ML20236X780

Nonproprietary Version of Point Beach Unit 2 Steam Generator Sleeving Rept (Mechanical Sleeves) Follow on Sleeving Effort

ML20236X780
Person / Time
Site:	Point Beach
Issue date:	09/30/1987
From:	WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML19302D135	List: ML20236X765 ML20236X780
References
S6-87-10-019, S6-87-10-19, WCAP-11574, NUDOCS 8712100264
Download: ML20236X780 (164)

Text

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WESTINGHOUSE CLASS'3 WCAP 11574 56-87-10-019 l

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POINT BEACH UNIT 2 l

STEAM GENERATOR SLEEVING REPORT (Mechanical Sleeves)

Follow on Sleeving Effort I

l September 1987

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PREPARED FOR WISCONSIN ELECTRIC POWER l

a WESTINGHOUSE ELECTRIC CORPORATION STEAM GENERATOR TECHNOLOGY DIVISION P.O. BOX 855 PITTSBURGH, PA 15230 1

I 4746M/102387:49 -l 8712100264 871204 l

PDR ADOCK 0500C301

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TABLE OF CONTENTS

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Section Title Page l.0 INTR 000ChkON 1-1 2.0 SLEEVING ORJECTIVES AND BOUNDARIES 2-1 2.1 Objectives 2-1 2.2 Sleeving Boundary 2-1 2.3 Report Applicability 2-2 3.0 DESIGN 3,j 3.1 Sleeve Design Documentation 3-1

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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"

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3.3.4.2 Description of Verification Tests for the Lower Joint 3-35 3.3.4.3 Leak Test Acceptance Criteria 3-37 3.3.4.4 Results of Verification Tests for Lower Joint 3-39 l

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4637M/102787:49 -2 l

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9 TABLE OF CONTENTS (Continued)

Section Title Pace 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 I

Upper HEJ 3-46 3.3.6 Test Program for the Fixed /Flxed Mockup 3-50 3.3.6.1 Description of the Fixed / Fixed Mockup 3-50 3.3.6.2 Description of Verification Tests for the Fixed / Fixed Mockup 3-61 3.3.6.3 Results of Verification Tests for the F1xed/ Fixed Mockup 3-61 i

i 3.3.7 Effects of Sleeving on Tube-to-Tubeshe'et Weld 3-63 1

3.4 Analytical Verification 3-65 3.4.1 Introduction 3-65 3.4.2 Component Description 3-65 3.4.3 Material Properties 3-67 3.4.4 Code Criteria 3-67 l

3.4.5 Loading Condltions Evaluated 3-67 3.4.6 Methods of Analysis 3-72

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l 3.4.6.1 Model Development 3-73 3.4.6.2 Thermal Analysis 3-75 3.4.6.3 Stress Analysis 3-76 s.

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TABLE OF CONTENTS (Continued)

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Section Title Page 3.4.h Results of Analyses

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3-78 3.4.7.1 Primary Stress Intensity 3-78 3.4.7.2 Range of Primary Plus Secondary Stress Intensities 3-79 3.4.7.3 Range of Total Stress Intensities 3-79 3.4.8 References 3-88 j

3.5 Special Considerations 3-89 3.5.1 Flow Slot Hourglassing 3-89 3.5.1.1 Effect on Burst Strength 3-89 3.5.1.2 Effect on Stress Corrosion Cracking,(SCC) Margin 3-89 3.5.1.3 Effect on Maximum Range of Stress Intensity and Fatigue Usage Factor 3-89 3.5.2 Tube Vibration Analysis 3-90 3.5.3 Sludge Height Thermal Effects 3-90 3.5.4 Allowab'le Sleeve Degradation 3-90 3.5.4.1 Minimum Required Sleeve Thickness 3-90 3.5.4.2 Determination of Plugging Limits 3-91 3.5.4.3 Apo!! cation of Plugging Limits 3-92 3.5.5 Effect of Tubesheet Interaction 3-97 3.5.6 Structural Analysis of the Lower Hybrid Expansion Joint 3-97 3.5.6.1 Primary Stress Intensity 3-9'?

3.5_6.2 Range of Primary Plus Secondary Stress In*ensities 3-97 til 4

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TABLE OF CONTENTS (Continued) i l

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Section Title Page

'3.5.6.3 Range of Total Stress Intensities 3-100 3.5.7 Evaluation of Operation with Flow Effects Due to Sleeving 3-102 3.5.7.1 One Sleeve Per Tube 3-102 l

3.5.7.2 Two Sleeves Per Tube 3-103

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3.5.7.3 Flow Effects Summary 3-104 1

3.5.8 Alternate Sleeve Materials 3-108

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3.5.9 Effect of an Axial Tube Lock-up on Fatigue Usage Factor 3-108 i

l 3.5.10 Minimum Sleeve Wall Thickness 3-109 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 Ory Honing 4-4 4.2 Sleeve Insertion and Expansion 4-4 4.3 Lower Joint Seal 4-5 l

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4637M/102787:49 -5

IABLt OF CONitNTS (Continued)

Section Title Page Uppir' Hybrid Expansion Joint (HEJ) 46 4.4 r

4.5 Process Inspection Sampling Plan 4-6 4.6 Establishment of Sleeve Joint Main Fabrica-tion Parameters 4-7.

4.6.1 Lower Joint 4-7 4.6.2 Upper HEJ 4-7 5.0 SLEEVE / TOOLING POSITIONING TECHNIQUE 5-1 6.0 NDE INSPECTABILITY 6-1 1

6.1 Eddy Current Inspections 6-1 6.2 Summary 6-5 j

7.0 ALARA CONSIDERATIONS FOR SLEEVING OPERATIONS 7-1 7.1 Nozzle Cover and Camera Installation / Removal 7-2 i

t 7.2 Platform Setup / Supervision 7-2 l

7.3 Radwaste Generation 7-3 7.4 Health Physics Practices and Procedures.

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4637M/102787:49 -6 i

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l TABLE OF CONTENTS (Continued)

Section Title g

7.5 Airborne Releases 76 7.6 Personnel. Exposure Estimate 7-7 8.0 INSERVICE INSPECTION PLAN FOR SLEEVED TUBES 8-1 i

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LIST OF TABLES Table Title Pag 3.1-1 ASME Code and Regulatory Requirements 3-2 3.3.2-1 Summary of Corrosion Comparison Data for 3-11 Thermally Treated Alloys 600 and 690 3.3.2-2 Effect of Oxidizing Species on the SCC Susceptt-3-12 bl11ty of Thermally Treated Alloy 600 and 690 C-rings in Deaerated Caustic j

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' 3.3.3-1 Design Verification Test Program - Corrosion 3-31 a

3.3.3-2 Residual Stresses at [

.l,c.e 3-32 I

3.3.3-3 Results of Magnesium Chloride Tests at [

3-33 a,c.e j

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3.3.3-4 Rasults of Magnesium Chloride Tests at (

3-34 a,c.e j

3.3.4.3-1 Maximum Allowable Leak Rates For Point Beach Unit 2 3-38 Generators l

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 (Fatigue 3-51 and Extend Operation Tests Incl.)

l 3.3.5.3-2' Test Results for HEJ's Formed Out of Sludge 3-53 l

(Static Axial Load Leak Test, SLB and Reverse

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Pressure Test Incl.)

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LIST OF TABLES (Continued)

Table Title Page 3.3.5.3-3 Test Results for HEJ's Formed In Sludge 3-55 4

(Fatigue and Roverse Pressure Tests Includ.)

I 3.3.5.3-4 Test Results for HEJ's Formed in Sludge (Axial 3-57 f

Load Leak Test and Post-SLB Test Included).

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3.3.5.3-5 Alloy 690 Limited Scope Test Results - Upper HEJ 3-58 l

3.3.6.3-1 T,est Results for Full Length Sleeves Formed and 3-62 Leak Tested in Fixed /Ftxed Hockup (In sludge and Out'of Sludge).

3.4.4-1 Criteria for Primary Stress Intensity Evaluation 3-68

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(Sleeve)

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l' 3.4.4-2 Criteria for Primary Stress Intensity Evaluation 3-69 (Tube)

I 3.4.4-3 Criteria for Primary Plus Secondary and Total 3-70 l

Stress Intensity Evaluation (Sleeve) l 3.4.4-4 Criteria for Primary Plus Secondary and Total 3-71 Stress Intensity Evaluation (Tube) l 3.4.7.1-1 Umbrella Pressure Loads for Design, Upset, 3-81 Faulted, and Test Conditions 3.4.7.1-2 Results of Primary Stress Intensity Evaluation 3-82 (Upper Hybrid Expansion Joint)

Primary Maabrane Stress Intensity, P, 3.4.7.1-3 Results of Primary Stress Intensity Evaluation 3-79 (Upper Hybrid Expansion Joint)

Primary Membrane Plus Bending Stress Intensity, Pg+P' b

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LIST OF TABLES (Continued) l Table Title Page 3.4.7.2-1 Pressure and Temperature Loadings for Maximum 3-80 Range of Stress Intensity and Fatigue Evaluations 3.4.7.2-2 Results of Maximum Range of Stress Intensity 3-8b Evaluation (Upper Hybrid Expansion Joint) 3.4.7.3-1 Results of Fatigue Evaluation (Upper Hybrid 3-87 Expansion Joint) 3.5.4-1 Regulatory Guide 1.121 Criteria 3-95 3.5.6.1-1 Results of Primary Stress Intensity Evaluation 3-98 (Lower Hybrid Expansion Joint) Primary Membrane Stress Intensity, P, 3.5.6.1-2 Results of Primary Stress Intensity Evaluation 3-99 i

(Lower Hybrid Expansion Joint) Primary Membrane Plus Bending Stress Intensity, Pg+Pb 3.5.6.2-1 Results of Maximum Range of Stress Intensity 3-101 Evaluation (Lower Hybrid Expansion Joint) 3.5.7-1 Allowable Sleeving Parsmeters Under Normal 3-106 Conditions (One Sleeve Per Tube) 3.5.7-2 Allowable Sleeving Parameters Under Normal 3-107 Conditions (Two Sleeves Per Tube) 3.5.9-1 Results of Maximum Range of Stress Intensity 3-110 Evaluation. Axial Tube Lockup 3.5.9-2 Results of Fatigue Evaluation. Axial Tube Lockup 3-111 ix 4637M/102787:49 -10 t.

