ML20210D283
| ML20210D283 | |
| Person / Time | |
|---|---|
| Site: | Farley  |
| Issue date: | 04/30/1987 |
| From: | WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19292H205 | List: |
| References | |
| WCAP-11179, WCAP-11179-R01, WCAP-11179-R1, NUDOCS 8705070110 | |
| Download: ML20210D283 (161) | |
Text
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l WESTINGHOUSE CLASS 3 SG-87-04-024 l WCAP-11179 '
REVISION 1 i \
l J. M. FARLEY UNITS 1 Ale 2 STEAM GENERATOR SLEEVING REPORT (Mechanical Sleeves)
April 1987 PREPARED FOR ALABAMA POWER COMPANY l ..
l l
WESTINGHOUSE ELECTRIC CORPORATION STEAM GENERATOR TECHNOLOGY DIVISION P.O. 80X 855 PITTSBURGH, PA 15230 aa45M/;aZa!7.43 I e 50Jgy$$$80 P
TABLE OF CONTENTS Section Title Page
1.0 INTRODUCTION
1-1 2.0 SLEEVING OBJECTIVES AND BOUNDARIES 2-1 2.1 Objectives 2-1 2.2 Sleeving Boundary 2-1 3.0 DESIGN 3-1 3.1 Sleeve Design Documentation 3-1 3.2 Sleeve Design Description 3-1 3.3 Design Verification: Test Programs 3-6 3.3.1 Design Verification Test Program Summary 3-6 3.3.2 Corrosion and Metallurgical Evaluation 3-7 3.3.3 Upper and Lower Joints 3-17 3.3.4 Test Program for the Lower Joint 3-35 3.3.4.1 Description of Lower Joint Test Specimens 3-35 3.3.4.2 Description of Verification Tests for 3-35 the Lower Joint 3.3.4.3 Leak Test Acceptance Criteria 3-37 3.3.4.4 Results of Verification Tests for Lower Joint 3-39 1
3592M/060986: 49-2
l TABLE OF CONTENTS (Continued)
Section Title Page 3.3.5 Test Program for the Upper Hybrid Expansion Joint (HEJ) 3-44 3.3.5.1 Description of the Upper HEJ Test Specimens 3-44 3.3.5.2 Description of Verification Tests for the Upper HEJ 3-46 3.3.5.3 Results of Verification Tests for the Upper HEJ 3-46 3.3.6 Test Program for the Fixed-Fixed Mockup 3-50 3.3.6.1 Description of the Fixed-Flxed Mockup 3-50 3.3.6.2 Description of Verification Test.s for 3-51 the Fixed-Fixed Mockup 3.3.6.3 Results of Verification Tests for the 3-51 Fixed-Flxed Mockup 3.3.7 Effects of Sleeving on Tube-to-Tubesheet Weld 3-61 3.4 Analytical Verification 3-63 3.4.1 Introduction 3-63 3.4.2 Component Description 3-63 3.4.3 Material Properties 3-65 3.4.4 Code Crit.eria 3-65 3.4.5 Loading Conditions Evaluated 3-65 3.4.6 Methods of Analysis 3-68 3.4.6.1 Model Development 3-69 3.4.6.2 Thermal Analysis 3-71 l 3.4.6.3 Stress Analysis 3-72 l
11 3592M/060986:49-3
TABLE OF CONTENTS (Continued)
Section Title Page 3.4.7 Results of Analyses 3-74
. 3.4.7.1 Primary Stress Intensity 3-75 3.4.7.2 Range of Primary Plus Secondary Stress Intensities 3-75 3.4.7.3 Range of Total Stress Intensities 3-76 3.4.8 References 3-84 3.5 Special Considerations 3-85 3.5.1 Flow Slot Hourglassing 3-85 3.5.1.1 Effects on Burst Strength 3-85 3.5.1.2 Effects on Stress Corrosion Cracking Margin 3-85 3.5.1.3 Effect on Maximum Range of Stress Intensity and Fatigue Usage Factor 3-85 3.5.2 Tube Vibration Analysis 3-86 3.5.3 Studge Height Thermal Effects 3-86 3.5.4 Allowable Sleeve Degradation 3-86 3.5.4.1 Minimum Required Sleeve Thickness 3-86 3.5.4.2 Determination of Plugging Limits 3-87 3.5.4.3 Application of Plugging Limits 3-89 3.5.5 Effect of Tubesheet Plate Interaction 3-93 3.5.6 Structural Analysis of the Lower 3-93 Hybrid Expansion Joint 3.5.6.1 Primary Stress Intensity 3-93 3.5.6.2 Range of Primary Plus Secondary
~
Stress Intensities 3-93 111 11:59/C42 37;43 4
TABLE OF CONTENTS (Continued) i Section Title Page 3.5.6.3 Range of Total Stress Intensities 3-96 3.5.7 Evaluation of Operation with Flow Effects Oue to Sleeving 3-98 3.5.7.1 One Sleeve Per Tube 3-99 3.5.7.2 Two Sleeves Per Tube 3-100 3.5.7.3 Flow Effects Summary 3-101 3.5.8 Alternate Sleeve Materials 3-105 3.5.9 Effect of an Axial Tube Lock-up on Fatigue Usage Factor 3-105 3.5.10 Minimum Sleeve Hall Thickness 3-106 4.0 PROCESS DESCRIPTION 4-1 4.1 Tube Preparation 4-1 4.1.1 Tube End Rolling (Contingency) 4-1 4.1.2 Tube Honing 4-3 4.1.2.1 Wet Honing 4-3 4.1.2.2 Dry Honing 4-4 i
4.2 Sleeve Insertion and Expansion 4-4 4.3 Lower Joint Seal 4-6 l Iv 4445M/C42487:49-5
TABLE OF CONTENTS (Continued)
Section Title Page 4.4 Upper Hybrid Expansion Joint (HEJ) 4-7 4.5 Process Inspection Sampling Plan 4-7 4.6 Establishment of Sleeve Joint Main Fabrica- 4-9 tion Parameters 4.6.1 Lower Joint 4-9 4.6.2 Upper HEJ 4-9 5.0 SLEEVE / TOOLING POSITIONING TECHNIQUE 5-1 6.0 NDE INSPECTABILITY 6-1 6.1 Eddy Current Inspections 6-1 6.2 Summary 6-4 7.0 ALARA CONSIDERATIONS FOR SLEEVING OPERATIONS 7-1 7.1 Nozzle Cover and Camera Installation / Removal 7-2 7.2 Platform Setup / Supervision 7-2 l
7.3 Radwaste Generation 7-3 7.4 Health Physics Practices and Procedures 7-5 i
v 3592M/060986:49-6
)
TABLE OF CONTENTS (Continued) a Section 7, Page 7.5 Airborne Releases 7-6 7.6 Personnel Exposure Estimate 7-7 8.0 INSERVICE INSPECTION PLAN FOR SLEEVED TUBES 8-1 i
l 1
l l
! vi 3592M/060986:49-7
LIST OF TABLES Table Title Page 3.1-1 ASME Code and Regulatory Requirements 3-2 3.3.2-1 Summary of Corrosion Comparison Data for 3-11 Thermally Treated Inconel Alloys 600 and 690 3.3.2-2 Effect of Oxidizing Species on the SCC Suscepti- 3-12 bility of Thermally Treated I-600 and I-690 C-rings in Deaerated Caustic 3.3.3-1 Design verification Test Program - Corrosion 3-31 3.3.3-2 Residual Stresses at [ l"'C 3-32 3.3.3-3 Results of Magnesium Chloride Tests at [ 3-32 j .c.e a
3.3.3-4 Results of Magnesium Chloride Tests at [ 3-34 j .c.e a
3.3.4.3-1 Allowable Leak Rates For J. M. Farley Steam 3-38 Generators 3.3.4.4-1 Test Results for the As-rolled Lower Joints 3-41 3.3.5.3-1 Test Results for HEJ's Formed Out of Sludge 3-51 3.3.5.3-2 Test Results for HEJ's Formed Out of Sludge 3-53 (State Axial Load Leak Test, SLB and Reverse Pressure Test) l l
1 i
vit i
3592M/060986:49-8
LIST OF TABLES (Continued)
Table Title Page 3.3.5.3-3 Test Results for HEJ's Formed In Studge 3-55 (Fatigue and Reverse Pressure Tests) 3.3.5.3-4 Test Results for HEJ's Formed in Sludge (Axial 3-57 Load Leak Test and Post-SLB Test).
3.3.5.3-5 I-690 Limited Scope Test Results 3-58 3.3.6.3-1 Test Results for Full Length Sleeves 3-62 3.4.4-1 Criteria for Primary Stress Intensity Evaluation 3-66 (Sleeve) 3.4.4-2 Criteria for Primary Stress Intensity Eva,1uation 3-67 (Tube) 3.4.7.1-1 Umbrella Pressure Loads for Design. Upset, 3-77 Faulted, and Test Conditions 3.4.7.1-2 Results of Primary Stress Intensity Evaluation 3-78 Primary Membrane Stress Intensity, P, 3.4.7.1-3 Results of Primary Stress Intensity Evaluation 3-79 Primary Membrane Plus 8ending Stress Intensity, Pg+P' b 3.4.7.2-1 Pressure and Temperature Loadings for Maximum 3-80 Range of Stress Intensity and Fatigue Evaluations vill 3592M/060986:49-9
LIST OF TABLES (Continued)
Table Titie Pace 3.4.7.2-2 Results of Maximum Range of Stress Intensity 3-82 Evaluation 3.4.7.3-1 Results of Fatigue Evaluation 3-83 3.5.4-1 Regulatory Guide 1.121 Criteria 3-91 3.5.6.1-1 Results of Primary Stress Intensity Evaluation 3-94 (Lower Hybrid Expansion Joint) Primary Membrane Stress Intensity, P, 3.5.6.1-2 Results of Primary Stress Intensity Evaluation 3-95 ,
(Lower Hybrid Expansion Joint) Primary Membrane Plus Bending Stress Intensity, P( + Pb 3.5.6.2-1 Results of Maximum Range of Stress Intensity 3-97 f Evaluation (Lower Hybrid Expansion Jolat) i I
3.5.7-1 Allowable Sleeving Parameters (One Sleeve Per Tube) 3-103 1 3.5.7-2 Allowable Sleeving Parameters (Two Sleeves Per Tube) 3-104 3.5.9-1 Results of Maximum Range of Stress Intensity 3-107 !
Evaluation. Axial Tube Lockup 3.5.9-2 Results of Fatigue Evaluation. Axial Tube Lockup 3-108 in 4:15Mi0:2487:49 'O
LIST OF TABLES (Continued)
Table Title Page 4.0-1 Sleeve Process Sequence Sumary 4-2 7.5-1 Estimate of Radioactive Concentration in 7-4 Water per Tube Hones (Typical) x 4445M/042387: 43 11
LIST OF FIGURES Fiqure Title Pace 2.2-1 Sleeving Boundary ( ]C Sleeves 2-3 3.2-1 Installed Sleeve with Hybrid Expansion 3-4 Upper Joint Configuration 3.2-2 Sleeve Lower Joint Configuration 3-5 3.3.2-1 SCC Growth Rate for C-rings (150 percent YS and 3-13 TLT) in 10 percent NaOH 3.3.2-2 Light Photo micrographs illustrating IGA 3-14 3.3.2-3 SCC Depth for C-Rings (150 percent YS) in 3-15 8 percent Na 50 2 4 3.3.2-4 Reverse U-bend Tests at 360*C (680*F) 3-16 3.3.3-1 Location and Relative Magnitude of Residual 3-25
, Stresses Induced by Expansion 3.3.3-2 Schematic of HEJ Section of Sleeve 3-26 3.3.3-3 Residual Stresses Determined By Corrosion Tests 3-27 in MgC1 2
.(Stainless Steel) or Polythlonic Acid (Inconel 600) l 3.3.3-4 Results of C-Ring Tests of Type 304 Heat 3-28 No. 605947 in Boiling MgC1 2
l l
x1 4445M/C42357:43 '2
LIST OF FIGURES (Continued)
Floure Title Page 3.3.3-5 Axial Residual Stresses in Tube / Sleeve Assembly 3-29 3.3.3-6 Circumferential Residual Stresses in Tube / Sleeve 3-30 Assembly 3.3.4.1-1 Lower Joint As-rolled Test Specimen 3-36 3.3.5.1-1 Hybrid Expansion Joint (HEJ) Test Specimen 3-45 3.3.5.1-2 HEJ Specimens for the Reverse Pressure Tests 3-47 3.3.6.1-1 fixed-Flxed Mockup - HEJ 3-60 3.4.2-1 Hybrid Expansion Upper Joint / Roll Expanded Lower Joint Sleeve Configuration 3-64 3.5.4-1 Appilcation of Plugging Limits 3-90 6.1-1 Absolute Eddy Current Signals for SLV-026 at 6-6 400 kHz (Front and Rear Colls) Rear Coll Give Saturated Signal 6.1-2 ( ]*
- C d Calibration Curve 6-7 6.1-3 Eddy Current Signals from the ASTM Standard, 6-8 Machined on the Sleeve 0.0. of the Sleeve / Tube Assembly Without Expansion ( Cross Wound Coil Probe )
xil 4445M/042457: 49-;3
LIST OF FIGURES (Continued)
Fiqure Title Pace 6.1-4 Eddy Current Signals from the ASTM Standard, 6-9 Machined on the Tube 0.0. of the Sleeve / Tube Assembly Without Expansion ( Cross Wound Coll Probe )
6.1-5 Eddy Current Signals from the Expansion Transition 6-10 Region of the Sleeve / Tube Assembly (Cross Wound Coil Probe )
6.1-6 Eddy Current Calibration Curve for ASME Tube 6-11 a
Standard at C l .c.e and a Mix Using the Cross Wound Coll Probe 6.1-7 Eddy Current Signal from a 20 Percent Deep Hole, 6-12 Half the Volume of ASTM Standard, Machined on the Sleeve 0.D. In the Expansion Transition Region of the Sleeve / Tube Assembly (Cross Wound Coil Probe) 6.1-8 Eddy Current Signal from a 40 Percent ASTM 6-13 Standard, Machined on the Tube 0.0. in Expansion Transition Region of the Sleeve / Tube Assembly (Cross Wound Coll Probe) 6.1-9 Eddy Current Response of the ASTM Tube Standard 6-14 at the End of the Sleeve Using the Cross Wound Coll Probe and Multifrequency Combination xill 4445w/042437: 43-14
1.0 INTRODUCTI0m The document herein contains the necessary technical information to support the sleeving repair process as applied to the J. M. Farley Units 1 and 2 (ALA, APR) Model 51 steam generators. As a result of extensive development programs in steam generator repair, Westinghouse has successfully developed the capability to restore degraded steam generator tubes by means of a sleeve.
, Sleeving itself is a technique in which a slightly smaller diameter tube (a sleeve) is inserted into a degraded steam generator tube. The sleeve bridges and isolates the degraded section of the original tube and is joined to sound sections of the original tube at each end. As installed in the steam generators at J. M. Farley Units 1 and 2, this repair process will allow numerous tubes to remain in service therefore maintaining the life of the steam generator and the efficiency of the entire nuclear steam supply system.
To date, over 18,500 steam generator tubes at six operating nuclear power plants world-wide have been successfully sleeved, tested, and returned to service by Westinghouse. Both mechanical-joint and brazed-joint sleeves of i Inconel 600, Inconel 690, and bimetallic Inconel 625/Inconel 690 have been installed by a variety of techniques - hands-on (manual) installation, Coordinate Transport (CT) system installation, and Remotely Operated Service Arm (ROSA) robotic installation. Westinghouse sleeving programs have been successfully implemented af ter approval by licensing authorities in the U.S.
(NRC - Nuclear Regulatory Commission), Sweden (SKI - Swedish Nuclear Power Inspectorate), and Japan (MITI - Japanese Ministry of International Trade and Industry).
The sleeving technology was originally developed to sleeve 6.929 degraded tubes (including leakers) in a plant with Westinghouse Model 27 series steam generators. Process improvements and a remote sleeve delivery system (CT) were subsequently develooed and adapted to Westinghouse Model 44 series steam generators in large scale programs at two operating plants (2971 and 3000 l sleeves). This technology has also been modified to facilitate installation of sleeves in a plant with non-Westinghouse steam generators. A total of 3592M/060986:49 l-1
4.962 sleeves were installed in two successive programs of 2.036 and 2,926 sleeves utilizing CT and ROSA delivery systems, respectively. Also completed was a 17 sleeve ROSA delivery program in a Model 51 steam generator overseas and most recently a 635 sleeve manual installation program in a previously sleeved plant.
