The heavy-section steel technology pressurized-thermal-shock experiment, PTSE-1

Enginrwmp Frrrcrurr M~~honrts Printed in Great Britain.

Vol. 23, No. I, pp. 81-97.

1986

Wl3-7944/X6 53 00 + .oO Pergamon Press Ltd.

THE HEAVY-SECTION STEEL TECHNOLOGY PRESSURIZED-THERMAL-SHOCK EXPERIMENT, PTSE-l’f R. H. BRYAN, B. R. BASS, J. G. MERKLE, C. E. PUGH, G. C. ROBINSON and G. D. WHITMAN Oak Ridge National Laboratory, Oak Ridge, TN 37831, U.S.A. Abstract-A pressurized-thermal-shock (PTS) facility was developed in the Heavy-Section Steel Technology Program at Oak Ridge National Laboratory for performing experiments that challenge predictions of analytical methods applicable to full-scale reactor pressure vessels under combined loading. The first experiment (PTSE-1) was designed to address three principal issues: (1) warm-prestressing phenomena; (2) crack propagation from brittle to ductile regions; and (3) transient crack stabilization in ductile regions. The paper presents a description of the PTS facility at ORNL and a review of the objectives and results of the first test. Also included are elastodynamic finite-element analyses of the two crack run-arrest events that occurred in the second and third phases of the test. Finally, some conclusions and recommendations are presented based on the outcome of the first experiment.

INTRODUCTION THE FIRSTpressurized-thermal-shock

experiment (PTSE-I) in the Heavy-Section Steel Technology (HSST) Program at the Oak Ridge National Laboratory (ORNL) is the most recent of a long succession of fracture-mechanics experiments that are on a scale that allows important aspects of the fracture behavior of nuclear reactor pressure vessels to be simulated. Such experiments serve as a means by which theoretical models of fracture behavior can be evaluated for possible application to fracture analysis of vessels in nuclear plants. The principal issues of concern in the pressurized-thermal-shock experiments are (1) warm-prestressing phenomena, (2) crack propagation from brittle to ductile regions, (3) transient crack stabilization in ductile regions, and (4) crack shape changes in bimetallic zones of clad vessels. PTSE-1 was designed to investigate the first three of these issues. A plan and facility were developed for performing pressurized-thermal-shock experiments that challenge the predictions of analytical methods applicable to full-scale reactor pressure vessels (RPVs) under combined loadings. The scale of the tests was chosen to be large enough to attain effectively full-scale behavior of the flawed region. Tests conditions and materials were selected to produce stress fields and gradients around the flaw that are characteristic of RPVs and to give realistic fracture-toughness conditions. The test facility was designed to be capable of producing a variety of fracture possibilities: cleavage initiation of small flaws, cleavage initiation and arrest below the upper shelf, cleavage initiation with arrest on the upper shelf, arrest in a high K1 gradient, warm and antiwarm prestressing states in succession, and progressive (upper-shelf) tearing, tearing instability, and restabilization. In the first test, PTSE-I, the crack was long, sharp, and shallow, as is assumed in regulatory evaluations. The material properties were typical of those for vessel steel after moderate neutron embrittlement (such as a high RT NET). The stress levels and gradients around the outside surface flaw in the test vessel were chosen to approximate those that would occur in a pressurized water reactor (PWR) vessel with a flaw on the inside surface during a severe pressurizedthermal-shock transient. The following sections present a description of the pressurized-thermal-shock test facility at ORNL and a review of the objectives and results of the first test. Also included are elast Research sponsored by the Office of Nuclear Regulatory Research, U.S. Nuclear Regulatory Commission under Interagency Agreements 40-551-75 and 40-552-75 with the U.S. Department of Energy under Contract DE-COS840R21400 with Martin Marietta Energy Systems, Inc. 81

