ML20207S726

Rev 3 to Technical Justification for Continued Operation of Ja Fitzpatrick Nuclear Power Plant W/Existing Recirculation Sys Piping

ML20207S726
Person / Time
Site:	FitzPatrick
Issue date:	01/13/1987
From:	Gustin H, Riccardelli P STRUCTURAL INTEGRITY ASSOCIATES, INC.
To:
Shared Package
ML20207S724	List: JPN-87-008, Forwards Technical Justification for Continued Operation of Ja Fitzpatrick Nuclear Power Plant W/Existing Recirculation Sys Piping. Weld Overlay Repairs Meet Guidelines of Generic Ltr 84-11 & Draft NUREG-0313,Rev 2 ML20207S726
References
SIR-86-008, SIR-86-008-R03, SIR-86-8, SIR-86-8-R3, NUDOCS 8703200218
Download: ML20207S726 (122)

Text

{{#Wiki_filter:r 4 d Report No. SIR-86-008 Revision 3 51 Project No. PASNY-10 January 13, 1987 s TECHNICAL JUSTIFICATION FOR CONTINUED OPERATION OF JAMES A. FITZPATRICK NUCLEAR POWER PLANT WITH EXISTING RECIRCULATION SYSTEM PIPING wm Prepared by Structural Integrity Associates Prepared for New York Power Authority 4l -e / Date: /-l3-8 f Prepared by: ~ H. L. ustin / Reviewed and '/ [ 1[M%7 Approved by: 8'[f l/1[ Date: i d P. C. Riccardella hoh33 DOC P pon M ASSOCIATESINC

t SIR-86-008 REVISION CONTROL SHEET 9 SECTION PARAGRAPH (S) DATE REVISION REMARKS 4 i - vi all 6-5-86 0 Initial Issue 1 1-3 ~! 2 2-61 3 3-36 4 4-8 5 5-2 6 6-3 all 6-5-86 0 Initial Issue ii - ix all 8-22-86 1 Incorporated Client 1 1-3 Comments 'I. 2 2-60 3 3-37 i 4 4 8 5 5-3 6 6-2 all 8-22-86 1 Incorporated Client Conments 11 - ix all 10-20-86 2 Incorporated Client fr 1 1-3 Comments 'j 2 2 57 11 3 3-37 4 4-8 5 5-3 6 6-2 all 10-20-86 2 Incorporated Client Comments N-L pgs. ii-iv all 1-13-87 3 Revised to incorporate client comments. p.. all 1-13-87 Updated Table of Contents v 2-1 57 all 1-13-87 Revised to address use of B31.1 Allowable Stress Values. 3 Added Section 2.7 (pg. 2-20) L 4 4 i

i v EXECUTIVE

SUMMARY

Flaws believed to be attributable to intergranular stress corrosion cracking (IGSCC) have been observed at 9 locations in the recirculation system of the James A. FitzPatrick Nuclear Power Plant. Weld overlay repairs were applied to 6 of these locations. The remaining 3 locations contained only very small indications, which were treated with the induction heating stress improve-ment (IHSI) process, and shown to be acceptable by analysis. i This report provides documentation of the technical justification for continued long term operation of the FitzPatrick recirculation system without replacement or further repair of these flawed locations. The justification for continued operation of the weld overlay-repaired locations is based upon experimental and analytical studies which demonstrate the l following items: ul 1. The weld metal used for overlays (308L) is highly resistant to IGSCC,so degradation of the repair due to crack propagation into the overlay is not expected to occur. hi 2. The residual stress distribution which results from weld overlay application is highly compressive on the inside surface of re-paired piping. This will inhibit further crack propagation and initiation of new cracks in the original pipe weld. 3. The overlays are shown to meet recent NRC guidance (as sumarized in NUREG-0313, Revision 2 Draft), issued subsequent to the time of the FitzPatrick repairs. 7 4. Flaws can be effectively monitored beneath an overlay using ultra-sonic inspection techniques developed recently by EPRI. 5. Although not addressed by this report, New York Power Authority intends to implement a hydrogen water chemistry system in 1987. ii

e The system will result in recirculation water chemistry which is expected to be very effective in arresting future IGSCC. This report documents review of the original weld overlay design bases, which included consideration of the ratio of the sum of pressure, deadweight, seismic and thermal expansion stresses to the B31.1 Code allowable stress values. A re-evaluation of the as-built weld overlay designs to current criteria is also included. Current criteria do not require consideration of thermal expansion stresses in design of standard overlays (as defined in Reference 3) such as are applied at FitzPatrick. Evaluation to current criteria results in demonstration of greater margin on n-built overlay thickness, compared to the original design criteria. Because it may be necessary to mechanically smooth the surf ace of the existing weld overlays to enhance inspectability, the new criteria-required thicknesses will be taken as the acceptance criteria for re-evaluation of weld overlays following surface finishing. The justification for continued operation with the three IHSI-treated flaws is based upon the following arguments: 1. The IHSI process produces a strongly compressive residual stress field on the inside portion of treated pipes. This results in com-plete arrest of shallow flaws cr.d inhibition of new flaw initi-ation. 1 P Q 2. The flaws which were treated with IHSI are very short and shallow, ( and are acceptable by a large margin with respect to ASME Section XI standards. They also satisfy the recent NRC position on IHSI of P l flawed pipe welds contained in NUREG 0313 Revision 2 (Draft). 3. The implementation of hydrogen water chemistry will make further propagation of the flaws even more unlikely. In addition to the technical arguments in support of continued operation, recent industry experience in receiving licensing approval for continued operation with weld overlays and IHSI-treated small shallow flaws has been iii

)e o favorable. This experience is reflected in NUREG-0313, Rev. 2 (Draft). Consequently, continued operation of FitzPatrick with its recirculation system in the present configuration is technically justified. 2 6 b L. Il I. l l l iv i i i

a TABLE OF CONTENTS -Section Page 1.0 Introduction 1-1 i

1.1 Background

1-1 1.2 U.S. Regulatory Position on Weld Overlay Life Extension. 1-1 1.3 Purpose of Report 1-2 2.0 Description of Repairs and Flaw Evaluations 2-1 2.1 Flaw Indications 2-1 2.2 Weld Overlay Repairs 2-2 7 3 2.2.1 Weld Overlay Design Methodology 2-2 2.2.2 Weld Overlay Designs. 2-3 2.3 Application Recommendations. 2-6 2.4 As-Built Weld Overlays 2-7 2.5 Shrinkage Stresses. 2-7 2.6 Flaw Evaluation. 2-9 2.6.1 Flaw Evaluation Methodology 2-9 2.6.2 Evaluations and Results. 2-15 g 2.7 Fatigue Evaluation of Flaws Under Weld Overlays 2-20 3.0 Discussion of Major Technical Issues 3-1 1 ps 3.1 Weld Metal IGSCC Resistance. 3-1 3.1.1 Field Experience 3-2 3.1.2 Laboratory Experience 3-4 3.1.3 Modelling Studies. 3-8 3.1.4 Model Results 3-10 3.2 Residual Stress Benefits. 3-11 .a 3.2.1 Georgia Power Company / Structural Integrity Associates (SIA)/ Welding Services Incorporated (WSI) 28-Inch Notched Pipe Test. 3-11 3.2.2 EPRI/GE Residual Stress Results 3-13 3.2.3 Nutech/ Georgia Power Company 12-Inch Weld Overlay Mockups 3-14 3.2.4 EPRI/J.A. Jones 24-Inch Weld Overlay Mockup 3-14 3.2.5 EPRI/BWROG II Pipe Tests 3-15 3.2.6 Destructive Assay of Hatch Unit 2 Overlay Specimens at Argonne National Laboratory (ANL) 3-15 I v

D e TABLE OF CONTENTS (continued) Section Page 3.3 Non-Destructive Examination. 3-16 3.3.1 Recent Developments at the EPRI NDE Center Workshop on Weld Overlay Inspections (RP1570-2). 3-17 3.3.2 Argonne National Laboratory (ANL) UT Inspection Workshops 3-18 4.0 Weld Metal Fracture Toughness. 4-1 4.1 Battelle/NRC Degraded Pipe Tests 4-2 4.1.1 Test Objectives 4-2 4.2 Application of Test Data to FitzPatrick Weld Overlays 4-3 4.3 Sumary 4-3 5.0 ' Sumary and Conclusions. 5-1 5.1 Sumary 5-1 5.2 Conclusions 5-2 6.0 References 6-1 Y 1 4 . } T t i I vi

LIST OF TABLES Table Page i 2-1 Weld Summary, James A. FitzPatrick - Results of IGSCC 2-21 Inspection 2-2 Sumary of Piping Stresses 2-22 23 Weld Overlay Sizing for Weld 12-12 2-23 2-4 Weld Overlay Sizing for Weld 12-23 2-24 2-5 Weld Overlay Sizing for Weld 12-64 2-25 2-6 Weld Overlay Sizing for Weld 12-69 2-26 2-7 Weld Overlay Sizing for Weld 12-70 2-27 2-8 Weld Overlay Sizing for Weld 22-22 2-28 2-9 As-Built Weld Overlay Dimensions 2-29 2-9 (Continued) Definition of Parameters 2-29a 2-10 Sumary of Shrinkage Stresses at Overlay Locations 2-30 2-11 Sumary of Shrinkage Stresses at Sweep-o-let Locations 2-31 2-12 Weld Overlay Sizing for Weld 12-69 Considering Shrinkage Effects.................... 2-32 2-13 Weld Overlay Sizing for Weld 12-70 Considering Shrinkage Effects.................... 2-32 2-14 Allowable End-of-Evaluation Period Flaw Depth to Thick-ness Ratio for Circumferential Flaws - Normal Operating f (Including Upset and Test) Conditions [From ASME Sec-4 tion XI, IWB-3640] 2-33 v. 2-15 Allowable End-of-Evaluation Period Flaw Depth to l Thickness Ratio for Circumferential Flaws in / Shielded Metal Arc and Submerged Arc Welds Normal Operating (Including Upset and Test) Conditions 2-34 3-1 Four Inch Pipe Weld Overlay Residual Stress Mockup Parameters................ 3-20 4-1 Sumary of Differences Between Battelle/NRC Weld Overlay Test Pipes and FitzPatrick Weld Overlays 4-4 i vi l l

LIST OF FIGURES Figure Page i 2-1 Isometric of Recirculation System - Loop A 2-35 2-2 Isometric of Recirculation System - Loop B 2-36 2-3 Recirculation System Piping Model - Loop A [Ref. 4].... 2-37 2-4 Recirculation System Piping Model - Loop B [Ref. 4].... 2-38 2-5 Weld Overlay Designs - 12 Inch Risers........... 2-39 ~ 2-6 Weld Overlay Design - Weld 22-22 2-40 1 2-7 Finite Element Model of Riser in Recirculation 2-41 2-8 (a) Through-Wall Distribution of Axial Residual Stress in a Large Diameter Weldment [Ref. 2] 2-42 2-8 (b) Assumed Through-Wall Welding Residual Stress Distri-bution in Small Diameter Weldments (<12 in.) [Ref. 2]... 2-42 2-9 Through-Wall Residual Stress Profile for a Welded and IHSI Treated 26-Inch Schedule 80 Pipe at a Cross-Section 0.12 Inch (0.3cm) from the Weld Centerline [Ref. 6] 2-43 r 2-10 Through-Wall Residual Stresses Computed at a Cross-Section in the Sensitized Zone, 0.12 Inch (0.3cm) from Weld Centerline, of a Welded and IHSI Treated 10-Inch Schedule 80 Pipe [Ref. 6] 2-44 2-11 Magnification Factors of Circumferential Crack in Cylinder (a/t=0.1).......................... 2-45 7 l 2-12 Stress Corrosion Crack Growth Rate Data for Sensitized ^ Stainless Steel in BWR Environment [Ref. 2]......... 2-46 i( 2-13 Common Assumptions Used to Estimate Circumferential Crack Growth 2-47 2 2-14 Average Effective Circumferential Crack Growth Rate as a [ Function of Operation Periods Used in Calculation of Time Between Inspection................... 2-48 2-15 Stress Intensity Factor Versus Crack Growth Depth for J. A. FitzPatrick Recirculation System Weld 12-4 2-49 2-16 Predicted Stress Corrosion Crack Growth for Observed Ultrasonic Flaw Indication, Weld 12-4............ 2-50 vit

LIST OF FIGURES (continued) Figure Page 2-17 Comparison of Predicted Crack Growth with Allowable Flaw Size Limits - Weld 12-4 2-51 2-18 Stress Intensity Factor Versus Crack Depth for J. A. FitzPatrick Recirculation System - Weld 28-53....... 2-52' 2-19 Predicted Stress Corrosion Crack Growth for Observed Ultrasonic Flaw Indication - Weld 28-53.......... 2-53 ) 2-20 Comparison of Predicted Crack Growth with Allowable Flaw Size Limits - Weld 28-53............... 2-54 2-21 Stress Intensity Factor Versus Crack Depth for J. A. FitzPatrick Recirculation System - Weld 28-112...... 2-55 2-22 Predicted Stress Corrosion Crack Growth for Observed Ultrasonic Flaw Indication - Weld 28-112 2-56 2-23 Comparison of Predicted Crack Growth with Allowable Flaw Size Limits - Weld 28-112.............. 2-57 3-1 Cracking in Weld Metal of NMP-1 Recirculation Line.... 3-21 3-2 Weld Metal Cracking in NMP-1 3-22 3-3 Subsurface Crack Present in Weld Metal in Quad Cities Core Spray Line...................... 3-23 1 3-4 Cracking Morphology of Bolt-Loaded WOL Specimen EA1 of 316L Stainless Steel 3-24 3-5 Cracking Morphology of Bolt-Loaded WOL Specimen EB1 of Type 316NG Stainless Steel 3-25 .} M 3-6 Cracking Morphology of Bolt-Loaded l WOL Specimen EB2 Type 316NG Stainless Steel........ 3-25 7 3-7 Cracking Morphology of Bolt-Loaded WOL Specimen EJ1 of Type 30A Stainless Steel 3-26 viii

b LIST OF FIGURES (continued) Figure Page 3-8 Weld Overlay Arrest of IGSCC Specimen RSP-14 [Ref. 19].. 3-27 Influence of N "-Y on the Intergranular Corrosion 3-9 L Behavior of Aged Samples of Wrought and Weld-Deposited Type 308 Stainless Steel. Open Symbols Indicate IGSCC per ASTM A262 Practice E Testing; Closed Symbols Indicate No IGSCC [Ref. 3-7]............... 3-28 3-10 Number of Intercepts of a Random Test Line with Austenite-Ferrite Boundaries per Unit Length of Test Line, N, Versus L ? Volume % Ferrite for Type 308 Compositions........ 3-29 3 M CL2 Test Set-Up Using 28" Pipe as Vessel......... 3-30 3-11 g f 3-12 Circumferential and Axial Notch Sizes and Locations (Bottom Half is Mirror Image of Top Half of Drawing) - Showing Stainless Steel Baffle to Permit Separate Testing with M Cl2 of Top and Bottom Halves of Pipe (Seal Against M C12 Fumes) 3-31 3-13 Metallographic Sections (100X) of Moderate Depth, Circumferential Notch Tips from GPC/51/WSI 28-Inch Notched Pipe Test [Ref. 22]................ 3-32 3-14 Through-Wall Residual Stresses [Ref. 23] 3-33 3-15 Through-Wall Residual Stresses [Ref. 23] 3-34 3-16 ID Stress for 24 Inch Overlay [Ref. 23].......... 3-35 3-17 Calculated Through Wall Stresses after the First and Second Weld Overlay Layers for a 24 Inch Pipe l. with 1.48 inch Wall. Overlay contains five weld layers for a total thickness of 0.35 inch) [Ref. 23]....... 3-36 i H 3-18 Calculated Through Wall Stresses after the Fifth and l Final Weld Overlay layer for a 24 inch Pipe with 1,48 Inch Wall. (Overlay contains five weld layers for a total thickness of 0.35 inch) [Ref. 23]....... 3-37 4-1 Illustration of Cracked Pipe and Weld Overlay Configuration Used in Battelle/U.S. NRC Experiments [Ref. 26]...... 4-5 4-2 Schematic Illustration of Test Setup Used in Battelle/ U.S. NRC Weld Overlay Experiments 4-6 1 l 4-3 Comparison of Recent Battelle/USNRC Degraded Piping Program Weld Overlay Tests With Overlay Design Basis Calculations. 4-7 l 4-4 Comparison of Fitzpatrick Design and As-Built Weld Overlay l Data to ASME Design Limits and Recent Pipe Failure Data.. 4-8 l ix ~

1.0 INTRODUCTION

1.1 Background

During the fall 1984 and spring 1985 outages at the James A. FitzPatrick Nuclear Power Plant, ultrasonic (UT) examination of the recirculation system piping identified indications at twelve weld joints. Several of these indications are believed to result from intergranular stress corrosion cracking (IGSCC). l Six of these welds have been repaired by application of weld overlays. The weld overlays were designed in accordance with recommendations of the U. S. Nuclear Regulatory Commission Generic Letter 84-11 [1]. As a result of the conservative design approach employed, the FitzPatrick weld overlays are considered sufficient to act as long term repairs of the flawed welds. Details of the weld overlay design and as-built dimensions are discussed in W Section 2 of this report. l Of the remaining indications, three were attributed to reflectors due to counterbore and weld root. Indications in the remaining three locations have been evaluated to demonstrate their acceptability in accordance with ASME Section XI requirements, supplemented by the reconnendations of NRC Generic Letter 84-11. The unrepaired welds were also treated by induction heating 't" stress improvement (IHSI) to inhibit further IGSCC propagation or initia-tion. 1.2 U.S. Regulatory Position on Weld Overlay Life Extension The published documents which contain the U.S. regulatory position for 7 repair of IGSCC fiaws in operating piants inciude Generic tetter 84-11 [1] and NUREG-1061 (Volumes 1 & 3) [2]. In addition, the draft of Revision 2 of NUREG-0313 [3] consolidates the regulatory position represented by the l previous documents, and includes recent regulatory actions and research results. 1-1 l

I i Although the position stated in Generic Letter 84-11 is that weld overlays on circumferential1y flawed pipe are acceptable for a maximum of two fuel cycles, the USNRC has recently been receptive to extension of that limit on a case by case basis. To date, two plants have received positive responses from the NRC on their plans for weld overlay life extension. These plants have as many as 35 weld overlays in place in their recirculation and associated piping systems. Some of these overlays will be going into operation for a third fuel cycle, and will have seen a total of five years of o service at the end of this cycle. 1 NUREG-0313, Revision 2 (Draft) provides new guidelines for crack evaluation 6' and repair, including specific provisions for extended use of weld overlay repairs [3). It recognizes that, with proper attention to overlay surface finish, effective ultrasonic inspection methods are now available which can detect cracking in or near the overlay. It permits long term operation of weld overlay repaired welds, with the provision that they be inspected ultrasonically, using qualified NDE examiners and procedures, every two I refueling cycles. Similar provisions will apply to unrepaired cracked welds, except that the inspection frequency is each refueling cycle, unless they have been treated with a stress improvement remedy such as IHSI. These provisions of NUREG-0313, Revision 2 more closely reflect what has been the NRC approach in case by case licensing reviews. 1.3 Purpose of Report d The purpose of the present report is to provide technical justification for p continued operation of the FitzPatrick recirculation system with the system M in the present condition (6 weld overlays and 3 flawed welds repaired with IHSI). The justification is based on a detailed review of weld overlay designs and fracture mechanics analyses for the observed indications at FitzPatrick. It is also based on analytical and experimental results developed by the industry regarding veld overlay resistance to IGSCC crack propagation, the residual stress benefits of weld overlay application in crack arrest, and the strength and toughness of weld overlay material. 1-2 I

