given to each of these points: Because resins are thermally unstable, they have to be installed at the cool end of the feed train; to take m axim um advantage o f the resins, all recycled water (e.g. reheater drains) has to be piped back to the condenser hot well, im posing constraints on the feed train design; impurities introduced beyond the resins (notably corrosion products from the feed water heaters) are not removed: constant vigilance is required to ensure that resins are effective (e.g. if a large condenser leak occurred when a resin bed was nearly depleted, breakthrough would occur); impurities m ay be introduced by the resins themselves (e.g. acids and bases used for regeneration); even new resins (or properly regenerated resins) will leach sodium ions (a phenomenon known as “sodium throw”); the significance of sodium throw to caustic cracking is a subject on which very strong positions are taken, for and against; there is little evidence concerning the significance o f this effect.
5.2. Hot Filtration
There is growing support for the use of filters at the hot end of the feed train to remove particulate material before it enters the boiler. Com pared with condensate polishing, hot filtration has the advantage of being able to remove all the solids present in the feed water. Its disadvantage is that it will not remove soluble ions which may form precipitates within the boiler. M ost of the emphasis at the moment is on magnetic filtration.
5.3. Combined AVT/Phosphate Treatment
The traditional procedure in controlling power station chemistry is to take samples manually from the feed water or boiler blowdown according to a predetermined frequency, analyse them in a laboratory, and report the results to a 1 8 4 HODGE et al. senior member of the operating staff for interpretation and decision. Clearly, this procedure is too slow and inflexible to give adequate control of phosphate tre a tm e n t. We recognized this deficiency several years ago and devised a scheme of autom atic sampling, analysis and control [ 1 ]. Samples are taken and analysed automatically. The results are fed (as electrical signals) to a computer which activates chemical addition pum ps to make appropriate corrections to the chemistry. Samples may be taken much more frequently than by manual sampling, and sampling continues 24 hours a day. The intention is to operate with A V T control when the feed water is pure, but to switch automatically to phosphate additions at times when sea water in-leakage occurs. Maxim um blowdown will be employed at this time. The section of the condenser containing the leak will be located and isolated as soon as possible. Once this has been done, the feed water can return to A V T control. A demonstration of the system at a coal-fired power station has proven the reliability and versatility of the autom atic control concept. With good boiler design and adequate, automatic control of water chemistry, phosphate additions can be used safely.
6 . THERMAL-HYDRAULIC ANALYSIS PROGRAM .
We have developed a computer program (BO SS), principally to analyse the distribution of flow on the secondary side of steam generators. It also computes steaming capacity, recirculation rate, primary flow distribution and primary outlet temperature spread. In the past two years it has been used regularly in appraisal of manufacturers’ design proposals, to establish adequate margin for repair of prim ary to secondary leaks by plugging defective tubes, and to provide flow velocity and quality distribution for further analyses of tube vibration, thermal stress and tendency to concentrate soluble salts in various locations. Recently it has been used more directly in design engineering, to assess the value of design options, for example of the preheater thermal plate, or of the windows in the shroud. The BO SS program is based on one developed several years ago for Atom ic Energy of Canada Limited by C H A M Ltd. in the U.K. BO SS retains the basic methods introduced by C H A M Ltd. to solve the coupled thermal and fluid- dynam ic equations in a large matrix of cells, but has been extensively developed in its treatment o f the U-bend region, of com plex entries and of preheaters, also in its iteration procedures, and in its input and output routines. B O SS results for the design operating condition of a Pickering steam generator (see Fig.2) are presented as an example in Figs 3 and 4. There is a plane of sym m etry in the boiler (that of Fig.2) parallel to the planes of the U-tubes IAEA-SM-218/25 185
FIG.2. Pickering-type steam generator.
and cutting the hot (primary flow up) and cold (which includes the preheater) zones in halves. The radial scale has been expanded to twice the axial scale for clarity of vector presentation. Design operating conditions are:
steam pressure 4.1 MPa steam flow 67.8 kg/s primary inlet temperature 293°C primary flow 697 kg/s
The BO SS calculation at the design steam pressure and primary flow and inlet temperature indicates that 6 8 . 6 kg/s steam will be raised and that the primary outlet temperature spread is from 250° C to 255°C. Correct steam rate can then be obtained by m inor reduction of primary inlet temperature, and there is a margin for some plugging of tubes. Some 24% of the downcomer flow 186 HODGE et al.
I m /s VECTOR
FIG.3. Mixture velocity vectors. IAEA-SM-218/25 187
VERTICAL SECTION HORIZONTAL SECTIONS
FIG.4. Mixture quality contours (numbers indicate percentage). 188 HODGE et al. enters the hot-side boiling zone by way of a space under the preheater floor or thermal plate. This is supplemented by a leakage of about 33% of the feed water from the preheater through the floor. The resulting velocity distribution is shown on the left-hand side o f the lowest diametral section in Fig.3. The recirculation ratio (downcom er flow divided by design steam flow) is 6.76, and the average quality at the prim ary separators is 13.2%. Figures 3 and 4 show that the distributions of flow velocity and quality entering the annular duct leading to the primary separators are somewhat uneven. Presumably improvement in separator effectiveness could be obtained by rendering these distributions more uniform. The lowest diametral section on Fig.4 shows that boiling can occur on the tube-sheet surface but that 1% quality is not attained; in fact the m aximum quality is ~ 0.4%. This is considerably less than found in other boilers. The vertical section shows that evaporation will occur in the preheater, with about 2 % quality in the flow out into the boiling zone. Boiling within the preheater should be avoided at design operating conditions because of the possibility of concentrating solubles in the baffle holes; in this case the effect of feed water leakage was underestimated in the original design. However, the tube material, the control o f the feed water chemistry and the purity o f the condenser cooling water give confidence that this flow condition will not prove troublesome. We would like to acknowledge that R.H. Shill of Chalk River Nuclear Laboratories is wholly responsible for the developments described in this section. The proprietary nature of his reports on his work prevents us from referencing th e m .
7. TUBE VIBRATION ANALYSIS
Our previous paper [ 1 ] described the various ways in which steam generator tubes can be excited into oscillation by the external (secondary) flow; internal flow is insignificant in this respect. Since 1974 we have greatly improved the empirical knowledge essential to the calculation of tube response by a variety of observations ranging from direct laboratory experimentation to field tests of production heat exchangers. A computer code (PIPEA U ) has been developed to calculate m odel responses of ideally-supported multispan tubes to distributed random forces (e.g. resulting from turbulence) and to periodic forces (e.g. resulting from wake shedding): it also calculates the critical cross-flow velocity at which fluid-elastic instability can occur. M.J. Pettigrew has prepared a detailed review o f this work recently [16]. Tube vibration analysis should include consideration of any departures from the general layout of the tube bundle, for example: baffle support rods differing in diameter from the neighbouring tubes; tubes adjacent to a partition wall or bundle edge seal; missing tubes; tube-free lanes. IAEA-SM-218/25 189
Д
0.2
E
>-
о о LU > 0.1 < О U- <Г ÜJ CL ••• SEAL TYPE-EDGE 3 (Л и а -GAP □ □□-RADIUS ч
0.0 2 3 4 CLEARANCE (mm)
FIG.5. Chatter velocity for various seal types.
A careful study of m any such special situations in liquid cross-flow has been conducted at the University of Ottawa [17], under contract from AECL. An example is given in Fig.5. Here the mean superficial velocity (i.e. calculated on duct dimensions, ignoring tubes) at the commencement of chatter (i.e. audible contact between tubes or tube and seal) is plotted against seal clearance for 3 different types. The tubes in this test apparatus are 12.7 mm diameter, have a natural frequency immersed o f about 30 Hz, and are arranged in an equilateral triangular array of pitch: diameter ratio 1.36, the flow being parallel to the pitch. It is evident that the radius seal is particularly vulnerable to chatter, particularly at small clearance. The norm al bundle experiences general fluid-elastic instability at about 0.28 m/s so evidently the edge strip acts very m uch like a neighbouring row o f tubes. Canadian steam generators have various designs of tube support plates for the boiling zone which have the feature of non-circular holes in common: these designs include lattice bars, three-lobed broached holes, and welded corrugated 1 9 0 HODGE et al.
FIG. 6. Tube fretting in interrupted-hole supports.
strip assemblies. Currently we are examining the dynam ic interaction between tube and support in such designs, which m ay lead us into m odification of our current linear response analysis.
8. FRETTING OF TUBES AND SUPPORT PLATES
A s mentioned in our previous paper [1 ], we are concerned with fretting rates less than 5 X 10"9 m/h, at which level tube penetration could occur within 30 years o f operation. Sustained fretting at such low levels is possible in the steam generator environment at contact forces of about 1 N or less, depending upon material and geometry, etc., and which can result from the hydraulic excitations mentioned above. IAEA-SM-218/25 191
P.L. Ко of the Chalk River Nuclear Laboratories has been responsible for laboratory studies both in house and under contract at Westinghouse Canada Lim ited [ 18]. Standardized equipment has been developed to study the relation ship between low level fretting and various com binations of orthogonal periodic excitation forces, tube acceleration and support reaction force, for various typical arrangements of tube and support geometry and material, at room temperature and at boiler operating temperature, in water and in steam. Reproducible test results have been elusive and the building up o f trends and confidence lim its has proved laborious. Our steam generator manufacturers have avoided plain holes in tube support plates in the boiling region. We have therefore made particular study of the fretting propensity of the various form s of ventilated or interrupted geometries that they have proposed. M ost of this information is proprietary. However we release here for the first time some basic data related to interrupted circular holes. Figure 6 gives the average local penetration (derived from overall weight loss) over the support area versus the fraction of perimeter of the original circular hole remaining. The test specimens were made from 12.5 m m carbon steel plate, drilled to give a central hole of ~ 12.5 m m diameter together with three other equispaced holes on the circumference of the first. A n Inconel 600 tube specimen was oscillated at 30.5 Hz in each plate specimen under saturated water at ~ 260°C. The excitation force was constant, about 7.6 N rms with an orthogonal force ratio of 1.73 applied to the external end of the beam passing through the autoclave gland and carrying the tube specimen (i.e. an unrestrained tube would oscillate in an elliptical orbit). The rms support reaction force varies with support area, with clearance and with excitation frequency, in any particular a u to c la v e . To assist in applying the empirical fretting correlations to the vibratory response calculated using PIPEA U , we have studied the nature of the tube-plate contact both theoretically at the University of Waterloo under contract [19] and experimentally at the Chalk River Nuclear Laboratories. Figure 7 shows results of both kinds, with the theoretical study m odelling the operating conditions of the room-temperature cantilever test rig. The computer model is being extended to handle multiply-supported beams, which should provide the final step required in the overall flow-vibration response-fretting appraisal technique. The computer model predicts abrupt but not complete loss of contact at about 1 m m diametral clearance, which is the major diameter of the elliptical orbit that the unrestrained tube would follow. These tests were carried out at room temperature using an Inconel 600 tube of ~ 12.7 mm diameter and plain circular hole support plates in carbon steel; excitation forces are normalized to approximately 1.9 N rms. W ith reaction force increasing with clearance, and fretting rate proportional to clearance, one would expect constant excitation fretting rate to increase with 192 HODGE et al.
RELATIVE EXCITATION FORCE
DIAMETRAL CLEARANCE (mm)
FIG. 7. Reaction force in plain circular supports.
time of exposure. Periodic eddy current monitoring of a test in progress has shown an increase to occur in the first thousand hours or so but a steady rate was attained at an unexpectedly low value (~ 150 pm) of total wear (i.e. both support and tube).
9. IN-SERVICE INSPECTION
World-wide problems with nuclear power station steam generators have stimulated the development of improved equipment and procedures for inspecting tubes during plant shutdowns. Specifically, several tube-sheet walking devices have been developed which carry eddy current probes from tube to tube without manual intervention. However, the equipment is still located in the boiler room and the output signals are recorded on paper charts and analogue tape recorders. Record keeping and analysis are for the most part laborious manual tasks. Although C A N D U steam generator tubes have performed extremely well to date and have not followed the world trend, our major utility user asked us to provide a reliable high speed system for inspecting the Pickering steam generator tubes. This new eddy current inspection system is called C A N SC A N , and is being jointly developed by the Chalk River Nuclear Laboratories and Bristol Aerospace in Winnipeg. Delivery is scheduled for December 1978. C A N SC A N differs from other tube inspection systems in two important respects. Firstly, control and data processing are both digital; a microprocessor is used for automatic control of the mechanical components and a mini-computer is used for processing the eddy current data. Secondly, the equipment in the IAEA-SM-218/25 193
FIG. 8. Block diagram of the CANSCAN system. boiler room does not require full time operators. After the tube-sheet walking device has been installed inside the head of the steam generator, the entire system is controlled remotely from a trailer outside the reactor building. A single coaxial cable 300 m long, carries all of the control signals and inspection data for two eddy-current probes operating simultaneously.
Figure 8 is a block diagram of the C A N SC A N system, showing the basic components inside and outside the reactor building. Figure 9 shows the tube- 1 9 4 HODGE et al.
FIG.9. Tube-sheet walking probe carrier. F IG .l 0. P robe drive.
sheet walker, designed to traverse the tube-sheet lattice in a step-by-step manner carrying two eddy current probes and cables. It is powered by an air m otor and controlled automatically by a microprocessor. The device weighs approxi mately 1 kg and requires only one manual intervention to cover 97% of the tubes. Because reliability of the tube-sheet walker is absolutely essential, a more robust device weighing 5 kg has been developed as a back-up. The probe drive system was developed by the Teleflex Corp. USA; it consists o f a hollow core flexible steel cable wrapped with a wire helix which engages a hobbed gear. In Fig. 10 the cable has been displaced to reveal the hobbed gear. In Fig.l 0 the cable has been displaced to reveal the hobbed gear. IÄEA-SM-218/25 195
The 30 m long cable is motor-driven at a maximum speed of 50 cm/s, and when constrained in a steam generator tube, can exert a thrust of 180 N. The eddy- current signals are carried on insulated wires inside the drive cable, which is stored on a rotating drum equipped with slip rings. The eddy-current probe incorporates a permanent magnet to saturate the tube material (Monel-400), and an internal reference coil to eliminate temperature effects. Operating in the absolute mode, it can detect flaws equivalent to a 0.3 mm hole. The eddy-current instrument is a commercial model modified for improved frequency response and remote control. Tw o test frequencies will be used for routine inspection, 50 kHz and 100 kHz. Analogue output signals from the four eddy-current instruments (two for each probe) are converted to digital pulses in a portable instrument package located near the steam generator being in sp e c te d .
The trailer contains a control console, a mini-computer and output devices for graphic display and printing. Data processing and system m onitoring are the main functions of the computer. It comprises a dual central processing unit with 96 К semi-conductor memory, dual discs, dual tape decks, C R T display and electrostatic printer plotter. Ideally the computer output would be a printed list, identifying defective tubes and giving the location and nature of each defect. However, the present goal is less ambitious, namely, to identify and display all eddy current anomalies in a convenient form for visual analysis, assuming that anomalies are a small fraction of the data (e.g. < 1%). The analyst then decides which anomalies are unacceptable defects requiring immediate action. He can also decide how much data should be permanently stored on magnetic tape. A t present, calibration signals are obtained from artificial defects (i.e. holes, machined notches, etc.) rather than real defects. After a significant number of real defects has been observed, it should be possible to program the computer to classify defects according to severity. Ultim ately it m ay be possible to estimate the operating lifetime o f doubtful tubes, thereby enabling the station operator to schedule tube plugging in a convenient and rational manner.
10. SUMMARY
The work described in this paper covers our efforts over the past three years to maintain our current record o f reliability in the face of design economies, particularly those o f increased size and heat rating. This effort has enlarged our analytical and empirical knowledge of the thermal, fluid, chemical and mechanical behaviour. A fully automated system for in-service inspection of steam generator tubing significantly faster than equipment available elsewhere is in an advanced stage of development. 196 HODGE et al.
REFERENCES
[1] HODGE, R.I., et al., “Steam generator reliability - the Canadian approach”, XIX Nuclear Congress of Rome, Italy, March 1974. Also available as Atomic Energy of Canada Limited Report No.AECL-4771. [2] HARE, M.G., “Steam generator tube failures: World experience in water-cooled nuclear power reactors in 1975”, AECL-5625 (1976). [3] LeSURF, J.E., et al., “Radiation dose reduction programs for CANDU nuclear power stations”, paper presented at the 10th World Energy Conference, Istanbul, Turkey (19-24 Sept. 1977). [4] LISTER, D.H., “The accumulation of radioactive corrosion products in nuclear steam generators”, Paper 102, CORROSION/76, Houston, Texas (March 1976). Also available as AECL-5422. [5] LeSURF, J.E., Control of radiation exposures at CANDU nuclear power stations, J. BNES 16 1 (Jan. 1977) 53-61. Also available as AECL-5643. [6] SCHENK, H.J., “Investigation of tube failures in Inconel-600 steam generator tubing at KWO Obrigheim”, CORROSION/75, Toronto, Canada. Published in Materials Performance 15 3 (March 1976) 25. [7] BERGE, Ph., et al., “Caustic stress corrosion of Fe-Cr-Ni austenitic alloys”, 6th Inter national Congress on Metallic Corrosion, Sydney, Australia (Dec. 1975). [8] BLANCHET, J., et al., “Corrosion et fissuration sous contrainte d’alliages du type Incoloy-800 dans l’eau et la vapeur ä haute temperature”, NUCLEX’75, Basel, Switzerland (October 1975). [9] PATHANIA, R.S., “Caustic cracking of steam generator tube materials”, Paper 98, CORROSION/76, Houston, Texas (March 1976). Also available as AECL-5453. [10] HARE, M.G., STEVENS-GUILLE, P.D., “Steam generators: the long road to maturity”, European Nuclear Conference on Nuclear Energy Maturity, Paris, France (1975). Also available as AECL-5089. [11] SMITH, R., “Steam generators to be replaced at Surrey, Turkey Point nuclear units”, Nucleonics Week 18 (14 July 1977) 24. [12] BALAKRISHNAN, P.V., McVEY, E.G., “Model boiler studies on deposition and corrosion”, presented at NACE Symposium on Corrosion in the Power Industry, Montreal, Canada (27—29 September 1977). Also available as AECL-5801. [ 13] “Denting in Ringhals has led to the plugging of 72 tubes”,Nucleonics Week 18 (12 July 1977) 24. [14] BALAKRISHNAN, P.V., “Effect of condenser water in-leakage on steam generator water chemistry”, presented as Paper IWC-77-12 at the 38th Annual International Water Conference, Pittsburgh, USA (1—3 Nov. 1977). Also available as AECL-5849. [15] BALAKRISHNAN, P.V., “A study of phosphate hide-out from boiling water”, accepted for publication by Canadian Journal of Chemical Engineering. [16] PETTIGREW, M.J., et al., “Flow-induced vibration analysis of heat exchanger and steam generator designs”, 4th International Conference on Structural Mechanics in Reactor Technology, San Francisco,California, 15-19 August 1977. Also available as AECL-5826. [17] GORMAN, D.J., “Experimental study of peripheral problems related to liquid flow induced vibration in heat exchangers and steam generators” , Structural Mechanics in Reactor Technology (Proc. 4th Int. Conf. San Francisco, 1977) Paper F4/g.
[ 18] KO, P.L., “Fundamental studies of tube fretting in steam generators and heat exchangers, 6th Canadian Congress of Applied Mechanics, Vancouver, B.C.,May 30 — June 3, 1977. IAEA-SM-218/25 197
[19] ROGERS, R.J., PICK, R.J., “Experimental and simulated studies of a 3-dimensional heat exchanger tube with baffle interaction”, 4th International Conference on Structural Mechanics in Reactor Technology, San Francisco, California, 15—19 August 1977.
DISCUSSION
M. de HES: It is important for reliability engineers to have extensive inform ation on failure rates for steam generator tube material and tube-to-tube plate welds. However, there is a lack of such information for the liquid metal fast breeder reactor (LM F B R ), so that the failure rates obtained for water reactor steam generators will for the time being have to be used for reference purposes in connection with LM FBRs. Information reported by M.G. Hare gives relatively unfavourable tube and weld failure rates. However, the data in your paper relating to Canadian plants show m uch more favourable failure rates. M ay I ask you what mathematical model you used to define the value of one failure in 3.1 X 107 m years mentioned in your Introduction? What is the confidence level for this value? I should also like to know why you mention three faults in your abstract, whereas you discuss only one fault in your failure rate analysis. And lastly, can you tell me what the failure rate (X) is with a 90% confidence level for Canadian steam generators — that is, for both tube material and tube-to-tube welds? R.I. H O D G E: For specific answers to your questions I would refer you to our 1974 paper (Ref.[ 1 ]). The three failures we have experienced — one fretted tube in the N P D boiler, several fretted tubes (one penetrated) in a Douglas Point boiler and one tube which was though to contain a manufacturing defect causing a leak in a Pickering boiler — are all described in Ref.[ 1 ]. We have experienced no further leaks since January 1974. I have not attempted a statistical analysis of these failures, but the value I quoted is the product of operating hours and tube length in the Pickering steam generators. M. de HES: Do you think we could use Canadian figures for preliminary L M F B R analysis, or should we use more optimistic figures? R.I. H O D G E: I would say that the faults we have experienced have little relevance to an L M F B R heat exchanger since their origin is in the design and operating environment.
IAEA-SM-218/27
RELIABILITY CONSIDERATIONS FOR LMFBR STEAM GENERATORS
G.A. de BO ER, M. de HES, P.W.P.H. LUDW IG
B.V. NERATOOM, T h e H a g u e , The Netherlands
A b stra c t
RELIABILITY CONSIDERATIONS FOR LMFBR STEAM GENERATORS. LMFBRs require a high availability and thus high reliability of the components. The steam generator seems to be a critical component with respect to availability. Since sodium reacts violently with water, even a small leak, which can grow to a large leak within a short time, is unacceptable. It is shown that the sodium/water reaction has only availability consequences; it is not a nuclear safety problem. An optimum number of steam generators per loop is calculated in order to improve availability. Only a few statistical data are available on the reliability of sodium-heated steam generators. These data relate to prototype and first-of-a-kind steam generators. A survey is given of the causes of leaks and the measures to prevent the leaks. It includes system and leak detection considerations, component design and fabrication aspects. The manufacturing of tubes and tube connections is detailed, as well as material selection, including corrosion effects. It is concluded that steam generators can be designed, built and operated to ensure a high degree of reliability.
1. I N T R O D U C T I O N
Liquid Metal Fast Breeder Reactors (LM FBRs), with their high investment costs and low fuel cycle costs, require high availability and thus high reliability of the components. Experience with steam generators in several countries shows that the steam generators are critical pressure components with respect to availability. In the case of a tube leak the steam generator has to be repaired, replaced or isolated; this is a tim e-consum ing operation. In several cases it has proved very difficult even to locate the leak; the leak can open during operation and close again when the steam generator cools down, m aking location and repair impossible. In the case of a large leak, extensive cleaning of the secondary circuit will be necessary, during which the plant or at least the circuit cannot be operated, thus causing a loss in production. Even after successful leak localization and repair the question remains whether secondary damage has occurred or not. Requalification is necessary before operation can be resumed. A major design objective is thus the prevention of leaks in steam generators. To improve the
199 200 de BOER et al. availability of the steam generators N E R A T O O M has developed a design philosophy for the steam generator and the steam generator system and is carrying out R and D programmes. These activities are described in the following sections. In our opinion steam generators can be designed and manufactured to the same standards of reliability as the other main com ponents of the heat transfer system of LM FBRs.
2. MAIN FACTORS AFFECTING THE RELIABILITY OF STEAM GENERATORS FOR LMFBRs
The steam generators for L M F B R s are usually operated at a so-called “conventional” pressure and temperature of 170 bar and 500°C. The sodium is at low pressure (10 bar) and its excellent heat transfer properties result in compact designs. Tube wall thickness is between 2 and 4.5 mm. As usual in heat exchangers, leaks from one medium to the other are the cause of unavailability. In this sodium/water heat exchanger the leaks are accompanied by a violent chemical reaction, which in turn can cause other defects, for example deformations caused by the high pressure or corrosion due to the chemicals formed. Leaks in the outer shell of the steam generator result only in a sodium leakage. The probability of this kind of leak is estimated to be extremely low, since the wall thickness is high, the surface is relatively small, and sodium corrosion rates are very low. These leaks can be compared with leaks in the piping system and are not discussed in this paper.
Causes of leaks in steam generator tubes are:
opening of inherent defects of the tubes opening of inherent defects in the tube connections rupture due to high stresses rupture due to fatigue tube erosion by sodium or water tube corrosion by sodium or water tube fretting due to vibration contact in tube supports.
These causes can result in small, m edium or large leaks. Sm all leaks may grow by self-wastage to larger sizes or cause leaks in adjacent tubes by impingement wastage. Medium-size leaks cause a mild sodium/water reaction, which does not burst the rupture disc. A large leak will be followed by burst of the rupture disc. It m ay be concluded from this that there is a major difference from, for example, steam generators in PW Rs. IAEA-SM-218/27 201
3. NUMBER OF LOOPS, REPAIR PHILOSOPHY AND NUMBER OF STEAM
GENERATORS PER LOOP
The availability of the nuclear power plant is not only affected by the reliability of the steam generators, but also by the configuration in which they are used. The main factors influencing the configuration and thus the availability are discussed below.
3.1. The number of loops
The choice of an L M F B R ’s optimal number of loops depends, amongst other factors, on:
thermal power (e.g. SN R 300 with ~ 750 MW (th) and SN R 2 with
~ 3500 MW(th) decay — heat removal scheme (via or not via secondary loops) plant symmetry max. flow velocity in sodium pipes; 6 n rs-1 seems at the moment to be the generally adopted m axim um velocity determining the diameter of the main piping (SN R 300 with 500 to 600 mm and SN R 2 with 900 to 1000 mm/diam.).
Based on these requirements the number of primary and secondary loops of SN R 2 has been chosen as four.
3.2. The repair philosophy
For SN R 300 the repair philosophy is simple and straightforward: repair of a possibly faulty component in situ is not considered; the component which has failed will be replaced by a spare component. For SN R 2 the repair philosophy has not yet been decided but the following three possibilities are now being studied.
(a) The faulty component is repaired in situ with the plant in 3-loop operation. (b) The faulty component is replaced by a spare one, with the plant in 3-loop o p e r a tio n .
(c) The faulty component is isolated from the sodium and steam system by isolation valves. The plant stays in 4-loop operation. The defective unit is replaced by a spare one during the following shut-down period.
In the case of a large leak repair procedure (a) will not be used, since secondary damage is then extensive and repair and requalification impossible. 202 de BOER et al.
a d d i t i o n a l g e n e r a t i n g
FIG.l. Influence of repair concept, leak probability and unit size on the power-generating costs.
3.3. Number of steam generators per loop
The choice of the number of steam generators per loop depends primarily on the repair philosophy which is adopted for the plant. In calculating the optimal number of steam generators for SN R 2 the following points were considered:
Cost analysis shows a rather low decrease of the cost per unit power by increasing the steam generators’ unit size.
H 2 leak detection and countermeasures are conducted so fast that no damage to neighbouring tubes occurs. Thus the possibility of a water/sodium reaction which actuates a rupture disc is not considered in the evaluation.
Since hydrogen leak detection is a concentration measurement there is a sodium mass flow above which it is not easy to recognize a micro-leak. As a result of a cost optimization with respect to m inim um plant kW 'h costs, the economic number of steam generators per loop for SN R 2 should be 4 to 8. This applies to all 3 main repair philosophies and is nearly independent of the assumed IAEA-SM-218/27 2 0 3 number of steam generator failures per year. It is apparent from Fig. 1 (see also Ref. [1 ]) that little or nothing is gained by increasing the unit size, even when the reliability is assumed to be very high.
4. DESIGNING FOR RELIABILITY
4.1. General
There are several reasons for unavailability of steam generators. The unavailability results from:
Repair + requalification or replacement or isolation + later replacement of steam generator. Here the leak is detected and located.
Indication of leaks which cannot be located, for example due to the fact that
the leak is not continuously open; the leak does not exist; the leak detection signal was false.
The following measures have to be taken to improve the availability of the steam generator:
The number of leaks during the steam generator life has to be minimal.
It has to be guaranteed that nuclear safety is not affected by any large leak in the steam generator.
The necessity to guarantee that during their working life only a very low number of leaks will occur has its consequences on the design, the fabrication and the operation of steam generators. The design of steam generators could be simplified if a guarantee could be given that nuclear safety is not affected by any large leak in the steam generator. This allows the designer to concentrate on the reliability of the steam generators. Extensive studies in the last few years have proved that it is possible to protect the intermediate heat exchanger against any pressure transient com ing from the steam generator with an adequate system design. Basically, two relief tanks in the hot and cold leg of the secondary system are needed. One of these tanks can be the expansion tank, the other has to be isolated from the system by a rupture disc [1]. 2 0 4 de BOER et al.
4.2. Design of steam generators
4.2.1. Prototype components
The enlargement of scale makes it difficult to acquire experience step by step, as was done in the past with “conventional” equipment. This was recognized in the SN R 300 project and early in the development programme it was decided to design, build and test full size prototypical components. The main heat transfer com ponents were tested for at least 3000 h before they were dismantled and inspected. In doing this, valuable information was obtained for the designer and manufacturer [see Appendix]. The structural integrity was demonstrated. The testing time was in the order of magnitude of the first part of an estimated “bath-tub” curve.
4.2.2. System consideration
It will be clear from the aforementioned facts that leak detection in steam generators is a major design criterion. The leak detection must take place using a reliable method. False signals are probably worse than true leak signals. In designing the steam generators system im portant criteria are derived from the capabilities of the leak detection instruments. Several instruments are being d e v e lo p e d :
chemical ones: hydrogen and oxygen concentration measurements;
acoustical ones which are triggered by the sound of the flow through the leak and/or by the noise of the reaction.
The possibility of locating a leak acoustically is also being studied. A currently em ployed criterion, based on the sensitivity of the H 2 detecting system (a concentra tion measurement) is that a unit of about 200 M W can be monitored with one H 2 detecting device. However, it is possible to build several such 200 M W units in one shell. Further system considerations deal with items such as:
blow down system for reaction products;
interaction of units during a sodium/water reactions;
easy clean up after a large sodium/water reaction. IAEA-SM-218/27 2 0 5
BRAZED INTERNAL BORE WELD (nicro braze)
FIG.2. Tube to tubeplate joints.
4.3. Component considerations
In the design of the steam generator units it is of great importance to avoid the causes of leaks mentioned in Section 2, because they affect the availability of a nuclear generating station and its energy costs. Principle design criteria s h o u ld be:
availability determined by reliability and ease of maintenance
s a fe ty
c o st
performance depending on factors concerning hydraulics, heat transfer, fabrication, inspection and material properties such as strength and corrosion resistance.
It will be clear that none of these factors can be optimized without affecting the others. The designer has to weigh all aspects before m aking his final choice. The optim um steam generator includes several near-optimum solutions for these fa c to rs. 2 0 6 de BOER et al.
^ 7 Т 7 7 Ш /Л
У//////////Л У У / У У /л S i ______L film
TZZZZZZZZZZE
FIG.3. Micro-focus anode ray rod.
All thermal-hydraulic aspects must be fully understood and must be accessible for computer modelling. This is particularly true for the boiling process. Experim ental verification has to be used to validate the computer models, thereby increasing the confidence one can have in the computer prediction of for example operational stresses during transients. Transient analysis is a major design activity since transients can cause high loads during operation [3]. By properly elaborating the design, failures due to high stresses, fatigue, vibration, fretting or erosion can be avoided. It is beyond the scope of this paper to analyse current designs as given in Fig. 5 on the above mentioned points.
4.4. Fabrication
Fabrication includes all aspects from the manufacture of raw materials, such as tubes and forged pieces, up to delivery of the component. Only two aspects will be detailed here: tubes and tube-connections.
4.4.1. Tubes
Tube manufacturing demands high standards [4]. However, development work in cooperation with tube manufacturers will be necessary to achieve guaranteed reliability of the tubes. Until now little long-term experience is available and it will take some time until it will be available. Strengthening of the acceptance criteria is not the only way; analytical studies must give a better understanding of the origin of possible defects and the way to prevent them. Also studies of the behaviour of tubes under working conditions, with varying temperatures and stresses, must give inform ation on the possible influence of m inor defects which have been accepted until now. IAEA-SM-218/27 2 0 7
FIG.4. Thermal sleeve penetration.
The data generated during the acceptance test on defects from the manufacturing process indicate a high quality, but the tests which are carried out cannot exclude a micro-failure. The tube-manufacturing process, from the raw material to the finished tubes, must be studied in detail in order to exclude every possible cause for defects. Tube deformation during the fabrication process, which is necessary for example to fabricate helical coil tube bundles, m ay affect the structural integrity of tubes. There is little inform ation on the influence of the plastic deform ation on tube integrity, since it is difficult to repeat all the tests which are done on the straight tube. It can be understood that some designers prefer straight tubes in their designs. However, excellent operational results are also achieved with helical tube steam generators. Som e designers of steam generators favour duplex tubes consisting of two concentric tubes with metallic or mechanical bonding between the tubes. With respect to the E B R -II reactor there is already more than 12 years’ experience with this type of steam generator [5]. From the point of view of availability it is clear that a leak detection system which would detect leaks through one of the tubes and cause plant shut-down too is not desirable. False signals will cause undesired outages. Within the R and D programme the applicability of duplex tube steam generators with respect to economics and availability is being studied. Heat transfer tests on non-bonded duplex tubes have already been carried out by N ER A T O O M , showing rather good heat transfer if the gap between the tubes is filled with gas.
4.4.2. Tube connections
Tube joints in the sodium can be entirely omitted, as is done in some designs. So there are no leaks from defective joints, but this can also lower the 2 0 8 de BOER et al.
FIG.5a. Steam generator, straight-tube design.
degree of freedom of the designer to such an extent that, for example, thermal hydraulic solutions have to be adopted that are far from optimum. In straight tube designs as well as in U-tube and helical tube designs the tube-to-tube sheet connection is often used. Several variations are practised (Figs 2, 4) [6 ]. NERATOO M and the two factories STORK and DE SCHELDE have over the last year built up wide experience with the so-called internal bore weld [7]. This weld is made fully automatically and its inspection is easily possible by a special developed micro-focus X-ray tube (Fig. 3) [8 ]. A p a r tic u la r feature is that the image of the weld is enlarged by this process, resulting in a better possibility of inspection. IAEA-SM-218/27 2 0 9
FIG.5b. Steam generator, helical-coiled design.
Not only radiography is used as a non-destructive inspection technique. Others are: dye penetrant, visual inspection using horoscopes and helium leak testing. The destructive tests, which are executed on specimens during the fabrication, complete this extensive testing programme [9]. Apart from welding, brazing can be used to join tubes to the tube sheet. It has been used for the prototype straight tube superheater, which is still in operation in the test facility at Hengelo. This paper does not allow for a comparison of welded and brazed tube-to-tube sheet connections. The unfavourable heat treatment of the tube bundle during the brazing process was the reason that this method was not adopted for the SN R 300 components. 210 de BOER et al.
4.5. Material selection
4.5.1. Selection criteria
The material selection criteria are divided into two broad categories, based on the desired performance objectives and desired material properties. Basically the prim ary selection criteria cannot be influenced by any conservatism in design. Prim ary selection criteria are:
requirements and specifications of the nuclear plant licensing authority;
significant resistance to stress corrosion cracking in caustic- and chloride- contaminated environments;
possibility of fabrication, especially good welding properties.
The secondary selection criteria are either less critical to the overall component performance or involve uncertainties which can be accommodated by conservative design. Exam ples of secondary criteria are:
mechanical properties;
wastage resistance;
corrosion resistance to water, steam and sodium.
Apart from the 2£Cr-lM oN i stabilized ferritic steel which is used for the steam generators of the SN R 300 two candidate alternative materials are considered for the next generation steam generators: 9% and 12% chromium steel with addition of stabilizing elements. These groups of alloys are considerably stronger than the lower chrom ium steels and comparable with austenitic steel 316. A difficulty with the 9% and 12% chromium steels is the fact that an accurate pre- and post-heat treatment is necessary in the welding procedures to prevent structure hardening and, related to this, cracking of these steels is a potential p r o b le m .
4.5.2. Corrosion
Test programmes have shown that sodium corrosion is not a problem. The behaviour of steels in flow ing sodium can accurately be described. Waterside corrosion is more complex. The data from waterside corrosion in the recirculation type steam generator of LW Rs are self-explanatory. However, the situation in L M F B R steam generators is different. The medium outside the IAEA-SM-218/27 211 tubes is sodium with well-known and easily controllable corrosion properties. The steam generators are preferably of the once-through type. The feedwater quality requirements can easily be met, as is demonstrated by several fossil- fired once-through boilers in operation. However, short-term deviations in the quality prescribed can occur and m ay have unwanted consequences. To be able to guarantee adequate steam generator operation, it will be necessary to prescribe that the feedwater treatment installation meets certain reliability criteria and is, for example, designed so as to be able to cope with a certain number of tube ruptures in the condenser.
5. CONCLUSIONS
We are confident that it is possible to minimize the number of failures either by obviating or by counteracting the causes. The operation of prototype steam generators has given the necessary opportunities for “learning”. The operation of the steam generators in the SN R 300 can be looked forward to with confidence.
REFERENCES
[1] LIEVENSE, K., et al., Design considerations and cost analysis of the secondary system of SNR 2, BNES Conf. London, 30 Oct. - 3. Nov. 1977. [2] van WESTENBRUGGE, J.K., et al., Results of the test and inspection program of the prototype SNR 300 steam generators, BNES Conference, London, 30 May - 2 June 1977. [3] van LEEUWEN, N.J.M., et al., Verification of dynamic steam generator model, Nat. Heat Transfer Conf., Salt Lake City, 1977 . [4] LARSON, B., et al., Manufacture of ferritic steel tubes for fast reactor steam generators, BNES Conference, London, 30 May - 2 June 1977. [5] CHOPRA, P.S., On the fail-safety of LMFBR steam generator systems, 4th Int. Conf. Structural Mechanics in Reactor Technology, San Francisco, 15-19 Aug. 1977, Paper E5/3. [6] de RAAD, J.A., Ultrasonic and other non-destructive testing methods for tube joints used . in heat exchangers, Int. Symp. Non-Destructive Testing of Nuclear Power Reactor Components. Rotterdam, February 1970. [7] de BLIECK, T., et al., “Non-destructive examination of tube to tubeplate connections”, 6th Int. Conf. on Nondestructive Testing, Hanover, 1970, Rep.No. F6. [8] de BLIECK, T., Methods of inspection adapted to new construction details in highly stressed pressure vessels, Conf. Research Sodium Cooled Fast Breeder Reactors, Petten, Oct. 3, 1969. [9] BEIJER, N.H.H., Welding experience with tube-to-tubesheet joints of steam generators in 2 \ Cr — 1 Mo — Ni — Nb ferritic steel, Int. Conf. on Ferritic Steels for Fast Reactor Steam Generators, London, 30 May — 2 June 1977. [10] de CLERCQ, W.J.C., et al., The development of sodium-heated steam generators in the Netherlands, BNES Conf., London, March 1974, Specialist Session I. 2 1 2 de BOER et al.
[11 ] MANTE, A.H.J., et al., Operational experience with main circuit components for the SNR 300 Power Plant, BNES Conf., March 1974, London, Specialist Session I. [12] de CLERCQ, W.J.C., Sodium-water reaction in the reheater of the 50 MW test facility at Hengelo, Conf. Engineering of Fast Reactors for Safe and Reliable Operation, Karlsruhe, Oct. 9-13, 1972, 3 (1972) 1305. [13] de BOER, G.A., et al., Choice of steam cycle for sodium-cooled reactor , Proceedings IAEA Symp. Sodium-cooled Fast Reactor Engineering, Monaco, March 1970, IAEA, Vienna (1970) 130. [14] BRASZ, J., Experimental determination of reliability data, these Proceedings, Paper IAEA-SM-218/57.
APPENDIX
TEST OF PROTOTYPE COMPONENTS [2, 10-14]
G e n e r a l
Rather early in the design stage of SN R 300 it was decided to build and test full-size prototypical components. The steam generator components were tested for at least 3000 h, and some for much longer periods. All tests took place under simulated reactor conditions including several transients, e.g. a reactor scram simulation. The testing gave valuable design information, but hardly any statistical data which can be used in a reliability analysis.
Prototype of sodium-heated 50 M W straight-tube steam generator
This steam generator consists of a 26 M W evaporator, a 16 M W superheater and a 7 M W reheater. All these components are of the straight-tube type with welded tube-to-tube sheet connections in the evaporator and reheater, and brazed ones in the superheater.