R. H. BRYAN etal.

82

todynamic finite-element analyses of the two crack run-arrest events that occurred in the second and third phases of the test; the dynamic results are compared with those from a quasi-static method. Finally, some conclusions and recommendations are presented based on the outcome of the first experiment. DESCRIPTION

OF THE EXPERIMENT

The flawed vessel was enclosed in an outer vessel as shown schematically in Fig. 1. The outer vessel is electrically heated to bring the vessel to the desired uniform initial temperature of about 290°C. A thermal transient is initiated by suddenly injecting chilled water or a methanol-water mixture into the outer vessel. The annulus between the cylindrical surfaces of the two vessels was designed to permit coolant velocities that would produce the appropriate convective heat transfer from the test vessel for a period of about 10 min. Pressurization of the test vessel is controlled independently by a system capable of pressures up to about 100 MPa. The plan for PTSE-I was to initiate and arrest a fast-running crack, make the arrested crack supercritical (Ki > Ki,) while in a warm prestressed state, and subsequently reinitiate the crack, allowing it to advance toward the completely ductile material deep into the wall. The necessity to preserve evidence of crack geometry precluded the deliberate bursting of the test vessel. Extensive material properties tests and fracture analyses preceded the transient test of the PTSE-1 vessel. The initial l-m-long by 12-mm-deep flaw was axially oriented on the outside (cooled) surface of the 148-mm-thick vessel. The transient test was performed in three phases; in each phase the vessel was initially in an isothermal state (-290°C). Each phase consisted of a pressure transient and a thermal transient, which were coordinated to produce an evolution of stress and toughness states that would fulfill the objectives of the plan. Fracture analyses performed to define the transients were based on fracture-toughness data from tests of small specimens. Much of the run-arrest portions of the expected crack jumps in the experiment would take place in a temperature range above that for which the small specimens could provide data; consequently, the transients were selected to attain the desired objectives in the presence of uncertainty. The ORMGENiADINAiORVIRT system[ l-31 of finite-element computer programs was used in conjunction with the OCAKJSA program[4] to define fracture properties and transients that would meet PTSE-1 objectives. The details of test vessel and flaw geometry are shown in Fig. 2 and Table 1. HSST intermediate test vessel V-8 was prepared with a plug of specially tempered SA508, Class 2 steel welded into the region to be flawed. The l-m-long sharp flaw was implanted by cracking a shallow electron-beam weld under the influence of hydrogen charging. The vessel was extensively instrumented to give direct measurements of crack-mouth opening displacement,

INITIAL TEMPERATURE Tel

INTERNAL PRESSURE P It)

Fig.

I, Schematic

of a pressurized-thermal-shock

test vessel

enclosed

in an outer

vessel.

Pressurized-thermal-shock

experiment

83

PTSE-1

/HEAD AND ACCESS NOZZLE SUBASSEMBLY

h

f ‘y-_

/

7

381 mm ID

686 mm ID

FLAW

1

1372 mm

I -

LTEST REGION WELD INSERT

Fig. 2. Geometry of PTSE-1 test vessel.

temperature profiles through the vessel wall, and internal pressure during the transient (see Fig. 3). Pretest fracture analyses were based on computed temperature profiles and postulated pressure transients, while posttest analyses employed measured temperatures and pressures. Properties of the test material are given in Table 2 and Figs. 4 and 5. Fracture initiation and arrest toughnesses were determined from tests of 25 mm and 37 mm compact specimens, respectively. Figures 4 and 5 show the K, and K, data together with size-effect-adjusted data[5] and the curves used in fracture analyses. The A curve in Fig. 4 was used in analyses made prior to execution of the first transient (PTSE-lA), and curve B was used subsequently. The initiation-toughness data in Fig. 4 were obtained by the Babcock and Wilcox Research Center and the arrest-toughness data in Fig. 5 by the Battelle Columbus Laboratory in support of this program. In any persistent transient there is generally a time tr beyond which an arrested crack in the upper-shelf regime would not be stable, either because of a tearing instability or a net ligament tensile instability. Because it was considered essential to the interpretation of the experiment to preserve evidence of the arrested-crack geometry, it is necessary that rI be predictable and that the transient be terminated at some time prior to tr.

