Text

Ja FitzPatrick Nuclear Power Plant Core Shroud Vertical & Top Guide Support Ring Radial Weld Flaw Evaluation Screening Criteria
	ML20081K079
Person / Time
Site:	FitzPatrick
Issue date:	03/31/1995
From:	Marrone N, Mccurdy H, Swanner C; MPR ASSOCIATES, INC.
To:	;
Shared Package
ML20081K059	List: JPN-95-016, Forwards Repts Re Core Shroud Insp & Preemptive Repair Performed During 1994-95 Refueling Outage of Plant, Nonproprietary Version of XM-19 Matl Constant Extension Rate Testing Rept Requested by NRC; ML20081K069; ML20081K079;
References
	MPR-1580, MPR-1580-R, MPR-1580-R00, NUDOCS 9503280391
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{{#Wiki_filter:,. - ,o f.* ATTACHMENT 2 to JPN-95-016 i I r MPR Report 1580 James A. FitzPatrick Nuclear Power Plant Core Shroud Vertical and Top Guide Support Ring Radial Weld Flaw Evaluation Screening Criteria March 1995 t i New York Power Authority JAMES A. FITZPATRICK NUCLEAR POWER PLANT Docket No. 50-333 DPR-59 i 9503280391 950321 PDR ADOCK 05000333 p PDR

EMPR ASSOCI ATES INC ENG4NEERS Revision 0 March 1995 b James A. FitzPatrick Nuclear Power Plant Core Shroud Vertical and Top Guide Support Ring Radial Weld Flaw Evaluation Screening Criteria I a Prepared For New York Power Authority 123 Main Street White Plains, NY 10601

7._ MMPR A S S O CI AT E S IN C. 3 ENG1NEERS James A. FitzPatrick Nuclear Power Plant Core Shroud Vertical and Top Guide Support Ring Radial Weld Flaw Evaluation Screening Criteria 1 MPR-1580 Revision 0 March 1995 Principal Contributors H. Wm. McCurdy N. J. Marrone C. B. Swanner S. J. Weems Prepared for New York Power Authority i 123 Main Street White Plains, NY 10601 320 KING STREET ALE X ANDRI A. VA 22314-3238 703 510-0200 FAX' 703 510-0224

t .ggpg-A ssOCI ATES IN C. L fNGiNEER5 j i CONTENTS Section Eage i 1 INTRODUCTION 1-1 1.1 Purpose 1-1 1.2 Background and Scope 1-1 i 2

SUMMARY

2-1 2.1 Screening Criteria Design Basis Considerations 2-1 l 2.2 Screening Criteria 2-2 j 2.2.1 Screening Criteria Applicability 2-2 2.2.2 Flaw Characterization 2-3 i 2.2.3 Vertical Weld Criteria 2-3. f 2.2.4 Radial Weld Criteria 2-3 2.3 Additional Capability of Tie-Rod Repair Assemblies 2-4 3 DISCUSSION 3-1 3.1 Flaw Evaluation Approach 3-1 { 3.1.1 Overall Approach 3-1 3.1.2 Assumptions 3-2 { 3.2 Flaw Characterization 3-2 t 3.3 Venical Weld Evaluation Approach 3-3 l 3.4 Radial Weld Evaluation Approach 3-5 4 REFERENCES 4-1 i 5 TABLES AND FIGURES 5-1 l t ii l

r t CONTENTE (Continued) i Section East APPENDICES APPENDIX A MPR Calculation No. 2919401-801," Required Intact Vertical Weld Area Based Upon the Limiting Applied Loads", Revision 0 A-1 APPENDIX B MPR Calculation No. 2919401-803," Initial Acceptance Criteria for Radial Welds with Indications", Revision 0 B-1 i APPENDIX C MPR Calculation No. 2919401-804," Critical Flaw Size for Indications in Vertical Welds of the FitzPatrick Shroud", i Revision 0 C-1 APPENDIX D MPR Calculation No. 2919401-805," Critical Flaw Size Revision 0 D-1 for Indications in Circumferential Welds H2 (or H3)," i 1 --l

MMPR A S SO CI ATE S INC. ENGINEERS Section 1 INTRODUCTION 1 1.1 PURPOSE The purpose of this report is to provide initial screening criteria for the evaluation of flaws which may be identified in the vertical and top guide support ring radial welds in the James A. FitzPatrick Nuclear Power Plant core shroud. The screening criteria apply to the FitzPatrick shroud with the core shroud repair assemblies installed per JAF Modification No. F1-94-036. i

1.2 BACKGROUND

AND SCOPE The core shroud at FitzPatrick was fabricated from a number of segments joined by circumferential, vertical and radial welds. The location of these welds is shown in l Figures 1,2 and 3. Stainless steel circumferential welds H1 through H7 join the cylindrical segments which make up the shroud. In response to industry wide concerns regarding potential cracking of these welds (and the associated heat-affected zones), a shroud repair was developed and installed to structurally replace any or all of these circumferential welds. The repair installed to address potential degradation of circumferential welds was not specifically designed to replace the shroud vertical and radial welds. The ability of the shroud to satisfy its functional requirements with the repair installed and failed vertical or radial welds was not addressed in the circumferential weld repair design because no significant cracking of these welds has been found thus far. However, as a contingency, the New York Power Authority (NYPA) has decided to develop conservative screening criteria for the initial evaluation of indications. I i [ B k 1-1

1 1 /'MPR A S SOCI Af f 5 IN C. ENGINEERS Section 2 i

SUMMARY

l l 2.1 SCREENING CRITERIA DESIGN BASIS CONSIDERATIONS The vertical and radial shroud welds allow the shroud to carry differential pressure and lateral loads as ring or cylindrical sections. Failure of the vertical or radial welds could potentially affect the following: Shroud stresses, I Shroud displacements, and Leakage across the shroud through the degraded welds. i The screening criteria developed herein, and summarized in Table 1, are conservatively established so that the vertical and radial welds will not fail. Accordingly, there will be no significant effect on shroud stresses, displacements or leakage. No credit is taken in the development of the screening criteria for the ability of the radial j restraints to maintain the basic shape of the shroud. (See Section 2.3 for a brief discussion on this capability.) Instead it is conservatively assumed that sufficient load carrying capability of the vertical and radial welds is required to maintain shroud stresses within allowables. In establishing the vertical weld screening criteria, it is also assumed that the circumferential welds have failed. With respect to radial welds, it is assumed that radial welds fail completely and the screening criteria establishes requirements on the adjacent circumferential welds to ensure that the load carrying capability of the radial weld is maintained. A number of other conservative assumptions are made in the development of the screening criteria. These assumptions are summarized in Section 3.1. t The screening criteria were developed using both Limit load and Linear Elastic Fracture Mechanics (LEFM) evaluations of the stresses in the shroud to ensure that the required load l carrying capability of the welds is maintained. The minimum length ofintact weld required to carry the applied loads is calculated based on Limit Load analysis with stresses limited to the design stress allowables from the ASME Code (Reference 3). Flaws in close proximity are suitably characterized per ASME Section XI criteria. The potential growth of the flaw is addressed by (1) increasing the calculated minimum weld size by a conservative estimate of the potential crack growth over the next operating cycle, and (2) limiting the maximum flaw size to the critical flaw size conservatively calculated by LEFM techniques. 2-1

~ i-b It should be noted that the load cases and allowable stresses addressed in the shroud repair ) design report (R.Lw,cc 4) are considered in the development of the screening criteria. i Accordingly, all pertinent requirements of the BWROG Vessel Internals Project Specification (Reference 6) and the ASME Code are met. 2.2 SCREENING CRITERIA i -i 2.2.1 Screening Criteria Applicability i 2.2.1.1 Ver+ leal==d Ton G=td= EW D=dt=1 WM-Screening criteria are provided here for the evaluation of indications in the stainless steel vertical and top guide support ring radial welds in the FitzPatrick core shroud. Specifically, criteria are provided for vertical ] welds SV2, SV4 through SV6, SV8, and SV9, and radial welds in the top guide support ring i (SV3A and SV3F). 'Ihis screening criteria is applicable to the FitzPatrick core shroud with ) the shroud repair (i.e. tie rods) installed per JAF Modification No. F1-94-036. The criteria i i also applies to both the current operating and uprated operating conditions addressed in Reference 4. ) 1 As discussed in Section 3, a number of conservative assumptions are made in the development of the screening criteria. As a result, shroud cracking which is more severe j than allowed by the screening criteria may be acce;&ble. Since the screening evaluation represents the first step in the evaluation of inspection results, the criteria are by definition conservative. Additional evaluations can be performed for any indications which exceed the screening criteria and the indications may be quite acceptable based on such a focused evaluation. 2.2.1.2 Other Radial Welds-As discussed above, this study considered only the radial welds in the top guide support ring. Although not explicitly considered in this study, it is expected that substantial cracking in the radial welds in the other shroud rings (see Figures 1 and 3) would be arcar*=hle based on the following: The upper shroud flange ring between circumferential weld H1 and the top of the shroud (contains radial welds SVI A through SV1F) is connected to the shroud separator head assembly by guides posts and 36 preloaded shroud separator bolts. This flange ring is also held down by the tie-rod preload and its radial motion is limited by the tie-rod radial restraints. As a result, the flange ring segments are expected to be constrained from significant displacement, even with substantial cracking of its radial welds. The lower core plate assembly support ring between circumferential welds H6A and H6B (contains radial welds SV7A through SV7F) is connected to the core plate assembly by 72 preloaded bolts. As a result, the support ring segments are constrained by: (1) the preload between the ring and the core plate assembly, (2) the tie-rod preload, and (3) the tie rod radial restraints. Therefore, the support 2-2

x L ring is W to be constrained from significant displamments, even with substantial cracking of its radial welds. Based on the above, it is W that a detaUed evaluation would show that the structural integrity of the shroud for design loads could be maintained with significant flaws in the radial welds in the upper shroud flange and core assembly support rings. 2.2.2 Flaw Characterization Flaws in close proximity to each other may effectively act as a single larger flaw due to crack growth in-service and stress concentration effects. lAs a result, indications in close. proximity to each other are treated as a single larger flaw. 'Ihe effective length of the combined flaw is determined using pid.w.t flaw proximity rules. The effective flaw length is compared to the screening criteria to determine the acwpi.1,3ity of a given flaw. The flaw proximity rules to be used with the screening criteria provided here are the same as those developed for the screening evaluation of indications in unrepaired shrouds (see Sections 2 and 3.2 of Reference 1). The screening criteria provided below conservatively assume that flaws in vertical welds are through-wall over the effective flaw length. The criteria also assume that flaws identified in the radial welds in the top guide support ring are through-wall over the entire weld length. i 2.2.3 Ventical Weld Criteria i The screening criteria for indications in the vertical welds of each shroud segimt are shown - in Table 1. An -aptahle vertical weld for a given shroud segment must have the tabulated minimum length of intact weld and the maximum effective flaw size (for any single flaws within the weld in question) must be within the maximum flaw size listed in Table 1. In j addition, if substantial cracking of the vertical welds is identified (i.e., the sum of the length of all failed vertical welds is greater than about 210 inches), then the impact of shroud i . leakage on plant operation should be evaluated. The minimum length of intact weld is calculated based on a limit load analysis, and the maximum flaw size is determined using LEFM. Because of the inherent ductility of the austenitic stainless steel, an LEFM analysis provides very conservative results. Further, the j maximum flaw size is determined for the limiting location and highest service loading l condition. The resulting critical flaw size is applied as the maximum flaw length for any shroud vertical welds. Consequently, the actual maximum flaw length before failure is much closer to the total weld length minus the minimum intact weld length predicted by the limit load analysis, than the maximum flaw size calculated by the LEFM analysis. 2.2.4 Radial Weld Criteria The radial welds in the top guide support ring (welds SV3A through SV3F) are only necessary if both circumferential welds H2 and H3 have failed. If H2 gr H3 are intact, then j any indications identified in the radial welds of the top guide support ring are acceptable. j 2-3 l .m_ .,-4 4

2 .2 m. e If significant indications are identified in both H2 and H3, then the initial =@nce criterion for indications in the radial welds of the top guide support ring is as follows: I With an unlimited flaw length in H3, a maximum single flaw length in circumferential weld H2 of 115 inches is permitted and H2 must be intact for a minimum of 20 inches directly adjacent to both sides of the radial weld containing t the flaw, et With an unlimited flaw length in H2, a maximum single flaw length in circumferential weld H3 of 115 inches is permitted and H3 must be intact for a minimum of 20 inches directly adjacent to both sides of the radial we'd containing the flaw. 2.3 ADDITIONAL CAPABILITY OF TIE-ROD REPAIR ASSEMBLTES As indicated in Figures 6A through 6E, the tie-rod assemblies installed to address potential cracking of circumferential welds and the other existing reactor internals provide a number of l lateral supports for all sections of the shroud. By providing lateral support to the shroud at various elevations and azimuthal locations the repair provides an alternate means of maintaining the basic shape of the shroud and meeting displacement limits, for all design loads, with the failure of many of the circumferential and vertical welds. In addition, the design is such that more radial restraints could be added at some later date to provide further capability to accommodate cracking of vertical welds, if desired. The existing tie-rod design also permits changes to the elevation and location of the lateral supports which are on the replaceable sleeves which surround the bracket. With this approech, all plant safety considerations could likely be met, even with the complete failure of many of the vertical welds. However, the effects of shroud leakage on plant operation would need to be evaluated. j i a 2-4 -~

s Table 1 James A. FitzPatrick Nuclear Power Plant i Shroud Vertical Weld Inspection Acceptance Criteria Cylindrical Shroud Total Required length Shroud Vertical Weld ofIntact Weld, Segment Weld length, in. in. (Note 2) Additional Requirements

- - I' #
H1-H2 32.0 14 SV2B SV4A 36.0 16

. Note 3 H3-H4 tes 1,4 aM 5 SV4B 36.0 16 + Note 6 b + Notes 1,4 and 5 H4-H5 98.1 11 SV5B + Note 6 tes 1,4 and 5 H5-H6A 35.9 16 SV6B + Note 3 SV8A Notes 1,4 and 5 H6B-H7 SV8B 52.8 13 SV8C r H7-H8 SV9B 16.0 6 7 SV9C Notes: 1. The proximity rules described in Sections 2 and 3.2 of Reference 1 (GENE-523-154-1093) should be used to characterize all indications identified in weld inspections. 2. De required intact weld length is based on a limit load evaluation of the shroud repair design loads and the ASME i Code design allowable stresses. Appropriate allowances for crack growth during one 2 year fuel cycle are included. l For longer operating periods, the muumum required weld length should be increased by 0.876 in for each additional 2 year operating period. j 3. The required intact weld length must be along the lower porti3n of the weld. This requirement tesults from the need to preserve shroud load-carrying capability in the vicinity of the shroud repair radial supports; as shown in Figure 2, this weld is within about 5" of a radial support. l 4 4. The maximum allowable sire of any single flaw in the weld is 22 inches. This limit is based on a very con:ervative Linear Elastic Fracture Mechanics (1.EFM) evaluation for the limiting location with the highest service loading condition. Appropriate allowances for crack growth during one 2 year fuel cycle are included. For longer operating periods, the maximum allowable flaw size should be reduced by 0.876 in for each additional 2 year operating period. l 5. If substantial cracking of the vertical welds (Lc., the sum of the length of all failed vertical welds is greater than about 210 inches) is identified, then the impact of shroud leakage on plant operation should be evaluated. 6. Distortion or "ovalling" of this shroud segment is not prevented by the three radial restraints provided at this elevation. The required weld length results form the need to carry the resulting bending loads around the shroud. 7. The circumferential weld at H8 is not subject to the same failure mechanism (i.e., intergranular stress corrosion cracking) as the other circumferential welds. i 1 2-5 (

