ML20064K429
| ML20064K429 | |
| Person / Time | |
|---|---|
| Site: | Peach Bottom  |
| Issue date: | 12/03/1993 |
| From: | Marisa Herrera, Mehta H, Ranganath S GENERAL ELECTRIC CO. |
| To: | |
| Shared Package | |
| ML20064K422 | List: |
| References | |
| GENE-523-141-10, GENE-523-141-1093-R1, NUDOCS 9403230107 | |
| Download: ML20064K429 (93) | |
Text
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GENE-523-141-1093, Rev; 1 DRF 137-0010-6 i
Evaluation and Screening Criteria for the Peach Bottom Unit-3 Shroud Indications i
December 3,1993
/
Prepared by: ' Am va- "
1
'qafcos L. IIerrera, Principal Engineer Structural Mechanics Projects ;
vW Y Dr. Hardayal Mehta, Principal Engineer .
Structural Mechanics Projects Approved By: -
Dr. Sampath Ih(nganath, Manager Structural Mechanics Projects l
GE Nuclear Energy San Jose, CA l P
9403230107 940314 PDR ADOCK 0500027B P PDR L a
4i 4 IMPORTANTNOTICE REGARDING CONTENTS OF TIIIS REPORT Please Read Carefully The only undertakings of the General Electric Company (GE) respecting infc>rmation 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 infonnation contained in this document, or that its use may not infringe privately owned rights.
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Table of Contents EXECUTIVE
SUMMARY
iv
1.0 INTRODUCTION
.. . . . .. .. . . . . .I 1.1 Background . . . .3 1.2 Screening Criteria.. . . . . . .. . .. 3 1.3 References ... . . . . . . .. . . . . .. .6 ,
2.0 FABRICATION IllSTORY . . . . . . ... . . . . . . . 11 3.0 CIIEMISTRY AND FLUENCE CONSIDERATIONS.. . .. . .23 3.1 Water Chemistry History. . . .. .23 3.2 Fluence Considerations.. . . . . 25 3.3 References .. . . . ... . . . .. 26 4.0 IN-VESSEL VISUAL INSPECTION.. . . . . . . . .36 5.0 FLAW EVALUATION . . . . . .. .39 5.1 Structural Analysis. . . . . . . . .. . .. . .. . 39 '
5.2 Allowable Through-Wall Flaws . . . . .43 5.3 3creening Criteria.. .. .. . . . .. .46 5.4 Summary of Screening Criteria.. . . . . . . z. . . . .48 5.5 Application of Screening Criteria.. .. . . . . . .. .49 5.6 References . . ... .. . . . . . . . . . .50 6.0
SUMMARY
AND CONCLUSIONS. . . . . . . . .57 DETERMINATION OF THE EFFECTIVE FLAW LENGTH [
APPENDIX A APPENDIX B BASIS FOR CRACK GROWTH RATE k, O
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GENuclear Energy GENE-583-141-1M3, Rev.1 EXECUTIVE
SUMMARY
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 showed that both circumferential and axial indications 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, circumferentialindications were observed in the shrou, plate associated with a venical 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 disposition the indications by demonstrating that the structural integrity of the shroud is maintained for the next fuel cycle (two year cycle with power rerate conditions). In addition, the report documents material, water chemistry and fluence information which are additional variables which may have contributed to the shroud condition.
The primary focus of this report is to demonstrate that even with several conservatisms in the evaluation, the stmetural integrity of the shroud is maintained during a limiting event.
This was performed by developing conservative screening criteria, assuming throughwall indications, which can determine the acceptability of the flaws based solely on the IVVI results. The assumption of through-wall indications removes any 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-wall indication lengths. The limiting flaw length based on either LEFM or limit load was used for the screening criteria.
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 1ips was considered for the LEFM allowable flaw size calculation. The resulting effective flaw lengths were compared against the j screening criteria to determine if the structural integrity of the shroud was maintained. l iv l
GENuclear EnerO GENE 523141 1093, Rev,1 ,
a' ,
Based on the results of the application of the screening criteria to the observed indications, !
it is concluded that the structural integrity of the shroud is maintained for the next fuel cycle. All effective indication lengths were shown to be less than the allowable flaw size. t i
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GENuclear Emergy GENE-333.H1-1093, Rev.1
1.0 INTRODUCTION
The objective of this report is to document the conditions found on the Peach Bottom Unit-3 shroud, and evaluate these conditions based on GE SIL 572, Rev.1 (Reference 1-
- 1) recommendations, in order to validate the structural margins of the shroud.
Recently, in-vessel visual inspection (IVVI) of the Peach Bottom Unit-3 shroud revealed indications in the inside surface heat a!Tected zones (HAZ) at veld locations H3 and H4.
Figure 1-1 is a schematic illustrating the general locations of the shroud welds in Unit-3.
Figurcs 1-2 and 1-3 are the shroud maps which show the locations where indications were found. Figure 1-4 is a plan view which indicates the locations referred to in Figures 1-2 and 1-3. Fig' ire 1-2 shows the inside surface shroud map and Figure 13 shows the outside surface shroud map. Horizontal and verticalindications were seen associated with the H3 and H4 welds. Circumferentialindications were also observed associated with one of the vertical welds (V3). However, these indications were relatively short compared to those associated with the horizontal welds.
In addition to the H3 and H4 welds, IVVI of the H1, H2, H5, H6, H7 and H8 welds was performed on the outside surface. Only a few short indications were observed on the outside surface of H1 and H4. It should also be noted that the area adjacent to the H9 weld was visually inspected as part of the access hole cover inspection (AHC) at this outage. The inspection did not reveal any indications. Additional detail of the IVVI ,
results is presented in Section 4.0. .
GE SIL 572, Rev.1, provides the following recommendations based on the observed indications and evaluations performed to date:
Plant Fabrication and Oocrational History Review plants fabrication and operational histories for the core shroud, including the materials of construction. ,
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 1
, , GENcclear Enerv GEVEJ23141 1093, Rev.1 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 during each refueling outage. The SIL also provides a recommended examination process.
Destructive Testing A boat or core sample may be necessary depending on the results of the examination.
.Smtclural Margin Analysis Perform a stmetural 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 astessed using cumulative flaw length stmetural margin criteria.
.Cprrective ActinD Based on the results of the structural margin evaluation, determine if continued operation isjustified 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-3 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 i
. StructuralMargin Analysis
. Screening Criteria for Application to IVVI results. ,
l It is noted that the loads used in this evaluation correspond to those for power rerate. A I two-year operating cycle was used in the determination of crack growth.
2 i
GENuclear Energy GENE.333-H1-1H3, Rev.1
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 app oximately 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 l.8x1020 nyt (E>lMev). .
In addition to the H3 weld HAZ cracking, indications were visually observed at the H1, H2, H4, HS (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 weld,s.
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-wall j 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 allowa6te flaw size and selection ofindications for more detailed evaluation. The determination of effective flaw length is based on ASME i Code,Section XI, Subanicle IWA-3300 (1986 Edition) proximity criteria. These criteria 3
i
GENcclear Energy GENE-333-141-1M3, Rev.1
- 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.
6 Flaws are considered in the same plane if the perpendicular distance between the planes is 4" or less. Any flaws 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 efTective 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 compared against the allowable through-wall flaw size.
The report covers the limiting stresses for all the shroud welds (H1 through H8 welds).
Therefore, the screening criteria developed here cover all shroud weld indications. A list of conservatisms used in this evaluation is summarized in Table 1-1. ,
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., GENulcar Eurg GEN &583.In.100, Rev.1 Table 1-1 Conservatisms Included In Screening Evaluation i
- 1. All surface indications were assumed to be through-wall for analysis.
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- 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. k
- 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. j
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). .
- 11. A proximity rule to account for perpendicular flaws was applied, although not required by Section XI. k
- 12. Power rerate conditions were used although it will not be in effect during the next fuel cycle.