Table 1. Geometric parameters of PTSE- 1 vessel Parameter Inside radius, mm Wall thickness (w), mm Flaw length, mm Flaw depth (a), mm alw

Value 343 147.6 1000 12.2 0.083

R. H. BRYAN

Fig. 3. Instrumented

Table

PTSE-I

2. Properties

test vessel

et al.

with crack-mouth-opening along the flaw.

of PTSE-I vessel special tempering

material (A508, treatment)

Property

displacement

Cl 2 steel with

Value

Ja Parameters?

&et of Charpy upper shelf, “C Ductile threshold temperature, “C RTND~, “C Yield stress, MPa Young’s modulus, GPa pretest posttestS Coefficient of thermal expansion, Km pretest posttestS Poisson’s ratio Thermal conductivity, W.m-‘,K-’ Heat capacity, J.kgg’.K-’ Density, kg/m3 + Ja = c(Aa)“; Ja in MJ/m’, $ These average values are E(r) and a( 7) for the vessel 1% of the values based upon

2.60 0.359 150 175 91 600 200 and 209.6 202.3

’ 1.3 x lo-‘and 1.445 x lo-’ 1.441 x 10 -5 0.3 41.54 502.4 7833

Aa in m. based on experimental measurements of material, and they give values of K1 within temperature-dependent properties.

gages

vlisible

Pressurized-thermal-shock

_

0

experiment

85

PTSE-1

ITCS DATA

I

&-ADJUSTED

V VALID

BAND

K,,

0

CURVES K,, = a + beCT A a b c

0 -50

51.276 1.83311 0.036

6 51.276 2.2 0.036

I

I

I

I

I

0

50

100

150

200

TEMPERATURE

250

(‘C)

Fig. 4. Fracture initiation toughness values obtained from laboratory specimen tests. (Curve A used in pretest analyses for PTSE-IA and Curve B used subsequent to performing PTSE-IA.)

The pressure and thermal transients for the three phases (A, B, and C) of the experiment are given in Fig. 6 and Table 3. For the A test, the KI trajectory reconstructed from experimental data is shown in Fig. 7. Since temperature (on the abscissa) decreases monotonically with time, one can discern from this plot two episodes of simple warm prestressing (KI < 0), each followed by simple antiwarm prestressing (KI > O), while K1 is greater than K1,. Because of the inhibiting influences of warm prestressing, the crack did not propagate during the A transient. Plans for the B and C transients were based upon the evidence from test phase A that the vessel was tougher than originally estimated and that, to overcome warm prestressing, a higher KI value would have to be attained. Accordingly, the B curve of Fig. 4 was adopted for further Table 3. Conditions for PTSE-IA, -lB, and -lC transients Test transient

Thermal transient parameters Initial vessel temperature, “C Coolant temperature T(r), “C h(t), W.m-*.K-’ Pressure transient (planned) Initial flaw depth a, mm a/w

t Initial and final (t = 300 s) values.

PTSE-IA

PTSE-1B

PTSE-1C

277.6 1%34t SOOO-6000t Curve A, Fig. 6

290.7 -22-o-F 5500-6500t Curve B, Fig. 6

287.4 - 29-14t 4000-5500t Curve C, Fig. 6

12.2 0.083

12.2 0.083

24.4 0.165

86

R. H. BRYAN et al.

I _

I

0

CS K, DATA

A

!+ADJUSTED

CURVE:

0

DATA

K,, = at beCT a b c

-50

I

= 35.0 = 4.0177 = 0.02408

50

150

100 TEMPERATURE

200

250

(‘C)

