ML20072L963
| ML20072L963 | |
| Person / Time | |
|---|---|
| Site: | Peach Bottom  |
| Issue date: | 12/13/1993 |
| From: | Marisa Herrera, Ranganath S GENERAL ELECTRIC CO. |
| To: | |
| Shared Package | |
| ML20072L561 | List: |
| References | |
| GENE-523-176-12, GENE-523-176-1293, NUDOCS 9409010150 | |
| Download: ML20072L963 (91) | |
Text
{{#Wiki_filter:_. . ._ GENE-523-176-1193 DRF 137-0010-6 Evaluation and Screening Criteria for'the Peach Bottom Unit-2 Shroud December 13,1993 l l
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Prepared by: /sw/8 plarcos L. Herrera, Principal Engineer l l Stmetural Mechanics Projects l Approved By: # [* %O Dr. Sampath Ranganath, Manager Stmetural Mechanics Projects GE Nuclear Energy San Jose, CA 9409010150 941213 PDR ADOCK 05000277 PDR p u.
GENE-sal-i76-1293 GENuclear Enero lbfPORTANTNOTICE REGARDING CONTENTS OF TIIIS REPORT Please Read Carefully The only widertakings of the General Electric Company (GE) respecting infortnation in this document are contained in the contract between Philadelphia Electric Co. and GE, and nothing contained in this document shall be construed as changing the contract. The use of this information by anyone other than PECo, orfor any purpose other than thatfor which it is intended under such contract is not authori:ed; and with respect to any unauthori:ed use, GE makes no representation or warranty, and assumes no liability as to the completeness, accuracy, or usefidness of the information contained in this document, or that its use may not infringe privately owned rights. l
5 GENE.32317&l293 GE Nuclear Energy Table of Contents EXECUTIVE
SUMMARY
.. 1 l
1.0 INTRODUCTION
1.1 Background . .. . . . . .3 1.2 Screening Criteria... . .
. .. .. . . , .3
. .6 1.3 References . . .
I1 2.0 FABRICATION HISTORY . 12 2.1 Summary.. . .. .
.24 i 3.0 CHEMISTRY AND FLUENCE CONSIDERATIONS.
3.1 Water Chemistry History.. . . .24
.26 3.2 Fluence Considerations.. .
28 3.4 References . . . 38 4.0 IN-VESSEL VISUAL INSPECTION OF PEACH BOTTOM UNIT-3.
. .. 40 5.0 JUSTIFICATION FOR CONTINUED OPERATION.
40 5.1 Structural Analysis. .. . .. . aa l 5.2 Allowable Through-Wall Flaws. . . .. . . . .
. 47 5.3 Screening Criteria.. .
49 5.4 Summary of Screening Criteria... . 50 5.5 Justification for Continued Operation. . . 51 5.6 References .
. . 5:
6.0
SUMMARY
AND CONCLUSIONS. . . APPENDIX A DETERMINATION OF THE EFFECTIVE FLAW LENGTH APPENDIX B B ASIS FOR CRACK GROWTH RATE
GENuclear Enugy_ _ GENE 523-n6.nn i EXECUTIVE
SUMMARY
I I This report documents the evaluation of the Peach Bottom Unit-2 shroud based on GE l SIL 572, Rev,1 recommendations and comparison with the Peach Bottom Unit-3 shroud. 1 Justification for continued operation of Unit-2 was based on evaluating Unit-2 against i Unit-3 water chemistry, fluence and fabrication information. The comparison showed that the Unit-2 conditions were less severe than that at Unit-3 from a stress corrosion viewpoint. Therefore, it is reasonable to use the Unit-3 IVVI results for Unit-2 as bounding. Based on this, continued operation of Unit-2 for one more cycle isjustified. Indications have been observed in the Peach Bottom Unit 3 core shroud. Indications were ' seen during in-vessel visual inspection (IVVI) of the various shroud welds as recommended by GE SIL 572, Rev.1, Results of the Unit-3 shroud IVVI showed that both circumferential and axialindications were present at the H3 and H4 welds. H3 corresponds to the weld between the top guide support ring and core shroud cylinder, and the H4 weld is located at approximately the mid-height of the fuel. In addition, circumferential indications were observed in the shroud plate associated with a vertical ! weld. The lengths of the indications associated with the vertical welds were short (22.5" ) max.) compared to those associated with the horizontal welds. This evaluation was performed to demonstrate that the stmeturalintegrity of the Unit-2 shroud is maintained for the current fuel cycle (two year cycle with power rerate . conditions) even with postulated indications similar to that in Unit-3. In addition, the J report documents material, water chemistry and fluence information which show that Unit-2 is likely to be better than Unit-3 from the stress corrosion viewpoint. i This report demonstrates that even with several conservatisms in the evaluation, the : structuralintegrity of the shroud is maintained during a limiting event, even with the j postulated indications. This was performed by developing conservative screening criteria, assuming throughwallindications, which can determine the acceptability of the flaws based i solely on the IVVI results. The assumption of through-wallindications removes any 'l uncertainty regarding sizing and the need to further characterize the indications. By
)
meeting the screening criteria, the ASME Code Section XI safety margins are satisfied- ' The screening criteria use both linear elastic fracture mechanics (LEFM) and limit load concepts to determine acceptable through-wallindication lengths. The limiting flaw length-based on either LEFM or limit load was used for the screening criteria. , iv l
]
J
1. _ GENE-523-1761:e3 GENuclear Encru The screening criteria also use the ASME Code Section XI criteria for combining flaws based on the proximity ofindications. In addition, a second method for including the interaction between neighboring indication tips was considered for the LEFM allowable flaw size calculation. The resulting effective flaw lengths were compared against the screening criteria to determine if the structural integrity of the shroud was maintained. Based on the results of the application of the screening criteria to the observed indications, it is concluded that the structural integrity of the Unit-2 shroud is maintained for the current fuel cycle even with postulated flaws similar to that in Unit-3. All effective indication lengths were shown to be less than the allowable flaw size. The use of the Unit-3 indications was justified based on a comparison ofwater chemistry, fluence and fabrication features between the two plants. V
GENE 523-176-1293 GENuclear Energy
1.0 INTRODUCTION
This report documents the evaluation of the Peach Bottom Unit-2 shroud based on GE SIL ',72, Rev.1 (Reference 1-1) recommendations. The objective of this evaluation is to justify continued operation for the Peach Bottom Unit-2 plant. Visualinspection of the Peach Bottom Unit-2 shroud has not been performed. Thus, justification for continued operation of Unit-2 will be based on evaluating Peach Bottom Unit-2 against the recommendations provided in Reference 1-1 along with a comparison with Unit-3 water chemistry, material and fabrication information, and application of the Unit-3 IVVI results to Unit-2. Due to the many similarities between Units 2 and 3, it is reasonable to use this comparison as a basis for assessing the condition of the Unit-2 shroud. l l Recently, visualinspection of the Peach Bottom Unit-3 plant revealed indications at the j H1, H3, H4, and V3 weld locations. Figure 1-1 is a schematic illustrating the locations of , the shroud welds in the Unit-2 shroud which are the same for Unit-3. Details of the Peach Bottom Unit-3 IVVI results are discussed in Section 4.0 and Reference 1-2. Figures 1-2 through 1-4 are schematics of the observed indications in the Unit-3 shroud. GE SIL 572, Rev.1, provides the following recommendations based on the observed indications and evaluations performed to date: ) 1 l Plant Fabrication and Operational Historv Review plants fabrication and operational histories for the core shroud, including the materials of constmetion. Non-Destructive Examination Actions Visual examinations of accessible areas should be performed on the shroud ID and OD surface at the next scheduled refueling outage for all plants with Type 304 stainless steel shrouds with six or more years of power operation, and for all plants with L-grade stainless steel shrouds with eight or more years of power operation. These examinations should be performed with an enhanced VT-1 system or a qualified UT examination from the outer surface. Ifindications are not observed, examination should be performed at every second refueling outage. Ifindications are observed, the shroud should be examined and lengths measured 1
}
g , ,. ,g , aExe-s231161293 l during each refueling outage. The SIL also provides a reconunended examination process. l Destructive Testine A boat or core sample may be necessary depending on the results of the examination. Structural Marcin Analysis Perform a structural margin analysis using the results from the NDE, and, if performed, the destructive analysis. If numerous indications are observed, the need for corrective action can be assessed using cumulative flaw length structural margin criteria. Corrective Action Based on the results of the structural margin evaluation, determine if continued operation is justified for another cycle without repair. If cracking is found and sufficient structural margin remains, examine the shroud during each subsequent refueling outage. This report provides the pertinent information required to demonstrate that continued operation of Peach Bottom Unit-2 is justified based on the SIL recommendations noted above. Specifically, the report presents the following information:
. Fabrication history of the shroud.
. Water Chemistry and Fluence Considerations
. In-Vessel Visual Inspection
. Stmetural Margin Analysis
. Screening Criteria for Application to IVVI results.
A It is noted that the loads used in this evaluation correspond to those for power rerate two-year operating cycle was used in the determination of crack growth. 1 l 2 I
GENE-5:31761393 GE Nuclear Enerv 1.1 Background Indications have been observed in the shrouds of three plants to date (including Peach Bottom Unit 3). Cracking was observed in a BWR/4 located outside the United States in 1990. The cracking was confined to the heat affected zone (HAZ) of a circumferential weld. In 1993, the second occurrence of cracking in a shroud was reported. Cracking was observed on the inside surface (ID) of the top guide support ring near the H3 weld. The cracking was approximately 360 around the circumference, in the weld heat affected zone (HAZ), in a material with carbon content of 0.06%. The fluence was estimated as 1.8x10 20 nyt (E>lMev). In addition to the H3 weld HAZ cracking, indications were visually observed at the H1, H2, H4, H5 (shroud cylinder) and, H6a weld HAZ (at core plate support ring). Indications were seen mostly on the inner surface at H3, H4 and H5. Indications were seen on the outer surface at the H1, H2 and H6 welds. 1.2 Screening Criteria IVVI provides the length characterization of any present indications. Given that non-destructive examination (NDE) of every visually detected indication could be difficult and time consuming, a method of screening indications for subsequent evaluation is required This report presents such a screening criterion. The guiding parameter used for the selection of the indications for further evaluation is the allowable through-wall flaw size, which already includes the safety factors. If all of the visually detected indications are assumed to be through-wall, then the longest flaws, or combination of flaws, would have the limiting margin against the allowable through-w all flaw size. In reality, the indications are likely not through-wall, and therefore, the criteria and methods presented in this report are conservative. The result of this procedure will be the determination of the effective flaw lengths which will be used to compare against the allowable flaw size and selection ofindications for more detailed evaluation. The determination of effective flaw length is based on ASME Code, Section XI, Subarticle IWA-3300 (1986 Edition) proximity criteria. These entena 3
GENE 3D.176.In) GE Nuclear Ewgy provide the basis for the combination of neighboring indications depending on various geometric dimensions. Crack growth over a subsequent two year operating and power rerate cycle is factored into the criteria. This is conservative since power rerate will not be in effect during the next fuel cycle. The proximity rules described here also conservatively assume that there is interaction between two perpendicular flaws. It is assumed that circumferential and axialindications could increase the effective flaw length depending on the unflawed distance between them, This effective circumferential flaw length must be compared against the allowable circumferential flaw length. The axial flaw would be compared against the allowable axial flaw length. Flaws are considered in the same plane if the perpendicular distance between the planes is 4" or less. Any tiaws which lie at an angle to the horizontal plane should be separated into a circumferential and axial component. These components can then be used separately in the determination of effective flaw lengths The selection ofindications for further investigation can be performed by evaluating the resulting effective flaw lengths. Indications with effective flaw lengths greater than the allowable flaw sizes would require further characterization by NDE or more detailed analysis. The procedure described here is conservative since all of the indications are assumed through-wall and are being cornpared against the allowable through-wall flaw size. The report covers the limiting stresses for all the Unit-2 shroud welds (H1 through H8 welds). Therefore, the screening criteria developed here is applicable to all shroud locations where indications are postulated. A list of conservatisms used in this evaluation is summarized in Table 1-1. 4
i GENE-323 1*6-1293 GE Nudear Enerv Table 1-1 Conservatisms Included In Screening Evaluation
- 1. All surface indications were assumed to be through-wall for analysis.