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LIST OF TABLES (Continued)

Table Title Page 4.0-1 Sleeve Process Sequence Sumary 4-2 1

7.3-1 Estimate of Radioactive Concentration in 7-4 Water per Tube Honed (Typical) l l

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4637M/102787:49 -11 m

LIST OF FIGURES Figure Title Page J.c.e Sleeves 2-3

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a 2.2-1 Sleeving Boundary [

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 After 3-14 5000 Hours Exposure of Alloy 600 and 690 C-Rings to 10% NaOH at 332'C (650*F) 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 MgCl 2 (Stainless Steel) or Polythlonic Acid I

(Alloy 600) 3.3.3-4 Results of C-Ring Tests of Type 304 Heat 3-28 No. 605947 in Belling HgCl 2

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4637M/102787: 49 -12

LIST OF FIGURES (Continued)

Flag Title Page 3.3.3-5 Axial Residual Stresses in Tube / Sleeve Assembly 3-29 at a Depth of 0.001 g 0.0004 in at Five locations Along Length of Transition 3.3.3-6 Circumferential Residual Stresses in Tube / Sleeve 3-30 Assembly at Depth of 0.001 1 0.0004 in. at Five Locations Along Length of Transition 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 / Fixed Mockup - HEJ 3-64 3.4.2-1 Hybrid Expansion Upper Joint / Roll Expanded Lower Joint Sleeve Configuration 3-66 3.5.4-1 Application of Plugging Limits 3-94 6.1-1 Absolute Eddy Current Signals at 6-6 400 kHz (Front and Rear Colls) a 6.1-2

[

l,c.e Calibration Curve 6-7 6.1-3a

.610 Diameter Probe Response to ASME Tube 6-8 Standards Using MIZ-18 Data Acquisition l

6.1-3b

.720 diameter probe Response to ASME Tube 6-9 Standards with HIZ-18 Data Acquisition I

xit 4637M/102787:49 -13 L

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LIST OF FIGURES (Continued)

Flaure Title Pace 6.1-4 Eddy Current Signals from the ASTM Standard, 6-10 Machined on the Sleeve 0.0. of the Sleeve / Tube Assembly Without Expansion ( Cross Wound Coll Probe )

6.1-5 Eddy Current Signals from the ASTM Stendard, 6-11 Machined on the Tube 0.D. of the Sleeve / Tube Assembly Without Expansion ( Cross Wound Coll Probe )

6.1-6 Eddy Current Signals from the Expansion Transition 6-12 Region of the Sleeve / Tube Assembly (Cross Wound Coll Probe )

1 6.1-7 Eddy Current Calibration Curve for ASME Tube 6-13 Standard at [

]"'C and a Mix Using the Cross Wound Coll Probe Eddy 'urrent Signal from a 20 Percent Deep Hole, 6-14 6.1-8 C

Half the Volume of ASTM Standard, Machined on the Sleeve 0.D. In the Expansion Transition Region of the Siseve/ Tube Assembly (Cross Wound l

Coll Probe)

I 6.1-9 Eddy Current Signal from a 40 Percent ASTM 6-15 Standard, Machined on the Tube 0.D. In the Expansion Transition Region of the Sleeve / Tube Assembly (Cross Wound Coil Probe) 6.1-10 Eddy Current Response of the ASME Tube Standard 6-16 at the End of tne Sleeve Using the Cross Wound Coil Probe and Multifrequency Combination xilt 4637M/102787:49 -14 l

1.0 INTRODUCTION

l" The document herein contains the necessary technical information to support the sleeving repair process as applied to the Point Beach Unit 2 (WIS) Model 44 steam generators. As a result of development programs in steam generator j

repair, Westinghouse has developed the capability to restore degraded steam generator tubes by means of a sleeve.

Sleeving at Point Beach Unit 2 is not a novel approach to tube wall degradation.

It was previously implemented at Point Beach Unit 2 to restore the pressure boundary integrity of a number of degraded tubes.

The previously performed work is documented in WCAP-9960 Revision i dated February, 1982 titled Point Beach Steam Generator Sleeving Report.

The sleeving technique and instal-lation process utilized in the previous sleeving effort is essentially the same as that described herein. While the reports are consistent with each other, this document has been modified to reflect advancement in the sleeving process as well as more recent test data. Of all the changes, the most significant are 1) the use of thermally treated Alloy 690 sleeves instead of sleeves fabrichted from thermally treated Alloy 600 material and 2) sleeving both the hot and cold legs of the same tube in the same generator as opposed to the prior report which considered only cold leg sleeving.

Items 1 and 2 are addressed more fully in Section 3.0 of this report.

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Sleeving 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 Point Beach Unit 2, this repair process is expected to allow numerous tubes to remain.in service thereby helping to maintain the design life of the steam generator and the efficiency of the entire nuclear steam

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supply system.

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To date, approximately 22,000 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

Alloy 600, 690, and bimetallic 625 and 690 have been installed by a variety of techniques - hands-on (manual) installation, Coordinate Transport (CT) system installation, and Remotely Onerated Service Arm (ROSA) robotic installation.

Westinghouse sleeving programs have been successfully implemented after 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 enhancements and a remote sleeve delivery system (CT) were subsequently developed and adapted to Westinghouse Model 44 series steam generators for large scale programs at two operating plants (2971 and 3000 sleeves).

This technology has also been modified to facilitate installation of sleeves in a plant with non-Westinghouse steam generators. A total of 5,187 sleeves were installed in three successive programs of 2,036, 2.926, and 225 sleeves utilizing CT and ROSA delivery systems. Also completed was a 17 sleeve ROSA delivery program in a Model 51 steal generator overseas and a 635 sleeve manual installation prCgram in a previously sleeved plant.

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4746M/102387:49 -16 I-2 l

i 2.0 SLEEVING OBJECTIVES AND BOUNDARIES 2.1 OBJECTIVES' Point Beach Unit 2 (HIS) is a Westinghouse-designed 2 loop pressurized water j

reactor rated at 1519 MHt.

The unit utilizes two vertical U-tube steam

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generators.

The steam generators are Westinghouse Model 44 Series containing heat transfer tubes with dimensions of 0.875 inch nominal OD by 0.050 inch nominal wall thickness.

l The sleeving concept and design are based on observations to date that the-tube degradation due to o,perating environmental conditions has occurred near the tubesheet areas of the tube bundle.

The sleeve has been designed to span f

the degraded region in order to maintain these tubes in service.

j The sleeving program has two primary objectives:

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1.

To sleeve tubes in the region of known or potential tube degradation.

i 2.

To minimize the radiation exposure to all working personnel (ALARA) l 2.2 SLEEVING BOUNDARY f

1 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 elevation tolerance of one inch will be employed to allow for any potenz:lal 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 i

define the sleeving boundaries.

Figure 2.2-1 shows an estimated radial sleeving boundary for a (

]C sleeve as determined by a geometric radius computed from the channelhead surface-to-tubesheet primary face clearance distance minus the tooling clearance distance.

(The actual "as I

l is" bowl ge@etry will be slightly different in certain areas.) This is the 1

l sleeving boundary for a generic Westinghouse series 44 steam generator and J.c.e Sj,,y,,

j a

represents the maximum sleeving potential with a [

l 4746M/102387:49 -17 2-1

i Tubes within the sleeving boundary that are degraded beyond the plugging limit J

'C sleeve or A

but not within the axial restrictions of the [

not within the radial sleeving boundary will be plugged.

The actual sleevable region may be modified based on tool length or other variables.

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The actual tube plugging / sleeving map for each steam generator will be l

provided as part of the software deliverables at the conclusion of the' i

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 the sleeve pressure boundary which are greater than the plugging limit.

2.

Concurrenc'e on the eddy current analysis of the extent and location of the j

degradation.

I 2.3 REPORT APPLICABILITY

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3.0- DESIGN 3.1 SLEEVE DESIGN DOCUMENTATION The Point Beach Unit 2 steam generators were built to the 1965 edition of Section III of the ASME Boller and Pressure Vessel Code, however, the sleeves have been designed and analyzed to the 1983 edition of Section III of the Code.

I through'the winter 1983 addenda as well as applicable Regulatory Guides.

The associated materials and processes also meet the requirements of the Code.

The specific 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 Figure 3.2-1.

C' aj.c.e At the upper end, the sleeve configuration (see Figure 3.2-1) consists of a section which is [

]"'C

This joint configuration is known as a hybrid expansion joint (HEJ).

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In the process of sleeve length optimization and allowing for axial tolerance in locating defects by eddy current inspection, the guideline was the lower most elevation of the hard roll region to be positioned a minimum of 1 inch above the degraded area of the tube.

4746M/102387:49 -20 3-1

TABLE 3.1-1 l

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ASME CODE AND REGULATORY REQUIREMENTS Item Applicable Criteria-Requirement Sleeve Design Section III NB-3200, Analysis l-NS-3300, Hall Thick-ness l

Operating Requireme"ts Analysis Conditions n

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Reg. Guide 1.83 S/G Tubing Inspec-tibility l

Reg. Guide 1.121 Plugging Margin j

Sleeve Material Sdction II Material Composition Section III NB-2000, Identifica-tion Tests and Examinations 1

1 Code Case N-20 Mechanical Proper-ties Sleeve Joint 10CFR100 Plant Total Primary-l Secondary Leak Rate Technica.1 Specifications Plant Leak Pate l

I 4746M/102387:49 -21 3-2 l

i At the lower end, the sleeve configuration (Figure 3.2-2) consists of a section which is (~

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.J,c.e The lower end of the a

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 l

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 Boller 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 t 41 sting tube were to become severed just above the upper edge of the mechanical joint, the tube would be restrained by the sleeve and lateral and axial motion, and subsequent leakage would be limited. Restrictied lateral motion could 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 i

1 current signal interpretation.

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To minimize stress concentrations and enhance inspectability in the area of I

the upper expanded region, (

),a,c.e,f The sleeve material, thermally treated Alloy 690, is selected to provide l

additional resistance to stress corrosion cracking. (See Section 3.3.2 for further details on the selection of thermally treated Alloy 690).

4637M/102787:49 -22 3-3 l

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Installed Sleeve with Hybrid Expansion Upper Joint Configuration 4746H/101587:49 -23 3-4 l

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Sleeve Lower Joint Configuration 4746M/101587:49 -24 3-5 t.

l 3.3 DESIGN 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.

l A substantial data base exists from previous test programs which verifles the l

adequacy of the sleeve design and process.

The results of much of this j

testing is directly applicable to the present sleev",g program.

The sleeve material is Alloy 690 (UNS 066900) manufactured t requirements of ASME l

58-163 with supplemental requirements of Code Case N-20.

The material has been heat treated to anhance its resistance to corrosion in steam generator j

primary water and many secondary-side water environments.

This material in i

the thermal treated condition has been,used in prev'ious sleeving programs.

Previous testing of the sleeve design has been for sleeves to be installed into Model 44 steam generators.

The installation of the sleeves by the l

combination of [

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lC is the same as that verified and used in previous sleeving programs.

Rigorous mechanical testing programs were conducted to verify the sleeve design for various steam generator models.

I The objectives of the mechanical testing programs included:

I Verify the leak resistance of the upper and lower sleeve to tube joints.

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Verify the structural strength of the sleeved tube under normal and accident conditions.

6 4746M/102387:49 -25

Verify the fatigue strength of the sleeved tube under transient loads representing the remaining design life of the plant.

Confirm capability for installation of sleeves in tubes with conditions

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such as. deep secondtry 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, I

temperatures, and loads.

The testing also included evaluation of joints fabricated using Alloy 600 sleeves as well as Alloy 690 sleeves in Alloy 600 tubes.

While the bulk of the original qualification data.is centered on Alloy 600 sleeves, a series of limited scope verification tests were run using i

Alloy 690 fleeves 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 j

and disclosed no areas which would result in a change of joint performance.

The sections that follow describe those portions of the corrosion (sections i

3.3.2-3.3.3) and mechanical (sections 3.3.4-3.3.6) verification programs that are relevant to this sleeving program.

I 3.3.2 CORROSION AND METALLURGICAL EVALUATION The objectives of the corrosion evaluations are (1) to verify that thermally treated Alloy 690 is a suitable material for use in steam generator environments and (2) to verify that sleeving does not have a detrimental effect on the serviceability of the existing tube or the sleeve components.

The material of constri.ction for the steam generator tubes of the Westinghouse f

design, including the steam generators at the Point Beach site, is Alloy 600 l

4746M/102387:49 -26 3-7

in the mill annealed (MA) and Thermally Treated (TT) condition. Alloy 600 is a high nickel austenitic alloy that is nominally 72 percent nickel, 14-17 percent chromium, and 6-10 percent iron.

The sleeving material proposed for sleeving the Point Beach steam generators is Alloy 690 in the thermal treated (TT) condition. Alloy 690 is also a high nickel austenitic material but contains a higher chromium content and a correspondingly lower nickel content and has a nominal composition of 60 percent nickel, 30 percent chromium, and 9 percent iron.

Alloy 690 TT is recomniended in lieu of Alloy 600 MA or TT because laboratory testing has shown the alloy to have a resistance to corrosion in steam generator environments that is equal or better than Alloy 600 in either heat treated condition.

The higher chromium content of Alloy 690 is responsible for this greater corrosion resistance.

In addition, the alloy is thermally treated to enhance its stress corrosion cracking (SCC) resistance.

Alloy 690 TT is the current tubing material of construction recommended by.

Westinghouse for ' steam generator applications.