Fig. 5. Crack-arrest toughness values obtained from laboratory specimen tests. (CS) data are from 37-mm-thick compact specimens, and the curve is a least-squares fit to the size-effectadjusted data with the highest point at 67°C excluded.

analysis, lower coolant temperatures were specified for the thermal transient (Table 3), and a higher pressure transient was selected (curve B, Fig. 6). A two-step pressure transient was not performed during PTSE-1B test because a second pressure increase of a useful magnitude was not within the capabilities of the pressurization system. The B transient resulted in a crack jump to a depth of 24.4 mm. The conditions of initiation and arrest from an analysis performed with OCA/USA are shown in Fig. 8. The final transient, PTSE-IC, was performed under the conditions given in Table 3 and with the planned pressure transient described by curve C in Fig. 6. The crack jumped to a depth of 41 mm under conditions presented in Fig. 9. The vessel was examined visually and ultrasonically after the C transient. At the outside surface the crack extended axially about 110 mm at the upper end of the vessel and about 120 mm at the lower end (Fig. 10). The crack branched at the lower end, as shown in Fig. Il. Test instrumentation indicated that all of the axial extensions occurred in transient PTSE-IB. The flawed region was cut from the vessel, chilled i,n liquid nitrogen, and broken apart to reveal the fracture surfaces which are shown in Fig. 12. Details of a segment of one surface are shown in Fig. 13. Fractographic examination of the surfaces and measurement of the flaw geometry indicated that the initial flaw tore slightly prior to the initial cleavage fracture. The initial crack extension was essentially a pure cleavage fracture throughout the first half of the extension and predominantly cleavage (-90%) with finely dispersed ductile tearing in the remaining extension. The crack extension in the second crack jump was mixed mode throughout with about 85% cleavage. At the ends of the two crack extensions there were no coherent regions of ductile tearing, contrary to predictions based on the measured tearing resistance JR of the material.

Pressurized-thermal-shock

experiment

87

PTSE-1 3

200

100

TIME FROM START

Fig. 6. Planned

pressure

OF THERMAL

transients

TRANSIENT

for PTSE-lA,

(5)

-lB,

and -lC.

250 225 200 175

m %

125

k+ - 100 it?75 50 25

50

75

100

125

150

175

200

225

CRACK TIP TEMPERATURE Fig. 7. Results

250

275

300

(‘C)

of OCA/USA analysis of PTSE-IA based on measured temperature, and flaw depth. The KI, expression is curve B of Fig. 4.

pressure,

88

R. H. BRYAN et al.

..-.*--

75

100

CRACK JUMP

125 150 175 200 225 250 CRACK TIP TEMPERATURE F’C)

275

Fig. 8. Results of OCAlUSA analysis of PTSE-1B based on measured temperature, flaw depth, and time of crack jump.

450 -

:I : :

400-

: :

500

0

I

I 50

I

1

I

1

1 . /

I

I

100

150

CRACK

I’

II

I

I

INITIATION

pressure,

’ -

A ARREST -...w CRACK JUMP -

I 200

I

I 250

I

4 300

TIP TEMPERATURE tot)

Fig. 9. Results of OCAKJSA analysis of PTSE-IC based on measured temperature, flaw depth, and time of crack jump.

pressure,

Pressurized-thermal-shock

experiment

89

PTSE-1

\

AXIAL

/

CRACK

EXTENSION

FRACTURE AFTER

LINE

CHILLING

IN LIOUID

N2

HORIZONTAL INITIAL F LAW

FLAME

SOLID

LINES ACROSS FLAW ARE SAW CUTS

CUT

/ 4 38

4 58

AXIAL

CRACK

EXTENSIONS >B J/BRANCH

BRANCH

1

21

Fig. 10. Diagram of the intersection of the flaw with the outside surface of the test vessel. Locations of segments of the flaw and axial extensions of the flaw are shown.