- 2. The screening criteria limit one-fourth of allowable circumferential flaws to any arbitrary 90 sector.
- 3. All indications are assumed to be grouped together for the limit load calculation and no credit is taken for the spacing between indications.
- 4. ASME Code primary pressure boundary safety margins were applied even though -
the shroud is not a primary pressure boundary. )
- 5. ASME Code, Section XI proximity rules were applied.
- 6. An additional proximity rule which accounts for fracture mechanics interaction between adjacent flaws was used (See Appendix A). ,
- 7. The highest stress computed for any single location was used for all locations.
- 8. Both LEFM and li.mit lead analysis were applied, even though LEFM underestimates allowable flaw size for austenitic materials and is not required per ASME Code Section XI procedures.
- 9. Fracture toughness measured for similar materials having a higher fluence was used.
- 10. The bounding crack growth estimated for the next fuel cycle was included in flaw lengths used for evaluation (See Appendix B).
I1. A proximity rule to account for perpendicular flaws was applied, although not required by Section XL l l
- 12. Power rerate conditions were used although it will not be in effect during the next j l
fuel cycle. i 4 5
GENuclear Enerxy GENE-523-1w1203 1.3 References 1-1 GE Services Information Letter (SIL) 572, Rev.1, October,1993 1-2 " Evaluation and Screening Criteria for the Peach Bottom Unit-3 Shroud Indications," GENE-523-141-1093, Rev.1, DRF 137-0010-6, December 3,1993 l 6
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GE Nuclear EnerO GENE-H3-176-12H 2.0 FABRICATION HISTORY This section describes the fabrication history of the Peach Bottom Unit-2 shroud. Of key interest is the material composition and any activities which could have possibly contributed to the increase ofintergrannular stress corrosion susceptibility (IGSCC). Quality assurance records received from the vessel vendor (Rotterdam) were examined in detail to determine the appropriate information. Table 2-1 shows the material data for the Unit-2 shroud. The part numbers are identified in the schematic shown as part of Table 2-1. Also shown in the table is the number of pieces for each part, material designation, heat numbers, and carbon content. For comparison, Table 2-2 shows the material data for the Unit-3 shroud. As can be seen from this comparison, the Unit-2 and Unit-3 shrouds are very similar, differing slightly only in carbon content and hea'kertificate number. However, the very small differences are not significant in terms of response to cracking mechanisms. Figure 2-1 shows the assembly of the shroud All weld locations are identified including vertical and horizontal welds. Figure 2-2 through 2-9 show the details of the shroud welds as labeled in Figure 2-1. The Unit-2 and Unit-3 shroud geometry are identical. The upper, central and lower rings (part numbers 1,3 and 6) are austenitic stainless steel seamless rolled forging. The materialis ASTM A182 - F304. The heat treatment of these rings consisted of heating to 1100 C, holding for 6 hours and followed by water quenching to below 100 C. The carbon contents of the rings range from 0.028% to 0.035% max. Hardness measurements on completion of solution heat treatment and rough machining of the rings ranged from Brinell hardness of 148 to 159. Each cylinder is made of 2 plate segments formed and welded to drawing requirements. Plate materialis austenitic stainless steel made to ASTM A240 Type 304 specifications The carbon contents of the plate materials range from 0.056% to 0.062% max. The hardness of the plate material range from Brinell Hardness of 145 to 159. All welding was performed by submerged are welding except H7. The procedure and welder qualification was performed to ASME Section IX requirements. The filler metal 11
GE Nasitar Energ GENE-533-176-1293 met ASME A-371 Type ER 308 requirements with required carbon content of 0.08% max. The welded joints did not use backing strips but utilized 3 to 4 hand weld passes. The maximum interpass temperature was 350 F. Weld prep surfaces of the base metal were prepared by machining. The backside of the groove welding was prepared by grinding or gouging followed by liquid penetrant inspection. Final surfaces of the welds were inspected by liquid penetrant examination. The H7 weld was performed using metalinert gas with Alloy 82 wire. In addition,100% ultrasonic examination of weld H7 was performed. Based on GE Quality Assurance records received from Rotterdam, no abnormal fabrication history was found. General practice during assembly and shipment of the of the shroud, bracing, temporary welds, and supports are used to help in meeting the joining of the various components and to meet geometric tolerances. Although there is no record documentation of these practices, it is likely that they were present during fabrication. These actions result in a local effect on material behavior and stress. For example, the welding of temporary pads would result in a local area of weld residual stress and perhaps some grinding (cold work). If these local effects contribute to SCC, it is likely that the cracking would be oflesser concern than cracks near the horizontal welds. 2.1 Summary Based on a comparison of the Unit-2 and Unit-3 shroud fabrication history,it is concluded that the shrouds are essentially similar. Thus, the similarity of material and fabrication information supports the use of the Peach Bottom Unit-3 IVVI results in the evaluation of the Peach Bottom Unit-2 shroud. 12 l i
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GE Nuclear Enerv 3.0 CHEMISTRY AND FLUENCE CONSIDERAT Water Chemistry History 3.1 d Unit-3 operated with For the first decade of hot operation, both Peach Bottom Unit-2f relatively high primary water conductivity,d with Un t- exceeded i ity actually
- 3. As can be seen in Figure 3-1, Unit-3's ii arithmetic ery high and mean con uct 1.0 pS/cm in 1976 and 1977. The arithmetic mean conduct v exceeded 0.4 pS/cm for both units until 1986. S/cm for Units-2 and Unit-3, decreased and were <0.1 pS/cm (0.078 pS/cm
' conductivity valuesand are 0.089 respectively) during 1992 and 1993. These last two year s considered world class performance. -
l ki condensers and The high conductivity during the first half oflife was partly due resin change out problems. However, i besides the high ear yd at Unit-3 i there was also one relatively severe transient exper ence be seen in . 3-1, a summary of BWR fleet severe transients k 13) duringthrough abo Table 3-1, Unit-3 suffered at least one power resin intrusion which the conductivity reached 23.6 pS/cm due to poss RPV. Since intrusion. This intrusion type results in the dinjection nickelbase of sulfate IGSCC initiation and propagation f in i s sensitized austenitic sta such as dissolved oxygen,- alloys are controlled by the rate of cathodic reduction fd io spec ntal e hydrogen peroxide and/or variousdhoxyanions, ccelerate anodicthen the oxyanion such as sulfate, would increase the cathodic Unit-2 had current j dissolution at the crack tip,i.e., IGSCC (Reference 3-1). On only a mild transient (1.0 S/cm, Incident Rank 69) where an org] ! u well. I h 24 1
4 GENE-5231%1:g3 GE Nuclear Energy 3.1.1 Effects ofImpurities on IGSCC An example of the effects of sulfate / conductivity on crack initiation in uncreviced mater is presented in Figure 3-2. It is clear that an increase in sulfate / conductivity results in acceleration in crack initiation as measured by the constant extension rate test (CERT) (References 3-1 through 3-4). A specific Peach Bottom example of an acceleration crack propagation rate (creviced) with sulfate is shown in Figure 3-3. Figure 3-3 disp June 1986 (not included in Table 3-1) Unit-3 on-line crack monitoring data for sensitized Type 304 stainless steel. The results clearly illustrate the change in crack grow after two closely linked water chemistry transients of 4-5 S/cm, i.e., increases in water conductivity due to intrusions of demineralizer resin material (Reference 3-5). This fig i demonstrates the dramatic increase in crack growth rate (2X) with conductivity. Simi!ar on-line crack monitoring results with sulfate have also been documented in the laboratory, Figure 3-4 (Reference 3-6). Other anions such as chloride, carbonate, etc. have sil kinetic effects on IGSCC initiation and propagation (References 3-7 and 3-8). This high conductivity crack initiation and propagation acceleration factor is consisten with the relatively high incidence ofIGSCC observed at Unit-3 in creviced Alloy 600 shroud head bolts (15 of 25 bolts examined cracked) and access hole covers. No crack of these two components has been identified in Unit-2. Both units have suffered IGSCC of creviced safe ends. Additional documentation on the strong correlation ofIGSCC susceptibility with actual BWR plant water chemistry history for creviced BWR components has been published (Reference 3-9). 3.1.2 IGSCC Modeling Finally, the effect of conductivity on crack propagation has also been quantifie Research and Development Center based on a "first principles" model of crack advance known as the film rupture / slip dissolution model(Reference 3-10). Predictions from PLEDGE (Plant Life Extension Diagnosis by GE) model have been extensively 25
