Fracture mechanics investigations on cylindrical large scale specimens under thermal shock loading

Nuclear Engineering and Design 144 (1993) 31-44 North-Holland

31

Fracture mechanics investigations on cylindrical large scale specimens under thermal shock loading L. S t u m p f r o c k a, E. Roos b, H. H u b e r a a n d U. W e b e r a

a Staatliche Materialpriifungsanstalt (MPA), University of Stuttgart, Pfaffenwaldring 32, D-70569 Stuttgart, Germany b Energie-I/ersorgung Schwaben, Stuttgart, Germany Received 30 September 1991, revised version 12 March 1993

To investigate the crack growth and crack arrest behaviour of primary circuit materials large scale experiments were conducted on component-like specimens under pressurized thermal shock loading at MPA Stuttgart. The material characteristics varied from high tough material to low tough material with higher nil ductility transition temperature to simulate EOL or beyond EOL-state. All tests started from in-service conditions and were cooled down to room temperature. The specimens showed both stable and unstable crack growth and partly crack arrest. The crack growth behaviour was verified by post test calculations and could be explained with the help of the multiaxiality of the stress state.

1. Introduction The pressurized components of the primary loop have to resist all loads during service, emergency and faulted conditions even in the case of postulated flaws of limited size which may occur in case of accidents. Appropriate material characteristics are required for the fracture mechanics assessment. They are specified as lower bound curves, e.g. in the ASME Boiler and Pressure Vessel Code, Sects. III and XI. Prevention against brittle fracture has been stated in the code for a postulated flaw of depth a ffi 0.25 times the wall thickness t and length 2c = 1.5 t = 6a for all loads occurring in accidents, emergency and faulted conditions. The appropriate fracture mechanics characteristics can be found in the code from the curves related to the reference nil ductility transition temperature RTNDr [1,2,3]. The curves indicated correspond to the reference fracture toughness curve in [1] representing the lower bound curve of critical static and dynamic K I initiation-values (KIc , Kld) as well as critical crack arrest values (Kia). It was derived from experimental values which were determined on steels SA-533 Grade B, Class 1, SA-508-1, SA-508-2 and SA-508-3 [4]. It has not been considered that the material properties are dependent on direction with respect to the crack orien-

tation. Furthermore, no difference between base material, welding material and heat affected zone could be found in the case of welded joints, i.e. the above mentioned fracture toughness curves apply to those material conditions as well. In the case of cracks detected by non-destructive examination it has to be decided whether the service may be continued with regard to the different load levels. For this purpose, the fracture toughness KIc as initiation value and the crack arrest toughness Kia are used. They are given with reference to RT~DT see, e.g. [2] and [5]. The K1c-CUrve is the lower bound of the K1c-values based on the same materials as the KIRcurve. The Kla-distribution is identical with the KIRcurve except for the upper bound, which amounts to 187 M P a v ~ , Fig. 1, in the case of the KiR-curve. Emergency cooling processes are safety relevant loading cases, for the material is loaded in the toughness spectrum from the upper shelf to the transitional regime. Due to thermal shock events high thermal stresses occur in the wall areas adjacent to the inner surface and great thermal stress gradients arise over the wall thickness. Especially zones with existing cracks are critical in case of pressurized thermal shock. These secondary thermal stresses may considerably increase the crack tip load resulting from primary mechanical

0 0 2 9 - 5 4 9 3 / 9 3 / $ 0 6 . 0 0 © 1993 - Elsevier Science P u b l i s h e r s B.V. All rights reserved

32

L. Sttanpfrock et al. 20o E

"~

1"50

~

lao-

! ,

•

i .I

• •

/ Fracture mechanics investigations 200

i

0 NKS3 TN~ 600C

]

• NKS5 TNDT= ~°C

-iscc~lerbandKS22.B ~

•

S3NIWoTN~= l -50°C

CRACKlIP IEMPERATURE Iso'¢ 175] 2go'c ATCRACKINITIATION ~ HgSl ANDSTABLECRACK6ROWTH / -f~$3 CRACKTtPTEMPERAIURE / 16~i'c II.O) ~2N'c ATCRACKINITIATION / 1.._ ~;;'.'.'.'.;C;%%; I~s2 ARKCRACKEXTENSION / / /

I

................................ 0

-3oo-2oo

-Ioo

o

T - TNDT /

,oo

200

=¢:

~oo_.~.(__As~ ix)__. . . . .

