ML20198G449
| ML20198G449 | |
| Person / Time | |
|---|---|
| Site: | San Onofre  |
| Issue date: | 07/31/1998 |
| From: | WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML20138M341 | List: |
| References | |
| SG-98-07-005, SG-98-7-5, WCAP-15015, WCAP-15015-R01, WCAP-15015-R1, NUDOCS 9812290081 | |
| Download: ML20198G449 (95) | |
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- Westinghouse Non-Proprietary Class 3 WCAP-15015 Revision 1
++++++++
Specific Application of
. Laser Welded Sleeves for l Southern California Edison.
San Onofre Units 2 and 3 Steam Generators t~. l l' y. e l s. W e.s t i n g h o u s e E n e r g y S y s t e m s W
=
D2 ) )( I d b1 p PDH !. -t ,
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WESTINGHOUSE CLASS 3 WCAP-15015 SG-98-07-005 Revision 1 SPECIFIC APPLICATION OF LASER WELDED SLEEVES FOR SOUTHERN CALIFORNIA EDISON SAN ONOFRE UNITS 2 AND 3 STEAM GENERATORS July 1998 i i e i J i WESTINGHOUSE ELECTRIC COMPANY NUCLEAR SERVICES DIVISION P.O. BOX 158 MADISON, PENNSYLVANIA 15663 C1998 WESTINGHOUSE ELXCTRIC COMPANY ALL RIGHTS RESEttVED 4 4
l l ABSTRACT Under Plant Technical Specification requirements, steam generator tubes are periodically inspected for degradation using non-destructive examination techniques. If established inspection criteria are exceeded, the tube must be removed from service by plugging, or the I tube must be brought back into compliance with the Technical Specification Criteria. Tube sleeving is one technique used to return the tube to an operable condition. The purpose of this evaluation is to establish the applicability of a generic laser welding sleeving analysis for 3/4 inch diameter tube, Asea Brown Boveri-Combustion Engineering feedring-type and Westinghouse preheater steam generators (WCAP-13698, Rev. 2) to the San Onofre Units 2 and 3 steam generators. All of the sleeve design, structural evaluation, mechanical testing, stress corrosion resistance testing, installation processes and nondestructive examination discussed in the generic report, as modified or superserled by this document, apply to San Onofre Units 2 and 3. This document has been revised in four areas, weld width, removal of conservatism in hydraulic equivalency of sleeves to plugs, removal of restriction on sleeve installation elevation due to seismic loads and evaluation of steam generator conditions for the combination of 1,000 tubes plugged per steam generator and a Tcora reduction to 540 'F. Based on the combined results of this evaluation, the generic evaluation (WCAP-13698, Rev. 2) and Addendum 1 to WCAP-13698, Rev. 2, the laser welded sleeves are concluded to meet applicable ASME Boiler and Pressure Vessel Code and regulatory requirements for San Onofre Units 2 and 3. The allowable plugging margin for sleeve degradation is 39%. Changes from the initial issue denoted by line in margin or, for more numerous changes in a section, a footnote on the first page of the section. ii WCP15015-cO951an712. doc
l TABLE OF CONTENTS i l l I Section Title Page l 1 l
- 1. 0 I NT R O D U CTI O N . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1 - 1 i'
2.0 SLEEVE DESIGN AND DESCRIPTION . .... ....... .. .. ..... .. ...... ...... . ....... . .. ....... 2 1 3.0 ANALYTICAL VERIFICATION . ... ...... . . . . .. . . . .. . . . . . .... .... .. . . .. . .. . . . . . .. . . . . . . .. .. . .. .. . . . . . . . 3 1 1-3.1 Applicable Loading Conditions ..... . ...... .. . .. ........ ... .. ... ........ ...... .... ... ...... ... 3- 1 3.2 Steam Generator Geometry Effects . ........ ... .. . ...............................3-1 i 3.3 Sleeverfube Contact Pressures Elevated Tubesheet Sleeves..... ....... .. . .. .. .. 3 2 3.3.1 Normal Operation. . ..........................................................3-3 j 3.3.2 Faulted Condition . .... .... . . .. . .... .. ... . . . . . . . . . . . . . . . . .. ........3-4 , 3.3.3 Summary of Results.. .... ....................................................3-5 3.4 Sleeve and Tube Primary Stress Limits ...... . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. .. . 3-5 i 3.5 Sleeve and Tube Maximum Range of Stress Intensity and Fatigue . ... . . . .. . 3-5 3.6 Weld Sizing for Pressure Loads.. . .. .. .... ...... . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. . . 3 6 ! 3.7 Burst Strength of Laser Weld.. ......... .......... .. . .............................37 i 3.8 Fatigue Evaluation of Laser Weld., ...... ........ .... .. .. .. .. ... ... . ... . . .. ...... 3 7 3.9 Minimum Wall Thickness Requirements... .. . . . . . . . ... ..... .... .. ... .... 3-9 i 3.10 Use of Preheat SG Sleeve in Feedring SG ......... . . . . . . . . . . . . . . . . . . . . . . ...... . ... 3-9 3.11 Applicability of Thermal Hydraulic Analysis, Section 3.2 of Ref. I to S O N G S 2 & 3 : . . .. . . . . . . .. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. ...3-10 3.12 Applicability of Flow Induced Vibration Analysis, Section 3.3 of Ref.13-i to SONGS 2 & 3:........ . ... .... . . . . . ..... ..... . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3 13 4 3.13 Conclusions ... ........ ... ..... . .. . ........................................... .. 3 13 4.0 MECHANICAL TESTS . ........ . ..... .. . ...................................................4-1 4.1 Applicability of FSG Processes for FLTS Lower Joint to SONGS 2 & 3.... ...... . 41 4.2 Applicability of FSG Processes for ETS Lower Joint to SONGS 2 & 3. ... . . .... 4-2 4 4.2.1 SONGS 2 & 3 ETS MIF Joint Leak Test Equivalence ... .... . ....... ..... . . 4-3 4.2.2 Justi5 cation for Using the Previous FSG MIF Lower Joint Pullout Resistance for SONGS 2 & 3 Based on CP Evaluation ..... .... . .. . ... . 4-5 : 4.2.3 Justification for Using Previous FSG MIF Joint Leakage Resistance Results as Bounding Results for SONGS 2 & 3... ........... ...........4-6 4.2.4 MIF Joint Pullout and Leakage Resistance Tests......... .. ... . ..... ... . .. . 4-7 4.2.5 MIF Joint Acceptance Criteria ....... ............ .. .................. ... .. ... ... . ... 4-7 l iii WCP15015-cO95lan712. doc
i l 4.2.6 Results and Conclusions of MIF Joint Verification Tests-Previous FSGs 4-8
- 4.2.7 MIF Joint Conclusions. ...... . ... .. ... ..... .. .. ............. . . . . . . . . . . . . . .4-9 1 4.3 Applicability of 7/8 inch Tube LWS Test Results to SONGS 2 & 3. ..... .............. 4-9 i 5.0 STRESS CORROSION TESTING OF LASER WELDED SLEEVE JOINTS.... .. .. 5-1 4
5.1 LWS Process and SG Design Variables... .... . . . . . . . . . . . . . . . .................52 ! 5.2 Residual Stresses vs. Stress Relief Temperature in LWS Sleeve Repairs..... . . 5-3 5.3 Corrosion Test Description. ... . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .................5-3 5.4 Corrosion Resistance of Free-Span Laser Weld-Repaired Tubes As-Welded Condition . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .....................5-4 5.5 Corrosion Resistance of Free-Span Laser Weld-Repaired Tubes - with Post Weld Stress Relief... . .... . .... ... . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. ... 5-5 5.6 Corrosion Resistance of Free-Span Laser Weld-Repaired Tubes with Post Weld Stress Relief and Conditions of Axial Load During Test.... . . . . . . . . . . .... ... 5-5 5.7 Estimated Sleeve Performance.. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 5-8 5.8 0 uter Diameter Surface Condition . . . ... .. ... . ..... . . . . . . . . . . . 5-10 6.0 INSTALLATION PROCESS DESCRIPTION. . ..... ... .. . . . . . ...........6-1 7.0 NONDESTRUCTIVE EXAMINATION (NDE) INSPECTABILITY... . . .. . .7-1 7.1 Inspection Plan Logic for Installation.. . . . . . . . . . . . . . . . . . . . . . . . .. . 7- 1 , 7.1.1 Tubesheet Sleeve Basic Inspection Plan. . ..... .. ... . . .. . .. 7-1 7.1.2 Tube Support Sleeve Basic Inspection Plan ... ... . . ..... 7-2 i 7.2 General Process Overview of Ultrasonic Examination...... ... .... ... ...... ... ...... 7-2 7.2.1 Principle of Operation and Data Processing of Ultrasonic Examination. 7-3 7.2.2 Laser Weld Test Sample Results .... .. ... .. . . . . . . . . . . . . . . . . . . .. ... 7-4 7.2.3 Ultrasonic Inspection Equipment and Tooling. . . .. . . . . . .. .7-4 7.3 Eddy Current Inspection. . . . ... . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7- 5 7.4 Cecco-5/ Bobbin Principles of Operation. . .. ... . ... . ..... ..... . . . . . . . . . . . . . . 7- 5 7.5 Alternate Post Installation Acceptance Methods... ........ .. . . . . . . . . . . . . . 7-6 l 7.5.1 Bounding Inspections .. ... ... . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .7-6 7.5.2 Workmanship Samples. . . .. .. .... . . . . . . . . . . . . . 7-7 7.5.3 Other Advanced Examination Techniques ....... .. .. . . . . . .7-7 7.6 Inservice Inspection Plan for Sleeved Tubes.... . . . . . . . . . . . . . .... 7-7
8.0 REFERENCES
. . ......... .......... . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. . .. 8- 1 I'v WCP15015-c!395lan712. doc
..-. - . _ - _-.- ~.-.- - - - _ - - - . - - - . . . - . - _ . - - - --
C l' LIST OF TABLES j Table Title Page , i Table' 31 Tubesheet Comparisons for CE Steam Generators.... . .............. .................. 3 14 ! Table 3 2 . Minimum Sleeve / Tube Contact Pressures for Elevated Tubesheet Sleeve ... 3-15 1 j Table 3 3 Enveloping Pressure Loads for Design, Faulted, and Test Conditions for i Generic and SCE Unit 2 & 3 . ..... . . . . ..... . .. . .. . . . ... .. .. ...... ... ....... ... . . . . . . . . .. . . .. ... . .. 3 16 i i . j Table 3 4 Summary of Maximum Primary Stress Intensity for Sleeve and Tube for Full ~ . Length Tubesheet Laser Welded Sleeve [ l a ,c . . . . . . . . . . . . . . . . . . . . . . .. . . . . 3 - 17 i ! Table 3 5 Comparison of Transient Cycles for 3/4 Inch OD Tube Sleeving Generic
- vs. SCE San Onofre Units 2 and 3......... . ... ... . ......................................3-18
! . Table 3-6 Maximum Range of Stress Intensity and Fatigue Usage in Sleeve l andTube.......................................................................................3-19 I
- Table 3-7 Maximum AP Loads Used to Size Laser Welds i AP = primary to secondary pressure differential (psi) .... .......................3-20
- Table 3-8 Comparison of Minimum Wall Thickness Calculations for Sleeves 3/4 inch l OD Tube Generic and SCE Unit 2 & 3 Steam Generators ........ . ........... .. . 3 21 I
i-Table 3 9 Sleeve Total Equivalency per Steam Generator. . ..............................322
- Table 4-1 Applicability of FSG LW* Elevated Tubesheet Sleeve Lower Joint Process l to SONGS 2 & 3 Normal Operation.. ..... .. ...... . .. ... .. ... .... ............... ... 4-11 i Table 4 2 Applicability of FSG LW Elevated Tubesheet Sleeve
- Lower Joint Process i to SONGS 2 & 3 Faulted Condition. ... . ............. . ..... .. . . . . . . . . . . . . . . . . 4 - 12 1
j Table 4-3 Application of Previous FSG . Roll Last ETS Lower Joint for SONGS 2 & 3
- Roll Last ETS Lower Joint.. . . . . ..... ....... . . . . . . . . . . . . . . . . . . . . . . . . ..... 4-13 Table 4 4 Verification Test Results Mechanical Interference Fit Lower Joint Elevated Tubesheet Sleeve Alloy 690 Sleeve for 3/4 Inch Tube (Feedring SG ) ... ..... 414 j Table 5-1 Summary ofImpact of Laser Welded Sleeve Operations on Stresses .... . .. . 5-13 i Table 5 2 Far-Field Stress as a Function of Stress Relief Temperature ........... ...... .... 5-14 Table 5-3 Results of 750 F Doped Steam Tests for Nd
- YAG Laser Weld-Repaired
; Mockups...........................................................................................5-15 l Table 5 4 Doped Steam Corrosion Test Results for 3/4 inch Tube-Sleeve Mockups -
Tested Without Axial Load .. . . . . . . . .. .... . . . .. . . . . . . ... . . ..... .. .. .. .. . .. . . . . . . .. . .. . .... .. .. . 5 16 Table 5-5 Summary of Fabrication Parameters, Temperatures, Stresses, and Corrosion
- Test Results for % Inch Sleeve Mockups Tested With Applied Axial Load... 517 Table 5 6 Summary of Temperatures, Stresses and Corrosion Test Results for 7/8 inch Sleeve Mockups Tested with Applied Axial Load .. . .. .......................... ... . 5-18 i
v WCP150154:1395 tan 712. doc
- = _ - . . -- . . - _ . .- .- - . - . - - -
t e LIST OF FIGURES 4 1 Fisture TiLi.g Eage Figure 3-1 Finite Element Model of CE Channel Head /Tubesheet/Shell...... ...... ........ 3-23
- Figure 3-2 Contact Pressures for Normal Conditions with an Intact Tube ............ . . 3-24 Figure 3 3 Contact Pressuree for Normal Conditions with a Separated Tube.... . ......... 3-25 l Figure 3 4 Contact Pressurea for Faulted Conditions with Intact or Separated Tube... 3-26
. Figure 3 5 Hydraulic Equivalency of Alternate Sleeve Configurations.. .......... ....... .... 3-27 ! Figure 3 6 Thermal Equivalencly of Ahernate Sleeve Con 6gurations ............. ...... . . . 3 28 Figure 5-1 Corrosion Test Specimen for Doped Steam Testing of an LWS Joint without .
Axial Load .. . . .. .......... ... . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ........ .... 5-19 ! Figure 5-2 Test Stand used for Fabrication of LWS Mockups under Locked Tube i Conditions. .. .. .. . . .... .. . . . . . . . . . . . . . . . . . . .. . . . . . . . . . . . . . . . . . . . . . .. 5 20 Figure 5-3 Schematic of Test Assembly used for Doped Steam Testing of Tube /LWS Mockups under Conditions of Applied Axi.a Loading. .. . . . . . . . . . . . . .. . 5 21 Figure 5-4 IGSCC in Alloy 609 Tube of YAG Laser Welded Sleeve Joint after 109 Hours in 7500F Steam Accelerated Corrosion Test.. . . . . . . . . . . . . . . . . . . . . . . .5-22 Figure 71 Ultrasonic Inspection of Welded Sleeve Joint. .. . . . . . . . . . . . . . . . . . . . ..79 Figure 7-2 Typical Digitized UT Waveform.... .. . . . . .. .. . . . . . . . . . .7 10 Figure 7-3 A, B, C, and Combined C-Scan Display for Weld in UT Calibration Standard .. .. . . .... . . .. . . . . . . . . . . . . . . .. . . . . . . . .7-11 1. ! Figure 7 4 UT Calibration Standard. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7 12 Figure 7 5 Cecco-5 Sleeve Calibration Standard.. .. . . .... .. . .. . . . . . . . . . 7-13 I Figure 7 6 Strip Chart Display for Cecco/ Bobbin Data... . . .. . . . . . . . . . .. 7-14 4 Figure 7 7 Response of Cecco-5 Probe to 60% OD Axial Notch in Parent Tube Located
- at Expansion Transition. . . ....... ... . .. . . .. .. ......... . ... . . .. 7- 15 Figure 7-8 Response of Cecco 5 Probe to 60% OD Circumferential Notch in Parent
. Tube located at Expansion Transition.. ... . .. . . . . ... . . . . . . . . . . . . . . . . .... 7-16 i i l I
]
l i. vi WCP15015-cl395 tan 712. doc
i 1 e : i i l' l l l l r I t 9
i-
1.0 INTRODUCTION
This report documents the results of an analysis to evaluate the applicabiEty of the generic laser welded sleeving analysis for 3/4 inch diameter tube, Asea Brown Boveri Combustion Engineering (ABB-CE) feedring-type and Westinghouse preheater steam generators (WCAP-13698, Rev. 2) to the San Onofre Units 2 and 3 steam generators. For the generic analysis, transient loads are used that umbrella these steam generators (SGs). Included in the generic analysis are calculations to determine minimum wall thickness requirements for the sleeves. These calculations are a function of plant operating parameters which vary from plant to plant and which can change with the implementation of operating or system modifications. The purpose of this evaluation then is to compare the current set of transient and operating parameters for San Onofre Units 2 and 3 (SONGS 2 & 3) to those used in the generic analysis, with the intent of confirming '. hat the generic analysis provides a bounding analysis for San Onofre Units 2 and 3, and to also remove any conservatism in the generic analysis for sleeve minimum wall thickness, if possible. The results of this analysis are based on transient data supplied by Southern California Edison (SCE) in References 3 and 4. In establishing the structural adequacy of the laser welded sleeves in the generic analysis criteria were evaluated for primary stress limits, maximum range of stress intensity and fatigue, and minimum wall thickness requirements. The load conditions applicable to each of these areas are reviewed in this analysis to establish the applicability of the generic analysis. In general, the discussions to follow provide only a brief overview of each area. More in-depth discussions are contained in Reference 1 for the generic analysis. The statenients contained in this report are based on engineering analyses and tests and cannot be relied upon to state the contractual responsibilities of the parties. This section supersedes or clarifies Section 1 of Ref.1 in several areas. These areas are: Page No. of Section ChangelClarification Comment Ref, i No. 12 1.1 Tube support sleeve (TSS) for feednng steam Shorter TSS, for example, generators (FSGs), i.e., ABB-CE SGs, changed increases number of tubes which to 12 to 15 inches length, from 15 inches. can be sleeved in vicinity of periphery of bundle 1-2 1.1 Discontinuance of laser welding of sleeve-to- Weld is not required due to tube mechanical interference fit (MIF) joints redundancy in MIF joint (within tubesheet) as a redundancy 1-2 1.1 Discontinuance of option of "no post weld heat PWHT more generally accepted treatment (PWHT)" of laser welds for shorter as important in the industry than a term operation few years ago 11 WCP15015-c2395 tan. doc 07/14/98
2.0 SLEEVE DESIGN AND DESCRIPTION The sleeve design, as shown in Reference 1, and as modified and updated below, including the design documentation, is directly applicable to San Onofre Units 2 and 3. Based on the very conservative assumption that the bending stress range in the tube for the OBE is at the [
.]*.c Therefore, sleeves may be installed in the straight leg of the tube bundle anywhere below the top support. Additionally, application of the LWSs to tubes at degraded eggerate locations is beyond the scope of this documentation.
The following information supersedes the designated information in Ref.1: This Section replaces Section 2.1.1.2 Elevated Tubesheet Sleeve,in Ref.1: Elevated Tubesheet Sleeve The sleeve elevated in the tubesheet, or elevated tubesheet sleeve (ETS), is illustrated in Figure 2-2 of Ref.1. It is applicable to the steam generators in which the tubes were installed in the tubesheet by the [
]* c". The ETS upper joint is identical to other free span joints, i.e., the upper joint of the full length tubesheet sleeve (FLTS) and the tube support sleeve (TSS). The ETS I lower joint is fabricated by the same types of processes which are used to fabricate the l FLTS lower joints, i.e., [ .]*c4 The ETS is similar to I the FLTS in that it is designed to address tube degradation in the tube free span and in the vicinity of the tubesheet top. However, unlike the FLTS, it is limited to these applications !
and is not designed to address degradation in the remainder of the tube within the tubesheet. l [ l l l
.]*ce As in the case of the FLTS free span joint, the ETS inboard-most weld, i.e., either the weld at the initial elevation or the reweld at the specified lower elevation, must be separated from the extreme extent of the degradation by a [ l j a.c.e 2-1 i WCP15015-c!3951an. doc 07/13/98 l
i l i [
,]..e The separation distance for degradatir 1 in the vicinity of, and above, the ETS, MIF lower .
joint roll is determined the same way as it is determined for the FLTS, MIF lower joint. As discussed for the FLTS, the ECT process used to determine the extreme extent of l degradation will be appropriate for the type of degradation expected and it will meet.the required uncertainty in elevation of the degradstion. The extent of [ i i i i i 1
, j ..e..
