Methods for the evaluation of crack resistance of reactor P.V. materials under conditions of thermal shock

Int. J. Pres. Ves. & Piping 52 (1992) 159-175

Methods for The Evaluation of Crack Resistance of Reactor P . V . Materials under Conditions of Thermal Shock B. T. Timofeev, V. A. Petrov, G. P. Karzov & V. V. Anikovsky Central Research Institute of Structural Materials 'Prometey' St Petersburg, Russia 193167 (Received 15 October 1991; accepted 23 November 1991)

ABSTRACT This paper presents the results of thermal-shock tests on large-scale specimens produced from steels and their welded joints used for pressure vessels of the types VVER-440 and VVER-IO00; seven large-scale specimens with mechanical notches of various depths and shapes were tested. The neutron-flux effect on the materials studied was simulated by a special heat treatment, which caused metal embrittlement of the same degree as that resulting from the effect of operating conditions during the whole period of service life. Test results on large-scale specimens obtained by thermo-force loading (mechanical load + thermal shock) are compared with those on standard specimens, performed with general methods by using material of the same steel melts and the same welding wire and flux. Satisfactory correlation in calculations and experiments were achieved. By tests on large-scale welded specimens, two-stage fracture modes were found. In conclusion, there is a discussion of the possibility of using these results in the evaluation of the crack resistance of reactor cases in accident situations.

INTRODUCTION During the assessment of nuclear-power e q u i p m e n t , the determination of allowable loading conditions for pressure vessels to prevent brittle-fracture risk is a necessary step. The reactor vessels of powerful 159

160

B. T. Timofeev et al.

power plants with water coolant are large-sized welded high-pressure vessels operating under conditions that could include pressurized thermal shocks. They must work not only in the ductile but also in the quasi-brittle and brittle temperature ranges of loading. The danger of brittle failure of such structures is associated with their large wall thickness together with the reduction in brittle failure resistance of vessel metal after operating time under neutron flux. In accordance with the existence in U S S R of certain standards or norms ~ the calculation of brittle-failure resistance in all operating ranges is based on crack-initiation criteria under static loading (Fig. 1). A more conservative approach to the evaluation of all the allowable structure states is available; the operation of the structures is allowed in a temperature and stress range when the metal can arrest or retard any

12x2 MFA, 15x2 MFA,15x 2 MFAAI 3 200 -base metal for VVER-440 160 [ K t ] I = 17"5 + 2 2 " 5 . e °'re
120

[ K I ] 2 = 23"5 + 3 0 . e °'°ztr-r~)

80

[KI]3 = 35 + 45 . e °'°2(r-tO

40 [

~ 0_

I

[

I

I

200 160

[Kx]l = 37 + 5 " 5 . e 3"85 lo 2(r-Tk) [KI]2 = 50 + 5"1 . e 4 " 1 ' 10-2(T--Tk)

120

[KI]3 = 74 + 11 . e 3"~5 ,o-Z~r r~)

4O

I I J [ I Weldments reactor steel for /1 2OO -VVER-440 and VVER-IOOO/ _circumferential welds / 2 160 [ K i ] l = 17.5 + 2 6 . 5 . e °°217(r rk) [ K I ] 2 = 25 + 2 7 . e 2"35" 10-2(T Tk) [ K I ] 3 = 35 + 53 . e O'0217(T Tk)

120 80 4O 0 -160

I

-1OO

I -40

I

O

I

40

I

80

[T - Td,°C

A l l o w a b l e values of stress-intensity factors for calculation of the brittlefracture resistance of reactor pressure vessels in the f o l l o w i n g operating regimes: 1 - - n o r m a l operating conditions ( n s a f = 2.0; A T = 3 0 ° C ) , 2--hydrostatic-pressure test (nsaf = 1"5; A T = 30 ° C ) , 3 - - a c c i d e n t situation (nsa f = 1 " 0 ; A T = 0 ° C ) . F i g . 1.