~

/

13O"~ u ]

/

/

~lJ

._1

325'C

.a'a~._.~.) =n.c

so

K

Fig. 1. Fracture toughness of the materials. 1111/I

TEMPERATURE

stresses caused by internal pressure. The fracture toughness of the ferritic material may decrease simultaneously due to cooling. Extensive thermal shock investigations were carried out on the decommissioned hot steam reactor (HDR) pressure vessel in Germany [6]. A variation of material parameters is not possible in a wide spectrum on this real component, especially with regard to the material toughness. Therefore, such parameters were studied within the scope of the research projects "RDBNotkiihlsimulation (NKS)" [7] and "KS 22" [8] at MPA Stuttgart. The specimens investigated were thick-walled hollow cylinders ( D i = 400 mm) with an original wall thickness (s -- 200 mm). There was a variation in both the material conditions and the crack configurations. The material conditions in terms of notch impact energy vs. temperature investigated within the scope of NKS are shown in Fig. 2. Circumferential and partial circumferential surface cracks were tested.

2. Materials Methods of determining fracture mechanics material characteristics in the linear-elastic and elastic-plas-

Fig. 2. Notch impact energy spectrum of the NKs-program.

tic application have been delineated and discussed in detail in ref. [9]. Specimens NKS3, NKS5 and NKS6 dealt with in this paper were manufactured from various heats of the material 22 NiMoCr 37 [9] and a speciaily heat-treated material based on 17 MoV 84 (KS22). An external ring made of the high-tough material S 3 NiMo I was shape-welded for specimens NKS5 and NKS6. The material characteristics are shown in Table 1, the notch impact energy-temperature curves in Fig. 3. The JR-curves of the materials for various temperatures are compiled in Figs. 4 and 5. The initiation values Ji determined from these JRcurves were converted into Kjc-values and plotted in Fig. 1. This representation shows dearly that none of the Kjc-Values determined for the upper transition regime or in the upper shelf are above the limit value of 220 MPavr~ stated in the Kic-CUrve. Consequently, the area of the upper shelf of the Kzc-CUrve provides too high initiation values. Regarding fracture mechanics analyses in the upper shelf regime, the values from the curves indicated in the code are not applicable.

Table 1 Material characteristics Specimen

NKS 3 NKS 5 NKS 6 Weld mat.

Material

22 NiMoCr 3 7 (KS 05 mod.) 22 NiMoCr 3 7 (KS 07 mod.) 17 MoV 8 4 mod. (KS 22) S 3 NiMo 1

Test temperature (°C)

R e/Rpo,2

Rm

A5

Z

(MPa)

(MPa)

(%)

(%)

Young's module (GPa)

RT 220 RT 240 RT 280 RT 220

563 504 441 430 1092 997 469 428

723 679 679 597 1165 1080 617 534

20 18 18 16 5 7 29 24

60 55 45 47 6 19 71 72

210 194 210 189 207 178 207 200

TND T

(*C) 60 75 250 - 30

L. Stumpfrock et aL / Fracture mechanics investigations

33

400.0.

[3 22 Nil~oCr 37 0 22 NIMoCr 37 A +

,-),

53 NiI~o 1 17 MoV 8 4

NKS3 NKS5 weld NKS 6

L-S L-S L-S S-L

3000

300

i

>, OD i

1220 °c, Bz4 N/mm ....." .. .....

...,........

2o0.0. . . . .

/

240 °C,37.7 N/mm

250 A

0

s

H 200 D.

>

I

NKS 3 NKS 5

A

A-G

A

looo

/

/

....;. ;::-~]SIII]I

i 200 °C,33.1 N/ram

! .,::.-'"

. ............