J This section supersedes or clarifies Section 2 of Ref.1 in these areas: Page No. of Section Change / Clarification Comment
- Ref.1 No. I l
2-2 2.1.2 e Tube support sleeve length changed from . Shoiter TSS, for example, 2 (TSS) 15 inches to 12 to15 inches increases number of tubes I which can be sleeved in I vicinity of periphery of bundle ) e Tube support sleeve - discontinuance of "no e PvVHT more generally PWHT option" accepted as important in the industry than a few years ago ) l 2-4 2.2.1 . Ref.1 stated that the % inch tube weld Ter's of thousands of sivs in %,7/8 procedures for FSGs are being qualified; the and 11/16 inch tubes have been procedures were qualified and were used installed. over 30,000 times in the field. Laser weld within tubesheet discont.wei
- Discontinuance of "no PWHT option" 2-6 Fig. 2-1 Laser weld within tubesheet discontinued Weld is not required there
~
2-7 Fig. 2-1 Laser weld within tubesheet discontinued Weld is not required there a 4 22 WCP15015-cl395 tan. doc 07/13/98
3.0 ANALYTICAL VERIFICATION 3.1 Applicable Loading Conditions ! The umbrella loading conditions for the generic analysis are defined in Reference 2, and are based on a review of the applicable design specifications for ABB-CE feedring steam generators and for Westinghouse Model DS, D4, D5 and E1/E2 steam generators. The applicable loading conditions for San Onofre Units 2 and 3 steam generators are those . supplied by SCE for use in this evaluation (Reftrences 3 - 6). SCE had requested that j Westinghouse incorporate operating conditions for a planned reduction in Teola to 540'F (Reference 5). SCE provided conditions for the Teota reduction with 432/425 plugged tubes
- (Reference 6), but did not have the information for 1000 plugged tubes. The operating conditions for the Teoid reduction with 1000 plugged tubes have been estimated by Westinghouse. The most conservative set of conditions (highest AP across the tubesheet, lowest primary temperature) was selected for this evaluation.
Normal Operation Parameter Ranges 4 Parameters Operating Range Primary Side Conditions: Primary Side Pressure 2235 psig i Hot Leg Temperature Thot 611/594.4* aF ; { Cold Leg Temperature Teota 553/540* *F l Secondary Side Conditions: j Steam Pressure 885/761* psig j Steam Temperature 532/514.8* 'F Note:
- Indicates assumed conditions at 1000 tubes plugged with Teota Reduction.
I i 3.2 Steam Generator Geometry Effects The ABB-CE stean ,merators at San Onofre Units 2 and 3 are slightly different from the CE steam generator that was modeled in Reference 1. The only significant differences are that the tubesheet is slightly thicker and has a larger diameter in the San Onofre Units 2 and 3 steam generators. Table 3-1 provides a comparison of key geometric parameters for the tubesheets for the two designs. 1 This section has a general revision for weld width, removal of conservatism in hydraulic equivalency and SG condition changes due to combination of 1,000 tubes plugged and T oia reduction to 540 'F 31 WCP15015<!3951an712. doc
i , i l Plate bending equations from Reference 7 (Table 24, Case 2b) were used to compare the , stresses in the perfcrated part of the tubesheet for the two geometries. 4 As can be seen in Table 3-1, deflections and stresses calculated for the San Onofre Units 2 and 3 steam generators are less than those obtained in Reference 1 for the CE steam generators with thinner tubesheets. Thus, the results obtained for the CE steam generator ) in Reference 1 are conservative when applied to tLe CE steam generators at San Onofre i Units 2 and 3. l i 3.3 Sleeve / Tube Contact Pressures - Elevated Tubesheet Sleeves i
- Some of the sleeves, i.e., the elevated tubesheet sleeves (ETSs) for San Onofre Units 2 and 3 are to be installed with the sleeve lower joint in the upper half of the tubesheet, where
, tubesheet bow upward during operation tends to increase the diameter of the holes drilled ' in the tubesheet. This diameter increase will result in a decrease in the contact pressures between the sleeve / tube and tube /tubesheet produced by system pressures and differential thermal expansions among the sleeve, tuba, and tubesheet. This section determines the effect of tubesheet rotations on the sleeve / tube and tube /tubesheet contact pressures. 4 Loads are imposed on the sleeve as a result of tubesheet rotations under pressure and ! temperature conditions. A 2-D axisymmetric finite element analysis of an ABB-CE tubesheet, channel head, and lower shell was performed in Reference 1. The previous section established that the results from that analysis are conservative when applied to the . San Onofre Units 2 and 3 steam generators. The model is shown in Figure 3-1. This j yielded displacements throughout the tubesheet for two pressure and three thermal unit loads. The three temperature loadings consist of applying a uniform thermal expansion to i each of the three component members, one at a time, while the other two remain at j ambient conditions i
- The radial deflection at any point within the tubesheet is found by scaling and combining
- the unit load radial deflections at that location according to
= (UMPrim (Primary Pressure /1000)
Un
+ (Ugs,c(Secondary Pressure /1000)
, + (UMTubesheet[(Tubesheet Temperature 70)/500]
+ (Ug)Shell[(Shell Temperature - 70)/500]
l + (UdChannel Head [(Channel Head Tcnperature 70)/500] This expression is used to determine the radial deflections along a line of nodes at a constant axial elevation (e.g. neutral axis) within the perforated area of the tubesheet 3-2 WCP15015-ct3951an712. doc
The expansion of a hole of diamder D in the tubesheet at a radius R is given by: Radial: AD = D { dun (R)/dR} Circumferential: AD = D {Un(R)/R} Un is available directly from the finite element results, dug /dR may be obtained by numerical differentiation. The maximum expansion of a hole in the tubesheet is in either the radial or circumferential direction. Typically, these two values are within 5% of each other. Since the analysis for calculating contact pressures is based on the assumption of axisymmetric deformations with respect to the centerline of the hole, a representative value for the hole expansion must be used that is consistent with the assumption of axisymmetric behavior. A study was performed to determine the effect of hole out-of-roundness on the contact pressures between the sleeve and tube, and between the tube and tubesheet. The equation used for the hole AD is: AD = (SF)(ADm,x) + (1 - SF)(Admin) where SF is a scale factor between zero and one. For the eccentricities typically encountered during tubesheet rotations, SF is usually between [ ]a,c. I This hole expansion includes the effects of tubesheet rotations and deformations caused by 8 the system pressures and temperatures. It does not include local effects produced by interactions between the sleeve, tube, and tubesheet hole. Thick shell equations from Reference 8 in combination with the hole expansions calculated from the finite element model displacements are used to calculate the contact pressures between the sleeve and tube, and between the tube and tubesheet. For a given set of primary and secondary side pressures and temperatures, the contact pressure equations are solved for selected elevations in the tubesheet to obtain the contact pressures as a function of radius between the steeve and tube and the tube and tubesheet. The elevations selected were the neutral axis of the tubesheet and three elevations
, spanning the section from the bottom of the ETS to two inches from the top surface of the tubesheet.
3.3.1 Normal Operation Two sets of operating parameters for normal operation at San Onofre Units 2 and 3 were considered, both for the original operating conditions and for the changes due to licensed 4 plugging conditions for 1000 tubes (10% SGTP) with a Ta reduction to 540 F. The 3-3 WCP15o15-ct3951an712. doc
I l temperatures and pressures for normal operating conditions at San Onofre Units 2 and 3 are therefore: , i Normal Operating Parameter 0% SGTP 10% SGTP Primary Pressun 2235 psig 2235 psig Secondary Punure 885 psig 761 psig Primary Fluid Temperature (Thot) 611'F 594.4 'F Secondary Fluid Temperature 532 F 514.8 'F For this set of primary and secondary side pressures and temperatures, the contact pressures between the sleeve and tube and the tube and tubesheet are obtained as functions of radius for selected elevations in the tubesheet for both intact tubes and tubes j separated above the tubesheet. The primary pressure acts on the inside of the tube and, for conservatism, the secondary pressure is assumed to act on the outside of the tube and the inside of the tubesheet hole. With separated tubes, the secondary pressure is assumed to act on the outside of the sleeve and inside of the tube as well as between the tube and tubesheet. 3.3.2 Faulted C(>ndition Conditions at three times during the Steam Line Break transient and the time of maximum AP during the Feed System Pipe Break transient were considered. The limiting faulted condition for contact pressure occurred during the Steam Line Break transient at 200 seconds for 10% SGTP with a Tcoid reduction to 540 F. The temperatures and pressures for this condition are: 1 Primary Pressure = 790 prig Secondary Pressure = 0 psig Primary Fluid Temperature (Li) = 402.5 F Secondary Fluid Temperature = 514.8 F For this set of primary and secondary side pressures and temperatures, the contact pressures between the sleeve and tube and the tube and tubesheet are obtained as functions of radius for selected elevations in the tubesheet for both intact tubes and tubes separated above the tubesheet. 3-4 WCP15015<,0951 ant 12. doc
f 3.3.3 Sununary of Results l The contact pressures between the sleeve and tube, and between the tube and tubesheet are plotted versus radius in Figures 3-2 through 3 4. Results from these figures are
- summarized in Table 3-2.
These pressures are for the elevation [
]ted inches below the tubesheet top.) They are conservative for any lower l
t elevation in the tubesheet. 4
- Note that, in all cases, the net effect of the tubesheet rotations, thermal expansions, and pressures is an increase in the contact pressure between the sleeve and tube. This contact pressure is in addition to the interference pressures between the sleeve and tube and tube
! and tubesheet produced during installation of the sleeves. J I 3.4 Sleeve and Tube Primary Stress Limits i
- The limiting pressure loads used in the San Onofre Units 2 and 3 analyais for evaluating the sleeve primary stresses are summarized in Table 3 3. In comparing the plant specific
) loads to the generic loads found in References 1 and 2, with respect to primary stress j limits, it has been determined that all of the San Onofre Units 2 and 3 loads are enveloped
- by the generic loads for the evaluation of the sleeve and tube.
) The limiting stresses for primary stress evaluations are summarized in Table 3-4. Note that for primary stresses, the limiting condition for the sleeve is with the [ e l .]*e i 3.5 Sleeve and Tube Maximum Range of Stress Intensity and Fatigue In evaluating the maximum range of stress intensity and fatigue in the sleeve and tube, the number of transients, as well as the temperature and pressure fluctuations are l significant. A comparison of the transient cycles considered in the generic analysis to the applicable transients for San Onofre Units 2 and S is provided in Table 3-5. This comparison shows that the generic analysis considers a larger number of transients, and in general, more trmient cycles, than are applicable to San Onofre Units 2 and 3. Relative to the temperature and pressure fluctuations, the transient definitions for both
, the generic and San Onofre Units 2 and 3 design specifications are defined in terms of 3-5 WCP15015-cl3951an712. doc
changes in applicable parameters from an initial starting point, typically normal operation. Comparing the generic and plant specific transient definitions shows the San Onofre Units 2 and 3 transients differ slightly from those defined in the generic design specification in terms of the some of the magnitudes. In general, the overall magnitude of the changes are typically bounded by the generic transients. However there are some notable exceptions. [ i i 1 i j ]= c. l j A summary of the limiting stress ranges for the sleeve and tube is shown in Table 3-6, along with the accumulated fatigue usage for the limiting cross section, which is [
] ]=c compared to the limit of 1.0.
l 1 3.6 Weld Sizing for Pressure Loads
- I
+ j The LWS installation quality procedures ultrasonically (UT) scan the weld thickness over the full 360 of circumference. The UT NDE process is described in Section 7 of the generic 3/4 inch LWS, WCAP-13698, Reference L The UT laser weld inspection system can
- confirm the presence of a metallurgical bond between the sleeve and host tube, and is used to determine the existence of a leak path across the [
l I 4 I
.] *.c 4
i Table 3 7 lists the maximum primary to secondary pressure differential (AP) loads acting 1 on the LWS weld at San Onofre and the comparable design, upset, test, emergency, and ! faulted AP loads considered in the generic pressure load evaluation of the 3/4 inch ABB-CE i LWS, Reference 13. The generic pressure load evaluation is based on burst testing of LWS " assemblies for 7/8 inch OD tubes. These burst tests indicate that [
- ]=c design AP acting on the minimum 0.015 inch weld width with the minimum ASME
, Code strength properties from Reference 11 for Alloy 690, the autogenous laser weld material. I I k l 1 36 WCP15015-c!3951an712. doc
l l It is shown in Table 3 2 of Reference 13 that the above [ l
.] =.e The " enveloping" upset transient maximum AP of (
.]* c As chown in Table 3 7, the generic 3/4 inch AP loads bound the SCE San Onofre Units 2 and 3 AP loads for design, upset, test, emergency .nd faulted conditions. Therefore, the 0.015 inch minimum laser weld width satisSes all ASME Code pressure limits for LWS installation in SCE San Onofre Units 2 and 3.
3.7 Burst Strength of Laser Weld 4
- l
[ 1 1 ! ,]= e which satisSes the j minimum requirement of 3.0 in Regulatory Guide 1.121, Reference 9. Since the weld is j not expected, based upon current analysis, to experience significant degradation in service, l l this ratio applies over the life of the LWS installation. l ! l
- l I 3.8 Fatigue Evaluation of Laser Weld ,
\
- In evaluating cyclic loading, the ASME Code, Reference 10, requires that the elastically
- calculated fatigue cumulative usage fraction, summed over all specified normal, upset, and test loads, must be less than one.
,1 The normal, upset and test condition cyclic loads, specified for LWS in San Onofre, are given in Table 3 5. The generic structural evaluation for LWS in 3/4 inch tubes, given in Reference 13, include a fatigue evaluation of the weld based on the generic cyclic loads j specified in Reference 2, which bound both Westinghouse Model D-E steam generators and 3-7 WCP15015-cl3951an712 doc i
. - .- -- _ - . - . . . . - . . - . - - - - . - ~ . - . - - - .-
j also ABB-CE cWam generators. Therefore, the conservative generic cyclic loads in j Reference 2, which are also listed in Table 3-5, may be used as a conservative basis for the fatigue evaluation of the LWS repair of the San Onofre 3/4 inch tubes. This follows since
! the generic transient cyclic loads listed in Table 3 5 are based on the full power parameters j shown below, which bound the San Onofre [
,]a c also shown below. The San Onofre parameters conservatively mix the
]
overall maximum AT and maximum AP from the thermal and mechanical design flows,
- respectively.
Full Power Operating Parameters Generic [Ref 13] Versus San Onofre i ! From the above comparison, the [ i l I 1 l f Ja.c
- The purpose of the fatigue evaluation is to demonstrate that the minimum required laser
] weld axial engagement length of 0.015 inch also satisfies the fatigue limits of the ASME l Code, Reference 10, (Eu s 1) with respect to the specified normal, upset and test loads.
- Based on the above cyclic load comparisons, [
4 I- .] *.c The generic fatigue evaluation considers all types of laser welded sleeving, full length tubesheet sleeves (FLTS) [ ,]a elevated tubesheet sleeves (ETS) [ 1 , jac and tube support sleeves (TSS) [ ,]a c
- l all limiting host tube states (fully severed or fully intact), and all potential conditions at
)
the tube support plates or egg crates (locked or free). The overall maximum calculated { cumulative usage factor for the bounding cyclic loads in Table 3 5 (for 40 years) was 3-8 WCP15015<,095Lan712. doc
[ ,]* c with intact host tube and free conditions at the egg crat s. The overall maximum cumulative fatigue usage for the LWS installation in San Onofre is
.]"
It should be noted that the above [ ]=c generic cumulative usage factor is very conservatively based on [ .]u Based on actual plant operation, a much more realistic, yet conservative, number of (
)]" would give a cumulative fatigue usage of ( '
Jac usage cited above. Thus, the corresponding more realistic cumulative usage factor for San Onofre LWS is estimated to be [ .]a.c All of the above calculated LWS cumulative usage factors are substantially less than the ASME Code limit of one 3.9 Minimum Wall Thickness Requirements For computing t.,o, the pressure stress equation NB-3324.1 of the ASME Code is used. That is, AP,xit, 4 t**= P, - 0.5(P, + P,) l 4
- Where
- P. = Primary side pressure i Po = Secondary side pressure j AP = Primary to-secondary pressure differential Ri = Sleeve inside radius l Pm = Allowable stress i
i A comparison of the calculations to determine the minimum required thickness for the j generic analysis and for San Onofre Units 2 and 3 is provided in Table 3-7. These results l show the plugging margin to be 39% of the nominal sleeve thickness for San Onofre Units 2
- and 3, versus 37% for the generic analysis.
j 3.10_ Use of Preheat SG Sleeve in Feedring SG 1 An assessment has been made relative to using a sleeve designed for preheat steam
- ' generators, which has a [ ]ac inch outer diameter, in the San Onofre 2 and 3 SG tubes, i.e., in the tubes cf a feedriag type steam generator, which preferentially has sleeves of a
[ ]" inch OD. Relative to structural issues, use of the larger sleeve in this
- feedring type steam generator, is bounded by the analysis of the [ ]" inch outer diameter sleeve in the preheat steam generators.
l I 3-9 WCP15015-cl3951an712. doc i
~ _
l The critical region of the sleeve is the unexpanded zone, and this is unaffected by placing the larger sleeve in the preheat steam generator. The sleeve geometry in the expanded region will be essentially the same as when using the [ ]" inch diameter sleeve, so the tube / sleeve interface conditions will not be significantly affected. Finally, the transition region between the expanded and unexpanded regions of the sleeve will involve less of a discontinuity due to the smaller amount of expansion and it is therefore less limiting than - I the reference cases. Thus, it is concluded that use of a ( ]" inch sleeve in feed ring steam generators ! (FSGs) such as the SONGS 2 & 3 FSGs is bounded relative to structural issues by the analysis for the ;eference geometries. This (structural) section supersedes or clarifies Section 2 of Ref.1 in these areas: Page Section, Change / Clarification No. of Tab.Or Ref.1 Fig. No. 3-32 Fig. 3-1 Remove sleeve-to-tube weld at tube mouth and term " lower weld" label and
" upper" from term " upper weld". (Weld within tubesheet has been discontinued) 3-33 Figs. 3-2 Remove " Upper" terms; a TS weld has only one weld joint; the term " upper" is ,
through through unnecessary l 3-35 3-4 3-40 Figs. 3-9 Remove sleeve-to-tube weld at tube mouth and term " lower weld" label and through through " upper" from term " upper weld". (Weld within tubesheet has been discontinued) 3-48 3-17 3.11 Applicability of Thermal Hydraulic Analysis, Section 3.2 of Ref.1 to SONGS 2 & 3: The Ref.1 information applies to SONGS 2 & 3. The following additional information, is also provided for clarity. The generic sleeving report for % inch LWS sleeves, Reference 1, presented conservative values of the hydraulic equivalency numbers, Nhyd, for this type of sleeve. Since that report was issued, two factors in the assessment of total sleeve equivalency have changed. Measurements of sleeve hydraulic equivalency have shown that the calculation of hydraulic equivalency contains significant conservatism. The inherent conservatism in the calculation eliminates the need to impose additional conservatism in the analysis by assuming the most conservative sleeve configurations and plant operating conditions. Also, the possibility that large numbers of tubesheet and tube support sleeves will be installed in the San Onofre Units 2 & 3, requires that more realistic assessments of hydraulic equivalency to be used. 3-10 WCP15015-ct3951an712. doc
j Test Measurement of Hydraulic Eauivalency While the SLEEVE Code, used to calculate the hydraulic equivalency, is designed to be conservative, the degree of conservatism was established through testing in 1997 (Reference 12). The test results show that hot leg tubesheet sleeves have measured equivalency values a factor of two to four bigger than the SLEEVE Code calculation. Though cold leg sleeves are rarely used, tests of these sleeves were used to confirm the conservatism of the SLEEVE Code. Measured cold leg sleeve hydraulic equivalencies were 25 to 60% larger than those calculated by the SLEEVE Code. Although these results are not used to adjust values calculated by the Code, they justify the use of best estimate sleeve configurations and operating conditions. Best Estimate Eauivalency Calculations Calculations using the SLEEVE Code were made for several hot leg tubesheet sleeve configurations:
- 1. A 30" Elevated Sleeve
- 2. A 20" Elevated Sleeve
- 3. A 12" Elevated Sleeve
- 4. No Tubesheet Sleeve j
Each of the tubesheet sleeve configurations was run with 0,1, 2, 3 and 4 hot leg tube ; support plate sleeves. Figure 3-5 presents the results of the SLEEVE Code calculation of hydraulic equivalency for all of these configurations. The primary application of hydraulic equivalency numbers is to estimate where a plant stands with respect to its analyzed plugging limit. When the total equivalent plugging due to sleeving is added to the number of plugged tubes, the total can be compared to the plugging limit to confirm adequate primary flow will be measured. The SLEEVE Code also calculates a thermal equivalency number, Nth, which is defined as the number of sleeved i tubes which cause the same thermal performance (steam pressure) change as a single plugged tube. Thermal equivalency for the sleeved configurations evaluated are presented in Figure 3-0. The thermal equivalency numbers are consistently higher than the hydraulic equivalency numbers. Total Eauivalent Tubes Plucced The total equivalent plugging is determined by dividing the number of tubes with each sleeve configuration by the configuration hydraulic equivalency number, then summing the results for all configurations. Equivalent plugging, P., for a generator is given by P, = P, + f' h,w/,i 3 11 WCP15015-ct395 tan 712. doc
4 where P. = Number of tubes actually plugged Si = Number of tubes with the "i" sleeve configuration Ndya. = Hydraulic equivalence number of"i" sleeve configuration l To illustrate the plugging equivalency calculation and the effect of best estimate equivalency numbers on that calculation, an arbitrary sleeving configuration for San Onofre Unita 2 & 3 was assumed. Twenty percent of the 9350 tubes in each steam generator were assumed to have been sleeved. The 1870 sleeved tubes were spread evenly among the following array of 15 configurations. Tube Sheet Number of Tube Sleeve Suonort Plate Sleeves 30" Elevated 0,1, 2, 3 20" Elevated 0,1, 2, 3 12" Elevated 0,1, 2, 3 None 1, 2, 3 A total of 2992 tube support plate sleeves are included in the array. Table 3-9 lists the equivalent plugging for each of the 15 configurations and the total equivalent plugging for the generator. The equivalent plugging is calculated for three equivalency number assumptions,
- 1. Conservative hydraulic equivalency numbers from the generic report, Reference 1.