Section 2 of this report summarizes the inspection results, reviews the weld overlay design analysis for the six repaired welds and flaw evaluations for the welds which were not repaired, and compares them to current practices and requirements. Section 3 discusses the pertinent technical issues involved in extended use of weld overlay repairs, including experimental demon-strations of the IGSCC resistance of weld metal, and residual stress analyses and experimental results for weld overlay repairs and IHSI remedies. Discussion of weld overlay strength and toughness considerations, including g recent test data from an NRC sponsored degraded pipe program at Battelle l Columbus Laboratories, is also presented in Section 4. 4 l (L. f i I 1-3 l l t

2.0 DESCRIPTION

OF REPAIRS AND FLAW EVALUATIONS 2.1 Flaw Indications In-service inspections of the FitzPatrick recirculation and associated stainless steel piping systems performed in 1983, 1984, and 1985 revealed indications in seven 12-inch diameter recirculation riser welds, one 22-inch end cap weld, and four 28-inch diameter welds. All unrepaired flaw locations identified in the earlier inspections were re-examined in 1985. All identified flaws in the FitzPatrick recirculation system have been weld 1 overlay repaired or treated with IHSI. Table 2-1 provides a weld-by-weld summary of these indications including indication sizes and corrective actions taken. All the indications are circumferential1y oriented, and have been conservatively assumed to be cracks for purposes of this evaluation. The weld locations containing indications are also identified N in the isometric drawings of Figures 2-1 and 2-2. Since all of the welds under discussion are in the recirculation system, an abbreviated weld numbering system, in which the system identification has oeen deleted from the weld numbers, is used in these figures and throughout the report. The original overlay designs and flaw evaluations for these indications are contained in [4]. These designs contain conservatisms which were prevalent at the time of Reference 4. Since that time, techniques have been slightly modified in current repair design and evaluation practice for indications b" of this type. In the following paragraphs, the original design and analysis methods of [4] are reviewed, and the results obtained for the FitzPatrick d welds are compared to what would be obtained by current, state-of-the-art methodology. In general, the comparison is very favorable. In addition, 7 the overlays have been determined to meet the regtirements of the recently issued NUREG-0313, Revision 2 (Draft) [3]. 2-1 i

1 I. 2.2 Weld Overlay Repairs As a result of deep and/or long UT indications, welds 12-12, 12-23, 12-64, 12-69, 12-70, and 22-22 were repaired by depositing weld overlay material 360 degrees around and to either side of the existing weld. Weld overlay repairs of this type have been used in operating BWRs to increase pressure boundary thickness, restore the original design basis structural margins of the piping, and produce a favorable compressive residual stress pattern in g cases where relatively large IGSCC-like indications have been found. T pe 308L weld metal with controlled (high) ferrite J IGSCC-resistant, 3 content was used for all overlay repairs at FitzPatrick, to prevent the propagation of the IGSCC into the repair itself. 2.2.1 Weld Overlay Design Methodology Piping stresses in the J. A. FitzPatrick recirculation system obtained from the General Electric Company (GE) stress report [5] were used in the weld overlay design calculations of [4]. Figures 2-3 and 2-4 illustrate the piping model used by GE. Stresses taken from [5] for the twelve affected c weld joints are summarized in Table 2-2. I Weld overlay thickness was determined iteratively using the applied stresses in Table 2-2 and the allowable flaw size table for wrought, l austenitic material from ASME Section XI, IWB-3640 [6] (Table 2-14 of this report). The allowable stress value used in this procedure was taken from p ANSI B31.1 [29], rather than using Sm from Section III of the ASME Code, as t-specified in Section XI. This leads to greater conservatism in weld overlay f. design. Minimum weld overlay length was chosen to be 1.5V1TCwhere R and t are outside radius and wall thickness of the pipe, respectively. It is noteworthy that, although not required by the ASME Section XI, thermal stresses were included in the original overlay design calculations to account for possible effects of lo~w toughness weld materials. Referring to Table 2-2, the stresses due to (pressure + deadioad + thermal expansion) were used for crack growth evaluations and the stresses due to (pressure + deadioad + thermal expansion + seismic) were used for weld overlay design 2-2 e m. . w y.=

y 4 - l;, jg + .] n / . n and critical-flaw < size evaluatio'ns. The stresses due to weld overlay e '}. shrinkagepre also' incl jded in tra'ck growtincOculations. Since the applicatien of FitzPatrick eld overlay repairt in 1984 ed 1985, the ASME Ssction XI rules for evalusti6a of flaws in.austenitic materish have been revised tc adhess ratt. $tidl iow toughness concerns with certaih weldtrent materials (Ninter, 1985 Addendum. [6])], '.This Code revision prov' ides more restricO a cliewabli flaw dze lim,its for' flux-shielded ). weldments I,snielded metal and st.bmerged arc ~ weldir.g). The nbw allov,able, flaw shes are based on elastic-plasti!: tearing analysit of the lower toughn3ss materials, including consideration of thermal expansion and other global secondary stress terms whi:h were excluded by the etiilier require-ments. However, the Fitzpitrick weld overlays were all applied using a gas-shielded (GTAW) welr' ig process, which is not si.bject to the low toughness concerns discussed above. The Winter,1985 Codi Addendum explicitly states that the wrought material tables (IWB-3641-1) remain applicable to GTAW weldments, githout any consideration of thermal expension orWher second-i ary stress terms. Thus, the inclusion of thermal stresses in 'ths. design basis for-i.he subject. overlays is seen to br/ highly conservative, and to lead to larger overlay thirknesses than would be specified if the overlays were designed under-the current requirements. These observations are described in dei; ail in the following paragraphs. 2.2.2 Weld Overlay Designs J i ( For the five flawed welds in the 12-tr.ch risers of the recirculation. system (12-12,12-23,12-64,12-69 and 12-7d), weld overlay designs have the sr.me /- minimum overlay length of 1.5 / E = 3.0 inches where R and t were assunied to 'be 6.331' and 0.69" (nominal values) respectively.. 1 i The minimum overlay length for weld 22-22 is 4.87 inches using R and t values of 11 inches anc 0.96 inch, respectively. 2-3 ~

y

/ y/ '

./i ' E' 3 Results of the overlay thickness calculations follow. h; j. ' 2.2.2.1 Weld 12-12 n i ' Indication Length / Pipe Circumference = 100% ,'p Indication Depth / Wall Thickness = 5% to 100% l? [ Applied Stresses (Table 2-2) p' - Pressure + Dead Weight + Thermal = 13747 psi Pressure + Dead Weight + Thermal + Seismic = 18479 psi Pressure + Dead Weight + Seismic = 11079 psi Overlay Design (Figure 2-5) Thickness = 0.49 inch (original design basis) See Table 2-3 Thickness = 0.32 inch current design basis w/o thermal 3 'gl, Length = 3.0 inches 2.2.2.2 Weld 12-23 f}, Indication Length / Pipe Circumference = 100% Indication Depth / Wall Thickness = 5% to 75% Applied Stresses (Table 2-2) h Pressure + Dead Weight + Thermal = 10531 psi L, Pressure + Dead Weight + Thermal + Seismic = 11483 psi Pressure + Dead Weight + Seismic = 5993 psi Overlay Design (Figure 2-5) 7' 3 Thickness = 0.34 inch (original design basis) See Table 2-4 Thickness = 0.24 inch (current design basis) Length = 3.0 inches il t 2.2.2.3 Weld 12-64 l i Indication Length / Pipe Circumference = 100% Indication Depth / Wall Thickness = 5% to 100% 2-4 l L

i \\ 'l Applied Stresses (Table 2-2) Pressure + Dead Weight + Thermal = 13079 psi Pressure + Dead Weight + Thermal.+ Seismic = 16367 psi Pressure + Dead Weight + Seismic = 8176 psi t. I Overlay Design (Figure 2-5) Thickngss = 0.48 inch (original design basis) See Table 2-5 Thickness = 0.28 inch (current design basis) Length = 3.0 inches i 2.2.2.4 Weld 12-69 Indication Length / Pipe Circumference = 100% Indication Depth / Wall Thickness = Spots 25% to 100%: Assumed to be 100% 0 Applied Stresses (Table 2-2) Pressure + Dead Weight + Thermal = 11875 psi Pressure + Dead Weight + Thermal + Seismic = 15625 psi Pressure + Dead Weight + Seismic = 10446 psi e i Overlay Design (Figure 2-5) Thickness = 0.40 inch joriginal design basis) a See Table 2-6 Thickness = 0.29 inch (current design basis) Length = 3.0 inches 2.2.2.5 Weld 12-70 p- 'h Indication Length / Pipe Circumference = 12.6% f-Indication Depth / Wall Thickness = 45% Applied Stresses (Table 2-2) Pressure + Dead Weight + Thermal = 8434 psi y Pressure + Dead Weight + Thermal + Seismic = 9619 psi Pressure + Dead Weight + Seismic = 6121 psi 2-5

Overlay Design (Figure 2-5) Thickness = 0.29 inch (original design basis) See Table 2-7 Thickness = 0.24 inch (current design basis) Length = 3.0 inches 2.2.2.6 Held 22-22 Indication Length / Pipe Circumference = 76% Indication Depth / Wall Thickness = 27% to 100% b Applied Stresses (see Table 2-2) Pressure + Dead Weight + Thermal = 6480 psi Pressure + Dead Weight + Thermal + Seismic = S480 psi Pressure + Dead Weight + Seismic = 6480 psi 9 Overlay Design (Figure 2-6) h Thickness = 0.35 inch (original design basis) See Table 2-8 Thickness = 0.35 inch (current design basis) j Length = 4.87 inches 2.3 Application Recommendations i-j The following application recommendations were employed in conjunction with the above weld overlay designs: As-deposited weld overlay ferrite levels were measured after each layer, s.. with a target minimum ferrite number (FN) of 10 FN. The weld overlay material was also specified to a maximum carbon level of 0.02% to ensure no propagation of the observed IGSCC into the weld overlay. f The minimum overlay thickness was measured after the first PT clear layer. This allows for the possibility of some propagation of the observed IGSCC into the ferrite diluted first layer, without reducing overlay design margins. 2-6 i

i l e The welding was performed using a controlled heat input process (<40 kJ/in) with water in the pipe to minimize any further sensitization of the underlaying piping material, and to ensure a favorable residual stress pattern. The as-installed designs will be re-examined during the 1987 refueling outage using EPRI-qualified techniques. The Authority has acquired weld overlay repair specimens from an operating plant after actual service and D laboratory weldments fabricated to procedures typical of field practice for h use in procedure development and training prior to that outage. 2.4 As-Built Weld Overlay l As-built dimensions and shrinkages were measured before, during, and after the application of the overlays. Results of these measurements are sumarized in Table 2-9. The as-built overlay thicknesses and lengths exceed the original design basis by a comfortable margin in all cases, and the current design basis by an even larger margin. 2.5 Shrinkage Stresses q l As illustrated in Table 2-9, average axial weld overlay shrinkages from 0.1 I inch to 0.2 inch were observed after the weld overlay repairs. Finite element analyses have been performed to calculate the increase in piping stress due to this weld overlay shrinkage. '7 Figure 2-7 illustrates the finite element model used for a typical riser in 5 the recirculation system. Both ends of the riser were conservatively assumed to be fixed. Shrinkage stresses predicted by this model are thus p l expected to be higher than those which would result from a model of the whole recirculation piping system, since such a model would account for struc-l tural flexibilities at the reactor pressure vessel nozzle and sweep-o-let. As-built weld overlay lengths were input as the distances from nodes 9 and ( 13 and from nodes 14 to 18 in Figure 2-7. Weld overlay shrinkages were then 1 2-7 l

\\ simulated as temperature differences between the overlay nodes and other nodes. The resulting shrinkage stresses are summarized in Table 2-10 for stresses at the weld overlay locations and in Table 2-11 for stresses at the sweep-o-lets, where shrinkage stresses are the highest. It is seen from Tcble 2-10 that, except for the riser with two overlays (12-69 and 12-70), the shrinkage stresses at the overlay locations are negligible (1.2 ksi). Even for the two-overlay riser, the maximum shrinkage stress is only 6.6 ksi at weld 12-69, and 4.9 ksi at weld 12-70. The total primary + secondary stresses, including these shrinkage effects, also listed in Table 2-10, are shown to be below the ANSI 8331.1 [29] allowable by a large margin. Additional overlay sizing calculations are performed for welds 12-69 and 12-70 in Tables 2-12 and 2-13, which conservatively include shrinkage stress along with thermal expansion as a primary stress. It is seen from these tables that the required overlay thickness does increase somewhat for q b these welds, but review of the as-built overlay thicknesses in Table 2-9 shows that sufficient extra thickness was applied to these welds to accommodate even this very conservative case. Review of Table 2-11 indicates that the shrinkage stresses at the sweep-o-let locations, f although somewhat higher than at the overlay locations, also do not significantly reduce the margins with respect to Code Allowable. An average axial shrinkage of.116 inch was observed after the weld overlay repair of weld 22-22. Since weld 22-22 is an end cap weld, there is no axial constraint, and the observed axial shrinkage can be accomodated with no additional stresses developed in the recirculation system. J. On the basis of the above evaluation, it is concluded that weld overlay f shrinkage stresses, calculated in a conservative manner, do not alter the conclusions of this report regarding the acceptability of the repairs and flaw evaluations. 2-8

e 2.6 Flaw Evaluation Six of the welds with UT indications, 12-4,12-17, 28-48, 28-53, 28-112, and 28-113, were evaluated in [4] in accordance with ASME Section XI, IWB-3600, supplemented by the recommendations of NRC Generic Letter 84-11 [1]. Both pre-and post-IHSI conditions were considered. All of these welds were re-inspected in spring of 1985. As a result of this inspection, three welds (12-17, 28-48 and 28-113) were determined to contain only geometric reflectors due to counterbore and weld root geometry. 2.6.1 Flaw Evaluation Methodology Induction Heating Stress Improvement (IHSI) has been applied to IGSCC-susceptible welds at FitzPatrick. Test and analytical data have shown that the IHSI will suppress not only crack initiation but also crack propagation for small cracks in both the length and depth directions. Consequently, the crack extension rate presented below must be considered a highly censerva-tive bound on the extension of FitzPatrick flaws. 2.6.1.1 Input Stresses As in weld overlay design, stre'sses used for flaw evaluation were obtained from [5] utilizing the piping models shown in Figures 2-3 and 2-4. These j stresses are sumarized in Table 2-2. The original critical flaw size b evaluations of [4] used stresses due to pressure + dead load + thermal p-expansion + seismic loadings in conjunction with the ASME Section XI L allowable flaw size tables for wrought, austenitic material (Table 2.14), 2 using allowable stress values from [29]. These calculations are discussed here, and updated to reflect recent Code changes to address low toughness weld metal concerns. Crack growth evaluations do not include the seismic components but do include thermal expansion, residual stress effects, and weld overlay shrinkage-induced stresses. A wide body of experimental and analytical data exists on the residual stresses in austenitic pipe welds. A summary of such data for axial residual 2-9

a stresses in large diameter pipes (20 inch nominal diameter and greater) taken from [2] is given in Figure 2-8(a). Note that there is considerable variability, but that the majority of the data show a U-shape trend varying from tension on the inside surface to compression in the central region of the pipe, and back to tension again near the outside surface. Also shown in this figure is an " evaluation" curve recommended in [2] for use in IGSCC crack growth evaluations of large pipe. This curve provides a reasonably conservative representation of the available data for large diameter weldments. Thus, this " evaluation" curve was used in the present IGSCC crack growth analyses for welds 28-53, and 28-112. This curve can be expressed as follows: e = 30 - 242.05(x/t) + 394.20(x/t)2 - 174.30(x/t)3 (1) where ais axial stress in the units of ksi, and x and t, are the distance from the inside surface and the wall thickness, respectively. l Also shown as Figure 2-8(b) is a linear through-wall residual distribution I which is widely accepted as a reasonable representation of the residual stresses in moderate to small diameter piping (12 inch nominal diameter and less). The small pipe residual stress pattern of Figure 2-8(b) generally results in more rapid crack propagation than the large pipe distribution l' discussed previously. The linear curve of Figure 2-8(b) can be expressed as e = 30 - 60(x/t) (2) c Li where a, x, and t are defined in the same way as in equation (1). A large body of laboratory data and analytical solutions also exists on post-IHSI residual stresses in austenitic pipe welds. These data are summarized in [7]. Typical post-IHSI axial stress distributions are illustrated in 2-10

O 6 Figures 2-9 and 2-10 for large and small diameter pipes respectively. These stress distributions were also curvefit by third order polynomials for use in the analysis, and the resulting equations are given in Figures 2-9 and 2-10. 2.6.1.2 Stress Intensity Factors Actual pipe dimensions were used in this analysis. An analytical model of a L 3600 circumferential crack in a cylinder of radius to thickness ratio of 10:1 [8]wasusedforthefracturemechanicsevaluation. For the pre-IHSI case, applied loading is the superposition of piping stresses tabulated in Table 2-2, and the as-welded residual stress defined by equation (1) for 28-inch pipe or equation (2) for 12-inch pipe. For the post-IHSI case, applied loading is the sum of the same piping stresses from Table 2-2 and the post-IHSI residual stress distributions given in Figure 2-9 for 28-inch pipe or Figure 2-10 for 12-inch pipe. For purposes of the fracture mechanics analysis, the axial stress dis-tributions of piping stress, pre-IHSI residual stress, and post-IHSI residual stress have all been expressed in terms of third degree polynomials of the form: g ) 0 + A x + A x2 + A x3 (3) e=A 1 2 3 where e and x are defined the same as in equation (1) and Ao - A3 are the coefficients resulting from the curvefit. 7 The stress intensity factor for a circumferential crack in a cylinder of J. radius to thickness ratio of 10:1 can be expressed as follows [8]: 7 I = [da (A F1 + 2_a A F a2 AF 4a3 K AF) (4) 0 I 2 + Ti 2 3 + Jit 34 n where F, F, F, and F4 are magnification factors and a is crack depth as 1 2 3 shown in Figure 2-11. 2-11

( For linear elastic fracture mechanics evaluations, stress intensity factors can be calculated independently for piping stress and pre-and post-IHSI residual stress distributions. The resultant stress intensity factor is the superposition of the appropriate loading cases. f 2.6.1.3 Crack Growth A large body of laboratory data also exists on stress corrosion crack growth ). 3 rates for sensitized stainless steels in simulated BWR environments. These data are summarized in Figure 2-12, also taken from [2]. These data were obtained using fracture mechanics type specimens with different crack sizes and loadings which can be characterized by the crack tip stress intensity factor K. The data represent a wide variation in both material sensitization and levels of dissolved oxygen in the water. The widely used power law "best estimate" curve in Figure 2-12 is believed to provide a representative crack propagation rate for plant crack growth assessments. The "best estimate" curve can be described by a power law representation of the form: i da/dt = 2.27 X 10-8(K)2.26 (5) h where a is the crack depth in units of inches, t is time in units of hours, andKisthestressintensityfactorinunitsofksi/13. [ Crack growth analyses typically make use of one of the two assumptions l ~. illustrated in Figure 2-13 regarding crack length extension, self-similar i crack growth or constant aspect ratio crack growth. The former assumes that a the incremental crack extension is the same at all points on the crack front, A while the latter assumes that the ratio of depth to length remains constant during crack extension. Considering field and laboratory experience with circumferential crack extension, it appears that the self-similar assumption may underpredict crack length versus time, while the constant aspect ratio assumption overpredicts. 2-12

