ML20151N782

Fracture Sensitivity Study of Girth Weld 6 Repaired Configuration,Indian Point Unit 2

ML20151N782
Person / Time
Site:	Indian Point
Issue date:	05/31/1988
From:	Bamford W, Lee Y WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML100330987	List: ML20151N782 ML20151N788
References
WCAP-11850, NUDOCS 8808090090
Download: ML20151N782 (24)

Text

WESTINGH USE CLASS 3 CUSTOMER DES!GNATED CISTRIBUTION WCAP-11850 FRACTURE SENSITIVITY STUDY OF GIRTH WELD NO. 6 REPAIRED CONFIGURATION INDIAN POINT UNIT 2 May 1988 W. H. Bamford Y. S. Lee Verified by: C8 '

/M C. B. Bond Approved by: .e M

5. S. Pa)dsamf,' Rah'ager Structytal Materials Engineering Although information contained in this report is nonproprietary, no distribu-tion shall be made outside Westinghouse or its licensees without the customer's approval.

WESTINGHOUSE ELECTRIC CORPORATION Power Systems Business Unit P.O. Box 355 Pittsburgh, Pennsylvania 15230 8808090090 880729 PDR ADOCK 03000247 O PNU _

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TABLE OF CONTENTS Section Title Page EXECUTIVE

SUMMARY

iii

1.0 INTRODUCTION

1-1 2.0 LOAD CONDITIONS, FRACTURE ANALYSIS METHODS, 2-1 AND MATERIAL PROPERTIES 2.1 Transients for the Steam Generator 2.2 Stress Intensity Factor Calculations 2.3 Fracture Toughness 2.4 Critical Flaw Size Determination 3.0 FRACTURE ANALYSIS RESULTS AND CONCLUSIONS 3-1

4.0 REFERENCES

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EXECUTIVE SUMARY In the fall of 1987 a series of surface flaws were found in the upper shell to cone weld of the Indian Point Unit 2 steam generators. These flaws were removed by grinding, appropriate evaluations performed, and the vessels returned to service.

During the discussions which took place with the Nuclear Regulatory Commission on this topic, a question was raised concerning the fracture toughness of the heat-affected zone (HAZ) region of the weld. Specifically, the high hardness l of the HAZ region (as determined from tests on boat samples) was suspected to be an indication of low fracture toughness. As a result of this concern this l

sensitivity study was undertaken, in conjunction with an experimental program (reported separately) to characterize the HAZ properties.

The sensitivity study reported here shows that the fracture toughness of the upper shell to cone weld region is sufficient to ensure its integrity during future operation, even with a very long surface flaw. This conclusion results from the fracture calculations reported here as well as the experimental findings of the companion program, m e. w e e m io $gj

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SECTION 1 INTRODUCTION The indications found in girth weld number six of the Indian Point Unit 2 steam generators have been removed by grinding. Boat samples removed during the repair process have shown that there are regions within the heat affected zone with relatively high hardness, and this has raised concerns relative to the integrity of the steam generator vessel.

This work has been carried out to investigate the integrity of the vessel, and to determine the sensitivity of the repaired vessel girth weld region to the presence of cracks, even though it is not expected that cracks would reinitiate in this region. Specifically, analyses were done to determine what level of fracture toughness or RT NDT w uld be required to maintain the integrity of the vessel for a postulated surface flaw.

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SECTION 2 LOAD CONDITIONS, FRACTURE ANALYSIS METHODS AND MATERIAL PROPERTIES 2.1 TRANSIENTS FOR THE STEAM GENERATOR The design transients for the Indian Point Unit 2 steam generators are listed by umbrella groupings in table 2-1. The critical flaw sizes under normal operating conditions, or under faulted conditions, and the stress intensity factors, K g, are a function of the stresses caused by these transients at the cross-section where the flaw of interest is located, and the material properties. Therefore, the first step for a fracture evaluation is to determine the appropriate limiting load conditions for the location of interest.

For the region of interest, the upper chell to cone weld, the full range of design transients was considered. Transients such as pressure tests, including both hydrostatic and leakage tests, can be controlled by setting the test temperature. Therefore, the operational transients were considered in one analysis and a separate determination was made of the toughness required for the pressure tests. On this basis, the governing operational transient was found to be the reactor trip condition, which is even more severe than emergency and faulted transients in the steam generator.