- 2. Best estimate hydraulic equivalency numbers from Figure 3 5.
- 3. Best estimate thermal equivalency numbers from Figure 3-6.
Though the present calculations are termed "best estimate," the test results noted above suggest th9t there may still be significant conservatism in these numbers. The degree of conservatism is difficult to quantify since a limited number of sleeve configurations were tested. In summary, Table 3-9 shows the conservative equivalent plugging, for 20% of the tubes sleeved, to be 2.3% hydraulic equivalency. The best estimate numbers, still judged to have conservative margin, reduce the total equivalency to 1.4%. If steam generator thermal performance, rather than approaching the analyzed plugging limit, is of primary interest, then the thermal equivalency is relevant. This total equivalent plugging is just 1.2% for 20% of the tubes sleeved. I l 3 12 WCP15015-cl3951an712. doc
l 4 l Conclusions l The total equivalent plugging which results from large numbers of sleeves remains fairly small. When calculated using the same methodology, variations of equivalent tube plugging which result from different sleeve designs are likely to be a small fraction of 1% plugging. This small difference should be a secondary factor in sleeve design selection. 1 ! 3.12 Applicability of Flow Induced Vibration Analysis, Section 3.3 of Ref. I to SONGS 2 & 3: The Section 3.3 information of Reference 1 applies to SONGS 2 & 3. I j 3.13 Conclusions i ! Based on the combined results of this analysis and the generic evaluation, the following conclusions are made regarding the structural requirements for laser welded sleeves for ] the San Onofre Units 2 and 3 steam generators.
- 1. The laser welded sleeves are concluded to meet applicable code and regulatory j requirements.
I
- 2. The allowable plugging margin for sleeve degradation is 39% for the San Onofre Units l 2 and 3 steam generators. ;
j 1 1 j f i 1 l 3-13 WCP15015-cl3951an712. doc
l I I Table 3-1 ; Tubesheet Comparisons for CE Steam Generators ;
)
3,C ' f l l 3-14 WCP15015-cl3951an712. doc
Table 3-2 Minimum Sleeverrube Contact Pressures ! for Elevated Tubesheet Sleeve a,c I a f
- Contact pressures shown result, at the limiting location in the tubesheet, i.e.,
at a radius from the tubesheet vertical centerline of approx. [ Jac.', in going from the elevated tubesheet sleeve as installed condition to the condition shown, i.e., to either normal operation or to the limiting faulted condition. f l 3 15 WCP15015-cJ395 tan 712. doc
Table 3-3 Enveloping Pressure Loads for Design, Faulted, and Test Conditions for Generic and SCE Unit 2 & 3 , a,c 1 4 I l I l l (*7 -
,)a c WCP15015-ct395lan712. doc
1 l 1 l l l Table 3-4 l 2 Summary of Maximum Primary Stress Intensity for Sleeve and Tube for l Full Length Tubesheet Laser Welded Sleeve l [ J a,c l a,c l l l This table replaces Table 317 of Ref.1. 3 17 WCP15015-ct3951an712. doc
Table 3-5 Comparison of Transient Cycles for 3/4 Inch OD Tube Sleeving Generic vs. SCE San Onofre Units 2 and 3 __a,c l
- W SCE normal plant variation is enveloped by generie small step load increase / decrease.
i
.. 3-18
( WCP15015-cl395lan711 doc l
Table 3-6 Maximum Range of Stress Intensity and Fatigue Usage in Sleeve and Tube a,C i 4 i 8 This table replaces Table 318 of Ref.1. l WCP15015-cl3951an712. doc
Table 3 7 Maximum AP Loads Used to Size Laser Welds AP = primary to secondary pressure differential (psi) a,c i l WCP15015-c!395 tan 712. doc l l
i Table 3-8 Comparison of Minimum Wall Thickness Calculations for Sleeves 3/4 inch OD Tube Generic and SCE Unit 2 & 3 Steam Generators a,c
-Notes:
(1) Enveloping Upset Transient (2) Steam Line Break Transient (3) Conservative value used 3-21 WCP15015-cl3951an712. doc
1 l Table 3-9. Sleeve Total Equivalency per Steam Generator l l l sleeve Gonliguration Hycraulic equivalence - Sest tstimate Hyaraulic; Best estimate inermal Generic Report ? Equivalence. - Equivalence Hot Leg Numoer traction Number Equivalence total equivalence total Equivalence total Tubesheet of TSP of Total of Tubes Number Equivailent Number Equivailent Number Equivailent Sleeve Sleeves Sleeved Sleeved # of Tubes # of Tubes # of Tubes _ n,c 20% of Total Tubes Steeved 3-22 WCP15015-d3951an712. doc
a,c Figure 3-1. Finite Eleinent Model of CE Channel Headfrubesheet/Shell 3-23 WCP15015-ct3951an712. doc
a,c d i 1 l l 1 1 4 3 l . 1 l Figure 3-2. Contact Pressures for Norm,.1 Conditions with an Intact Tube WCP15015<J3951an712. doc
a,c I n l l l h Figure 3-3. Contact Pressures for Normal Conditions with a Separated Tube 3-25 WCP15015-ct395 fan 712 doc
a,c i l l i l l l 2 a Figure 3-4' Contact Pres,sures for Faulted Conditions with Intact or Separated Tube 3 26 WCP15015-d3951an712. doc
a,c Figure 3-5. Hydraulic Equivalency of Alternate Sleeve Configurations WCP15015-ct3951an712. doc
a,c 1 l l 1 I i l i Figure 3-6. Thermal Equivalency of Alternate Sleeve Configurations ! 3-28 i WCP15015-cl395 tan 712.aoc
f; l ' 4.0 MECHANICAL TESTS j The results of Section 4.0 of Reference 1 apply directly to SONGS 2 & 3. However, two i changes are made to the Reference 1 text and one clarification is made to a Reference 1
- table. '1) Change No.1
- . the term for the mechanical interference fit (MIF) joint, at the
[ lower ends.of- full length tubesheet sleeves (FLTSs) and of elevated tubesheet sleeves i j (ETSs), is referred to as "HEJ" throughout Reference 1. In this site speci6c document, it is : referred to throughout as a MIF joint; the MIF joint term should be used. 2) Change No. 2: ! the acronym "WSG" in Reference 1 is corrected to "FSG" in the titles of Sectione 4.1.1.1 and ' 4.1.1.2 of Ref.1, 3) A clarification on P. 4-6 of Ref.1 involves the note at the end of the table (Tab. 4-2). The note should be changed from "** Based on installation of 2000 l tubesheet sleeves with non welded lower joints - for plant" to "** Based on an ' example of ' ; installation of 2000 tubesheet sleeves with MIF lower joints for plant. (Note: These allowable leak rates are applied to the MIF joints only; the laser weld joints are hermetic.)" e 4.1 Applicability of FSG Processes for FLTS Lower Joint to SONGS 2 & 3 The mechanical tests documented in Section 4.0 of Reference 1 for the lower joints of full length tubesheet sleeves (FLTSs) apply, with the minor changes discussed in Section 4.0 above, directly to SONGS 2 & 3. A portion of these mechanical tests determined the resistance to primary-to-secondary leakage for the mechanicalinterference fit (MIF) lower joints of ETSs and the applicable leakage criteria. The criteria for primary-to-secondary . permissible leakage per SG at normal operation (N.Op.) is listed as 500 gpd per SG in Table 4-2 of Reference 1 and the maximum leakage per SG for SONGS 2 & 3 is 150 gpd at that condition. However, it is seen that for the two cases of FLTS tests for previous ABB- 1 CE feedring steam generators (FSGs) and which are applicable to the SONGS 2 & 3 FLTS lower joint configuration, (Sections 4.1.1 and 4.1.2 of Reference 1), the leakage is essentially nil. Therefore, on the basis ofleakage, the existing process for the mechanical interference fit joint at the lower end of the FLTS is suitable for SONGS 2 & 3 at (N.Op.). The leakage limit for the limiting postulated accident condition, SLB, is listed in Table 4-2 i of Reference 1 as 1.0 gpm for the plant. This is also the limit for SONGS 2 & 3. The SLB leakage results shown for the two FSG cases in Reference 1 also apply to SONGS 2 & 3; l the leakage per sleeve at this condition, per the tests was also essentially nil. l l As shown in Section 4.2, below, the bounding sleeve resistance to pullout due to endca. l loads for N.Op. for SONGS 2 & 3 is 1605 lb. The pullout test data of Tables 4 3,4-4 and 4- ; 5 of Reference 1 conservatively show a margin of approximately a factor of two over the l 1605 lb. Compared with the corresponding sleeve required pullout resistance to the endcap i load for the limiting faulted condition, SLB, the N.Op. load U limiting. The bounding SLB
- load for previous FSGs and applied to SONGS 2 & 3,1286 lb, is exceeded by the N.Op. load i
I
-41 "
. WCP15o15-cO951an712. doc
1 i i of 1605 lb. Therefore, the existing FSG process for the lower joint of the FLTS is ! completely suitable for SONGS 2 & 3 for N.Op. and for the limiting faulted condition, SLB. i
- 4.2 Applicability of FSG Processes for ETS Lower Joint to SONGS 2 & 3 l
1 - A sleeve elevated in the tubesheet (ETS) is shown in Figure 2 2 of Reference 1. The lower , joint of this sleeve is a mechanical, non-welded joint known as a mechanicalinterference fit l ;
; (MIF) joint. Sleeve pullout resistance and leakage resistance.at this lower joint [
f j ,]..c4 4 I The ETS for the SONGS 2 & 3 SGs consists of a (
- ,]m installed in a 3/4 inch (nominal) outside diameter x 0.048 j inch (nc minal) wall thickness, Alloy 600 tube, explosively expanded in the tubesheet in the ;
j factory. 1 l
- Although the FLTS lower joints for FSG 3/4 inch tube sleeves shown in Reference 1 have been completely satisfactory, the ETS lower joints for FSG sleeves have been developed
)
, separately. Both types of joints must meet the same pullout and leakage resistance l l requirements for the respective applications. One of the reasons for separate developments l in the past was that [ ] l r i i f i 1 i 1 4 1
,]..e4 The longest ETS which has been generically evaluated in Reference 1 is [ lace and this applies to SONGS 2 & 3. The maximum length is stated because tlio maximum length is 4-2 WCP15015-cO95 tan 712. doc
4 bounding in terms of stress in the sleeve / tube structure; sleeves shorter than these involve } lower stresses in the sleeve / tube structure. The previous ETS lower joint was used in a 100 percent sleeving program. Consisting of the same sleeve, as well as the factory explosive expanded tube as the SONGS 2 & 3, it was i developed for the 3/4 inch x 0.048 inch (nom.) wall, Alloy 600 tube, in 1995. The tubesheet
- unit cells, the quantity of tubesheet material which is considered part of the tube joint
! surrounding a single tube, are the same for the previous FSG and SONGS 2 & 3 SGs; the ! pressure boundary materials, including the sleeve material, are also the same. The sleeve lower joint fabrication, including the roll expansion and hydraulic expansion processes, roll cxpander type and torque are the came for sleeves for the two FSGs. ! The[ 1
,j u .
The thermalhydraulic conditions and some of the tubesheet dimensions differ slightly between the previous FSGs and the SONGS 2 & 3 FSGs. Therefore, the sleeve lower joint ' undergoes different contact pressure changes for the SGs; the effects of these changes on leakage and pullout resistance for SONGS are shown below and in Tables 4-1 and 4-2 for N.Op and faulted conditions, respectively. The changes for the previous FSG are also shown in these tables. l 4.2.1 SONGS 2 & 3 ETS MIF Joint Leak Test Equivalence Leak testing at room temperature for the previous FSG, along with minor adjustments of the results to relate to the plant, is demonstrated to be applicable for N.Op. and accident conditions for SONGS 2 & 3 on the following basis. Normal Operation I l The change in CP in going to N.Op. from the au-instalJed condition, for SONGS 2 & 3 was determined for the most stringent conditions. This is for a tube location in the tubesheet at a radius of approximately [ l J a.c.e 4-3 WCP15015-d3951an712. doc l U- ._ _ _ _- I
1 i 4
,) ..c..
' Applied Load Effects on ETS Sleeve-to-Tube Lower Joint Roll Expansion - N.Op. a,c,e 1 The previous laboratory coom temperature (RT) leak testing results for the FSGs were applied to the SONGS 2 & 3 SGs. In this lab. test application, the [
.]m Refer to Table 41.
This is over an order of magnitude below the criterion of 5.61 dpm from Table 4-2 of Reference 1. This criterion was determined by adjusting the permissible leakage in the table from 500 to 150 gpd, taking only % of the 150 gpd for this source and assuming 1000 sleeves in the EG, in keeping with the SONGS 2 & 3 limit. Unless the tube is perforated within the length spanned by the sleeve, any leakage through the MlF joint is expected to be contained within the sleeve-to tube annular space; the weld joint at the sleeve top is hermetic, i.e., it is taken as having no leakage. For a hypothetical, conservative number of sleeved tubes with 100 percent throughwall degradation in the tube in the cleeve-spanned portion and 100 tubes per SG, the leakage during N.Op. would be insignificant. WCP15015-cO951an?t Joe
l l Faulted I An evaluation for the faulted test conditions, similar to that made for the N.Op. conditions ' above, was made and it indicates a [ .] *." I Applied Load Effects on ETS Sleeve-to-Tube Lower Joint Roll Expansion - SLB ' l a,c,e . J i i i
- The laboratory RT leak test results for the ETS lower joint from the previous FSG program were applied to the SONGS 2 & 3 FSGs. In this application, the [
l i
.]m Refer to Table 4 2. This is over an order of magnitude less than the limit of 18.7 dpm per sleeve shown in Table 4-2 of Ref.1, assuming % of the 1 gpm for this potential source, f
Unless the tube is perforated within the length spanned by the sleeve, any leakage through the MIF joint is expected to be contained within the sleeve-to-tube annular space; the weld joint is hermetic, i.e., it is taken as having no leakage. For a hypothetical, conservative number of sleeved tubes with 100 percent throughwall degradation in the tube in the sleeve-spanned portion,100 tubes per SG, the leakage during SLB would be insignificant. 4.2.2 Justification for Using the Previous FSG, MIF Lower Joint Pullout Resistance for SONGS 2 & 3 i 2d on CP Evaluation 4-5 WCP15015-cl3951an712. doc
- s b
f The previous FSG lower joint process pullout resistance at N.Op. applies to SONGS 2 & 3 because the as-opert , limiting, sleeve-to tube contact pressures (cps) will be approximately equal, 7._.5 (FSG) vs. 7113 psi (SONGS 2 & 3) at N.Op. (This is based on the top of the roll expansion of the MIF joint being 4.00 inches below the top of the tubesheet and the roll being 1.251.75 inches in length axially.) Previous FSG lower joint process pullout resistance at the limiting faulted condition applies to SONGS 2 & 3 i because the cps will be approximately equal. Refer to Tables 41 through 4-3. 4.2.3 Justification for Using Previous FSG MIF Joint Leakage Resistance
! Results as Bounding Results for SONGS 2 & 3 i
The change from the [ i
,] a.c.e As discussed in Section 4.2.1, above, bounding criteria for permissible primary-to-secondary side leakage during N.Op. for SONGS 2 & 3 were determined by allocating one-half of the 150 gpd limit, i.e., 75 gpd, to the lower joints of laser weld sleeves and by conservatively assuming that all sleeved tubes develop throughwall degradation. This allocation permits the other 75 gpd to be assigned to other potential primary-to-secondary leakage. This flow, averaged over 1000 hot leg tube joints, as an example calculation, results in a per-sleeve leakage criterion of approximately 3.91 drops per minute (dpm) per sleeved tube. With the per-sleeve leakage projected at {
j a.c.e [
,]a.c.e 1
Similar assumptions, i.e., allocation of one half of the permissible flow, for 1000 ETSs, i.e., i 18.7 dpm per sleeve for the faulted condition, (based on the 1 gpm (1440 gpd) primary to-secondary leakage for the plant) results in an average permissible flow of approximately 18.7 1 1 i l WCP15o15-c!3951an712. doc
It is estimated that, for a SONGS 2 & 3 FSG with 1000 ETSs, [
.]m This leakage would be insignificant.
L2.4 MIF Joint Pullout and Leakage Resistance Tests The pullout resistance calculated for the qualification of new ETS configurations requires the determination of the CP between the sleeve OD and tube ID in the as-installed condition. In conjunction with an appropriate coefficient of friction, from previous pullout tests involving the same materials and other conditions, the sleeve pullout resistance can be calculated. After determination of the as-installed interference fit pressure, the changes from that condition, both increasing (beneficial) and decreasing (detrimental) to joint integrity are calculated, based on tubesheet deformation and temperature and fluid pressure changes. It has been determined that for ETS MIF lower joints, [
.]m Similar to the method used in the pullout resistance calculation, the changes from the as-installed CP, as discussed above, were also used to project leakage for the RT leak test. The result of the beneficial and detrimental changes to the as-installed CP were used to bound the projected leakage at elevated temperatures.
I 4.2.5 MIF Joint Acceptance Criteria l I l Primary-to-Secondary Leakage ' The acceptance criterion for average leakrate at normal operating conditions in the previous FSG test for ETS lower joint development was 0.25 drops per minute (dpm) per sleeve. This was based on a fraction of an administrative permissible leakrate of 50 gallons per day (gpd) per SG for another ABB-CE plant. A similar limit for faulted conditions was not specified for that plant. However, the 0.25 dpm per sleeve limit was 4-7 WCP15015-ct395 tan 712. doc
I conservatively taken for the faulted condition. (Note: There are approximately 75,000 drops in one gallon.) Pullout Resistance - Normal Operation 1 A pullout resistance of three times the maximum primary-to-secondary pressure differential, times the tube cross sectional area, i.e., the "endcap" load, for normal 1 operation has been used as the requirement for sleeve MIF lower joints and it is consistent I with the ASME B&PV Code. Based on this approach, the limiting required resistance to pullout, upward, for the FSG in the previous case was 1605 lbs for a 1500 psi pressure differential, and the largest tube ID, resulting from installation of a tube in the largest hole in the tubesheet. This criterion bounds the SONGS 2 & 3 overall maximum operating j differential pressure condition of 1474 psi for Tem reduction anu 1000 plugged tubes.