IBMPR ASSOCIATES INC ENGINEER $ Section 3 DISCUSSION 3.1 FLAW EVALUATION APPROACH 3.1.1 Overall Anoroach 4 There are several alternative approaches to evaluating the acceptability of flaws in shroud vertical and radial welds. These alternatives include but are not limited to the following approaches: " Failed Weld Approach" - The ability of the shroud to perform its functional requirements is evaluated assuming that the welds have failed; the basic shroud configuration is maintained by the radial support provided by vessel internals (e.g., the shroud radial supports installed with the repair for circumferential weld failures and the jet pump supports). For this approach, the effects of leakage must be evaluated due to potential opening up of failed welds. " Intact Weld Approach" - This approach assumes the weld ioints do not fail and that sufficient weld load carrying capability is required for the shroud to perform its functional requirements (i.e., maintain the geometry of the shroud and limit leakage). The acceptability of flaws is evaluated based on a combination of factors (e.g., stresses in the remaining weld, crack growth and stabihty). With regard to the " Failed Weld Approach", the approach would likely meet the design basis requirements with regard to shroud stresses, shroud displacements, and leakage across the shroud following a design basis accident event, i.e., all plant safety concerns and considerations would be satisfied. However, the shroud leakage with the failed horizontal and vertical welds, if such failures are extensive, could be greater than allowable for normal plant operation conditions and plant shutdown could be required. The screening criteria developed here are based on the Intact Weld Approach. In addition, as discussed below a number of conservative assumptions are made in the implementation of this approach for screening criteria development. As a result, shroud cracking which is more severe than allowed by the screening criteria may be acceptable. Additional evaluations can be performed to evaluate the acceptability of any such indications. 3-1

9 3.1.2 Assumotions As discussed above, the screening criteria developed here assume that at least some weld load carrying capability is required for the shroud to perform its functional requirements. The major assumptions used in implementing this evaluation approach are listed below: The minimum required weld size is based on a limit load analyses in which the stresses in the welds are limited to ASME Code design stress allowables (Reference 3). The maximum permitted flaw size is conservatively estimated using Linear Elastic Fracture Mechanics (LEFM) approach. (Note that Section XI of the ASME Code requires only a limit load analysis for evaluation of indications in austenitic piping (Reference 5)). Accordingly, this is not required per the ASME Code, however, for completeness, it is included herein. Potential crack growth over a two year operating cycle is considered in determining both minimum required weld size and the maximum permitted flaw size. A bounding crack growth rate of 5 x 10-5 in/hr is used which is consistent with the crack growth rate used in the screening criteria for an unrepaired shroud (Refereace 1). Flaws are assumed to be through-wall over the effective flaw length. The screening criteria for vertical weld failures assumes that all circumferential welds have failed. The tie-rod assemblies for repairing circumferential welds have been installed per JAF Modification No. F1-94-036. The evaluations are performed using the limiting tie rod and seismic restraint loads on the shroud from the shroud repair design report (Reference 4, Appendix G-1). l The differential pressure across the shroud is based on the reactor core differential i pressures for uprated flow conditions (Reference 4, Appendix G-3-see Table 2). Additional assumptions regarding the evaluation of leakage across the shroud are discussed in Section 3.3. 3.2 FLAW CHARACTERIZATION As a result of crack growth in-service and stress concentration effects, flaws in close proximity to each other effectively act as a single flaw. Consequently, proximity rules must be applied to characterize the effective flaw length of flaws which are in close proximity to each other. 3-2

Flaws in vertical welds are assumed to be through-wall over the indicated flaw length. The flaw proximity rules to be used with the screening criteria provided here for flaws in the vertical shroud welds are the same as those developed for the screening evaluation of indications in unrepaired shrouds (see Sections 2 and 3.2 of Reference 1). These rules consider crack growth in conjunction with the ASME Code flaw proximity rules for flaw characterization provided in Code Subsection IWA-3300 (Reference 2). If indications are identified in the radial welds in the top guide support ring, the flaw is conservatively assumed to be through-wall over the entire weld length (i.e., the weld is assumed to fail). 3.3 VERTICAL WELD EVALUATION APPROACII The vertical welds react the hoop stress in the shroud caused by the differential pressure between the reactor core and downcomer annulus. During seismic and recirculation-line break events, the lateral load on the shroud is also reacted through the vertical welds. A more detailed discussion of loadings and load paths through the shroud vertical welds is provided in Appendix A. Minimum Reauired Weld Size. The required weld size is determined using limit load techniques. In this approach, the weld area required to provide stresses below the ASME Code design allowable stress level is determined. The required intact weld length for a through-wall crack is calculated in Appendix A. The potential growth of the crack over an assumed subsequent operating cycle of two years is considered in the calculation. Maximum Permitted Flaw Size. The shroud material (austenitic stainless steel) is inherently ductile and it can be argued that the structural integrity analysis can be performed entirely on the basis oflimit load. The ASME Code recognizes this fact in using limit load techniques in Section XI, Subarticle IWB-3640. Flaw stability can be evaluated using Elastic-Plastic Fracture Mechanics or Linear Elastic Fracture Mechanics (LEFM) techniques. While LEFM techniques provide a very conservative estimate of the maximum permitted flaw size they are simple to perform and provide an acceptable basis for screening criteria. Accordingly, the maximum permitted flaw size is conservatively estimated using Linear Elastic Fracture Mechanics (LEFM) techniques. LEFM predicts the critical flaw size at which a crack becomes unstable. When the flaw size is greater than the critical size, the opening stress causes the crack to run uncontrollably through the remainder of the weld ligament. LEFM relates the mean stress that opens a crack to the crack size and the failure criterion called the critical stress intensity factor. The mean stress is further modified by a stress concentration factor to determine the stress at the crack tip. The stress concentration factor depends on the crack size and geometry. The critical stress intensity factor is material specific and depends on a number of factors. 3-3

i, i -V 7, L l 'Ihe critical flaw size for vertical welds in the FitzPatrick shroud is determined i usir g this approach in Appendix C for the limiting location and service loading i (i.e., the shroud section between H1 ant H2 during an OBE event). In addition, the differential pressure across the shroud for this case is conservatively taken as i the pressure drop between the downcomer annulus and the lower plenum. The resulting critical flaw size is applied as the maximum flaw length for all shroud ' i vertical welds. Imakage Considerations. Typical operating conditians include approximately 0.5% (of core flow) bypoas leakage and current safety analyses are 'voveding for 1-2% i leakage conditions (Reinence 7). With the minimum intact weld lengths and maximum flaw sizes discussed above, and <==hing the tortuous nature of the j flow path through a jagged crack, the leakage flow rate across the shroud during normal operation through a cracked vertical weld would be small. However, for a shroud with substantial cracking in its vertical and circumferential welds, shroud leakage should be confirmed as being acceptable. l Recame of the effect on effective crack leakage paths and leakage rates of the i many possible combinations of vertical and horizontal weld failures, such ~ evaluatioas are not included in this report. 'Ihese evaluations can be done on a case sped.fic basis. The following assumptions are used here to develop a conservative worst case l screening criteria. I Assume that cracks open under normal operating conditions about l 0.006 inches. - (This is estimated from scoping computations not included in this report for flaws about 24 inches in length.) Assume that the average leakage flow rate per linear inch of cracked i circumferential weld is 0.5 gpm. (This is conservatively estimated from the leakage flow rates through vertical welds calculated m i Reference 8. In Reference 8 it is estimated that if all nine circumferential welds fail and open up to 0.001 inches, then the total leakage rate across the shroud under normal operating c>nditions would be about 135 gpm.) i ) Assume that the allowable leakage through the vertical welds during normal operation is 0.05% of core flow (about 105 gpm). (This is estimated to be a reasonably small fraction of the typical bypass leakage of 0.5% mentioned above.) i Racerl on these assumptions, a conservative estimate for the total allowable length of vertical weld cracking is 210 inches. Therefore, if the sum of the lengths of all i flaws in the vertical welds is greater than 210 inches, then an evaluation should be performed to verify that shroud leakage is acceptable. ] 3-4

1 l t Radial Support Proximity Considerations. Most of the vertical welds are far from 1 the radial restraints, i.e., about three times (Rt)" (where R is the shroud radius j and t is the thickness), or about 36 inches'away. However, three welds are within about eight inches of a radial restraint (welds SV4A, SV6A and SB6B). For these -l welds, the screening criteria requires that the lower portion of the weld be intact in the vicinity of the support, where local h=lia: stresses are highest. (See App =Hu A.) Shell Deflection Considerations. Distortion or "ovalling" of the shroud rs prevented at most elevations by multiple radial restraints. However, in the l intermediate portion of the shroud only three radial restraints are provided at each elevation. For welds in these shroud segments (welds SV4B, SV5A and SV5B), i the screening criteria requires that additional weld length be intact in order to carry the resulting bending moment. (See App =lir A.) 3.4 RADIAL WELD EVALUATION APPROACH I The shroud ring between circumferential welds H2 and H3 supports the core top guide and i carries the load from the steam separators and tie rod preload from the top section of the [ shroud down to the lower barrel section of the shroud which rests on the shroud support plate (see Figures 1 and 3). j i The ability of the ring to transfer the load from the upper to the lower part of the shroud j depends on the condition of the circumferential welds at H2 and H3 and the radial welds in the ring (SV3A through SV3F). These welds (e.g., H2 and H3) provide redundant support for the ring radial welds. If either or both H2 and H3 are intact, then stresses in the shroud ring remain well below allowables and the radial welds are not required for the loads to be carried across the ring. However, if both H2 and H3 are failed 360' through-wall, the stresses and deflection of the shroud ring become significant and the radial welds are required to maintain the structural integrity of the ring and to carry the loads across the ring. t In developing the screening criteria, any indications in the radial welds are conservatively assumed to indicate a through-wall crack over the entire radial length of the weld (i.e., the radial weld is completely failed). Thus, to enable the shroud ring to perform its function, j the bending moment carried by the ring must be transferred across the failed radial weld i through an alternate load path. As a result, a portion of either the H2 or H3 weld must be i intact for a given length on both sides of the failed radial weld. This allows the bending moment to be transferred into the shroud cylinder, around the radial weld, and then back into the ring. This alternate load path is shown in Figure 5. } The length of circumferential weld required in the alternate load path is determined using a i limit load analysis. 'Ihe weld area required to transfer the moment is calculated using the l ASME Code allowable stresses. The limit load analysis is consistent with the analytic i evaluation procedure and acceptance criteria for austenitic piping described in the ASME Code, Subsection IWB-3640 (Reference 5). The calculation of the required weld length is contained in Appendix B and described below. 3-5 i

__. ~__ i [ a-i To transfer the bending moment into the shroud cylinder, the section of intact weld must be l i capable of carrying the stresses produced. The limiting cross sections shown in Figure 4 are as follows. - De ring cmss aection bounded by the length of intact circumferential weld and the ring thickness (Section 2 in Figure 5). De weld cms: section bounded by the length of intact circumferential weld and the shroud cylinder thickness (Section 3 in Figure 5). l The required length of intact weld is calcidatad by limiting the stresses in each cross section to be less than the ASME Code allowable stresses and solving for the required length. To ensure that the required length of weld'is available throughout the subsequent operating I cycle, two additional requirements on the condition of the weld at H2 (or H3) are considered. First, flaws in the H2 (or H3) weld are limited to the critical flaw size calculated in a Linear Elastic Fracture Mechanics evaluation of the welds at H2 and H3. i This evaluation, contained in Appendix D, considers the vertical loading produced by the core shroud repair (e.g., tie-rod preload effects). Second, the minimum weld length calculated in the limit load analyses and the critical flaw size are adjusted to account for crack growth during an assumed cpsiedig cycle of two years; bounding value for crack l growth is chosen, and linear crack growth with operating time is assumed. The resulting l screening criteria are summarized in Section 2 of this report. i i t 1 [ l 1 f l 3-6 i . ~. -., _ - _ _. ~..

1 MMPR ASS O CI AT E S IN C. ENG1NEERS Section 4 l i REFERENCES 1. General Electric Nuclear Energy Report GENE-523-154-1093, " Evaluation and Screening Criteria for the FitzPatrick Shmud Indications", October 1993. 2. ASME Boiler and Pressure Vessel Code, Section XI,1980 Edition with Addenda through Winter 1981. 3. ASME Boiler and Pressure Vessel Code, Section III, Subsection NG,1983 Edition with Addenda through Summer 1984. 4. MPR Report 1560, " James A. FitzPatrick Nuclear Power Plant Core Shroud Repair-Design Report", October 1994, Revision 0. 5. ASME Boiler and Pressure Vessel Code, Section XI,1992 Edition with no Addenda. 6. BWRVIP letter (J. T. Beckham) to U.S. Nuclear Regulatory Commission dated September 13, 1994, forwarding "BWR Core Shroud Repair Design Criteria," Rev.1, September 12, 1994. I 7. GE letter AO94-001, FitzPatrick Shroud Analysis Questions, dated October 27,1994. 8. MPR Calculation No. 2919401-409, Evaluation Of Shroud Leakage Through Failed l Circumferential Welds, Revision 0. l ? i f I 4-1

MMPR ~ A S S O CI ATE S IN C. ENGlNEERS Section 5 TABLES AND FIGURES 5-1

f Table 1 James A. FitzPatrick Nuclear Power Plant Shroud Vertical Weld Inspection Acceptance Criteria Note: ' Ibis table is provided in the report summary at the end of Section 2. L 4 I l 5-2

i Table 2 j Pressure Differences Across FitzPatrick Core Shroud Pressure Difference (psid) Shroud Elevation Normal Upset Steam-Line Break From top of shroud to core support 7.79 11.69 29.00 plate (i.e., from shroud top down to weld H6B) i From core plate to bottom of 33.93 36.33 55.50 shroud (i.e., from weld H6B to weld H8) l t h i I 5-3 i r- - ~t -a m m a

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Radiol Weld i L!,[j [p Broken / ~~Lj,' [J s s 's 'ld H2 Weld a 's ([a, s\\, Broken p t + s f N l ' Radial Moment N s l j in Ring (Mr) N s b s /%. o H3 Weld ) I Broken / ,/ Tongential Ring Bending Moment Moment in A Lood Path } Ring (Mt) h Bending Moment in Ring / / - / @ Torsion in Ring Section N Under intact Portion of Weld g h Bending in intact Weld N A N 'y N / h Through Shroud Cylinder h Bending Moment Returned Through intact Weld into Shroud Ring / ~ Section A-A Figure 5. Alternative Load Path Around Failed Radial Weld In Shroud Ring i MMPEt 97le%'U"' j )

COI TIE ROD A3SEMBLY 7 @l994 WPR ASSOCIATES H A SU % RTS / GO3 TYP 3 PLACES PATENT PENDING o f /4// / I / ~l A ^ Y l '/ [ s GO3 l / l i .f} i / j / j / j / ~ I / I e, j f GO2 TIE ROD ASSEMBLY / WITH UPPER & LOWER '4 j / l ./ RADIAL SUPPORTS ONLY TYP 3 PLACES T !./ !1 !i j ! ~/ ll l / ij i / I / / vo -/ . __._/ j l l\\ / /- j s - / s j / 's, g ! / N I / '\\, ~ l / l / j -g l '\\ l / I i ' i i

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OGO1 TIE ROD A3SEMBLY WITH 3 RADLAL SUPPORTS / @ t994 UPR ASSOCIATES GO3 TYP 3 PLACES l PATENT PENDtNG o 7 1 m / - / / \\ 1 l / cO3 i / c2 /. i r 'a . I / A I ' 3 [ CO2 TIE ROD ASSEUBLY A / wtTH UPPER & LOWER / RADtAL SUPPORTS ONLY -/ TYP 3 PLACES / c: j : / !/ ll I / jI i / / l l / / vo _/ .. _. _ _ _/ j f _,o / .s i l 'y ./ j 's, .+ j ',, s i 's, 9 ! ( l N I / j '\\. i / \\ i / j -s l / GO l . _ _L / C / ! / i / l ! / j l / cO3 j j /A// / ,ja TIE ROD ASSEMBLY l / WITH NO RADtAL SUPPORTS j cO2 TYP 4 PLACES /, SECTION A-A - f / A FITZPATRICK SHROUD REPAIR INTERMEDIATE SEISMIC RADIAL SUPPORT AREA em 87/J.To" FIGURE 6B

7 @1990 MPR ASSOCIAfts / i PAttni Ptnowc o ' / ,)/, / -/ \\1 / 1 l, N-I / j i / 7 i / i I i / j l / s. lv / l / i-t: i - / 7 l A l ~/ A j '! l / j i / / I / / j -w / va- / 4 / / I i ! / I ! / l I / x. .g j y / i / i / g: ! / i - N-r ! / i I i ! / i i I / i / i /A// l / do l / I .SECTION A-A FITZPATRICK SHROUD REPAIR - JET PUMP SUPPORT BRACKET e;;",'.%%" FIGURE BC

G01 V!E ROD ASSEuBLY @1994 MPR ASSOCIATES GO3 WITH 3 RADLAL SUPPORTS TYP 3 PLACES PATENT PENDING 0 r p ' / / /// j / / / \\' G03 h!/ GO2 [ l / j / l / 4 / /, GO2 TIE ROD ASSEMBLY yp / WITH UPPER & LOWER r ;- l RADIAL SUPPORTS ONLY / j / / / TYP 3 PLACES g / t j ! / / j ! l ~/ / i / i / / / ( / 2 ro.-1 _,o l \\ '/ I / / j N, 4 ! / j \\ / j / I 'N / t i / \\ N. _. i / G01 l

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/ j / / 7 F l i Aq G01 / A / e j l ! / \\ G03 i l /A// l l / l / ,jo GO3 TIE ROD ASSEMBLY WITH NO RADLAL SUPPORTS / GO2 SECTION A-A I / -/ / ~ FITZPATRICK SHROUD REPAIR I., / apara LOWER SEISMIC RADIAL SUPPORT AREA U O N'" FIGURE BD

O @1994 MPR ASSOCIATES /fjf 7 PATENT PEN 0tNG l / I i / X i n~ ../ X i { -a l \\ l Jg]$ -s: [@ l i l $ no-L...... -. _. -. -......:_L l / -,o '@t l l l! l / / k.- j .g' 4 ! / s j / j / j g ! / -/ I / \\ 7- + i / 'h-F i / X '+ / .A. I A, i / i E i ' / i n A// l l / do j l / l / !] i / _SECTION A-A j / / f' ./ 7 A / A DI ,/ l esatra FITZPATRICK SHROUD REPAIR - SHROUD SUPPORT GUSSET PLATES '7.'#"iW" FIGURE 6E ~ -...