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1.3 References
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- 7E Nuclear Energy GENE-583-141-1M3, Rev.1
- 2.0 FABRICATION HISTORY This section describes the fabrication history of the Peach Bottom Unit-3 shroud. Of key interest is the material composition and any activities which could have possibly contributed to the increase ofintergranular stress corrosion cracking (IGSCC) susceptibility. 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-3 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. Figure 2-1 shows the assembly of the shroud. All welds 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 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, followed by water quenching to below 100 C. The carbon content of the rings ranges from 0.03% to 0.035% max. Hardness measurements upon completion of solution heat treatment and rough machining of the rings ranged from Brinell Hardness of 137 to 153. 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. l The carbon content of the plate material ranges from 0.057% to 0.062% max. The hardness of the plate material ranges from Brinell Hardness of 137 to 155. All welding was performed by submerged arc-welding except H7. The procedure and welder qualification was performed to ASME Section IX requirements. The filler metal met ASTM 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. 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 surfac:s of the welds were inspected by liquid penetrant examination. l 11
4 A
, . , GENxclear Energy GENE-523-141 1093, Rev.1 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 efTects contribute to SCC, it is likely that the cracking would be oflesser concern than cracks near the horizontal welds. t 9 I g 0 L O 12
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4 s 4 Peach Bottom 3 Shroud Data 3 a s Part Part Name Quantity Meterial lleet / Certincate Carbon Comments on l Number Number Content - % Material / Process 1 Upper Ring I Piece AtB2 - F304 F1343-68.168 0 035 2 Upper Cylinder 2 Placee *240 Type 304 3562-E9967 0.062 ( 3 Central Ring I Piece AtR2 - F304 F1399-65.587 0 030 4 Centra Cyunder 2 Pims A240 Type 304 5727-Ell 4 0.057 5 l 5637-Ette 0.060 S Centrol Cynnder 2 Piece A240 - Type 304 2170-E212 0.057 { l 2818-E15 0.060 - 6 Lower Ring i Piece A152-F304 F1400-66.t S9 0.035 [ 7 Lower Cyttader 2 Pieces A240 Type 304 3586-E40 0 059 2784-Ell 0 059 s a
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j NOTES
- l. CORE SPRAY SPARCER ASSEMBLY. INTERNAL AND EXTERNAL
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* *i GE Nuclear Energy GENE 523-141-1093, Ren 1 X l
-i 3.0 CHEMISTRY AND FLUENCE CONSIDERATIONS j i
I 3.1 Water Chemistry History l For the first decade of hot operation, Peach Bottom Unit-3 operated with relatively high primary water conductivity. As can be seen in Figure 3-1, Unit-3's arithmetic mean conductivity actually exceeded 1.0 S/cm in 1976 and 1977. The arithmetic mean L conductivity was very high and exceeded 0.4 S/cm through 1986. Subsequently, conductivity values steadily decreased, and was <0.1 S/cm (0.089 S/cm) during 1992 , and 1993. These last two year's conductivity values are considered world class performance. The high conductivity during the first half oflife was partly due to leaking condensers and resin change out problems. However, besides the high early life steady state conductivity, there was also one relatively severe transient experienced at Unit-3 ~asl presented in Table 3-1, a summary of BWR tleet severe transients through about 1983. As can be seen in Table 3-1, Unit-3 sufTered at least one power resin intrusion (Incident Rank 13) during . which the conductivity reached 23.6 pS/cm due to possible condensate demineralizer resin intrusion. This intrusion type ofincident results in the injection of sulfate into the RPV. Since IGSCC initiation and propagation in sensitized austenitic stainless steel and nickel P base alloys are controlled by the rate of cathodic reduction of species such as dissolved - oxygen, hydrogen peroxide and/or various oxyanions, then the additional presence of a - detrimental oxyanion, such as sulfate, would increase the cathodic current and thus accelerate anodic dissolution at the crack tip, i.e., IGSCC (Reference 3-1).
-t 3.1.1 E Tects ofImpurities on IGSCC ,
An example of the effects of sulfate / conductivity on crack initiation in uncreviced material is presented in Figure 3-2. It is clear that an increase in sulfate / conductivity results in an < 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 acceleration in crack , ; propagation rate (creviced) with sulfate is down in Figure 3-3. . Figure 3-3 displays June 1986 (not included in Table 3-1) Unit-3 oi-iine crack' monitoring data for sensitized Type : I j 1 304 stainless steel. The results clearly illustrate the change in crack growth observed aller two closely linked water chemistry transients of 4-5 S/cm, i.e., increases in water. ,
'1 23 ,
a
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. GENuclear Energy GENE-583-Hi-1093, Rev.1 conductivity due to intmsions of demineralizer resin material (Reference 3-5). This figure demonstrates the dramatic increase in crack growth rate (2X) with conductivity. Similar 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 similar kinetic effects on IGSCC initiation and propagation (References 3-7 and 3-8).
This high conductivity crack initiation and propagation acceleration factor is consistent 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 cracking 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 oflGSCC 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 quantified at the GE Research and Development Center based on a "first principles" model of crack advance known as the film mpture/ slip dissolution model (Reference 3-10). Predictions from the , film mpture/ slip dissolution model, PLEDGE (Plant Life Extension Diagnosis by GE), have been extensively 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 confidence over a range in observed crack growth rate of more than six orders of magriitude. , Likewise, it provides a very reasonable mean value and can accurately bound the observed crack growth rate in stainless steel 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) indicates the strong effect of conductivity on crack growth rate and by inference crack initiation .
~
Figure 3-5 presents a schematic estimation of 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 (0.752 pS/cm) were compared to those averages for the last two 24
l
- GENcclear Iterg GENE.5331411H3, Rev. I years. A value of 200 mV[SHE] was used for the electrochemical potential (ECP) in these calculations. As noted in Figure 3-5, a factor of approximately nine decrease in _
crack growth rate is obtained with the unit's decrease in conductivity. Thus, crack growth over the past few years has been significantly reduced by proper control ofwater chemistry. 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 panicular location. Shroud peak fluence was calculated by multiplying peak flux at the shroud location by the effective tbli power seconds of operation. The peak fluence in the Unit-3 shroud at the end of the next - fuel cycle will be approximately 7.9x1020 n/cm2 (E>lMev). Typically, the fluence varies with shroud azimuthal location and elevation. A though peak shroud flux may vary significantly from cycle to cycle, available flux results : aie 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 assut f oni that flux remains constant throughout the life of the plant. Further evaluation would be needed to quantify the uncertainty associated with this assumption. 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. The impact ofirradiation on core materials including crack growth rates has been
~
studied and is discussed in References 3-11 and 3-12. i f
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~
25 l l
GENcclear Fargy GENE-323-141-1093, Rev. I 3.3 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 M. Trolle, " Effects o.* Impurities on the IGSCC of Stainless Steel in High Temperature Water," Corrosion, Vol. 44, No. 2, February 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 ofType 304 Stainless Steel," paper 102 presented at Corrosion 85, Boston, MA, NACE, March 1985, published in Corrosion, Vol. 44, No. I1, November 1988. 3-5 D.A. Hale and C. G. Diehl, "Real Time Monitoring of Er,vironn.cntal Crack Growth in BWRs", paper 455 presented at Corrosion 88, St. Louis, MO, NACE, March 1988. 3-6 B.M. Gordon, Corrosion and Corrosion Control in BWRs, NEDE-30637, p. 6-22, December 1984. 3-7 R.B. Davis and M. E. Indig, "The Effect of Aqueous Impurities on the Stress Corrosion Cracking of Austenitic Stainless Steel in High Temperature Water," paper 128 presented at Corrosion 83, Anaheim, CA, NACE, April 1983. 3-8 P.L. Andresen, "A Mechanism 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, MA, NACE, March 1985 3-9 K.S. Brown and G. M. Gordon, " Effects of BWR Coolant Chemistry on the Propensity for IGSCC Initiation and Growth in Creviced Reactor Internals Components," paper presented at the Third Int. Symp of Environmental Degradation of Materials in Nuclear Power Systems-Water Reactors, Traverse City, MI, August 1987, published in proceedings of same, TMS-AIME, Warrendale, PA,1988. 3-10 F.P. Ford et al, " Prediction and Control of Stress Corrosion Cracking in the - Sensitized Stainless Steet/ Water System," paper 352 presented at Corrosion 85,
~
Boston, MA, NACE, March 1985. 26 4
P
* '
- GENE-333-141.It193, Rev.1
, , GENcclee Enery 3.3 References (cont'd) 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. t 3-12 P.L. Andresen, F.P. Ford, and A.M. Murphy, " State of Knowledge of Radiation EfTects on Environmental Cracking in Light Water Reactor Core Materials,"
Proceedings of the Fourth International Conference on Environmental Degradation of Materials in Nuclear Power Systems - Water Reactors, Jekyl Island, GA, August ; 1989, NACE,1990 t O 27
. . . .- .n - .- . ++ 2 .
