ML20095J189

Design & Analysis of Weld Overlay Repair for Sequoyah Unit 1 CRDM Lower Canopy Seal Welds

ML20095J189
Person / Time
Site:	Sequoyah
Issue date:	10/05/1995
From:	Gustin H STRUCTURAL INTEGRITY ASSOCIATES, INC.
To:
Shared Package
ML20095J187	List: ML20095J185 ML20095J189
References
SIR-95-105, SIR-95-105-R, SIR-95-105-R00, NUDOCS 9512260365
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{{#Wiki_filter:'. B88 951025 8 0.0 QAR8 Cord Report No. SIR-95-105 RevisionNo. O APPR0VEO c $.*Jf,'Ud *' ;*YY'ni.*,t Project No. WSI-20Q .n, lZfoM*.' A* l%.'n*0.%." d ProjectFileNo. WSI-20Q-401 October 1995 Letter No. 30M246 Cat.: October 23 1005 TENNESSEE VALLEY AurnomTY SOEP (N) BY u.J.surzynsu 1 a .5 DESIGN AND ANALYSIS OF A 5 WELD OVERLAYREPAIRFOR THE .5 SEQUOYAHUNIT 1 CRDM 5 LOWER CANOPY SEAL WELDS o 3 8 8* a ,, 3 a 6.$ $ 6 Prepared by: q$ u g g 2 na StructuralIntegrity Associates,Inc. 3 6 E bs 5-l hd Prepared for: N !E h E AAGA Tennessee Valley Authority Date: /0 9f Prepared by: J H. L. Gustin, P.E. I Reviewed by: Date: N.G. Cofie Approved by: Date: M if H. L. Gustin, P.E. RIMS, WT 3B-K 951226d365 951219 PDR ADOCK 05000327 Structurallategrity Associates, Inc. P .PDR

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i I 4 l i Table of Contents 1 I Section P_ggg 3 -42 1.0 IN1'RODUCTION AND

SUMMARY

1-1. 1.1 B ackground................................................. 1-1 1.2 S u mm ary................................................... 1 -3 7 2.0 WELD OVERLAY DESIGN.......................................... 2-1 j 2.1 Dimensions.................................................. 2-1 2.2 Weld Procedure and Material.................................... 2-3 2.3 Code Reconciliation................................... '........ 2-4 ~ 3.0 WELD RESIDUAL STRESS AND STRESS INTENSITY FACTOR ANALYSIS. 3-1 3.1 Method ology................................................ 3-1 3.2 Residua 1 Stress Results......................................... 3-1 3.3 Stress Intensity Factor......................................... 3 - 4.0 CRACK GROWTH RATE EVALUATIONS............................. 4-1 4.1 Crack Growth Law............................................ 4-1 l 4.2 Remaining Life Estimate........................................ 4-3 4.3 Fatigue Crack Growth......................................... 4-4 5.0 C ONCLUSIONS....................................'............... 5-1 i

6.0 REFERENCES

6-1 APPENDIX A Detailed Description ofWeld Residual Stress and Stress Intensity Factor Analyses for the Sequoyah Lower CRDM Canopy SealWeld f Overlay Rep air.......................................... A-1 ~~ I I Ia o JL SIR-95-105, Rev. O i Structural lategrity Associates, Inc.

4 5 0-j' t. = 1 List ofTables i Table g 2-1 Structural Reinforcement Sizing Evaluation............................... 2-6 A-1 Summary of Thermal and Mechanical Properties ' Used in the WELD 3 Weld Overlay Simulations............................ A-18 4

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1 1 J l 3 1 .J .ha 6 SIR-95-105, Rev. O StructuralIntegrity Associates, Inc.

g,.. 1 List ofFigures Figure hgn l 1-1. Geometry of Spare CRDM Head Adapter Used for Zion Canopy Seal Weld ? Overlay Model Development.......................................... 1-4 3 1-2. Geometry of Spare CRDM Head Adapter Plug Used for Zion Canopy Seal Weld Overlay Model Development...................................... 1-5 2-1. Blustration ofNet Section Collapse Criterion (NSCC) for Circumferential Crack of Depth a and Length 20 in a Pipe................................ 2-7 2-2. Canopy Seal Weld Overlay Design and Dimensions Used in Analysis............ 2-8 t-3. Canopy Seal Weld Overlay Design...................................... 2-9 3 3-1. WELD 3 Axisymmetric Finite Element Grid Used for Overlay Weld Modeling with Blustration ofBoundary Conditions......................... :....... 3-3 3-2. Definition of Overlay Weld Model Segments for the Overlay Design Analysis..... 3-4 ~ 3-3. Stresses at Sections A, B, C, and D After the Last Layer of the 3-Layer Overlay Design and After Being Heated to 550*F.......................... 3-5 3-4. Meridional Stresses at Sections A, B, C, and D After Each Layer of the 3-Layer Overlay Design at 100*F....................................... 3-6 3-5. Comparison,of Stresses at Section D for the As-Welded Condition at 100'F and After Heating to S 5 0 *F........................................... 3-7 3-6. WELD 3 Axisymmetric Finite Element Grid Used for Stress Intensity Factor l Calculations with llustration ofBoundary................................. 3-8 3-7. Fracture Mechanics Crack Model A - Continuous Surface Crack in Half Space.... 3-9 3-8. Comparison ofWELD3 Stress Intensity Factor Calculations with those of p c-CRA CK...................................................... 3 - 10 4-1. NRC Crack Growth Rate Data and Curve for Type 304 Stainless Steel with Data and Estimated Curves for Fe-Cr-Ni Alloys........................ 4-5 A-1. Geometry of Spare CRDM Head Adapter Used for Canopy Seal Weld i' Overlay Model Development.......................................... A-19 A-2. Geometry of Spare CRDM Head Adapter Plug Used for Canopy Seal Weld. ,j Overlay Model Development.......................................... A-20 A-3. WELD 3 Axisymmetric Finite Element Grid Used for Overlay Weld Modeling with Blustration of Boundary Conditions......................... A-21 j 3 A-4. Definition of Overlay Weld Model Segments for the Overlay Design Analysis..... A-22 A-5. Isotherms for Simulation of a Weld Pass Near the Center of the Seal ,} and in the First Weld Layer....... s................................... A-23 J A-6. Isotherms for Simulation of a Weld Pass Near the Center of the Seal and in the Second Weld Layer......................................... A-24 1 A-7. Isothenns for Simulation of a Weld Pass Near the Center of the Seal and in the Third Weld Layer........................................... A-25 A-8. Isotherms for Simulation of the Weld Pass of the Last Weld Layer............ A-2 6 l A-9. Meridional Stresses at Sections A, B, C, and D After Each Layer of the 3-layer Overlay Design........................................... A-27 3 f StructuralIntegrityAssociates,Inc. SIR-95-105, Rev. O III a

i y,. t i List ofFigures (continued) Fiere P_ags A-10. Hoop Stresses at Sections A, B, C, and D After Each Layer of the 3-Layer ? O verl ay D e sign.................................................... A-2 8 I A-11. Stresses at Sections A, B, C, and D After the Last Layer of the 3-Layer Overlay Design and After Being Heated to 550*F.......................... A-29 A-12. Comparison of Stresses at Section D for the As-welded Condition at 100*F and After Heating to 5 50

• F........................................... A-3 0 A-13. WELD 3 Axisymmetric Finite Element Grid Used for Stress Intensity Factor j

Calculations with Illustration ofBoundary Conditions and the Crack Plane (S ection D)................................................... A-31 A-14. An Illustration of the Accuracy of the Finite Element (WELD 3), Energy-Based Method for Stress Intensity Factor Determinations.......................... A-32 A-15. Illustration of pc-CRACK Third Degree Polynomial Curve Fits for Meridional Stresses at Sections C and D of the Overlay Design................ A-33 A-16. Comparison of WELD 3 Stress Intensity Factor Calculations with Tho se of p c-CRA CK............................................... A-3 4 1 1 1 .] I ~1 .]

• "1 SIR-95-105, Rev. O iv structuralIntegrity Associates, Inc.

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1.0 INTRODUCTION

AND

SUMMARY

A weld overlay repair design was developed for appl cat on to potent a y ea ng contro rod drive i i i ll l ki l 4 mechanism (CRDM) lower canopy seal welds (CSWs) at the Sequoyah Nuclear Power Plant Unit 1. An evaluation of the predicted weld residual stress distributions resulting from this repair was performed. The weld residual stress and the applied stresses on the CSWs were used to determine r the design life of the overlay repaired CSWs, assuming the mechanism for repair degradation to be i stress corrosion cracking (SCC). This repair uses Alloy 625 fdler metal, applied by the GTAW process. Previous work performed by StructuralIntegrity Associates (CSW thld overlay repairs at other plants) was used as part of the basis for the overlay design for Sequoyah.

1.1 Background

Potentialleakage has been detected in three Sequoyah CRDM lower CSWs. The overlay design in this report was prepared to repair these potentially leaking canopy seals. 1 In the past, through-wall defects were detected in the CRDM CSWs, between the head adapter and head adapter plug, on spare CRDMs at Zion, Unit 1, and weld overlay repairs using Alloy 625 were designed and implemented. The geometry and dimensions of the CRDM lower canopy seal at Sequoyah [1] are essentially identical to those at the Zion Nuclear Power Plant [2], for the purpose of this analysis, and are shown in Figures 1-1 and 1-2. Failure analysis results on the Zion-1 CSWs, concluded that transgranular stress corrosion cracking -j (TGSCC) was the mode of failure, resulting in local through-wall cracks and leakage, similar to that 7 at Diablo Canyon [2,3]. TGSCC can be caused by local contamination with impurities such as chloride ions or caustic [2,3,4], although such impurities were not conclusively identified at the time J of this report. 1 ~ SIR-95-105, Rev. 0 1-1 g,,,ggy,,,,gg,g,,gy,gggg,gggg,,gg,

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i(, In spare CRDMs at Zion, a major factor in promoting the cracking and leakage appeared to have been the lack of weld penetration and poor quality of these CSWs. Such crevices can exacerbate off-chemistry conditions and lead to local concentrations ofimpurity ions which result in SCC initiation. a lo It is believed that SCC led to local through-wall cracking and leakage, with the time of SCC initiation not being certain. At least one of the Sequoyah CRDM lower CSWs in question was on an active p l CRDM. It is expected that SCC has been present in the active CRDMs as in the spare CRDMs at y Zion. The weld overlay repair method for the lower CSWs at Sequoyah is discussed in this report. In particular, these repairs were designed for leaking CSWs observed at Sequoyah during the Fall 1995 outage. The weld overlay repair concept [5] has been employed many times for the repair of intergranular stress corrosion cracks (IGSCC) in boiling water reactor (BWR) stainless steel piping. j welds. It was also used for CRDM canopy seal repairs at the Zion, Diablo Canyon, Callaway, and Prairie Island pressurized water reactor (PWR) plants among others. Weld overlay repair concepts f similar to those used at these plants are used to design CRDM canopy seal repairs for Sequoyah. Details of the overlay design are in accordance with ASME Section XIIWB-3642 [6] and References

w 5 and 7. ASME Code CaseN-504-1 provides additional guidance. The overlay design is based upon 2

assumed stresses equal to ASME Code allowable stress limits. Details of the design considerations i are presented in the following section of this report. The service life of the Sequoyah repairs is predicted to be in excess of 40 effective full power years. ..I

.J The failure analysis of the CSWs from Zion-1 concluded that SCC was the mode of cracking and this i

mode of cracking is expected to be operative at Sequoyah. Therefore, the SCC resistance of the f repair was evaluated to predict the remaining 1ife. In order to perform this evaluation, including a 4 - fracture mechanics crack growth analysis, the weld residual stress distributions from the repair were

• 1 estimated by using the WELD 3 computer program [8], an elastic-plastic, thermo-mechanical Snite element program. This residual stress distribution is much more significant than the sustained loads l

[2,3] in providing the driving force for SCC. The residual stress is used as a principal input to the ' fracture mechanics SCC crack growth law to predict the remaining life of the repair. 4 ~ STR-95-105, Rev. 0 1-2 g,,,ggy,,,,,,,,,,,,,gg0gygggg, jag,

e 1 ? l.2 Summary A weld overlay repair was designed for the Sequoyah CRDM lower CSWs. This design is based on meeting the requirements of ASME Section XI, IWB-3640 [6] and the NRC requirements outlined 7 3 inNUREG-0313, Rev. 2 [5] for the repair of SCC flaws. Based on an assumed maximum stress of 1.0 S,(16.2 ksi), for Type 304 at 650*F for primary membrane stresses in accordance with ASME ] Section III, NB-3227-7 [6], the minimum required overlay thickness is 0.0511 in. The width of the ] overlay is such that it should be blended into the thicker sections of the head adapter and CRDM, pressure housing, as shown in this report. In order to provide SCC resistance, the repair weld } material was selected as Alloy 625 filler, using the GTAW weld process [9]. ll Weld residual stresses were computed with the WELD 3 computer program [8] for the repair model. The model consisted of depositing three layers of Alloy 625 weld metal, each layer being 0.12 inch thick and welded starting from the top. This model was confumed on a weld mock-up performed for the Zion spare CRDM CSW overlay [2]. A fracture mechanics SCC crack growth law, a factor of ten slower than the upper bound NRC curve forIGSCC of 304 [5] was used for the remaining life predictions. This law is based on the superior SCC resistance of Alloy 625 (similar to I-82 filler, which is judged to be SCC resistant [5]), and a literature survey (discussed in this report). This crack growth law, combined with the 3-layer overlay residual stress distribution, and a bounding 2 ksi applied membrane stress, gave a predicted remaining ld SCC life in excess of 50 years. l i em 51R-95-105, Rev. 0 1-3 gg,gggy,gg pgggg,,gy pgggg;gggg, fac_

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A* i e <y 3 I 5 g \\ l 3 I 3 1 i o l Jg y I V ACOT ctA. w OETAIL A MATCMINC Aeut. v UiatAo - st: shin 2 I !!T. OCxt. *3* c w f / ~ .ne h$T ' } [ . --]f a oIA. / M ca. // ../_ } ~ 1 l 8 j / r / /l u //. l ./ DETAll 8 e 9 A SECTION OF AD A0 6 OIMENSiONAL DATA .] A =r 5.66* I F = 2.504" iK = 1.370" lP = 1.0" lU = 0.160" B = 4.0" l G = 6.0" l L = 5.701" l Q = 6.45" lV = 4 743'i.n j C = 2.75" i H = 5.177" iM = 4.000" lR = 5.5 5" lW = 30* 3 D= 17.75 to 59.251 I = 3.875" IN = 45' iS = 0.075" lX = 0.25" E = 11.0" lJ = 0.78" 1 0 = 30* IT = 0.160" I Y = 0.50" 1 T, Figure 1-1. Geometry of Spare CRDM Head Adapter Used for Zion Canopy Seal Weld Overlay ModelDevelopment SIR-95-105, Rev. 0 14 f StructuralIntegrityAssociates Inc. 4 l

o n j.,, t /> / c n 6 //////A i / wuca cw lacg og e V////1 t t l y t l, 2 l ENLARGED ME% o I a.cco-s Acut-2c .i ( E/f, J I l r i F l t. ,,,,,_A WATI?JA!.: Asut sAts: Tr/t sc4 SECTION OF HEAD ADAFTER PLUG m DIMEN5tCNAL DATA A = 7.59" iH = 1.25' IO = 6.995" S = 5.125" lI = 5.71" lP = 5.05" C = 3.00" lJ = 15' IO = 5.50' 0 = 1.30' IK = 0.075" iR = 6.02" .) - E = 0.485' lL = 6.122" l 3 F = 0.40" lM = 6.440" l-0 = 0.40" lN = 8.00" I l f i Figure 1-2. Geometry of Spare CRDM Head Adapter Plug Used for Zion Canopy Seal Weld Overlay Model Development J i ' SIR-95-105, Rev. O l-5 f StructuralIntegrityAssociates,Inc.