Recent work by Gerber [9] under contract to EPRI provides a new approach for addressing circumferential crack extension which is more technically de-fensible than the self-similar or constant aspect ratto approaches. This approach utilizes data generated in a laboratory stress corrosion test of a 26-inch diameter welded pipe specimen at Battelle Pacific Northwest Labora-tories [10]. IGSCC was induced in this pipe through loading to a high applied stress in a simulated BWR environments. Crack initiation was accelerated by the use of graphite wool to create an artificial crevice. Crack growth occurred and was monitored both during operation and at several scheduled shutdown intervals for the test. At the completion of the test, there were a total of 63 cracks with a combined length of 32.57 inches. f f The average effective circumferential crack extension observed in this test } is presented in Figure 2-14. This rate includes both growth of existing cracks as well as new defects initiating and contributing to the effective crack growth rate in each inspection interval. Examination of Figure 2-14 ki suggests that an average effective circumferential crack growth rate of 0.5 mils / hour should give a reasonably conservative estimate. Thus, 0.5 mils / hour was used as the crack length growth rate in this report. It should be pointed out, however, that although this is an average effective rate, it [1 is based on a laboratory test in which the local environment, load and cycles were all intentionally modified to accelerate IGSCC relative to actual plant i conditions. a 2.6.1.4 Allowable Flaw Size l Based on detailed calculations presented in [11], allowable flaw sizes for j wrought austenitic piping have been specified in ASME Section XI, IWB-3640. A tabulation of allowable flaw sizes as a function of applied load is given [ in Table 2-14, which is taken directly from Section XI, Table IWB-3641-1 [6]. Note that this table permits very large defects in some cases (as great as 75% I of pipe wall) and does not include consideration of any stress other than primary. Notably secondary and peak stresses from the design stress report, as well as any weld residual stresses or misalignment / fit-up stresses which might exist from construction are not addressed. The argument for this 2-13

O exclusion is that, given the extremely high ductility of austenitic stainless steel, these strain-controlled effects will self relieve after a small amount of plastic deformation and/or stable crack extension and will 'have little or no impact on the loads and flaw sizes needed to cause unstable crack propagation or pipe rupture. After publication of the above-referenced Code rules for austenitic piping, some fracture toughness data were discovered which tend to invalidate the above argument for some classes of custenitic weld metal [3]. The data in ) question suggest that welds made with flux-shielded processes (SAW and SMAW) may possess much lower toughness than the wrought base material. To account for the possibility of low ductility weld metal, secondary stresses from the stress report were included in the original flaw evaluations performed in [4], but the wrought material table was still used for flaw evaluation. Since that time, ASME Section XI has been revised (Winter,1985 Edition) to address the concerns for low toughness weld metals explicitly. This revision p included a new table, IWB-3641-5, which is to be used in lieu of the wrought L material table for flux type weldments (SAW and SMAW). This table, included for reference as Table 2-15, requires the inclusion of thermal expansion and other global secondary stresses, but with a reduced safety factor relative to the primary stresses. The allowable flaw sizes are lower than the values in ~ Table 2-14 to account for the lower toughness material. Allowable flaw size calculations are presented in this report using both the original method-ology of [4], as well as the more up-to-date methodology provided in the new Code revision. It is important to note that the very low measured toughness was observed only in a small percentage of the materials studied and may be of only limited concern from a probabilistic viewpoint. Indeed, most IGSCC observed to date has been restricted to weld heat affected zones, which should exhibit the high toughness attributed to wrought stainless steel base material. Never-theless, to address these possible concerns, the analysis procedure used throughout this report includes thermal expansion and weld overlay shrinkage effects as a primary stress condition whenever the allowable flaw size is determined from Table 2-14. 2-14

2.6.2 Evaluations and Results ( 2.6.2.1 Weld 12-4 Input to the flaw evaluation for this weld was as follows: Indication Length - 1.8 inch Indication Depth - 0.046 inch Pipe 0.D. - 12.662 inches Pipe I.D. - 11.442 inches f. Pipe Wall Thickness - 0.61 inch Applied Stresses (Table 2-2) Pressure + Dead Weight + Thermal = 14145 psi Pressure + Dead Weight + Thermal + Seismic = 17629 psi Pressure + Dead Weight + Seismic + Thermal /2.77 = 12439 psi Residual Stresses Figure 2-8(b) Pre-IHSI j Post-IHSI - Figure 2-10 Figure 2-15 provides applied stress intensity factor versus crack depth data for the three load cases used in the evaluation. Assuming the indication to be IGSCC, these stress intensity curves were used to perform 1 IGSCC crack growth estimates for both as-welded and post-IHSI residual L.- stress conditions. The resulting crack growth prediction is illustrated in j Figure 2-16 for the as-welded case. The post-IHSI case is also shown, and results in no predicted crack growth for the balance of plant life. The allowable end-of-cycle flaw size was determined both from the original methodology of [4] and in accordance with the more recent ASME Section XI, Table IWB-3641-5. The results are illustrated in Figure 2-17 in terms of allowable flaw depth versus length. Also, in accordance with the 2-15

recomendations of NRC Generic letter 84-11, [1] a maximum allowable flaw size of 2/3 of the IWB-3641-1 limit (shown as a dashed line in Figure 2-17) was used to allow for uncertainty in flaw depth sizing, when the original methodology of [4] was used. The factor of 2/3 is not required when the new methodology of Table IWB-3641-5 is used. Referring to Figure 2-17, it is seen that the 2/3 of IWB-3641-1 limit is more restrictive than the new ASME Code methodology. This limit is predicted to be exceeded in a fairly short time in the as-welded case, but is satisfied indefinitely in.the post-IHSI case, since no crack propagation is pre-dicted. To add further assurance in the post-IHSI case, the IGSCC crack growth analysis has been repeated assuming various initial flaw sizes ranging upward from the observed UT depth. No crack propagation is predicted in the post-IHSI condition for initial crack depths up to 0.488 inch, or 80% of the pipe wall. It is also noteworthy that, given the relatively short length of the observed indication (2.8% of circumference), it would not lead to failure of the pipe joint even if the above crack growth or initial flaw size estimates are significantly in error. Leak-before-break is clearly the expected, hypothetical failure mode for this indica-tion. ql I On the basis of the above evaluation, it is concluded that continued h operation of the plant with this weld, considering the observed indication and the IHSI treatment which has been applied, will not lead to a reduction in plant safety margins, or a plant operational concern for the balance of the plant lifetime. h1 2.6.2.2 Weld 28-53 Input to the flaw evaluation for this weld was as follows: Indication Length - 0.4 inch l Indication Depth - 0.062 inch i l l 2-16

Pipe 0.D. - 28.363 inches Pipe I.D. - 25.867 inches Pipe Wall Thickness - 1.24 inch Applied Stresses (Table 2-2) Pressure + Dead Weight + Thermal = 7177 psi Pressure + Dead Weight + Thermal + Seismic = 11,211 psi Pressure + Dead Weight + Seismic + Thermal /2.77 = 10,399 psi p ) Residual Stresses Figure 2-8(a) Pre-IHSI Post-IHSI - Figure 2-9 Figure 2-18 provides applied stress intensity factor versus crack depth data for the three load cases used in the evaluation. Assuming the indication to be IGSCC, these stress intensity curves were used to perform u l IGSCC crack growth estimates for both as-welded and post-IHSI residual stress conditions. The resulting crack growth prediction is illustrated in Figure 2-19 for the as-welded case. The post-IHSI case is also shown, and results in no predicted crack growth for the balance of plant life. q The allowable end-of-cycle flaw size was determined ooth from the original 7 methodology of [4] and in accordance with the more recent ASME Section XI, Table IWB-3641-5, and is illustrated in Figure 2-20 in terms of allowable flaw depth versus length. Also, in accordance with the recommendations of [1], a maximum allowable flaw size of 2/3 of the IWB-3641-1 limit (shown as a dashed line in Figure 2-20) was used to allow for uncertainty in flaw depth sizing, when the original methodology of [4] was used. The factor of 2/3 is 7 not required when the new methodology of Table IWB-3641-5 is used. Referring to Figure 2-20, it is seen that the 2/3 of IWB-3641-1 limit is more restrictive than the new ASME Code methodology. This limit is not predicted to be exceeded for greater than 36 months in the subject weld, even considering as-welded residual stress conditions. If IHSI residual stress 2-17

benefits are accounted for, the flaw will remain at its present size indefinitely, and thus satisfy the allowable flaw size limit by a large margin for the balance of plant life. On the basis of the above evaluation, it is concluded that continued operation of the plant with this weld, considering the observed indication, will not lead to a reduction in plant safety margins, or a plant operational concern for the balance of the plant lifetime. F I 2.6.2.3 Weld 28-112 Input to the flaw evaluation for this weld was as follows: Indication Length - 4.72 inch Indication Depth - 0.198 inch Pipe 0.D. - 28.363 inches Pipe I.D. - 25.867 inches Pipe Wall Thickness - 1.24 inches Applied Stresses (Table 2-2) Pressure + Dead Weight + Thermal = 7146 psi i Pressure + Dead Weight + Thermal + Seismic = 8126 psi Pressure + Dead Weight + Seismic + Thermal /2.77 = 7688 psi Residual Stresses Figure 2-8(a) Pre-IHSI p G Post-IHSI - Figure 2-9 r Figure 2-21 provides applied stress intensity factor versus crack depth data for the three load cases used in the evaluation. Assuming the indication to be IGSCC, these stress intensity curves were used to perform IGSCC crack growth estimates for both as-welded and post-IHSI residual stress conditions. The resulting crack growth prediction is illustrated in Figure 2-22 for the as-welded case. The post-IHSI case is also shown, and results in no predicted crack growth for the balance of plant life. 2-18

9 The allowable end-of-cycle flaw size was determined both from the original methodology of [4] and in accordance with the more recent ASME Section XI, Table IWB-3641-5. The results are illustrated in Figure 2-23 in terms of allowable flaw depth versus length. Also, in accordance with recom-mendations of NRC Generic Letter 84-11, [1] a maximum allowable flaw size of 2/3 of the IWB-3641-1 limit (shown as a dashed line in Figure 2-23) was used to allow for uncertainty in flaw depth sizing, when the original methodology of [4] was used. The factor of 2/3 is not required when the new methodology 8 of Table IWB-3641-5 is used. i Referring to Figure 2-23, it is seen that the 2/3 of IWB-3641-1 limit is more i restrictive than the new ASME Code methodology. This limit is not predicted to be exceeded for greater than 36 months in the subject weld, even considering as-welded residual stress conditions. If IHSI residual stress benefits are accounted for, the flaw will remain at its present size indefinitely, and thus satisfy the allowable flaw size limit by a large t 3 margin for the balance of plant life. On the basis of the above evaluation, it is concluded that continued operation of the plant with this weld, considering the observed indication, will not lead to a reduction in plant safety margins, or a plant operational concern for the balance of the plant lifetime. 2.6.2.4 Welds 12-17, 28-48, and 28-113 k The 1985 in-service inspection showed that ultrasonic indications in 12-17, [ 28-48, and 28-113, which were identified as smal.1 IGSCC-type flaws during l' the Fall 1984 151, were actually due to geometry effects. Prior analysis f [4] showed that even without IHSI, none of the observed indications (con-servatively assuming the indication to be cracks) would exceed 2/3 of the ASME Section XI, Article IWB-3641-1 allowable indication size for at least 36 months. With IHSI, the flaws were not predicted to propagate at all. Operation of the plant with only the geometric reflectors present in these welds does not jeopardize safety margins or present a plant operational J concern for the balance of the plant lifetime. 1 2-19 L

All of the 6 welds discussed in this se; tion (12-4,12-17, 28-48, 28-53, 28-112, and 28-113) were inspected in October 1984 and again in April, 1985. 1 Although the indications in 3 of these welds were classified as IGSCC in 1984, the indications were reclassified as being due to counterbore or geometric effects in 1985, based upon re-inspection and re-evaluation of the earlier inspection results. No significant change in the raw inspection data for any of the 6 welds discussed in this section was observed, between 1984 and 1985. This supports the analytical predictions 3 of no growth for these welds. 3 2.7 Fatigue Evaluation of Flaws Under Weld Overlays 1 Section 3 of this report discusses evidence that a weld overlay-repaired location is unlikely to degrade further by an IGSCC mechanism. The possibility of degradation of weld overlay design margins by growth of IGSCC g flaws due to fatigue has also been considered. Taking as a bounding case the highest-stressed weld overlay repaired weld (12-12) and assuming that the original flaw was through the original pipe wall, analysis has shown that the flaw would not propagate significantly into the overlay due to fatigue, over the life of the plant (less than 0.001 inch). >O F 2-20

~- ~ - VW !L in T 1_; ~ TABLE 2-1 WELD

SUMMARY

JAMES A. FITZPATRICK - RESULTS OF IGSCC INSPECTIONS Weld No. Loop Weld Location Crack length Thru Depth IHSI Discovery Previous Corrective Remarks Type Wall Var. Method Inspection Action 12-4 A Pipe to C 4.5% No 7% Yes UT 10/84 IHSI & See Note sweep-o-let 9/84 Analysis 3 12-12 A Pipe to C 100% Yes Avg. 50% Yes PT, Visual, Weld See Note safe end 1 Max. 100% 9/84 UT, Post IHSI Overlay 1 Min. 5% 12-17 A Pipe to C N/A No N/A Yes UT 10/84 IHSI & See Note safe end 9/84 Analysis 2 (counterbore) 12-23 A Pipe to C 100% No Avg. 40% Yes UT, Post IHSI 6/83 Weld See Note safe end int. Max. 75% 9/84 Overlay 1 Min. 5% 12-64 B Pipe to C 100% Yes Avg. 30% Yes PT, Visual, Weld See Note

• f safe end int.

2 Max. 100% 9/84 UT, Post IHSI Overlay 1 ta Min. 5% 12-69 B Pipe to C 100% Yes Not avail Yes PT, Visual, 12/81 Weld See Note safe end 2 9/84 UT, Post IHSI Overlay 1 12-70 B Elbow to pipe C 12.6% No 45% Yes UT, Pre-IHSI 12/81 Weld 9/84 Overlay 28-48 A Pipe to None N/A N/A N/A Yes N/A 10/84 IHSI & See Note safe end (counter bore) 3/84 Analysis 2 28-53 A Elbow to C 1% No 5% Yes UT 10/84 IHSI & See Note valve 9/84 Analysis 3 28-112 B Elbow to C 5.5% No 16% Yes UT 10/84 IHSI & See Note valve 9/84 Analysis 3 28-113 B Valve to pipe None N/A N/A N/A Yes N/A 10/84 IHSI & See Note (counter bore) 9/84 Analysis 2 22-22 8 End Cap to C 76% No Max. 27% Yes UT 10/84 Weld to Manifold 3/84 Overlay Note 1 - No pre-IHSI UT examination performed. Note 2 - Fall 1984 Examination gave IGSCC-type indications,1985 ISI showed no reportable indications Note 3 - These locations were examined and IHSI treated in 1984, and re-examined in 1985. There was no significant difference in the 1984 and 1985 inspection results.

T TABLE 2-2

SUMMARY

OF PIPING STRESSES

• STRESS, PSI **

THERMAL WELD NO. JOINT NO.** PRESSURE DEAD WEIGHT EXPANSION SEISMIC 12-4 258 5672 350 8123 3484 a-12-12 194 4921 1426 7400 4732 7 12-17 178 5672 599 7807 4711 i 12-23 215 4778 263 5490 952 12-64 24 4528 360 8191 3288 12-69 64 5210 1486 5179 3750 12-70 65 4726 210 3498 1185 22-22 NA 6480 0 0 0 28-48 324 5674 783 847 1083 23-53 331 5675 232 1270 4034 28-112 175 5674 786 686 980 28-113 176 5871 820 587 1192 b Nodalnumbersfrom[4] Based upon stress from [4]; normalized to measured wall thickness ne u ~ 9 1 2-22 l

3 TABLE 2-3 Weld Overlay Sizing for Weld 12-12 WELD OVERLAY SIZING EVALUATION WELD OVERLAY S121NG FOR CIRCUMF. CRACK,Wh0!!GHT/ CAST STAINLESS WELD 12-12 W/O THERMAL WALL THICKNESS: 0.7030 inch MEMBRANE STRESS = 4921.0000 psi BENDING STRESS: 6158.0000 psi STRESS RATIO: 0.6968 ALLOWABLE STRESS =15900.0000 psi FLOW STRESS =47700.0000 psi L/ CIRCUM 0.00 0.10 0.20 0.30 0.40 0.50 FINAL A/T 0.7500 0.7500 0.7500 0.7500 0.7500 0.6895 OVERLAY THICKNESS 0.2343-0.2343 0.2343 0.2343 0.2343 0.3166 (inch) WELD OVERLAY SIZING FOR CIRCUMF. CRACK, WROUGHT / CAST STAINLESS WELD 12-12 W/ THERMAL WALL TH3CKNESS: 0.7030 inch MEMBRANE STRESS: 4921.0000 psi BENDING STktSS: 13588.0000 psi STRESS RATIO: 1.1641 ALLOWABLE STRESS:15900.0000 psi FLOW STRESS =47700.0000 psi L/ CIRCUM 0.00 0.10 0.20 0.30 0.40 0.50

7 FINAL A/T 0.7500 0.7500 0.7500 0.7378 0.6831 0.5874 OVERLAY THICKNESS 0.2343 0.2343 0.2343 0.2498 0.3261 0.4938 (inch) 2-23

[

TABLE 2-4 Weld Overlay Sizing for Weld 12-23 WELD OVEhbAY S3Z3NG FOR CIRCUMF. CRACK,Wh00GHT/CALT STAINLEds } WELD 12-23 W/O THERMAL WALL THICKNESS: 0.7240 inch MEMBRANE SThESS: 4778.0000 psi BENDING STRESS: 1215.0000 psi STRESS RATIO: 0.3769 ALLOWABLE STRESS =15900.0000 psi FLOW STRESS:47700.0000 psi L/ CIRCUM 0.00 0.10 0.20 0.30 0.40 0.50 FINAL A/T 0 7500 0.7500 0.7500 0.7500 0.7500 0.7500 OVERLAY THICKNESS 0.2413 0.2413 0.2413 0.2413 0.2413 0.2413 WELD OVERL(AY SIZING FOR CIRCUMF. CRACK, WROUGHT / CAST STAINLESS inch) WELD 12-23 W/ THERMAL WALL THICKNESS: 0.7240 inch MEMBRANE STRESS: 4778.0000 psi BENDING STRESS: 6705.0000 psi STRESS RATIO: 0.7222 ALLOWABLE STRESS =15900.0000 psi FLOW SThESS:47700.0000 psi L/ CIRCUM 0.00 0.10 0.20 0.30 0.40 0.50 FINAL A/T 0.7500 0.7500 0.7500 0.7500 0.7500 0.6826 OVERLAY THICKNESS 0.2413 0.2413 0.2413 0.2413 0.2413 0.3366 (inch) 4 i 2-24