Test with the straight-tube reheater
The testing of the reheater lasted only 350 hours. A (medium) sodium/water reaction was detected with the instrumentation (hydrogen monitors and flow meters) during the start up and thus low-pressure operation. Plant shut-down was realized before the rupture discs burst. When this unit was dismantled, it was found that the medium leak ( 1 —2 m m) was caused by impingement tube wastage from a micro-leak in the weld of a neighbouring tube. The original leak was a defect in the weld that was not found during inspection. N o further tests with sodium-heated reheaters were carried out since they are no longer incorporated
in the SN R 300 system. IAEA-SM-218/27 2 1 3 test with the straight-tube evaporator
The test of the evaporator lasted 3659 h. During the test period a plant shut-down was performed, due to a rise in the measured hydrogen level. During several m onths an intensive leak investigation was carried out using m any techniques. However, no leak was found. A detailed analysis of the 50 M W test circuit’s behaviour during the weeks before the leak indication showed that hydrogen disappeared in the loop at the same time as a rapid decrease in temperature of the air cooler occurred (below the stalling temperature, 100°C). In this process condition the air cooler could serve as a kind of diffusion cold trap. It was found also that the air cooler’s temperature increased during the plant’s hydrogen increase. Tests of the 50 M W test circuit under the same conditions proved the validity of the hide-out theory. After these events the steam generator was operated successfully over more than 1700 hours, up to a total of 3659 hours. After finishing the defined test programme the component was dismantled for inspection. A s a result of the operation of this prototype some m inor m odifications were incorporated in the SN R 300 straight-tube steam generator design.
Test with the straight-tube superheater
This com ponent was tested during 3659 hours as a superheater and after the dism antling of the straight-tube evaporator the superheater was operated also as an evaporator. Till 16.11.1975 this component was tested for 3337 hours as an evaporator, giving a total of 6996 hours of testing. This superheater com ponent is still in the test circuit and is used occasionally for additional tests.
Test with 50 M W steam generator (helical tube design)
A prototype of the 50 M W (th) helical tube steam generator was tested for 3379 hours under normal evaporator conditions. After carrying out the tests, the steam generator was dismantled for inspection purposes. Hereafter it was reinstalled in the 50 M W test facility for continuing therm ohydraulic tests. In this second test period the com ponent was also operated under once-through conditions with 500°C steam at the outlet. As per 1 July 1977 this component has been tested a total o f 9577 hours.
M ain test results
The main test results and statistics are summarized in Tables I and II. Extrapolation of the test results to an L M F B R is difficult. The SN R 300 life time is 200 000 h or 30 times the test period and the SN R rated power is 4 1 2 TABLE I. DATA OBTAINED FROM TEST RESULTS PROTOTYPES FOR STRAIGHT-TUBE SG AND HELICAL- C O I L E D S G
50 MW straight-tube steam generator Straight-tube Straight-tube Straight-tube 50 MW helical evaporator superheater reheater coiled-tube SG
100% power (MW) 26.35 16.5 7.15 52.75 T primary in (°C) 441.7 529.7 529.7 445 T primary out (°C) 343.5 442.1 440.6 334 T secondary in (°C) 288.8 353.9 325.0 250 T secondary out (°C) 355.2 510.0 1198.9 390 al. et BOER de M primary (kg’s '1) 210.5 Na 147.9 Na 62.6 Na 371 Na M secondary (kg's'1) 21.85 H20 20.72 H20 16.9 H20 30.33 H20 P outlet (bar) 176.3 168.8 41.5 179.7 x outlet (%) 95 100 100 100
Total test time (h) 3659 3659 (6996*) 350 3379 (9577**) Test time nominal (h) 3003 3003 (5573*) - 3268 (9209**) Data per 28.6.74 28.6.74 31.8.72 * including added ** tests are continued test as evaporator! in 1976-1977 Data per * 16.11.75 ** 1.7.77 No. of tubes 139 139 139 100 total tubes length (m) 2688 2026 1246 4200 No. of welds 278 278 278 400 TABLE II. STATISTICS EVALUATED FOR TESTED STEAM GENERATORS OF PROTOTYPES (details given in Table I ) Upper bounc s for X T = time of N = number A! = a 2 - T.Ai = T.A2 = X 1 X 2 testing of failures tube 1 number of h.m. h. weld h-1 • m“1 h“1 'weld-1 (h) (m) welds IAEA-SM-218/27
Helical EVA 9577 0 4200 400 40.2 X 106 3.83 X 106 0.06 X 10~6 0.6 X 10~6 Straight EVA 3659 0 2688 278 9.8 X 106 1.0 X 106 0.23 X 10'6 2.3 X 10-6 Straight SUP 6996 0 2026 278 14.2 X 106 1.9 X 106 0.16 X 10‘6 1.2 X 10-6 Straight REH 350 1 1246 278 0.4 X 106 0.10 X 106 8.9 X 10“6 40 X 10~6 DT.Ai and 2T.A 2 64.6 X 106 6.83 X 106 Mean value 1 and 2 0.036 X 10~6 0.57 X 10~6
The given X are calculated with a one-side 90% confidence interval (according to the chi-squared test).
GO 216 de BOER et al.
~ 750 MW (th) compared with the 50 MW (th) of the test-facility. The ratio: (lifetime X heated surface of SN R 300) to (test period X heated surface 50 M W loop) is 450. (SN R 300 components design lifetime: 10s h). The only steam generator failure is a failed weld in the reheater, a component not used currently for SN R 300. Also an improvement in inspection techniques guarantees that such a defect will easily be detected, and will not occur again. On the problem of extrapolating relatively few data from tests of prototype components, to L M F B R data, see Ref. [14].
DISCUSSION
S.T. N A ISH : You mentioned leak testing as a method of testing welds. Can you say what technique is used and how effective it is in detecting weld faults? Very small leaks are easily blocked and may not be detected unless measures are taken to ensure high standards of cleanliness. G.A. de BO ER: Shortly after a limited number of tubes have been welded they are individually leak tested by the helium leak testing method. Special equipment is used to seal the bore of the tube and also the hole in the tube sheet. The volume between the two plugs is connected to the leak testing equipment and is evacuated. Helium is applied to the outside of the weld. The sensitivity is 1СГ 8 torr/litre/second. So far no leaks have been found. By carrying out the leak test shortly after welding one can be sure that the weld is still very clean. D. van RO O YEN : Do you anticipate any corrosion problems? Also, are you carrying out any specific tests for corrosion behaviour at the m om ent? G.A. de BO ER: Corrosion on the sodium side of the tubes is very slight and it is well understood. It is not expected that there will be more corrosion in the sodium-heated steam generators than in conventional once-through boilers. A magnetite, layer will be form ed on the inside of the ferritic tubes which will give adequate protection. The quality of the feed water will have to meet require ments equal to those which are specified for conventional once-through boilers, but experience has shown that these requirements can be met. Special attention will have to be paid to condenser m onitoring and any leak in the condenser must be detected quickly and repaired in a short time. Corrosion behaviour has been studied during the operation of the prototype components and in special loops such as one-tube steam generators. The form ation of the magnetite layer can be studied by measuring the hydrogen concentration in the steam and in the sodium. After 3000 h of operation the magnetite layer has been examined destructively and has been found to be sound. The testing of the prototype helical tube steam generator will be continued for more than 10 0 0 0 h in order to obtain long-term data. IAEA-SM-218/26
PWR STEAM GENERATOR TUBING
Corrosion problem s
D. van RO OYEN Department of Applied Science, Brookhaven National Laboratory, Upton, New York, United States of America
A b stra c t
PWR STEAM GENERATOR TUBING: CORROSION PROBLEMS. The corrosion problems that have been experienced with tubing in recirculating steam generators are summarized, indicating the extent to which phosphate wastage, caustic cracking, fatigue, denting, and primary side stress corrosion cracking have contributed to tube degradation in operating plants. Observations of the two most widespread types of attack, i.e. phosphate wastage and denting, are described in detail. Wastage seems to be consistent with a mechanism in which solid sodium ortho-phosphate is deposited on the Inconel tubes, followed by conversion to highly corrosive pyro-phosphate. Reasons for differences between laboratory results and the rates of wastage found in service are discussed. Denting is ascribed to the formation of an acid chloride environment in tube support plate crevices, which attacks the carbon steel and produces non-protective, bulky magnetite, leading to deformation of tubes and support plates. Impurities from condenser leaks, corrosion products from steel, Inconel, or copper base alloys and phosphates may trigger or accelerate denting; improved condenser performance and/or demineralization are needed to prevent problems caused by contamination of the secondary water. Severe denting has led to intergranular stress corrosion cracking of Inconel 600 in primary water, and deformation and cracking of the support plates. It is believed that the cracking in Inconel occurs in primary water according to predictions made many years ago, and that hydrogen is a likely cause of cracking in the highly deformed steel. A few alternate materials for steam generator tubing are mentioned together with some of their properties.
1. GENERAL SUMMARY OF PROBLEMS At this time there are 37 operating PWR units in the U.S.A., of which 19 have at least one form of tube degradation, as shown in Table X. No degradation has been found in 8 others, and 10 have only operated for a short time. The two most widespread of the problems listed in Table I are phosphate wastage of tubing (12 plants) and denting (14 plants). The major cause of excessive corrosion of the Inconel in phosphate-treated plants was the phos phate itself, i.e., a deliberate addition, although impurities may possibly have added to the rate of attack. The solution to this problem was simply
2 1 7 TABLE I
SUMMARY OF U.S. STEAM GENERATOR PROBLEMS ROOYEN van 8 1 2
OPERATING PHOSPHATE CAUSTIC FATIGUE DENTING HISTORY WASTAGE CRACKING* DEGREE PRIMARY SIDE SUPPORT PLATE (MONTHS) CRACKING P°4 ALL VOLA TILE U-BENDS DRILLED HOUR- CRACKS HOLES GLASSING Westinghouse: ' Zion 1 21 25 X - Yankee Rowe 0 185 (SCO of Stainless Steel Tubes) - H.B. Robinson 2 81 0 X X minor (!) Conn. Yankee 96 22 X X minor tubesheet Point Beach 1 47 30 X X moderate X (!) Point Beach 2 28 30 X moderate X(!) Ginna 1 60 29 X X moderate X (!) Indian Point 2 33 25 moderate X
Babcock & Wilcox* Oconee 1 X - Oconee 2 X - Oconee 3 X -
I Nature to be confirmed - no tubes removed for inspection. (Verify if wastage contributed at San OnofreJ ? May have been interrupted for some period(s). * Once-through steam generators. Note absence of corrosion problems. This may be related to support plate design and to demineralizers which would give minimal levels of copper and iron corrosion products and dissolved minerals in the feedwater. Also, low copper specifications preclude use of copper tubes in feedwater heaters. IAEA-SM-218/26 2 1 9 to omit the phosphate. Denting, by contrast, is the end result of rapid crevice corrosion caused by inleakage of impurities that form acid chloride in the drilled hole crevices of the tube support plates. Improvement of con denser performance, therefore, is a very urgent and critical item, although it is not discussed in this paper. Secondary effects of severe denting are primary side intergranular stress corrosion cracking of Inconel 600, as well as deformation and cracking of support plates. These forms of damage have been confirmed only at Surry and Turkey Point, but cracking of the In conel 600 is suspected at some other plants, where detailed examination is still pending. M echanistically, p ra c tic a lly a l l of the phenonema involved in the cor rosion that has been encountered were earlier described in the literature, but equivalent conditions were not anticipated in operating steam generators. Once the problems arose, however, the available information was helpful in diagnosing the problems and trying to find solutions.
2. PHOSPHATE WASTAGE OF STEAM GENERATOR TUBES
2.1 Observation of Attack
Congruent phosphate treatment is intended to eliminate hard scale and to control pH. Some believe (as in the U.S.) that it also inhibits corrosion in boilers, but there are other authorities^ who describe the use of phosphate as perilous because it can render the magnetite coating on steel boiler sur faces nonprotective. Before 1973, very low phosphate levels at relatively high Na/P ratios were used in secondary water of PWR plants, with good re sults when well controlled. Examples are trouble-free operation for 5 years or more a t Ginna, Connecticut Yankee and o th e rs. U nfortunately, a general increase in phosphate levels with a reduction in Na/P ratio was implemented because caustic cracking occurred in a few plants where deviations from usual procedures were made. Combined w ith the high power output of commercial plants it promoted the concentration of phosphates in heated crevice regions, causing corrosion and thinning of steam generator tubes. This so-called wastage appears to be least for stainless steel, and worst for Inconel 600, with Incoloy 800 in between.
2.2 Mechanism and Laboratory Tests
It is generally known that corrosion of Ni-Cr-Fe alloys in dilute aqueous phosphate with Na/P ra tio s in the range 2-3 is extremely low, even a t high tem perature. In service the combination of repeated evaporation in crevices and retrograde solubility of sodium phosphate can lead to precipitation of solid 220 van ROOYEN phosphate in these occluded regions. The phase diagrams of the H20-Na20- system, for various Na/P ratios have been summarized in an excellent c _ [2] review by Marcel Poubaix,L including the temperature region of 240-300°C, which is of interest to PWR steam generators. At operating temperatures mono- or disodium phosphate can lose one molecule of water to form pyrophos phate in the presence of liquid water:
2 NaH P0„ 2 4 = Na.H„P2 2 2 7
2 Na-HPO„ * Na.P-0- + 2 4 4 2 7
This happens somewhat more rea d ily a t low Na/P. Meta-phosphates CNaPO^) w ill be stable at higher temperatures above 343°C. All of the glassy phosphates can be highly corrosive to a variety of metals, including Ni, because of com- [2] plex ion formation. This is a well known phenomenon, described as long ago as the "blow-pipe analysis" which dates back to the 19th century.^ The degree to which these glassy phosphates actually account for the wastage of steam generator tubes remains to be clarified. Analysis of the green corrosion products on Inconel tubes has shown ortho-phosphate high in Ni and Fe, [4] while Fe-ortho-phosphate appears to be the product on Inconel 800. However, these analyses were made after cycling back to room tempera ture, and compositions may have been different at the corrosion temperature. In fact, in another excellent recent paper, [4] it was shown in capsule tests that a saturated solution of Na2HP0^, in contact with excess solid remained dissolved as the ortho-phosphate at 300°C while the solid was converted to Na^P^^, which slowly started to change back to the ortho form on cooling down. This la tte r work [4] also showed th a t the high temperature corrosion would increase as the ortho-phosphate concentration (e.g., HP0~) increased in the aqueous phase where no pyro-phosphate conversion occurred. The rates for Inconel 600 and 690, Incoloy 800 and sta in le s s ste e l showed about the same response. However, the maximum in these rates did not approach the severe wastage that can occur in service, such as about 1.25 mm per year at Mihama. This is listed in Table II, together with some of the differences between the laboratory and field conditions. It is believed that an essential factor is the absence of pyro-phosphate in contact with the metal in the laboratory tests, which would account for the lower rates in general when HPO^ is the main corrodent. In service, it is believed that retrograde solubility would cause the precipitation of solid phosphate on tube surfaces in hot, low flow regions where the pyro-phosphate transformation would occur and cause metal wastage at the rates observed, without more information, it is impossible to IAEA-SM-218/26 221
TABLE II
COMPARISON OF SERVICE AND LABORATORY CONDITIONS
ITEM LAB SERVICE
Temperature Same Range
Rates 0.05 mm/a (max) Up to 1.25 mm/a
Heat Flux No Yes
Max. Concentration Saturated Solution Solid may preci p ita te on metal
A lt. Wet/Dry No Yes
Galvanic Couples No Yes
Waterline. Constant or Fluctuating absent
Pyro-phosphate Probably not May contact on metal metal surface
Im purities None added From condenser in-leakage
FIG.l. Possible adverse results of contact between Inconel and carbon steel. 222 van ROOYEN discuss any contribution of impurities to tube wastage in operating steam generators, except to point out that they may well have played a part.
2.3 Future Work and Conclusions on Phosphate
The possible contribution of impurities to wastage needs clarification in future work.. If anything, they would promote the formation of pyro-phos- phate if they increased acidity in crevice regions because this would lower the Na/P ratio. However, they could increase the corrosion rate even in the absence of pyro-phosphate formation. Galvanic effects have not been mentioned in detail. In service, the tubing would be in contact with carbon steel, giving a good couple because the solution in crevices would have good conductivity at least when it is highly corrosive. Such contact may not necessarily protect the tube, as shown in Figure 1. Meaningful corrosion test procedures with phosphate do not exist in any laboratory. While it is recognized that it would be difficult to convince the utilities to return to phosphate treatment, it is suggested that some testing be continued to develop a good alloy screening test, for four reasons, despite the drawbacks of PO^ treatment in service: 1. Some units still use P0„. 4 2. There has been some difficulty with AVT* 1(denting), and in sufficient experience with AVT to show that it will be satisfactory in long term use (pitting? SCC? scaling?)2. 3. Early experience showed excellent performance with low PO^ in secondary water. 4. Existing tests do not distinguish between Alloys 600, 690, 800 and stainless steel, while in-service results (e.g., low wastage of Alloy 800 at Stade) show that there may be such differences, and more candidate alloys should be screened.
3. DENTING
3.1 Description
Because of the phosphate wastage problem in several recirculating steam generators, many plants were converted from PO^ to AVT. Within a relatively short time thereafter, denting was observed at Turkey Point and Surry in the
1 AVT = All Volatile Treatment. 2 SCC = Stress Corrosion Cracking. IAEA-SM-218/26 2 2 3
U.S. The source of denting was corrosion of the carbon steel to form massive non-protective magnetite (Fe^O^) in the tube-to-support-plate crevices, but for some reason not in the tube sheets. At first it was thought that only plants with a long P04 history followed by AVT were subject to this damage. Table III gives a list of plants affected, from which it is evident that denting has now occurred: 1, After prolonged PO^ operation + AVT conversion 2, After short P04 operation + AVT 3, Without PO^, sta rte d on AVT 4, While still operating on P04 The degree of denting varies from severe, where extensive tube and sup port plate deformation and leaks have been found, to minor, where only about 1 mil tube deformation is found. In addition, abnormal eddy current signals have been found at the top of the tube sheets in 4 plants; these have not been fully documented, but if they indicate tube deformation it is very s lig h t. As already noted in Table I, there is a noticeable absence of severe corrosion in once-through steam generators. This is most likely related to the fact that full flow condensate polishing is practiced, together with a specification for low copper in the secondary water, which precludes the use of copper-base alloys in the condensers. It is also to be noted that the support plate holes in these steam generators have a different design.
3.2 Mechanism
Carbon steel in high temperature water normally acquires a thin protec tiv e layer of Fe 0 , from the reaction 3Fe + 4H О Fe 0 + 4H,. However, aggressive ions such as chloride, and perhaps phosphate also, can damage the film. When this happens in the presence of acid Cl , there is a rapid linear growth of porous Fe^04, which can exert a tremendous outward pressure if formed in a confined space. This part of the mechanism is very straight forward, but it is difficult to ascertain exactly how the acid chloride is formed under service conditions. A quantitative understanding of this phe nomenon is still being developed as searches continue for ways to stop denting. All of the plants that were converted to AVT had earlier suffered some degree of tube damage as a result of reaction between Inconel and high con centrations of ortho- or pyro-phosphate. Physically, the corrosion products would create a very tight crevice, although it is uncertain if this makes denting any worse. Chemically, the unbuffered hydrolysis of sparingly sol uble Fe, Ni or Cr phosphates would lead to a pH drop inside the crevices 2 2 4 van ROOYEN
TABLE I II INCIDENCE OF DENTING
PLANT MONTHS MONTHS DENTING LEAKAGE CONDENSER ON ON AT COOLING AVT P04 DENTS WATER SURRY 1 29 27 Extensive Yes Brackish SURRY 2 23 26 Extensive Yes Brackish TURKEY POINT 3 23 31 Extensive Yes Sea TURKEY POINT 4 14 31 Extensive Yes Sea
SAN ONOFRE 107 0 (?) Moderate (?) * Sea INDIAN POINT 2 33 25 Moderate No Brackish POINT BEACH 1 47 30 Moderate No Lake POINT BEACH 2 28 30 Moderate No Lake RINGHALS 2 4 29 Moderate No Sea MILLSTONE 2 0 23 Moderate No Sea
PALISADES 48 42 Minor No Lake ROBINSON 2 81 0 Minor or none No Lake BEZNAU 2 34 32 Minor No River GINNA 60 29 Minor No Lake CONN YANKEE 96 22 Minor No River MAINE YANKEE 0 53 Minor NO Brackish.
DOEL 1 3 30 None** No Sea (FFCP) DOEL 2 0 21 None** NO Sea (FFCP) D.C. COOK 1 HF 27 None** No Lake TIHANGE 1 0 26 None** No River
BEZNAU 1 33 33 None No River ZORITA 115 0 None No River MIHAMA 2 21 17 None No Sea MIHAMA 3 0 15 - No Sea TAKAHAMA 1 6 23 None No Sea TAKAHAMA 2 HF 28 None No Sea GENKAI 1 0 27 None No Sea KEWAUNEE 6 30 None No Lake ZION 1 21 25 None No Lake ZION 2 11 29 None No Lake PRAIRIE ISLAND 1 10 30 None No River PRAIRIE ISLAND 2 HF 28 None NO River INDIAN POINT 3 0 13 - NO Brackish FESSENHEIM 1 0 1 - No River TROJAN 0 16 - No River (FFCP) SALEM 1 0 5 - No Brackish BEAVER VALLEY 1 0 12 NO River
Not y et reported. To be examined. Incompletely interpreted abnormal signals at tube sheets - see text. HF Hot functional te s t only FFCP Full flow condensate polishing. IAEA-SM-218/26 2 2 5 when the sodium phosphate treatment is discontinued and much of the sodium compounds have d iffused away. This could trig g e r corrosion of the carbon steel, so increasing the concentration of Fe"14" in the crevice, accompanied by the accumulation of high concentrations of chloride (if present even in small amounts in the bulk solution) in order to preserve charge neutrality. In addition, the subsequent hydrolysis of FeCl^ would keep the pH down. The above mechanism may be possible after PO^ to AVT conversion. How ever, prior phosphate treatment is not needed for rapid Fe 0 growth, accord- 151 3 4 ing to the work of Potter and Mann, who examined the behavior of carbon steel coupons in solutions containing FeCl^, NiCl2, CuCl., and others. This was done without crevices, and did not involve denting as such. They re ported sporadic "run-away" Fe3C>4 reactions with FeCl2 or CuCl2, but consistent results in solutions containing NiCl . More recent data from the Westing- house laboratories1 now show that, when crevices are present, copper ions together with chloride produce denting consistently. In related tests, dent ing was produced at Combustion Engineering^ in model boilers, in the presence of Cl and crevices packed w ith Ni, Cu, CuO, Fe^O^. The adverse effects of crevices combined with CuCl^ or Ni-salts ' such as may also be introduced in service through in-leakage and corrosion of Inconel and Cu or Cu-base alloy condenser tubes, are thus indicated. A combination of tight crevices and Cl" alone does not seem to be enough for denting, since lOx seawater by itself, i.e., a non-acid and rela tively non-oxidizing medium, did not give denting at Westinghouse, ^ even with "frozen" crevices. Therefore, C u ^ , Cu(OH) + , or Ni++ seem to be needed together with Cl to set off denting under neutral conditions. In-service implications remain to be clarified, such as the relationship between the time to denting in AVT units, and (a), the accumulation of inert sludge and (b), the accumulation of active "denting" ingredients, i.e., hydrolyzable chlorides and oxidizing ions. The role of an external oxidizing ion, such as Cu44" or Ni*"*" can be th reefo ld : 1. At reasonably high concentrations they can depolarize the cathode and so produce the initial acid chloride in the crevice that leads to denting; in conjunction, there could be a catalytic effect, as in "2” below. 2. At lower concentrations where their oxidizing potential may be low, they could simply catalyze the ferrous [8] hydroxide conversion to magnetite. 3. The metal that is deposited after reaction with electrons could form a low hydrogen overvoltage site, and so affect the kinetics of corrosion. 2 2 6 van ROOYEN
FIG.2. Illustration of “protection ” potential.
In the depolarizing sense, which probably applies to laboratory tests with high Cu++ levels, the triggering of denting would be analogous to pitting of stainless steel in Cl , which can be started at an elevated electrochemical potential, but then, on lowering this potential, continues because of the local acidity and higher Cl inside the pit, and only ceases at the "protection" potential which can be considerably lower than the ori ginal pitting potential in the bulk solution. (See schematic representation in Figure 2.) The "protection" level of denting, once started, would then be below the potential at which H^O polarization curve intersects the anodic curve. The "triggering" cathodic reactions, Ni++ + 2e ->■ Ni or Cu** + 2e -*• Cu would be replaced by 4H 0 + Fe -*■ Fe,0 + 4H_, because there is not nearly ++ ++ * J 4 2 enough Cu or Ni to account for denting and also it is known that copious H2 evolution accompanies denting. It is yet not certain that Cu++ or Ni++ are present at concentrations high enough to fu lfill such a role under service conditions, but they may play a triple role, the second of which, is to cata lyze the conversion of the initial corrosion product, ferrous hydroxide, to [8] magnetite. If this should happen under conditions where the magnetite is deposited a small distance away from the steel surface, because the Cl may give a more soluble initial corrosion product, then it would be non-protective and by speeding its formation Cu++ or Ni++ would accelerate the corrosion reaction. This probably happens in the laboratory as well as in service, but IAEA-SM-218/26 2 2 7 it remains to be established to what degree it aggravates denting, since the ++ , . ++ Fe^0„ will also form in simple acid chlorides without Cu or N1 3 4 c Г91 + additions, i.e., where "triggering" is done artifically by adding the H . In the th ird ro le , the deposited Ni or Cu may form su itab le low over potential sites for the cathodic H2 evolution. This may permit more rapid Fe^O^ formation, and accelerate denting, but does not appear essential for the denting reaction to proceed, for the same reason given above.
4. CRACKING OF INCONEL TUBING
4.1 Observation
Several small leaks have been found in Inconel 600 tubes of 4 or perhaps 5 dented plants, (Table I) , and one larger leak of3001/min occurred at Surry 2 in 1976. Where examinations have been made, these were longitudinal cracks starting from the primary side in the support plate region as well as in U- bends in some first tube row (smallest radius) where deformation was greatest. The nature of the cracks suggest only one cause, that is intergranular SCC in primary water. Two cracks have also occurred in tubes at points just above and at the tube sheet region of Ginna, which is a plant with, minor denting; the nature of these cracks have yet to be examined in detail.
4.2 Mechanisms
Increased tensile stress was associated with SCC. In the support plate regions, the. origin of the stress is obviously the. pressure of Fe^O^. Inci dentally, the presence of Fe^O^ will tend to seal the outside of cracks in the support plate regions when they penetrate the wall and keep the leak small. In U-bends, the stresses were caused by inward movement of the tube legs due to closure of flow slots in the support plate. As the U-bends be came "tighter", ovality resulted at the apices, so that the stress pattern became complex, including hoop, axial and radial components. From the axial direction of the cracks, it is evident that the hoop stress due to ovality is of major importance in this region. It is equally important in the sup port plate regions because those cracks have the. same direction. The penetration of the wall, when it occurs, is usually confined to a short length, of the crack, so that a very small primary to secondary leak re s u lts . Unresolved questions are the degree of strain before SCC starts, and the contribution of dynamic straining during denting. Usually, the strain 2 2 8 van ROOYEN
FIG.3. Strain rate effect on SCC.
ra te is re la te d to crack v elo city as in Figure 3. However, since SCC in pure water has not occurred in tests in the region of 10 ^ to 10 7 s-1 for Inconel, it is suggested that the corresponding curve for Inconel lies in an extremely low strain rate region, which may be very difficult to distinguish from creep during "ordinary" U-bend tests. Coriou and his co-workers first demonstrated the susceptibility of Inconel 600 to SCC in primary water. Other workers have confirmed this, as [12] reviewed earlier. Before the denting era the only known in-service pri mary side crack was in the rolled region at Obrigheim, so that it was assumed that operating plant conditions were so much different that the SCC data from laboratory tests did not apply. It is now evident that an event that aggra vates the stress in tubing will cause SCC, and it should be considered equally possible that there could perhaps be other as yet unidentified fac tors that could be equally capable of causing SCC in service if introduced inadvertently. Consequently, there is a need to develop quantitative data to establish, relationships between the factors contributing to SCC in Inconel so as to be able to assess the likelihood of cracking of Inconel 600 under a variety of possible conditions. Environmental variables should include effects of boric acid, as well as other primary water ingredients, secondary water, stress, strain, strain rate, strain type, temperature and temperature cycles, small amounts of 0^ and probably other test variables. Also included as a variable in this work should be processing parameters and microstructure since there are strong indications that they are important. For instance, when Inconel was "sensitized" at about 60Q-65Q°C, it was found to be much more r e s is ta n t to SCC in pure water (no 0^) than the annealed alloy, con- IAEA-SM-218/26 2 2 9 trasting with normally expected behavior. Also, additional microstructural features have been related to the history of the material and its SCC r 1 4 1 resistance. It is felt that this work is only the beginning of a pattern of inter-relationships that will emerge, and which is ready for further re search. Meanwhile, for those who have the. courage to do so, deliberate "sensitization" or, as sometimes erroneously called "stress relief" (although it may precede the application of stress) could be applied in an attempt to increase primary side SCC resistance. Of course., secondary side effects should always be considered also.
5 CRACKING OF CARBON STEEL SUPPORT PLATES
Distortion of tube support plates occurred in severely dented plants, resulting in partial closing of the flow slots, expansion at the periphery and cracking of some of the ligaments between flow holes and other drilled holes. It is felt that the cracking is a result of the high stress in the carbon steel, combined with the large amounts of that is formed during the "run-away" Fe О formation: H O + M -*■ M O + H . A similar mechanism 3 4 г 3 4 Z [15] for cracks in carbon s te e l was e a r lie r put forward by P o tte r, whose work with Mann laid the basis for explaining the overall process in which non- protective magnetite can be formed on carbon steel surfaces.
6. PREVENTION
6.1 General
Current problems are closely related to impurities in secondary water. Existing plants, therefore, would have to take precautions to eliminate in leakage or to remove impurities as they are detected. Sea water cooled plants may need demineralizers. Future plants have the advantage of choosing alter nate materials for various components. They can also change the design of the support plates so as to minimize crevices, as is already being undertaken by U.S. manufacturers, and is used in once-through units.
6.2 Secondary Water Control
Operating p la n ts require some c r ite r ia for eliminating corrosion problems. Chloride for instance is very adverse, but setting operating tar gets is difficult for two reasons: 1. Fixing a bulk concentration does not necessarily limit crevice conditions, and it is the crevice liquid concen tration that determines whether denting will start. 2 3 0 van ROOYEN
Increases of several orders of magnitude above the bulk water levels have been reported for Cl”. Also, it is believed that hydrolyzable salts and cathodic depolar ization or catalysis by metal ions will alter the level of Cl that can be tolerated before denting sets in. In support of this, the Westinghouse lOx seawater ' te s t showed th a t q u ite high Cl is not bad by i t s e l f . 2. All the species and interactions that lead to denting may not yet be identified. With those that are known, little quantitative work has been done. For instance, laboratory tests so far have been in quite high bulk Cl solutions, in order to identify adverse species and to reproduce denting quickly. Two of the extremes in Cl are (a) lOx seawater, without denting, i.e., 15Q,Q00 - 200,000 ppm Cl in a Westinghouse test, and (b) 1 ppm Cl with severe denting in the presence of a mixture of Ni, Fe^04, CuO, and Cu at Combustion Engineering. These two examples indicate that there could be very important interactions or synergystic effects in denting.
If a general criterion has to be fixed for chloride, if would have to be conservative, and the assumption will have to be made that other adverse species are always present. For this purpose, it is useful to review the experience at a few selected plants, as in Table IV. Obviously, the maximum level of Cl seems to be quite low to avoid dent ing. Also, this applies to the steam generator blowdown, and not to the con densate, since the latter already contains considerable inleakage at the detection limit of 0.05 ppm Cl of usual monitors. It would be an advantage if the sensitivity of monitoring the condensate could be improved, so that leaks could be stopped before there is any serious accumulation of impuri ties in the steam generator itself.
6.3 Inhibitors
Inhibitors that would raise the local pH in the crevice regions are under study, since such additions may stop the rapid magnetite growth.
6.4 Chemical Cleaning
Regular cleaning of secondary sides of steam generators would also be valuable in controlling corrosion problems such as denting; no proven pro cedures are available at present, but considerable research progress has been IAEA-SM-218/26 2 3 1
TABLE IV
COMPARISON OF CHLORIDE LEVELS IN SOME OPERATING PLANTS
PLANT MONTHS Cl ppm ON P0„ TOLERATED ACHIEVMENT RESULT 4
Maine Yankee 0 0.05 ppm 60 days > 1 ppm. Minor denting Few excursions in 1975-77 about 30 ppm period.
Ringhals 2 4 < 0.1 ppm No dents 1975. Moderate dent Since then, 5 ing in 1975-77 excursions > 0.1 ppm. 19 days out of spec. All < 1 ppm.
Point Beach 47, 28 <0.15 condensers Moderate 1 and 2 leaked dents
Robinson* All Not Expected to Minor dents? 81 av ailab le be very low, Perhaps none but lake in view of water re cooling water ported composition to be < 1 ppm
Japanese Reported meticulous care to No denting maintain "clean" secondary systems.
Reference 1 warns that phosphate could be directly involved in forming non-protective oxide on steels in boilers.
made and a "demonstration" effort is planned for Indian Point 1 this year. It stands to reason that a reliable chemical cleaning procedure, cannot be indiscriminately used, and that there will have to be assurance that the re moval of solid products on badly dented plants, for instance, will not re move the support for cracked tube support plates or tubes. Other factors, such as effects of the "solvent" on all components, and the. possibility of creating excessive clearances and causing tube vibration, are among the items being considered in the process of developing good cleaning procedures. 2 3 2 van ROOYEN
6.5 Alternative Materials
For condenser tubing, Ti is being considered so that copper can be bet ter controlled. For steam generator tube support plates, stainless steel is a strong candidate to eliminate the denting reaction. Of course, tile latter will have to be tested to make sure that impurities will not lead to corrosion product build-up or to stress corrosion cracking. Steam generator tubing for PWR applications also deserves a thorough review, because of the cracking that has occurred in the U.S, from the pri mary side, as well as the fact that Incoloy 800, which does not have as much Ni as Alloy 600, is not entirely immune to chloride SCC. In addition, the increased risk of pitting corrosion in AVT secondary water should be taken into account. Based on the above requirements, and the absence to date of chloride in European operating plants of Alloy 800, there appears to be 3 main candidates as alternative m aterials in the U.S. They should be included in tests to obtain more information as well as in-service evaluation where p o ssible: Alloy 800 - Excellent resistance to SCC in primary water, otherwise approximately similar to Alloy 600. This alloy has increased Cr content that may be helpful, and satisfactory service experi ence in Europe supports its usefulness in spite of chloride SCC that can be produced in high temperature laboratory experiments. Alloy 625 - It has excellent pitting and crevice corro sion resistance, because of the higher Cr content and substantial level of Mo. It will resist chloride SCC, because of high Ni, but its resistance to primary water SCC must be established. Also, its heat transfer properties may require improvement. Alloy 690 - This 60Ni-30Cr-lQFe alloy is probably the major candidate to replace Alloy 600 as steam generator tubing. It was tailor- made for this application, with laboratory tests done to ensure resistance to SCC in high temperature water with and without 0^. and also chloride SCC resistance. It has excellent resistance to pickling acid, therefore may be superior to Alloy 600 to crevice corrosion. Pitting corrosion in IAEA-SM-218/26 2 3 3
chlorides is also in?>roved over Alloy 600, as is SCC in aerated caustic. In deaerated caustic it may be a little less resistant than Alloy 600. It is difficult to "sensi tize" this alloy because of its high Cr content.
REFERENCES
[1] STANISAVLIEVICI, L., The use of phosphate in high-pressure boilers, Combustion, A pril (1967) 41.
[2] POURBAIX, M., On the Chemistry of Phosphate Solutions Under the Opera tive Conditions of Pressurized Water Reactors PWR, Cebelcor Report MP.2437/cc, 21 March (1977).
[3] ELDERHORST, W., Manual of Blowpipe A nalysis, Van Nostrand, New York (1867).
[4] PESSAL, N., et a l., The corrosion resistance of Inconel in high tempera ture phosphate so lutions, Corrosion 21 (1970) 130.
[5] POTTER, E.C.,and MANN, G.M.W., The fa st lin e a r growth of magnetite on mild steel in high-temperature aqueous conditions, Br. Corros. J. (1965) 26.
[6] LINDSAY, W.T., P rivate Communication and ECONOMY, G., e t a l ., Labora tory Investigations Related to Denting in Nuclear Steam Generators, Paper presented at N.A.C.E. Corrosion Research Conference, San Francisco, March (1977).
[7] BRYANT, P.E.C ., P rivate Communication.
[8] BUTLER, G., and ISON, H.C.K., Corrosion and its Prevention in Waters, Reinhold, New York (1966) 176.
[9] GARNSEY, R ., P rivate Communication.
[10] VAN ROOYEN, D., Unpublished Work at Brookhaven National Laboratory.
[11] C0RI0U, H., et al.. Corrosion Fissurante Sous Contrainte De L'Inconel DansL’Eau A Haute Temperature (3e Colloque de Metallurgie Corrosion, Centre d'Etudes Nucl^aires de Saclay, France) North Holland, Amsterdam (1959) 161.
[12] VAN ROOYEN, D., Review of the stre s s corrosion cracking of Inconel 600, Corrosion (1975) 327.
[13] BLANCHET, J., et a l., Influences of Various Parameters on Intergranular Cracking of Inconel 600 and X750 in Pure Water at Elevated Temperature, P reprint G13, S tress Corrosion Cracking and Hydrogen Embrittlement of Iron Base Alloys, Firminy, France) June 1973 .
[14] DOMIAN, H.A., e t a l . , E ffect of m icrostructure on stre s s corrosion cracking of allo y 600 in high p u rity water, Corrosion 22. (1977) 26.
[15] POTTER, E.C., Trans. Am. Soc. Mech. Eng. J. (Eng. for Power) 86_ (1964) 320.
Session VII
QUALITY ASSURANCE AND QUALITY CONTROL C h a ir m e n
J. F O R S T E N F in la n d
G . Ö S T B E R G S w e d e n IAEA-SM-218/51
THE CONTRIBUTION OF QUALITY ASSURANCE TO SAFETY AND RELIABILITY IN NUCLEAR TOWER PLANTS
N . R A I § l C International Atom ic Energy Agency, V ie n n a
A b stra c t
THE CONTRIBUTION OF QUALITY ASSURANCE TO SAFETY AND RELIABILITY IN NUCLEAR POWER PLANTS. This paper analyses the potential contribution of quality assurance to nuclear power plant safety and reliability. In the first part of the paper an attempt is made to establish a relationship between quality and reliability. It is assumed that reliability may be considered as quality over the course of time. The reliability may be expressed in quantitative terms as “the probability that an item will perform a required function for a stated period of time”. Quality, however, cannot be expressed in simple quantitative terms but only as a set of required properties which an item should have for a specific application. The achievement of quality and additional reliability objectives is a task of project activities such as design, construction, installation, operation, etc. Objectives of quality assurance are to give confidence to the plant owner and to the regulatory organization that the quality of the plant as specified by regulatory requirements, codes, standards and specifications, has been achieved and main tained. These objectives can be attained either through inspections and testing by an indepen dent inspection body of all items and activities affecting the quality of the plant, or through the establishment and operation of a quality assurance system which provides for a planned and systematic approach to all activities. The elements of a quality assurance system and its functions in nuclear power projects are presented in some detail in the paper. Confidence in plant quality, which should be a basis for the regulatory body issuing the construction permit or operation licence, should be based on the capability of quality assurance activities to prevent errors and correct deficiencies in nuclear power plants. An analysis is made of those errors in plant design, manufacture, construction and operation which contribute most frequently to plant outages. Most of the failures have in fact common causes and may be categorized as common-cause failures. It is concluded that these errors can be avoided or corrected by strict adherence to quality assurance principles and by the efficient functioning of quality assurance systems. In fact, quality assurance may be considered an effective defence against common cause failures originating in errors in the design, manufacture, installation or operation of a nuclear power plant.