GE Nudeu Ewty GENE-3D-M.HM compared with laboratory and field data and has provided validation of the technique. For example, PLEDGE predicts the crack growth rate in stainless steel and low alloy steel within a factor of approximately two for a 70% statistical conndence over a range in observed crack growth rate of more than six orders of magnitude. Likewise, it provides a very reasonable mean value and can accurately bound the observed crack growth rate in stainless piping and other components. Aside from piping predictions, PLEDGE has been successfully used for on-line crack growth monitoring data, safe ends (avoiding mid-cycle plant shutdowns), non-sensitized (stabilized) stainless steels and reactor internals such as the core shroud, top guide, access hole cover and in-core monitor housing. The PLEDGE Model ofIGSCC and more recently IASCC (Reference 3-11) indicate the strong effect of conductivity on crack growth rate and by inference crack initiation. Figure 3-5 presents a schematic estimation of Unit-2 and Unit-3 crack growth rates as a function of conductivity using PLEDGE. Crack growth rates based on actual conductivity averages for the first ten years (Unit-2: 0.593, Unit-3: 0.752) were compared to those averages for the last two years. A value of 200mV [SHE] was used for the electrochemical potential (ECP) in these calculations. As noted in Figure 3-5, a factor of improvement (FOI) of approximately 20 decrease in crack growth rate is obtained with the Unit-2 decrease in conductivity. This compares against an FOI of 11 for Unit-3. 3.1.3 Water Chemistry Summary The above results and current discussion clearly indicate that Unit-2 IGSCC performance is superior to Unit-3's. Unit-3 is characterized by higher average coolant conductivity, a more severe water chemistry transient and cracking of shroud head bolts and access hole covers. Therefore, Unit-3 can be utilized as a " worst case" scenario for the current shroud cracking concern at Unit-2. 3,2 Fluence Considerations An important parameter which helps in the evaluation of the cracking mechanism is fluence. The fluence is the time integrated flux at a particular location. Shroud peak fluence was calculated by multiplying peak flux at the shroud location by the effective full power seconds of operation. The peak fluence in the Unit-2 and Unit-3 shrouds at the end 20 2 of the next cycle is expected to be approximately 7.9x10 n/cm (E>lMev). This fluence corresponds to the mid-beltline region of the shroud at the location where the 26
GENE 323-176-1293 GE Nuslear EnenD' azimuthal variation is also maximum. Although peak shroud flux may vary significantly from cycle to cycle, available flux results are generally limited to one operating cycle per plant due to substantial resource requirements for vessel flux analysis. Shroud fluence estimates were therefore calculated based on the assumption that flux remains constant throughout the life of the plant. Further evaluation would be needed to quantify the uncertainty associated with this asrumption. However, the method of determining fluence is considered to be sufficient to obtain an estimate of the overall condition of the material with respect to irradiation effects. 3.2.1 Fluence Considerations Summary The fluence at the peak shroud location is essentially the same for Units 2 and 3. Thus, from a fluence standpoint, the use of the Unit-3 IVVI results in the Unit-2 evaluation is justified. 1 l I l i 1 l l 27
GENE-5831761393 GE Naslear Energy 3.4 References 3-1 W.J. Shack, et al, " Environmentally Assisted Cracking in Light Water Reactors: Semiannual Report April- September 1985," NUREG/CR-4667, ANL-86-31, June 1986. 3-2 W.J. Shack, et al, " Environmentally Assisted Cracking in Light Water Reactors: Annual Report October 1983 - September 1984," NUREG/CR-4287, ANL-85-33, June 1985. 3-3 L.G. Ljungberg, D. Cubicciotti and hi. Trolle, " Effects ofImpurities on the IGSCC of Stainless Steel in High Temperature Water," Corrosion, Vol. 44, No. 2, Febmary 1988. 3-4 W.E. Ruther, W. K. Soppet and T. F. Kassner, "Effect of Temperature and Ionic Impurities at Very Low Concentrations on Stress Corrosion Cracking of Type 304 Stainless Steel," paper 102 presented at Corrosion 85, Boston, hiA, NACE, h1 arch 1985, published in Corrosion, Vol. 44,'No.11, November 1988. 3-5 D.A. Hale and C. G. Diehl, "Real Time hionitoring of Environmental Crack Growth in BWRs", paper 455 presented at Corrosion 88, St. Louis, hiO, NACE, hiarch 1988. 3-6 B.hi. Gordon, Corrosion and Corrosion Control in BWRs, NEDE-30637, p. 6-22, December 1984. 3 7 R.B. Davis and hi. E. Indig, "The Effect of Aqueous Impurities on the Stress Corrosion Cracking of Austenitic Stainless Steelin High Temperature Water," paper 128 presented at Corrosion 83, Anaheim, CA, NACE, April 1983. 3-8 P.L. Andresen, " A hiechanism for the Effects ofIonic Impurities on SCC of ' Austenitic Iron and Nickel Base Alloys in High Temperature Water," paper 101 presented at Corrosion 85, Boston, hiA, NACE, hf arch 1985 3-9 K.S. Brown and G. hi. Gordon, " Effects of BWR Coolant Chemistry on the l Propensity for IGSCC Initiation and Growth in Creviced Reactor Internals Components," paper presented at the Third Int. Symp. of Environmental Degradadon of hiaterials in Nuclear Power Systems-Water Reactors, Traverse City, hil, August 1987, published in proceedings of same, ThiS-AIhiE, Warrendale, PA,1988. l l 3-10 F.P. Ford et al, " Prediction and Control of Stress Corrosion Cracking in the Sensitized Stainless Steel / Water System," paper 352 presented at Corrosion 85. Boston, hiA, NACE, hiarch 1985. i 28 l J i
GE Nuclear Ewty GENE-323-176-12H 3-11 P.L. Andresen and F. P. Ford, "Modeling ofIrradiation Effects on Stress Corrosion Cracking Growth Rates," paper 497 presented at Corrosion 89, New Orleans, LA, NACE, April 1989. S k 29
GENuclear Enagpr ggyg.3gj.j76.jgg3 Table 3-1 Severe Water Chemistry Transients in BWRs MAX MAA 10/07/93 Cl, PCV OATE DATA CONO. pH, pco LEV y-m-d CCwMENTS REFERENCE POINT SANK PLANT uS/cm on ... ............. ...... ..... ..... ... ...... ==..==............................................ ............. .,.. 95.0 4.5 100 P 780307 COND OEMlN RESIN BLEE0THROUGH PC&RT 82LCA01 35 1 AG 9 800802 CONDENSATE CEMIN PESIN INTRUSICN EPRI NP 4134 46 2 AG 88.0 3.2 14500 P 720901 CCNCENSER LEAK, CEMIN 00EPLETED PC&RT 821.0A01 6 3 At 84.0 560 P 660820 PCLRT 82LDA01 1 4 OZ 72.0 PCSRT 82LDA01 30 5 AG 70.0 4.6 198 P 771116 CRUO & CONDENSATE CENIN RES!N INTRUSICN NECE 13405 13 54.0 3.8 P 740804 RESIN BEAD INTRUSION
- 6N 11 40.5 3.9 P 740608 A!R/A!R RESIN MIXTURE INJECTED INTO Rx FRCM RVCU PC&RT 82LDA01 7 A8 NEDE 13405 10 8N 33.0 4.0 P 740426 RESIN BEAD INTRUSION
- PCLRT 82 LOA 01 3 30.0 P 710903 HIGH CONOUCTIVITY VATER IN CST 9 AC PC&RT 82LCA01 2 10 02 28.5 P 661130 PC&RT 821.0A01 28 11 8 25.6 4.1 50 P 770601 RESIN INTRUSION 2500 p 810412 CONDENSER LEAX PMET 81-688-45 52 12 8 25.0 EPRI NP 4134 80 P 800205 POS$18LE CON 0ENSATE DEMIN RESIN INTRUSION 13 PEACH 80TTCH 3 23.6 42 14 02 23.0 3000 P 790407 LEAKAGE OF CCCLING VATER INTO RPV VIA CCRE SPRAY PC&RT 82LCA01 PCLRT 82LCA01 7 15 V 23.0 30 P 730406 AIR INJECTED INTO Rx FROM RVCU PC&RT 82LCA01 32 22.0 P 771212 COND DEMIN RESIN INTRUSION 16 AG PC&RT 82LCA01- 53 17 K 21.0 4.6 2500 P 820428 TRICHl.0R0 ETHANE FROM RADVASTE AND CST EPRI NP 4134 94 20.0 4.5 p 821004 POSSIBLE CONDENSATE DEMIN RESIN INTRUSICN 18 AG PC&RT 82 LOA 01 49 17.0 P 801001 CCNCENSER TUBE LEAXS 19 F PCLRT 82 LOA 01 20 14.0 P 750605 RVCU OUT OF SERVICE 20 C 39 4.7 100 P 781110 CRGANIC INTRUSION VI A CONDENSATE, DECON DETER /CILS PC&RT 82 LOA 01 21 7 13.8 PCSRT 821.0A01 14 22 A8 13.5 P. 740925 HIGH CONO VATER EPRI NP 4134 81 23 A8 13.0 100 P 800428 UNKNOVN (LONG SHUTOCVN)