- Pullout Resistance - Faulted Condition A pullout resistance of 1.43 times the "endcap" load for the corresponding primary-to-secondary pressure differential, for the limiting faulted condition, SLB, was used as the requirement for sleeve MIF lower joints and it is consistent with the ASME Code. For the FSG in the previous case, the maximum pressure differential for the limiting faulted condition was 2520 psi; this enveloped the SONGS i & 3 maximum differential pressure of 790 psi. The largest tube ID, [
.]=ce Pullout Resistance - Upset and Test Conditions These conditions had been beunded by normal operation and SLB conditions in previous evaluations and it was assumed that the same was true for the SONGS 2 & 3 case.
Pullout Resistance Limiting Condition - Conclusion 1 The limiting axial load for the joint design is the greater of the "3 AP" endcap load for the ' normal operation condition and the "1.43 times the largest faulted endcap load". In this ! case, the normal operation condition endcap load,1605 lbs. caused the largest load and was the limiting condition. 4.2.6 Results and Conclusions of MIF Joint Verification Tests - Previous FSGs Leakage Resistance - Normal Operation 4-8 WCP15015-cL395lan712. doc
! The leak test results are shown in Table 11. The average leakage under normal operation conditions, i.e.,1900 psi AP and RT, was [
,] tc..
i Leakage Resistance -SLB i The same sleeve / tube /tubesheet simulant samples used for the normal operation testing j were also used for SLB testing at a AP of 2650 psi. The average leakage under SLB conditions, i.e., 2650 psi AP was [ 4 J a.c.e Pullout Resistance l As stated above, the objective of the secondary to-primary side differential pressure testing j was to determine the sleeve-to-tube interference fit CP. The pressure at which the leakage for the previous FSG became significant was a conservative measurement of the contact pressure; the actual contact pressure was higher than the leakage initiation pressure. The l test pressures were [ i t
,} &.C.e 4.2.7 MIF Joint Conclusions The existing ETS MIF joint processes for 3/4 inch x 0.048 inch wall thickness tubes for previous FSGs will provide acceptable pullout resistances for N.Op., faulted, test and upset conditions in the [ ]cc' sequence for the ABB-CE FSGs of SONGS 2 & 3.
The FSG ETS lower joint also provides acceptable leakage resistance for application in SONGS 2 & 3; the leakage will be negligible assuming that [
]*c'in the tube portion spanned by the sleeve.
49 WCP15015-cl3951an712. doc
l 4.8 Applicability of 7/8 inch Tube LWS Test Results to SONGS 2 & 3 - The upper joint testing in Section 4.4.2 of Ref. I applies to the free span, i.e., laser weld, joints of the SONGS 2 & 3 sleeves. l i I f 9 l l l l l l I 4-10 WCP15015-ci395 tan 712. doc ___2 __. ___ _____ _ _ _ _ _ _ _ _ _ _ _ _ _
Table 4-1 Applicability of FSG LW* Elevated Tubesheet Sleeve Lower Joint Process to SONGS 2 & 3 Normal Operation ; a,c e i i i i j _ As conservatively modified by Previous FSG ETS Lower Joint results
" Note: Roll-first CPmwas 6,000 psi; CP was reduced based on previous testing for a PSG for roll-first and ,
l roll-last fabrication sequence i f Nomenclature: m; leakrate; ~: Approximately; Lbs.: #; Pressure: P (psi); inches: "; dpm: drops per minute, TGM: Thermal growth mismatch; SGTP: steam generator tube plugging I l l i 4-11 WCP15015-c!3951an712. doc w -v+-. ,w
i , i , Table 4-2
- Applicability of FSG LW Elevated Tubesheet Sleeve l 1
- Lower Joint Process to SONGS 2 & 3 Faulted Condition i a,c.e a
i ) 1 1 i i l l s As conservatively modified by Previous FSG ETS Lower Joint results
" Note: Roll-first CPwwwas 6,000 psi; CP was reduced based on previous testing fo. a PSG for roll-first and roll-last fabrication sequence Nomenclature: m; leakrate; -: Approximately; Lbs.: #; Pressure: P (psi); Inches: "; dpm: drops per minute, TGM: Thermal growth mismatch; SGTP: steam generator tube plugging 4 12 WCP15015-d3951an712. doc t.
i l l Table 4-3 Application of Previous FSG Roll-Last ETS Lower Joint for SONGS 2 & 3 Roll-Last ETS Lower Joint a,c,e 1 4-13 WCP15015-ct395hn712. doc f
! Table 4-4
- Verification Test Results - Mechanical Interference Fit Lower Joint 4 Elevated Tubesheet Sleeve i Alloy 690 Sleeve for 3/4 Inch Tube (Feedring SG )
$ 8,C,e } . l i 4 4 I 4 i i i i l 1 i j l l i l 1
' AWT = Apparent Wall Thinning 4 14 WCP15015-ct3951an712. doc
. . . . . _ . . . - . . _ _ . . .- _ -_ . ._ l
. . ~.. _ ... - _ _ _ . _ _ _ . . _ _ . . _ . . . _ _ _ _ . - . . . _ _ _ . _ _ . . . . . _ . _ _ . _ _ _ . . _ _ . _ .
Table 4-4 (Cont.) Verification Test Results - MechanicalInterference Fit Lower Joint Elevated Tubesheet Sleeve Alloy 690 Sleeve for 3/4 Inch Tube (Feedring SG ) a c.e i l J ! l i t i' l -
- ' AWT = Apparent Wall Thinning i
I. I 1 I I 4
- WCP15015<!3954an712. doc i . . _ . _ . . . - _ -.._._ _ ___._. -
1 Table 4-4 (Cont.) i Verification Test Results - Mechanical Interference Fit Lower Joint i Elevated Tubesheet Sleeve i
- Alloy 690 Sleeve for 3/4 Inch Tube (Feedring SG )
a,c,e i-L
'AWT = Apparent WallThinning i
WCP15015-c!396 tan 712. doc
5.0 STRESS CORROSION TESTING OF LASER WELDED SLEEVE JOINTS
- The information in this section supersedes the information in Section 5.0 of Reference 1. ;
I The Alloy 690 'IT (thermally treated) sleeve material exhibits exceptional resistance to stress corrosion cracking in steam generator environments. Based on all available corrosion test results, Alloy 690 TT appears immune to stress corrosion cracking in i primary water (PWSCC), and offers advantages over other candidate SG tube alloys in l 4 faulted secondary side environments. It has been the preferred alloy for heat transfer tubing in new and replacement SGs since approximately 1988; its use for sleeving extends
- back as far as 1984.
l The resistance, therefore, of the laser weld repaired sleeve joint is dictated by the resistance of the Alloy 600 tubing at the repair elevation. Hence, the major threat to the operational integrity of laser welded sleeve repairs is the magnitude of the stresses residual to the sleeve installation process. These stresses are the combined results of: (a) the hydraulic expansion of the sleeve and tube, (b) the stresses associated with the welding process, and (c) the far-field stresses that develop during post-weld thermal stress relief. l The purpose of the thermal stress relief operation is to reduce the peak residual stresses in
- the fusion weld and, for certain installation geometries, the peak stresses in the uppermost (free span) hydraulic expansion transitions. However, under conditions where the tubes are axially restrained by locking and/or denting at the tube support plates, the thermal stress relief can elevate substantially the far-field stresses that develop in the tubing.
These stresses would be additive to any remaining unrelaxed stress at the laser weld / hydraulic expansion locations. As discussed in subsequent paragraphs of this section, the role of these stresses on the corrosion resistance of tube-sleeve assemblies has been recognized and an attempt made to evaluate their effects. In view of the role played by the stress level in datermining the service performance of weld repaired SG tubing, a discussion is presented in the next subsection of the influence of the LWS process parameters and SG design variables on stress. A subsequent section reviews briefly the effects of thermal stress relief on stress levels and is followed by a summary of the results of corrosion tests performed to evaluate the resistance to PWSCC
't oflaser welded sleeve-repaired tube mockups. Included in this summary are the results of tests on as-welded mockups (i.e., without post weld stress relief), stress relieved mockups tested under conditions without applied axial loads, and stress relieved mockups tested under conditions believed to reflect the conditions which might exist under conditions of axial restraint. The laser welding processes used to prepare the test specimens are representative of the field processes using the necdymium YAG (Nd:YAG) pulsed laser currently in use by Westinghouse for sleeve welding.
5-1 WCP15015-cG951an712. doc
l l l Not all data in this section are for steam generators with M inch OD tubing. Where the intent is merely to illustrate the effect(s) of certain variables or parameters on far-field l stress or corrosion resistance, data from test programs or mockup fabrications of other size tube sleeve assemblies are also used. 5.1 LWS Process and SG Design Variables i The influence of the sleeve process parameters and steam generator design features or tube conditions on stress levels is summarized in Table 51. As installed, i.e., prior to thermal stress relief or final hard rolling, the far field stresses are generally low, on the order of a few ksi. The peak residual stresses at the laser weld, however, are quite high; they have been estimated as approaching 80 - 85% of the tensile yield strength of the sleeve-tube assembly. When corrosion tested in this condition, the failures occur in the parent Alloy 600 tube near the weld-tube interface at times [ ) ..ex. The residual stresses at the upper hydraulic expansion are somewhat lower - [
]* c#; however, this region may also be subject to relatively early corrosion failures if these stresses are not reduced.
The most practical means to relax these peak residual stresses is by thermal stress relief. For conditions in which the tube is free to expand axially, i.e., no fixity or restraint at the support plate locations, stress reliefis an efficient process. and has a negligible impact on far field stresses. However, recent experience with operating steam generators suggests this condi6n may not always exist, and it is useful to assume for conservatism that the tubes may in fact be locked at the tube support plate (s); most recent corrosion tests have thus been performed under the more conservative assumption, i.e., under conditions of applied axial stress. The consequence of a locked tube condition is that thermal stress relief, while lowering peak residual stresses at the laser weld and at the hydraulic expansion, may increase the l far-field axial stresses in the tube and may lead to bulging distortion of the tube at and above the elevation of the weld. Since this response is a consequence of the thermal expansion of the tube, the higher the stress relief temperature or the greater the axial extent of the region being stress relieved, the greater the axial far-field stresses. l Hence, the thermal stress relief process must be carefully tailored to achieve a trade-off between reduction of the peak stresses at the weld and hydraulic expansion transition while at the same time minimizing the far-field stresses. 5-2 WCP15015-ct395 fan 712. doc
In view of the influence of the tube-tube support plate span length on the magnitude of the far field stresses, optimization of the sleeve installation and stress relief process must be defined on a plant (or SG design) specific basis. 5.2 Residual Stresses vs. Stress Relief Temperature in LWS Sleeve Repairs Table 5 2 summarizes the expected range of far-field stresses that result as a function of the stress relief process. These are conservative stress values from strain gage measurements above and below the laser weld location and are for temperatures measured at the weld and upper hydraulic expansions of sleeve mockups. The data shown were measured during preparation of pre-heat model SG sleeve mockups and are directly applicable to Model D or E SGs [ j a.c.e, These data show the substantial reduction of far-field stress that can be realized in LWS. repaired SG tubing by controlling the stress relief temperature to be in the lower portion of the allowable range. They also show the general trend that separate stress relief of the upper hydraulic expansion (UHE) region tends to increase the far-field stress, and final hard rolling for the roll last sequence contributes to a small reduction in far field stress. 5.3 Corrosion Test Description Since approximately 1988, Westinghouse has used the doped steam corrosion test to evaluate the resistance of test mockups or repair assemblies to primary water stress corrosion cracking (PWSCC). This test is conducted in dense steam in an autoclave operating at 750 F (400 C). [
]a c.e, This test simulates an extreme acceleration of the corrosion process relative to that which occurs in an operating steam generator. In some respects, the doped steam test can be viewed as a stress indexing test; failure times in the doped steam test can generally be analyzed in terms of the stresses (residual and pressure) present in the test articles. In view of the dominant role stress plays in PWSCC of Alloy 600, this is a particularly valuable feature of the test.
The acceleration of the corrosion process provides the opportunity to evaluate the corrosion resistance of configurations appropriate to the repair process of interest, and avoids the need to rely on such stress-indexing tests as the stainless steel MgCh or Alloy 600 sodium 53 WCP15015-cl3951an712. doc 1
l i 1 tetrathionate tests which require surrogate materials or nonrepresentative { microstructures. I l As mentioned above, corrosion tests have been performed on tube-sleeve mockups in the i as-welded condition, and for conditions representing weld stress relief with and without I the addition of axial loading. Generally, two types of specimens have been tested. The first of these, illustrated in Figure 51, has been used to test laser weld joints in the.as-welded condition, or in tle condition following thermal stress relief of the joint, but without additional axial load. The second c.onfiguration is somewhat more complex. In this mockup test, the specimen is fabricated using a test stand as shown in Figure 5 2. The purpose of the test stand is to permit the sleeve inctallation, hydraulic expansion, welding, and post weld thermal stress relief under locked tube conditions. The nominal span length between supporting plates is varied to simulate the appropriate values for the SG model/ design ofinterest. The stresses that result from the several stages of fabrication are measured by placing strain gages above the weld location. Temperatures are recorded throughout the stress relief process. Following all specimen fabrication steps, the specimens are unloaded and prepared for corrosion testing. The configuration of the test assembly used for these tests is shown in Figure 5 3. By means of the threaded end fitting at the top of the assembly and the compression cylinder /Belleville washer assembly at the bottom, the axial load is established and maintained on the sleeve joint throughout the corrosion test. I To facilitate interpretation of the corrosion test results and to provide verification of the l aggressiveness of the test erwironment, roll expansion transition mockups, prepared of Alloy 600 tubing with known low resistance to cracking, are included in the test autoclaves I i l 5.4 Corrosion Resistance of Free-Span Laser Weld-Repaired Tubes - As-Welded , Condition Corrosion tests have been performed on laser weld repaired tube assemblies prepared j using both the CO2 and the Nd:YAG laser processes. The former process is no longer of interest and willlikely not be used for field operations; hence, data ase presented here only for the Nd:YAG process. I The corrosion tests on as-welded mockups have been performed on specimens of the configuration shown in Figure 5-1; i.e., without added axial load. The doped steam test results for these tests are summarized in Table 5-3. (Table 5 3 also includes some data for l stress relieved Nd:YAG welds.) While the data exhibit typical scatter in the results from different test sets, the results generally show that the as-welded joints [ ]a.c l l 5-4 l WCP15015-cG951an712. doc
[ ]*ce, while the specimens s+ress relieved at 1400 F exhibit much greater resistance to cracking. A hmited number of as-welded 3/4 inch tube-sleeve mockups have also been tested (ca. 1994) to support a field sleeving campaign. For these tests, failures [
]*ce, encompassing the range observed previously.
Figure 5-4 is a micrograph showing the typical failure location in these test specimens. The failures invariably occurred in the Alloy 600 base metal adjacent to the weld. The cracking is intergranular, typical of PWSCC, and is circumferential in orientation. This failure mode hae been observed in essentially all laser weld-repair mockups tested, irrespective of whether or not the specimen was stress relieved, or subjected to additional axialload during the test. 5.5 Corrosion Resistance of Free-Span Laser Weld-Repaired Tubes - with Post Weld Stress Relief In addition to the results presented in Table 5-3 referred to in the previous subsection, doped steam corrosion tests were performed on 3/4 inch tube-sleeve mockups to support the 1994 field sleeving campaign. These specimens were tested without the imposition of axial loading. One of the objectives of this test program was to evaluate the effectiveness of the post weld thermal stress relief over the temperature range [
)]* c*. The results of these doped steam tests are presented in Table 5-4.
1 These tests were, for the most part, terminated at [ ]*ce, a time period i agreed hpon with the utility as sufficient to demonstrate adequate resistance to in service l degradation through the remaining service performance of the steam generators. All specimens were post-test destructively examined by splitting and flattening. Only specimens that were intentionally stress relieved well above the Seld process maximum i temperature [ j a.c.e, ) 5.6 Corrosion Resistance of Free-Span Laser Weld-Repaired Tubes - with Post Weld Stress Relief and Conditions of Axial Load During Test Experience related to a field sleeved-tube inspection campaign indicated that restraint to axial expansion due to locking of the tube at the tube support plate (TSP) elevations could lead to " bulging" of the tube above the sleeve, and the introduction oflarge axial "far-field" 5-5 WCP15015-cO951an712. doc
stresses. This provided the incentive to include conditions of restraint both during fabrication of mockups for testing and during corrosion testing. The degree of axial restraint varies (see discussion in Subsections 5.2 and 5.3) with span length (e.g., the distance from tubesheet to TSP) and installation / fabrication parameters - in particular, with the thermal stress relief. Hence, most recent tests have used conditions which recognize these factors for the specific plant or sleeve application ofinterest. Doped steam corrosion tests have been performed on % inch tube LWS repaired mockups prepared to simulate sleeving installations for two different models of operating steam generators. A summary of the fabrication parameters, pertinent measurements made during the mock up fabication, and the results of corrosion tests is provided in Table 5-5. For each of these mockups, only the laser welded regions were stress relieved. This minimizes the increase in far-field axial stresses while providing a more efficient field installation by avoiding the need to separately stress relieve the upper hydraulic expansion (UHE) transitions. This is only practical, however, when the distance between the weld and the UHE is sufficiently small that the stress relief of the weld provides a measure of stress relief to the UHE as well. [
.ju Nine 7/8 inch sleeved tube mockups were prepared for doped steam corrosion testing using parameters appropriate to tubesheet sleeves in Series 51 SGs. These mockups were prepared in sets of three to represent the following variations in sleeve installation:
Mockuos APR-1 thru 3 - The post-weld stress relief was performed only at the laser weld elevation. [ ,
). ~
Mockuus APR-4 thru Post-weld stress relief was performed at the laser weld elevation, the tube was permitted to cool to near room temperature, and the UHE transition was subsequently stress relieved. The distance from the LW to the UHE transition was [ jm, Mockuos APR 7. -8.11 - This set of mockups was fabricated using modified tooling which establishes the distance from the laser weld to the UHE transition at [
]=.
This modification was performed to permit a single post weld stress relief, performed just above the laser weld elevation, to effect at least partial stress relief of the UHE transition as well. WCP15015-cl3951an712. doc
A suinmary of the stress relief temperatures realized during mockup fabrication, the stresse 3 used for corrosion testing and the results of testing are provided in Table 5-5. In order to test under the most conservative conditions, i.e., under conditions of maximum far-field stress, the test stress values reported in Table 5 5 include: the residual stress (from strain gages) measured durir.g mockup fabrication, plus two standard deviations; an "end cap" stress term due to internal pressurization; and an additional 10% allowance.for stress relaxaticn during testing at 750 F. These latter additional stress terms add significantly to the total stresses used during testing. For example, the following table summarizes the various stress components for each set of mockups. Avg. Stress Additional Stresses Added, ksi Mockups During Total Test Fabrication, Stress, ksi ksi 2 std. End Cap Stress dev'ns Stress Relaxation APR-1 thru -3 8,703 2,220 982 1,191 13,096 APR-4 thru -6 10,483 2,660 982 1,413 15,538 APR-7, 8, -11 8,964 2,220 982 1,217 13,383 The above test data Oows that mockups APR-4 through -6 have higher. fabrication stresses due to two stress reliefs as compared to the other mockups which experienced only one l stress relief. j The experience accrued in the fabrication and testing of tube-sleeve mockups has been used to optimize the field sleeving process so as to minimize field installation time while at l the san,e time arriving at a configuration in which the local weld stresses and far-field tube I stresses are controlled so as to maximize field service performance of the sleeve repairs. This optimization involves modifying the equipment such that the distance between the laser weld and the UHE is kept to a practical minimum, thereby permitting effective stress relief of both regions at the same time. The other important parameter that contributes to process optimization is the use of a nominal stress relief temperature in the mid range of the [
]a% Both of these factors were recognized in the test matrix described in Table 5 5.