TABLE 2-5 Weld Overlay Sizing for Weld 12-64 WhLD OVEhLAY SIZING FOh CIRCUHF. CRACK, WROUGHT /CALT STAINLESS WELD 12-64-W/O THERMAL i WALL THICKNESS = 0.7640 inch MEMBRANE STRESS: 4528.0000 psi BEND]NG STRESS: 3648.0000 psi STRESS RATIO: 0.5142 ALLOWABLE STRESS =15900.0000 psi FLOW STRESS =47700.0000 psi L/ CIRCUM 0.00 0.10 0.20 0.30 0.40 0.50 FINAL A/T 0.7500 0.7500 0.7500 0.7500 0.7500 0.7344 E OVERLAY TH1CKNESS 0,.2547 0.2547 0.2547 0.2547 0.2647 0.2763 (inch) WELD OVERLAY SIZING FOR CIRCUHF. CRACK, WROUGHT / CAST STAINLESS = WELD 12-64 W/ THERMAL I WALL THICKNESS = 0.7640 inch MEMBRANE STRESS: 4528.0000 psi BENDING STRESS: 11836.0000 psi STRESS RATIO: 1.0294 i ALLOWABLE STRESS:15900.0000 psi P FLOW STRESS =47700.0000 psi j~ L/ CIRCUM 0.00 0.10 0.20 0.30 0.40 0.50 FINAL A/T 0.7500 0.7500 0.7500 0.7500 0.7202 0.6138 OVERLAY THICKNESS 0.2547 0.2547 0.2547 0.2547 0.2968 0.4808 (inch) [ l 1 2-25 L

O TABLE 2-6 Weld Overlay Sizing for Weld 12-69 WELD 12-69 W/O THERMAL ) WALL THICKNESS: 0.6040 inch MEMERANE STkESS: 5210.0000 psi. BENDING STRESS: 5236.0000 psi STRESS RATIO: 0.6570 ALLOWABLE STRESS:15900.0000 psi FLOW STRESS =47700.0000 psi L/ CIRCUM 0.00 0.10 0.20 0.30 0.40 0.50 FINAL A/T 0.7500 0.7500 0.7500 0.7500 0.7500 0.6997 OVERLAY THICKNESS 0.2213 0.2213 0.2213 0.2213 0.2213 0.2850 (inch) WELD OVERLAY SIZING FOR CIRCUHF. CRACK, WROUGHT / CAST STAINLESS WELD 12-69 W/ THERMAL WALL THICKNESS = 0.6640 inch MEMERANE STRESS = 5210.0000 psi BENDING STRESS: 10416.0000 psi STRESS RATIO: 0.9527 i ALLOWABLE STRESS:15900.0000 psi 'l FLOW STRESS:47700.0000 psi L/ CIRCUM 0.00 0.10 0.20 0.30 0.40 0.50 FINAL A/T 0.7500 0.7500 0.7500 0.7500 0.7344 0.6230 OVERLAY THICKNESS 0.2213 0.2213 0.2213 0.2213 0.2402 0.401Y (inch) m 2-26

e o TABLE 2-7 Weld Overlay Sizing for Weld 12-70 WELD OYEhLAY SIZING FOR CIRCUMF. CRACK.WROUGIC/CACT STAINLESS WELD 12-70 W/O THERMAL 3 e WALL THICKNESS-0.7320 inch HEMBRANE STRESS: 4726.0000 psi BENDING STRESS: 1395.0000 psi STRESS RATIO: 0.3850 ALLOWABLE STRESS =15900.0000 psi FLOW STRESS =47700.0000 psi L/ CIRCUM 0.00 0.10 0.20 0.30 0.40 0.50 FINAL A/T 0.7500 0.7500 0.7500 0.7500 0.7500 0.7500 OVERLAY THICKNESS 0.2440 0.2440 0.2440 0.2440 0.2440 0.2440 WELD OVERL(AY IZING FOR CIRCUMF. CRACK, WROUGHT / CAST STAINLESS inch WELD 12-70 W/ THERMAL WALL THICKNESS: 0.7320 inch se MEMBRANE STRESS = 4726.0000 psi BENDING STRESS: 4893.0000 psi STRESS RATIO: 0.6050 ALLOWABLE STRESS =15900.0000 psi FLOW STRESS =47700.0000 L/ CIRCUM 0.00 0.10 0.20 0.30 0.40 0.50 FINAL A/T 0.7500 0.7500 0.7500 0.7500 0.7500 0.7139 OVERLAY THICKNESS 0.2440 0.2440 0.2440 0.2440 0.2440 0.2934 (inch) 3., 3 t l l I I 2-27 / 9 l

TABLE 2-8 Weld Overlay Sizing for Weld 22-22 o I f WELD OVERLAY SIZING FOR CIRCUMF. CRACK,WROUOHT/ CAST STAINLESS WELD 22-22 (END CAF SEES ONLY PRESSURE STREGS) WALL THICKNESS: 1.0360 inch n MEMBkANE STRESS: 6480.0000 psi BENDING STRESS: 0.0000 psi k STRESS RATIO: 0.4075 ALLOWABLE STRESS =15900.0000 psi I FLOW STRESS =47700.0000 psi L/ CIRCUM 0.00 0.10 0.20 0.30 0.40 0.50 g FINAL A/T 0.7500 0.7500 0.7500 0.7500 0.7500'O.7480 OVERLAY THICKNESS 0.3453 0.3453 0.3453 0.3453 0.3453 0.3489 (inch) s, U3 1 2-28 i I i I L

TABLE 2-9 AS-BUILT WELD OVERLAY DIMENSIONS i WELD 12-12 12-23 12-64 12-69 12-70 22-22 (t)o inches .703 .724 .764 .664 .732 1.036 (D)o inches 5.363 5.286 5.172 5.976 5.701 8.002 L (t)1 inches .846 .834 .797 .778 .851 1.138 (t)w inches 1.352 1.4 04 1.469 1.268 1.236 1.583 7 (D)w inches 5.26 5.106 5.022 5.851 5.544 7.886 (E)w inches 3.071 3.455 3.115 3.093 3.225 5.062 (F)w inches 3.9% 4.054 4.21 4.064 3.838 5.834 (h)1 inches .143 .11 .033 .114 .119 .102 g (h)w inches .499 .57 .672 .482 .385 .444 (h)t inches .642 .68 .705 .5% .504 .546 SHRINKAGE inches .103 .18 .15 .125 .157 .116 f THETA degrees 54.525 66.229 52.167 51.207 58.695 55.220 p F 8. D D-A.

7..

F. E., r, b. p 3 x., nur .,x o 7 i e uni s e e 2-29 l 6

f TABLE 2-9 (CONTINUED) DEFINITION OF PARAMETERS PARAPETER DEflNITION to average original pipe wall thickness Do distance between punch marks before overlay tj average thickness including first overlay layer i t average thickness after overlay w D, distance between punch marks af ter overlay E, full thickness overlay length 8 F, toe-to-toe overlay length neglecting first layer h; average thickness of first overlay layer m h average thickness of overlay neglecting first layer w average total overlay thickness i. ht iI SHRINKAGE average overlay shrinkage-Do-D, THETA weld overlay transttlon angle 4, 4 2-29a ...,.. ~ _ _ _ _ _. _

q y- .c ~ j s TABLE 2-10 SUPetARY OF SHRINKAGE STRESSES AT'9VERLAY t0 CATIONS 4 g e, I WELD 12-12 12-23 12-64 12-g 12-70 22-22 MEASLGED WALL THICK'iESS .703" .724" .164!'n-.664d __732" 1.036" SHRINKAGE .103" .180" .150" .125" .157" .116" f. ' OVERLAY LENGTH 5.363" 5.286" 5.172" "S;976" 5.701" 8.002" ' i SHRINKAGE STRESS (ksi) 1.160 2.032 1.701 6.605 4.920 0 ~ 1 TOTAL STRESS * (ksi) 18.479 11.483 16.357 15.625 9.619 6.480 (PRE-WELDOVERLAY) 1 l TOTAL STRESS * (ksi) 19.639 13.515' 13.068 22.230 14.530 6.480 i (POST-WELDOVERLAY) j"7' ALLOWABLE STRESS (ksi) 43.25 43.25 . 43.25 43,25 43.25 43.25 n (fromB31.1[293) Total Stress = Pressure + Dead Weisht + TNrmal Expansion + Seismic from Tatle 2-2 for pre-weld overlay, for. post-weld overlay, shrinkage stress is included. r .o ) s y L. i' s s 2-30 i, 3

TABLE 2-11

SUMMARY

OF SHRINKAGE STRESSES AT SWEEP-0-LET LOCATIONS 7_ OVERLAY AT 12-12 12-23 12-64 12-69 & 12-70 4 STRESS AT WELD 12-15 12-26 12-67 12-72 i' SHRINKAGE STRESS (KSI) 3.578 6.251 5.210 11.993 TOTAL STRESS * (KSI) 18.422 23.935 18.302 14.102 (PRE-WELD OVERLAY) TOTAL STRESS * (KSI) 22.000 30.186 23.512 26.095 (POST-WELD OVERLAY) / ALLOWABLE STRESS (KSI)** 43.25 43.25 43.25 43.25 (from B31.1 [29]) h' I Total Stress = Pressure + Dead Weight + Thermal Expansion + Seismic (' from Table 2-2 for pre-weld overlay; for post-weld overlay, shrinkage stress is included.

• • Allowable Stress = Sh + Sa = 1.25Sh + 1.25Sc

)

with Sh = 15.9 and Sa = 27.35 a Sc = 18.7 i C. m i D I J l 2-31

TABLE 2-12 Weld Overlay Sizing for Weld 12-69 Considering Shrinkage WELii OVERLAY SIZING FOh CIR00MF. CRACK, WROUGHT / CAST STAINLESS WE!.D OVERLAY DESIGN FOR WELD 12-69 CONSIDERING SHRIt'KAGE WALL THICKNESS:

0. 664 0 inch MEMBRANE STRESS: 5210.0000 psi BENDING STRESS = 11841.0000 psi STRESS RATIO:

1.0724 l ALLOWABLE STRESS =15900.0000 psi FLOW STRESS 47700.0000 psi L/ CIRCUM 0.00 0.10 0.20 0 30. 0.40 0.50 FINAL A/T 0.7500 0.7500 0.7500 0.7490 0.7080 0.6050 OVEhLAY THICKNESS 0.2213 0.2213 0.2213 0.2225 0.2738 0.4336 (inch) TABLE 2-13 Weld Overlay Sizing for Weld 12-70 Considering Shrinkage l. WELD OVERLAY SIZING FOR CIRCUMF. CRACK, WROUGHT /CAS'I STAINLESS WELD OVERLAY DESIGN FOR WELD 12-70 CONSIDERING SHRINKAGE l WALL THICKNESS: 0.7320 inch a MEMBRANE STRESS:- 4726.0000 psi BENDING STRESS: 6315.0000 psi g. STRESS RATIO: 0.6944 ))' ALLOWABLE STRESS =15900.0000 psi FLOW STRESS =47700.0000 psi L/ CIRCUM g 0.00 0.10 0.20 0.30 0.40 0.50 FINAL A/T 0.7500 0.7500 0.7500 0.7500 0.7500 0.6899 OVERLAY THICKNESS 0.2440 0.2440 0.2440 0.2440 0.2440 0.3290 (inch) 2-32

i l i 1 e l TABLE 2-14 (TABLE IWB-3641-1 from ASME Section XI, 1983 Edition, Winter 1985 Addenda [31] ALLOWABLE END-OF EVALUATION PERIOD FLAW DEPTH 2 TO THICKNESS RATIO FOR CIRCUMFERENTIAL FLAWS - NORMAL OPERATING (INCLUDING UPSET AND TEST) CONDITIONS Ratio of Fla Length. /,, to Poe Cartumference lhote (3)) P, + P, 0.5 5, (Note (2)) 0.0 0.1 0.2 0.3 04 er More 3.5 tel (4) (43 141 (al (4) 1.4 0 75 0 40 0.21 0 15 tal del 3.) 075 0 75 0 39 0 27 0 22 0 19 1.2 0 75 0 75 0 56 0 40 0 32 0 27 1.1 075 0.75 0 73 0 51 C 42 0 34 1.0 0 75 0 75 0 75 0 63 0 51 0 41 09 0 75 0 75 0 75 0 73 0 59 0 47 08 0.75 0.75 0 75 0 75 0 68 0 53 0.7 0 75 0.75 0 75 0 75 0 75 0 58 g06 075 0.75 075 075 0 ?$ 0 63 C OT L 5. (1) Flaw septh e a for a surface fian Ia,for a subsurface fiam f = nom.naltNckness r~ Linear nte'pciation is perm'ssible (2) P, = primary membrane stress P, a pr.rra f encing st*ess t 5, = allomatN desge. stress eettasip t.n etto'Cance ete settic911D (3) Ce*tumferente based en nominat p.pe 6 ameter. l 841IWB 3514 3 shall be used. Note: for the FitzPatrick weld overlay design, the allowable stress, identified as Sm above, was taken from ANSI-B31.1, the original .a CodeofConstruction[29]. S r" 7 l l l 2-33 t

1 TABLE 2-15 (TABLE IWB-3641-5 from ASME Section XI, 1983 Edition, Winter 1985 Addenda [31] ALLOWABLE END-0F EVALUATION PERIOD FLAW DEPTH 1 TO THICKNESS RATID FOR CIRCUMFERENTIAL FLAWS IN SHIELDED METAL ARC AND SUBMERGED ARC WELDS NORMAL OPERATING UNCLUDING UPSET AND TEST) CONDITIONS A-Ratio of Flow Length, t,, to Pipe Circumference INote (3)) O Stress h Ratie 0.5 (Note (2)) 0.0 0.1 0.2 0.3 0.4 er Greater j 1.05 (4) (4) (4) (4) (4) (4) V 1.00 0.60 0.25 0.16 0.13 (4) (4) 0.95 0.60 0.47 0.25 0.10 0.13 0 12 0.90 0.60 0.60 0.39 0.27 0.22 0.19 0.05 0.60 0.60 031 0.36 0.29 0.25 0.00 0.60 0.60 0.60 0.45 0.36 0.30 O.75 0.60 0.60 0.60 0.52 0.43 C.35 0.70 0.60 0.60 0.60 0.60 0.51 0.41 0.65 0.60 0.60 0.60 0.60 0.55 0.45 l< 50 to 0.60 0.60 0.60 0.60 0.60 0.49 t L NOTES: (1) Fhw depth = a,for a surface flaw 24.for a subsurface flaw t = nominalth.ckness Linear interpolation is permissible. (2) Stress ratio = M(P,+ P,+ P/2.77)/5, Where l rl 5,= allowable design stress intersity (in accordance wtth Section III) e P,= primary longitudinal enembrant stress (P. s 0.55.) ~ P.= primary bending stress P,= espansion stresses resuking from restraint of free end displacement M = 1.0 for sheided rnetat arc weids when 0.D. 4 24 in. = 1.0 + 0.01(0.D. - 24) for shielded metal arc wetes where 0.0. > 24 in. = 1.00 fri submerged are welds where 0.D. 5 24 in. k-s.1.00 + 0.009(0.D. - 24) for submerged arc weeds where 0.D. > 24 in. { (3) Circumference based on nominal pipe diameter. J.. (4)!WS 3514.3 shaft be used. n Note: For the FitzPatrick weld overlay design, the allowable stress, identified as Sm above, was taken from ANSI-B31.1, i the original Code of Construction [29]. I 2-3d l l ll.

It 2 w-th*St ,23

• • I M 23 st et.,

N 88 5E m-3c 5E p- , rg., y ',p ',g 2- ~ =; " n.. 5 4.,..., 3 2. a se ,,.4, It.25 '8' '8 L rt.is 1:1 -19 Itt 84 ozs sea in as as as 33=28 o 26 58 tt'05 bpt Je N a.e l O m.se za.si " ~ * ' ~ - Welds with UT indica' tion 1. Spring, 1985 2. Fall, 1984 3. Fall, 19E2 ] - Weld overlay repairs r L 29-33 j.. 4-46 re sz 2,3 i ze.54 < xs-u ,u l so se r so-st n s re s6

• -45 29 94 4 41 4* 41 4.M 4 48 Fiqure 2-1. Isometric of Recirculation System - Loop A Note:

Abbreviated 151 Weld Numbres 2-3S

7 bi ??) ,g, g l 4' 2. stes\\ ,,. l ,,... Q ** u.ss.3, n ,,.,, m ' 78 22 0a N 2F sg it H 33 13 28 1s El se. gg, 21 01 " s' it.es >a= ..... P 82 61 i)28 49 22 fat 26-iet

28. +o s u,oe

{ el.se 23=st s.94 i'

2-s3 E*'**

s-ses

• • M 4-ame O *'88 4.yy

(... ,<,e u..,, a q a..., I is 25 a,o su I i I l-Welds With UT indications 1. Spr2ng, 1955 2. Fall, 1984 l.., 3. Fall, 1983 - Weld overlay repairs 20 sig to :n-s.es f $ **"* al yze-ne les-ea s 1,2,3 Figure 2-2. Isometric of Retirculation System - Loop B Note: Abbreviated 151 Weld Numbers 2-36

y p,- -s's';,4 N., ss i r" 's

s...

Ill l . E ' a c '. "/ g / 2 no d.i.s 21,! i ,,, \\ / ./ we s-e, s 6-ers i ,.s"* s ' s.I's] "[q ;" \\* ~ se / / 4a /~ D *" i N "$N

)..h.. D,* W l

,','.<.f m - ;' L",,,, ? ~ es N.a. s-3 e as{-),.., l N, N.... x.. ,,,I, ....k*' [, ',,, ..' ~ N 'p,,,,,",,1 v b .n s .n s "j, ds - )$,- ., 'd $d ;,. " < -oX." _., ,.c.

• %,y N<*r c %'
• = Gs

..s l-r ns .m \\ A

• • t a3 a

sei g.., J. v ,w g /' n< in / [ e'/ - PRIMARY C0 ORDINATE AUXILIARY COORDINATE gW~,,s/ 1 \\, ,, k :,,,.',',.,,, s's" s's" j ..,6 F ,r ele, i Figure 2-3. Re,.irculat inn sy, tem Pipinq Moile! - Loop A [4]

915

• • ~

A As / sa .,s l'5 s

1. 1.

