ML20077N073
| ML20077N073 | |
| Person / Time | |
|---|---|
| Site: | Prairie Island  |
| Issue date: | 12/31/1994 |
| From: | WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19311B656 | List: |
| References | |
| SC-94-11-007, SC-94-11-7, WCAP-14226, NUDOCS 9501130287 | |
| Download: ML20077N073 (110) | |
Text
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i Exhibit E Prairie Island Nuclear Generating Plant License Amendment Request Dated January 9, 1995 WCAP-14226 4
F* and L* Tube Plugging Criteria For Tubes j
.With Degradation in the Tubesheet Roll Expansion Region of the Prairie Island Units 1 and 2 Steam Generators 1
NON-PROPRIETARY Prepared By Westinghouse Electric Corporation December 1994 i
501130287 9503o9 pDR p
ADOCK 05000282 DM
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WESTINGHOUSE NON-PROPRIETARY CLASS 3 WCAP-14226 SG-94-12-007 i
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F* AND L* 1UBE PLUGGING CRITERIA FOR TUBES WITH DEGkADATION IN THE TUBESHEET ROLL EXPANSION REGION OF THE PRAIRIE ISLAND UNITS 1 AND 2 STEAM GENERATORS December 1994 Work Performed Under Contract:
Purchase Order No. PG 4571 SQ Westinghouse General Order No. MP,27634 l
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l WESTINGHOUSE ELECTRIC CORPORATION i
NUCLEAR SERVICES DIVISION P.O. BOX 158 MADISON, PENNSYLVANIA 15663 i
C1994 WESTINGHOUSE ELECTRIC CORPORATION l
ALL RIGHTS RESERVED l
WPF/NSP/120994 l
ABSTRACT
. An evaluation has been perforiaed to develop repair criteria for tubes with degradation in the partial depth hardroll expanded region of the Prairie Island Units 1 and 2 steam generator j
tubes. A criterion based on maintaining structural adequacy of the tube to resist tube pullout forces, called F*, is developed. This criterion specifies a distance below the bottom of the.
1 roll transition, designated F*, in which no degradation is permissible; below the F* length, all 1
types of degradation may be left in service independent of the number or na;ure of eddy current indications..It is determined that the required F* length is 1.07 inches plus an allowance for NDE uncertainty.
Since tubes may.be found with degradation within the F* region, it is desirable to develop a j
criterion which will permit a significant number of such tubes to remain in service. ' A
-l criterion based on adequate axial strength and potential leakage (with the applicable length designated as L*) is developed. An L* value of 0.50 inch is established, such that certain types of tube degradation below the L* distance are acceptable for continued plant operation.
For application of L*, at least 1.34 inches (plus allowances for eddy current uncertainties) of sound, undegraded tubing are required below the bottom of the roll transition, with no more j
than two bands of axial degradation (crack angles s; 30 from tube axis) separating the nund portions of tubing. Below the required undegraded region, all types of degradation may be
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left in service independent of the number or nature of eddy current indications.
l The evaluation demonstrates that application of the F* and L* criteria for indications of tube degradation within the roll expansion affords a level of plant protection commensurate with that provided by Regulatory Guide 1.121 for degradation located outside of the tubesheet region.
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TABLE OF CONTENTS i
j Section.