The stress corrosion cracking performance of thermally treated Alloys 600 and 690 in both off-chemistry secondary side and primary side environments has been extensively investigated. Results have continually demonstrated the additional st.'ess corrosion cracking resistance of thermally-treated Alloys 600 and 690 as compared to mill annealed Alloy 600 material. 01 rect comparison of thermally treated Alloys 600 and 690 has further indicated an additional margin of SCC resistance for thermally treated Alloy 690. (Tab'le 3.3.2-1).

l l

The caustic SCC performance of mill annealed and thermally treated 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 from 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.

l 4746M/102387:49 -27 3-8

The crack growth rates presented in Figure 3.3.2-1 indicate that thermally I

treated Alloys 600 and 690 have enhanced caustic SCC resistance compared to that of Alloy 600 in the mill annealed condition.

The performance of thermally treated Alloys 600 and 690 are approximately equal at temperatures l

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 Alloy 690 at higher temperatures is a result of a lesser temperature I

dependency.

C-ring specimens were tested in 10 percent NaOH solution at 332*C to index the.

relative intergranular attack (IGA) resistance of Alloys 600 and 690.

Comparison of the IGA morphology for these C-rings stressed to 150 precent of the 0.2 percent yield strength is presented in Figure 3.3.2-2.

Mill annealed Alloy 600 is characterized by branching intergranular SCC extending from a 2009 front of uniform IGA.

Thermally treated Alloy 600 exhibited less SCC.

and an IGA front limited to less than a few grains deep.

Thermally treated Alloy 690 exhibited no SCC and only occasional areas of intergranular oxide penetrations that were 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 Alloy 600 C-rings exposed to a deaerated sodium hydroxide environment.is the formation of intergranular SCC with uniform grain boundary corrosion (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 i

is a controlling factor.

Thermal treatment of Alloy 600 provides additional grain boundary corrosion resistance along with additional SCC resistance.

In the case of Alloy 690, the composition provides an additional margin of resistance to IGA and the thermal treatment enhances the SCC resistance.

The adaition of oxidizing species to deaerated sodium hydroxide environments resdits in either a deleterious effect or no effect on the SCC resistance of thermally treated Alloys 600 and 690 and depends on the specific oxidizing 4746M/102387:49 -28 3-9

specie and concentration-(Table 3.3.2-2).

The addition of 10 percent copper oxide to 10 percent sodium hydroxide decreases the SCC resistance of thermally treated Alloys 600 and 690, and also modifies the SCC morphology with the presence of transgranular cracks in the case of Alloy 690.

The exact mechanism responsible for this change is not well understood, but may be related to an increase in'the specimen potential that corresponds to a transpassive potentia!', which may result in an alternate cracking regime.

The specific oxidizing specie and the ratio of oxidizing specie to sodium hydroxide concentration appear to effect the cracking mode. The apparent deleterious effect on SCC resistance is eliminated by lowering the copper oxide or sodium hydroxide concentration.

Mill annealed and thermally-treated Alloys 600 and 690 were tiso evaluated in a number of 8 percent sodium sulfate environments.

The room temperature pH value at the beginning of th'e test was adjusted using either sulfuric acid and j

ammonia. As the pH is lowered, the SCC resistance for mill annealed and

)

thermaliy-treated' Alloy 600 is decreased.

In compariton, thermally treated Alloy 690 did not crack even at a pH of 2, the lowest tested (Figure 3.3.2-3).

.I The primary water SCC test data are presented in Figure 3.3.?-4.

For the beginning of fuel cycle water chemistries,10 of 10 specitaens of mill annealed Alloy 600 exhibited SCC, while 1 of 10 specimens of thermally-treated Alloy 600 had. cracked in exposure times of about 12,000 hours0 days <br />0 hours <br />0 weeks <br />0 months <br />.

In the end of the fuel cycle water chemistries., 9 of 10 specimens of mill annealed Alloy 600 exhibited SCC, while 3 of 10 specimens of thermally-treated Alloy 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 la the mill-annealed or thermally-treated Alloy 690 specimens in either test environment.

~

Continuing investigation of the SCC resistance of Alloys 600 and 690 in primary water environments has shown mill annealed Alloy 600 to be susceptible to cracking at high levels of strain and/or stress.

Therral treatment of Alloy 600 in the carbide precipitation region greatly enhances its SCC l

resistance.

The performance of Alloy 690, both mill annealed and thermhlly treated, demonstrates primary water SCC resistance and is presumably due to alloy composition.

4746M/102387:49 -29 3-10 1

• m

Table 3.3.2-1

SUMMARY

OF CORROSION COMPARISON DATA FOR THERMALLY TREATED ALLOYS 600 AND 690 1.

Thermally treated Alloy 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 Alloy 690 tubing exhibits additional SCC resistance ccmpared to thermal treated Alloy 600 in caustic acid sulfate, and primary water environments.

3.

The alloy composition of Alloy 690 along with a thermal treatment provides additional resistance to caustic induced IGA.

4.

The addition of 10 percent Cu0 to a 10 percent deaarated NaOH environment reduces the SCC resistance of both thermal treated Alloys 600 and 690.

Lower concentrations of either Cu0 or NaOH had no effect, nor did additions of Fe 0 and S10 '

34 2

I I

S.

Alloy 690 is less susceptible to sensitization than Alloy 600.

I l

4746M/102387:49 -30 3 11 l

l Table 3.3.2-2 EFFECT OF OXIDIZING SPECIES ON THE SCC SUSCEPTIBILITY OF THERMALLY TREATED ALLOY 600 AND 690 C-RINGS IN DEAERATED CAUSTIC Temperature Exposure Alloy Alloy j

Environment

(*C)

Time (Hrs) 600 TT 690 TT j

10 Percent NaOH +

316 4000 Increased Increased l

l 10 Percent Cu0 Susceptibility

• Susceptibility

• l l

10 Percent NaOH +

332 2000 NC effect No effect i Percent Cu0 l

1 Percent NaOH +

534 4000 No effect No effect 1 Percent Cu0 10 Percent NaOH +

316 4000 No ' ffect No effect e

10 Percent Fe 0 34 10 Percent NaOH +

316 4000 No effect No effect 10 Percent SiO2

• Intergranular and transgranular SC'.

4716H/102387 : 49 -31 3-12

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i Figure 3 3.2 2. Lignt ?notomicrographs illustraung IG A after 5000 Hours Esposure of inconel Alloy 600 ano 690 C. Rings to 10% NaOH at 3320C (630*F).

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3.3.3 UPPER AND LOWER JOINTS l

l.

All the data presented in Section 3.3.2 relative to the corrosion and stress corrosion cracking resistance of thermally treated 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.

Taole 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 (

l I

I l,c.e The stresses in the l

a sleeve, based on tube to tubesheet data, should be as shown schematically at B 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 expansion technique.

For an outer tube expansion produced solely by l

[

l l

t a3ce 4637M/102787: 49 -36 3-17

4 The specimen design is shown in Figure 3.3.3-2 and the test parameters are listed in Table.3.3.3-2.

C.

m

) a ', c. e '

[.

aj,c e No cracking was detected on the 00 surface of any specimen. These results indicate that the OD stresses are below the threshold required to cause cracking in the stainless steel (less than 10 to 15 ksi).

To summarize the results of this test:

o

[

o aj,c,e 4637M/102787:49 -37 3-18 L_

o

[

j,c.e a

Confirmation that the 00 stresses cx1 the parent tubing are very low tensile or compressive was obtained by X-ray diffraction analysis of an Alloy 600 tube expanded 30 mils and by the parting / layer removal technique, as shown below:

l X-RAY RESIDUAL STRESS MEASUREMENTS OF HEJ JOINT: 0D Of TUBE a,c.e l

l 1

l j

(a) in un-expanded tube above upper most transition j

(b) in un-expanded tube below lower most transition f

CONCLUSION:

Residual stresses on 00 of tube are compressive and results are consistent with MgCl test findings.

2 l

I 4746M/101587: 49 -38 3-19 l

The residual stresses in a HE] with an Alloy 600 MA tube / Alloy 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 testr, are summarized in Figures 3.3.3-5 and 3.3.34.

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 I

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, were me.asured on the tube ID in the fully hydraulic expanded region and on the 00 in the unexpanded tube near the hydraulic expansion transition.

The 0D surface of the sleeve was also in compression in the axial and circumferential directions except for one measurement that was in tension l

(about 5 ksi) in the axial direction in the [

l l,c.e The 10 surface of the sleeve had areas where the stresses were as a

high as about 25 ksi 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.

l l

4746M/101587:49 -39 3-20 L

l

Polythlonic Acid Tests l

j,c,e a

Primary Water Tests

~

Tro tests to confirm the primary water stress corrosion cracking resistance of HE]'s have been conducted. A summary of the results of these tests is as follows:

4746M/102387: 49 -40 3-21

680'F Primary Water Tests:

Material -

a.

Alloy 600 mill annealed tubing with known susceptibility to primary water stress corrosion cracking.

b.

Alloy 600 special thermally treated sleeves.

)

Expansion Matrix:

a,c,e 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 prim'ry to secondary leakage.

a Destructive examination detected no tube wall degradation.

750*F Steam Tests:

Material =

a.

Alloy 600 mill annealed tubing with known susceptibility to primary and pure water.

b.

Alloy 600 special thermally treated sleeves.

4637H/101287:49 -41 3-22

I i

i Expansion Matrix:

a,c.e -

l 1

l 1

l I

l Test Environment:

Temperature:

750*F Pressure:

Secondary and Primary at the same pressure Chemistry:

Hydrogenated pure water f

i 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. NDE

{

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 (PHSCC). Calculations have been made using an equation suggested by the Brookhaven National I

Laboratory } for the prediction of PHSCC.

[

1 l

)

aj,c e i

1)

R. Bandy and D. van Rooyen, A Model for Predicting the Initiation and Propagation of Stress Corrosion Cracking of Alloy 600 in High Temperature Hater.

1 l

l l

4637H/102787:49 -42 3-23 l

lt- - - - - - - - -- - - _ _ _ _ _ _ _ _ - - - _ - _ - _ - _ _ - _

[

I l

y J.C,e a

l 4746M/102387:49 43 3-24 l

l

Iiqurg ).3.3 1 Locaten and ReWtrve Maratude of Resadual Stresses insuced try EssNJ y,e B

9 I

m

~

i l

l 3-25

Figure 3.3.3-2 1 C,9 i

6 5

I 1d I*

M t

l-M M'

i 3-26 It

)

i Figure 3.3.3-3 RESIDUAL STRESSES DETERMINED BY CORROSION TESTS IN HgCL2 (STAINLESS STEEL) OR POLYTHIONIC ACID (ALLOY-600) 3,C,e l

4 l

n t

I i

4746M/101587:49 46 3-27 l

1 j

l Ffgure 3.3.3 3 i

i I

A4.e l

M I

l 1

i i

i i

4 l

Results of C-ring tests of T pe 304 heat its. 405947 in boiling 3

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t Table 3.3.3-1 DESIGN VERIFICATION TEST PROGRAM - CORROSION ISSUE FINDINGS 1.

CORROSION AND STRESS CORROSION

-I~~

2.

CORROSION AND STRESS CORROSION L

CRACKING OF LOWER SLEEVE JOINT 3.

CORROSION AND STRESS CORROSION CRACKING OF UPPER JOINTS f

l.

~

i I

l l

l 4

CORROSION AND STRESS CORROSION l

CRACKING IN ANNULUS I

4746M/101587:49 -50 3-31

Table 3.3.3-2 al c.e RESIOUAL STRESSES AT [

a,c,e 1

4 1

I I

4746h/101587:49 -51 3-32 1

IL-

i Table 3.3.3-3 3.c.e a

RESULTS OF MAGNESIUM CHLORIDE TESTS AT [

&aC.e

~

~

.m 9

0 l

l 3-33

l l

l l

i i

l 1

1 Table 3.3.3-4 i

RESULTS OF MAGNESIUM CHLORIDE TESTS AT (

']A'C

A.C.e j

i

}

t a*

O e

h e

l 3-34

3.3.4 TEST PROGRAM FOR THE LOWER JOINT 3.3.

4.1 DESCRIPTION

OF LOWER JOINT TEST SPECIMENS

~

Th? tube /tubesheet mockup was manufactured so'that it was representative of the partially rolled tube to tubesheet joint (Figure 3.3.4.1-1) of the Model 44 steam generators.