~~ 44 M I i g,6 -l I' 'i* S88 - 311 451 WYE / .I. N. >:f" s *. g, o. (15.2M\\ 1 88 f "'9 4 SIO 7,0 - 88 4 L E!!' ' ~4 v ' y.5 /

. [-

, we 33 1" ' , N geg,sA.474 rye - eyv

33. -,

,,,, j ~71,. ** 11/g .,.-i ,,e c W""" g im ' u pas I)..,'a,'j I N kN 's X g SW N.3:e gi sig es4 is 1 .s. , 41 - '-359 its t g e m y t_t**wetf=2 &}$ ,,e g een 6 ~g ~,,, ,gg,,,,. ... 44r W s /.. %y g/> y't ,.. ~ e.... s 4 3,,, y 4, ~ ,',s /y i "Y' ' sie N o/ M &a [ **' 332. M a*T / / 4 ,,o d c Di DINATE PR MARY C0 ORDINATE AUXILI% SYSTEM 1 Figure 2-4. Recirculation System Piping Model Loop B [4]

v- - r o 1-p m v -- WELD k 1.5" l_ 1.5" r I WELD OVERLAY THICKNESS ,i AFTER FIRST LAYER 450 MAX 450 MAX t" f t ]. n r-l I 0.69" .i L I '? O WELD NO. t" (INCH) 12-12 0.47 12-23 0.32 12-64 0.46 12-69 0.38 12-70 0.28 Figure 2-5. Weld Overlay Designs Inch Risers

c - . mg., a l WELD l 4. 2.5" 2.5" i 0.323* AFTER FIRST LAYE 50 MAX f50 MAX f, { \\ I l 0.96" l l 0 MANIFOLD ? END CAP' e-I NOT TO SCALE Note: 1. Put first pass from toe to toe of the overlay. Put second pass on the pipe side to equalize and smooth out I 2. overlay surface. l Items 1 and 2 together represent the first layer to be excluded 3. l from overlay effective thickness (per Generic Letter 84-11). .l j I Figure 2-6. Weld Overlay Design for Weld 22-22 i

50" OVERLAY AT OVERLAY AT WELDS 12-12, 12-23, WELD 12-70 s m 12-64 & 12-69

7,,", -

RPV naa a p 's 4( Db r, a 1. T I l m; 169.6" g P-P. t 0 1. ~ SWEEP-0-LET y 'fil/If RING HEADER Figure 2-7. Finite Element Model of Riser in Recirculation Piping 2-41

q i ";s to le d h -20 Evaluation Curve g -40 T I 64 02 04 0.6 De e s. MM/THICKG Figure 2-8(a). Through-Wall Distribution of Axial Residual Stress in a Large-Diameter Weldment. [2] ri l lI W l insiot unu. outsoc unu. 80 g j l i i i i 4o - no ~ l, ao A. { j s no - lo ~' E -io P -se _ -se _ l l l I i i i i i

1. o o.1 as o.s oe eD all Figure 2-8(b).

Assumed Through-Wall Welding Residbal Stress Distribution in Small-Diameter Weldments (<12 in.). [2] l l 2-42 l

e M Pe M Ps -300 -200 -10 0 0 10 0 200 300 -300 -200 -m o 200 300 400 g W W W W W W W W l3 W OUTER SunpacE / 1 x / MLD .s 3 1 tss eemt g _3 f M LO + 1NW g f h ) f u e0- . I.0 - t t w -l 3W3.3cm) M j i k / - M*I3 O W33.0te=1 5 e i /*5 i N'/ l / t' / 1 ( E / [ { OS-OSa a N A ~' E 1 6 6 j a a a a a a a -00 -40 -20 0 to 40 -40 -40 -20 0 20 40 IESIOUAL ANtat. STIESS, het RESmuel CIRCURAFERENTIAL STRESS, bei cr = -19.22 - 228.40(k) + 856.07(k)2-616.80(k)3 i Figure 2-9. Through-Wall Residual Stress Profile for a Welded and IHSI . Treated 26-Inch Schedule 80 Pipe at a Cross-Section 0.12 inch I (0.3 cm) from the Weld Centerline [6] l 4 )

e M Pa M Ps -300 -200 -soo o soo zoo 300 -300 -200 -soo o mo 200 300 y os-,,,N OUTER SURFACE g os. g N l N WELDifeG (IS.2cm) f, / [ geeG s-o.s-t i O I n / T we w. __r 4 5 N * *" io/ __o3.jj7'"4* "1030cm) / $'i2. ~'O O (I Sicm) E / h E - R S.375m [p#~ lI (13.6Scm) 03 ( E-Jb 02- _oS 02- -os I hs E f o. - y oe-I J O t -40 -20 0 70 40 -40 -20 0 20 40 RESIDUAL AMIAL STRESS, hel RES400AL CtftCUMFERENTIAL STRESS,hel er = - 25.02 - 199.09({} + 922.74({)2-493.05({)3 l Figure 2-10. Through-Wall Residual Stresses Computed at a Cross-Section in the Sensitized Zone, 0.12 Inch (0.3 cm) from Weld Centerline, of a Welded and IHSI Treated 10-Inch Schedule 80 Pipe [6] i l l

i I I I l f \\ e'*A g*As*A2' *b'3 g 3. E F1 t- ) l B d~ F2 y ,2

• ah

,e D l +-- t + 1 t c ] }= 2 s g a,-f77 p, A,. p, t p A,.rty)A,.7,tgiA3; 3 ~ 0 I f O .2 4 .o .6 Fractier.a1 Distance Through Wall (a/t) f. h f L! Figure 2-11. Magnification Factors of Circumferential Crack in a Cylinder (a/t = 0.1) 2 2-45

ee-8 k Upper Bound (Furnace Sensitiza d) da/dt = 5.65x10-9(K)3.07 l / [@ Best Estimate (Weld Sensitizec ) O da/dt = 2.27x10-8(g)2.26 l 'A g ~ g ~ / y sensitized at 1150*F, 2h, 0.2 ppm j 10*8 2 (Heat 04904)(GE - T118-1) Asensitized at 1150*F, 2h, 0.2 ppa 9 0 (Heat 03580) (GE - T118-1) V g 2 5 Y asensitized at 1150*r, 24h, 0.2 ppm { y V 0 (GE - T118-1) 2 QSensitizedseverely0.2 ppm 0 (GE - RP 1332.2, Ref. H-35) d' 2 Osensitized at 1150*F. 24 h, 8 ppm te GE-T118-1 Wang, Clarke-GENED Solomon-GECRD Hasaoki-Hitachi Research ,l @ Lab ~ Park-Argonne Nat. Lab. (Ref. H-36) A Sensitized by welding LTS at [. 932*F, 24h, 8 ppm 0 (SRI Ref. H-37) 2 i i i i so se as se oc ?* _f 8Tmt381NTgasssTV,Eihal 6 8 Figure 2-12. Stress Corrosion Crack Growth Data for Sensitized Stainless Steel in BWR Environment [2] I 2-46

j e" 's I a' ,AN 1WI 1r D L, s' ( a) Self.Similar Assug tion; l'-f = 2(a'-a) i 1 '~ p g g. / N a i b) Constant Aspect Ratio Assugtion; f/a = f'/a' .i r-l 2 Figure 2-13. Comon Assumptions Used to Estimate Circumferential Crack Growth 2-47

1.0 - .9 - .8 _ .? 5 n e .6 5.--

• h vg

.5 0 2

• e ME 0

w$ ~ 1 g= .3 a .2 O'O I .1 t i i i l I l-1 2 3 4 5 6 7 flua6cr of Operat ion Per cepts inn leided in the Average Crack Growth Hate I.alsialation 1 I I _1 1 I I ?,000 4,000 6,000 8,000 10,000 12,000 14,000 Approximate Time Between Inspection (Hours) Figure 2-14. Average Effective Circumferential Crack Growth Rate As a Function of Operation Periods Used in Calculation of Time Between Inspections 4

i FITZPATRICK WELD 12-4 0 STRESS INTENSITY vs. CRACK DEPTH 50 5 o a. 30 - 20 - e 10 - !f 0 a 2 i 2 i ' P l -M-l - t _q. ll -50 0 0.2 0.4 CRACK DEPTH, Inches o P+DW+THERWAL + AS-WELDED o WITH IHS1 g k Figure 2-15. Stress Intensity Factor Versus Crack Depth for J. A. Fitzpatrick Recirculation System Weld 12-4. 2-49

FITZPATRICK WELD 12-4 0.6 m 0.5-0.4-s E g 0.3-o 5 0.2 - 0.1 - l 0 a 0 2 4 5 (Thousands) ,p HOURS I D AS-WELDED + WITH IHSI ? Figure 2-16. Predicted Stress Corrosion Crack Growth for Observed Ultrasonic Flaw Indication, Weld 12-4. ~ 2-50 L

o 1 L-0.9 - 0.8-IWB-3641-1 Limit 3 0.7 - IWB-3641-5 Limit ,m 0.6 ~ 2/3 of IWB-3641-1 Limit .M 0.5 n 1 Cycle (Ass As-Welded Stresses) N ^ _E. 0.4 - t p {0.3- ] 8 g 0.2 - u. a 0.1 - Initial Flaw Size & Post IHSI Case (No Growth) i L l 0 0 0.2 0.4 Flaw Length / Pipe Circumference (f/n0m) Figure 2-17. Comparison of Predicted Crack Growth With Allowable Flaw Size Limits - Weld 12-4 l 2-51 l I c.., ,r

FITZPATRICK WELD 28-53 4-STRESS INTENSITY vs. CRACK DEPTH 4o. m } 30 - 20 - l 10 - 0-e s I h I j i 1 l ' -70 0 0.2 0.4 0.6 0.8 1 CRACK DEPTHInches D P+DW+TNERWAL + AS-WELDED o WITH IHSI f Fioure 2-18. Stress Intensity Factor Versus Crack Depth for l J.A. Fitzpatrick Recirculation System - Weld 28-53 2-52 l

FITZPATRICK WELD 28-53 ) ~ 0.6 w 0.5 - e y 0.4 - 3 i g 0.3 - O Eo4E 0.2 - U 0.1 - 0 [- 0 4 8 12 16 20 24 i (Thousands) F HOURS D AS-WELDED + WITH IHSI l Figure 2-19. Predicted Stress Corrosion Crack Growth for Observed Ultrasonic Flaw Indication - Weld 28-53 2-53

b. t 1 0.9-r l 0.8-IWB-3641-1 Limit }0.7.- IWB-3641-5 Limit L vi i U 0.6 C t 2/3 of IWB-3641-1 Limit 5 y0.5 0 e [0.4-l 1 o a p0.3-3 Cycles (As-welded) i 5 0.2 - 2 Cycles (As-welded) fi E l 0.1 - 1 Cycle (Assumes As-welded Stresses) Initial Flaw Size & Post IHSI Case (No Growth) 0 0 0.2 0.4 Flaw Length / Pipe Circumference (f/nDm) (7 Figure 2-20. Comparison of Predicted Crack Growth With Allowable Flaw Size Limits - Weld 28-53 2-54

FITZPATRICK WELD 28-112 STRESS MTENSITY vs. CRACK DEPTH 40 l 30 - 20 - [ 10 - 0 e ~ e' E j i F -H-P _4o. ';00' -70 3 0 0.2 0.4 0.6 0.8 1 L, CRACK DEPTH. inches A D P+DW+THERWAL + AS-WEl.DED + WITH IHSI Figure 2-21. Stress Intensity Factor Versus Crack Depth for J. A. Fitzpatrick Recirculation System - Weld 28-112 2-55 1 " " " ~ = ' -, - - - _.

r FITZPATRICK WELD 28-112 0.8 0.5-0.4-s i v I 0.3-a x 0 g 0.2-e r; 0.1 - I l l l: 0 i i i i i i i i i i i LI 0 4 8 12 15 20 24 (Thousands) h HOURS o AS-WELDED + WITH IHS1 1 N Figure 2-22. Predicted Stress Corrosion Crack Growth for l Observed Ultrasonic Flaw Indication - Weld 28-112 I i 2-56

e-L 1 0.9 - 0.8 - IWB-3641-1 Limit $0.7- ~ L. IWB-3641-5 Limit $ 0.6 o5 2/3 of IWB-3641-1 Limit ^ u G.5 ^ ^ I _EO.4-3 Cycles (As-welded) R 2 Cycles (As-welded) {0.3- !.} 1 Cycle (Assumes As-welded Stresses) ~ g0.2-y C Initial Flaw Size & Post IHSI Case (No Growth) O.1 - 0 0 0.2 0.4 p Flaw Length / Pipe Circumference (f/nDm) 5. N Figure 2-23. Comparison of Predicted Crack Growth With Allowable Flaw Size Limits - Weld 28-112 2-57

7.= n != U h, i~ l 3.0~ DISCUSSION OF MAJOR TECHNICAL ISSUES 3.1 Weld Metal IGSCC Resistance s l Operating experience with Type 308 and 308L weld metsi in BWR service has indicated that these materials possess inherently high resistance to IGSCC. ) ~ Despite the fact that residual stresses are generally higher in theIweld s ,i. itself than in the heat-affected zone (HAZ) of the pipe wall, no lea'kage has ever been observed to result from cracks propagating through weld netal.

• Recently, however, the intended use of weld overlays for extended plant rc

\\ f l service has promp.ted a more 'congrehensive examination of weld metd1 ' resistance to IGSCC. " The results 'of industry-sponsored ' lab 5ratory in-f ( vestigations have added considerable confidence in the behavicr of weid metal as a crack arrest barrier in the BWR service environment. Addi tionally, these recent test results have provided a more quantitative P3 ' understanding of the relationship between weld metal microstructure and the observed cracking behavior in both field and laboratory examples. r-The recent data confirm that Type 308L weldaetals (less than 0.02 wt% fj carbon) typically used,in weld overlay reprirs are immune to IGSCC when they [ have minimum ferrite contents of approximately 5 or 6 FN. Type 308 weld { metal, on the other hand, would require approximately twice this ferrite I-content for a similar level of resistance to cracking, based rpon limited a. E laboratory test data. Field experience as well as laboratory, data show that 'r f.a g Type -308L weld metal with appropriate ferrite content will consistently [, arrest propagating IGSCC, even under severe ' load and env}ronmental con-ditions. The low-carbon weld :TEtal possesses far greater resistance to cracking than the weld metal chemistries typically used in original plant construction. l Welds at FitzPatrick were weld overlay repaired by the automatic gas tungsten arc welding,tschnique using bare wire ER308 stainless steel electrodes containing 0.02 wt% carbon maximum and 8FN ferrite minimum. In addition, the r l l (. 3-1 1

first overlay layer was not included in the estimate as an IGSCC resistant layer because of the possibility of weld metal dilution and compositional , l changes in this layer. This approach very conservatively provides that a full structural weld overlay consisting of Type 308L stainless steel j containing 0.02 wt% carbon maximum and 8FN ferrite minimum is available to arrest a growing IGSCC. All laboratory and field data developed to date support the engineering judgement that this material is able to arrest O growing IGSCC which may have propagated to the overlay / base metal interface. f 1. A review of the recent weld metal cracking experience, both field and laboratory, is summarized in the following subsections. In this discussion, j the results of research by several different organizations is presented. Some of these studies present ferrite content in terms of %, while others use the term FN or ferrite number. These terms are approximately equivalent. I 3.1.1 Field Experience (A) Weld Metal Cracking in Recirculation Piping at Nine Mile Point Unit 1 I Following the removal of the recirculation system piping at Nine Mile Point Unit 1 in late 1982, metallurgical analyses were performed to characterize the depth and mode of cracking in the Type 316 stainless [ steel base metal. Surprisingly, these analyses revealed that in two of the 28 inch diameter girth weld samples, cracking had penetrated into I the weld metal. Figures 3-1 and 3-2 [12] illustrate cracking which li initiated in the pipe material and propagated into the weld metal. Also shown in Figures 3-1 and 3-2 are the respective ferrite measure-ments in the welds measured in both the horizontal and vertical orientations using a ferritescope. It is seen that the weld metal regions through which the crack propagated in Figure 3-1 were of l relatively low ferrite (3.8% to 4.2%). Figure 3-2 presents photo-micrographs from the second NMP-1 specimen. Again, in this specimen, 3-2 l

n '9 ID, initiated IGSCC in the parent metal appears to have propagated into n weld metal with measured; ferrite levels between 3% and 6%. No data on j the carbon content of these welds is available at this time.

[

It can be seen in Figure 3-1 that the crack has propagated through the approximate mid-plane of a repair weld volume, thus posing questions about the possible contributory role of hot cracking in this weld . i.- defect. Since weld metal microfissuring or hot tearing tendencies are C usually increased in such repair geometries, the extent and location of p the cracking are definitely suggestive of a preferential crack path. Nonetheless, th'e crack appears to have propagated in an interdendritic , f* manner from ID-initiated IGSCC through a substantial amount of weld metal. It, therefore, has the train characteristics of an environ-4 mentally-assisted crack. yb (B) Weld Metal Cracking in Quad Cities Core Spray Line. Metallurgical analysis of a cracked core spray line from Quad Cities Unit 2 [13] revealed axially-oriented IGSCC that had propagated transversely into weld metal. Figure 3-3 shows the interdendritic

l morphology of this weld metal cracking.

Analysis of the weld indicated that the material was Type 308 stainless steel with about 5% ferrite. The analysis further revealed that the carbon content of this material was 0.064 wt%. This observation b provides additional ev m %e that carbon content is also an important factor in the IGSCC n P.ance of Type 308 weld metals with ferrite .j contents of up u 3 e. J C. As will be discussed in the following sections, these examples of field experience are fully consistent with the results of laboratory cracking tests. With the exception of cracking in low ferrite, high carbon weld metal such as the cases above, destructive metallurgical examinations and field experlience fully support the premise of the IGSCC resistance of 308L weld r.atal. s ll t 3-3 )

3.1.2 Laboratory Experience (A) General Electric Weld Metal Tests As part of a test study to evaluate the structural stability of large diameter pipes containing intergranular stress corrosion cracks [14], fracture mechanics (IT-WOL) specimens were fabricated from Type 304 stainless steel plates welded with Type 308 and Type 308L electrodes of varying ferrite levels. The specimens were load cycled in high o temperature water containing. 6 ppm 02 with an initial 4K of 26 ksi (in)1/2 at an R of 0.05, where R is the ratio of minimum to maximum i1 cyclic load. The specimens were on test for 5448 hours. { Failure analyses performed at the conclusion of the tests revealed that intergranular stress corrosion cracks which had initiated in the base metal penetrated the weld metal in six of the seven specimens. In all O but one case, the crack had arrested in the weld metal following some penetration. For the Type 308L specimens containing from 5.5 to 11.5% ferrite and 0.025 wt% carbon, the penetration into the weld was a maximum of 0.031 inches before crack arrest. Branches of the primary M ff crack continued to propagate in the wrought Type 304 along the weld HAZ, parallel to the weld / base metal interface. f.. For the Type 308 welds, the low (1.9 - 3.3%) ferrite and high (9.5 - E 11.5%) ferrite welds exhibited an average penetration of 0.104 inch and 5 0.045 inch respectively, followed by crack arrest. The crack in the j medium ferrite content Type 308 specimen (containing 7.0 to 8.5% i ferrite) penetrated 0.101 inch into the weld metal but. showed no evidence of arresting. The carbon level for the Type 308 SS weld metal f^ was 0.053 wt%. These test results are in agreement with the field experience summar-ized above. Carbon content is seen to be a very significant factor in i 3-4

weld metal cracking resistance in addition to ferrite level. Type 308L weld metal exhibits markedly better IGSCC resistance than the higher-f carbon Type 308. In addition;to the General Electric crack propagation work for weld metal described in the paragraphs above, the Alternative Materials Program [15] provides striking evidence of the ability of Type 308 i stainless steel weld metal to resist IGSCC propagation even in creviced GWR-like environments. i In the Alternative Materials program, plate welded (IT-WOL) fracture j mechanics specimens of the candidate alternative materials (including nuclear grades Type 304 and Type 316 stainless steels) were fatigued and bolt loaded in an autoclave and tested for more than 10,000 hours at 5500F.in high purity oxygenated water. The alternative materials all exhibited extensive IGSCC extension in the -test. Several photo-i K micrographs illustrating the extensive crevice corrosion crack growth in the alternative alloys are presented in Figures 3-4 through 3-6. Note in Figure 3-4 that although the original fatigue terminated in weld metal, no significant crack extension was observed in the weld 'y metal, whereas substantial IGSCC crack extension occurred in the Type 316L SS sample (containing 0.026 wt% carbon). Figure 3-5 also shows that crack penetration into weld metal was minor, whereas substantial IGSCC growth occurred in the Type 316NG SS material. In Figure 3-7, a crack which inadvertently was grown into weld metal was observed to arrest with no measurable crack extension. Although the initial stress 7 intensity on this specimen was only one-half of that for the alternative material specimens, 25 ksi(in)1/2 versus 45 ksi (in)1/2, the fact that no crack growth was observed provides striking evidence of the excellent resistance of weld metal to IGSCC growth. It is believed that several of the weld filler samples were Type 308L SS with 8% ferrite minimum, as that was the weld filler specified by General Electric at that time for plant piping. This has been confirmed by General Electric Company [16]. 3-5