{ WMPR ~ A680C4 ATES INC ENGINEERS APPENDIX A MPR Calculation No. 2919401401, " Required Intact Vedical Weld Ana Based Upon the Ti=l* lag Applied Imds", Revision 0 i l l l l 1 t I L ? 4 6 b A-1 4

a MMPR si! #; s,*e' t' ' ' " ' Alexandria, VA 22314 CALCULATION TITLE PAGE Client Niw f0fk howcf fythgt, Page 1 of 2f, Project Task No. .5krov) k/cl} Bspe C e,'t<<,a - h ase E 241-9401-046-t n _ Title Rep,',,/ 3,,isci y(,j,,, j ygy Asea Based Caucusation No.

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40'~10' Preparer /Dat,e Checker /Date Revie.wer/Date Rev. No.

f f j 1%y e,, e. s - 2. -- ll3l45 Il1Bl9[ e

MPR Associates, Inc. 320 King Street Alexandria, VA 22314 RECORD OF REVISIONS Calculation No. Pre red By Checked By 898 2919401-801 2 Revision DDription 0 Original Issue 1 i

9 MPR Associates, Inc. 320 King Street Alexandria, VA 22314 Calculation No. Pre red By Checked By / Page s-um 3 2919401 801 [ # _ _'y r m m TABLE OF CONTENTS Section Egage 1. PURPOSE 4 2. RESULTS 5 3. DISCUSSION 6 3.1 Shroud Configuration 6 3.2 Applied Loads 6 3.3 Service Loadings 8 3.4 Stress Components 9. 3.5 Calculation Method and Assumptions 10 4. CALCULATION 11 - 4.1 Primary Membrane Stress 11 4.2 Required Weld Length 15 4.3 Bending Stresses in Vertical Welds 17 4.4 Crack Growth 25 5. REFERENCES 26

1 MPR Associates, Inc. 320 King Street Alexandria, VA 22314 s Calculation No. Prep ed By Checked By Page ~ 2919401-801 e 6f A s 4 ~ .. / L PURPOSE The purpose of this calculation is to determine the area of intact vertical weld required in each shroud segment in the FitzPatrick plant. The required weld area is calculated based upon a limiting load analysis where the stresses in the vertical weld are limited to the ASME Code design allowable stresses (Reference 6). The loading and load paths used in this calculation are consistent with the design analysis of the FitzPatrick Core Shroud Repair (Reference 8). i e e 1

O EMPR 72"^"d'e':t ~ Alexandria, VA 22314 Calculation No. Pre red By Checked By Page v h 5 2919401-801 l' % : _7

2. RESULTS Results are reported in Table 1 for the required weld length for each shroud segment assuming all flaws are through-wall. The required length is based on limiting the primary membrane stress in the vertical welds to the ASME Code design allowable stress. The additional requirements listed in Table 1 are based upon limiting the primary membrane plus bending stresses in the shroud vertical welds to the ASME Code allowables.

Table 1. Required Vertical Weld Length k Shroud Length of Shroud Length for Ve m Venid Additional Requirements Segment I Weld Weld (in) Flaws (in) ^ 13 0 H1-H2 32.00 SV2B 3.4 15.5" of the lower portion of weld SV4A bv#^ is required to be intact (Note 1).15.5" H3-H4 36.00 SV4B of weld SV4B is required to be intact (Note 2). ^ 6.1 10.9" of welds SV5A and SYSB is H4-H5 98.13 SV5B required to be intact (Note 2). 1.9 f the lower portbn of welds SV6A H5-H6A 35.88 SV6A and SV6B is required to be intact SV6B (Note 1). SV8A 12.5 H6B-H7 SV8B 52.81 SV8C SV9A 5.8 H7.H8 SV9B 16.00 SV9C Notes to Table 1: 1. As shown in Figure 1, these welds are within about 5' of a radial restraint. The required weld length results from the need to preserve shroud load-carrying capability in the vicinity of the radial restraint contact where local bending stresses may be significant. 2. Distortion or "ovalling" of this shroud segment is not prevented by the three radial restraints provided at this elevation. The required length is based on the need to carry the resulting bending loads around the shroud.

MPR Associates, Inc. 320 King Street Alexandria VA 22314 Calculation No. Pre red By Ch ked By Page 2919401-801 .keh.,_.-w s 6

3. DISCUSSION 3.1 Shroud Configuration A developed view of the FitzPatrick shroud is shown in Figure 1. Circumferential welds H1 through H8 separate the shroud into a number of cylindrical segments. Each of these segments has two or more vertical welds which join individual plates to form the cylinders. The vertical shroud welds are identified as SV1 through SV9 in Figure 1.

The shroud configuration is further modified by installation of the core shroud repair. The core shroud repair consists of ten tie rod restraint assemblies that structurally replace all the circumferential welds (H1 through H8). Figure 1 shows the location and positioning of each tie rod assembly. Note that each shroud segment between two adjacent circumferential welds has radial restraints at a single elevation to react possible lateral loads. 3.2 Applied Loads The vertical welds in the shroud segments must have enough intact weld area to transmit circumferentialloads across the weld. The maximum loads transmitted through a vertical weld in any given shroud segment are taken to occur when the circumferential welds - above and below the shroud segment have failed 360" through-wall. The following circumferential loads are considered: Pressure-the pressure difference across the shroud creates a hoop load (stress) in the shroud which must be reacted through vertical welds in each shroud cylinder. The differential pressure for uprated flow conditions from Appendix A of Reference 4 is used in this calculation. 1.ateral leads-seismic events as well as a recirculation line break cause lateral loads on the shroud. Lateral loads are transmitted into the shroud barrel when the shroud contacts the radial restraints. The lateralload in each shroud segment is the radial restraint load for the restraint adjacent to the shroud segment. Also, each radial restraint may suppon more than one shroud segment (i.e., the restraint is located at a circumferential weld). For purposes of calculating the membrane stress, the entire radial restraint load is conservatively assumed to be reacted through each shroud segment supponed by the restraint. In determining the bending stresses around a restraint contact, only the load reacted through a given shroud segment is considered. i

MMPR %"a;s'"e'*P'"- Alexandria, VA 22314 Calculation No. Prepa d By Checked By 2'919401-801 vj" e tw 7 I II e f 1 18 Il 3 1 i! 4 3 3 88 aa a en g s ne se 5._._f 7 -. if a a ~ 3. s s _ _..._ _._ t _._g i_ _._._ I T: r --- - - ---t N i;l.s e . _. _4._. _. _. _.. . _._ n i_.___..._._. 1 1 o e e* ... _ g i is. .c. g_._._._._.._._ i 11 1 g lg _.__._ h _._;g --- ! j 8 Hu - -- g r _._ g -s y g x i____ E ,Y .w T 2 g .._g w 3.3 m ._c g-..._.__. . _.- I y u t E. 2 f W e ,--- a g 27 --- : 5 z:ssm s_._._.__.. _ _ c 4 m _._ i g l N_._ ! i,s, 3_._._.,,...._._,,... _. _.... _ _ _. _. _ _ _ v, 3 g I, m r;- i_._._.__.._._. ii s 8-- e. 63 _ec 5 ,_.__..,e... _. _,,. _ _... _. _. _. _ _. _ _.... _. _.e 4 2 1 t - - --- --t e ua M 4 y a[ e bf3 jg En h i h45 8* ~ ~k ~K ~ h 3 i

MMPR ??? ^" Oe':t' ' " " ~ Alexandria, VA 22314 Calculation No. Prep ed By Ch d By Page 8 2919401 801 %hm / 33 Service Leadines The core shroud repair is designed to react loads arising from the following shroud service conditions (see Reference 4). Table 2. ASME Code Service Limits ASME Allowable S mee Load Combination ay embrane Load Case Stress Intensity Limit (Note 1) Upset Level B Upset differential pressure (AP) Sm OBE Level B Operating basis earthquake (OBE) loads Sm plus normal AP DBE Level C Design basis earthquake (DBE) loads 1.5 Sm ) plus normal AP ~ DBE + Level D Main steam line break (MSLB) AP plus Lesser of 2.4 Sm or 1 MSLB DBE loads 0.7 Su DBE + Level D Recirculation line break (RLB) loads Lesser of 2.4 Sm or RLB plus DBE loads 0.7 Su Notes to Table 2: 1. The allowable stresses are from Section NG-3220 of Reference 6. Specifically, Level B limits are from NG-3223, Level C limits are from NG-3224, and Level D limits are from NG-3225 and F-1331. Also, note that Sm is the allowable stress of the material at design temperature and Su is the ultimate tensile strength of the material at design temperature. The controlling service loadings for comparison with primary membrane stress allowables can be determined by examining the load components for each service condition. The' ptessure load is proportional to the differential pressure across the shroud. As shown in Appendix A of Reference 4, the differential pressure increases by a factor of 3.7 between normal and MSLB conditions, and the ASME Code allowable stress only increases by a factor of 2.4. Consequently, the least margin between the stress produced by the pressure load and the design allowables is for the DBE + MSLB Level D load case.

l [ Gi3MPR = 0; s,*e':t j Alexandria, VA 22314 Calculation No. Prep ed By Ch ed By Page f 9 A swm 2919401 801 j,. The lateral load component is proportional to the radial restraint load produced during a seismic event or a RLB accident. The worst case lateral loads occur at the lower restraint during either an OBE event or a DBE event concurren; with a RLB. As shown in Table 1 of Reference 5, the maximum restraint load increases by a factor of 2.2 between OBE and DBE + RLB conditions, and the ASME Code allowable stress increases by a factor of 2.4. Consequently, the least margin between the stress produced by the lateral load and the design allowables is for the OBE Level B load case. (Note that the difference between OBE (Level B) and faulted (Level D) conditions is more limiting than the difference between OBE (Level B) and DBE (Level C) conditions.) The required weld length to withstand the primary membrane stresses is only calculated below for these two limiting service loadings (i.e., for the OBE and DBE + MSLB load cases). 3.4 Stress Components The pressure load and lateral load components produce the following stress components in the shroud segments. Shear Stress-the radial restraint load produces a pure shear stress as the restraint tries to punch through the shroud at the contact point. Membrane Stress-both the differential pressure and the radial restraint load produce uniform membrane stresses. The differential pressure across the shroud produces a tensile hoop stress that is constant around the entire i circumference of the shroud. The radial restraint load produces a compressive membrane stress which is greatest 90* from the contact point. Bending Stress-the multiple restraints provided around each shroud section prevent any significant distortion or "ovalling" of the shroud due to restraint loads. As a result, the bending stresses in the shroud occur largely at the radial restraints. The primary membrane stress intensity for Level A/B is limited to Sm (Reference 6, NG-3222). The membrane stress intensity varies around the shroud due to the shear distribution of the lateral load. The critical points are at the contact point of the radial restraint and 90" around the shroud from the contact point. The membrane stress intensity at the contact point is calculated from the shear stress produced by the restraint load and the hoop stress produced by the contact load. The membrane stress intensity 90* from the contact point is conservatively calculated by adding the membrane stresses due to the radial restraint and pressure loads. 1

EMPR Til ^: L'"e':t Alexandria, VA 22314 Calculation No. Pre red By ked By Page 2919401-801 6Vh

s e-10 q

The ASME Code also specifies an allowable for pure shear for Level A/B loadings of 0.6 Sm (Reference 6, NG-3227). The shear stress produced by the radial restraint load at the restraint contact must satisfy this limit. However, this limit is less restrictive than the Sm limit on the primary membrane stress intensity. Consequently, the weld area required to satisfy the pure shear stress criteria is not explicitly calculated in Section 4. Bending stresses in the shroud drop off quickly away from the radial restraint contact point. For distances greater than about 3 (R t)05 or approximately 23* around the shroud, the bending stresses become insignificant in shroud segments where ovalhng is prevented by multiple restraints. (Note that R is the shroud mean radius, and t is the shroud wall thickness.) At locations where only three radial restraints are provided some general deflection or "ovalling" of the shroud may occur. Tlie bending stresses in these shroud segments may be significant at distances greater than 23* away from the radial restraint contact point. 3.5 Calculation Method and Assumptions Stresses in the vertical welds of the shroud segments are limited to the ASME Subsection NG allowable limits for each service level. As discussed above, for welds located away from the radial restraints, the minimum weld length is governed by the primary ~ membrane stress limit:. Additional requirements are imposed on vertical welds close to radial restraints due to local bending stresses near the restraint contact and on welds in i shroud segments which may " oval". In particular, the bending plus membrane stresses in the shroud are limited to 1.5 times the primary membrane allowable stress. The average load reacted through an intact vertical weld is determined by multiplying the primary membrane stress intensity by the total weld area. Assuming all flaws are through-wall, the required length of weld can be determined by dividing the total load by the shroud thickness and the ASME Code allowable stress for the appropriate service level. This weld length is the intact weld requirement at the end of an operating cycle. In order for a flawed weld to be acceptable at the beginning of the operating cycle (i.e., during inspection), the weld must have the required area to withstand the limiting applied loads throughout the subsequent operating cycle. Consequently, crack growth during an assumed operating cycle of two years is considered. A bounding value for crack growth is applied to the minimum required weld length assuming flaw extension at both crack tips. j i i

n MPR Associates, Inc. 320 King Street Alexandria, VA 22314 Calculation No. Pre red By Checked By 2919401 801 /w % m 11