Table 4. Summary of fracture conditions in PTSE-1

Experiment PTSE-IA

PTSE-1B PTSE-1C

Event

Crack depth (mm)

1st max Kr 12.2 (at KI = KI,) 2nd max Kr 12.2 3rd max Kr 12.2 Initiation 12.2 Arrest 24.4 Subsequent max KI 24.4 Initiation 24.4 Arrest 41 Subsequent max K, 41

Crack depth ratio WV)

Crack tip temperature (“Cl

0.083

105

152

0.083 0.083 0.083 0.165 0.165 0.165 0.278 0.278

78 57 104 163 118 125 179 156

154 139 177 201 247 254 299 340

Fig. 11. Photograph of the branched lower end of the PTSE-1 flaw. Strain gages XE53 and XE54 were located 10 mm and 100 mm, respectively, from the end of the initial flaw. M&C PHOTO Y196788A

)-OUTSOE

SURFACE OF VfS!EL

LOWER EN0 OF INITIAL UPPER ENDOF

INITIAL

-TOP

FLAW

FLAW OF VESSEL

Fig. 12. Montage of fracture surfaces from PTSE-1. Only branch 1 surfaces are shown in pieces 6A and 6B. (See Fig. 10 for key to location and branch number.)

----EXTENSION

/

FIRST CRACK . EXTENSION

-8

FIRST I~tTtATtON StTE

ARREST AND SECOND lNtTlATlON SITE

ARREST SITE

Fig. 13. Typical portion of fracture surface from PTSE-1 (surface A in segment 3).

Pressurized-thermal-shock

experiment

91

PTSE-1

Initiation and arrest toughness from quasi-static calculations are summarized in Table 4 for the three phases of the experiment. The values of Ki, and Kla inferred from test data are shown in Fig. 14 in comparison with the pretest estimates and with the Ki, and Ki, relationships suggested in Section XI of the ASME Boiler and Pressure Vessel Code. Pretest estimates of fracture toughness are reasonably close to the PTSE-1 values. The arrest values in PTSE-1 are consistent with arrest measurements made in the wide-plate tests reported in Refs. [6] and [7] and illustrated in Fig. 15. ELASTODYNAMIC

ANALYSES OF THE EXPERIMENT

For purposes of comparison with the quasi-static posttest analyses described above, elastodynamic analyses of PTSE-1 were carried out with the SWIDAC[8] dynamic crack-analysis code. The salient features of the crack-propagation technique used in SWIDAC are reviewed briefly in this section, followed by applications of the technique to the B and C phases of PTSE1. The SWIDAC code utilizes a displacement-based finite-element formulation with quadratic isoparametric elements employed in the discretization of a two-dimensional (2-D) domain. Integration of the equations of motion is carried out with the implicit Newmark-Beta[9] time

400

SMALL-SPECIMEN PTSE-1 DATA -0OK,,

I

’

I I

Klc Klc

------SECTION

’

I Kla

XI K,, AND K,, I

BASED ON FIT,,, PTSE-1

RESULTS

OCAIUSA 3-D INFLUENCE CALCULATIONS

= 91.3’C f

BASED ON I COEFFICIENT I

0 I

l Klc

I

UPPER SHELF SHEAR FA) l

-50

0

50

100

TEMPERATURE Fig. 14. Comparison

150

(1005

200

(‘Cl

of PTSE-1B and -lC results with curves representing and KI, data and ASME Section XI curves.

small-specimen

K,,

92

R. H. BRYAN et al.

400

I

I 0 K,a DATA

l JAPANESE ESSO RTNDl W

300

*

= -3oOc

PTSE-1 = 91.3Oc

8

*:::I RTNDT = -23OC (K, = o[na sec(sa/2w)l

“1

IT ‘2 g

200 m

J

100

ONSET OF CHARPY UPPER SHELF (100% SHEAR FA) WP-1.1

-100

0

100

T-RTNDT

zuo

(K)