Use of the conservatively hirh applied stresses should ensure that the corrosion data are l appropriate to the earliest field failures, i.e., those sleeved-tube assemblies with maximum residual-plus-applied stresses. 57 WCP15015-cI3951an712. doc l w-
~ Note re. Current Field Installed Laser Welded Sleeves The performance of laser welded sleeve repairs in operating steam generators has been excellent. Tubesheet and TSP sleeves have been in service in a domestic nuclear power plant for over four years with no indications of degradation. These sleeves are in tubes known to have some degree oflock up at the TSPs; far-field stresses are estimated [
]m. Stress relief was limited to the weld region and was performed. at
[ ]ut
- In a non-domestic plant, approximately 5 years of operation had been attained with LWS-repaired tubes at the time the repaired SGs were replaced, again with no incidents of degradation. In another non-domestic plant, over 11,000 elevated tubesheet sleeves were in service for approximately 24 months at the time the SGs were replaced. After approximately 10 m'onths of operation, NDE of all sleeved tubes and destructive examination of ten pulled tube-sleeve assemblies revealed no in-service corrosion degradation of the laser welds, the hydraulic expansion regions, or the tube bulges that resulted from stress relief under locked tube conditions.
In the following subsection an estimate is provided of the service performance that might reasonably be expected for sleeve installations in steam generators with % inch tubing. 5.7 Estimated Sleeve Performance An estimate of the sleeve performance was performed using experience from previous prot, rams and extrapolating the results where necessary. Two conditions were considered. These were: (a) the tubes are completely free to expand axially upon sleevinc and thermal stress relief; and (b) the tubes are rigidly fixed at the first tube support plate (TSP). Four cases were considered for the fixity,1) one of them being with 47 inch tube span length, 2) with the minimum tube span length that would result in maximum far field stresses,3) with a 15 inch tube support sleeve at TSP 1 and a 20 inch tubesheet sleeve, and 4) with 15 inch tube support sleeves at TSP 1 and TSP 2. In all cases, the stress relief was performed for the weld and not for the hydraulic expansion. Tubes Free to Exnand Axially In this case, following thermal stress relief at [ ]m of the laser weld region, the primary stresses acting on the tube-sleeve assembly .n the steam generator are the remaining residual weld stress and the operating pressure stress. Doped steam accelerated corrosion tests on mockups prepared under the condition of no axial fixity have 58 WCP15015<.!395lan712. doc
run for periods [ ]ac4 (average of five mockups no failures in test). Based on comparison with roll transition mockups, prepared of the same Alloy 600 material and tested at the same time, the sleeved tubes are projected to exhibit resistance to PWSCC for periods [ J a c". Tubes Fixed at the First Tube Suoport Plate The average failure time of the mockups with applied axial stress of 15.8 ksi (span length = 47 inches), included in Table 5-5, during corrosion testing was [ ]a c". The average failure time for the roll transitions was [ ]=ce. Based on comparison of the failure times of the mockups and the roll transitions, the laser welded tubes will exhibit resistance to PWSCC for periods greater than 20 times the operating time required to crack roll transitions in primary water. The stress relief operation in this case was carried out at a temperature of 1450 to 1500 F. The present practice is to perform the stress relief operation at a temperature of 1350 to 1400 F as indicated in Table 5-6. The lower stress relief temperature will reduce the far field stresses and correspondingly increase the failure time of the laser welded sleeved tubes. In addition, the life of the laser welded sleeves was estimated for the worst case of far field stresses. Considering the various 0.75 inch tubed SGs, the minimum distance between the top of the tubesheet and TSP 1 is 29 inches for Series D3. If the tubes are locked at the TSP 1 for this SG, it will result in maximum far field stresses. From the extrapolation of I the existing far field stresses data, the stresses for this span length were determined as [
]a c*. Considering that the time to crack the roll j transitions is [ ]ac* s s indicated in the previous example, the laser welded joints are projected to exhibit a corrosion life more than 7 times that required to crack the roll transitions in primary water A limited avaluation was performed of the situation when both a 15-inch tube support plate sleeve (at TSP 1) and a 20-inch tubesheet sleeve were installed in the same tube. In this case, the far field stress in the tubesheet-to-TSP 1 span was judged to be [
] a c' alone. The far-field stress in the span above the TSP sleeve is [
] a.c.e 5-9 WCP15015-ct3951an712. doc
i [ ,
,j u.
A summary of the estimates for the service performance of laser welded sleeve-repaired tubes discussed above, for the different conditions assumed for tube fixity, is provided in - the table below. In order to make a conservative estimate of the service performance, the l time to crack the roll transition was considered as 1.5 efpy. Roll transitions have not l cracked earlier than this time in any of the plants with 0.75 inch tubing. i l I 1 l l 1 l l j i l l d ) I i l 1 i 5-10 WCP1501543951sn712. doc -
a,c,e i Notes: (a) Ratios from doped steam results expressed as "equiv." ratios in primary water , (b) Based on earliest roll transition cracking of tubes with 0.75 inch OD > 1.5 efpy 5.8 Outer Diameter Surface Condition l Because the sleeving involves operations only on the primary side, no aspect of the sleeve installation directly involves the tube OD surfaces. In operating SGs, however, the OD surfaces undergo surface corrosion and may collect deposits. These are typically oxides or related minerals in the thermodynamically stable form of the constituent elements; in PWR secondary water, magnetite is the most prominent oxide that forms. At the temperatures experienced during sleeve welding and thermal stress relief, these compounds are stable and do not thermally decompose. All such compounds have crystal structures that are too large to permit diffusien into the lattice of the Alloy 600. Reactions j between these stable oxides and minerals and the alloying elements of Alloy 600 are ; thermodynamically unfavorable. Consequently, their presence dunng sleeve installation is I not expected to produce deleterious tube-sludge / scale interactions. This judgment has been evaluated by installing and laser welding sleeves into tubes l removed from operating plants. Following the sleeving operations, microanalytical examinations were performed to verify the lack ofinteractions. Prior to welding, the tubes had oxide deposits which contained Cu, Ti, Al, 2n, P and Ca as measured by EDAX analyses on an SEM. Following welding and stress relief the maximum penetrations of the OD surfaces were on the order of 7 to 8 m (less than a grain depth). s 5-11 WCP15015-c!395 tan 712 doc
)
Additional evaluations were performed on three areas cf an Alloy 600 U-bend section which was coated with sludge and heat treated in air for 10 minutes at 1350 F. The sludge was a i simulant of SG secondary side sludge (FeaO4, Cu, CuO, ZnO, CaSO4 and MgCh) and was 4
. applied to.the U bend using acrylic paint as a binder. Post-thermal exposure evaluations ;
indicated no general or intergranular corrosion had occurred. l The conclusions reached in Reference 1 for performance of free-span laser welded joints i apply directly to San Onofre Units 2 and 3. I 1 l l I l l I I
\
l. I i 4 1 e i i l 5 l i i 3 1 1 i 5 12 WCP15015 cf3951an712. doc
~. .. . . . ... . - . . . . - . . _ .. Table 5-1 Summary ofImpact of Laser Welded Sleeve Operations on Stresses a,c,e (a) Magnitude also depends on temperature and length of tube stress relieved. 5-13 WCP15015-d395 tan 712. doc
Table 5-2 Far-Field Stress as a Function of Stress Relief Temperature [ Data shown are for 3/4 inch OD Tube-Sleeve Mockups] a,c.e 5-14 WCP15015-d395Ian712. doc
l l Table 5-3 l l Results of 750 F Doped Steam Tests for Nd:YAG Laser Weld-Repaired Mockups a,c,e 1 l I
)
l i d l W l
)
l l l l l l 1 1 5-15 WCP15015-ct395 tan 712. doc
I l l Table 5-4 Doped Steam Corrosion Test Results for 3/4 inch Tube-Sleeve Mockups -Tested Without Axial Load a,c.e D .
.\ '3 I
L e - Stress relief was well above 1600 F max. of the field process. W 5-16 WCP15015-c!3951an712.de
Table 5-5 Sununary of Fabrication Parameters, Temperatures. Stresses, and Corrosion Test Results for % Inch Sleeve Mockups Tested With Applied Axial Load a,c,e a - l t Mean cracking time for roll transitions tested with mockup MIIEs was 78 hours WCP15015-d395lan712. doc 5-17 u._ ._ _ . _ _ _ _ _ _ _ __ _ _ _ _ . . . _ . . ,
__ __ .__ __ . _ _ . . . __. _ _ ... _ . ..- .. _ _ . . . _ _ _ . . . _ . _ - _ . . _ . _m. . ..- m ._ . _. . _ - . . . _ . - _ . Table 5-6 Summary of Temperatures, Stresses and Corrosion Test Results for 4 7/8 irach Sleeve Mockups -Tested with Applied Axial Load a,c.e i I (a) Data not recorded 1 CE = Circumferential extent 2 WT = Wall thicknets and cracking time for five roll transitions = 74 hours. (One roll transition did not crack in 244 hours which was inconsistent with the failure time of other roll transitions. This sample was not considered while calculating mean cracking time of the roll transitions.) t i 5-18 WCP15015-c!3951an712. doc
1 _._ _ a,c l 1 l i i i j Figure 5-1. Corrosion Test Specimen for Doped Steam Testing of an LWS Joint without Arial Load 5-19 WCPiS015-ct395Lan712. doc
_ a,e l 3
)
I i i j i
- h Figure 5-2 Test Stand used for Fabrication of LWS Mockups under Locked Tube Conditions WCP15015-ct3051an712. doc
t i
- a,c l i -
1 1 e E d R l
- l 1
.i. i 6 i l I 1 o a l t 1 1 4 i e 1 J 4 d 1 1
- f .._ _
1 Figure 5-3 Schematic of Test Assembly used for Doped Steam Testing of Tube /LWS Mockups under Conditions of Applied Axial Lor ding 5 21
- WCP15015-ct3951an712. doc
M t 4 a,c l e 4 f 4 a 4 i 1 } 4 l i 3 i i 4. 4 l' e 1 w A i 1 i 4 I
-' Figure 5-4 IGSCC in Alloy 600 Tube of YAG Laser Welded Sleeve Joint after 109 Hours in 750'F Steam Accelerated Corrosion Test
~ s 5-22
;- EP15015.c0951an712. doc w- - - , , - - ,
r 4-,, .,
G.0 INSTALLATION PROCESS DESCRIPTION The outline of the installation processes in Ref.1 applies to San Onofre Units 2 and 3. The discussion in Ref. I was intended to provide an overview; the detailed installation process verification steps are all speci6ed in the individual applicable Geld service procedures that will be provided by Westinghouse as part of the job. This document supersedes or clarifies Ref.1 in several areas of Section 6. These areas are shown below: Changes and Clarifications to Reference 1 Page No. of Section ChangeiClarification Ref.1 No.of Ref.1 6-1 2* Para. In Ref.1 it is stated that LWS process for % inch tube SGs will be qualified; of Sect. this was performed and the process was used over 30,000 times for % inch 6.0 tubes 6-1 4* Para. The sequence of the installation steps for tubesheet sleeves will be optimized of 6.0 & .o minimize tube far-field stresses. Therefcio, it is planned that the rol! Tab. 6-1 expansion of the FLTS and ETS lower joints be performed after welding and stress relief. 6-4 6.3 The term "HEJ lower joint" appears twice in this section; it should be clarified by using the term "MIF lower joint." in place of the term HEJ. (The hybrid , expansion joint (HEJ) term was used for a )LV sleeve which had a free span mechanical joint; it may be confusing to use the term in a document about laser welded sleeves.) 6-5 6.4 Reference is made to the laser welding of the tubesheet " upper... joint". This term is misleading; a tubesheet sleeve has only one laser weld joint; a laser weld at the lowerjoint has been discontinued. 6.4 The last two sentences in the section are ambiguous. These two sentences are replaced with the senter.ce: "The welding parameters, qualified to the rules of the ASME Code, are computer controlled at the weld operators station." 6-5 6.5 Retain only the first sentence. The following information replaces the Reweld. remaining sentences in the para.: "Also, the ultrasonic test (UT) examination of ing a completed initial weld may result in a weld being rejected. In these cases, if the sleeve / tube has not been perforated by the initial weld, an additional weld, having the same nominal characteristics as the initial weld, wil! be made at the same nominal elevation of the initial weld. If the sleeve / tube has not been perforated by the interrupted weld, or if a perforation of the sleeve is suspected at the initial weld elevation, up to two rewelds, having the same nominal characteristics as the original weld, will be made in the expansion zone inboard of the initial weld. If the sleeve / tube were perforated during interruption of the initial weld, the tube would be removed from service." 6-1 WCP15015-cl395 tan 712. doc
l l \ Changes and Clarifications to Reference 1 (Cont'd) ) I Page No. of Section Change / Clarification Ref.1 No. of Ref.1 I
\
6-6 through Section Replace Sect. 6.6 of Ref.1 with this information: I 6-8 6.6 6.6 Post-Weld Heat Treatment t Based on the results of corrosion tests of as-welded laser weld-repaired l mockups, it has been clearly established that optimum resistance to corrosion l requires the use of a post-weld thermal stress relief. The effect of the stress ! relief is to reduce the high peak stresses at the laser weld and hydraulic
- expansion locations while minimizing the far-field stresses that may develop in
! the parent tube. These effects and means to rninimize them, were discussed I 1 in Section 5. The data presented there clearly support the prudence of post- l j weld thermal stress relief. I 1
Since stress corrosion cracking is related to a large extent to residual j stresses, a reduction in the residual welding stress level will enhance the ; corrosion resistance of the LWS. The Westinghouse development program
]
determined that a stress relief in the [ j reduces the level of residual stress without significant microstruct'iral changes. Accordingly, upon completion of all of the welds, the weld sites are stress relieved using a quartz l lamp with sufficient power to maintain the tube temperature in the desired j range for [ )*
- The value of the heating power is established based on tube surface emissivity information derived from pulled tubes, operating history, visual observation, and prior heat treat programs. This post-weld heat treatment (PWHT) is effective over a tube length sufficient to a cover the actual weld length as well as any heat affected zone.
- 6-9 Figure 6-1 There are two changes to this figure
e if a reweld at an elevation different from the elevation of the initial weld is required, it will always be inboard, i.e., closer to the degradation which
- caused the tube to be sleeved, of the initial weld.
i
- l. The optional laser weld within the tubesheet has been discontinued.
l Refer to Table 6-1 of this document. The following is also applicable and is included for )i clarification: j The length of the tube heated to temperatures above the minimum stress relief
, temperature is approximately [ ]*c inches. Consideration of positioning tolerances of the j weld and the heater, and the extent of the zone to be stress relieved, has been included in heater design.
4 4 6-2 WCP15015-ct395 tan 712. doc
7.0 NONDESTRUCTIVE EXAMINATION (NDE) INSPECTABILITY Section 7.0 of Reference 1 is replaced entirely by the following Section 7.0
- The welding parameters are computer controlled at the weld operator's station. The essential
; variables, per ASME Code Case N 395, produce repeatability of the weld process. In .
l addition, two non-destructive examination (NDE) capabilities have been developed to l evaluate the success of the sleeving process. One method is used to confirm that the laser welds meet critical process dimensions related to structural requirements. The second method is then applied to provide the necessary baseline data to facilitate subsequent routine in-service inspection capability. i .)
- 7.1 Inspection Plan Logic for Installation ,
e i 7.1.1 Tubesheet Sleeve Basic Inspection Plan
; The basic tubesheet sleeve inspection plan shall consist of: )
1 A. Ultrasonic Inspection (Section 7.2) [ ]*.ce or alternate methods (Section 7.4). i 1. Verify minimum required weld width. 4 B. Eddy current examination (Section 7.3) [ ].= ce i is
- 1. Demonstrate presence of upper and lower hydraulic expansions.
j 2. Demonstrate lower rolljoint presence.
- 3. Verify weld is located within the hydraulic expansion.
4
- 4. Verify presence of a post weld heat treatment.
i
- 5. Record baseline volumetric inspection of the sleeve, the sleeve / tube joint, i
and t'ae parent tube in the vicinity of the welded sleeve joint for
- future inspections.
l l 7-1 WCP15015-cl395 tan 712. doc
l C. Weld Process Control [ ).=
- 1. Demonstrate weld process parameters comply with qualified weld process specifications.
7.1.2 Tube Support Sleeve Basic Inspection Plan . i The basic tube support sleeve inspection plan shall consist of: A. Ultrasonic Inspection (Section 7.2) [ ]*c' or alternate ] methods (Section 7.4).
- 1. Verify minimum required weld width.
B. Eddy current examination (Section 7.3) [ }.* l
- 1. Demonstrate presence of upper and lower hydraulic expansions.
- 2. Demonstrate lower joint presence
- 3. Verify welds are located widiin the hydraulic expansions.
- 4. Verify presence of a post weld heat treatment.
- 2. Record baseline volumetric inspection of the sleeve, the sleeve / tube joint, and the parent tube in the vicinity of the welded sleeve joint for future inspections.
C. Weld Process Control [ ]a.c' , 1
- 3. Demonstrate weld process parameters comply with qualified weld I process speci5 cations. l I
7.2 General Process Overview of Ultrasonic Examination. l l The ultrasonic inspection process is based upon field proven techniques which have been ! used on laser welded sleeves installed by Westinghouse. The inspection process developed for application to the laser welds uses the transmission of ultrasound to the interface region (the sleeve OD / tube ID boundary) and analysis of the amount of reflected energy from that region. An acceptable weld joint should present no acoustic reflectors from this interface above a predetermined threshold. 7-2 WCP1501%ct3951an712. doc
I 1 l Appropriate transducer, instrumentation and delivery systems have been designed and j techniques established to demonstrate the ability to identify welds with widths below the structural requirements. The entire weld interface (100 per cent of the axial and i circumferential extent) will be examined. Acceptance of welds is based upon application of criteria which is qualified by destructive examination of marginal welds. l l 7.2.1 Principle of Operation and Data Processing of Ultrasonic Examination. The ultrasonic examination of a laser-weld is schematically outlined in Figure 7-1. An ultrasonic wave is launched by application of an electrical pulse to a piezoelectric 1 transducer. The wave propagates in the couplant medium (water) until it strikes the ID of the sleeve. Ultrasonic energy is both transmitted and reflected at the boundary. The reflected wave returns to the transducer where it is converted back into an electrical signal l which is amplified and displayed on the UT display. ! The transmitted wave propagates in the sleeve until it reaches the sleeve OD. If fusion between the sleeve and tube exists, the wave continues to propagate through the weld joint into the tube. This wave then reaches the outer wall (backwall) of the tube and is reflected back to the transducer. The raulting UT display from a sound weld joint is a large signal from the sleeve ID, followed by a tube backwall " echo" spaced by the time of travel in the sleeve tube weld assembly (Ti.2.3). If no fusion between the sleeve and the tube exists, another pattern is observed with a large signal from the sleeve ID followed by a reflection from the sleeve OD. The spacing of these echoes depends on the time of travelin the sleeve alone (Tu). Additional reflections after the sleeve OD reflections are considered " multiples" of the sleeve
)
OD reflection. These are caused as the sound energy reflected off the sleeve OD bounces back and forth between the sleeve ID and OD, and decays over time. 4 [
,j a.c.e Criteria for the acceptance of a laser weld is based upon combination of the observed ultrasonic response at the weld surface, the sleeve / tube interface, and the tube OD.
An automated system is used for digitizing and storing the UT wave forms (A-Scans). l
.] *.ce The ultrasonic response from the weld is then digitized for each pulse. A typical digitized A-scan is shown in Figure 7-2.
7-3 WCP15015-ct3951an712. doc
Time intervals known as " gates" are set up over the signals of interest in the A-Scan so that an output known as a "C-Scan" can be generated. The C-Scan is a developed view of j the inspection area which maps the amplitude of the signals of interest as a function of j
- position in the tube. A combined C-scan which shows the logical combinations conditions of !
l signals in two gates with respect to predetermined threshold values can also be displayed. } Figure 7-3 shows the A, B, C, and combined C-scan display for a weld in a calibration
- standard.
t b j 7.2.2 Laser Weld Test Sarnple Results 1 l j Ultrasonic test process criteria are developed by [ I l 1 < l j ,J a.c.e
\
j Field application requires calibration to establish that the system essential variables are , l set per the same process which was quali6ed. Elements of the calibration are to: ; I i
. Set system sensitivity (gain).- Provide time of flight reference for sleeve ID, OD and '
tube OD signals.