4. CALCULATION The primary membrane stresses for each shroud segment are calculated in Section 4.1.

The required weld length to satisfy the ASME Code primary membrane allowable stress requirement is determined for each weld length in ~Section 4.2. Additional weld requirements based upon the bending stresses in 1) welds near a radial restraint contact point and 2) welds in shroud segments supported by only three radial restraints are calculated in Section 4.3. Crack growth is applied to all intact vertical weld length requirements in Section 4.4. 4.1 Primary Membrane Stress 4.1.1 Membrane Stress at Restraint Contact. The primary membrane stress intensity at the restraint contact is made up of a shear component from the lateral load and a normal component from the differential pressure. The membrane stress intensity is given by: SI,e = /o} + 4Ti [Eq.1] Where: Slac = Membrane stress intensity st the restraint contact point (psi) op = Normal stress in the hoop direction due to the differential pressure (psi) rR = Shear stress due to the radial restraint load at the contact point (psi) A normal circumferential stress is developed in the shroud wall due to the differential pressure between the reactor core and downcomer annulus. Using a thin-walled cylinder model, the hoop stress in the shroud is calculated r, follows: AP R / t [Eq. 2] op = Where: op Normal stress in the hoop direction due to the differential pressure (psi) = AP Pressure drop across a shroud segment for a given service condition (psi) = R Inside radius of the shroud segment (in) = Thickness of the shroud segment (in) t = The radial restraint load causes a shear stress in the shroud wall where the restraint pushes against the shroud wall. A free body of the shroud wall at the contact point is shown below:

m l l MPR Associates, Inc. WMPR 320 King Street Alexandria, VA 22314 Calculation No. Prep ed By he ked By Pm 2919401-801 aea_ 12 ./ fy ? a[2 'A The shear stress becomes: rR = (Shear Force) / Area rR = Fgt / (2 H t) [Eq. 3] Where: rg = Shear stress due to the radial restraint load (psi) Fgt = Totallateral load applied to the shroud segment (Ibf) H = Height of the shroud segment (in) t = Thickness of the shroud segment (in) 4.1.2 Membrane Stress 90* From Contact. The membrane stress intensity in the shroud wall can be bounded by adding the normal stress components due to pressure and lateral loads 90* around the shroud from the contact point. Note that there is only a normal stress component from the radial restraint load at this point along the shroud. The membrane stress intensity becomes: SIm=op+og [Eq. 4] Where: SIm = Membrane stress intensity 90* from the contact (psi) op Normal stress in the hoop direction due to the differential pressure (psi) = on Normal stress due to the radial restraint load 90" from the contact (psi) = The pressure hoop stress is the same as calculated previous!y. The normal stress due to the radial restraint load can be calculated as follows: OR (Force) / (Area) = Fat / (2 H t) [Eq. 5] og = Where: oR Normal stress due to the radial restraint load 90" from the contact (psi) = Fut Total lateral load applied to the shroud segment (Ibf) = H Height of the shroud segment (in) = Thickness of the shroud segment (in) t =

MPR Associot:;s, Inc. 320 King Street Alexandria, VA 22314 l Calculation No. Pre red By Checked By N-13 2919401-801 [ J@/ - - ~ .. / 4.1.3 Calculation of Primary Membrane Stress. The geometric data for each shroud segment are shown in Table 3. Table 3. Shroud Segment Geometric Data Elevation Elevation Inside Height of I Shroud of Upper ofIxnver Radius of Shroud Segment Weld, m Weld, m Shroud, m b"3"#8*' " (Note 3) (Note 1) (Note 2) (Note 1) (Note 1) (Note 1) H1-H2 389.00 357.00 93.25 1.50 32.00 304 SS H3-H4 354.00 318.00 87.25 1.50 36.00 304 SS H4.H5 318.00 219.88 87.25' 1.50 98.13 304 SS H5.H6A 219.88 184.00 87.25 1.50 35.88 304 SS H6B.H7 180.00 127.19 87.25 1.50 52.81 304 SS (Note 4) H7-H8 127.19 111.19 83.75 2.375 16.00 Inconel 600 (Note 5) Notes to Table 3: 1. Taken from Reference 1. 2. The height of each shroud segment is equal to the difference in the upper and lower circumferential weld elevations. j 3. The shroud above H7 is manufactured from type 304 stainless steel (Ref. 2), and the shroud support plate below H7 is manufactured from Inconel 600 (Ref. 3). 4. As shown on Reference 1, the H6B-H7 shroud segment is not a right circular cylinder. However for calculation purposes, it is assumed that the H6B-H7 shroud segment is a right circular cylinder with an inner radius of 87.25" and a thickness of 1.5" (dimensions of shroud just below H6B). This approach results in a longer (i.e., more conservative) required intact weld length for this cylinder. 5. Thickness of the shroud between H7 and H8 taken from Reference 3. The stress components and the membrane stress intensities at the restraint contact point and 90* from the contact point are calculated in Table 4.

Table 4. Primary Membrane Stresses p s> 0 _o ,b o hh Shear Stress Normal Stress Membrane Stress P Lateral at Cbntact 90* from Intensity (psi) Differential S du Shroud Service Pressure, psi lead,Ibf due to Contact due to s t Pressure Segment Condition (Note 1) (Note 2) Restraint Restraint imad at Restraint 90* from M z (P5 ) Imad (psi) (psi) Contact Cbntact O o OBE 7.79 484.3 293,000 3052.1 3052.1 6123.4 3536.4 DBE + MSLB 29.00 1802.8 390,000 4062.5 4062.5 8322.6 5865.3 OBE 7.79 453.1 15,000 138.9 138.9 531.5 592.0 ~U 113 114 DDE + MSLB 29.00 1686.8 20,000 185.2 185.2 1727.0 1872.0 e OBE 7.79 453.1 15,000 51.0 5 t.0 464.4 504.1 4 a., s 114-115 DBE + MSLB 29.00 1686.8 20,000 67.9 67.9 16923 1754.8 b a L to OBE 7.79 453.1 218,000 2025.6 2025.6 4076.4 2478.7 x Il5.Il6A I DBE + MSLB 29.00 1686.8 290,000 2694.5 2694.5 5646.9 4381.4 OBE 33.93 1973.6 218,000 1375.9 1375.9 3386.4 3349.5 i 116B-117 DBE + MSLB 55.50 3228.3 290,000 1830.4 1830.4 4880.9 5058.6 O OBE 33.93 1196.5 218,000 2868.4 2868.4 58603 4064.9 117-118 DBE + MSLB 55.50 1957.1 290,000 3815.8 3815.8 7878.5 5772.9 ox > ra t m g to.n Notes to Table 4: sO 2 W 1. Differential pressures from Reference 4, Appendix A. Note that the pressure in the downcomer region is assumed to be the same as $[# the pressure above the steam separators. ?g o %4s 2. The total lateral load applied to a shroud segment is the limiting component seismic load shown in Table 3 of Reference 5. The

g Ill-H2 shroud segment is contacted by the upper radial restraint load, the 113114 and 114115 shroud segments are contacted by the g

y~y intermediate radial restraint, and the II5.II6A and 116B-117 shroud segments are contacted by the lower radial restraint. The su ca maximum lateralload in the 117118 shroud segment is conservatively taken to be the limiting load in the lower radial restraint from Reference 5. N + w a

MMPR sit a; n':t Alexandria, VA 22314 Calculation No. Prep ed By Checked By 2919401-801 cre &Sd4 e _ 15 4.2 Reauired Weld Length The load reacted through a vertical weld is determined by multiplying the primary membrane stress intensity by the intact weld area. Assummg all flaws are through-wall, the required length of weld can be determined by dividing the total load by the ASME Code allowable stress for the appropriate service level. The required weld area can be calculated as follows: (Imad through a vertical weld) / Sa A = req (Membrane stress in weld) (intact weld area) / Sa A = req (SIQ (H. t) / Sa [Eq.6) A = req Where: 2 Required vertical weld area (in ) A = SI q Maximum membrane stress intensity from Table 4 (psi) re = Height of the shroud segment (in) H = Thickness of the shroud segment (in) t = ASME Code allowable stress for a given service level (psi) Sa = Assuming all vertical flaws are through-wa!!, the requhcd length of weld is calculated as: L Areg / t (Eq. 7] = req Where: L,,q = Required length of intact vertical weld (in) Substituting, L (SIQ H / Sa [Eq.8] = req The required length of intact weld is calculated below in Table 5.

l MPR Associates, ine, R 32o xi#a sir > Alexandria, VA 22314 Calculation No. Pre red By Ch eked By Page tw 16 2919401-801 Table 5. Required Length of Intact Weld l D* 'I "" *" #8"* Shroud Semce Required Length Membrane Stress, psi Stress Intenstry, pst ofIntact Weld (in) Segment Condition (Note 1) (Note 2) OBE 16,675 6123.4 11.75 H1-H2 DBE + MSLB 40,020 8322.6 6.66 OBE 16,675 592.0 1.28 H3-H4 DBE + MSLB 40,020 1872.0 1.68 OBE 16,675 504.1 2.97 H4-H5 DBE + MSLB 40,020 1754.8 4.30 i OBE 16,675 4076.4 8.77 HS-H6A DBE + MSLB 40,020 5646.9 5.06 OBE 16,675 3386.4 10.73 H6B-H7 DBE + MSLB 40,020 5058.6 6.68 OBE 23,300 5860.3 4.02 H7.H8 DBE + MSLB 55,920 7878.5 2.25 Notes to Table 5: 1. From Section NG-3220 of Reference 6, the allowable membrane stresses for each service limit are as follows: Level B Sm (NG-3222/3223) Level D Lesser of 2.4 Sm or 0.7 Su (NG-3225/F-1331) Note that Sm is the allowable stress of the material at design temperature and Su is the ultimate tensile strength of the material at design temperature. From Table 3, the shroud segments between H1 and H7 are manufactured from 304 SS. For a design temperature of 575"F (Ref. 4), the allowable stress for 304 stainless steel is 16,675 psi and the ultimate strength is 63,500 psi (Ref. 7, Table 4). The H7-H8 shroud segment is Inconel 600. For a design temperature of 575'F (Ref. 4), the allowable stress for Inconel 600 is 23,300 psi and the ultimate strength is 80,000 psi (Ref. 7, Table 4). 2. Maximum membrane stress intensity in the shroud segment from Table 4. j

MPR Associctas, Inc. ~ 320 King Street Alexandria, VA 22314 Calculation No. Prep ed By C-ed By Page m-- 17 g,6 0'e. 2919401-801 g / 43 Bendine Stresses in Vertical Welds The lateral loads applied by the radial restraints can cause significant bending in the shroud in the vicinity of the restraint contact. Bending stresses may also be signi5 cant in shroud segments which are not prevented from ovalling by the radial restraints. Consequently, a longer length of intact weld may be required for vertical welds at these locations. As shown on Figure 1, the only vertical welds near a radial restraint (within 23') are the SV4A, SV6A and SV6B welds. In addition, the ovalling effect may not be prevented by the three intermediate restraints that support the H3-H4 and H4-H5 shroud segments. Consequently, the primary membrane plus bending allowable stress limits may impose further intact weld length requirements on the SV4B, SV5A, and SV5B welds. All of the above welds will be analyzed separately for bending in the following subsections. 43.1 Bending Stresses in SV4A and SV4B. The circumferential bending moment in the SV4A and SV4B welds can be determined by approximating the shroud as a ring supported at its base as shown below.

$tlfE5%"'N 7%" '

W ZERw The maximum bending moment is at the restraint contact and is given by: 2 M, = 1.5 o R [Eq. 9] (Reference 10, Table VIII, Case 18) Where: M = Maximum circumferential bending moment in the shroud (in-lbf) Total load per inch circumference (lbf/in) u = R Inside radius of the shroud at the H3-H4 shroud segment (in) = The lateral load is related to the total load per inch circumference as follows: 2nRu [Eq.10] (Reference 10, Table VIII, Case 18) Fw = Conservatively applying the maximum bending moment at each vertical weld, the bending 1 stress in the weld is given by: c3 M. c/I [Eq.11] =

a MMPR 72%';"n'!? '"*- Alexandria, VA 22314 Calculation No. Pre red By Ch ed By Page 18 2919401-801 1

2k% _

7 Where: ab Circumferential bending stress in the weld (psi) = Distance from neutral axis to extreme fiber (in) d) c = Moment ofinertia ofintact weld cross section (in I = The distance to the extreme fiber is: t/2 [Eq.12] c = And the moment of inertia of the intact weld cross section is: 3 I t /12 [Eq.13] I = y Where: Shroud thickness (in) t = L,q Length of intact weld required (in) = Additional stress components due to the restraint contact load include a membrane stress due to the restraint contact and a shear stress due to the restraint contact. However, these components result in stresses that are negligible when compared to the bending ~ ' stress. Consequently, these stress components are not considered when determining the stress intensity. As discussed previously, the differential pressure across the shroud wall causes a membrane stress in the hoop direction. The pressure hoop stress is given by: AP R / t [Eq.14] op = Where: op Normal stress in the hoop direction due to the differential pressure (ps0 = AP Pressure drop across a shroud segment for a given service condition (psi) = R Inside radius of shroud segment (in) = Thickness of the shroud segment (in) t = The hoop stress due to differential pressure is calculated above assuming the entire vertical weld is intact. If some amount of the weld is failed, then the pressure hoop stress at the weld will increase since the entire hoop load must be carried by a smaller area. The stress increases by the ratio of the area of a totally intact weld to the remaining area of a flawed weld. Assuming all flaws are through-wall, the thickness term in the each area can be eliminated, which yields:

MMPR 727 #; s'e':t* ' " * - Alexandria, VA 22314 ~ Calculation No. Prep ed By Che ked By 19 2919401-801 -wkh e p (H / I,q) [Eq.15] o = o g Where: Pressure hoop stress at a vertical weld (psi) a = 1[ = Height of shroud segment (i.e., length of vertical weld) (in) Lrq = Length of intact vertical weld required (in) The maximum stress intensity at the weld due to bending is given by: TEq.16] SIb ab + am The above equations can be solved for the required weld length by setting the stress intensity equal to the allowable stress. It can be shown that the most limiting service condition in the SV4A and SV4B weld (i.e., the H3-H4 shroud segment) for bending is the OBE load case. Considering the maximum bending stress intensity at the weld a primary stress, the allowable stress for primary membrane plus primary bending for Service Level B is: Sa 1.5 Sm (Reference 6, NG-3223) = Where: Sa Allowable stress for Level B loads (psi) = Substituting for equations 9 through 16 and limiting the maximum stress intensity to the allowable stress, the required length becomes: f ( ' 'f R (g, 37) 1,y a + q For an OBE event, the design basis static seismic load applied to the H3-H4 shroud segment is: Fw 5880 lbf (Reference 12, Table 2-1) = Note that the design basis static loads bound the actual dynamic loads that occur during an OBE event (Reference 5, page 3).

MMPR l 4" s'et Alexandria, VA 22314 Calculation No. Prepa d By Checked By 20 /4/VAm-w- 2919401-801 ,/- The other input parameters are: R = 87.25" (Reference 1) t = 1.50" (Reference 1) H = 36.00" (Reference 1) AP = 7.79 psi (Reference 4, Appendix A) Sm = 16,675 psi (See Note 1 of Table 5) From equation 17, the solution for the required length of intact weld at SV4A and SV4B is 13.71". (The primary bending and primary pressure hoop stresses corresponding to the required intact weld length are 23,820 psi and 1190 ps!. respectively.) 43.2 Bending Stresses in SV5A and SV5B. The circumferential bending moment in the SVSA and SV5B welds can be determined by approximating the shroud as a ring supported at its base. The bending moment at each weld can be calculated using the following equation from Reference 10, Table VIII, Case 18: 2 M = u R [ 1 + 0.5 cos x + (x - n) sin x ) [Eq.18] Where: o = Total load per inch circumference (Ibf/in) R = Inside radius of the shroud at SV5A and SV5B (in) x = Angle around shroud where bending moment is determined (rad) From Figure 1, the SV5A and SV5B welds are 35' and 25' away from a radial restraint, respectively. The maximum bending moment at either SV5 weld occurs at SV5B (i.e., x = 25') and is: 2 M,= u R { 1 + 0.5 cos 25' + [(25' n /180") - n] sin 25' } 2 M = 0.3099 o R [Eq.19] Where: Mm= Maximum bending moment in the SV5 vertical welds (in-Ibf) The lateral load is related to the total load per inch circumference as follows: Fw =2nRu [Eq.20] The maximum stress intensity as well as the bending stresses and pressure hoop stresses can be determined using the same approach as in Section 4.3.1 with equations 11 through 16.

MMPR lis #; C.t Alexandria, VA 22314 Calculation No. Prep ed By ked By Pm 21 2919401-801 2269 n _ e The equations can be solved for the required weld length by setting the stress intemity equal to the allowable stress. It can be t,hown that the most limiting service cond; tion in the SV5A and SV5B weld (i.e., the H4-H5 shroud wgment) for bending is the OBE load case. Considering the maximum bending stress intensity at the weld a primary nress, the i allowable stress for primary membrane plus primary bending for Service Ievel B is: 1.5 Sm (Reference 5, NG-3223) Sa = Where: Allowable stress for Level B loads (psi) Sa = Combining equations 18 through 20 and 11 through 16 while limiting the maximum stress intensity to the allowable stress, the required length becomes: R [Eq. 21] L,,, 2 + Where: Q = Length of intact vertical weld required (in) -~ Lateral load applied to the shroud segment (Ibf) Fai = Pressure drop across a shroud segment for a given service condition (psi) AP = Inside radius of shruad segment (in) R = Thickness of the shroud segment (in) t = Height of.shrcud segment (i.e., length of vertical weld) (in) H = Sa Allowable stress (psi) = For an OBE event, the design basis static seismic load applied to the H3-H4 shroud segment is: Fgt 16,020 lbf (Reference 12, Table 2-1) = Note that the design basis static loads bound the actual dynamic loads that occur during an OBE event (Reference 5, page 3). The other input parameters are: R 87.25" (Reference 1) = 1.50" (Reference 1) t = H 98.125" (Reference 1) = AP 7.79 psi (Reference 4, Appendix A) = Sm 16,675 psi (See Note 1 of Table 5) =

EMPR

2:s'.*:,"c-Alexandria, VA 22314 Calculation No.