The Point Beach Unit 2 steam generator tubes are partial depth rolled inside the tubesheet.

The formation of lower mechanical rolled joint of tube / sleeve is identical to the above mockup.

The tube was examined J c.e cleaned by j

a with a fiberscope, C swabbing, and re-examined with the fiberscope.

Then the preformed sleeve j

.(Thermally Treated Alloy 600 or 690) was inserted into the tube and the lower joint formed.

[

l a3,c.e.

j

~

3.3,

4.2 DESCRIPTION

OF VERIFICATION' TESTS FOR THE LOWER JOINT The as-fabricated specimens for the Model 44 tests were. tested in the sequence described below. Note that the tests of the Alloy 690 sleeve are similar to those performed on the Alloy 600 sleeve except that the Steam Line Break (SLB) i and Extended Operation Period (EOP) tests were not considered necessary based on previous results.

j 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.

2.

The specimens were fatigue loaded for 5000 cycles.

3.

The specimens were temperatt.re cycled for 25 cycles.

i 4746M/102387:49 -54 3-35

t L

..J. C, e l

m j

i l

)

}

l A

4 I

rigure 3.3.4.1-1 Lower Joint A>4elles Test Sessivnen 3-36

4.

The specimens were leak tested at 3110 pst 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 EOP test was performed after Step 4 for three as-rolled specimens.

3.3.4.3 LEAK TEST ACCEPTANCE CRITERIA Site specific or bounding analyses have been performed to determine the allowable leakage during normal operation and the limiting postulated accident 1

condition.

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 Point Beat:h dnit 2 steam generators. These criteria can be compared to the actual leak test results to provide verification that the mechanical sleeve exhibits no leakage under what would j

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 volumetric measurement is based on assuming 19.8 drops per

)

milliliter.

l 4637M/102787:49 -56 3-37 l

L

l i

TABLE 3.3.4.3-1 MAXIMUM ALLOWABLE LEAK RATE $ FOR POINT BEACH UNIT 2 STEAM GENERATORS 1

I Allewable Leak Allowable Leak j

Condition Rate +

Rate per Sleeve *

.j a,c,e l

I 1

)

~

j j

)

Based on [

Ja,c.e sleeves per steam generator.

1 Standard Technical Spec.ification Limit for i steam generator.

l

+

.. t Ja,c.e The analysis assumes primary and secondary coolant initial inventories of l

19Cl/gm and 0.1HC1/gm of Dose Equivalent I-131, respectively.

In addition, as a result of the reactor trip, an todine. spike is initiated which increases the iodine appearance rate in the primary coolant to a value equal to 500 times the equilibrium appearance rate.

I 4746H/101587:49 -57 3-38

i 3.3.4.4 RESULTS CF VERIFICATION TESTS FOR LOWER JOINT It should be fatee dat in many cases reference is made to "simuleted" l

conditions.

In fact these test conditions simulate only one key aspect of operation. For example, in the case of tne fatigue testing, 5000 cycles were

)

1 used.

This numb r does not represent the number of cycles expected in one I

year, it actually represents the number of expected yearly cycles multiplied by a suitable factor to estcblish an ac:elerated 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 acquire test data within an abbreviated time frame. Con $quentlythetestresultsobtainedanddiscussed throughout the rest of this report are those of accelerated conditions designed to test the sleeve at its endurance limit. Sleeving qualification tests demonstrate that under extreme..celerated test conditions leakage is minimal and the sleeves perform with; acceptable leakage margin.

Additionally, by using that same test ;eries for a?1 sleeve designs it is possible to measure consistency in process modification and or small changes in the overall design to facilitate an assessment of their effect on total sleeve performance.

Reference is occasionally made to the leakage reduction phenomenon of the mechanical joint design.

This is in reference to the phenomena (observed in the test deta) 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 observ6d consistently in all mechanical joint testing.

Another consistent characteristic observed in the testing of mechanical joints is that the leakage, when observed, is gentrally higher at room temperature conditions and, as in the case of the in d.ge reduction phenomena, decreases as the temperature is elevated.

This characteristic has led to the almost exclusive use of the rocm temperature hydrostatic test in the process, tooling, personnel, procedure and demonstration phases associated with a plant specific sleeving operation.

While not a specific part of this report, this additional process verification data exists for review.

4746M/102387:49 -58

The test results for the Model 44/51 lower jotnt 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 [

1

'C

All of the three as-rolled speciraens were leak-tight 8

during the Extended Operating Period (EOP) test.

For the Alloy 690 sleeve tests the following were noted:

Specteens MS-2 (Alloy 690 Sleeve):

Initial leak rates at all pressures and at normal operating pressure following thermal cycling were [

j,b,c.e a

4637M/102787:49 -59 3-40

1 1

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.4

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3-41

{

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F"--------------________._______

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+

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EL w

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44 M

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f 3-43

Specimen MS-3 (Alloy 690 Sleeve):

[

j,b,c.e a

Specimen MS-7 (Alloy 690 Sleeve):

C j,b,c.e a

3.3.5 TEST PROGRAM FOR THE UPPER HYBRID EXPANSION JOINT (HEJ) 4 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.

{

l 3.3.

5.1 DESCRIPTION

OF THE UPPER HEJ TEST SPECIMENS

'{

^

Two types of HEJ test specimens were fabricated for the Model 44 testing (

]"'

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 l

conditions.

[

J., b c Any leakage was collected a

and measured as it issued from the annulus between the tube and sleeve.

This type of specimen was used in the majority of the tests.

q The second type of test specimen was a modification of the first type.

It was utilized in tne reverse pressure tests, i.e., for LOCA and secondary side 4637M/102787: 49 -63 3-44

Me a

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1 Hyorid Isoansson Jaim Test Sessimen 1

1 Tigure 3.3.5.1-1 3-45

hydrostatic pressure tests. As shown in Floure 3.3.5.1-2, the specimen was l

modified by [

)

J,b,c The possible reverse pressura-test leak path is shown in a

l Figure 3.3.5.1-2.

Only specimens like Figure 3.3.5.1-1 (excluding the sludge conditions) were used in the Alloy 690 HEJ specimen fabrication as the effects of sludge had been established in the ear?ier Model 44 tests.

1 1

j 3.3.

5.2 DESCRIPTION

OF VERIFICATION TESTS FOR THE UPPER HEJ l

The verification test program for the HEJ was similar to that for the lower l

joint.

l The HEJ was subjected to fatigue loading cycles and temperature cycles to l

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 1

the fatigue load used in testing was that which was caused by loading /

I 1

1 unloading. Hence, it was judged necessary to determine that the leak rate at.

static and fatigue conditions were comparable.

The upper HEJ specimens were 1

also subjected to the loadings / deflections caused by & 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 sittulated a secondary hydrostatic test. An extended operation period test was also

)

performed.

]

l 3.3.5.3 RESULTS OF VERIFICATION TESTS FOR THE UPPER HEJ l

l l

~

The test results are presented in Tables 3.3.5.3-1 to 3.3.5.3-5.

I i

4637M/102787:49 -65 3-46

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MEJ Sessimons Mr the Am % Tm Tigure 3.3.5.1-2 3-47 l

a

I As can be seen from Table.3.3.5.3-1, the HEJ's formed out-of-sludge, i.e.,

in air, had an average initial leak rate of approximately [

]b,c.e at the normal operating. condition of 600*F and 1600 psi. After five years of simulated normal operation due to 5000 fatigue cycles and 29 to 32 temperature cycles, the leak rate was [

]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 g

),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 operating conditions was measured for four out-of-sludge HEJs.

[

]b,c.e These same four s'pecimens were then subjected to the SLB temperature, pressure and axial load conditions.

[

]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. C

)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 3.3.5.3-4, was the reference process per Table 4.0-1 except that the l

l 4637H/102787:49 -57 3-48

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,j 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 (

.]b,c.e After exposure of the specimens to five years of simulated normal oneration due to fatigue and temperature cycling, the average leak rate remained very low, I

bl c.e at the 600*F and 1600 psi condition.

(

The results of the reverse pressure test for the in-sludge upper HEJs are also shown in' Table 3.3.5.3-3.

['

.]D'C It was also zero for the simulated secondary side hydrostatic pressure test.

]

Table 3.3.5.3-4 also contains data for HEJs formed in-sludge.

It includes the i

l 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 Tables 3.3.5.3-1 and 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 untti a compressive load of 2950 lbs was reached. However, the two leak rates at 2950 lbs were low, [

]b c.e for specimens Number PTSP-23 and PTSP-33, respectively.

In 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 l

4637M/102787:49 -68 3-49

~

As shown in Table 3.3.5.3-4, the average leak rate for four in-sludge specimens during the SLB test was [

l 3,c.e a

The test data generated for the Alloy 690 samples is presented in Table 3.3.5.3-5.

The following observations were noted:

3,b,c were found at initial leak a

Specimen S-5 (Alloy.690):

[

testing at room temperature (R.T.).

At 600*F, the leak rates reduced J,b c during a a

significantly and remained below [

subsequent thermal cycling test.

This specimen was formed with a. tube diametral bulge that was smaller than will probably be used in the field.

Specimens S-8 (Alloy 690); B-4, B-6, and B-7 (Alloy 625/690 - 0.740 in.

l Sleeve Ola.), and BA-11 (Alloy 625/690- 0.630 in. Sleeve Dia.): These five specimens all exhibited moderate to small or very smal' leaks, mostly during the initial leak testing at R. T.

In all cases, by the end of the l

testing, including thermcl cycling and fatigue in some cases, the leak rates had reduced to zero (or near zero), illustrating the self healing characteristic of rolled joints.

Specimen BA-1 (Alloy 625/690, 0.630 Sleeve Dia.):

This specimen exhibited zero leak rate at initial testing, both R.T. and 600 F.

Small leak rates were found at R.T. after fatigue testing; however, they reduced to very small values, less than 0.5 drops / min, after testing.

This specimen was formed with a tube diametral bulge at the 1.ow end of the probable i

field range.

l 3.3.6 TEST PROGRAM FOR THE FIXED / FIXED MOCKUP 3'.3.

6.1 DESCRIPTION

OF THE FIXED / FIXE 0 MOCKUP 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 j

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TABLE 3.3.5.3-~3 (cont)

TEST RESULTS FOR HEJ'S FORMED IN SLUDGE (FATIGUE AND REVERSE PRESSURE TESTS INCL.)

(CONT) 1

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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 locations.

There were thirty-two tubes in two clusters of sixteen. A sludge simulant composed of alumina was formed around or,e cluster of sixteen.

Sleeves thirty inches long were installed in the tubes by [

3.a.c.e 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 Alloy 690 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 1.:stallation 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 /FlXED 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. All of the room temperature initial leak tests produced [

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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 duplicated the most stringent structural loading conditions for sleeves.

Therefore, it was concluded that all of the testing with short spec! mens 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, l

l 3.3.7 EFFECTS OF SLEEVING ON TUBE-TO-TUBESHEET HELD 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 l

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.

l Evaluation of the 0.750 inch tubes showed no tearing or other degrading effects on the we.ld after hard rolling.

Therefore, no significant effect on the tube-to-tubesheet weld is expected for the larger 0.875 inch 00 tube l

configuration.

l I

J 4746M/102387:49 -82 3-63

i a,c.e f

Figure 3.3.6.1-1

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Fixed Fixed Mockuo - HEJ 4746M/101587: 49 -83 J

1 4

3.4 ANALYTICAL VERIFICATION j

l i

3.