Fig. 15. Comparison of PTSE-1 arrest toughness results with Japanese and ORNL wide-plate test results (Refs. 6 and 7) and ASME Section XI curves.

integration scheme. In the crack-growth modeling technique of SWIDAC, the finite element immediately ahead of the crack tip is divided into N subelements (typically, N = 3-6). During propagation, the tip is moved through these subelements along the crack plane in discrete jumps. The position of the crack tip relative to these subelement divisions is determined from restraining forces which are placed on the crack-plane nodes of the element adjacent to the crack tip; these forces are released incrementally as the tip propagates through the element. In SWIDAC, the restraining forces are postulated to vary linearly with the crack-tip location according to the relation

(1) where Fi is the force at node i, Foi is the force at node i just prior to node release, and a is the length of the crack in the finite element of length Ax. Prior to node release, the crack-plane nodes are restrained normal to the crack plane by stiff spring elements. For the analyses in the present study, the crack tip is propagated incrementally according to the following relations: if Ki < Kib(b, T) then no propagation, if KI = Kdiz,

T) then propagation;

(2)

Pressurized-thermal-shock

experiment

PTSE-I

93

here K, is the dynamically computed stress-intensity factor and Ku, is a dynamic fracturetoughness relation that is taken to be a function of crack velocity Lr and temperature T. The dynamic stress-intensity factor Ki is determined in each time step from the dynamic J integral[ IO] containing the appropriate inertial and thermal terms. Additional details concerning the crack-growth modeling scheme of the type used in SWIDAC are described in a paper by Jung et a1.[8]. In applications of the above formulation to PTSE-1, a 2-D plane-strain finite-element formulation was utilized to model the test vessel of Fig. 2 with initial crack depth ratio ah = 0.083. The finite-element model is depicted in Fig. 16 and consists of 696 nodes and 208 eightnoded isoparametric elements. A total of 13 spring elements are used in the crack plane to model propagation of the crack tip. Constant material properties employed in the analysis are given by Young’s modulus E = 202.3 GPa, Poisson’s ratio v = 0.3, thermal expansion coefficient (Y = 1.44 x 10p5/“K, and density p = 7850 kg/m3. The dynamic fracture toughness relation is KD = KI, + A(b)*,

(3)

where KI, = 35.0 + 4.0177e0.02408’;

(4)

for T - RTNDT > - 13.9”C, A(T)

= [329.7

+ 16.25(T

- RT~ur)l

(5) x lo-‘j;

and for T - RT NDT 5 - 13.9”c, A(T)

= [121.71

+ 1.2962(T

- RTNDT~I X lVh,

where RTN,,-r = 91.3”C.

Fig. 16. Finite-element

ri = 0.343 m t = 0.147 m model for posttest elastodynamic fracture analysis of PTSE-I (696 nodes and 208 elements).

(6)

R. H. BRYAN et al.

94

Units for K, A, ic, and T are MPa.m”*, MPa*s2*m-3’2, m/s and “C, respectively. Measured conditions of the radial temperature profile and internal pressure at the time of crack propagation in transients B and C are given in Table 5 and are assumed to be constant during the run-arrest events. Also included in the table are the pressure and initial (uniform) temperature conditions. Initial-displacement conditions for the dynamic analysis were obtained from a quasi-static thermoelastic analysis performed with ADINA* and with the thermal and mechanical loading of Table 5. The ADINA program was employed for this calculation because of the absence of a consistent pressure element formulation in SWIDAC. For the dynamic analysis, the internal

Table 5. Pressure and temperature conditionst at time of crack initiation in tests PTSE-1B and -1C Temperature

(“C)

Radius (m)

transient B

transient C

.49060 .48714 .48347 .47953 .47584 .46846 .46125 .44741 .43173 .41680 .40226 .38745 .37274 .36514 .35798 .34300

20.45 44.65 70.25 96.4 118.3 154.1 182.6 229.0 263.3 278.85 285.7 288.95 290.5 290.9 291.2 291.4

18.45 34.75 52.05 70.3s 86.85 117.05 142.75 185.15 223.45 248.7 264.9 275.65 282.45 284.8 286.3 287.3

t Pressure at time of crack jump and initial uniform temperature: Transient B--P = 24.91 MPa and T = 290.7”C; transient C--P = 75.12 MPa and 7 = 287.3”C.