1. I N T R O D U C T I O N
The planned and systematic implementation of quality assurance in nuclear power projects dates back only a few years. However, m any activities which are now regarded as quality assurance have been implemented in the past both in
2 3 7 2 3 8 RAISIC nuclear and conventional power projects, although they were not previously identified as such. The major difference lies in the planned and systematic way in which quality assurance activities are implemented at present. A set of specific and well-established criteria is used as a basis for those activities, and form al requirements for their im plementation in all nuclear power project activities have become the concern of governmental regulatory bodies and the general p u b lic . The ultimate goal of quality assurance activities is usually defined in terms of “obtaining confidence that an item or facility will perform satisfactorily in service” [ 1 ] . Obviously, satisfactory performance in service is used here as a synonym for quality. However, the satisfactory performance of an item in service is a reliability goal. This illustrates that the m eaning of quality is intim ately related to the meaning of reliability. In order to make a clear distinction between these two terms, we m ight attempt to define quality and reliability, although the questions of term inology are not determinant; neither in quality assurance, nor in reliability technology. One can define the quality of an item as: a set of attributes pertinent to its properties which determine the degree of aptitude of this item to perform its intended function. Reliability, on the other hand, m ay be defined as the ability of an item to conserve its quality over time when operated under stated conditions. In other words, reliability m ay be considered as quality over time. It is not possible to express the quality of an item in simple quantitative terms. However, reliability can be expressed by “the probability that an item will perform a required function under stated conditions for a stated period of t im e ” [2]. This quantitative definition of reliability makes it possible to establish reliability objectives for an item in precise quantitative terms such as the acceptable rate of failure, and to undertake appropriate measures for achieving this objective w ith a given level of confidence. Since quality cannot be expressed in simple quantitative terms, quality objectives should be specified as the required properties which an item shall have for a specific application. Wherever possible, these properties should be defined so as to allow for the verification of an item ’s conformance to the requirements through inspection and testing. When such verification is not convenient, or when it is impossible, quality requirements m ay be expressed in addition to the specification of properties also in terms of the procedures by which the item shall be designed, constructed, installed, tested and operated. In this way, quality is indirectly identified by those attributes which are obtained by the conformance of an item to pre-established requirements related to proper ties, and by conformance to those methods which are to be utilized so as to acquire those properties. The relation between quality requirements expressed in this way and “satisfactory performance of the item in service” is not very direct, although such a relationship m ay be established in the long run through IAEA-SM-218/51 2 3 9 the collection of field experience with the items concerned, feedback inform ation for the process of specification and standard preparation, and the establishment of other procedures for the disciplined approach to activities. For the purpose of establishing a relationship between quality and reliability, one m ay consider the possible m ethods for achieving the established reliability requirement for complex systems. It is to be expected that increasing complexity of a system will tend to decrease its reliability. On the other hand, the require ments for satisfactory functioning of such systems are becoming more severe when they are of particular importance to safety or the satisfactory functioning of the facility to which they belong. To solve this problem, two different approaches are possible:
The first is to increase the quality of each com ponent of the system. This is norm ally accomplished by providing detailed quality specifications for the items, which include: the specific requirements for design, selec tion of material proved by experience, specification of manufacturing processes, and strict quality control procedures.
The second is to develop a special design for the system using standard components. Points of low reliability are determined by reliability analysis of the system, and should be eliminated by incidence of redundancies, and by the elaboration of periodic testing and maintenance schemes to be implemented during operation of the system or by other methods. This approach is intimately related to the application of a new engineering discipline, “reliability engineering”, which includes reliability a n a ly s is .
The first approach, increasing the reliability of a system by ensuring the high quality of its components, is more readily used in nuclear power plants. For m ost of the systems, such as pressure systems, engineered safety systems, etc., this is the m ost efficient way of achieving high reliability and high overall efficiency of those systems in operation. One would expect that a quality component would have a low failure rate. For a quality product causes of potential failures such as design, manufacture or installation error should be eliminated and the failures should be reduced to typical random events.
2. QUALITY ASSURANCE AND QUALITY CONTROL
The task of achievement of quality and reliability objectives is not within the scope of quality assurance activities. The responsibility for achieving quality 2 4 0 RAISIC rests entirely with those individuals or organizations assigned to those tasks. The objectives of quality assurance can be restricted to:
Obtaining confidence through quality planning and verification activities that an item or a facility has been designed, constructed, manufactured, installed, inspected, tested and operated in accordance with given standards and procedures, and that the properties of the item conform to pre- established quality requirements.
In this definition, the key word is “confidence”. The objectives of quality assurance are, therefore, not to achieve or improve the quality of items, but to gain confidence that the required quality has been achieved. In the scope of defined objectives of quality assurance, it may be useful to define also the role of quality control. The difference between these two disciplines is sometimes confusing; the reason for this being the rather recent use of the term quality assurance in the nuclear industry and the meaning of the term “control”, which is not uniform in all countries. For the purpose of clarity in this paper, the term “quality control” m ay be defined as:
Those actions which provide for the means to determine the characteristics of an item, process or facility, and to control those characteristics in relation to established requirements.
In this definition, the emphasis is more on “control” than on “quality”. Whereas control is a part of any organized activity, quality control is more a part of project activities (manufacture, construction, installation, etc.) than of quality assurance. The relationship between quality assurance and quality control is practically the same as that between quality assurance and any other project activities, except that quality assurance m ay sometimes utilize the same m ethodology and techniques as quality control. The objectives of quality control are, therefore, to control the product specification. The satisfactory performance of this task might be just one of the contributing factors towards gaining confidence in a product’s ability to meet all requirements.
3. QUALITY ASSURANCE OPTIONS
The m ethodology for gaining confidence is not uniform and may be depen dent on several factors, which vary from one nuclear power plant to another, or from one country to another. It m ay be assumed that the selection of adequate methodology will depend on the following: IAEA-SM-218/51 2 4 1
National regulations with respect to safety in nuclear power plants.
The organization seeking to gain confidence (constructor, plant owner, regulatory organization).
The existing body of technical documents, such as engineering codes, standards and specifications which will be used as the technical basis for the activities.
The size of the nuclear power programme which will be subject to quality assurance activities.
The availability and capability of qualified and skilled personnel for perform ing the planning and verification activities.
All these factors are taken into account when selecting an optimal approach to quality assurance. However, the concept of quality assurance will always reflect national regulations and the adopted approach to regulatory controls for nuclear power plant safety, and practices used in exercising these controls.
3.1. Confidence through independent review, inspection and testing
The independent review, inspection and testing of products or activities performed by competent and authorized individuals or organizations has been a frequently employed method in the industry of gaining confidence in the quality of products or activities. Taking into consideration that reviews, inspections and tests are also employed as methods of quality control, performance of such inspec tions and tests for the purpose of gaining confidence in the achieved conformance to requirements means, in fact, a redundancy of such reviews, inspections and tests. This approach is sometimes referred to as “product-oriented quality assurance”, because the measures employed to gain confidence are directly oriented to the system or component which is to be manufactured, built, or installed. The independent inspection organization is engaged on a full-time basis, and undertakes review and verification functions during design, construc tion, fabrication and com m issioning of the nuclear power plant. The main characteristics of “product-oriented” quality assurance m ay be summarized as follows:
Existence of a body of technical standards is not of primary importance in performing quality assurance functions. The independent inspection organization is expected to establish a level of control which it deems adequate, and it has a considerable latitude in the interpretation of regulatory and other safety requirements. 2 4 2 RAJSIC
The quality control/assurance systems of manufacturers or other task performers are not necessarily subject to review or audits, and the m ethod ology of their control and verification is not prescribed.
The qualifications and experience of independent inspectors strongly influence the effectiveness of the “product-oriented quality assurance” a p p ro a c h .
The existence of a well organized and skilled inspection organization is re q u ire d .
In this way, design, manufacturing, installation and quality controls performed through inspection and tests are subject to verification by independent inspec tion agencies. The objects of this verification are the timely detection of non conform ances relative to established requirements, and subsequent corrective actions aimed at correcting the deficiencies. A confidence in the future satis factory performance of the items in service is established through redundancy in review, inspection and testing activities, which eliminates the potential errors which m ight occur in any of these activities.
3.2. Quality assurance system
Another method of gaining confidence that items and services have conformed to established requirements is through the functioning of the quality assurance system established by each im portant organization participating in the nuclear power project.
From the regulatory point of view, the responsibility for establishing such a quality assurance system in a nuclear power project, rests with the licensee or organization having the overall financial and technical responsibility for the nuclear power plant. The regulatory requirements for quality assurance are implemented by the licensee through contractual agreement with other construc tors of the plant. To maintain regulatory control over the establishment and operation of the quality assurance system, the general criteria which should be satisfied by the system are established as regulatory requirements. The purpose of these criteria is to oblige all designers and constructors of the plant to plan, conduct and document their work in a systematic way. The immediate objectives of the functioning of the quality assurance system are:
To assure the establishment and implementation of procedures and instruc tions for all activities which are im portant to the quality of nuclear power plant, as they m ay be needed to ensure conformance to existing regulatory require ments, codes and specifications. IAEA-SM-218/51 2 4 3
To control and verify im plementation of these procedures in all activities.
To assure the generation of objective evidence of the conformance to established requirements.
To assure the timely detection of all non-conformances and the prompt undertaking of corrective actions to eliminate the deficiencies and preclude
the repetition of errors.
To keep the quality assurance system under periodic review and assessment in order to assure its effectiveness.
M ost of these objectives should be achieved in the course of project activities, by individuals and groups responsible for direct performance of these activities. Quality assurance system functions are, therefore, more the responsibility of all participants in project activities than the sole responsibility of quality assurance groups or departments. Aside from the regulatory aspect, which requires literal compliance with the specified requirements of quality assurance criteria, the concept of quality assurance might be considered a set of recommended principles for good management of the nuclear power project. These principles should be implemented by project management and all participants in project activities. Functions of quality assurance personnel serve to assure that the system is operating efficiently, and through sam pling inspection or other direct verification methods they create confidence in product quality.
4. DESCRIPTION OF THE QUALITY ASSURANCE SYSTEM
The quality assurance system is sometimes identified as a management system of an organization, which functions for the purpose of ensuring the planned, systematic and controlled approach to all activities affecting quality. In order to perform its functions satisfactorily, the quality assurance system should be organized and operated on the basis of established criteria which prescribe the system elements and allow for the selection of system functions in accordance with the responsibility of the corresponding organization. The International Atom ic Energy Agency, within the framework of its programme of preparation of Safety Codes of Practices and Safety Guides, has prepared the Code of Practice on Quality Assurance for Nuclear Power Plants [3], which contains principles and guidelines for the establishment of a quality assurance system in nuclear power plant projects. This system is briefly described in the following paragraphs. 2 4 4 RAISIC
4.1. Quality assurance system elements
As in any management system, the quality assurance system embodies constituent elements which provide the basis for system operation. The following elements are considered necessary for the efficient implementation of system fu n c t io n s :
Quality Assurance Programme: Represents a documented basis for imple menting quality assurance system functions. It consists of written procedures and instructions for perform ing activities, to ensure conformance with existing requirements, (regulatory requirements, codes, standards and specifications), and of plans for a systematic implementation of all of these procedures.
Quality Assurance Organization: Implementation of system functions is the responsibility of all individuals and groups participating in activities. However, as a general rule, quality control and verification functions are performed by individuals and groups who have not directly participated in the activity being verified and who are independent of other groups. The function of supervising the quality assurance system and its efficiency is assigned to management, supported as necessary by a quality assurance group or organizations.
Quality Assurance Records System: A system of records, including record preparation, collection, maintenance and conservation, is a basis for the evaluation of quality assurance system effectiveness. Records represent the objective evidence of quality, and are used for documenting the conformance of performed activities with established specifications and procedures.
4.2. Quality assurance system functions
The quality assurance system shall provide for a planned and systematic approach to all activities which are im portant to safety. This can be achieved through the control and verification of project activities which are considered to influence product quality. Quality assurance system functions may vary from organization to organization, and m ay depend on types of products or on activities engaged in. Typical functions of the quality assurance system, with respect to the plant owner, are the following:
Document Control: Docum ent preparation, review, approval, issue and distribution are subject to control. Control measures shall ensure that only correct and approved docum ents are used in activities. IAEA-SM-218/51 2 4 5
Design Control: Design process and design documents are subject to control in order to assure that applicable regulatory requirements, codes and stan dards are correctly translated into specifications, drawings, procedures and instructions.
Procurement Control: Regulatory requirements, design bases, and other quality requirements shall be included or referenced in procurement docu ments. Purchasers shall implement regulatory requirements regarding quality through contractual arrangements with the suppliers.
Control o f Purchased Material, Equipment and Services: Purchased material and equipment shall conform to quality requirements specified in procure ment documents. Control measures shall include both the examination of products upon delivery and examination of objective evidence concerning quality during the process of procurement source evaluation and surveillance.
Control o f Processes: Special processes such as welding, heat treatment, non-destructive examination, etc. shall be accomplished under controlled conditions using qualified procedures and by qualified and skilled personnel.
Inspection and Test Control: Inspection and testing of items as a part of quality control activities shall also be subject to control to assure that these activities have been adequately and responsibly administered.
Control o f Test and Measuring Equipment: This area of control may be considered a part of inspection and test control, and has the objective of ensuring that equipment has been regularly checked, calibrated and adjusted.
Identification and Control o f Materials, Parts and Components: Identification and control of items shall be ensured through appropriate markings on the items, or on records traceable to the item. These measures should also provide means for tracing items back to the materials and ahead to their location within an assembly.
Non-Conformance Control: Non-conform ing materials should be subject to control to prevent their inadvertent use or installation.
Corrective Actions: Conditions identified as adverse to quality shall be corrected and appropriate measures taken to prevent recurrence.
Quality Assurance System Audits: Regular and unscheduled audits of the quality assurance system within the organization (internal audits) and of suppliers and other contractors (external audits) shall provide confidence in the efficiency and adequacy of the quality assurance system. 2 4 6 RAlSlC
5. CONTRIBUTION OF QUALITY ASSURANCE TO COMPONENTS RELIABILITY IN NUCLEAR POWER PLANT
Confidence that the plant will operate safely must be based on evidence that quality assurance activities are reducing the probability of some categories of failures. This reduction of failure probability will derive from the following effects of quality assurance activities:
Prevention of component failures through a planned and systematic approach to all activities. This involves an assurance of correct performance of all basic functions in building nuclear power plant, such as in design, manufacturing, construction and operation.
Tim ely detection of non-conformances through verification activities such as inspections and testing and subsequent corrective actions.
System-oriented quality assurance more emphasizes the failure prevention aspect of quality assurance, while product-oriented quality assurance more relies on detection and correction of deficiencies. One can analyse which categories of failures will be m ostly affected by quality assurance. It is believed that quality assurance can reduce or eliminate those failures which are not random in nature and which m ay be associated with com m on causes such as deficiencies in design, manufacture or operation. These failures are m ainly categorized as com m on cause failures. According to some analyses they are the main contributors to plant outages and abnormal occurrences in nuclear power plants [4]. Typical random failures, those which are m ost frequently handled by reliability engineering, m ay not be affected by quality assurance activities. Categories of failures originating in design errors m ay be caused by the lack of specification of functional requirements, design realization faults, engineering errors, etc. An effective defence against these causes of failures is quality assurance in general and quality assurance for design in particular. The function of a quality assurance system is to prevent those causes by assuring that functional require ments, regulatory requirements, codes and standards, are correctly translated into design, that the design process is performed under permanent administrative control and that design verification is strictly performed including, when necessary, testing of prototypes, under m ost adverse design conditions [5]. Categories of failures with an origin in manufacturing and installation errors have their causes m ainly in the application of inadequate standards and procedures which are used in activities, and in deficiencies in quality control. The role of quality assurance during procurement, manufacture, construction and installation is to eliminate these causes of failure. Quality assurance system functions such as IAEA-SM-218/51 2 4 7 procurement control, control of purchased materials, equipment and services, control of processes, test and inspection control, including control of test and measuring equipment etc., should provide for correct performance of activities and should assure that the products are adequately controlled [6,7]. Failures of components during operation which are caused by human error are in a category of failures for which effectively deterrents are administrative control and quality assurance in the operational phase of nuclear power plant. Hum an error in operation may be either an action performed by an operator or an incorrect procedure used in an activity. In the first group one can include the incorrect calibration of components, inadequate testing and other human errors. This could be overcome by adherence to strict calibration and test procedures and by administration and supervision of activities. The second group of errors is caused by the use of incorrect operation procedures either through a direct error or as a consequence of deterioration of standards for operational actions. In both cases a periodic review of operational procedure and subsequent changes can be considered as a requirement of quality assurance for operation of
the plant [ 8 ]. Corrective actions implemented in parallel to all other functions of quality assurance systems are not only a method to correct deficiencies discovered in the quality of the plant components but also as a feedback mechanism for improving the effectiveness of the quality assurance system and to preclude the repetition of errors. Feedback of information from operating experience and in particular of reliability data to design and construction m ay be of particular importance to quality assurance. A n analysis of failure causes m ay discover deficiencies in the quality assurance system and provide the means for corrective actions.
6 . CONCLUSIONS
The reliability of a com ponent is intimately related to its quality. In fact reliability m ay be considered as a quality demonstrated over the course of time. To achieve high reliability in system operation one of the possible methods could be by improving the quality of each component in the system. This method is frequently used to achieve reliable operation of systems im portant to safety in nuclear power plant, such as coolant systems, engineered safety systems, etc. Quality is determined by detailed specification of properties and of methods and techniques for achieving this specification. Quality assurance cannot achieve or improve quality but it can assure that quality objectives are attained through a planned and systematic approach to all activities which m ay affect the product quality. Confidence in achieved quality must be based on the fact that quality assurance can contribute to the elimination or diminished frequency of some categories of com ponent failures. In fact quality assurance can be considered 2 4 8 RAISIC as the best defence against com m on mode failure with causes in design errors, manufacturing and installation errors and in human errors in operation, m ain tenance and repair. Through analysis of quality assurance system functions one can identify all actions directed towards the elimination of com m on causes of failure and indirectly contributing to the decrease of frequency of failure in nuclear power plant systems.
REFERENCES
[1 ] Quality Assurance Terms and Definitions, ANSI-N-45.2.10 (1973). [2] General Principles for Reliability Analysis of Nuclear Power Generating Station Protection Systems, IEEE Std. 352 (1975). [3] Code of Practice on Quality Assurance for Nuclear Power Plants, IAEA (to be published). [4] TAYLOR, J.R., “A study of failure causes based on U.S. power reactor abnormal occurrence reports”, Reliability of Nuclear Power Plants (Proc. Symp. Innsbruck, 1975), IAEA, Vienna (1975) 119. [5] Quality Assurance Requirements for Design of Nuclear Power Plants, ANSI-N-45.2.11. [6] Safety Guide on Quality Assurance for Procurement of Items and Services for Nuclear Power Plants, SG-QA3, IAEA (1977) (Draft). [7] Safety Guide on Quality Assurance for Manufacture of Items for Nuclear Power Plants, SG-QA8, IAEA (1977) (Draft). [8] Safety Guide on Administrative Control and Quality Assurance for Operation of Nuclear Power Plants, SG-QA5, IAEA (1977) (Draft). IAEA-SM-218/49
COMPARISON OF QUALITY SYSTEMS
STANDARDS
J.L. FOW LER Central Electricity Generating Board, Walden House, L o n d o n , United Kingdom
Abstract
COMPARISON OF QUALITY SYSTEMS STANDARDS. Over the years, a variety of Quality Systems Standards have been prepared internationally, specifying customer requirements for quality assurance and quality control systems and procedures which must be in force as a prerequisite to contract placing. These systems standards reflect national requirements and national legislation, and the standards writer’s own interpretation of legislation. The result is that even experienced quality assurance personnel sometimes have difficulty in comparing the various national requirements. The problems developing countries experience in comparing one standard with another must, therefore, be almost insurmountable. This paper discusses and compares the various international legislative requirements, standards and guides and attempts to isolate detailed differences. The documents are compared on a matrix basis. The case is argued for the preparation of international legislation, quality assurance systems standards and guide documents as the only sbnsible means of ensuring satisfactory quality for nuclear power plant throughout the world.
This paper considers the variety of Quality Systems Standards that have evolved nationally and internationally over the last few years and which are still being generated at a rate that is frankly unacceptable. It would seem that as countries become aware of the need for quality assurance requirements, so they find it necessary to identify themselves nationally by writing new standards even though perfectly adequate proven ones already exist. It has been stated that the basic criteria in the selection of a reactor type are safety, costs, reliability and acceptability, and it is a m oot question how far we have yet to advance technically before we can be confident that the m ost im portant requirement of safety can be balanced against the other three criteria. The IA E A and ISO Standards are compared here with the U SA Federal Code 10 C F R 50 Appendix B, the American National Standards Institute A N SI N45.2, the Canadian Standards Association, C SA Z299, the U K Ministry of Defence Standard Mil. Q 9858A as well as other standards mentioned in both the American Society of Quality Control Matrix [1 ] and in the C EG B Matrix [2].
2 4 9 2 5 0 FOWLER
Italy has produced a series of similar standards and so has France, and it seems that before long m atrix comparisons of these standards will be so complex that they will be almost unworkable. There is no doubt that, to the purist, these standards are different and carry different interpretations, but to the practical person who wishes to understand and implement the principles and sound logic that all of these documents try to convey, for all practical purposes there is no difference between these standards. This observation m ay be criticized by the purist, but experience within the United Kingdom , Europe, and in some parts of North America suggests that similar results can be achieved by sensible interpretation by experienced personnel, irrespective of which current systems document is applied — it is all a matter of interpretation and understanding. W hy are these docum ents generated if they are all attempting to specify the same 18 criteria of 10 C F R 50 Appendix B? The problem is one of national legislation and individual requirements, the variation in contractual needs and perhaps the thought that having an identifiable domestic standard that has credibility m ay be to some commercial advantage. The motive m ay even be the very basic desire to have one’s own document. These are the sort of reasons why standards proliferate, so what is the answer to this wasteful effort? There are m any sound national approaches being made: an example is the American National Standard, A N S I N45.2, which has achieved credibility not only throughout the United States, but also internationally. The Canadians have a National Standard C SA Z299 and the United Kingdom has an accepted Utility Standard C EG B QA42 which will in due course be replaced by a revised British Standard BS 5179. It should be noted that these last two standards are applicable to both nuclear and fossil-fired power plants. M any other European countries are also developing standards, so it seems that we are now progressing towards rationalization of standards on a national basis, but this is giving rise to a further problem in that more countries are becoming aware of the need for Quality Assurance Standards and therefore the number of national standards is rapidly increasing. The answer to this problem must lie in an international approach. There are currently two approaches being made. Firstly by the IA E A , which has produced a Code of Practice for Quality Assurance supported by Safety Guides. Secondly, The International Organization for Standardization (ISO ) has produced a W orking Draft Docum ent for Quality Assurance for Nuclear Power Plants, which has sought compatibility of Quality Assurance principles between 10 C F R 50 Appendix B, A N SI N45.2 and the IA E A Code of Practice. One m ight quite reasonably ask the question, therefore, w hy are these two approaches being made, is there not true com patibility between the two documents, and which of the organizations has sufficient international standing to influence international legislation for implementation. IAEA-SM-218/49 2 5 1
In June 1976 a joint meeting [3] was held between the Am erican Nuclear Society (A N S) and the Canadian Nuclear Association (CN A) to discuss the International Im plication of Nuclear Quality Assurance Standards. The require ments and objectives of both the ISO and IA E A documents were discussed and the meeting concluded that both were necessary. It appears however just as irrational to have a dual international approach as it does to have a m ultiple national approach and here we have a situation that will have to be resolved between these two international groups. A s to the relative merits of each international document so far produced, there is no doubt that the IA E A Code of Practice and supporting Safety Guides are more comprehensive. There is also an interchangeability between Codes of Practice on Quality Assurance and Codes of Practice on other aspects of Nuclear Power Plant Construction and Operation, and I am inclined therefore to the view that the IA E A is the correct vehicle, and whilst the Code of Practice is being prepared as a ‘hard’ document, using the verb ‘shall’, it is in fact a statement of Q A principles that can be adapted to suit most, if not all, circumstances. Let us now take from 10 C F R 50 Appendix В one of the 18 criteria and discuss over a relatively broad band the difference within the various systems standards, demonstrating by this means the point made earlier, that m any of the problems that are usually discussed at length when writing these different standards are not necessarily valid. The criterion here selected is “Organization”, but before it is examined in detail, here are a few of the definitions of ‘Quality Assurance’, which after all is the fundam ental title of the total discipline:
(1) The American Code of Federal Regulations 10 C F R 50 Appendix В states:
A ll those planned and systematic actions necessary to provide adequate confidence that a structure system or component will perform satisfactorily in service. Quality Assurance includes Quality Control.
(2) American National Standard A N SI N45.2:
A ll those planned and systematic actions necessary to provide adequate confidence that a structure system or component will perform satisfactorily in service. Quality Assurance includes Quality Control.
(3) ASME III NA 4000:
A ll those planned and systematic actions necessary to provide adequate confidence that all items are manufactured or installed in accordance with the rules of Section II, etc. 2 5 2 FOWLER
(4) The Canadian Standards Association C SA Z299:
A planned and systematic pattern of all means and actions designed to provide adequate confidence that items or services meet contract and jurisdictional requirements and will perform satisfactorily in service.
(5) British Standard 4779 Glossary of General Terms used in Quality Assurance:
A ll activities and functions concerned with the attainment of quality.
( 6 ) United Kingdom CEGB QA42 Series:
A ll those planned and systematic actions necessary to provide adequate confidence that an item or a facility will perform satisfactorily in service.
(7) International Atom ic Energy Agency Code of Practice for Quality Assurance:
A ll planned and systematic actions necessary to provide adequate confidence that an item or a facility w ill perform satisfactorily in service.
( 8 ) International Organization for Standardization W orking Draft for Quality Assurance for Nuclear Power Plants:
A ll those planned and systematic actions necessary to provide adequate confidence that a structure system or component will perform satisfactorily in service.
Here I have demonstrated that in eight recognized and proven documents variations exist, although to the experienced Quality Assurance Engineer the interpretation he places upon them does not differ, unless he so wishes it. It is in this interpretation that difficulties arise, and if we cannot find a means of being consistent with the definition of the title of our profession then it does not stimulate confidence in the interpretation of the 18 criteria. Organization as a leading criterion is fairly consistent within the various documents, and the order in which the criteria are set down is again consistent with 10 C F R 50 Appendix B, which again brings one back to the earlier conclusion that it is perhaps all a matter of national legislation and identity. Let us now consider the single criterion “Organization” : 10 C F R 50 Appendix В is extremely instructive, but rather long and detailed; it assigns responsibilities, authority and duties. It also states that the persons and organizations performing quality assurance functions shall have sufficient authority and organizational freedom to identify quality problems, to initiate, recommend, or provide solutions and to verify implementation of solutions. It goes on to say m uch more about organization, and this particular point is here emphasized because it, or variations of it, form a com m on theme in other documents. IAEA-SM-218/49 2 5 3
A N SI N45.2 Consolidation is less detailed than 10 C F R 50 Appendix B, but in slightly different wording it repeats the requirements and here again the theme is the same: “Persons or Organizations responsible for assuring that an appropriate Quality Assurance Programme is established, and for verifying that activities affecting quality have been correctly performed shall have sufficient authority and organizational freedom to ( 1) identify quality problems, ( 2 ) initiate, recommend or provide solutions to quality problems through designated channels, (3) verify implementation of solutions, and (4) control further processing delivery or installation o f a non-conforming item, deficiency or unsatisfactory condition until proper dispositioning has occurred, etc.”. Although it is very similar to 10 C F R 50 Appendix В we see the introduction of two new ideas: (a) reference is made here to a Quality Assurance Programme, (b) a ‘stop w ork’ situation and control of further processing. A SM E N A 4000 is very similar to A N SI N45.2, which is written differently but emphasizes three of the main points of the com m on theme: (a) Identify Quality Assurance Problems, (b) Initiate actions which result in solutions, (c) Verify implementation of solutions to those problems. It does not directly make reference to control of further processing. The Canadian Standards Association Docum ent Z299.1 draws on the requirements of A N S I N45.2 and includes the requirements for control of further processing. The ISO working draft is very similar in wording to 10 C F R 50 Appendix В and A N S I N45.2, for example, “shall have sufficient authority and organizational freedom to identify quality problems, to initiate, recommend or provide solutions” etc. It is seemingly identical, but ISO , like A N S I N45.2, slips in a further require ment — “and where necessary control, further processing, delivery or installation of a non-conform ing item, deficiency, or unsatisfactory condition until proper dispositioning has occurred”. This is a specific ‘stop w ork’ clause and ‘Advantage Server’ in tennis parlance to ISO , but the devotees of Q A42 for example will quickly reply that this is an unnecessary clause because their docum ent refers to mandatory hold points beyond which work shall not proceed without consent — ( ‘D e u c e ’).
British Standard BS 5179 Part 3 is currently being revised to take into account international requirements, so will not be commented upon here. C EG B Standard QA42 is very brief. It makes the point that “effective management for quality shall be clearly presented in writing by the contractor and that the contractor shall delegate to those persons perform ing quality functions both the responsibility and the authority to identify and evaluate quality problems, and to initiate, recommend or provide solutions during all phases of the contract. The contractor shall appoint a management representative, independent of other functions, with authority to resolve matters pertaining to quality to the satisfaction of the engineer”. 2 5 4 FOWLER
Here again we have this com m on theme plus a Quality Assurance Manager with authority and independence. The United Kingdom Ministry of Defence Docum ent 05.21, which is also applicable to N A T O countries, is almost identical with Q A42 except that ‘independent’ is qualified by ‘preferably’. What is the difference apart from endless discussion by the experts debating the differences between a m andatory require ment or a matter of interpretation?
The IA E A Code of Practice on Quality Assurance has a very comprehensive section on organization. It covers all the requirements of the documents so far mentioned, and in addition it is to be supplemented by detailed Safety Guides currently under preparation. In summary, what conclusions can be drawn from an analysis of the documents? Are they essentially the same and are we not in effect indulging in semantics, i.e. the significance and meanings of words rather than the setting down in simple terms the basic principles of Quality Assurance for sensible interpretation by competent persons? I would recommend to any organization that is about to embark on the extremely costly task of preparing Quality Assurance Standards that it very carefully considers the existing standards, adopts one, and spends the m oney on training staff to interpret it carefully, because it does not matter how good the Quality Assurance Programme is, it is the quality of the people who interpret it that counts.
REFERENCES
[ 1 ] Matrix of Nuclear Quality Assurance Programme Requirements, 2nd Edn, American Society for Quality Control, Milwaukee (1973). [2] Matrix Comparison of Quality Assurance Programme Requirements CEGB QA Publications, London, 1 1 (1975). [3] MARASH, S.A., Introduction to Panel Session on the International Implication of Nuclear Quality Assurance Standards, ANS-CNA Joint Meeting, Toronto, 14 June 1976, Nucl. Standard News Special Report (July 1976). DISCUSSION
on papers IA EA -SM -218/49 and 51
G. Ö ST B ER G (Chairman): Before asking Dr. Raisic a question there is one category of failure cause am ong those listed by him on which I would like to comment, and that is the category of so-called “human errors” or rather “systems errors”, as I prefer to call them. These errors have considerable bearing on confi dence, which, according to Dr. Raisic, is what quality assurance is aimed at. By confidence I mean the confidence not only of manufacturers and utilities but also that of the public. In Sweden we have made a study of so-called “inconceivable” events occurring in connection with the handling of materials in a heavy mechanical engineering industry. I would like to summarize the results of this study, which will also be published in the form of a report in November 1977. The work to be described in the report is the introductory part of a project on certain risks of catastrophic failure in welded nuclear pressure vessels. The probability of such a failure occurring is generally considered to be of the order of 1 0 " 5 t o 1 0 ~ 6 p e r reactor-year. This probability estimate is based either on statistics for actual failures of similar pressure vessels or on calculations made on the basis of proba bilistic fracture mechanics. A ll causes are of course included in the failure statistics, while the probabilistic fracture mechanics are based on statistical data for stress, properties of materials and defects. In addition to these factors, however, we m ay also have to reckon with events that cannot be treated by conventional statistical methods. In principle such so-called “inconceivable” events should be covered by quality assurance systems which are set up to prevent, as far as possible, fractures occurring as a result of defective design, manufacture and control of nuclear pressure vessels. Nevertheless, deficiencies in the reliability and effectiveness of such quality assurance systems cannot yet be excluded, and it was for this reason that the present study of inconceivable events was initiated in a special project. Part of the problem lies in the fact that it cannot be treated solely by the usual technical and scientific methods, otherwise it would already have been solved by normal fracture mechanics calculations. Instead we have to approach the incident we call “inconceivable” from a broader industrial and psychological point of view. This will make special demands, not only with regard to the planning and performance of the work itself, but also with regard to the way in which the results and conclusions of the study are to be interpreted. Since an industrial and psychological analysis of the manufacture of nuclear pressure vessels can be considered a new approach embarked upon both by those involved in this work and by the pressure vessel manufacturer, we felt it would be advisable to make a preparatory study in a comparable area of the heavy mechanical engineering industry.
2 5 5 2 5 6 DISCUSSION
A team of specialists in materials technology, reliability technology and industrial psychology performed a field study on the handling of certain material components of a company of this type. Different stages in the handling and processing of the material used by the com pany were observed, particular emphasis being placed on deviations from or exceptions to the form al quality assurance systems. In these studies the management of the company and those staff who were involved co-operated in a very positive and committed manner. To give examples of the observations in a summary like this would be difficult, since there is a risk that attention will be drawn too m uch to specific technical situations, while the principal features of the incidents m ay not be understood as clearly. Another difficulty lies in the fact that the language and concepts which have to be used in order to describe and analyse human actions and psychological relationships differ from com m only used technical language and concepts. Interesting observations made about inconceivable events can be characterized according to the following categories, which are defined in “hard” technological te rm s:
Undetected failures in the material;
Lack of knowledge of working instructions;
Occasional failure on the part of personnel to observe routine in complicated working situations;
Unregulated delegation of control.
It should be pointed out that none of the events which we observed could in any way jeopardize the functioning of the com pany’s products. Yet they were of definite value for this project in that they provided a basis for conclusions to be made concerning inconceivable events which m ight occur during pressure vessel manufacture and which could cause catastrophic failure. In the light of the psychological concepts established for the events studied, certain inferences could be drawn about how a system designed to ensure the quality of processes and products would function and how individual workers would act. In principle, these conclusions would also be true for the form al safety system in its practical application. A s intended, the project yielded valuable knowledge and experience on the methods of study. Am ong the other achievements I should like to m ention is, first, the very fact that the w orking team succeeded at all in winning over the com pany to the idea of studying topical problems and in getting the staff concerned to com m it themselves to co-operating. Secondly, the project group managed to obtain information by holding discussions and m aking observations. Thirdly, experience was gained in the analysis and systemati zation of inform ation concerning the reasons for the differences between the quality system and its practical application. DISCUSSION 2 5 7
On the basis of the study performed so far it is not possible to establish any precise figures showing the effect that inconceivable events can have on the risk of catastrophic failure of nuclear pressure vessels. But this was not our intention; the knowledge and experience obtained with regard both to inconceivable events as such and to methods of studying them were such that the aim of the preliminary study can definitely be considered to have been achieved. And now I would like to ask Dr. Raisic to enlarge on the meaning of “confidence” in relation to reliability. N. RAlSiC: I have tried to show in my paper that confidence in the quality of a nuclear power plant should be based on the effectiveness of quality assurance activities in preventing errors in design, construction and operation and in discovering and correcting whatever does not conform with design in nuclear power plants. Quality assurance can contribute to the elimination or the reduction of some categories of failure in nuclear power plants, and especially common-cause failures. It is the regulatory organization which needs to have confidence in plant quality (quality is used here to mean conform ity with regulatory requirements, codes, standards and specifications). Regulatory organizations act on behalf of governments and represent the public interest. Confidence in quality is a pre requisite for the acceptance of a nuclear ро\лег plant or, more precisely, it is on the basis of confidence that the construction permit or an operating licence can be issu e d . S.M. BH U TTA: Operators have too much paperwork to do for quality assurance, and this affects availability and plant factors. Is it possible to find a solution or a compromise for this problem? N. RA I§IC: Excessive documentation in connection with a quality assurance programme is more likely to be a consequence of the misinterpretation of existing regulatory requirements and industrial standards than an inherent characteristic of quality assurance systems. S.M. BH U TTA: There appears to be some controversy about quality assurance procedures during plant design. Can you comment on this? N. RA I§IC: The IA EA is preparing a volume with the title “Safety Guide on Quality Assurance for the Design of Nuclear Power Plants”. This document will be available early in 1978. However, there is a United States standards docu ment entitled “Quality Assurance Requirements for the Design of Nuclear Power Plants” (AN SI-N 45.2.11), which covers the same subject. L. ALVAREZ DE BUERGO: Mr. Raisic and Mr. Fowler have discussed in their papers two of the most im portant questions: that of international standardization and that of defining what is mean by quality assurance and reliability. Countries like Spain, for example, which have large-scale nuclear programmes and which receive components from various sources, need to have standards which are of a compatible nature. We usually use the IA E A standards as laid down in the Code of Practice and believe that they should be the basic 2 5 8 DISCUSSION
standards used, but we are very concerned about the apparent diversity in ways of implementing these standards. Some standards are designed for individual countries or for industrially developed countries as a whole, while there are others which are accepted internationally but which need updating and approval in terms of international consensus. I think the IA E A could help to ensure that these standards are in fact compatible. J.L. FO W LER: In m y paper I tried to put across the idea that, although different countries and different organizations, such as the IA E A and the Inter national Organization for Standardization (ISO ) write different standards, these standards are basically all aimed at the same things. So it does not matter much which standard you use; what is important, however, is the training and degree of competence of staff, because the key question is that of whether staff are capable of satisfactorily interpreting the standard used. L. ALVAREZ DE BUERGO: According to Mr. Raisic, quality assurance has nothing to do with the establishment of the level of quality to be achieved. I do not believe that this definition goes far enough, since one of the aims of quality assurance should, through feedback, etc., be actually to improve the quality so that the overall objectives laid down can be reached. J.L. FO W LER: In m y mind quality assurance clearly does not lay down the level of quality to be achieved, but specifies the discipline that must be maintained in order for the desired level of quality to be achieved. In short, quality assurance means sim ply the assurance that we are all doing our jobs properly. IAEA-SM-218/35
QUALITY ASSURANCE REQUIREMENTS FOR THE
RELIABILITY OF NUCLEAR POW ER PLANTS
IN DEVELOPING COUNTRIES
S.M. BHUTTA Chashma Nuclear Power Project, Pakistan Atom ic Energy Commission, Is la m a b a d , P a k is t a n
Abstract
QUALITY ASSURANCE REQUIREMENTS FOR THE RELIABILITY OF NUCLEAR POWER PLANTS IN DEVELOPING COUNTRIES. Quality Assurance (QA) has not been taken very seriously in developing countries despite their having significant nuclear power programmes. This may affect plant reliability and demands that urgent and serious attention be given to the development and implementation of Quality Assurance programmes according to local conditions. Confusion has been created by the differences in definitions and interpretations of terminologies of Quality Assurance and Control. Problems have been aggravated by the lack of clearly defined responsibilities and accountabilities during the projects execution phases of site selection, data collection, design, equipment fabrication and construction, etc. Therefore,reliability of nuclear power plants in developing countries is relatively low. But whenever some programme of QA has been imple mented it has helped to improve the plant performance. This is highlighted in this paper by the practical examples from the experience of Karachi Nuclear Power Plant whose forced outages have been reduced by over a half within a period of 3 years. In view of the benefits,a QA pro gramme for Chashma Nuclear Power Project has also been initiated. In this paper urgency is emphasized for the establishment and implementation of a formal QA programme in the develop ing countries if the reactor suppliers and purchasers both want to ensure higher reliability of their plants. The best way that management can play its role effectively is by setting up a strong QA organization with local personnel, thus helping to attain self-reliance and higher reliability during plant operation.
1. I N T R O D U C T I O N
The energy crisis and the continuously rising oil price spiral has forced developing countries to pay more attention to the installation of nuclear power plants. Although at present there are only 11 nuclear power plants in operation in 5 developing countries, in several developing countries many nuclear power plants are either under construction or in an advanced stage of planning. It is expected that by 1990 there will be about 54 developing countries which would have 356 nuclear power reactors with a total installed capacity of 220-GW (e) [1].
2 5 9 2 6 0 BHUTTA
Although the time schedule m ay be uncertain it is now a well-recognized fact that nuclear power is bound to play an increasingly im portant role in the electricity production of developing countries. But so far the concept of Quality Control and Assurance has not attracted any serious attention in the developing world. This could effect the reliability of nuclear power plants and thus might hamper the power programmes of m any countries. Therefore,there is an urgent need for the development and implementation of Quality Assurance programmes according to local conditions.