P 780225 RVCU RESIN INTRUSICN PC&RT 82LCA01 34 24 T 12.1 PC&RT 82LCA01 21 12.0 P 750702 CCNCENSER TUBE LEAK 25 Q 50 P 761025 RVCU RES!N TRAP, RVCU INCPER;8LE PC&RT 82LCA01 25 26 0 12.0 4.8 EPRI NP.4134 82 i 27 T 11.8 P 800812 CRGANIC INTRUSION 11.5 4.8 60 P 750127 RVCU RESIN INTRUSION PC&RT 82LCA01 17 l 23 0 95 l 830106 POS$1BLE CONDENSATE DEMIN RESIN INTRUSION EPRI No 4134 29 A8 11.3 4.7 p 4.5 50 P 750601 RESIN FROM FLUFFING CONCENSATE OF/0 PC&RT 82 LOA 01 19 f 30 8 10.8 PC&RT 82 LOA 01 8 l 31 AG 10.6 4.5 100 P 730507 RVCU RESIN INTRUSION 10.0 P 741208 CONDENSER LEAK PC&RT 82LCA01 15 ) 32 0 EPRI NP 4134 92 10.0 p 820818 RVCU RES!N INTRUSION 33 AG 74 EPRI NP 4134 34 H 9.2 7.4 57 P 780211 CONOENSATE DEMIN RESIN INTRUSION EPR1 NP 4134 63 35 8 8.2 4.3 50 P 751210 VASHOUT OF IMPURITIES FRCH TURBINE PMET 81-688-45 44 l 36 8 8.0 5.0 500 P 790516 EPRI No 4134 86 l 7.5 p 810411 RVCU RES!N INTRUSION 37 8 89 7.1 100 P 811010 DECCMPOSITION OF RADVASTE RESINS OUE TO HOT VATER EPRI NP 4134 l 38 C 84 EPRI NP 4134 39 C 6.5 P 810210 CAUSTIC INTRUSICN VIA CONDENSATE STORAGE EPRI NP 4134 57 40 K 6.2 4.8 20 P 741118 SUSPECTED.RES!N INTRUSION EPRI NP 4134 54 41 0 5.8 4.5 50 P 730812 SUSPECTED RESIN INTRUSION P 700123 RESIN INTRUSION VHEN C/0 RETURNE0 TO SERYlCE JMS CC 930717 98 42 AA 5.6 64 EPRI NP 4134 43 8 5.4 4.7 50 P 751218 PROBABLE RVCU RESIN INTRUSION EPRI NP 4134 85 5.1 p 810220 CRGANIC INTRUSION VIA RADVASTE 44 8 ESRI N8 4134 70 5.1 68 P 770727 CCN0ENSATE DEMIN RESIN INTRUSION, AN!ON RICH 45 H 67 EPRI NP 4134 46 0 5.1 4.8 50 P 760806 CCNDENSATE DEMIN RESIN INTRUSICN 30
GENuclear Energy GENE-323-176-1293 Table 3-1 (cont'd) Severe Water Chemistry Transients in BWRs N 10/07/93 MAX DATA COND. pH, C1, POV CATE uS/cm min pcb LEV y-m-d COMMENTS REFERENCE POINT RANK PLANT EPRI NP 4134
....79 47 C 5.0 100 P 750309 POSSIBLE CONDENSATE CEMIN RES!N INTRUSlCN '
EPRI NP 4134 66 48 0 4.9 4.9 50 P 760522 SUSPECTED RESIN INTRUSION EPRI NP 4134 77 49 i 4.5 5.0 50 P 781227 ORGANIC INTRUSION VIA CONDENSATE SYSTEM EPRI NP 4134 65 50 Q 4.3 4.9 48 P 760221 SUSPECTED RESIN [NTRUSION EPRI NP 4134 56 51 K 4.1 5.1 80 P 741015 RVCU RES!N INTRUSION EPRI NP 4134 68 52 AL 3.3 5.4 50 P 770126 [MPROPER RINSE OF CONDENSATE DEMIN EPRI NP 4134 59 53 C 3.3 5.2 38 P 750309 POSSIBLE RESIN INTRUSION EPRI NP 4134 88 3.2 p 810715 RVCU RESIN INTRUSION 54 8 73 EPRI NP 4134 55 S 3.2 4.7 495 P 780131 RESIN INTRUSION EPRI NP 4134 61 56 H 3.0 5.6 96 P 750902 SUSPECTED RESIN INTRUSION EPRI NP 4134 62 57 8 2.9 5.4 50 P 751126 PROBABLE RVCU RES!N INTRUSION EPRI NP 4134 71 58 T 2.8 5.2 65 P 770912 IMPROPER RINSE OF CONDENSATE DEMIN EPRI NP 4134 60 59 8 2.7 7.6 P 750626 RESIN [NTRUSION EPR[ NP 4134 83 2.3 P 800824 ORGANIC INTRUSION 60 T 78 l EPRI NP 4134-61 T 2.2 5.5 50 P 790108 SUSPECTED CRGANICS IN CONDENSATE STORAGE EPRI NP 4134 . 76 l 62 Y 1.8 5.4 355 P 781208 CONDENSATE DEMIN RESIN INTRUSION ) EPRI NP 4134 58 63 AC 1.4 5.6 83 P 741125 VALVING ERROR DURING RESIN TRANSFER EPRI NP 4134 72 64 H 1.4 8.1 38 P 780112 CONOENSATE DEMIN RESIN INTRUSION 65 AJ 1.4 P 750906 SUSPECTED FLOC / FILTER A10/SURFACT FROM RAD WASTE JMS QC 930717 99 EPRI NP 4134 69 ; 66 0 1.1 5.6 30 P. 770225 SUSPECTED RESIN INTRUSION EPRI NP 4134 75 l 67 H 1.1 8.8 72 P 780511 CONOENSATE DEMIN RESIN INTRUSION ' EPRI NP 4134 90 68 V 1.0 P 811030 GYLCOL [NTRUSION VIA RA0VASTE EPRI NP 4134 87 l 69 PEACH BOTTOM 2 1.0 P 810622 OIL INTRUSION INTO HOTVELL 1 PC&RT 82 LOA 01 4 70 AS 725 P 711113 HIGH FEE 0 VATER CONDUCTIVITY PC&RT 82LDA01 33 71 8 540 P 780129 PC&RT 82LCA01 24 72 8 1200 P 760708 CONDENSATE SYSTEM MCMENTARILY BYPASSED PC&RT 82LDA01 16 73 AR 600 P 750103 RVCU CUT OF SERVICE 43 641.0 3.5 87000 5 790426 COOLING VATER INGRESS FROM RHR, RPV H2O TO HOTVELL PC&RT 82LDA01 74 8 PC&RT 82LDA01 12 75 8 423.0 3.2 5 740801 ACID INTO RPV FROM DEMIN STORAGE TANK PC&RT 82LDA01 22 i 76 H 140.0 $ 760519 PC&RT 82LDA01 26 l 77 T 45.9 3.8 244 5 761103 TORUS VATER PUMPED INTO RPV PRIOR TO STARTUP PC&RT 82LDA01 23 78 8 13.3 1800 $ 760520 PC&RT 82 LOA 01 36 79 A 13.0 5 780801 LEAX IN RHR HEAT EXCMANGER PC&RT 82LDA01 29 80 8 12.9 5 770917 RVCU OUT OF SERVICE PC&RT 82LCA01 47 81 T 12.1 5 800815 PC&RT 82LCA01 9 82 AL 11.6 5 730603 RVCU OUT OF SERY1CE PC&RT 82LCA01 48 83 8 11.2 S 800822 PC&RT 82LDA01 51 84 0 11.2 5 801219 EPRI NP 4134 91 85 F 10.5 5 820427 POSSIBLE ORGANIC INTRUSION EPRI NP 4134 37 86 8 10.5 5.6 140 5 780923 RVCU RESIN INTRUSION PC&RT 82LCA01 19 87 H 10.3 5 750405 PVCU OUT OF SERVICE PC&RT 82 LOA 01 5 88 AS 10.0 730 $ 720604 DEPLETED RVCU CEMIN EPRI NP 4134 55 89 0 5.0 5.5 60 5 740829 CONDENSATE DEMIN RESIN INTRUSION EPRI NP 4134 97 90 AX 4.5 5.2 220 $ 830505 CRGANIC INTRUSION VI A RADVASTE 38 4.2 5.3 600 $ 781110 CRGAN!C INTRUSION VIA CONDENSATE, DECON DETER /0!LS PC&RT 82LCA01 91 8 EPRI NP 4134 93 1.0 S 820900 GYLCOL INTRUSION VlA RADVASTE 92 AP 31
GENuclear Energy . GENE-323-176-1293 Table 3-1 (cont'd) Severe Water Chemistry Transients in BWRs
*
- 10/07/93 C1, POV DATE CATA CONO. pH, uS/cm min ppb LEV y-m-d COMMENTS REFERENCE POINT RANK PLANT
... ............. ...... ...., ..... ... ...... .................................................. ............. .... j PC&RT 82LDA01 45 93 F 700 $ 800305 PC&RT 82LDA01 41 I 1300 i 790329 CON 0ENSER LEAK, CCNDENSATE BYPASSED, RVCU CUT 94 T PC&RT 82LDA01 50 l 95 0 500 5 801017 PC&RT 82LCA01 27 l 96 AC 683 5 770309 96 f 97 K 5 830213 GYLCOL INTO RA0VASTE. DETECTED PRICR TO COND STOR EPRI NP 4134 1 1200 5 790316 CONDENSER LEAK, CONDENSATE DEPLETED, Cl INTO CST PC&RT 82LCA01 40 98 T PC&RT 82 LOA 01 31 99 8 800 5 771206 RVCU OUT OF SERV!CE I
)
NOTE: 8WRS RANKED IN THE FOLLOVING ORDER:
- 1. POWER (P) OR SHUT 00VN (S)
- 2. CONOUCTIVITY OTHER NOTES: * = RESIN BEADS PROV!0E LONG TERM LOV pH 32
e a Arithmetic Vlean Concuctivity a Peach Bottom 2 anc 3 lD Conductivity, pS/cm __ _ _ _ _ 1.4 - -G- Peach Bottom 2
+ Peach Bottom 3 1.2 - .
O Peach Bottom 2 all years - 0.409
/
/ Peach Bottom 3 all years - 0.536 '
g 0.8 - f Og n / hs 0.6 -
"J 0.4 -
U V 3 ,_ e $g
'- ' '- ~~~-
'------i-------'---'-----'-
5-0 - --- 90 92 94 g 76 78 80 82 84 86 88 74 :: Fuel year Figure 31 Arithmetic Mean Cotuluctivity Ibr Peach Bottom Unit-2 and Unit-3
Effect of Concentration anc. Conductivity e y on IGSCC Initiation - FS Type 304 i m R Acceleration Factor _ _ . 5 O ANL Data- V EPRI Data 4 -
/ ( >
4- O 3 - V 2 - O Conductivity (pS/cm) ao Hs ei or oso.o. o ra n o so Nas 'Q o rs no o, o r as oa os i i -r , rrrr a i i , r rr li i i i i i i iii k' 1& i i i iiiiii 1000 10000 t 10 10 0 1 Sulfate (ppb) {
. Crack initiation da ta based on CERT Figure 3-2 litTect of Concentration and Conductivity on IGSCC Initiation - Furnace Sensitized
n Peach Bottom 3 Response to June 1986 . m ife Water Chemistry Transient !! Crack length, mm . . . - . . .
- lm 19.16 Resin Resin intrusion 1 Intrusion 2 ..
19.14 - . f'
. .f'. . '
~
19.12 -
.'. da/dt = 1.09 mm/y w
u,
~ ' -
- J.-
da/dt = 0.55 mm/y , 19.08 -
' ' ' ' ' ' - ~ ~~
I 19.06 2500 2600 2700 2800 $ 210 0 2200 2300 2400 4 E Hours :D , Figure 3 3 Peach Bottom Unit-3 Response to June 1986 Water Chemistry Transient iw ir-.g- e +--w + v w a w lw -F-_Y - b
- - - - - - - ~ ~ ~ -
MM IN.SCL1 COMST A NT L O A D - - - - - -- - - - - - - - - - - - - - - - - - - - - 20.8 i - 0.82- - 2 1.0 pS/CM Q 2- <0.1 pS/cm Na2SO4 h,- O 45 pS/cm Na2SO4 l
! <0.1 pS/cm
- 0.81 - O.5 pS/cm 7 W 20.5 -
O~ g.
- a. t-i 3 = m=
z Z 4
- < 4 S
- us K J >-
20.3 - 0.80 - - o t 2
~
. --- 20 PPB OXYGEN ---
~
p --- -200 PPB OXYGEN - -- -- 20.1 . 0.79
~
---- -- -2OO PPB OXYGEN --- - - - - - - - - - - - - - __
I
~
19.8 - 0.78 -
,a 7 :
- 0.77 ~-
19.6 -
~
g
# # ~
19.3 - 0.76 ~- - # 1.5 pM/h O.1-0.03* pM/h 0.09 9M /h 1.1 pM/h (4.1-1.2
- pin./h) (60* pin./hl
- (45 pin./h)
(3.8 pin./h) 0.75 Q 19.0 - E'1
~
~
l w b ' ' ' ' 2500 Q '
' ' ' 2000 18 8 - 0.74 1000 1500 t O 500 E
nouns *NO COMPLlANCE UNLO ADING liigure 3-4 Crack Length vs. Time, Sensitized 'lype-304 Stainless Steel,286C Oxygenated Water Environment K = 28.6 to 30.5 MPa-sqrt(m) (26 to 28 ksi-sqrt(in))
--.----___--,__---____--____----,--_--_-_-_____-_--_.-----_----___-__---_--a- - . , _ _ _ _ _ - - - , _ - - _ _ _ - - _ - - _ _ - - a _ - _ - _- - - - - - _ - - _ - - - _ - - - _ __------- - - - - -- -- -
GENE PLEDGE Model Prediction fdr PB-2/3 .