(B) -Inverse IHSI Pipe Tests q]. As part of recent EPRI efforts to detect and size IGSCC in austenitic stainless steel pipe ' welds, a group of 12-inch pipe samples of Type 304 material were fabricated by Ishikawajima Harima Heavy Industries [17]. .These specimens contained girth welds and were inverse-IHSI treated so q - as to produce deep IGSCC when exposed to high purity, oxygenated, 5500F [ water. One of these samples developed an intergranular stress corrosion crack which penetrated the pipe wall and extended several millimeters into the weld. The pipe specimen was metallurgically examined for level of sensitization and for ferrite content. The examination revealed that the weld metal was highly sensitized (prob-i ably due to a 5000C/24 hour LTS treatment). Further, the weld metal was determined to be Type 308. The weld metal cracking was observed to terminate when the direction of the dendrites made an abrupt change. The initial weld metal crack propagation occurred in approximately 5% [E-ferrite material and appeared to terminate in approximately 9% ferrite material. (C) Weld Overlay Repair Large Diameter Pipe IGSCC Crack Growth Test .Ll The BWR Owners Group (BWROG-II) and EPRI (EPRI Project T303-2) [18] are i' sponsoring an IGSCC pipe test program examining the effectiveness of residual stress remedies in retarding or arresting IGSCC crack growth ]< in large diameter Type 304 stainless steel pipe welds. A total of two J 24-inch diameter 1.2 inch wall thickness, Type 304 stainless steel 7 pipes each containing two test welds were tested. Dr.e of the pipes was ,J IGSC cracked by loading to an axial load of approximately 18 ksi in 5500F high purity water containing 6 ppm 0. The IGSCC pre-cracking 2 required approximately 4000 hours under test at load. Following the 4000 hour pre-crack exposure, this pipe was returned to test and crack growth occurred in both joints over an additional 1 approximately 6000 hours on test. The pipe was removed from test, the 3-6

l crack locations were documented by UT and PT, and a full structural weld overlay was applied to one of the joints. The weld overlay was designed ] to be approximately 0.29 inch thick. j At the present time, the weld overlay tested joint has been on test for 3000 hours since the weld overlay application. No additional IGSCC initiation has occurred in this weld since the application of the weld E overlay, as observed by UT and PT. Furthermore, no apparent change in crack depth has occurred in the existing IGSCC since the weld overlay was applied. The companion as-welded reference joint has accumulated approximately 9000 hours on test since the IGSCC was first observed. During that period of time, the deepest cracks have grown to approximately 250 to 300 mils in depth. The crack growth rate is slowing measurably as determined by UT and acoustic emission and the deepest cracks appear to be arresting. Additional ID crack initiation and lengthening of previously initiated cracks is taking place in this reference weld as determined by UT and liquid penetrant measurement. (D) EPRI/ General Electric Pipe Tests Another part of the BWROG-II remedies and repairs program (T302-1) [19] is being conducted at the GE pipe test facility. Pipes of 4 and 12-inch 1 diameters are being pre-cracked under exposure to high stress in simulated BWR conditions. The resultant pre-remedy IGSCC defects ranged from 10% to through-wall penetration. Various specimens are

L,t then treated with weld overlay, IHSI or LPHSW remedies. The tests are designed to measure the effectiveness of these remedies in arresting the growth of pre-existing IGSCC.

Two of the 4-inch Schedule 80 pipes have received weld overlay repairs to several joints following pre-cracking by IGSCC in a simulated BWR environment containing 0.2 ppm 02 at 5500F. The pre-cracks were sized 3-7

I. 1 Il ~ by UT, and weld overlay repairs were performed in both pipes as presented in Table 3-1. The pipes were then returned to test in 9 ppm t-l[ 0, 1.5 pS/cm water at 5500F, and loaded to 1.3 Sm following the weld 2 overlay application. The two pipes have accumulated test exposures of l 11,112 and 8,406 hours on test respectively with no failures having been observed in any specimen in either pipe. (A failure is defined as a leak at a specimen location). L It is noteworthy that during application of the weld overlay repair to IGSCC pre-cracked joints in both pipes, through-wall cracks were observed at two joints. One joint containing a through-wall crack and a weld overlay repair (pipe RSP-14, 0.5T overlay in Table 3-1), was removed from test following 1000 additional hours on test with the weld j overlay repair in place. Optical metallography performed on this d specimen revealed that the weld overlay effectively arrested the through-wall IGSCC at the overlay / weld metal interface as is observed E inFigure3-8[19]. In addition, two 12-inch pipe specimens have been IGSCC pre-cracked and have received weld overlay repairs. A total of nine weld overlays have R vi been applied to the two pipes. The pipes have accumulated 1,321 and 1,057 hours on test since' the weld overlay repair applications. No [ evidence of crack growth has been observed as of the publication of [20]. Although these tests have been performed primarily to address the 3 issue of residual stress benefit of weld overlays, they are included L nere for compieteness, l 3.1.3 Modelling Studies Based upon the preceding in-reactor and laboratory IGSCC studies, it is clear that Type 308 and 308L stainless steel weld metal exhibit superior IGSCC resistance compared to wrought Type 304 or Type 316 stainless steels. However, the field and laboratory data also iilustrate that Type 308 stainless steel weld metal is not immune to IGSCC. The data from NMP-1 and ~ Quad Cities 2 and the data at GE, IHI and Battelle Northwest suggest that 3-8

ferrite levels of 5 - 9FN alone may not be sufficient to eliminate IGSCC propagation into weld metal. However, the crack arrest data on Type 308L SS weld metal suggest that combinations of carbon and ferrite level (and I probably ferrite distribution) do exist where IGSCC crack propagation in austeno-ferritic weld metal is either extremely slow or nonexistent. Devine [21] performed a laboratory study investigating the interaction effects among carbon level, ferrite level and ferrite distribution on the IGSCC susceptibility and sensitization immunity of Type 308 SS weld metal. This work involved not only the SCC testing and microstructural char-acterization of Type 308 welds, but also included such studies on wrought Type 308 compositions. In this context, it is helpful to review these studies and the mechanism by which increasing ferrite mitigates IGSCC in austenitite stainless steel. In [21], the beneficial effects of ferrite content in Types 308 and 308L wrought and weld-deposited compositions are discussed with regard to IGSCC susceptibility and sensitization immunity. As is generally believed, zones depleted of chromium due to the precipitation of chromium carbides during welding or furnace sensitization act as sites for potential IGSCC. Whereas f chromium carbide precipitation occurs intergranularly during aging of austenitic Type 308 stainless steel, no such precipitation occurs along austenite-austenite grain boundaries in duplex Type 308 containing suitable amounts and distributions of ferrite. Instead, the precipitation occurs exclusively along austenite-ferrite phase boundaries. Since chromium .J diffusivity is approximately 1000 times greater in ferrite than in austenite at 11000F, the chromium for this precipitation is supplied principally from ,j the chromium-rich ferrite phase. A small zone of chromium depletion in the austenite is subsequently replenished by chromium diffusing from the f interior of the austenite. Af ter this " healing", the material is immune to intergranular corrosion in ASTM A262 Practice E and IGSCC in air-saturated water at 5500F. Models have been developed [21] based on the above mechanism to describe IGSCC as a function of carbon content and the amount of ferrite-austenite 3-9

l. i (a-y) boundary area. The critical distribution of boundary area for rapid healing is that amount which is sufficient to tie-up all of the available I carbon as chromium carbide exclusively along ferrite-austenite boundaries. I Both the amount and distribution of ferrite-austenite boundary area can be expressed as a function of the metallographic parameter, N " . This is a L I measure of the number of intercepts a random test line makes with a-7 3 boundaries per unit length of test line. 1 Figure 3-9 shows the model predictions of NL , as functions of %C to maintain a critical amount of a-Y boundary area (line Sy*Y) and a critical Y 8-distribution of a-Y boundary area (curve A ). The value of NL for rapid I healing is the higher of the two curves; however, the straight line model for Sya-y has been used [21] to describe IGSCC resistance over the complete range of carbon content for Type 308 compositions. ASTM A262 Practice E results shown in Figure 3-9 verify the model predictions. At 0.03%C, the maximum for of 100 to 200 cm-1 is required for rapid healing and Type 308L, NL W immunity to intergranular corrosion. This translates to a ferrite level of about 3 wt% as shown in Figure 3-10. At 0.015% carbon, or less, essentially no ferrite is required to confer imunity to intergranular corrosion. f 3.1.4 Model Results The Devine model, in combination with field and laboratory data on the IGSCC behavior of Type 308 stainless steel weld metal, has been used as a guide to specify the maximum carbon level and minimum ferrite level necessary in Type u 308L stainless steel weld overlay material to resist IGSCC growth into the weld metal. The model indicates that 0.03 wt% carbon and 3% ferrite should ij be sufficient. The laboratory test data confirms the model result and the l field IGSCC data in Type 308 stainless steel weld metal are not in conflict f with the model estimate. However, to provide additional conservatism and to allow for some weld metal dilution or compositional estimating error, the BWR industry has generally specified that the weld overlay material Type 308L stainless steel contain no more than 0.02 wt% carbon and 8FN ferrite (approx-( imately 8%). 3-10

f L. 3.2 Residual Stress Benefits Since their initial use on BWR pipe welds, weld overlays have been analytically shown to produce beneficial residual stresses in a variety of pipe sizes and joint configurations. Such analyses typically employ finite element thermal / stress modeling techniques to predict the behavior of the pipe material undergoing repair. Analytically, the application of a weld overlay repair is shown to produce highly compressive residual stresses through a major portion of the original pipe wall, thus effectively arresting the growth of pre-existing IGSCC in the pipe material. 1 More recently, a number of laboratory programs have been undertaken in order to verify experimentally the effectiveness of weld overlays in arresting the growth of pre-existing cracks under BWR conditions. Furthermore, several weld-overlay repaired pipes have been removed from plants and destructively analyzed after operational plant service with the overlays. The results of [4 these new programs, coupled with previously reported data, provide over-whelming evidence that weld overlay repairs, in addition to being resistant to IGSCC crack propagation, will also arrest further crack propagation in the original pipe weld, both in the through-wall depth direction, and in the R crack length direction. l The following subsections summarize some of the recent developments on the topic of weld overlay residual stress / crack arrest capability in the original piping material. L 3.2.1 Georgia Power Company / Structural Integrity Associates (SIA)/ Welding p J; Services Incorporated (WSI) 28-Inch Notched Pipe Test The objective of this project [22] was to varify the analytically-predicted residual stress benefits of an overlay repair on a large-diameter pipe weld joint with pre-existing defects. The test piece included a number of crack-like defects of axial and circumferential orientation, in order to examine the post-remedy residual stress state at the extremities of pre-existing flaws. I 3-11

F Two sections of a 28-inch diameter,1.5-inch thick Type 316 stainless steel i pipe were welded together using a joint configuration and welding procedures [ typical of those used in the original recirculation system piping fabrica-tion at Hatch Unit 1. Following the butt weld, a bottom plate of stainless steel was fillet welded to the pipe, so that the pipe could be used as a self-contained boiling magnesium chloride (M C1 ) residual stress test (Figure 3-g 2 11). A stainless steel baffle plate was fillet welded to the bottom plate and b to the inside surface of the test pipe, so as to divide the test pipe into two 7 equal halves. Axial and circumferential notches of varying depth were ground into the ID of the pipe at various locations near the girth weld. The notches were introduced syninetrically in both halves of the pipe (Figure 3-12). One half of the pipe was exposed to the boiling M C12 following the g introduction of the notches. A structural weld overlay was then deposited over the outside surface of the entire girth weld and the entire pipe was re-exposed to the M Cl2 solution. The pipe was liquid penetrant inspected and g sections were removed for metallurgical analysis following the liquid penetrant examination. A typical result of this testing is illustrated in Figure 3-13. Figure 3-13a shows a metallographic section of the tip of a moderate depth circumferential l notch which was exposed to M Cl2 testing prior to weld overlay. The g extensive cracking indicates the high level of tensile residual stresses l present at this location. Figure 3-13b shows a similar metallograph of the corresponding notch in the section of the pipe tested following weld overlay. No M Cl2 cracking is apparent at the second notch tip, indicating that the g weld overlay process reduced the notch tip residual stress to near-zero or compression. Essentially identical results were observed at every notch g illustrated in Figure 3-12, both axial and circumferential, deep and shallow (i.e., extensive cracking in the notches tested prior to weld overlay and no cracking in the notches tested after weld overlay). These results and the f entire test program are described in detail in [22]. The tests thus confirmed that tensile residual stresses were present in the vicinity of the axial and circumferential notches in the as-welded pipe. The tests also l I l ( 3-12 1

I I; provided striking evidence of the effectiveness of weld overlays in producing compressive ID surface and through-wall residual stresses. The l weld overlay produced compressive residual stresses at the tip of both axial and circumferential defects; this was true in the case of both shallow and j deep notches. In addition, the overlay was shown to produce ID-surface compressive residual stresses that would prevent any length extension of ( pre-existing IGSCC. I i 3.2.2 EPRI/GE Residual Stress Results The EPRI/GE Degraded Pipe Program, examining the effect of residual stress li remedies on the crack propagation behavior and residual stress distribution of IGSCC pre-cracked 4-inch and 12-inch Schedule 80 pipe, has been underway for approximately two years [19, 20]. The program has consisted of pipe tests, residual stress measurements by boiling M Cl2 and finite element g analysis examining the effect of a weld overlay repair on the IGSCC crack growth in these pipes. The pipe test results were discussed in Section 3.1.2

1. D of this report. The residual stress results are described below.

g A total of four 4-inch pre-cracked specimens with girth welds and containing IGSCC pre-cracks which were estimated to have depths as great as 60% of the wall thickness were weld overlay repaired using the parameters presented in Table 3-1. The IGSCC pre-cracks were produced by exposure of the welded pipe specimens to 5500F, 200 ppb 02 simulated BWR water. Following the introduction of the pre-crack and the depth measurement by ultrasonic techniques, the pipe samples were weld overlay repaired and exposed to { boiling M Cl2 to determine the residual stress state. Following the M Cl2 g g '1 tests, liquid penetrant measurements were performed on all pipe samples exposed, including those which had not received a weld overlay repair. The unrepaired welds exhibited extensive M Cl2 cracking, while no liquid g penetrant indications whatsoever were observed on the ID surface of the weld overlay repaired joints. Not even the IGSCC pre-cracks were observed by penetrant examination following the weld overlay application. l t 3-13

i 3.2.3 Nutech/ Georgia Power Company 12-Inch Weld Overlay Mockups Georgia Power Company, in conjunction with Nutech Engineers [23], fabricated two weld overlay test specimens in conjunction with 1983 repair activities at Plant Hatch. A total of three specimens were fabricated, one each for a 0.20 inch overlay, a 0.23 inch overlay and a last pass heat sink weld (LPHSW). The weld overlay lengths were 4 inches. The weld overlays were applied to butt welds in short sections of 12-inch, Schedule 100, Type 304 stainless steel pipe using the same procedures, operators and equipment as were used for the in-plant repair work. Analytical predictions of residual stress based upon finite element studies were compared with measurements on the actual samples. The calculated and measured axial residual stresses are presented in Figures 3-14 and 3-15 for both weld overlay repaired pipes. One observes from these data that both the calculated and measured results indicate that the inner half of the repaired sections are in axial compression. The calculated residual stress results, however, are less compressive in general N than the measured results. 3.2.4 EPRI/J. A. Jones 24-Inch Weld Overlay Mock-up A 24-inch Type 304 stainless steel pipe having a wall thickness of 1.48 inches received a weld overlay repair at the J. A. Jones Applied Research Center as part of an EPRI/BWROG Il sponsored effort to examine the effectiveness of the weld overlay in providing favorable residual stresses in large diameter pipe. The overlay consisted of a total of five weld layers L constituting a total thickness of 0.35 inch. The overlay process was modelled by Nutech Engineers using the WELDS-II elastic-plastic finite element program [23]. The experimental ID residual stress measurements performed on this pipe following the weld overlay repair are presented in Figure 3-16. The through thickness analytical results are presented in Figures 3-17 and 3-18. The results of this residual stress analysis and measurement project illustrate that both axial and circumferential residual stresses are compressive at the pipe inside surface following a weld overlay repair of this thickness to this pipe. Further, the analytical results show 3-14 j

i-that the residual stresses are expected to remain compressive to a depth of 50% to 70% of the composite wall thickness. 1 3.2.5 -EPRI/BWR0G II Pipe Tests The EPRI-funded projects RP T302-1 and T302-2 were discussed in the previous section [18 - 20] in regard to the evidence they produced in support of weld metal cracking resistance. These two laboratory programs are also mentioned here, because of the significance of some of their results in terms of s residual stress benefits: f T302-1 In general, weld overlays have been shown to be very effective in arresting the growth of pre-existing IGSCC in 4-inch and 12-inch specimens. In one case, a through-wall crack was effectively arrested during a 1000 hour test; no increase in crack length or depth was 7] observed during the test. These tests, including 12-inch specimens, N are continuing. T302-2 The 24-inch weld overlay specimens with pre-existing IGSCC has shown no detectable crack growth after more than 3000 hours under test. al These tests are important in that they provide some of the most realistic { weld overlay residual stress test data available, outside of actual BWR in-plant repairs. The data indicate that weld overlays are very effective crack-arrest remedies, even under the severe stress and environmental s conditions of these tests. 3 3.2.6 Destructive Assay of Hatch Unit 2 Overlay Specimens at Argonne National i Laboratory (ANL) Two weld overlay-repaired pipe-to-elbow welds were destructively examined at ANL [24]. The welds had been overlay-repaired as a result of UT indications during ISI at the Hatch Unit 2 facility, and were then returned to service for l approximately one fuel cycle before removal from the plant. As will be 3-15 L l - =