1 3592M/060986:49 1-2
2.0 SLEEVING OBJECTIVES AND SLEEVING BOUNDARIES 2.1 CBJECTIVES J. M. Farley Units 1 and 2 (ALA, APR) are Westinghouse-designed 3 loop pressurized water reactors rated at 2660 MHt. The two units utilize a t0tal of three vertical U-tube steam generators each. The steam generators are Westinghouse Model 51 Series containing heat transfer tubes with dimensions of 0.875 inch nominal 00 by 0.050 incn nominal wall thickness.
The sleeving concept and design are based on observations to date that the tube degradation due to operating environmental attack has occurred near the tubesheet areas of the tube bundle. The sleeve has been designed to span the degraded region in order to maintain these tubes in service.
The sleeving program has two primary objectives:
- 1. To sleeve tubes in the region of known or potential tube degradation.
- 2. To minimize the radiation exposure to all working personnel (ALARA) 2.2 SLEEVING SCUNDARIES Tubes to be sleeved will be selected by radial location, tooling access (due to channel head geometric constraints), and eddy current indication elevations and size. An axial e,levation tolerance of one inch will be employed to allow for any potential eddy current testing position indication inaccuracies and degradation growth. Tube location on the tubesheet face, sleeve length, tooling dimensions, and tooling access permitted by channelhead bowl geometry define the sleeving boundaries. Figure 2.2-1 shows an estimated radial sleeving boundary for a ( Ja c.e sleeve as determined by a geometric radius comcuted from the channelhead surface-to-tubesheet primary face clearance distance minus the tooling clearance distance. (The actual "as is' Ocwl geometry will be slightly different in cartain areas.) This is the sleeving boundary for a generic Westingnouse series 51 steam generator and represents the maximum sleeving potential with a.[ ]#'C sleeve.
II)y- .,;;" . 2-1
Tubes within the sleeving boundary that are degraded beyond the plugging limit but not within the axial restrictions of the [ ]C sleeve or not
- within the radial sleeving boundary will be plugged. The actual sleevable region may be modified based on tool length or other variables.
The actual tube plugging / sleeving map for each steam generator will be provided as part of the software deliverables at the conclusion of the sleeving effort.
The specific tubes to be sleeved in each steam generator will be determined based on the following parameters:
- 1. No indications beyond an elevation spanned by tne sleeve pressure boundary j which are greater than the plugging limit.
- 2. Concurrence on the eddy current analysis of the extent and location of the degradation.
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, 3.0 DESIGN 3.1 SLEEVE DESIGN DOCUMENTATION The J. A.' Farley steam generators were built to the 1968 editten Summer 1970 Addenda of Section III of the ASME Botier and Pressure Vessel Code, however,
! '1 the sleeves have been designed and analyzed to the 1983 edition of Section III of the Code through the winter 1983 addenda as well as applicable Regulatory Guides. The associated materials and processes also meet the requirements of the Code. The specifle. documentation applicable to this program are listed in Table 3.1-1.
3.2 SLEEVE DESIGN DESCRIPTION The reference design of the sleeve, as installed, is illustrated in Fiqure 3.2-1. [
ya .c.e At the uccer end, the sleeve configuratien (see Figure 3.2-1) consists of a i
section which is [
]A*C This joint configuration is known as a hybrid expansion joint (HEJ). [~
s ya .c.e In the process of sleeve length optimization and allowing for axial tolerance in locating defect by eddy current inspection, the guideline was the lower most elevation of the hard roll region to be ;;ositioned a minimum of 1 inen ,
above the eddy current detected defect area of the tube.
~
' i 's . .' 13 i ' ! )
TABLE 3.1-1 ASME CODE AND REGULATORY REQUIREMENTS Item Appilcable Criteria Requi remen t 1
Sleeve Design Section III NB-3200, Analysis '
NS-3300, Wall Thick- . k. ,
ness Operating Requirements Analysis Conditions ,
4 Reg. Guide 1.83 S/G Tubing Inspec- ,
a tibility Reg. Guide 1.121 Plugging Margin Sleeve Material Section II Haterial Composition Section III NB-2000, Identifica-tion, Tests and Examinations Code Case N-20 Mechanical Proper- l ties Sleeve Joint 10CFR100 Plant Total Primary-Secondary Leak Rate Technical Specifications Plant Leak Rate 0
35924/060986: 49 3-2 ,
At the lower end, the sleeve configuration (Figure 3.2-2) consists of a section which is C l A 'C The lower end of the sleeve has a preformed section to facilitate the seal formation and to reduce residual stresses in the sleeve.
The sleeve, after installation, extends above the top of the tubesheet and spans the degraded region of the original tube. Its length is controlled by the insertion clearance between the channel head inside surface and the primary side of the tubesheet, and the tube degradation location above the tubesheet. The remaining design parameters such as wall thickness and material are selected to enhance design margins and corrosion resistance and/or to meet ASME Boiler and Pressure Vessel Code requirements. The upper joint is located so as to provide a length of free sleeve above it. This length is added so that if in the unlikely event the existing tube were to become severed just above the upper edge of the mechanical joint within free sleeve region, the tube would be restrained by the sleeve for lateral motion and therefore axial motion, and subsequent leakage, would be limited.
Restrictied lateral mo, tion would also protect adjacent tubes from impact by the severed tube. The upper end of the sleeve is tapered in the thickness to reduce the effect of double wall in eddy current signal Interpretation.
To minimize stress concentrations and enhance inspectability in the area of the upper expanded region, (
) a,c.e,f The sleeve material, thermally treated Inconel 600 or 690, is selected to -
provide additional resistance to stress corrosion cracking. (See Section 3.3.2 for further details on the selection of thermally treated Ir.conel 600 and 690).
3592M/060986: 49 3-3
_ a. c. of i
4 Flqure 3.2-1 Installed Sleeve with Hybrid Expansion Upper Joint Configura:'en 3-4 i
_ a.c.e _'
I i
I Figure 3.2-2 Sleeve Lcwer Joint Configuration 3-5
3.3 OESIGN VERIFICATION: TEST PROGRAMS 3.3.1 DESIGN VERIFICATION TEST PROGRAM
SUMMARY
The following sections describe the material and design verification test
! programs. The purpose of these programs is to verify the ability of the sleeve concept to produce a sleeve capable of spanning a degraded region in a steam generator tube and maintain the steam generator tubing primary-to-secondary pressure boundary under normal and accident conditions. This program includes assessment of the structural integrity and corrosion resistance of sleeved tubes.
A substantial data base exists from previous test programs which verifies the sleeve design and process adequacy. Much of this testir.g is applicable to this sleeving program. The sleeve materials to be used, thermally treated nickel chromium tron alloy (Inconel 600 or 690), are identical to those used in prior sleeving programs. The standardized mechanical sleeve design is the same as that used in prior sleeving programs. The fabrication of sleeve / tube joint: by the :snbination of ( l a.c.e at both ends of the sleeve was verified and applied on all programs. r,1gorous mechanical testing programs were conducted to verify the sleeve design for various steam generator models. The Model 44, and the Model 51 steam generatcrs have the same tube dimensions. In addition to being dimensionally similar, analysis has demonstrated that the operating conditions of a Model 44 steam generator are considered to be similar to tho'se of a Model 51.
Therefora, the Model 44 testing is directly applicable to the Model 51 steam generators.
The objec tives of tne mechanical testing programs included:
Verify the leak resistance of the upper and lower sleeve to tube joints.
Verify the structural strength of the sleeved tube under normal and accident conditions.
3542W/0609A6 44 1.C
Verify the fatigue strength of the sleeved tube under transient loads representing the remaining ilfe of the reactor plant.
Confirm capability for installat'en of sleeves in tubes with conditions such as deep secondary side hard sludge and tubesheet denting.
Establish the process parameters required to achieve satisfactory installation and performance. These parameters are discussed in Section 4.6.
The acceptance criteria used to evaluate the sleeve performance are leak rates based on the plant technical specifications. Over 100 test specimens were used in the various test programs to verify the design and to establish process parameters. Testing encompassed static and cyclic pressures, temperatures, and loads. The testing also included evaluation of joints fabricated using Inconel 600 sleeves as well as Inconel 690 sleeves in Inconel 600 tubes. While the bulk of the original qualification data is centered on Inconel 600 sleeves, a series of limited scope verification tests were run using Inconel 690 sleeves to demonstrate the effectiveness of the joint formation process and design with either material. Additionally an engineering evaluation of those properties which would affect joint performance was made and disclosed no areas which would result in a change of joint performance.
The sections that follow describe those portions of the corrosion (sections 3.3.2-3.3.3) and mechanical (sections 3.3.4-3.3.6) verification programs that are relevant to this s'leeving program.
3.3.2 CORROSION AND METALLURGICAL EVALUATION The basic objective of the corrosion and metallurgical evaluation programs conducted was to verify that the sleeving concepts and procedures employed did not introduce any new mechanism that could result in premature tube or sleeve degradation. -
3592M/060986: 49 37
l Inconel Alloy 600 and Inconel Alloy 690 (I-600 and I-690) are austentic nickel-base alloys. I-600 has been extensively used, originally in the mill annealed condition, as steam generator tubing in pressurized water reactors.
In recent years, attempts to enhance the IGA-SCC (Intergranular Attack-Stress Corrosion Cracking) resistance of I-600 have focused on the application of a U.ermal treatment in the carbide precipitation temperature range (593* to 760'C). Microstructural modifications have concentrated on the grain boundary region since the SCC morphology in I-600 is predominantly intergranular. The maximum enhancement in caustic and primary water SCC performance was correlated with the presence of a semicontinuous grain-boundary carbide precipitate.
Inconel Alloy 690, which contains a higher chromium content (30 percent) than Inconel Alloy 600 has also indicated the ability to exhibit improved IGSCC resistance when thermally treated in the carbide precipitation region.
The stress corrosion cracking performance of thermally treated Inconel Alloys 600 and 690 in both off-chemistry secondary side and primary side environments has been extensively investigated. Results have continually demonstrated the additional stress corrosion cracking resistance of thermally-treated Inconel Alloys 600 and 690 compared to mill annealed Inconel Alloy 600 material.
Direct comparison of thermally treated Inconel Alloys 600 and 690 has further indicated an increased margin of SCC resistance for thermally treated Inconel Alloy 690. (Table 3.3.2-1).
The caustic SCC performance of mill annealed and thermally treated Inconel Alloys 600 and 690 were evaluated in a 10 percent NaOH solution as a function of temperature from 288'C to 343'C. Since the test data were obtained over various exposure intervals ranging frcm 2000 to 8000 hours0.0926 days <br />2.222 hours <br />0.0132 weeks <br />0.00304 months <br />, the test data were normalized in terms of average crack growth rate determined from destructive examination of the C-ring test specimens. No attempt was made to distinguish between initiation and propagation rates.
3592M/060986: 49 3-8
The crack growth rates presented in Figure 3.3.2-1 indicate that thermally treated I-600 and I-690 have enhanced caustic SCC resistance compared to that of I-600 in the mill annealed condition. The performance of thermally treated I-600 and I-690 are approximately equal at temperatures of 316*C and below.
At 332*C and 343*C, the additional SCC resistance of thermally treated Inconel Alloy 690 is observed. In all instances the SCC morphology was intergranular in nature. The superior performance of thermally treated I-690 at higher temperatures is a result of a lesser temperature dependency.
Testing in 10 percent NaOH solution at 332*C was performed to index the relative intergranular attack (IGA) resistance of I-600 and I-690. Comparison of the IGA morphology for I-600 and I-690 rings stressed to 150 precent of the 0.2 percent yield strength is presented in Figure 3.3.2-4. Mill annealed I-600 is characterized by branching intergranular SCC extending from a 200u front of uniform IGA. Thermally treated I-600 exhibited less SCC and an IGA front limited to less than a few grains deep. Thermally treated I-690 exhibited no SCC and only occasional areas of intergranular oxide penetrations, limited to less than a grain deep.
The enhancement in IGA resistance can be attributed to two factors; heat treatment and alloy composition. A characteristic of mill annealed I-600 C-rings exposed to deaerated sodium hydroxide environment is the presence of intergranular SCC along with uniform grain boundary corrosion referred to as intergranular attack (IGA). The relationship between SCC and IGA is not well established but it does appear that IGA occurs at low or intermediate stress levels and at electrochemical potentials where the general corrosion resistance of the grain boundary area is a controlling factor. Thermal treatment of I-600 provides additional grain boundary corrosion resistance along with additional SCC resistance. In the case of I-690, the composition provides an additional margin of resistance to IGA and the thermal treatment enhances the SCC resistance.
The addition of oxidizing species to deaerated sodlum hydroxide environments results in either a deleterious effect or no effect on the SCC resistance of -
thermally treated I-600 and I-690 depending on the specific oxidizing specie 3592M/060986: 49 3-9
and concentration (Table 3.3.2-2). The addition of 10 percent copper oxide to 10 percent sodlum hydroxide decreases the SCC resistance of thermally treated I-600 and I-690, and also modifies the SCC morphology with the presence of transgranular cracks in the case of I-690. The exact mechanism responsible for this change is not well understood, but it is believed to be related to an increase in the specimen potential, corresponding to a transpassive potential, which results in an alternate cracking regime. The specific oxidizing specie and the ratio of oxidizing specie to sodium hydroxide concentration appear to play an important role. By lowering the copper oxide or sodium hydroxide concentration, the apparent deleterious effect on SCC resistance is eliminated.
Mill annealed and thermally-treated I-600 and I-690 were also evaluated in a number of 8 percent sodium sulfate environments. The room temperature pH value, at the beginning of the test, was adjusted using sulfuric acid and ammonia. Test results are presented in Figure 3.3.2-3. As the pH is lowered, decreased SCC resistance for mill annealed and thermally-treated I-600 is observed, but thermally treated I-690 material did not crack even at a pH of 2, the lowest tested.
l l
The primary water SCC test data are presented in Figure 3.3.2-4. For the beginning of fuel cycle water chemistries, 10 of 10 specimens of mill annealed I-600 exhibited SCC, while 1 of 10 specimens of thermally-treated I-600 had cracked. In the end of the fuel cycle water chemistries, 7 of 10 specimens of mill annealed I-600 exhibited SCC, while 3 of 10 specimens of thermally-treated I-600 had cracked. After 13,000 hours0 days <br />0 hours <br />0 weeks <br />0 months <br /> of testing, no SCC has been observed in the mill annealed or thermally-treated I-690 specimens in either test environment.
Continuing Investigation of the SCC resistance of I-600 and I-690 in primary water environments has shown mill annealed I-600 to be susceptible to cracking at high levels of strain and/or stress. Thermal treatment of I-600 in 6 e carbide precipitation region greatly improves its SCC resistance. The performance of I-690, both mill annealed and thermally treated, demonstrates highly desirable primary water SCC resistance, presumably due to alloy ccmcositlon.
3S92M/060986:43 3-10
Table 3.3.2-1
SUMMARY
OF CORROSION COMPARISON DATA FOR THERMALLY TREATED INCONEL ALLOYS 600 AND 690
- 1. Thermally treated Inconel 600 tubing exhibits enhanced SCC and IGA resistance in both secondary-side and primary-side environments when compared to the mill annealed condition.
- 2. Thermally treated Inconel 690 tubing exhibits additional SCC esistance compared to thernal treated Inconel 600 in caustic, acid sulfate, and primary water environments.
- 3. The alloy composition of Inconel 690 along with a thermal treatment provides additional resistance to caustic induced IGA. .
4 The addition of 10 percent Cu0 to a 10 percent deaerated NaOH environment reduces the SCC resistance of both thermal treated I-600 and I-690. Lower concentrations of either Cu0 or NaOH had no effect, nor did additions of Fe340 and S102 '
- 5. Inconel Alloy 690 is less susceptible to sensitization than Inconel Alloy 600.
3592M/060986: 49 3-11 I
Table 3.3.2-2 EFFECT OF OXIDIZING SPECIES ON THE SCC SUSCEPTIBILITY OF THERMALLY TREATED I-600 AND I-690 C-RINGS IN DEAERATED CAUSTIC Temperature Exposure Environment ('C) Time (Hrs) I-600 TT I-690 TT 10 Percent NaOH + 316 4000 Increased Increased 10 Percent Cu0 Susceptibility
- Susceptibility
- 10 Percent NaOH + 332 2000 No effect No effect 1 Percent Cu0 1 Percerit NaOH + 332 4000 No effect No effect 1 Percent Cu0 10 Percent NaOH + 316 4000 No effect No effect 10 Percent Fe34 0 10 Percent NaOH + 316 4000 No effect No effect 10 Percent 510 2
- Intergranular and transgranular SCC.
l l
l i 3592M/060986: 49 3-12 m.v_., - - _ - ._..~__,.-
-- --_ - --_-- --.- . m.--- -._v- ,_,m-_. _ _ . , , . - - . - - , .-----,_-y,7
l SCC GROWTH RATE FOR C-RINGS (150% YS AND TLT)IN 10% NaOH .
j Average Crack Growth
] Rate ( gm/Hr.)