2.2 STRESS INTENSITY FACTOR CALCULATIONS One of the key elements of the critical flaw size calculations is the determination of the driving force or stress intensity factor (K ). g This was done using expressicns available from the literature. In all cases the stress intensity factor for the critical flaw size calculations utilized a representation of the actual stress profile rather than a linearization. This was necessary to provide the most accurate determination possible of the critical flaw size, and is particularly important where the stress profile is generally nonlinear and often very steep. The stress profile was represented by a cubic polynomial:

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o(x) = 0A + A1{+A2({} +A3(k) where x is the coordinate distance into the wall t = wall thickness a = stress perpendicular to the plane of the crack In the study of sensitivity to the presence of flaws (section 3) three flaw shapes were used, near-continuous (a/t = 0.05) semielliptical with length six times the depth (a/t = 0.167) and semi circular (a/t = 0.5). As will be seen in Section 3, the study covers the full range of shapes between these values. All the postulated flaws were circumferentially oriented, and were presumed to be in the region of maximum griding depth.

For the surface flaw with length six times its depth (a/t a 0.167),the stress intensity factor expression of McGowan and Raymund [2] was used. The stress intensity factor K; (4) can be calculated anywhere along the crack front, where & is the angular position, as defined in figure 2-1. The point of maximum crack depth is represented by 9 = 0. The following expression is used for calculating K; (4):

0.5 2 1/4 K;(#) = ((] (cos2 , , a, sin 29) (A,H,+ 2 AyH1 c

l

+

A2 N2+ A3 H) 3 l The magnification factors H 0

(*)' H 1(#), H (9) 2 and H 3(4) were obtained by the procedure outlined in reference [2].

l The stress intensity factor calculation for a semi-circular surface flaw, (a/t = 0.5) was carried out using the expressions developed by Raju and Newman (3). Their expression utilizes the same cubic representation of the stress profile and gives precisely the same result as the expression of McGowan and Raymund for the flaw with a/t = 0.167, and the form of the 2819:4440s-cs244410 22

equation is similar to that of McGowan and Raymund above. The stress intensity factor expression used for a very long surface flaw (a/t = 0.05) was also carried out using the expression of Raju and Newman (3).

2.3 FRACTURE TOUGHNESS The other key element in the determination of critical flaw sizes is the fracture toughness of the material. The fracture toughness has been taken directly from the reference curves of Appendix A,Section XI. In the transition temperature region, these curves can be represented by the following equations:

KIc = 33.2 + 2.806 exp. [0.02 (T-RTNDT + 100'F))

Kg , = 26.8 + 1.233 exp. [0.0145 (T-RTNDT + 160*F)]

where K Ic and K g , are in ksi / in.

i The upper shelf temperature regime requires utilization of a shelf toughness which is not specified in the ASME Code. A value of 200 ksi/in has been used here. This value is consistent with general practice in such evaluations, as shown for example in reference (4), which provides the l background and technical basis of Appendix A of Section XI.

The fracture toughness of steam generator materials has been well character-ized, since A3028 steel was used to fabricate many reactor vessels as well.

This material was used in developing and verifying the reference toughness curves of the ASME code. Fracture toughness tests were conducted on base metal, weldments, and heat-affected zones, and were all found to be bounded by the ASME gK , curve for dynamic and arrest tests, and the KIc curve for static tests.

The other key element in the determination of the fracture toughness is the value of RTNDT, which is a parameter determined from Charpy V-notch and drop weight tests, me.co. ems to 2-3 i .. . _ - - - -. - -. - - -

To allow determination of RT NDT f r the upper shell and cone materials, a compilation was made of the properties listed on the original material test certificates. The materials used in the steam generators were tested after a post-weld heat treatment cycle of 1050-1150'F for 18 to 28 hours3.240741e-4 days <br />0.00778 hours <br />4.62963e-5 weeks <br />1.0654e-5 months <br />, as shown in table 2-2. The Charpy impact properties of these materials are listed in tables 2-3 and 2-4.

The U.S. Nuclear Regulatory Commission has established guidelines for estimating the value of RTNDT from Charpy properties in their Standard Review Plan [5). Review of table 2-3 shows that in general the materials in the shell and cone region have excellent Charpy properties, and therefore the value of RT NDT is equal to the test temperature, which is 10*F for all the base materials and 30*F for the welds.