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j i;.. g-2.0 WELD OVERLAY DESIGN j Details of the CSW overlay design are in accordance with ASME Section XIIWB-3640 [6] and References 5 and 7. Since the overlay weld process is GTAW, a relatively tough and ductile repair B results, and secondary stresses (such as thermal and residual stresses) are not required [6] to be considered in the design of overlay thickness. Thus, the design is conservatively based upon assumed 1 primary stresses equal to the ASME Sectbn III Code allowable values. u 7 l The overlay thickness required for stmetural reinforcement to satisfy the preceding mies of ASME Section XI [6] and the NRC requirements [5] is significantly less than that needed for a practical remaining life, based on residual stress and SCC crack growth predictions. Because the failure analysis of the Zion-1 CSWs resulted in the conclusion that stress corrosion cracking was the mode of failure [2], the methods outlined in NUREG-0313, Revision 2 [5] were considered in designing and evaluating the remaining life of the CSW repairs in this report, in order to appropriately consider resistance to stress corrosion cracking. The present analysis conservatively 1 assumes that a crack is already present in the original CSW, and considers crack growth by the stress corrosion mechanism per NUREG-0313, Revision 2 [5]. Dimensional and weld material requirements are described below. ij 2.1 Dimensions

- j The weld overlay repair thickness was derived in accordance with References 5 and 7, in order to reduce both the nominal stresses and the ratio of flaw through-wall dimension to wall thickness (a/t) i to meet the requirements of ASME Section,XI, IWB-3640 [6]. This repair is sufficient for the h

maximum possible postulated defect (360* by through-wall) in the original canopy seal weld, thus ] meeting the requirements of a standard weld overlay [5]. Although the guidelines ofReference 5 are, strictly speaking, applicable to BWR stainless steel pipe welds, the subject CRDM repair is close j enough in material and design basis, such that the weld overlay guidelines are considered to be j directly applicable. l'r I SIR-95-105, Rev. 0 2-1 Structural Integrity Associates, Inc. LJ

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} The methodology ofReference 7, the net section collapse criterion (NSCC), serves as a basis for flaw acceptance and repair in austenitic stainless steels, to satisfy ASME Section XIIWB-3640 [6] and 4 j NRC requirements in NUREG-0313, Rev. 2 [5]. A brief description of the NSCC methodology follows. u The NSCC, as illustrated in Figure 2-1, assumes that the pipe with a crack of depth "a" and a length p of "20" (mcluded angle) will fail by plastic collapse when the net section reaches the flow stress of. n I-This criterion considers the fact that when a circumferentially cracked pipe is placed under an axial 1' load, due to the primary membrane and bending stresses P, and P, a shift in the neutral axis of the l b )] pipe will occur by the angle p. iJ The equations to predict the acceptable flaw size as a function of the applied stresses P, and P are b givenin Reference 7 as: Case 1: 0+psx P = 2a [2Sinp - (a/t)*(Sine)] / n 3 f = [(n-Ba/t) - (P,,/o)n] / 2 f Case 2: 0 + p > x, then P = 2c)(2-a/t)Sinp] / n 3 p = n[1-a/t-P,,/o] / [2-a/t] f ~1 The flow stress, o, for use in the above equations is estimated as three times the S, value (16.2 ksi ,j r at 650*F) from Section III of the ASME Code for Type 304, the material to be repaired. i e SIR-95-105, Rev. 0 2-2 StructuralIntegrity Associates, Inc. J

3.. Y ' The weld overlay repair thickness as shown on the attached design sketch (Figures 2-2 and 2-3) was computed with the Structural Integrity Associates computer program pc-CRACK [11] (based on the above NSCC methodology) for an applied stress equal to the ASME Code allowable stress level. b This allowable stress level is provided in ASME Section III, NB-3227.7 [6] wher'ein the CSW shall be designed to meet the pressure induced general primary stress intensity limit ofP, = S, as given in Figure NB-3221-1 of ASME Section III. The value of S,is taken as 16.2 ksi at 650*F, from l Appendix I of ASME Section III Appendices (6]. .I A safety factor of 3 is placed on the above stress ratio, for design of the overlay, in accordance with IWB-3640 [6] and Reference 5. The overlay design thickness resulting from the above stress ratio 'is 0.0511 in., as shown in the attached pc-CRACK output (Table 2-1). The number ofweld passes to meet these dimensions should be determined by the procedure qualification and the dimensions should be verified on the actual welds. A final overlay thickness of 0.36 inch is specified to account for stress corrosion cracking mechanisms. The minimum width of the 360' circumferential weld overlay repair is also illustrated in the attached l design sketch (Figure 2-2). When overlay thickness requirements are computed, as above, an iterative computation is performed to both decrease the ratio of flaw depth to wall thickness (a/t) and decrease the nominal stress in the wall as the overlay thickness is increased. The assumed decrease in nominal stress in proportion to the increased wall thickness is valid only if the overlay width is sufficient to effectively transfer stresses across the discontinuity (flawed region). j 2.2 Weld Procedure and Material il The combination of the Alloy 625 weld metal and the selected welding process (GTAW) assures - 7 adequate weld toughness. Therefore, per [5], secondary (thermal expansion) stresses do not need 1 to be considered in the weld overlay design. Alloy 625 also provides a sound weld with resistance i-' to pitting corrosion, crevice corrosion and stress corrosion cracking (SCC), as these are potential

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I failure mechanisms for the postulated defect. It is assumed that this overlay repair is performed

' subsequent to drying, to avoid the possibility of blow-outs during welding. s J 4 The following measurements and controls for the repair are recommended, when practical. 1. Perform preweld " bake", or drying to avoid blow-outs during welding. 4 a )~' 2. Measurement of final weld overlay repair thickness and width. 4 j" 3. Measurement of axial (vertical), diametral (at the canopy seal region), and angular (from vertical) distortions produced by the weld repair. Although all the above measurements may not be practical in the field application, it is recommended [ that at least a mock-up of the repair be performed to verify acceptable results in this respect, before the repairis implemented. l i 2.3 Code Reconciliation l

lT The CSW repair design and remaining life analyses of this report were performed in accordance with the 1989 Edition of ASME Sections III and XI [6], in addition to NUREG-0313, Revision 2 [5].

Additional guidance was taken from ASME Code Case N-504-1. Paragraph 4.1 of the NUREG [5] j-states that the 1986 Edition of Section XI now provides appropriate criteria for the flaw evaluation and repair of all types of welds, in contrast to earlier editions for some cases. This comment specifically applies to Section XI paragraphs IWB-3641 and IWB-3642, which were used for the CSW analysis. There is no difference between the 1986 and 1989 Editions of Section XI which is pertinent to the present analysis. The 1989 Edition of Section XIis approved in 10CFR50.55. l The residual stress analysis documented in Appendix A was performed using material properties from J i the 1968 Edition of ASME Section III with the Winter 1968 Addenum. The Code of record for the . Sequoyah reactor vessel is the 1968 Edition of ASME Section III [12]. The CRDM lower CSW i ~ l SIR-95-105, Rev. 0 2-4 Structural Integrity Associates, Inc. l \\ i

+ {y q I overlay repair work is specified to be done to the 1989 Edition of ASME Section XI. The 1980 Edition of Section XI, with Addenda through Winter 1981, is the Edition / Addenda to which the Sequoyah Unit-1 FSARis committed. A Code search was performed to determine any significant technical changes which could affect . reconciliation of the 1989 Edition analyses with the above Codes of record. It was noted that y paragraphIWA-4120 of Section XI was revised in the Winter 1983 Addenda to permit an Owner to use later Editions of Section XI when performing repairs, subject to approval by the regulatory 7 authority (the 1986 Edition has been approved by the NRC). Section XI previously had permitted use oflater editions of Section m for repairs, in this paragraph. In addition to the above guidance by NUREG-0313, Revision 2 [5] to use the 1986 Edition of Section XI, and paragraph IWA-4120 of Section XI permitting later editions of Sections E and XI ~ 'to be used when performing repairs (as discussed above), specific Code paragraphs used in relation to this current study were searched for significant technical changes. Speci6cally, it was found that j NB-3227.7, and NB-4360, added provisions for specially designed welded seals, such as omega and canopy seals, in the Winter 1971 Addenda of Section III. Thus, later Code editions give more specific, improved rules for CSW design. In earlier Editions [eg.1968] of Section M it was found that S, at 650'F was 15.3 ksi, rather than the value of 16.2 ksi in the 1989 Edition, for SA-182, Grade F304, the head adapter and plug material. Since the present analysis for designing the CSW 3 repair thickness assumes general membrane stresses in the CSW to be at the Code limit, use of the .j higher value of 16.2 ksi is conservative. In view of the above discussions, as well as considering other less significant Code changes, it is concluded that the 1989 Editions of Sections E and XI would be permitted by current rules for 4 Section XI repairs, and also provide more specific and complete rules for the repairs of CSWs, as

q designed in this report. Thus, the current analysis is enhanced by use of the 1989 Edition of the i:J Code, The integrity of the repair design is improved relative to that which would result from the Section 11I Code of record for Sequoyah. The 1989 Editions of both Section m and Section XI are approved by the NRC in 10CFR50.55. More recently, Code Case N-504 [13] was published. This SIR-95-105, Rev. 0 2-5 structural Integrity Associates, Inc.

c. I Code Case specifically deals with weld overlay repairs of austenitic stainless steel pipes. This Code . Case offers additional guidance for the overlay repair design for CSWs at Sequoyah. l-. 2

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?' (C) COPYRIGHT 1984, 1990 STRUCTURAL INTEGRITY ASSOCIATES, INC. SAN JOSE, CA (408)978-8200 VERSION 2.1 Date:-23-Sep-1995 Time:-15: 2:55.15 STRUCTURAL REINFORCEMENT SIZING EVALUATION 5 STRUCTURAL REINFORCEMENT SIZING USING SOURCE EQUATIONS FOR CIRCUMFERENTIAL CRACK WSI-200: SEQUOYAH CRDM CANOPY SEAL WELD OVERLAY REPAIR WALL THICKNESS = 0.0750 MEMBRANE STRESS = 16.2000 SAFETY FACTOR = 3.0000 BENDING STRESS = 0.0000 SAFETY FACTOR = 1.0000 STRESS RATIO = 3.0000 ALLOWABLE STRESS = 16.2000 FLOW STRESS = 48.6000 i L/ CIRCUM 0.50 0.60 0.70 0.80 0.90 1.00 FINAL A/T 0.6475 0.6182 0.6025 0.5967 0.5947 0.5947 ] REINFORCEMENT THICK. 0.0408 0.'0463 0.0495 0.0507 0.0511 0.0511 1..i A 1 SIR-95 105, gey. o 2-7 Structural Integrity Associates, Inc.

].,. L 4 ^ !'l '3 S R '.e Nominal stress in the uncracked section of pipe .,{ P, + P --+{ F-3 ( l i ~ I { { a +-

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jw-0.155 to 160"R (( gg gy %,,wm y I" 5.660 0.D. 0.155" to 0.160"R (re0 (reo mmet Into nicker Arc +1= As Shown NOT TO SCAIE l } Figure 2-2. Canopy Seal Weld Overlay Design and Dimensions Used in Analysis 7 t 1 i 1 SIR-95-105, Rev. 0 3-

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~_ 1 2 ) ~ i '? J 11 I 3 12 2 13 s u J. J 1 5 16 0.075' -.-, 26 6 I 7 17 27 A0 10 Each weld layer = 28 0.120 teen 19 8 gg J 30 .] J 3 I Jytrr m s c u m 1 ~' Figure 2-3. Canopy Seal Weld Overlay Design Bead Sequence SIR-95-105, Rev. 0 2-10 Structural Integrity Associates, Inc.

b... D 9 3.0 WELD RESIDUAL STRESS AND STRESS INTENSITY FACTOR ANALYSIS i Weld overlay repair residual stress and stress intensity factor analyses for the Sequoyah CRDM lower canopy seal were performed, for use in fracture mechanics crack growth evaluations of the repair. 3.1 Methodology These stress predictions are made through the use of a thermal-elastic-plastic finite element program ? WELD 3 [8]. Details of the methodology, including assumptions, are given in Appendix A to this report. Dimensions of the component (Figures 1-1 and 1-2) are similar to those at Zion and Diablo Canyon [2,3], and welding parameters are given in Appendix A, as used in this analysis. It can be seen in Appendix A that this analysis consists of essentially two parts; a thermal analysis and a stress analysis, to model the welding process in both thermal and mechanical respects. Fracture mechanics stress intensity factors (K) were also computed for the design repair, using a finite element model and the concept of elastic energy release rate [14], as detailed in Appendix A. This result was used to check the more general fracture mechanics crack model [11] results used in crack growth remaining life predictions. Such veri 5 cation ofK is considered important for the CRDM lower CSW geometry, because of the boundary conditions and constraints imposed by the adjacent thicker sections of the component. J 3.2 Residual Stress Analysis and Results Detailed residual stress results are given in Appendix A. The results of this analysis are to be used to develop fracture mechanics stress intensity factor (K) solutions for use in a fracture mechanics crack growth law, f 1j ~ SIR-95-105, Rev, 0 3-1 Structural Integrity Associates, Inc. i

f' i ' 3.2.1 Model l The overall finite element model used for the WELD 3 analysis is shown in Figure 3-1. 1 Figure 3-2 shows the sequence of welding the repair overlay. This model was developed to provide j the capability to analyze up to three 0.12 inch thick layers of Alloy 625 weld overlay, with each layer bead sequence starting at the top as shown in Figure 3-2. Note that the bead sequences run from top I to bottom in all cases, except for segments 4 and 5 in the first layer. This model is in accordance with the Zion repair mock-up results [9]. The meridional stresses at Section C are ofprimary interest for the crack growth analysis. However, a study of stress contours resulting from the WELD 3 analysis (Appendix A) revealed that the highest meridional stresses occur at Section D, which is immediately adjacent to the original lower CSW ] location. The meridional stress results for Sections A, B, C and D are shown after each of the three layers are deposited for the final design in Figures 3-3 through 3-5. These residual stresses are those resulting from welding, and after the componer.t is cooled to 100*F. Note that the stresses after the third layer are significantly lower in the region at the inner surface of the CSW. Figure 3-3 shows the resultant residual stresses after the 3-layer repair, for Sections A, B, C and D, when heated to a service temperature of 550*F. It can be seen in Figure 3-3 that the stresses at the inner surface, where a crack could exist, are lower at 550*F than at 100*F in Figure 3-4. This is further illustrated c in Figure 3-5. Thus, the results at 100*F will be used as the bounding case for the crack growth remaining life predictions. Additional verification of this conclusion was found in stress intensity factor calculations which showed that the crack growth rates for cracks with initial depths ofup to 0.075 inches are smaller for the 550*F stress distribution than for the 100*F distribution. i 1 3.3 Stress Intensity Factor 7i The pc-CRACK Model A is shown in Figure 3-7. This model represents a continuous surface crack } in half space [11]. This crack model realistically models the decrease in stress. field due to the _t secondary nature (and thick-section constraints) of this stress, as the crack extends. 4 a ) h StructuralIntegr!!yAssociates,Inc. SIR-95-105, Rev. 0 3-2

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p b,$'. ', 4.0 CRACK GROWTHRATE EVALUATIONS i The fracture mechanics program pc-CRACK [11] was employed with the preceding calculated residual stress distributions and the following crack growth law to estimate SCC crack growth rates in the repairs. The initial flaw for such an evaluation is conservatively assumed to be completely. through the wall of the original CSW (0.075 inch thick wall). 4.1 Crack Growth Law 0% } The crack growth law for SCC has the following form: I n = cr dt I where da/dt is the crack growth rate in in/hr, K is the stress intensity factor in ksi gri, and C and n are constants which depend on the material and environment. The stress intensity factor K is ]' computed as a function of stress and crack size for a given geometry [11]. ] The objective here is to estimate an SCC growth law for the fracture mechanics analysis of remaining life of the Alloy 625 weld repair to the Sequoyah lower canopy seal weld on the CRDM. Several references [4,5,15-20] have disclosed relevant data, for both high oxygen coolant (as in a crevice) and off-chemistry conditions, including sulfates, caustic and chlorides. Fracture mechanics test crack growth rates from [15], in water with 7 ppm. 0 and 1 ppm. H SO 2 2 4 1g at 550*F, are shown plotted in Figure 4-1 for various Inconel alloys. This figure also shows data on

u Type 304 stainless steel which was used by the NRC in NUREG-0313, Rev. 2 to derive an SCC law

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4 metal /weldments and Inconel alloys. Thus the exponent of 2.161 is also used for Inconel alloys.