i I , GENuclear Energy GENE-523141-1093, Rev. J
, }
r i Table 3-1 Severe Water Chemistry Transients in BWRS MAX "AX 10/07/93 CCND. ed. 01 PCW OATE OATA RANC PLANT uS/cm mm ;;:e gEV y.m d CCMMENTS REFERENCE POINT ! 1 AG 35 0 45 :00 7 780307 CONO OE=!N RE5!N 8LE!0THROUGH PC&RT 82L% 01 35 , 2 AG ae P BCC802 C040EN54TE DEM!N RE5 N INTRUSION EPRI NP 4134 46 3 At 84 0 3.2 ;4500 # 720901 C0h0ENSER LEAK. OEm!N 00EPLETED PC&RT 82LDA01 5 l 4 OZ 72 0 560 P 660820 POSRT 82LCA01 1 j 5 AG 70.0 4.6 198 P 771116 CRUD & CCNDENSATE DEMIN RES!N INTRUSION PC&RT 82 LOA 01 30
$4.0 38 p 740804 RE5th BEA0 INTRU510N
- NECE 13405 13 !
6N 7 AB 40.5 3.9 P 740608 A!R/A!R RESIN MIXTURE INJECTED INTO R FRCM RWCU PCMT 82LCA01 11 8N 33.0 4.0 P 740426 RESIN 8EAD INTRU5tch
- NECE 13405 10 .
30 0 P 710903 HIGH CONDUCTiv!TY WATER IN CST PC&RT 82LDA01 3 l 9 AC ' 10 02 28.5 P 661130 PC M T 82LCA01 2 11 8 25.6 41 50 P 770601 RE51N INTRus!DN PCMT 82LCA01 28 12 8 25.0 2500 0 810412 CONOENSER LEAK PMET 81-688-45 52 P 800205 POSSIBLE CONOENSATE DEMIN RES!N INTRUSICN EPRI NP 4134 80 13 PEACH 80TTOM 3 23 6 14 CZ 23.0 3000 P 790407_ LEAKAGE OF COOLING WATER INTO RPV VIA CORE SPRAY PC&RT 82LDA01 42 23 0 30 P 730406 AIR INJECTED INTO Rx FROM RWCU PC&RT 82LDA01 7 15 V 16 AG 22.0 P 771212 COND DEMIN RESIN INTRU5104 PCMT 82LCA01 32 17 K 21.0 46 2500 P 820428 TRICHLOR0 ETHANE FROM RA0WA5fE AND CST PCMT 82LCA01 53 20.0 4.5 o 821004 POSSISLE CONDENSATE DEMIN RESIN INTRU510N EPRI NP 4134 94 18 AG 17.0 P 801001 CONCEN5ER TUBE LEAKS PCMT 82LCA01 49 ? 19 F 14.0 P 750605 RWCU OUT OF SERVICE PCMT 82 LOA 01 20 , 20 C 21 T 13,8 4.7 100 P 781110 CRGAMIC INTRUSION VIA CONDENSA*E DECON DETER /CILS PCMT 82LCA01 39 P 740925 HIGH COND WATER PCMT 82 LOA 01 14 22 A8 13.5 ' 100 P 800428 UNKNOWN (LONG 5 HUT 00VN) EPRI NP 4134 81 23 A8 13.0 12.1 P 780225 RWCU RESIN INTRU5104 PC&RT 82LCA01 34 . 24 T P 750702 CON 0(NSER TUBE LEAK PC&RT 82LDA01 21 25 0 12.0 12.0 4.8 50 P 761025 RVCU RESIN TRAP RWCU INCPERABLE- PCMT 82LDA01 25 26 0 P 800812 ORGANIC INTRU5 ION EPRI NP 4134 82 27 T 11.8 4.8 60 P 750121 RWCU RE5!N INTRU$10N PCMT 82LDA01 17 28 0 11.5 , p 830106 POSSISLE CONDENSATE DEMIN RESIN INTRUSION EPRI NP 4134 95 29 A8 11.3 4.7 i PCMT 82LDA01 19 30 8 10.8 4.5 50 P 150601 RE5!N FROM FLUFFING CONDENSATE OF/0 PCMT 82LCA01 8 -; 31 AG 10.6 4.5 100 P 730507 RWCU RE5!N INTRU5104 PCMT 82 LOA 01 15 ! 32 0 10,0 P 741208 CONDEN5ER LEAK EPRI NP 4134 92 33 AG 10.0 p 820818 RWCU RES!N INTRU510N ' EPRI W 4134 74 34 H 9.2 1.4 57 P 780211 CONDENSATE DEMIN RE514 INTRU$10N EPRI hp 4134 63 5 35 6 8,2 4.3 50 P 7kOJ WASHOUT OF IMPURITIES FROM TlatBINE PMET 81 684 45 44 ~ > 36 8 8.0 5.0 500 P 790516 EPRI W 4134 88 7.5 p 810411 RWCU RE11N INTRUSION 37 8 89 5 38 C 7.1 100 P 811010 DECOMPOSITION OF RA0 WASTE RE51NS DUE TO HOT WATER EPRI NP 4134 ' EPRI W 4134 84 39 C 6.5 P 810210 CAUSTIC INTRUS10N VIA CONDENSATE STORAGE EPR! W 4134 57 40 K 6.2 4.8 20 P 741118 $USPECTED RE51N INTRUSION EPRI W A134'~ 54 41 0 5.8 4.5 50 P 730812 SUSPECTED RE51N INTRUS10N P 700123 RESIN INTRU1104 WEN C/D RET 1stNED TO SERV!CE JM5 QC 930717 98 t 42 AA 5.6 ' EPRI NP 4134 64 43 8 5.4 4.7 50 P 751218 PROSAall RWCU RE51N INTRUSION ' EPRI W 4134 85 44 8 5.1 p 810220 ORGANIC INTRU$10N VIA RADWASTE EPRI W 4134 70 5.1 68 P 770727 CONDENSATE DEN!N RE5!N INTRU5!0N, ANION RICH 45 N 67 EPRI W 4134
^ 46 0 5.1 4.8 50 P 760806 C0a0ENSATE DEMIN RESIN INTRU$10N 28
GENE-3D.!41 1093, Rev.1
, ggg g g ,y Table 3-1 Severe Water Chemistry Transients in BWRS MAX "AX 10/07/93 CONO. cH. C1 POV OATE cATA RANK PLANT uS/cm, min Dc0 LEV y-m-d CCMMENTS- REFERENCE POINT 47 C 5.0 '. 0 0 P 750309 POSSIBLE CONCENSATE CEMIN RE51N [NTRUS[CN gpgg gp gg34 79 48 0 4.9 49 50 P 760522 SUSPECTED RE51N INTRU5!0N [PRI NP 4134 66 49 T 4.5 50 50 P 781227 CRGANIC INTRU$1CN VIA CONDENSATE $75 FEN EPRI NP 4134 77 4.3 49 48 P 760221 SUSPECTED RESIN INTRU510N Epqg Np 4134 65 50 0 4.1 51 80 P 741015 RWCU RESIN INTRUSICN (pgg gp gg34 gg . '