4.1 INTRODUCTION

This section contains the structural evaluation of the sleeve and tube I

a assembly with HEJ, sleeve material Alloy 690 and sleeve length [

J.c.e in relation to the requirements of the.ASME Boiler 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-676380, 9/20/66, Revision 1 (Reference 2).

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. A detailed stress evaluation for the upper joint is addressed in this section.

Structural analysis of the lower joint is presented in Section 3.5.

The tolerances used in developing the models were-such th't the maximum sleeve and tube outside diameters were evaluated in a

combination with the minimum sleeve wall thickness.

This allowed maximum stress levels to be developed in the roll transition regions.

1

1) Sleeve Material Alloy 600 is considered in Section 3.5.

3-65 4746M/102387:49 -84

l l

1

~

a,c.e

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i Figure 3.4.2-1 Hybrid Expansion Upper Joint / Roll Expanded Lower Joint Sleeve Configuration-4746M/101587: 49 -85

l 3.4.3 MATERIAL PROPERTIES I

The sleeve material is Alloy 690 described in ASME Code Case N-20 (Reference 3).

The tube material is S8-163 (Alloy 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 Boiler and Pressure Vessel Code,Section III, Appendix 1 (Reference 4) and Code Cases (Reference 3).

3.4.4 CODE CRITERIA The ASME Code Stress Criteria which must be satisfied are given in Tables 3.4.4-1 through 3.4.4-4.

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 Alloy 600 is considered in Section 3.5.

4746M/102387: 49 -86 3-67

Table 3.4.4-1 CRITERIA FOR PRIMARY STRESS INTENSITY EVALUATION (SLEEVE) a c.e 1

3-68 4746M/101587:49 -87 J

I Table 3.4.4-2 CRITERIA FOR PRIMARY STRESS INTENSITY EVALUATION (TUBE)

.a.c.e i

l l

l 4746M/101587:49 -88 3-69

TABLE 3.4.4-3 CRITERIA FOR PRIMARY PLUS SECONDARY AND TOTAL STRESS INTENSITY EVALUATION (SLEEVE) a,c.e i

i i

l 4746M/iO2387:49 -89 3-70

TABLE:3'4.4-4 CRITERIA FOR PRIMARY-PLUS SECONDARY-~

AND TOTAL STRESS INTENSITY EVALUATION (TUBE) a,c.e i

4746M/102387:49 -90 3 71 I.

t

d.

Maximum secondary to primary pressure differential - 670 psig, T = 650*F-2.

Full'Ioad steady state conditions are:

Primary side pressure 2235 psig Hot leg temperature - 616.8'F Cold leg temperature - 552.3*F Secondary side pressure - 705 psig-Feedwater temperature - 427.3*F Steam temperature - 506.3*F Zero load reactor coolant temper 3ture - 547.0*F Other operating conditions are specified in Tables 3.4.7.1-1 and 3.4.7.2-1.

3.4.6 METHODS OF 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 5) performs thermal and stress analyses of the structure. Pressure stress is calculated separately for a 1000 psi 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 determ16e the total pressure stress distribution.

Thermal analysis provides the temperature distribution needed for thermal stress calculations.

Thermal stress calculations are performed for fixed 4637M/102787: 49 -91 3-72

times under thermal transients.

These times for the total pressure and thermal analysis are chosen for the anticipated maximum and minimum total 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 6).

HECEVAL is a multi-purpose code which performs ASME Code,Section III, Subsection NB stress evaluations.

At any given point or section of the model, the program HECEVAL is used to determine 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 NB-3216.2.

3.4.6.1 H00EL DEVELOPMENT

{

A finite element model was developed for evaluating the sleeve design.

Some j

significant considerations in developing the model are:

J 1.

The model has been divided in two parts: upper model and lower.

1 model.

Structural it;tegrity 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 regicns were achieved by coupling the interface raodes in the radial direction.

For conservatism, locations of contact in tne 4637M/102787:49 -92 3-73

s' 1

I' sleeve-tube interfaces along the upper hard roll region contain I

elements which share nodes. This approximates a rigid fix by the rolling process involved. Additional axial coupling was effected also for the lower sleeve-tube and tube-tubesheet interface nodes.

I 3.

The interface nodes along the upper and lower hydraulic expansion regions of the HEJ were coupled in the radial dira: tion for temperature and thermal strest runs.

In the cases when pressure may I

penetrate into the interface, the interface nodes along these areas were disconnected for pressure stress runs.

4.

By varying the boundary conditions at a specified region of the model, conditions of either intact tube or discontinuous tube were simulated.

1 The element types chosen for the finite element analysis were the following HECAN (Reference 53 elements:

a,c.e l

4 i

All'the element types are quadratic, having a node placed in the l

center of each surface in addition to nodes at each corner.

{

3 l

1 I

l 462?M/101287:49 -93 3-74

)

3.4.6.2 THERMAL ANALYSIS 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:

1 Small step load increase 1

Small step load decrease Large step load decrease Hot standby operations i.ossof'1 cad Loss of power Loss of secondary flow Reactor trip from full power 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.

In 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 l

and secondary fluids is a maximum) and therefor.e, only the hot leg s'leeves were considered.

This condition bounds 'the thermal stresses on the cold leg.

4637M/102787:49 -98 3-75

7 i

t 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 close to the periphery of the bundle that it

. experiences the water temperature exiting the downcomer.

I Special hydraulic and thermal analysis was performed to define the primary and secondary side fluid temperatures and film coefficients as a function of time. Both boiling and convective heat transfer correlations were taken into consideration.

3.4.6.3 STRESS ANALYSIS

'A WECAN (Reference 5) 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 demonstrat'ing the applicability of a linear elastic an,alysis, thermally induced and pressure induced stresses were. calculated separately and ther, combined to determine the total stress distribution using the WECEVAL computer program (Reference 6).

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 to the lo'ading condition considered in order to determine the total pressure

~

stress distribution.

4637M/102787:49 -95 3-76

1

-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:

i Primary pressure - tube intact Primary pressure - tube discontinuous Secondhry 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 stresses were determined for each steady state solution as well as for the thermal transient solutions at those times during the thermal trcnslent which were anticipated to be limiting from a stress stan'dpoint.

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:

Pori total " 100' ' I } unit primary pressure P

sec. I } unit secondary pressure

+

1000

+(

) thermal This procedure was performed with the program HECEVAL (Reference 6).

l 4637M/102787:49 -96 3-77

I s

Stress and Fatique Evaluation

)

I Stress and fatigue evaluation were completed using the program HECEVAL The program HECEVAL performed the primary stress intensity l

(Reference 6).

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 HECEVAL determined the total. stress distribution for a loading condition considered and categorized that total distribution per the Subsection NB requirements.

That is, the

. total stress for a given cross section through t% thickness is categorized i

into membrane, linear bending, and non-linear components.'

These categorized stresses were then compare:. to the Subsection NB allowables.

1 In addition, complete transient histories at given locations on the model were used to calculate the total cumulative fatigue usage factor per Code Paragraph NB-3216.2.

For the fatigue evaluation, the effect of local discontinuities was considered.

~

O 3.4.7 RESULTS OF ANALYSES Analyses were performed for both intact and discontinuous tubes.

Design 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 effect and frequency of occurrence.

Fatigue and stress analyses of the sleeved tube assembly have been completed in accordance with the requirements I

of the ASME Boller and Pressure Vessel Code,Section III.

I 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.

4637M/102787:49 -97 3-78

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 thel sleeved tube assembly are well within allowable ASME Code limits.

U The largest value of the ratio " Calculated Stress Intensity / Allowable Stress Intensity" of (

j,b,c a

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 va. lues for the sleeved tube assemblies are summarized in Table 3.4.7.2-2.

The requirements of-the ASME Code, Paragraph NB-3222.2, were met at all' locations.

3.4.7.3 RANGE OF TOTAL STRESS INTENSITIES Based on the sleeve design criteria, the fatigue analysis considered a design I

life objective of 40 years for.the sleeved tube assemblies.

Table 3.4.7.2-1, describes the umbrella transient conditions used in the fatigue analysis.

Because of-possible opening of the interface between the sleeve and the tube along the hydraulic expansion regions, the maximum fatigue strength reduction factor of 5.0 (NB-3222.4(3)) was applied in the radial direction at the " root" interface nodes of the hard roll region.

s, 4637M/102787:49 -98 3-79

The results of the fatigue analysis for the sleeved tube assemblies are summarized 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 i

i l

1 1

l i

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l 4637M/102787:49 -99 3-80

TABLE 3.4.7.1 1 UMBRELLA PRESSURE LOADS FOR DESIGN, FAULTED, AND TEST CONDITIONS a.c.e

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3.4.8. REFERENCES 1.

ASME Boller and Pressure Vessel Code,Section III, Subsection NB, 1983 Edi tio,n, July ' 1, 1983.

2.

Equipment '.gecification G-676380, Westinghouse, Revision 1, September 20, 1966.

3.

ASME Boiler and Pressure Vessel Code, Code Cases, Case N-20, 1983 Edition, July 1, 1983.

4.

ASHE Boller and Pressure Vessel Code,Section III, Appendix 1, 1983 Edition, July 1, 1983.

5.

WECAN, WAPPP and FIGURES II, F. J. Bogden Editor, Second Edition, May

'1981. Westinghouse Advanced System Technology, Pittsburgh,' PA 15235.

l1

'6.

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.

l 1

I i

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'4638M:49/102787-1 3-88

3.5 SPECIAL CONSIDERATIONS 3.5.1 FLOW SLOT HOURGLASSING Along the tube-lane, the tube support plate has several long rectangular flow l

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 ON 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,e.f a

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 Inter Granular Attack (IGA) was detected by destructive examination.

It is to be noted that thermally treated Alloy 600 and 690 have additional SCC resistance compared to the resistance of mill annealed Alloy 600 tubing.

3.5.1.3 EFFECT ON MAXIMUM RANGE OF STRESS INTENSITY AND FATIGUE USAGE FACTOR In addition to the above two considerations, one should also consider the effect of the hourglassing induced bending stresses on max ~qum range of itress intensity and fatigue usage factor of the sleeve.

Taking into account the neurglassing induced bending stress along with the transient oressure and 4638M:49/102787-2 3-89

1' i

l thermal str'ess, the largest value of maximum stress intensity would be 59.70 KSI (allowable 79.80 KSI), fatigue usage factor is negligible.

I 3 5.2 TUBE 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 cn the natural frequencies, amplitude of vibration, and bending stress due to installation of various lengths of sleaves.

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 l

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, th6 tubesheet is isothermal at the bulk temperature of the primary fluid.

The net effect of the sludge is to reduce tube /tubesheet thermal effects.

3.5.4 ALLOWABLE SLEEVE DEGRADATION i

3.5.4.1 MINIMUM REQUIRED SLEEVE THICKNESS 1

The minimum required sleeve wall thickness, t, to sustain normal and r

accident condition loads is calculated in accordance with the guidelines of j

Regulatory Guide 1.121, as outlined ln Table 3.5.4-1.

In this evaluation, the surrounding tube is assumed to be completely degraded; that is, no design credit is taken for the residual strength of the tube.

The sleeve material may be either thermally treated Alloy 600 or-thermally treated Alloy 690.

It has been shown that the properties of Alloy 600 are

)

very similar to those of Alloy 690.

In particular, the yield strength'and ultimate strength are very similar.

l l

4638M:49/102787-3 3-90

Since Regulatory Guide 1.121 is to be. addressed, it is permissible to derive the allowable stress limits based on expected lower bound material properties, as opposed to the Code minimum values.

Ex'pected 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 greater than LTL.

l' L

3.5.4.2 DETERMINATION OF PLUGGING LIMITS l

{

p The minimum acceptable wall thick. ness and other practices in Regulatory Guide 1.121 are used to determine a plugging limit for the sleeve.