O0

lso

0,0

-O.5

0.5

1.0

1.5

2.0

2.5

[mm]

I

Fig. 4. Notch impact energy spectrum of the NKS-program. 100

400,0 -r

o

~ -200

-10C

,-rE I ~ ~ L-+°=Y-~+ 0

100

~

]-"NKS 61

/ I +-+~ ~ + +'" [ 200

temperature

300

' ;

I

I.//

oool--weld

]

.~'leo0c,185 wmm I ,~:~°C.lSa'N/mm

~.00

[OC]

Fig. 3. Notch impact energy spectrum of current materials.

g, , '~

The load paths for infinitely long axial cracks of various depths in the RPV-wall and the transients resulting in case of a hypothetical rupture of the feedwater line (LOCA accident) are shown in Fig. 6. It can be assumed that smaller cracks will not expand provided the K 1 loading path will not tangent or intersect the appropriate material characteristic curve. However, in the previous paragraph it was stated that crack initiation may appear in the upper shelf of the material toughness on reaching the appropriate material characteristics. Crack initiation may occur for Ji-values (or formally calculated Kae-values) which will be less than the cut off values of 220 MPat/-m and 187 MPav/-m, resp., of the codes.

' mo.o--:

3 ~ 0 0C,47 N/mm - ........:~:::==:.;t::.......................1................................... i........ T300 °c,33 wmi"n ]

i:" i"

o.o .....

,

-0.5

0.0

- - I 0.5

1,0

I

I

2.0

1.6

Fig. 5. Crack resistance curves of the materials.

6O00 Nmm-31!

i W=160

O = &0tglfn~.&0

mrn

. i

5000

"\

I¢;.-",.r',\ 11

o

HPnf~

A~. B I III//

- 100

'

\.l

3. Description of the examinations

The test loop installed at MPA Stuttgart is shown in Fig. 7. The thick-walled hollow cylinder (D i = 400 mm, t = 200 ram) has been welded in at both ends to the grips of a 100 MN-tensile testing machine. In addition to an axial tensile load the specimen was loaded by internal pressure (pressurized water of p = 30 MPa =

- 150

•,

&O0O

3000

3.1. Experimental and numerical investigations

2.0

Aa [mm]

:!i

2000

50 1000

- / # / 2o .c . 5o

0 - 200

i -10O F i g . 6. T o a d

,~ ~ "/" I 0

100

path for [OCA

200

"C 300

accident.

L. Stumpfrock et aL / Fracture mechanics investigations

00

J (N/mm) 1000|

• K u (NKS3) ,,,E sso- ,......... Kio (A8ME) ~sooKI. (ASME) 2so- ¢ K u (mat•Hal) .2oo .............. :, p.) 160 .

.

|

.

/

.o~. ~

,

.... .... .

.:

.

~- 1oov

/

-

~ 50

-/ I"

T

150

o

I

....... .table crack

o

I

~

""

.00;,.-o-- 1

~

100

/

....

.

500

35

T 200

T 250

i 300

350

30mln2'o

temperature / °C

10

tlme

Fig. 10. K[ vs. crack tip temperature in the NKS 3 specimen.

0

1

2

3

mm

crack growth Aa

Fig. 11. Comparison of crack driving force and crack resistance in the NKS 3 specimen.

over the entire specimen length could be obtained as in previous tests [12,13]. The distribution of the stress intensity factor Kj (formally calculated from the J-integral) as function of the crack tip temperature, Fig. 10, was determined from the measured temperature distribution using an elastic plastic finite element analysis. The distance to the Kzc-Curve can clearly be seen. Owing to the correlations described in section 2 a

cleavage initiation should not be expected. However, the Kjc-value of the material calculated from the effective Ji-value of the material is clearly lower than the load path with K-values, Fig. 10. Considering this fact the crack should have initiated by tear. This correlation becomes obvious in another representation showing the crack tip loading by the J-integral. The course of the J-integral in the left part of Fig. 11 shows that the crack has reached its maximum

80.0

A

r =:

•Ib

average measured c r a c k e x t e n s i o n (:1.6 a m )

118.0 (

1

•q

a v e r a g e Initial c r a c k depth 62.8 m m

ultra sonic c r i c k depth m e a s u r e m e n t

i 0

1

2

3

4

5

6

7

8

c i r c u m f e r e n t i a l position, h

9

10

11

0 " b e f o r e test A - after test

12 f r a c t o g r e p h l c c r a c k depth m e a s u r e m e n t • - before test • - after test

Fig. 12. Crack extension measured and calculated in the NKS 3 specimen.