Evaluation of crack resistance of reactor P.V. materials

161

randomly initiated and rapidly propagating brittle crack. In this case, the conditions of fracture occurrence in the range of local stresses are not analyzed, conditions in terms of nominal stresses for retardation of any random crack are considered. If non-steady thermal stresses of the vessel are present, a more detailed analysis of crack-retardation condition can be performed in terms of stress-intensity factors as the American standards or norms provide. 2 This approach is reasonable if the temperature gradient throughout the wall thickness is sufficiently high, i.e. when there is a great inhomogeneity of the stress field across the part section. These conditions for a reactor vessel can occur when it operates in emergency when cold coolant is supplied for rapid heat removal. In this case, the non-steady thermal fields and high thermal stresses greatly exceeding the allowable nominal stresses (as defined by the standards ~) occur in the various parts of the heavy vessel i.e. in base metal and the weld of the cylindrical case (core) and nozzle area. In this case, non-uniform temperature stresses are occurring in the vessel wall, and these stresses may considerably exceed the level of stresses allowable by standards. L2 In other words, the operation of this emergency cooling system results in thermal shock followed by longterm cooling of vessel walls under the action of the continuously supplied cold coolant. This situation has been considered many times by Russian 3"4 and foreign ~-7 scientists in terms of design and experiment, including pressure-vessel tests. 8,9 Under the conditions of various cooling cases the m a x i m u m tensile stresses occur in the cooler areas, which are related to the most dangerous situations from the brittle fracture viewpoint. The application of design methods developed for isothermal loading for such a case does not give adequate results. Bearing in mind that the evaluation of crack-retardation conditions and of the jump-like propagation of crack as considered in standards, are not available, we aimed to investigate reactor-material behaviour under non-steady pressurized thermal shock and under the simultaneous influence of a large-sized sample tension. At the same time, j u m p crack propagation occurs mainly in welded joints (weld and H A Z ) and embrittled irradiated materials. The detailed condition of the cooling of reactor vessels, greatly depends on the structure of units supplying cold coolant so that the conditions for mixing cold and hot water were simulated in the test by reproduced high-temperature gradients and high mechanical stresses. For the vessel of VVER-440, the emergency can occur in two cases: (i) the vessel is cooled coaxially when the cold coolant is uniformly supplied to the perimeter of the cylindrical part by non-uniform temperature distribution and by stresses across the wall thickness only:

162

B. T. T i m o f e e v et al.

(ii)

the occurrence of cold tongues of boride solution on the vessel wall and an assymetric stress field across the thickness and perimeter of the shell.

THE MATERIALS AND EXPERIMENTAL PROCEDURE The heat-resistant 15 x 2 M 0 A and 15 x 2 HM~pA steels used in Russia to produce vessels of VVER-440 and VVER-1000 reactors and their welded joints, performed by automatic submerged welding by wire of the Cb-10 x M~pT type and Cb-10 x M~pTY type and flux of AH-42M type and wire of Cb-08 x F H M T A type under flux of HO - 18M type,

4

i I

"

/1 T

I.I

IPt

692

1

932

No

Dimensions(mm) B

~

a

c

1

400

127

15

60

2

420

130

14

420

3

450

140

14

450

4

380

162

26

380

5

410

147

15

410

6

470

150

15

470

7

480

120

12

480

Fig. 2.

Full-scale sketch of the specimen.

Evaluation o f crack resistance o f reactor P.V. materials

163

TABLE 1 Typical Chemical Analyses of the Materials Studied Test sample 1 2 3 4

Material

15x2M~A Cb-10xM~T Cb-10xM~pTY 15 ×2Mq~A with cladding 15 x 2HMq~A Cb-08x FHMTA 15 × 2HM~pAA

5 6 7

C

Si

Mn

Cr

Ni

Mo

V

Cu

S

P

0-17 0-06 0-06 0-16

0.31 0.32 0.48 0.22

0-35 1.15 1.08 0.38

2-25 0.87 1.35 2.50

0-14 0.16 0.21 0.16

0-64 0-34 0.48 0-72

0.29 0-15 0-23 0.30

0.16 0.15 0-04 0.14

0.014 0.014 0-010 0-017

0.016 0.036 0.020 0.011

0 - 1 7 0 . 2 0 0 . 4 7 2 . 1 5 1 . 2 7 0 . 5 7 0-11 0 . 0 5 0.007 0-009 0.05 0.30 0 - 6 5 1 . 5 2 1 . 3 2 0-63 0 - 2 0 0 . 0 3 0.010 0-008 0.18

0 . 2 8 0.49

1 . 8 9 1 . 0 8 0 - 4 4 0 - 1 3 0 . 0 8 0-005 0.013

were investigated. Figure 2 shows a sketch of the sample. The geometric parameters of the samples studied 1° are presented in the table given with this figure. The chemical compositions of the materials studied are shown in Table 1. Mechanical properties of the materials studied are presented in Table 2. The embrittlement of the material, to give an adequate simulation of TABLE 2 Mechanical Properties of the Materials Studied Test sample