This Regulatory Guide was written to provide guidarce for the' determination of a plugging limit for steam generator tubes undergoing localized tube wall thinning and l

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, a

~

As recommended in paragraph C.9..b. of the Regulatory Guide, an additional 4

thickness degradation allowance must be added to the minimum acceptable tube wall thickness to establish the operational tube thickness acceptance for I

continued service. Paragraph C.3.f. of the Regulatory Guide specifies that the basis used in setting the operational degradation allowance include the I

method and data used in predicting the continuing degradation and consideration of eddy current measurement errors and other significant eddy current testing parameters.

I 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 determination of the operational tube thickness acceptable for continued service and thus determination of the plugging limit.

4638Mt49/102787-4 3-91 L

-.__._.___-___-_____.m

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Paragraph C.3.f of the Reg. Guide specified that the basis used in setting the l

operational degradation analysis include the method ano data used in

~

predicting the continuing degradation.

To develop a value for cont.nuing i

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 relat've to the tube and the lower heat flux due to the double wall in the sleeved region.

It is the position of Westinghouse Electric that since no degradation has been detected in the sleeves, presently any allowance for continuing degradation

[

]c,e would be an arbitrary value not supported by the data and would represent a conversatism in addition to the safety factors implicit in the determination of minimum acceptable tube wall thickness using Reg.

Guide 1.121 recommendations.

In summary, the operational tube t^nickness acceptable for continued service includes the minimum acceptable tube wall thickness ([

]D'C of wall thickness, see Table 3.5.4-1), the combined allowance for eddy current a

uncertainty and operational degradation ([

J,c of wall thickness as recommended by Westinghouse). These terms total to 59% resulting in a plugging limit as determined by Regulatory Guide 1.121 guidelines of 41% of the tube wall thickness.

The plugging limit 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.

3.5.4.3 APPLICATION OF PLUGGING LIMITS Sleeves or tubes which have eddy current indications of degradation in excess of the plagging limits must be repaired or plugged.

Those portions of the l

tube and the sleeve (shown in Figure 3.5.4-1) for which indications of wall

'~

degradation must be evaluated cre summarized as follows:

4638M:49/102787-5 3-92

n 1)

Indications of degradation in the entire. length of the sleeve must be j

evaluated against the sleeve plugging limit.

)

2)

Indication of tube degradation of any type including a complete gullio, tine 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 I

sleeve plugging limit applies to these areas also.

4)

The tube plugging limit continues to apply to that portion of the tube above the top of the upper joint.

l 4638M:49/102787-6 3-93

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r p 3.5.4-1 Appucaten of Mugeng Lamets a The plugging timtt for the tube is 40% of tube watt, the olvgging timJt for the sleeve is 31% with a conservative allowance for continued degradation, see Section 3.5.4.2.

3-94

I i

Table 3.5.4-1 REGULATORY GUIDE 1.121 CRITERIA 1.

Normal'and Voset Condition loadings

~

Normal Operations Criterion: Su 1 90.58 ksi Loading:

P - 2250 psia

!p P - 720 psia AP - 1530 psi j

3 Hence, miminum required sleeve wall thickness t IS r

AP. R'

~

r"S

=[

] inch C

t

[-0.5(Pp+P) 3 which is [" ) a,c.e percent of the nominal wall thickness.

{

Upset Conditions l,

Criterion: S - 39.59 ksi j

P - 2775 psia p

P 1180 psia AP - 1595 psi 3

I aP. R

-[

] inch C

lience, tr"S - 0.5 Pp+P) y s

which is [

la.c.e percent of the nominal wall thickness.

L 2.

Accident Condition Loadings a.

LOCA + SSE The major contribution of LOCA and SSE loads is the bending stresses at the top tube support plate due to a combination of the support motion, inertial loadings, and the pressure differential across the tube U-bend resulting from the rarefaction wave during LOCA. Since the sleeve is located below the first support, the LOCA + SSE bending stresses in the sleeve are quite small. The governing even*

1

~

for the sleeve therefore is a postulated secondary side blowdown.

4638M:49/102787-8 3-95 1

a

1

.I Table 3.5.4-1 (cont.)

~

b.

FLB + SSE T.he maximum primary-to-secondary pressure differential occurs during a postulated feedline break (FLB) 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 membrane stresses.

Criterion:

P, 1 smaller of 0.75 r 2.45,i.e. 63.4 ksi u

Loadings:

P - 2650 psig p

P

• O AP - 2650 3

Hence, [

]

8'C

or, [.

] a,c.e percent of nominal wall.

The. required sleeve wall thickness is [

lC.

51 percent minus growth and uncertainty, could be the

~

plugging criteria with confirmation of leak-before-break. A 40 percent criteria would permit 11 per cent for growth and uncertainty.

3.

Leak-Before-Break Verification The leak-before-break evaluation for the sleeve is based on leak rate and burst pressure test data obtained on 7/8 inch OD x 0.050 inch wall and 11/16 inch 00 x 0.040 inch wall cracked tubing with various imounts of uniform thinning simulated by machining on the tube 00. The margins to burst during a postulated SLB (Steamline Break Accident) con'dition 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 factor of 9.5 for the nominal sleeve, the current Technical Specifications allowable leak rate of.35 gpm, a SLB pressure differential of 2560 psi, and the nominal leak and nominal 4638M:49/102787-9 3-%

burst curves, a 29.8 percent margin exists between the burst crack length and the leak crack length.

For a sleeve thinned 51 percent through wall

~

l over a 1.0 inch axial length, a 24.8 percent margin to burst is demonstrated.

Thus the leak-before-break behavior is confirmed for unthinned and thinned conditions.

3.5.5 EFFECT OF TUBESHEET INTERACTION i

Since the pressure is normally higher on the primary side of the tubesheet than on the secondary side, the tubesheet becomes convex upward. Under these conditions, the tubes protruding from the top of the tuDesheet will rotate

~ from the vertical.

This rotation develops stresses in the sleeved tube assembly. Analysis performed showed that these stresses are not large enough to affect significantly the fatigue usage factors already found.

l 3.5.6 STRUCTURAL ANALYSIS OF THE LOWER HYBRID EXPANSION JOINT 3.5.6.1 Primary Stress Intensity The results of primary stress intensity evaluation for the analysis sections located at the lower hybrid expansion joint are summarized in Tables 3.5.6.1-1 and 3.5.6.1-2.

All primary stress intensities for the sleeved tube assembly at the lower hybrid expansion joint meet the ASME code limits.

3.5.6.2 Range of Primary Plus Secondary Stress Intensities Primary plus secondary stress at the Lower Hybrid Expansion Joint are developed by the pressure acting on the sleeve, tuce and tubesheet ligament surfaces (primary stress), and by thermal stress and deformations imposed by the tubesheet motion (secondary stress).

The tubesheet motion results from the primary and secondary side pressure and l

interactions among the tubetheet, support ring, channel head, and the stub barre 1.

4638M:49/102787-10 3-97

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tube intact was analyzed.

The~ maximum range of stress i

I intensity values for the sleeved tube assembly are summarized in Table 1

3.5.6 2-1.

The requirements of the ASME Code, paragraph NB-3222.2 were satisfied.

3.5.6.3 Range of Total Stress Intensities I

The fatigue analysis considered a design life objective of 40 years for the sleeved tube assemblies.

The maximum 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 negligible, hence, they are below the allowable value of 1.0 specified in the ASME Code.

4638M:49/102787-13 3-100

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3.5.7 EVALUATION OF OPERATION WITH FLOW EFFECTS DUE TO SLEEVING An ECCS performance analysis has been completed for Point Beach Unit 2.

This safety analysis assumed a specified level of Model 44 steam generator tube plugging (SGTP) Ir. the steam generators and established a 2.21 value of total core peaking factor for Unit 2.

The corresponding non-LOCA safety analysis defines a more limiting (lower) tube plugging level in order to maintain Thermal Design Flow (TDF).

Therefore, the limiting plugging condition which 1s applicable to this evaluation of the flow effects due to sleeving is the non-LOCA value which has been established to maintain thermal design flow.

The accidents evaluated in this report include LOCA and non-LOCA transients as

~

well as consideration of the effects on the nuclear design and thermal-hydraulic performance witn the existing 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, DNBR, RCS pressure, etc.).

For the Model 44 steam generators in Unit 2, maintenance of thermal design flow has been determined to bi c hievable with up to 391.2 tubes plugged in any one steam generator.

Inserting a sleeve into a steam gen.rator tube results in a reduction of i

primary coolant flow.

For the purposes of this discussion, it is assumed that up to 1750 slesves will be installed in any steam generator channel head.

This implies that an additional 250 sleeves can be installed in the hot legs, and the full 1750 sleeves car, be installed in the cold legs. The selected sleeving program at Point Beach assumes the use of [

]"'C inch long sleeves which are presumed to be long enough to span the degraded areas in the i

tubesheet region and to be above the sludge pile in either the hot or cold leg side of the steam generators.

l 3.5.7.1 ONE SLEEVE PER TUBE c

3,c.e inch sleeve installed in the hot leg of a tube, the a

For a single [

J.c.e j

a

~

primary coolant flow reduction per tube is approximately equal to [

percent of normal flow under normal conditions.

This reduction in primary 4638M:49/102787-15 3-102

]

L

)

coolant flow equates to a hydraulic equivalency ratio of C 3,c.e sleeved a

tubes to one plugged tube under normal conditions.

a Using this [

J,c.e to I ratio while maintaining TDF for Unit 2 (Model 44 steam generator), Table 3.5.7-1 can be used to determine the allowable sleeving parameters assuming installation of up to [

J,c.e sleeves per a

J,c.e sleeved a

steam generator under normal conditions. Note that C a3.c.e plugged tubes.

tubes are equivalent to (

For the condition presented above for Unit 2, the most limiting equivalent plugged tube condition in the two steam generators occurs in Steam Generator B where 147 tubes, including 89 sleeved tubes, are currently plugged.

It is seen in Table 3.5.7-1 that with (

]"'C tubes sleeved there would be a margin of (

J c.e tubes (391 minus [

a J,c.e) available for a

additional plugging before violating the TDF basis of the non-LOCA analysis.

With a smaller number of sleeved tubes, the margin to TDF would be larger.

3.5.7.2 TWO SLEEVES PER TUBE 3,c.e inch sleeve on the hot-leg side and a a

When a single tube has one [

second [

J.c,e inch sleeve on the cold-leg side, the primary coolant flow a

aJ.c.e percent of normal flow.

loss per tube is approximately equal to C This reduction in primary coolant flow equates to a hydraulle equivalency a

ratio of C J,c.e double sleeved tubes to.one plugged tube under normal conditions.

J.c.e to I ratio while maintaining TDF for Unit 2, a

Using this C Table 3.5.7-2 can be used to determine the allowable sleeving parameters assuming installation of up to C J c.e sleeves for [

a J,c,e tubes a

J c.e sleeves a

per steam generator under normal conditions.

Note that [

'a.c,e are equivalent to C J

plugs.

I l*

For the condition presented acove for Unit 2, the most limiting equivalent plugged tube condition in the two steam generators occurs in Steam Generator 8 where 147 tubes, including 89 sleeved tubes, are currently plugged.

It is A

seen in Table 3.5.7-2 that with [

J

'C

tubes sleeved, there would be a margin of (

J c.e tubes (391 minus C J,c.e) available for additional a

a 4638M:49/102787-16 3-103

plugging before violating the TOF basis of the non-LOCA analysis. With a smaller number of double sleeved tubes, the margin to TOF would be larger.

3.5.7.3-FLOW EFFECTS

SUMMARY

o' The effect of sleeving on the non-LOCA transient analyses has been reviewed.

Analyses of the-level of sleeving and plugging discussed in this report have shown that the Reactor Coolant System flow rate will not be less than.the Thermal Design Flow rate.

The Thermal Design Flow rate is the value used In the non-LOCA safety analyses and is designed to be less than the minimum RCS l

flow rate that occurt under normal or degraded conditions. Since the reduced RCS flow rate is not less than the assumed flow rate (Thermal Design Flow),

the non-LOCA safety analyses'are bounded by the anticipated maximum amount of l.c.e sleeves per steam generator) and a

steam generator tube sleeving ([

cause no safety concerns. Any smaller number of sleeves would have less of an effect.