2. THE IMPORTANCE OF QUALITY ASSURANCE
According, to the definition of American National Standards Quality Assurance (Q A ) includes “All those planned or systematic actions necessary to provide adequate confidence that an item or a facility will perform satisfactorily in service” [2]. Q A is applicable to complex projects which require original and extensive design, development, manufacturing, construction and installation works. The failure of such a project m ay result in undue risk to health and safety or total loss. Therefore, the Q A programme applies to all activities affecting the safety- related functions of all those structures, systems and components. These activities include designing, purchasing, fabricating, handling, shipping, cleaning, erecting, installing, testing, operating, maintaining, repairing, refuelling and modifying,etc. To all these activities there are m inim um mandatory requirements of quality assurance as specified in the codes, regulations and specific criteria. This way, when a Q A programme is properly implemented it would ensure that nuclear power plants would operate reliably and safely.
3. RESPONSIBILITY FOR QUALITY ASSURANCE
It is the responsibility of the nuclear power plant owner to establish and execute the Q A programme [3]. However, the basic responsibility for achieving quality in perform ing a particular task lies with the individual or individuals assigned the task and not only with those seeking to assure by means of quality assurance verification that it has been achieved. The national regulatory body has responsibilities for the development of codes, guides and criteria for quality and safety requirements. It has also to review, approve and audit the Q A programmes of the organizations. The industry and owners have the responsibility to ensure that Q A programmes encompassing these codes and guides are developed and implemented properly. In most cases developing countries are at the receiving end, so it is urged that the suppliers of nuclear power plants and components make sure that an appropriate Q A pro gramme is implemented by the local personnel. IAEA-SM-218/35 2 6 1
4. THE STATUS OF QUALITY ASSURANCE IN DEVELOPING COUNTRIES
Unfortunately most of the developing countries lack quality consciousness. The situation has been made worse by the confusion created by the different definitions and interpretations of Quality Assurance and its terminology. In most developing countries there are no independent organizations and regulatory bodies having full competence and authority for the establishment and imple mentation of Q A programmes. Usually the nuclear power plant owners them selves are the regulatory bodies, authorized by the respective governments of these developing countries. Experience has shown that in some of the cases safety rules and codes are not fully observed because of cost and schedule considerations. Developing countries also lack technical know-how and have a poor industrial infrastructure. Therefore supply and construction of nuclear power plants are usually on turn-key contracts awarded to the industrially advanced countries. In some cases non-turn-key contracts are also tried out but even then the active participation of the owners in the engineering w ork is limited. Therefore, transfer of sophisticated technology has been quite lim ited,thus creating difficulties in the operation and maintenance of these plants because the contractors lose interest after handing over. Difficulties arise because of lack of back-up technical information, trained specialist man-power, spare parts, documents, as-built drawings and communication with industry and the promoters and suppliers o f the reactor. Design inputs, especially the site characteristics, are not available with any great degree of accuracy and do not cover a sufficient period of time for us to have full confidence in the reliability of plants over their lifetime. Inform ation related to geology, seismology, meteorology, hydrology, chemistry of site and even temperature variations of the cooling water requires thorough investigations and so does the manner in which these would be used in plant design. In some of the developing countries management considers that Q A can be ensured by hiring a foreign expert and two or three local people and referring to them as the Q A staff. At the same time m ost of the developing countries desire to participate fully in design, fabrication, construction and other engineering services in order to achieve m axim um self-reliance during operation and main tenance of their power plants.
5. COMPARISON OF RELIABILITY OF NUCLEAR POWER PLANTS IN DEVELOPING AND INDUSTRIALIZED COUNTRIES
The need for Q A programmes in developing countries may be emphasized by illustrating that the availability and load factors of the same type o f nuclear power plants are lower in developing countries than those in industrialized 2 6 2 BHUTTA countries. This can be illustrated by the Canadian built C A N D U type PHW R Pressurized Heavy Water Reactors of K A N U P P (Pakistan), Rajasthan-I (India) and Pickering-I (Canada) for the years 1974 and 1975 as this information is published by the IA E A [4]. During 1974 the Operation Factor (O F) of Pickering-I was 79.9% and Load Factor (LF) 72%. It had only 6 outages, of which 3 were due to equipment failures and 2 due to operation error. The operation factor of Rajasthan-I during 1974 was only 53.5% and LF 36.8%, because it had 25 outages. O f these 4 were due to operators’ error while the other 21 outages were due to equipment failures. During that year the O F of K A N U PP was 77% and L F 52.8% as it had 17 outages, 13 due to equipment failures and 3 due to operators’ error. During 1975 at Pickering-I there were only 7 outages, of which 3 were of conventional equipment, that is due to hydrogen leaks into the generators’ stator cooling system, and 3 were due to control equipment failures. This way its LF was 80.2% and O F 82.8%. The Load Factor of Rajasthan-I during 1975 was only 32.2% and O F 43.6%. The main reason for such a poor performance was 15 outages, of which 3 were due to operators’ error and the rest due to equipment failures. The Operation Factor of KA N U PP during 1975 was 72.8% but its LF only 44.8%. It had 14 outages, of which 2 were due to problems in the grid system, 11 were due to equipment failure and one due to operators’ error [4,5]. The above analysis shows that reactors of the same origin and type built in developing countries have more equipment failures and operators’ errors resulting in a poorer performance than that of reactors in industrialized countries. This performance improved when steps were taken to apply quality control, as explained in the next section.
6 . THE IMPACT OF QA PROGRAMMES IN DEVELOPING COUNTRIES
Although Q A programmes are not form ally prepared and organized in most of the developing countries they are being implemented by local engineers directly or indirectly. Their impact has been felt in improved reliability of the nuclear power plant during operation and maintenance. Some examples are quoted here to highlight the achievements whenever there was any necessity and in c e n tiv e . Karachi Nuclear Power Plant (KAN UPP), which went into commercial operation in late 1972, has shown an interesting performance up to 1976, as reported in Table I [5,6]. The total number of forced outages, mainly due to equipment failure, has steadily decreased, while outages due to grid problems have not decreased much and outages due to operators’ error have been changing IAEA-SM-218/35 2 6 3
TABLE I. TOTAL OUTAGES OF КANUPP (see Refs [ 5 ,6 ])
Causes of outages 1973 1974 1975 1976
Equipment failures 18 13 11 5
Operators’ error 3 3 1 4
Grid system problems 3 1 2 2
Other 2 - - -
Total outages 26 17 14 11
as explained below. It may be kept in mind that no formal Quality Assurance programme was followed either during design, fabrication, construction or during commissioning and initial operation of KANUPP. 18 forced outages due to equipment failures in 1973 were steadily reduced, with only 5 in 1976. This is because the initial teething problems were resolved by extensive modifications, preventive maintenance and in-service inspection. The major problems were leakages of heavy water and faults in controlling computers because of the m alfunctioning of various items of equipment. Frequent failures of transistors in the analog input-hardware, breakdown of parallel communication links between the two control computers and mal functioning of high-gain multiplex amplifiers are some of the faults which were identified. Trips of K A N U PP because of computer breakdowns were overcome by applying quality control actions and carrying out improvements in its hardware as well as in software programmes in phases. Heavy water leakage from the reactor system was one of the major factors affecting plant availability. During 1973 K A N U P P experienced 5 outages of comparatively longer duration because of abnormal leakage of heavy water. There are 208 end-fittings of the fuel channels of the reactor having removable screw-in type end-plugs to facilitate on-power fuelling. These end-fittings are fixed on one end of the reactor and left floating at the other end to allow for expansion of the pressure tubes. A large number of the end-fittings started leaking on experiencing power transients and cycling. The failure rate was higher at the floating end of the fuel channels. Since m ost leaks of this type are very minor, appearing in the form of mist, and since there was no detection equipment for such leaks, their remote detection during on-power operation was quite a problem. Pakistani engineers, faced with this challenge, applied quality control techniques and developed a new instrument for the detection of such m inor leaks. During 2 6 4 BHUTTA the early part of 1974 a leak detection device was developed and installed on each of the fuelling machines to provide routine scanning of the reactor end-fittings and accordingly to enable corrective measures to be taken [10]. Since then plant outages due to uncontrolled leakage from reactor faces have been reduced. Impressed by such quality control actions, the maintenance management have established a quality control section and in-service inspection facilities at KANU PP. A fully-fledged Q A department is desirable: this would be able to provide m uch better service and increase the reliability of plant during operation. Operators’ errors resulted in three costly outages during the initial years of plant operation. This was m ainly because the operators did not have adequate experience and when management paid attention to this deficiency this problem was solved in 1975. But when most of the trained man-power left plant operation the problem again arose in 1976, resulting in 36% outages of K A N U P P because of inadvertent operator errors. A rigorous programme of in-plant training of technicians and engineers has been started at K A N U P P to help overcome both the hand- and brain-drain.
7. THE QUALITY ASSURANCE PROGRAMME FOR CHASNUPP
Pakistan has undertaken site development work for the construction of her 2nd nuclear power plant of 600 MW (e) at a place called Chashma. Most of the suppliers of the nuclear power plant and the IA E A have strongly recommended that every applicant for nuclear plant construction permits should include in its preliminary safety analysis report a description of the Quality Assurance pro gramme. Faced with this requirement and also appreciating the benefits of QA, the Pakistan Atom ic Energy Commission (PAEC) has appointed a group in CHASHM A NUCLEAR POW ER PROJECT (CHASNUPP) for the development of its Quality Control and Quality Assurance programme. Even though this group is limited in staff several positive actions have been taken in order to prepare and implement a Q A programme. Surveys of the local industry and engineering firms have been carried out with a view to finding out about their Q A programmes. It is noted that in most of the firms such formal programmes do not exist. Therefore, efforts are made to familiarize the management and to establish Q A groups in important engineering organizations. The response was quite encouraging after we had delivered lectures and held discussions about the importance and implementation of Q A programmes in fulfilling the nuclear power programme of Pakistan [7]. A seminar about Q A and training programme for local engineers and technicians is planned by CHASN U PP. In the bidding documents prepared for C H A SN U PP the bidders’ Quality Assurance programme is required to be governed by the U S Code of Federal Regulations, 10 CFR-50 Appendix В [8]. The requirements of this Appendix, IAEA-SM-218/35 2 6 5 known as the “Eighteen Criteria”, are recognized as providing the m inim um basis for a sound and standardized Quality Assurance programme. It is required of contractors that their Q A programme should meet each of these eighteen criteria. The Bidder may, if he so elects, observe alternative rules, regulations and standards in his Quality Assurance programme, provided that the Bidder proves that the rules, regulations, and standards observed in his Quality Assurance programme are equivalent to those quoted in the C H A SN U PP specification. Before award of contract, the contractor’s Q A programme would be reviewed by PAEC. This review may involve audit, surveillance and inspection of the con tractor’s and sub-contractors’ facilities or engineering services, etc.
8 . QA REQUIREMENTS IN DEVELOPING COUNTRIES
Although industrialized countries have developed and are carrying out Q A programmes, because of local conditions and constituent activity levels, these will vary from country to country and organization to organization. Q A encompasses m any functions and activities and extends to various levels in all participating organizations, from the top executive to all the workers. Therefore a Q A pro gramme can be properly and effectively implemented only if it is carried out with the m axim um participation of local staff. Because the developing countries have largely to depend on industrial nations for the supply of design, equipment, and technical know-how, top management can play its role most effectively by setting up its own strong Q A organization. Local staff should have easy and direct access to responsible management at a level where appropriate action can be taken. Quality Assurance is an essential aspect of “Good Management” [9] because it contributes to the achievement of quality through analysis of the task to be performed, identification of the skill required, the selection and training of appro priate personnel and the creation of a satisfactory environment in which activity can be performed. Briefly stated,a Q A programme can provide for a disciplined approach to all activities affecting quality. This way a Q A programme may appear to be the key element in the success of nuclear power programmes in the develop ing countries. It is also realised that a Q A programme when carried out by local personnel would help to reduce the level of dependence,making the whole pro gramme lead to the achievement of self-reliance in time. The following actions require immediate attention in developing countries so as to help these countries achieve the goal of safe and reliable operation of their nuclear power plants.
(1) Setting up of an independent group for Q A activities, with well defined responsibility and authority of the personnel and organizations involved. 2 6 6 BHUTTA
(2) Developing and documenting the quality assurance programme for nuclear power plant, keeping local conditions in view.
(3) Indoctrination and training of personnel performing activities affecting q u a lit y .
(4) Preparation of codes, regulations, standards, procedures, specifications, criteria and other applicable principles.
(5) Verification of design inputs and bases such as: (a) Loads, such as seismic, wind, thermal and dynamic, etc. (b) Environmental and site conditions, such as temperature, humidity, pressure, site elevation, wind direction, chemical analysis, radiation background, etc.
(6) Establishment of standard laboratories for measuring, testing and inspection instruments, tools, gauges and equipment for calibration and control purposes.
9. CONCLUSION
In developing countries the need for Q A is not fully appreciated. A ll persons from top management to ordinary workers and engineers have to be made quality conscious and indoctrinated. The Q A programme has not only to be written or preached but it has to be implemented and practised in its true sense. Management in developing countries can play its role m ost effectively by setting up strong Q A organizations with local talent in the fields of data collection, design, procurement, fabrication and construction. In this way the developing countries could attain self-reliance in maintenance and operation. A Q A pro gramme,when fully implemented by local staff, would also help to improve the quality of all local products and services. This would boost the econom y of these c o u n trie s. Thus Q A represents a vital management tool in maintaining a disciplined approach towards more safe and efficient plant construction and operation. This is the only way to fulfil effectively and economically the programmes calling for rapid growth of the nuclear power industry which have coincided with the awakening of public interest in the issues of reliability, safety and the environment.
ACKNOWLEDGEMENT
The author is grateful to Mr. Muslim Kidwai, Assistant Engineer, for the help he has provided in the preparation of this paper. IAEA-SM-218/35 2 6 7
REFERENCES
[1] INTERNATIONAL ATOMIC ENERGY AGENCY, IAEA Bulletin Vol.17, No.3, June 1975. [2] Quality Assurance Terms and Definitions ANSI N.45.2.10.1973, published by the American Society of Mechanical Engineers. [3] USA Code of Federal Regulations 10-CFR-50 Appendix B. [4] Power and Research Reactors in Member States, IAEA, Vienna (1974); Power Reactors in Member States, IAEA, Vienna (1975). [5] Safety Aspects of KANUPP Operation for years 1973, 1974, 1975. [6] JAFRI, M.N., Production Capability Report of KANUPP for 1976. [7] BHUTTA, S.M., lectures delivered on “Importance and implementation of QA programme for nuclear plants” (unpublished). [8] Specifications for Chashma Nuclear Power Project, Volume III, Appendix A. [9] Safety Code of Practice on Quality Assurance, IAEA, Vienna (to be published). [10] MAHMOOD, S.B., Detection, location and estimation of light water and heavy water leaks in nuclear reactors, Ref. No. KANUPP STR-74-4, Patent No. 12457.
DISCUSSION
J. FORSTfiN (Chairman): What measures on the quality assurance side have been taken to reduce reactor outages? Have you analysed equipment failures in order to make improvements in the quality assurance system used during equipment manufacturing? S.M. BH U TTA : In-service inspection, preventive maintenance and extensive modifications were carried out to reduce equipment failures. Engineers and technicians are being trained to reduce human error. Equipm ent failure causes have been analysed and this has helped to improve the design and fabrication techniques for reducing failures (e.g. leakage of D 20, heat exchangers and condensers). A. K A K O D K A R : In the section of your paper entitled “Comparison of Reliability of Nuclear Power Plants in Developing and Industrialized Countries” you compare the Pickering-I and Rajasthan-I reactors. While it is true that both are pressurized heavy water reactors there are very significant differences in the hardware of the two systems. If a comparison of this nature has to be made at all, it would be more appropriate to compare the Rajasthan-I reactor with the Douglas Point reactor. Furthermore, a major factor in reliability seems to relate to the conventional system and the grid system. S.M. BH UTTA: Pickering-I, Rajasthan-I and KA N U PP were selected for comparison in order to illustrate the impact of quality assurance on reactors which went into operation at almost the same time and which were of the same type and had the same origin, but which were located in different countries. Data on the reasons for outages of these plants are available in the IA E A ’s series entitled “Operating Experience with Nuclear Power Stations in Member States”.
IAEA-SM-218/36
SOME ASPECTS OF QUALITY
CONTROL PROBLEMS EXPERIENCED IN
NUCLEAR COMPONENT MANUFACTURE
V . S . G . R A O Power Projects Engineering Division, Department of Atom ic Energy, B o m b a y , In d ia
Presented by A. Kakodkar
Abstract
SOME ASPECTS OF QUALITY CONTROL PROBLEMS EXPERIENCED IN NUCLEAR COMPONENT MANUFACTURE. The Pressurised Heavy Water Reactor (PHWR) has been adopted by India for all its first-generation nuclear power stations.. This type of reactor contains a number of massive and sophisticated reactor components which have to be built to the stringent quality require ments of nuclear specifications. In the Indian nuclear programme great emphasis has been laid on achieving self-sufficiency in the manufacture of nuclear components. Due to sustained efforts over the past few years this goal has almost been reached. It has been a challenging task for Indian industry to take up the manufacture of sophisticated nuclear components and as a result a new outlook on quality has emerged. Many problems were faced in achieving the desired quality standards during the manufacture of nuclear components. Some aspects of these problems relating to raw material, fabrication, weld quality and non-destructive examination are discussed. As an illustration, quality control problems experienced during the manufacture of the reactor vessel are highlighted. Indian industry has demonstrated that the stringent quality standards required by nuclear specifications can be achieved by a developing country with limited resources. While the efforts made to date are commendable, still greater efforts are required to strengthen quality control departments in order to maintain effective quality control.
1. INTRODUCTION
From the early stages of development of nuclear power stations in India, great emphasis has been laid on achieving national self-sufficiency in the manu facture of nuclear components. Over the past few years rapid strides have been taken to achieve this goal and a stage has now been reached when practically all the reactor components required for nuclear power stations are being made in the country. Although many industries had previous experience in the manufacture of heavy and specialized equipment, it was indeed a challenging
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FIG.l. Simplified flow diagram. IAEA-SM-218/36 2 7 1 task for them to gear up their manpower and facilities to undertake the manu facture of sophisticated and complex nuclear components. In order to meet the stringent quality requirements of the nuclear specifications and codes it was soon realized by these industries that the existing quality control practices, which were confined to control of hardware, were not adequate and a more comprehensive outlook with the accent on control over “activities affecting quality” was needed. This led to the following improvements in the organizational set-up of some big industrial plants.
• A separate shop was set up for segregating nuclear work. Closer control was maintained over shop floor activities and cleanliness standards. The shop was manned by a group of well-qualified and experienced engineers who could devote their full time to the manufacture of nuclear components. Only highly skilled and experienced technicians and welders were employed on the nuclear jobs. In this shop, development of special welding procedures, jigs, tools and fixture was also taken up. • A separate Quality Assurance Department having independent authority to enforce quality control measures in the shops was established. This group was responsible for the preparation of quality assurance manuals, manufacturing procedures, non-destructive examination and testing procedures, documentation and qualification of welding and non-destructive-examination personnel.
M any problems were faced in achieving the required quality standards for the nuclear com ponents and some aspects of these problems, with special reference to Indian conditions, are discussed in this paper.
2. DESCRIPTION OF REACTOR SYSTEM
Although the first nuclear power station to be built in India is of boiling water enriched uranium type, India has adopted the natural uranium, heavy water moderated and cooled, horizontal pressure tube type of reactor (referred to as PH W R) for all its first-generation nuclear power stations. A schematic diagram of such a nuclear power plant is shown in Fig. 1. The reactor essentially consists of 306 zircaloy horizontal pressure tube assemblies housed in a stainless steel reactor vessel called the calandria and supported at the two ends in end shields. Just below the calandria and connected to it by an expansion joint is the dump tank. At the bottom of the calandria is a set of U-shaped dum p ports which provide a gas-liquid interface to support the heavy water moderator in the calandria by helium gas pressure and also facilitate fast dumping of moderator into the dump tank on reactor shutdown. The pressure tubes contain the natural uranium fuel bundles over which pressurized heavy water is pumped to remove the fission heat. The heavy water coolant is then passed through the tube side of inverted-U-type steam generators to produce 2 7 2 RAO saturated steam on the shell side, which is then led to the turbines to generate electricity. The reactor power is controlled by adjuster rods m oving vertically up and down in between the pressure tubes inside the calandria. On-power fuelling is accomplished by means of remotely controlled fuelling machines located at either end of the pressure tube.
The PH W R is characterized by the following important features:
(1) The reactor uses substantial quantities of heavy water, which is a very expensive fluid, as moderator and primary coolant. A ny appreciable leakage of this fluid m ay result not only in the shutdown of the plant but also release of hazardous radioactive material endangering public safety. Therefore the integrity and reliability of the pressure boundary components is m ost vital.
(2) The satisfactory operation of the reactor is dependent upon the precise alignment of main reactor components such as the calandria and end shields and reliable function of fuelling machines and numerous other com ponents in the pressure tubes under severe operating conditions.
(3) The PHW R unit is very compact, with the main reactor components installed close to one another. Once the reactor goes into operation it is difficult to carry out any inspection or repairs on these components due to their inaccessibility and the high radiation field.
3. QUALITY CONTROL PROBLEMS EXPERIENCED DURING MANUFACTURE OF REACTOR COMPONENTS
For the safe, reliable operation of a nuclear plant it is m ost essential to have adequate control of the quality of the components that go into it. The various activities affecting quality are design, procurement, manufacture, in spection, testing, shipment, storage, erection, commissioning, operation and maintenance. O f these, manufacturing m ay be considered to have the greatest effect on quality and therefore quality control at this stage assumes great significance. In view of the fact that m ost of the components are not easily accessible for inspection and repair once in service, any m anufacturing deficiencies will pose serious problem s for the safety of the plant. Problem areas associated with the manufacture of nuclear components are now discussed.
3.1. Raw material
Quality control at the raw material stage is very im portant not only in achieving the required quality of the component but also in helping avoid problems IAEA-SM-218/36 2 7 3 during fabrication. It is necessary to ensure at this stage, by appropriate testing, inspection and examination, that the material conforms to specification require ments as regards its chemistry, physical properties, heat treatment, metallurgical structure, surface treatment, freedom from flaws, etc. Proper documentation in the form of certified m ill test certificates of testing and inspection carried out on the material, with proper correlation, is to be maintained in order to ensure that the correct material is used on the job. It is im portant that testing and acceptance standards for raw materials required for critical com ponents are based on the end use rather than follow routine codes and standards which m ay not take into account all the service conditions. It is quite often necessary to conduct tests over and above those specified in the code to meet product requirements. The acceptance standards m ay also have to be more restrictive. This aspect is very im portant in non destructive examination and is dealt with separately. There are m any components, such as end fittings and sealing plugs, which contain seal faces form ing part of the primary pressure boundary. These seal faces are not only required to have close dimensional tolerances and a high degree of finish but should also be absolutely free from flaws such as cracks, inclusions and scratches. A very high degree of metallurgical cleanliness of the raw material is required. In order to achieve this, it is necessary to exercise stringent quality control right from the m elting practice to various stages of processing till the final product is obtained. At appropriate stages macro- and micro-etch tests have to be conducted to assess the quality of the material.
3.2. Fabrication
The main reactor components, such as calandria, end shields, the dump tank, and the shield tank, are massive and com plex structures, built to the stringent requirements of the nuclear code and product specifications. Their manufacture involves not only extensive fabrication and welding but also precision manufacturing to close dimensional tolerance and high finish. In view of their critical application, stringent quality control measures have to be exercised during all stages of manufacture. Since m any of these components are too big and heavy to be accom modated in standard machines, very careful pre-planning of the sequence of fabrication has to be done so that m achining can be carried out with the avail able facilities. In order to illustrate some of the problems mentioned above, we m ay consider fabrication of the calandria, which is one of the most complex and critical reactor components. A schematic diagram of the calandria, showing various sub-assemblies, critical dim ensions and alignment requirements, is given in Fig. 2. The major quality control problems encountered in the fabrication of calandria were: distortion control, precision boring and grooving operations. 4 7 2 RAO
FIG.2. Calandria (weight: 50 tonnes). Critical tolerances. all dimensions in inches, except where otherwise indicated IAEA-SM-218/36 2 7 5
3.2.1. Distortion control
The major portion of the calandria is made out of A ST M A240 Type 304L stainless steel. The austenitic stainless steels have a m uch higher coefficient of thermal expansion and lower thermal conductivity as compared to carbon steels. Therefore distortion during welding is also correspondingly higher. Distortion control to achieve the close tolerances required in fabrication has posed consider able problems. The problems encountered in the fabrication of the important sub-assemblies and the remedial measures taken to control distortion are highlighted.
3.2.1.1. Tube sheet and small shell sub-assembly
Since the tube sheet was not available as a single plate it had to be fabricated by butt welding of two plates. Some of the measures taken to exercise close control over the flatness consisted of presetting the plates and sequencing of welding. The welding of the tube sheet to the small shell also posed distortion problems and had to be controlled by adopting balanced welding.
3.2.1.2. Welding of stiffener rings on main shell
Welding of the thick box section stiffener ring on the main shell caused opening out of the shell ends during welding and this had to be corrected by heating and quenching.
3.2.1.3. Dum p port box
The dum p-port box is a critical sub-assembly requiring very careful distortion control to achieve close tolerances. The dump-port box consists of 3 box sections each containing 14 numbers of S-shaped bends and bulbs which are welded to side plates of the dump-port box by full penetration weld. Detail A of Fig. 2 shows the close tolerances required on the port gaps. Before starting fabrication of the actual dump-port box, prototypes using the same materials and methods of construction were prepared to provide an accurate estimate of weld shrinkages and to arrive at the proper welding sequence. Although valuable information was obtained from the mock-up test it was quite a difficult task to control distortion on the actual job, which was on a m uch bigger scale than the mock-up. The sequence of welding had to be carefully planned to control bowing of side plates to within the machining tolerance. Very careful control over the quality of welding also had to be exercised. The whole fabrication had to be done in a dust-free enclosure since the inside of the dump- port box is not easily accessible for cleaning at a later stage. 2 7 6 RAO
3.2.1.4. Transition piece
This piece provides transition from a rectangular opening to a circular opening. This was a very difficult piece to fabricate due to its continuously varying contour, which resulted in unequal deposition of weld metal. This gave rise to serious problems of distortion control and heavy fixtures had to be employed to overcome this problem.
3.2.1.5. Assem bly and welding of dump-port box and transition piece to main shell
This was a very difficult assembly operation in view of the close tolerances to be achieved for the whole assembly. Very skilful control had to be exercised to balance the welding and keep the distortion to a minimum. Precision optical survey instruments were used to level and align the sub-assemblies before starting welding and all stages of welding were closely monitored.
3.2.2. Precision boring and grooving o f tube sheets
After completion of welding, the boring and grooving of 306 lattice holes had to be carried out in each tube sheet. The accuracy, alignment and finish required for these lattice holes is given in Fig. 2. This operation had to be executed in a very careful manner so as to get “zero error”, since rectification of damaged holes was very difficult. So extensive work towards development of proper tooling and a mock-up had to be done before satisfactory results could be obtained. In order to achieve the locational tolerances, a special stainless steel jig, carefully aligned with respect to the vertical and horizontal central lines, was employed during boring operations.
3.2.2.1. Machining of adjuster nozzles
Six sets of top adjuster nozzles located in the top of the calandria shell and corresponding bottom nozzles housed in the dump-port box had to be machined to the alignment accuracies shown in Fig. 2. Ideally this operation should be carried out with a single setting for each top and bottom set of nozzles. A s this was beyond the capacity of standard machines, the boring operation was carried out on individual nozzles, with the required alignment being obtained by using precision optical instruments.
3.2.3. Cleanliness
A very high standard of cleanliness was maintained during all stages of fabrication and machining. In order to avoid carbon steel contamination, efforts IAEA-SM-218/36 2 7 7 were made to segregate stainless steel fabrication from carbon steel work. The free halogen content of all cleaning solutions, dye penetrants, paints, etc. was restricted to 25 ppm to avoid risk of stress corrosion cracking. Tests for chloride contamination were carried out on a sample basis on the finished calandria and areas showing contamination were thoroughly passivated.
3.3. Weld quality
In view of its versatility, safety, reliability and econom y, welding is the most widely used joining method in the fabrication of nuclear components and thus plays a crucial role in determining the safety and integrity of these components [ 1, 2 ]. In a nuclear reactor a weld joint is subjected to the m ost severe environmental conditions. In addition to high temperature, pressure, corrosion, erosion and cyclic stresses, it is subjected to severe effects of embrittlement due to radiation. A s stress concentrations arise at welded joints and are aggravated by the presence of defects, the risk of failure whether by creep, fatigue or brittle fracture is greatest in and around a welded joint. Therefore the properties of welded joints and the risk of defects present in them are critical factors affecting the safety of the component. It is therefore most important that in the construction of nuclear com ponents careful consideration is given to the selection of material, welding procedures, processes and consumables to ensure that sound weld joints are produced. For effective control of weld quality an understanding of the nature of welding problems is essential. Welding problems m ay be generally classified into the following types:
3.3.1. Technological welding problems
These are problems which arise due to faulty technique on the part of welders, lack of proper control over welding variables such as voltage, current, polarity, use of faulty equipment or consumables. Although these problems appear to be easy enough to correct by proper training and qualification of welders and adequate shop floor supervision, they are quite frequently encountered even in well-run shops and have been responsible for the occurrence of some m ost serious welding defects, which required extensive repairs, thereby causing major delay in the completion of the equipment. Therefore the need for adequate shop control and inspection to avoid such problems cannot be over-emphasized. In this connection a “daily welder’s test” which consists of a 180° bend test conducted on a test plate of same specification as the job, welded by each welder before commencement of work and using the production welding machines has been found to be a useful quality control measure. In the case of multi-pass welded joints it is useful to carry out an intermediate stage radiographic examina- 2 7 8 RAO tion so that any defects present in a partially completed weld, especially in the root area, can be detected and repaired before proceeding with further welding. This can save costly repairs after the completion of the weld.
3.3.1.1. Quality control of tube-to-tube sheet welding of heat exchangers
The tube-to-tube sheet joint in the case of heat exchangers used for nuclear service is a very critical joint and, in order to achieve the desired quality, requires stringent measures to control the welding variables as well as the performance of the welders. Generally, the semi-automatic tungsten-inert-gas process is employed and welding is accomplished either by fusing the projected end of the tube to the tube sheet to form a sm ooth weld bead or as a fillet weld by depositing external weld metal in 2 to 3 passes. The latter technique is more com m only employed. The quality requirements for the weld are that it should be not only sound (free from cracks, slag inclusions, porosity, weld craters) but must meet the m inim um leak path requirement. The m inim um leak path is defined as the minimum dimension of sound weld metal. The quality control measures adopted generally consist of a sampling procedure as outlined below. Each welder before commencing welding on the job does mock-up welding on a Quality Control Block (Q CB) which is prepared by the identical procedure and using the same spacings, hole size and materials as used in the heat exchanger. Each Q CB consists of rows of five tubes each, the number of rows being the same as the number of weld passes on the job. The first pass is welded on all the rows and the first row is sectioned and etched to check the soundness of the weld joint and to measure the m inim um leak path. The second and the subsequent rows, if any, are used for checking the second and further passes. The welder is allowed to proceed on the job only after checking the Q CB and ensuring that the m inimum leak paths are within the control lim its which are based on statistical quality control principles. Each Q C B is taken as a sample to represent a certain number of tubes in the heat e x c h a n g e r.
3.3.2. Metallurgical welding problems
These problems arise from factors such as impurities, inclusions and alloy content of materials, undue strain on the weld, inadequacy of welding procedures or wrong choice of welding processes. These problems m ay lead to cracking in the weld metal or heat-affected zones. Some of the com m on types of cracking phenomena and remedial measures are discussed below. IAEA-SM-218/36 2 7 9
3.3.2.1. Solidification cracking
This is a cracking phenomenon which develops in weld metal during the final stages of solidification and is due firstly to segregation of impurities such as sulphur, phosphorus to the grain boundaries in the form of liquid films and secondly, to the presence of shrinkage stresses. This is a com m on problem faced in welded fabrication. It can be avoided by:
(a) Controlling the sulphur and phosphorus levels in the material and keeping the weld joints clean. (b) Using welding consumables and procedures which can accomodate im purity levels in the plate. (c) Im proving fit-up to get proper weld bead shape. (d) Im proving assembly procedures to avoid undue restraint. (e) Proper sequen cing of welds.
This type of cracking is known to occur in fully austenitic stainless steel welds and can be prevented by providing a small percentage (3 to 5%) of delta ferrite, which dissolves m uch of the impurities. The delta ferrite is generally provided through welding consumables.
3.3.2.2. Underbead cracking
This form of cracking is quite a serious problem and occurs m ainly in hardenable steels due to hydrogen embrittlement in the heat-affected zone and under the influence of welding stresses. This manifests itself quite often as toe crack in a fillet weld. But it is also observed in butt welds where the restraint is high. Thicker materials are more prone to these cracks than thinner ones. This type of cracking can be avoided by preventing the hardening of the heat-affected zone and weld by the proper choice of material and welding processes, by using correct preheat and by use of properly dried basic coated low-hydrogen electrodes.
3.3.2.3. Lamellar tearing
This is essentially a cracking phenomenon found to occur in rolled plates having low through-thickness ductility and subjected to welding or design stresses transmitted through the thickness. This stress situation is found to occur in certain applications such as nozzles and stiffeners involving full-penetration T-joints. These cracks have a characteristic stepped appearance and are oriented parallel to the plate surface. They originate in planar inclusions such as laminations in the plate. Susceptibility to lamellar tearing is difficult to predict either by destructive or non-destructive test. A through-thickness tensile test is considered useful, although the m inim um acceptable reduction of area for different steels 2 8 0 RAO has yet to be determined. Remedial measures consist of improving joint design, the use of forged parts in place of rolled plates and in some cases buttering the susceptible plate area with softer weld metal before welding with the other member.
3.3.2.4. Heat treatment cracking
This form of cracking is found to occur in certain weld joints during thermal stress-relieving operation or during service and is due to low creep ductility of the weld metal or heat-affected zone. It is generally found to occur in alloy steels containing strong carbide-forming elements such as chromium, molybdenum and vanadium and it is more likely in thicker materials than in thin. The remedial measures consist of reducing strain concentration in the weld by proper weld profiling, using weld metal of improved creep ductility.
3.4. Non-destructive examination
Non-destructive examination (N D E) is an essential constituent of the quality control programme in checking the soundness of the components and providing adequate confidence that the components will perform satisfactorily in service [3]. Reliability of N D E for detection of flaws is of vital significance in the case of nuclear components, due to the fact that any failure of pressure-retaining com ponents m ay lead to the release of dangerous radioactive material. The reliable detection and assessment of flaws is dependent upon the correct selection of the N D E technique employed. The selection of N D E technique is in turn governed by factors such as accessibility, metallurgical structure, magnetic and electrical properties of the material, geometry and shape of the components, likely type and orientation of defects. It is most important that adequate pro vision is made at the design stage to ensure that there is sufficient accessibility and that the fabrication sequence, weld joint design and finish are suitable for carrying out N D E satisfactorily. It is well known that no single N D E technique is sensitive enough to detect all types of flaws. It is therefore im portant to know the advantages and limitations of each technique and the conditions under which faulty readings and failures may occur. As an example, in ultrasonic testing, defects of rounded shape which have less reflecting surface are hard to find. Even if they are detected the size is likely to be very much underestimated. In the case of radiographic examination the above type of flaws are ideal for detection and size is correctly estimated. In the light of the above facts it is therefore essential to complement one N D E technique with another to make a reliable assessment of location and size of flaws.
3.4.1. Acceptance standards for non-destructive examination
Acceptance standards for non-destructive examination of material and components m ust be specified on the basis of the final product rather than just IAEA-SM-218/36 2 8 1 follow the routine codes and standards. Although A SM E code Section III d iv is io n 1 is generally adopted for the construction of nuclear components, there are many components in PH W R which do not fit into the code. It is generally found necessary to specify more stringent requirements than called for in the code in order to meet the product requirements and assure its reliability in service. A s an example we m ay consider the case of the end fitting, which is a critical component in the primary pressure boundary. Each end fitting is machined out of a solid forging of 180 mm diameter and 2275 mm length which is required to be 1 0 0 % ultrasonically examined, using the straight beam technique. The acceptance standard given in the A S M E Code for Class 1 forgings is that a forging shall be unacceptable if straight beam examination results show one or more reflectors which produce indications accompanied by complete loss of back reflection not associated with or attributable to the geometric configuration. If a probe of 25 m m diameter is used then it will be seen that the m axim um defect which will be accepted as per the code is very big in relation to the size and product requirement of the end fitting.
4. CONCLUSION
Indian industries have demonstrated that the capability for manufacture of sophisticated nuclear components to meet the stringent quality requirements of nuclear specifications and codes can be achieved by a developing country with limited resources. Although foreign know-how has been beneficial, the Indian industries had to make their own efforts to develop technical expertise and make improvements to their existing facilities. These developmental efforts have been beneficial in im proving their expertise in conventional areas also. This has been an incentive to m any industries to take up more nuclear jobs and is an encouraging sign for the future of nuclear power in India. While the efforts made by Indian industries so far to meet the high quality standards are commendable, still greater efforts are required to strengthen their quality assurance/control departments in order to maintain effective quality c o n tr o l.
REFERENCES
[1] BAKER, R.G., The Welding of Pressure Vessel Steels, Climax Molybdenum Co. Ltd., London, UK. [2] The Procedure Handbook of Arc Welding, The Lincoln Electric Company, Cleveland, Ohio, USA. [3] NIXON, E.G., The NDT requirements for CANDU, Non-Destructive Testing, December 1974, IPC Science and Technology Press, UK. 2 8 2 RAO
DISCUSSION
S.M. BH U TTA: What have been the main problems with your fuelling machines and what solutions have you found from the quality control point o f v ie w ? A. K A K O D K A R : There have been no major problems associated with the manufacture of fuelling machines from the quality control point of view. The performance of Indian-built machines has been quite satisfactory. S.M. BH U TTA : What are the reasons for major outages of Rajasthan-1? W hat are the quality assurance solutions that have been offered in order to over come these problems? A. K A K O D K A R : The main reasons for the outage of the Rajasthan - Unit 1 Station have been grid fluctuations and turbine blade failure. W ith better stabili zation of the grid, failures have been minimized. The turbine blade failure problem has been overcome with certain design m odifications affecting the thickness and fit of the blades. G. Ö STBERG (Chairman): I would like to comment on an important aspect of both the principles and the application of quality assurance. It sometimes seems to be assumed that so-called “hum an errors” could be eliminated by the development of a “perfect” quality assurance system. It is also said that if we all work properly no quality assurance should be needed. I am not so sure. Nuclear power plants are systems of such com plexity that synergetic interactions of a “systems” nature will always occur to an extent which lim its the reliability to something less than perfect. Deviations from “norm al” behaviour will occur and will cause what statisti cians call “values exceeding the extremes” or “outliers”. This has in fact been observed in a study we have recently done on the distribution of the mechanical properties of structural steels in samples of up to 10 000 tests. This point should be remembered both by those concerned with the effectiveness of quality control and by those concerned with the assessment of the probabilistic aspects of safety. Session VIII
INSPECTION AND TESTING C h a ir m e n
R . S A G L I O F r a n c e
R. RODRIGUEZ SOLANO S p a in IAEA-SM-218/39
A TENTATIVE APPROACH TO A MORE RATIONAL PREPARATION OF
IN-SERVICE INSPECTION PROGRAMMES
C.B. BUCHALET, G. M ARTIN F r a m a t o m e , C o u r b e v o ie
M. VAUTERIN Tour EDF-GDF, Paris-la-Defense, F ra n c e
Abstract
A TENTATIVE APPROACH TO A MORE RATIONAL PREPARATION OF IN-SERVICE INSPECTION PROGRAMMES. A method is proposed by Framatome and Electricite de France (EDF) to: (a) establish a rational preparation of in-service inspection programmes of the primary coolant system (PCS) of a pressurized water reactor (PWR); (b) recommend design changes to components of the PCS to facilitate or improve current examinations and/or to decrease the vulnerability level of certain zones of the PCS in order to reduce the necessity for high efficiency in-service exami nations of these zones. The proposed method is based on the three following parameters: (1) the “vulnerability” of a given zone of the PCS, the level of which depends upon the potential damage to which this zone could be subjected during the plant’s life; (2) the “importance of the safety hazard” created by a crack or a rupture occurring in the preceding zone; (3) the “efficiency level” of current in-service inspection programmes for the zone considered. The application of the proposed method is illustrated with an example for which recommendations are made as a result of the global evaluation performed.