- e Verify proper system function by scanning of the standard i l I
I { Figure 7-4 depicts a calibration standard for the sleeve weld UT exam. i j 7.2.3 Ultrasonic Inspection Equipment and Tooling i ! l [ The probe is delivered with the Westinghouse ROSA III zero entry system. The various i subsystems include the water couplant, UT, motor control, and data display / storage. , The probe motion is accomplished via rotary and axial drives which allow a range of speeds
- and axial advances per 360' scan of the transducer head (pitch). The pitch provides a high j degree of overlapping coverage without sacrificing resolution or sensitivity.
)
i 7-4
! WCP15015-c!395 tan 712. doc
_~ _ _ _ . _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ . The controls and displays are configured for remote location in a trailer outside of containment. The system also provides for periodic calibration of the UT system on the steam generator platform. 7.3 Eddy Current Inspection. Upon conclusion of the sleeve installation process, a final eddy current inspection is performed on every installed sleeve to meet the process verification and baseline inspection requirements outlined in Sections 7.1.1 B and 7.1.2 B. The combined Cecco-5/ bobbin probe is utilized towards this end to provide an enhanced baseline inspection without sacrificing data acquisition speed. The bobbin probe provides the inspection to verify the presence and location of the expansions, as well as weld location. The Cecco-5 probe provides baseline examination of the sleeve and tube. 7.4 Cecco-5/ Bobbin Principles of Operation. The standard bobbin probe configuration consists of two circumferentially wound coils which are displaced axially along the probe body. The coils are connected in the differential mode; that is, the system responds only when there is a difference in the properties of the materials surrounding the two coils. The Cecco-5 (C5) design operates as a transmit-receive probe. The C5 configuration is designed to provide detection of both circumferential and axial degradation. There are two bracelets of coils, each consisting of an array of transmit-receive sets. Each bracelet is l capable of achieving 50 percent coverage of the circumference of the tube. This is due to the l fact that there is no coverage directly underneath the coils of a transmit receive probe. For i this reason, the second bracelet is offset relative to the first to achieve full coverage. J l Transmit-receive probes are, by nature of their operational principles, less sensitive to lift- ; off effects than a comparable impedance coil. By virtue of this feature, probes can be < designed in such a fashion that the coils do not have to ride the surface of the tube in order , to achieve a reasonable level of detectability in a region of geometric change. The coupling ! of the probes with instrumentation and software designed to take advantage of their specific design features makes transmit-receive probes an attractive technology for the l inspection of sleeved tubes. l l The calibration standard used for Cecco-5 sleeve inspection includes various axial and ; circumferential notches as depicted on Figure 7-5. Notches are located in the expansion transitions as well as in the tube and sleeve free span. Figure 7-6 depicts a 20 channel strip chart plot of the calibration standard. The analysis software allows the data from the two bracelets and the bobbin coil to be displayed in an aligned fashion. The channels may 7-5 WCP15015-c!395lan712. doc J
l 1 ! I be selected so that data from each sensing point is viewed, enabling viewing of an entire tube circumference on a single screen. ] Figures 7 7 and 7-8 show the response of the Cecco 5 probe to I
] a.c.. notches in the parent tube at the sleeve expansion transition.
2 Cecco-5 Probes have been qualified to EPRI Appendix H requirements for detection in .3/4
- inch and 7/8 inch sleeved tubing.
I i 7.5 Alternate Post Installation Acceptance Methods ,
- Ultrasonic and eddy current inspection are the prime methods for post-installation weld
). quality evaluation, with eddy current examination being used as the prime in-service 1 examination technique. However, there are cases, due [
- ,]a.c.e 4
I i [ l 1 : J , i l 1 i i l l ,) a.c.e j i I In support of accepting UT indeterminate welds, several alternate strategies will be l applied, as agreed to by the implementing utility and Westinghouse. While this summary ] is not meant to preclude other methods, it is included to provide an indication of the rigor i of the alternate methods, i j 7.5.1 Bounding Inspections f [ l < 3 i l j a.c.e I
} a.c.e 7-6 WCP15015<.13951an712 doc
~
i l I 1 l 7.5.2 Workmanship Samples [ j a.o i i 7.5.3 Other Advanced Examination Techniques i As other advanced techniques become available and are proven suitable, Westinghouse may elect, with utility concurrence, to alter its post-installation inspection program. [
]b
[ 4 0
.]b In summary, Westinghouse proposes to apply alternate inspection techniques with utility concurrence as they become available. It is intending that this licensing report not preclude the use of these inspections as long as they can be demonstrated to provide the same degree or greater ofinspection rigor as the initial use methods identified in this report.
7.6 Inservice Inspection Plan for Sleeved Tubes The need exists to perform periodic inspections of the supplemented pressure boundary. The inservice inspection program will consist of the following:
- a. The sleeve v>ill be eddy current inspected upon completion of installation to obtain a baseline signature to which all subsequent inspections will be compared.
- b. Periodic inspections will be performed to monitor sleeve and tube wall conditions in accordance with the inspection section of the individual plant Technical Specifications.
77 WCP15015-cl3951an712. doc
i The inspection of sleeves will necessitate the use of an eddy current probe that can pass j through the sleeve ID For the tube span between sleeves, this will result in a reduced fill factor. The possibility for tube degradation in free span lengths is extremely small, as plant data have shown that this area is less susceptible than other locations. Any tube indication in this region will require further inspection by alternate techniques (i.e.,
)
surface riding probes) pdor to acceptance of that indication. Otherwise the tube shall be l removed from service by plugging. Any change in the eddy current signature of the sleeve and sleeve / tube joint region will require further inspection by alternate techniques prior to ecceptance. Otherwise the tube containing the sleeve in question shall be removed from ; service by plugging. ' l ) I i l 4 i a i 7-8 i WCP15015-c!395lan712. doc i m e- _. - , , . . ~ . . _ _. _ . _ .
u,c e
)
l 1 i i i j l i l l\ l l I l l t Figure 7-1 Ultrasonic Inspection of Welded Sleeve Joint 7-9 WCP15015-c!395 tan 712. doc
.- .___-______ - --___- --_--_-_-__-_-__-___ ____ _ _ _ _ - . ~
il1 c, - a m r f o e v a W 2T-7U ed 0 r e 1 uz gi 7 i t i Fg i D l a i c p y T c o d. 2 ~ 1 7 n ta 5 9 3 tc-5 1 0 5 1
^ P
^ C W
. ll ll ll1lllllllll1 l I <
a,c,e l l l Figure 7-3 A, B, C, and Combined C-Scan Display for Weld in UT Calibration Standard 7-11 WCP15015-d3951an712. doc l
_ aj,e i l Figure 7-4 UT Calibration Standard 7-12 WCP15015-c!3951an712. doc
a,c,e 1 i i I i i Figure 7-5 Cecco-5 Sleeve Calibration Standard i i 7 13 i i WCP15015.ct395 tan 712. doc 1 U
a,c.e 1 I l l l i F Figure 7-6 Strip Chart Display for Cecco/ Bobbin Data 7-14 WCP1501&d3951an712. doc
a,c,e l Figure 7-7 i Response of Cecco-5 Probe to 60% OD Axial Notch in Parent Tube Located at Expansion Transition j 7-15 : WCP15015-ct3951an712 doc
3
- h l
t 5 a,c,e - i I' l' 1 i a d 4 l i
-i J
a , i
!, l . ; i j l l'
j I 1 i I a q s Figure 7-8 Response of Cecco-5 Probe to 60% OD Circumferential Notch in Parent Tube Located at Expansion Transition 7 16 WCP15015<0951an712. doc
8.0 REFERENCES
- 1. WCAP 13698, Rev. 2, " Laser Welded Sleeves for 3/4 inch Diameter Tube Feedring-Type and Westinghouse Preheater Steam Generators Generic Sleeving Report",
Westinghouse NSD, Pittsburgh, PA, April, 1995, Addendum 1, June 1998 (Proprietary report)
- 2. Design Speci5 cation 412A24, " Laser Welded Sleeves for 3/4 inch O.D. Tubes of '
Combustion Engineering Feedring Steam Generators and for Westinghouse Model D3, D4, D5, and E1/2 Steam Generators", Dated 4/30/93 (Proprietary report)
- 3. Letter, SCE 97 200, W. Rice, lY, to M.Mihalik, Southern California Company, Assumed Loading Conditions for Laser Welded Sleeving of ABB-CE Steam Generator Tubes, 12/11/97
- 4. Letter, NSD E-SGDA 98 216, To: Steam Generator Design and Analysis File, July 10,1998: Electronic Meil. M.Mihalik, Southern California Edison, to L. Nelson, Westinghouse, Comments to Westinghouse, Topical Report,1/21/98
- 5. Letter, NSD E-SGDA 98 216, To: Steam Generator Design and Analysis File, July 10,1998: Mihalik, M. S. (SCE), e-mail to L. A. Nelson Re: Westinghouse Sleeving, dated June 5,1998. '
G. Letter, NSD.E-SGDA-98-216, To: Steam Generator Design and Analysis File, July 10,1998: Mihalik, M. S., Fax to L. A. Nelson of SONGS Teota Reduction Nominal Operating Point Data, June 9,1998.
- 7. Roark, R. J. and Young, W. C., Formulas for Stress and Strain, Fifth Edition, McGraw Hill Book Company, New York, NY,1975
- 8. Timoshenko, S., Strength of Materials, Part II, Third Edition, Van Nostrand Company, Princeton, NJ,1956
- 9. USNRC Regulatory Guide 1.121, " Basis for Plugging Degraded PWR Steam Generator Tubes (For Comment)", August,1976.
- 10. ASME Boiler and Pressure Vessel Code, Section III, " Rules for Construction of Nuclear Power Plant Components", the American Society of Mechanical Engineers, New York., NY,1989 Edition.
WCP15045-cl3951an712. doc -
'l
. 11. Code Case N 20 3, to Reference 10.
- 12. NSD.E-SGDA-97 071. " Hydraulic Equivalency Test Program for SLEEVE Code l Calibration", August 28,1997. (Proprietary report) i l
- 13. Addendum 1 to Reference 1, WCAP-13698, Rev. 2, " Laser Welded Sleeves for 3/4 inch Diameter Tube Feedring-Type and Westinghouse Preheater Steam Generators i Generic Sleeving Report", Westinghouse NSD, Pittsburgh, PA, June 1998 l (Proprietary report) i 4
i l 1 i i i l n I t 8-2 . WCP15015-d3951an712. doc - f
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- Westinghouse Non-Propt J ry Class 3 -
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I L Laser Welded Sleeves. For 3/4 Inch Diameter Tube Feedring-Type and Westinghouse Preheater-Steam Generators Generic Sleeving Report Westinghouse Energy Systems [ _[' l _ _-
Westinghoure Non-Proprietary Class 3 WCAP-13699 Revision 2 Addendum 1 SG-98-06-009 f f LASER WELDED SLEEVES FOR 3/4 INCH DIAMETER TUBE FEEDRING-TYPE AND WESTINGHOUSE PREHEATER STEAM GENERATORS 4 GENERIC SLEEVING REPORT i June,1998 1
, @ 1998 Westinghouse Electric Company All Rights Reserved WESTINGHOUSE ELECTRIC COMPANY ENERGY SYSTEM BUSINESS UNIT NUCLEAR SERVICE DIVISION P.O. BOX 355 PITTSBURGH, PENNSYLVANIA 15230 i_ _ _ _ _ _ _ _ _ _ _ _ _ . - -
.- . ._. . - - . - = _ _ - - - . - _ . - . . .. . - - - - -- . .__. - - - --
6 ABSTRACT i { 'Under Plant Technical Specification requirements, steam generator tubes are Periodically i inspected for degradation using non-destructive examination techniques. If established inspection criteria are exceeded, the tube must be removed from service by plugging, or the tube must be brought into compliance with the Technical Specification Criteria. Tube i sleeving is one technique used to return the tube to an operable condition. Tube sleeving is 4 a process in which a smaller diameter tube, or sleeve, is positioned to span the area of
- degradation. It is subsequently secured to the tube using a laser weld, forming a new pressure boundary.
4 l Recent analysis to evaluate revised operating conditions for steam generators relative to , j the integrity of laser welded sleeves determined that the finite element model used to
] initially qualify the minimum acceptable weld width of 0.015 inch under-predicted the j shear stress in the welds. This report documents analysis and test results used to verify
- the acceptability of the 0.015 inch minimum weld width for laser welded sleeves. The analysis addresses tubesheet, both fulllength and elevated, and tube support plate (i.e.
l plate and/or egg-crate for CE steam generators) sleeves. This analysis applies only to the l qualification of the 0.015 inch laser weld. The qualification of the sleeve and tube j documented in earlier revisions of the generic analysis, and in plant specific analyses, still applies. Overall, it is verified that the tube / sleeve minimum weld width of 0.015 inch is
- acceptable and meets the requirements of the ASME Code. -I 1
2 l i 4 1 1 1 i i d a I _ _ _ _ _ . _- ,,__i- .~ y, - . , . . _ , _ . . . , _ _ _ _ , , ,
_ - _ _ . - .. . . . - . . --_~ - -- - _ _ . - .- . . . . -. -I , s I i' TABLE OF CONTENTS I i l 1 j Section Title _ Pace I s
! 1 I NTRO D UCTI ON . , . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. . . . . . . . . . . . . . . . . . . . . . . . . . ........ 1-1 ; 2 AS ME CODE REQUIREM ENTS . . . . . .. . .. . .. . . . .... . ... . . . . .. . .. . .. . .. . . . . .. . . . . . . . . . . .. .. .. . 2-1 1
j 2.1 ....................2-1 I n trod u cti on . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. . . . . . . . . . . . . . . . . . . . . . . 1 2.2 Primary Stress Requirement .. ....... .......... .......... ... ..... .....................2-1 l 1 2.3 Fatigue Requirement. .... .... .. . ........................................................2-3 ' t 2.4 Maximum Range of Stress (38 ) Requirement ......... ...... .. ................... . 2-3
~
1 3 WELD ANALYTICAL VERIFICATION....... ...... ....... . ... . ....................31 3.1 Limiting LWS Host Tube Size For Pressure Testing.. . ..... . . . ........ . 3-1 3.2 Primary S tress Analysis ... . . . . . . . . . . . . . .... . . . . . . .... ... . .. . . .. . . . . . .. . .. . . . . . . . 3-3 ! 3.3 Fatigue Analysis....... . . .............................................3-7 I 4
SUMMARY
AND CONCLUSIONS.. . . . . . . . . . . . . . . . .... . . . . . . . . . . . . . . . 4-1
- 5 R E FE RE N C E S . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5 - 1 4
W a i 3 4 i 11 . a b
LIST OF TABLES Table Title Page 1 3-1 Maximum AP Loads (psi) Used to Size Laser Welds ............... ... ....... .. ... . . .. 3-18 3-2 Ratios of Maximum Calculated Average Shear Stress. .... .. . .... .. .... .... .. ... 3-19 to ASME Code Allowable Shear Stress 3-3 Summary of Test Sample Matrix ..... ....... . . .. ................................3-20 3-4 Summary of Pressure Test Results ... . ..... ... . ... .. ...... . .... ... . .. ... ... . .... .. ... 3-21 7/8 Inch OD Tubes 3-5 Loads Used in Fatigue Evaluation of .... ... .......... ........... ... . .... ........... ......... 3-22 l 3/4-W and 3/4-CE LWS from Reference (2) (NOC cycles are for 40 years.) 36 Calculated Shear Forces. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .... 3-24 on Sleeve / Tube Laser Welds Due to Indicated Unit Load Cases i 3-7 Calculation of Weld Forces / Stresses.. .... .... . ...... ..........................3-25 ; Generic 3/4 Inch OD Tube '! I l'e l l 1 3-8 Max Overall Calculated Cumulative Fatigue Usage Table. .. ... . . . . . . . . . . . 3-2 7 .l I }'e Min Weld Thickness of 0.015 inch & 40 Year Fatigue Design Life I l 1 l l iii ' L. _ ,_ -__ ,
h LIST OF FIGURES t
' Figure Title Pace 1-1 Schematic of Full Length Tubesheet Sleeve (FLTS)... . .............. .. .. . .... ..... 1-2 3- Schematic of Sleeve, Tube and Weld and ..... ................. . .. ..................3-28 Free Body Diagram of Severed Tube .
Showing Shear Force on Laser Weld 3-2 Schematic of Structural Models ... . .... .. ........... . .. .. ..... ... .... .... ...... . ... .. .. . .. 3-29 FLTS, ETS, and TSS Laser Welded Sleeve Types 3-3 Ki/ Kr as a Function of Notch Radius ...... ..... .... ...... . . . . . . . . . . . . . ... .. . .. . . 3-30 3-4 q*Kt as a Funetion of Notch Radius . ... .... .. ... . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-31 3-5 Finite Element Model - Element Layout . ...... .... . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-3 2 - 3-6 Transient Operating Cycle... .... . .. . . . . . . . . . . . . . . . . . . . . .. . . . . . . . . . . . . . . . . . . 3-33 Generic Series 3/4 Inch OD Tube : (Transient Numbers Refer to Loads Listed in Tables 3-5 and 3-8) 4 k i
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SECTION 1 INTRODUCTION Under Plant Technical Specification requirements, steam generator tubes are periodically inspected for degradation using non-destructive examination techniques. If established inspection criteria are exceeded, the tube must be removed from service by plugging, or the tube must be brought into compliance with the Technical Specification Criteria. Tube sleeving is one technique used to return the tube to an operable condition. Tube sleeving is a process in which a smaller diameter tube, or sleeve, is positioned to span the' area of degradation. It is subsequently secured to the tube using a laser weld, forming a new pressure boundary. A schematic showing an installed fulllength tubesheet sleeve (FLTS) is provided in Figure 1-1. Recent analysis to evaluate revised operating conditions for steam generators relative to the integrity oflaser welded sleeves determined that the finite element model used to initially qualify the minimum acceptable weld width of 0.015 inch under-predicted the shear stress in the welds. Calculations using a refined finite element modelin the weld region confirmed that the initial model had under-predicted the shear stress in the welds. It was subsequently determined that the shear stress for the 0.015 inch minimum weld width exceeded the ASME Code allowable. 1 The design of the laser welded sleeve is predicated on the design rules of Section III of the ASME Code. Section III of the ASME Code provides two alternative approaches to qualify l 1 a component: either analysis or experiment. Therefore, the acceptability of the 0.015 inch ' laser weld was evaluated using pressure tests to establish the pressure retaining capability of the welds (primary stress) and a revised finite element analysis to evaluate fatigue l usage in the welds. This report documents the analysis and test results used to verify the acceptability of the 0.015 inch weld width for laser welded sleeves. The analysis addresses ! l tubesheet, both full-depth and elevated, and tube support plate plate (i.e. plate and/or egg- 1 crate for CE steam generators) sleeves. This report applies only to the qualification of the 0.015 inch laser weld. The qualificatwn l of the sleeve and tube documented in earlier revisions of the generic analysis, and in plant specific analyses, still applies. Overall, this effort shows that a minimum width of I 0.015 inch for the tube / sleeve weld is acceptable and meets the requirements of the ASME Code. l 1 1-1
a,c 4 Figure 1-1 Schematic of Full Length Tubesheet Sleeve (FLTS) 1-2
SECTION 2 ASME CODE REQUIREMENTS 2.1- Introduction' - The steam generator tube sleeve repair is performed per the requirements in Section XI (IWA 4120) of the ASME Boiler and Pressure Vessel Code, Reference (1), which refer back to Section III (code of construction) as the preferred method of repair. Section III, paragraph NB 3649, of the ASME Code provides two alternative approaches to qualify a i component: analysis or experiment. Based on the weld sizing calculations in Section 3.1, the design pressure is the limiting condition, and the 7/8 inch tube is the governing geometry. The ASME Code requires a - margin of 3.0, or greater, on the ultimate tensile strength (S.) of the material, as seen in the Code lirait on primary membrane stress intensity (P.), namely: S P S S, where Sm=j-m Appendix II, Paragraph 1230, provides for tests to destruction to qualify piping, and the steam generator sleeve and tubing can be considered to be piping. The requirements that must be met using this approach are described below. l l 2.2 Primary Stress Requirement t l The evaluation of the laser weld to show compliance with ASME Code requirements for j primary stresses is based on Paragraph NB-3649. This paragraph requires that, in the case of burst testing, that the burst pressure "be equal to or greater than that of the I i weakest pipe to be attached to the piping product, where the burst pressure of the weakest ) pipe is calculated by the equation" below. ? I 2St n P=Do S = specified minimum tensile strength of pipe material I t = minimum speciSed wall thickness of pipe Do = outside diameter of pipe The minimum specified wall thickness of the bounding 7/8 OD tube is [ ]" inch. However, there is significant margin in the tube wall thickness. The minimum required wall thickness for the tube can be established using equation NB 3641.1 for piping, or 21
.-,_a . - . . . . - - - . , _ . . _ __... .-
NB 3324.1 for vessels. The equations used to calculate a minimum required wall thickness ! for the tube are shown below: NB-3641.1 P D* -- i t """ = 2(S, + Py) + A where, tmio = minimum required wall thickness P = internal design pressure Do = outside diameter of pipe S. = allowable stress intensity y = 0.4 A = 0. - NB-3324.1 PR P R,
!""" = S, - 0.5 P or S, + 0.5 P where, tmio = minimum required wall thickness P = design pressure l I
R = inside radius of pipe t Ro = outside radius of pipe Sm = allowable stress intensity J The governing minimum wall requirement for the bounding 7/8 OD tube is 0.026 inch. Substituting this value for the minimum wall thickness in the formulation for required test pressure above gives a required burst pressure for the test of 4754 psi. I In a more conservative manner, compliance with elastic ASME Code limits for design
- conditions requires a factor of safety of .nree, consistent with the allowable membrane stress for design conditions of Su/ 3, for the maximum test pressure to the maximum
! design pressure. For a design pressure of 1600 psi, this corresponds to a minimum test pressure of 4800 psi. This limit is independent of temperature, as it is based on the ASME Code specified material ultimate strength, which is constant over the range ofinterest. Both sets of criteria result in a minimum test pressure close to three times the design pressure. Since the 4800 psi is more conservative, it is used in assessing the structural ! ! integrity of the welds. 1 2-2 n 4 r - --- - ,w , r - - - --, cr ---
. _-. . . - .- . . - . _ - - - - . - - . _ ~ . . -
)
l 2.3 Fatigue Requirement The ASME Code, Reference (1), requires that the elastically calculated cumulative usage factor for the 0.015 inch minimum weld be less than unity when summed over all applied ! cyclic load conditions, namely: [uf 51, ! i / j
\
- where ui = the usage factor for the ith cyclic stress range = ru / Ni, nt = number of j occurrences for the ith cyclic load combination range, and N is the allowable number of occurrences at the stress amplitude for the ith cyclic load combination range, as read from l the ASME Code fatigue design (S-N) curves, Figures I 9.2.1 and I-9.2.2 of Reference (1), for j the autogenous weld material (Alloy 690).