Heat-treatment procedure

Tensile properties at 20 °C UTS

YS

A

(MPa) (Mea) (%) 1 2 3 4 5 6 7

Quenching 1000 °C, 3 h, water temp. 600 °C, 5 h, air Temp. 350 °C, 5 h, air Quenching 1000 °C, 3 h, water temp. 600 °C, 10 h, air Quenching 1000 °C, 3 h, water temp. 600 °C, 10 h, air Quenching 935 °C, 5 h, water temp. 550 °C, 10 h, air Quenching 980 °C, 6 h, water temp. 540 °C, 10 h, air Quenching 955 °C, 5 h, water temp. 550 °C, 10 h, air

Tcr (°C)

Z

(%)

1 140

1 050

17

62

140

820 980

720 770

14 16

55 60

90 130

1 050

990

18

65

140

1 100

950

18

67

100

995

820

17

53

110

960

840

20

60

65

164

B. T. Timofeev et al.

any irradiation effect at the termination of vessel service life, together with the associated increase of strength and decrease of plastic properties and increase of brittle critical temperature (see Table 2) were simulated by heat treatments. The test procedure includes heating of the sample up to 500 °C, and mechanical loading followed by rapid cooling of the notch area 200 m m in width by a cold-water spray at 20 °C. The thermometry of cooling was carried out by means of thermocouples installed at different depths. This allows one to plot temperature distribution at various time intervals after the start of the cooling. The mechanical loading by a ZZ-8000 machine at the test time of each sample was held constant or was increased in a definite manner in the 20-40-s interval after the start of cooling. This allows one to increase the range of the most dangerous combinations of stress patterns in the test. The basic characteristics measured were: temperature, stress, and crack opening. The fracture stresses were evaluated by determining temperature and stress changes. The plot of crack opening allows one to evaluate the individual stages of the fracture process. The start of pre-critical crack growth was determined. T H E R E S U L T S OF I N V E S T I G A T I O N S A N D T H E I R ANALYSIS The tests performed permit one to carry out the analysis of the main factors determining crack initiation and crack retardation in relation to the safety margin no = K~c/K~ in emergency conditions with regard to the co-operative effect of thermal and mechanical loads. The main factors are: • crack depth and shape a/t; • the character of mechanical load change and its cycling; • preliminary thermal loading; • material-structure heterogeneity; • notch-cooling method. At the same time, it should be noticed that the type 15 x 2MFA and 15 x 2NMFA steels gain a high level of toughness after quenching and high degrees of tempering, and it is impossible to induce the unstable fracture of specimens with cracks of various depths by cooling them in water with temperature head of T = 2 7 0 ° C . " In this study, the materials were therefore tested in embrittled condition after quenching and low degrees of tempering, providing the mechanical properties shown in Table 2, which is similar to that resulting from the effect of neutron fluence on the reactor-vessel wall at the end of its service life.

Evaluation of crack resistance of reactor P. V. materials

165

TABLE 3 Test Results on Large-scale Specimens

Sample number

Loading conditions

Test results T (°C)

Specimen heating up to 300 °C, cooling in water stream (T = 20 °C), displacement is set; • specimen heating up to 300 °C, cooling in water stream (T = 20 °C), loading set from 487 MPa to 772 MPa (a notch is thermally isolated); • specimen heating up to 300 °C, cooling in water stream (T = 20 °C), loading set from 487 MPa to 654 MPa (a notch is opened). 2

Specimen heating up to 300 °C, cooling in water stream (T = 20 °C), the loading is constant at 251 and 323 MPa, at load increase before 395 MPa jump took place (330 MPa, a notch is opened): • the first crack jump • isothermal loading to fracture without thermal shock. Specimen heating up to 300 °C, cooling in water stream (T = 20 °C), the loading is equal to o,. = 260 MPa (a notch is thermally isolated), thermal shock with the load increase from 260.5 to 618.6 MPa; • specimen heating up to 300 °C, cooling in water stream (T = 20 °C), the loading is equal to 260 and 358 MPa (a notch is opened, the thermal shock with the load increasing from 358 to 626 MPa). Specimen heating up to 300 °C, cooling in water stream (T = 20 °C) without load; • specimen heating up to 300 °C, cooling in water stream (T = 20 °C), preliminary loading is equal to o,. = 255 MPa, thermal shock with increasing load up to o,. = 485 MPa during 210 s.

o,,, (MPa) K,c (MPaVm)