51 K 3.3 5.4 50 P 770126 IMPROPER. RINSE OF CONCENSATE DEMIN EPRI NP 4134 68 52 AL 3.3 5.2 38 P 750309 POSSIBLE RESIN INTRUSION EPRI NP 4134 59 53 C p 810715 RWCU RESIN INTRua!ON Epgg Np ag34 gg 54 8 3.2 4.7 495 P 780131 RE5tN INTRUS104 EpRI Np 4t34 73 55 5 3.2 EpR1 mp 4134 61 56 H 3.0 5.6 96 P 750902 SUSPECTED RE51N INTRU$10N Epgl Mp 4134 62 57 8 2.9 5.4 50 P 751126 PROBABLE RWCU RES!N INTRU$10N 5.2 65 P 770912 !> PROPER RINSE OF CONDENSATE DEMIN EPR! NP 4134 71 52 T 2.8 P 750626 RESIN INTRUSION Epgg Np 4134 60 59 8 2.7 7.6 P 800824 CRGANIC INTRU$10N EpRJ Mp ag34 g3 60 T 2.3 EPRI NP 4134 78 61 T 2.2 5.5 50 P 790108 $USPECTED ORGANICS IN CONDEN$ ATE STORAGE 5.4 355 P 781208 CONCENSATE DEMIN RE$1N INTRU$10N EPRI NP 4134 76 62 Y 1.8 1.4 5.6 83 P 741125 VALVING ERROR DURING RE5!N TRANSFER EPRI NP 4134 $8 63 AC 1.4 8.1 38 P 780112 CONDENSATE DEMIN RE$1N INTRU$10N EPRI NP 4134 72 64 H 65 AJ 1.4 P 750906 $U$PECTED FLOC / FILTER A10/$URFACT FROM RAD WA5TE JMS OC 930717 99 5.6 30 P 770225 $U$PECTED RES!N INTRU$10N EPRI NP 4134 69 ' 66 0 1.1 8.8 72 P 180511 CON 0ENSATE DEMIN RE51N INTRUSION EPRI NP 4134 75 ~ 67 H 1.1 i P 811030 GTLCOL INTRU$104 V!A RADWA$TE EPRI NP 4134 90 68 V 1.0 EPRI Np 4134 87 69 PEACH BOTTOM 2 1.0 P 810622 OIL INTRU5104 INTO HOTVELL PCMT 82 LOA 01 4 ; 70 A$ 725 P 711113 HIGH FEEDWATER CONDUCTIVITY 54L P 180129 PCMT 82LDA01 33 11 8 1200 P 760708 CONDENSATE $75 TEM MOMENTARILY BTPA$$E0 ,, PCMT 82 LOA 01 24 72 8 PCMT 82LDA01 16 73 AR 600 P 150103 RVCU OUT OF SERVICE 3.5 87000 5 7904 . C001lM6 WATER INGRE$$ FROM RHR, RPV H20 TO HOTVELL PCMT 82LDA01 43 74 8 641.0 PCMT 82LDA01 12 75 8 423.0 3.2 $ 740401 ACIO INTO RPV FROM OENIN STORAGE TANK
$ 760519 PC MT 82LDA01 22 76 H 140.0 PCMT 82LDA01 26 77 T 45.9 3.8 244 5 761103 TORUS WATER PtMPED INTO RPV PRIOR TO STAATUP PCMT 82 LOA 01 23 78 8 13.3 1800 $ 760$20
$ 780401 LEAK in RMt HEAT EXCHANGER PCMT 82LDA01 - 36 {
79 A 13.0 l
$ 170917 RWCU OUT OF $ERVICE PCMT 82LDA01 29 80 8 12.9 PCMT 82LDA01 47 l 81 T 12.1 5 800815 PCMT 82LDA01 9 i 82 AL 11.6 5 130603 RWCU OUT OF $ERVICE '
5 800422 PCMT 82LDA01 48 83 8 11.2 PCMT 82LDA01 51 64 0 11.2 5 801219 EPRI NP 4134 91 85 F 10.5 $ 820427 POS$18LE ORGUIC INTRU$10N EPRI NP 4134 37 86 8 10.5 5.6 140 5 780923 R)CU RESIN INTRU510N PCMT 82 LOA 01 18 87 H 10.3 $ 750405 RWCU'0UT OF $ERVICE PCMT 82 LOA 01 5 88 AS 10.0 730 5 720604 DEPLETED RWCU DEMIN EPtl NP 4134 55 89 0 5.0 5.5 60 $ 740629 Con 0EN$ ATE OL'IN RESIN INTRUSION EPt! NP 4134 97 90 AX 4.5 5.2 220 $ 83050$ ORGANIC INTRU510N VIA RA0WA$TE 91 8 4.2 5.3 600 $ 781110 ORGANIC INTtuSION VIA CONDENSATE. DECON DETER 93 /DILS PC EPNI NP 4134 92 AP 1.0 $ 820900 GTLCOL INTRUSION VIA RA0 WASTE 29 l l
i i GE Nucl40! EMrgy GENE.3831411993, Rev. ] Table 31 Severe Water Chemistry Transients in BWRS
%R MAX 1 COND. D r' . C1 #CW OATE 4ANK 8 L A Ni- us/cm min sco LEV y-m-d COMMENTS ggggggggg pg gy 5 800305 PC&RT 82 LOA 01 45 .
34 I 1300 5 790329 CCNDEN5ER LEAK. CONCENSATE BYPASSED. RWCu OUT PC&RT 82 LOA 01 41 00 $ 801017 PC&RT 82 LOA 01 50 16 AC 683 5 770309 PC&RT 82 LOA 01 27 31 K $ 830213 GYLCOL INTO RA0VASTE. CETECTED PRIOR 70 COND STOR EPRI NP 4134 96 98 T 1200 $ 790316 CONCENSER LEAK. CONDENSATE CEPLETED. Cl INTO CST PC&AT 82LDA01 40 99 3 800 5 771206 RWCU OUT OF SERVICE PC&RT 82LCA01 31 NOTE: BW95 RANrE0 IN THE FOLLOWING CRDER:
- 1. POWER (P) CR SHUTOCWN (5)
- 2. CCNDUCT!v!TY OTHER NOTES:
- aESIN BEADS PRoviUE i,0*G TERM LOW pH t
P 6 30
. Arithmetic Mean Conductivity a Peach Bottom 3 l P Conductivity, pS/cm 1.4 - l 1.2 - 1
' Peach Bottom 3 all years - 0.536 g
0.8 - O.6 -
-\
O.4 w e N 0.2 - O .
&-A y
' ' ' ' t 0 90 92 94 .E 76 78 80 82 84 86 88
'74 E
Fuel year ~- W Figure 3-1.. Peach Bottom 3 Arithmetic Mean Conductivity
- 4 Effect of Concentration and Conductivity E
on IGSCC Initiation - FS Type 304 0 5
=
p , Acceleration Factor $ 5 O- ANL Data V EPRI Data 4_ (> O d 3 -
~
+ V. 2 - o Conductivity (pS/cm) I
. . g. g y
a* i I I' l I I i t i i i iiiii e i = i i i i i i ii i i i - i i iii-1Gs i i i i i i i si - 1000 - 10000 t 10 0 1 10 Sulfate (ppb) f{ Crack initiation data based on CERT FS 'Iype 304 Figure 3-2_. Effect:of Concentration and Conductivity on IGSCC Initiation
*- e _ -e. e_ . - _ - --*e *4--..mmem,=w-.r w=vr w , =m-m-e a w- + m.= =-e e e- w e+- -.. &-wn *-p-r- - - - + = *- ---. - 2 ---r%---r - - -4 e ,e#+--.wew--<- wa . . . + + +,2r---=e- . .-
~
Peach Bottom 3 Response to June 1986 Water Chemistry Transient @ t Crack length, mm ? 6 19.16 Resin Resin Intrusion 1 Intrusion 2 .- 19.14 - ~
. .g. .
19.12
.[ da/dt = 1.09 mm/y
.y..- .