Topic P.agm i
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1.0 INTRODUCTION
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1.1 Background
1-1 i
2.0 DEVELOPMENT OF F' CRITERION 2-1 2.1 Determination of Elastic Preload Between the Tube and Tubesheet 2-2-.
2.1.1 Radial Preload Test Configuration Description 2-2' i
2.1.2 Preload Test Results Discussion and Analysis 2-3 2.1.3 Residual Radial Preload During Plant Operation 2-4 j
2.1.4 Increase in Radial Preload Due to 'Ibermal Expansion Tightening 2-5 i
2.1.5 Increase in Radial Preload during N.O. and FLB Due to 2-5 Differential Pressure
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2.1.6 Change in Radial Preload due to Tubesheet Bow 2-6
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2.1.7 Net Preload in Roll Transition Region for N.O. and 2-6 i
FLB Conditions 2.2 Determination of Required Engagement Distance 2-6 l
2.2.1 Applied Loads 2-7 j
2.2.2 Coefficient of Friction at Tube-to-Tubesheet Interface 2-7 2.2.3 End Effects 2-7 2.2.4 Calculation of Engagement Distance Required, F*
2-8 2.3 Limitation of Primary to Secondary Leakage 2-9 2.3.1 Operating Condition Les Considerations 2-9 l
2.3.2 Postulated Accident Condition Leak Considerations 2-10 2.3.3 Operating Plant Leakage Experience for Within' Tubesheet
.2-11 Tube Cracks 2.4 Tube Integrity Under Postulated Limiting Conditions 2-11 2.5 Chemistry Considerations 2-11 2.5.1 Tubesheet Cormsion Testing 2-12 2.5.2 Tubesheet Corrosion Di:cussion 2-13 2.6 Summary of F* Evaluation 2-14 3.0 DEVELOPMENT OF L* CRITERIA 3-1 3.1 General 3-1 3.2 Leakage Considerations 3-1 3.3 Strength Considerations 3-2 3.4 Determination of Resistance to Leakage of Degraded Roll Expansions 3-3
- Prairie Island Units 1 and 2 3.4.1
Background
3-3 3.4.2 Applicability of 3/4 Inch Tube Tests 3-3 3.4.3 Resistance to Leakage 3-4 wrrusenam ii
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TABLE OF CONTENTS (continued) i Section Topic P. age i
3.0
. DEVELOPMENT OF L* CRITERIA' (continued) 3.5 Leakage Criterion and Calculated Leakage from Degraded Roll 3-5 j
Expansions -- Prairie Island Units 1 and 2 3.5.1 Leakage Criterion 3-5
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3.5.2 Leakage Projected for Prairie Island Units 1 and 2 -
3-7 i
3.6 General Application of the 3/4 Inch Tube Axial loadbearing Test :
3-9
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Results to Prairie Island Units 1 and 2 3.7 Application of 3/4 Inch Tube Test Results for the Axial Loadbearing 3-9.
Capability of Degraded Roll Expansions to Prairie Island Units 1 and 2
- 3.7.1 Maximum Axial Load Criterion 3-9 3.7.2 Evaluations 3-10 3.7.3 Results 3-12 3.7.4 Conclusion 3-13 4
3.8 Pullout Load Reaction and RPC Inspection Lengths -- Prairie Island 3-13 Units 1 and 2 -
3.8.1 General 3-13 1
3.8.2 Single Band Degradatian 3-13 3.8.3 Multiple Band Degradadon 3-14 3.8.4 Technical Specifications Requirements 3-14 3.9 Summary of L' Evaluation 3-14 3.9.1 General 3-14.
3.9.2 Leakage-3-14 3.9.3 Strength 3-15 i
l 3.9.4 Recommended L* Criteria 3-16 l
4.0 REFERENCES
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I WPr/NSPI!20904 kii
TABLE OF CONTENTS (continued) 4 i
Section Topic P. age 4
APPENDIX A L* TESTS FOR 3/4 INCH DIAMETER TUBE SGS A-1 5
j A.1 Test for the Determination of Resistance to Leakage of Degraded A-1 Roll Expansions - 3/4 Inch Tube S/G A.I.1 Background A-1 A.I.2 Objective A-1 A. I.3 Test Equipment A-2 j
A.I.4 ~ Test Major Steps
'A-2 A.l.5 Test Facility A-3 A.I.6 Test Procedure A-4 l
A.2 -
Test Results for Leakage from Degraded Roll Expansions - 3/4" Tubes A-7 A.2.1 Leakage Criterion A-7 A.2.2 Roll Torque Reduction Test Results A-8
-1 A.3
-Test for the Determination of Aximi Loadbearing Capability of A-ll Degraded Roll Expansions - 3/4 Inch Tubes I
A.3.1 Introduction A-11 A.3.2 Objective A-11 A.3.3 Test Equipment A-12 A.3.4 Test Major Steps A-12 A.3.5 Data Acquisition System A-13 A.3.6 Procedure A-13 A.4 Criterion and Test Results for the Axial Loadbearing Capability of A-20 Degraded Roll Expansions - 3/4 Inch Tubes A.4.1 Maximum Axial Load Criterion A-20 A.4.2 Strength Test Results A A.4.3 Evaluations and Conclusions A-21 l
A.5 Conclusion A-30 l
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LIST OF FIGURES -
- Fin. No.