55

654

140

100 295

330 440

87 148

45

626

235

165

485

110

166

B, T. Timofeev et al. TABLE 3----(Continued)

Sample number

Loading conditions

Test results T (°C)

Specimen heating up to 300 °C, cooling in water stream (T = 20 °C), a set of constant loadings crm= 260, 390, 482, 585,683 MPa (a notch is thermally isolated); • specimen heating up to 300 °C, cooling in water stream (T = 20 °C), the loading is equal to o,1 = 780.5 MPa (a notch is opened). Specimen heating up to 300 °C, cooling in water stream (T = 20 °C), displacement is set; • cooling in water stream by crm--278 MPa • cooling in water stream by Om----278 MPa, the loading is further increased up to Crm= 667 MPa (a notch is isolated); • cooling in water stream by or. = 278 MPa, the loading is further increased up to Om= 667 MPa (a notch is opened). Specimen heating up to 300 °C, cooling in water stream (T = 20 °C), displacement is set (a notch is thermally isolated); • specimen heating up to 300 °C, cooling in water stream (T = 20 °C), a set of constant loadings Om204, 306, 442, 679"5 MPa (a notch is opened).

o,,, (MPa) Kic (MPaVmm)

104

795

265

60

481

240

100

695

220

T e s t results o n l a r g e - s c a l e s p e c i m e n s a r e g i v e n in T a b l e 3. F i g u r e 3 shows the load versus c r a ck- opening displacement for the seven l a r g e - s c a l e s p e c i m e n s t e s t e d . It f o l l o w s t h a t , w i t h d e c r e a s e o f d e f e c t size t h e f r a c t u r e stress i n c r e a s e s ; w i t h t e m p e r a t u r e i n c r e a s e t h e c r a c k tip o p e n i n g v a l u e s at f r a c t u r e i n c r e a s e a n d a r e a c c o m p a n i e d b y h i g h e r p r e - c r i t i c a l c r a c k g r o w t h . A c o m b i n a t i o n o f h i g h t h e r m a l s t r e s s e s is a c c o m p a n i e d b y a m e c h a n i c a l - l o a d i n c r e a s e a n d l o w t e m p e r a t u r e in t h e tip o f a c r a c k . A n o t h e r a s p e c t s h o w n s c h e m a t i c a l l y in Fig. 3 is t h e occurrence of displacement jumps (pop-in) during the fracture process

Evaluation of crack resistance of reactor P.V. materials

167

4OO0

3000

v

13. 2000

1000

0

0'5

1"0

1'4

A (mm)

Fig. 3.

Load versus crack-opening displacement for the seven large-scale specimens

tested.

in welded joints (specimens 2 and 6). A combination of high temperature at the crack tip resulted in fracture. These results of experiments on large-scale specimens were compared with those of calculations relating to crack-initiation and retardation conditions in terms of the stress-intensity factor, thus permitting one to use them to analyze the various stages of the fracture process (Fig. 4). Calculation of the stress-intensity factor for large-scale specimens consists in non-stationary temperature-field analysis and calculation of the stress-strain condition in the elastic-plastic range. The calculation of the non-stationary temperature fields in such specimens has been performed within the temporal range of specimen cooling by using temperature values measured at different depths in the specimen-thickness direction. Temperature-field calculations on a specimen were made by solution of the inhomogeneous equation of thermal

B. T. Timofeevet al.

168

240

0

15X2 MFA

I ~., :0.1 vw I

~K1 (test 5) "~ I I X~/ / K1( t ~

I

p=~,y 160 --

• eTest 5 A oTest7 15X2 H M F A

o @ Test I • Test4

oI

o~ I •

/

o

/

I

o

c&

° o// o .c~v~==+ 4~ / / K,(test 4 / " ~1. . . . . . ;// ~ t c h ) ~ - - - ~ - ~

.....

BO tl n

~/o

v

I

"

24

8

u

I

(~ ¢ (test 3)1~/

-

1

K1

/~

~

I

I

& •Test2 o @Test 3 CblOxMFT

-

Test

6

P'W

-

//

CbO8x FHMTA

K1c ~ ~ •

o ]Kla

P: O.y/

~11

~

8C

P--95% -160

I

I

-160

I

I

I

I

1

-BO

0

80

160 -160

P:95% I

-80

I

0

1

80

T-Tk (*C)

Fig. 4. Test results of large-sized specimens, together with tests on notched/cracked specimens (marked by arrows). conductivity in a two-dimensional range by FEM with mixed boundary conditions of the first type (on a cooled surface) and the second type (on thermally isolated surface): OT ca.c =

a T / 02T (

027""]