L. Stumpfrock et al. / Fracture mechanics investigations

36

crack depth of 27 mm were put on the specimen NKS 5 (low toughness with a shape welded high tough external ring of 160 mm thickness made of S 3 NiMo 1). The inner diameter of the specimen was again 400 mm with a wall thickness of t = 200 mm. The internal pressure provides a stress intensity factor of 20 MPa~/-m. In contrary to test NKS 3 the additional axial load was raised to the maximum value of 100 MN with an increase rate of 3 M N / m i n 11 min after cooling had started and kept up until the end of the test. The maximum axial force was achieved after 38 min; subsequently the crack tip loading decreased. To commence the test the specimen temperature was 230°C at the inner wall. This reduced the distance to the linear-elastic range of the fracture mechanics characteristics. The purpose of the test was to attain unstable crack extension of both cracks in the transition regime of the material up to the tough external ring. The CMOD distribution of both cracks are plotted over the testing time in Fig. 13. There were no considerable changes in CMOD after approx. 8 min., i.e. the

load after a cooling time of approx. 7 min. Stable crack growth during permanently increasing transient loading occurs after crack initation up to the time of the maximum crack tip load. The crack resistance of the material used for an evaluation of this fact should be determined at a characteristic crack tip temperature. The temperature at the crack tip was between 220°C to 230°C, Fig. 9. Therefore, the JR-CUrVe plotted in the right part of Fig. 10 was determined accordingly at 220°C. The crack growth which was numerically determined in the post test analysis amounts to approx. 3.5 mm, Fig. 11. Following the test, the crack extension was measured on the crack surface using the electron microscope. The crack extension was about 3.6 mm [7] which was averaged via the circumference, Fig. 12, and was in good agreement with the analysis. 3.1.2. Test results N K S 5 Two prefatigued semi-elliptical cracks with the circumferential angle of 52° and each with a maximum

O,8 mm

O,6 a 0.~ 0 1111 SI'17

0 0,2 G3

0,2

20 T i m e

-20

,,o

60

min

80

/,0

60

min

80

gl1~ TUI-

0,8 mm

j

0.6 a0.~ 0 r,.p 03

-0,2

-20

f-

-

-

A

G6

20

Time

Fig. 13. Crack opening of the partial circumferential cracks in the NKS 5 specimen.

.r'll :IIIIZ

L. Stumpfrock et aL / Fracture mechanics investigations

most important occurrences of the test had already taken place. This is also shown in the course of the stress intensity factor as function of time, Fig. 14. The thermal and mechanical portion of the Krvalue which are separately shown elucidate the essential crack loading for the test to be due to thermal loading. Five minutes after starting the test the Krvalue reached the range of the ductile crack initiation value Kjc (cf. Fig. 1, Fig. 4) according to which ductile crack initiation should take place with subsequent ductile crack growth. At that time the temperature at the crack tip ranges between 140°C and 150°C, Fig. 15. This representation shows clearly that the course of the Krvalues as a function of the temperature reaches the ductile initiation values considerably earlier than the Kit-curve of the code which also characterizes the cleavage initiation, however, in the linear-elastic range. In Fig. 16 the stress intensity factor is represented along the crack front hardly indicating any differences in the Krdistribution except in the range 5 ° < ~o < 30°. In this representation a two-dimensional, axisymmetric analysis is compared with a 3D-Finite-Element Analysis. The deepest point of the crack (symmetry plane, 90°) agrees well in both calculations. The fracture behaviour can be described in good approximation by a plane analysis. The fracture surface of specimen NKS 5 is represented in a survey in Fig. 17. The analysis depicted in Fig. 15 shows that the crack has initiated. It has grown both in specimen circumferential direction over an area of 220° as well as in wall thickness direction where it has been arrested by the tough welded material at a crack depth of 40 ram, Fig. 18. The crack arrested due to the considerable higher initiation values of the welded material which could not be reached by the Kz-values at the crack tip (cf.