Based on this assumption, curves are drawn on Figure 4-1 through the data points for Alloys 600 and 690. The curve for crack growth a factor of10 lower than the NRC curve, which was derived based U I on canopy seal weld cracking ofType 308 at Diablo Canyon (3), is also shown. The values of C for ) these curves are derived as follows: l NRC/10: C = 3.59 x 104 4 I Alloy 600: C = 1.61 x 10 l Alloy 690: C = 4.95 x 10-10 e9 References 15 and 17 state that SCC resistance in Fe-Cr-Ni alloys is a strong function of Cr content. Furthermore, the Mo and Nb contents in Alloy 625 arejudged [18,19] to further mitigate SCC even 4 for I though their effect is not as strong as that of Cr content. Thus, C is computed as 1.11 x 10 Alloy 625, based on a Cr content of 21% Cr and linearly interpolating the factor ofimprovement of i 32.5 for C in going from Alloy 600 (15% Cr) to Alloy 690 (29% Cr). This curve for Alloy 625 is plotted as the dashed curve in Figure 4-1. The law for crack growth a factor of 20 slower than the I NRC curve for Type 304, and a factor of 2 slower than the NRC/10 curve used for Types 308L and i 316L weld metal at Diablo Canyon (3), is shown in Figure 4-1 to be just above the estimated Alloy .J 625 curve. The Diablo Canyon curve is based upon actual CSW estimated crack growth rates. Thus, the NRC/10 curve is used to bound the Alloy 625 estimated growth rates, and the following law is ' Ai employed for Sequoyah remaining life predictions for the Alloy 625 weld repair: i da, 3.59 x 10-' g2.ist 1 dt 10 ~ l

• s STR-95-105, Rev. 0 4

StructuralIntegrity Associates, Inc.

u. h a-The following non-quantitative test specimen (U-bend) results generally support the use of the above law for other postulated off-chemistry environments. Results [4] in 50% caustic show Alloy 625 to crack about 90 times slower than Type 304, justifying the NRC/20 curve for this environment. Results [16] in MgCt2 and with 500 ppm Ci show no cracking for Alloy 625, whereas Type 304 b cracked. Results [17] in high oxygen (6,20-100 ppm) with varying pH'(some induced by H S0), 2 show general agreement with the da/dt results in Figure 4-1. The da/dt results are 7 times to 50 times lower than for Type 304 when estimated from U-bend tests for Type 304 and Alloy 625. It should be noted, however, that the above data is generally from tests of Alloy 625 wrought material. Alloy 625 weld material may be slightly more susceptible to SCC propagation [20]. Because of the potential increase of susceptibility of A!!oy 625 weld metal as compared to wrought 625 metal, the i~ NRC/10 law will be employed for remaining life predictions in this study. 1 4.2 Remaining Life Estimate H The fracture mechanics program pc-CRACK [11] was employed to compute crack growth rates and remaining lives (to penetration of the crack to 75% of the repaired wall thickness) for the overlay repair design. The postulated initial. defect for these analyses is an infinitely long crack completely j through the original 0.075 inch wall of the CSW. A bounding 2 ksi applied stress [2,3] and the 4 f residual stress distributions shown in Section 3.0 of this report were used to compute K as input to the NRC/10 crack growth law. This law is as follows: ..] b = 3.59 x 104X" dt { where da/dt is in in/hr and K is in ksi fin. ~ l The meridional residual stress distributions for Sections C and D at 100*F, presented in Section 3.0 m for the 3-layer (each layer is 0.12 inch thick) design, were employed as discussed above (with a 2 ksi l applied stress) to compute crack growth rates. The remaining life for the repair is predicted to exceed a 57 years (the longest the analysis was run). m f StructuralIntegrity Associates, Inc. SIR-95-105, Rev. 0. 4-3

>>j,.. 7 4.3 Fatigue Crack Growth + 4 The increased section thickness in the canopy seal region reduces the stress significantly as compared i lt to the stresses in the original CSW as reported in [12). The cyclic stresses in the as-overlay-repaired j condition are predicted to be well below the endurance limit for the Alloy 625 material, so fatigue 'I usage for the unflawed materialis not a concern. Fatigue crack growth of existing defects is also not 'j a concern, since the as-repaired cyclic stresses are small and the predicted num'oer of cycles [12] is 'l also small. l -m .t '1 1 1

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?, f' y 1 5.0 ' CONCLUSIONS 'j The following conclusions resulted from this study. 1. The minimum required design thickness of the CSW overlay repair, in accordance with ,a ASME Se:tions III and XI, and NRC requirements for structural reinforcement is 0.0511 inch. This thickness accounts only for structural reinforcement, with no crack growth . considered. Further considerations of SCC growth rates follow below. This overlay should be blended into the thicker sections of the head adapter and CRDM pressure housing, as - shownin this report. In order to provide SCC resistance, the selected repair materialis Alloy 625 weld filler metal, using the GTAW (TIG) process. l 2. . A fra'cture mechanics, SCC growth law for Alloy 625 a factor of ten slower than the upper bound NRC curve forIGSCC in Type 304 isjustified for the remaining life predictions of the i repair. This law is based upon experience with CSW cracking rates with Type 308 welds, upon crack growth rate data for Alloys 600 and 690, and upon a comparison with the SCC resistance of Alloys 600 and 625 versus Type 304 in various off-normal chemistry i environments. 3. A Code reconciliation was performed establishing the acceptability of using the 1989 Edition of AS. E Sections III and XI for the overlay repair. M ,L ,q The overlay weld repair design employs three layers, each 0.12 inch thick, of alloy 625 4. , a overlay, with the bead sequence starting at the top of the component for each layer. The f predicted remaining lives for critical cross-sections of this repair, based on a residual stress j., analysis and a 2 ksi applied stress, are in excess of 57 years for the NRC/10 law. yn 5. The effects of service temperature and repair bead sequence completion were also examined ) _, in this analysis. Results suggest that residual stresses are improved for this design when welding heat inputs are kept low; thus, avoid high heat inputs. Furthermore, the significant addition of an SCC-resistant material during this repair serves to substantially mitigate further leakage. l !s-SIR-95-105, Rev. 0 5-1 StructuralIntegrity Associates, Inc. 4 ' ~~

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6.0 REFERENCES

1. Westinghouse Electric Corporation Drawing Number 541F469, Revision 2, Control Rod Mechanism Head Adapter Details, March 31,1977, SI File No. WSI-20Q-213. t 2. _Copeland, J. F., et al, " Design and Analysis of a Weld Overlay Repair for the Zion CRDM iJ Canopy Seal Welds", SI Report No. SIR-88-040, Revision 1, SI File No. WSI-11Q-200. )] 3. Copeland, J. F., et al, " Design and Analysis of a Weld Overlay Repair for the Diablo Canyon ii CRDM Canopy SealWelds", StructuralIntegrity Report No. SIR-85-029, September 1988, SIFile No. WSI-11Q-201. I 4. McIlree, A. R., and Michels, H..T., " Stress Corrosion Behavior of Fe-Cr-Ni and Other Alloys in High Temperature Caustic Solutions", Corrosion-NACE, Vol. 33, No. 2, February,1977, 4 pp. 60-67, SI File No. WSI-11Q-203-1. 5. NUREG-0313, Revision 2, " Technical Report on Material Selection and Processing Guidelines for BWR Coolant Pressure Boundary Piping", U.S. NRC, January,1988. 6. ASME Boiler and Pressure Vessel Code,1989 Edition, Sections III and XI with no Addenda. 7. Section XI Task Group for Piping Flaw Evaluation, ASME Code, " Evaluation of Flaws in l Austenitic Steel Piping", Journal ofPressure Vessel Technology, Vol.108, August,1986, pp. 4 352-366, SIFile No. WSI-11Q-203-2. 8. Stone'sifer, R., et al, " WELD 3 Computer program Verification Manual", April,1988. 9. Telecopy, Reed, L. (Welding Services, Inc.) to Gerber, D. A., (Structural Integrity), Welding ~ Parameters, SI File No. WSI-11Q-205. 5. 10. StructuralIntegrity Associates Calculation No. WSI-11Q-302, Revision 0, dated April 13, 1992, " Crack Growth Analysis of CRDM Lower CSW Overlays." j 11. Structural Integrity Associates pc-CRACK, Version 2.1,1992. 12. Westinghouse Electric Corporation, "High Speed Control Rod Drive Mechanism Stress Analysis", E3980, Addendum 4. y 13. ASME Code Case N-504-1, August 1993. 14. Rybicki, E. F., Stonesifer, R. B., and Olson, R. J., " Stress Intensity Factors Due to Residual Stresses in Thin-Walled Girth Welded Pipes", Journal of Pressure Vessel Technology, Vol. ~ l 103,1981, pp. 66-75. 3 1 I SIR-95-105, Rev. 0 6-1 ~ Structural Integrity Associates, Inc. J

j.. i 15. Page, R. H. and McMinn, A., " Stress Corrosion Cracking Resistance of Alloys 600 and 690 and Compatible Weld Metals in BWRs", EPRI Report No. NP-5882M, July,1988, SI File No. WSI-11Q-203-3. 16. Telecopy, Nelson, J. L. to Giannuzzi, A.J., Alloy 625 SCC Results, October 25,1988, SIFile No. WSI-11Q-203-4.

a 17.

Nelson, J. L. and Floreen, S., "An Evaluation of the SCC Behavior ofInconel Alloy 690 Weldments in a Simulated BWR Environment", SI File No. WSI-11Q-203-5. 18. Yamauchi, K., et al, "Effect ofNiobium Addition on Intergranular Stress Corrosion Cracking Resistance ofNi-Cr-Fe Alloy 600", Proceedings ofthe Corrosion Cracking Program, ASTM, ASME, NACE, etc., Salt Lake City, Utah, December 2-6,1985, pp.11-22, ST File No. WSI-11Q-203-6. o 19. Frank, R. B. and DeBold, T. A., " Properties of an Age-Hardenable Corrosion-Resistant "~ l Ni-Based Alloy", Materials Performance, NACE, September,1988, pp. 59-66, SI File No. WSI-11Q-203-7. ] 20. ABB Atom AB," Stress Corrosion Cracking of Alloys 600 and 182 in BWR Environments - ) l J Interim Report", TR-100658, EPRI, May 1992. ..q 4 1 hi '.) 4 d I a i '4 SIR-95-105, Rev. 0 6-2 Structural Integrity Associates, Inc.

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3-0- APPENDIX A i Detailed Description of Weld Residual Stress and Stress Intensity Factor Analyses for the Sequoyah Lower CRDM Canopy Seal Weld Overlay Repair r J 1 t. l J a . SIR-95-105, Rev. O A-1 StructuralIntegrity Associates, Inc.

- --~- - _ ,i 0' i;s INTRODUCTION i SI reported on the overlay designed for the Zion spare CRDM canopy seal [1]I. The Zion canopy l, seal and overlay design are identical to the Sequoyah CRDM lower canopy seal and overlay design. d A weld mock-up was prepared for the Zion design and verified against the welding residual stress model. This appendix provides the details of the residual stress analyses performed to support the weld overlay design at Zion and its application to the current overlay design for Sequoyah. 3 i1 i WELD MODELING BACKGROUND 4 i, The WELD 3 software for welding analysis consists of several computer programs which are used j for predicting temperature transients, predicting residual stress changes due to these temperature j; l transients, and then graphically presenting the results of the temperature and stress analyses. The stress predictions are made through the use of a thermal-clastic-plastic finite element program. The programs which are responsible for temperature predictions rely on a variety of analytical solutions l} to basic heat conduction problems associated with moving and stationary heat sources, on j semi-empirical formulas and approximations to other heating methods ofinterest such as induction lj heating, and on nonlinear heat transfer finite element methods. f', ] The methodology for modeling welds was initially developed at Battelle's Columbus Laboratories under funding first from the U.S. Nuclear Regulatory Commission and then from the Electric Power [j} i:j Research Institute [2-7]. Since this initial work, extensions and refinements have been made to the methodology through continued EPRI support [8] as well as through applications of the methodology l6 to a variety of electric power industry related problems [9-17]. l l. The WELDS II software evolved with the methodology [2-8], and as a result of the code evolution, lj and a low emphasis on computational efficiency, the WELDS II software became inefficient, highly J patched, and generally difficult to maintain. The solution to these problems has been the development i Reference numbers refer to those at the end of this appendix. SIR-95-105, Rev. 0 A-2 StructuralIntegrity Associates, Inc. .J

j l

I of a completely rewritten set of software tools known as WELD 3 [18,19]. Since the people who originally developed the methodology and the WELDS II software have developed WELD 3, all that l*

was learned from the evolution of the WELDS II software has been applied to the design, 29 development, and verification of the new software. The WELD 3 software is used for the weld modeling of this study, j OVERl AY WELDING TEMPERATURE CALCULATIONS t, The geometry of the lower CRDM canopy s'ealis defined in Figures A-1 and A-2. Figure A-3 shows i the axisymmetric finite element grid which was used for the WELD 3 simulations of th'e weld overlay. The model did not include pre-existing cracks. The assumed mechanical boundary conditions and the analysis coordinate system are also illustrated in Figure A-3. The boundary conditions did not allow rotation at either end of the modeled region, but did allow unconstrained radial or axial 3 shrinkage and expansion. For the thermal analyses, all model boundaries were assumed to be insulated. I I f Heat Input Rates. i i The overlay welding technique was assumed to result in an electrical input energy of 13.9 kJ/in for the passes of the first weld layer and 16.6 kJ/' for the passes of the remaining layers. The welding d m speed for alllayers was assumed to be 3.5 ipm. These energy rates resulted from actual amperage, [ voltage and weld speed used during construction of the Zion mock-up overlay repairs at Welding Services, Inc. [20]. The finite element thermal solutions assumed a weld heat input efficiency of 40%. This means that the thermalinput was 40% of the electrical input energy. This efficiency was based on the work of Christensen, et al [21] on TIG welding of steel as reported by Masubuchi [22]. It was found in the thermal modeling that this efficiency resulted in reasonable temperatures for the simulation of the first weld layer. A higher efficiency would have resulted in the model predicting burn-through. .l 2 i[ SIR-95-105, Rev. O A-3 StructuralIntegrity Associates, Inc. p 4

l '.