Temperature (*F)
J
' 550 675 800 830 860 l -I I I I I l
O I 800 MA .
e 1800 TT m 1690 TT 3 10-8 --
3
. _ m
" L.
I 7
~
10 '* --
l = = _
I I I I I I I 10'8 290 300 310 320 330 340 350 Temperature (*C) 0
J ,c.e 4
1 Figure 3.3.2 2. Lignt Photomicrograons illustrating IGA after 5000 Hours Enoosure of incones Alloy 600 and 690 C Rings to 10% NaOH at 332'C 1630*F).
3 14
SCC DEPTH FOR C-RINGS (150% YS)
IN 8% NA 2 SO 4 Maximum Crack D'apth, gm ,
Inches
^
i 1000 .040 l 80,000 ppm Na2SO4 i 332*C (830*F) 800 -
IN 800 MA
.032 l 5000 Hours ,,
j 600 -
C-rings,150% YS
.024 j
{ m
- y -
l = V m
j 400 - -
.018 IN 600 TT l 200 - -
.008 IN 890 TT I
i k:;: m
. m W
2 3 4 6 8 7 10 1 Room Temperature, pH i
i ,
i i
i l REVERSE U-BEND TESTS AT 360*C (680*F) l BEGINNING OF FUEL CYCLE PRIMARY WATER I Cumulative Number Crdcked % ,
10 - uA eco _
100 8 - -
^
60 2 -
n soo O ' TT see MA Sea - 0 2 END OF FUEL CYCLE PRIMARY WATER "
Cumulative Number Cracked % i A.
10 -
100 8' - [
1 l 6 -
W n .oo so 4 -
l #
2 -
O 0
1 1 1 1 1 1 1 0 2000 4000 6000 8000 10,000 12,000 14,000 Exposure Time (Hours)
3.3.3 UPPER AND LCHER JOINTS All the data presented in Section 3.3.2 relative to the corrosion and stress corrosion cracking resistance of thermally treated Inconel Alloys 600 and 690 are applicable to the sleeve.
A similar corrosion verification test program has been conducted to demonstrate that the residual stresses induced in the parent tubing by the expansion process does not degrade the integrity of the tubing. Table 3.3.3-1 identifies the various tests which have been performed and the findings. A more detailed discussion of the most significant tests follows.
The expansion processes for both the lower and upper joints involve a combination of (
lC The stresses in the sleeve, based on tube to tubesheet data, should be as shown at 8 and C on Figure 3.3.3-1, which are also judged acceptable, particularly in view of the superior corrosion resistance of the thermally treated sleeve material.
Stress levels in the. outer tube are also influenced by the exoansion technique. For an outer tube expansion produced solely by (
j a.c.e 35929/060936: 49 3-17
The specimen design is shown in Figure 3.3.3-2 and the test carameters are listed in Table 3.3.3-2. C j a.c.e i
5 i
j a.c.e No cracking was detected on the 00 surface of any specimen. These results indicate that the 00 stresses are belcw the threshold required to cause cracking in the stainless steel (less tnan to to 15 ksi).
l To summarize the results of this test:
(
i I
ja.c.e 3-18 ti::9 .::: ::
- - - - - - . .....,. - ~ - -
[
j a.c.e Confirmation that the 00 stresses on the parent tubing are very low tensile or compressive was obtained by X-ray diffraction analysis of an Inconel 600 tube expanded 30 mils and by the parting / layer removal technique, as shown below:
i X-RAY RESIDUAL STRESS MEASUREMENTS OF HEJ JOINT: 00 0F TU8E a,c.e i
I (a) in un-expanded tube above upper most transition (b) In un-espanded tube below lower most transition CONCLUSION: Restdual stresses on 00 of tube are ccmpressive and results are -
, consistent with MgCl y test findings.
3592M/060986: 49 3-19
The residual stresses in a HEJ with an Inconel 600 MA tube /Inconel 690 TT sleeve were measured using the parting / layer removal technique. The conditions of the joint were as follows:
o Nominal Tube 00 - 0.875 inch
~~
o Nominal Sleeve 00 - 0.740 inch a,c.e The results of these tests are summarized in Figures 3.3.3-5 and 3.3.3.6.
These results show an excellent correlation with the MgCl tests and the 2
results'of the x-ray measurements. The 00 surface of the tube was in compression in the axial direction at all locations along the expansion transitions. The 10 surface was in tension in the axial direction in the expansion transitions with the highest measured stress located at the hydraulic transition. In the circumferential direction, both surfaces of the tube were generally in compression although low tensile stresses, about 5 ksi or lower, measured on the tube 10 In the fully hydraulic expanded region and on the 00 in the unexpanded tube near the hydraulle expansion transition. The CD surface of the sleeve was also in compression in the axial and circumferential directions except for one measurement that was in tension (about 5 ksi) in the axial direction in the (
lC. The 10 surface of the sleeve had areas wnere the stresses were as high as about 25 kst in either the axial or circumferential direction. Residual stresses of this magnitude should not effect the service performance of the special thermally treated sleeve material.
3-20 tit:e :::!5 ::
Polythlonic Acid Tests C'
f.c.e ,
Primary Water Tests Two tests to confirm the primary water stress corrosion cracking resistance of HEJ's have been conducted. A summary of the results of these tests is as _
follows:
3532M/060336:43 3-21
680*F Primary Water Tests:
Material - a. Inconel 600 mill annealed tubing with known susceptibility to primary water stress corrosion cracking.
- b. Inconel 600 special thermally treated sleeves.
Expansion Matrix:
- Ada ,. '
1 Total Expansion, a0, inch - C lC
Test Environment:
Temperature: 680*F Pressure: Primary Side - 2850 psig Secondary Side - 1450 psig Chemistry: Primary Side - Hydrogenated Pure water Secondary Side - Pure water Results: 2000 hour0.0231 days <br />0.556 hours <br />0.00331 weeks <br />7.61e-4 months <br /> exposure with no primary to secondary leakage.
Destructive examination detected no tube wall degradation.
750*F Steam Tests:
Material - a. Inconel 600 mill annealed tubing with known susceptibility to
, primary and pure water,
- b. Inconel 600 special thermally treated sleeves.
l l
3592M/060986: 49 3-22
Expansion Matrix:
-- - A.C,e NOTE Total Expansion, C.
j ac.e Test Environment:
Temperature: 750*F Pressure: Secondary and Primary Chemistry: Hydrogenated pure water Results: 1700 hour0.0197 days <br />0.472 hours <br />0.00281 weeks <br />6.4685e-4 months <br /> exposure with no degradation of tube or sleeve defect by NOE including 10 ECT and 00 UT or by destructive examination.
In addition, both temperature and stress influence the time required to initiate primary water stress corrosion cracking (PWSCC). Calculations have been made using an equation suggested by the Brookhaven National II Laboratory forthepredictionofPWSCC.[
S j a.c.e
! 1) R. Bandy and D. van Rooyen, A Model for Predicting the Initiation and ProDagation of Stress Corrosion Cracking of Alloy 600 in High Temperature Water.
3592M/060986:49 3-23 m- + ,e* - - - ,
w-, , - - __ -- - - - -
,---,-,.---,c - -- w ---r --
o For MA I-600 in Primary Hate :
a,c.e o for Farley Unit I and 2:
a,c.e as Pressure Total Residual (Hoop) (Hoop)
Temp. Stress Stress Stress location *K ksi ksi ksi
~ '
Hard roll transition HEJ joint -
Postulation of PHSCC at the HEJ vs Hard Roll Transition:
a c.e l
o ThetimetoinitiatePHSCCattheHEJiscalculatedtobeafactorof(
j a.c.e 3592M/060986:49 3 24
i Figure 3.3.3-i Locaten and Retstrve Magnrtude of Messfusi Stresses Inouced by Expansson a.C,e
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i Figure 3.3.3 2 i
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l Results of C ring tests of Type 304 heat No. 605947 in boiling l MC1 2 l
3-28
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l Taele 3.L 3-1 ISSUE FI.NDINGS
~
- 1. CORROSION AND STRESS CORROSION '
CRACXING RES,ISTANCE OF SLEEVE
- 2. CORROSION AND STRESS CORROSION I CRACKING OF LOWER SLEEVE JOINT
- 3. CORROSION AND STRESS CORROSION CRACKING OF UPPER JOINTS
- 4. CORROSION AND STRESS CORROSION CRACKING IN ANNULUS w --
3-31
k Taele 3.3.3-2 RESIOUAL STitt$$t3 AT NtJ 5!.EEVE J0! m
~
Sc .e e
- s.
3-32
Table 1 1.1.1 RESULTS OF MAGNESIUM CHLORIDE TESTS AT ROLL EXPANCED SECTIONS
- A e
e 3-33
Tacle 3.3.3-4 RESULTS OF MAGNESIUM CHLORIDE TESTS AT HYDRAULIC EXPANDED SECTIONS a,,,;
e m
3-34
3.3.4 TEST PROGRAM FOR THE LOWER JOINT 3.3.
4.1 DESCRIPTION
OF LOWER JOINT TEST SPECIMENS The tube /tubesheet mockup was manufactured so that it was rtpresentative of the partially rolled tube to tubesheet joint (Figure 3.3.4.1-1) of the model 44/51 steam generators. The Farley steam generator tubes are full depth rolled / explosively expanded Inside the tubesheet. The formation of lower mechanical rolled joint of tube / sleeve is expected to be identical for both partially rolled tube and full depth rolled tube. The tube was then examined with a fiberscope, [ ]I'C cleaned by swabbing, and re-examined with the fiberscope. Then the preformed sleeve (Thermally Treated Inconel 600 or Thermally Treated Inconel 690) was inserted into the tube and the lower joint formed.
[
j a.c.e 3.3.
4.2 DESCRIPTION
OF VERIFICATION TESTS FOR THE LOWER JOINT The as-fabricated specimens for the Model 44/51 (as discussed in Section 3.3.1, Model 51 parameters and conditions are similar to those of Model 44 parameters and conditions) were tested in the sequence described below. Note that the tests of the Inconel 690 sleeve are similar to those performed on the Inconel 600 sleeve ex, cept that the Steam Line Break (SLB) and Extended Operation Period (EOP) tests were not considered necessary based on previous I
results.
- 1. Initial leak test: The leak rate was determined at room temperature, 3110 psi and at 600*F, 1600 psi. These tests established the leak rate of the lower joint after it has been installed in the steam generator and prior to long-term operation.
l 3592M/060986:49 3-35
- J,C e I
I f
Figure 3.3.4.1-1 Lower Joent As Mottee Test Soecimen 3-36
- 2. The specimens were fatigue leaded for 5000 cycles.
3 The specimens were temperature cycled for 25 cycles.
- 4. The specimens were leak tested at 3110 psi room temperature and at 1600 psi 600*F. This established the leak rate after 5 years of simulated normal operation (plant heatup/cooldown cycles) produced by steps 2 and 3.
Several specimens were removed from this test sequence at this point and were subjected to the EOP Test. See Step 7, below.
- 5. The specimens were leak tested while being subjected to SLB conditions.
- 6. The specimens were leak tested as in Step I to determine the post-accident leak rate.
- 7. The EGP test was performed after Step 4 for three as-rolled specimens.
3.3.4.3 'EAK
. TEST ACCEPTANCE CRITERIA Site specific or bounding analyses have been performed to determine the allowable leakage during normal operation and the limiting postulated accident condi tion. The leak rate criteria that have been established are based on Technical Specifications and Regulatory requirements. Table 3.3.4.3-1 shows the leak rate criteria for the Farley Units 1 and 2 steam generators. These criteria can be compared to the actual leak test results to provide eerification that the. mechanical sleeve exhibits no leakage under what would be considered normal operating conditions and only slight leakage under the umbrella test conditions used. It should be noted that any leakage experienced is well within the allowable limits. Leak rate measurement is based on counting the number of drops leaking during a 10-20 minute period.
Conversion to volumerric measurement is based on assuming 19.8 drops per milliliter.
3592M/060985:49 3-37
)
TABLE 3.3.4.3-1 MAXIMUM ALLCHABLE LEAK RATES FOR FARLEY STEAM GENERATORS Allowable Leak Allowable Leak Condition Rate
- Rate per Sleeve a,c.e k
Based on [ la.c.e sleeves per steam generator.
. Standard Technical Specification Limit for 1 steam generator.
i .. ( -
l
! Ja.c.e i
l The analysis assumes primary and secondary coolant initial inventories of l luCl/gm and 0.luCl/gm of Dose Equivalent I-131, respectively. In l addition, as a result of the reactor trip, an todine spike is initiated l which increases the lodine appearance rate in the primary coolant to a value equal tc 500 times the equilibrium appearance rate.
3592M/060986:49 3-38
3.3.4.4 RESULTS OF VERIFICATICN TESTS FOR LCHER JOINT 8t should be noted that in many cases reference is made to " simulated" conditions. In fact these test conditions simulate only one key aspect of operation. For example, in the case of the fatigue testing 5000 cycles were used. This number does not represent the number of cycles expected in one year, it actually represents the number of expected yearly cycles multiplied l by a suitable factor to achieve an accelerated test condition. On that basis the test results provide data which is conservative in nature and exceed the actual operating conditions. The other parameters associated with the thermal cycle test for example such as temperature ramp, hold time, temperature gradient are accelerated to achieve meaningful test results within an abbreviated time frame. Consequently the test results obtained and discussed 2hroughout the rest of this report are those of accelerated coaditions designed to test the sleeve at its endurance limit. The results do not imply that after a specific length of operating time the sleeves will begin to leak. Rather they demonstrate that under extreme accelerated test conditions leakage is small or zero providing assurance that in the actual operating case the sleeves will perform at a zero leakage base. Additionally by using that same test series for all sleeve designs it is possible to measure consistency in process modification and or small changes in the overall design to facilitate as assessment of their effect on total sleeve performance.
Reference is occasionally made to the "self healing" qualities of the mechanical joint design. This is in reference to the phenomena (observed in the test data) which shows that as the mechanical joints operate, if they exhibited leakage at the outset of the test, the rate of leakage decreases gradually with operation, to zero in most cases. This characteristic has been observed consistantly in all mechanical joint testing.
Another consistant characteristic observed in the testing of mechanical joints is that the leakage, when observed, is generally higher at room temcerature conditions and, as in the case of the self-healing phenomena, decreases as the temperature is elevated. This characteristic has lead to the almost exclusive _
3592M/060986: 49 3-39
use of the room temperature hydrostatic test in the process, tooling, personnsl, procedure and demonstration phases associated with a plant specific sleeving operation. While not a specific part of this report the process verification data exists for reference should the need arise.
The test results for the Model 44/51 lower joint specimens are presented in Table 3.3.4.4-1. The specimens did not leak before or during fatigue loading. After five years of simulated normal operation due to C
]"*C All of the three as-rolled specimens were leak-tight during the Extended Operating Period (EOP) test.
For the Inconel 690 sleeve tests the following were noted:
Specimens MS-2 (Inconel 690 Sleeve): Initial leak rates at all pressures and at normal operating pressure following thermal cycling were [
, j a,b,c.e 1
i 3592M/060986:49 3 40 i -__ . . . _ . .
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Specimen MS-3 (Inconel 690 Sleeve): C l
\
j a,b,c.e Specimen MS-7 (Inconel 690 Sleese): C j a,b,c.e 3.3.5 T'EST PROGRAM FOR THE UPPER HYBRID EXPANSION JOINT (HEJ)
The discussion contained in Section 3.3.4.4 is relevant to testing in general and applies in the following tests conducted on upper joints as well.
3.3.