300 ~

¢

:E 2 0 0 ~ * - m

/

.......

~

~E 150~ "-/

/

•

:~,~... • /

I

K (mech. + therm.) J

I

T 80

.........

/

crack Jump from I1 la/W O.lSS to IdW 0.21 I I

I100

0

~

:

T 150

I

~ 200

'~ 250

200

350

temperature / ° C

Fig. 15. Krcourse versus crack tip temperature in the NKS 5 specimen (crack centre).

also Fig. 15). Enlargement of the crack channel indicates a fast cleavage crack propagation with short straight main and secondary cracks, Fig. 19. This is especially shown in twin-formations, Fig. 20, which are significant for fast crack growth processes. The conclusion can be drawn that the crack expanded in a "jump" up to the tough weld material after initiating in the brittle regime. This corresponds with the time history of the COD-value and the J-integral according to which no significant increases appeared in the course of time history after initiating, Fig. 13. The little jumps in the crack opening values may indicate a slight crack growth in circumferential direction. However, the cracks did not grow together completely in circumferential direction until the end of the test, Fig. 17. These statements were confirmed by the appearance of the fracture surface showing cleavage fracture with little ductile portions. However, an exception is a seam of ductile fracture at the end of the fatigue crack

3D-ABAQUS

z

~-

.....

-10

.:

" ~'~;~:

0-~ 0

.........

.

OCA @

I

ot

50

/~ f

',4

K,(mech.)

~"-1 ~ K, (~rm.I -

• Ku ( 2 2 N I M o C r 3 7 ) l [ ~ K u (83 NIMo 1) I I

_ | • K I(mech.+therm.) "~ 2s0~ ......... K~ (ASME)

150

a°° /

37

50~

!

i

,

@

I @ t-Omln l i~ t = 6 mln [ ® t=38mln

I 25 ~

..........

:

i~ t = 60 mln 10

2o 20 time / rain

4o

so

eo

Fig. 14. Kx-course versus time in the NKS 5 specimen (crack centre).

0

~0

20

3o

40

50

eo

70

so

90

a n g l e ~ / degree

Fig. 16. Krcourse along the crack front in the NKS 5 specimcn.

38

L. Stumpfrock et aL / Fracture mechanics investigations

which can be interpreted as "stretched zone", Fig. 21, Aa ~ 0.05 mm.

The delineated time history of the crack expansion corresponds to the results of the acoustic emission measurement. 3.1.3. Test results o f specimen N K S 6

Specimen NKS 6 was tested within the scope of the research project KS 22 [8] (D i = 400 mm, t = 200 mm).

It was also manufactured with a shape-welded high-tough external ring made of S 3 NiMo I in 100 mm thickness. Prior to testing an axial load of 25 MN was applied in addition to the internal pressure of 300 bar. The combination of internal pressure and axial load results in Ki-values ( ~ 50 MPa~/-m) just below the scatter band of the Kit-, Kjc-values of material KS 22 out of which the NKS 6 specimen was manufactured. The transition temperature determined with the

Fig, 17. Fracture surface of specimen NKS 5, view.

L. Stumpfrock et al. / Fracture mechanics investigations

Fig. 18. Crack profile of the specimen NKS 5, polished.

Fig. 19. Crack profile NKS 5 with secondary cracks.

39

40

L. Stumpfrock et a L / Fracture mechanics investigations

Fig. 20. Crack profile of specimen NKS 5 with twin formation.

FATr50-criterion is about 250°C. The usual procedures for determination of RTNDT (PeUini) were not applicable.