0.30

0.25 7 2 0 G 0.20

a z 4 0.16 I 0 0.15 :: aQ 0 0.10 0.08

I-

0.05 0

-I I 20

40

60

80

100

120

140

160

180

I

200

TIME (/.Is)

Fig. 17. Crack-depth ratio (ah)

vs time for posttest

elastodynamic

analysis of PTSE-1.

Pressurized-thermal-shock

experiment

95

PTSE-1 1

c

330

300

270

PTSE-II3

/

3a

_

C

I

I

0.1

“.L

ah,

Fig. 18. Stress-intensity

-

^”

CRACK

DEPTH

u.3

RATIO

factor vs crack-depth ratio (a/w) for posttest elastodynamic of PTSE-I.

analysis

Table 6. Comparison of initiation and arrest parameters from posttest quasistatic and dynamic analyses of PTSE-1 Experiment phase

n/w

PTSE-IA OCAKJSA 0.083 PTSE-1B OCAKJSA 0.083 ADINA 0.08

a Temperature (mm) (“C)

KI

Event

(MPa.m”*) 154

Max KI; no initiation

103.7 101.9

177.4 164.3

Initiation

24.0 24.4 24.4

161.4 163.5 118.0

230.8 200.9 I 247.0

Subsequent max K,

0.165 0.16

24.4 24.0

125.3 123.7

254.3 234.8

Initiation

OCARJSA 0.278 SWIDAC 0.26

41.0 39.0

179.0 172.6

298.9 291.4

Arrest

0.278 41.0

156.0

340.0

Subsequent max KI

SWIDAC OCAKJSA OCA/USA PTSE-1C OCAiUSA ADINA

OCAiUSA

12.2

78

12.2 12.0

0.16 0.165 0.165

Arrest

pressure boundary nodes of the model were fixed at the initial static-displacement values from ADINA, and the nodal temperatures were interpolated from Table 5. The stress-intensity factor calculated for the initial condition is taken as the critical initiation factor Ki,. The time step At for the dynamic analyses was fixed at At = 5 ps. Results from the elastodynamic analyses of the B and C transients are depicted in Figs. 17 and 18 and in Table 6. The variation of crack-depth ratio, a/w, with time is given in Fig. 17 for the two transients, and good agreement is seen between measured data ((u~/w)~ = 0.165; = 0.278) and computed values ((af/w)s = 0.16; (udw)c = 0.26) at arrest when the (aflw)c crack length is uf. Figure 18 depicts the dynamic Ki vs u/w relations for the transients, and it indicates that the crack propagated into a rising Kr field for both run-arrest events. Table 6 compares selected results obtained from the quasi-static OCA/USA analyses and from the dynamic ADINASWIDAC analyses. The differences in computed Ki initiation values from OCAUSA (B: Ki, = 177.4 MPa*m”‘; C: Ki, = 254.3 MPa.m”‘) and from ADINASWIDAC

96

R. H. BRYAN

et al.