~
Sensitized Type 304 Crack Growth Rate [ t Crack Growth Rate, in/h [m 1975-1984 200 mV
^ M0_my_
1.000 E- 04
- 5.000E-05 ~1
-100 mV FOI y -
20 1.000 E - 05 2 3 / [ -200 mV 1992-1993
~
-300 mV 1.000 E- 06 g
~
O PB-3 A P B-2 ! S I ' (. 1.000E- 07 0.6 0.7 0.8 0.9 2-0 0.1 0.2 0.3 0.4- 0.5 Conductivity, pS/cm [, i . P LEDGE: 15 C/cm2, 20ksi/in Figure 3-5 GENE PLEDGE Model Prediction for Peach Bottom Unit-2/3 Sensitized . ___ ___ _ _____ ___ ____ ___-__ - ___ _ _= - _ ____ - _ __ - _.__ - _
GENE-523-176-1293 GENuden Emu 4.0 IN-VESSEL VISUAL INSPECTION OF PEACH BOTTOM UNIT-3 An important factor in demonstrating the justification for continued operation of Unit-2 is ' the validity of using the results of the IVVI of the Unit-3 shroud. Based on the similarity between the two units as discussed in Section 2 and 3, it is reasonable to expect that any potential shroud cracking in Peach Bottom Unit-2 is bounded by that observed in the Peach Bottom Unit-3 shroud. This observation is strengthened by the comparison between the Unit-2 and Unit-3 water chemistry which showed that Unit-2 water chemistry has been historically better than that at Unit-3. Following is a discussion of the IVVI l results for the Unit-3 shroud. 1 IVVI of welds H1 through H8 were performed during the recent Unit-3. The IVVI included both inside surface and outside surface examination. Figures 1-2 through 1-4 , show the indications associated with the H3 and H4 welds. Only a few short indications were observed on the outside surface of H1 and H4. Circumferentialindications were observed on the inside surface associated with venical weld V3 (See Figure 1-2).' It should also be noted that the area adjacent to the H9 weld was visually inspected as part of access hole cover (AHC) inspection at this outage. The inspection did not reveal any indications. The H9 weld in the vicinity of the access hole cover is considered a higher .! stressed location and therefore these IVVI results are considered to provide a reasonable assessment of the overall condition of the entire H9 weld. All indications associated with the H3 weld inside surface were in the HAZ of the shroud cylinder. No indications were found in the ring. The indications near H3 were all circumferentially oriented. As can be seen in Figure 1-2, most of the indication length is located between the azimuth of 146 and 360 The indications observed at the H4 inside surface HAZ were a mixture of circumferential' and axial indications as shown in Figure 1-2. Circumferential indications were observed emanating from the V3 weld on the inside surface. Eight indications were observed grouped together with a spacing of - approximately 2" between indications. All other IVV1 vertical scans found no indications 38
GE Nuc!<ar Energy ' GENE-523-176-1293 On the outside surface a limited number of short indications were observed associated with H1 and H4. Table 4-1 summatizes the indications found during the Unit-3 IVVI. Table 4-1 Summary of Unit 3 IVVI Indications Weld Inside Surface Indications Outside Surface Indications N/A 1 short vertical HI N/A None H2 Circumferentialin Shroud None H3 Cylinder HAZ Circumferential and Axial 2 short vertical H4 N/A None H5 N/A None H6 N/A None H7 N/A None H8 8 short circumferential N/A V3 None N/A PLATE l 1 ~ 39
GENE-5231w1293 GENuctw EnerKY 4 5.0 JUSTIFICATION FOR CONTINUED OPERATION This section provides the basis for the justification for continued operation of Peach Bottom Unit 2. The JCO is based on postulating that indications in the Unit-3 shroud are present in the Unit-2 shroud and that they meet the screening criteria presented in this section. As discussed earlier, Peach Bottom Unit-3 is considered to be more limiting in regards to extent of cracking. 5.1 Structural Analysis This section describes the details and the results of the structural analysis performed to i determine the allowable flaw lengths. The structural analysis consists of two steps: the determination of axial and circumferential stress magnitudes in the shroud, and the ' calculation of the allowable flaw lengths. Both the fracture mechanics (LEFM) and limit load methods are used in the calculation of allowable flaw lengths. i 5.1.1 Applied Loads and Calculated Stresses The applied loads on the shroud consist ofinternal differential pressure, weight and seismic. The seismic loads consist of a horizontal shear force at the top of the shroud and an overturning bending moment. The shear force produces a shear stress ofinsignificant. magnitude, and is not considered. The bending moment stress at a shroud cross-section varies as a function ofits vertical distance from the top of the shroud. Because of the inherent ductility of the material, residual stresses and other secondary stresses do not l affect structural margin. Thus, they need not be considered in the analysis. The magnitudes of the applied loads were obtained from the seismic stress analysis and system information reports. The nominal shroud radius and thickness (2.0 in.) were used to calculate the stresses from the applied loads. The stresses are essentially based on the strength of materials formulas. Since the bending stra due to seismic shear force varies with the elevation of a location, two conservative vain'es of this stress were calculated: one applicable to shroud sections above the core plate (H1, H2, H3, H4, and H5) and the other for sections below the core plate (H6, H7 and H8). Figure 5-1 shows the weld designation and relative locations in the shroud. 40
GENE-323-176-1293 l GE Nuclear Energy l 1 1 Table 5-1 shows the calculated seismic stress magnitudes for both the upset (Design , Earthquake - DE) and faulted conditions (Maximum Credible Earthquake - MCE). The j appropriate pressure differences for the upset and faulted conditions are shown in Table 5-2. Table 5-1 Seismic Axial Stresses at Shroud Welds Weld MCE Stress (ksi) ( Moment Designation (ft-kips) MCE DE H1 1104.7 0.18 0.08 q 1438.6 0.23 0.11 j H2 j H3 1479.1 0.27 0.13 H4 2995.8 0.54 0.24 i HS 4583.8 0.83 0.37 H6 4679.7 0.90 0.40 H7 5697.6 1.10 0.49 H8 6749.7 1.30 0.58 i Table 5-2 Pressure Differences l Pressure Differences (psi) Component Faulted Condition Upset Condition l 32.9 14.12 Shroud Head and Upper Shrcud 54.8 35.68 l Core Plate Support Ring and Lower Shroud The structural analysis for the indications uses two methods; linear elastic fracture mechanics (LEFM) and limit load analysis. Both the limit load and the LEFM methods were used in determining the allowable flaw sizes in the shroud. Since the limit load is concerned with the gross failure of the section, the allowable flaw length based on this approach may be used for comparison with the sum of the lengths of all the flaws at a cross-section. On the other hand, the LEFM approach considers the flaw tip fracture toughness and thus, the allowable flaw length based on this approach may be used for comparison with the largest effective flaw length at a cross-section. The technical approach for the two methods is described below. i l 41
GE Nuclear Energ GENE-523-1%133 5.1.2 Fracture Mechanics Analysis 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. In fact, J-R curve measurements (Figure 5-2) made on a core shroud sample taken from an 20 n/cm2) showed stable crack extension and overseas plant having higher fluence (8x10 ductile failure. The ASME Code recognizes this fact in using only limit load techniques in Section XI, Subsubarticle IWB-3640 analysis. Nevertheless, a conservative fracture mechanics evaluation was performed using an equivalent Kjc corresponding to the materialI J c. The Kjc for the overseas plant shroud was approximately 150 ksidin. Use of this equivalence is conservative since: i) The calculated fluence for Peach Bottom Unit-2 is lower than that for the overseas plant from which J-R curves were obtained. values well above the JI c, confirming that there is load ii) The J-R curves show Jmax capability well beyond crack initiation (See Figure 5-2). Using the ASME Code safety factor of 3, which is applicable for normal and upset conditions of pressure boundary components, the allowable K cI value becomes 50 ksiVin For faulted conditions the allowable K Ic is 107 ksidin using the ASME Code safety factor of 42. For the analysis presented here, the LEFM analysis is confined to the H4 weld and above. The fluence corresponding to welds at and below the core plate elevation is an order of magnitude lower and the associated fracture toughness is comparable to that of I the unirradiated material. For those locations, limit load analysis is used. An additional consideration that applies only to the fracture mechanics analysis is the question, "When is a flaw independent of an adjacent flaw?" The AShE Code proximity rule considers how flaws can link up and become a single flaw as a result of proximity However, even when two flaws are separated by a ligament that exceeds the criterion. they may not be considered totally independent of each other. That is, the flaw tip stress intensity factor may be affected by the presence of the adjacent flaw. This can be accounted for by using the finite width correction factor for a flaw in a finite plate. For a through-wall flaw in an " infinite" plate, the stress intensity factor is: K = c4(na) 42
GE Nuleap Euror_ GENE-383176-1293 For a finite plate, the K value is higher as determined by the finite width correction factor, F. In this screening evaluation it is assumed that the plate is " infinite" if the correction factor F is less than 1.1. As seen in Figure 5-3, if the width of the plate exceeds 2.5L (or a/b less than 0.4), then there would be no interaction due to plate end edge effects. If this ame condition is applied to two neighboring flaws, then there will be no interaction between the two indications if the tips are at least 0.75(Ll+L2) apart. If the distance between indications is greater than 0.75(L1+L2), then they are considered as two separate ) flaws. However, if they are closer, for the purpose of fracture analysis, the equivalent flaw length is the sum of the two individual flaws. 5.1.3 Limit Load Analysis A through-wall circumferential flaw was assumed in this calculation. Limit load calculations were conducted using the approach outlined in Subsubarticle IWB-3640 and Appendix C of Section XI of the ASME Code. The flow stress was taken as 3Sm. The value for the shroud material (Type 304 stainless steel)is 16.9 ksi at the normal Sm operating temperature of 550 F. Safety factors similar to that used in the ASME Code (2.8 for normal and upset and i 4 for emergency and faulted) were used in the analysis. The highest seismic stress was used for the limit load calculations and is shown in Table 5-1. Similarly, the highest axial pressure stress corresponding to the lower shroud was used. Thus, the analytical re are applicable for all welds since limiting values are used. 43
GENuclear Energy GENE-323-176-1293 5.2 Allowable Through Wall Flaws Allowable through-wall flaw sizes were determined using both fracture mechanics and limit load techniques for both circumferential and axial flaws. It should be emphasized that the allowable through-wall flaws are based on many conservative assumptions and are intended for use only in the screening criteria. More detailed analysis can be performed to justify larger flaws (both through-wall or part through when measured flaw depths are available). However, since the intent of the screening criteria is to determine when additional evaluation or NDE characterization is needed, a conservative bounding approach is utilized. 5.2.1 Allowable Through-Wall Circumferential Flaw Size Both the LEFM and limit load methods were used to evaluate the allowable through-wall flaws. Above the core plate, LEFM and limit load analysis methods were used. Since this is a screening criteria, single allowable flaw size criteria (limiting location) was used for all weld locations. It should be noted that the H7 and H8 welds involve Alloy 600 which has higher Sm values and therefore has higher limit load capability. Fracture Mechanics Analysis The total axial pressure and seismic stress corresponding to the upset condition is 0.61 ksi, and 1.39 ksi for the faulted condition. Using the ASME Code safety factors for fracture analysis, the faulted condition is limiting. To determine the allowable flaw size based on LEFM methods, the conservatively estimated irradiated material fracture toughness K cI value of 150 ksiVin was used. Applying a safety factor of 1.4 for the faulted condition, the allowable KI of 107 ksiVin was obtained. The allowable flaw size was calculated using the following equation: K I= Gm *o*V(na) where Gmis a curvature correction factor as defined in Figure 5-4 (Reference 5-1), o is the axial stress, and 'a'is the half flaw length. The allowable through-wall circumferential flaw length (2a) was determined as 2 344 inches. 44