1. I N T R O D U C T I O N
In 1976, in order to optimize in-service examinations of the prim ary coolant system (PCS), Framatome and Electricite de France (ED F ) created a joint task g r o u p .
One of the goals aimed at by the task group was the improvement of current in-service inspection programmes and maintenance operations. In order to achieve this goal, the task group embarked on the following programme of analysis:
Evaluate the potential damage to which the m ost critical zones of this PCS could be subjected during plant operation.
2 8 5 2 8 6 B U C H A L E T et al.
Evaluate the consequences of a postulated rupture occurring in one of the preceding zones.
Evaluate the degree of efficiency of current in-service examinations of the zones considered.
Taking into account the preceding factors, make recommendations for improving the mechanical behaviour, the design, fabrication and in-service examinations of the m ost critical zones.
A method was proposed by the task group to permit, as much as possible, a qualitative approach to the above evaluations and a classification of the most critical zones of the PCS. This method, which is presented here, was applied to evaluate 300 zones of the PCS. A n example follows, illustrating the method.
2. PROPOSED METHOD
The proposed method is based upon the three following parameters:
The “vulnerability” of a given zone of the PCS, the level of which depends upon the potential damage to which this zone could be subjected during the plant’s life.
The “importance of the safety hazard” created by a rupture occurring in a given zone of the PCS.
The “efficiency level” of current in-service inspection programmes for a given zone of the PCS.
2.1. Vulnerability level
The vulnerability of a given zone of the PCS is determined by estimating the potential damage to which this zone could be subjected during the plant’s life. A series of factors that could result in potential damage is derived. These factors are those that could, in the short, medium or long term, cause local weakness points in the PCS. These factors are classified in the three follow ing categories.
(a) Category 1 — Factors related to mechanical and thermal loads
(al) Geometrical discontinuities
They are subjected to secondary and peak stresses whose importance depends upon the importance of the discontinuity, the variation in shape or thickness. IAEA-SM-218/39 2 8 7
They are very often located close to a weld, the presence of which is not taken into account here. Example: Nozzles.
(a2) External loads, reactions, vibrations
These loads lead to stresses which could be locally high. In this case, the usage factor as calculated in the stress reports of the com ponents allows evaluation of the resulting potential damage.
(a3) Metallurgical discontinuities
These are essentially weld discontinuities between two different materials. They are subjected to secondary thermal stresses due to the differences in thermal expansion coefficients between the two materials. Example: Welds between ferritic and austenitic steels.
(a4) Thermal gradients
They result from thermal transients and lead to cyclic secondary stresses, analysed in the stress reports of the components. These gradients can be fairly large in certain zones of the PCS and could result in relatively large usage factors.
(a5) Residual stresses
They result from fabrication processes and are difficult to estimate with accuracy. They are generally reduced to a low value by the stress relief heat treatments usually applied. In some cases when stress relief treatments are not applied, they can be as high as the yield stress of the material (e.g. rolling of steam generator tubes). These stresses are never cyclic.
(b) Category 2 — Factors related to fabrication and examinations
(bl) Volumetric defects
These are defects left during fabrication in accordance with the fabrication acceptance criteria and therefore estimated innocuous. They are always present, in various degrees, in cast parts and welds but the risk of potential damage due to these defects always remains low.
(b2) Planar defects
Sm all planar defects could be missed by the examinations during fabrication. This risk exists, to various degrees, in welds, especially socket welds and in parts of complex geometry. 2 8 8 B U C H A L E T et al.
EVALUATION OF POTENTIAL DAMAGE
N” ai 32 аз ЗА as bi b2 Ьз b4 b5 C1 C2 C3 c< cs C6 C7 C8 c? C-IO TOTAL 1 3 3 3 3 2 1 1 2 1 8
2 3 3 3 1 2 12 3 3 3 3 3 2 1 1 1 2 19
FIG.l. Nozzle region.
(b3) Metallurgical defects
These defects could result from a local m odification of the material metallurgical structure such as in bimetallic welds.
(b4) Difficult examinations during fabrication
Exam inations during fabrication are difficult in geometrically complex zones such as junctions, penetrations and nozzles. These examinations are essentially ultrasonic, non-destructive examinations. Another difficulty exists in ultrasonic examinations of stainless steel parts and bimetallic welds.
(c) Category 3 — Factors related to environment
(cl) Corrosion due to low flow rate
This type of corrosion is due to a concentration of corrosive elements in zones of low water-flow rate and in corners where water remains still. IAEA-SM-218/39 2 8 9
(c2) Corrosion due to liquid-vapour interface
This pertains essentially to the zone of the pressurizer corresponding to the level of liquid-vapour interface.
(c3) Corrosion due to local boiling
This pertains to the external surface of the steam generator tubes.
(c4) External corrosion of ferritic steel due to lack of paint
This type of damage is only associated with external surfaces of the ferritic steel welds which remain free of paint because of in-service ultrasonic examinations. Example: Spherical dome to flange weld.
(c5) Corrosion due to possible leaks
This type of corrosion could occur on the ferritic steel external surfaces of the PCS when they are located close to zones where leakage could occur. In this case, the corrosion is due to a local high concentration of boric acid.
( c 6 ) Erosion, wearing
Erosion could occur in zones of high water-flow rate (pum p casing, pipe elbows, nozzles, etc.). Mechanical wearing could be significant for surfaces in friction under high loads (threads of bolts or nuts subjected to frequent removal).
(c7) Mating
Mating could occur only on surfaces such as threads of bolts, bolt holes and sealing surfaces. M ating could initiate stress corrosion.
(c 8 ) T e a r in g
This type of damage only concerns the PCS internal surfaces’ stainless steel cladding. It could occur in highly stressed zones (e.g. partition plate of the steam generators) and/or in regions of high water-flow rates.
(c9) Impact, shock
This type of damage could result from possible circulation of loosened parts from various origins. 0 9 2 he number of . a s ven to ever e n o ry e v e o t n e iv g is r e b m u n a , e n o z h c a e r o F s. e n o z f se a o cre r in e b n a m u d n n a e s th s e d n n h a g u y r o t t e l m ia r o e e t g a m e th f o se a re c e d a in g in lt u s e r , t n e m d r a b m o b n t ti emper ur . re tu ra e p m te n io it s n a r t e th in rradi on io t ia d a r Ir ) O l c ( ri he potenti damage f or It i obvi hat he numbers r e b m u n e th t a th s u io v b o is t I . rs to c fa e g a m a d l ia t n e t o p e th o t d e t u ib tr t a s r e b m u n s rie a v r e b m u n is h T . e v o b a d e fin e d s rie o g te a c e re th e th o t g in g n lo e m b o s r r f o t c a f e th f o gh probabii . y ilit b a b o r p h ig h = ) 3 ( acteri ng t evel vul lty shoul dered a qualtatve i i d in e tiv a lit a u q as d e r e id s n o s. c te e a b stim ld e u o e h iv s t a y it t ilit n a b u a r q e n ln a u h t v f r e o th l a e r v le s n e io th t a g c in iz r e t c ra a h c ( ( IMPORTANCE OF SAFETY HAZARD-» 2 1 = l lty of ng thi es i he zone consi ed re e id s n o c e n o z e th in t n se re p r o t c a f e g a m a d is h t g in v a h f o y ilit b a b o r p w lo = ) = l um probabii y ilit b a b o r p m iu d e m o t w lo = ) i cal hels t ti r ect n o r t u e n o t d te c je b u s are e ltlin e b r o t c a e r e th f o lls e sh l a ic r d lin y c e h T el ed regi t t e i est i ch shows t e th s w o h s h ic h w d e h lis b ta s e is le b ta a , S C P e th f o n io g e r d te c le se h c a e r o F t evel vul lty i ned by addi l he th ll a g in d d a y b d e in a t b o n e h t is y ilit b a r e ln u v f o l e v le e th , e n o z h c a e r o F 1 3: o t I 2 Glbleauto tepi r coa ytm zones. system t coolan ary prim the f o evaluation lobal G .2. FIG T et E L A H al. C U B 2, ECO VSE NOZZLES VESSEL REACTOR - ,3 ,2 1 URN EAINTOS F MEDIUM OF NATIONS EXAMI CURRENT ■ ( IMPROVEMENT IS RECOMMENDED) RECOMMENDED) IS IMPROVEMENT ( y c n e i c i f f e
IAEA-SM-218/39 2 9 1
Let us consider for example the nozzle region of a reactor pressure vessel (Fig.l). This region is divided into three zones, the vulnerability levels of these zones being respectively 17,11 and 18. From the table, it appears that the zones constituted by the stainless steel buttering and the safe-end to nozzle weld are more vulnerable than the pipe to safe-end weld zone. Further, the table shows that the vulnerability of zones 1 and 3 is m ainly due to factors related to mechanical and thermal loads.
2.2. Importance of safety hazards
The importance of the safety hazards created by the appearance of a rupture in a zone of the PCS is estimated at three levels, low, medium and high. This estimation is based upon the consequences, in terms o f safety, of an assumed rupture of the zone considered. If we consider the previous example of the reactor vessel nozzle region, the im portance of the safety hazard resulting from an assumed rupture in this zone is estimated as being high.
2.3. Efficiency level of current in-service inspection programmes
The efficiency of current in-service inspection programmes for each zone selected is estimated using four levels:
not examined low efficiency medium efficiency high efficiency.
This efficiency level is estimated from the ability of the current inspection programme to detect or not the potential damage introduced in the vulnerability estimate of a given zone (see subsection 2.1 ). In our previous example, the efficiency of the current in-service examinations is estimated to be medium.
3. CONCLUSIONS AND RECOMMENDATIONS
The determination of the three parameters described above permits an evaluation for every zone analysed showing whether it is necessary or not to improve the design, manufacture and/or current inspection programmes. Figure 2 summarizes this global evaluation. 2 9 2 B U C H A L E T et al.
The figure is divided into five regions:
R e g io n I examination not required R e g io n I I examination of low efficiency required R e g io n I I I examination of medium efficiency required R e g io n I V examination of high efficiency required R e g i o n V design changes necessary to decrease vulnerability level.
Every region of the PCS analysed can be represented by a dot on the figure. If current examinations appear to be insufficient (closed dots), two types of actions are recommended: to change the design and/or to improve the examinations. In the example considered before, the three zones are represented by closed dots in regions IV and V on the figure. From this evaluation, it is recommended that both design changes and improvements of in-service examinations be implemented. The evaluation described in this paper has been performed for 300 zones of the PCS.
DISCUSSION
R.I. H O D G E: I am impressed by your logical appraisal of the requirements for in-service inspection. However, there is no discussion in your paper of data handling, that is, recording, assessment, storage and retrieval for comparison with later inspections. This is a difficult problem and its solution essential to successful inspection. Could you please indicate what steps are being taken for this aspect of in-service inspection? C.B. B U C H A L E T : The purpose of the paper was merely to describe a method of preparing in-service inspection programmes. Obviously data from actual inspections should be used to improve the preparation of such programmes but this aspect was not covered in the present analysis. M. KITTNER: What methods of inspection should be applied in regions I—IV shown in Fig.2? C.B. BU C H A LET : Methods will be selected to correct defects associated with the potential damage factors pertinent to each zone. For instance, if corrosion from leaks is recognized to be a significant factor for a given zone, a leak detection system will be recommended. IAEA-SM-218/37
CONTROLES NON DESTRUCTIFS ET METHODE D’EVALUATION DES DEFAUTS EN TANT QUE MOYEN POUR AMELIORER LA FIABILITE DES COMPOSANTS DE REACTEURS
A.C. PROT*, R. SAGLIO **, M. ASTY**, M. PIGEON** CEA , Centre d’etudes nucleates de Saclay, Gif-sur-Yvette, F r a n c e
Abstract-Rdsumd
NON-DESTRUCTIVE TESTING AND FLAW EVALUATION AS A MEANS OF IMPROVING THE RELIABILITY OF REACTOR COMPONENTS. The paper reports on developments in those non-destructive testing techniques which can contribute to the reliability of nuclear components, especially in techniques to be used for periodic inspection. A description is given of the most recent improvements in ultrasonic testing, eddy currents and acoustic emission, emphasis being placed on what they can contribute to reliability. For example, it is shown how the systematic use of a flaw sizing technique can improve techniques for analysing the harmful effects of flaws (e.g. fracture mechanics analysis). There is also a description of the problem of the large differences between tests carried out in fabrication and the more sophisticated tests used in periodic inspection and it is shown what effect this can have on the reliability of facilities. The importance of problems associated with the testing of austenitic steels and welds of dissimilar metals is stressed, as is the need for finding solutions to these problems quickly.
CONTROLES NON DESTRUCTIFS ET METHODE D’EVALUATION DES DEFAUTS EN TANT QUE MOYEN POUR AMELIORER LA FIABILITE DES COMPOSANTS DE REACTEURS. On presente les developpements des methodes d’examen non destructif qui sont susceptibles de contribuer ä la fiabilite des composants nucleaires, particulierement dans l’optique de leur utilisation ä l’inspection periodique. Les plus recentes ameliorations dans le domaine des ultrasons, des courants de Foucault et de l’emission acoustique sont exposees, l’accent etant mis sur ce qu’elles apportent au niveau de la fiabilite. On montre par exemple comment l’utihsation systematique d’une methode de dimensionnement des defauts peut ameliorer l’analyse de leur nocivite par des methodes telles que la mecanique de la rupture. On aborde egalement le Probleme des differences profondes qui existent entre les contröles pratiques en fabrication et ceux plus sophistiques utilises en inspection periodique, et l’on montre quelle peut etre l’incidence sur la fiabilite des installations. L’importance des problemes lies au contröle des aciers austenitiques et des soudures de metaux dissemblables est egalement soulignee, comme la necessite de trouver rapidement des solutions ä ces problemes.
* Institut de protection et de sürete nucleaire. ** Departement de technologie, Section des techniques avancees.
2 9 3 2 9 4 P ROT et al.
INTRODUCTION
La fiabilite peut etre definie comme la caracteristique d’un materiel, exprimee en terme de probabilite, ä accom plir la fonction pour laquelle il a ete conqu, dans des conditions determinees et pendant une duree determinee. Une telle definition rend compte immediatement des rapports etroits qui existent entre fiabilite et qualite et done entre fiabilite et contröle. Ceci est particulierement vrai pour les com posants de centrales nucleaires, soum is ä un environnement severe (temperature, pression, irradiation), et ce pendant plusieurs d iz a in e s d ’an n ee s. II est done particulierement important de pouvoir disposer de methodes de contröles non destructifs dont la sensibilite, la fidelite soient adaptees ä l’objectif recherche, et dont les resultats soient denues de toute ambigui'te. L ’expose qui va suivre se propose de montrer les resultats d’ores et dejä acquis pour differentes methodes de contröle en precisant comment les ameliorations apportees contribuent ä la fiabilite des composants soumis ä ces e x a m e n s. Une precedente com m unication [ 1 ] a dejä fait etat des etudes menees dans ce sens par le Com m issariat ä l’energie atomique. Le present memoire s’attachera ä montrer que les buts recherches sont soit atteints, soit en bonne voie d’etre a tte in ts.
ULTRASONS
Rappel des developpements
L ’inspection periodique par l’interieur des cuves de reacteurs ä eau sous pression (PW R) a necessite la recherche d’une methode apte ä resoudre le Probleme de la traversee du revetement en acier inoxydable. On sait que la solution consiste ä travailler ä basse frequence (1 M H z) avec des traducteurs de grand diametre (jusqu’ä 150 m m ) et fortement amortis; la resolution laterale est obtenue en focalisant le faisceau. Ces traducteurs sont prevus pour le contröle en immersion. En corollaire de la solution au Probleme du revetement, ces traducteurs a ssu re n t: — une reproductibilite ä long terme due ä la possibilite de former le faisceau utile dans la zone focale, par usinage correct des lentilles; — un rapport signal/bruit tres ameliore par rapport aux traducteurs conventionnels (l’energie emise est fonction de la surface emettrice); — une sensibilite constante pour tout le faisceau utile, quelle que soit la profondeur traversee; — un pouvoir de resolution constant, independant egalement de la profondeur traversee. 295 j/ 2 9 6 P R O T et al.
FIG.2. Traducteurs ultrasonores - Equipement type.
Ces deux derniers avantages sont acquis au prix de l’utilisation de plusieurs traducteurs pour les epaisseurs importantes (par exemple, trois sont necessaires pour le contröle des parties courantes de la cuve d’un reacteur PW R). Ces traducteurs sont actuellement utilises sur toutes les machines d’inspection en service developpees en France (fig. 1 et 2), tant pour les preinspections que pour les inspections de reacteurs chauds.
METHODE DE D1MENSIONNEMENT
L ’analyse des parametres enregistres par ces machines d’inspection est particulierement aisee puisque celles-ci permettent l’enregistrement de carto graphies type В ou C et la comparaison immediate d’une inspection ä l’autre par simple superposition. II est done facile de noter les variations eventuelles. Toutefois la presence d’un defaut ou de l’accroissement d’un defaut preexistant n’est en soi pas significative. Elle ne prend tout son sens que si Гоп peut quantifier ce qui a ete enregistre. C ’est effectivement ce que permet l’utilisation des traducteurs focalises. On a pu montrer experimentalement et verifier, tant sur defauts artificiels que sur defauts reels, que dans le cas general la dimension d’un defaut est donnee, avec une precision egale ä un diametre d du faisceau utile au foyer, par celle de la cartographie type C effectuee ä un gain superieur de 6 db ä celui pour lequel le defaut est juste visible; ceci est verifie de la maniere suivante [2]: — si Гоп augmente encore le gain de 6 db, l’image ne s’accroit que d’une valeur egale ä d/2; IAEA-SM-218/37 2 9 7
Image au niveau N1
DEFAUT SUPERIEUR i d
Image au niveau N1
APPARITION DE NOUVELLE ZONE REFLECHISSANTE
FIG.3. Dimensionnement des defauts -D efaut superieur au faisceau utile - Ondes L.
— si l’accroissement de l’image est superieur, cela signifie que de nouvelles zones m oins reflechissantes du defaut sont mises en evidence ä ce nouveau gain; on repete alors le processus d’augmentation du gain de 6 db (fig. 3 et 4).
Im portance pour la fiabilite
Le Code A SM E, Section XI, propose, pour l’analyse des resultats de l’inspection en service par ultrasons, l’utilisation de la mecanique lineaire de la rupture. Bien evidemment, ceci ne suppose pas seulement une connaissance de l’existence des defauts, ni meme de leurs dimensions relatives ou comparees ä un etalon arbitraire. L ’application des formules de la mecanique de la rupture
(suite du texte p. 303) 8 9 2 PROT et al. et PROT
FIG.4. Dimensionnement des defauts — Defaut superieur au faisceau utile - O ndes T. IAEA-SM-218/37 2 9 9
CARTOGRAPHIES ULTRASONORES
TRADUCTEURS FOCALISSS : 4MHZ
0 fo y e r : 2 , 7 m
COUPES MACROGRAPH IQUES
FIG. 5. Correlation cartographies-macrographies - Defaut A. 3 0 0 P R O T et al.
CARTOGRAPHIES ULTRASONORES
TRADUCTEURS FO CALIStS : 4MHz
0 FOYER : 2 , 7 MM
COUPES MACROGRAPHIQUES
Tl
+ 1 2 d b
^lbi s
FIG.6. Correlation cartographies-macrographies - Defaut B. IAEA-SM-218/37 3 0 1
CARTOGRAPHIES ULTRASONORES
TRADUCTEURS FO CALIStS : 4MHz
0 FOYER : 2 , 7 MM
COUPES MACROGRAPH IQUES
FIG. 7. Correlation cartographies-macrographies - Defaut C. 3 0 2 P ROT et al.
FIG.8. Appareil d courants de Foucault ä trois frequences. IAEA-SM-218/37 303
(ou de toute autre methode equivalente) necessite la connaissance des dimensions reelles des defauts et notamment de la dimension mesuree suivant l’epaisseur de la piece. Ainsi que le montrent les figures 5 a 7, relatives ä trois defauts de soudure (inclusions de laitier), seuls les enregistrements type C obtenus avec les traducteurs focalises ont permis de determiner les trois dimensions avec une precision chiffree ä l’avance: le diametre du faisceau utile au foyer. Dans ce cas particulier cette precision atteignait 3 mm. Afin de verifier sur un grand nombre de defauts cette importante propriete, plusieurs contrats sont actuellement lances en France avec la Societe Creusot- Loire. Ainsi une eprouvette de 800 mm X 1000 mm, d’epaisseur 250 mm, comporte un joint soude dans sa longueur; dix defauts ont ete volontairement introduits et seront detectes et dimensionnes selon toutes les methodes a disposition. Puis la piece sera sectionnee pour faire des correlations avec les defauts reels. Des examens du т ё т е type seront egalement pratiques sur des eprouvettes de fatigue de grande dimension. Une telle methode de dimensionnement valorise les contröles par ultrasons en les rendant quantitätifs. La connaissance des dimensions reelles des defauts permet alors de juger de leur innocuite ou de leur danger et, par consequent, compte tenu par ailleurs des sollicitations prevues de la structure, permet de determiner soit la capacite du composant ä continuer sa fonction soit la necessite de le reparer, ou le cas echeant de le changer. Elle concourt de ce fait ä accrofre la fiabilite du composant, done de l’installation. La methode proposee a ete developpee pour le contröle des cuves et composants en acier ferritique. Elle a ete egalement utilisee pour le contröle des soudures d’acier inoxydable austenitique et plus particulierement des soudures de metaux dissemblables du type embouts de sdcurite. Les resultats obtenus constituent un progres tres important: la sensibilite est tres supdrieure ä celle des methodes utilisant des traducteurs traditionnels. On a emis l’hypothese que la tres forte absorption rencontree dans ce type de materiau etait due principalement au phenomene d’interferences creees par les dephasages introduits au niveau des joints de grains par les variations importantes d’impedance acoustique. La focalisation qui est une operation de remise en phase aurait alors un effet favorable. II est souhaitable que ces resultats puissent etre compares objectivement avec ceux obtenus dans d’autres pays (pour tenir compte par exemple des conditions de soudage); e’est pourquoi un groupe de travail a ete сгёё avec la Republique federate d’Allemagne: ainsi en echangeant etalons et experiences sera-t-il possible d’ameliorer encore le contröle de ces aciers austenitiques. II s’agit d’un probteme important, dans la mesure oü de nombreux constituants (tuyauteries, pompes notamment) sont realises dans ce type d’acier et comportent un grand nombre de soudures. 3 0 4 PROT et al. PROTet 4 0 3
FIGS. Machines de contröle de goujons et ecrous. IAEA-SM-218/37 305
CONTROLES DE FABRICATION ET INSPECTIONS PERIODIQUES
II est interessant de noter que la fiabilite d’un composant depend de l’aptitude de la structure ä s’accommoder des sollicitations auxquelles eile est soumise et, par consequent, de l’aptitude ä evoluer des defauts existant dans cette structure. C’est done non seulement la connaissance de la presence des defauts deceles qui est utile pour Evaluation de la fiabilite, mais egalement la connaissance de to u s les defauts et de lew s caracteristiques. Dans cette optique les contröles de fabrication et les inspections periodiques jouent des roles tres differents. — les contröles de fabrication assurent un niveau de qualite defini, meme si ce niveau est arbitraire, comme c’est le cas generalement avec les contröles ultrasonores; — les inspections periodiques doivent fournir des donnees quantitatives sur les defauts deceles et ne sauraient, par consequent, se contenter de definitions arbitraires. II semble cependant que la definition arbitraire (quoique sanctionnee par une experience non negligeable) d’un niveau de qualite en fabrication ne soit pas entierement satisfaisante pour l’esprit. Les soucis de fiabilite ameneront sans nul doute ä plus ou moins long terme ä harmoniser les contröles en fabrication et en inspection periodique, dans le sens d’une connaissance plus approfondie des caracteristiques des defauts. En outre, ainsi que Га montre clairement Jordan [3], l’effet des inspections periodiques sur la probability de rupture d’une cuve est particulierement sensible, dans la mesure oil ces inspections sont suivies des dispositions que leurs resultats suggerent. Or ces dispositions, pour etre efficaces, ne peuvent etre prises qu’ä partir de donnees objectives qui ne peuvent etre obtenues que par des methodes plus sophistiquees que celles actuellement utilisees en fabrication.
COURANTS DE FOUCAULT
Dans une centrale ä eau sous pression du type PWR, le generateur de vapeur constitue pour la sürete et pour la disponibilite de Finstallation un facteur important. La fiabilite de ce composant est done essentielle: les nombreux incidents qui ont marque et marquent encore le fonctionnement des centrales de ce type en est une preuve. Les divers incidents qui peuvent affecter ces generateurs de vapeur, en ce qui concerne plus particulierement les faisceaux tubulaires, sont les suivants: - corrosion interne ou externe, - depot de boues eventuellement magnetiques, - chocs repetes dus aux vibrations au niveau des plaques supports, 306 P R O T et al.
— fissuration par fatigue, — d e n t in g , phenomene relativement complexe associe ä la chimie de l’eau. Le precede normalement utilise lors des inspections periodiques est celui des courants de Foucault avec bobine interne. Jusqu’ä maintenant ces contröles presentaient un certain nombre d’inconvenients: — difficult^ de localiser les defauts dans l’epaisseur (defaut de surface interne, de surface externe, ou dans l’epaisseur); — tres grande difficult^ ä detecter les ddfauts au droit de la plaque tubulaire et des plaques supports intermediaries. Le bruit de fond, dü ä la geometrie du tube (fabrication par le procede du pas de pelerin) ou d’origine magnetique (composition du tube ou presence de boues magnetiques), posait egalement des problemes. Sans entrer dans les details, on peut dire que l’appareillage mis au point par le Commissariat ä l’energie atomique a permis de resoudre la plupart de ces problemes. II utilise pour ce faire les dispositions suivantes: — deux ou trois frequences differentes simultanees; — une sonde de saturation magnetique; — l’analyse en phase et amplitude qui permet notamment: l’enregistrement des defauts internes et externes sur deux voies separees, avec la meme sensibilite; par combinaison de deux frequences, l’elimination d’un parametre parasite (trois frequences permettent done d’en eliminer trois); l’elimination des parametres magnetiques parasites; la mesure des epaisseurs par utilisation d’une voie absolue. Ces diverses possibilites sont acquises sans sacrifier la simplicite d’utilisation et de rdglage. C’est ainsi que 1’echange d’une bobine usee ne demande que quelques minutes. Des bobines ont ete developpees qui permettent le contröle de toute la longueur des epingles en U, meme celles ä tres faible rayon de courbure. Ce materiel a dejä la sanction de nombreuses utilisations sur le site, tant pour les preinspections que pour les inspections (fig. 8). II est ä noter que les signaux bruts sont enregistres sur bande magnetique, ce qui permet ä la restitution toutes les analyses dont il a ete question plus haut. Un tel dispositif contribue tres largement ä la fiabilite des generateurs de vapeur. II permet en effet de deceler ä temps — meme dans des endroits reputes jusqu’alors difficiles, sinon impossibles - tous les types de defauts actuellement repertories dans les centrales en fonctionnement, de les localiser en epaisseur et, dans une certaine measure, de les qualifier. II permet egalement de quantifier, avec une precision süffisante pour repondre aux exigences des codes, certains types de defauts: corrosion, fissures localisees, etc. L’utilisation des courants de Foucault n’est pas restreinte aux tubes de generateurs de vapeur. Divers equipements ont ete developpes utilisant les memes principes, notamment pour les contröles de goujons et d’ecrous, tant pour la cuve que pour des composants de plus petite dimension (fig. 9). IAEA-SM-218/37 307
Les progres realises grace ä l’utilisation des courants de Foucault multi- frequences constituent done une contribution tres notable ä l’amelioration de la fiabilite de composants tres importants pour la sürete et la disponibilite des centrales nucleaires. On peut souhaiter que, dans ce domaine, les techniques de traitement de signal judicieusement appliquees permettent de soulager quelque peu les equipes chargees du depouillement et de l’interpretation. C’est dans ce sens que les efforts sont actuellement poursuivis.
EMISSION ACOUSTIQUE
Le Commissariat ä l’energie atomique etudie egalement depuis plusieurs annees les diverses possibilites de remission acoustique. L’idee essentielle est la suivante: dans toutes les methodes de controle non destructif par balayage (ultrasons, courants de Foucault, . . .), il est necessaire d’amener le traducteur au voisinage tres immediat du defaut ä deceler de telle sorte que l’energie mise en jeu dans le procede atteigne le defaut et subisse une alteration revelatrice de sa presence. L’emission acoustique est, par contraste, une methode globale en ce que le defaut signale lui-meme sa presence et en ce qu’un reseau plus ou moins läche de detectcurs permet de le signaler. Dans l’optique de la fiabilite des centrales nucleaires, cette difference est importante: eile signifie en effet que, dans les methodes classiques, la probability de d6celer un defaut sur lequel le detecteur n’a pas passe est nulle, alors qu’elle ne Test jamais dans le cas de remission acoustique. Cet aspect justifie ä lui seul les etudes dans ce domaine, quels que soient les succes ou les echecs que l’on rencontre actuellement. Deux applications distinctes quant au type de materiel mis en oeuvre sont possibles: — surveillance par un capteur d’une zone particuliere, par exemple ä proximite d’un defaut connu dont on veut suivre Г eventuelle evolution; — surveillance avec localisation des sources grace ä un appareillage multicapteur. L’application ä un circuit de centrale nucleaire necessite la resolution prealable de la tenue des capteurs ä l’environnement: temperature, irradiation. Deux solutions sont а Г etude: — les capteurs chauds, susceptibles de supporter l’environnement severe, — les capteurs froids, montes ä l’extremite de guides d’ondes. Au vu des resultats acquis tant sur le reacteur prototype CAP que sur le pressuriseur de Fessenheim I, il semble que la seconde solution soit preferable, en particulier lorsqu’il est possible de souder directement les guides d’ondes ä la surface du composant: on elimine ainsi, et le Probleme de perte de sensibilite du 308 P R O T et al. capteur avec le temps, et le Probleme de tenue mecanique du couplant eventuel, surtout en presence de cyclages thermiques. En ce qui concerne la surveillance par un capteur d’une zone particuliere, les essais dejä realises ä Fessenheim montrent que l’interpretation des resultats (en taux de comptage ou en comptage cumule) est impossible manuellement par suite de la quantite prohibitive d’informations en sortie: l’utilisation de moyens avances en informatique (microprocesseurs en particulier) permet de diminuer considerablement le volume d’informations en eliminant les bruits propres d’origine hydraulique (ecoulement de fluides) ou mecanique (machines en rotation). Un deuxieme interet de l’utilisation de l’informatique reside dans la possibility de mise en memoire des bruits de fonctionnement (par exemple: variation de debit dans une canalisation ou de vitesse de rotation d’une machine tournante). Dans le domaine de la localisation, un appareil ä 32 voies a ete conqu et realise pour repondre aux problemes specifiques de la surveillance en continu, en particulier grace ä un dispositif ä seuil automatique asservi au niveau de bruit de fond — la detection des evenements se fait alors dans les meilleures conditions quelle que soit la valeur du bruit de fond — et ä un programme de gestion et de depouillement des informations en temps reel. Accessoirement, le materiel developpe peut etre utilise ä la detection des objets migrants ou ä la detection de fuites.
CONCLUSIONS
L’ensemble des developpements qui ont ete presentes ci-dessus dans divers domaines des contröles non destructifs constitue un appoint interessant dans la recherche d’une fiabilite toujours plus grande des composants de reacteurs. Certaines methodes ont acquis une maturite qui leur permet d’etre utilisables immediatement. D’autres, telle remission acoustique, necessitent que des efforts importants soient entrepris avant que les espoirs qu’elles ont fait naftre se concretisent. Nul doute qu’une utilisation rationnelle de l’informatique avancee ne constitue ä l’avenir une aide precieuse et indispensable pour que l’ensemble des methodes etudiees fournissent le plein des informations qu’elles sont susceptibles de delivrer.
REFERENCES
[1] PROT, A.C., SAGLIO, R., «Contröles non destructifs et examens periodiques des circuits primaires de reacteurs», Reliability of Nuclear Power Plants (C.R. Coli. Innsbruck, 14-18 avril 1975), AIEA, Vienne (1975) 533-50. IAEA-SM-218/37 309
SAGLIO, R., TOUFFAIT, A.M., PROT, A.C., «Determination des caracteristiques des defauts de soudure ä l’aide de traducteurs focalises», XXX th Annual Assembly, International Institute of Welding, Copenhague, 4 -9 juillet 1977. JORDAN, G.M., «The influence of frequency and reliability of in-service inspection on reactor pressure vessel disruptive failure probability», 3rd Conf. on Periodic Inspection of Pressurized Components, London, 1976.
IAEA-SM-218/38
INSPECTION EN SERVICE DU CIRCUIT PRIMAIRE DES REACTEURS A EAU SOUS PRESSION
M.T. DESTRIBATS, A.M. TOUFFAIT, M. ROULE, M. PIGEON, M. ASTY, A. SAMOEL, R. SAGLIO Departement de technologie, Section des techniques avancees, CEA, Centre d’etudes nucleaires de Saclay, Gif-sur-Yvette, France
Abstract-R^sumd
IN-SERVICE INSPECTION OF THE PRIMARY CIRCUIT IN PWRs. The Commissariat ä l’energie atomique, at the request of Electricite de France and Framatome, has developed original techniques for testing critical components of the primary circuit in PWRs. These techniques involve a very wide range of non-destructive testing methods, namely: eddy currents, in particular, with a multiple frequency device for testing steam generator pipes, studs and nuts; focused ultrasonic beams for testing all the welds of the reactor vessel and of its lid, the mixed welds of vessels and of steam generators, pressurizer lower head welds and studs from the inside; gamma radiography of the mixed welds of the vessel; televisual examination and examination by sweating of the rust-proof lining of the vessel and of its lid together. To these tests should be added acoustic emission studies which have been carried out during the hydraulic trial of the primary circuit and which are continuing at present during operation. These techniques are already in use both on reactors which have been in operation and on those which have not yet gone critical. So far, the two 900-MW Fessenheim reactor units and the Chooz reactor have been inspected successfully.
INSPECTION EN SERVICE DU CIRCUIT PRIMAIRE DES REACTEURS A EAU SOUS PRESSION. Le Commissariat ä Tenergie atomique, ä la demande d’Electricite de France et de Framatome, a developpe des methodes et des techniques originales de contröle des composants delicats du circuit primaire des reacteurs ä eau sous pression. Ces techniques mettent en oeuvre un eventail tres ouvert de methodes de contröles non destructifs: les courants de Foucault, notamment multifrequences, pour le contröle des tubes des generateurs de vapeur, des goujons et des ecrous; les ultrasons focalises pour le contröle de l’ensemble des soudures de la cuve du reacteur et de son couvercle, des soudures mixtes des cuves et des generateurs de vapeur, des soudures basses du pressuriseur et des goujons par l’interieur; la gammagraphie des soudures mixtes de la cuve; l’examen televisuel et par ressuage du revete- ment inoxydable de l’ensemble de la cuve et de son couvercle. A ces examens, il faut ajouter
311 312 DESTRIBATS et al. les etudes d’ecoute acoustique qui se sont deroulees pendant l’epreuve hydraulique du circuit primaire et qui se poursuivent actuellement pendant le fonctionnement. Ces methodes sont d’ores et dejä operationnelles, aussi bien sur les r6acteurs ayant fonctionne que sur ceux qui n ’ont pas encore diverge. A ce jour, les deux tranches de 900 MW de Fessenheim et le reacteur de Chooz ont ete inspectes avec succes.
1. INTRODUCTION
La sürete d’une installation nucleaire, et particulierement d’un reacteur electrogene de puissance, est Нее ä de nombreux facteurs dont Tun des plus importants est la fiabilite que Гоп peut accorder aux divers composants. Le Commissariat ä l’energie atomique, ä la demande d’Electricite de France et de Framatome, a developpe des methodes et des techniques originales de contröle des zones delicates du circuit primaire des reacteurs ä eau sous pression. Un large programme de recherche a permis le developpement de dispositifs dont on donne ici la description, ainsi que les resultats qu’ils permettent d’obtenir.
2. DEVELOPPEMENT DES METHODES
La figure 1 presente toutes les zones du circuit primaire oü des developpe- ments de methodes specifiques ont ete necessaires pour satisfaire aux exigences des reglementations et accroitre la fiabilite. Les avantages des differentes methodes utilisees sont les suivantes:
2.1. Ultrasons
Les etudes dejä presentees ailleurs [ 1—3 ] ont conduit le Commissariat ä l’energie atomique ä etudier et developper les traducteurs ä ultrasons focalises. Les avantages resultant de l’utilisation de ces traducteurs peuvent etre resumes comme suit.
Existence d ’une zone focale ayant les proprietes suivantes: — diametre constant pour une chute a 6 dB ä l’interieur de cette zone, — sensibilite constante (ä 6 dB pres) tout au long de cette zone.
Aucune perte de sensibilite n’est observee ä la traversee du revetement d’acier inoxydable.
Reproductibilite ä long terme aisement obtenue
Ceci est considere en France comme un point tres important; en effet, l’experience a montre qu’avec les traducteurs classiques il etait strictement difficile IAEA-SM-218/38 313
d’assurer cette reproductibilite, alors qu’avec les traducteurs focalises celle-ci etait aisee ä obtenir.
Dimensionnement possible des indications avec une precision connue
La precision du dimensionnement est directement liee ä la dimension de la tache focale. La procedure ä utiliser est explicitee dans [4].
Prise en com pte des form es geometriques
Un traducteur focalise comporte toujours une lentille; de ce fait il est toujours possible de realiser cette lentille de telle maniere qu’elle corrige les aberrations dues ä la courbure ou ä l’inclinaison de l’interface. Dans ces conditions, le faisceau acoustique est homogene au niveau de la zone ä contröler, ce qui constitue un progres considerable.
Probabilite plus elevee de detecter les defauts mal Orientes
Lors d’un contröle par ultrasons, la probabilite de detecter un defaut est directement liee ä sa dimension et ä son orientation. Les etudes ont montre [5] que les capteurs focalises donnent la probabilite la plus grande de detecter un defaut. Sur le plan de la sürete, ce point est important. 314 DESTRIBATS et al.
FIG.2. Appareil a courants de Foucault a trois frequences.
2.2. Courants de Foucault
Le contröle par courants de Foucault est la seule methode actuellement utilisee dans le monde pour l’examen des tubes des generateurs de vapeur. Les principaux defauts recherches sont: — l’amincissernent par corrosion — les depots — les fissures de fatigue IAEA-SM-218/38 315
— les fissures par corrosion — les chocs sur les plaques supports — le d e n tin g . Les appareils monofrequences generalement utilises ne permettent pas de detecter toutes ces indications, principalement en presence des plaques supports, des barres antivibratoires, des depots, etc. La solution reside dans l’utilisation d’un appareil multifrequence. Dans l’appareil developpe en France (fig-2), trois frequences sont utilisees; il est alors possible de: — detecter des defauts sur la surface externe des tubes dans la plaque ä tubes, meme dans la zone dudgeonnee, et en presence d’une variation dimensionnelle interne telle qu’on peut la trouver sur les tubes fabriques suivant le procede du pas de pelerin; — supprimer les signaux parasites dus aux inclusions magnetiques dans le tube ou aux depots magnetiques situes ä l’exterieur du tube; — mesurer la hauteur des boues au-dessus de la plaque ä tubes; — mesurer l’epaisseur moyenne restante du tube grace ä une voie absolue. De plus revaluation de la profondeur des defauts est aisee en raison de la richesse de l’information obtenue. Si l’appareillage a trois frequences constitue le developpement le plus important, celui-ci ne constitue qu’une partie de l’effort entrepris sur les courants de Foucault, et la description des outils le montrera.
2.3. Emission acoustique
Cette technique en pleine expansion a fait l’objet d’un developpement important qui a conduit ä la realisation d’un appareil de localisation dont on peut resumer ainsi les principales qualites: — detecteurs differentiels; — amplificateur large bande equipe d’un discriminateur ä seuil automatique (asservi au niveau du bruit de fond) et d’un dispositif de mesure de l’energie d’un evenement; — independance totale des mailles, chacune d’elle etant equipee d’un tiroir de comptage; les evenements re$us par une maille n’empechent pas l’acquisition par les autres mailles; — programme de traitement en temps reel permettant la visualisation des donnees acquises maille par maille, en histogrammes des differences de temps de parcours, et de beaucoup d’autres parametres qui facilitent le jugement de l’operateur. Ce materiel de localisation est bien sür Tun des elements du materiel deve- loppe qui comprend d’autres dispositifs dont Tun, equipe d’un micropresseur, permet la surveillance en continu. 316 DESTRIBATS et al.
NIVEAU mini
FIG.3. Representation schematique de la machine de contröle de la cuve.