.y .. 19.1 - -
,j, , , ,,
. , ' , da/dt = 0.55 mm/y 19.08 -
c D 6 19.06 2100 2200 2300 2400 2500 2600 2700 2800 E a Hours 2
~
Figure 3-3 Peach Bottom 3 Response to June 1986 Water Chemistry Transient
MM
!N- SCL1 CONST A N T LO AD - - - - - --- - - - - - - - ~ ~ - - - - - - - - - - - - - - - - -- - - - - - - - - - - - '
20.8 l - 0 82 , -
! 1.0 pS/CM ,
Na2SO4 h,
! <0.t pS/cm >
<0.1 pS/cm 045 pS/cm Na2SO4 0 5 pS/cm z W
;20.5 - 0 81 -- ,
9 *
~
g t
- a a z m=
- z <
- < sr $
ue
. s p-
. 0.80 -- o 20.3 -
4 2 2
. --- 20 PPE OXYGEN ---
# --- -200 PPS OXYGEN --- --
20.1 - 0.79 p-- -- --- -200 PP8 OXYGEN -. 19.8 - 0.78 -- I , s
*s :
i A . 13.6 - 0.77 ~-
~
19.3 - 0.76 2- p # 1.5 pM/h O.1-0 03* pM/h O 09 pM/h 1.1 MM/h 180* pin./h)
; (46 pin./h) 14.1-i.2' annJhl 13.6 pinJhl 0
y vi 19.0 - 0.75 , e
- ' 5-
- g I l ' ' ' ' I ' - ' '
?
2500 f
- i i I '
2000 te a w 0 74 1000 ~1500 500 M O Q HOURS *NO COMPLI ANCE UNLOADING % Figure 3-4 Crack Length Versus Time, Sensitized' Type-304 Stainless ~ Steel 286*C Oxygenated ' Water Environment, K = 28.6-30.5 MPalhi(26-28 ksi 4IN) _.-__ _ ______ _ _ _ _ _ . . _ _ . _ _ . - _ ~, .. .._ -.__ __ ._ _._ _ _ _ _ _ _ . . _ _ _ . _ . _ . _ . . _ . . . _ _ . . . _ _ _ . _. _
a GENE PLEDGE Model Prediction for PB-3 Sensitized Type 304 Crack Growth Rate [D . Crack Growth Rate, in/h 1975-1984 00 mV 4 000E-04 FOF -
.000E-05
' 100 mV O
1.000E 1992-1993 [ '
-200 mV
/
~ -300 mV 1.000E-06 -
b N 1.000E-07 f 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 { Conductivity, pS/cm & P LEDGE: 15 C/cm2, 20ksi/in Figure 3-5 GENE Pledge Model Prediction for Peach Iloitom 3 Type 304 Crack Growtli
'
- GE Nucitar Energy GENE-523-141-1993, Rev. I 4.0 IN-VESSEL VISUAL INSPECTION L
This section summarizes the IVVI resnits of the Peach Bottom Unit-3 core shroud. IVVI of welds ill through H8 were performed during this outage. The IVVIincluded both j inside surface and outside surface examination. Figures 1-2 shows the indications associated with the 113 and H4 welds. Only a few short indications were observed on the outside surface of H1 and H4. Circumferential indications were observed on the inside surface associated with vertical 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 119 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 3606 The indications observed at the H4 inside surface HAZ were a mixture of circumferential j 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 IVVI venical scans found no indications. , On the outside surface a limited number of short indications were observed associated l with H1 and H4. Table 4-1 is the IVV' plan which indicates the original planned inspections. Due to the observation ofindications at H3 and H4, the inspection scope was expanded and is ako described in Table 4-1. A summary of the IVVI results is shown in Table 4-2. See a mre ' l-2,1-3 and 1-4 for further details. - 36 t P
+
. GE Nmcle.sr Energy GDT 523141-1093, Re 1 ,
Table 4-1 P Peach Bottom Unit 3 Core Shroud Enam Plan (3R09) ORIGINAL PLAN (Prior to issuance of SIL 572)
- 1. Perform sample examination "ID" at (8) cell locations of the "H3" and "H4" welds.
1
- 2. Perform sample examination "OD" at (8) locations in the high flux areas at welds "H1, "H2", and "H5" EXPANDED PLAN (Following identification ofindications on the "H3" and "114")
- 1. Perform 100% examination of the "H3" and "H4" welds from the ID.
- 2. Perform 100% examination of accessible areas of the "H4" weld from the
- OD.
- 3. Perform examinations of the "H3" weld, "OD", where cracks were not identified from the "ID"
- 4. Perform an examination of the "H3" weld "OD", including significant corresponding areas of cracking identified on "ID".
S. Perform a sample examination on the "OD", at (8) locations,of the "H6" weld.
- 6. Perform a sample examination of the "OD" at (2) locations of the "H7" and "H8" welds t
- 7. Perform an examination of(l) vertical weld between the "H3" and "H4" weld.
- 8. Perform a sample examination of the plate to include:
(1) 8" area at the vertical weld. (1) 8" area between "H3" and "H4" welds. ' (1) 2" area between "H3" and "H4" welds. ' NOTE: Consideration was given to high neutron fiux, stress, and repair areas for selection of the sample locations 37
p 1]
, *, GENuclear Energ_ GEhE-333-141-1993, Rev.1 i
Table 4-2 ; Summary ofIVVIIndications Weld Inside Surface Indications Outside Surface Indications .; Hl. N/A 1 shon vertical : H2 N/A None H3 Circumferentialin Shroud None Cylinder HAZ H4 Circumferential and Axial 2 short vertical H5 N/A None H6 N/A None H7 N/A None H8 N/A None ,. V3 8 short circumferential N/A PLATE None N/A i k f 9 l
.t 5
l 38 , i
. GENewlear Enny _GEVE-333-141-1093, Rev. I 5.0 FLAW EVALUATION This section provides the flaw evaluation and application of the screening criteria to the Peach Bottom Unit-3 indications. Included in this section is the structural analysis, allowable flaw size determination, and screening criteria.
5.1 Structural Analysis This section describes the details and the results of the structural analysis performed to 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. 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 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 stresse.s from the applied loads. The stresses are essentially based on the strength of materials formulas. Since the bending stress due to seismic shear force varies with the elevation of a location, two conservative values 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. 39
, . GFNuclear Energy GEVE-583-141 1093, En.1 Table 5-1 shows the calculated seismic stress magnitudes for both the upset (Design i Earthquake - DE) and faulted conditions (Maximum Credible Earthquake - MCE). The 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 , H2 1438.6 0.23 0.1 I H3 1479.1 0.27 0.13 H4 2995.8 0.54 0.24 H5 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 Table 5-2 Pressure Differences Pressure DifTerences (psi) Component Faulted Condition Upset Condition Shroud Head and 32.9 14.12 Upper Shroud ; Core Plate Support Ring 54.8 35.68 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. 40 i
O
*
- GENuclear Energy GENE-323141-1093, Rev.1 l 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 overseas plant having higher fluence (8x10 20 n/cm2) showed stable crack extension and 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 K;c corresponding to the material JI c. The Kjc for the overseas plant shroud was approximately 150 ksiVin. Use of this equivalence is conservative since:
i) The calculated fluence for Peach Bottom Unit-3 is lower than that for the overseas plant from which J-R curves were obtained. ii) The J-R curves show Jmax values well above the JI c, confirming that there is load 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 ksiVin using the ASME Code safety factor of 42. For the analysis presented here, the LEFM analysis is confmed 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 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 ASME 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 sep ted 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 afTected 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 = oV(na) 41
- GENaclear Energy GENE-583-141 1093, Rev. I 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 same 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(Ll+L2), then they are considered as two separate flaws. Hov.ever, 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 3S m . The Sm value for the shroud material (Type 304 stainless steel) is 16.9 ksi at the normal operating temperature of 550 F. Safety factors similar to that used in the ASME Code (2.8 for normal and upset and 1.4 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 results are applicable for all welds since limiting values are used. , F e l I 42 l l
. GENselear Energy GENE-323141-1093, Rev.1 l 5.2 Allowable Through-Wall Flaws Allowable through-walj 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). IIowever, 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 t1 e 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 wclds involve Alloy 600 which has higher Smvalues 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]c value of150 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: KI= Gm *o*V(xa) 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 = 344 inches. 43
+ -
l l ^
* . GCNuclear Faergy GFNE-533-141-1993, Ren 1 Limit Load Analysis 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 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. 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 govemed 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 , is 1.85 ksi and the allowable K = 150/3 (where 150 ksiVin represents a conservative estimate of the material toughness and 3 is the safety factor). , The allowable flaw length was conservatively determined to be 2a = 59 in. 44
GENaclear Energy GEVE.533141 1093, Rev.1 Limit 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) where Mi is a curvature correction factor (which is a function of the flaw length (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.