GENustrar Energy GENE 323-1761293 Limit Load Analysis ! {. I A through-wall circumferential flaw was assumed in this calculation. The limit load calculations were conducted using the approach outlined in Subsubarticle IWB-3640 and Appendix C of Section XI of the ASME Code. The flow stress was taken as 3Sm. The Sm value for the shroud material is 16.9 ksi at the normal operating temperature of 1 550 F. The stresses for the limit load analysis for the upset condition consisted of an axial force stress of 0.71 ksi, and a bending moment stress of 0.49 ksi. i Similarly for the faulted condition, the axial force stress was 1.21 ksi, and the bending moment stress was 1.1 ksi. The allowable flaw length was approximately 430 in. including the ASME Code, Section XI safety factors. 5.2.2 Allowable Through-Wall Axial Flaw Size Fracture Mechanics Analysis The allowable axial flaw size is governed entirely by the pressure hoop stress. Similar to the circumferential flaw case, the allowable axial flaw size was determined assuming a through-wall flaw. For a through-wall flaw oflength 2a in the shroud, the applied stress intensity factor is given by: K = M
- oh
- V(na) where M is the curvature correction factor. M is given by:
M=Gm+Gb (Figure 5-5, from Reference 5-1) In the above expression, the allowable flaw length,2a, can be determined by equating the ; calculated K to the fracture toughness divided by the safety factor of 3. The hoop stress l is 1.85 ksi and the allowable K = 150/3 (where 150 ksiVin represents a conservative j estimate of the material toughness and 3 is the safety factor). The allowable flaw length was conservatively detemiined to be 2a = 59 in. j i 45 1 i
GENustear Energy GENE-5231761293 Linst Load An alternate approach to determining the allowable flaw size is to use limit load techniques. The allowable flaw length is given by the equation: ch " of/ (M1* SF) is a curvature correction factor (which is a function of the flaw length where M1 (Reference 5-2)), or= 3Sm is the flow stress, SF is the safety factor of 2,8 for upset conditions, and ch = the hoop stress corresponding to the upset AP of 35.68 psi. The allowable flaw length based on the limit analysis is 200 in, which exceeds that determined by LEFM. Thus, the allowable axial through-wall flaw length is 59 in. 46
GENudear Energy GENE-523-1761293 5.3 Screening Criteria The determination of the allowable through-wall flaws has been described in Section 5.2. The objective was to use the allowable flaw size as the basis for the screening criteria. Since the screening rules represent the first step in the evaluation, they are by defmition conservative. If the criteria are exceeded, the option of doing further detailed evaluation or performing additional NDE remains. The effective flaw lengths (Lleft, L2 ef r, etc.) determined by combining indications using the proximity and interaction rules, are used in the comparison with the allowable flaw sizes. The determination of effective flaw sizes are discussed in detailin Appendix A. The allowable through-wall flaws were:
- Circumferential Flaws
- 344 in. using LEFM
- 430 in. using limit load
. Axial Flaws
- 59 in. using LEFM
- 200 in. using limit load A conservative approach in developing the screening rule is to include both the LEFM and limit load analysis. For axial flaws, the allowable flaw length based on the LEFM controls, and the screening limit is 59 in.
For circumferential flaws the fracture mechanics based limit for a single flaw is 344 in. I This in itselfis not sufficient since there could be several flaws (each less than 344 in.) in a circumferential plane that cumulatively add up to greater than 430 in. (the a!!owable circumferential flaw size based on limit load analysis). Thus, the cumulative flaw length should be less than 430 inches. While this fully assures the ASME Code margins, an additional conservatism is included in the screening. This states that the cumulative flaw length cannot be more than 430/4 = 107.5 in. in any 90 degree sector of the shroud. This is a conservative restriction that assures that long continuous flaws are not admissible. With the provision that the cumulative flaw length cannot exceed 107.5 in. in any 90 sector of the shroud, this criterion becomes more limiting than the fracture l mechanics limit of 344 in. The approach used here for the 107.5 inch limit for circumferential flaws is to assume a template with a moving window equal to the 90 sector. The cumulative length of flaws that appear in the window should be less than 47
GENE-523-176-1293 GENustear Energy 107.5 in. A similar restriction based on limit loads is not needed for axial flaws since they are associated only with circumferential welds and are unlikely to be aligned in the same plane. The removal or reduction of the factor of 4 isjustified if the IVVI results indicate that the remaining ligament is spread around the circumference of the shroud circumference. l Evaluation of the IVVI results indicate that this is the case for the Peach Bottom Unit 3 shroud (See Figure 1-2 and 1-3). It should be noted that when considering LEFM based evaluations, the crack interaction criteria described in Appendix A, must be applied in comparing against the allowable lengths. For example, the adjacent flaws where the spacing S is less than 0.75 (Ll ef y+ L2 y), the length L=L1efr+ L2 eft si used for comparison with the LEFM based ef allowable flaw length. 1 I i 48 4
GENuclear Energy GENE-523-176-1293 l l 5.4 Summary of Screening Criteria The screening criteria is schematically shown in Figure 5-6. The first step is to map the flaw indications observed by IVVI. Next the prodmity mies are applied to the flaw map to develop effective flaw lengths (Appendix A provides the details for determining effective lengths). The results of the effective flaw lengths are also mapped. For axial flaws located in a vertical plane, two neighboring flaws must be summed if S < 0.75(L1 eff+L2 ef r). If the longest resulting flaw is less than 59 inches, then the screening limit is met for axial flaws. For circumferential flaws, all flaws are summed in any 90 sector using a template. The total flaw length in the 90 window must be less then 107.5 inches to meet the screening criteria. The next step is the LEFM based comparison using the interaction criteria. If S <0.75 (Llefr+L2 rr), e then the length L = L1 efr+ L2 ef rshould be compared with the LEFM limit of 344 in for circumferential flaws. If significant ligament remains around the shroud circumference, the factor of 4 may be removed or reduced. The application of this factor would be considered conservative since the presence of the ligament around the circumference assures that extremely long indications are not present. The removal or reduction of the factor of 4 would be considered as part of the "Further Evaluation" box in Figure 5-6. 49 s
GE Nuclear Energ_ GENE-323-176-1293 5.5 Justification for Continued Operation The screening criteria was used to evaluate the Unit-2 shroud assuming that the Unit-2 shroud contains the indications found by IVVIin the Unit-3 shroud. The structural integrity of the core shroud is assured if the screening criteria are met for all of the indications. All axial effective indication lengths are significantly less than the allowable flaw size based either on LEFM or limit load methods. Thus the axial indications seen by IVVI are acceptable per the screening criteria developed for Unit-2. The effective indication lengths were determined for all of the circumferential indications as shown in Figure 1-2. The calculation took-into consideration the detailed geometric information such as each indications length, the azimuth of each crack tip and the spacing I between indication planes in order to properly determine the effective length. J l The resulting effective indication lengths were then compared against the allowable i indication length. Since all effective indication lengths satisfy the screening criteria, the structural integrity of the Unit-2 shroud (postulating that indications found in Unit-3 shroud are present in the Unit-2 shroud)is assured for the current two year cycle with power rerate conditions. l l As mentioned in Section 5.3, for the indications observed in the Peach Bottom Unit 3 shroud, it is justified to remove or reduce the factor of 4 when detennining the allowable throughwall circumferential flaw length. Again, this is justified since there was remaining ligament spread around the entire shroud circumference. If this is done, the observed indications, and resulting effective flaw lengths are well below the allowable throughwall circumferential flaw lengths. It should be noted that the limiting allowable throughwall circumferential flaw is now governed by the LEFM method and not the limit load method. Since the Unit-3 shroud is expected to be limiting, the removal of the factor of four for this Unit-2 evaluation is also considered justified. 50
GENE-3231% ]293 GE Nuclw Emu 5.6 References 5-1. Rooke, D.P. and Cartwright, D.J., " Compendium of Stress Intensity Factors," The Hillingdon Press (1976). 5-2. Ranganath, S., Mehta, H.S. and Norris, D.M., " Structural Evaluation of Flaws in Power Plant Piping," AShE PVP Volume No. 94 (1984). l 1 1 I I I I l 51 l
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GENE 5231%2193 GE Nuclear Energy __ IWI Y Map All Flow Indications lf f
> M p Effective Flow Lengths - Circumferentic Fi Flows I
II Axial Flows No Sum of Flows in any 90 0 Sector <107.5"* Y Yes No j <0.75(Llef t + L2,gr) For Adjacent Flows? j No , y S<0.75(L1ett +L2ett) I T for Adjacent flows? L= L1,gg or L2 egg l Yes ' Yes No y,3
= Llegg +L2eff<275" L= Lleff +Geff l Y Y Further
> L < 59"? ? Evoluotion ' /
Yes Y t Continued Operation m Justified Figure 5-6 SCHEMATIC OF SCREENING CRITE=iA 57
GENE-583-176-1293 GE Noclear Energy 1 6.0
SUMMARY
AND CONCLUSIONS An evaluation was performed to evaluate the Peach Bottom Unit 2 shroud based on GE SIL 572, Rev. I recommendations and comparison with the Peach Bottom Unit 3 shroud. Justification for continued operation of Unit-2 was based on evaluating Unit-2 against Unit-3 water chemistry, fluence and fabrication information. The comparison showed that the Unit-2 conditions were less severe than that at Unit-3 from a stress corrosion viewpoint. Therefore, it is reasonable to use the Unit-3 IVVI results for Unit-2 as bounding. Based on this, continued operation of Unit-2 for one more cycle is justified. Indications have been observed in the Peach Bottom Unit 3 core shroud. Indications were i seen during in-vessel visual inspection (IVVI) of the various shroud welds as recommended by GE SIL 572, Rev.1. This evaluation was performed to demonstrate that the structural integrity of the ' Unit-2 shroud is maintained for the current fuel cycle (two year cycle with power rerate conditions) even with postulated indications similar to that in Unit-3. In addition, the repon documented material, water chemistry and fluence I information which show that Unit-2 is likely to be better than Unit-3 from the stress corrosion viewpoint. This report demonstrated that even with several conservatisms in the evaluation, the structural integrity of the shroud is maintained during a limiting event, even with the postulated indications. This was performed by developing conservative screening criteria, , l assuming throughwallindications, which can determine the acceptability of the flaws based I solely on the IVVI results. The assumption of through-wallindications removes any uncertainty regarding sizing and the need to funher characterize the indications. By meeting the screening criteria, the ASME Code Section XI safety margins are satisfied Based on the results of the application of the screening criteria to the observed indications, it is concluded that the structural integrity of the Unit-2 shroud is maintained for the current fuel cycle even with postulated flaws similar to that in Unit-3. All effective indication lengths were shown to be less than the allowable flaw size. The use of the Unit-3 indications was justified based on a comparison of water chemistry, fluence and fabrication features between the two plants. 1 58
. . - ~ . - . . .- . . - . . - - - - - .