The measured temperature values represent the typical behaviour of the thermal shock test over the wall thickness and over the time, cf. Fig. 9. In this case

SR

Thermoschockril~

Laborbruch

Fatigue crack

Thermal shock crack

lab. Induced brittle fracture

L. Stumpfrock et aL / Fracture mechanics investigations

Thermoschockril~

SR _i ~atigue Icrack f

Thermal shook crack

41

t "~i

Laborbruch lab. Induced brittle f r a c t u r e

Fig. 22. Fracture surface of specimen NKS 6.

it was also possible to obtain constant cooling over the total specimen length. The fracture surface of a broken segment of the specimen can be seen in Fig. 22, and in Fig. 23 the crack contour is demonstrated in a longitudinal cut. Two regions in different fracture modes are visible during fractographic examinations. The fracture surface immediately following the fatigue crack demonstrates predominantly cleavage fracture (area 1) which turns into a complete ductile fracture mode (area 2) after about 13-20 mm crack extension. In area 1 twin formation could be observed, however, it was not so

pronounced as in specimen NKS 5. The ductile fracture decelerated. This is supported by the acoustic emission results. Therefore, a crack extension in two phases up to crack arrest in the up-welded external ring can be derived from these findings. On the basis of the load measured an axi-symmetric finite element analysis was carried out including crack propagation. In accordance with the fracture surface findings and the acoustic emission measurement results the crack growth was simulated during the analysis in two phases. Following initiation at t = 35 s the first phase was a cleavage crack jump of A a = 20 mm. Following a holding phase of A t = 17 s, in the second crack growth phase the crack propagation continued up to the tough welded material ( A a ffi 41 mm, A t = 24

IEmqmqmM mll,nlmil,,,I,,li,Sl,lll,,li,ll,,IH,l,, ll"q"lll" 112 IQ

L~. eD

i,-

L%'l.2L,,[,,tlil.h.J..h.,L,i,. J.,,h,~.,~J,,~,, ,, r

E E a o o

1.6 1,4 1.2 1 0,8- i

0.8-

0.4-

..... • calculation m measurement (G6)

0.21 0 0

50

i

i i 100 150 200 250 300 350 400

time / s

Fig. 23. Crack profile of specimen NKS 6.

Fig. 24. Crack opening versus time comparison measurement - calculation.

42

L. Stumpfrock et aL / Fracture mechanics investigations

s). The comparison between measurement and calculation in Fig. 24 explains the good assessment of the crack growth by the calculation. The course of the crack tip load as a function of the temperature is given in Fig. 25. As a result of the crack jump and the subsequent crack growth the crack tip grew into the range of higher temperature. Due to this the loading first decreases and later increases again during further cooling. On reaching the ductile weld material no more crack growth occurred, Fig. 23.

olt

CT2G

•

NK86 (before crack Jump: t - 35 s)

® NK83

•

•

NK86 (lifter orack jump: t - 3G s)

e

0

1

NKI~

2

® NKH(t-Ha)

3 4 5 6 r 8 9 10 11 12 13 14 15 distance from crack tip / mm

3.2. Explanation of the test performance by means of the multiaxiality of the stress state

Fig. 26. Multiaxiality degree q in the ligament.

The degree of multiaxiality of the stress state may quantitatively be covered by the multiaxiality quotient q defined as

This is the case in specimen NKS 5, Fig. 26, so we have to expect ductile initiation with following cleavage crack extension. In addition a flat crack resistance curve, Fig. 4, indicates relatively little crack resistance of the material. Furthermore, the material is clearly in the transition area of the toughness even at a temperature of 150°C. In the case of specimen NKS 3 the course of the quotient of multiaxiality q in the ligament cannot be compared with the distribution of q in specimen NKS 5, Fig. 26. With the exception of the crack tip area the q-course obtained in the NKS 3 specimen did not attain the low values which occur in the NKS 5 specimen over larger ligament areas, and in fact, it increases continuously. The course of q in the ligament of a CT-25 specimen at maximum load is entered as well and can be compared with the q-course in the ligament of specimen NKS 3. This means that according to [15] the degree of multiaxiality of the stress state and also the crack resistance behaviour is comparable, i.e., the JR-Curve of the CT-specimen reproduces the crack resistance behaviour of the NKS 3 specimen. Thus, the good agreement of the numerically and experimentally determined stable crack growth is explained. The NKS 3 specimen, in contrary to the NKS 5 specimen, did not show any fast crack propagation. The reason can be found in the toughness of the material on reaching the critical test phase (initiation). In the ease of NKS 3 specimen the crack tip with approx. 220"C was in the upper shelf at a notch impact energy of approx. 90 J. The NKS 5 specimen was noticeable in the transition regime of the notch impact energy at the crack tip with approx. 140-151YC and a notch impact energy of about 50 J, This influence of toughness is also documented in the course of the crack resistance curve running steeper