fa 4 f discussed in detail in the next section, the ANL work was largely concerned with the NDE aspects of the weld overlays. However, several observations f that relate to this section were made during the ANL examinations. During metallographic sectioning of the welds, it was discovered that the application of the weld overlay had " blunted" deep cracks. There was no evidence of tearing or extension of the crack beyond the blunted region, which marks the crack depth at the time of the application of the overlay. The ANL report further states that finite element analyses predict that crack growth will be inhibited by the overlay application. This example of crack arrest (including the case of a very deep crack) was established by destructive assay of an operational pressure boundary repair. As such, both the reliability of this data and its importance in the technical discussion is great. 3.3 Non-Destructive Examination As noted in the foregoing sections, the preponderance of both field and laboratory data with stainless steel pipe welds has shown that weld overlays are very effective in establishing favorable material and residual stress conditions that arrest further growth of IGSCC in the BWR service environ-ment. The material properties and configuration of the weld overlay, however, pose additional challenges for effective volumetric inspection of the original pipe weld joint. )L Recent developments in inspection technology have yielded significant p li improvements in through-overlay detection and sizing capabilities. In addition to repeatable through-overlay detection of deep cracks, these recent developments appear to offer the capacity to examine reliably the integrity of the overlay volume itself. These inspection developments have recently been successfully applied to inspection of weld overlays in several BWRs. 3-16

t a n The following subsections summarize some of the more significant new developments. 3.3.1 Recent Developments at the EPRI NDE Center Workshop on Weld Overlay j Inspections (RP1570-2 [25]) A number of 12-inch, Schedules 80 and 100 pipes containing IGSCC as i well as machined notches were repaired with weld overlays and then used for UT procedure development. In addition to standard cali-l bration notches, several types of deliberate defects were induced in the weld overlays, including lack of fusion, porosity, and cop-per/Inconel contamination-induced cracking. Automatic (Intraspect and UDRPS) as well as manual scanning techniques were included in the tests. It is clear that significant enhancements in UT weld overlay inspection capabilities were achieved during the development testing. Among other things, the tests demonstrated that defects in the outer pipe wall (i.e., outer 1/4 T) as well as the overlay volume itself were reliably detectable. Longitudinal-wave UT demonstrated the ability to overcome much of the ultrasound attenuation and scattering characteristics seen in weld metal microstructures. i The tests showed that deep cracks (in the outer 1/4 T) could be manually or automatically detected and sized through overlays with Li excellent repeatability. Further, contamination-induced cracking and lack of fusion defects in the overlay itself produced indications with high signal-to-noise ratio and were detected. The test results indicated that ASME Section XI Code-rejectable defects could rou-F tinely be detected in the overlay itself. Conversely, the tests indicated that detection and length sizing of shallow cracks or the crack mouth-ID/ surface interface of deep flaws do not appear to be reliable at present. The compressive strains induced by the overlay are such that even 0.015-inch width machined slots are closed after 3-17 t

weld overlay repair application. Such powerful crack closure forces l appear to change the UT response of defects in areas with high levels of compressive residual stress. FitzPatrick is planning to re-inspect the weld overlays during the l 1987 refueling outage using the recently developed techniques. In addition, to develop effective inspection procedures and train personnel in their use, the Authority has procured weld overlay I' specimens from an operating plant, as well as laboratory-prepared specimens which were fabricated using procedures and processes ~ similar to field practice tests. The procedures will be demonstrated and personnel will be trained on these samples prior to the actual inspection. 3.3.2 Argonne National Laboratory (ANL) UT Inspection Workshops Two informal NDE workshops were conducted at ANL during May 1984 and N-January 1985, using weld overlay repaired pipe samples removed from the Hatch Unit 2 recirculation piping system [24 & 26]. The samples I included two 12-inch diameter pipe-to-elbow weldments and two 22-inch diameter pipe-to-end cap weldments. All four of these weldments had h! been weld overlay repaired and then returned to service for approx-imately one fuel cycle before removal from the plant. The samples were subjected to a wide variety of tests including RT, i UT, PT, and destructive metallurgical examination. RT and UT proved ~ difficult to apply, but the PT and metallurgical examinations g indicated only a limited amount of cracking in two of the four lL weldments (one of each type identified above). As a result of the I limited cracking, the emphasis of the NDE workshops was on trying to ~ understand the nature of overcalling cracks and the distortion of I ultrasonic waves due to the presence of weld overlays. f Various inspection teams were involved in the workshops, and the i tests included a wide variety of UT techniques. The results of the li 3-18 j (! l b h

.[

l* t t I tests led to a number of observations / conclusions by the workshop participants, the most salient of which are summarized below. It was shown that cracks present before overlaying the pipes could be relocated by UT. The tests confirmed the advantages, discussed above, of longitudinal over shear waves for inspection of pipes with overlays, and also yielded recommendations on transducer frequencies L in order improve signal to noise ratios. Finally, the destructive r examinations yielded the conclusions, noted in Section 3.2.6 above, regarding crack tip blunting and crack arrest following weld overlay application. All in all, the ANL program is not in serious technical disagreement with the EPRI NDE Center tests summarized above, particularly in the case of cracks that have significant depth. The latest longitudinal- ' f.., wave techniques can penetrate the dendritic overlay microstructure c') into the outer pipe wall without unacceptable attenuation or scatter. 'l 4, ..y ia i 3-19

i TABLE 3-1 j FOUR-INCH PIPE WELD OVERLAY PARAETERS PIPE RSP-14 Am RESIDUAL STRESS MOCKUPS PlPE RSP-14 Four-inch Pine Par==atars Bead Bead Overlay Overlay 5-Weld 12-inch Thickness Width Thickness Length hfald Position Overlav Thicknamn M M ( I n. )( a) (In.)(b) A 5G 0.37t' = 0.254 0.03 0.37 0.125 2.0 8 5G 0.37t = 0.254 0.03 0.37 0.125 1.0 C 5G 0.37t = 0.254 0.03 0.37 0.125 2.0 D 5G 0.5t = 0.343 0.03 0.37 0.169 2.0 E 2G 0.5t = 0.343 0.03 0.18 0.169 2.0 t F 2G 0.37t = 0.254 0.03 0.18 0.125 2.0 G 2G 0.37t = 0.254 0.03 0.18 0.125 1.0 H 2G 0.37t = 0.254 0.03 0.18 0.125 2.0 MOCKUPS I Four-inch Pine Par---tars HAZ 12-inch Pipe Bead Bead Overlay Overlay R PTL Pre-Weld Overlay Thickness Width Thickness Length Speciman Mald Cracks Position Thicknats (In.) ( In. ) (In.)(a) (In.)(b) AWC-3 H H1,H2 5G 0.37t=0.25 0.03 0.37 0.125 2.0 l AWC-3 K K1 5G 0.37t=0.25 0.03 0.37 0.125 2.0 DE-8 F F1,F2 5G 0.5t =0.34 0.03 0.37 0.125 2.0 DE-8 A A2 5G 0.37t=0.25 0.03 0.37 0.125 1.0 [.. it (a) After grinding final overlay surface smooth l; (b) Length not including 3 to 1 taper at each end b. 3-20 i I

._.7,.. o E2 '375 O 7' O "Ci 'l 23 4 l llll illlilimlill illi i lu lliiiullnini lii niiniiinulinill l ii ot:p' &. V n o 'M/.k'D- ~* ~ @ ',i ,, -st.. n tf : 'Tg ~ b L. ^ y ^) j ,( ~ FAR g Lr Nos. Hor. Vert. I 1 0.7 0.85 2 2.1 3.0 3 3.6 4.1 4 4.8 5.2 5 8.3 8.4 6 9.0 7.3 7 3.8 3.4 l l,., 8 4.2 4.0 i 9 5.9 6.1 [ figure 3-1. Cracking in Weld Metal of NMP-1 Recirculation Line. Ferrite Levels Are as Presented in figure. [12] l l 3-21 l l l \\.

1 i

_.04

-2: 3 +

.l A 'c$...4 f t' I '. t. ,4 9:-$. 'T

• E5. f

~ .s-. 4 p. ?q h &,,uT %.? .,;j - ' ~ - , :, 'k.3., a-e 'y l T, t, :- d ,.4 ri .?'

• 'N i; A f:

s. 2 Y

.o FA Nos. Hor. Vert. q 1 1.9 3.2 2 5.4 6.3 3 5.5 6.0 4 5.5 5.8 5 5.0 5.5 6 7.3 6.9 7 6.5 7.0 l l. I 8 9.2 8.8 9 8.6 7.2 1 Figure 3-2. Weld Metal Cracking in NMP-1. Ferrite Levels Are as Presented in Figure. [12] 3-22 1 i

.j o i i i 6 ( ....- sg <=.,.. ;. <,.7.: :2 4*;-{..-y,/. i. ~.Y1 ~.g-( ~.9 ' 'r 47,:. py. .,. 3., s j q T, p . s. ~..:.,.#.f.\\ w, */ e< i...

.,.. s,.*

. r - - *.u ~,.........e). y <. 1.. ~.,.~~,,.. !a.. - s. .,a ,k[,,. 4[<'...,.l',..h'y\\'7-v.. ~.a.... ;... :....... :., .7 .p. ? .p-.. ;;....s. . o. c... ;.,... i, .., :.., s .=. - .- ;( ~. e s' ..w,. y.. .,~ e.:. o. s..- (. <.;...~.,- ,s* . s. ?..a... ;. o. L: x...;., s . a.. .n e% . *p. 4.... ~p. .c s. e s.,.

• o W ;. h. ;..m.,., '. :~.100 p m W

.'.

• Q, ?> A..'..

s *,< g.'..... ;

.,~.,....

>>c.. I 4 2 i Figure 3-3. Subsurface Crack Present in Weld Metal in Quad Cities Core Spray Line D3] 1 I 9 f i 1 I 3-23 r l 1 r

3 j 3 .I I ..kM ' ' :NI% ' S .k

• d d, k((.

.,.g. (A, ,.', ),,f', A. [;*,b'- ~ N ~ e, 4 geni- ... p, t, 4.,,.. ..\\ g ,W.a"' s ,..s l .3 ~ , l* j_ ' ' %,'s.f'Q

• '(

\\ ' ~ - s, lt x. ~ s y E . \\ s",- y o ~,(* O,? 3 , y,_ j, TE57 ENVIRONeetNT 590'F (200'O.8 puun D15501.VEO ONYCEN DEIONIZED WATER Figure 3-4. Cratking Morphology of Bolt-Loaded WOL Specimen EAl of 316L Stainless Steel (26X), Heat 9662 L15]

lo ii .Id.df is

y * 'y,..k.

• Q i

' *r a 4 b ..f+$ '.d,G.i./ 15g '.q+7.%,P;w *e "' *,a,* ', O ^ s .4 s. y.t /.i . }f' Y

g ' { f f,_ A,' g. y l

q- , s i l-f x,. 4 e . <c f

• b

/ ,....; e, . /.,. y m1 s s. ir gY s.. y. . ~ 4, g _' _.,(jg, k. -3 ),*' (., _- c.y :._

. cw

• y. f -

I test tNvinoNu ~T uo'r iree'ct :- oissotvro oxvGe, osioNizio n Avsn Figure 3-5. Cracking Morohology of Bolt-Loaded WOL Specimen EB1 of Type 316NG Stainless Steel, Heat TV0076 (26X) j [- [15] 'l W i.'. M 7 e'T. ,%s R s

• . '\\

gr f 4s r lg &, _. ~

4.QX #

4 .,, ~ 1 ~~ g ess. i r .? e ,J

k.

• 4

../' g l 64 6 'f TEST ENvinohutNT po'F (250'C).e opm DissOLvtD ONYGEN D(10Ni2f D W Af tn 1 Figure 3-6. Cracking Morphology of Bolt-Loaded l WOL Specimen EB2 of Type 316NG Stainless Steel, Heat TV0076 (26X) [15] l i 3-25 l 1 I i

1-i l i l i l T.i

• *p 7,,

./) 2 '

] .\\ g.. ,.. i.. (,M., k% e. _

s. -

1 , - i ;.';.h .s. ' '.' 4 e * ..y!.,y 'I M

f. +' *

, >' ',. 6 g y ;"..%. Q.<. .gt .c. lg .a .s ' s- ., E.h }' ' g> 'Y$.'- ~ " W.2 A.x u.-' ~

8..,. _r.,

I TasT suvinomutur. soo'r im*ci. e oissotvto oxyct= ossomno utta L.. Figure 3-7. Cracking Morphology of Bolt-Loaded WOL Specimen EJ1 of Type 304 Stainless Steel (26X), Heat 46436 [15]

• 9

,3 l l 3-26

, y i 'i 'm',>:,W*' x )1 v.>.r },,'Qy', } ; t* ts<. t r 1. p' '*A-3 ,q ,.. s. o . $ W.,3 n g. t/,,g.t* ,.;e. ~ L,1. (> &_.....,,g. e $.?,. \\. s,'.gs,,' fb-X?& y'.'-./ 5...., .g. \\ ? ..u ?. ', (, ~j ' & ; e.,, 3

.~

Q'. . v ~., ..c

; '.n

,\\ .:- u. i. ' i,2. ; ', l_ < M. 4, '.% -f.~....r .r u t s - [,,,j, '.. V~ jJ } b 4. .. = *

z. !{-

9,, ' . f' ' t g,..,. "%]p.sj 4gs.,,.! '.e s g g ,.., \\., .M'li. - s

l.

"i.) -G4,$,,,_... ~ ' xf'? t J. s,- s n-i h:,LL.hl?.. L W l l !d l ' yt u l p, e ^1 i .2

r i

\\ l ~ Figure 3-8. Weld Overlay Arrest of IGSCC Specimen RSP-14 [19] l l i. 3-27 1 l 1

i S,"*7 s= - A G 8 O [ h " CRIT 1 e r t e 'j 200 O O k O *accat O WELD 100 I I I I O 0.02 0 04 0 06 0 08 0 10 CARBONf%) m The Influence of N a-y on the Intergranular Corrosion Figure 3-9. t I. Behavior of Aged Samples of Wrought and Weld-Deposited I Type 308 Stainless Steel. Open Symbols Indicate IGSCC per ASTM A262 Practice E Testing; Closed Symbols Indicate No IGSCC [21] L. 1 ir 1 3-28

t-s ~e t t i ,c t t ^ o 1.3 0 i 1.2-1.1-c - As-Deposited-i t Ty'pe 308 Weld 1.0 - 4,.' Solution Heat Treated Wroug'it ] 0.9-Heats of Type 308 Composition 1 0.8-r1 e NC 0.7-I wa,$ 08-Ob-0.4-e 0.3-o ? 0.2- '~ /s e, 0.1 - r-m 0.0 t.. 0 4 E 12 18 20 24 Off ,32 RMTE DCLWE 5) i i1 7 Figure 3-10 Numt,0r, of Intercepts of a Random Test I.ine with Aur,tenite-Fers ite Boundi. ries per Unit Length of Test lire, N, Versus L Volume t Ferrite for Type 308 Compositions 3-29 I eA,

1-

)

4 Sealed 5.5. Lines For Cold ,f' Baffle Sheet to / City Water To Separate Halves of / Condense MgCl2 Funes n Pipe Against M Cl2 g i Fumes Stainless Steel (5.5) Condenser Plate (Leave N t Being ested Hole For Pressure Relief) g g Molten McCl2 Heaters fl \\ 'l ; ' ~ ~ Around Outside si o I [O Pipe Butt Weld ll ~, O j o O si o ~, L' w., % 5.5. Fillet-Welded 'l'- O ( 6 Bottom Plate to Support i

• ='r.'.0=~#[

g f M Cl2 Molten Salt ei g i i 1.'. 3 !l 0 Hot Plate At 150 C L.a lI l1 Figure 3-11. MgCl2 Test Set-Up using 28" Pipe As Vessel. One Side of Baffle (Half of Pipe) Tested Before Weld Overlay, Other Side After Weld Overlay l-3-30 ,-n, c, a ~

e

~ --

y ap-* j Cire. Asial notches.51 ~ ~N th Circ. Notch s.m = a t es o.25-Iong fcenter on ,,,,,,ot<, + 0.1* from 5,, e w 1 D usion line W, M g r klI J \\ Circ. Notch / Circ. - .351 .6t - Notch.15t .35t Pipe Butt Weld I i .15t .I \\ E I \\ ID.. - \\ l Asial Notch \\ of. '\\, i "E \\ g .51 deco q f 2t long q mwamiMMEMB ....~~to e rz '-w file o separate top & / N, g,ggy, bottom flalves of pipe from \\ N-Sh MgC12 fumes (Case stainless '\\ N,eet g / e / Corners Can steel sheet) W N \\ / \\ be Rounded ] \\ / / \\ f.. \\ \\ x q ( \\lllT A I i "l 2 /f __j, /. Side @ Side view (s) Cnd View of 28" of Pipe Dia. Pipe (5 thematic) s Figure 312. Circumferential and Antal Notch Stres and Locations (Bottom Half is Mirror Image of Top Half of Drawing) - Showing Stainless Steel Baffle to Permit Separate Testing With M Cl2 of Top and Bottom Halves of Pipe g (Seal Against M C12 Fumes) g

y a g. p, ] i t,9 .A% ei 8 '.f.ph j (, -f f t t y l i

j

( A j 0m,'h,.h}YA,,.' _ :. ,3 \\ . h' 4 wl 4 - f i. r ql1, f...,,{ s d l r o e q g i,I \\ No MgC12 Crack ing i c- }.s hA - j' i l - [' . &l ',' V 'T;j' - I<. i T' 'CEickiId# ',,'s ; i M Cl2 ,g w f, b 5 8 r 4.~.,l ) .L l y e }s. 1 l-I I e A a) Natch Tested Before Weld Overlay b) Notch Tested After Weld Overlay Figure 3-13. Meta 11ographic Sections (100X) of Moderate Depth, Circumferential Notch Tips from GPC/SI/WSI 28-Inch Notched Pipe Test [22] e .n n n

1-i I WELD OVERLAY TESTS OUTSIDE DIAMETER .i O .l g,e.. e ee e g3 O e e 0.7 - - e e e* e ge e s e e *8 e e 1 g -ee w g ( 1 O FA/M DtpprExcur~ .o e t At/AL /193/THAlf > g,3 g

• ese as..

e 4e 0.1 - - a 6 - i e i i i a i e i m .ao 70 e 40 m 30 30 10 0 to ao 30 40 50 so 70 80 STRESS (KSil itsSIDE DIAMETER t l ...} l .20" OVERLAY HATCH-1 lLONG) WELD PREP T CALCULATED AXIAL RESIDUAL STRESS THROUGH-WALL OF 12" SCH.100 PIPE Figure 3-14. Through-Wall Residual Stresses (23] 3-33

.I* il WELD OVERLAY TESTS (i OUTSIDE DIAMETER .i u. e e SJ- - o es e e S ee e

1. 1. G e

G ee %#-e y g

• f*-

(MEASLttcp Ad72 \\ t ^ " " " " " ' ' * " ~ e (AxteLfostr>aM:r} u-y{~ e l e . u. _ eee* 0.1 - - ?"". I e 70 e 40 .a0 -30 20 10 ~ o 10 30 30 40 so ao 70 so STRESS (KSI) MSIDE DIAMETER iV 'iL. l l ~ .23" OVERLAY HATCH-1 (LONG) WELD PREP l CALCULATED AXIAL RESIDUAL STRESS l THROUGH-WALL OF 12" KH.100 PIPE l Figure 3-15. Through-Wall Residual Stresses [23] 3-34 1

'i s-AXIAL STESS

g o.