1
- 2.4 Maximum Range Of Stress (3S ) Requirement The purpose of the 3S. limit in the ASME Code is to prevent incremental distortion in structures under successive thermal cycles. Thermal stresses develop in the weld due to differential thermal expansion between the tube and sleeve. There are two scenarios that
; lead to thermal stresses in the weld. The first is the case where the tubes are assumed to be dented at the first tube support plate, with the parent tube below the laser welded joint assumed to be fully severed. In this case the sleeve / tube combination tries to grow axially relative to the wrapper, which supports the tube support plate, and compressive loads
, develop that pass through the welded joint. The second scenario is for the case where the j parent tube below the weld is assumed to maintain its axial load carrying capability. In
- this case, the tube and sleeve are at different temperatures, and a shear load is developed
, across the welded joint due to differential thermal expansion of the tube and sleeve. It has been shown that for both of the above scenarios, that the combined primary plus j secondary stresses in the tube and sleeve meet the 3S. limit. The 0.015 inch minimum (to about 0.040 inch maximum) laser weld engagement length represents a highly localized region that is subjected to essentially pure shear stresses due to both pressure (end cap) and thermalloads. The weld is surrounded by elastic members (tube and sleeve) which are j within the 3S. Code limit. Even if plastic strains occur in the weld, large deformations cannot occur due to the geometry of the welded joint and the constraint of the surrounding elastic members. 1, Based on the above, the combined pressure and thermal stresses in the weld due to normal, upset and test loadings need only be considered from a fatigue standpoint in order j to show compliance with the ASME Code for cyclic operation. 2-3
4 SECTION 3 WELD ANALYTICAL VERIFICATION
, 3.1 Limiting LWS Host Tube Size For Pressure Testing
] The evaluation of the laser weld to show compliance with ASME Code requirements for l primary pressure stress is based on an experimental stress analysis as discussed in l Section 3.2. This section determines the limiting or bounding host tube size for burst pressure testing of the LWS installation. The limiting condition for the laser weld in terms of primary membrane pressure stress I occurs when the host tube is assumed to be fully severed inboard (below) the weld. + l Assuming the host tube is not locked to the tube support plates, the shear force in the weld ' must be in force equilibrium with the end cap load on the host tube, as shown in } Figure 31. Since the weld is also the pressure seal, the maximum tube inside radius j defines both the pressure drop end cap load and the shear area of the weld. Applying force : ) equilibrium to the free body of the tube portion above the assumed fully severed section, i gives: I ? End Cap Load Due to AP = Shear Load on the Laser Weld xR AP = 2 x R w r , where: R = inner radius of host tube (inch), j JP = primary to secondary pressure differential (psi), w = weld axial engagement length (0.015 inch minimum), I r= direct average shear stress on weld (psi). Solving for the average shear stress on the weld gives: R AP r= 2w The maximum average shear stress is: 4 Rmax APmax
*** ~
2 wmin Using a refined finite element model analysis of the tube / sleeve weld interface it has been shown that the stress intensity across the weld throat is very nearly twice the average shear stress, which is essentially a pure shear condition. The allowable maximum average shear stress (ra) is the ASME Code limiting value for pure shear given in NB-3227.2 of Reference (1). The ratio of calculated maximum shear stress to allowable shear stress is: 3-1 w
j L I i _T max E max O max r, 2 wm in fa l . The ratio determines the limiting host tube size and loading condition for testing. . ~ Per NB-3227.2, the allowable shear stress in the weld (ra)is: i
- Design Allowable
- ;
i i ra = 0.6 Sm = 0.6 (26.6) = 15.96 ksi . i ! Test Allowable: I l , j ra = 0.6 (0.9 Sy) = 0.6 (0.9) (35.3) = 19.06 ksi i Emergency Allowable (largest of the following): , i i j ra = 0.6 (1.2 S ) = 0.6 (1.2) (26.6) = 19.15 kai i i ra = 0.6 Sy = 0.6 (35.3) = 21.18 ksi. (use as allowable) i l Faulted Allowable: l l Ta= 0.42 Su = 0.42 (80) = 33.6 ksi i l l A summary of the maximum primary to secondary pressure differentials (dPmar) is i provided in Table 31. Loads are based on the deSned transient duty cycles in , Reference (2) for various steam generator models, identified by the corresponding nominal ; j outside diameter of the host tubest . Table 3-2 lists the calculated values for the ratio for j j each nominal host tube size and load condition. It is observed that the maximum value of '
- [ ]u occurs for the 7/8 inch OD host tube under Design Condition loads. Therefore, )
{ experimental justification of the 0.015 inch minimum weld width is performed for the ! 7/8 inch OD tube under design loads as a bounding set of parameters for the 3/4-W, 3/4 CE, $ l and 11/16 OD host tube sizes. l ! l 1 i 4 For CE steam generators, a primary to-secondary AP of 1600 psiis used for Design Conditions. I The original design qualification for the laser welds conservatively used 2250 psi for CE Design Conditions. This, however, is judged to be unrealistic. For any CE plant where a primary-to-j- secondary AP higher than 1600 psi is speciSed for Design Conditions, additional calculations j are required to qualify the 0.015 inch minimum weld. 3-2
~
l 3.2 Prirnary Stress Analysis ! i 3.2.1 Introduction i The evaluation of the laser weld to show compliance with ASME Code requirements for f primary stress is based on an experimental stress analysis following the guidelines of j Paragraph NB-3649 of the Code. Based on the results in Section 3.1 and Table 3-2, the ! } limiting geometry is the 7/8 inch OD tube under design pressure conditions. Test samples ' l were prepared for a 7/8 inch OD tube, pressure tested, and then using the test results, a j failure pressure is calculated that corresponds to a 0.015 inch minimum weld widt.h, i maxi.num tube / sleeve interface radius, and with ASME Code minimum strength ; 2 properties. The resulting failure pressure is then compared to the design pressure load of 1600 psi to determine if a factor of safety of three or greater exists, satisfying the ASME I Code requirements for primary stress. ' !. q
- 3.2.2 Pressure Test Rel.uha l Two sets of test samples were prepared. In both cases, the test samples were prepared i consistent with field installation procedures, with one exception. 'In an attempt to achieve l l an average width of 0.015 inch, the power setting was deliberately set outside the bounds ,
i of the field process (below the low end) for the first set of samples, sample numbers 1 to 15. r j For the second set of samples, sample numbers 16 to 27, the power setting was within the . l allowable power range, but at the low end. i } In the course of the sample preparation, the welds were UT inspected using the same t
- technique as in the field to determine acceptability. As a result of these UT inspections, i j several of the welds, samples (1,2,6, 7,11 - 15), were rejected and are not included as part i of this evaluation. There is also a requirement for the field process that the tube / sleeve l weld be more than 0.50 inch away from the end of the expansion zone in the sleeve. Three samples (25,26, and 27) did not meet this requirement and are also not included as part of l the evaluation. A complete summary of the test sample matrix is provided in Table 3-3, j indicating which samples have been included in the evaluation of the weld. It should be i noted that although samples 9 and 20 did not pass the UT inspection, the results are ,
i conservatively included in the evaluation. Finally, sample 19 was not included because
- accurate measurements of the weld thickness could not be made due to scarring of the ;
- sleeve OD surface.
j A summary of the average weld widths for each of the samples that were pressure tested is ; i provided in Table 3 4. Also provided is a summary of the resulting failure pressure for ; )4 each test. In a number of cases, particularly for the second set of samples (15 - 27), the { welds did not fail. Rather the tubes experienced a fish-mouthed burst failure. In order to i measure the weld size for the samples where a burst failure of the tube occurred, it was necessary to expose the weld by failing the samples in tension. A summary of the pull ! 4 forces necessary to cause tensile failure for these specimens is also listed in Table 3-4. For [ i 3-3 . l 1 4 ,w 4 --+. _ s - ,, .. -
1 l i those samples where the weld failed during the pressure test, "N/A"is listed for the pull i force. 3.2.3 ASME Code EvaluaEon As discussed in Section 2.2, compliance with elastic ASME Code limits for design conditions requires a factor of safety of three, consistent with the allowable membrane stress for design conditions of So/3, between the maximum test pressure and the mnximum design pressure. For a design pressure of 1600 psi, this corresponds to a minimum test pressure of 4800 psi. i in the course of evaluating tha test results, the failure pressures [ l 4 ja.c I ! a I ( ja.c 8 Since the field acceptance criteria for the welds is based on a go / no go relative to the minimum weld width anywhere around the tube circumference, the measured minimum weld width is used to have a consistent set of criteria. 3-4
Substituting in the above equation, [ J te The test samples satisfy the 4800 psi preasure requirement as long as the calculated Safety Factor (SFi) is 2 3.0. Using the above formulation, safety factors are calculated for each of the test samples where the sample failure occurred in the weld joint (samples 3,8, and 10). In calculating the safety factors for the test specimens, the average of the weld width readings is used as a conservative upper bound of the minimum weld width to account for variability and uncertainty in the optical readings. In the formulation above for the safety factor, use of a larget weld size is conservative, in that it results in a smaller safety factor. The resulting safety factors are shown in Table 3-4. Note that for samples,5 and 9, assuming the welds to have failed at the reported maximum test pressure is conservative, and the above algorithm is conservatively applied to the test results, with the safety factor representing a minimum value for these tuts. For the samples where the tube burst, an estimate of the weld strength is made by comparing the weld geometry and failure pressure for the ith sample to the same parameters for the samples where the w.,ld failed. The pressure that would have failed the weld in these samples is higher than the maximum test pressure that was achieved for the test. For those samples where the tube burst, an estimated failure pressure is calculated using the results of the pull force tests. Comparison of the pull force to the measured weld width shows that the samples have consistently greater strength with increasing weld size. Thus, it would be expected that these samples would also have a failure pressure for the welds that increases with increasing weld size. In order to estimate the failure pressure for the welds, a relationship is needed between the pull force and weld failure pressure. Since a pull force does not exist for samples where the weld failed under pressure, a direct one-to-one relationship cannot be established between pull force and failure pressure. For the available test data, sample 5 is judged to give the best approximation of the ratio between pull force and burst pressure, simply because it is one of the samples with a relatively small weld, and the expected failure pressure for the weld is expected to be close to the pressure at which the tube failed (based on the weld failure pressure for samples 3, 8, and 10). Thus, for each of the samples where tube burst occurred, an estimated failure pressure is calculated by scaling the failure pressure for sample 5 by the ratio of the pull force for the ith test to the pull force for sample 5. The resulting estimated failure pressures are sutamarized in Table 3-4. The estimated failure pressures are then 3-5
i ; 1 substituted in the algorithm above to calculate SF for the samples where tube burst l occurred. An overall summary of the resulting safety factors is provided in Table 3-4. For the samples where the weld failed (3,8,10), the minimum safety factor is 4.0. For the remaining samples, where a fish mouthed failure of the tube occurred, the minimum calculated safety factor is 3.5. Overall, these results show that significant margin exists relative to the strength of the 0.015 inch minimum weld width. i 3.3.4 Primary Stress Analysis Conclusions l 1 The following conclusions, regarding the pressure retaining capability of the 0.015 inch i minimum laser weld width for the 7/8,3/4-W,3/4 CE, and 11/16 inch OD host tube sizes, are based on the recults of the experimental stress analysis: ;
- 1. Scaling the results for each of the test samples to minimum strength and weld !
width requirements shows that the ASME Code factor of safety of 3.0 for design ! conditions is met for each test sample, with a minimum safety factor of 4.0 for the , I samples where the weld failed under pressure. '
- 2. Welds performed using a power setting at the low end of the field weld process specification have weld widths well above the minimum specified weld width of 0.015 inch.
- 3. Given that strength properties for representative tube and sleeve materialis well above the ASME Code minimum specified values, significant additional margin exists for field welds.
- 4. The safety factors calculated in this analysis are judged to represent a lowei bound for welds in the field. The probability of having a weld with the minimum 1
acceptable weld width of 0.015 inch, a sleeve with minimum ASME Code strength l properties, a tube with minimum section properties (maximum inside radius), and a ! tube that is completely severed below the welded joint is judged to be extremcly low. Even if all of these highly unlikely events would occur simultaneously, the results of the experimental analysis show that the ASME Code minimum margin of 3.0 for design loads is satisfied.
- 5. The pressure terts show that for welds made within the weld process specifications, using a tube and sleeve with prototypic strength characteristics, that the welded joint is stronger than the tube itself, in that the tube burst in twelve out of twelve l
cases. Overall, it is conduded that a minimum weld width of 0.015 inch for the tube / sleeve weld is structurally acceptable for all speci6ed pressure loads acting on the FLTS, ETS, and TSS laser welded sleeves for each of the host tube sizer summarized above. 3-6
3.3 Fatigue Analysis In the laser welded sleeve (LWS) repair of steam generator tubes, the Alloy 690 sleeve is hydraulically expanded against the host Alloy 600 tube and a laser inside the sleeve is used to autogenously fuse the Alloy 690 sleeve and Alloy 600 tube over the full 360 circumference, The evaluation presented in this section demonstrates that the current minimum required laser weld axial engagement (fused) length of 0.015 inch satisfies the fatigue requirement of the ASME Code, Reference (1). with respect to the specified generic cyclic loads given in Reference (2). The ASME Code fatigue requirement is given in Section 2.3. The evaluation considers the following sleeve types for installation in
)
Westinghouse steam generators with 3/4 inch OD x [ ]ax inch wall host tubes and ABB CE steam generators with 3/4 inch OD x [ Jax inch wall host tubes: Full Length Tubesheet Sleeve (FLTS), l Elevated Tubesheet Sleeve (ETS), { Tube Support Plate Sleeve (TSS). For each of the above sleeve types, the fatigue evaluation considers all possible combinations of [ lax , t 3.3.1 Geometry and Materials The weld is located at approximately the center of the 2.5 inch hydraulically expanded length of sleeve. The residual stresses due to the hydraulic expansion and the laser welding process are removed by heat treatment prior to returning the sleeved f.ube to service. Therefore, it is reasonable to assume that the strength properties (including the S-N fatigue curves) of Alloy 690 in the ASME Code also apply to the weld. Geometry The fatigue evaluation of the laser weld considers the following sleeve types: Full Length Tubesheet Sleeve (FLTS), 36 inch length, Elevated Tubesheet Sleeve (ETS),12 inch length, Tube Support Plate Sleeve (TSS),12 inch length for 3/4 Westinghouse Host Tubes. Tube Support Plate Sleeve (TSS),15 inch length for 3/4 ABB-CE Host Tubes . All hydraulic expansion lengths are assumed to be 2.5 inches with 0.25 inch long transitions at each end. Most geometric parameters are assumed to be at their nominal values, except that the minimum weld engagement length of 0.015 inch is assumed for all laser welds. The nominal outer diameter and wall thickness of the sleeve and tube are: 3-7
Westinghouse 3/4 Host Tube: ( ]*e inch wall, Sleeve: [ ]=e inch wall. ABB-CE 3/4 Host Tube: [ ]s inch wall, Sleeve: [ la c inch wall. Material Data The following material properties, taken from the ASME Code, Reference (1), are used in the fatigue evaluation: Allov 600 Tubes at .550 E E = elastic modulus = 28.85x108 psi
= Poisson's ratio = 0.3 a = mean coefficient of thermal expansion from 70 F = 7.77x10-8 ( F) 2 K = thermal conductivity = 2.57x10 4 BTU / (sec-in *F)
Allov 690 Sleeve at 600 E E = castic modulus = 27.8x108 psi
= Poisson's ratio = 0.3 a = mean coefficient of thermal expansion from 70 F = 8.16x10-8 ( M 1 K = thermal conductivity = 2.31x10-4 BTU / (sec-in *F)
SA-508 Class 2 or 2a Tubesheet at 600 E E = elastic modulus = 26.4x108 psi
= Poisson's ratio = 0.3 l
a = mean coefficient of thermal expansion from 70 F = 7.42x10-8 (*F)1 Air at 600 F (assumed in the eso between the unexoanded sleeve and host tube) K = thermal conductivity = 5.78x10 7 BTU / (sec-in *F) ; 3.3.2 Cvelic Load Conditions i The cyclic loads used in this evaluation are defined in the generic laser welded sleeve (LWS) design specification for the 3/4 inch OD host tubes, Reference (2). The normal, upset, and test transients are generally given in terms of transient changes relative to an initial steady state condition, usually at full power operation. The full power conditions, which are assumed in the generic LWS fatigue evaluation, are listed below: 3-8
l 100% Full Power Steady State Conditions Assumed in the Generic 3/4 LWS Fatigue Evaluation (Conditions Apply to Both Westinghouse and ABB-CE 3/4 LWS) _ a,c All of the pressure and thermalloads are deSned in terms of the transient values of the four parameters, Pp, P.t , Th, and T.im, whose full power steady state values are defined above. The variations in these load parameters, and the specified number of occurrences (cycles) for each transient, are obtained from the design specification, Reference (2). Table 3 5 lists values of Pp, P.tm, Th, and T t and the number of occurrences (NOC) for each of the specified transients for a 40 year fatigue design life. The transient load values in Table 3 5 are relative to the above assumed generic full power conditions, for which the 4 primary to secondary pressure differential is [ ]*c psi, and the hot leg (primary) to steam (secondary) temperature differentialis [ ]* c F. These generic full power load differentials and the number of occurrences (NOC in Table 3-5) conservatively bound the various plants with Westinghouse and ABB-CE 3/4 inch OD host tubes. For the primary and secondary hydro test load events in Table 3-5, the pressure drops are limited to the design pressure (see footnotes on page 3-2 and to Table 3-1). 3.3.3 Thermal Analysis The purpose of the thermal analysis is to estimate the tube and sleeve temperatures, as functions of the primary side hot leg temperature (Th) and the secondary side steam temperature (T tm), to calculate the thermal-structuralloading on the laser welds for the fatigue evaluation. The hot leg is limiting structurally since the largest difference between the tube and sleeve temperature occurs in the hotleg. The heat transfer coefficients (inside the sleeve and tube and also outside the tube) are very high compared to the heat capacity and the conduction of the thin-walled metals. Since the thermal transients are relatively slow, it is reasonable to expect the rather thin tube and sleeve wall temperatures to follow the transients without significant thermallag. In addition, the axial or longitudinal temperature gradients are small compared to the radial gradients which transfer heat from the primary water inside the tube at about 600 F to the steam outside the tube at about 500 F. Therefore, axisymmetric " slab" finite element thermal models of the 7/8 LWS were employed to calculate the radial gradients in the sleeve and tube above the tubesheet. (The 3/4 LWS sleeve and tube will exhibit similar thermal responses. Therefore, the results of the 7/8 LWS thermal analysis also apply to the 3/4 LWS.) Three 3-9
sections, or slabs, were simulated to estimate the radial temperature gradients through the walls of the sleeve and tube above the tubesheet: (1) a section through the unexpanded sleeve, gap and tube; (2) a section through the expanded sleeve and tube; (3) a section through only the tube at the far field above the sleeve. In section (1) of the thermal model, the gap between the unexpanded sleeve and tube was filled with thermally conducting static air, plus nonlinear thermal radiation link elements were also used to thermally connect the sleeve and tube finite element nodes across the gap. The resulting calculated tube and sleeve wall temperatures lead to the following conclusions: j (1) [ ]ac (2) ]ac (3) [ lac {
]ae Thus, the average temperature distribution of both the tube and sleeve may be determined for any of the transient cyclic loads specified in Table 3-5 using the listed values of Th and T.tm and the above conclusions.