Prep ed By hecked By 2919401-801 4,,/,f,;'/. 22 c From equation 21, the solution for the required length of intact weld at SVSA and SV5B is 9.13". (The primary bending and primary pressure hoop stresses corresponding to the required intact weld length are 20,140 psi and 4870 psi, respectively.) 4.3.3 Bending Stresses in SV6A and SV6B. The circumferential bending moment in the SV6A and SV6B weld can be determined by approximating the shroud as a symmetrically supported ring.

USUd g3_u.a j

1 w 60 EavR EsoR The bending moment in the shroud can be determined by using formulas from ~ Reference 10, Table VIII, Case 19: 2 M M - T R (1 --cos x) + u R (x sin x + cos x - 1) [Eq. 22] = 1 1 2 2 M 1 u R [ 0.5 + cos e + (e - n) sin e + sin e ] [Eq. 23] = 2 T1 u R (sin e - 0.5) [Eq.24] = Where: Total load per inch circumference (Ibf/in) o = R Inside radius of the shroud at SV6A and SV6B (in) = Angle around shroud where bending moment is determined (rad) x = Half angle between the lower radial restraints (rad) e = 4 The half angle between the lower radial restraints is at 30" (see Figure 1). It can be shown that the maximum bending moment occurs at the restraint contact (i.e., x = 30"). Substituting these values into equations 22 through 24 and conservatively assummg that the bending moment at the vertical welds is the same as at the contact (approximately 5" away), the bending moment at the vertical welds becomes: 2 2 M 1 u R { 0.5 + cos (n / 6) + [(n / 6) - n] sin (n / 6) + sin (n / 6) } = 0.3070 u R2 M = 1 1 2 T1 u R [ sin (n / 6) - 0.5] = T - 0.25 u R = 1 2 2 M = 0.3070 u R + 0.25 o R [ 1 - cos (n / 6) ] + 2 u R [ (n / 6) sin (n / 6) + cos (n / 6) - 1 ] M = 0.4683 u R2 [Eq. 25]

EMPR 72T2702*'""- Alexandria, VA 22314 Calculation No. Prep d By Checked By Page 23 2919401.801 sgu -n _ - _ Where: M,= Maximum bending moment at vertical weld SV6A or SV6B (in-lbf) The total lateral load is related to the total load per inch circumference as follows: F =2nRu [Eq.26] ht The maximum stress intensity as well as the bending stresses and pressure hoop stresses can be determined using the same approach as in Section 4.3.1 with equations 11 through 16. The equations can be solved for the required weld length by setting the stress intensity equal to the allowable stress. It can be shown that the most limiting service condition in the SV6A and SV6B weld (i.e., the H5-H6A shroud segment) for bending is the DBE + RLB load case. Considering the maximum bending stress intensity at the weld a pnmary stress, the allowable stress for membrane plus bending fer Service Level D is: Sa = Lesser of 3.6 Sm or 1.05 Su Where: Sa = Allowable stress for Level D loads (psi) Combining equations 25 and 26 with equations 11 through 16 while limiting the maximum stress intensity to the allowable stress, the required length becomes: AP R f [Eq. 27] L,y a + Where: Iy = Length of intact vertical weld required (in) Fut - I.ateralload applied to the shroud segment (Ibf) AP = Pressure drop across a shroud segment for a given service condition (psi) R = Inside radius of shroud segment (in) t = Thickness of the shroud segment (in) H = Height of shroud segment (i.e., length of vertical weld) (in) Sa = Allowable stress (psi) The lateral load for the DBE + RLB load case is: Fui = Fogg + Fats [Eq.28)

EMPR Tis ^*; L'"e':t 2 Alexandria, VA 22314 Calculation No. Prep ed By C ed By Page D.j 24 2919401-801 @9 u Where: FDBE= Total lateral load during a DBE seismic event (Ibf) Fm = Lateral load on the H5-H6A shroud segment due to a RLB (Ibf) For a DBE event, the design basis static seismic load applied to the H5-H6A shroud segment is 7810 lbf (Reference 12, Table 2-1). Note that the design basis static loads bound the actual dynamic loads that occur during a DBE (Reference 5, page 3). The lateral load due to a RLB can be determined from the force diagram in Figure 2-3 of Reference 11. The RLB lateralload is the difference in the force between the H5 and H6A elevation, which is 39,000 lbf. (Note that it is assumed that circumferential welds H5 and H6A are not intact.) Using these values, the total lateral load becomes: Fui = 7810 lbf + 39,000 lbf Fu, = 46,810 lbf The other input parameters are: R = 87.25" (Reference 1) t = 1.50" (Reference 1) H = 35.88" (Reference 1) AP = 7.79 psi (Reference 4, Appendix A) Sm = 16,675 psi (See Note 1 of Table 5) Su = 63,500 psi (See Note 1 of Table 5) From equation 27, the. solution for the required length of intact weld at SV6A or SV6B is 13.79". (The Level D bending and pressure hoop stresses corresponding to the required intact weld length are 58,850 psi and 1180 psi, respectively.) i l

MMPR If" ^"; s'e' t* ' ' " ' Alexandria, VA 22314 Calculation No. Pre red By C d By Page 2919401-801 r wh _ 25 4.4 Crack Growth Assuming the same bounding crack growth rate as in the GE Screening Criteria for 4 FitzPatrick (Ref. 9, p. 4 and Figure 7-1), the crack growth rate is 5.0 x 10 in/hr. For a iinear growth over a two year operating cycle (17,520 hr), the maximum crack extension in a single direction is 0.876 in. The required intact weld length due to primary membrane stress is calculated in Section 4.2. The minimum intact weld length is determined in Table 6 assuming crack growth is linear with time over the entire operating cycle. Table 6. Intact Vertical Weld Requirements Required Length at Minimum Length Shroud End of Cycle,in of Intact Weld, in

c8* * "I (Note 1)

(Note 2) [ H1-H2 11.75 13.50 H3-H4 1.68 3.43 H4-H5 4.30 6.05 H5-H6A 8.77 10.52 H6B-H7 10.73 12.48 H-7-H8 4.02 5.77 Notes to Table 6: 1. From Table 5. 2. Calculated assuming crack growth at both ends of through-wall flaw. Crack growth over an operating cycle must also be applied to welds in close proximity to radial restraints. The minimum length of intact weld required for the SV4A or SV4B weld is 15.46", the intact weld length required for the SV5A or SV5B weld is 10.88", and the minimum length of weld required for SV6A and SV6B is 15.55".

~ ) MPR Associates, Inc. p 320 King Street Alexandria, VA 22314 d By C ed By [ Prep Calculation No. Page 2919401-801 W2,. m ~ _ - 26

5. REFERENCES 1.

MPR Drawing No. 1291-001-17, " James A. FitzPatrick Nuclear Power Plant, Shroud Contingency Repair, Shroud Restraint Assembly", Rev. A. 2. Sun Shipbuilding and Drydock Co. Drawing No. 42361-2, " Assembly and Finish Machining - Shroud", Rev. 2, Sheet 1. 3. Combustion Engineering Drawing No. E233-254, " Shroud Support Assembly and Details", Rev. 4. 4. MPR Specification No. 291001-001, " Design Specification for James A. FitzPatrick Nuclear Power Plant (JAF) Core Shroud Repairs", Rev. O. 5. MPR Calculation No. 2919401-741, ' Tie-Rod Restraint Limiting Reactions Using Dynamic Analysis Seismic Loads", Rev. 0 (provided in Appendix G of Reference 8). 6. ASME Boiler and Pressure Vessel Code, Section III, Subsection NG,1983 Edition with Addenda through Summer 1984. 7. MPR Calculation No. 2919401-202, ' Tie Rod Assembly Stress Evaluation", Rev.1. 8. MPR Report 1560, " James A. FitzPatrick Nuclear Power Plant Core Shroud Repair - Design Report", Rev. O, October 1994. ~ 9. GENE-523-154-1093, " Evaluation and Screening Criteria for the FitzPatrick Shroud Indications", GE Nuclear Energy, October 1993.

10. Roark, R. J. Formulas for Stress and Strain. 3d ed. McGraw-Hill Book Company, New York,1954.

11. GENE-523-A136-0994, "FitzPatrick Nuclear Station Shroud Safety Assessment",

GE Nuclear Energy, October 27,1994 with revised pages in GE Nuclear Energy Letter A094-001 from R. DuBord (GENE) to W. Birely (NYPA), "FitzPatrick Shroud Analysis Questions".

12. MPR Calculation No. 2919401-104, " Equivalent Seismic Loads on the FitzPatrick Shroud", Rev.1 (provided in Appendix G of Reference 8).

i 1 MMPR A SS O CI AT E S IN C. ENGINEER 5 APPENDIX B MPR Calculation No. 2919401-803, " Initial Acceptance Criteria for Radial Welds with Indications", Revision 0 i B-1 I

r MPR Associates, Inc. 320 King Street f Alexandria, VA 22314 CALCULATION TITLE PAGE Client 9' YGw' fof{ fowr kut$ er 1 Project Task No. bhrov/ Weld Snspec (r,leria - fhase [ 3 A l

4401- 04G -l n

Title fn,4,,( M ec</f c. ce Or,her.'4 h Ae/,11 Calculation No. WeIOS w kh Znd eeben 2.4I940I-803 Pieparer/Date Checker /Date Reviewer /Date Rev. No. Dd UO ~ A. g NePe haAu^m n/w/H C f 6. 5, _,, /2//2/4y 12/19/94 l

MPR Associates, Inc. P PR 32o ximo street Alexandria, VA 22314 RECORD OF REVISIONS Calculation No. Prep d By C eked By Page 2919401-803 g 3 Revision [ Description 0 OriginalIssue

MPR Associr,tss. Inc. Alexa dr a, A 2314 Calculation No. Prep ed By Checked By 2919401-803 .f#s% 3 L PURPOSE The purpose of this calculation is to determine screening criteria for initial acceptance of flaws discovered during inspections of the radial welds in the ring between the H2 and H3 circumferential welds in the FitzPatrick core shroud. The screening criteria are based upon the limiting loads on the shroud ring and the amount of crack growth expected in a two year operating cycle. The loading and load paths used in this calculation are consistent with the design analysis of the FitzPatrick Core Shroud Repair (Reference 11).

2. RESULTS The screening criteria for acceptance of indications in the radial welds of the shroud ring between H2 and H3 based on the limit load analysis are as follows:

a mmimum length of 20 inches of intact circumferential weld at H2 immediately e adjacent and on both sides of the radial flaw, E a minimum length of 20 inches of intact circumferential weld at H3 immediately e adjacent anden both sides of the radial flaw.

MPR Ass 0cintos, Inc. ~ 320 King Street Alexandria, VA 22314 Calculation No. Prep d By C eked By Page 2919401-803 W,,, _ 4

3. DISCUSSION 3.1 Shroud Confiruration A developed view of the FitzPatrick core shroud is.shown in Figure 1. Circumferential welds H1 through H8 separate the shroud into a number of cylindrical segments. Each of these segments has two or more vertical or radial welds which join individual plates.

The radial welds to be inspected are located on the ring between the H2 and H3 welds. The shroud configuration is further modified by installation of the core shroud repair. The core shroud repair consists of ten tie rod restraint assemblies that structurally replace all the circumferential wem H1 through H8. Reference 1 shows the location and positioning of each tie rod assembly. Note that each shroud segment between two adjacent circumferential welds has a radial restraint to react possible lateralloads. 3.2 Service Imadings The following conditions are evaluated for each shroud segr:ent. The ASME B&PV Code, Section III service limits which corresponds to the specific load cases are also shown. These service loadings are consistent with those specified in the design specification for the core shroud repair (Reference 7). ASME Imad Case Service Load Combination Limit Upset Level B Upset differential pressure (AP) OBE Level B Operating basis earthquake (OBE) loads plus normal AP DBE Level C Design basis earthquake (DBE) loads plus normal AP ~ DBE + MSLB Level D Main stedm line break (MSLB) AP plus'DBE loi~ds- ~ ~ DBE + RLB Level D Recirculatic, line break (RLB) loads plus DBE loads

MPR Associates, Inc. ~ Q 320 King Street Alexandria, VA 22314 Calculation No. Pre red By Checked By Page .. ~- 2919401 803 e% 5 i n n i i It la 1 I I e a. e e e

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MMPR lis a; L'"e'*t Alexandria, VA 22314 Calculation No. Prep ed By Checked By 2919401-803 - M./W;m 6 3.3 Function of Shroud Ring The shroud ring between circumferential welds H2 and H3 supports the core top guide and carries the load from the steam separators and tie rod preload from the top section of the shroud down to the lower barrel section of the shroud. i The ability of the ring to transfer the load from the upper section to the lower part of the shroud depends on the condition of the circumferential welds at H2 and H3 and the radial welds in the ring (SV3A through SV3F). These welds provide redundant support for the ring and not all of the welds are required to transfer the load. If either or both H2 and H3 are intact, then stresses in the shroud ring remain well below allowables and the radial welds are not required for the loads to be carried-across the ring. However, if both H2 and H3 are failed 360" through-wall, the stresses and deflection of the shroud ring can become significant and the radial welds are required to maintain the structural integrity of the ring and to carry the loads across the ring (Reference 11, Sections 3 and 4). I 3.4 Calculation Method and Assumptions As the preload from the tie rod assemblies is applied, the vertical load transmitted through the shroud ring by a vertical couple acting between the H2 and H3 welds produces a uniformly-distributed, tangential moment on the ring. The distnbuted moment tends to turn the ring inside-out. As shown in Figure 2, two bending stresses are produced by this action: a tangential bending stress produced by the bending of the entire ring about its e centroid. a radial bending stress produced by the distributed tangential moment on the e ring (in the region where weld H2 or H3 is intact). The radial welds in the shroud ring must be able to carry the radial bending moment. However, when a radial weld is completely failed, the ring no longer has the capability to react the bending moment through the radial weld. To enable the throud ring to perform its function without significant additional deflection, the bending stresses must be transferred across the weld through an alternate load path. A portion of either the H2 or H3 weld must be intact for a given length on either side of the failed radial weld. This allows the bending moment to be transferred into the shroud. cylinder, around the radial weld, and then back into the ring. The length of circumferential weld required in the alternate load path is determined by calculating the required area to transfer the moment and not overstress any portion of the weld or shroud.

l MPR Associates, Inc. 320 King Street ~ Alexandria, VA 22314 Calculation No. Prep d By Checked By Page 7 2919401-803 ffMw-To transfer the stresses through the alternate load path, the section of intact weld must be capable of carrymg the stresses produced. The limiting cross sections shown in Figure 2 are as follows: The ring section bounded by the length of intact circumferential weld and the ring thickness transfers the bending stresses by a combination of torsion and bending (see Section 2 in Figure 2). The weld section bounded by the length of intact circumferential weld and the shroud cylinder thickness transfers the bending stresses to the shroud (see Section 3 in Figure 2). The required length of intact weld can then be calculated by constraining the maximum stress intensity in each cross section to be less than the ASME Code allowable stresses and solving for the required length. The acceptance criteria for flaw indications discovered during inspections of the radial welds in the shroud ring is the minimum length ofintact circumferential weld on each side of the cracked radial weld. In order for a flaw to be acceptable, the circumferential weld must have the required area to transfer the tangential bending moment across the weld throughout the subsequent operating cycle. Consequently, crack growth during an assumed operating cycle of two years is also considered. A bounding value for crack growth is chosen, and linear crack growth with operating time is assumed. m .e I

MPR Associates, Inc. lexa dr a, VA 2314 Calculation No. Pr ed By Ch eked By 2919401 803 -M h s ./ A Z M( z 2( l f M I Length of Intact r l I Weld Required ) 1% 2c -Radict Weld 1,[j 2c [ Broken /- - rs \\ q;N 's fd N H2 Weld 's N N Broken g Redici Moment / s . \\ N s in Ring (M ) r s l N wt-H3 Weid Broken 7 Tcngentici Rinc Bending Moment i Moment in Loce Peth g Ring (M t) @ Bencing Moment in Ring [ / @ Uncer intcet Por*Jon of Weld@y i Torsion in Ring Section h / @ Bending in inteet Weld X\\ @ Through Shroud Cylinder 2b @ Bending Moment Retumed Through intact Weld into Shroud Ring / L l Seeflon A-A Figure 3. Alternative Lead Path Around Failed Radial Weld In Shroud Ring r4 M P R 'i720:T ?