Titl_s Eags 1-1 Configuration for Tubesheet Region F* and L* Alternate Plugging Criteria for
' 1 -3 Partial Depth Roll-Expanded Steam Generator Tubes, Single Band Degradation 2-1 "End-Affected" Regions in a Tube Postulated to be Severed at the' Bottom 2-22 of F* Length i
3-1 Configuration for Tubesheet Region L' Alternate Plugging' Criteria for
~3-26
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Partial Depth Roll-Expanded Steam Generator Tubes, Multiple Band Degradation i
3-2 Postulated Leakage Configuration for Prairie Island Units 1 and 2 SG -
3-27 i
Normal Operation and FLB Conditions 3-3 Postulated Leakage Configuration for Steam Generator LOCA Conditions 3-28 l
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3-4 Model for Plastic Collapse 7/8" Diameter, Mill Annealed, Alloy 600 Tubing 3-29 l
t 3-5 Comparison cf Load-Displacement Records, Computed 3-30 vs. Measured for 30 Slots at $=45*,3/4 Inch Tubes 3-6 Comparison of Load-Displacement Records, Computed 3-31 l
vs. Measurad for 30 Slots at $=30*,3/4 Inch Tubes l
3-7 Ultimate Pull Force as a Function of Slot Angle for Collared and ~
3-32 i
Decollared Tubes,3/4 Inch Tubes 3-8 Pull Force Design Curve for 7/8 Inch Tubes at Prairie Island Units 1 and 2 3-33 L
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A.1-1 Roll Expansion Leakage Test fiample for 3/4 Inch Tube Steam Generator A-5 Normal Operation and FLB Conditions A.1-2 Roll Expansion Leakage Test Sample for 3/4 Inch Steam Generator LOCA A-6 I
Conditions l
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WPF/NSP/120094 v
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l LIST OF FIGURES (continued)
Fin. No.
Iids Eags A.3-1 ' Geometry of Slots in Collared Tube Specimen - Ultimate Strength Test.
A-16 A.3-2 Geometry of Slots in Decollared Tube Specimen - Ultimate Strength Test A-17 A.3-3 Geometry of Slots in Unexpanded, Never-Collared Tube Specimen -
. A-18 Ultimate Strength Test A.3-4 Sketch of Zetac Tension and Compression Testing Machine A-19 A.4 l roll Force as a Function of Pull Displacement for a Collared, A-23 Non-degraded Tube (Specimen No.11)
A.4-2 Pull Force as a Function of Pull Displacement for a Collared Tube.
A-24 (Specimen No. 3), with (15)-30 Degree Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition A.4-3 Ultimate Pull Force as a Function of Slot Angle for Expanded, Collared -
A-25 and Decollared Tubes,15 Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transitica; - 3/4 Inch Tubes A.4-4 Model For Plastic Collapse A-26 l
A.4-5 Comparison of Load-Displacement Records, Computed vs Measured A-27 for Slots at & = 45* -- 3/4 Inch Tubes A.4-6 Comparison of Load-Displacement Records, Computed vs Measured A-28 for 30 Slots at $ = 30 -- 3/4 Inch Tubes I
A.4-7 3/4 Inch Tube Steam Generator Degraded Tube Pull Strength Design A-29 l
Curve for L' l
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l WPF/NSP/120994 v1
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- I j-LIST OF TABLES '
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2-1 SG Tube Roll Preload Test - Test Data 2-15
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2-2 SG Tube Roll Preload Test - Stress Analysis Results 2-18 2-3 Calculation of F* Length for T/TS Hardroll Interface - Prairie Island I and 2 2-19 2-4 Preload Analysis Summary 2-20 3-1 Technical Approach to Meet Regulatory Requirements 3-17 3-2 Basis for Use of 3/4 Inch Tube Joint Leakage Data for Projected Leakage 3-18 for Prairie Island Units 1 and 2 Degraded Roll Expansions 3-3 3/4 Inch Tube Roll Expansion Leakage Test Results and Projected Leakage 3-19 for Prairie Island Units 1 and 2 (7/8 Inch Tubes) 3-4 Basis for Use of 3/4 Inch Tube Joint loadbearing Test Data for Projected 3-21 1
Loadbearing Capability for Prairie Island Units 1 and 2 Degraded Tube Joints 3-5 Calculation of Pullout Load Reaction Length and Tube Length 3-22 l
to be RPC Inspected l
A.2-1 Roll Expansion Leakage Test Results - 3/4 Inch Tube S/G -
A-9 A.3-1 Selection of Bounding Features for Axial Loadbearing Test of Degraded A-14 Roll Expansions A.3-2 Axial Loadbearing Strength Test of Degraded Roll Expansions
- A-15
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I WPFINSPn20004
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NOMENCLATURE j
(Terms are listed in the approximate order of appearance in the report) i RT Roll transition BRT Bottom of roll transition RE Roll expansion-NDD No Detectible Degradation N.O.