7r q ----- Orm

1 crv ~

O"m

with o,v as von Mises equivalent stress and crm as the hydrostatic stress portion [14]. By means of FE-analyses q can easily be determined for discrete points of the ligament. In [15] it was possible to explain the different slopes of the crack resistance curve of specimens different in geometry and size with the distribution of q in the ligament. However, quantitative dependence could not be obtained. Furthermore, spontaneous fracture of the specimen or component without previous stable crack extension has to be expected if the q-value attains or is below qe in front of the crack tip. A critical qe-value was defined within the scope qc < 0.3 in [15]. Additionally, the effective crack initiation value Ji must be reached or exceeded.

300

.•

250-

m ~; 200-

• Koj (NKS6) .......... K~ (ASME)

I

..... K m (ASME)

1

eee,,

j

.i

_l 150,v 100-

i ~,'

50-

0

50

100

150

200

250

300

350

tenlperature / °C

Fig. 25. Ki-course over the time in the NKS 6 specimen.

L. Stumpfrock et aL / Fracture mechanics investigations •

CT2§ NK83 ;~ NK85

• NK86 (before Grack lump: t = 35 s) I • NK86 (after crack Jump: t - 35 s) ® NKS6 (t ,: 52 s)

1 0.80.6o

~r

0.4 0.2 0

o.o

~.1

~.~

o.~

o.,

normalized ligament

Fig. 27. Multiaxiality degree q in the normalized ligament. at the relevant temperature of NKS 3 material. It has been enhanced to higher J-integral values than the appropriate JR-curves of the NKS 5 material. Furthermore, the initiation value of the NKS 3 material is nearly twice as high as the one of the NKS 5 material. Consequently, the initiation in the NKS 3 specimen takes place under considerably higher loading. Higher loading increments are required as well for driving the crack. The distribution of q over the normalized ligament * provides results which confirm qualitatively the above representation, Fig. 27. However, the NKS 3 specimen should have shown more stable crack extension than it was estimated in the calculation. This is due to higher multi-axiality (smaller q-value) which is present over larger areas of the remaining ligament. The stable crack growth, in the fracture mechanics sense, represents under the microscope a coalescense of voids which develop mainly at nonmetallic inclusions in front of the crack tip. However, the distribution of these inclusions is material-specific and thus geometry-independent. Therefore, the representation in Fig. 26 which is not normalized is to be used when assessing the stable crack extension. The test run of the NKS 6 specimen is not comprehensible with the q-courses, Figs. 26 and 27. At the beginning of the test a temperature of 285°C, which is in the upper transition area of the notch-impact energy temperature curve, can be found at the crack tip. The cooling procedure is the reason for the further dislocation of the crack tip temperature into the transition area of the toughness. It was not possible to determine crack resistance curves on CT-specimens in this temperature region. However, even valid Kie-values could * The ligament is standardized to the initial one.

43

be determined, Fig. 25. The first initiation at t = 35 s causes a brittle fracture on reaching the Kic-Vahie. This will be represented again with the characteristic values even if the material scattering is relatively large. It is remarkable that the Ki-value releasing the brittle fracture nearly meets with the Kle-CUrve according to [2] and [5], Fig. 25. The crack jump should have been caused up to the Kla-Value from the Kia-curve. This was not the case but is not surprising because of the large material scattering. The holding phase of the crack which can also be gathered from the acoustic emission measurement is plausible, since the second initiation (t = 52 s) took place on reaching (accidentally) the Kit-curve at a higher Ki-value than in the case of the first initiation. As described previously, the fracture was a ductile one which did not jump. The crack extension velocity, however, was larger than the cooling velocity in the ligament so that the crack tip was arrested at a temperature in the upper shelf of the toughness. This behaviour can now be explained by means of the course of the multi-axiality quotient q. The multiaxiality of the stress state and the risk of reduced embrittlement fracture decreases with increasing material toughness. Thus, there is a possibility of ductile fracture, cf. Figs. 26 and 27. It is not surprising that the Kic controlled brittle fracture phase (first initiation) cannot be explained by means of the multi-axiality quotient q, because q was developed for the assessment of the cleavage fracture susceptibility in the ductile regime [14].