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b Definition of Model Weld Seements d One major simplification of the weld overlay modeling methodology is the grouping of several weld 1 passes into a weld segment [4,9,10]. This reduces the input preparation and computational expense in proportion to the number ofpasses which are grouped into a modeled segment. Figure A-4 shows b how the finite elements used to represent the overlay weld material were grouped into weld segments. o Sections A, B, C, and D ofFigure A-4 are the sections for which through thickness residual stresses 5 ' and crack growth rates were considered. 1 ' Figure A-4 shows the model which reflects the overlay design. As can be seen, this'model assumed up to three weld layers with eac!, layer being 0.120 inches thick. The welding sequence was that passes were generally put down stating from the top of the canopy seal. This sequence and the f "end-in" sequence of the first layer was adopted in an attempt to minimize the possibility of [ burn-through. i In grouping passes into model segments for the present am ysis, it was assumed that as thinner j j sections are modeled, the number of passes that can be reasonably included in a segment becomes smaller. Therefore, while some heavy section weld overlays have been modeled by grouping as many l as 10 passes into a segment (more typically 4 or 5), the models of this study grouped approximately two passes into a weld segment. i !d In butt weld and overlay simulations of heavier sections, it has generally been found necessary to I modify the thermal analysis to account for modeling more than one pass per model segment [5,10]. When using analytical solutions, this is usually accomplished by artificially widening the computed

'q-isotherms to match the width of the weld segment. Due to the tendency for weld passes to overlap, due to the grouping of only two passes into a weld segment, and due to the difficulty of adopting the isotherm widening technique when using a numerical thermal analysis approach, no such 1

modifications were applied to the thermal analyses of this study. The credibility of the isotherms from ,'{ the resulting finite element thermal analyses and the fact that sufficient heat input existed to raise the ,a l-1s Ld i ~ STR-9.%105, Rev. O A-4 Structural Integrity Associates, Inc. t

s,,. a-weld segments and neighboring material to temperatures at wh!ch stresses are largely relieved, ) support the reasonableness of the adopted analysis procedures.

Weld Deoosition Modelina Anoroach '

] )y The WELD 3 finite element thennal analysis simulated the sequential deposition of weld metal by l sequentially activating appropriate elements and nodes as each weld segment was " deposited". The ~ output from the thermal analysis which was used as input for the stress analysis consisted of a file of l element temperatures. Insctive elements were given a teinperature which was above the stress-free l' temperature defmed in the stress analysis input. Elements above the stress-free temperature have an ~ insignificant stiffness and are assumed (i.e., forced) to have zero stress and strain and thus do not significantly affect the stresses or displacements in the r aterial which does exist. The stress free temperature was selected so as to be consistent with the input mechanical property data. In the ] present analyses, the stress-free temperature was assumed to be 2000"F. It can be seen from'the

o assumed temperature dependent properties ofTable A-1 that this temperature resulted in the stiffness j

jj of the stress-free or " inactive" elements being small compared to the stiffness of" active" elements. i j Transient ThermalModeling Anoroach t The combination of thin sections and curved boundaries of the canopy seal made the usual p., temperature calculation procedures for welded overlays inappropriate. Typically, weld overlays have been applied to sections which could be accurately represented as being an insulated half space, or l as being a unifonn thickness plate or pipe wall. The wall is generally assumed to be insulated at the outer surface and the inner surface is either assumed to be insulated or to have a water heat sink.

o Transient thermal solutions for a steadily moving point heat source and such idealized boundary conditions are easily obtained using analytical solutions. Temperatures for the canopy seal weld overlay simulations were computed using the finite element method. This numerical method was better able to model the curved boundades of the canopy seal geometry than simple analytical models p'

and also made it possible to include the effects of temperature dependent thermal properties. While i .f it would have been possible to include other factors such as convective or radiant heat losses at the [ +, SIR-95-105, Rev. O A-5 StructuralIntegrity Associates, Inc.

,p... surfaces, the present analyses assumed all surfaces were insulated (i.e., heat loss at the surfaces was included in the welding efficiency factor). While the finite element model has apparent advantages over the analytical models, the movement of the welding arc in the circumferential direction cannot be exactly represented using a 2D axisymmetric finite element model. I J Since the numencal thermal model simulated a 3D moving heat source problem using a 2D T axisymmetric assumption, some approximations were inherent [18]. The heat input to the 2D numerical model was computed from the welding heat input rate (Bru/in) by multiplying this input ~ rate by the circumferential length of the pass or segment being modeled. This amount of heat was ~ input to the model at nodes which were interior to the weld segment being deposited using a ramp function (with respect to time) which started from zero at -1 second, rose linearly to a peak value at t=0, and then decreased linearly to zero at +1 second. Using this approach, t=0 approximately represented the time when the welding are passed the plane for which the thermal transient was being i modeled. This approach has been found to produce results which are relatively insensitive to the length of time used for the ramp function heat input [18]. 1 Using the 2D axisymmetric fmite element approach for modeling the thermal transient due to a moving point heat source has been found to predict peak temperatures which tend to be on the order of 10% higher then those from a comparable 3D analytical solution [18]. This tendency is apparently 7 the result of heat flow parallel to the heat source path in the 3D solution that is not reflected in the 2D solution. That is, the cool material ahead of the heat source acts as a heat sink for times shortly after the passing of the heat source. For times when the heat source is several thicknesses beyond l the modeled section, the 2D and 3D solutions are in near perfect agreement (assuming material and geometric representations are consistent for both solutions). '1 Another factor which enters the thermal analysis when using the numedcal solution approach rather } than an analytical approachis the effect of time step size and the number of time-steps used to obtain the transient thermal solution. For the present ana' lyses,10 time-steps of 0.2 seconds were used for the two-second heat input phase of the solution. Then time-step sizes were increased to values i ranging from 0.25 to 1.0 seconds depending on the cooling rate (i.e., thickness of the segment being !h, SIR-95-105, Rev. O A-6 StructuralIntegrity Associates, Inc.

4, l modeled and the time). A sensitivity study on time step size showed that changing time-step size by a factor of two did not significantly affect the results. It was found that the combination of using temperature dependent properties and finite time-step sizes tended to result in some integration-related error. The method for quantifying this error was to compute the heat content of the entire model and then to compare this to the specified heat input. t . It was found that the heat content could become 5 to 10% greater than the specified heat nput. The i 7 origin of this discrepancy wa determined to be the temperature-dependent properties and the solution I procedure which uses temperatures at the beginning of the increment to interpolate material properties. Since temperatures generally increase during the time period that was modeled and heat capacity increases with temperature, the solution tended to underestimate heat capacity and thus [ ' over-predicted temperatures. i The combined effect of the time-step integration error and the 2D modeling effect discussed previously, could have produced temperatures on the order of 15% above those that would be predicted using a 3D model and very small time-step sizes. While this may sound like a rather large error, it should be kep,t in mind that the wdding heat efficiency has an uncertainty of at least 10% (more likely 20%). It is possible to adjust the heat input to the model so that calculated heat contents are more in line with the desired welding heat input, thus correcting to some extent the integration related error. The overlay analysis used this correction approach by reducing nodal heat inputs by !q 5%. Therefore, heat content errors were reduced to O to 5% above the desired heat input. To the !.J extent that higher heat input rates are believed to create less favorable residual stresses for the canopy seal overlay geometry, any tendency of the model and modeling assumptions to over-predict temperatures is believed to introduce conservatism into the analysis. Therefore, with the uncertainty regarding actual welding efficiency, the approach was to err on the side of higher temperatures. The analysis did not predict unreasonably high temperatures during the welding of the first overlay layer l *] (when errors in heat input would be most apparent); therefore, it seems that the analysis did not 1 - t produce overly conservative results. ~ .,e STR-95-105, Rev. 0 A-7 gg,oggg,,y jaggy,ggy ggocygggg, gag, p

1

2 1

I?iL TEMPERATURE RESULTS F Since no thermocouple measurements were taken during the Zion mock-up welding, the primary basis i for judging the reasonableness of the thermal analysis results was in terms of peak inner surface temperatures during the welding of the first overlay layer, and in terms of temperatures in the newly deposited material for timesjust after " passing" of the arc. Figures A-5 through A-8 contain isotherm i l plots for four of the 16 modeled weld segments of the analysis which assumed three weld layers at Il a thickness of 0.120 inches per layer. The isotherms are for the times at which temperatures were used for input to the residual stress anaiysis. The first plot for each segment is for the time when the new material has cooled to about 2000*F. The second plot is for the time when the n'ew material has. s 5 cooled to about 1100*F. 4 Figure A-5 shows isotherms for the weld segment of the first layer which produced the highest maximum inner surface temperature (about 2200*F). The plot is for a time shortly after this peak temperature was reached and thus has an inner surface temperature of about 2000*F. The isotherms clearly illustrate the effects of the thinner section below the weld. 1 Figure A-6 shows the isotherms for Segment 8 (which is in the second weld layer). Though the heat j input for this layer is greater than for the first layer, the additional thickness results in lower peak inner surface temperatures than computed for the first layer. The peak computed inner surface temperature for the second weld layer was about 1125'F and occurred for Segment 9. q Figure A-7 shows the isotherms for Segment 14 which is the third from last segment of the third and b last layer. The peak computed inner surface temperature for this layer of about 650*F occurred for j this segment. The isotherm plot at 7.0 seconds corresponds approximately to the time of this peak temperature. !mb Figure A-8 shows the isotherms for the last segment of the last layer. This segment has the fastest l} cooling rate of all modeled segments due to the large volume of material surrounding the weld. The 4 9!

s6 h StructuralIntegrityAssociates,Inc.

~ l ' SIR-95-105, Rev. 0 A-8

l f. s peak inner surface temperature for this segment was about 375'F and did not occur until a time of about 8 seconds. By this time the weld region had cooled to about 550*F. OVERLAY WELDING STRESS CALCULATIONS i The welding thermal transient for each weld segment was represented in the residual stress simulation } by three temperature distributions. The first distribution corresponded to the point in time when the newly deposited material had cooled to about 2000*F. The second distribution corresponded to the point in time when the newly deposited material had cooled to about 1100*F. The third distribution was a uniform interpass temperature of 100*F (which was also assumed as the initial and fmal temperature). The above described temperature analyses produced 48 temperature distributions for the 3-layer simulation. To determine the residual stresses, these temperature distributions were input to a 2D, incremental, thermal-elastic-plastic finite element stress analysis using the WELD 3 software. The stress analysis assumed axisymmetdc behavior as is usually done when modeling circumferential welds of axisymmetric geometdes. Essentially, all verification of the weld modeling methodology has involved showing that 2D models can provide reasonable estimates of residual stresses due to welding. The ability to use 2D models is fortuitous since the cost of doing 3D analyses would limit weld simulations to welds oflittle n practicalinterest. For example, a recent study did a 3D anaiysis of a one-inch single pass weld on a 4 x 4 x 0.5 inch plate [23]. Extrapolating the cost of this analysis, a 3D analysis of a typical butt weld on a 24 inch pipe would take about one hour of Cray 2 CPU time for each 2 inches ofweld. A 36 pass weld would, therefore, require about 1300 hours of Cray 2 CPU time (typically costing in ercess of $1000.00 per hour). ~I Each temperature change associated with going from one temperature distribution to the next was simulated by subdividing the change into ten increments. This number ofincrements has been found to provide a good balance between solution stability / equilibrium / convergence and cost. l h StructuralIntegrily Associates, Inc SIR-95-105, Rev. O A-9

i a.. 4 '1 Comparisons of predicted overlay induced stresses with measured overlay stresses have only been made for relatively heavy section overlays (typically with initial wall thicknesses of about an inch). The primary basis for expecting the methodology to provide reasonable predictions for the canopy seal is that the overlay welding methodology is a modification of an e,arlier developed butt weld t 1 modeling methodology which has been shown to reasonably predict stresses in sections as thin as T 0.180 inches [2,3,18]. ' 1 RESIDUAL STRESS RESULTS I Unlike overlays on heavier sections, the residual stress state of the canopy sea! changed significantly as additionallayers were added to the overlay. Residual stresses also sometimes changed significantly from the welding of one pass to the next. To detemtine the sensitivity of the final predicted residual stresses to overlay design parameters such as number oflayers, layer thickness, welding direction, and weld bead position, stresses have been plotted and examined at various stages of overlay completion for the overlay design. Stress Results for the Overlav Design Figures A-9 and A-10 show the stresses at Sections A, B, C, and D after completion of each layer of the 3-layer design. It is seen that there is a significant improvement in the inner surface stresses H with each additional layer. Examination of stress contour plots covering the entire seal region .I;J showed that the peak inner surface meridional stress occurred at Section D. Preliminary crack growth calculations further suggested that the through-thickness meridional stress distribution at f Section D results in the most rapid crack growth rates of the Sections A through D. ~ Effect of Ooeratine Temoerature on Residual Stresses ii ?~ All of the residual stress plots presented above were for the overlay welding reference temperature 'I of 100*F. Since the overlay design calls for Inconel 625 weld metal to be used on a 304/308L i Jt 4 A i 4 .i a e STR-95-105, Rev. O A-10 StructuralIntegrity Associates, Inc.