Concern has been expressed relative to the RTNDT estimation procedures being applied to the heat-affected zone material, and to answer this question completely further experimental work has been carried out (6). The properties of the girth weld were expected to be quite good. Charpy tests made from the plug removed from the Indian Point Unit 3 girth weld number six showed very good results, as shown in table 2-5. The maximum hardness of the heat-affected zone of this plug was found to be very similar to that of boat samples removed from unit 2 during the recent investigation.

Results of the experimental study of the high hardness HAZ material (6) showed that RT NDT was much lower than the original estimate of 30*F. Drawing a lower bound curve under the Charpy data of figure 3.53 of reference 6, the 50 f t-lb energy level was reached at 28'F, as reproduced here in figure 2-2.

Using the procedure of the ASME Code Section III, paragraph NB-2300, the RT NDT value is 60'F below this 50 ft-1b temperature, or -32'F. This is believed to be a reliable estimate of RT NDT f r the material, even though drop weight tests were not performed, since the Charpy curve is very steep and the transition is well-defined. The Charpy results from the experimental program [6] at room temperature (52-64 ft-lb) are similar to the results of the tosts from the Indian Point Unit 3 material reported in table 2-5 and we,wes-osms to g.4

discussed above. The results of (6) are also comparable to the original weld qualification test results for Indian Point 2, in table 2-4, since the HAZ results at 0*F were approximately 50 ft-lb.

Once the value of RTNDT is established, the reference toughness curves of the ASME Code discussed above may be used directly, since the materials are SA302 grade B which has a minimum specified yield strength of 50 k:1. These toughness curves were used in the critical flaw size determinations to be discussed below.

2.4 CRITICAL FLAW SIZE DETERMINATION, AND SENSITIVITY APPROACH The applied stress intensity factor (K )y and the material fracture toughness values (Kg , and KIc) were used to determine the allowable flaw size values i

used to construct the handbook charts. For this study, the critical flaw size was determined as the depth at which the applied stress intensity factor Kg exceeds the fracture toughness K Ic*

l In this study, since the fracture toughness of the material is in question the critical flaw size determination method was used in reverse, to decide what fracture toughness is necessary to maintain the integrity of the steam generator girth weld region in the presence of a postulated flaw. The results of these calculations will be discussed in Section 3.

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m e. c o. e n n io 2-5

TABLE 2-1 TRANSIENT GROUPING FOR FATIGUE CRACK GROWTH ANALYSIS [7]

Transient Cycles 1 Cold Shutdown 200 2 No Load 200 3 100% Power (Plant Load / Unload) 14500 4 Small Step Load Decrease 2000 5 Steady-State Fluctuations (+) 1.0E+06 6 Steady-State Fluctuations (-) 1.0E+06 7 Large Step Load Decrease 2200 Small Step Load Decrease (2000) 8 Loss of Power 205 LossofLoad(40)

Loss of Flow (80)

Secondary Side Leak Test (5) 9 Reactor Trip 400 10 Feedwater Cycling 25000 11 Secondary Hydrotest (Init.) 1 12 Secondary Hydrotest (Subs.) 50 m e.w..-mm in 2-6

i TABLE 2-2 POSTWELD HEAT TREATMENT OF UPPER SHELL - CONE WELDS INDIAN POINT UNIT 2 Heatup to 1050*F > 7 hours8.101852e-5 days <br />0.00194 hours <br />1.157407e-5 weeks <br />2.6635e-6 months <br /> Soak at 1050-1150'F SG #1 - 26 hrs SG #2 - 27 hrs SG #3 - 18 hrs SG #4 - 28 hrs Cooldown in Air l

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TABLE 2-3 MATERIAL PROPERTIES OF UPPER SHELL-CONE REGION INDIAN POINT UNIT 2 Charpy Values (10'F)