b 4 O 45
=
GEN: clear Energy GENE-523-1411093, Rev.1 5.3 Screening Critoria The determination of the allowable through-wall flaws has been described in S& tion 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 definition conservative. If the criteria are exceeded, the option of doing further detailed evaluation or performing additional NDE remains. The effective flaw lengths (Ll e g L2e m 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 detail in Appendix A. The allowable through-wall flaws were:
. Circumferential Flaws
- 344 in. using LEFM I
- 430 in. using limit load
. Axial Flaws
- 59 in. using LEFM
- 200 in. using limit load A conservative approach in developing the screening mle 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 flawls 344 in. 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 allowable 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 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 46 . l
I
* , GENuclear Energy GENE-323141-1093, Rev.1 107.5 in. A similar restriction based on limit loads is not needed for axial flaws since they l are associated only with circumferential welds and are unlikely to be aligned in the same plane.
The removal or ieduction of the factor of 4 is justified if the IVVI results indicate that the remaining ligament is spread around the circumference of the shroud circumference. Evaluation of the IVVI results indicate that this is the case for the Peach Bottom Unit 3 k shroud (See Figure 1-2 and 1-3). l 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 (Llefr+ L2 efr), the length L=L1efr+ L2 rctis used for comparison with the LEFM based allowable flaw length. W f 4 I I r a l 1 1 47 1 j
. GENuclear Energy GENE-583-1411093, Ren 1 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 proximity rules 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 efpFL2efr). If the longest resulting flaw is less than 59 inches, then the screening limit is met for axial flaws. For circumferential f'aws, 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 (L1efr+L2 r), cf then the length L = L1 ef r+ L2 gshould e 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. L e O 48
i
. . GENaclear Energy .
GENE 323-141-1093, Ren 1 5.5 Application of Screening Criteria The screening criteria was used to evaluate the indications found by IVVI. The structural integrity of the core shroud is assured if the screerdng 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-3. The effective indication lengths were determined for all of the circumfereniial 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 between indication planes in order to properly determine the effective length. i The resulting effective indication lengths were then compared against the allowable indication length. Since all effective indication lengths satisfy the screening criteria, the structural integrity of the Unit-3 shroud is assured for the next two year cycle with power rerate conditions. . 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 determining 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 f'aw is now governed by the LEFM method and not the limit load method. 49
1
-1
. GENuclear Energy GENE-533-H1-1093, Rev.1 l
l S.6 References l l 5-1. Rooke, D.P. and Cartwright, D.J., " Compendium of Stress Intensity Factors," The j Hillingdon Press (1976). l 5-2. Ranganath, S., Mehta, H.S. and Norris, D.M., " Structural Evaluation of Flaws in Power Plant Piping," ASME PVP Volume No. 94 (1984). e b 50
GENE 323141 1093, Rev, ! , GENuclear Enerv Shroud Head Flange .
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r GENE.52).141.jppy, p,, j GENuclear Energy . i f Ao A a- I l m r-- o/b F 0.0 1.0 2a 0.1 1.006
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t GENuclear Energy GENE-523141 1093, Rev.1 IW1 . Y Map All Flow Indications lI ox ules Ma Effective Flow Lengths Circumferentic Flows i lf ' Axial Flows No Sum of Flows in any 900 f sector <107.5"?/ lI
/ Yes No <0,75(Ll gg e + L2 e gg lf ,
l For Adjacent Flows? No 1 S<0.75(L1ett +L2ett for Adjacent flows? L= Llegg or Regg , Yes lf NO Yes L= Ll gg e +L2 gg e - L1 egg +L2 e gg <275" - - l i If lI i
> ' Further L < 59"? '
s Evoluotion / Yes Y ' Continued Operation m Justified , Figure 5-6 SCHEMATIC OF SCREENING CRITERIA 56
GENuclear Energy GENE 523-141-1093, Rev. I 6.0
SUMMARY
AND CONCLUSIONS An evaluation of the indications in the Peach Bottom Unit-3 core shroud has been performed to demonstrate that the structural integrity of the shroud is assured for the next , two year cycle. In addition, the report documents material, water chemistry and fluence information which provides additional insight to the shroud condition. , The primary focus of this report was to demonstrate that even with several conservatisms , in the evaluation, the structural integrity of the shroud is maintained during a limiting event. This was performed by developing a screening criteria, assuming throughwall indications, which can determine the acceptability of the flaws based solely on the IVV1 results. The assumption of through-wall indications removes any 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 mening criteria uses both linear elastic fracture mechanics (LEFM) and limit load
.oncepts te determine acceptable through-wall indica. tion lengths. The limiting flaw length based on ei her LEFM or limit load was used for tne screening criteria.
The screening criteria also uses 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 shroud is maintained for the next fuel cycle. All effective indication lengths were shown to be less than the allowable flaw size. b 57
GENc: clear Energy GENE 333-1411093, Rev.1 APPENDIX A DETERMINATION OF TIIE EFFECTIVE FLAW LENGTII The effective flaw lengths are based on ASME Code, Section XI proximity criteris as presented in Subarticle IWA-3300. The procedure addresses both circumferential and axial flaws. Indications are ensidered to be in the same plane if the perpendicular distance between 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.1 Proximity Rules The flaw combination methodology used here is similar to the ASME Code, Section XI proximity rules 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. In Figure A-1, two adjacent flaws L1 and L2 are separated by a ligament S. Crack growth 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 Appendix B for crack growth rate discussion). Therefore, combining the crack growth 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 thickness of 2.0 in., this bounding ligament is 5.6 in. Thus, if the ligament is less than 5.6 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 adding the projected tip growth to each end of the flaw. For this example, ef Ll r= LI + 1.6", and L2eff = L2 + 1.6" 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 A-1
. , GENaclear Energy GENE-583-141-1093, Rev.1 the effective flaw length for the circumferential flaw is Lefr= Ll+S40.8" (the bounding ligament far 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 of the effective flaws in the shroud can be made, and the effective flaw length can be used for subsequent fracture mechanics analysis. In order to demonstrate the proximity criteria, three examples are shown in Table A-1 and described below. Table A-1 Flaw Combinations Considered in Proximity Criteria Case Circumferential Flaw Axial Flaw A Yes No B Yes Yes C No Yes , A.I.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 l indications must be combined such that the efTective length is (See Figure A-2b): L efy= L1 + S + L2 + 1.6" where: L1 = length of first circumferentialindication L2 = length of second circumferential indication S = distance between two indications , A-2 l
t
, GENcclear Energy GENE-523 141-1993, Rev.1 If the distance between the two tips is greater than 5.6", the effective flaw lengths are (See Figure A-2c):
L1 egy= L1 + 1.6" L2 egy= L2 + 1.6" A.I.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 circumferentia! l flaw is considered. If the distance between the circumferentialindication tip and the axial indication is less than 4.8", then the effective circumferential flaw length is (See Figure A- - 3b): Leff = 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): Lefy= 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: i L1egy= L1 + 1.6" (for circumferential flaw) L2ety= L2 + 1.6" (for axial flaw) i j A.I.3 Case C: No Circumferential Flaw - Axial Flaw This case applies 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. l l 1 A-3 i j
i i
- GE'Nuslear Energy GENE-533-141-1991, Rev. I 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.
. Ths adication is indication 2.