(B: K,, = 164.3 MPa.m”*; C: Krc = 234.8 MPa.m”*) are due in part to a relatively coarser radial refinement and to the slightly shallower initial crack depth of the dynamic finite-element model. The K, arrest values given in Table 6 for the dynamic analyses are obtained from the K1, fracture-toughness relation of eqn (4) evaluated at the temperature of the crack-tip location predicted by SWIDAC (interpolated from Table 5) at arrest. These K,, values (B: K1,(7’ = 161.4) = 230.8 MPa.m”*; C: K1,(T = 172.6) = 291.4 MPa.m”*) can be compared with the dynamic Kl values computed by SWIDAC in the time step immediately preceding crack arrest (B: Kl = 207.1 MPa.m”‘: C: Kl = 277.9 MPa.m”‘) and not given in Table 6. In summary, the comparisons presented in Table 6 support the perception that quasi-static methods and dynamic methods of fracture analysis should be in good agreement for the relatively short run-arrest events observed in PTSE- 1. CONCLUSIONS AND RECOMMENDATIONS The first pressurized-thermal-shock experiment is a basis for quantitative conclusions regarding initiation and arrest toughnesses for one reactor vessel grade steel. Results from test data indicate that the ASME Boiler and Pressure Vessel Code Section XI toughness relations are conservative relative to actual material characteristics. The experiment demonstrated that arrest toughness substantially above the 220 MPa.m “’ cutoff of Section XI could be realized. Furthermore, the highest PTSE-1 value of arrest occurred at a temperature above 30°K above the onset of the Charpy upper shelf. This is believed to be very close to the threshold temperature, above which cleavage fracture cannot persist. This result also suggests that the methods of linear elastic fracture mechanics have an important role in fracture evaluation at high (upper-shelf) temperatures. The PTSE-IA and -lB transients were a demonstration that simple warm prestressing (K’ < 0) strongly inhibits crack initiation. With allowance for uncertainty in the true Kl, values it was evident that Kl exceeded KI, during warm prestressing by 50-90%. In transient A, simple antiwarm prestressing (K, > 0) prevailed during two periods of 40 s and 60 s duration without crack initiation, although Kl exceeded K,, by 30-50%. Clearly simple antiwarm prestressing (K1 > 0) is not a sufficient condition to alleviate the effects of warm prestressing. A narrow band of ductile tearing formed ahead of the initial cleavage fracture. This was not unexpected, since analysis as well as prior intermediate vessel tests[l l-l 31 indicated the potential for stable tearing prior to cleavage. The complete absence of ductile tearing after crack arrest is not consistent with tearing analysis based on pretest data on tearing resistance. This result suggests that the data or the method of analysis or both are very conservative. The conclusions drawn from PTSE-I suggest that procedures used for evaluating overcooling accidents in pressurized-water reactors should realistically take into consideration the following fracture mechanisms that have been clearly demonstrated but not yet all generally accepted. Account should be taken for the inhibiting effect of simple warm prestressing. Furthermore, it is not premature to allow consideration of crack-arrest toughness values above the ceiling suggested in Section XI of the ASME Code. These two measures would make evaluations less conservative without being unrealistic. In a change toward conservatism, the phenomenon of ductile tearing below dynamic upper-shelf temperatures should be explicitly considered in vessel evaluations to ensure that the procedure is never nonconservative. This is because ductile tearing can precede the onset of cleavage if the value of toughness calculated from d\/EJ’, is less than Kl,. Some allowance for a decrease in initiation toughness with cracktip strain rate caused by crack-front motion may also be necessary under these circumstances. Viscoplastic constitutive models are presently being considered for this purpose. Elastodynamic analyses have shown good agreement between data and predictions for the short crack runs of these experiments, and static and dynamic analyses showed little difference. It is reasonable to expect dynamic effects to be more pronounced in longer crack-run events that could accompany axial extension of a short flaw in a pressure vessel. A&no&dgments-_The authors gratefully

acknowledge the essential contributions of H. A. Domian et al. of the Babcock & Wilcox Alliance Research Center; A. R. Rosenfield et al. of the Battelle Columbus Laboratories; J. W. Bryson, T. M. Cate, D. p. Edmonds, F. R. Gibson, R. W. McCulloch, and K. R. Thorns of the Oak Ridge National Laboratory; and D. A. Steiner? of the Computing and Telecommunications Division.

Pressurized-thermal-shock

experiment

PTSE-1

97

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