It should be noted that any combination of sleeving and plugging may be utilized at Point Beach Unit 2 as long as the TOF is not violated.

Tables 3.5.7-1 and 3.5.7-2 give the number of tubes which may be sleeved up to

[

l"'C tubes and the number of additional plugs per steam generator that could be installed at the present plugging levels of Unit 2 without violating TDF, In addition, as a result of tube plugging and sleeving, primary side fluid velocities in the steam generator tubes will increase.

The effect of this velocity increase on the sleeve and tube has been evaluated assuming a conservative plugging condition which tends to increase flow velocity within a tube. As a reference, normal flow velocity through a tube is approximately

[

l c..e ft/sec, for the unplugged condition. Assuming the conservative a

plugging condition, the fluid velocity through an unplugged and unsleeved tube a

is [

l.c.e ft/sec, and for a tube with a sleeve, the local fluid velocity in the sleeve region is estimated at [

J"'C ft/sec.

Because n

these fluid velocities are less than the inception velocities for fluid impacting, cavitation, and erosion-corrosion, the potential for tube degradation due to tnese mechanisms is low, f

l 4638M:49/102787-17 3-104

)

_ - - _ _ =

\\

1 Accordingly, using the assumptions stated i' Section 3.5.7,.no ECCS results n

more adverse than those in the existing Westinghouse safety analyses are l

anticipated for equivalent tube plugging projected to occur at Point Beach' Uni t 2 wi th up to. [

]T,c.e tubes sleeved per steam generator using

[

]"',C sleeves.

I l

'l l

l i

l

.i i

i I

I I

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4638H:49/102787-18

.{

TABLE 3.5.7-1

~ ALLOWABLE SLEEVING PARAMETERS UNDER NORMAL CONDITIONS-(ONE SLEEVE PER TUBE) a,c.e 4

e,*

S e

e.

f..

I I

i l

l 4638M:49/101287-19 3-106 l.

Ar

l TABLE 3.5.7-2 ALLOWABLE SLEEVING PARAMETERS UNDER NORMAL CONDITIONS (TWO SLEEVES PER TUBE) a.c.e I,

1 1

l i

~

l 4

f--

J ia E

I.

1 l

l 3-107 4638M:49/101287-20

3.5.8 ALTERNATE SLEEVE MATERIALS I

As mentioned above in Section 3.4.3, mechanical properties of Alloy 690 are considered by the ASME Code Case N-20.

The design stress intensity value S, fo,r both materials Alloy 690 and 600 in the Case N-20 is identical (S,- 26.6 ksi). Therefore, for these materials, Primary Stress Intensity and Maximum Range of Stress Intensity allowables are equivalent. 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.

Therefore, because of the similarity of design stress intensity values, the ratio " Calculated Maximum Primary Membrane Stress Intensity / Allowable Stress Intensity" is not a function of the sleeve material (Alloy 690 or 600).

Thermal stress, and hence maximum range of stress intensity and fatigue usage factor, could depend upon the mechanical properties of the sleeve material.

The modulus cf elasticity of Alloy 600 is higher than that of Alloy 690 by approximately 7 percent.

However, the coefficient of thermal expansion of Alloy 690 is highar than that of Alloy 600.

Past analytical experience indicates that the coefficient of thermal expansion. mismatch between Alloy 69p and 600 materials significantly effects the thermal stress.

Therefore, the results of the Maximum Range of Stress Intensity and Fatigue Evaluations for the sleeve and tubs assembly with Alloy 690 as the sleeve material are conservative relative to those which q

would be calculated with Alloy 600 as a sleevs material.

3.5.9 EFFECT OF AN AXIAL TUBE LOCK-UP ON FATIGUE USAGE FACTOR i

In this analysis, only one tube is considered to be locked-up at the first i

tube support plate under 100 percent power conditions.

i The following effects on the stress components of the locked-up tube were

]

analyged:

j effect of primary and secondary pressure stresses effect of thermal stresses in the assembly l

effect of tubesheet rotations

~

effect of axial thermal displacements in tube, tube / sleeve, and wrapper /shell regions 4638M:49/102787-21 3-108

1' l

effect of pressure drop through the tubesheet and tube support slates 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 L.

outermost tubes.

The results of maximum range of stress intensity and fatigue evaluations are given in Tables 3.5.9-1 and 3.5.9-2 L

'For the central locked-up tubes, only the sleeve for the worst case, i.e.,

tube discontinuous, was considered.

It is seen that the requirements of the ASME Code are satisfied for both outermost and central axial locked-up sleeved tubes.

3.5.10 Minimum Sleeve Wall Thickness i,

Nominal and minimum sleeve wall thickness'was knalyzed.

Taking into 1ccount plus 0.003 inches for corrosion /errosion, the recommended sleeve wal. ;hickness is:

Nominal Sleeve Wall Thickness 0.037 inches Minimum Local Sleeve Wall Thickness 0.0363 inches l

I*

I 463BM:49/102787-22 3-109

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4.0 PROCESS DESCRIPTION The sleeve installation consists of a series of steps starting with tube end prepart. tion (if required) and progressing through sleeve insertion, hydraulic

~

expansion at both the lower joint and upper Hybrid Expansion Joint (HEJ) regiores, 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 Thisre are two steps involved in preparing the steam generator tubes for the sleevin'g' operation.

These consist of light rolling (as required) at the tube end and tube honing.

4.1.1 TUBE END ROLLING (CONTINGENCY)

If gaging of tube inside diameter measurements indicate a need for tube end rolling to provide a uniforrn 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 a much higher torque showed no effect on the tube-to-tubesheet weld.

Because the raditi forces transmitted to the tube-to-tubesheet weld would be much lowe for a larger Model 44 sleeve than for the above test configuration no effect on the weld as a result of the light roll is expected.

4638M:49/102787-25 4~I

TABLE 4.0-1 SLEEVE PROCESS SEQUENCE

SUMMARY

?

a,c.e i

t l

4638M:49/101287-26 4-2

/

\\

4.1.2 TUBE HONING l

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 from 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 selection of the honing process used is dependent primarily on the installation technique utilized, the scale of the e

sleeving operation (small scale vs. large scale sleeving), and the customers site specific rad-waste requirements.

Evaluation has demonstrated that neither of these processes remove any significant fraction of the tube wall base material.

4.1.2.1 WET HONING Tube honing will be performed using a [

3,c,e a

A waste handling system may be used to collect the [

],a.c.e 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, 4638M:49/102787-27 4-3

4.1.2.2 DRY 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 utilizing a fluid (typically detonized 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 INSER', ION AND kXPANSION 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.

l j c.e a

4638M:49/102787-28 4_4

[

l l

1 aJ.c.e This process is repeated until all sleeves are installed and hydraulically expanded.

4.3 LOWER _ JOINT SEAL At the primary face of the tubesheet, the sleeve is joined to the tube by a

[

),a,c.e The contact forces between the sleeve and tube due to the initial hydraulic expansion are sufficient to keep the sleeve from rotating during the-[

e

},a Cee l

The: appropriate extent of hard roll expansion of the sleeve is attained by aJ.c.e The hard roller

[

torque is calibrated on a standard torque calibrator prior.to initial hard' rolling operations and subsequently recalibrates at the beginning of each shift for automatic tcoling.

This control and calibration process is a technique used thrcughout industry in the installation of tubes in heat exchangers.

1 4638M:49/102787-29 4-5

4.4 UPPER HYBRID EXPANSION JOINT (HEJ)

The HEJ first utilizes a C aJ.c.e 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 hydraulic and roll expansions.

The basic process check on 100 percent of the sleeved tubes will be:

1.

Verify presence of lower hydraulic expansion zone.

2.

Measure lower hydraulic expansion and roll average diameter and verify location within the lower hydraulic expansion.

3.

Verify presence of upper hydraulic expansion zone.

4.

Measur, 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 initial application of the eddy current profilometry may be performed to obtain sleeve ID data. As acceptable diameters are verified and the sleevi.ng process is proceeding as anticipated, this inspection may be eliminated.

These average diameters will be evaiuated versus tne expected tolerances established through the design requirements, laboratory testing results, and previous experience.

This evaluation will determine whether or not the equipment / tooling is performing satisfactorily.

If process data is determined to be outside of expected ranges, a non-conformance repcrt is i5 sued and further analysis performed, i

4638M:49/102787-30 4-6 l

L------------------------

If required, Diatest may be used in lieu of eddy current to perform sleeve installation acceptance and in-process monitoring evaluations. Unders12 ed diameters will be corrected by an additional expansion step to produce t.1e desired degree of expansion. Oversized diameters will be dispositioned bj 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.

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 (

J.a.c.e Sleeve (

l,c.e is determined by [

a J.a.c.e Accordingly, rolling torque was varied to J,c.e in the original Model 44 a

achieve the desired sleeve (

program (also applicable to the model 51).

(

J.c.e was achieved was used throughout the program a

verification testing.

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 [

~

p,c.e 4638M:49/102787-31 4-7

[

3

'C

(Refer t 8

E Section 3.3.5.3 for an additional discussion of the roll expansion torque for the in-sludge case.)

l In the first sleeving project performed by Westinghouse, hydraulic expansion axial length was also evaluated.

[

3.C

Therefore, in later programs, the HEJ hydraulic expansion: axial length (

l 3,b,c.e a

4638M:49/102787-32 48

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 tooling is not critical so long as it does not affect the sleeve installation processes, i

Some techniques used to position the sleeve installation tooling are: fully robotic (ROSA and SM-10HS) and hands-on (manual), or the combination of two or more tooling installation modes.

The actual technique utilized is dependent upon many variables including what is mutually agreed to between the utility and Westinghouse.

a 4

l j

.g.

5-1 4638M:49/102787-33

I 6.0 NDE INSPECTABILITY I'

The Non-Destructive Examination (NDE) development effort has concentrated on l,

two aspects of the sleeve system.

First, a method of confirming that the joints meet critical process dimensions is required. Secondly, it must be l

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 relied upon eddy current technology.

l 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 EDDY 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 s' team generators in accordance with the plant's Technical Specifications.

The purpose of these inspections a

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 resuce the potential for significant primary-to-o e ndary leakage.

The standard inspection procedure involves the use of a bobbin eddy current probe, with two circumferentially wound coils which are dispiaced axially along the probe body.

The coils are connected in the so-called differential mode; that is, the system responds only when there is a difference in the properties of the material surrounding the two coils.

The coils are excited by using an eddy current instrument that displays :hanges 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.

I 6-1 4638M:49/102787-34

The outputs of the various frequencies are combined 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, enhanced inspectability of the tubing results (i.e., a higher signal-to-noise ratio).

Regions in the steam generator such as the tube supports, the tubesheet, and sleeve transition zones are examples of areas where multifrequency processing has proven. valuable in enhancing inspection of the tube.

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 l

all subsequent inspections will be compared.

While there are e number of probe configurations that lend themselves to enhancing the inspection of the tube / sleeve assembly in the regions of configuration transitions, the crosswound coil probe has been selected as offering a significant advancement over the conventional bobbin coil probe, yet retaining 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 l

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 illustrates the coil response and measurement technique fer typical sleeve / tube joint.

The inspection for degradation of the tube / sleeve assembly has typically been performed using crosswound coil orobes operated with multifrequency excitation.

For the straight length regions of the tube / sleeve assembly, the inspection of the sleeve and tube is consistant with normal tubing j

inspections.

In tube / sleeve assembly joint regions, data evaluation becomes

~

more complex.

The results ciscussed below suggest the limits on the volume of degradation that can be dete'cted in the vicinity of geometry changes.

4638M:49/102787-35 l

I i

The detection and quantification of degradation at the transition regions of the sleeve / tube assembly depends upon the signal-to-noise ratio between the

~

degradation response and the transition response. As a general rule, lower.

frequencies tend to suppress the transition signal relative to the degradation

]

signal at the expense of the ability to quantify. Similarly, the inspection of the tube through the sleeve requires the use of low frequencies to achieve l

i detection with an associated loss in quantification.