. Apply proximity rules to the pair ofindications (indications I and 2). Combine the fnws if necessary (Ll+L2+S). Old indication 2 becomes new indication 1. ,
. Continue along pc szimuthal direction until the next indication is cacountered. Thi .es new indication 2.
. Apply proximity rules to new indications I and 2.
. Continue proximity rule evaluation until all indications along the subject weld or plane have been considered.
l I 1 g M l l A-4 i
~ GENuctrar Energy GENE 523141-1093, Res I t w >-
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i 1 1 8 i L1 re, i Combined l Flow i i i 1 V D1 i n 4 ? l L i i 1 5 D2 V 4 > A . I L2 no. 2 4, 1 1 I i I Y V l Figure A ASME Code Proximity Criteria A-5
GENE 5231411093, Ren i GE Nuclear Duergy FIQw9 MSumec Through-woll +
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** 'a N (2a) 4 L1-- > --d S ? :C - L25 Ocws usumed Through-woil S< S.6"
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I Figure A APPLICATION OF PROXIMITY PROCEDURE TO NElGHBORING AXIAL AND CIRCUMFERENTIAL FLAWS A-7 l
l GE Nuclear Encey GENE 523141-1093, Ren i Start at Theta =0 Move in + Theta Direction Y i=1 Y First now is Flow i I Y Next Flow is Flow i+1 lI Perform Effective Length Calculation lf Combine Raws if Necessary To Determine Effective Length i=1+1 If g Raw i+1 = Flow i if No / Last Flow? Yes
- Done Figure A PROCESS FOR DETERMINING EFFECTIVE CIRCUMFERENTIAL FLAW LENGTH A -8
- GENuclear Ecergy GENE-533-141-1093, Rev.1 APPENDIX B BASIS FOR TIIE CRACK GROWTH RATE The basis for the crack growth rate used in the screening criteria is provided in this section. The Peach Bottom Unit 3 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 laboratory and field measured values. This model was used to predict a Peach Bottom Unit-3 specific crack growth rate and a conservative value was then selected based on this value.
B.1 Slip-Dissolution Model Figure B-1 schematically shows the GE slip-dissolution film-mpture model (Reference B-1) for crack propagation. The crack propagation rate V tis defined 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 Potentiokinetic Reactivation, EPR) and the changes in the value of constant A as a function of neutron fluence (>1Mev) is given as the following: EPR = EPRo + 3.36x10-24 (fluence)l.17 (B-1) where, EPR is in units of C/cm2, fluence is in units of n/cm2and the calculated value of , EPR has an upper limit of 30. t B-1 i
, *.. GENuclear EnertY GENE-323-141-1093, Rev.1 The constant A is defmed as the following:
for fluence $ 1.4x1019 n/cm2: C = 4.1x10-14 (B-2a) for fluence > 1.4x10 19 n/cm2 but 5 3x1021 n/cm2:, (B-2b) C = 1.14x10-13 in(fluence)- 4.98x10-12 for fluence 51.4x1019 n/cm2: C = 4.1x10-14 (B-2c) The units of K to be used with the above expressions is MPaVm. B.2 Calculation of Parameters The parameters needed for the crack growth calculation are: stress state and stress intensity factor, effective EPR, water conductivity, and electro-chemical corrosion . potential (ECP). The stress state relevant to IGSCC growth rate is the steady state stress which consists of weld residual stress and the steady applied stress. Figure B-4 shows observed through-wall weld residual stress distributions 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 reachi:s a maximum , of approx. 25 ksiVin. The average value of K was estimated as 20 ksiVin and was used in ! I the crack growth rate calculations. The weld residual stress magnitude is expected to decrease as a result of relaxation j produced by irradiation-induced creep. Figure B-7 shows the stress relaxation behavior of Type 304 stainless steel due to irradiation at 550 F. Since most of the steady state stress l in the shroud comes from the weld residual stress,it was assumed that the K values shown l in Figure B-6 decrease in the same proportion as indicated by the stress relaxation !
. 1 l
behavior of Figure B-7. 1 l B-2 l l
k: F i GENuclear Energy GEVE-323-141-lM3, Rev. I j V ,
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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 l predicted increase in EPR value as a function of fluence is shown in Figure B-8. l The third parameter used in the GE predictive modelis the water canMivity. The reactor water conductivity at Peach Bottom Unit-3 has recently (1992-93) been good (approx. 0.1 pS/cm 2) compared to earlier operating period (1975-84) when the average 2 L conductivity was in excess of 0.7 S/cm . 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 oferack growth predictions by the GE model. l 1 The last parameter needed in the GE prediction modelis the ECP. Figure B-1I shows the - measured values of ECP at two locations in the core. Since Peach Bottom Unit-3 does not plan to use hydrogen injection, the ECP values at zero H2 injection are relevant in Figure B-11. It is seen that the ECP values at zero H2injection 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 discuaion in the preceding section, the crack growth rate calculations were conducted as a function ofiluence assuming the following values of parameters: Initial K = 20 ksiVin EPR<> = 15 C/cm2 Cond. = 0.1 S/cm2 ECP = 200 mV Figure B-12 shows the predicted crack growth rate as a function of fluence. It is seen that , l i the predicted crack growth rate initially increases with the fluence 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 4.5x10-5 n/hr at a fluence of 1x1020 n/cm2 The crack growth rate for the expected fluence at the end of the next operating cycle is B-3
- . - -u
L GENE-523 141-1093, Rev.1
*
- GENuclear EnerU approximately 2x10-5 n/hr. i Thus, a bounding value of 5x10-5 in/hr can be conservatively used in the structuralintegrity evaluations for the shroud.
This bounding crack growth rate is quite conservative as can be shown in Figure B-13 from NUREG-0313, Rev. 2. It is seen that the crack growth rate of 5x10-5 in/hr at 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 { I crack growth rate. B.4 Summary A crack growth rate calculation using the GE predictive model was conducted considering the steady state stress, EPR, conductivity and ECP values for the Peach Bottom Unit.3 shroud. The evaluation accounted for the efTects ofirradiation induced stress relaxation and the increase in efTective EPR. The evaluation showed that a bounding crack growth rate of 5x10-5 n/hr i may be conservatively used in the structural integrity evaluation of the j Peach Bottom Unit 3 shroud. i 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. , i h B-4
__ GENE-583-1411093, Reu 1 GE Nuclear Energ The GE PLEDGE Slip Dissolution - Film Rupture Model of Crack Propagation E CT m VT crack-t.ip advance by enhanced oxidation at strained crack tip e
.n VT = AE CT Where:
VT crack propagation rate A, n = constants, dependent on material and environmental conditions 6 CT = terms crack-tip st. rain rate, formulated in of stress, loading frequency, etc. Figure 0-1 B-5
3 SOLUTION RENEWAL
': RATE TO CRACK-TIP ,, ', N STRESS O
i
=
A ^ NOXiDE RUPTURE l
' ' TR PORT RATE AT
~e
# ENVIRON- CRACK-TIP
# MENT
( MICRO-j STRUCTURE M y-FIELD hI$ CRACK TIP $(A}_, pH RELAXATIOK
~
PASSIVATION RATE M - E AT CRACK-TIP N-FLUENCE GATION e d
- , 5
~ 2 E
~
Figure B-2 Effects of Fast Fluence, flux & Gamma Field on parameters affecting IGSCC of Type 304 Stainless Steel
*n g VT
= A6 #
ct I& i d ( M ' r et
- JA', JH2O g Ju +
g h- ( U i P S, Ni, Si NC h
- -n n3.6 g x ,g_iq g4 Vy = 7.8 x to-3 - .
f(EPR) _ g f(K) O E 4 5e te c,i , s 5 5 _w
- =
2 Figure B-4 Parameters of Fundamentalimportance to Slip Dissolution Mechanism of IGSCC in Sensitized Austenitic Stainless Steel
' I
. . _ GENE-523141 1093, Rev.1
)
GE Nuclear hergy I l OBSERVED RESIDUAL STRESS PM0 FILES I IN HAZ OF 24" 28" OI A. SCH. 80 PIPING
\
I I I I l l l i I
- +300
\
40 9 ~ Ug - +200 g d 20 - a g p/ +' d ( *
\ A[ =
E g \
' M ao
- k'e e OI 0 -A Nga g O . a
. a 8 -100 8
\0 gd / ,o -
g U / e p E -20 -
- - 200
. e-
-40 - - -300 I l I l I I I I .