Location Material Type (ft-lb) RT NDT Cone traterials, SG #1 heat A0058-3 SA 302-56 Gr B 64,67,63 10*F heat A0991-2 SA 302-56 Gr B 58,62,50 10*F heat A0042-2 SA 302-56 Gr B 61,74,77 10'F heat A0042-4 SA 302-56 Gr B 55,64,60 10*F Upper shell materials, SG #1 heat 85012-2 SA 302-56 Gr B 70,67,68 10*F heat 85012-3 SA 302-56 Gr B 73,83,48 10'F heat A0310-4 SA 302-56 Gr B 105,95,104 10*F heat B5012-1 SA 302-56 Gr B 78,78,74 10*F Cone materials, SG #2 heat C1108-4 SA 302-56 Gr B 124,120,125 10'F heat A9941-1 SA 302-56 Gr B 95,81,90 10*F heat A0091-1 SA 302-56 Gr B 72,60,59 10*F heat C1108-2 SA 302-5G Gr B 85,68,97 10*F heat A0042-4 SA 302-56 Gr B 55,64,60 10'F Upper shell materials

. SG #2

, heat A0126-1 SA 302-56 Gr B 60,80,82 10'F heat A0126-2 SA 302-56 Gr B 86,47,57 10*F heat A0126-3 SA 302-56 Gr B 105,97,91 10'F heat A0126-4 SA 302-56 Gr B 79,66,64 10*F l

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TABLE 2-3(continued)

MATERIAL PROPERTIES OF UPPER SHELL-CONE REGION INDIAN POINT UNIT 2 Charpy Values (10*F)

Location Material Type (ft-lb) RT NDT Cone materials, SG #3 B5010-2 SA 302-56 Gr B 81,143,85 10*F B5010-3 SA 302-56 Gr B 81,143,85 10*F C1108-1 SA 302-56 Gr B 81,95,76 10'F B4873-5 SA 302-56 Gr B 86,121,85 10*F B5010-1 SA 302-56 Gr B 93,73,113 10'F Upper Shell Materials, SG #3 A0902-4 SA 302-56 Gr B 70,67,67 10'F A0877-2 SA 302-56 Gr B 77,89,87 10*F A0877-1 SA 302-56 Gr B 94,70,55 10*F A0872-3 SA 302-56 Gr B 79,83,83 10'F Cone Materials, SG #4 heat C1488-4 SA 302-56 Gr B 82,88,89 10'F heat C1488-3 SA 302-56 Gr B 81,77,74 10*F heat B5387-1 SA 302-56 Gr B 84,84,88 10*F heat B5387-2 SA 302-56 Gr B 90,98,110 10*F Upper Shell Materials, SG #4 A0877-4 SA 302-56 Gr B 83,92,95 10*F A0902-2 SA 302-56 Gr B 51,78,45 10*F A0877-3 SA 302-56 Gr B 72,96,75 10*F B5973-2 SA 302-56 Gr B 66,74,70 10'F l

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TABLE 2-4 MECHANICAL PROPERTIES AND TOUGHNESS OF WELD QUALIFICATION SPECIMENS Charpy Yield Tensile Energy Strength Strength (ft-lb)

Heat No. (ksi) (ksi) at 10'F RT NDT B5012 67.7 83.4 30/37/29 30 50/53/80 10 A0265 70.9 86.9 38/30/47 30 37/111/48 10 C1108 64.6 83.4 41/40/41 30 43/48/98 10 84873 68.0 85.0 38/37/42 30 80/33/83 10

• • '"""" 2-10

TABLE 2-5 CHARPY RESULTS FROM MATERIAL REMOVED FROM THE INDIAN POINT UNIT 3 STEAM GENERATOR, TESTED AT 76*F Lateral Expansion Notch Location Cv (ft lbs) (Mils)  % Shear Trans cone 79 63 85 Trars cone 69 58 80 Upper shell 77 67 95 Upper shell 75 64 95 Weld cone 60 51 60 Repair weld 111 85 95 Crown to cone HAZ 83,74 59,59 90,85 Crown to upper shell 83,71 61,44 95,60 l

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SECTION 3 FRACTURE ANALYSIS RESULTS AND CONCLUSIONS Fracture analyses were carried out for both the reactor trip transient and a

pressure test, to determine the sensitivity of the girth weld in its repaired configuration to the presence of cracks. The worst case repaired configura-1 tion was used, which incorporates a grinding depth of 1.1 inches [7].