)t--~xZ+ ay2],

where a(T) is the material temperature conductivity. The stress-strain condition associated with thermal and force monotonic loadings of a specimen is determined in a two-dimensional range by solving the problem by FEM thermoelastic plasticity in accordance with leakage theory by the use of a technique involving variable elasticity parameters. A most interesting result was obtained on specimen 4, which had anti-corrosive cladding. Figures 5 and 6 show the fracture mode and temperature distribution in the specimen section, respectively. Stress distributions in this specimen at various moments of time are illustrated in Fig. 7. It should be mentioned that, earlier several experiments had been performed ~2 that estimated the effect of thermocyclic loading (without applying mechanical force) on crack

Evaluation of crack resistance of reactor P.V. materials

__1

35o~

/,

t

i

[,

II~

~

IIIIglh

a

~oW

I

~13C r

Fus,on ,,°e I

Fig. 5.

3c

;,--~,ipof ~tch

'If, 0

169

]

20

I

40

60

80 100 120 1 4 0 t (min)

160

Temperature distribution in depth of specimen 4 at various moments of time.

propagation through anti-corrosive cladding up to the fusion line with pearlitic steel. After 500 cycles within the regime at 300~--20 °C in 15 x 2MFA steel plate (400 × 400 × 149 mm), owing to the difference in pearlite and austenite coefficients of expansion, a crack growth of 1.5-2.0-mm depth along the whole crack length of 250 mm was found. The stress-intensity factor at the crack tip was determined for each Brittle fracture (fine)

Cladding • B r i t t l e fracture ( rough )

t

I

I

I

1

I

-'6_.___*c

..I- Heat treatment CraCk,

162 I~

-I Vl

Mechanical notch

Fig. 6.

Sketch of the fracture surface of specimen no. 4.

I

3

5

7

1~

t(sm)

stage ( 7 3 - A c e x ) Xlg stage (113-A) XTv stage (153-A) ~rv stage (Ig3-A) 3tln stage (233-A) 3tV11 stage (270-A)

3{1 stage ( 6 0 - A }

3ZI st age ( 3-A cex ) x~r stage ( 8 - A ) stage (13-A) 13r stage (33-A)

g

11

if)

i

13

I

[

15

I

~

~3Z

i

3Z~

x'r

Fig. 7. Stress distribution in the large-scale specimen with anti-corrosive cladding at various moments of time.

1~

i

[..notch _I

° LI

200Ii

300~X

400

:E 6 0 0

O'

100

20C

30(

40(

500~

=._

¢

Evaluation of crack resistance of reactor P.V. materials

171

time point in accordance with the methods given in 125-01-90.13 In this case, the calculation was carried out both for the mechanical notch (a = 1 5 m m ) and for the crack (a = 2 6 m m ) . Fractograms were developed that showed the crack generated by heat treatment of the specimen with a mechanical notch. The analysis of large-scale specimens was carried out with regard to isothermal tests on standard specimens, manufactured from pieces of the fractured large-scale specimens. In these standard specimens, the crack depth was equal to 0.5W. The specimens with a crack depth of a = 0.1 made from type 15 × 2MFA and 15 × 2 H M F A steels were also tested by three points bend. Figure 4 shows the fracture-toughness curves, constructed with a probability parameter equal to 50 and 95% obtained by using the available test results. Compact specimens of 50-mm thickness with lateral notches were tested using wedge-loading following the A S T M technique to estimate crack-retardation conditions) 4 The material of these compact specimens was in the same condition as that of the large-scale specimens. Some of the test results of such specimens manufactured from welded joints have also been added to Fig. 4. However for the base metals, the K = f(T - Tcr) curves coincide with the low (95%) boundary of fracture-toughness values obtained under crack-initiation conditions. These are therefore not shown in Fig. 4. For weld metal, these curves are moved in the direction of increase in temperature by 10-40°C (Fig. 4). As mentioned above, a spasmodic crack growth was observed in welded joints (Fig. 3). This process may be satisfactorily explained by the results obtained on Kia.15 The relation for K~ = f(T - T,), given in Figs. 3 and 4 make it clear that there is a difference in the fracture processes of the base metal and the weld metal. In the base metal of 15Z2MFA and 15 × 2 H M F A steels, the dynamic effects are absent, whereas in welded joints they are evident. In this case, the crack is growing in accordance with the scheme loading initiation and then j u m p retardation. The differences observed in the fracture process are connected with the effect of the cast structure and metal-heterogeneity and are characterized by the relation of the plastic zone to the c o m p o n e n t size of the structure. In particular the influence of this relation was observed by comparable fracture-toughness tests on specimens made from base metal and weld metal and having a short crack. 16 It should be pointed out that a crack initiation occurred in the area of the K~ = f ( T - Tr) relation that did not decrease (for all tests). In this connection, the calculated and experimental estimation of warm pressurized stress (WPS) on crack growth by thermoshock would be an important addition to the analysis described. We have performed such