1 -l O O.1 0.2 0.3 0.4 0.5' O.6 0.7 0.8 0.9 1.0 ' OUTSIDE-INSIDE FRACTION OF TM400GH-MLL - WALL WALL OlMENSION i f - Figure B-4 Throughwall longitudinal residual stress data adjacent to welds in 12 to 2 inch diameter stainless steci piping l O B-8 l l
GENuclear Energy GENE-523141 1093, Rey, j l 275.9 l CD 40 In 206.9 ; 30 - 137.9 20 - Total Stress i LM
- 68.9 y
.c 10 -
i
.$ f pplied A Load Stress ' O 3
' 2 2 o
. B.
2 05
- 68.9 j
- lo
- 137.E
- 20 -
- - 206.9
_ so -
' ' ' - 275.9
' 80.48 55.56
_4o _ 15.24 30.32 86.40 0 8.06 10.16 l DEPl"H (mm) Figure B-5 Shroud Total Throughwall Stress Profile B-9
GENuctw Energy GENE-523141 1093, Rev. I 26 27.26 24 - 22 - C 5 20 - - 21.81 t' 2: 7 18 - .
.a C' f) lo - _
16.36 E. 2 14 - G
.@ 12 - 4-
- s. ,
5 10 - - 10.91 3 E 2
~
8 E 2 ,.. o 3 6 E m - 6.45 t-4 - 3 -
' ' ' ' ' C 0 ao.s2 25.40 0 8.04 10.16 15.24 Crack Depth . A (mm)
Figure B-6 Shroud Throughwall Stress Intensity Factor B-10
e I ' GENE 323141-1093, Rev. I GE Nuclear EnerKY 1
. Stress Relaxation Behavior .
from Irradiation Creep. en 0.8 -
.5
.5, .
E c: 0.6 - ' m Average Data
$5 -
O 0.4 - e .
.O_ _
U S ' 0.2 Type 304 Stainless Steel ' at 288'C 0 3 gas _. 30 22 , 10 10ao Neutron Fluence, n/cm8 (E>1 MeV) Figure B 7 Stress Relaxation Behavior of Type 304 Stainless Steel Due To Irradiation at 28 B-11
I \ l n t". ' U l i!- R T 4 EPR VERSUS NEUTRON FLUENCE 30 f
,, i i __._
25 l si --- -- 20 ,, to I' 15 .
. -+
9
- D w
1E+21 0 i 10 1E+20 g 7 1E+19 FLUENCE g C I Figure B-8 EPR Versus Neutron Fluence _. - - - - - , a-~---- - u. - - - -
-- - , a . - -_ - - .--- . - - - _ - - - - - - - - - - -
1 a GENE PLEDGE Model Prediction for PB-3 Sensitized Type 304 Crack Growth Rate [ T Crack Growth Rate, in/h l 00.mv 1.000E-04 FOl 5.000E-05 _
- -100 mV
. 1992-1993 - - - - - -
1.000E- 05 -200 mV i f w 2
- /
7
~
-300 mV 1.000E- 06 :
O e p 4. 1- ' ' '
' 2- '- 2 '- f 1.000E- 07 0.6 0.7 0.8 0.9 .E 0 0.1 0.2 0.3 0.4 0.5 E
Conductivity, pS/cm ; PLEDGE: 15 C/cm2. 20ksi/in . Figure B-9 GENE PLEDGE Model Prediction for PB-3 Sensitized Type 304 Crack Growthb Rate
i m Effect of Conductivity on Sensitized 304 ~ Crack Growth Rate
!a IP Crack Growth Rate, in/h 1.000E ;
5.000E-05 _ 100 mV
~
1 1.000E-05 O on no l c oc o i
~~ AJ w
A AG ug 1.000E-06 goo o
- c
- ~
o M c g
- g Ao O h u
TMM M ' ' - E
*O 0.35 9 1.000E-07 = 0.25 0.3 3 0.1 0.15 0.2 s f 0.05 Conductivity. pS/cm :=
[ I l J PLEDGE: 20 ksi/in,15 C/cm2 CAV: 20-25 kai/in,13 C/cm2,100-160 mV Figure B-10 Effect of Conductivity on Sensitized Type 304 Crack Growth Rate
.N _
l g in-Core Bypass ECP vs Feedwater Hydrogen f for a BWR-4 4
?-
ECP, mV (SHE) , i 300 p 200 - U 100 - i O ~ Just Below Top Guide Level G
-100 -
v3
-200 -
\ 'ta s n N
h-t Just Above Core Plate Level
-300 -
S
- . -cy . .--a z
' ' t 2
-400 60 70 80 90 _E 20 30 40 50 0 10 Feedwater H2, SCFM f Figure B-11 In-Core Bypass ECP vs Feedwater Hydrogen for a BWR-4
_ _ _ _ - _ _ _ ._ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - _ _ _ _ _ ____--_____________-_a
d Figure B_12 g D k GROWTH RATE VERSUS FLUENCE l 1.00E-04 _ . . ._ _ ._ . . ._ __ _ _ _ _ . _~. _ _ _ . . _ _ _ . . _ _q 3 _.
* -l
.g_ _ . _ _. ._! l'
! O , o ; o
=
.'c_
v 5 E 1.OOE-05 ~ ' ' ' ' ~
..:_ ._.L.
'^'~: -.
h p . _ . _ _ _ m __. _ _ v 1.OOE-06 1E + 21 IE+20 1E+19 E Fluence (n/cm'2) $ E Stress Intensity = 20 Ksi/in, initial EPR = 15 C/cm2
- 6 GENuclear Energy GME-3231411093, Ren i
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10 - l l - NRC CURVE 4 - 1inlyr /, 10-4 - o - A f G -
/'
~
g ~
~
- O. Y 0.2 pena 02; sensttired at 1150'F/2 h 2
td (EPA
- 15 C/cm ) , gg ,
Q
~
A 0.2 ppm 0 ; senstttred at 150'F/2 h - 2 CZ" 2 (EPR
- 10 C/cm ) , gg z ~
O 0.2 90 0,; seasitized at tis 0'r/24 h H . C(
~
3 Q 0.2 ppa 0 ; severely sensittred 2 E 10-6 o 4 , ,0,; un io n. n n 50.,/24 h _ (S G( MITACHI _ GENIO AML M ~ CECAD - N HO H 8 ppe 0 3 2 "'II'd 'T "IdI"8I Cr O 0.04 inlyr u l .i i r r/24 h (5 i
@ 8 ppe 0 ; senstttred at g)1292*f/10 min)'
2 I
,,,,, + (932'F/24 h) (EPR = 4 C/cm )
O 8 ppe 0 ;2sensittred at (1292'F/10 min) ~ l 2
* (932'f/24 h) (EPR
- 4 C/cm )
l' 0~'7 D 8 ppe 0; sen'st ttred't 2 a (1212*f/10 min)
* (842*F/257 h) ((PR
- 15 C/caI )
l @ 4 pos 0 23 8'"Si"d " III'I*'/IO *i"I
+ (842*F/257 h) (CPR = 15 C/cm )
I $3T gg f = 0.1 Mt. t
- 0.14 0ATA l 9 8 pp. 0,; sensittred et itst'r/14 h I
I (EPRe 20 C/ce ) f + 0.006 Mt. A = 0.15 l X 8 ppe 0 ;2sensttised at 1292'F/14 h (IPE = 20 C/enI ) f
- 0.04 Ma, I R
- 0.15 s 1
0 10 20 30 40 50 60 70 STRESS INTENSITY,K (ksid.) . Figure B-13 Nureg 0313 Crack Growth Rate Data B-17
. - - - - - . . .-. . .}}