~ i MPR Associates, Inc. 320 King Street Alexandria, VA 22314 Calculation No. Prep ed By Checked By 2919401-803 9

4. CALCULATION 4.1 Applied Inad The applied verticalload through the shroud ring are determined in Reference 3. Using these results, Table 2 shows the vertical load transmitted through the shroud ring for each service condition assuming a maximum normal operating load of 32,350 lbf per tie rod (see Section 5.2 of Reference 4).

Table 2. Applied Load Through the Shroud Ledge Vertical Load Through Shroud Ledge (kips) Load Case load as a Function of ated bad Tie Rod Load (Note 3) P 10 Fp - 201.8 121.7 (Note 1) OBE 3.53 Fp + 255.2 369.4 DBE 3.53 Fp + 315.0 429.2 DBE + MSLB 409.4 409.4 E 3.53 F + 315.0 429.2 p gg Notes to Table 2: 1. For the upset pressure case, the vertical load is calculated using the same method as in Section 4.1 of Reference 3 using a shroud head differential pressure load of 319.3 lbf (Reference 5, Table 2). 2. Recirculation line break + DBE loads are conservatively assumed to be the same as the DBE loads in the vertical direction. 3. From Table 2-1 of Reference 3. Note that F is the load in a single tie-rod at p normal operating conditions. I J

MPR Assseictss, Inc. j 320 King Street Alexandria, VA 22314 Calculation No. Prepar By Checked By 10 2919401-803 qf67,jp 4.2 Bendine Moment in Rine The verticalload transmitted through the shroud ring produces a distributed tangential moment which tries to turn the ring inside-out (see Figure 2). 4.2.1 Tangential Bending Moment. The highest stresses in the ring are produced when both H2 and H3 are failed 360", through-wall. The tangential bending moment in the ring is determined in Reference 6 for this condition. The tangential bending moment per unit circumference is given by: T 0.00787 Fsg - 0.00525 (dPSR) (ASR) (Reference 6, p. 7) M = Where: '~ M = Tangential bending moment in the shroud ring per unit circumference T (in-lbffm) F = Verticalload transmitted through the shroud ring (Ibf) SR dP g = Differential pressure across the shroud ring (psi) 3 2 ASR = Pressure area of the shroud ring (in ) The tangential bending moment is calculated in Table 3. Table 3. Applied Tangential Bending Moment in Shroud Ring Imad Through Differential Pressure angen Moment 2 Load Case Shroud Ring, Pressure, psi Area, in Cir r ce' Ibf (Note 1) (Note 2) Mbffm ~'~ ~ ~ pset 121,700 11.69 749.0 OBE 369,400 7.79 2768.0._ DBE 429,200 7.'i9 3402 3238.6 %._m DBE + MSLB 409,400 29.00 270s.O DBE + RLB 429,200 7.79 3238.6 Notes to Table 3: n 1. From Appendix A of Reference 7. 2. From Table 2 of Reference 5.

MPR Associates, Inc. 320 King Street Alexandria, VA 22314 Calculation No. Prepa d By Checked By Page 11 2919401-803 4.2.2 Radial Bending Moment. If a vertical couple between H2 and H3 is applied at some point along a ring, the couple produces a radial bending stress at the ring's neutral axis. As the couple is distributed around the entire ring (i.e., the distntuted tangential moment), a radial bending stress is generated at every cross section along the ring. This radial bending stress is the primary stress produced in the ring due to a distntuted torque (see Figure 2). The case of a narrow ring under a distributed torque is shown in Roark (Reference 8, Section 10.9). The radial bending moment can be determined from the uniformly distributed tangential moment as follows: Afg =MR (Reference 8, p. 454) 7 Where: hig = Radial bending moment at each ring cross section (in-lbf) M = Tangential bending moment in the shroud ring per unit circumference T (in-lbffm) R = Radius of shroud at centroid of shroud ring (in) Table 4. Radial Bending Moment at Each Shroud Ring Cross Section Tangential Moment per Radius at Radial Moment Load Case - Unit Circumference, Centroid of at Each Ring in-lbffm Shroud Ring, in Cross Section, 1 (Note 1) (Note 2) in-Ibf Upset 749.0 68,160 OBE 2768.0 251,890 - DBE 3238.6 91.00 -294,720 DBE + MSLB 2704.0 246,070 DBE + RLB 3238.6 294,720 4.3 Beauired Circumferential Weld Length The applied tangential moment results in a tangential stress in the ring as the ring tries to turn inside-out. A radial stress in the ring also results from the radial bending

l MPR Associates, Inc. 320 King Street Alexandria, VA 22314 Calculation No. Prep d By Checked By 2919401-803 ' ist--w 12 produced at each ri g cross section. The radial welds in the shroud ring must be able to carry these stresses, or an alternate load path must be provided to carry the load around the radial weld. The alternate load path considered in this evaluatio~n is shown in Figure 2. The load is transferred around the radial weld through the shroud cylinder and back into the shroud ring. The transfer of the load into the shroud cylinder requires that some minimum length of the circumferential weld (either H2 or H3) along each side of the failed radial weld be intact. The required length of circumferential weld is determined by calculating the required area to transfer the bending moment and not overstress any cross section. There are two critical cross sections: Ring Cross Section-Ring section under the intact circumferential weld where e the radial moment in ring is transferred in torsion and the tangential moment is transferred in bending (Figure 2, Section 2) Weld Cross Section-Intact weld cross section in shroud cylinder where the e stresses are trasferred in bending (Figure 2) 4.3.1 Weld length Required for Ring Cross Section. The ring cross section is the area bounded by the length of intact circumferential weld and the thickness of the shroud ring (see Figure 2). Both the radial and tangential moments must be transferred across the ring cross section. The radial moment produces torsion in the ring cross section. Assuming that the cross section is rectangular, the manmum shear stress can be calculated using formulas from Roark (Reference 8, Table 20). The maximum shear stress is located at the midpoint of the longer side of the cross section and is determined as follows: 1 + 0.6095( ) + 0.8865( - 1.8023( + 0.9100( ) T = Where: = Maximum shear stress in the limiting cross section in torsion (psi) r hi = Radial bending moment at each ring cross section (in-lbf) g = Length of half the intact weld required on one side of the axial weld (in) a b = Half the thickness of the shroud ring (in) b = 1.50 in. (Reference 1) The distributed tangential moment produces bending in the ring cross section. The

3 MPR Associates, Inc. 320 King Street Alexandria, VA 22314 Calculation No. < Prep d By Checked By Page / 2919401-803 - m /V4e m 13 m maximum bending stress can be determined using the flexural formula for beams with rectangular cross sections: = M, c / I, a t t Where: a, = Bending stress in the ring cross section due to the tangential moment (psi) M, = Tangential bending moment along cross section length (in-lbf) M = (Tangential bending moment per circumference) (length of intact weld) t M, = 2 Af a T = Distance to extreme fiber (in) e c =b 4 I = Moment of inertia about the tangential axis of the ring cross section (in ) n I = (1/12)(base)(height)3 = (1/12) (2a) (2b)3 n 3 4ab /3 I = n Substituting the above equations, the bending stress due to the tangential moment becomes: = (2 Af a) (b)2)[ (1/12) (2a) (2b)3 ) / a t t = (3 Af ) / (2 b a T i The stress intensity at the location of maximum stress can be determined as follows: SI - lat + Kz Where: SI = Stress intensity (i.e the largest difference between principal stresses) (psi) The length of intact weld required on each side of a failed radial weld can be determined as follows: Setting the stress intensity of the cross section equal to the ASME Code allowable stress Solving the above equations for the length of intact weld required e Results for each service level are shown in Table 5. j l i

3 MPR AssocintGs, Inc. PR 32o x'"a S"e ' Alexandria, VA 22314 Calculation No. Prep ed By Checked By I4 2'919401-803 h furwm Table 5. Required Length of Intact Weld for the Ring Cross Section All wable Service Calculated Length of Intact ad II*88 T Level Value for a, Circumferential Combm. anon 304 SS, pst (Note 1) in Weld Required, in (Note 2), Upset B 16,675 2.018 4.04 OBE B 16,675 5.995 11.99 DBE C 25,013 4.844 9.69 DBE + MSLB D 40,020 2.824 5.65 DBE + RLB D 40,020 3.275 9.69 Notes to Table 5: 1. From Section NG-3220 of Reference 9, the allowable stresses for each service limit are as follows: Level A/B Sm (NG-3222/3223) Level C 1.5 Sm (NG-3224) level D Lesser of 2.4 Sm or 0.7 Su (NG-3225/F-1331) Note that Sm is the allowable stress of the material at design temperature and Su is the ultimate tensile strength of the material at design temperature. 2. For a design temperature of 575'F (Reference 7), the allowable stress for 304 stainless steel is 16,675 psi and the ultimate strength is 63,500 psi (Reference 4, Table 2). 4.3.2 Weld length Required for Weld Cross Section. The radial moment transmitted across the circumferential weld causes bending over the length of the intact weld, and the tangential bending moment causes bending about the thickness of the intact weld. The weld cross section is shown below:

MPR Associates, Inc. A exa dr a, A 2314 Calculation No. Prep d By cke By 15 2'919401-803 - jg,mh _ The bending stress due to the radial moment is determined as follows: Mg c / I o = rr r Where: Bending stress in shroud (psi) o = r Radial bending moment (i.e., bending moment) (in-Ibf) Mg = Distance from midpoint of weld to extreme fiber (in) c = c = a 4 Moment of inertia of the weld cross section about the r-r axis (in ) I = rr (1/12) (base) (height)3 = (1/12) t (2 a)3 I = n 3 2ta /3 I = y Shroud thickness (in) t = 1.50 in (Reference 1) t = Substituting the distance to the extreme fiber and the moment of inertia into the bending stress equation yields: 3 (Ma ) / ( 2 t a / 3) a o = r 2 (3 Mg) / (2 t a ) o = r The distributed tangential moment produces bending in the weld cross section. The maximum bending stress can be determined using the flexural formula for beams with rectangular cross sections: M c/I - a = t t n Where: a, Bending stress in the weld cross section due to the tangential moment (psi) = Tangential bending moment along cross section length (in-Ibf) M = t M (Tangential bending moment per circumference) (length of intact weld) = i M 2M a = 7 i Distance to extreme fiber (in) c = t/2 c = d I Moment of inertia about the tangential axis of the ring cross section (in ) = n (1/12) (base) (height)3 (1/12) (2a) (t)3 I = = tt 3 I at /6 = n

= l MPR Associates, Inc. i 320 King Street Alexandria, VA 22314 Calculation No. Prep ed By Checked By 2'919401 803 16 e mm Substituting the above equations, the bending stress due to the tangential moment becomes: 3 a (2 M, a) (t / 2) / (a t / 6) = 2 (6 M.g.) / (t ) = o, The stress intensity at the location of maximum stress can be determined as follows: SI =o+0r t Where: Stress intensity (i.e., the largest difference between principal stresses) (psi) SI = The length of intact weld required on each side of a failed radial weld can be determined as follows: Setting the stress intensity of the cross section equal to the AShE Code allowable stress Solving the above equations for the length of intact weld required e Results for each service level are shown in Table 6. Table 6. Required Length of Intact Weld for the Weld Cross Section Allowable Service Length of Intact Load Stress for Calculated g) Case 3 Circumferential Weld (Note 1) - Required, in - oe Upset B 16,675 2.155 431 OBE B 16,675 5.206 10.41 DBE C 25,013 4.242 8.48 ] DBE + D 40,020 2.739 5.48 MSLB ~ DBE + D 3.064 6.13 g 40,020

l MPR Associates, Inc 320 King Street Alexandria, VA 22314 ared By ecked By Prep / Page Calculation No. ~ 2'919401-803 ' -[#h 17 Notes to Table 6: 1. From Section NG-3220 of Reference 9, the allowable stresses for each service limit are as follows: Ievel NB Sm (NG-3222/3223) Izvel C 1.5 Sm (NG-3224) Ievel D Lesser of 2.4 Sm or 0.7 Su (NG-3225/F-1331) Note that Sm is the allowable stress of the material at design temperature and Su is the ultimate tensile strength of the material at design temperature. 2. For a design temperature of 575"F (Reference 7), the allowable stress for 304 stainless steel is 16,675 psi and the ultimate strength is 63,500 psi (Reference 4, Table 2). 433 Maximum Weld length Required. After comparing Tables 5 and 6, the most limiting cross section is the ring cross section during the OBE load combination. The required intact weld length is 11.99". I

I I MMPR 72M'; s'e,' '" - Alexandria, VA 22314 Calculation No. Prep d By Checked By Page 2919401 803 MKwh_ 18 4.4 Crack Growth Assuming the same bounding crack growth rate as in the GE Screening Criteria for FitzPatrick (Reference 10, p. 4 and Figure 7-1), the crack growth rate is 5.0 x 10-5 in/hr. For linear crack growth over a two year operating cycle (17,520 hr), the maximum crack extension in a single direction is 0.876 in. The intact circumferential weld requirements for continued operation of a radial flaw is calculated in Section 4.1 through 43 of this calculation. Assuming linear crack growth over the entire operating cycle, the minimum length of intact H2 or H3 weld on each side of a fully-cracked radial weld for initial acceptance is determined as follows: (Required weld length from Section 433) + Required Length = (Maximum crack extension) 11.99" + 0.876" Required Length = 12.87" Required Length = The required length of intact circumferential weld to structurally replace a radial weld in the shroud ring is 12.87". This result is based upon the limiting loads and ASME Code allowable stresses as well as crack growth over one operating cycle. Note that because of the structural continuity there will likely be additional stresses in the vicinity of the crack tip which have not been determined in the evaluation above. Accordingly to account for these additional stress components, a conservative estimate of 20" for the required weld length will be used as the screening criterion for acceptance of radial welds with flaws. 4+

6 ew

WMPR = 2 ; s:t ~ Alexandria, VA 22314 Calculation No. Prep d By Checked By Page 19 2919401 803 , ggh _

5. REFERENCES 1.

MPR Drawing No. 1291-001-17, " James A. FitzPatrick Nuclear Power Plant, Shroud Contingency Repair, Shroud Restraint Assembly", Rev. A. 2. New York Power Authority Drawing No. SK-1, " Core Shroud Layout - Basic Inspection Plan for 1994 Refuel Outage", Rev. 0A-0. 3. MPR Calculation No. 2919401-113, " Evaluation of Vertical Loads Transmitted Across the Shroud Ledge", Rev.1. 4. MPR Calculation No. 2919401-202, " Tie Rod Assembly Stress Evaluation", Rev. O. 5. MPR Calculation No. 2919401-101, " Pressure Drop Loads on FitzPatrick Core Shroud", Rev.1. 6. MPR Calculation No. 2919401-501, " Determination of Maximum Allowable Vertical Loads on t'Te Shroud Ledge Assuming Failure of Welds H2, H3, and H6B", Rev. 2. 7. MPR Specification No. 291-001-001, " Design Specification for James A. FitzPatrick Nuclear Power Plant Core Shroud Repair", Rev. O. 8. Young, Warren C. Roark's Formulas for Stress and Strain. 6* ed. McGraw-Hill Book Company, New York,1989. 9. AShE Boiler and Pressure Vessel Code, Section III,1983 Edition with Addenda through Summer 1984.

10. GENE-523-154-1093, " Evaluation and Screening Criteria for the FitzPatrick Shroud Indications", GE Nuclear Energy, October 1993.