).. 1 stainless steel component, and since the Inconel has a different coefHeient of thermal expansion from the stainless steel, it is necessary to consider the effect ofgoing to an operating temperature of 550*F. i

i Figure A-11 shows the residual stresses at the Sections A, B, C, and D at 550*F. Section D remains

.y ,J the critical section in terms of the least favorable meridional stresses as a result of the more tensile j m inner surface stresses. ,] Figure A-12 compares the Section D stresses for the 100*F and the 550*F conditions. The change in stress that is seen is due to the stainless steel having a larger coefBeient of thermal expansion than } - the Inconel. For the meridional stresses, there are two effects occurring. First, the whole seal region is being stretched axially due to the larger expansion of the head / cap assembly. This stretching tends j j to make the inner seal surface more tensile and the outer surface less tensile. However, the stainless i steel at the inner surface of the seal has its expansion constrained by the Inconel overlay so that Jl] stresses in the stainless steel portion of the seal actually become less tensile in spite of the overall i tendency of the seal to have increased tensile stress at the inner surface. Stress intensity factor

,a j

calculations showed that the crack growth rates for cracks with initial depths ofup to 0.075 inches l, are smaller for the 550*F, stress distribution than for the 100*F distribution. i STRESS INTENSITY FACTOR EVALUATIONS l l All crack growth calculations have been done using the stress corrosion crack growth option of the lJ pc-CRACK LEFM module (24]. However, the crack growth calculations for the overlay configuration were verified using stress intensity factor versus crack depth data developed with a finite element approach that accurately reflected the seal / overlay geometry. All crack growth calculations, were made using stress intensity factors calculated from an approximate geometric t representation provided within pc-CRACK known as "Model A". This model is for a long surface crack in a half space. J 7 l l 1 1 1 i i: SIR-95-105, Rev. O A-11 Structural Integrity Associates, Inc. ~ ,--.-4-. -r. -.. --4.,.- m.---- ---..-.-------.e-----.------m-

3.- 1 The following discussion compares stress intensity factors based on the finite element method with those based on the approximate pc-CRACK geometric model and shows that use of the pc-CRACK Model A provided realistic crack growth predictions. i i I The Finite Element Mesh and Boundary Conditions i The geometry used for the finite element based stress intensity factor calculations was the 3-layer, 120 mil per layer, overlay design. Since the finite element mesh that was used for the overlay welding sunulation old not provide for the introduction of a crack at the critical Section D, a special mesh was constructed for the stress intensity factor calculation. This mesh is shown in Figure A-13. The mesh differed from that of Figure A-3 in two respects. First, less of the head adapter was modeled. Second, the CRDM tube and the head adapter were not connected. Stress intensity factor results j were found to be insensitive (approximately 1%) to the type of boundary conditions used at boundaries B1 and B2, thusjustifying the first modeling change. No connectionbetween the CRDM tube and the head adapter was assumed since the resulting reduction in radial restraint of the plug was j believed to be more realistic and also produced slightly larger (approximately 5%), and thus more conservative, stress intensity factors. 1 The Stress Intensity Factor Calculation Method The method u' sed for stress intensity factor calculations.was based on the concept of elastic energy .q I release rate, and made use of the relationship between energy release rate and the stress intensity factor. Because the method was an energy method, accurate results could be obtained without resorting to highly refined finite element grids or special crack tip elements. While the method can L be used to calculate mode II as well as mode I stress intensity factors, the mode II component was not significant for the residual stresses of the present study, and therefore only mode I behavior was 7 considered. _} The method consisted ofgenerating an influence matrix relating nodal displacements and forces along the Section D crack plane. This matrix was constructed from a series of finite element analyses in j t f StructuralIntegrityAssociates, lac. SIR-95-105, Rev. 0 A-12 a

a.. .t which different crack lengths were simulated by uncoupling node pairs along the crack plane. For the mesh of Figure A-13, which has 16 elements through the thickness at the crack plane,15 crack lengths were modeled. Each crack length required a single elastic solution. The matrix was i 1 generated using unit nodal forces and, therefore, the resulting matrix could be used to calculate stress y 4 intensity factors for any stress distribution. T This method has been previously used to study clastic crack growth behavior in planar and l axisymmetric bodies [25]. Figure A-14 compares a stress intensity versus crack depth solution i obtained using this method to the exact solution given in [26]. Expedence suggests that the stress intensity factor solutions resulting from the use of this method should be accurate to within 10% for the mesh refinement used in this study. Stress Intensity Factor Results The elastic modulus used for the stress intensity factor calculations was 28.0 x 10 psi (Inconel 625 6 at 550*F) and the Poisson ratio was 0.3. In converting from energy release rate to stress intensity factor, the crack tip was assumed to be in a state of plane strain. Stress intensity factors for three stress distributions were considered. The first was the residual meridional stress at Section D after completion of the 3-layer overlay (120 mils / layer) and after heating to a uniform temperature of 550*F. The second stress distribution was the residual I meddional stress existing at Section C for the same conditions. The third stress distribution was an assumed uniform applied stress of 2 ksi. For the finite element based method, the residual stress distribution was represented in a piece-wise linear manner. Therefore, there was little if any a approximation involved in representing the stress distribution. u ] For the pc-CRACK-based method, it was necessary to fit the stress distributions with a third order polynomial. Therefore, in addition to the geometric approximation, there was also a stress ] distribution approximation. Only for the uniform applied stress was there no approximation of the stress distribution for the pc-CRACK calculations. Figure A-15 shows the quality of fit that could SIR-95-105, Rev. O A-13 Structural Integrity Associates, Inc.

1 [j,, ) be achieved for the two residual stress distributions using the third order polynomial assumed by pc CRACK. Since crack growth is not predicted to result in crack depths greater than 0.2 inches 5 over the lifetime of the overlay repair, the quality of the stress fit was improved by limiting the range of the fit to the first 0.2 inches of the wall thickness. Fitting stresses over the entire wall thickness o significantly reduced the quality of fit with the reduction in quality being most significant for the Section D stress distribution. Figure A-16a compares the WELD 3 and pc-CRACK stress intensity 1 factor behavior for the uniform applied stress of 2 ksi. Since there was no stress distribution ~ approximation in the calculations for this case, the difference in solutions is due to the difference in the assumed geometries and errors in the solution methods. The pc-CRACK results are about 10% above the WELD 3 results at a depth of 0.1 inches. As might be expected, the pc-CRACK model becomes nonconservative for crack depths greater than half the wall thickness. Figure A-16b shows j the pc-CRACK computed stress intensity factor distributions resulting from the polynomial fits of Figure A-15. Since the fits were for stresses within the first 0.2 inches of wall thickness, the stress l intensity factors are only valid for crack depths up to 0.2 inches. The WELD 3 based stress intensity factors and the pc-CRACK results are seen to be in good agreement over the range of crack depths up to 0.2 inches. 'J i .mia b

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] 1a SIR-95-105, Rev. 0 A-14 gg,yggy,,, fgggg,,,y pggggygggg,,gg. ~

j.. 3 I APPENDIX A-i REFERENCES J 1. Structural Integrity Associates Report SIR-88-040, Revision 0, " Design and Analysis of a Weld Overlay Repair for the Zion Station CRDM Canopy Seal Welds", December,1988. = h 2. Rybicki, E. F., et al, " Residual Stresses at Girth-Butt Welds in Pipes and Pressure Vessels", 1 Final Report to U.S. Nuclear Regulatory Commission, Division ofReactor Safety, Research under Contract No. AT (49-24)-0293, NUREG-0376, published November,1977. N 3. Rybicki, E. F., et al, "A Finite Element Model for Residual Stresses and Deflections in Girth Butt Welded Pipes", Joumal ofPressure Vessel Technology, Vol.100, No. 3, August,1978, pp. 256-262. 4. Rybicki, E. F., et al, " Residual Stresses Due to Weld Repairs, Cladding and Electron Beam Welds and Effect of Residual Stresses on Fracture Behavior", Final Report to U.S. Nuclear Regulatory Commission, Division of Reactor Safety, Research under Contract No. AT (49-24)-0293, NUREG-0559, published December,1978. 5. Rybicki, E. F. and Stonesifer, R. B., " Computation ofResidual Stresses Due to Multipass Welds in Piping Systems", Journal of Pressure Vessel Technology, Vol.101, No. 2, May 1979, pp.149-154. ~ 6. Rybicki, E. F. and Stonesifer, R. B., "An Analysis Procedure for Predicting Weld Repair j Residual Stresses in Thick-Walled Vessels", Journal ofPressure Vessel Technology, Vol.102, 1 No. 3,1980, pp. 323-331. ] 7. Bmst, F. W. and Stonesifer, R. B., "Effect of Weld Parameters on Residual Stresses in BWR J Piping Systems", Final Report to Electric Power Research Institute, NP-1743, Research Project 1174-1, March 1981. i I 8. Rybicki, E. F., et al, " Computational Residual Stress Analysis for Induction Heating of Welded BWR Pipes", Final Report prepared for the Electric Power Research Institute by the j University of Tulsa, EPRI NP-2662-LD, Project T113-6, December 1982. 9. Stonesifer, R. B. and McConaghy, J. M., " Residual Stress Analysis of Sleeve to Safe-end ) Welds" prepared forNuclear Technology Incorporated by Battelle Columbus Laboratories, Columbus, OH, December,1978. j 10. " Optimization ofWeld Overlay Repair for BWR Piping Phase 1", prepared for Electric Power Research Institute by NUTECH Engineers, Inc., San Jose, CA, June,1983. .a t sa STR-95-105, Rev. O A-15 StructuralIntegrity Associates, Inc.

3.. 11. Tang, S. S., Stonesifer, R. B., and Javid, A., " Design Report for Recirculation Piping Sweep-o-lets Repair and Flaw Evaluation Brown's Ferry Nuclear Plant Unit 1", prepared for i Tennessee Valley Authority by StmeturalIntegrity Associates, San Jose, California, October ) 1983. i d, 12. Stonesifer, R. B., Kuo, A. Y., and Gerber, T. L., " Analysis of Browns Ferry Unit 3 Jet Pump } Instrumentation Nozzle Safe-end Repair", prepared for Tennessee Valley Authority by StructuralIntegrity Associates, San Jose, California, July,1984. 13. Stonesifer, R. B., Tang, S. S., "An Evaluation ofIHSI Temperature and Heating Cycle Limitations", prepared forNUTECH Testing Corporation by StructuralIntegrity Associates, . San Jose, California, August 1984. 14. Stonesifer, R. B., " Browns Ferry Unit 2 Recirculation Inlet Safe-ends Determination of j Residual Stress Distribution for ' Modified Full Structural' Weld Overlay", prepared for Tennessee ' Valley Authority by Structural Integrity Associates, San Jose, California, September,1986. l 15. " Development ofInconel Weld Overlay Repair for Low Alloy Steel Nozzle to Safe-end i 1 Joint", prepared for Georgia Power Co. and Electric Power Research Institute by Georgia .!.4 Power Co. and StructuralIntegrity Associates, October,1986. j l 16. " Contingency Plan for Inconel Weld Overlay Repair to Carbon Steel to Cast Valve in RHR M, System at Fitzpatrick", prepared for New York Power Authority by Structural Integrity Associates, San Jose, California, December,1986, il i 17. Stonesifer, R. B., " WELDS II Residual Stress Analysis of Resistance Heating Stress Improvement of Welded Pipes", prepared for New York Power Authority by Structural ] Integrity Associates, San Jose, California, August,1987. 18. Stonesifer, R. B., " WELD 3 Verification Manual", prepared by Computational Mechanics, Inc., Julian, PA, November 1988. 19. Stonesifer, R. B., " Input Description and User Guide for WELD 3", prepared by I ij Computational Mechanics, Inc., Julian, PA, November 1988. II 20. Naughton, T., Telecopy of welding Technique Sheets from Welding Services Inc. to l5 StructuralIntegrity Associates on October 26,1988. 'i 21. Christensen, N., Davis, V. de L., and Gjermundsen, K., " Distribution of Temperatures in Arc J Welding", British Welding Journal, 12(2),1965.

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22. Masubuchi, K., Analysis of Welded Structures, Pergamon Press, New York,1980. SIR-95-105, Rev. O A.16 StructuralIntegrity Associates, Inc. J

1 g.. 1 I 23. Goldak, J., et al, " Progress in Computing Residual Stress and Strain in Welds", Advances in Welding Science and Technology, Ed. S. A. David, May,1986. 24, pc-CRACK User's Manual, Version 2.0 StructuralIntegrity Associates, August,1989. 4 25. Rybicki, E. F., Stonesifer, R. B., and Olson R. J., " Stress Intensity Factors due to Residual } Stresses in Thin-Walled Girth Welded Pipes", Journal of Pressure Vessel Technology, Vol, 103,1981, pp. 66-75. y 26. Tada, H., Paris, P. C., and Irwin, G. R., The Stress Analysis of Cracks Handbook, Del Research Corporation, Hellertown, PA,1973. 27. ASME Boiler & Pressure Vessel Code,1968 Edition with Winter,1968 Addendum. I e a t ad 1 1 1 1 1 Al I m w I am SIR-95-105, Rev. O A-17 StructuralIntegrity Associates, Inc. s

j,. t t Table A-1 Summary of Thermal and Mechanical Properties i Used in the WELD 3 Weld Overlay Simulations m Material 1 (304 stainless steel) Temp E Poisson CTE E Yield k c (F) (ksi) Ratio (1/F) (ksi) (ksi) (Btufm-sec-F) (Bturm^3-F) ~ 50 28,700 0.26 8.16E-6 539.7 36.0 1.82E-4 3.12E-2 1 300 27,100 0.28 8.94E-6 452.3 31.1 2.12E-4 3.46E-2 550 25,800 0.31 9.60E-6 364.8 25.9 2.42E-4 3.71E-2 m 750 24,200 0.32 10.03E-6 296.3 22.3 2.66E-4 3.81E-2 1000 22,500 0.30 10.56E-6 217.9 18.5 2.96E-4 3.89E-2 1300 20,200 0.28 11.41E-6 139.0 14.9 3.32E-4 3.99E-2 m 1600 16,000 0.24 12.63E-6 79.6 10.2 3.68E-4 4.19E-2 2100 10 0.22 14.88E-6 1.0 1.0 4.28E-4 4.58E-2 Material 2 (Inconel 625 weld) Temp E Poisson CTE, E Yield k c (F) (ksi) Ratio (1/F) (ksi) (ksi) (Bturm-sec-F) (Btu /in^3-F) 50 30,000 0.30 7.20E-6 539.7 53.0 1.32E-4 3.00E-2 300 28,800 0.30 7.20E-6 452.3 42.4 1.57E-4 3.22E-2 i 550 28,000 0.30 7.50E-6 364.8 37.8 1.83E-4 3.41E-2 j 750 27,000 0.30 7.90E-6 296.3 36.7 2.06E-4 3.63E-2 1000 25,100 0.30 9.40E-6 217.9 37.0 2.34E-4 3.82E-2 j 1300 23,400 0.30 11.20E-6 139.0 39,0 2.66E-4 4.09E-2 3 1600 18,500 0.30 13.00E-6 79.6 24.8 3.06E-4 4.35E-2 2100 10 0.30 16.00E-6 1.0 1.0 3.70E-4 4.80E-2 F 'JJ l Material Properties taken from ASME Section III (27] e l t W. SIR-95-105, Rev. O A-18 Structural Integrity Associates, Inc. -"m o a~-

.e g.* 1 y slt ociAst.*A* l e e i+ I ,I g / I I I /// s NJ = i c il I W I Gl V RCQr CIA. q J w OETAll A MATCMNC ACME, j w l rdREA0 - sr.I racIT 2 l l st cera:t. *a* e 4 3 h$t -l ] l J[ / o crA. ~ R ctA //_ / T I u / OETAIL 8 e' e !q A i SECBON OF HEAD A0 APTER DIMENSIONAL DATA A = 5.6 6" iF = 2. 5 04" i K = 1.370" iP = 1.0" iU = 0.160"

j 8 = 4.0" lG = 6.0" lL = 5.701" l Q = 6.45" lV = 8 7 38)$.n C = 2.75" iH = 5.177" iM = 4.300" lR = 5.5 5" l W = 30'

!g'g D= 27.75 to 59.2!!! l = 3.875" lN = 45' lS = 0.075" lX = 0.25" E = 11.0" IJ = 0.78" l0 = 30* lT = 0.160" lY = 0.50" l l I n J Figure A-1. Geometry of Spare CRDM Head Adapter Used for Canopy Seal Weld Overlay Model Development