GENE-523-176-1293 GENucleu EnerK7 APPENDIX A DETERMINATION OF THE EFFECTIVE FLAW LENGTII i The effective flaw lengths are based on ASME Code, Section XI proximity criteria as presented in Subarticle IWA-3300. The procedure addresses both circumferential and axial flaws. Indications are considered to be in the same plane if the perpendicular distance oetween the planes is less than 4" (2 times the shroud thickness). All flaws are considered to be through-wall. Therefore, indications on the inside and outside surface should be treated as if they are on the same surface. When two indications are close to each other, rules are established to combine them based on proximity. These rules are described here. A.I Proximity Rules The flaw combination methodology used here is similar to the ASME Code, Section XI proximity mies concerning neighboring indications. Under the rules, if two surface indications are in the same plane (perpendicular distance between flaw planes <4") and are within two times the depth of the deepest indication, then the two indications must be considered as one indication. i in Figure A-1, two adjacent flaws L1 and L2 are separated by a ligament S. Crack powth would cause the tips to be closer. Assuming a conservative crack growth rate of 5x10:5 in/hr, crack extension at each tip is 0.8 in, for 16,000 hours or one fuel cycle (See 1 Appendix B for crack growth rate discussion). Therefore, combining the crack growth i and proximity criteria, the flaws are assumed to be close enough to be considered as one. continuous flaw if the ligament is less than (2 x 0.8 + 2 x shroud thickness). For a shroud thiclo.ess of 2.0 in., this bounding ligament is 5.6 in. Thus, if the ligament is less than 5.6 1 inches, the effective length is (Ll+L2+S+1.6"). Note that the addition of 1.6 in. is to include crack growth at the other (non-adjacent) end of each flaw (See Figure A-2).. , if the ligament is greater than 5.6 in., then the effective flaw length is determined by l adding the projected tip growth to cad tna af the flaw. For this example, Llefy= L1 + j 1.6", and L2efy= L2 + 1.6" l 1 A similar approach is used to combine flaws when a circumferential flaw is close to an ) axial flaw (See Figure A-3). If the ligament between the flaws is less than 4.8 inches, then 1 A-1 i
+ , .
GENaclear EnnKY GENE-523-1761293 the effective flaw length for the circumferential flaw is Leff = Ll+S+0.8" (the bounding i ligament for these cases). If the ligament is greater than 4.8 in., then the flaws are treated separately. After the circumferential and axial flaws have been combined per the above criteria, a map I of the effective flaws in the shroud can be made, and the effective flaw length can be used l for subsequent fracture mechanics analysis. l In order to demonstrate the proximity criteria, three examples are shown in Table A-1 and described below. 1 Table A-1 Flaw Combinations Considered in Proximity Criteria i Case Circumferential Flaw Axial Flaw
\
A Yes No B Yes Yes l l No Yes C l l A.1.1 Case A: Circumferential Flaw - No Axial Crack This case applies when two circumferentialindications are considered. Figure A-2a shows this condition. If the distance between the two surface flaw tips is less than 5.6", the indications must be combined such that the effective length is (See Figure A-2b): Leff = L1 + S + L2 + 1.6" where. L1 = length of first circumferential indication L2 = length of second circurdrentialindication S = distance between two indications A-2
- , - - , , - . . ~ . - - . - . . -. . . .- - _ - - . . - - . .
GENE-523-176-1293 GE Nuclear EurKY If the distance between the two tips is greater than 5.6", the effective flaw lengths are (See Figure A-2c): Lleff = L1 + 1.6" ' . L2efr= L2 + 1.6" A.1.2 Case B: Circumferential Flaw - Axial Flaw This case applies when both a circumferential and an axial flaw are being considered. Figure A-3a demonstrates this condition. For this case, only growth of the circumferential flaw is considered. If the distance between the circumferential indication tip and the axial indication is less than 4.8", then the effective circumferential flaw length is (See Figure A-3b): 'l 4 ! L eg= L1 + S + 0.8" , where: L1 = length of circumferential indication ; S= distance between the circumferential tip and axial flaw. and the effective axial length is (Figure A-3b): L efr= L2 + 1.6" where: L2 = length of axialindication If the distance between the circumferential indication tip to the axialindication is greater than 4.8", then the flaws are not combined (See Figure A-3c) and the effective lengths are: Lleg= L1 + 1.6" (for circumferential flaw) L2eg= L2 + 1.6" (for axial flaw) , A.1.3 Case C: No Circumferential Flaw - Axial Flaw This case ap plies to when only axial flaws are being considered. The effective length is determined in a manner similar to that used for case A for circumferential flaws. 1 a A-3 ,
GENuclear Energy GENE-523-1%-1293 A.2 Application of Effective Flaw Length Criteria The application of the effective length criteria is applied to two adjacent indications at a time. Figure A-4 is a schematic which illustrates the process. For example, using the 0 azimuth as the starting location for a circumferential weld or plane, the general procedure would be as follows:
. Moving in the positive azimuthal direction, the first indication encountered is indication 1.
. The next indication is indication 2.
. Apply proximity rules to the pair ofindications (indications 1 and 2). Combine the flaws if necessary (Ll+L2+S). Old indication 2 becomes new indication 1.
. Continue along positive azimuthal direction until the next indication is encountered. This becomes new indication 2.
. Apply proximity mies to new indications 1 and 2.
. Continue proximity rule evaluation until all indications along the subject weld or plane have been considered.
l l l A-4 ,
GENE-523-176-1293 GE Nuclear Enervo t 4 >
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i i i I nc 1 Combined L1 l Flow I I I I if i D1 JL m -- l i L i t S D2 3r 4 > JL l L2 nc. 2 i l i 1 1 U T N Figure A ASME Code Proximity Critera A-5
. GENE-523-1761293 GE Nuclear Energy _.
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Figure A APPLICATION OF PROXIMITY PROCEDURE TO NEtGHBORING CIRCUMFERENTIAL FLAWS A6
GENE 323176-1293 g news Assomed Througn*cil ll g f y As-Fou n c l weio / 3> (3a) - Y
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ows Assumed Throughwcll
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GENE-523-176-1293 GE Nusicar Energy Start ct Theta =0 Move in + Theta Direction I l lf i= 1 Y Rrst Rcw l is Flow i l l lf p ,Next Figw is Flow i+ 1 Y l l Perform Effective Length Calculation l lf ( Combine Flows If Necessary To Determine Effective Length i=i+1 lI Rcw i+1 = Ficw i f I lf No
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+dl Done l Figure A4 - PROCESS FOR DE .'ERMINING EFFECTIVE CIRCUMFERENTIAL FLAW LENGTH l
A-8
GENachar EnerKY GENE-523176133 APPENDIX B BASIS FOR THE CRACK GROWTH RATE The basis for the crack growth rate used in the screening criteria is provided in this section The Peach Bottom Unit-2 shroud cylinder was fabricated from roll formed Type 304 stainless steel plate. Therefore, the weld heat-affected-zone (HAZ)is likely sensitized. , The shroud is also subjected to neutron fluence during the reactor operation which further increases the effective degree of sensitization. The other side-effect of neutron fluence induced irradiation is the relaxation of weld residual stresses. The slip-dissolution model developed by GE quantitatively considers the degree of sensitization, the stress state and the water environment parameters, in predicting a stress corrosion cracking (SCC) growth rate. The crack growth rate predictions of this model have shown good correlation with laboratony and field measured values. This model was used to predict a Peach Bottom , Unit-2 specific crack growth rate and a conservative value was then selected. B.1 Slip-Dissolution Model Figure B-1 schematically shows the GE slip-dissolution film-rupture model (Reference B-
- 1) for crack propagation. The crack propagation rate V tis defmed as a function of two constants (A and n) and the crack tip strain rate. The constants are dependent on material ,
and environmental conditions. The crack tip strain rate is formulated in terms of stress,- ; loading frequency, etc. When a radiation field, such as the case for the shroud, is present, there is additional interaction between the gamma field and the fundamental parameters . which affect intergranular stress corrosion cracking (IGSCC) of Type 304 stainless steel (see Figures B-2 and B-3). The increase in sensitization (i.e., Electrochemical Potentiokinematic Reactivation, EPR) and the changes in the value of constant A as a function of neutron fluence (>1MeV) is 4 l given as the following: i (B-1) EPR = EPRo + 3.36x10-24 (fluence)l.17 1 2 l where, EPR is in units of C/cm2, fluence is in units of n/cm and the calculated value of j EPR has an upper limit of 30. l B-1 ) 1
{ GE Natear Eurg GENE-323-176-1293 The constant A is defined as the following: 2 (B-2a) for fluence 5 1.4x10 19 n/cm . C = 4.1x10-14 i 2 2 for fluence > 1.4x1019 n/cm but 5 3x1021 n/cm :, (B-2b) C = 1.14x10-13 In(fluence)- 4.98x10-12 2 (B-2c) for fluence $ 1.4x1019 n/cm : C = 4.1x10-14 The units of K to be used with the above expressions is MPaYm.