5.1 DESCRIPTION
OF THE UPPER HEJ TEST SPECIMENS Two types of HEJ test specimens were fabricated for the Model 44 testing (as discussed in Section 3.3.1, Model 51 parameters and conditions are bounded by Model 44 parameters and conditions). The first type was a short specimen as shown in Figure 3.3.5.1-1. Some of these specimens were fitted with pots containing hard sludge to simulate the structural effects of sludge on the joint. The only type of sludge simulated in this program was hard sludge. Soft sludge effects were bounded by the hard sludge effects and by the out-of-sludge conditions. (
- Ja ,0,c Any leakage was collected and measured as it issued from the annulus between the tube and sleeve. This type of specimen was .' sed in the majority of the tests.
l The second type of test specimen was a modification of the first type. It was utilized in the reverse pressure tests, i.e., for LOCA and secondary side l
l 3592M/060986:49 3 44
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Hybrid Escension Joint Test Scocimen (The Leak Path, if Any Leakage Exists, is Shown DY tMe Cetted Lines) -
i i
! Figure 3.3.5.1-1 3-45 l
1,
~ , . , , . , --,. - -.- . a, - - , - - . , . , .. .-~---n,..-,n,-.,-,_.n-_,-----,n,n--nn....., , . , - -- _ - - - , _ _ ,,.c , , , , , , , - - , ,
hydrostatic pressure tests. As shown in Figure 3.3.5.1-2, the specimen was modified by [
l a,b c The possible reverse pressure test leak path is shown in Figure 3.3.5.1-2.
Only specimens like Figure 3.3.5.1-1 (excluding the sludge conditions) were used in the I690 HEJ specimen fabrication as the effects of sludge had been established in the earlier Model 44 tests.
3.3.5.2. DESCRIPTION OF VERIFICATION TESTS FOR THE UPPER HEJ The verification test program for the HEJ was similar to that for the lower joint.
The HEJ was subjected to fatigue loading cycles and temperature cycles to simulate five years of normal operation and the leak rate was determined before and after this simulated normal operation. For a number of the specimens, the leak rate was also determined as a function of static axial loads which were bounded by the fatigue load. It is important to note that the fatigue load used in testing was that which was caused by loading /
unloading. Hence. It was judged necessary to determine that the leak rate at static and fatigue conditions were comparable. The upper HEJ specimens were also subjected to the loadings / deflections caused by a steam line break (SLB) accident and the leak rate was determined during and after this simulated accident. The upper HEJ was also leak tested while being subjected to two reverse pressure conditions, a LOCA and a condition which simulated a i secondary hydrostatic test. An extended operation period test was also performed.
3.3.5.3 RESULTS OF VERIFICATION TESTS FOR THE UPPER HEJ The test results are presented in Tables 3.3.5.3-1 to 3.3.5.3-5.
3592M/060986:49 3-46
-- L . --
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i l MEJ Soecimens for tfie Reverse Pressure Tests. The Lean Pn-l if Any Exists, is Shown by tine Cetted Lines t
i i
Figure 3.3.5.1-2 l 3-47 i
l l
l f
1
-ev-- -m ,- rww---,--e-- , g,-w .,w,,ee.,w.,-,,e.-e-we.,,nw_ ,-,,ma___ _ . , , _ _ - ~ ~ . , - . , - - - - - - -
As can be seen from Table 3.3.5.3-1, the HEJ's formed out-of-sludge, i.e., in D 'C
air, had an average initial leak rate of approximately [ 3 at ;he normal operating condition of 600*F and 1600 psi. After five years of simr.'ited normal operation due to 5000 fatigue cycles and 29 to 32 temperature cycles, the leak rate was C ]b,c.e at the normal operating condition. Furthermore, for the E0P test, i.e., after thirty-five years of simulated normal operation due to at least 175 temperature cycles (208 were actually used) and a total of 35000 fatigue cycles, the leak rate was
( ),b,c.e Table 3.3.5.3-2 contains data for upper HEJ's formed out-of-sludge. It includes the same basic test data as Table 3.3.5.3-1, i.e., initial leak rate data. However, it includes static axial load leak tests SLB and reverse pressure tests in place of the fatigue and E0P tests included in Table 3.3.5.3-1. Five of the six specimens were leaktight at normal operating conditions during the initial leak test. The leak rate'during static axial sleeve loads, bounded by the fatigue load and caused by normal operatina conditions was measured for four out-of-sludge HEJs. C
]b.c.e This leak rate
- was a neglig hle fraction of the per-sleeve limit of [
l b,c.e The results for the post-SLB leak test, at the same temperature and pressure conditions, were similar to the during-SLB results.
( 3b ,c.e l This leak rate was a small fraction of the post-SLB, per-sleeve limit of l [ j b,c.e The results for the out-of-sludge HEJ reverse pressure test are shown in Table 3.3.5.3-2. For both the simulated LOCA and secondary side hydrostatic pressure test the leak rate was zero for the two specimens tested.
The process used for forming HEJ's in sludge, in Tables 3.3.5.3-3 and a,c.e 3.3.5.3-4, was the reference process, per Table 4.0-1 C ]
3592M/060986:49 3 43
~
^i - _ --
[
.)a.c.e The initial leak rate of the first group of upper HEJs formed in sludge was [
]b,c.e at the normal operating condition as is shown in Table 3.3.5.3-3. Only one specimen had a (
l b.c.e After exposure of the specimens to five years of simulated normal operation due to fatigue and gemperature cycling, the average leak rate remained very low, [
]b,c.e at the 6T*F and 1600 ps! condition.
The results of the reverse pressure test for the in-sludge upper HEJs are also shown in Table 3.3.5.3-3. [
l a,b.c It was also zero for the sireulated secondary side hydrostatic pressure test.
Table 3.3.5.3-4 also contains data for HEJs formed in-sludge. It includes the same basic initial leak tests as Table 3.3.5.3-3. However, it includes axial load leak test and post-SLB leak tests in place of the fatigue and reverse pressure tests included in Table 3.3.5.1-2. All of the four specimens were leaktight during the initial leak test, per Table 3.3.5.3-4 Two specimens did not leak at any static axial load and two others did not leak until a compressive load of 2950 lbs was reached. However, the two leak rates at 2950 lbs were low, ( l b,c.e for specimens Number PTSP-23 and PTSP-33, respectively.
En general, the leak rates for static loads were approximately the same as for dynamic (fatigue) loads of the same magnitude. However, a specific set of specimens was not subjected to both types of loads. -
l' 3592M/060986:49 3 49
As shown in Table 3.3.5.3-4, the average leak rate for four in-sludge specimens durina the SLB test wat [
j a.c.e The test data generated for the Inconel 690 samples is presented in Table 3.3.5.3-5. The following observations were noted:
Specimen S-5 (Inconel 690): [ l"'D'C were found at initial leak tasting at room temperature (R.T.). At 600*F, the leak rates reduced significantly and remained below [ 3a ,b.c during a subsequent thermal cyc1tng test. This specimen was formed with a tube diametral bulge that was smaller than will probably be used in the field.
Specimens 5-8 (Inconel 690); 255-8 and 255-13 (Inconel 625), B-4, 8-6, and B-7 (Inconel 625/690 - 0.740 in. Sleeve Ola.), and BA-11 (Inconel 625/690-0.630 in. Sleeve Dia.): These seven specimens all exhibited moderate to small or very small leaks, mostly during the initial leak testing at R.
T. In all cases, by the end of the testing, including thermal cycling and fatigue in some cases, the leak rates had reduced to zero (or near zero),
illustrating the self healing characteristic of rolled joints.
3.3.6 TEST PROGRAM FOR THE FIXE 0-FIXE 0 MOCXUP 3.3.
6.1 DESCRIPTION
OF THE FIXE 0-FIXE 0 MOCXUP The fixed-fixed full scale mockup is shown in Figure 3.3.6.1-1. This mockup simulated the section of the steam generator from the primary face of the tubesheet to the first support plate. The bottom plate of the mockup represented the bottom of the tubesheet, the middle plate simulated the top of the tubesheet and the upper plate simulated the first support plate. The tubes were roll expanded into the bottom plate to simulate the tube /tubesheet joint and into the upper plate to simulate a dented tube condition. The term
" fixed-fixed" was derived from the fact that the tubes were fixed at these two 3592M/060986:49 3-50
Tatste 3.3.5.3-1. I I[5I RESULIS FOR MODEL 44 HEJ'S FOR'8E0 OUI 0F SLUDGE (Page I of 2)
(FAIIGUE AND [XMINIED OPERATISH MSIS INCL.) a .he Specimen No.
PIA-46 PIA-47 Y PIA-48 5
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Awcrages:
th. e - Short specimens useil ior thls test. l e
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TABLE 3.3.5.3-3 (cont)
TEST RESULTS FOR H00EL 44 HEJ'S FORMED IN SLUDGE (FATIGUE AND REVERSE PRESSURE TESTS INCL.) (CONT)
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Iccations. There were thirty-two tubes in two clusters of sixteen. A sludge simulant composed of alumina was formed around one cluster of sixteen.
Sleeves thirty inches long were installed in the tubes by (
j.** Each tube was perforated between the upper and lower joints to simulate tube degradation and thereby provide a primary-to-secondary leak path. End plugs were welded to the tubes to permit pressurization with water.
No fixed fixed mockup tests were performed on the I690 samples based on the results of the earlier tests performed.
3.3.
6.2 DESCRIPTION
OF VERIFICATION TESTS FOR THE FIXED-FIXED MOCKUP The fixed-fixed mockup was used first to verify the full length sleeve installation parameters and tooling. It was then used to measure the leak rate of the lower joint and upper HEJ. This leak rate was determined with the sleeve installed in a tube fixed at the tubesheet and dented at the first support plate, i.e., for the fixed-fixed condition.
3.3.6.3 RESULTS OF VERIFICATION TESTS FOR THE FIXED-FIXED MOCKUP Table 3.3.6.3-1 contains leak test results recorded for full length sleeves formed and tested in-situ, in the fixed-fixed mockup, in-sludge and out-of-sludge (see Figure 3.3.5.3-3). All of the room temperature initial leak tests produced C ya ,b,c These initial leak rate results were similar to the initial leak rate results in which the short specimens were structurally unconstrained during forming of the upper HEJ. Therefore, it was concluded that the results of the other several tests performed only on short specimens would be similar if the test _
had been performed in-situ, in the fixed / fixed mockup. During the pre-test evaluation, it was determined that the fixed / fixed mockup dupilcated the most 3592M/060986:49 3-59
--, , . - - , , - - - - - - - , - - , - - ., -.-.,-r-- ,--
- w ::
- a,c.e Fixed-Fixed Mockuo - HEJ (For tne HEJ In-Situ Leek Te:ts.
the Leak Path,if Any Exists. is Shown cy tne Dettec Lines) rigure 3.2.E.1-1 3-60
stringent structural loading conditions for sleeves. Therefore, it was concluded that all of the testing with short specimens was valid. Because the model 44 loads envelope the model 51 loads, this testing is considered appilcable to model 51 units and consequently validates the results for both units.
3.3.7 EFFECTS OF SLEEVING ON TUBE-TO-TUBESHEET WELO The effect of hard rolling the sleeve over the tube-to-tubesheet weld was examined in the sleeving of 0.750 inch 00 tubes. Although the sleeve.
Installation roll torque used at in 0.750 inch tubes is less than a .875 inch 00 tube, the radial forces transmitted to the weld would be comparable.
Evaluation of the 0.750 inch tubes showed no tearing or other degrading effects on the weld after hard rolling. Therefore, no significant effect on the tube-to-tubesheet weld is expected for the larger 0.875 inch 00 tube configuration.
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3.4 ANALYTICAL VERIFICATION 3.
4.1 INTRODUCTION
This section contains the structural evaluation of the sleeve and tube assembly with HEJ, sleeve material I690 II and sleeve length ( ]A'C in relation to the requirements of the ASME Boller and Pressure Vessel Code,Section III Subsection NB, 1983 Edition (Reference 1]
The analyses include primary stress intensity evaluations, maximum range of stress intensity evaluations, and fatigue evaluations for various mechanical and thermal conditions which umbrella the loading conditions specified by the Westinghouse Equipment Specification G-677164, 12/18/69, Revision 1 (Reference 5], Westinghouse Equipment Specification Addendum No. 677439, 2/16/81, Revision 1 (Reference 6].
3.4.2 COMPONENT DESCRIPTION The general configuration of the sleeve-tube assembly with HEJ is presented in Figure 3.4.2-1.
The critical portions of the sleeve-tube assembly are two joints, the upper and lower Hybrid Expansion Joints (HEJ), and straight sections of the sleeve and tube between the two joints. The finite element model developed contains both upper and lower joints. In Section 3.4, a detailed stress evaluation was performed for the upper joint. Structural analysl: of the lower joint is presented in Section 3.5. The tolerances used in developing the models were such that the maximum sleeve and tube outside diameters were evaluated in combination with the minimum sleeve wall thickness. This allowed maximum stress levels to De developed in the roll transition regions.
- ec
- 1) Sleeve Material I600 is considered in Section 3.5.
3592M/060986:49 3-63
i Figure 3.4.2-1 l
Hyerro Excansion Uncer Joint Poll Excanced Lower Joint Steeve Configuration a.c.e.f
~
4 i
e 6
3 64 1
3.4.3 MATERIAL PROPERTIES I
The sleeve material is Inconel 690 ' described in ASK: Code Case N-20 (Reference 7]. The tube material is58-163 (Inconel 600).
An air gap was included between.the tube and sleeve below the HEJ as well as between the tube and the tubesheet. Although this space may be filled with secondary fluid, assuming the physical properties of air for these elements is conservative for the thermal analysis. Primary fluid physical properties were used for the gap medium above the HEJ.
All material properties used in the analyses were as specified in the ASME Boller and Pressure Vessel Code,Section III, Appendix 1 (Reference 4] and Code Cases (Reference 7].
3.4.4 CODE CRITERIA The ASME Code Stress Criteria which must be satisfied are given in Tables 3.4.4-1 and 3.4.4-2.
3.4.5 LOADING CONDITIONS EVALUATED The loading conditions are specified below:
- 1. Design conditions
- a. Primary side design conditions P - 2485 psig T - 650*F
- b. Secondary side design conditions P - 1085 psig T - 600*F
- c. Maximum primary to secondary pressure differential - 1600 psig, _
T - 650'F
- 1) Sleeve material Inconel 600 is considered in Section 3.5.
3592M/060986:49 3-65
TABLE 3.4.4 1 CRITERIAFORPRIMARYSTRESSINTENSITYEVAk.UATION nuno m 6
e l
I M
3-66 r - _ _ _ _ _ . , _ _ - . . . . , _ _ , _ . - - _ . _ . _ _ . - , . _ _ _ _ _ .
- TABLE 3.4.4-2 CRITERf A FOR PtiMARY Stat!s INTENSITY EVAt,UATICN (TuSE) a,c,
. I 3-67 W
- d. Maximum secondary to primary pressure differential - 670 psig .
T = 650*F
- 2. Full load steady state conditions are:
Primary side pressure = 2235 psig Hot leg temperature = 610.9'F Cold leg temperatura - 543.3*F Secondary side pressure - 778.0 psig Feedwater temperature - 437.3*F Steam temperature = 518.3*F Zero load reactor coolant temperature = 547.0'F Other operating conditions are specified in Tables 3.4.7.1-1 and 3.4.7.2-1.
3.4.6 METH005 0F ANALYSIS Structural analysis of the sleeve-tube assembly includes finite element model development, thermal, pressure stress and thermal stress calculations, primary membrane and primary membrane plus bending stress intensity evaluation, primary plus secondary stress intensity range evaluation, and fatigue evaluation for various mechanical and thermal conditions which umbrella the loading conditions specified by the appropriate Design and Equipment Specifications. Two basic computer programs, WECAN and WECEVAL, are used in structural analyses of the sleeved tubes.
The WECAN program (Reference 21 performs thermal and stress analyses of the structure. Pressure stress is calculated separately for a 1000 pst primary and a 1000 psi secondary pressure. The results of these " unit pressure" runs are then scaled to the actual primary side and secondary side pressures corresponding to the load conditions considered in order to determine the total pressure stress distribution.
Thermal analysts provides the temperature distribution needed for thermal stress calculations. Thermal stress calculations are performed for fixed 3592M/060986:49 3-68
times under thermal transients. These times for the total pressure and thermal analysis are chosen for the anticipated maximum and millmum total l stresses in critical regions of the structure.
Total stress distribution is determined by combining the pressure and thermal stress results.
Total stress calculations as well as stress evaluations are carried out by the WECEVAL computer program (Reference 3].
WECEVAL 15 a multi-purpose code which performs ASME Code,Section III, Subsection N8 stress evaluations.
At'any given point or section of the model, the program WECEVAL is used to datermine the total stress distribution per the Subsection NB requirements.
That is, the total stress at a given cross-section through the thickness, so-called analysis section, ASN, is categorized into membrane, linear bending, and non-linear components which are compared to Subsection NB allowables. In addition, complete transient histories at given locations on the model are used to calculate the total cumulative fatigue usage factor per Code Paragraph N8-3216.2.