2.4. Gammagraphie
Avant le demarrage du reacteur, il n’y a aucune radioactivity, mais il n’en est pas de т ё т е apres un certain temps de fonctionnement. La radioactivite developpee dans le bätiment du reacteur peut constituer un inconvenient majeur pour l’utilisation de la gammagraphie. Le premier pas effectue pour tenter de resoudre ce probleme a ete de choisir 192Ir comme source, en raison d’un spectre d’energie tres different de celui du 60Co present dans le circuit primaire. IAEA-SM-218/38 317
La prochaine etape sera d’utiliser des films insensibles au rayonnement de 60Co mais sensibles ä 192Ir. La dimension de la source constitue un compromis entre le temps d’exposition et le flou geometrique. Les differents dispositifs qui mettent en oeuvre les techniques qui viennent d’etre exposees sont decrits dans les paragraphes qui suivent.
3. DEVELOPPEMENT DE DISPOSITIFS
3.1. Machine de contröle de la cuve
La reference [3] decrit la premiere machine developpee en France et qui a dejä ete utilisee sur les cuves des reacteurs de Fessenheim 1, Fessenheim 2 et Bugey 2. La figure 3 est une representation schematique de la deuxieme machine realisee, qui est un modele integre ой tous les outillages sont fixes sur la machine. Cette machine effectue la totalite de l’inspection de la cuve ä partir du moment ou eile est mise en place sur la bride cuve. La commande est faite ä l’aide d’un ordinateur de processus et l’automatisme est total. La machine emet elle-meme un journal de bord. La figure 4 est une Photographie de cette machine lors de sa premiere utilisation sur la cuve du reacteur de Chooz en novembre 1976. La figure 5 est une photographie de la baie de commande. Les developpements actuellement en cours sur cette machine concernent un dispositif de traitement automatique des donnees (dispositif Stadus-Produs) pour faciliter le traitement des informations obtenues. II est ä signaler que l’ensemble de l’electronique necessaire au fonctionnement de cette machine peut etre situe hors de l’enceinte du reacteur et que ceci peut constituer un avantage considerable pour le personnel.
3.2. Machine de contröle de la soudure basse du pressuriseur
La figure 6 est un schema de principe de cette machine qui utilise un capteur focalise de 100 mm de diametre. Le parcours dans l’acier, en ondes transversales, est de 600 mm. En raison de cette distance et de l’epaisseur du fond spherique, un capteur plan classique ne peut en aucun cas assurer le contröle de cette soudure. Avec la machine realisee (fig.7), il est possible de detecter un defaut equivalent ä un trou cylindrique de 2 mm. La presentation des resultats peut etre faite en A-Scan, В-Scan ou С-Scan simultanement. L’interet d’un tel dispositif est evident dans le cadre de la sürete.
3.3. Machine de contröle des tuyauteries primaires
Les capteurs focalises constituent, a ce jour, la meilleure solution pour le contröle des soudures en acier inoxydable [6]; aussi une machine a ete developpee
IAEA-SM-218/38 319
FIG. 5. Baie de commande.
pour l’inspection des soudures du circuit primaire. La figure 8 est une Photo graphie de cette machine sur son stand d’essai. Elle est constitute de quatre elements — le rail, la mini-cuve et les capteurs, la motorisation, la baie de commande — dont chacun a un poids inferieur a 8 kg afin de faciliter les inter ventions. Equipee d’un dispositif а pantographe et ä ressorts, eile peut contröler sans Probleme un coude ou une canalisation ovale, le rail ne constitue qu’une reference et un guide independant. Cette machine prototype constitue le premier pas vers une automatisation du contröle des soudures par ultrasons des canalisations du circuit primaire.
3.4. Machine de contröle des goujons de cuves
Cette machine, presentee ä la figure 9, utilise les ultrasons et les courants de Foucault. Un traducteur ultrasonore avec une lentille correctrice a ete developpe pour detecter les fissures ä la fois a fond de filet et dans la partie courante des goujons. Une sonde speciale par courants de Foucault detecte egalement les fissures eventuelles ä fond de filets. Une fissure de 1 mm de profondeur est aisement detectee par les deux dispositifs. 320 DESTRIBATS et al.
F IG .6. Principe du con trdle de la soudure basse du pressuriseur.
FIG. 7. Machine de contrdle de la soudure basse du pressuriseur. IAEA-SM-218/38 321
FIG.8. Machine de contröle des soudures du circuit primaire.
3.5. Dispositifs de contröle des parties file tees
La figure 10 presente deux dispositifs, portables et automatiques, utilisant les courants de Foucault, qui permettent la detection des fissures dans les parties filetees des goujons de trou d’homme et des ecrous correspondants. Ces memes dispositifs sont utilises pour le contröle des filetages dans les brides. La rusticite et l’efficacite de ces dispositifs en font des outils remarquables. 322 DESTRIBATS et al
FIG.9. Machine de contröle des goujons de cuves. IAEA-SM-218/38 IAEA-SM-218/38 323
FIG. 10. Dispositifs de contröle des parties filetees. 324 DESTRIBATS et al.
FIG. 11. Dispositif automatique d ’introduction et d ’extraction de la sonde.
3.6. Machine d’inspection des tubes des generateurs de vapeur
La figure 2 montre l’appareil multifrequence, la figure 11 le dispositif automatique d’introduction et d’extraction de la sonde, la figure 12 le dispositif dit a r a ig n e e de positionnement de la sonde sur la plaque ä tubes. L’ensemble de ces dispositifs constitue un appareillage unique avec lequel plus de 50 000 tubes de generateurs de vapeur ont ete inspectes ä ce jour. Cet ensemble, entierement automatique, constitue un progres important par rapport aux autres dispositifs existants. Le principe general est decrit dans la reference [7].
3.7. Equipement de gammagraphie
Certaines soudures peuvent etre contrölees par gammagraphie, ä l’aide d’appareils classiques, s’il est possible de vider les circuits; c’est le cas de la soudure mixte entre la canalisation primaire et la boite ä eau du generateur de vapeur. Pour d’autres soudures cette vidange est impossible; aussi, pour controler IAEA-SM-218/38 325
FIG. 12. D isp o sitif araignee de positionnement de la sonde sur la plaque ä tubes.
la soudure mixte cuve-tuyauterie primaire, un appareil special, presente sur la figure 13, a ete developpe. II permet d’evacuer l’eau presente entre la source de 192Ir et la canalisation, par interposition d’une chambre ä air. Cet appareil equipe les machines d’inspection de cuve.
3.8. Localisation par emission acoustique
A l’aide du materiel developpe pour la localisation et afin d’en connaitre les possibilites dans le cadre de la surveillance en continu, des equipements ont ete installes ä titre permanent, d’une part sur le reacteur experimental de la Chaudiere avancee prototype (CAP) ä Cadarache (surveillance de la cuve), d’autre part sur le pressuriseur de Fessenheim 1. Les conclusions que Гоп peut en tirer 326 DESTRIBATS et al.
FIG.13. Equipement de gammagraphie pour le controle des embouts de securite.
а ce jour sont limitees a la reception des informations. En effet, les equipements comprennent ä lä fois des guides d’ondes associes ä des capteurs dits f r o id s , et des capteurs poses directement sur la cuve examinee, dits c h a u d s . Apres trois ans d’utilisation, les caracteristiques des capteurs f r o id s sont inchangees alors que les cap,teurs c h a u d s , apres avoir perdu environ la moitie de leur sensibilite, semblent s’etre stabilises.
4. COMPARAISON DES RESULTATS OBTENUS EN FABRICATION ET EN INSPECTION EN SERVICE
De par la difference des techniques et des moyens mis en oeuvre pendant les controles de fabrication et les inspections en preservice ou en service, les indications trouvees peuvent etre differentes. En ce sens un effort particulier a ete fait, ä la demande des organismes de sürete, pour etablir une correlation entre les resultats. Des essais tres importants sont en cours actuellement en collaboration avec Electricite de France et Framatome. Ces essais montrent une tres bonne correspondance entre les indications de defauts qui, dans le cadre du Cahier de prescriptions, de fabrication et de contröle (CPFC) (equivalent franqais du IAEA-SM-218/38 327 code ASME), sont considerees corame ä rejeter et cette correlation reste bonne meme pour les indications de faible amplitude. II faut rappeler que la severity du contröle par ultrasons en appliquant le CPFC est tres superieure a ce qui est obtenu en appliquant le code ASME, Section III. Les dispositifs utilises pour l’inspection en service assurent toujours une reproductibilite remarquable. La cuve du reacteur experimental CAP, examinee ä deux reprises a un an d’intervalle, a permis de montrer le tres grand interet de l’automatisme associe aux traducteurs focalises pour assurer la reproductibilite.
5. CONCLUSION
Les exigences de la fiabilite des composants de reacteurs ont conduit le Commissariat ä l’energie atomique ä developper de nouvelles techniques et de nouveaux outils pour le contröle non destructif. Ces methodes et ces outils represented un pas en avant dans la connaissance des defauts. On peut se demander si de tels outils ne devraient pas etre utilises durant la fabrication; les etudes et essais en cours apporteront une reponse ä cette question. II reste necessaire d’acquerir une experience de plus en plus grande avec ces nouveaux outils, qui constituent l’une des seules possibility d’obtenir une information correcte sur la nature des defauts. Les methodes exposees sont maintenant etudiees dans d’autres pays; aussi la naissance d’une cooperation internationale ä ce sujet est-elle hautement souhaitable pour une fiabilite et une sürete accrues des installations tant industrielles que nucleaires.
REFERENCES
[1 ] SAGLIO, R., «Routine checkout), Conf. on Periodic Inspection of Pressurized Components, Londres, 4—6 juin 1974, Institution of Mechanical Engineers, Londres, Conf. Publication 8 (1974) Cl 16/74. [2] SAGLIO, R., PROT, A., «Improvements in ultrasonic testing methods of welds specially in the presence of austenitic stainless steel cladding), 2nd Int. Conf. on Pressure Vessel Technology, San Antonio, 30 sept.—4 oct. 1973. [3] PROT, A., SAGLIO, R., «Survey of the French developments in the field of in-service inspection), IAEA Technical Meeting, Kobe, Japon, 25—27 avril 1977. [4] SAGLIO, R., TOUFFAIT, A.M., PROT, A.C,, «Determination des caracteristiques des defauts de soudure ä l’aide de traducteurs focalises), XXXth Annual Assembly, Int. Inst, of Welding, Copenhague, 4—9juillet 1977. [5] SAGLIO, R., Better detection of large poorly oriented plane defects by ultrasonics, Non-Destr. Test, (aoüt 1976). 328 DESTRIBATS et al.
[6] SAGLIO, R., ROULE, M., Progres realises dans le contröle automatique par ultrasons des soudures d’acier inoxydable austenitique ä l’aide des traducteurs focalises, Rev. Soudure (Bruxelles) 2 (1974). [7] PIGEON, M., SAGLIO, R., «Inspection en service des tubes de generateurs de vapeur ä l’aide d’un appareil ä courants de Foucault multifrequences», 8e Conf. mondiale sur les essais non destructifs, Cannes, 1976. DISCUSSION
on papers IAEA-SM-218/37 and 38
R.I. HODGE: The development of the equipment described in paper SM-218/37 is impressive. Can you comment on the problem of data interpretation in relation to in-service inspection? A.C. PROT: With the French device used, data interpretation is quite easy since C-scan presentation is involved. Only relevant flaws are considered and a comparison is made by superposition of successive cartographic displays. Never theless, we are also making a microprocessor study of automatic data processing systems in order to improve the interpretation time for data reduction. N.F. HAINES: It is interesting to note that both the French and German ultrasonic inspection systems now use a frequency of 1 MHz. However, when the frequency is reduced, cracks which are either corroded or contain water are known to transmit ultrasonic energy more intensively. Has any consideration been given to this problem in France? R. SAGLIO (Chairman): Studies done both in Germany and in France show clearly that the probability of detecting faults is increased considerably when the frequency is diminished. However, the situation is very different with cracks situated near the internal surface, and in this case we in France use a frequency of 4 MHz and an acoustic beam 2 mm in diameter; since 1971 we have published many papers on this subject. I should point out that the use of focused transducers also increases the probability of faults being detected, and therefore increases the reliability of inspection. M. KITTNER: Could you say what you consider to be acceptable failure sizes in in-service inspections as compared with acceptable failure sizes in pre- operational inspections? A.C. PROT: During preoperational inspection only two possibilities need to be considered, namely, whether the flaws discovered correlate with fabrication testing, and whether new flaws are discovered; the latter are to be treated as during fabrication. During periodic in-service inspections, the situation is totally different. Once a defect has been detected and measured, it needs to be analysed by fracture mechanics or similar techniques. This means that a number of parameters have to be considered, such as the position relative to the surface, the orientation or nature of the defect, the shape of the component in which the flaw was detected, the stress field around the flaw during operation or even during postulated incidental situations and the metallurgical history of the material. From the analysis performed one can get an idea of the harmfulness of the defect. During fabrication non-destructive testing (NDT) is only used to
329 330 maintain a constant quality level, and normal arbitrary specifications can be used. This is no longer true during in-service inspection, when a decision has to be taken as to whether the component may be put into operation again, or whether it needs to be repaired or replaced. IAEA-SM-218/40
A METHOD FOR ESTIMATING FLAW DETECTION PROBABILITY FROM INSPECTION DATA
D.H. SHAFFER Westinghouse R & D Center, Pittsburgh, Pennsylvania, United States of America
Abstract
A METHOD FOR ESTIMATING FLAW DETECTION PROBABILITY FROM INSPECTION DATA. A novel approach to estimating the probability of detecting (or not detecting) a flaw as a function of its size is proposed and developed. This approach does not require any associated destructive testing except to validate and normalize the method; it depends instead on some properties of extremal statistics and a willingness to collect non-destructive inspection data in a particular way. An extension of the method to allow the use of more conventional inspection data is also proposed. Let the structure for which the probability of flaw detection is required be divided into small sub-regions. Concern will be focused on the largest flaw in each region, and it is assumed that a suitable measure of equivalent flaw size, regardless of type, is available. Arguments can be constructed to the effect that the size distribution for the actual largest flaws is adequately represented by the first asymptotic extreme value distribution. A procedure for asymptotic fitting of the observed flaw inspection data is then given which allows the estimation of both the real and observed flaw size distributions. It is then shown that the solution of a simple integral equation yields the required information about the probability of flaw detection as a function of flaw size. If this relationship can be normalized at one flaw size, an estimate of flaw detection probability is available at each flaw size; otherwise, a comparison of flaw detection probability between any two flaw sizes is still available.
1. INTRODUCTION
In recent years a number of investigators have been developing methodology intended to estimate the probability of brittle rupture in heavy section structures. (See e.g. Refs [1,4, 5,7].) The primary concern of the above has been the long term integrity of a nuclear pressure vessel, considering both the mechanism of flaw growth during stress cycles from anticipated transients and the reduction of fracture toughness due to irradiation embrittlement. In view of the uncontrolled variability in both material properties and operating environment, the problem of integrity assessment has been formulated probabilistically, i.e. with several of the
331 332 SHAFFER governing parameters being interpreted as random variables with distributions inferred from available relevant data. The usual first result to come from such a formulation is a probability of brittle rupture conditional on a given existing flaw. Sometimes the probability is evaluated for the much-discussed Appendix G gT flaw; sometimes the probability is expressed parametrically over a range of flaw sizes. Whatever the form of the presentation, it is necessary to remove the “existing flaw size condition” in order to have a result of real practical importance. Although the question of the distribution of flaw sizes in a structure is complicated by such matters as variability in flaw growth rates both from differing material properties and fluctuating operating conditions, the heart of the problem goes back to the quality of the pre-service inspections: How well do the results of non-destructive testing correlate with the actual flaws that are present? A limited amount of properly paired data, relative to this question, is available. The high cost and effort needed to do destructive sectioning of material already subjected to NDT precludes the existence of massive amounts of data to develop definitive correlations. Nevertheless, some methodology is being constructed (see e.g. Ref. [6]) that will use some of this correlative information, as it becomes available, to estimate the probability of detecting (or not detecting) flaws. Once such estimates become available, they can be combined with conventional inspection data to generate meaningful initial flaw size distributions. The present paper offers a somewhat novel approach to the specific question: What is the probability of detecting (or not detecting) a flaw as a function of its size? The development of this approach does not depend explicitly on correlations between NDT results and destructive examinations. Rather, it uses some properties of the inspection data that are consequences of considerations in extreme value statistics. The application of the method, as it now stands, requires some modification of the way inspection data are collected, but this requirement seems not to be objectionable. Verification of the method, yet to be done, will call for some destructive/non-destructive correlations, but only as method verification. Meantime, its novelty and its employment of concepts from extremal statistics will hopefully encourage new ideas for other investigators.
2. FORMULATION
Consider the structure to be inspected and let its surface be divided into N equal regions. Let the inspection proceed and denote by yj the size of the largest 1AEA-SM-218/40 333 flaw that is found in the j-th region. By this process we can generate a sample {yj, j = 1,2, . . . N} of the largest observed flaws in the N regions. If we associate the random Y with these observations, we define the distribution function F(y) as
F(y) = Pr(Y < y)
Two additional probability functions need to be defined for our development. First, we conceive of the population of the largest flaw that exists in each region of the structure. Let X be a random variable associated with this population of existing flaws, and let G(x) be its distribution function, as
G(x) = Pr(X < x)
In addition, we define a probability p(x) as
p(x) = Pr (observing a flaw of size x | there exists a flaw of size x)
We now proceed to derive an expression for p(x) in terms of F(y) and G(x) and their related density functions f(y) and g(x). Note the following
number of observed flaws in size range (у, у + dy) f(y )d y = ------total number of existing flaws
number of existing flaws in size range (x, x + dx) g(x)d x total number of existing flaws
number of observed flaws of size x P(x) number of existing flaws of size x
number of observed flaws in size range (x, x + dx) number of existing flaws in size range (x, x + dx)
Then
Kf(x)dx =g(x)-p(x)dx where
total number of observed flaws total number of existing flaws 334 SHAFFER
Integrating both sides gives
К = / g(x)p(x)dx
0
But
f(x) P(x) = К - j — g(x) so
F '(x) p(x) = G'(x)p(x)dx G '(x) /
We note that, if p(x) satisfies the above equation, so does Xp(x). The multiplier X simply reflects a change in the fraction of existing flaws that are observable. The multiplier К may be replaced by renormalizing the expression for p(x). Let x0 be a value of flaw size such that it is equally likely that the flaw will be observed or not observed. Then
f ( x 0) ]_ p (x 0) = K g (x 0) 2
SO
1 g (x 0) f(x) p(x) = ------■------2 g(x) f(x0)
It remains to show that there are rational choices for the functions f(x) and g(x) that make this formulation useful.
3. EXTREMAL CONSIDERATIONS
Consider a sample of size n drawn from a population whose probability density function is 0(x). Let the sample values be ordered by increasing IAEA-SM-218/40 335
magnitude.and let z denote the largest observed value. It the sampling procedure just described were repeated a number of times, a set of sample values of z would be obtained, each z being the largest from a parent sample of n. The z values themselves are realizations of a random variable Z, whose probability distribution must be derivable from the parent density 0(x); let the probability density function for Z be denoted by £(z). The random variable Z, because of the way it has been generated, is known as an extremal or extreme value statistic. A discipline has been built up around extreme value statistics and has found diverse application to failure mechanisms of the “weakest link” type. (For the basic results in this discipline see Refs [3 and 2].) It can be shown that the form of £(z) is largely independent of 0(x) if n is large enough; more precisely, the £(z) has a limiting form as n -> °o. The asymptotic form for £(z) is dependent only on the gross tail behaviour of 0(x). Gumbel [3] develops three asymptotic forms for £(z), corresponding roughly to these conditions of 0(x):
I: 0(x) = o(e~x), x ->• °o
II: 0(x) = o(x~c), x -* ■ ° °
III: 0 (x ) = 0, x > a
We will confine our attention to the first of these for the balance of this paper. Recall that the flaw size data described earlier consisted of the largest observed flaw from each of the regions on the surface of our structure. Presumably, each region contains a number of flaws, most hopefully very small; we shall denote the probability density function for the existing flaw sizes by 7 (x); i.e. b
0
We also presume that 7 (x) applies throughout the structure. If our observational techniques were perfect, i.e. if we could detect, without error, each and every flaw present, then we might anticipate that, not only would g(x) and f(x), as previously defined, be equal to each other, but they could each be represented by one of the extreme value distributions. We now take, as a basis for further development, the following proposition: The density function g(x), related to the largest existing flaw in each region on the structure, is satisfactorily represented by the first extreme value distribution, namely
-(a + b x) G (x) = e~e , b > 0 336 SHAFFER where
g(x) = G'(x)
Furthermore, the distribution function F(x) for the observed largest flaws will be asymptotic to G(x) on the assumption that as the flaw size grows the distinction between “existing” and “observed” vanishes.
4. ESTIMATION
In accordance with the preceding, we take
c , , -(a + bx + w(x>) F (x) = e e , b > 0
where со -* ■ 0 as x -* ■ and со is differentiable for all x > 0.
A graphical check on the quality of the assumed forms for the distributions G(x) and F(x) is readily made, and should be carried out. If F(x) is as shown above,
lnF (x)= -e-(a + bx+cj(x)) and
In In ------= - (a + bx + со (x)) F (x)
Let an ordered sample of values {xj} be denoted by xj, xj, . . . , x^. Corresponding to xj we take F(xj) = i/(n + 1), and write
(n + Л qj = -In In ( — ;— I = + (a + bXj + co(Xj))
If the pairs of points (qi; xj) are now plotted, we would expect to see a curve that is becoming increasingly linear for the larger Xj and is concave upward over the entire range. A typical curve is shown in Fig. 1. This graphical display, in addition to verifying the distributional assumptions made, also suggests a fitting procedure to be followed: Select a form for co(x) that satisfies the asymptotic conditions imposed. Examples might be co(x) = с/х + d/x2, or c j ( x ) = ce_x. Estimate the parameters in a + bx + co(x)by IAEA-SM-218/40 337
Reduced Variate
FIG.l. Plot of points (q-v x{) showing a typical curve. whatever method is convenient. Let the estimates be represented by ä + bx + cb(x). Now we take
_p-(ä + bx + w(x)) F;(x) = e e and
e-(ä + b x) G (x) = e because of the asymptotic identity imposed on F and G. From these we calculate
f(x) = [b + g(x) = b e “(ä + 6x)G (x) Using the formula for p(x) (the probability that a crack will be observed) as derived in Section 2, we obtain G (x 0) F(x) P(x) = — ------2 F (x 0) G(x) b + w '(x 0) > where p (x 0) = 338 SHAFFER We comment in passing that it may sometimes be difficult to find a suitable analytic form for oo(x). It is still possible to carry out approximations to the indicated computations. If the data are plotted on extreme value paper, and a smooth curve interpolated between them, an asymptotic straight line can also be estimated visually and from these, estimates of f'(x) and g'(x) obtained. 5. CONCLUDING REMARKS To date, no data have been taken in the required form to allow a full application of the proposed methodology. Some data purporting to measure flaws in pressure vessel weld material have been studied, and the method looks promising, but no report can be made yet. There appear to be three important steps that are required before this method can be considered widely useful: 1. It must be carefully and thoroughly tried on data gathered for the purpose. Logic and reason can produce methodology that is then offered for consideration, but only successful applications can win it acceptance. 2. The method, if promising, should be extended so that it does not require specially collected data. In early attempts to investigate it, we divided the surface that received NDT into small enough regions that only one flaw (observed) registered per region. An observation of zero was recorded for the empty regions, and the resulting data were then considered as censored in those regions. 3. The method, if promising, needs considerable work in investigating candidate functions co(x). It can be seen that the large-value behaviour of p(x) is dependent on so a good choice of со is essential. REFERENCES [1] BECHER, P.E., PEDERSEN, A., “Application of statistical linear elastic fracture mechanics to pressure vessel reliability analysis”, Structural Mechanics in Reactor Technology (Proc. 2nd Int. Conf. Berlin, Sep. 1973, North-Holland, Amsterdam (1974). [2] EPSTEIN, B., Elements of the theory of extreme values, Technometrics (Feb. 1960). [3] GUMBEL, E.I., Statistics of Extremes, Columbia University Press (1958). IAEA-SM-218/40 339 [4] JOURIS, G.M., SHAFFER, D.H., “Probabilistic brittle fracture analysis for major thermal transients in pressure vessels”, Structural Mechanics in Reactor Technology (Proc. 3rd Int. Conf. London, Sep. 1975). [5] JOURIS, G.M., SHAFFER, D.H., Use of probability with linear elastic fracture mechanics in studying brittle failure in pressure vessels, Int. J. Pressure Vessels Piping (in press). [6] JOURIS, G.M., SHAFFER, D.H., “A procedure for estimating the probability of flaw nondetection”, Structural Mechanics in Reactor Technology (Proc. 4th Int. Conf. San Francisco, August 1977). [7] NILSSON, F., “A model for fracture mechanical estimation of the failure probability of reactor pressure vessels”, Pressure Vessel Technology (Proc. 3rd Int. Conf. Tokyo, April 1977). DISCUSSION G.I. SCHUELLER: I would like to congratulate the author on the method he has put forward. In my opinion this is the best way of linking inspection procedures and quality control, i.e. linking quality assurance and reliability in a quantitative manner. I am also in favour of the author’s way of applying extreme value theory. In view of the relative insensitivity of the various extreme value distributions relative to their parent distribution one can dispense with the endless discussion as to which distribution should be used, i.e. normal, log-normal, gamma, beta, etc. IAEA-SM-218/41 THE RELIABILITY OF ULTRASONIC INSPECTION N.F. HAINES Berkeley Nuclear Laboratories, CEGB, Berkeley, Gloucestershire, United Kingdom A b s t r a c t THE RELIABILITY OF ULTRASONIC INSPECTION. Ultrasonic techniques have become the preferred method of volumetric examination of the walls of thick section reactor pressure vessels. Different countries have devised methods of calibrating ultrasonic systems and establishing acceptance criteria; in general, these are based upon the amplitude of the reflected signal. Two particular examples are the American ASME XI code and the German Reactor Safety Commission guidelines. In order to assess the limitations of amplitude threshold criteria, work at Berkeley Nuclear Laboratories has been aimed at providing a theoretical model of pulse reflection from surfaces. Crack size, shape, orientation and surface roughness as well as the ultrasonic pulse shape are taken into account in the model. Particular geometries which have been included are elliptic, planar (rough and smooth), cylindrical and spherical. The results have been interpreted to demonstrate the limited range of orientation over which a crack of any size may be detected by applying either the American or German criteria using a common transmit and receive system. A discussion is given of two approaches to overcome the problems of detecting a crack growing directly into the wall of a vessel, either by the German Tandem system or by use of a В-scan presentation of data. The model presented only considers the physical limitations set by the geometry of the crack. No consideration has been given to attenuation caused by the surface condition of the wall of the vessel, presence of cladding, etc. The model, therefore, presents the optimum results that can be obtained with a simple pulse echo system. 1. INTRODUCTION Ultrasonic techniques have become the preferred method of volumetric examination of the walls of thick-section reactor pressure vessels. The inspection procedure is split into two stages; the first detects and positions a defect and the second measures its size. The reliability of a particular technique principally depends upon the first stage of detection although, without an adequate sizing technique, the whole inspection would become ineffective. It is assumed, for the purpose of this paper that adequate sizing procedures exist and, therefore, that the reliability of an inspection is simply the probability with which defects of a given size may be detected. The most direct method of assessing the reliability of a particular inspection procedure is from the statistical analysis of the number of successes and failures to locate defects in a population of similar components. 341 342 HAINES For example, from the comparison of the destructive and non-destructive examination of a large number of similar bolts it is possible to establish a figure for the probability of detecting defects. Clearly for economic reasons it is impossible to apply the same method to the inspection of pressure vessels. Exercises to examine large specimens into which real or artificial flaws have been introduced are also expensive and provide a very small sample from which statistically to assess reliabilities. The alter native approach is to attempt to construct a theoretical model of the physical processes involved in ultrasonic inspection and to deduce the con ditions under which a particular technique may be ineffective or, at least, unreliable. Some consideration has been given by Wustenberg [1] to the physical factors which affect ultrasonic inspection although the model presented is far from exhaustive in terms of the number of parameters which influence the 'visib ility' of defects. Orientation of a defect relative to the incident beam is recognised as the most important parameter in deter mining the reflected amplitude of an ultrasonic pulse; shape as well as size and surface roughness of the defect also affect the signal. Work at Berkeley Nuclear Laboratories has been aimed at providing a model of pulse reflection at surfaces which combines a ll of these parameters together with the pulse shape of the incident wave. Simple programs have been written which may be run on small computers (< 28K words of memory store) to predict the reflected wave form and hence peak amplitudes. Laboratory experiments have verified the model over a range of different shapes of reflectors and surface rough n e ss. The results presented in this paper have been interpreted to assess the capability of a common transmit and receive ultrasonic test system. The results have been compared with the American [2] and German [3] recommended amplitude threshold levels for this type of inspection and conclusions drawn on the range of defect size and orientation which may be located if these levels are applied. 2. THE DEPENDENCE OF SIGNAL AMPLITUDE ON DEFECT GEOMETRY AND SURFACE ROUGHNESS Three different classes of shape have been used to simulate both defect and calibration reflectors. The ellipse has been used to represent a general planar surface; by varying the axes the effect of width (direction parallel to the plane of the incident wave front) and length (direction making an angle 6 with the incident wavefront) on the signal amplitude may be examined. Surface roughness may be superimposed on the ellipse to simulate a more realistic defect. Two classes of roughness are identified; the first is a fine roughness as found on fatigue cracks and the second is an extremely angular surface composed of a number of discrete facets. This latter is more representative of some types of weld fracture. A lim iting case of the ellipse is a circular reflector normal to the incident wave front and hence represents a flat bottom hole (FBH) as used by the German Federal Republic as a calibration standard. The cylinder is recommended by the American ASME XI code as a calibration reflector and has therefore been included. The sphere has been used to simulate an included type of defect; although it is unlikely that a defect is exactly spherical, the results may be inter preted as the reflection from a convex surface directed towards the incident wave front. This point is discussed in more detail in Section 3. 2,1 General Theory The detailed theory w ill not be developed here since it has been described previously by Haines and Langston [4]. The model used is an extension of the work of Neubauer [5] which has been used by Johnson [6] to successfully account for the experimental observations of Whaley and Adler [7]. Neubauer assumed that the reflection from a target can be simply IAEA-SM-218/41 343 described by a diffraction limited model and that a Huygens' construction can be used to generate a complex frequency reflection coefficient for a surface. Since the results are based on a simple linear wave equation, each component frequency (f) of the incident pulse may be treated separately each being changed on reflection in both amplitude and phase. The reflected wave may then be expressed in the form ♦r(f) “ ^(f) G(f) (1) where ф (f) and ф.(f) are the complex Fourier transforms of the reflected and incident pulsls and G(f) is a complex reflection coefficient of the surface. If the system is calibrated against a small flat reflector lying normal to the direction of the incident wave and having an area A^,^ then following Johnson [6] ( 2 ) ♦r(f) ■ W f) F(f) _ 1 ___ - 4 tt j Azf where F ( f ) exp ) dA ' (3) c ACAL [ , ф (f) is the complex Fourier transform of the pulse reflected from the calibration reflector and c is the velocity of sound. Equation (3) shows that F(f) may be evaluated by integrating the exponential phase term over the projected area A* of the surface onto the plane of the incident wave front. Az is the distance of an elemental area 6A' from an arbitrary position of the wave front measured parallel to the direction of travel of the incident wave. Equation (3) has been evaluated for several geometries by both Johnson [6] and Haines and Langston [4]. A much simpler interpretation of equations (2) and (3) may be obtained by transforming into the time domain as shown by Haines and Langston [4]. In this case the equations predict the reflected pulse shape ф (t) in terms of the calibration pulse shape ф (t) and the impulse response function of the surface F(t) ( 4 ) *r(t) = *CAL(t) * F(t) __L__ 2Az where F ( t ) &( t - c ) dA ' (5) a c a l A' and * denotes the mathematical operation of convolution, 6 the Dirac delta function. The simple physical interpretation of equations (4) and (5) is that the reflected pulse from a surface can be considered as the summation of a number of elemental pulses each identical in shape to the calibration pulse and having an amplitude proportional to the projected elemental areas of the surface <5A'. The pulses are superposed with the correct time separation according to the parameter 2Az/c. Equation (5) has been used to obtain the impulse response functions of the ellipse, cylinder, sphere and for a surface having a fine scale roughness. The response functions are shown diagramatically in figure 1 and their explicit mathematical forms are given in appendix 1. Experimental results have been obtained for surfaces immersed in water using a 5MHz, 12mm 344 HAINES FIG.l. Diagrammatic impulse response function. 6mm diameter FBHsOdb -ASME Я compression ASME Я shear -3mm FBH T6 26 2i 28-Shear 3$Tilt (degs)—Compression FIG.2. Signal am plitude versus orientation - ellipse. IAEA-SM-218/41 345 diameter ultrasonic transducer and for flat bottom holes in blocks of aluminium. A ll immersion experiments have been performed at a range of 450mm giving a null to null width (2d) of the acoustic pressure distribution at 5MHz of approximately 26mm according to the equation 2d = 2.4 ~ (6) where R * range, D * diameter of transducer (see, for example, Krautkramer [ 8 ] ) . Results have been predicted for 25mm diameter, 2.25MHz compression and shear wave transducers at ranges of lOOram and 200mm respectively in steel. These figures have been chosen such that the null to null distance is the same for the immersion experiments as for the compression and shear waves in steel and hence the reflectors in all cases have the same incident acoustic pressure distribution. 2.2 Reflection from an Ellipse To simplify the analysis of the ellipse it is assumed that one axis lies parallel and the other at an angle 0 to the plane of the incident wave front. Little generality is lost since it is possible to explore practically all shapes of defect by varying the values of the two axes. From equation A1 in appendix 1 it may be seen that increasing the tilt of the ellipse (increasing 0) extends the width of the response function and decreases the height. The reflected pulse is, therefore, reduced in amplitude and in- creased in width with increasing orientation relative to the incident wave front. Increasing the dimension of the ellipse parallel to. the incident wave increases all ordinates of the response function by a constant factor and hence increases the amplitude of the reflected pulse in proportion. In creasing the other axis of the ellipse increases the amplitude at normal and near normal incidence but has a more complicated behaviour at larger angles. To illustrate these effects figure 2 shows the variation in ampli tude relative to a 6mm flat bottom hole for 3 different shapes of ellipse. To a reasonable approximation up to 30 Sin0 = Tan0 = 0 and hence it is possible to use a single graph with two scales of orientation, one for a shear wave and one for a compression wave. Experiments were performed on a range of sizes of disc reflector in water and flat bottom holes in aluminium. The results and comparison with the theoretically predicted behaviour are given in figures 3, 4 and 5. An important observation is the behaviour of the 12mm disc at normal incidence shown in figure 4. Theoretically the model assumes an infinite plane wave incident on the reflector and hence at normal incidence the signal amplitude should increase directly with the area of the reflector. However, because of the finite width of the beam from the transducer the experimentally observed signal amplitudes fall below those theoretically predicted. However, this only occurs at small angles, the agreement being good at higher orientations of the disc. The saturation effect of amplitude versus reflector area at normal in c id e n ce i s c l e a r ly im portant in lim it in g the maximum am plitude that can be received by an ultrasonic transducer. To investigate this further a number of different size reflectors were placed at normal incidence and the results are shown in figure 6. It is striking that the signal amplitude is within 2db of its saturation level for a disc of only 12mm in diameter. This corresponds to approximately half the null to null distance for the trans ducer field as calculated above in section 2.1. Since the ultrasonic field conditions are the same for the predicted results of figure 2 this means 346 HAINES 6 mm flat bottom holeoOdb Theoretically predicted curves for 5 MHz, 12mm transducer. Range 450 mm in water. 3mm FBH -4 /<5pm RMS j j / roughness 0 05 VO 10 20 20 3Ö 30 40 Tilt (degs) FIG.3. Experimental results on rough and smooth 6 mm diameter discs. FIG.4. Experimental results for a 12 mm diameter disc. IAEA-SM-218/41 347 FIG.5. Experimental results on a 6 mm diameter disc in aluminium. 20 6 mm diameter FBH = Odb 16 12 .a Saturation level *o 8 X CD 4 E Experimental results for E 0 a 5MHz, 12mm transducer. to Range 450mm water. a> 4) TJ D 25 50 75 100 125 150 175 200 Area of disc reflector mm2 FIG. 6. Experimental results for discs at normal incidence. that the 6mm x 6mm (half axes of ellipse) reflector is the largest that w ill give any increase in signal amplitude at normal incidence. Further experi ments were performed on discs larger than 12mm and little (2db) increase in signal amplitude was observed at normal or other orientations. To a good approximation, therefore, the curve for the 6mm x 6mm ellipse in figure 2 represents the behaviour of any reflector having semi-axes greater than these. 348 HAINES hear wave 6mm FBH =Odb Theoretically predicted curves for 2'25 MHz 25mm transducers. Compression wave range 100mm in steel. Shear wave range 2 0 0 mm in steel. 4 8 12 16 20 24 Diameter of cylinder (mm) FIG. 7. Signal amplitude versus diameter - cylinder. Compression Shear ASME Я compression Theoretically predicted curve for 2-25 MHz. 25mm transducer. Range :200mm in steel shear wave. 20 40 60 80 100 120 140 Diameter of sphere (mm) FIG.8. Signal amplitude versus diameter - sphere. 2.3 Reflection from Cylinders and Spheres The predicted behaviour of cylindrical and spherical holes in steel are given in figures 7 and 8 and the experimentally observed behaviour is shown in figures 9 and 10. The theoretical curves have been calculated for all cases using the impulse response functions given in appendix 1 and shown diagramatically in figure 1. For the cylinders the experimental results were obtained with long metal rods immersed in water, however, the length irradiated by the transducer was assumed for the purposes of the theoretical calculation to be 12mm in accordance with the behaviour of the discs discussed in section 2.2 above. The excellent agreement between experiment and theory again confirms the very restricted effective beam diameter of the transducer at the targets. IAEA-SM-218/41 349 .6 mm. diameter flat bottom hole ~ - и л Expt. observed *o results Theoretically . -6 predicted CD U. E -8 E 5 MHz 12mm transducer ш К range Л50 mm in water -S-12 э | - u о S;-i6 -20, 0 4 8 12 16 20 24 Diameter of cylinder (mm) FIG.9. Experimental results for cylinders. 6 mm flat bottom hole = 0 db FIG. 10. Experimen tal results for spheres. 