11. MPR Repon No.1560, " James A. FitzPatrick Nuclear Plant Core Shroud Repair - Design Repon", Revision 0, October 1994.

O ? MMPR A S SO CI ATE S IN C. ENGlNEERS APPENDIX C MPR Calculation No. 2919401-804, " Critical Flaw Size for Indications in Vertical Welds of the FitzPatrick Shroud", Revision 0 i C-1 1

72o""^"s'e'I'"' MMPR Alexandria, VA 22314 CALCULATION TITLE PAGE Client "**O Ale-Yod Po~e AviL,<.tj Project Task No. Shaol k&Id hpe t,a, Cr.L<l - A.<<.r 2 n 9401-o u -i Title {, [4,',,( g/,w ((, f,7 73 g,;, f,,,,, Calculation No. I ~$W YeAccel Veldt 0$ Ele b thcl<sc$ Sk (ovd Preparer /Date Checker /Date Reviewer /Date Rev. No. fs/ s y; - ^Q^ jff v_ a-m o. $ee 0. S u e r v'c f tQttolq4 12/15/N i

MPR Associates, Inc. PR 32o xi#o street Alexandria, VA 22314 RECORD OF REVISIONS Calculation No. Prep ed By Checked By age 2919401-804 y [ Description Revision 0 Original Issue

MPR Associstas, Inc. l 320 King Street Alexandria, VA 22314 Calculation No. Prep ed By C eked By Page b 2919401-804 3Mc a 3 ./ L PURPOSE The purpose of this calculation is to determine the critical flaw size in the vertical welds of the FitzPatrick core shroud based upon linear elastic fracture mechanics. The loading and load paths used in this calculation are consistent with the design analysis of the FitzPatrick Core Shroud Repair.

2. RESULTS The critical flaw size in any vertical weld in the FitzPatrick shroud is 22.73 in. This is a conservative estimate calculated using the worst case loading for any shroud segment.

Consequently, the result is applicable to all vertical welds in the FitzPatrick shroud.

MPR Associntos, Inc. 320 King Street Alexandria, VA 22314 Calculation No. Prep ed By Checked By Page 4 2'919401-804 sex 7

3. DISCUSSION 3.1 Shroud Confiruration The FitzPatrick shroud has eight circumferential welds designated H1 through H8, which separate the shroud into a number of cylindrical segments. Each of these segments has two or more vertical welds which join individual plates to form the cylinders.

The shroud configuration is further modified by installation of the core shroud repair. The core shroud repair consists of ten tie rod restraint assemblies that structurally replace all the circumferential welds H1 through H8. Reference 1 shows the location and positioning of each tie rod assembly. Note that each shroud segment between two adjacent circumferential welds has a radial restraint to react possible lateral loads. 3.2 Applied IAeQ The vertical web in the shroud segments react circumferential loads through the shroud barrel. The following circumferentialloads are considered: Pressure-the pressure difference across the shroud creates a hoop load (stress) in the shroud which is reacted through vertical welds in each shroud cylinder. Lateral Loads-seismic events as well as a recirculation line break cause lateral loads in the shroud. These loads are reacted through vertical welds before they are transmitted into the core shroud repair's radial restraints. 3.3 Calculation Method and Assumptions Linear clastic fracture mechanics (LEFM) predicts the critical flaw size at which a crack becomes unstable. When the flaw size is greater than the critical size, the opening stress causes the crack to run uncontrollably through the remainder of the weld ligament. LEFM will be employed to determine the critical flaw size for any vertical weld in the Fitz?atrick shroud. LEFM relates the mean stress that opens a crack to the crack size and the failure criterion called the critical stress intensity factor. The mean stress is further modified by a stress concentration factor to determine the stress at the crack tip. The stress concentration factor depends on the crack size and geometry. The critical stress intensity factor is material specific and depends on a number of factors. For a known mean stress and critical stress intensity factor the critical llaw size can be determined by iteration.

MPR Associstos, Inc. 320 King Street Alexandria, VA 22314 Calculation No. Pre red By eked By h_ 5 2919401 804 f~ For conservatism,"any flaw in the vertical shroud welds is assumed to be through-wall. The mean stress opening cracks is the hoop stress in each shroud segment. The hocp stress is conservatively calculated as the sum of the stress caused by the pressure load and the stress caused by the lateral load component. The worst case pressure stress is determined using the maximum differential pressure between the reactor core region and the downcomer annulus for a given loading condition. The stress from lateral loads is conservatively calculated using the maximum load reacted by any core shroud repair radial restraint. The entire lateral load is assumed to be reacted through the length of the vertical weld, so the stress produced is equal to the lateral load divided.by the total weld area. As shown in the calculation section of Reference 2, the dominant stress-producing load for the most highly loaded shroud segment is the lateral carthquake load. Comparing the lateral loads between upset and faulted conditions, the maximum load increases by a factor of 1.7 between these conditions (Reference 5, Table 1). However, the safety margin between upset and faulted conditions decreases by a factor of 2 from 2.25 to 1.125 (Reference 3). Therefore, the maximum stresses that open a crack occur during the worst case upset loading condition (i.e., OBE loading). As a result, the critical flaw size will be calculated only for an OBE. Note that this is consistent with the GE screening criteria for axial flaws. For flaws to remain stable throughout an entire operating cycle, the flaw must be less than the critical crack size at the end of the operating cycle. Consequently, crack growth - - during an assumed operating cycle of two years is also considered. A bounding value for crack growth is chosen, and linear crack growth with operating time is assumed.

EMPR 72T a"&'2' '"- 2 Alexandria, VA 22314 Calculation No. Prep ed By Checked By N 6 2919401-804 7_

4. CALCULATION 4.1 Membrane Stress The membrane stress is calculated as the sum of the pressure hoop stress and the stress from the lateral load component. The pressure hoop stress is determined as follows:

op (dP.R) / (1000 t) = Where: op Hoop stress due to differential pressure across the shroud (ksi) = Maximum differential pressure across the shroud for an OBE event (psi) dP = 33.93 psi (Reference 4, Appendix A) dP = Outside radius of shroud (in) R = 94.75" (Reference 1) R = t = Thickness of the shroud (in) t = 1.50" (Reference 1) Substituting, the pressure hoop stress becomes: op = [(33.93 psi) (94.75")) / [(1000 lbf/ kip) (1.50")] op = 2.143 ksi The hoop stress from lateral loading is based upon the worst case OBE radial restraint i load. From Reference 5, the largest radial load is in the upper radial restraint and is 169 kips. The upper radial restraint load is reacted along the H1-H2 shroud cylinder and is given by: = F / (H t) at g Where: at = Hoop stress in shroud due to lateral loading (ksi) ~ ~ Fa = Maximum load reacted by a radial restraint during an OBE event (kips) Fa = 169 kips (Reference 5, Table 1) H = Height of shroud segment reacting the load (in) H = 32.00" (Reference 1) t = Thickness of sluoud segment (in) t = 1.50" (Reference 1)

MMPR "f!#;s".':1 2 Alexandria, VA 22314 Calculation No. Pre red By Checked By 2919401 804 E; -- 7 Substituting, the lateral hoop stress becomes: at = (169 kips) / [(32.00") (1.50")) at = 3.521 ksi The total membrane stress is the sum of the pressure hoop stress and the lateral hoop stress and is: og =ap+on Where: og Maximum membrane stress in shroud which opens a vertical flaw (ksi) = Substituting: 2.143 ksi + 3.521 ksi ~ og = 5.664 ksi og = 4.2 Stress Intensity Factor The stress intensity factor for axial Daw in a thin walled cylinder under uniform membrane hoop stress is given by: (Gm + Gb) og (n a)o.5 K = i Where: K = Stress intensity factor for Mode I fracture (ksivin) i Gm = Stress concentration due to membrane stress Gb = Stress concentration due to bending stress og = Membrane stress that opens the crack (ksi) a = Half of the axial flaw size (in) Experimental correlations for the stress concentration at the crack tip are taken from Reference 6 for an axial flaw in a thin walled cylinder under uniform membrane hoop stress. The stress concentrations due to membrane and bending are shown below. l l

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MPR Associct;s, Inc. 320 King Street Alexandria, VA 22314 Calculation No. Pre red By eked Page - 7f,u/77w ~ 9 2919401-804 43 Critical Stress Intensity Factor The critical Daw size can be determined by setting the above equation for the stress intensity factor of a flaw to the critical stress intensity factor (K c). The critical stress I intensity factor is a material property which varies with the operating environment. For a shroud in a foreign GE BWR with an operational history that is similar to FitzPatrick, the critical stress intensity factor was determined to be 150 ksivin (Reference 3, Section 3.2). For an OBE event (an upset loading condition), a safety factor of 2.25 is applied to the critical stress intensity factor which reduces the allowable to 66.7 ksivin (Reference 3, Section 3.2). 4.4 Critical Flaw Size Using a hoop stress of 5.664 ksi and an allowable stress intensity factor of 66.7 ksivin, iteration on the equation for the stress intensity factor yields the critical flaw size as 24.48". This result can be veri 5ed by substituting these values into the original equations as follows: K = (Gm + Gb) om(ra)05 i The nondimensional crack size is: a / (R t)05 = (24.48"/2) / [(94.75") (1.50")]05 a / (R t)05 = 1.03 Using the nondimensional crack size, the stress concentrations for membrane and bending become: Gm = 1.572 Gb = 0.326 Substituting, ~ K = (1.572 + 0.326) (5.664 ksi) [n (24.48"/2)]05 i K 66.7 ksivin = i Note that the stress intensity factor is equal to the allowable stress intensity factor. Therefore, the critical flaw size for any vertical weld in the FitzPatrick core shroud is 24.48".

Eil5 M P R L" #; n'e':," '"- 2 Alexandria, VA 22314 Calculation No. Pre red By Checked By hh 10 2919401-804 4.5 Crack Growth Assuming the same bounding crack growth rate as in the GE Screening Criteria for FitzPatrick (Reference 3, p. 4 and Figure 7-1), the crack growth rate is 5.0 x 10-5 g, For linear crack growth over a two year operating cycle (17,520 hr), the maximum crack extension in a single direction is 0.876 in. The critical flaw size for axial flaws is calculated in Section 4.4 of this calculation. Assummg linear crack growth over the entire operating cycle and crack extension at both crack tips, the critical flaw size in the axial welds for continued operation is determined as follows: (Critical flaw size from Section 4.4) - ~ Required Length = 2 (Maximum crack extension) 24.48" - 2 (0.876") Required Length = 22.73" Required Length =

5. REFERENCES 1.

MPR Drawing No. 1291-001-17, " James A. FitzPatrick Nuclear Power Plant, Shroud Contingency Repair, Shroud Restraint Assembly", Rev. A. 2. MPR Calculation No. 2919401-801, " Required Intact Vertical Weld Area Based Upon the Limiting Applied Loads", Rev. O. 3. GE Nuclear Energy Report No. GENE-523-154-1093, " Evaluation and Screening Criteria for the FitzPatrick Shroud Indications", October 1993. 4. MPR Specification No. 291-001-001, " Design Specification for James A. FitzPatrick Nuclear Power Plant (JAF) Core Shroud Repair", Rev. O. 5. MPR Calculation No. 2919401-741, ' Tie Rod Resuaint Limiting Reactions Using Dynamic Analysis Seismic Imads", Rev. O. 6. Rooke, D. P. and Cartwright, D. J. Compendium of Stress Intensity Factors. Her Majesty's Stationery Office, London,1976, i i

BBMPR A S S O CI ATE S INC. ENGINEER $ e APPENDIX D MPR Calculation No. 2919401-805, " Critical Maw Size for Indications in Circumferential Welds H2 (or H3)", Revision 0 D-1

l MPR Associates, Inc. ^55 320 King Street Alexandria, VA 22314 ( N CALCULATION TITLE PAGE Client Ne-Yark Po ~ e, dudo<,0 Page 1 of lO y Project Task No. bhroud klcId NSfaYon (rttric-hhesc.~{ 2M-4 401 - 046-l Title [r,'4,'ge l F)c y 6,, e fy fy, Calculation No. 0'(Cum _hren f a l Wfl$ $1 (or 03) ~ Preparer /Date Checker /Date Reviewer /Date Rev. No. EEE%e P4MANATA h SCL 3, \\Ll%% l% p g ( Gr 'e P. M anner 2/27jqq '*l*9h'

MPR Associates, Inc. D 320 King Street ^ss C Alexandria, VA 22314 iNG t wa RECORD OF REVISIONS Calculation No. Pre red By Checked By age 2919401-805 2 ,,u, .,n ev ; xw ~ Revision Description 0 Original Issue l ~ l i l I L

MPR Associates, Inc. 320 King Street e55 Alexandria, VA 22314 ENG6NEER5 Calculation No. Pre ed By ecked By Page 2919401-805 //"- 3

1. PURPOSE The purpose of this calculation is to determine the critical flaw size in the circumferential welds of the top guide support ring in the FitzPatrick core shroud based upon linear clastic fracture mechanics.

The loading and load paths used in this calculation are consistent with the design analysis of the FitzPatrick Core Shroud Repair (Reference 1).

2. RESULTS The critical flaw size in circumferential welds in the top guide support ring in the FitzPatrick shroud is 115.74 in. This is a conservative estimate calculated using the worst case loading and weld conditions.

e oga = -p-

MPR Associates, Inc. E 320 King Street Alexandria, VA 22314 ^ Calculation No. Prep ed By Checked By 2919401-805 emu -_ 4

3. DISCUSSIO 3.1 Shmud Confiruration The FitzPatrick shroud has eight circumferential welds designated H1 through H8, which separate the shroud into a number of cylindrical segments. Each of these segments has two or more vertical welds which join individual plates to form the cylinders.

The shroud configuration is further modified by installation of the core shroud repair. The core shroud repair consists of ten tie rod restraint assemblies that structurally replace all the circumferential welds H1 through H8. Reference 2 shows the location and positioning of each tie rod assembly. Note that each shroud segment between two adjacent circumferential welds has a radial restraint to react possible lateral loads. 3.2 Applied Loads The circumferential welds in the top guide support ring react the vertical loads transmitted between the upper shroud region and the lower shroud region. Specifically, the H2 and H3 welds carry the loads from the steam separators and the tie rod preload. 33 Calculation Method and Assumptions Linear clastic fracture mechanics (LEFM) predicts the critical flaw size at which a crack becomes unstable. When the flaw size is greater than the critical size, the opening stress causes the crack to run uncontrollably through the remainder of the weld ligament. i LEFM will be employed to determine the critical flaw size for the H2 and H3 circumferential welds in the FitzPatrick shroud. i LEFM relates flaw size, geometry, and mean applied stress to the applied stress intensity factor._ The applied stress intensity factor is compared to the critical stress intensity. factor for the material. For a known mean stress and critical stress intensity. factor the _. critical flaw size can be determined by iteration. The preload as well as the loads due to the steam separators are transmitted across the top guide support ring through bending in the H2 and H3 circumferential welds. The bending stresses in the welds are determined in References 3 through 6 for all possible failure conditions. As discussed in Section 3 of Reference 7, a portion of the circumferential weld around a failed radial weld in the top guide support ring must be intact for acceptance of the failed radial weld. Consequently, any flaws in the required circumferential weld must be less than the critical flaw size for the circumferential weld. The highest stresses in an

J

20 K ng Street Alexandria, VA 22314 ENGINEERS Calculation No. Prep ed By Checked By Page 5 Afe- _ 2919401-805 j 7 intact circumferential weld occur when the opposite circumferential weld has failed completely. Further, after examination of References 4 and 5, the highest stresses occurring in either top guide support ring weld is at H3 after H2 has failed 360*, through-wall. For conservatism, any flaw in the circumferential welds is assumed to be through-wall. Comparing the vertical loading across the top guide support ring between upset and faulted conditions, the maximum load increases by a factor of 1.2 between these conditions (Reference 8, Table 2). However, the safety margin between upset and faulted conditions decreases by a factor of 2 from 2.25 to 1.125 (Reference 3). Therefore, the maximum stresses that open a crack occur during the worst case upset loading condition (i.e., OBE loading). As a result, the critical flaw size will be calculated only for an OBE. Note that this is consistent with the GE screening criteria for unrepaired shroud. For flaws to remain stable throughout an entire operating cycle, the flaw must be less than the critical crack size at the end of the operating cycle. Consequently, crack growth during an assumed operating cycle of two years is also considered. A bounding value for crack growth is chosen, and linear crack growth with operating time is assumed. o