3.4.6.1 MODEL OEVELOPMENT A finite element model was developed for evaluating the sleeve design. Some significant considerations in developing the model are:
- 1. The model has been divided in two parts: upper model and lower model. Structural integrity of the whole model was provided by all direction coupling of the nodes along the upper model and lower model interface.
- 2. Mechanical roll fixities between the sleeve and tube at the hard roll regions were achieved by coupling the interface nodes in the _
radial direction. For conservatism, locations of contact in the 3592M/060986:49 3-69
sleeve-tube interfaces along the upper hard roll region contain elements which share nodes. This approximates a rigid fix by the rolling process involved. Additional axial coupling was effected l
also for the lower sleeve-tube and tube-tubesheet interface nodes.
- 3. The interface nodes along the upper and lower hydraulic expansion regions of the HEJ were coupled in the radial direction for temperature and thermal stress runs. In the cases when pressure may penetrate into the interface, the interface nodes along these areas were disconnected for pressure stress runs.
i
- 4. 8y varying the boundary conditions at a specified region of the model, conditions of either intact tube or discontinuous tube were simulated.
The element types chosen for the finite element analysis were the following HECAN (Reference 2] elements:
I
_. a.c.e l
i All the element types are quadratic, having a node placed I the center of each surface in addition to nodes at each corner.
i 3592M/060986:49 3-70
_---__---,____----..-.-,__,-w,- - . . . . _ , . _ ,,. .- . --,m__, , , , . - , - - -- - - .._, , , --,---,.n., -
3.4.6.2 THERMAL ANALYSIS 4
The purpose of the thermal analysis is to provide the temperature distribution needed for thermal stress evaluation.
Thermal transient analyses were performed for the following events:
Small step load increase Small step load decrease Large step load decrease Hot standby operations Loss of load Loss of power Loss of secondary flow Reactor trip from full power i
The plant heatup/cooldown, plant loading / unloading and steady fluctuation events were considered under thermal steady state conditions.
The finite element types chosen for the thermal analysis were STIF58 and STIF68.
8n order to perform the WECAN thermal analysis, boundary conditions consisting
- of fluid temperatures and heat transfer coefficients (or film coefficients) for the corresponding element surfaces are necessary. The conditions
- considered in the thermal analysis are based on the following assumptions
- The temperature induced stresses are most pronounced for sleeves in the hot leg (where the temperature difference between the primary and secondary fluids is a maximum) and therefore, only the hot leg sleeves were considered. This condition bounds the thermal stresses on the cold leg. .
3592M/060986:49 3-71
l
- The sleeves may be installed in any tube in the generator. Thus, to be conservative, it is assumed that the sleeve to be evaluated is sufficiently : lose to the periphery of the bundle that it experiences the water temperature exiting the downcomer.
Special hydraulle and thermal analysis was performed to define the primary and i
secondary side fluid temperatures and film coefficients as a function of time. Both bolling and convective heat transfer correlations were taken into consideration.
l 3.4.6.3 STRESS ANALYSIS A WECAN'(Reference 21 finite element model was used to determine the stress levels in the tube / sleeve configuration.
Elements simulating the medium between the tube and the sleeve were considered as dummy elements. The element types employed were STIF53 and STIF56.
Based on the results demonstrating the applicability of a linear elastic analysis, thermally induced and pressure induced stresses were calculated separately and then combined to determine the total stress distribution using the dECEVAL computer program (Reference 3]. In addition, WECEVAL performs the stress categorization required for an ASME Boller and Pressure Vessel Code,Section III, stress analysis and for the complete fatigue evaluation.
1 The criteria in the evaluation were those specified in Subsection NB of the ASME Boller and Pressure Vessel Code (Reference 1].
Pressure Stress Analysis For superposition purposes, the WECAN model was used to determine stress distributions induced separately by a 1000 psi primary pressure and a 1000 psi secondary pressure. The results of these " unit pressure" runs were then scaled to the actual primary side and secondary side pressures corresponding 3592M/060986:49 3-72
to the loading condition considered in order to determine the total pressure stress distribution.
The two modeling considerations in determining the unit pressure load stress distributions were tube intact and tube discontinuous. Therefore, the following unit pressure loading conditions were evaluated to determine the maximum anticipated stress levels induced by primary and secondary pressures:
Primary pressure - tube intact Primary pressure - tube discontinuous Secondary pressure - tube intact Secondary pressure - tube discontinuous The end cap forces due to the axial pressure stress induced in the tube away from discontinuities were taken into consideration.
Thermal Stress Analysis The WECAN model was used to determine the thermal stress levels in the tube / sleeve configuration that were induced by the temperature distribution calculated by the thermal analysis. Thermal stress were determined for each steady state solution as well as for the thermal transient solutions at those times during the thermal transient which were anticipated to be limiting from a stress standpoint.
Combined Pressure Plus Thermal Stress Evaluation As mentioned previously, total stress distributions were determined by combining the unit pressure and thermal stress results as follows:
P ort .
' total
- 1000 unit primary pressure 3592M/060986:49 3-73
'sec . (
- 1000 unit secondary pressure I thermal This procedure was performed with the program WECEVAL (Reference 3]
Stress and Fatique Evaluation Stress and fatigue evaluation were completed using the program HECEVAL (Reference 31. The program WECEVAL performed primary stress intensity evaluation, primary plus secondary stress intensity range evaluation, and fatigue evaluation of the sleeved tube assembly.
At any given point or section of the model, the program WECEVAL determined the total stress distribution for a loading condition considered and categorized that total distributton per the Subsection N8 requirements. That is, the total stress for a given cross section through the thickness is categorized into membrane, linear bending, and non-linear components.
These categorized stresses were then compared to the Subsection N8 allowables.
In addition, complete transient histories at given locations on the model were used to caleviate the total cumulative fatigue usage factor per Code Paragraph N8-3216.2. For the fatigue evaluation, the effect of local discontinuities was considered. -
3.4.7 RESULTS OF ANALYSES Analyses were performed for both intact and discontinuous tubes. Cesign and operating transient parameters (pressure, temperature, etc.) were selected from the applicable Westinghouse Design Specifications for the Model 44 and 51 Series steam generators in such a manner as to be conservative in structural 3592M/060986:49 3-74
effect and frequency of occurrence. Fatigue and stress analyses of the sleeved tube assembly have been completed in accordance with the requirements of the ASME Soller and Pressure Vessel Code,Section III.
3.4.7.1 PRIMARY STRESS INTENSITY The umbrella loads for the primary stress intensity evaluation are given in Table 3.4.7.1-1.
The results Of primary stress intensity evaluation for the analysis sections are summarized in Tables 3.4.7.1-2 and 3.4.7.1-3.
All primary stress intensities for the sleeved tube assembly are well within allowable ASME Code limits.
The largest value of the ratio " Calculated Stress Intensity / Allowable Stress Intensity" of (
j a.o.c 3.4.7.2 RANGE OF PRIMARY PLUS SECONDARY STRESS INTENSITIES Table 3.4.7.2-1 contains the pressure and temperature loads for maximum range of stress intensity evaluations as well as for fatigue evaluation.
The maximum range of " stress intensity values for the sleeved tube assemblies are summarized in Table 3.4.7.2-2.
The requirements of the ASME Code Paragraph N8-3222.2, were met at all locations.
t 3592M/060986:49 3-75
3.4.7.3 RANGE OF TOTAL STRESS INTENSITIES Based on the sleeve design criteria, the fatigue anaiysis considered a design life objective of 40 years for the sleeved tube assemblies. Table 3.4.7.2-1, describes the transient conditions used in the fatique analysis.
Because of possible opening of the interface between the sleeve and the tube along the hydraulle expansion regions, the maximum fatigue strength reduction factor of 5.0 (N8-3222.4(3)) was applied in the radial direction at the " root" Interface nodes of the hard roll region.
The results of the fatigue analysis for the sleeved tube assemblies are
~
summarlzed in Table 3.4.7.3-1.
All of the cumulative usage factors are below the allowable value of 1.0 specified in the ASME Code.
I 3592M/060986:49 3-76
i TABLE 3.4.7. I-1. .
UM8RELLA PRESSURE LOA 05 FOR DESIGN UPSET. FAULTED, AND TEST CONDITIONS h
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~
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I 8
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3592M/060986:49 3,77
,m. m,----- ,_.-...y-,- - . . . . - . , - . . . _ - - , - -,y- -m-- - , , - - - - -, . . , , - . - - - - - , - y--,-- e 9, ,--_7-,_,m.-c.e--- ,-_.-.-n. -..e-
N e
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. I e t-e M e W
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>=
N E 3 -J W 0-W = me-s== W W e-o
. I "w" g g3m m = C g >= w g
- 9 w J M >=
- M Y b g -d 3 M g
, w - = m
'i .J M == g S b 4
M a
w h
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5 ls =
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I-3-78
TABLE 3.4.7.1-3 RESULTS OF PRIMARY STRESS INTENSITY EVALUATION (Upper Hybrid Expansion Joint)
PRIMARY MEMBRANE PLUS BENDING STRESS INTENSITY. Pg,Pb CALCULATED MAXIMUM ALLONABLE OF STRESS STRESS RATIO INTENSITY. INTENSITY. CALCULATED S.I.
LOCATION KS! KS! ALLONABLE S.I.
_ 4.C.e Y
- s l
l l .
l l
TABLE 3.4.7.2-1 PRESSURE AND TEMPERATURE LOADlWGS FOR MAXIMUM RANGE OF STRESS INTENSITY AND FATIGUE EVALUATIONS CASE PRESSURE, PSIG Time,sec/ Thermal CON 01110N NAME NO. CYCLES PRIMARY SECONDARY Conditlons Ambient Ambient i 200 0 0 NA/No Thermal Stress Plant Loading
- Plant Heatup IPLLD 2 18300 2235 1005 0/UF Plant Cooldown 2PLLD 3 18300 2235 705 3200/ST y Plant Unloading Small Step Load Decrease ISSLD 4 2000 2310 795 30/TR 2SSLD 5 2000 2160 760 150/TR Small Step Load Increase ISSLI 6 2000 2215 610 50/ f!!
2SSLI 7 2000 2230 660 ife . IR Large Step Load Decrease ILSLD 8 200 2335 100 36/TR 2LSLD 7 200 2160 830 480/TR Hot Standby Operations IHSIB 10 18300 2235 655 0/ST 2HSIB 11 18300 2235 925 400/ST
TABLE 3.4.7.2-1 (cont)
PRESSUREANDTEMPERATURkLOADINGSFORMAXIMUMRANGE Of STRESS INTENSITY AND FATIGUE EVALUATIONS CASE PRESSURE, PSIG Time sec/ Thermal CONDITION NAME NO. CYCLES PRIMARY SECONDARY Conditions Turbine Roll Test ITRT 12 10 2235 1035 O/No Thermal Stress 2TRT 13 10 1875 525 1680/No Thermal Stress Loss of Load ILLD 14 100 2585 1020 12/TR 2LLD 15 100 1600 1020 ,100/TR Loss of Power ILPH 16 50 2060 1065 125/TR 2LPH 17 50 2485 1065 2000/TR
, Loss of flow ILfH 18 100 1860 875 140/TR Reactor Trip from IRTR 19 500 1855 935 100/ST full Power
~
Steady State ISFL 20 106 2335 725 NA/ST fluctuations 2SFL 21 106 2135 690 NA/ST Tube Leak Test ITLT 22 800 0 840 NA/No Thermal Stress Primary Side Leak Test IPSLT 23 200 2485 885 NA/No Thermal Stress Secondary Side Leak Test ISSLT 24 80 415 1085 NA/Nd Thermal Stress
~
i
- Umbrella transient i Note: Thermal conditions: TR = transient, ST = steady state, Uf = Uniform temperature i
i l ll l
'O t.
C.
a I
. I _ -
S .
S D
L TA B AL A RU N C O N L L Et A L CA f
A U
L A
V E
Y T
I
)
t S
N i n
E T o I N J S I E 2 n LMf
- S o BUOI 2 S i AM S E s WI EK 7 R n OXG T a LAN 4 S p LMA x A R 3 f E E
O d i
L E B G br A N y T A H R
Mre U
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S UI EK T CXG L LAN U AMA S C R E
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1
- TABLE 3.4.7.3-1 RESULTS OF FATIGUE EVALUATION (Upper Hybrid Expansion Joint) i
! CUMULATIVE USAGE ALLONABLE USAGE
! LOCATION , FACTOR FACTOR TUBE INTACT
] - _ 4,C.e Sleeve 1.0 Tube 1.0
'i y i $
TUBE DISCONTINUOUS .
]- Sleeve 1.0 i
Tube 1.0 t
I l
I
]
3.
4.8 REFERENCES
- 1. ASME Boller and Pressure Vessel Code,Section III, Subsection NB, 1983 Edition, July 1, 1983.
- 2. WECAN, WAPPP and FIGURES II, F. J. Bogden Editor, Second Edition, May 1981, Westinghouse Advanced System Technology, Pittsburgh, PA 15235.
- 3. J. M. Hall, A. L. Thurman, "WECEVAL, A Computer Code to Perform ASME BPVC Evaluations Using Finite Element Model Generated Stress States,"
Westinghouse, April, 1985.
- 4. ASME Boller and Pressure Vessel Code,Section III, Appendix 1, 1983 Edition, July 1, 1983.
- 5. Equipment Specification G-677164, Westinghouse, July 10, 1969, Revision 1, December 18, 1969.
- 6. Equipment Specification Addendum No. 677439. Westinghouse, February 29, 1972, Revision 1 February 16, 1981.
- 7. ASME Boller and Pressure Vessel Code, Code Cases, Case N-20, 1983 Edition, July 1, 1983.
4 3592M/060986: 49 3-84
3.5 SPECIAL CONSIDERATIONS 3.5.1 FLOH SLOT HOURGLASSING Along the tube-lane, the tube support plate has several long rectangular flow slots that have the potential to deform into an " hourglass" shape with significant denting. The effect of flow-slot hourglassing is to move the neighboring tubes laterally inward to the tube lane from their initial
- positions. The maximum bending would occur on the innermost row of tubes in the center of the flow slots.
3.5.1.1 EFFECT OF BURST STRENGTH The effect of bending stresses on the burst strength of tubing has been studied. Both the axial and circumferential crack configurations were investigated. (
j a,e f 3.5.1.2 EFFECT ON STRESS CORROSION CRACKING (SCC) MARGIN Based on the results of a caustic corrosion test program on mill-annealed tubing. the bending stress magnitude due to flow-slot hourglassing is judged to have only a small effect, if any, on the SCC resistance margins. Two long term modular model boiler tests have been conducted to address the effect of bending stresses on SCC. No SCC or IGA was detected by destructive examination. It is to be noted that thermally treated Inconel 600 and Inconel 690 have additional SCC resistance compared to mill annealed Inconel 600 tubing. -
3.5.1.3 EFFECT ON MAXIMUM RANGE OF STRESS INTENSITY AND FATIGUE USAGE FACTCR In addition to the above two considerations, one should also consider the effect of the hourglassing induced bending stresses on maximum range of stress intensity and fatigue usage factor of the sleeve. Taking into account the hourglassing induced bending stress along with the transient pressure and -
2425M / ~.4223 7 : 23 3-35
thermal stress, the largest value of maximum stress intensity would be 59.70 KSI (allowable 79.80 KSI), fatigue usage factor is negligible.
3.5.2 TU8E VIBRATION ANALYSIS Analytical assessments have been performed to predict nodal natural frequencies and related dynamic bending stresses attributed to flow-induced vibration for sleeved tubes. The purpose of the assessment was to evaluate the effect on the natural frequencies, amplitude of vibration, and bending stress due to installation of various lengths of sleeves.
Since the level of stress is significantly below the endurance limit for the tube material and higher natural frequencies result from the use of a sleeve / tube versus an unsleeved-tube, the sleeving modification does not contribute to cyclic fatigue.
3.5.3 SLUDGE HEIGHT THERMAL EFFECTS l In general, with at least 2.0 inches of sludge, the tubesheet is isothermal at the bulk temperature of the primary fluid. The not effect of the sludge is to reduce tube /tubesheet thermal effects.
3.5.4 ALLOWABLE SLEEVE DEGRADATION l
3.5.4.1 MINIMUM REQUIRED SLEEVE THICKNESS The minimum required sleeve wall thickness, rt , to sustain normal and accident condition loads is calculated in accordance with the guidelines of
! Regulatory Guide 1.121, as outilned in Table 3.5.4-1. In this evaluation, the surrounding tube is assumed to be comoletely degraded; that is, no design credit is taken for the residual strength of the tube.