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E StructuralIntegrity Associates, Inc. STR-95-105, Rev. O A-19

4 N 7 0 i -) .b V///// l f MWCR 0% u[eg aw VHM t i 1 I l: \\ I ENLARGED VIE % l I s.oco-s Acut-2c I c I l t. K WAmrAu u ASMC 1Ata: Tf7C Joe SECTION OF HEAD ADAPTER PLUG .1-DIMENSIONAL DATA A = 7.59" lH = 1.25" l0 = 6.995" 8 ' = 5.125" lI = 5.71" iP = 6.06" i C = 3.00" lJ = 15' l0 -- 5.8 0"

,j-0 = 1.30" iK = 0.075" lR = 6.02" E = 0.485" lL = 6.122" l

J F = 0.40" lM = 6.440" l 3 G = 0.40" lN = 8.00" l 1 i i '1 t - Figure A-2. Geometry of Spare CRDM Head Adapter Plug Used for Canopy Seal Weld Overlay Model Development 4 . SIR-95-105, Rev, o A-20 StructuralIntegrity Associates, Inc. -v-=-- m* m.. - ~,,. -, m ..m

e,, I i i i = 4 g-tg"; ga l / / A l l l i l i ,c,a I n 4 A\\ A / . li p I lll 1 - fh \\/ I V \\I IIIl 5 IIII \\lll ] lI// \\IIl 3 1 Q s v i '1 aL 'i Figure A-3. WELD 3 Axisymmetric Finite Element Grid Used for Overlay Weld Modeling with Illustration of Boundary Conditions i a

6. N SIR-95-105, Rev. O A-21 a

g. b 7 l /\\ l /A /\\ Illl ') ll\\\\\\ I tt l4 y sj ml e i I E y ' 11f t t U D' "l "a' in e 7t i f i i ie .i \\ e ~* _e' afni \\i I/ .6e do! in W. 8 il --\\ ',, -C ll \\ l l l Y s / N O ~ ~..., LJH !'E 4 J: p i i t i l l et -_ I i i,it 15 \\

g lI iN wo B

E l l I fl l l l A l 'l a l Figure A-4. Definition of Overlay Weld Model Segments for the Overlay Design Analysis ..? 1 a e StructuralIntegrity Associates, Inc. $}g.C)5 105, Rev. O A-22

L__.. U W E i - [. C M haast hous bbl 3 L___J I ._) Loyar 1 Segmant 3, t= 11.O cec Layor 1 Segment 3, t - 2.O sec 6 'dymisems-- i y Temperature w (F) NE .NO '8 s-- a 4 .o Temperaturo C 600 D 800 (F) i E 1000 f F g__ O O A 200 6So ^"'%. r / ! s$oo mva-/ -3 2000 o 9,o_,_,,,,,,_ s 'S - e""~ F+ in 1600 -E'e.s-e I 1000 gp_p. E-E-T eft 3R aq-TEM 3D D i tj i 'r t T qk ni 1( a 3 6 co? n. D E. &E Dg Figure A-5. Isotherms for Simulation of a Weld Pass Near the Center of the Seal and in the First Weld Layer x$ 8 m-5-* N 9

d . L; _. L_. I.__: is==d Inn.a t L_.J id bei W W U .y-Layar 2, Sagment 8, t== 2. 0 s e c. Layer 2, Sagm:nt'O, t = 0.0 cac Vt un MIMLII I Tearporature ( E'sD'B~ 3~ vi n 200 () 400 gy I'd$iggiumEREGl ti - s - B-8-8-g_ g 'y c soo ,;r a D 000 Temperature (U F 1200 g(% g." ~% 7 a zoo a-1.aoo 0 400 tt 1600 c soo y 1 1800 ^D - ~ 10 g F 1200 0 TEHOA G 1400 lt il 1600 a 1 il I 1s00 y __3 20_0_0_ a ygggg. ~ t 'l y i 8 v n O L ~ i l Sa. l n E i Q t l

• = = ~

y l EE Figure A-6. Isotherms for Simulation of a Weld Pass Near the Center of the Seal and in the Second Weld Layer. i 9 ka 8-l W I N 9

b.. [ .y oJ C M g W Q t I I w .-J _s I I HEE _LIibrahhi l C/'2 Toroperature Temperatura (F) (p3 8i. A. 200 8 200 .L u 400 u 400 o c soo c 600 .t^ D 000 D 000 W E !$$$ E !$$$ cW"- ^ om 4 G 1400 G 1400 0 18 1600 ll 1600 I 1800 I 1800 3 2000 3 2000 o"B-O T fl 146 (120 mil) I j [pa.a-a w y cs* a f'f. - C-c. C D-33-D- .h V f f" ?' \\, o ppt o v<- 2. i' t i, 1 >h J sn I y Layer 3 Sagment 14, t = 2.O ooc Layer 3 Sagment 14, t = 7.O sac = EL = E.w k Figure A-7. Isotherms for Simulation of a Weld Pass Near the Center of the Seal and in the Third Weld Layer 4 D I a 8-W -M E 9

L i_ w C w amme home hJ t L-I .a L---- ~J 5 :- Layer 3, Segmont 16, t= 1.25 sac Lays.r 3, Segment 16, t - 3.5 cac i ltis5rtir:11FM l'etasitii:24m Ternpa rature Temperature 6 (F) (F) Y a 200 a 200 8 U 400 a 400 m C 600 C 600 D 000 D 800 [ E 1000 g 1000 F 1200 F 1200 "o G 1400 G 1400 si 1600 13 1600 I 1800 I 3000 3 2000 3 2000 TEI168 (120 mil) 1 TE 16n (120 mill) I s 1B+s s /gnm, g -o-o sw. + o-0 nf s M [ j o 8f i i -a l' h \\ 0 ) ,e S 5E d 4 (C3 k Figure A-8. Isotherms for Simulation of the Weld Pass of the 12st Weld 12yer w E .N E c2

L. L L._; W L' lJ tuomi bemid W 'a b . ;.1 y 70 73 a0. (100 F) a0 O 50 0 50 y 40 f 40 M 30 30 20 20 10 10 o a O M.. u O A - -10 - -10 { l-20 Ie Layer 1 l-20 g ~30 2e Layer ~ 2 G -30 Ie Layer ~ l "L -40 2e Layer ~ 2 "y -40 y; - Sa Layer ~ 3 o 1 -50 Se Layer-3 1 -50 e -80 -e0 (100 F) -70 O 0.1

0. 2

0. 3

0. 4

0. 5 0

0.1

0. 2

0. 3

0. 4

0. 5

-70 Dietance from Imer Surface (in) Oletance from Inner Surfooo (in) Sqction A Hersdional Strose (120 mile / layer) Section S Heridional Strose (120 mile / layer). 1e Layer ~ l Ie Layer i 70 2e Layer.2 70 ' ' 2e Layer 2 60 3e Layer 3 60 3e Layer 3 O 50 O S0 J 40 J 40 30 30 20 e 20 p. f to p '* 10 - -1 -1 f-20 ~20 g -30 y -30 40

• -40 h_50

-50 eo. (100 F) _ea (100 F). _gg 0 0.1

0. 2

0. 3

0. 4

0. 5 0

0.1

0. 2

0. 3

0. 4

0. 5

_,g m Otetance from Imer Surface (in) Q Oletance from Imer Surface (s n) Section C Hertdional Strees (120 male /loyer) Section 0 Heridional Strees (120 mile / layer) oEDC =E g Figure A-9.. Meridional Stresses at Sections A, B, C, and D After Each Layer of the 3-layer Overlay Design w l w 8 m- -D N 9

E-O l.~

h. _$

M M l__ 'd .-.df 70 70 a0 7; 80 m 50 50 40 40 Q b c 3D 3D m y o 20 20 h ,u y 10 10 a p p f -10 a -30 o s -20 Ie Loyar 1 _yo 2e Layar-2 0 -30 2s Layer 2, -30 2e Log 2 -40 Loyar 3 -40 3s Layer-3 -50 -50 e -a0 (100 F) -a0 (100 F) -70 -70 O D.1

0. 2

0. 3

0. 4

0. 5 D

D. ! D. 2 D. 3

0. 4

0. 5 Distance from Inner Surface (in)

Distance from Inner Sur foos (in) "Seatten A Iloop Strame (120 mile / layer) Saatton 8 Hoop Straea (120 mile /layse) 70 70 - 60 60 50 50 H -~~I 40 40 0 e 0 p D p D i. .: -10

• 8 -10

-20 2e Layer J _go 1e Layer J 2s Layer-2 -30 2e Layas-2 ] -30 3e Layer 3 ~- -40 ~40 3e Layer 3 -50 -50 -en (100 F) -so (100 F) _yo D 0 D. ! D. 2

0. 3

0. 4

0. 5 0

0.1

0. 2

0. 3

0. 4

0. 5

_,a E Distance from Inner Surface (in) Distanos from Inner Eur-face (in) s y section C Iloop Strase (120 mile /loyar) Saotion D lloop Strees (120 mile /loyer) W Figure A-10. Hoop Stresses at Sections A, B, C, and D After Each Layer of the 3-Layer Overlay Design 4C3 x$ 8sr 5 ~% N 9 1

y ( [;j W g [ W h.)

---

8

-C E' ~ U3 D Y Aa Saotiers A 5 60 80 Be Saotion B Ca Saotion C NO ~ ~ Dn Saotion D 1 3 30 30 Y h n 20 8" 20 10 to \\ ~ 0 0 l Ae Seat t or A G -10 -10 y 8s Seatters B

! -20

-20 Ce Snottors C -30 -30 De Saotion D (120 mite /Joyer) ~#* ~O ~ (120 nIta/Jayer-) -50 -50 O D.1

0. 2

0. 3

0. 4

0. 5 0

0.1

0. 2

0. 3

0. 4

0. 5 0 etance from Inner Surface (in)

Distance from Irwier Sur-face (in) Hacidsonal Strama ofter Loyer 3 at 550 F floop St! ase af ter Layer 3 at 550 F co E n E, E E Stresses at Sections A, B, C, and D. After the Last Layer of the 3-Layer Overlay Design and After Being Figure A-11. 4 Heated to 550*F w m R m-E-" E 9

I- . -a led L iheseid b i J InsulE besid sJ I -.A n+3 1 .ww ~ 015 e8 SD fa 1e 200 F L. Js 200 F 8 50 so 5a 350 F 5a 550 F y 40 40 y 3d 3D 3 J 20 20 ID 10 g D - I-en l D a 't. j -10 } -10 dx -20 -20 -30 ~30 La -40 (120 atJe/ layer) -40 (120 atle/ layer) o -50 -50 D D.1

0. 2

0. 3

0. 4

0. 5 0

0.1 P. 2

0. 3

0. 4

0. 5 Diatonoe from Inner Surfaica (in)

Dietence from Inner Surface (In) Section D Heridional Strees afi.c Loyer 3 Section D lloop Strees after Layer 3 i w E nEQ T 4 EE Figure A-12. Comparison of Stresses at Section D for the As-welded Condition at 100*F and After Heating to 550*F 9 1 8 i a-l E-" N 9

31 t. a t u w w t_. u_a w w i u e i .a B3 ,uvuw r us F ( ( l l I \\\\\\ / e a \\) \\ I k \\,\\ \\ t\\ A /t I b E 11 1 f a L J_i_ s D% E-22: P :. :(-L. _L 81 \\ ~ N \\\\MV\\W _i _n_ \\ D ~ x

s. '

N ~ t s - f -w , _c / ? s w y / ~ I /

/

d 3 E a / 5 62 Ok E ~ f Figure A-13. WELD 3 Axisymmetric Finite Element Grid Used for Stress Intensity j l o a 4 82 p Factor Calculations with Illustration of Boundary Conditions and the c Crack Plane (Section D) 8 E-w ~* N 9

..... _.... ~. - -<M L 4 t-J W h I J W lessa + . - ~ s u)h $20 e 'Y +S30MPs g (y gy fy yglga eOO Y s, 7 .8 = p l. C 1 f h 11 AL13 M lbIA LI I) (b o = wei.ne..idwer t.se = (r l 40 y 20 = LJ 2 0 1 m ' m -40 y O - Exact Solution [243 -i=-- X Finita Element Computation = .iro 8' l B B 3 3 a a R R 00 08 0.2 03 0.4 03 06 Or 04 e/T vs 4 R EQ ~ i y, Figure A-14. An Illustration of the Accuracy of the Finite Element (WELD 3), Energy-Based Method for Stress Intensity.' T Factor Determinations b .t N 8 m .M N 9

L_, M L (___] lisend

Misia, L__ _

i U

• • • • • W. -.

.~. -d }' en h 40 30 - u .L Y VELD 3 Section D Raoulte es WELD 3 Section C Raoulte e o N 30 20 o a n u t s 20 I 1 .1 .c m to w

• ii

~ 8 c o 10 0, 3 a h h x x 0 I D-0 l l a a Third Degree Polynomial Fit Third Degree Polynamtal FIL -10 O 0.1

0. 2

0. 3

0. 4

0. 5 0

0.1

0. 2

0. 3

0. 4

0. 5 i

-10 Croch Depth (in) Crack Depth Cin) f 52 pc-CRACK Curve Fit of Section D Straea (550 F) pc-CRACK Curva Fit of Section C Strees (550 F) E nED C n .[= Figure A-15. Illustration of pc-CRACK Third Degree Polynomial Curve Fits for Meridional Stresses at Sections C and Q D of the Overlay Design k e: 11 EiT -D N 9

L L_ L_! aissui Masse L liestd W L M -- -3 3.0 i4 De Section D Reeldval C Ce Section C Residual 0 u 12 o o

2. 5 u

D. C e VELD 3 c 10 g po-CRACK (model A) O n' ~s C O Q 8 d d a 'c 'c u a e r

1. 5 7,,

a o u o 3 D u c 4 x u x

1. 0 u

11 : VELD 3 c 2 c u po-CRACK (modal A) -e p

0. 5 3

0 l A ". '05 '.'5 -2 0 0 0.10 0.15 0.20 02 0.30 0.35 0 0.05 0.1 D 0.15 0.20 0.25 0.30 0.35 .0 Crocis Depth (in) Crocit Depth Cin) a. Uniform 2 itsi applied' stress b. Section D and C residual stresses y n$ N Comparison of WELD 3 Stress Intensity Factor Calculations with Those of pc-CRACK Figure A-16. el E ET-" 59

4.

1. s.