4. Summary An analysis of the failure behaviour with Kic- resp. KiR-curve the specimens NKS 3 and NKS 5 would have provided results predicting an entirely different specimen behaviour than that realised in the experiment. According to this the N'KS 3 specimen should have been without crack growth. The NKS 5 specimen should have initiated at a time when considerable events in the test had already been finished. However, the processes can correctly be described with the analyses using ductile fracture mechanics. The examples demonstrate that the effective initiation values Ji of the J-integral determined on small CT-specimens (thickness B = 25 mm) with extended plastic deformation in the elastic-plastic loading range may be transferred to large, thick-walled components and delineate the same physical effects (crack initiation). This applies to the examples discussed in spite of the fact that the plastic deformation of the large scale speci-

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L. Stumpfrock et al. / Fracture mechanics investigations

mens was distinctly less than in the CT-specimens where the relevant characteristic values were determined. In comparison the NKS 6 specimen would have been safely calculated with the Kic-CUrve.

References [1] ASME Boiler and Pressure vessel Code Section III, Division 1, Appendices, Article G (1988). [2] KTA 3201.2, Fassung 3/84: Sicherheitstechnische Regel des KTA. Komponenten des Prim~kreises von Leichtwasserreaktoren, Teil 2: Auslegung, Konstruktion und Berechnung. [3] R~gles de Conception et de Construction des Materiels M6caniques des Ilots Nucl~aires PWR, RCC-M, Annexe ZG (Edition 1988). [4] PVRC Recommendations on Toughness Requirements for Ferritic Materials, PVRC Ad Hoc Group on Toughness Requirements, WRC Bulletin 175 (Aug. 1972). [5] ASME Boiler and Pressure vessel Code, Section XI, Division 1, Appendix A (1988). [6] G. Katzenmeier and W. Miiller-Dietsche, HDR-Safety Programme, Phase II: PHDR-Report No. 05, 19/84, Kernforschungszentrum Karlsruhe (1984). [7] BMFT-Project 150618, Experimental and numerical fracture analyses for reactor pressure vessel subjected to pressurized thermal shock loadings, MPA Stuttgart, final report (1990). [8] BMVI'-Project-150787, Experimental and numerical in-

[9]

[10]

[11]

[12]

[13]

[14]

[15]

vestigation of a low toughness vessel material at unstable crack propagation and arrest under combined mechanical and thermal loading, MPA Stuttgart reports in preparation. E. Roos, T. Demler, U. Eisele and R. GiUot, Fracture mechanics safety assessment based on mechanics of materials, Steel Research 61 (1990) 172-180. K. Kussmaul and A. Sauter, Anlagensimulation zur Behandlung der Notkiihlproblematik beim Leichtwasserreaktor-Druckbeh~lter, 10. MPA-Seminar, Universit~it Stuttgart, Oktober 1984. K. Kussmaul, A. Sauter, R. Stegmeyer and W. Guth, Fracture evaluation of surface cracks in light water reactor vessels under pressurized thermal shock loading, One day discussion meeting on Thermal Shock; NPTEC, Leatherhead, UK, 28 Sept., 1989. A. Sauter, T. Nguyen-Huy and H. Huber, Experimentelle und analytische Ergebnisse eines GroBzugversuchs zur Notkiihlsimulation, 11. MPA-Seminar, Universit~it Stuttgart, Oktober 1985. K. Kussmaul and A. Sauter, Ductile crack behaviour under Pressurized Thermal Shock, ANS/ENS Topical Meeting on Thermal Reactor Safety, San Diego, USA, February 1986. H. Clausmeyer, l]ber die Beanspruchung yon Stahl bei mehrachsigen Spannungszusfiinden, Konstruktion 20 (1968) 395-401. H. Clausmeyer, K. Kussmanl and E. Roos, Der EinfluB des Spannungszustandes auf den Versagensablanf angerissener Bauteile aus Stahl, Materialwissenschaften u. Werkstofftechnik 20 (1989) 101-117.