Thus, the search for an optimum eddy current inspection represents a trade-off between detection and i

quantification.

With the crosswound coil type inspection, this optimization leads to a primary inspection frequency for the sleeve on the order of [

l,c.e and for the tube and transitio'n regions on the order of (

a 3,c.e a

a Figure 6.1-2 shows a typical [

J,c.e phase angle versus degradation depth curve for the sleeve from which OD sleeve penetration.s can be assessed.

In the regions of the parent tube above the sleeve, conventional bobbin coil or crosswound coil inspections will continue to be used. However, since the l

diameter of the sleeve'is smaller than that of the tube, a probe inserted through the sleeve will have a reduced fill factor.

Eddy current inspection for the unsleeved portion of the steam generator tube having sleeves at both ends is accomplished by utilizing a.610 dia. bobbin test probe. This probe size has been successfully utilized to inspect dented tubes, small radius U-bend tubes, and tube sleeve assemblies, both baseline and in service inspecitons. Calibration is performed as per ASME B & PV code Section V Article 8 - Appendix I to allow for a phase analysis correlating a percent of wall loss to a specific phase angle. Utilization of the MIZ-18 Digital Data Acquisition equipment provides the necessary sensitivity for achieving the required signal amplitude from the 207. flat bottom hole on the ASHE Defect Standard.

The phase analysis technique structures the' signal such that responses from tube degradation result in the vertical plane, and signals such as probe motion lie in the horizontal plane.

This signal separation decreases the effects of a reduced fill factor when utilizing a.610 dia.

probe.

Test configurations incorporate both the differential and absolute 4638M:49/102787-36 6-3

capabilities of the eddy current probe.

Figures 6.1.3a and b demonstrate the acceptance response of the 0.610 dia probe to the ASME Tube Standard when utilizing the MIZ-18 Digital Data Acquisition equipment.

For the tub.e 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 OD tube wall penetrations with acceptable signal-to-noise ratios at the transitions when the volume of metal removed is

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equivalent to the ASME calibration standard.

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The response from the tube / sleeve assembly transitions with the crosswound I

coil is shown in Figures 6.1-4, 6.1-5, and 6.1-6 for the sleeve standards, tube standards and transitions, respectively. Detectability in transitions is enhanced by the combination of the various frequencies.

For the cross-wound probe, two frequency combinations are shown; [

provides s

J,b,c.e Figure 6.1-7 shows the phase / depth curve a

for the tube using this combination. As examples of the detection capability at the transitions, Figures 6.1-8 and 6.1-9 show the responses of a 20 percent 00 penetration in the sleeve and 40 percent 00 penetration in the tube, respectively.

For inscection 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 additional inspectability has been gained by tapering the wall thickness at the top end of the sleeve.

This redu es the end-of-sleeve signal by a factor of approximately two.

The crosswound coil, howaver, again significantly reduces the response of the sleeve end.

Figure 6.1-10 shows the response of various ASME tube calibration standards placed at 1

'0

8 the end of the sleeve using the cross-wound coil and the (

frequency combination.

Note that under these conditions, degradation at the top end of the sleeve / tube assembly can be detected.

4638H:49/102787-37 6-4

6.2

SUMMARY

Conventional eddy current techniques have been modified to incorporate the more recent technolog~y in the inspection of the sleeve / tube assembly.

The resultant inspection of the sleeve / tube assembly involves the use of a cross-wound coil for the straight regions of the sleeve / tube assembly and for the transition regions.

The advent of MIZ-18 digital E/C instrumentation and its attendant' increased dynamic range and the availability of 8 channels for four raw frequencies has expanded the use of the crosswound coil for sleeve inspection. While there is a significant enhancement in the inspection of portions of the issembly using the cross-wound coil over conventional bobbin coils, efforts continue to advance the. state-of-the-art in eddy current inspection techniques. As advanced state-of-the-art techniques 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.

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4638M:49/102787-38 j

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o Ff9 m 6.1-1 Absolute eddy current signals at 400 khr (front and rear coils)

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a,c,e manusua t

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f Figure 6,1 2 i

(

3 Calibration Curve 6-7 1

AStf STANDARD 2-831 PRGE A-720 SF/RMF i

4638M:49/100)S7-41 Figure 6.I ~b 6-8

/&E ST/K#D Z-831 PROEE A-610 LC/RWGM a.c.e

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l 4638M:49/100187-42 Figure 6.1-3a 6-9

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84, '

• F1qure 6.1-4 1

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s E.C. $'ensis from tee ASTM Standartt Macninae en tee Slee' O.D. of tee StewTune Anemoty witnout Escansion (Ct:n Wouna Coil Proce) 4638M:49/100187 43 6-10

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Figure 6.1-5 3.c,e l

E.C. Signate %m the ASTM Stancare, Macnined on tMe T co C.Q. of tne $lowe Twee Anemely Witnout Essansson (C.cu Wound Coit Procol 4638M:49/100187-44 6-11

i Figure 6.1-6 a,c.e 1

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E.C. 5;gnals from the Escansen Transtion Region of t-e Twee Sieeve Assemoty (Crom Woune Cod Procol 4638M:49/100187-45 6-12

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l Figure 6.1-7

a. c. e 3

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Eddv Carrent Calibration Curve for ASME Tuee Stancare at

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! a.c.e see a Mix Usang tne Crom Woune Coil P'cce l

4638M:49/100187-46 6-13

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Figure 6.1-8 age l

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E.C. Signal frem a 20% Qese Hole. Half the Volums of ASI

Stancare. Macamec on tne Steere 0.0. in tne Excanoon Transition Region of tne Slee<* Tuoe Asserreiv :cres.veu :

r Coil Prece) 4638M:49/100187 47 6-14

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Figure 6.1-9 l

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s E.C. Signal from a 40% ASTM Stancare, Maoiined on tes Tuce 0.0. in tne Escanven Transstion Region of Sleev>Twee Asemety (Crem Woune Casi Prece) s4 4

A a638H:49/100187-48

Figure 6.1 -10 a,c.e l

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Eddy Current Response of the ASME Tube Standard at the End of the Sleeve Using the Cross Wound Coll Probe and Mult1 frequency Combination 4747M:49/101587-49 6-16

7.0 ALARA CONSIDERATIONS FOR SLEEVING OPERATIONS

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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 advancement of remote and semi-remote tooling, including state-of-the-art robotics; decontamination of steam generators; the use of shleiding 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 is 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, i. 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 t,een insta'11ed. A control station operator (CS0).then manually operate controls to guide the m nipulator arm through the manway and attach a baseplate to the tubesheet.

The insicllation 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 station trailer. 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.

i The control of personnel exposures can also be effected by careful planning, training, and preparation of maintenance procedures for the jcb.

This form of i

administrative control can help to provide that the minimum 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 I

i 4747M:49/102387-50 71

platform and channel head operations and the monitoring of personnel exposure to identify high exposure areas. Local shielding will be used whenever possible to reduce the general area background radiation levels at the work stations inside containment.

7.1 N0Z7LE' COVER AND CAMERA INSTALLATION / REMOVAL The installation of temporary nozzle covers in the reactor coolant pipe nozzles in c. reparation of the steam generators for sleeving operations may require u.annel 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 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 Robotic 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 potentiais requirements for channel head entries during the sleeving project.

7.2 PLATFORM SETUP /5UPERVISION I

The majority of the radiation exposures recorded for the sleeving program is expected to result primarily from personnel working on or near the steam generator platforms and in the channel head for hands-on operations.

The I

setup and checkout of equipment for the various sleeving processes, installation / removal of teoling, and the operation of the tooling are the 4747M:49/102387-51 7-2

setup and. checkout of equipment for the various sleeving processes, installation / removal of tooling, and the operation of the tooling are the major sources of radiation exposure.

In addition to channel head video j

monitoring systems, visual monitoring and supervision by one or more workers l

on the platform will be required for a major part of the sleeving schedule.

l 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 is spent on the platforms potentially giving rise to personnel exposures.' An evaluation of radiation surveys around the steam generators should indicate appropriate standby stations.

7.3 RADWASTE GENERATION The surface preparation of tubes for the installation of sleeves requires that the oxide film tc 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 l.c.e and the a

assemblies (optional), [

normal anti-C consumables associated with steam generator maintenance.

The anti-C consumables are the utility's responsibility and will not be addressed in this report.

For the [

]**C'" approximately thirty tubes can be honed before the hone is changed for process control and [

).a,c.e A typical estimate of the l

radioactive concentration from a honed tube transported by the [

8 1

'C

is given in Table 7.3-1.

These concentrations are based on a general area radiation level of 4R/HR.

The tube hones as well as the tubes J,c.e Consequently, a

[.

radiation levels of the spent hores are normally 1-2 r/hr based on field measurements in previous sleeving projects, l

4638M:49/102787-52 7-3

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TABLE 7.3-1 ESTIMATE OF RADI0 ACTIVE CONCENTRATION IN WATER PER TUBE HONED (TYPICAL) a,C,e l

ASSUMPTIONS

1) Tube honed 45 inches (in length)

2) Water flow rate of 0.6 gallons per tube honed

3) EssentiaT1y all radioactivity removed from tubes honed.

4747M:49/102387-53 74

The flexible honing cable used to rotate the hone inside the tubes is also flushed during the honing process.

However, the construction of the stainless steel cable will'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.

l 7.4 HEALTH PHYSICS PRACTICES AND PROCEDURES The Health Physics (HP) requirements for sleeving will be those estabished 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 maintainea so that appropriate 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 beneficial changes to the techniques are. mutually recognized but not covered a

in these HP procedures, appropriate changes will be made according to established change procedures.

The field service procedures which are prepared by Westinghouse for the comolete setup of equipment and subsequent sleeving operations include the specific radiologically related responsibilities, prerequisites and precautions.

These are expected to minimize further exposure and control contamination.

1 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.

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Identification of high radiation zones on the work platform and 1

emphasis of minimizing stay times, i

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l 4747M:49/102387-54 7-5

o' Handling of contaminated tools and changeout of contaminated mandrels.

o Location and use of waste disposal containers.

Westinghouse,e implements an extensive training and qualification program to

);

prepare supervisory, maintenance and operations personnel for field l

Implementation'of the sleeving process.

Satisfactory completion of this training program verifles that the trained personnel are qualified to perform all assigned operations from a technical as well as radiological 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 to demonstrate that a trainee has knowledge of the material presented.

This examination is written and all trainees tested. A minimum grade of 80 percent is required.

The test results are retained for audit.

Phase II - Consists of hands-on and mockup sleeving training during which the trainee must demonstrate 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.

fhe major operations include [

3"'C and sleeve installation.

4747M:49/102387-55 7-6

Experience has shown that these sleeving processes do not contribute 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 as"follows:

P = ((N 0)+S).N 3

3 g

g P - Project total exposure (MAN-REM)

N - Number of sleeves install'ed/ steam generator 3

D - Exposure / sleeve installed 3

S - Equipment setup / removal exposure per steam generator g

N - Number of steam generators to be sleeved g

l.

This equation and appropriate variations are used in estimating the total personnel exposures for the sleeving project.

1 474?M:49/102387-56 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 Service Arm (ROSA) or SM-10NS.

e.

Man-rem exposure results obtained from recent Westinghouse steam generator manual sleeving operations show approximately 300 man-rem for sleeving of 650 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.

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4747M:49/102387-57 7-8

8.0 INSERVICE INSPECTION PLAN FOR SLEEVED TUBES

.1 In addressing current NRC requirements, periodic inspections of the supplemented pressure boundary must be performed.

This new pressure boundary s

consists of,the sleeve with a joint at the primary face of the tubesheet and a joint at the opposite end of the sleeve.

The inservice in'spection program w111 consist of the following.

Each sleeved tube 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.

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Periodic pressure testing of the steam generator, similar to that performed following tube plugging will be performed as recommended in the technical manual.

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I 4747M:49/102387-58 8-1 J