.2 ..t. a e APPENDIXB 1 . pcCRACK SCC Crack Growth Results 1 1 O I 4 e. i i 1 4 I Y l' _ i 4 ; 3 i e ' e 4 d' g h StructuralIntegrity Associates, Inc. SIR-95-105, Rev. 0 - B-1 ~ 9

l ) ) L,, 1 6, - 8 tm i pc-CRACK -(C) COPYRIGHT 1984, 1990 ~~ STRUCTURAL INTEGRITY ASSOCIATES, INC. SAN JOSE, CA (408)978-8200 VERSION 2.1 Date: 25-Sept-1995 Times 12:55: 8.65 STRESS CORROSION CRACK GROWTH ANALYSIS l WSI-20Q: SEQUOYAH CRDM LOWER CSW OVERLAY, SECTION C INITIAL CRACK SIEE= 0.0750 MAX CRACK SIZE FOR SCCG= 0.8000 STRESS CORROSION CRACK GROWTH LAW LAW ID C N Kthres K1C NRC/10 3.590E-09 2.1610 0.0000 200.0000 STRESS COEFFICIENTS CASE ID CO C1 C2 C3 ~- BEND 1.7128 -5.8723 0.0000 0.0000 l MEMBRANE 0.7100 0.0000 0.0000 0.0000 SECC 24.9810 -618.1000 3314.0000 -4183.0000 SECD 40.9851 -788.8973 3968.8642 -5013.4105 APPLIED 10.0000 0.0000 0.0000 0.0000 Kmax CASE ID SCALE FACTOR SECC 1.0000 APPLIED 0.2000

j TIME PRINT

{ 4 TIME INCREMENT INCREMENT 500000.0 1000.0 1000.0 1 (. i crack model: CONTINUOUS SURFACE CRACK IN HALF SPACE l } 1 .. I CRACK ---------------STRESS INTENSITY FACTOR---------------- j -I SIZE CASE CASE CASE CASE CASE i BEND MEMBRANE SECC SECD APPLIED i - 0.0160 0.416 0.178 4.861 8.481 2.511 1 0.0320 0.568 0.252 5.145 9.756 3.551 1 0.0480 0.670 0.309 4.494 9.578 4.349 l 0.0640 0.745 0.357 3.441 8.728 5.022 0.0800 0.801 0.399

• 2.249 7.578 5.615 L

0.0960' O.843 0.437 1.076 6.348 6.151 0.1120 0.872 0.472 0.031 5.187 6.644 0.1280 0.892 0.504 -0.812 4.195 7.102 ..q 0.1440 0.903 0.535 -1.403 3.443 7.533 1 Ll 0.1600 0.906 0.564 -1.709 2.977 7.941 i O.1760 0.903 0.591 -1.708 2.828 8.328 i 0.1920 0.893 0.618 -1.391 3.010 8.698 0.2080 0.878 0.643- -0.760 3.526 9.054 i' St l 0.2240 0.858 0.667 0.179 4.368 9.395 0.2400 0.832 0.690 1.407 5.522 9.725 j. 0.2560 0.802 0.713 2.905 6.961 10.044 ,' q ,.m l-SIR-95-105, Rev. O B-2 StructuralIntegrity Associates, Inc.

'j.' J. I:s .I ,1 0.2720 0.767. 0.735 4.642 8.653 10.353 0.2880 0.729 'O.756 6.587 10.561 10.653 0.3040 0.686 0.777 8.699 12.639 10.945 j 0.3200 0.640 0.797 10.935 14.835 11.230 1 0.3360 0.590 0.817 13.248 17.094 11.507 0.3520 0.537 0.836 15.586 19.354 11.778 0.3680 0.480 0.855 17.891 21.550 12.043 3 I .0.3840 0.420 0.873 20.106 23.610 12.302 !.I 0.4000 0.357 0.891 22.166 25.459 12.555 0.4160 0.291 0.909 24.005 27.018 12.804 0.4320 0.222 0.926 25.552 28.202 13.048 ,e i 0.4480 0.150-0.943 26.734 28.926 13.287 0.4640 0.075 0.960 27.474 29.097 13.522 0.4800 -0.002 0.976 27.693 28.621 13.754 ij 0.4960 -0.082 0.993 27.309 27.400 13.981 l 0.5120 -0.165 1.009 26.237 25.333 14.205 .3 0.5280 -0.250 1.024 24.388 22.313 14.425 0.5440 -0.337 1.040 21.672 18.233 14.642 j., 0.5600 -0.427 1.055 17.996 12.982 14.855 0.5760 -0.519 1.070 13.265 6.445 15.066 0.5920 -0.613 1.084 7.379 -1.495 15.274 5 0.6080 -0.710 1.099 0.239 -10.958 15.479 1 0.6240 -0.809 1.113 -8.259 -22.068 15.681 0.6400 -0.910 1.128 -18.219 -34.949 15.881 0.6560 -1.013 1.142 -29.750 -49.732 16.078 0.6720 -1.118 1.155 -42.961 -66.547 16.273 0.6880 -1.226 1.169 -57.964 -85.528 16.466 i 0.7040 -1.335 1.183 -74.874 -106.812 16.656 0.7200 -1.447 1.196- -93.805 -130.538 16.845 O.7360 -1.560 1.209- -114.878 -156.847 17.031 i 0.7520 -1.675 1.222 -138.211 -185.884 17.215 0.7680 -1.792 1.235 -163.928 -217.794 17.397 0.7840 -1.911 1.248 -192.152 -252.727 17.577 0.8000 -2.032 1.261 -223.011 -290.834 17.756 i ~ 5, TIME KMAX DA/DT DA A A/THK 1000.0 3.71 6.093E-08 0.0001 0.0751 0.000 2000.0 3.70 6.078E-08 0.0001 0.0751 0.000 3000.0 3.70 6.064E-08 0.0001 0.0752 0.000 _i 4000.0 3.69 6.049E-08 0.0001 0.0752 0.000 5000.0 3.69 6.035E-08 0.0001 0.0753 0.000 6000.0 3.69 6.021E-08 0.0001 0.0754 0.000 3}j 7000.0 3.68 6.006E-08 0.0001 0.0754 0.000

E 8000.0 3.68 5.992E-08 0.0001 0.0755 0.000 9000.0 3.67 5.978E-08 0.0001 0.0755 0.000 10000.0 3.67 5.964E-08 0.0001 0.0756 0.000

,.} 11000.0 3.67 5.950E-08 0.0001 0.0757 0.000 .4 12000.0 3.66 5.936E-08 0.0001 0.0757 0.000 13000.0 3.66 5.922E-08 0.0001 0.0758 0.000 14000.0 3.65 5.908E-08 0.0001 0.0758 0.000 'f 15000.0 3.65 5.894E-08 0.0001 0.0759 0.000 '.J 16000.0 3.65 5.880E-08 0.0001 0.0760 0.000 17000.0 3.64 5.867E-08 0.0001 0.0760 0.000 18000.0 3.64 5.853E-08 0.0001 0.0761 0.000 .f. 19000.0 3.64 5.839E-08 0.0001 0.0761 0.000 3 20000.0 3.63 5.826E-08 0.0001 0.0762 0.000 21000.0 3.63 5.812E-08 0.0001 0.0762 0.000 22000.0 3.62 5.799E-08 0.0001 0.0763 0.000 h StructuralIntegrityAssociates,Inc. 51R-95-105, Rev. O B-3

3 23000.0 3.62~ 5.785E 0.0001 0.0764 0.000 24000.0 3.62 5.772E-08 0.0001 0.0764 0.000 25000.0 3.61 5.758E-08 0.0001. 0.0765 0.000 26000.0 3.61 5.745E-08 0.0001 0.0765 0.000 l 27000.0 3.60 5.732E-08' O.0001 0.0766 0.000 j l-28000.0= 3.60 5.719E-08 0.0001 0.0767 0.000 29000.0 3.60 5.705E-08 0.0001 0.0767 0.000 30000.0 3.59 5.692E-08 0.0001 0.0768 0.000 j i 31000.0 3.59 5.679E-08 0.0001 0.0768 0.000-i- 32000.0-3.58 5.666E-08. 0.0001 0.0769 0.000 33000.0 3.58 5.653E-08 0.0001 0.0769 0.000

i l

34000.0 3.58 5.640E-08 0.0001 0.0770 0.000 35000.0 3.57 5.627E-08 0.0001 0.0770 0.000 36000.0 3.57 5.615E-08 0.0001 0.0771 0.000-37000.0 3.57 5.602E-08 0.0001 0.0772 0.000 j 38000.0 3.56 5.589E-08 0.0001 0.0772 0.000 39000.0 3.56 5.576E-08 0.0001 0.0773 0.000 40000.0 3.55 5.564E-08 0.0001 0.0773 0.000 41000.0 3.55 5.551E-08 0.0001 0.0774 0.000 m l 42000.0 3.55 5.538E-08 0.0001 0.0774 0.000 43000.0 3.54 5.526E-08 0.0001 0.0775 0.000 44000.0 3.54 5.513E-08 0.0001 0.0776 0.000 45000.0 3.54 5.501E-08 0.0001 0.0776 0.000 i-46000.0 3.53 5.489E-08 0.0001 0.0777 0.000 2 47000,0 3.53 5.476E-08 0.0001 0.0777 0.000 48000.0 3.52 E.464E-08 0.0001 0.0778 0.000 i 49000.0 3.52 5.452E-08 0.0001 0.0778 0.000 50000.0 3.52 5.439E-08 0.0001 0.0779 0.000 ) .51000.0 3.51 5.427E-08 0.0001 0.0779 0.000 52000.0 3.51 5.415E-08 0.0001 0.0780 0.000 53000.0 3.51 5.403E-08 0.0001 0.0780 0.000 54000.0 3.50 5.391E-08 0.0001 0.0781 0.000 ,' l + 55000.0 3.50 5.379E-08 0.0001 0.0781 0.000 l 56000.0 3.50 5.367E-08 0.0001 0.0782 0.000 l., 57000.0 3.49 5.355E-08 0.0001 0.0783 0.000 58000.0 3.49 5.343E-08 0.0001 0.0783 0.000

• ~

59000.0 3.49 5.331E-08 0.0001 0.0784 0.000 1 60000.0 3.48 5.319E-08 0.0001 0.0784 0.000 61000.0 3.48 5.307E-08 0.0001 0.0785 0.000 J

l 62000.0 3.47 5.296E-08 0.0001 0.0785 0.000 II 63000.0 3.47 5.284E-08 0.0001 0.0786 0.000

) j-64000.0 3.47 5.272E-08 0.0001 0.0786 0.000 65000.0 3.46 5.261E-08 0.0001 0.0787 0.000 l 66000.0 3.46 5.249E-08 0.0001 0.0787 0.000 J 67000.0 3.46 5.238E-08 0.0001 0.0788 0.000 68000.0 3.45 5.226E-08 0.0001 0.0788 0.000 69000.0 3.45. 5.215E-08 0.0001 0.0789 0.000 70000.0 3.45 5.203E-08 0.0001 0.0789 0.000 ! m 71000.0 3.44 5.192E-08 0.0001 0.0790 0.000 I 72000.0 3.44 5.180E-08 0.0001 0.0790 0.000 j 73000.0 3.44 5.169E-08 0.0001 0.0791 0.000 ]- 74000.0 3.43 5.158E-08 0.0001 0.0791 0.000 l 75000.0 3.43 5.147E-08 0.0001 0.0792 0.000 76000.0 3.43 5.135E-08 0.0001 0.0793 0.000 77000.0 3.42 5.124E-08 0.0001 0.0793 0.000 2 78000.0 3.42 5.113E-08 0.0001 0'.0794 0.000 ) 1 79000.0 3.41 5.102E-08 0.0001 0.0794 0.000 i ) t 80000.0 3.41 5.091E-08 0.0001 0.0795 0.000 j 81000.0 3.41 5.080E-08 0.0001 0.0795 0.000 j - 82000.0 3.40 5.069E-08 0.0001 0.0796 0.000 a 83000.0 3.40 5.058E-08 0.0001 0.0796 0.000 J-84000.0 3.40 5.047E-08 0.0001 0.0797 0.000 85000.0 3.39 5.036E-08 0.0001 0.0797 0.000 4 f StructuralIntegrityAssociates,Inc. ) SIR-95-105, Rev. O B-A L

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l-338000.0 2.73 3.133E-08 0.0000 0.0897 0.000 339000.0 2.72 3.128E-08 0.0000 0.0898 0.000 340000.0 2.72 3.123E-08 0.0000 0.0898 0.000 341000.0 2.72 3.118E-08 0.0000 0.0898 0.000 342000.0 2.72 3.113E-08 0.0000 0.0899 0.000 343000.0 2.71 3.107E-08 0.0000 0.0899 0.000 344000.0 2.71 3.102E-08 0.0000 0.0899 0.000 m 345000.0 2.71 3.097E-08 0.0000 0.0900 0.'000 JI 343000.0 2.71 3.092E-08 0.0000 0.0900 0.000 347000.0 2.71 3.087E-08 0.0000 0.0900 0.000 348000.0 2.70 3.082E-08 0.0000 0.0901 0.000 p 349000.0 2.70 3.077E-08 0.0000 0.0901 0.000 350000.0 2.70 3.072E-08 0.0000 0.0901 0.000 351000.0 2.70 3.067E-08 0.0000 0.0901 0.000 r7 352000.0 2.70 3.062E-08 0.0000 0.0902 0.000 353000.0 2.69 3.057E-08 0.0000 0.0902 0.000 354000.0 2.69 3.052E-08 0.0000 0.0902 0.000 1 355000.0 2.69 3.047E-08 0.0000 0.0903 0.000 356000.0 2.69 3.042E-08 0.0000 0.0903 0.000 357000.0 2.69 3.037E-08 0.0000 0.0903 0.000 358000.0 2.68 3.032E-08 0.0000 0.0904 0.000 s 359000.0 2.68 3.027E-08 0.0000 0.0904 0.000 j 360000.0 2.68 3.022E-08 0.0000 0.0904 0.000 361000.0 2.68 3.017E-08 0.0000 0.0904 0.000 362000.0 2.68 3.012E-08 0.0000 0.0905 0.000 363000.0 2.67 3.008E-08 0.0000 0.0905 0.000 364000.0 2.67 3.003E-08 0.0000 0.0905 0.000 365000.0 2.67 2.998E-08 0.0000 0.0906 0.000 366000.0 2.67 2.993E-08 0.0000 0.0906 0.000 367000.0 2.67 2.988E-08 0.0000 0.0906 0.000 368000.0 2.66 2.983E-08 0.0000 0.0907 0.000 l 369000.0 2.66 2.979E-08 0.0000 0.0907 0.000 1 370000.0 2.66 2.974E-08 0.0000 0.0907 0.000 371000.0 2.66 2.969E-08 0.0000 0.0907 0.000 372000.0 2.66 2.964E-08 0.0000 0.0908 0.000 373000.0 2.65 2.959E-08 0.0000 0.0908 0.000 374000.0 2.65 2.955E-08 0.0000 0.0908 0.000 375000.0 2.65 2.950E-08 0.0000 0.0909 0.000 376000.0 2.65 2.945E-08 0.0000 0.0909 0.000 ] 377000.0 2.65 2.941E-08 0.0000 0.0909 0.000 378000.0 2.64 2.936E-08 0.0000 0.0910 0.000 379000.0 2.64 2.931E-08 0.0000 0.0910 0.000 380000.0 2.64 2.927E-08 0.0000 0.0910 0.000 } 381000.0 2.64 2.922E-08 0.0000 0.0910 0.000 2 382000.0 2.64 2.917E-08 0.0000 0.0911 0.000 383000.0 2.63 2.913E-08 0.0000 0.0911 0.000 384000.0 2.63 2.908E-08 0.0000 0.0911 0.000 j 385000.0 2.63 2.903E-08 0.0000 0.0912 0.000 S 386000.0 2.63 2.899E-08 0.0000 0.0912 0.000 l 387000.0 2.63 2.894E-08 0.0000 0.0912 0.000 l 388000.0 2.62 2.890E-08 0.0000 0.0912 0.000 389000.0 2.62 2.885E-08 'O.0000 0.0913 0.000 2 390000.0 2.62 2.880E-08 0.0000 0.0913 0.000 391000.0 2.62 2.876E-08 0.0000 0.0913 0.000 392000.0 2.62 2.871E-08 0.0000 0.0914 0.000 393000.0 2.62 2.867E-08 0.0000 0.0914 0.000 j .I 394000.0 2.61 2.862E-08 0.0000 0.0914 0.000 395000.0 2.61 2.858E-08 0.0000 0.0914 0.000 396000.0 2.61 2.853E-08 0.0000 0.0915 0.000 n 397000.0 2.61 2.849E-08 0.0000 0.0915 0.000 JL 398000.0 2.61 2.844E-08 0.0000 0.0915 0.000 3 399000.0 2.60 2.840E-08 0.0000 0.0916 0.000 400000.0 2.60 2.835E-08 0.0000 0.0916 0.000 d i SIR-95-105, Rev. O B-9 Structural Integrity Associates, Inc.