2.4 Effect of Surface Roughness For a fine scale of roughness where the surface is undulating rapidly over the area of the target illuminated by the transducer it is possible to assume a Gaussian distribution of heights of the surface about its mean plane. Haines and Langston [4] have shown that this leads to a response function for a rough surface which only depends on the rms height distribution of the surface о and that the response function is a Gaussian of standard deviation 2o fc where c is the velocity of sound in the surrounding medium. This is shown diagramatically in figure 1 and the explicit mathematical form 350 HAINES FBHsOdb smooth roughness Theoretically predicted curves for 2‘25MHz, 25mm transducer о 'S -32 Ql. Range :100 mm in steel compression wave -36 8 12 16 20 24 28 32 Tilt (degs.) FIG.ll. Effect of roughness on signal amplitude (compression wave). Theoretically predicted curves for 2'25MHz, 25mm transducer 0 4 8 12 16 20 24 28 T ilt (d e g s) FIG.12. Effect of roughness on signal amplitude (shear wave). IAEA-SM-218/41 35 is given in appendix 1. The effect on the signal amplitude may than be simply calculated from the equation ♦ r where F (t) is the shape response of the smooth surface (e.g. the ellipse of equatioS Al) and R(o ,t) is the roughness response of equation A4. For a reflector at other tßan normal incidence a small correction must be made by replacing a by a cos0. The theoretically predicted behaviour for com pression an§ sheaf waves is shown in figures 11 and 12 and the experimentally observed behaviour is shown in the lower curve of figure 3. It is interest ing to note that for the surface having a 41ym rms roughness in figure 3 the attenuation at normal incidence relative to a smooth reflector of the same size is some 10 - 12db. The wave length at 5MHz in water is 300ym hence the ratio of а /X is only0.14. Further calculations showed that surface rough ness may only be neglected for а /X <0.05 in contrast to the suggestion of Wustenberg [1]. S For an extremely angular surface it is not possible to assume a Gaussian distribution of surface heights over a small area of the surface and hence not possible to obtain an analytic mathematical expression for the response function. For any given section of surface it is possible to geometrically construct a response function as shown in figure 13. It is rapidly seen from such a construction how large, preferentially, oriented facets of a surface give spikes in the response function which in turn give a reflected pulse proportional to their height. It is possible, therefore, to a first approximation to consider a surface of this type as. a large number of facets each reflecting independently. Hence the probability of detecting such a defect can be enhanced over a smooth defect by the occurrence of the occasional facet at near normal incidence to the incident beam. However, the probability of a large facet occurring at a high angle to the general direction of the crack would be small compared with the probability of a small facet at a similar angle. It appears, therefore, that it is necessary to study crack topographies with a view to establishing probability density functions for the occurrences of facets as a function of both size and orientation relative to the mean direction of the crack. No such exercise has been performed as part of the present work although the relation ship between facet size, orientation and signal amplitude is shown in f ig u r e 14. 3. INTERPRETATION OF RESULTS The ASME XI code recommends that the reporting threshold for a defect at a depth of 100mm is 50% of the amplitude from a J" diameter (6.35mm) hole. From figure 7 this corresponds to a level of -7.2db and -9.9db for compression and shear waves respectively relative to the 6mm diameter reference flat bottom hole assumed in all the figures. For a 60° shear wave transducer the range in steel to a defect 100mm below the surface is 200mm which is also the range in steel assumed in all the figures. The German code recommends a 3mm flat bottom hole at normal incidence as the reporting threshold for defects. This has an area of 25% of the 6mm disc reference used in the present work and, therefore, corresponds to a level of -12db. Experimentally, figure 7 confirms that a 3mm disc 7mm2) does indeed give a signal 12db below the 6mm (28mm2) disc. The two reporting levels have been added to all the figures predicting the behaviour of reflectors in steel. In figure 2 the ASME XI level corres ponds to an orientation of the 6mm x 6mm (semi-axes) ellipse of 11° and the 3mm flat bottom hole to 12.5 for an incident compression wave and the 352 HAINES Mean direction of crack Crack surface t = 2 4z/c Response function r ' Amplitude «. to projected area of facets on to incident 1 wavefront. FIG. 13. Schematic geometric construction of response function for a faceted crack surface. 6 m m F B H = 0d b Theoretically predicted curves for 245MHz, 25mm transducer. In steel. RmgeitOOmm compression 200mm shear Тб 20 2i 28— Shear 32 Tilt (degs)— Compression FIG.14. Signal amplitude versus facet size and orientation. corresponding figures for а 60° shear wave are 6 to 7° for both codes. As discussed in section 2.2 the 6iran x 6iran ellipse is the largest reflector that would give any increase in signal for the 2.25MHz, 25mm transducer assumed in this work at the given ranges. Hence it is possible to say that no defect greater than this size would be 'visib le' at the ASME XI threshold if it were at an angle greater than 11 for a compression wave and 6.5 for a shear wave. The figures for the German code at the respective ranges are similar. For the particular case of using a 60° shear wave transducer to search for a defect growing directly into the wall normal to the scanning surface the angle of incidence would be 30 . From figure 2 a large smooth defect (> 6mm x 6mm half axes) would only be detected at a threshold level approximately 20 to 22db lower than either of the codes. For a similar defect having a faceted surface a decrease in threshold of between 13 and 15db would allow facets greater than 2mm in diameter making an angle of greater IAEA-SM-218/41 than 16° to the mean direction of the crack to be detected. However, it may be that only a few such facets exist and, therefore, only a few relatively unconnected indications would be observed. At an increase of 22 - 24db all facets greater than 1.5mm in diameter and making an angle of greater than 8 to the mean direction of the crack would be detected. At this level it is likely that a number of such facets may exist and the number of indications along the crack length could possibly increase sufficiently to make the crack appear continuous. It is interesting to note that whether the crack is smooth or faceted it is unlikely to be detected as a continuous crack unless the gain is increased over the two recommended thresholds by some 20 - 24db. The behaviour of spherical reflectors in steel is shown in figure 8, to which the respective recommended threshold levels have been added. Analysis of the reflection from a spherical cap rather than a complete sphere confirmed that only the part of the sphere close to the nearest point of approach to the transducer actually affects the reflected signal amplitude The result on a sphere may be interpreted, therefore, as the signal amplitude from a convex surface having the same radius of curvature as the sphere. In figure 8 the 3mm flat bottom hole threshold corresponds to a radius of curva ture of 6mm and the ASHE XI threshold to a radius of 10mm for a compression wave. The corresponding figures for a shear wave are 11 and 14ram respective!; Hence the detection of included defects composed of materials of low acoustic impedance relative to that of steel is possible,provided the radius of curvature directed towards the incident wave front is not less than these f i g u r e s . 4. DISCUS'S ION The ranges of 100mm and 200mm for compression and shear waves in steel were specifically chosen so that the acoustic pressure distributions across the reflector were the same as in the laboratory immersion experiments. If the ranges were increased the curves of signal amplitude against orientation would remain unchanged relative to the flat bottom hole reflectors. However, for the cylindrical calibration reflector of ASHE XI the length of the cylinder receiving energy would increase relative to the flat bottom hole. For example, doubling the range would double the length of cylinder being isonified and, therefore, (see equation A2) double the amplitude of the signal received back at the transducer. Hence at ranges of 200mm and 400ram the two ASME XI reporting levels would be raised by 6db relative to the German 3mm flat bottom holes. A second effect of increasing the range is that the size of reflector that causes saturation of the reflected signal w ill also increase. Again at double the range for the 2.25MHz, 25mm trans ducer the saturation size would be a disc of radius 12mm. However, the amplitude would fall very quickly with increasing orientation and no in crease in signal would be observed for a shear wave above 2 relative to the 6mm x 6mm ellipse in figure 2. Decreasing the frequency of the ultrasonic transducer would appear the obvious answer to lessen the dependence on orientation, since by halving the frequency the angle may be doubled for the same signal amplitude. However, the major risk of reducing the frequency is that a crack w ill transmit more and reflect less energy, particularly if it is filled with either a solid or liquid or is a very tight gas-filled crack. For cracks growing directly into the wall of a vessel it is apparent from the above that either the reporting threshold must be decreased considerably or that energy must be received by a second transducer as used by the German Tandem technique (for example, Wustenberg [9]). Decreasing the threshold level by 20db or more would be impractical using an A scan system in which each in dividual reflector above this level would need further investigation. (It is interesting to note that a reflector of only lmm in diameter at normal incidence would be above the threshold.) The number of indications would 354 HAINES probably be so enormous that the time taken to perform an inspection would become uneconomic. However, a В-scan presentationjin which the spacial relationship between the positions of reflectors may be rapidly assessed and hence discrimination made against small isolated indications, would appear to offer a considerable advantage. At present, the interpretation of B-scan data must be made by direct observation. However, it is not beyond the capability of computers to perform this simple type of pattern recognition. The German Tandem system uses a second transducer to receive signals reflected from a defect onto the back wall of the vessel and then back to the scanning Surface. Again the limited range of orientations over which the defect may be detected is still small and Wustenberg [1], therefore, suggests that it is necessary to use a frequency of 1MHz. Complete coverage is then achieved by combining 45 , 60 common transmit and receive systems, together with a 45 tandem system. The additional complication of equipment would appear an expensive price to pay if it were at all possible to use a В-scan presentation system with a reduced threshold level. 5. CONCLUSION The signal amplitude received from reflectors in steel has been examined as a function of the size, shape, orientation and surface roughness. The predicted behaviour has been compared with the presently proposed threshold levels of both the American ASME XI code and German codes for the inspection of nuclear pressure vessels. For a frequency of 2.25MHz and a 25mm transducer it is demonstrated that, using the ASME XI threshold, all smooth defects greater than 12mm in diameter would be missed if they were at an angle of incidence greater than 11 for a compression wave and 6.5 for a shear wave. The corresponding figures for the German codes are 12,5° and 7 . These figures are for a range of 100mm for a compression wave and 200mm for a shear wave. At greater ranges for the same transducer the angles would remain the same for the German code but would decrease for the American code. To dgtect a crack running directly into the wall normal to the surface using a 60 shear wave probe would require a threshold at least 20db below the requirements of either code. Two alternative approaches to this problem have been discussed; the first is the possibility of reducing the threshold by 20db and using a В-scan form of presentation to look for spatial correlation of indications. The second is the German tandem system which attempts to solve the problem by maintaining a higher threshold but using secondary receiving transducers to detect signals reflected from the crack onto the far wall of the vessel and back to the scanning surface. Both methods have their merits but the successful operation of the В-scan system would require much less instrumentation to enter the reactor vessel. The theoretical model used in the paper represents the physical limitations set by the nature of the crack surface. Further work is in hand using a similar 'response function' approach to estimate the signal attenua tions caused by rough or corroded surface of the vessel and the effect of corrosion on crack surfaces. Since all such processes cause a further reduc tion in signal amplitude the model presented here gives the maximum amplitude of signal that may be achieved using a pulse echo ultrasonic system. REFERENCES [1 ] WUSTENBERG, H ., e t . a l . , P ro c. E ig h th World Conference on Non- Destructive Testing, Cannes, 1976, C2] ASME Boiler and Pressure Vessel Code, 1974, Section XI. [3] TRUMPFHELLER, R., Proc. of Conf. on Periodic Inspection of Pressurized Components, 1974, Inst, of Mech. Eng., London, C113/74. [4] HAINES, N. F., and LANGSTON, D. B., C.E.G.B. Report RD/B/N4115, 1977. [5] NEUBAUER, W. G., J. Acoust. Soc. Am., 35, (1963), 279. IAEA-SM-218/41 355 [6] JOHNSON, D. M., J. Acoust, Soc. Am., 59, (1976), 1319. [7 ] WHALEY, H. L., and ADLER, L ., M ater. E v a l. , 29, (Aug. 1971), 182. [8] KRAUTKRAMER, J., and KRAUTKRAMER, H., Ultrasonic Testing of Materials, Springer-Verlag, Berlin, Heidelberg, New York, 1969. [9 ] WUSTENBERG, H. and MUNDRY, E ., P ro c. o f Conf. on P e r io d ic In s p e c tio n of Pressurized Components, 1974. Inst, of Mech. Eng., London, C112/74. ACKNOWLEDGEMENT This paper is published by permission of the Central Electricity Generating Board. APPENDIX 1 Explicit Forms of Response Functions for Surfaces See Figure 1 for Diagramatic Presentation 1. E ll ip s e be r 2 , ctcosec0 v о -.4 F ( t ) = . — L a* ( ------~------a ) 2 У AI ^ A L Tan6 4asin0 0 < t< a = semi-axis of ellipse making an angle 0 with the plane of the incident wave front. b = semi-axis of ellipse parallel to incident wave front, c = velocity of sound. C y lin d e r Lc (R - ^ ) F ( t ) A2 ACAL (R2 - (R - ^ O 2)* 1 2R R = radius of cylinder 0 3. Sphere F ( t ) A3 a c a l 2 2R R » radius of sphere 0 4. Roughness n2 r 2 F ( t ) 2o 727 exP ( - frrr ) A4 where a 1 ® a cos0 (Jg *» rms height of surface above its шеап plane. 356 HAINES DISCUSSION M. de HES: Since your paper is entitled “The Reliability of Ultrasonic Inspection”, I would be interested to know how reliable your ultrasonic technique is. For example, what chance is there of a particular defect not being detected? N. F. HAINES: In our present work we are concerned with the reliability of ultrasonic inspection; however, so far we have only been able to define the capability of ultrasonic techniques when an amplitude threshold criterion is imposed. In the future we shall be looking at variable effects, such as surface coupling and attenuation within components. When this work is done it may be possible to make a probabilistic assessment of a particular ultrasonic technique. M. de HES: Is there not also a chance that the inspection system will give a faulty signal as a result of noise or of an electronic failure, so that a fault may be registered in the material being examined when in fact there is none? N. F. HAINES: I cannot give you a detailed answer on the question of the reliability of electronic equipment, but it is obvious that some form of safeguards by which failure of equipment can be detected must necessarily be part of any inspection system. M. de HES: Have you managed to correlate the reliability of actual ultra sonic tests with your theory? N. F. HAINES: We have compared our theoretical results with those from actual inspection tests in a small number of cases. So far agreement has been good and I hope that a report on this work will be produced in the near future. R.W. NICHOLS: Can your model be extended to include either departures from the simple ultrasonic wave form assumed (e.g. non-parallel or finite diameter waves) or effects such as attenuation and displacement of the reflected waves by metallurgical features? N.F. HAINES: The model already takes the finite beam width from the transducer into account. It does not allow for a non-planar wave fault, although it could be extended to cover spherical or cylindrical wave faults if this was thought necessary. The variable effects of attenuation caused by metallurgical features are the additional physical effects which cause a variability in signal amplitude and hence introduce an element of chance into the reliability of inspection. R. SAGLIO (Chairm an): It was very interesting to hear in your presentation that the choice of transducer (i.e. choice of dimensions and frequency) influences the probability of a fault being detected. Obviously, if this probability is nil, the transducer in question should not be used. In any case, one should use the transducer which gives the highest probability. For this reason the importance of calculating probability is all the greater. Your paper describes an interesting line of research and confirms the value of focused-beam transducers with a beam of small dimensions. S. H. BUSH: Your technique seems to use more signal information than some systems but less than others. For example, it may be closer to acoustic spectro IAEA-SM-218/41 357 scopy than an adaptive learning technique which analyses and discriminates between signals. It would be interesting to see your method applied to a plate such as PVRC-202 on which extensive acoustic spectroscopy results are available so that the relative levels of prediction of the flaw shape, size and orientation could be established. N.F. HAINES: We are not preparing a new technique but are attempting to develop a quantitative understanding of ultrasonic methods already in use. In particular, studies have been carried out on the viability of amplitude threshold criteria and the limitations they impose. ROUND TABLE DISCUSSION C h a i r m a n : R.W. NICHOLS (United Kingdom) P articipants: A.B. LIDIARD (United Kingdom) D.G. DALRYMPLE (Canada) K. MAZANEC (Czechoslovakia) C. B. BUCHALET (France) G. MARCI (Federal Republic of Germany) Y. ANDO (Japan) D. G.H. LATZKO (Netherlands) G. ÖSTBERG (Sweden) T. VARGA (Switzerland) R. RODRIGUEZ SOLANO (Spain) S. H. BUSH (United States of America) SUBJECT AREAS (1) The most profitable areas of work for improving reliability (2) The improvement of reliability by the improvement of design (3) The improvement of reliability by the improvement of materials (4) The improvement of reliability by inspection (5) General aspects R.W. NICHOLS (C hairm an): In this Round Table discussion we shall try to review the main points that have arisen during the Symposium by dealing with them in five different categories. I shall ask two or three of the panelists to speak on each of four subjects and then propose a discussion on general aspects of reliability. I suggest that the first subject we tackle should be the question of how we find out the most profitable areas of work on the improve ment of reliability, for example, whether that should be from sensitivity calculations, from records of failures in service or from operational experience. The second subject I propose is that of how to improve reliability by improve ment of design; the third is that of the improvement of materials and the fourth is that of reliability through inspection. After these four topics have been examined, we could consider some more general aspects of reliability. I would now like to ask Dr. Lidiard to introduce the first subject, that of the most profitable areas of work on the improvement of reliability, since he has shown that by using his calculations one can indicate which are the most sensitive areas where reliability is concerned. 359 360 ROUND TABLE DISCUSSION A.B. LIDIARD: I would like to sum up the usefulness of our probability calculations by saying that they show us the sensitivity of reliability to various factors (material, defects, loading factors, etc.). Obviously, the extent to which cracks occur is important and in this case analysis merely indicates more pre cisely what we know from common sense (e.g. that ultrasonic testing during fabrication improves reliability). A slightly less obvious point is that fatigue crack growth is important. This is particularly true for our calculations because our model is based on the assumption that the systems analysed have passed a pre-service pressure test. Temporal variations in failure probability result from the combination of the effects of the pre-service pressure test and fatigue crack growth, and such variations show the importance of defining the conditions of the test precisely. R.W. NICHOLS (C hairm an): Of course, there is another way of finding out what the important factors are, that of building up a case history from operating experience. Obviously, this method should be seen as being com plementary to, not a replacement of, calculations. May I ask Professor Latzko to comment on this? D.G.H. LATZKO: What I would like to talk about in this connection is data collection and, specifically, that relating to com ponent failures and m al f u n c t i o n s and secondly actual operating loads. In both cases data would have to be collected by the utilities, and here I would like to point out that fewer than 15% of the participants in this Symposium work for utilities, a fact which should give rise to some reflection. The task of collecting data for storage in data banks accessible to the organizations concerned in all IAEA Member States appears well suited to the Agency. I use the term “data bank” delibera tely in order to indicate the need for computerized access and retrieval, a need which is demonstrated by the fact that the Licensing Event Reports (LERs) collected by the United States Nuclear Regulatory Commission at present cover some 1300 abnormal events per year. Data banks on com ponent failures and m alfunctions would serve at least three purposes. The first is that of avoiding generic defects originating in the design stage. Here it might be worthwhile to recall a fact noted by Dr. Crellin in his paper (IAEA-SM-218/4) on valves and valve actuators, namely, that the percentage of systematic defects was found to increase with the increasing complexity of the hardware. The second purpose is illustrated by the concept of “vulnerability” described by Mr. Buchalet (paper IAEA-SM-218/39) and is that of defining the damage that might be caused by a defect. To be useful, this concept requires an extensive data base on failures or defects found in the past. The third purpose is that of enhancing the credibility of — some might say of making credible — the statistical reliability analyses presented at this Symposium, for example, by Dr. Lidiard and Dr. Marriott (papers IAEA-SM-218 /14 and 33). An essential problem which must be solved prior to the proposed ROUND TABLE DISCUSSION 361 build-up of data banks is that of precise definitions for the various failure modes and causes, as stated in the papers by Dr. Bush and Mr. Chockie (papers IAEA-SM-218/11 and 29). The need for operational data on loads is borne out in the list made by Mr. Schulz of the various load cases according to ASME III; he gives a total of about 35 cases, many of which refer to transients potentially damaging to the system. An example of the probability assessment for such transients was shown by the empirical formula on pressure transients in Dr. Vesely’s paper on the OCTAVIA code (IAEA-SM-218/7), which gave a constant for the number of occurrences per year, illustrating the need for statistically useful load data as a prerequisite for the achievement of a high level of confidence in such a code. Similarly, refined analyses of fatigue crack growth, such as those presented at this Symposium by Messrs. Antalovsky, Boissenot and Marci (papers IAEA-SM-218/46, 43 and 15), are hardly worth the effort required for their implementation unless reliable input data on load fluctuations can be provided. S.H. BUSH: I too should like to mention a few areas of work which need to be investigated and to attempt to say what can and what cannot be done. First, as far as passive techniques are concerned, there is a need for improve ments in volumetric techniques, especially ultrasonics, in order to be able to define sensitivity limits, reliability and those factors associated with equipment, materials, and geometric joint design which may affect reliability. In the USA, adaptive learning is one process which, it is hoped, will provide some of the answers, if not all. Secondly, there is the problem of the reliability of active components such as pumps and valves. Important considerations are those of the adequacy of testing, the validity of tests, the detectability of passive failures (and I am thinking of valves here), and the implications for safety that testing has. The safety aspect has not been discussed at this Symposium, and I believe it should be given more attention. In the USA, for example, there is a requirement that tests be carried out on valves and pumps where safety is involved, but utilities have often requested and been granted exemption from these tests because of their effect on the safety of the system. Thirdly, the reliability of components depends very much on statistical data being collected on the same or comparable components so that inferences can be made about the reliability of such components in systems of interest. Analytical analyses are of marginal value and if used alone may lead to incorrect conclusions being made, e.g. about anticipated transients without trip (ATWT). So a balance between statistics and analytical methods is preferable. In this respect I would emphasize the vital need for collection and correlation of relevant data on an international basis. Data relating both to nuclear and to non-nuclear systems would be relevant; in addition, it will be necessary to use common sense to omit irrelevant information. 362 ROUND TABLE DISCUSSION Fourthly, where systems and components are concerned, it should be remembered that the whole is not necessarily equal to the sum of its parts. Component reliability is not a true indicator of system reliability and this should always be borne in mind when evaluating single effects data. Since integral tests are usually quite expensive (some will cost US $10-15 million), it is necessary to achieve a delicate balance in the combination of the two classes of test. W. REDPATH: I would like to ask the panel to what extent mathematical analysts are confident of their ability to estimate the probability of rare events occurring, i.e. events of the order of 10~4 or less. R.W. NICHOLS ( C h a i r m a n ) : Perhaps I could add to that question and ask whether it is still possible to make a sensitivity analysis if one is not confident of estimating such probabilities. A.B. LIDIARD: I think there are two ways in which these probability calculations are obviously useful. First, they give quantitative expression to ideas which may otherwise only be expressed qualitatively. Secondly, they allow us to assess quantitatively the effects of changes in uncertainties in the input information. The calculations should, however, always be assessed in terms of what physical concepts and effects they cover (or do not cover) rather than by the degree of their mathematical or computational sophistication alone. I do not think that the smallness of the calculated failure probabilities by itself presents difficulties for the assessment of the calculations. C. B. BUCHALET: With respect to the question of how to deal with unanti cipated transients I would say that some of the transients defined for analysis (emergency and faulted) are very likely to cover the unanticipated transient. It has been my experience with operating plants that every time an unanticipated transient has occurred, it has been possible to find a design transient that would cover it and therefore the analyses performed with the design transients remained valid. A. JACOBI: The Symposium has been concerned mainly with the influence of design, the selection of materials and procedures, quality control during manu facturing and pre-service and in-service testing of the reliability of the product. My question deals more with situations arising after commissioning of the product. What data must be recorded and stored during the life of the plant to ensure that reliability studies can be done later if required? R.W. NICHOLS (C hairm an): In this connection we should perhaps also be asking what mechanism we should use to set up a world data bank. D. G.H. LATZKO: I have not really given much thought yet to the organiza tional implications. I do not foresee any problems at the utilities level, as most countries with a sizeable nuclear programme either have one national utility (France, Italy, the United Kingdom and the Soviet Union) or a co-operative centre for the individual utilities (such as the Edison Electric Institute (EEI) ROUND TABLE DISCUSSION 363 and the Electric Power Research Institute (EPRI) in the United States or the Vereinigung deutscher Elektrizitätswerke (VDEW) in the Federal Republic of Germany). Not being a politician I have no idea whether or how a body like the IAEA could co-operate directly with the utilities. As far as Western Europe is concerned, an intermediate step might be to work through the International Union of Producers and Distributors of Electrical Energy (UNIPEDE). In the United States all relevant data, at least those concerning abnormal events, are compiled by the Nuclear Regulatory Commission (NRC) which is a federal government body. I am confident that, if the idea is taken up by the IAEA, proper organizational channels will also be found. J.L. FOWLER: Although the idea of an international data collection system is of course a good one, similar ideas have proved difficult to carry out in practice. It is essential to collect data during manufacture, erection and commissioning as well as operation, to categorize the defect correctly in order to be sure that the conclusions reached are meaningful. The Central Electricity Generating Board (CEGB), through its reliability group, as well as the United Kingdom Atomic Energy Authority (UKAEA) and the United States National Aeronautics and Space Administration (NASA) are attempting to collect meaningful reliability data in order to establish effective and useful feedback to improve design, construction and operation. R.W. NICHOLS (C hairm an): A major problem is of course that of providing a suitable classification system. D. G.H. LATZKO: I fully agree that there are practical problems in orga nizing failure data bases. If the costs should seem prohibitively high, however, let me remind you that the cost of the replacement power incurred in the case of an outage of a large nuclear unit, which has been mentioned several times during this Symposium, is between US $250 000 and 500 000 per day. Any data basis leading to a significant reduction in outage rates would therefore be well worth the money. E. E. HADDAD: I would like to mention one parameter which should be included in the reliability data base system discussed by Professor Latzko and Dr. Bush, namely, that of root cause failure in the data base system. D.H. GLICK: As this Symposium is on reliability it is as well that any confusion regarding hydraulic overpressure testing should be forestalled. The Cockenzie boiler drum failure has been mentioned twice. It should be remem bered that, apart from the works tests, the tests which have been referred to as “repeat tests” took place after erection in the boiler and full overpressure was not reached. Also, ambient temperature at the time of the failure was low. There is much that could be said about overpressure testing. I will, however, confine myself to pointing out that we must differentiate between vessels operating above and below the creep range. For vessels designed on the basis of 2/3 yield stress, any stress intensification factor greater than 1.5 will lead to 364 ROUND TABLE DISCUSSION local yielding. This need not worry us because we use ductile steels. On the other hand we do not want the vessel to be pressurized so that it will yield for the first time when under operating conditions. We therefore perform the overpressure test prior to service and follow it with a non-destructive examina tion. Since full pre-service overpressure testing constitutes a considerable fatigue cycle and will propagate defects we should perform only as many tests as are absolutely necessary — preferably only one. R.W. NICHOLS (C hairm an): The previous comment in fact leads us to the second topic I suggested we talk about and I would like to ask: can we improve reliability significantly by improving design assumptions, design methods or detailed design features? T. VARGA: First of all, additional fracture safety checks could be per formed; after the application of ASME Code linear-elastic fracture mechanics (LEFM) methods a second fracture mechanics (FM) evaluation could be used. With this second FM step all effective stresses, actual minimum material proper ties (representing different parts of the vessel at appropriate temperatures) and actual maximum defect dimensions could be dealt with. My second point is that nowadays all construction codes demand structures with no crack-like defects. In all pressure vessels indications of small defects are found and it is not always possible at this time to tell whether they are crack like or not. This means that strict application of the codes would lead to a need for excessive amounts of repair. But excessive repair would result in a greater loss of reliability than would small defects and therefore appropriate codes have to be adopted. This applies also to inclusions whose tolerable length is strictly limited for different wall thickness classes. Thirdly, I would say that at present even in vessels of nuclear quality there are large differences in that quality. For example, vessels which have only been X-ray tested may have considerable cracks; on the other hand, pressure vessel manufacturers with a good ultrasonic testing technique have to repair extremely small defects. And lastly, a prerequisite for better probabilistic evaluation is a reduction in quality scatter. R.W. NICHOLS (Chairm an): One of the most expensive failures experienced in the United Kingdom was that of a storage tank at Fawley which originated with a weld repair, the repair having been carried out at a point where a test sample had been removed. Ironically, the test had proved the original material to be satisfactory. D.G. DALRYMPLE: Like Professor Latzko I believe it worth commenting on the fact that the world’s utilities are underrepresented here, since they play a key role in questions of reliability and maintainability. Furthermore, it should be borne in mind that the aircraft industry is far ahead of the nuclear industry where these two criteria are concerned. However, this state of affairs can no longer be allowed to continue; the world must realize that it cannot do without ROUND TABLE DISCUSSION 365 nuclear power stations and that they cannot continue to be operated at load factors of 55—65%. The question of whether reliability can be improved signi ficantly by improvements in design can be examined in two ways, from the point of view of system design and from that of component design. Improvements in system design would start typically with fault tree ana lysis and availability analysis which could be based on “best guess” data, since for these analyses data accurate to, say, ± 1—2% are not necessary. If the ana lyses are approached in a straightforward fashion, they will yield information which is useful for making decisions about the cost-effectiveness of system design and for reducing redundancy. Further improvements in system design can be brought about with a maintainability audit; a key element of this is what we call a man • rem audit. For the man • rem audit a detailed assessment is made of what maintenance will be required over the lifetime of the plant and what fields of radiation there are likely to be in the area. These extremely extensive calculations give designers information about what changes in the system will be most cost-effective. Preliminary results of the man • rem audit indicate that it is paying off with a threefold reduction in exposure to radiation in connection with maintenance. Where component design is concerned, Canadian specifications for most components require that manufacturers submit reliability and maintainability estimates. These are based mainly on manufacturers’ past experience; eventually they will be supplemented by data from operating experience supplied by utilities. It is too early for feedback from manufacturers to have made an impression on component design leading to significant improvements in reliability but the indications are that it will have this effect. R.L. ROCHE: Fracture mechanics (FM) can be considered one of the most powerful tools for pressure vessel reliability analysis. As a general rule the FM approach is deterministic. Perhaps this is due to the difficulty of assessing either the accuracy (or rather the physical validity) of the FM rules applied, such as the linear-elastic initiation criterion, the propagation or the error proba bility in the stress analysis needed for FM studies. I should therefore like to raise the question whether it is now possible to assess the reliability of the FM rule applied and to estimate the error probability in the stress analysis. T. VARGA: With extensive testing a statistical evaluation of material properties becomes possible. Established confidence limits are a prerequisite for defining structure reliability figures. A full-scale spherical containment vessel consisting of more than 80 heats has been investigated with tensile, impact and compact tension specimens. With appropriate regrouping of the results a statistical evaluation can be made and such an evaluation is now under way. With regard to stress analysis, a check with external and internal strain gauges should be made. However, the possibility of making such checks is limited; for example, the usual hydro test measurements provide no proof with respect to service stresses. 