The sleeve material may be either thermally treated Inconel 600 or thermally treated Inconel 690. It has been shown that the properties of Inconel 600 are very similar to those of Inconel 690. In particular, the yield strength and ultimate strength are very similar.
I l
i 4445.w/042 37: 43 3-35
Since Regulatory Guide 1.121 constitutes an operating criterton, it is permissible to derive the allowable stress ilmits based on expected lower bound material properties, as opposed to the Code minimum values. Expected strength properties were obtained from statistical' analyses of tensile test data of actual production tubing. These data were used for the lower tolerance limits of material. Lower tolerance limit, LTL, means there is 95 percent of confidence that 95 percent of the sleeve / tubes will have strength grea*er than LTL.
3.5.4.2 OETERMINATION OF PLUGGING LIMITS The minimum acceptable wall thickness and other recommended practices in I Regulatory Guide 1.121 are used to determine a plugging limit for the sleeve.
This Regulatory Guide was written to provide guidance for the determination of a plugging limit for steam generator tubes undergoing localized tube wall thinning and can be conservatively applied to sleeves. Tubes with sleeves which are determined to have indication of degradation of the sleeve in excess of the plugging limit would have to be repaired or removed from service.
l As recommended in paragraph C.2.b. of the Reg. Guide, an additional thickness degradation allowance must be added to the minimum acceptable tube wall thickness to establish the operational tube thickness acceptance for continued service. Paragraph C.3.f. of the Reg. Guide specifles that the basis used in setting the operational degradation allowance include the method and data used in predicting the continuing degradation and consideration of eddy current measurement errors and other significant eddy current testing parameters.
As outlined in Section 6.0 of this report, the capability of eddy current inspection of the sleeve and tube in the sleeve area has been demonstrated.
The [ ]C eddy current measurement uncertainty value of (
lC of the tube wall thickness is appropriate for use in the 6etermination of the operational tube thickness acceptable for continued service and thus determination of the plugging limit.
6445M/042387: 49-106
P 1
i Paragraph C.3.f of the Reg. Guide specified that the basis used in setting the operational degradation analysis include th'e method and data used in predicting the continuing degradation. To develop a value for continuing degradation sleeve experience must be reviewed. No degradation has been detected to date on Westinghouse designed sleeves and no sleeved tube has been ,
removed from service due to degradation of any portion of the sleeve. This result would be expected due in part to the changes in the sleeve material relative to the tube and the lower heat flux due to the double wall in the sleeved region. At the direction of Alabama Power, for the application of sleeving in the Farley Units I and 2 steam generators a conservative allowance of C lC of tube wall for continued operational degradation will be used
! In determining the plugging limit for the tubes. Although past expertence has not identified any sleeve degradation, Alabama Power considers a combined j i lC allowance, in addition to the minimum required sleeve wall th'ickness, to be adequate compensation for eddy current measurement uncertaintles and continued sleeve degradation.
, It is the position of Westinghouse Electric that since no degradation has been 1
! detected in the sleeves, presently any allowance for continuing degradation
(' Jwouldbeanarbitraryvaluenotsupportedbythedata C I l and would represent a conversatism in addition to the safety factors impilcit '
in the determination of minimum acceptable tube wall thickness using Reg.
- Guide 1.121 recommendations.
In summary, the operational tube thickness acceptable for continued service
! Includes the minimum acceptable tube wall thickness (C l a,b c of ,,jj l thickness, see Table 3 5.4-1), the combined allowance for eddy current l uncertainty and operational degradation ([ lC of wall thickness). These
! terms total to 69% resulting in a plugging limit as determined by Regulatory ;
l Guide 1.121 recommendations of 31% of the tube wall thickness. The combined i C lC for operational degradation and eddy current uncertainty as recommended by Westingnouse would result in a plugging limit of C l #*D'C The plugging limtt for the tube, where applicable as defined below is as specified in the Technical Specifications for the non-sleeved portions of the
! tube, currently 40% of the tube wall thickness.
i 2245WO42787:49- 3-88 i
._ . . - . - _ _ _ . _ _- - _ _ _ _ _ _ _ . _ _ _ _ _ , _ _ . _ _ ._,,,, _ _ _ ,. _ _ _ , _ i
3.5.4.3 APPLICATION OF PLUGGING LIMITS Sleeves or tubes which have eddy current indications of degradation in excess of the plugging limits must be repaired or plugged. Those portions of the tube and the sleeve (shown in Figure 3.5.4-1) for which indications of wall degradation must be evaluated are summarized as follows:
- 1) Indications of degradation in the entire length of the sleeve must be evaluated against the sleeve plugging limit.
- 2) Indication of tube degradation of any type including a complete guillotine break in the tube between the bottom of the upper joint and the top of the lower roll expansion does not require that the tube be removed from service.
- 3) The tube plugging limit continues to apply to the portion of the tube in the upper joint and in the lower roll expansion. As noted above the sleeve plugging limit applies to these areas also, i l
i
- 4) The tube plugging limit continues to apply to that portion of the tube !
above the top of the upper joint.
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j Fleure 3.5.4-1 Appucaten of Plugems Lnts a The plugging limit for the tube is 40% of tube well, the
- pluggmg limit for the sleeve is 31% with a conservative i allowance for continued degradation, see Section 3.5.4.2. ,
3-90 1
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Table 3.5.4-1 REGULATORY GUIDE 1.121 CRITERIA
- 1. Normal Operation Determine rt . alnlaus required sleeve wall thickness.
- 4. Yield Celterton: Sj i 39.59 ksi Loading: P, = 2250 psia .
P, . 720 psia AP . 1530 pst AP . R Hence, t *E I I"Ch r *5 y - 0.5 Pp+P) s which is ( 14.C.' percent of the nominal well thickness.
- b. Ultimate Criterton: S, 1 90.54 ksi Loading: P, . 2250 psia P, . 720 psia AP . 1530 pst aP . R'
- Hence, tr"S *I I I"Ch
[-0.5(P,+P,)
which is ( ) a.c.e percent of the nominal wall thickness.
- 2. Accident Condition Lo4 dines
- a. LOCA SSE The major contribution of LOCA and SSE loads is the bending stresses at the too tube support plate due to a combination of the support motion. Inertial loadings and the pressure differential across the tube U-bend resulting from tne rarefraction' wave during LOCA. Since j the sleeve is located below the first support, the LOCA . SSE ,f l bending stresses in the sleeve are quite small. The governing event ,'
for the sleeve therefore is a postulated secondary side blowdown.
359N/060986: 49 3-91 I
i Table 3.5.4 1 (cont.)
- b. FL5 + SSE The maulmum primary-to-secondary pressure differential occurs during a postulated feedline break (FL8) accident. Again, because of the sleeve location, the SSE bending stresses are small. Thus, the governing stresses for the minimum wall thickness requirement are the pressure momerane stresses.
Criterton: P, i smaller of 0.75u or 2.45 ,l.a. 63.4 ksi Loadings: P, = 2650 psig P, s 0 AP - 2650 AP . R N'"C I r
0.7 S u - 0. I 5 ( P, + P 3 ) i or, ( )
C percent of nominal wall.
- 3. t.eak-8efore-8reak Verification The leak-before-break evaluation for the sleeve is based on leak rate and burst pressure test data obtained on 7/8 inch 00 x 0.050 inch wall and 11/16 inch 00 x 0.040 inch wall cracked tubing with various amounts of uniform thinning simulated by machining on the tube 00. The margins to burst during a postulated SLS (Steamline Break Accident) condition are a function of the mean radius to thickness ratio, based on a maximum permissible leak rate of 0.35 gpm due to a normal operating pressure differential of 1530 psi.
Using a mean radius to thickness farter of 9.5 for the nominal sleeve, the current Technical Specifications ai'owable leak rate of 0.35 gom. a SLB pressure differential of 2560 psi, and the nominal leak and ncminal burst curves, a 25 percent margin e:Ists between the burst crack lengen and the leak crack length. For a sleeve thinned 54 percent through wall over a 1.0 inch antal length, a 17 percent margin to burst is demonstrated. Thus the leak before break benavlor is confirmed for untninned and thinned conditions.
- ea u,.$e 34..n 3-92
3.5.5 EFFECT OF TUBESHEET/ INTERACTION Since the pressure is normally higher on the primary side of the tubesheet than on the secondary side, the tubesheet becomes concave upward. Under these conditions, the tubes protruding from the top of the tubesheet will rotate from the vertical. This rotation develops stresses in the sleeved tube assembly. Analysts performed showed that these stresses are not large enough to affect significantly the fatigue usage factors already found.
1.5.6 STRUCTURAL ANALYSIS OF THE LONER HYORIO EXPANSION J0!NT 3.5.6.1 Primary Stress Intensity The results of primary stress intensity evaluation for the analysis sections lccated at the lower hybrid expansion joint are summartzed in Tables 3.5.6.1-1 and 3.5.6.1-2.
All primary stress intensttles for the sleeved tube assembly at the lower hybrid encansion joint meet the ASME code limits.
3.5.6.2 Range of Primary Plus Secondary Stress Intensteles 8cimary plus secondary stress at the Lower Hybrid Escansion Joint are develoced ey the pressure acting on tne sleeve, tube and tubesheet ligament surfaces (artmary stress), and by thermal stress and deformations Imposed Dy the tutesheet motion (secondary stress).
The tutesneet motion result from the crimary and secondary side Dressure and Interactions among the tube $neet, succort ring, channel nead, and the stub
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l The worst case: tube intact was analyzed. The maximum range of stress intensity salues for the sleeved tube assembly are summarized in Table 3.5.6.2-1.
The requirements of the ASME Code, paragraph Ns.3222.2 were satisfied.
3.5.6.3 Range of Total Stress Intensities The fatigue analysis considered a design life objective of 40 years for the sleeved tube assemelles. The mastnum fatigue strength reduction factor of 5.0 was applied in the radial direction at the " root" Interface nodes of the hard roll region.
All of the cumulative usage factors are negingible, hence, they are below the allowable value of 1.0 specified in the ASME Code.'
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3.5.7 EVALUATION OF CPERATICN WITH FLOW EFFECTS DUE TO SLEEVING An ECCS performance analysts has been completed for Farley Units 1 and 2 This safety analysis was performed using the 1981 Evaluation Model with BART and assumed up to 10 percent steam generator tube plugging (SGTP) i the steam generators of either plant and estabilshed a ( l
value of total core pe4 king factor for Units I and 2. This study and tG corresponding non-LOCA study are considered applicable for the steam generator sleeving program with regard to the Westinghouse supplied fuel. The accidents evaluated include LOCA and non-LOCA transients as well as consideration of the effects on the nuclear cesign and thermal-hydraulle performance with the esisting plant reactor vessel internals. For the accidents considered in that study, the core and system parameters remained within their proper limits (i.e., peak clad temperature. ON8R, RCS pressure, etc.).
For the steam generators in either of the Farley Units,10 percent of tc total tubes (3388 tubes per S/G) equals 338.8 tubes plugged in any one steam generator. The ECCS analysis model is such that a uniform steam generator tube plugging condition is modeled. The NRC staff has recutr(
that the LOCA analysis for a plant with steam generator tube plugging model the ma:Imum tube plugging level present in any of the plant steam generators.
Inserting a sleeve into a steam generator tace results in a reduction :(
primary coo,lant flow. It is conservatively assumed that up to
( l tubes per steam generator will be sleeved. The selected slesving program at Farley assumes the ( l inch long sleeves whlCh are presumed to De long enough to span the degraded areas in tne tutesheet region and to be 4 Dove the sludge pile in either the not or cold leg side of the steam generators.
3 98
3.5.7.1 ONE SLEEVE PER TUBE For a single ( l i d inch sleeved tube, the primary coolant flow reduction per tube is approsimately equal to ( }a.c.e Dercent of normal flow under normal conditions. This reduction in primary coolant flow equates to a hydraulic equivalency ratto of ( l sleeved tubes to one plugged tube under normal conditions.
Using this ( l to I ratto and the assumed 10 percent tube plugging Ilcit for both units, Table 3.5.7-1 can be used to determine the mailmum number of sleeves possible up to ( l' d sleeves per steam generator under normal conditions. Note that C l sleeved tubes are equivalent to ( la.c.e plugged tubes, or ( l' d percent plugging.
For the condition presented above for farley Unit 1, the most Ilmiting equivalent plugged tube condition in the three steam generators occurs in steam generator 8 where 103 tubes are currently plugged. It is seen in Table 3.5.7.1 that with ( 1****' tubes sleeved there would be a margin of
( l tubes (334.8 minus ( . )' d ) available for additional plugging before escoeding the basis of the LOCA analysis with 10 percent SGTP. With a smaller numeer of sleeves, the margin of tubes available for additional plugging would be larger.
For Farley Unit 2, the most limiting equivalent plugged tube conditten in the taree steam generators occurs in steam generator C which currently nas 138 tuDes plugged. It It seen in Table 3.5.7.1 that with ( l tutes sleeved in SG C there would be a margin of ( l tubes availaD'e for addit onal plugging before esteedtng the basis of the LOCA analysts with 10 percent SGTP. With a smaller number of sleeves, tne margin of tubes avallaele for additional plugging (or any equivalent comolnation of sleeves and plugs) would be larger, i .
3 99 v .,.nn ,
3.5.1.2 TWO SLEEVES PER TUBE when a single tube nas one I l inch sleeve on tne hot-leg side and a second C l inch sleeve on the cold-leg side, tne primary coolant flow loss per tube is appro Imately equal to ( l percent of normal flow, this reduction in primary coolant flow equates to a hydraulle equivalency ratto of ( l sleeves to one plugged tube under normal conditions.
Using this ( l d to I ratto and the assumed 10 percent tube plugging limit for both units. Table 3.5.7-2 can to used to determine the ma:Imum numeer of sleeves possible up to ( lC sleeves for C l tubes per steam generator under normal conditions. Note that (
I l sleeves are equivalent to ( l plugs, or ( J
'C
percent plugging under normal conditions.
For the condition presented above for Farley Unit 1. the most limiting ecutvalent plugged tube condition in the three steam generators occurs in steam generator 8 where 103 tuces are currently plugged. It is seen in fable 3.5.7.2 that with ( .., l tubes sleeved, there would be a margin of
( l d tubes (338.8 minus ( lC) available for additional slugging before escoeding the basis of the LOCA analysis with 10 percent SafP. ditn a smaller numeer of sleeves, the margin of tubes avallacle for a:dttional plugging would be larger.
For Farley Unit 2. t,he most limiting equivalent plugged tube condition in t9e t9ree steam generators occurs in steam generator C wnich currently nas 138 ts00s cluqqed. It is seen in fable 3.5.7.2 that wt tn (' 1 4.c.e tubes sleeved in SG C there would to a margin of (' l tubes avallacle for additional clugging esfore esteeding tne easts of tne LOCA analysts with 10
- ercent SGTP. nitn a smaller numcor of sleeves, tne margin of tuees avatIsole for additional clugging (or anj eculvalent comotnation of sleeves and plugs)
.culd te larger.
J 100
....., ....gg ..
i fable 3.5.7-2 snows the allowable sleeving / plugging mis under the present steam generator condittens with no additional plugging margin before escoeding tneeastsoftNLOCAanalystswith10percentSGTP.
3.5.7.3 FLOW EFFECTS
SUMMARY
1 The effect of sleeving on the non-LOCA transient analyses has been reviewed.
Analyses of the level of sleeving and plugging discussed in this retart have shown that the Reactor Coolant System flow rate will not ce less tnan the Thermal Design Flow rate. The Thermal Design Flow rate is the value'uted in '
the non-LOCA safety analyses and is designed to be less than the alnlaum RCS flow rate that occurs under normal or degraded conditions. Since the reduced RCS flow rate is not less than the assumed flow rate (Thermal Design Flow).
tne non-LOCA safety analyses are bounded by the anticipated manimum amount of steam generator tube sleeving (( l sleeves per steam generator) and i
cause no safety concerns. Any smaller numeer of sleeves would have less of an effect.
It snould be noted that any comelnation of sleeving and plugging may be I
utilized at Farley Units I and 2 as long as tne effective SGTP of to percent t1 not onceeded. Tables 3.5.7-1 and 3.5.7 2 give tne mastmum numcer of tuees enten ma/ es sleeved up to ( l tuees and the numcer of additional ciugs oor steam ;enerator that could be installed at the present plugging levels of Farley Urlts I and 2 without onceeding tne 10 percent $GTP.
ln addition, as a result of tube plugging and sleeving, primary side fluid selocities In the steam generator tuces will increase. The effect of this -
velocity increase on tne sleeve and tuce nas been evaluated assuming a conservative limiting condition in which to percent of the tuees are plugged.