)
B.2 Calculation of Parameters i The parameters needed for the crack growth calculation by the GE model are: stress state 1 and stress intensity factor, effective EPR, water conductivity, and electro-chemical l corrosion potential (ECP). : l i The stress state relevant to IGSCC growth rate is the steady state stress which consists of l weld residual stress and the steady applied stress. Figure B-4 shows observed through-wall weld residual stress distribution for large diameter pipes. This distribution is ' expected to be representative for the shroud welds also. The maximum stress at the surface was nominally assumed as 35 ksi. The steady applied stress on the shroud is due to core differential pressure and its magnitude is small compared to the weld residual stress magnitude. Figure B-5 shows the assumed total stress profile used in the evaluation. Figure B-6 shows the calculated values of stress intensity factor (K) assuming a 360 circumferential crack. It is seen that the calculated value of K reaches a maximum of approx. 25 ksiVin. The average value of K was estimated as 20 ksidin and was used in the crack growth rate calculations. The weld residual stress magnitude is expected to decrease as a result of relaxation produced by irradiation-induced creep. Figure B-7 shows the stress relaxation behavior of Type 304 stainless steel due to inadiation at 550 F. Since most of the steady stress in the shroud comes from the weld residual stress, it was assumed that the K values shown in Figure B-6 decrease in the same proportion as indicated by the stress relaxation behavior of Figure B-7. B-2
. -~. .- . . . - . . ~-- _ . ~ _ . . ..
I GE Nuclear Energy GENE-523176-1293 F The second parameter needed in the evaluation is the EPR. 'In the model, the initial EPR value is assumed as 15 for the weld sensitized condition. Using Equation (B-1), the ; predicted increase in EPR value as a function of fluence is shown in Figure B-8, The third parameter used in the GE predictive model is the water conductivity. - The . reactor water conductivity at Peach Bottom Unit-2 has recently (1992-93) been good 2 (approx. 0.1 S/cm compared to earlier operating period (1975-84). This has a significant impact on the predicted crack growth rate by the GE model as seen in Figure ' B-9. To demonstrate that the GE model conservatively reflects the effect of conductivity, Figure B-10 shows a comparison of the GE model predictions with the measured crack growth rates in the crack advance verification system (CAVS) units installed at several BWRs. The comparison with CAVS data in Figure B-10 also demonstrates the conservative nature of crack growth predictidns by the GE model. The last parameter needed in the GE prediction ~modelis the ECP. Figure B-11 shows the measured values of ECP at two locations in the core. Since Unit-2 does not plan to use hydrogen injection, the ECP values at zero H 2 injection are relevant in Figure B-11. It is seen that the ECP values at zero H2 injection rate range from 150 mV to 225 mV. Therefore, a value of 200 mV was used in the calculation. B.3 Crack Growth Prediction Based on the discussion in the preceding section, the crack growth rate calculations were conducted as a function of fluence assuming the following values of parameters: Initial K = 20 ksiVin EPRo = 15 C/cm2 Cond. = 0.1 S/cm2
= 200 mV l ECP U
Figure B-12 shows the predicted crack growth rate as a function of fluence. It is seen that the predicted crack growth rate initially increases with the fluece value but decreases later as a result of significant reduction in the K value due to irradiation induced stress relaxation. The crack growth rate peaks at 4.5x10-5i n/hr at a fluence of lx1020 n/cm2 Thus, a bounding The crack growth rate for the expected fluence at the end of the current operating cycle is 2x10-5 i n/hr. Thus, a bounding value of 5x10-5 in/hr can be conservatively used in the structural integrity evaluation for the shroud. B-3
l GENuclear Energy GENM23-1W1293 l l This bounding crack growth rate is quite conservative as can be shown in Figure B-13 l from NUREG-0313, Rev. 2. It is seen that the crack growth rate of 5x10-5 in/hr at l 20 ksiVin is considerably higher than what would be predicted by using the NRC curve. This further demonstrates the conservatism inherent in the assumed bounding value of l crack growth rate. l B.4 Conclusion A crack growth rate calculation using the GE predictive model was conducted considering ) the steady state stress, EPR, conductivity and ECP values for the Unit-2 shroud. The evaluation accounted for the effects ofirradiation induced stress relaxation and the increase in effective EPR. The evaluation showed that a bounding crack growth rate of 5x10-5 .in/hr may be conservatively used in the structural integrity evaluation of the Unit-2 sluoud. B.5 Reference B-1 F.P. Ford et al, " Prediction and Control of Stress Corrosion Cracking in the Sensitized Stainless Steel / Water System," paper 352 presented at Corrosion 85, Boston, MA, NACE, March 1985 1 l 1 l l l
)
B-4 ;
GENE-323176-1293 GE Nucim Ewv_ The GE PLEDGE Slip Dissolution - l Film Rupture Model of Crack Propagation l I i 6 CT e VT Crack-t.ip advance by enhanced oxidation at strained crack tip i o
)
.n VT = AE CT Where:
V T = crack propagation rate 4 A, n = constants, dependent on material and environmental conditions
= crack-tip strain rate, formulated in /
E CT terms of stress, loading frequency, etc.' Figure B 1 B5
SOLUTION RENEWAL RATE TO CRACK-TIP R. N STRESS F-E A p ANIONIC N OXIDE RUPTURE f
,, TRANSPORT RATE AT ENVIRON-CRACK-TIP MICRO-STRUCTURE y-FIELD -
AN CRACK TIP $(A)~, pH PASSIVATION RATE 9,
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GENE-5331%]193 GE Nuc!<u EntrO OBSERVED RESIOUAL STRESS PMOFILES l IN H AZ OF 24"-28" DI A. SCH. 80 PIPING l l l l l 1 I I l 40 -
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gggg 333 375.,ggy GENukar Dupe __ OD 40 - ID - 206.9 50 - Total Stress .W i
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_4e 10.16 15.24 10.32 O 5.04 DEPTH (mm) Figure B-5 Shroud Total Throughwall Stress Profile B-9
GENE-523-176-1293 GE Nuclear Energy s 26 ~- 27.25 24 - 22 -
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9 GENE-523-176-1293 GE Nuclear Energ_ 1 i Stress Relaxation Behavior . from Irradiauon Creep cn 0.8 -
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22 0 2 102 ' 10 10 10 ' 2 Neutron Fluence, n/cm (E>1 MeV) Figure B 7 Stress Relaxation Behavior of Type 304 Stainless Steel Du B-11
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11992-1993
-300 mV 1.000 E- 06 __
- A PB-2
~ O PB-3
...i
/'
" ' "--l- (i
'- 2 --
2J 1.000E- 07 0.6 0.7 0.8 0.9 E 0.1 0.2 0.3 0.4 0.5 .:. 0 f Conductivity, pS/cm E , PLEDGE: 15 C/cm2, 20ksi/in GliNii PLliDGli Mottel Pretliction l'or Peach Botom Unit-2/3 Sensitizeti 'Iype 304 Crack Gro Figure H-9
a Effect of Conductivity on Sensitized 304 = Crack Growth Rate = D k Crack Growth Rate, in/h 200 mV 1.000E-04 5.000E-05 100 mV 1.000E- 05 ; on NO 3 oc o AJ to A AG no , 3ao o 1.000E-06 - c o
- u C
_ g AG O Q u TMM M ' '
' h !
i 0 0 0.35 g 0.25 0.3 1.000E-07 -- 0.15 0.2 - 0.05 0.1 f Conductivity, pS/cm E PLEDGE: 20 ksi/in,15 C/cm2 . CAV: 20-25 ksi/in,13 C/cm2,100-160 mV Firure 11-10 liffect of Cominctivity on Sensitized Type 304 Crack Growth Rate
a ' In-Core Bypass ECP vs Feedwater Hydrogen 54 for a BWR-4 1 ECP, mV (SHE) 300 O 200-fl 100 - . 0 - Just Below Top Guide Level G
-100 -
~'
wp
\ ' tit
-200 -
sx.
' 't t Above Core Plate Level m
-300 -
N --- . g . ._ .a g
' '- J g 80 90 hv
-400 50 60 70 20 30 40 y 0 10 Feedwater H2, SCFM Figure 11-1i in-Core 13ypass ECP vs Feedwater I'lydrogen for a BWR-4 w--____-__-_________-_-__________---_-- . _ _ _ _ _ _ _ _ _ _ _ . . , _ - _ _ - - _ - _ - _ _ _ _ _
o m Figure 11-12 E R-R m j ' GROWTH RATE VERSUS FLUENCE
- - -- J !
i.. 4 . :. -.
- _.-l- -- - ;_
1.00E-04 i _
~ _ _
- . ~ -
. . . }
^
6.-
~
E ZTT J'~~ J TZ ~-- Z Z~ ~
.v ~l~
~~ -
Z. Z
- Z - - --
~~ ~~~ Z_T. ~ ~_ __T Z'._~_ _E. _ _- -
E 1.00E 05
~
Z Z _" Z Z -- 2 - _ - . . - . t _ ___ _ _. - - _ _ _ . __ -~- i .
. _ _ ~ - - --
g . __ ~-- -- --
~ ~-~" ~ "-~ ~ 1E + 21
" {
1.OOE-06 i 1E+20 9 1E+19 Fluence (n/cm* 2) F M e Stress intensity = 20 Ksi/in, initial EPR = 15 C/cm2
GENuclear Enerv GENE 523-1761293
~
10 " _ _
- NRC CURVE e -
1inlyr ,
/,
~4 "
10
/)6 e -
i - -
,s.q'/
10
/.d _ $ O!/ y 0.2 ppe 0 ;2 sensttlaed at 1150*f/2 h _
w (EPR
- 15 C/cm ) GC -
g - A 0.2 ppa 0 2; senslttred at 1150'F/2 h (tPt 10 C/cm ) Gt C "
- 0 0.2 99 Or t 5'a5d " 5 2'
~
~
F- . CE 3: Q 0.2 ppa 02 I '"IY *'i*td o 8 , , 0,; sensitized .i iiS0 r/24 h g 0-6 _ GE (D MifACHI _ GEMED G ANL -~ M - GECAD y HOH8ppe0 2 "'IIid tI ** di"i - e - 0.04 in /yr tr$ .i in r/2. n (Sti) 3 8 ppe 0 ; sensttired at (1292'F/10 min)S U 2
* (932'F/24 h) (L7t
- 4 C/cm ) ~~
~
6 8 ppa 0 ;2sensttired at (1292*F/10 min) l
+ (132'f/24 h) (Eft
- 4 C/cm )
l r 0.1 wi. t 0.54 - 10 7
- 8 ,p. 0,; sensitized .i oZs2 in0 mini i . (842'f/257 h) (CPR = 15 C/cm#)
@ 8 ppe 0 ;2 seasttized at (1292*F/10 min) NT l t
+ (842*r/257 h) (CPR
- 15 C/ca ) AN t, 1 f
- 0.1 M r . R
- 0. 9 4 DATA 9 8 ppe 0 2; sensitized at l!92'F/14 h l
(Ett a 20 C/cat ) g , g,gog gg, I t
- 0.15 l X 8 e 0,; seaststied it 12$2*r/14 h I
((72 = 20 C/cm ) f
- 0.08 Hz.
l t = 0.15 s 8 1 0 10 20 30 40 50 60 70 l
)
STRESS INTENSITY,K (ksid.) l l Figure B-13 Nureg 0313 Crack Growth Rate Data l l l 1 B-17 1}}