The results of a typical pressure test are presented in figure 3-1, for three j flaw shapes. The 1000 psi pressure was chosen as representative of the secondary side operational pressure tests, which generally range from 750-1085

psi. It can be seen from the figure that the girth weld is not very sensitive to the presence of flaws under a pressure loading, since even the most elongated flaw has a stress intensity factor of less than 100 ksi/in for a flaw one-half inch deep beyond the 1.10 inch grinding depth. The assumption here is that RT NDT is a maximum of 30'F. This assumption was proven conservative by the Charpy tests of referen

e (6). The pressure tests could be accomplished with no difficulty at 80*F (RTNDT + 50*F) for a 1000 psi test, and at 60'F (RTNDT + 30*F) for a 770 psi pressure test. These values were obtained from use of the K Ic curve from section XI, which is reproduced in figure 3-3, and a reference flaw depth of one half inch deep, beyond the 1.1 inch grinding depth. The assumption here was that RTNDT is no higher than 30*F. Since the tests [6] have shown that RT<DT i is much lower than 30'F, the base metal properties become governing (RTNDT = 10*F) and the pressure tests can be accomplished'at room temperature with a generous safety margin.

The results of the calculation for the reactor trip transient are shown in figure 3-2, for three different flaw shapes. The applied stress inteesity factor is highest for the longest flaw (a/t = 0.05), as expected. Aftar increasing rather steeply with crack depth, the sensitivity to flaws decreases, as seen by the decreasing slope. The stress intensity factor for a flaw one half inch deep was found to be 160 ksi/in for the most elongated flaw shape, and 132 ksi / in for a six-to-one elliptical flaw (a/1 = 0.167).

me. .couune 31

The implication of this finding is that a half-inch deep surface flaw would remain unaffected during the most severe operational transient (reactor trip) as long as the fracture toughness exceeds the aforementioned values. The minimum inside surface temperature at the girth weld during the reactor trip was found to be 260'F, so this toughness requirement trar; slates into a required RT NDT f 190*F or less for ;e six-to-one elliptical flaw (a/t =

0.167). Likewise, a maximum RTNDT f 173'F is obtained for the very long flaw (a/t = 0.05). Since the RT NDT is in the range of - *F [6), a very large margin of safety can be shown to exist.

The above assessment is very conservative, because it is based on the assumption that the ground region of 1.1 inches deep extends around the entire circumference of the steam generator girth weld. The grind depth of 1.1 inches deep is actually very localized. Only at one location is the actual i maximum grinding depth of 1.07 inch reached. Further, the assumption of a flaw 0.5 inch deep is very conservative, since the girth weld area has been carefully ground to be defect-free, as proven by magnetic particle exams.

Lower bound toughness curves have been used in the fracture analysis, which adds further conservatism.

As a result, even with a very long, 0.5 inch deep surface flaw in the repaired girth weld region, the structural integrity of the upper shell to cone girth welds in the Indian Point 2 steam generators will not be affected, m e.co..oma in 3-2

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SECTION 4 REFERENCES

1. ASME Code Section XI, "Rules for Inservice Inspection of Nuclear Power Plant Components."

2. McGowan, J. J. and Raymund, M., "Stress Intensity Factor Solutions for Internal Longitudinal Semi-elliptic Surface Flaw in a Cylinder Under Arbitrary Loading", ASTM STP 677, 1979, pp. 365-380.

3. Newman, J. C. Jr. and Raju, I. S., "Stress Intensity Factors for Internal Surface Cracks in Cylindrical Pressure Vessels", ASME Trans., Journal of Pressure Vessel Technology, Vol. 102, 1980, pp. 342-346.

4. Marston, ' U., et. al.,Section XI Flaw Evaluation: Technical Basis and Sample . . h.ns, EPRI NP 7195R, August 1978.

)

5. U.S. Nuclear Regulatory Commission Standard Review Plan, NUREG-0800.

6. Rao, G. V., "Charpy Toughness and Brittle Transition Temperature Characterization of the HAZ High Hardness Zone of the IP-2 Steam Generator Girth Weld by Gleeble Weld Thermal Cycle Simulation,"

Westinghouse Electric Report WCAP-11831, May 1988.

7. "Indian Point Unit 2 - Steam Generator Girth Weld Repair," Westinghouse

- Electric Report STD 7.3.4-8105, February 1988 (to be published as WCAP-11730). [ Westinghouse Proprietary Class 2]

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