366 ROUND TABLE DISCUSSION G. MARCI: With fracture mechanics fatigue analysis the input parameters are usually selected so that the fatigue crack growth rates calculated represent the worst conditions. Fatigue crack growth rates could in principle be ordered equally well in a statistical manner and the expected crack growth could be associated with probability and confidence values. As far as crack size is con cerned, there are major difficulties involved in determining the crack configura tion from non-destructive testing data and putting probability and confidence values on the detectability of flaws. G. PACKMAN: In order to assess a particular topic and to exert an influence in connection with it, we must first recognize its existence. We need to answer questions like the following: What assurance do we have that all relevant mechanisms of failure have now been recognized? Is it possible that a systematic effort virtually to exclude failure due to factors which are already known may still leave an important number of failures from some other cause or causes? D.G. DALRYMPLE: Perhaps one could consider denting to be an example of a failure mechanism that has only recently been recognized? G.I. SCHUELLER: I would like to ask whether the panel has adopted a probabilistic definition of reliability, or in other words, whether reliability is quantifiable. Would the panel agree with the definition that the reliability of a component or a system is its probability of survival within a given period of time, i.e. design life? R.W. NICHOLS (C hairm an): I think we are using the term reliability both in the sense of the reliability of the component itself and in the sense of the extent to which techniques used to assess the reliability of the component are themselves reliable. A.B. LIDIARD: I think Dr. Schueller has made us aware that at this Symposium we have been using the word reliability in two ways; one is the qualitative and common usage, while the other is the precise mathematical usage, for which we have in mind something definite, such as a model or statistics. N.J. ADAMS: Dr. Varga has given examples of repairs to very small defects and Dr. Lidiard has claimed that proof tests remove certain vessels from a popula tion and thus reduce failure probability in service. However, the ASME proof test conditions, for example, are such that large defects could exist. On the other hand, the fabricators’ claim that defect sizes which exist after ultrasonic examina tion are very small. The severity of the proof test will not be such that the fabricators’ claims can be shown to be correct. The question therefore arises whether the fabricators would be prepared to accept a much more stringent proof test to back up their claims and thus obviate the need to carry out many repairs of small defects. T. VARGA: The hydro test has, among other things, the following purposes to check whether a critical (or bigger) defect has been left undetected and to ROUND TABLE DISCUSSION 3 6 7 reduce high local stresses. For the latter purpose conditions of ductile material behaviour are necessary, so a test above 50°C will be advisable. A “cold” pressure test, such as we have done several times in the past, could follow. But there is another method of indicating defects which may have been overlooked, that of acoustic emission localization, which detects sources where localized separation and/or plastic deformation generate stress waves. The first Austrian reactor had to undergo this test at the request of the licensing department, and it passed. This is perhaps the first time a new BWR pressure vessel had to undergo this type of test at the same time as the cold-hydro test. R.W. NICHOLS (Chairman): Let us now move on to the third subject for discussion and ask whether we can significantly improve reliability by improving our knowledge of materials or by actually improving their properties. Finding answers to this question involves the elaboration of special research and develop ment programmes and I will ask Professor Östberg to speak first since I know he is concerned with this aspect. G. ÖSTBERG: I would like to say a few words about a report prepared for the Swedish National Board for Technical Development a year ago con cerning the need for research on nuclear materials, in order to assist the Board in taking decisions on priorities. The original Swedish version describes the future in terms of scenarios and also deals with various technical problems in some detail. The conclusions arrived at as regards reliability are the same as for materials research in general. There is all too often a lack of correspondence between the effort on the part of the research workers on the one hand and the application of the results by the end-users on the other. I do not think I need to illustrate this by giving specific examples. There are several different types of action that should be taken to improve the situation. Research projects have to be considered in terms of their relevance to the end user, which means that the latter has to take a more active part and take on more responsibility than hitherto. Furthermore, the value of carrying out the research work, that is, of achieving the objectives stated in the project proposals should, as far as possible, be assessed beforehand. The product of the relevance of problems and the efficiency of the research work can be considered an index of its profitability. But more important with respect to the issues discussed at this Symposium is the implementation of the results of research on reliability. A large propor tion of the failures that have occurred in power plants have not been unexpected. Earlier research has in many cases provided the information necessary to avoid them. Obviously there are considerable and serious difficulties involved in making use of research. Often this problem is discussed only in terms of dissemination of information. I think it is a much more complicated matter, which has to do 3 6 8 ROUND TABLE DISCUSSION with the ways in which a complex system like a power station and a large utility operate. This problem is at present under investigation in Sweden. We do not yet know what the results of this project will be but we do know that it is a worth while undertaking because many of the failures which occur are in fact unnecessary: the data needed to prevent them are already available. G. MARCI: I would like to describe our views on reliability and indicate where we place our emphasis in the Federal Republic of Germany. Structural defects are recognized as being the sole initiators of the structural degradation of nuclear components. There is general agreement that the severity of degrada tion increases with increasing defect size. In research and engineering work we place the main emphasis on the prevention of microscopic and macroscopic defects and on the detection of flaws in the size range 1—5 mm. This attitude is exemplified in the following regulatory requirements: (a) In the Federal Republic, every nuclear pressure vessel to be manufactured has to be laid out in such a way that the beltline region does not receive a neutron irradiation dose greater than 1019 ; (b) On all primary pressure-retaining components, three independent ultrasonic inspections are carried out during manufacture; (c) The structural materials used for the primary pressure-retaining boundary have been optimized with respect to stress relief cracking by placing strict limits on certain trace elements; (d) Furthermore, the quality of the welding procedure with respect to the width of the coarse-grain region and stress relief cracking is checked for every unit manufactured by means of tangential cuttings. Since it is realized that the nuclear pressure vessel does not lend itself to a “fail-safe” design, great efforts are made to improve the toughness of the materials, to optimize design and manufacture, to improve the reliability of non-destructive testing and to quantify the safety margin. The goal is to obtain a pressure vessel which is as perfect as possible with regard to present scientific knowledge and manufacturing capabilities. The component safety programme in the Federal Republic was initiated with this goal in view. This is a joint programme in which the Federal Government, the utilities, the component manufacturers and the system suppliers participate. With respect to reliability, the most important question that needs to be asked is what margin safety components have. The answer can only be supplied by non-destructive testing. Where ultrasonic examination is concerned, a signal may be received indicating either that there is a flaw or that there is not. Either type of signal might be wrong. But for purposes of reliability the only important signal is that which indicates no flaw when there is a flaw. This depends on the ROUND TABLE DISCUSSION 3 6 9 flaw size and shape, but also on human factors and equipment. So a large part of our component safety programme involves analysis of the reliability of the nondestructive method used. One operator might give us completely different results for flaw detectability and reliability. Therefore we place emphasis first of all on preventing small defects occurring from the beginning, and subsequently on making sure we find them if they exist. K. MAZANEC: May I say a few words about the relationship between metallurgy characteristics and evaluation conditions? I should like to draw your attention to the problem of effective and representative taking of specimens from the cross-section of thick forges, plates, etc. An important question is that of choosing suitable specimens, that is, a type of specimen with relatively small dimensions, which would be sufficiently representative at least from the point of view of the matrix properties of the whole cross-section and which would produce a value equal to those obtained in large full-section specimens. I shall not discuss here the well-known problems of the effect of struc tural inhomogeneity, the processes of segregation in large sections, etc., but I would like to draw your attention to some questions which may be of interest from the point of view of the effect of metallurgical characteristics on the relia bility of products. In our experience the integral fracture toughness values obtained on large-section specimens are approximately equal to the values determined on small specimens machined out from one quarter of the cross- section (of the thickness of plates, for example). The choice of the specimen location across the thickness reflects both the microstructural effect and the metallurgical parameters involved (e.g. degree of inclusion content, form and distribution of inclusion, temperature at end of rolling or forging, and deforma tion technology). These problems have obviously not been completely solved yet, but where the thickness of plates is 150—250 mm it is possible to achieve a good correlation with the small specimens taken from one quarter of the thickness. In recent research on the same problems we have made an evaluation of the kinetics of fatigue crack growth. Those problems are being studied in various grades of steel (ferritic, ferritic-bainitic and bainitic structures). Finally, I would also like to comment briefly on the vanadium alloy content in steels. The beneficial effect of this alloying (the higher level of cry is held at higher temperatures) is in practice cancelled out by the need for a higher level of production control because the applicable region for annealing or tempering is reduced (for the temperature range 690—720°C). D.G.H. LATZKO: In his presentation of the work by Ewald and Kussmaul (Session Via) Mr. Maurer showed a slide according to which vanadium-alloyed steels are by far the most susceptible to cracking in heat-affected zones. Could Mr. Mazanec comment on this point? K. MAZANEC: Alloying with vanadium is technically feasible although it makes heavy demands on production control. The applicable tempering and 3 7 0 ROUND TABLE DISCUSSION annealing temperature range is about 30-40°C because it is dangerous to enter the region of influence of secondary hardening, which is characterized by lower ductility and toughness, since there is G-precipitation of carbides or carbo- nitrides. The temperature at which this secondary hardening peak occurs depends on the tempered structure (i.e. ferritic-bainitic, bainitic, etc.). W. REDPATH: I have become increasingly concerned, as the Symposium has proceeded, about the question of the acquisition and application of data truly relevant to the mathematical models used in calculating probabilities of failure. In particular, it appears that welds and their associated heat-affected zones are sensitive regions of potential failure, with the geometry of such regions also being of major importance. Would the panel please comment on the way in which experimental data from relevant specimens should be sought? R.W. NICHOLS (Chairman): I note from some of the papers that there is evidence that materials and welding processes can be so chosen that these are not the regions which should give most concern. But to provide this, one needs tests that are very specific with respect to the point of crack initiation. Very useful work with a view to the development of such tests has been done in the United States at Oak Ridge National Laboratory (electron beam brittle line weld), at MPA Stuttgart in Germany, and in the United Kingdom with a round- robin study done for Commission X of the International Institute of Welding. May we now move on to the fourth subject for discussion by asking: can we improve reliability by operational monitoring, surveillance and in-service inspection, or indeed by other aspects of inspection and quality assurance? R. RODRIGUEZ SOLANO: First of all, I would just like to say that I quite agree with Professor Latzko about the need for having a data bank available; this is a matter of great importance for countries with nuclear programmes. In answer to the question on inspection, I would say that this Symposium has shown that there are three conditions which have to be fulfilled if we are to achieve a high degree of safety in nuclear power stations, namely, a good design, appropriate manufacturing processes and pre-established in-service inspection procedures. Since this third condition is not fulfilled, there is a gap in know ledge about reliability. It would be very helpful if this subject were to be dealt with specifically by the International Working Group on Reliability of Reactor Pressure Components. With regard to future power stations, I think it is very important to know what the reliability of all systems and components is in order to be able to plan the frequency of inspection more efficiently for each case. C.B. BUCHALET: The first question one has to ask is what the aims of in- service inspection are. I think these are four, namely, to detect defects present in the component when it is put into service, to detect defects that could originate during service, to check that detected defects which are left in place do not grow during service and to evaluate the flaws detected in order to decide whether or not to repair. ROUND TABLE DISCUSSION 3 7 1 Then we must ask how these aims can be achieved. This can be done in various ways. First, we should select the zones to be inspected and determine how often they should be inspected. This selection of zones should be made having regard to the probability of finding a defect in them (in paper IAEA-SM-218/39 I presented a possible method of making such a selection). Also, the zones should be selected in relation to the risk that could result from a rupture in a given zone. Secondly, the design should be improved so as to reduce the necessity for inspection (self-safe design). Thirdly, the design should be improved in order to facilitate inspection (we have seen that there are many problems of accessibility with CANDU reactors). Fourthly, use should be made of information from actual in-service inspections performed. Fifthly, appropriate surveillance programmes should be devised to monitor the variation in material toughness with time. Sixthly, the material and techniques should be selected in such a way that they permit reasonably good determination of the size and orientation of defects. In addition, when defining the aims mentioned above and ways of achieving them I think we should bear in mind certain factors which are important in real life, such as economics, maintenance times, irradia tion of personnel and performance of the component. D.G. DALRYMPLE: Dr. Buchalet, you mentioned that there was difficulty with in-service inspection in CANDU reactors. You should realize, however, that the CANDU reactors you referred to were designed well before in-service inspec tion requirements were written into the ASME code. One of the features of pressure tube reactors is that the pressure vessel is of manageable size. In my paper (IAEA-SM-218/42) I noted that the major occurrences of lost production in Pickering NGS in 1976 were in each case due to a 2.5 week planned outage for in-service inspection. In any event, in-service inspection of CANDU pressure tubes is well in hand. C.B. BUCHALET: I used the CANDU reactor as an example because the difficulty of gaining access to the components of this very compact type of reactor has already been mentioned. However, the same comment could be made on pressurized water reactors. R.W. NICHOLS (Chairman): We learnt at the Kobe Conference about the changes in pressure vessel design to reduce the number of welds, in particular, the amount of nozzle welding. This is, I think, another example of the way in which design methods are used to reduce the need for inspection. J.S. MACLEOD: As far as inspection procedures for items other than pressure vessels are concerned, it is known that disturbance in some instances can affect reliability. What is the panel’s opinion of the maxim “leave well alone”, which might be the best alternative in some cases? R.W. NICHOLS (Chairman): I regret that time does not permit a detailed discussion of this important question. I personally have considerable sympathy for your point of view, although each case must be considered on its merits. 3 7 2 ROUND TABLE DISCUSSION Having discussed the four subjects proposed, may we now turn to some more general aspects? First I would like to ask Professor Ando to tell us about work in Japan. Y. ANDO: I would like first of all to emphasize one general point — namely that it is the integrated co-operation of people in various fields which increases the reliability of nuclear power stations. By way of an example, I shall describe work being done on the stress corrosion cracking (SCC) problem in BWRs. SCC is associated with a combination of three major factors: material sensitization by welding or heat treatment, environmental effect of coolant water, and tensile stress level. When the 304 stainless steel material of piping is replaced by 304L, for instance, SCC will decrease. However, the allowable stress intensity of 304L is some 20% lower, so in order for it to withstand the internal pressure, the wall thickness of piping should be increased. The load due to seismic vibration is proportional to the mass multiplied by the accelera tion, so from the standpoint of design, thicker material with a low allowable stress is not appropriate. In such a case, a stronger 304L with some modifica tion might be desirable. The ID-cooling welding procedure, or Induction Heating Stress Improvement (IHSI), is applied as a fabrication feature to improve the stress condition of the inside surface of piping before operation. Besides these, the weld buttering method is used to improve the metallurgical condition of welded joints. All these procedures have been applied to actual plants. With regard to operation, the improvement of the water environment by degassing through a condenser vacuum ejector before heating up has recently been achieved. In this way, the saturated dissolved oxygen content of primary coolant water (8 ppm) is much reduced. Research workers have developed an elastic-plastic fracture-elastic mechanics (EPFEM) code to calculate local stress in a short time. In this code three- dimensional elements are used in the main part and the other parts are divided by two- or one-dimensional elements. A systematic, large-scale research programme was started in 1977 in Japan; its budget for the next five years is about US $25 million. This programme is concerned not only with metallurgical tests but also with the study of anti fracture properties. It is by co-operation of this kind between people in the fields I have mentioned that the reliability of nuclear power stations can be improved. Another example is that of the work being done on reactor pressure vessels. Large vacuum-degassed steel ingots weighing up to 500 tons each are now available in Japan. From these, heavy material of high fracture toughness is produced for components such as PWR nozzle belt ring forgings and very wide heavy section steel plates. All the CRD stub tubes are located on a single dished plate. In this case, it is free from the shell weld defects. In this way the length of welded joints is much reduced and ISI can be carried out much more easily. As a result, reliability is much improved. ROUND TABLE DISCUSSION 3 7 3 G. ÖSTBERG: I have a comment on unknown causes of failure. This may by definition be an area which provokes questions which are impossible to answer. Nevertheless it might be worthwhile discussing methods of identifying risks. This problem has recently been considered in the United Kingdom by the Council of Science and Society. A working group of the Council has recom mended the setting up of an independent body with the task of monitoring advanced technologies with respect to the possible risks involved. The Council has issued an interesting report on this subject with the title “Superstar Technologies”. W. REDPATH: Reliability studies originate with the two powerful incentives of improving availability and improving safety. However, during this Symposium little has been said about the reliability targets which should be aimed at for pressure vessels and other pressure-containing components. Recognizing that any figure would imply a measure of judgement of acceptable risk to society in so far as safety is concerned, I wonder if anyone would care to estimate such a target value? A. JACOBI: Target values for avoiding pressure vessel failures with com plete release of radioactivity have been considered by the South African Licensing Branch and, judging by their curve showing probability versus risk, this should not happen more than “once in 10~7” vessel-years. This is, as stated, a target value! J.S. MACLEOD: We dealt with this question in paper IAEA-SM-218/2 when we described the work of the Nuclear Installations Inspectorate. We attempt to relate reliability aims to acceptable levels of radiological exposure, having due regard to frequency. We recognize that quantitative analyses can be an aid in the decision-making process, but for the determination of safety the Inspectorate’s policy is based on the view that quantitative analysis cannot be relied upon alone as a basis for safety acceptance. The approach we have adopted consists in the application of engineering judgements based upon well established principles and the provision of multi-level protection when this is feasible. S.H. BUSH: In th e past the United States Atomic Energy Commission accepted the value of > 1СГ7 for severe accidents leading to major releases of radioactivity. The cumulative probability of all such severe accidents should be no more than 10~6. The possibility of further types of accident is taken into consideration and so this value may not be achieved. The problem still requires evaluation. G. ACKERMANN: Is it possible that in the future methods could be found to monitor the reliability of pressure components of the primary circuit during operation? S.H. BUSH: Neutron noise yields excellent results in the reactor in so far as movements of components such as the thermal shield are concerned. Moni 3 7 4 ROUND TABLE DISCUSSION toring for loose parts, in my estimation, can now be considered an accepted procedure. This is probably also true for leaks. However, I doubt whether we can detect vibration of components without strain gauging or some similar process. With regard to acoustic emission as an on-line tool, I have reservations concerning the possibility of consistent separation of signal from noise. In this context I might mention a programme established by the Energy Research and Development Administration at operating reactors to determine acoustic emission reliability. I hope that the results will prove to be better than they have been so far. G. MARCI: The problem with acoustic emission is that at the moment one cannot characterize the source of acoustic emission from its signal. It is hoped that in the near future it will be possible for different sources of emission to be isolated by analysis of their composite signal. We need a method of analysis that works like the human ear, which can pick out a particular voice even if many sounds are being made or other voices are speaking at the same time. R.W. NICHOLS (Chairman): It is interesting that work somewhat similar to that done by the Central Electricity Generating Board on the theoretical assessment of the maximum efficiency of ultrasonics in a particular case is at present being carried out at Harwell. From the Harwell programme it has been found that not enough information is obtained to make it possible to develop the technique further for the particular application, although this would not apply to other areas such as stress corrosion cracking. A. JACOBI: In the course of the Symposium a lot has been said about how to profit from the present state of the art as regards the improvement of the reliability of reactor pressure vessels, steam generators and the piping of reactor systems. Little has been said about the reliability of valves and virtually nothing about pumps and sealing parts such as gaskets. Although failure or faulty operation of the latter is not likely to cause catastrophic accidents, it is due to the unreliability of these “small” components that the availability of nuclear power stations is reduced drastically. In the next symposium organized by the IAEA on this subject it would be desirable for more emphasis to be put on this fact and for some of the items I mentioned to be included on the agenda. R.W. NICHOLS (Chairman): Perhaps the International Working Group on Reliability of Reactor Pressure Components could take note of this point at its meetings starting on 14 October 1977. Before this Round Table discussion comes to a close, I would like to emphasize that the present Symposium was organized not as a result of serious worries about nuclear plants, since many of the papers and discussions have shown that reliability is very high. Indeed, in the United Kingdom I know of gas-cooled reactors that show higher reliability than conventional power plants. CHAIRMEN OF SESSIONS Session I S. HAVEL Czechoslovakia S.H. BUSH United States of America Session II R. O’NEIL United Kingdom H. SCHULZ Federal Republic of Germany Session III D.G.H. LATZKO The Netherlands Y. ANDO Japan Session IV L.E. STEELE United States of America K. MAZANEC Czechoslovakia Session V T. VARGA Switzerland P.M. PETREQUIN France Session Via R.L. ROCHE France K. BECKER Federal Republic of Germany Session VIb G. ÖSTBERG Sweden D.G. DALRYMPLE Canada Session VII J. FORSTEN Finland G. ÖSTBERG Sweden Session VIII R. SAGLIO France R. RODRIGUEZ SOLANO Spain Round Table R.W. NICHOLS United Kingdom Discussion 3 7 5 SECRETARIAT OF THE SYMPOSIUM Scientific I. K. TERENTIEV Division of Nuclear Power and Secretary: Reactors, IAEA Administrative Edith PILLER Division of External Relations, Secretary: IAEA Editor: R. PENISTON-BIRD Division of Publications, IAEA Records H. ORMEROD Division of Languages, IAEA Officer: 3 7 6 LIST OF PARTICIPANTS AUSTRALIA McDonald, N.R. Australian Atomic Energy Commission, Mattiellistrasse 2/4, A-1040 Vienna AUSTRIA Fritz, K. Österreichische Studiengesellschaft für Atomenergie Ges.m.b.H., Lenaugasse 10, A-1082 Vienna Halada, W. Bundesministerium für Handel, Gewerbe und Industrie, Sektion V (Energie und Grundstoffe), Schwarzenbergplatz 1, A-1015 Vienna Hirsch, H. Bundesministerium für Handel, Gewerbe und Industrie, Sektion V (Energie und Grundstoffe), Schwarzenbergplatz 1, A-1015 Vienna Holzer, H. Österreichische Studiengesellschaft für Atom energie Ges.m.b.H., Institut für Reaktorsicherheit, Lenaugasse 10, A-1082 Vienna Matulla, H. Österreichische Studiengesellschaft für Atom energie Ges.m.b.H., Institut für Reaktorsicherheit, Lenaugasse 10, A-1082 Vienna Obermair, G.E. Bundesministerium für Handel, Gewerbe und Industrie, Sektion V (Energie und Grundstoffe), Schwarzenbergplatz 1, A-1015 Vienna Struszkiewicz, R. VOEST-ALPINE AG, Lindengasse 48—50, A-1070 Vienna Wischin, K.G. Bundesministerium für Bauten und Technik, Stubenring 1,A-1011 Vienna BELGIUM Rottenberg, G. CEN/SCK, Boeretang 200, B-2400 Mol 3 7 7 3 7 8 LIST OF PARTICIPANTS CANADA Atchison, R.J. Atomic Energy Control Board, 270 Albert St., Ottawa Burns, D.J. Department of Mechanical Engineering, University of Waterloo, Ontario Dalrymple, D.G. Chalk River Nuclear Laboratories, Chalk River, KOJ 110, Ontario Hodge, R.I. Chalk River Nuclear Laboratories, Chalk River, KOJ 1 JO, Ontario CHILE Thoma Bollito, E. Comisiön Chilena de Energia Nuclear, Los Jesuitas 645, Casilla 188-D, Santiago CZECHOSLOVAKIA Antalovsky, S. Faculty of Nuclear Sciences and Physical Engineering, Section of Materials, Kremencova 10, 110 00 Prague 1 Becvar, J. Czechoslovak AEC, Slezska 9, 12000 Prague Brumovsky, M. Skoda Works, Nuclear Power Plants Division, Research and Development Centre, 316 00 Plzen Dugat, J. Czechoslovak AEC, VUJE Jaslovske Bohunice Havel, S. Nuclear Research Institute, Rez Mazanec, K. Institute of Metals Science, Technical University, tr. Vitezneho unora, 70833 Ostrava-Poruba Vacek, M. Nuclear Research Institute, Rez LIST OF PARTICIPANTS 3 7 9 DENMARK Hagen, H.H. Danish Inspectorate of Nuclear Installations, PO Box 217, DK-4000 Roskilde Kongs0, H.E. Ris$ National Laboratory, Department of Reactor Technology, DK-4000 Roskilde Lykke, E.R. Elsam Kraftverksgruppen, Skaerbaek, DK-7000 Fredericia Nielsen, A. Risф National Laboratory, DK-4000 Roskilde FINLAND Forsten, J. Technical Research Centre of Finland, Metallimiehenkuja 6, SF-02150 Espoo 15 Mankamo, T. Technical Research Centre of Finland, Electrical Engineering Laboratory, SAH/LTR, Otakaari 5, SF-02150 Espoo 15 Rahka, K.A. Technical Research Centre of Finland, Reactor Materials Research, Metallimiehenkuja 6, SF-02150 Espoo 15 Ranta-Maunus, A.K. Institute of Radiation Protection, Department of Reactor Safety, PO Box 268, SF-00101 Helsinki 10 FRANCE Baron, J.L. Service technique des constructions et armes navales, 8 bd Victor, F-75015 Paris Barrachin, B. CEA, Centre d’etudes nucleaires de Saclay, B.P. 2, F-91190 Gif-sur-Yvette Bernard, J.L. Framatome, Tour Fiat, F-92084 Paris la Defense Cedex 16 Boissenot, J.M. CETIM, 52 av. Felix Louat, F-60304 Senlis Bruna, J.G. CEA, Centre d’etudes nucleaires de Cadarache, B.P. 1,F-13115 Saint-Paul-lez-Durance 3 8 0 LIST OF PARTICIPANTS FRANCE (cont.) Buchalet, C.B. Framatome, Tour Fiat, F-92084 Paris la Defense Cedex 16 Capel, R. EDF, Centrale nucleaire du Bugey, B.P. 14, Loyettes, F-01800 Meximieux Cohu, P.J. Electricite de France, 3 rue de Messine, F-75 008 France Debray, J.M. Novatome Industries, 90 av. Edouard Herriot, F-92 350 Le-Plessis Robinson Duchemin, B.H.M.L. CEA, Centre d’etudes nucleates de Saclay, В .P. 2, F-91190 Gif-sur-Yvette Jubault, G.F.M. Novatome Industries, 90 av. Edouard Herriot, F-92 350 Le-Plessis Robinson Lebey, J. CEA, Centre d’etudes nucleates de Saclay, B.P. 2, F-91190 Gif-sur-Yvette Martin, G.A. Framatome, Tour Fiat, F-92084 Paris la Defense Cedex 16 Mencarelli, E.V. Service technique des constructions et armes navales, 8 bd Victor, F-75 015 Paris Morgand, P. CEA, Centre d’etudes nucleaires de Grenoble, 85 X, F-38041 Grenoble Cedex Orio, R. Dresser Industries, 527 av. du General de Gaulle, F-92140 Clamart Peffau, L. CEA, Centre d’etudes nucleaires de Saclay, B.P. 2, F-91190 Gif-sur-Yvette Petrequin, P. CEA, Centre d’etudes nucleaires de Saclay, B.P. 2, F-91190 Gif-sur-Yvette Prot, A.C. CEA, Centre d’etudes nucleaires de Saclay, B.P. 2, F-91190 Gif-sur-Yvette Roche, R.L. CEA, Centre d’etudes nucleaires de Saclay, B.P. 2, F-91190 Gif-sur-Yvette LIST OF PARTICIPANTS 3 8 1 Saglio, R. CEA, Centre d’etudes nucleates de Saclay, B.P. 2, F-91190 Gif-sur-Yvette Signoret, J.-P. CEA, Centre d’etudes nucleaires de Saclay, B.P. 2, F-91190 Gif-sur-Yvette Soulat, P. CEA, Centre d’etudes nucleaires de Saclay, B.P. 2, F-91190 Gif-sur-Yvette Toume, J.R. APPAVE, 13-17 av. Salneuve, F-75017 Paris Valibus, L.F. EDF, Centre des Renardieres, Departement EMA, F-77250 Moret-sur-Loing Verdeau, J.J.N. CEA/Technicatome, B.P. 18, F-91190 Gif-sur-Yvette GERMAN DEMOCRATIC REPUBLIC Ackermann, G. Ingenieurhochschule Zittau, Strasse der jungen Pioniere 2, DDR-88 Zittau Kittner, M. VEB Kombinat, Kraftwerksanlagenbau, Hans Beimler Strasse 91—94, DDR-1017 Berlin Rabold, H. Staatliches Amt für Atomsicherheit und Strahlenschutz, Waldowallee 117, DDR-1157 Berlin Schimmel, W. Staatliches Amt für Atomsicherheit und Strahlenschutz, Waldowallee 117, DDR-1157 Berlin Winkler, R. VEB Kernkraftwerk “Bruno Leuschner”, DDR-22 Greifswald GERMANY, FEDERAL REPUBLIC OF Bazant, E. Babcock Brown Boveri Reaktor GmbH, PO Box 410323, D-6800 Mannheim 31 Becker, K.H. - DIN Normenausschuss Kerntechnik, Unter den Eichen 87, D-1000 Berlin 45 Belda, W. Kraftwerk Union AG, Hammerbacher Str. 14, D-8520 Erlangen 3 8 2 LIST OF PARTICIPANTS GERMANY, FEDERAL REPUBLIC OF (cont.) Bosten, F. Technischer Überwachungs-Verein, Baden e.V., Dudenstrasse 28, D-6800 Mannheim-Wohlgelegen Brockmann, G. TÜV Rheinland, Am Grauen Stein, D-5000 Cologne 31 Döhier, J. Kernforschungsanlage Jülich, Postfach 365, D-517 Jülich 1 Dressier, E. Kraftwerk Union AG, Hammerbacherstrasse 14, D-8520 Erlangen Fickel, O.F. Babcock Brown Boveri Reaktor GmbH, PO Box 410323, D-6800 Mannheim 31 Fritz, R. Hochtemperatur-Reaktorbau GmbH, Gottlieb-Daimler-Strasse 8, D-6800 Mannheim 1 Kautz, H.R. Babcock Brown Boveri Reaktor GmbH, PO Box 410323, D-6800 Mannheim 31 Marci, G. Babcock Brown Boveri Reaktor GmbH, PO Box 410323, D-6800 Mannheim 31 Peter, K. Rheinisch-Westfälischer Technischer Überwachungs-Verein e.V., Feithstrasse 188, D-5800 Hagen Philippi, G. Technischer Überwachungs-Verein Baden e.V., Dudenstrasse 28, D-6800 Mannheim-Wohlgelegen Schallopp, B. Ringstrasse 71, D-1000 Berlin 45 Schatz, A.К. Babcock Brown Boveri Reaktor GmbH, Heppenheimerstrasse 27-29, D-6800 Mannheim 42 Schueller, G.I. Arcisstrasse 21, D-8 Munich 2 Schulz, H. Gesellschaft für Reaktorsicherheit mbH, Glockengasse 2, D-5000 Cologne 1 Wellein, R. Kraftwerk Union AG, Hammerbacherstrasse 14, D-8520 Erlangen HUNGARY Raszl, K. Hungarian Atomic Energy Commission. Postafiök 14, Budapest 5 LIST OF PARTICIPANTS 3 8 3 INDIA Kakodkar, A. Bhabha Atomic Research Centre, Reactor Engineering Division, Trombay, Bombay 400 085 ISRAEL Alter, J. Israel Atomic Energy Commission, PO Box 17120, Tel Aviv Markowitch, M.C. Nuclear Research Centre Negev, PO Box 9001, Beer Sheva Ramati, P. Israel Atomic Energy Commission, PO Box 17120, Tel Aviv ITALY Fizzotti, C. CNEN - RIT, CSN Casaccia, C.P. 2400, S. Maria di Galeria, Rome Mezzanotte, R. Comitato Nazionale per l’Energia Nucleare, Viale Regina Margherita 125, 1-00149 Rome Perria, N. FIAT TTG, Termomeccanica Nucleare e Turbogas S.p.A. Divisione Nucleare, Via Cuneo 20, Turin Scabbia, G. NUCLITAL, Via Federico Avio 2,1-16100 Genoa Tomassetti, G. CNEN — RAD-RSI/Serv. Ingegneria Impianto, C.S.N. della Casaccia, C.P. 2400, S. Maria di Galeria, Rome Turricchia, A. ENEL — Ente Nazionale per l’Energia Elettrica, Via Cardano 10, Milan JAPAN Ando, Y. Department of Nuclear Engineering, University of Tokyo, 7-3-1, Hongo, Bunkyo-ku, Tokyo 3 8 4 LIST OF PARTICIPANTS Shinohara, Y. Japan Atomic Energy Research Institute, 1-1-13 Shinbashi, Minato-ku, Tokyo 105 Tomono, K. The Tokyo Electric Power Co., Inc., No. 1—3, 1-Chome Uchisaiwai-Cho, Chiyoda-ku, Tokyo KOREA, REPUBLIC OF Min, K.S. Korea Electric Company, 5, 2-KA, Namdaemoon-Ro, Chung-ku, Seoul NETHERLANDS Brasz, J. TNO 50 MW, Sodium Component Test Facility, Petroleumhavenstraat 16, 7553 GS/Hengelo de Boer, G.A. B.V. Neratoom, Laan van N.O. Indie 129—135, The Hague de Hes, M. B.V. Neratoom, Laan van N.O. Indie 129—135, The Hague Latzko, D.G.H. Dept, of Mechanical Engineering, Geist University of Technology, Rotterdamseweg 139A Prij, J. Netherlands Energy Research Foundation (ECN), Westerduinweg 3, Petten Smit, C. Central Technical Institute, Petroleumhavenstraat 16, Hengelo NORWAY Bokalrud, T. Det Norske Veritas, Veritasveien 1, 1322 H$vik Föyen, L. Det Norske Veritas, Veritasveien 1, 1322 H PAKISTAN Bhutta, S.M. Chasma Nuclear Power Project, PO Box 1133, Islamabad LIST OF PARTICIPANTS 3 8 5 POLAND Aleksandrowicz, J. Ministry of Power and Atomic Energy, United Industries of Machines and Equipment for. Power Engineering, Pulawska Str. 18, Warsaw Kowalski, W.R. U.D.T. Technical Supervision Authority, Szczesliwicka 34, 02-353 Warsaw Zmystowski, A.S. Ministry of Power and Atomic Energy, ul. Mysia 2, 00-955 Warsaw SOUTH AFRICA Hudson, J.M. A.E.B., Licensing Branch, Private Bag X256, Pretoria Marriott, D.L. A.E.B., Licensing Branch, Private Bag X256, Pretoria Siebert, J.F. Electricity Supply Commission, PO Box 1091, Johannesburg SPAIN Alvarez de Buergo, L. Junta de Energia Nuclear, Ciudad Universitaria, Madrid Rodriguez Solano, R. Junta de Energfa Nuclear, Ciudad Universitaria, Madrid Santoma Juncadella, L. Junta de Energia Nuclear, Ciudad Universitaria, Madrid SWEDEN Östberg, G. Lund Institute of Technology, Box 725, S-220 07 Lund 7 Östensson, L.B. AB Atomenergi, Fack, 61101 Nyköping Tuxen-Meyer, H. AB Atomenergi, Section for Dynamics and Control, Fack, 61101 Nyköping 3 8 6 LIST OF PARTICIPANTS SWITZERLAND Brandstätter, K.W.L. BBC Aktiengesellschaft, Brown Boveri & Cie., Abt. TGD-4, CH-5401 Baden Dobrosavljevic, N. Electrowatt, Engineering Services Ltd., Bellerivestrasse 36,CH-8022 Zürich Jacobi, A. Electrowatt, Engineering Services Ltd., Bellerivestrasse 36, CH-8022 Zürich Njo, D.H. Swiss Nuclear Safety Division (ASK), CH-5303 Würenlingen Prantl, G.A.F. Federal Institute for Reactor Research (EIR). CH-5303 Würenlingen Varga, T. Sulzer Brothers Ltd., CH-8401 Winterthur UNITED KINGDOM Adams, N. Nuclear Installations Inspectorate, Thames House North, Millbank, London SW1 Butler, J.C. Rolls Royce & Associates Ltd., PO Box 31, DE2 8BJ, Derby Cowan, A. United Kingdom Atomic Energy Authority, Risley Engineering & Materials Laboratory, Risley, Warrington, WA3 6AT Cox, R.F. United Kingdom Atomic Energy Authority, SRD, Wigshaw Lane, Culcheth, Warrington, Cheshire WA3 4NE Darlaston, B.J.L. Central Electricity Generating Board, Berkeley Nuclear Laboratories, Berkeley, Gloucestershire Fowler, J.L. Central Electricity Generating Board, Walden House, 24 Cathedral Place, London EC4P 4EB Glick, D.H. Central Electricity Generating Board, Courtenay House, 18 Warwick Lane, London EC 4P 4EB LIST OF PARTICIPANTS 3 8 7 Haines, N.F. Central Electricity Generating Board, Berkeley Nuclear Laboratories, Berkeley, Gloucestershire Lidiard, A.B. AERE Harwell, Theoretical Physics Division, Harwell, Oxon 0X11 ORA Macleod, J.S. Nuclear Installations Inspectorate, Health and Safety Executive, Thames House North, Millbank, London SW1 Naish, S.T. Systems Reliability Service, United Kingdom Atomic Energy Authority, Wigshaw Lane, Culcheth, Warrington, Cheshire WA3 4NE Nichols, R.W. United Kingdom Atomic Energy Authority, Risley Engineering & Materials Laboratory, Risley, Warrington, Cheshire WA3 6AT O’Neil, R. United Kingdom Atomic Energy Authority, Safety & Reliability Directorate, Wigshaw Lane, Culcheth, Warrington, Cheshire WA3 4NE Packman, G. South of Scotland Electricity Board, Cathcart House, Inverlair Avenue, Glasgow G44 4BE Redpath, W. United Kingdom Atomic Energy Authority, Atomic Energy Establishment, Dorchester, Dorset DT2 8DH Smith, T.A. United Kingdom Atomic Energy Authority, Wigshaw Lane, Culcheth, Warrington, Cheshire WA3 4NE UNITED STATES OF AMERICA Bush, S.H. US Nuclear Regulatory Commission, Advisory Committee on Reactor Safeguards, 1717 H Street N.W., Washington, DC 20555 Cordovi, M.A. Inco United States, Inc., 1 New York Plaza, New York, NY 10004 Crellin, G.L. General Electric Co. Fast Breeder Reactor Dept., 310 DeGuigne Avenue, PO Box 5020, Sunnyvale, California 94086 3 8 8 LIST OF PARTICIPANTS UNITED STATES OF AMERICA (cont.) Haddad, E.E. Boston Edison Co., Boston, Massachusetts Loss, F.J. Naval Research Laboratory, Attn: Code 6390, Washington, DC 20375 Lowe, A.L., Jr. Babcock & Wilcox Company, Nuclear Power Generation Division, PO Box 1260, Lynchburg, Virginia 24505 Okrent, D. 5523 Boelter Hall, University of California at Los Angeles, California 90024 Perrin, J.S. Columbus Laboratories, Battelle Memorial Institute, 505 King Ave., Columbus, Ohio 43201 Shaffer, D.H. Westinghouse R & D Center, 1310 Beulah Rd., Pittsburgh, Pennsylvania 15235 Steele, L.E. Naval Research Laboratory, Attn: Code 6390, 4555 Overlook Av., Washington, DC 20375 van Rooyen, D. Brookhaven National Laboratory, Upton, New York 11973 Vesely, W.E. US Nuclear Regulatory Commission, Probabilistic Analysis Branch, Office of Nuclear Regulatory Research, US Nuclear Regulatory Commission, Washington, DC YUGOSLAVIA Susnik, J. Josef Stefan Institut, Jamova 39, PO Box 199/1V, 61001 Ljubljana ZAMBIA Mumba, N.K. Radioisotopes Research Unit, National Council for Scientific Research, PO Box CH 158, Chelston, Lusaka LIST OF PARTICIPANTS 3 8 9 ORGANIZATIONS COMMISSION OF THE EUROPEAN COMMUNITIES (CEC) Haas, R. Buero В 4 003, GD XIII/B, Jean Monnet Gebäude, Kirchberg - Luxemburg Maurer, H.A. 200, rue de la Loi, B-1049 Brussels INTERNATIONAL ATOMIC ENERGY AGENCY (IAEA) Raisic, N. Department of Technical Operations, Division of Nuclear Power & Reactors, Kärntnerring 11, A-1010 Vienna INTERNATIONAL ORGANIZATION FOR STANDARDIZATION (ISO) Becker, K. (also under FRG) Normenausschuss Kerntechnik, Unter den Eichen 87, D-1000 Berlin 45 (West) ORGANIZATION FOR ECONOMIC CO-OPERATION AND DEVELOPMENT (OECD) Oliver, P. Nuclear Safety Division, 38 bd. Suchet, F-75016 Paris AUTHOR INDEX Roman numerals are volume numbers. Italic numerals refer to the first page of a paper by the author concerned. Upright numerals denote comments and questions in discussions. Literature references are not indexed. Ackermann, G.: II 373 Darlaston, B.J.L.: I 193,345; Adams, N.J.: I 4 9 ; II 366 II 6 5 ,8 1 Alvarez de Buergo, L.: I 46,258 de Boer, G.A.: I 46; II 199, 216 Anderson, D.R.: I 21 deH es, M.: II 1 9 7 ,199, 356 Ando, Y.: II 359, 372 Destribats, M.T.: II 311 Antalovsky, S.: I 89, 103 Forsten, J.: II 17, 1 1 4 ,2 6 7 Asty, M : II 293, 311 Fowler, J.L.: II 2 4 9 , 258, 363 Barrachin, В.: I 4 3 9 Gamier, С.: I 4 3 9 Bazant, E.: I 4 29, 436; II 3 Glick, G.H.: I 122; II 363 Becker, К.: II 133, 142 Goldberg, F.F.: I 1 0 5 Bernard, J.L.: I 2 5 1 , 287 Haddad, E.E.: I 63, 87; II 363 Bhutta, S.M.: I 8 7 ,2 3 1 ,2 4 9 ; Haines, N.F.: II 113,329,547, II 92, 257, 2 59, 267, 282 3 5 6 ,3 5 7 Boissenot, J.M.: II 33, 49, 50, 51 Havel, S.: I 2 8 9 Brandstätter, K.W.L.: II 50 Hawthorne, J.R.: I 555 Brasz, J.: I 161 Herrmann, C.R.: I 55 Brumovsky, M.: I 1 2 2 ,1 8 1 , 1 9 3 ,1 9 4 Hilborn, J.W.: II 1 7 7 195, 317, 329, 393, 4 1 5 ; II 173 Hodge, R.I.: II 177, 197,292,329 Buchalet, C.B.: I 122,231,248,314, Houssin, В.: I 251 436; II 49, 2 8 5 , 292, 359, 362, Hudson, J.M.: II 95 370, 371 Hyspeckä, L.: I 2 8 9 Burns, D.J.: I 248,459; II 50 Ingram, G.E.: I 1 3 9 Bush, S.H.: I 43, 45,46, 199, 231, Jacobi, I.: I 13 7 ,3 9 4 ; II 362, 232, 287; II 52, 115, 119, 132, 3 7 3 ,3 7 4 356,359,361,373 Jacobs, I.M.: I 55 Chockie, L.J.: II 1 1 9 Kakodkar, A.: II 83, 92,93,267, Class, I.: I 4 2 9 2 6 9 ,2 8 2 Crellin, G.L.: I 33, 43, 44, 45, Karim, J.A.: I 3 7 9 4 6 , 1 3 9 Kautz, H.R.: I 4 2 9 ; II 5 Dalrymple, D.G.: I 21, 3 2 ,4 4 ; Kittner, M.: II 2 9 2 ,3 2 9 II 359, 364, 366, 371 Кирка, P.: I 3 1 7 3 9 1 3 9 2 AUTHOR INDEX Kussmaul, К.: II 174 Petrequin, P.: I 344, 3 47, 393, 394 Latzko, D.G.H.: I 45,231,301; Pigeon, M.: II 2 93, 311 II 51, 173, 359, 360, Prantl, G.: II 5 J, 6 1 ,6 2 362, 363, 369 Prot, A.C.: 11 2 9 3 , 3 2 9 Lebey, J.: I 4 39, 459 Rahka, K.A.: II 17, 52 LeSurf, J.E.: II 1 7 7 Raiäic, N.: II 2 37, 257 Lidiard, A.B.: I 123, 2 3 3 , 248, Rao, V.S.G.: II 2 6 9 249, 315; II 61, 114, 173, Redpath, W.: I 46,249; II 81, 362, 359, 360, 362, 366 370, 373 Loss, F.J.: I 393,43 6 ,4 6 1 Roche, R.: I 3 1 4 ,3 4 5 ,4 3 9 ; Lowe, A.L., Jr.: I 3 97, 436; II 5 0 ,3 6 5 II 174 Rodriguez Solano, R.: II 359,370 Ludwig, P.W.P.H.: II 1 9 9 Roule, M.: II 311 Lynn, E.K.: I 1 0 5 Saglio, R.: II 2 93, 3 11, 329, 356 Macleod, J.S.: I 49, 60, 103,344; Samoel, A.: II 311 II 93,371,373 Santoma Juncadella, L.: I 122 Mankamo, T.: I 1 2 5 Schimmel, W.: II 49 Marci, G.: I 102, 315, 345, 429, Schueller, G.I.: II 49,61,339,366 436; II 3, 49, 51, 359, 366, Schulz, H.: I 4 3 ,2 3 2 ,2 8 7 368, 374 Shaffer, D.H.: II 331 Markowitch, M.C.: I 394 Slama, G.: I 251 Marriott, D.L.: II 95, 113, 114 Smit, C.C.: I 161 Martin, G.: II 2 8 5 Smith, A.M.: I 3 3 Maurer, H.A.: II 174 Smith, T.A.: II 132 Mazanec, К.: I 2 89, 303, 345; Soulat, P.: I 3 4 7 II 359, 369 Steele, L.E.: I 194, 333, 344, 345, Michel, D.J.: I 3 3 3 393, 394, 429, 436, 437, 4 6 1 ; Mrkous, P.: I 3 1 7 II 145, 173, 174 Naish, S.T.: I 43; II 216 SuSnik, J.: I 149, 159 Nämec, J,: I 89 Touffait, A.M.: II 311 Nichols, R.W.: II 142, 356, 359, 360, Tvrdy, M.: I 2 8 9 362, 363, 364, 366, 367, 370, Vacek, M.: I 379, 393, 394 3 7 1 ,3 7 4 van Rooyen, D.: I 329; II 216, 2 1 7 Nielsen, A.: II 174 Varga, T.: I 59, 194, 195, 231, Njo, D.H.: II 5 3 303, 459; II 53, 113, 359, O’Neil, R.: I 3, 32, 44, 81, 103, 159 364, 365, 366 Östberg, G.: II 255,282,359, Vauterin, M.: II 2 8 5 3 6 7 ,3 7 3 Vesely, W.E.: I 105, 122, 123 Östensson, В.: I JO J, 314, 315 Whit marsh, C.L.: I 3 9 7 Packman, G.: I 44; II 1 3 2 ,3 6 6 Whooley, J.P.: I 6 3 Palme, H.S.: I 3 9 7 Williams, Mair: I 2 3 3 Perrin, J.S.: II 161, 173, 174 Zurlippe, C. F.: I 3 9 7 The following conversion table is provided for the convenience of readers and to encourage the use of SI units. FACTORS FOR CONVERTING UNITS TO SI SYSTEM EQUIVALENTS* SI base units are the metre (m), kilogram (kg), second (s), ampere (A), kelvin (K), candela (cd) and mole (mol). [For further information, see International Standards ISO 1000 (1973), and ISO 31/0 (1974) and its several parts] Multiply by to obtain Mass pound mass (avoirdupois) 1 Ibm = 4.536 X 1 0 '1 kg ounce mass (avoirdupois) 1 ozm = 2.835 X 101 g ton (long) (= 2240 Ibm) 1 ton = 1.016 X 103 kg ton (short) (= 2000 Ibm) 1 short ton = 9.072 X 102 kg tonne (= metric ton) 1 1 = 1.00 X 103 kg Length statute mile 1 mile = 1.609 X 10° km yard 1 yd = 9.144 X 10_1 m foot i ft = 3.048 X 10_1 m inch 1 in = 2.54 X 10“2 m mil (= 10-3 in) 1 mil = 2.54 X 10“2 mm Area hectare 1 ha = 1.00 X 104 m2 (statute mile)2 1 mile2 2.590 X 10° km2 acre 1 acre = 4.047 X 103 m2 yard2 1 yd2 - 8.361 X 10_1 m2 foot2 1 ft2 = 9.290 X 10~2 m2 inch2 1 in2 = 6.452 X 102 mm2 Volume yard3 1 yd3 = 7.646 X 1 0 '1 m3 foot3 1 ft 3 = 2.832 X 10-2 m3 inch3 1 in3 = 1.639 X 104 mm gallon (Brit, or Imp.) 1 gal (Brit) = 4.546 X 10-3 m3 gallon (US liquid) 1 gal (US) = 3.785 X 10~3 m3 litre 1 l - 1.00 X 10-3 m3 Force dyne 1 dyn = 1.00 X 10 '5 N kilogram force 1 kgf 9.807 X 10° N poundal 1 pdl = 1.383 X 10“ ‘ N pound force (avoirdupois) 1 Ibf = 4.448 X 10° N ounce force (avoirdupois) 1 ozf = 2.780 X 10 '1 N Power British thermal unit/second 1 Btu/s = 1.054 X 103 w calorie/second 1 cal/s = 4.184 X 10° w foot-pound force/second 1 ftlb f/s = 1.356 X 10° w horsepower (electric) 1 hp = 7.46 X 102 w horsepower (metric) (= ps) 1 ps 7.355 X 102 w horsepower (550 ft-lbf/s) 1 hp = 7.457 X 102 w Factors are given exactly or to a maximum of 4 significant figures Multiply by fo obtain Density pound mass/inch3 1 lbm/in3 = 2.768 X 104 kg/m 3 pound mass/foot3 1 lbm/ft3 = 1.602 X 101 kg/m 3 Energy British thermal unit 1 Btu 1.054 X 103 J calorie 1 cal 4.184 X 10° j electron-volt 1 eV 1.602 X 10“ 19 j erg 1 erg = 1.00 X 1 0 '7 J foot-pound force 1 ft- lb f 1.356 X 10° J kilowatt-hour 1 kW-h 3.60 X 106 J Pressure newtons/metre2 1 N /m 2 1.00 Pa atmosphere* 1 atm 1.013 X 105 Pa bar 1 bar 1.00 X 10s Pa centimetres of mercury (0°C) 1 cmHg = 1.333 X 103 Pa dyne/centimetre2 1 dyn/cm2 = 1.00 X 10_1 Pa feet of water (4°C) 1 ftH aO = 2.989 X 103 Pa inches o f mercury (0°C) 1 inHg 3.386 X 103 Pa inches of water (4°C) 1 inH 20 2.491 X 102 Pa kilogram force/centimetre2 1 kgf/cm2 = 9.807 X 104 Pa pound force/foot2 1 I b f/ft2 = 4.788 X 101 Pa pound force/inch2 (= p$i)b 1 lbf/in2 = 6.895 X 103 Pa torr (0°C) (= mmHg) 1 torr 1.333 X 102 Pa Velocity, acceleration inch/second 1 in/s = 2.54 X 101 mm/s foot/second {= fps) 1 ft/s 3.048 X 10~‘ m/s foot/minute 1 ft/m in = 5.08 X 10“3 m/s [4.470 X 1 0 '1 m/s m ile/hour (= mph) 1 mile/h [ l .609 X 10° km/h knot 1 knot 1.852 X 10° km /h free fall, standard (= g) = 9.807 X 10° m/s2 foot/second2 1 ft/s 2 3.048 X 10~‘ m/s2 Temperature, thermal conductivity, energy/area- time Fahrenheit, degrees — 32 °F — 321 5 f c Rankine °R I 9 1 к 1 B tu -in /ft2-s- °F 5.189 X 102 W /m -К 1 Btu/ft-s- °F = 6.226 X 10‘ W/m К 1 cal/cm-s-°C = 4.184 X 102 W/m К 1 B tu/ft2 -s 1.135 X 104 W /m2 1 cal/cm2*min = 6.973 X 102 W /m 2 Miscellaneous fo o t3 /second 1 f t 3/s 2.832 X 10"2 m3/s foot3/minute 1 ft3 /min - 4.719 X 10-4 m3/s rad rad = 1.00 X 1 0 '2 J/kg roentgen R 2.580 X 10-4 C/kg curie Ci 3.70 X 10‘° disintegration/s atm abs: atmospheres,absolute; 6 lb f/in 2 (g) (= psig): gauge pressure; atm (g): atmospheres gauge. lb f/in 2 abs (= psia): absolute pressure. 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