As a eference, normal flow velocity enrough a tube Is approntmately
( l ft/sec. for the unplugged condition. With 10 percent of tne P
i l
3-101
tubes clugged, the fluid velocity tnrougn an unplugged tuce Is ( JC
ft/sec, and for a tune with a sleeve, the local flutd velocity in the sleeve region is estimated at ( lC ft/sec. Because these fluid velocities are less than tne inception velocities for flutd impacting, cavitation, and eroston.corroston, the potential for tube degradation due to these mechanisms is lee.
Accordingly, using the assumotions stated in paragraph I of Section 3.5.7, no ECCS results more adverse than those in the entsting 14stinghouse fuel safety analysis are anticipated for equivalent tube plugging projected to occur at the Farley Untts wtth up to ( l a.c.e tubes sleeved per steam generator using ( lE sleeves.
e 3 102
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TA8LE 3.5.7 1 AL,LONASLE SLity!NG PARAMETERS UNDER NORMAL CON 0!TIONS (0NE SLitVE PER TU8E)
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TA8LE 3.5.7-2 ALLONASLE SLitV!NG PARAMETERS UNDER NCAMAL CON 0!TIONS (TWO SLitVES PER TU8t) ai O
M 3 104 1 ;)1* :::!!i t '.
3.5.8 ALTERNATE SLEEVE MATERIALS As mentioned aoove In Section 3.4.3. mechanical properties of Inconel 690 are considered by the ASME Code Case N-20. The design stress intensity value 5, for both materials 1690 and 1600 in the Case N-20 is identical (Sg = 26.6 kst). Therefore, for these materials, Primary Stress Intensity and Mastmum Range of Stress Intensity allowables are stellar. Only pressure stresses were considered when calculating Primary Stress Intensity. The maximum Primary Membrane Stress Intensity was found at the analysis section on the straight portion of the tube. Hence, the ratto " Calculated Maximum .
Primary Membrane Stress Intensity / Allowable Stress Intensity" does not depend upon the sleeve material (Inconel 690 or Inconel 600)., Thermal stress, and hence mastmum range of stress Intensity and fatigue usage factor, could depend upon the mechanical properties of the sleeve material. The modulus of elasticity of Inconel 600 is higher than that of Inconel 690 by -7 percent.
However, the coefflClent of thermal espansion of Inconel 690 is higher than that of Inconel 600. Past analytical experience indicates that the e-mismaten between 1690 and 1600 as sleeve material effects significantly the enormal stress. Therefore, the results of the Mastmum Range of Stress Intensity and Fatigue Evaluations for the sleeve and tube assemely with 1690 as the sleeve j material are conservative relative to those which would be calculated with I600 as a sleeve material.
3.5.9 EFFECT OF AN AX!AL TU8E LOCX-UP ON FATIGUE USAGE FACTOR In this analysis, only orie tube is considered to be locked-up at the first tuce succort plate under 100 percent power conditions.
The following effects on the stress Components of the locked-up tube were analyzed:
- effect of thermal conditions in the tube and wrapper /shell regions
- effect of pressure drop across the tubesheet
~
- effect of pressure drops across the tuDe support plates
- effect of Interaction among the tubesneet, tute succort plates, snell/wraccer, stayrods, and scacer rods.
Itsi e s.,oaa aa 3-105
The effects of pressure drops across the tubesheet and the tube support plates as well as the tubesheet-tube support plate assembly interactions were taken
! Into account for central locked-up tubes while they were neglected for the outermost tubes. The results of mantmum range of stress intensity and fatigue evaluations are given in Tables 3.5.9-1 and 3.5.9-2 For the central locked-up tubes, only the sleeve for the worst case, t.e.,
tube discontinuous, was considered.
It is seen that the requirements of the ASME Code are satisfied for both outermost and central antal locked-up sleeved tubes.
3.5.10 Minimum Sleeva Wall Thtekness Nominal and minimum sleeve wall thickness was analyzed.
T& King into account plus 0.003 inches for corroston/erroston, the recommended sleeve wall thickness is:
Nominal Sleeve Wall Thickness 0.037 inches i
Minimum Local Sleeve Wall Thickness 0.0363 inches 3-106
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TABLE 3.5.9-2 j RESULTS Of FATIGUE EVALUATION AXIAL IUSE LOCK-UP l
l I CUMULATIVE USAGE ALLOMASLE USAGE l 1OCAiION -
FACTOR FACTOR Outermost Tubes l
i TUBE INIACI Sleeve Negligible 1.0 i
iube Negingible 1.0 Y
1 o TUBE DISCONTINUOUS l
, Sleeve Negingible 1.0
, Tube Negligible 1.0 J
I Central Tubes TUBE DISCONilNUQUS .
Sleeve Negilglble 1.0 l
j
1 i
1 1
4.0 PROCESS DESCRIPTION The sleeve installation consists of a series of steps starting with tube end preparation (if required) and progressing through sleeve insertion, hydraulic expansion at both the lower joint and upper Hybrid Expansion Joint (HEJ) regions, hard roll joining at both joint locations, and joint inspection. The sleeving sequence and process are outlined in Table 4.0-1. All these steps are described in the following sections.
4.1 TUBE PREPARATION There are two steps involved in preparing the steam generator tubes for the sle'eving operatten. These consist of light rolling (as required) at the tube end and tube hering.
4.1.1 TUBE END RGLLING (CONTINGENCY) 8f gaging or tube inside diameter measurements indicate a need for tube end rolling to provide a uniform tube opening for sleeve Insertion, a light mechanical rolling operation will be performed. This is sufficient to prepare the mouth of the tube for sleeve insertion without adversely affecting the original tube hard roll or the tube-to-tubesheet weld. Tube end rolling will be performed only as a contingency.
Testing of similar lower joint configurations in Model 27 steam generator sleeving programs at 1 much higher torque showed no effect on the tube-to-tubesheet weld. Because the radial forces transmitted to the tube-to-tubesheet weld would be much lower for a larger Model 51 sleeve than for the above test configuration no effect on the weld as a result of the light roll is expected.
3593M/060986:49 41
TABLE 4.3-1 SLEEVE PROCESS SEQUENCE
SUMMARY
a,c.e e
3593M/060986: 49 4-2
4.1.2 TUBE HONING The sleeving process includes honing the inside diameter area of tubes to be sleeved to prepare the tube surface for the hybrid expansion joint and the lower joint by removing loose oxide and foreign material. Honing also reduces the radiation shine frcm the tube inside diameter, thus contributing to reducing man-rem exposure.
Tube honing may be accomplished by either wet or dry methods. Both processes have been shown to provide tube inside diameter surfaces compatible with mechanical joint installation. The s' election of the honing process used is dependent primarily on the installation technique utilized, the scale of the sleeving operation (small scale vs. large scale sleeving), and the customers site specific rad-waste requirements. Evaluation has demonstrated that neither of these processe's remove any significant fraction of the tube wall base material.
4.1.2.1 WET HONING Tube honing will be performed using a [
. j a.c.e A waste handling system may be used to collect the [ ],d' the hone debris, and the oxide removed from the tube ID. [
]A'C There may also be an inlet to the , _
suction pump which subsequently pumps the debris and water directly to the plant waste disposal system.
3593M/060986:49 4-3
4.1.2.2 ORY HONING The dry hone process is similar to the wet hone process with the notable exception that the water jet and the attendant systems needed to handle the effluent are omitted. The dry hone process is typically more applicable to hands-on (manual) or small scale sleeving operations.
In order to remove loose oxide debris produced by the dry honing operation, the tube interior is swabbed utt11 zing a fluid (typically delonized water or isopropyl alcohol) soaked felt pad to an elevation slightly less than the honed length, but above the top of the installed Sleeve.
4.2 SLEEVE INSERTION AND EXPANSION The following paragraphs describe the insertion of the sleeves and mandrels and the hydraulic expansion of the sleeves at both the lower joint and upper HEJ locations.
The sleeves are fabricated under controlled conditions, serialized, machined, cleaned, and inspected. They are typically placed in plastic bags, and packaged in protective styrofoam trays inside wood boxes. Upon receipt at the site, the boxed sleeves are stored in a controlled area outside containment and as required moved to a low radiation, controlled region inside containment. Here the sealed sleeve box is opened and the sleeve removed, inspected and placed in a protective sleeve carrying case for transport to the steam generator platform. , Note that the sleeve packaging specification is extremely stringent and, if unopened, the sleeve package is suitable for long term storage.
The bladder style hydraulle expansion mandrel is connected to the high l pressure fluid source, e.g., Lightweight Expansion Unit (LEU), via high pressure flexible stainless tubing. The mandrel / sleeve assembly is then positioned either manually or by use of the delivery system at the designated tube location.
3593M/060986:49 44
l l
i
]C This process is repeated until all sleeves are installed and hydraulically expanded.
3593M/060986:49 4-5
4.3 LOWER JOINT SEAL d
At the primary face of the tubesheet, the sleeve is joined to the tube by a mechanical hard roll (following the hydraulic expansion) performed with a powered roll expander which extends approximately 2 inches into the tube. The control of the mechanical expansion is maintained through a torque setting.
The tool automatically shuts off when it reaches a preset torque value.
The contact forces between the sleeve and tube due to the initial hydraulle expansion are sufficient to keep the sleeve from rotating during the [
j,a.c.e The appropriate extent of hard roll expansion of the sleeve is attained by
(- ]"' C # The hard roller torque is calibrated on a standard torque calibrator prior to initial hard rolling operations and subsequently recalibrated at the beginning of each shift for automatic tooling. This control and calibration process is a proven technique used throughout industry in the installation of tubes in heat exchangers.
3593M/060986: 49 4-6
4.4 UPPER HYBRIO EXPANSION JOINT (HEJ)
The HEJ first utilizes a ('
lC An upper hard roller is inserted into the sleeve until it is positioned at the prescribed axial location. The hard roller is then operated for a fixed time. At the end of this time the roller will have expanded to its set diameter and the total tube diametral expansion will have been accomplished. The maximum torque of the hydraulic or air operated drive motor is set at a value which is sufficient to achieve the desired tube expansion.
4.5 PROCESS INSPECTION SAMPLING PLAN In order to verify the final sleeve installation, an eddy current inspection will be performed on all sleeved tubes to verify that all sleeves received the required hydraulle and roll expansions. The basic process check on 100 percent of tne sleeved tubes will be:
- 1. Verify presence of lower hydraulle expansion zone.
- 2. Measure lower hydraulic expansion and roll average diameter and verify location within the lower hydraulle expansion.
- 3. Verify presence of upper hydraulic expansion zone.
4 Measure upper hydraulic expansion and roll average diameter and verify location within the upper hydraulic expansion.
- 5. Check for the presence of any anomalies.
In order to monitor the sleeving process after each operation frcm each lot, eddy current data on a percentage of the sleeves may be performed to obtain sleeve ID data. As confidence is gained that the sleeving process is proceeding as anticipated, the lot sizes may be increased and percentages reduced. These average diameters will be evaluated versus the expected tolerancis established through the design requirements, laboratory testing results Iand previous experience. This evaluation will determine whether or
.not the e?Jipment/ tooling it performing satisfactorily. If process data is
! determinej to be ;utside of expected ranges, a non-conformance report is issued and further analysis performed.
3593M/060986: 49 4-7
If recuired, Olatest may be used in lieu of eddy current to perform sleeve installation acceptance and in-process monitoring evaluations. Undersized diameters will be corrected by an additional expansion step to produce the desired degree of espansion. Oversized diameters will be dispositioned by a specific evaluation process on an individual tube basis, to determine their acceptability with respect to specified sleeving parameters.
If it is necessary to remove a sleeved tube from service as judged by an evaluation of a specific sleeve / tube configuration, tooling and processes will be available to plug the sleeve or the lower portion of the sleeve will be removed and the tube will be plugged.
As mentioned previously, the basic process dimensional verification will be completed and evaluated for 100 percent of all installed sleeves.
35939/060986: 49 4-8
m 4.6 ESTABLISHMENT OF SLEEVE JOINT MAIN FABRICATION PARAMETERS .
4.6.1 LOWER JOINT The main parameter for fabrication of acceptable lower joints is sleeve [
].C Sleeve [ l # 'C is determined by [ ,
].a.c.e Accordingly, rolling torque was varied to _
achieve the desired sleeve [ lC in the original Model 44 program (also applicable to the model 51). ['
a l .c.e was achieved was used throughout the program verification testing.
t,-
4.6.2 UPPER HEJ The main parameter for fabrication of HEJ's (in-sludge and out-of-sludge) which met the leak rate acceptance criteria was [
l a.c.e (Refer to Section 3.3.5.3 for an additional discussion of the roll expansion torque for the in-sludge case.)
In the first sleeving project performed by Westinchouse. hydraulic ornancinn asial length was also evaluated. [
1.a.c.e Therefore, in later programs, the HEJ hydraulic expansion axial length [ _
,1, :: , ; . a j
3593M/060986: 49 4-3
L j a,b.c.e 1
o e
3593M/060986: 49 4-10
5.0 SLEEVE / TOOLING POSITIONING TECHNIQUE With all positioning techniques, the process actually used to install the -
sleeves (hydraulic expansion, mechanical rolling, etc.) will not be changed due to the use of any sleeve / tooling positioning technique. It is the processes which the sleeves are subjected to that are critical to their successful installation; the technique used to position the sleeves and ',
c tooling is not critical so long as it does not affect the sleeve installation processes.
Some techniques used to position the sleeve installation tooling are: fully
~~ '
robotic (ROSA), alternate robotic or semi-remotely-operated equipmenc, (SN-10W) hands-on (manual), or the combination of two or more tooling L_ _
installation mode utilized is dependent upon many variables and mutually decided between the utility and Westinghouse.
3593M/060986.49 5-1
i 6.0 NDE INSPECTABILITY The Non Oestructive Examination (NDE) development effort has concentrated on two aspects of the sleeve system. First, a method of confirming that the joints meet critical process dimensions is required. Secondly, it must be shown that the tube / sleeve assembly is capable of being evaluated through subsequent routine in-service inspection. In both of these efforts, the inspection process has relled upon eddy current technology.
Previous sleeve installations have had baseline and subsequent in-service inspections of the sleeved tubes. Presently, no change has been observed in any of the in-service eddy current inspections compared to the baseline inspections.
6.1 E00Y CURRENT INSPECTIONS The eddy current inspection equipment, techniques, and results presented herein apply to the proposed Westinghouse sleeving process. Eddy current inspections are routinely carried out on the steam generators in accordance with the plant's Technical Specifications. The purpose of these inspections is to detect at an early state tube degradation that may have occurred during plant operation so that corrective action can be taken to minimize further degradation and reduce the potential for significant primary-to-secondary leakage.
The standard inspection procedure involves the use of a bobbin eddy current probe, with two circumferentially wound coils which are displaced axially along the probe body. The coils are connected in the so-called differential mode; that is, the system resacnds only when there is a difference in tne properties of the material surrounding the two coils. The coils are excited by using an eddy current instrument that displays changes in the material surrounding the coils by measuring the electrical impedance of the coils.
Presently, this involves simultaneous excitation of the coils with several different test frequencies. -
6-1 20:7u/:2:337.:;.23
The outputs of the various frequencies are ccmoined and recorded. The combined data yield an output in which signals resulting from conditions that do not affect the integrity of the tube are reduced. By reducing unwanted signals, improved inspectability of the tubing results (i.e., a higher signal-to-noise ratto). Regions in the steam generator such as the tube sucports, the tubesheet, and sleeve transition zones are examples of areas where mult1 frequency processing has proven valuable in providing improved l
Inspectability.
After sleeve installation, all sleeved tubes are subjected to an eddy current inspection which includes a verification of correct sleeve installation for process control and a degradation inspection for baseline purposes to which all subsequent inspections will be compared.
While there are a number of probe configurations that lend themselves to Improving the inspection of the tube / sleeve assembly in the regions of I configuration transitions, the crosswound coil probe has been selected as i
offering a significant improvement over the conventional bobbin coil probe,
, yet retalning the simplicity of the inspection procedure.
Verification of proper sleeve installation is of critical importance in the sleeving process. The process control eddy current verification is conducted utilizing one frequency in the absolute mode with a crosswound coil probe.
The purpose is to provide "in-process" verification of the existence of proper hydraulic expansion and hard roll configurations and also to allow determination of the sleeve process dimensions both axially and radially.