~- .o. E m -so - ~ t .ss - -a- .m. ELD WEALAY -s.o 4.0 4.0 -Lo -i.e i.e i.e i.e i.e i.o Atlk 905. IN. L s-N00P STESS o. ~ .g. b m -so-I- E.sg. m l .m-a s-3 W ERLAy -N.o 4.0 4.o -4.0 4.0 0.0 2.0 e.o s.o a.o r Arl E PDS. l#. Figure 3-16. ID Stress for 24 Inch Overlay (23] 3-35

~. - STRESS TtROUGH THICKPESS AFTER FIRST LAYER STRESS THROUGH THICKNESS AFTER SECOND LAYER ANI M SI8tESS -* *** M00P SfftESS ANIE $101E55

---

HOOP STESS g 3y 2

12 37 c At ANig t*ISTAsCE { Al AN84 OISTAseCE l MR 911tFA M 9 WIFA M 0300 tNDES e y pept OSM Iwots y , p,,( l n-7 7 Y W i 1153-l MS3-l g 2 2 i w b I ',I z =o 5' f z = o.s " tirs-l n is - i iwa sw.a su,,,,u 8 Of PIPC ) c a. I or PIPE = -en -so - o -so -ro -n o e to ao .o so e _.o _ io _ ;, _; _ }, _'. ; la j, jo g, g,., = STRESS (MSI) STRESS (KSI) Figure 3-17. Calculated Through Wall Stresses after the First and Second Weld Overlay Layers for a 24 Inch Pipe with 1.48 Inch Wall. (Overlay contains five weld layers for a total thickness of 0.35 inch) [23]

I i t ,i

I-1 STRESS THROUGH THICKNESS AFTER FIFTH LAYER AulAL SIRESS

---

Hoop STRESS a 37

-= OUTER Af AFI AL OISTANCE I- ,f' osco ino<s. u-r I M O t me3-l ~2 = ....s" m l i m is - ' R . man Suars a e se e or PIPE -so -so -so -so -ro -m o e no ao eo so so STRESS (KSI) l Figure 3-18. Calculated Through Wall Stresses after the Fifth and f Final Weld Overlay Layer for a 24 Inch Pipe with 1.48 Inch Wall. (Overlay contains five weld layers for a l. total thickness of 0.35 inch.) (23] e 3-37 l

f 4.0 WELD METAL FRACTURE TOUGHNESS The flaw evaluations and weld overlay repairs at FitzPatrick were conducted in accordance with ASME, Section XI rules for evaluation of flaws in austenitic piping, IWB-3640 (Winter,1983 Addendum). These rules provide I - - allowable flaw depths for axial and circumferential flaws based on the net section collapse criterion (NSCC), which assumes that the material has [ sufficient toughness that the only effect of the cracking is to reduce the load carrying cross-sectional area of the pipe. This method is well supported by test data and analysis for materials exhibiting toughness properties typical of the wrought stainless steels used in nuclear reactor piping systems. Recent fracture toughness data for stainless steel weld metal, however, have indicated some flux-type weldments (SAW/SMAW) to have significantly lower toughness than wrought stainless steel. This has led to revision of Section XI, IWB-3640 (Winter, 1985 Addendum), to provide more restrictive allowable flaw size limits for flux weldments, based on elastic-i d plastic tearing instability analysis of the lower toughness materials. This issue has relatively little impact on the current consideration of extended service of the FitzPatrick weld overlay repairs, since they were applied using a gas-shielded (GTAW) welding process, which has demonstrated sufficient toughness in all tests to justify the use of net section collapse ] methodology. The 1985 Code revisions discussed above explicitly state that the earlier net section collapse based criteria are still applicable to GTAW weldments. As detailed in Section 2.0, the weld overlay repairs at FitzPatrick were designed with a thickness which requires no credit for the .2 original pipe wall in maintaining design basis safety margins. Thus, the design basis of the overlays is maintained regardless of potential low toughness of any original, flux-type weld joints. Recent proof testing of Ib weld overlay repaired pipe, described below, further confirms this point. 4-1 i

.+ 5 3" 'i> =4'.1-' Battel'le/NRC Degraded Pipe Tests 4.1.1. Test Objecti/es An' experimental program to confirm the effectiveness of the weld overlay.

u f

, repair method for repairing cracked pipe is currently being conducted by-Battelle Columbus Laboratories on behalf of the U. S. Nuclear Regulatory Consnission [27, 28]. The purpose of the program is to evaluate' the accuracy

rt of the assumed safety margins of the Section XI net section collapse

] methodology which has been used as the basis for weld overlay design-in the } U.S. - and elsewhere. In particular, weld overlays were applied to pipes ji' containing deep flaws, the pipes were loaded to failure, and the actual failure stresses were compared to the Section XI predicted values. An h assessment of the actual margins of safety of the Code approach compared to -.the predicted margins can thus be obtained from the test results. 14.1.1.2 Experimental Approach Three experiments were performed. The test pipe specimens were 6-inch . Schedule 120 Type 304 stainless steel pipe. Each pipe had a flaw introduced l? which was through-wall and which extended circumferential1y approximately l 50% (Figure 4-1). Flaws 50% of wall depth, extending approximately 17% of v" circumference were introduced by electric discharge machining. - These flawed specimens were cycled in three-point bending to grow fatigue cracks through- [ wall, extending 50% of circumference. The flawed pipes were then weld overlay-repaired (weld overlays were 0.31 inches thick on the average) using l, techniques typical of field practice. Each pipe was then pressurized at a temperature of 5500F to different levels of internal pressure. The internal pressure was kept constant, and each sample was loaded in bending under displacement control to failure (Figure 4-2). ~ 4.1.1.3 Test Results l' Although final test data were not available at the time of this report, preliminary results presented in [27] and [28] suggest that the experimental 4-2 l . ~..

1 I failure data are in good agreement with the theoretical NSCC failure predictions which serve as the basis for weld overlay design (see Figure 4- .i. 3). Differences between the Battelle test data and the FitzPatrick weld overlays are summarized in Table 4-1. I 4.2 Application of Test Data to FitzPatrick Weld Overlays ~[ The FitzPatrick weld overlays have been plotted directly on a load margin diagram (Figure 4-4) in which equivalent theoretical failure predictions and design curves have been developed for the 12 inch riser overlays, using both the source equations and IWB-3641 tables, with safety factors of 3 and 2.773, respectively, as well as the no safety factor cases. The points shown on Figure 4-4 represent the as-built weld overlays for each 1 of the 12 inch risers, neglecting the thickness of the first weld layer. The 'I data demonstrate a factor of safety to failure in excess of 2.773 by the IWB- 'bl 3641 tables and 3.0 by use of the source equations. 4.3 Sumary m dj The weld overlays applied to six recirculation system welds at FitzPatrick were re-evaluated, considering the potential impact of recent concerns regarding weld metal toughness on ASME Section XI, IWB-3641 flaw evalu-ations. The factors of safety to f ailure for the FitzPatrick weld overlays l l were evaluated by comparison with Code Table and source equation values. i3 I" Since the weld overlay metal at FitzPatrick was applied using the gas ] tungsten arc welding process (GTAW), the low weld metal toughness data (- available for flux shielded processes (SAW, SMAW) does not apply to the FitzPatrick situation. The net section collapse approach applied in weld y overlay design for FitzPatrick is sufficient to guarantee adequate margin to failure. The above conclusion is supported by recent degraded pipe failure data developed by Battelle on behalf of the Nuclear Regulatory Comission. l 4-3 1

m; 6: I' TABLE 4-1 )

SUMMARY

OF DIFFERENCES BETWEEN BATTELLE/NRC WELD OVERLAY TEST PIPES AND FITZPATRICK WELD OVERLAYS ig; i L Battelle/NRC Fitzpatrick Tests Risers (As-Built) Pipe Diameter, in. 6.625 12.75 Pipe. Thickness,. in.(avg.-) - 0.56 0.717 0verlay Thickness, in. 0.314 .385 .672(1,2) Crack Length / Circumference 0.5 .126 - 1.0(3) Crack Depth / Thickness 0.65 53 - 66% Design Allowable Load .239 0.120 - 0.233(6) Ratio (Pm + P )/Sf .193 .b 1. Average As-Built Thickness at Crack location ] 2. Beyond 1st Weld Overlay Layer 3. Assumed for Design. Actual l/ Circumference Ranges. 126 to 1.0. -4. Average Design Load for 4 Tests by Source Equations 5. Average Design Load for 4 Tests by IWB-3641 Tables ip-6. Does Not' Include Thermal Expansion Stress l[ i l l.a l r i 4-4

~fac%~ m 6{ inch O.D

• A Internal 1

(168 mm) Surface Crock f----- l,/ 3 1 'f I ~---_ n r ts Weld Overlay Pipe Thickness Repair

• A

~O.562 inch (14.3 mm) Thickness of Weld j.. Overlay ~0.314 inch (8.0 mm) o O C [' Internal Surfoce l Crock r' = = Clip goge locotion i

• = d-c EP locations g

\\ {.; \\ / \\ j ,/ '~_ r SECTION A-A t l Figure 4-1. Illustration of Cracked Pipe and Weld Overlay Configuration Used in Battelle/USNRC Experiments l [28] 4-5 l i

l l IIli 48 inches 37 inches-e 37 inches = = = = (939 mm) (1,220 mm) (939 mm) l Weld Overley, Instrumentation l Line Return Internal (Pressure Pipe for To e-Pressure Line Surfoce Crock Line Moment Arms \\ \\ Accumulotor pd lt N W /End

1. Support R

A -Displacement [ Controlled Actuators h s 1l l Strongback I 1 i ,.I Q P l Figure 4-2. Schematic Illustration of Test Setup Used in d Battelle/USNRC Weld Overlay Experiments [28] l r 4-6 t

4 -l \\ + 'i 0.7 Failure Predictions (S.F.=1) 0.6-I NSC Source Equations g a IWB-3641 Tables 0.5-t S 0.4 - x2 X3b u g 0.3-a j Design Limits 3 0.2-(5.F.=2.773) ip; z NSCC Source Equations IWB-3641 Tables 't i \\ O s s e e e 0 u 0.4 0.6 L WEW8RANE STESS/FLDW STESS ii THEORY x BATTULE DATA l1 l'F Figure 4-3. Comparison of Recent Battelle/USNRC Degraded Piping Program Weld Overlay 1 Tests With Overlay Design Basis Calculations 4-7

f-- o. Based upon 12" weld overlay repaired pipe 3 with a/t = 58% j j 0.8 NSCC (Safety Factor = 1.0) 0.7 - l-E IWB-3641 (Safety Factor 0.6 - y = 1,0) E 3 05-S t g g 0.4 - !k NSC (Safety Factor = 3.0) 0.3-z Design Basis IWB-3641 (Safety Factor = 2.773) y z 0.2-12-12 a f j. 12-69 5 0.1-12 -' '/12-70 l o 23 0 L 0 0.02 0.04 0.06 0.08 0.1 0.12 0.14 0.16 0.18 l l WCW8RANE STRESS / FLOW STRESS Figure 4-4. Comparison of Fitzpatrick Design and As-Built T Weld Overlay Data to ASME Design Limits and l Recent Pipe Failure Data 4-8 i L

1

• o 5.0 SUMARY AND CONCLUSIONS 5.1 Summary During the Fall 1984 outage at the James A. FitzPatrick Nuclear Power Plant,

.l ultrasonic (UT) examination of the recirculation system piping revealed indications at eleven weld joints which were believed to result from l intergranular stress corrosion cracking (IGSCC). Subsequent examination in 1985 gave improved flaw sizing information on these defects (three were no - longer considered significant following this reinspection) and also revealed the existence of indications in another joint. Six of the welds were repaired by weld overlays. The weld overlays were I

l designed in accordance with IWB-3640 using conservative methods.

For ~ example, thermal stress was included in the original overlay design sizing calculations. The thickness of the first overlay layer was not included in sizing calculations. All Type 308L weld metal was deposited by the GTAW process with a heat input of 40 kJ/in., specified ferrite number of 10, and >c maximum carbon content of 0.02%. All recirculation system welds were IHSI treated (and inspected) in 1984, except for two inaccessible 28" welds (for which mechanical stress improve-ment or resistance heating stress improvement is presently planned), and the welds in the 4" bypass lines (which are expected to be removed). Crack growth analyses were performed for the three welds with small IGSCC indications which were not overlay repaired. With the residual stress pattern imparted by IHSI, no crack growth was predicted for any of these welds. Because of the small size of identified flaws, IHSI is considered to be a permanent mitigation of IGSCC initiation and propagation at these locations. .F The resistance of weld metal to IGSCC crack propagation, residual stress j benefits of IHSI and weld overlays and weld toughness as demonstrated by I experiment were reviewed. This extensive data base provides pertinent background regarding the effectiveness of weld overlays and IHSI at FitzPatrick against any further IGSCC crack propagation. 1 1 5-1 i

i O }{ 5.2 Conclusions

j

. The six weld overlays on the FitzPatrick recirculation system were a designed and constructed in a conservative manner. The overlays meet the requirements of NUREG-0313, Revision 2 (Draft) for long term operation. I( . Precedents exist for extended operation (i.e. beyond two fuel cycles) with weld overlays. . Weld overlays at FitzPatrick are adequately sized for all operating conditions.

l,

. No growth of existing flaws is predicted for any of the four flawed ~[ welds which were not overlay repaired. The three welds were inspected in October, 1984, and again in April, 1985. The inspection results ih for these welds were not substantially different in these inspec-tions. Tnis supports the analytical prediction of no crack growth. . The favorable residual stress pattern imparted by IHSI prevents any ] further crack growth. . Recent developments in UT inspection through weld overlays support p i[ qualification of weld overlays as long term repairs in accordance i with Reference 3 requirements. l . For the unrepaired welds, the possibility of low toughness weld metal { does not present a concern since the flaws are so small (i.e. well below even 2/3 of the IWB-3640 flaw size limits). l !r l l . All weld overlays at FitzPatrick were applied by the GTAW process; low j toughness of flux welds is, therefore, not an issue for the overlay l repaired welds. Local repairs to leaks using SMAW do not invalidate this conclusion. 5-2

'l W I V Continued operation with the six weld overlay repaired welds' does not q'. compromise safety margins or present operational concerns. 1'l ~ Continued operation of FitzPatrick with the recirculation system in the present as-repaired' condition is justified. IL, m ,! j! Ti !?l i l' ,N i L l' 5-3

[L,' w

6.0 REFERENCES

l 1. U.S. Nuclear Regulatory Commission Generic Letter 84-11, " Inspection of BWR Stainless Steel Piping," April 19, 1984. ( 2. U.S. Nuclear Regulatory Commission, NUREG-1061, " Report of the U.S. l Nuclear Regulatory Consnission Piping Review Committee," a. Vol.1, " Investigation and Evaluation of Stress Corrosion Cracking in Piping of Boiler Water Reactor Plants," Second Draft, April 1984. b. Vol. 3, " Evaluation of Potential for Pipe Breaks," November 1984. F 3. NUREG-0313, Revision 2 (Draft), " Technical Report on Material Selection and Processing Guidelines for BWR Coolant Pressure Boundary Piping," issued for Public comment July 11, 1986. ~ 4. " Fracture Mechanics Evaluation and Weld Overlay Design for Recircu1ation System Piping in the James A. FitzPatrick Nuclear Power Plant," Structural Integrity Associates, SIR-85-015, May 7, 1985. 5. GE Report 22A2622, Rev.1. " Design Report, Recirculation System for James u. A. FitzPatrick Nuclear Power Station, ANSI B31.1 Calculations," December 6, 1976. 6. ASME Section XI 1983 Edition, Winter 1985 Addenda. 7. EPRI Report NP-2622-LD, " Computational Residual Stress Analysis for Induction Heating of Welded BWR Pipes," EPRI Project T113-8, Final Report, December 1982.

8. Buchalet, C.B.

and Bamford, W.H., " ASTM 8th National Symposium on Fracture Mechanics, 1974",' ASTM-STP-590, pp. 385-402, 1975. ] 9. " Guidelines for Flaw Evaluation and Remedial Actions for Stainless Steel 3 Piping Susceptible to IGSCC," Final Report for EPRI Project T303-1, Report No. SIR-84-005, April 13, 1984.

10. EPRI Report NP-81-4-LD, " Residual Stress Improvement by Means of Induction Heating," March 1981.

'] 11. Ranganath, S., Mehta, H.S., and Norris, D.M., " Structural Evaluation of Flaws in Power Plant Piping," ASME PVP-Vol. 94, Circumferential Cracks

p-in Pressure Vessels and Piping - Vcl. I, pp. 91-116, 1984.

12. " Weld Metal Cracking in Nine Mile Point Unit 1 Recirculation Pipe Joints," Letter, R.E. Smith to D. Norris (EPRI), February 23, 1984. '. 3. Diercks, D.R. and Gaitonde, S.M., " Analysis of Cracked Core Spray from Quad Cities Unit 2 Boiling Water Reactor," Materials in Nuclear Energy, 1983. i

14. Horn, R.M., et al., "The Growth and Stability of Stress Corrosion Cracks in Large Diameter BWR Piping," Final Report, EPRI NP-2472, July 1982.

6-1

w* s 1 o 15. " Alternative Alloys for BWR Pipe Applications", EPRI NP-2671-LD, October 1982 1

16. General Electric Company, " Third Party Review of the Technical Justifi-cation for Continued Operation of James A. FitzPatrick, Nuclear Power Plant with Existing Recirculation System Piping", Transmitted by letter j

from J. Silva (GE) to T. Dougherty (NYPA) dated April 21, 1986: JS 0421-1. 17. " Assessment of the Feasibility of Producing Pipe Samples with Tight Through-Wall IGSCC, EPRI NP-2241-LD, February 1982. 18. " Verification of Intergranular Stress Corrosion Crack Resistance in Boiling Water Reactor Large-Diameter Pipe," Final Report, EPRI NP-3650-LD, July 1984.

19. Pickett, A.E., "AsF ssment of Remedies for Degraded Piping - First Semi-l Annual Progress Rt. port," NEDC-30712-1, September 1984.

20. Pickett, A.E.,

" Assessment of Remedies for Degraded Piping - Second Semi-Annual Progress Report," NEDC-30712-2, August 1984 - August 1985. 21. N.R. Hughes and A.J. Giannuzzi, " Evaluation of Near-Term BWR Piping Remedies, Vol. 1 & 2", EPRI NP-1222, Nov. 1979. 22. " Extended Lifetime Test Program for Weld Overlays at Hatch, Unit 1", Structural Integrity Associates, SIR-84-030, September 1984. 23. " Continued Service Justification for Weld Overlay Pipe Repairs," EPRI, BWROG Ad Hoc Committee, May 25, 1984. 5 24. J. Park, D. Kupperman, W. Shack, " Examination of Overlay Pipe Weldments Removed from Hatch-2 Reactor," Argonne National Laboratory, September 1984. r;L 25. L. Becker, et al., " Examination of Weld Overlayed Pipe Joints," EPRI NDE l Center Report RP-1570-2, April 1985. 26. J. Park and D. Kupperman, " Ultrasonic and Metallurgical Examination of a g Cracked Type 304 Stainless Steel BWR Pipe Weldment," ANL-84-1, January

1984, i.

27. G.M. Wilkowski, et al., " Degraded Piping Program - Phase II," NUREG/CR-4082, BMI-2120, Semi-Annual Report, 10-84 to 3-85. !i 28. G.M. Wilkowski, et al., " Degraded Piping Program - Phase II," NUREG/CR-4082, BMI-2120, Semi-Annual Report, 3-85 to 10-85 (Draft).

29. USAS B31.1

" Power Piping" 1957 Edition With Addenda Through 1969. I}}