MPR Associates, Inc. 320 King Street Alexandria, VA 22314 r N s i u e is k Calculation No. - Prep d By C eked By Page 2919401-805 -s/<h-6

4. CALCULATION 4.1 Bending Stress The bending stress in the circumferential welds is calculated for OBE loading using the same approach as described in Reference 4. The bending stress is given as:

(0.0245 F - 0.0140 AP. A) (1 kip /1000 lbf) (Reference 4, p. 7) og = 3 Where: og Uniform bending stress which opens a crack in the H2 or H3 weld (ksi) = Applied vertical load across the top guide support ring (Ibf) F = 3 369,400 lbf (Reference 8, Table 2) F = g Maximum differential pressure across the top guide support ring AP = 7.79 psi (Reference 9, Appendix A) AP = Area where differential pressure acts A = 2 3402 in (Reference 4, p. 6) A = Substituting, the maximum bending stress in either H2 or H3 during an OBE event becomes: 2 og [ (0.0245) (369,400 lbf) - (0.0140) (7.79 psi) (3402 in )] (1 kip /1000 lbf) = 8.679 ksi og = 4.2 Stress Intensity Factor The stress intensity factor for a flaw in a thin walled cylinder under a uniform bending stress is given by: (Gm + Gb) og (n a)05 (1 + v) / (3 + v) (Reference 11, Figure 202) -- - K = i Where: Ki Stress intensity factor for Mode I fracture (ksivin) = Gm Stress concentration due to membrane stress = Gb Stress concentration due to bending stress = og Bending stress that opens the crack (ksi) = Half of the axial flaw size (in) a = Poisson's ratio y = 0.3 (assumed) y =

MPR Associates, Inc. R 32o xi#a street Ass Alexandria, VA 22314 i n s,,4 eaas Calculation No. Pre ed By ecked By Page 7 291!T401-805 ~* mf Experimental correlations for the stress concentration at the crack tip are taken from Reference 10 for a flaw in a thin walled sphere under a uniform bending stress. Note that the stress concentration for a thin walled cylinder would be appronmately the same. The stress concentrations due to membrane and bending are shown below. _ -.~. g..+_,,,._. f... p o. _ p~~ +- 1-... - -3 . ~ _ . ~.. .l..,... ~~ . ~.,e. &.. *. 4. o

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+4++

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eWar 4.-., .~.{v... ..~{.. a ~{.. u T ...j 00 04 04 1-2 14 2-0 24 2.s 32 t/MI 04 ym pm A of a md es a $mm e@psuno a mfam W mm D N 0 O h [hd[ ( Qk 4 l l l 1 k

MPR Associates, Inc. 320 King Street y, Alexandria, VA 22314 ,u ,yggy3 Calculation No. Pr red By Checked By Page 2919401-805 . - ygm 8 43 Critical Stress Intensity Factor The critical flaw size can be determined by setting the above equation for the stress intensity factor of a flaw to the critical stress intensity factor (K c). The critical stress I intensity factor is a material property which varies with the operating environment. For a-shroud in a foreign GE BWR with an operational history that is similar to FitzPatrick, the critical stress intensity factor was determined to be 150 ksivin (Reference 7, Section i 3.2). For an OBE event (an upset loading condition), a safety factor of 2.25 is applied to the critical stress intensity factor which reduces the allowable to 66.7 ksivin (Reference 7, Section 3.2). 4.4 Critical Flaw Size Using a bending stress of 8.679 ksi and an allowable stress intensity factor of 66.7 ksivin, iteration on the equation for the stress intensity factor yields the critical flaw size as 117.49". This result can be veri 5ed by substituting these values into the original equations as follows: K = (Gm + Gb) om (n a)05 (1 + v) / (3 + v) i For a outer radius at H3 of 88.75" and a shroud thickness of 1.50" (Reference 2), the nondimensional crack size is: a / (R t)05 = (117.49"/2) / ((88.75") (1.50")]05 a / (R t)05 = 5.09 Conservatively assuming that the stress concentrations increase linearly past a nondimensional crack length of 3.2, the membrane and bending stress concentrations become: 1.263 ~ ~ ~ Gm = Gb = 0.173 ~ Substituting, K = (1.263 + 0.173) (8.679 ksi) [n (117.49"/2)]05 [(1 + 0.3) / (3 + 03)] i K = 66.7 ksivin i Note that the stre:S intensity factor is equal to the allowable stress intensity factor. Therefore, the critical flaw size for weld H2 or H3 in the FitzPatrick core shroud is 117.49". i

I MPR Associates, Inc. {) 320 King Street A (N 6 Alexandria, VA 22314 Calculation No. Pre red By Checked By gu_ 9 2919401-805 4.5 Crack Growth Assuming the same bounding crack growth rate as in the GE Screening Criteria for FitzPatrick (Reference 7, p. 4 and Figure 7-1), the crack growth rate is 5.0 x 10-5 g, For linear crack growth over a two year operating cycle (17,520 hr), the maximum crack extension in a single direction is 0.876 in. The critical flaw size for circumferential flaws in either H2 or H3 is calculated in Section 4.4 of this calculation. Assuming linear crack growth over the entire operating cycle and crack extension at both crack tips, the critical flaw size in the H2 or H3 weld for continued operation is determined as follows: (Critical flaw size from Section 4.4) - Required Length = 2 (Maximum crack extension) 117.49" - 2 (0.876") Required Length = Required Length 115.74" = l i ee w

-p es.

-ae L.

I a MPR Associates, Inc. R "'" a S* *' Alexandria, VA 22314 s n c.. u e e a s Calculation No. Prep ed By Checked By p Awm_ V 10 2919401-805 7

5. REFERENCF3 1.

MPR Report 1560, " James A. FitzPatrick Nuclear Power Plant Core Shroud Repair - Design Report", October 1994, Rev. O. 2. MPR Drawing No. 1291-001-17, " James A. FitzPatrick Nuclear Power Plant, Shroud Contingency Repair, Shroud Restraint Assembly", Rev. A. 3. MPR Calculation No. 2919401-501, " Determination of Maximum Allowable Vertical Imads on the Shroud Ledge Assuming Failure of H2, H3, and H6B", Rev. 2. 4. MPR Calculation No. 2919401-502, " Determination of Maximum Allowable Vertical Loads on the Shroud Ledge Assummg Failure of H2 and H6B", Rev.1. 5. MPR Calculation No. 2919401-503, " Determination of Maximum Allowable Vertical Loads on the Shroud I. edge Assummg Failure of H3 and H6B", Rev.1. 6. MPR Calculation No. 2919401-504, " Determination of Maximum Allowable Vertical Imads on the Shroud Ledge Assuming Failure of Weld H6B Only", Rev.1. 7. GE Nuclear Energy Report No. GENE-523-154-1093, " Evaluation and Screening Criteria for the FitzPatrick Shroud Indications", October 1993. 8. MPR Calculation No. 2919401-803, " Initial Acceptance Criteria for Radial Welds with Indications", Rev. O. 9. MPR Speci5 cation No. 291-001-001, " Design Specification for James A. FitzPatrick Nuclear Power Plant (JAF) Core Shroud Repair", Rev. O.

10. Rooke, D. P. and Cartwright, D. J. Compendium of Stress Intensity Factors.

Her Majesty's Stationery Office, lendon,1976. i.

i ATTACHMENT 3 to JPN-95-016 J. A. FitzPatrick Plant EDM Surfaces for Shroud Repair New York Power Authority JAMES A. FITZPATRICK NUCLEAR POWER PLANT Docket No. 50-333 DPR-59 L

r h EMPR A SS O CI AT E S IN C. ENGINEERS J. A. FITZPATRICK PLANT i EDM SURFACES FOR SHROUD REPAIR r t PURPOSE MPR has reviewed photomicrographs of electro-discharge machined (EDM) surfaces representing the upper (shroud flange) and lower (gusset plate) tie rod repair attachment surfaces in the FitzPatrick plant. The gusset plate surfaces are ball honed after EDM. This document records our evaluation of these surfaces for long-term BWR service. i DISCUSSION The upper attachment surface contains a seating depression and central pin trepanned by EDM out of the existing shroud flange. The material is solution annealed Type 304 stainless steel. The lower attachment surface is a hole in the existing NiCrFe Alloy 600 gusset plate that is produced by EDM followed by ball honing with 120 grit silicon carbide balls. Between 1 and 2 mils is removed from the surface by the honing pr, ess. Based on the vessel fabrication history, the Alloy 600 base matedal is furnace sensitized. The EDM surfaces were sectioned, etched and examined at 10X in accordance with the draft requirements of the ASME Code, Section XI Working Group on Welding and Other Special Processes. No cracking was observed by BWNT. Subsequent inspections were performed at 200X (not required by the draft code guidelines), and results are described below. A. Upper Attachment Surface The EDM process produces a surface, when examined at about 200X magnification, that contains the following features: 1. A variable thickness recast layer averaging 1.5 mils thick but containing isolated, local areas up to 10 mils thick. Isolated globules of recast material up to 10 mils in diameter are sometimes observed to be loosely attached to the 1 surface. The globules appear to be portions of the recast layer that have spalled from the surface. 2. Isolated craters up to 4 mils deep penetrating into the base metal. Craters, unlike fissures, are round-bottomed and have aspect ratios that vary from narrow " wormhole" types to broad depressions. 1 \\

t:* I i 3. Isolated fissures in the recast layer up to 6 mils deep. Fissures do not ' i penetrate into the base metal. i The acceptability of each of these features is discussed in the following section. I t 4 Recast Imer 'Ibickness We judge the thickness of the recast layer and the variability of the thickness will l have no significant impact on the serviceability of EDM surfaces in the upper i attachment area. The recast layer itself has a cast or weld metal microstructure which, for actual welds, has been shown to be no more susceptible to IGSCC than i solution annealed base n etal. The recast material is likely to have retained some i ferrite in its microstructure, which dramatically increases its stress corrosion cracking resistance compared to base metal The excellent record of stainlen steel weld. metal performance in SWR service supports the expectation that the recast layer will be highly stress concosion cracking resistant. L Besides having an acceptable microstructure, the EDM recast layer is likely to impose low residual stresses and no cold work on the base metal beneath it. The layer was formed by melting rather than by working and cutting as in normal machining, and its extreme thinness limits the ability of the recast layer to impose - i significant stresses on the base metal Local areas of the recast layer that develop high stresses are likely to fail by fissuring upon solidi 6 cation or thereafter, thereby l relieving the stresses. Furthermore, the rapid quenching of the recast metallimits i the degree of sensitization of the recast layer and the base metal heat affected zone to very low values. We judge the recast surface susceptibility to IGSCC to be less than or equal to a well-machined surface. I Crateriar j Isolated craters in the EDM surface will create local crevice conditions on free surfaces, but these are no different in principle than any other crevice in tlm reactor j pressure vessel. Since craters are without high residual stresses that penetrate the -l ~ base metal significantly (more than a few mils) or surface cold work, they should not be cause for enhanced IGSCC susceptibility. On this basis, we judge that isolated j cratering is acceptable in an EDM surface that is not near welds (which cause j residual stresses and sensitization) or areas of high applied stress. ' j i Fissures Observed fissures in the recast layer appear to have resulted from shrinkage stresses f that appeared when the layer solidified. The fissures are isolated and do not fully penetrate the recast layer. The fissures are crevices, and could be expected to propagate in service very slowly by crevice corrosion.. In the event a fissure penetrated the recast layer, it is unlikely to align with a grain boundary in the base metal. Therefore most fissures, even if they propagate, are expected to arrest near the base metal interface., i L

.. s' l While it is unhkely for IGSCC-resistant recast material, fissures could cause the l initiation of IGSCC due to the local high stress at the fissure tip. 'Ihe parameter characterizing the crack tip stress field is the stress intensity factor, which depends on i part geometry, fissure size and orientation, and applied stresses. l 1 For the upper attachment geometry, the most significant tensile stresses result from. the radially outward load placed by the tie rod on the attachment post, causing a net moment and bending stren at the base of the post. Stress corrosion cracking at this location could cause post failure, but would not cause the tie rod to separate from the shroud due to the redundant retention provided by the edge of the EDM hole { surrounding the post. However, for purposes of calculation, estimates of streu intensity factors due to fissures in the post surface and applied post stresses will be .i used to estimate the acceptability of the fissures. l For a post with a shallow circumferential crack on its OD surface and loaded in bending, the stress intensity factor is: { i i ab /rd 1.122, Reference (1) l K = i where fissure depth (inches) d = ab nominal bending stress (ksi) { = t The nominal, steady state bending stress in the post is 13.1 ksi, Reference (2). For a l 20 mil fissure extending through the thickest recast layer observed and wellinto the base metal, the stress intensity factor is: l 13.1/n0.020 1.122 Kg = 3.7 ksi/in l = This value is far below the stress intensity factor needed to initiate IGSCC in Type 304 stainless steel under BWR conditions. A value of 8 ksi/in is needed for initiation of IGSCC, as reported in NUREG 0313, Rev. 2, Reference (3). B. Imer Attachment Surthce l The EDM surface and when viewed at about 200X after honing has the following j features: 1. Where it exists, a very shallow recast layer averaging about 0.1 mils in thickness with isolated, local areas 4 mils in thickness. No fissures have been observed in the recast layer. \\

i N z,= 2.- Shallow, broad depressions ranging from 1 mil to about 5 mils in depth. The depressions often contain a recast layer 0.1 mils thick. t 3. Shallow intergranular attack in the base metal in the bottom of isolated ( depressions where honing could not reach. The attack appears to be the result D of chemical action during the EDM process and is quite shallow, less than 1 mil deep. 4. Most surfaces show no evidence of cold work due to honing. A few areas / showed grain distortion about 0.3 mils deep. This is comparable to the cold work depth in a carefully machined surface. l 7 iL i With the possible exception of the intergranular attack, the EDM surface on the. Alloy 600 materialis of excellent quality, comparable to a machined surface. He i isolated intergranular attack is judged to be acceptable on the basis that it is below 1 mil in depth, a value which is believed by the industry to be inconsequential At thh i depth, the streu intensity factor at the crack tips is considerably below the threshold. l for s'.ress corrosion cracking based on an evaluation provided by BWNT, j Reierence (4). In Reference (4), BWNT shows a steady operating stress intensity factor of only 0.% ksi/in, far below our estimated threshold for stress corrosion cracking, about 8 ksi/in for NiCrFe Alloy 600. CONCLUSIONS l 1. Fissures observed in the recast layer of the stainless steel EDM surface are not expected to propagate during long-term BWR service. The variable thickness of the recast layer and the presence oflocal cratering or pits are also not expected to degrade the stress corrosion performance of the upper attachment because of the lack of residual stresses that penetrate to a signi5 cant depth or cold work in the EDM surface. These conclusions are based on our engineering judgement and a fracture mechanics evaluation. It is also our opinion that the alternative.of using a j honed EDM surface for the complicated upper attachment geometry may not result j in any significant improvement in the performance of the surface, and if not carefully i controlled, may result in increased corrosion concerns due to the presence of cold j work. Therefore, we consider the EDM surface to be preferred, without honing, for i the upper attachment surface. 2. The honed EDM surface of the NiCrFe Alloy 600 is, in our judgement, of acceptable quality and presents no special concern for stress corrosion cracking during long-term BWR service. We consider that the use of light honing, as presently planned, is prudent since it appears to remove intergranular attack from most surfaces without introducing significant cold work. We would not recommend increasing the amount of metal removal by honing with coarse grits and increased metal removal rates because of the increased risk of causing generalized surface cold work with the resulting increased sensitivity to stress corrosion cracking.

v REFERENCES 1.

H. Tada, P. Paris and G. Irwin, The Analysis of Cracks Handbook (Del Research Corporation, St. Louis) 1985., p. 27.2. 2. MPR Calculation 2919401-204, " Bending Stress in Shroud Flange Bracket Attachment Posts," EW. McCurdy, December 29,1994. 3. NUREG 0313, Rev. 2, ' Technical Report on Material Selection and Processing Guidelines for BWR Coolant Pressure Boundary Piping," January 1988. 4. BWNT Calculation 32-1235178-00, "FitzPatrick Gusset Plate Hole, FM Evaluation," December 1994. 5 l l I 4 i i.}}

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