Figure 6.1-1 tilustrates the coil response and measurement technique for typical sleeve / tube joint.
The inspection for degradation of the tube / sleeve assembly has typically been performed using crosswound coil prcces operated with multifrequency excitation. For the straight lengtn regions of the tube / sleeve assemoly, the g inspection of the sleeve and tube is consistant witn normal tubing
{ inspections. In tube / sleeve assembly joint regions, data evaluation becomes l more complex. The results discussec celow suggest the limits on the volume of degradation tnat can de detected in :ne vicinity of geometry cnanges.
I 11:m::u8n a;-n 6'2
.I The detection and quantification of degradation at the transition regions of ,f the sleeve / tube assembly depends upon tne signal-to-noise ratio between the l degradation response and the transition response. As a general rule, lower f frequencies tend to suppress the transition signal relative to the degradation !
I signal at the expense of the ability to quantify. Similarly, the inspection l of the tube through the sleeve requires the use of low frequencies to achieve l detection with an associated loss in quantificaton. Thus, the search for an optimum eddy current inspection represents a trade-off between detection and -
quantification. With the crosswound coil type inspection, ttils optimization leads to a primary inspection frequency for the sleeve oil the order of [
l a.c.e and for the tube and transition regions on the order of (
3a .c.e ,
Figure 6.1-2 shows a typical [~. .]A'C phase angle versus degradation depth curve for the sleeve from which 00 sleeve penetrations can be assessed.
In the regions of the parent tube above the sleeve, conveational bobbin coil or crosswound coil inspections will continue to be used. However, since the diameter of the sleeve is smaller than that of the tube, the fill factor of a probe inserted through the sleeve may result in a decreased detection capability for tubing degradation. Thus, it may be necessary to inspect the unsleeved portion of the tube above the sleeve by inserting a standard size probe through the U-bend from the unsleeved leg of the tube.
For the tube sleeve combination, the use of the crosswound probe, coupled with a multifrequency mixing technique for further reduction of the remaining noise signals significantly, reduces the interference from all discontinuities (e.g.
transition) which have 360-degree symmetry, providing improved visibility for discrete discontinuities. As is shown in the accompanying figures, in the ,
laboratory this technique can detect 00 tube wall penetrations with acceptacle signal-to-noise ratios at the transitions when the volume of metal removed is '
equivalent to tne ASME calibration standard. .
- 479/04:387: 49-3C 6-3
The response from the tube / sleeve assembly transitions with the crosswound coil is shown in Figures 6.1-3, 6.1 4, and 6.1-5 for the sleeve standards, tube standards and transitions, respectively. Detectability in transitions is improved by the combination of the various frequencies. For the cross-wound probe, two frequency combinations are shown; ( *
.]D'C Figure 6.1-6 shows the phase / depth curve for the tube using this combination. As examples of the detection capability j at the transitions, Figures 6.1-7 and 6.1-8 show the responses of a 20 percent 00 penetration in the sleeve and 40 percent 00 penetration in the tube, respectively.
For inspection of the region at the top end of the sleeve, the transition response signal-to-noise ratio is about a factor of four less sensitive than that of the expansions. Some improvement has been gained by tapering the wall' thickness at the top end of the sleeve. This reduces the end-of-sleeve signal
- by a factor of approximately two. The crosswound coll, however, again significantly reduces the response of the sleeve end. Figure 6.1-9 shows the i
response of various ASME tube calibration standards placed at the end of the sleeve using the cross-wound coil and the C. lC frequency combination. Note that under these condttions, degradation at the top end of i the sleeve / tube assembly can be detected.
i i
i 6.2
SUMMARY
i Conventional eddy cy rent techniques have been modified to incorporate the f most recent technology in the inspection of the sleeve / tube assemoly. The i
resultant inspection of the sleeve / tube assembly involves the use of a l
i cross-wound coil for the straight regions of the sleeve / tube assemoly and for l the transition regions. Tne advent of MIZ-18 digital E/C Instrumentation and I
its attendant increased dynamic range and the availability of 8 channels for
- four raw frequencies nas expanded the use of the crosswound coil for sleeve inscection. While there is a significant improvement in the inspection of portions of the assemoly using the cross-wound coil over conventional bcbeln coils efforts continue to advance t9e state-of-the-art in eddy current 4447M/ a'337.a} ia
inspection techniques. As improved tecnniques are developed and verified, '
they will be utilized. For the present, the cross-wound coil probe represents '
an inspection technique that provides additional sensitivity and support for ,
eddy current techniques as a viable means of assessing the tube / sleeve assembly.
l i
k
l.
I Figure 6.1-1 Absolute eddy current signals at 400 khz (front and rear coils)
G-6
M ,e t -
. h I
Figure 6.1-2
[ Ja.c.e Calibration Curve l
6-7 l
f Figure 6.1-3 I
E.C. Signals from the ASTM Standard. Machined on tne 5: ewe 0.0. of the Steeve Twee Anemely Without Excansion IC.on 3
Wound Coil Proce) l 6-3 k
.,LJ , e w -
Figure 6.1- 4 E.C. Signals from tne ASTM Stancare. Macnined on tne Twee 0.0. of the Slene-Tube Assemoty Witnout Excension (Crou -
Wound Coil Procal i
I 6-9 ,
f I
- a l
l Figure 6.1-5 l' E.C. Signals from the Ezoansion Transition Region of tre Tuce Sleeve Assemoly (Cross Wound Coil Proce 6-10
le.e Figure 6.1- 6 Ecdy Cw:nnt Calibration Curve for ASME Tune Stancare at .
( ) a.c.e anc a Mix Using the Cross Wounc Coil Pecce 6-11
i a
Figure 6.1- 7 E.C. Signal from a 20% Deen Hole. Half the Volume of AST Stancare Macninec on tne Steeve 0.0. in tne Excansion Transition Region of tne Sleeve-Tube Assemeiy (Cross Woec 6-12
- ge.e Figure 6.1-8 ,
E.C. Signal from a 40% ASTM Stancare. Machined on tre Tuos -
0.0. in the Excension Transition Region of Stece Twee Anemoly (Crom Wounc Coil Proce) 6-13 .
l
~~
Eddy Current Rescense of the ASME Tube Standare at the E-of the Steeve using the Cross Wound Coil Prcoe and
, Multifrecuency Comeination rigure 6.'-9 6-14
i 7.0 ALARA CONSIDERATIONS FOR SLEEVING OPERATIONS The repair of steam generators in operating nuclear plants requires the utilization of appropriate dose reduction techniques to keep radiation exposures As Low As Reasonably Achievable (ALARA). Westinghouse maintains an extensive ALARA program to minimize radiation exposure to personnel. This program includes: design and improvement of remote and semi-remote tooling, including state-of-the-art robotics; decontamination of steam generators; the use of shielding to minimize radiation exposure; extensive personnel training utilizing mock-ups; dry runs; and strict qualification procedures. In addition, computer programs (REMS) exist which can accurately track radiation exposure accumulation.
The ALARA aspect of the tool design program 15 to develop specialized remote tooling to reduce the exposure that sleeving personnel receive from high radiation fields. A design objective of a remote delivery sleeving system is to eliminate channel head entries and to complete the sleeving project with total exposures kept to a minimum, 1. e., ALARA. A manipulator arm can be Installed on a fixture attached to the steam generator manway after video cameras and temporary nozzle covers have been installed. A control station operator (CS0) then manually operate controls to guide the manipulator arm through the manway and attach a baseplate to the tubesheet. The installation of the arm requires only one platform operator to provide visual observation and assistance with cable handling from the platform. The control station for the remote delivery system is located outside containment in a specially designed control statton traller. As previously indicated, under some conditions positioning of sleeve / tooling with the base Robotic system may not be practical. In these circumstances alternate techniques may be utilized.
such as hands-on (manual position, alternate robotic or semi-remotely operated equipment or a combination of the two.
The control of personnel exposures can also be effected by careful planning, training, and preparation of maintenance procedures for the job. This form of -
administrative control can ensure that the minimu:n number of personnel will be used to perform the various tasks. Additional methods of minimizing exposure include the use of remote TV and radio surveillance of all platform and 3593Me0'0986: 49 7-1
l l Channel head operations and the monitoring of personnel exposure to identify ,
high exposure areas for timely improvement. Local shielding will be used whenever possible to reduce the general area background radiation levels at the work stations inside containment.
7.1 N0ZZLE COVER AND CAMERA INSTALLATION / REMOVAL l
~
The installation of temporary nozzle covers in the reactor coolant pipe nozzles in preparation of the steam generators for sleeving operations may -
require channel head entries. The covers are installed to prevent the accidental dropping of any foreign objects (i.e., tools, nuts, bolts, debris, etc.) into the reactor coolant loops during sleeving operations. In the event that an accident did occur, an inspection of the loop would be required and any foreign objects or debris found would be retrieved. The impact on k=
schedule and radiation exposures associated with these recovery operations would far exceed the time and exposures expended to install or remove loop nozzle covers. Consequently, it is considered an ALARA-efficient procedure to utilize temporary nozzle covers during sleeving operations.
The use of video monitoring systems to observe ROSA operations in the channel head may require manual installation. The installation of overview cameras to monitor sleeving operations may require a full or partial channel head entry.
The installation and removal of this equipment in the steam generators are the only anticipated potentials requirements for channel head entries during the sleeving project.
7.2 PLATFORM SETUP / SUPERVISION The majority of the radiation exposures recorded for the sleeving program is etDetted to result primarily from personnel working on or near the steam generator platforms and in the channel head for hands-on operations. The 35339/060986:49 7-0
setup and checkout of equipment for the various sleeving processes, installationNemovaloftooling,andtheoperationofthetoolingarethe major sources of radiation exposure. In addition to channel head video monitoring systems, visual monitoring and supervision by one or more workers on the platform will be required for a major part of the sleeving schedule.
Experience has shown that rapid response to equipment adjustment requirements is efficiently accomplished by having a platform worker standing by in a relatively low radiation area during operations. Worker standby stations have ranged from the low radiation fields behind the biological shield to lead blanket shielding installed on the platform. Even though radiation levels on the platform are much lower than channel head levels, a substantially larger amount of time will be spent on the platforms giving rise to personnel oxposures. An evaluation of radiation surveys around the steam generators should indicate appropriate standby stations.
7.3 RA0 HASTE GENERATION The surface preparation of tubes for the installation of sleeves requires that the ontde film be removed by a honing process. A flexihone attached to a flexible rotating cable will be used to remove the oxide film on the inside surface of the steam generator tubes. The volume of solid radwaste is expected to consist of spent hones, flexible honing cables, hone filter assemulles (optional), [ ]A*C and the normal anti-C consumables associated with steam generator maintenance. The anti-C consumables are the customer's responsibility and will not be addressed in this report. .
For the [ '
]A*C approximately thirty tubes can be honed before the hone is changed for process control and [
]."'C A typical estimate of the radioactive concentration from a honed tube transported by the [
]#
- C d is given in Table 7.5-1. These concentrations area based on a general area radiation level of 4R/HR. The tube hones as well as the tubes
[ ] onsequently, radiation levels of the spent hones are normally 1-2 r/hr based on field measurements in previous sleeving projects.
3593H/060986: 49 7-3
TABLE 7.5-1 ESTIMATE OF RADIOACTIVE CONCENTRATION IN WATER PER TUBE HONE 0 (TYPICAL) a.c.e ASSUMPTIONS
- 1) Tube honed 45 inches (in length)
- 2) Hater flow rate of 0.6 gallons per tube honed
- 3) Essentially all radioactivity removed from tubes honed.
3593M/060986:49 74 i _, - - . _ . . . _ _ , _ _ ,.
The flexible honing cable used to rotate the hone inside the tubes 's also flushed during the honing process. However, the construction of the stainless I steel cable ulli cause radioactivity to build up over the course of the project. Radiation levels on segments of the cable could reach 5-10 R/Hr contact dose rates for major sleeving jobs. It is expected that an average of one cable per steam generator will be used during the sleeving project. The cables are consumables and are drummed as solid radwaste.
7.4 HEALTH PHYSICS PRACTICES AND PROCEDURES 4 The Health Physics (HP) requirements for sleeving will be those establshed by the licensee. Westinghouse will provide radiological engineering assistance, as needed, to assist in coordination of the radiological aspects of the
! Westinghouse activities. Open communications between involved parties will be maintained so that the best possible health physics practices can be established for the sleeving program. The HP procedures of the utility will be the guidelines followed during the sleeving operation. However, in specific instances where improvements are mutually recognized but not covered in these HP procedures, appropriate changes will be made according to established change procedures.
The fleid service procedures which are prepared by Westinghouse for the complete setup of equipment and subsequent sleeving operations include the specific radiologically related responsibilities, prerequisites and precautions. These will further n.inimize exposure and control contamination.
Mockup training at the Westinghouse Waltz Mill Training Center includes the following radiological practices:
o Technical skill training while dressed in full Anti-C clothing including bubble hoods.
o Identification of high radiation zones on the work platform and _
emphasis of minimizing stay times.
3593M/060986:49 7-5
o Handling of contaminated tools and changeout of contaminated mandrels.
o Location and use of waste disposal containers.
Westinghouse implements an extensive training and qualification program to prepare supervisory, maintenance and operations personnel for fleid Implementation of the sleeving process. Satisfactory completion of this training program verifles that the personnel addressed are quallfled to perform all assigned operations from a technical as well as radiological l aspect in keeping with the ALARA principals.
The qualification program consists of two phases:
Phase I - classroom Phase II - mockup Phase I - Consists of classroom training and addresses subject material that is related to the overall sleeving program. The Phase I instructors generate and administer an examination for Phase I training of sufficient difficulty to demonstrate that a trainee has sufficient knowledge of the material presented. This examination is written. All trainees will be tested. A minimum grade of 80 percent is required. The test results shall be documented are retained for audit. 1 Phase II - Consists of hands-on and mockup sleeving training during which the trainee must demostrate a' capability to perform a function or operation in a limited amount of time. If team training is required, each trainee must be able to perform all tasks required of the team.
7.5 AIRBORNE RELEASES The implementation of the proposed sleeving processes in operating nuclear plants has indicated that the potential for airborne releases is minimal. The major operations include C lC and sleeve installation, t
3593M/060986: 49 7-6
Experience has shown that these sleeving processes do not contrioute to -
airborne releases.
7.6 PERSONNEL EXPOSURE ESTIMATE The total personnel exposures for steam generator sleeving operations will depend on several plant dependant and process related factors. These may include, but not be limited to; the scope of work (quantity of sleeves, etc).
plant radiation levels, Ingress / egress to the work stations, equipment performance and overall cognizance of ALARA principles. Consequently, the projection of personnel exposures for each specific plant must be performed at the completion of mockup training when process times for each operation have been recorded. The availability of plant radiation levels and worker process times in the various radiation fields will provide the necessary data to project personnel exposure for the sleeving project. ,
The calculation of the total MAN-REM exposure for completing a sleeving project may typically be expressed asfollows:
P - ((N 3
.0) 3
+5) .N g 9
P = Project total exposure (MAN-REM)
N Number of sleeves installed / steam generator 0 Exposure / sleeve installed 3
S g
Equipment setup / removal exposure per steam generator Ng - Number of steam generators to be sleeved This equation and appropriate variations are used in estimating the total personnel exposures for the sleeving project. _
3593M/060986:49 7-7
Man-rem exposure results obtained during a recent Westinghouse steam generator sleeving operation showed approximately 50 to 100 millirem / tube, using the Remote Operating ServiceArm (ROSA).
- Man-rem exposure results obtained from recent Westinghouse steam generator manual sleeving operations show approximately 300 man-rem for sleeving of 650 l tubes. This estimate is based on chemical decontamination of the steam generator channel heads including approximately 4 feet inside the steam generator tubes with a resulting field of approximately 4 R/HR.
t 6
3593M/060985:49 7-8
8.0 INSERVICE INSPECTION 'LAN FOR SLEEVED TUBES
\
In addressing current NRC requirements, the need exists to perform periodic inspections of the supplemented pressure boundary. This new pressure boundary consists of the sleeve with a joint at the primary face of the tubesheet and a joint at the opposite end of the sleeve.
l The inservice inspection program will consist of the following. Each sleeved tubs will be eddy current inspected on completion of installation to obtain a baseline signature to which all subsequent inspections will be compared.
Periodic inspections to monitor sleeve wall conditions will be performed in accordance with the inspection section of the plant Technical Specifications.
This inspection will be performed with multi-frequency eddy current equipment.
Periodic pressure testing of the steam generator, similar to that performed following tube plugging, will be performed as recommended in the technical manual.
4447M/042387: 49 3-1
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