jf'g, ~. v 401000.0 2.60 2.831E-08 0.0000 0.0916 0.000 402000.0 2.60 2.826E-08 0.0000 0.0916 0 000 403000.0 2.60 2.822E-08 0.0000 0.0917 0.000 404000.0. 2.59 2.818E-08 0.0000 0.0917 0.000 405000.0 2.59 -2.813E-08 0.0000 0.0917 0.000 406000.0 2.59 2.809E-08 0.0000 0.0918 0.000 407000.0 2.59 2.804E-08 0.0000 0.0918 0.000 408000.0 2.59 2.800E-08 0.0000 .0.0918 0.000 409000.0 2.59 2.796E-08 0.0000 0.0918 0.000 410000.0 2.58 2.791E-08 0.0000 0.0919 0.000 411000.0 2.58 2.787E-08 0.0000 0.0919 0.000 412000.0 2.58 2.783E-08 0.0000 0.0919 0.000 413000.0 2.58 2.778E-08 0.0000 0.0920 0.000 414000.0 2.58 2.774E-08 0.0000 0.0920 0.000 415000.0 2.57 2.770E-08 0.0000 0.0920 0.000 416000.0 2.57 '2.766E-08 0.0000 0.0920 0.000 417000.0 2.57 2.761E-08 0.0000 0.0921 0.000 418000.0 2.57 2.757E-08 0.0000 0.0921 0.000 419000.0 2.57 2.753E-08 0.0000 0.0921 0.000 420000.0 2.56 2.748E-08 0.0000 0.0921 0.000 421000.0 2.56 2.744E-08 0.0000 0.0922 0.000 422000.0 2.56 2.740E-08 0.0000 0.0922 0.000 423000.0 2.56 2.736E-08 0.0000 0.0922 0.000 424000.0 2.56 2.732E-08 0.0000 0.0923 0.000 425000.0 2.56 2.727E-08 0.0000 0.0923 0.000 426000.0 2.55 '2.723E-08 0.0000 0.0923 0.000 427000.0 2.55 2.719E-08 0.0000 0.0923 0.000 428000.0 2.55 2.715E-08 0.0000 0.0924 0.000 429000,0 2.55 2.711E-08 0.0000 0.0924 0.000 430000.0 2.55 2.707E-08 0.0000 0.0924 0.000 431000.0 2.54 2.702E-08 0.0000 0.0924 0.000 432000.0 2.54 2.698E-08 0.0000 0.0925 0.000 433000.0 2.54 2.694E-08 0.0000 0.0925 0.000 434000.0 2.54 2.690E-08 0.0000 0.0925 0.000 435000.0 2.54 2.686E-08 0.0000 0.0926 0.000 436000.0 2.54 2.682E-08 0.0000 0.0926 0.000 437000.0 2.53 2.678E-08 0.0000 0.0926 0.000 438000.0 2.53 2.674E-08 0.0000 0.0926 0.000 439000.0 2.53 2.670E-08 0.0000 0.0927 0.000 440000.0 2.53 2.666E-08 0.0000 0.0927 0.000 441000.0 2.53 2.662E-08 0.0000 0.0927 0.000 442000.0 2.53 2.658E-08 0.0000 0.0927 0.000 -5 443000.0 2.52 2.654E-08 0.0000 0.0928 0.000 1 444000.0 2.52 2.650E-08 0.0000 0.0928 0.000 3 445000.0 2.52 2.646E-08 0.0000 0.0928 0.000 446000.0 2.52 2.642E-08 0.0000 0.0928 0.000 447000.0 2.52 2.638E-08 0.0000 0.0929 0.000 448000.0. 2.51 2.634E-08 0.0000 0.0929 0.000 449000.0 2.51 2.630E-08 0.0000 0.0929 0.000 450000.0 2.51 2.626E-08 0.0000 0.0930 0.000 451000.0 2.51 2.622E-08 0.0000 0.0930 0.000 j 452000.0 2.51 2.618E-08 'O.0000 0.0930 0.000 453000.0 2.51 2.614E-08 0.0000 0.0930 0.000 454000.0 2.50 .2.610E-08 0.0000 0.0931 0.000 455000.0 2.50 2.606E-08 0.0000 0.0931 0.000 >1 456000.0 2.50 2.602E-08 0.0000 0.0931 0.000 j a; 457000.0 2.50 2.598E-08 0.0000 0.0931 0.000 458000.0 2.50 2.594E-08 0.0000 0.0932 0.000 459000.0 2.50 2.590E-08 0.0000 0.0932 0.000 460000.0 2.49 2.587E-08 0.0000 0.0932 0.000 j ]' l 461000.0 2.49 2.583E-08 0.0000 0.0932 0.000 462000.0-2.49 2.579E-08 0.0000 0.0933 0.000 it 463000.0 2.49 2.575E-08 0.0000 0.0933 0.000 f StructuralIntegrityAssociates,Inc. SIR-95-105, Rev. O B-10

.h*, I 7 i I 464000.0 2.49 2.571E-08 0.0000 0.0933 0.000 465000.0 2.49 2.567E-08 0.0000 0.0933 0.000 466000.0 2.48 2.564E-08 0.0000 0.0934 0.000 l 467000.0 2.48 2.560E-08 0.0000 0.0934 0.000 468000.0 2.48 2.556E-08 0.0000 0.0934 0.000 469000,0 2.48 2.552E-08 0.0000 0.0934 0.000 470000.0 2.48 2.548E-08 0.0000 0.0935 0.000 471000.0 2.48 2.545E-08 0.0000 0.0935 0.000 472000.0 2.47 2.541E-08 0.0000 0.0935 0.000 473000.0 2.47 2.537E-08 0.0000 0.0935 0.000 ry 474000.0 2.47 2.533E-08 0.0000 0.0936 0.000 3 475000.0 2.47 2.530E-08 0.0000 0.0936 0.000 'I 476000.0 2.47 2.526E-08 0.0000 0.0936 0.000 477000.0 2.46 2.522E-08 0.0000 0.0936 0.000 478000.0 2.46 2.518E-08 0.0000 0.0937 0.000 479000.0 2.46 2.515E-08 0.0000 0.0937 0.000 480000.0 2.46 2.511E-08 0.0000 0.0937 0.000 481000.0 2.46 2.507E-08 0.0000 0.0937 0.000 482000.0 2.46 2.504E-08 0.0000 0.0938 0.000 483000.0 2.45 2.500E-08 0.0000 0.0938 0.000 484000.0 2.45 2.496E-08 0.0000 0.0938 0.000 485000.0 2.45 2.493E-08 0.0000 0.0938 0.000 486000.0 2.45 2.489E-08 0.0000 0.0939 0.000 487000.0 2.45 2.485E-08 0.0000 0.0939 0.000 488000.0 2.45 2.482E-08 0.0000 0.0939 0.000 489000.0 2.44 2.478E-08 0.0000 0.0939 0.000 490000.0 2.44 2.475E-08 0.0000 0.0940 0.000 491000.0 2.44 2.471E-08 0.0000 0.0940 0.000 492000.0 2.44 2.467E-08 0.0000 0.0940 0.000 493000.0 2.44 2.464E-08 0.0000 0.0940 0.000 494000.0 2.44 2.460E-08 0.0000 0.0941 0.000 495000.0 2.44 2.457E-08 0.0000 0.0941 0.000 i 496000.0 2.43 2.453E-08 0.0000 0.0941 0.000 497000.0 2.43 2.450E-08 0.0000 0.0941 0.000 498000.0 2.43 2.446E-08 0.0000 0.0942 0.000 j 499000.0 2.43 2.442E-08 0.0000 0.0942 0.000 500000.0 2.43 2.439E-08 0.0000 0.0942.0.000 .a } w-em STR-95-105, Rev. O B.I1 StructuralIntegrity Associates, Inc.

i- ,f 7 k tm pc-CRACK (C) COPYRIGHT 1984, 1990 STRUCTURAL INTEGRITY ASSOCIATES, INC. SAN JOSE, CA (408)978-8200 VERSION 2.1 y 3 Date: 25-Sept-1995 Times 12:57:47.66 m STRESS CORROSION CRACK GROWTH ANALYSIS a - WSI-20Q: SEQUOYAH CRDM LOWER CSW OVERLAY, SECTION D INITIAL CRACK SIZE = 0.0750 MAX CRACK SIZE FOR SCCG= 0.8000 STRESS CORROSION CRACK GROWTH LAW LAW ID C N Kthres K1C NRC/10 3.590E-09 2.1610 0.0000 200.0000 STRESS COEFFICIENTS CASE ID CO C1 C2 C3 BEND 1.7128 ~5.8723 0.0000 0.0000 MEMBRANE 0.7100 0.0000 0.0000 0.0000 l SECC 24.9810 -618.1000 3314.0000 -4183.0000 SECD 40.9851 -788.8973 3968.8642 -5013.4105 APPLIED 10.0000 0.0000 0.0000 0.0000 1 Kmax CASE ID SCALE FACTOR SECD 1.0000 APPLIED 0.2000 3 TIME PRINT TIME INCREMENT INCREMENT 500000.0 1000.0 1000.0 crack model: CONTINUOUS SURFACE CRACK IN HALF SPACE .) CRACK ---------------STRESS INTENSITY FACTOR---------------- SIZE CASE CASE CASE CASE CASE BEND MEMBRANE SECC SECD APPLIED j 0.0160 0.416 0.178 4.861 8.481 2.511 O.0320 0.568 0.252 5.145 9.756 3.551 0.0480 0.670 0.309 4.494 9.578 4.349 a 0.0640 0.745 0.357 3.441 8.728 5.022 0.0800 0.801 0.399 2.249 7.578 5.615 0.0960 0.843 0.437 1.076 6.348 6.151 0.1120 0.872 0.472 0.031 5.187 6.644 1 0.1280 0.892 0.504 -0.812 4.195 7.102 0.1440 0.903 0.535 -1.403 3.443 7.533 0.1600 0.906 0.564 -1.709 2.977 7.941 0.1760 0.903 0.591 -1.708 2.828 8.328 9 f 0.1920 0.893 0.618 -1.391 3.010 8.698 0.2080 0.878 0.643 -0.760 3.526 9.054 0.2240 0.858 0.667 0.179 4.368 9.395 0.2400 0.832 0.690 1.407 5.522 9.725 i 1 f StructuralIntegrityAssociates,Inc. SIR-95-105, Rev. O B-12

P,*. b 0.2560 0.802 0.713 2.905 6.961 10.044 0.2720 0.767 0.735 4.642 8.653 10.353 0.2880 0.729 0.756 6.587 10.561 10.653 0.3040 0.686 0.777 8.699 12.638 10.945 4 0.3200 0.640 0.797 10.935 14.835 11.230 0.3360 0.590 0.817 13.248 17.094 11.507 0.3520 0.537 0.836 15.586 19.354 11.778 ]- 0.3680 0.480 0.855 17.891 21.550 12.043 l 0.3840 0.420 0.873 20.106 23.610 12.302 0.4000 0.357 0.891 22.166 25.459 12.555 0.4160 0.291 0.909 24.005 27.018 12.804 0.4320 0.222 0.926 25.552 28.202 13.048 0.4480 0.150 0.943 26.734 28.926 13.287 0.4640 0.075 0.960 27.474 29.097 13.522 0.4800 -0.002 0.976 27.693 28.621 13.754 0.4960 -0.082 0.993 27.309 27.400 13.981 i 0.5120 -0.165 1.009 26.237 25.333 14.205 0.5280 -0.250 1.024 24.388 22.313 14.425 0.5440 -0.337 1.040 21.672 18.233 14.642-l 0.5600 -0.427 1.055 17.996 12.982 14.855 I 0.5760 -0.519 1.070 13.265 6.445 15.066 0.5920 -0.613 1.084 7.379 -1.495 15.274 i 0.6080 -0.710 1.099 0.239 -10.958 15.479 0.6240 -0.809 1.113 -8.259 -22.068 15.681 0.6400 -0.910 1.128 -18.219 -34.949 15.881 0.6560 -1.013 1.142 -29.750 -49.732 16.078 O.6720 -1.118 1.155 -42.961 -66.547 16.273 0.6880 -1.226 1.169 -57.964 -85.528 16.466 0.7040 -1.335 1.183 -74.874 -106.812 16.656 0.7200 -1.447 1.196 -93.805 -130.538 16.845 0.7360 -1.560 1.209 -114.878 -156.847 17.031 0.7520 -1.675 1.222 -138.211 -185.884 17.215 0.7680 -1.792 1.235 -163.928 -217.794 17.397 0.7840 -1.911 1.248 -192.152 -252.727 17.577 0.8000 -2.032 1.261 -223.011 -290.834 17.756 ] TIME KMAX DA/DT DA A A/THK 1000.0 9.02 4.165E-07 0.0004 0.0754 0.000 2000.0 9.00 4.138E-07 0.0004 0.0758 0.000 .2 3000.0 8.97 4.112E-07 0.0004 0.0762 0.000 4000.0 8.94 4.086E-07 0.0004 0.0767 0.000 5000.0 8.92 4.060E-07 0.0004 0.0771 0.000 j 6000.0 8.89 4.034E-07 0.0004 0.0775 0.000 E 7000.0 8.86 4.008E-07 0.0004 0.0779 0.000 8000.0 8.84 3.983E-07 0.0004 0.0783 0.000 9000.0 8.81 3.958E-07 0.0004 0.0787 0.000 } 10000.0 8.79 3.933E-07 0.0004 0.0790 0.000

5 11000.0 8.76 3.909E-07 0.0004 0.0794 0.000 12000.0 8.74 3.885E-07 0.0004 0.0798 0.000 13000.0 8.71 3.861E-07 0.0004 0.0802 0.000

] 14000.0 8.69 3.836E-07 0.0004 0.0806 0.000 15000.0 8.66 3.810E-07 0.0004 0.0810 0.000 a 16000.0 8.63 3.785E-07 0.0004 0.0814 0.000 17000.0 8.61 3.760E-07 0.0004 0.0817 0.000 f 18000.0 8.58 3.735E-07 0.0004 0.0821 0.000 .3 '19000.0 8.55 3.710E-07 0.0004 0.0825 0.000 20000.0 8.53 3.686E-07 0.0004 0.0828 0.000 21000.0 8.50 3.662E-07 0.0004 0.0832 0.000 ,1, f StructuralIntegrity Associates, Inc. STR-95-1OS, Rev, o B-13 )

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