3.3.4 Structural Analysis Models Used in Faticue Evaluation , Based on the thermal analysis, [ j d 4 d ja.c As stated in Section 3.1, a state of essentially pure shear exists (on the average) at the weld, and the average stress intensity across the weld is essentially given by 2r. Therefore, ' l the radial, hoop, and axial stress components at the sleeve-tube laser welded interface are ! 3 - 10
. _ _ _ . _ _ _ ._ _._._ __ _ _. _ _ _ _ . _ . _ _ _ _ . _ ~ .._.__ _ _ _ _ _.. _ _ _._._.
not required to calculate the average stress intensity on the weld for the fatigue evaluation. ! The shear force F acting on the laser weld is calculated by finite element simulation of the various sleeve types, host tubes, adjacent tube bundle, tubesheet, and the welds. Since only [ j a.c The effect of the relatively " rigid" tubesheet on the tube thermal expansion is simulated using a very stiff spar element that has the material properties of the tubesheet (SA-508 Class 2) and spans the distance from the top of the hard roll to the top of the tubesheet. The tubesheet, host tube, and adjacent tube bundle nodes (located at the top of the tubesheet) are coupled in the vertical direction to force the tubesheet longitudinal expansion onto the tubes. This approximates a fulllength hard rolled tube-tubesheet i l condition and is conservative compared to a partial hard roll, since free tubes (inside the ' tubesheet) would expand more due to the higher expansion coefficient of the tube compared to the tubesheet, resulting in less of a thermal expansion mismatch with the sleeve. In all of the [ ja.c j In all of the ( 1 4 6 l 8 Ja.c 3 - 11 i ' l
P The end cap loading due to pressure is applied to the uppermost nodes at the 2nd TSP in each model. If the pressure load on the [ 4 4 i i Ja.c i Thermalloads are simulated [ 1 ja.c I 3.3.5 Shear Forces on Weld Due to Unit Loads The LWS structural finite element models discussed in Section 3.3.4 and shown in i Figure 3 2 were used to calculate the shear forces acting in the various longitudinal spring elements simulating each laser weld. The following 24 possible model combinations were simulated: ' l 2 [ !
- i l
l l
)..e The following four unit load cases were run for each of the 24 above model combinations:
- a,c y
i j i d _
- All thermal expansion cases are relative to a reference temperature of 70 F. Table 3-6 lists
) the resulting finite element calculated shear forces due to the above unit loads for each of the 24 possible combinations of[
- 1*c 1
3 - 12
. . ... . - -....~ - ...- ---. - . _ . _ _ . . . . _ - _ . . . . - . --- -
3.3.6 Stress Ranee Calculations ! For any of the 24 combinations of sleeve type, host tube state, and TSP condition listed in l Table 3-6 (say the kth combination), the shear force F on the weld may be found for any of l the 71 load events (say the ith load event) listed in Table 3-5 as follows: i a,c ! i i 4
- whers(P,, P.,m, TA, T,,m), for the particular ith load event are defined in Table 3 5, and -
- i (Fe, iooo, Fr.,..iooo, Fra sto, Fr.,m.sro)& are the finite element calculated shear forces due to l the unit loads in "'able 3-6 for the kth combination. The shear force F4 (in Ibf) is divided by ;
j the weld's assumed minimum shear area [
]*c The typical j average shear stress range for the ith and jth load events is:
e 4 Ary= rf - rj
= ,
- w 1
j where Fj is the shear force for the loads (P,, P.,m, Tu, T.,m)j as given by an expression i similar to the above expression for F4 1 j For pure shear conditions at the weld, the average stress intensity range is: (3,)y = 2 Ary = 2 r, - rj ] The calculated [ I I 1 j-4 J a.c i The number of applied cycles ny for the i j stress range is: . 4 i
"# " " #1 *"0#
- 1. USE_ MIN VALUE
- 3 - 13 ,
e
, _ _ _ ~ .. . - _ _ . _ . _ _ ._ .__ _. .__ - . . . _ __ ._. - _ i l where: noes = current number of unused cycles for ith load event, i nocj = current number of unused cycles for jth load event, l i The fatigue usage factor up for the i-j stress range is: i . "U " y . . ' ny
]
v i r k where Ng = allowable cycles from ASME Code S-N curve for (S.)v. ! The number of applied cycles remaining for the next usage calculation, involving either the i I the ith or jth load events, is reduced by no, as shown below: i (nocj)FOR_NEXT_ USAGE _ CALC = MOC/ - ng , 0 , l
)
(nocj)FOR_NEXT_ USAGE _ CALC = nocj - ny ,0 .
- i in forming the average shear stress range on the weld Anj. n and n are selected such that the absolute value of the range Anj is always the maximum of all currently active n and n, i subject to the order in which the various transients can logically combine, as discussed in -
Section 3.3.8. When a load event's current number of cycles (noc) reaches zero, that load event has been used up, and it is removed from the process, which is repeated until all cycles for all load events are used up. i 3.3.7 Faticue Strenzth Reduction Factor e l Calculation of the stress amplitude (Sa)v requires definition of the fatigue strength reduction factor Kf, which gives the combined effect of the local stress field at the surface of a notch like section, and the material's microstructure relative to the average stress over the section. ( l 1 3 - 14 I i i l
e-fie--L.dy - Jude w,=l' A4c,y 4 J M 4, Ja. 4 44,.s 48.vh.2. 5 4. - J A .5- MJ Auwdif4 4 agre*2.-4 M A. h 4Am.-W-MD* Wd 5-84-d. -%M.e- E u.M 4->=4J-.m
=
_ e4 8 8 4 4 #4a .-ew4wsh4-- _ .a ,t,. 4._e 4_ 4A4r-- arm *e f i 1 e i Y 4
^
i ' c I' e d i' 8 e D i i k a B y i f I l I ! i i 4 4 i l f 1 4 4 1* r a
?
e 1 e 1 4 k j
, i I
i r 4 I l I, I l 3 - 15
.. - - ~~
lae. This then, is the fatigue strength reduction factor, Kr, used in the fatigue calculations. 1 3.3.8 Calculated Cumulative Usace Factors Cumulative fatigue usage factors are calculated for each of the [ Jae listed 'n Table 3-6 to assure that the maximum cumulative usage factor has been determined. All calculated cumulative usage 4 factors are substantially less than the ASME Code limit of one. The overall largest calculated fatigue usage factor of [ ]=e occurs for the [
)e The overall maximum cumulative fatigue usage factor of [ ]=e was calculated using the pr ocess discussed in Section 3.3.6 to obtain the stress ranges, subject to a logical ordering of the transients. Using the FE calculated weld shear forces for the pressure and temperature unit load cases from Table 3-6 for the maximum cumulative fatigue usage condition, the shear force transferred across the weld is calculated for each of the 71 transient load conditions specified in Table 3-5. A summary of the resulting calculated
- weld shear forces and average shear stresses is provided in Table 3-7. Using the transient conditions summarized in Table 3-7, a load histogram is developed to logically order the transients, as shown in Figure 3-6. This histogram is then used to establish fatigue cycles for the various transient conditions. A summary of the resulting fatigue cycles, stress ranges, stress amplitudes, allowable cycles and fatigue usage is contained in Table 3-8, for which the calculated cumulative usage factor is [ ]=e.
Note that load range combination 9 (load 5 in Table 3-5 minus load 19 in Table 3-5), which accounts for the majority of the usage [ lae, involves a conservatively large number ofload/ unload transients. The number ofload/ unload transients ([ ]=e in 1 Table 3-5) from the generic specification, Reference (2), assumes a load ramp from full power to hot standby approximately once a day, everyday, throughout the design lifetime. { This represents a very conservative generic bounding value. In actuality, most plants base load (i.e., remain at 100% power), and the actual number ofload/ unload transients over 40 l years is a fraction of the [ ]=e cycles. Assuming a more realistic, yet conservative, number ofload/ unload transients of [ ]*e rather than the [ ' Jae for u9, results in a cumulative usage fexor of [ lae compared to [ lae for the reference number of transients. ! i 3.3.9 Faticue Analysis Conclusions A conservative analytical fatigue evaluation of the 0.015 inch minimum weld engagement length for laser welded sleeves has been completed. The fatigue evaluation considered [ 3-16
]=2 All calculated cumulative fatigue usage factors are less than the ASME Code allowable of one. The fatigue evaluation is conservative for the following reasons:
- 1. In the fatigue evaluation all welds are assumed to be at the minimum thickness of 0.015 inch around a full 360' of circumference. In practice, due to the QA inspection procedure, the weld thickness (averaged over the circumference) is significantly larger than 0.015 inches (see Table 3-4).
j
- 2. The number ofload/ unload transients in the generic specification, Reference (2), is very conservatively based on operating the plant usir g load follow. Using the maximum number ofload cycles from the generic specifications, without modification, gives an overall maximum r.umulative usage factor of [ la c.
Actually, all Westinghouse plants base load (i.e., remain at 100%), and the actual number ofload/ unload transients are a fraction of the number given in the generic specifications. Assuming a much more realistic number of [ ]" load / unloading cycles over a 40 year fatigue design life, results in an overall maximum cumulative i factor of [ ]". Based on the analytical evaluations discussed in Section 3.3, it is concluded that the 0.015 inch minimum thickness weld satisfies the ASME Code fatigue cumulative usage 3 factor limit of one or less for a 40 year fatigue design life, for all 3/4 inch OD host tubes, i Westinghouse and ABB.CE, for all sleeve types [
]" and all combinations of specified normal, upset, and test service conditions and loads.
i 1 i Y 3 17
Table 3-1 Maximum AP Loads (psi) Used to Size Laser Welds a,c a
- Section XI (IWA 4700, IWA 5000, and IWB 5000) specifies that hydrostatic test loads will be run on these repair welds that will be below the design AP load at a maximum test temperature of 600 F.
l l l 1 1 3 - 18
Table 3-2 Ratios of Maximum Calculated Average Shear Stress to the ASME Code Allowable Shear Stress a,c emumpeu 1 1 l l l l 1 i 1 l l 1 3 - 19 1 1 1
- - . . . - . - -._ . - . - . . = _ - _ _ -- ._- ._
Table 3-3 , Summary of Test Sample Matrix , Sample UT Results") Pull Tested Data Used
- Tested 1 1 I%3V No --- ---
j 2 1%3V No --- -- 3 ACC Yes Yes Yes
~
j 4 Not Evaluated o --- -
- 5 ACC Yes N/A Yes
- 6 IWTV Yes N/A No(2)
! 7 IWW No --- --- 8 ACC Yes N/A Yes
- 9 IWW Yes Yes Yes' > ~
- 10 ACC Yes N[A Yes -
11 IWW No -- --- 12 IWW Yes N/A No(2) 13 1%TV No -- --- l 14 IWW No --- ---
~~
) 15 IWW .. . No ! 16 ACC _ Yes Yes Yes 17 ACC . Yes Yes Yes 18 ACC Yes Yes Yes 19 ACC . Yes Yes No* 20 SHI Yes Yes Yes'" j 21 ACC Yes Yes Yes 22 ACC Yes Yes Yes 23 ACC Yes Yes Yes 24 ACC Yes Yes Yes 25 ACC Yes Yes No* 26 ACC Yes Yes No* 27 ACC Yes Yes { No* (1) IWW Insufficient Weld Width / ACC-Accept / SHI-Surface Indication (2) Weld did not pass UT requirement I (3) Although weld did not pass UT requirements, results conservatively 4 included in weld evaluation (4) Scarred OD sleeve surface prevented accurate weld measurements (5) Tube was cut too close to weld (outside field tolerance for weld location) 3-20
__ - . . -. . - - - . . - . = _ . . . . - . . _ . _ - ! 1 1 1-3 i Table 3-4 j Summary of Pressure Test Results
- 7/8 Inch OD Tube
l i.
- _ a,b,c i i
i l l ) , I i
- i r
i 3-i i i 1 i i , ! i l ! i i d' i 4
- Notes
- (a) Estimated based on results of pull test.
} (b) In an attempt to achieve an average width of 0.015 inch, the power setting was deliberately set outside the bounds of the field process t (below the low end) for the first set of samples, sample numbers 1 to 15. 1 For the second set of samples, sample numbers 16 to 27, the power i setting was within the allowable power range, but at the low end.
- . l i
d 4 4 l J l w l 1 I ,! I 3 - 21
Table 3-5 (part 1 of 2) Loads Used in Fatigue Evaluation of 3/4-W and 3/4- CE LWS from Reference (2) (NOC cycles are for 40 years.) as 4 T e m 3 - 22 1
Table 3-5 (part 2 of 2) Loads Used in Fatigue Evaluation of 3/4-W and 3/4- CE LWS from Reference (2) (NOC cycles are for 40 years.) a,c i 1 af 5 3 23
Table 3-6 Calculated Shear Forces (Ibf) on Sleeverfube Laser Welds Due to Indicated Unit Load Cases a,e 4 W i l l l 1 l 1
)
m 3 - 24
Table 3-7 (part 1 of 2) Calcul- tion of Weld Forces / Stresses Generic 3/4 Inch OD Tube
)e e
[ a,e 4 l l l l l l i A l l 3 - 25
Table 3-7 (part 2 of 2) Calculation of Weld Forces / Stresses Generic 3/4 Inch OD Tube [ ) ..e 8,C 1 ~ 3-26
T Table 3-8 Max Overall Calculated Cumulative Fatigue Usage Table [ ]=e Min Weld Thk of 0.015 inch & 40 Year Fatigue Design Life "f ! w m
-e- e----9 -, - e n - m a n_. e _ i m -r x 2 *
- eme d
a,c b i Figure 3-1 Schematic of Sleeve, Tube and Weld and Free Body Diagram of Severed Tube Showing Shear Force on Laser Weld 3 28
a,c d 4 i l Figure 3-2 Schematic of Structural Models FLTS, ETS, and TSS Laser Welded Sleeve Types 3 - 29 i
a,c l 2 J Figure 3-3 K / Kr as a Function of Notch Radius 3-30
a,C
~
4 1 1 4 l l 1 Figure 3-4 4
- Ki as a Function of Notch Radius 3-31
3,C 1 .a s 4 I r 1 f l .. l I i 1 l Figure 3-5 Finite Element Model- Element Layout 3 -32
a,e c Figure 3-6 Transient Operating Cycle Generic 3/4 Inch OD Tube (Transient Numbers Refer to Loads Listed in Tables 3-5 and 3-8)
.- . . _ . ~ - . . . - - .
1 i t SECTION 4
SUMMARY
AND CONCLUSIONS j , 4.1 Primary Stress Analysis Summary and Conclusions The following anclusions, regarding the pressure retaining capability of the 0.015 inch i minimum laser weld width for the 7/8,3/4-W,3/4-CE, and 11/16 inch OD host tube sizes, are based on the results of the experimental stress analysis documented in Section 3.2:
- 1. Scaling the results for each of the test samples to minimum strength and welil
) width requirements shows that the ASME Code factor of safety of 3.0 for des;gn l conditions is met for each test sample, with a minimum safety factor of 4.0 for the samples where the weld failed under pressure.
- 2. Welds performed using a power setting at the low end of the field weld process specification have weld widths well above the minimum specified weld width of
) 0.015 inch. j 3. Given that strength properties for representative tube and sleeve material is well above the ASME Code minimum specified values, significant additional margin exists for field welds. f
- 4. The safety factors calculated in this analysis are judged to represent a lower bound d
for welds in the field. The probability of having a weld with the minimum acceptable weld width of 0.015 inch, a sleeve with minimum ASME Code strength properties, a tube with minimum section properties (maximum inside radius), and a j tube that is completely severed below the welded joint is judged to be extremely low.
- Even if all of these highly unlikely events would occur simultaneously, the results of the experimental analysis show that the ASME Code minimum margin of 3.0 for
-I design loads is satisfied. *
- 5. The pressure tests show that for welds made within the weld process specifications, using a tube and sleeve with prototypic strength characteristics, that the welded joint is stronger than the tube itself, in that the tube burst in twelve out of twelve cases.
Overall, it is concluded that a minimum weld width of 0.015 inch for the tube / sleeve weld is structurally acceptable for all specified pressure loads acting on the FLTS, ETS, and TSS laser welded sleeves for each of the host tube sizes summarized above. 4-1
l 4.2 Fatigue Analysis Summary and Conclusions Based on the results of the fatigue analysis in Section 3.3, the following conclusions are made regarding the cyclic loads on the welds.
- 1. The overall maximum calculated cumulative fatigue usage factor, for the two 3/4 host tubes (Westinghouse and ABB-CE) and the three types laser welded sleeves (FLTS, ETS, and TSS), is [ }*.c for a 40 year fatigue design life and is less than the ASME Code allowable cumulative usage factor of one.
- 2. A maximum overall cumulative fatigue usage factor calculated using the number of-load-unload cycles based on expected plant operation would result in a maximum fatigue usage factor of about [ ], much less than the above [ ]*.e value, which {
is based on the very conservative number ofload-unload cycles specified in i Reference (2), 4.3 Overall Structural Conclusions for 0.015 inch Minimum Laser Weld I Overall, it is concluded that the currently licensed minimum weld width of 0.015 inch for the sleeve-tube laser weld is structurally acceptable and meets the requirements of the ASME Code for the specified pressure and cyclic normal, upset and test transient loads. I 4-2
SECTION 5 REFERENCES
- 1. ASME Boiler and Pressure Vessel Code, Section III, " Rules for the Construction of Nuclear Power Plant Components," and Section XI, " Rules for Inservice Inspection of Nuclear Power Plant Components," American Society of Mechanical Engineers, New York, NY,1989 Edition.
- 2. Design Specification 412A24, Rev. O," Laser Welded Sleeves for 3/4 inch O.D. Tubes of Combustion Engineering Feedring Steam Generators and for Westinghouse Model D3, D4, D5, E1/2 Steam Generators." April 30,1993 (Westinghouse Proprietary)
- 3. R. E. Peterson, Stress Concentration Design Factors, John Wiley & Sons, Inc., New York, NY,1953.
- 4. R. C. Juvinal, Engineering Considerations ofStress, Strain, and Strength, McGraw Hill Book Company, New York, NY 1967.
- 5. H. J. Gough, Engineering Steels Under Combined Cyclic and Static Stresses, The Institution of Mechanical Engineers - Proceedings, January - December 1949, Volume 160.
l 5-1
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