172

B.T. Timofeev et al.

investigations. The experimental results thus obtained showed that, in this study, this effect due to the available concepts 17 manifested itself only weakly and it is therefore not discussed in this paper. The analysis of correlation of the K~, KIC, Kla (Fig. 4) shows that, for base-metal specimens, it is not feasible to describe the conditions of over-all unstable fracture by K~ = K~c crack-initiation conditions. In this case the elastic energy of the system provided specimen -ZZ-8000 machine (for brittle material with UTS of 1000 MPa at 20 °C) appears to be insufficient to propagate a brittle crack through the whole specimen section. In the welded specimen however, uneven two-sided crack propagation is observed. The analysis relates to fracture conditions in the quasi-brittle range of large-scale specimens (a/t=O.1) by the co-operative action of thermo-shock and mechanical load. For the type 15 x 2MFA and 15 × 2HMFA steel base metals in brittle conditions, such a change in AT, obtained experimentally on specimens with a/w =0.1, is equal to 50°C. Thus, in the range of realization of the quasi-brittle fracture mechanism, it becomes necessary to apply approaches by non-linear fracture mechanics (including the two-criteria approach) to estimate the ultimate conditions of the reactor vessel with a crack. It is interesting to compare the results simulating the fracture conditions for the vessel by thermo-shock (obtained in this investigation) with PWR brittle-fracture calculation and safety evaluation. For the shell of PWR core and nozzle zones under conditions of large and small leakage by the emergency-cooling-system operation, the crackresistance estimation is presented in Fig. 8, which has the co-ordinates of K versus (T - TK). The calculation of K~ = f(T - TK) was carried out for the depths of the inner-axial-surface crack with 0.1 <-a/t in the elastic range. Here T is the temperature at the crack tip. By calculation of the stress-intensity factor, it was assumed that KI = K~p) + K~t) + K~*) where K~p), K~t), and K~*) are the stress-intensity factors from the inner pressure p, the temperature gradient in thickness, and the azimuthal direction. In conformity with the condition of the shell material in the core zone to the end of its service life, it is necessary to consider the dependence of the degree of neutron embrittlement, characterized by the shift of the critical brittle temperature ATr in relation to the crack depth, a. The latter is induced by the neutron-fluence decrease (F) through the vessel-wall thickness from the inner to the outer surface [at the outer surface, it was decreased 10 times)]. TM Moreover, by the case-crackresistance analysis, it is important to take into consideration the

173

Evaluation of crack resistance of reactor P.V. materials

200-180 160

TK° : - 3 0 / _ 2 0 1

~ 140

~

~

20~.~ a:0.1 0 T:O'I

120 ~ 100 8 ~ 8O 6O

40 20

m

o

-50

I 0

I 5o T(oc)

I ~00

I 150

Fig. 8. Diagram showing an example of calculated brittle-fracture resistance of a reactor vessel for 15 x 2MFA steel with surface semi-eliptical crack (a/c = 1/3).

upper-shelf decrease of the K~c = f ( T - TK) curve, which depends on neutron fluence. The range of allowable values of pressure variation, p ( r ) , by the emergency- system operation is determined by the condition of KI = K~c(T - T~:) initiation.

CONCLUSIONS (1)

(2)

(3)

The analysis showed that, in the conditions of non-stationary temperature-force changes, fracture initiation was possible only in the case of the intersection of the temperature relation K~c and the change relation K~ by the loading process only in increasing area. For the base metal, there is an additional safety margin in the condition of fracture initiation for a short crack, a/t -- 0.1. It was found that, by the applied mechanical loads equal to those permissible [o], the thermal-shock attack with thermal head (pressure) of AT = Tm~x- Tmi. did not result in fracture initiation. The typical character of reactor-vessel fracture of welded joints is two-stage fracture. The crack-retardation condition is determined by the KJKIa relation.

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B. T. Timofeev et al. REFERENCES

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