Normal Operation,
LOCA Loss of Coolant Accident i
Indication length TS Tubesheet T/TS Tube to tubesheet ECT Eddy current test X
Distance from top of indication to BRT Number of linear indications per band n
Linear indication inclination angle from the tube axis, a.k.a., crack angle
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s TTS Top of tubesheet -
1 RPC Rotating pancake coil S/G Steam generator ID Inside diameter OD Outside diameter l
dpm Drops per minute l
ECI Eddy current indication LTL Lower tolerance limit FLB Feedline break SLB Steamline break
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s, T/fS interfacial radial contact pressure T
Torque MA Mill-annealed SRE Sound roll expansion PLRL Pullout load reaction length SBD Single band degradation N
Number of sound roll expansion portions RCS Reactor coolant system
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WPTNSPnm94 Ykkk
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INTRODUCTION i
l This report documents the development of criteria for repairing partial depth hardroll l
expanded steam generator tubes for degradation in the expanded region of the tube within the j
tubesheet. Existing Prairie Island Units 1 and 2 Technical Specification tube i
repairing / plugging criteria apply throughout the tube length, but do not take into account the
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reinforcing effect of the tubesheet on the external surface of the tube. Two separate repair' i
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criteria, called F* and L*, are developed for tubes with degradation in the expanded region below the bottom of the roll transition.
j First, an evaluation is performed to determine the length of hardroll engagement required to resist tube pullout forces during normal operation, test, upset and faulted conditions. The evaluation uses the results of previously conducted analyses and tests aimed at quantifying the l
residual radial preload of Westinghouse 51 Series steam generator tubes hardrolled into the tubesheet holes. It is shown that the radial preload would be sufficient to significantly restrict i-leakage during normal operation and faulted conditions. The necessary length of undegraded :
i tubing, called F*, required to resist tube pullout forces is calculated. On this basis, tubes in
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the Prairie Island Units 1 and 2 steam generators with no degradation within the F* distance,.
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measured from the bottom of the roll transition, are sufficient for continued plant operation,-
1 regardless of the extent or nature of tube degradation below F*.
1 In addition, leakage-based repair criteria are also developed for tubes with degradation within f
the partial depth hardroll expansion zone. These criteria, designated L*, are based on the fact that leaks from throughwall cracks will be substantially reduced by a length of tube expanded 1
l into the tubesheet which is shorter than the F* length required for structural integrity. The
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evaluation accounts for the fact that axially oriented cracks within the roll expansion region will still contribute to the structural integrity of the tube /tubesheet joint, even when the cracks
~j have propagated throughwall. Although the L* plugging criteria are limited in their application to tubes with axial or near-axial cracks, they can permit additional tubes with i
degradation within the F* zone of the expansion region to remain in service. Finally, this i
evaluation demonstrates that the F* and L* criteria for tube degradation within the tubesheet
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afford a level of plant protection commensurate with that provided by Reg. Guide (RG) L121 j.
(Reference 1) for degradation located outside of the tubesheet region.
1.1 Background
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l Existing plant Technical Specification tube repair / plugging criteria which have been applied
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throughout the tube length do not take into account the reinforcing effect of the tubesheet
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(TS) on the external surface of the expanded portion of the tube. The presence of the TS will constrain the tube and will complement tube integrity in that region by essentially precluding tube deformation beyond the expanded outside diameter. The resistance to both tube rupture i
and tube collapse is significantly strengthened by the TS. In addition, the proximity of the j
TS significantly affects the leak behavior of throughwall tube cracks in this region. Based on these considerations, the establishment of alternate, plugging criteria specific to the roll region 1
of the tubes is justified.
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1 WPF/NSP/121394 I-}
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l-j The roll expanded length of the tube can be considered to consist of two zones, the roll 1
transition (RT) region and the roll region. The roll transition region is defined as that portion I
of the tube where the roll expanded length transitions to the unexpanded length. The RT is l
approximately [
]* inch above the tube end at the bottom of the TS (i.e, above the i
initial tack roll applied during tube bundle assembly). The roll region is also referred to as j
the roll expansion (RE) or hardroll. Refer to Figure 1-1.
4 Taking credit for the reinforcing effect of the tubesheet and the radial contact load between the expanded region of the tube and the tubesheet, the F* and L* plugging criteria are i
j-developed. The F* criterion permits operation with any amount of tube degradation below a calculated distance, F*, below the bottom of the roll transition (BRT). Within the F*
distance, no degradation is permissible. The F* distance is shown to provide sufficient frictional force between the expanded tube and tubesheet to resist pullout due to normal j
operation and postulated accident conditions. The use of the F* criterion does not require any l
assessment of tube degradation other than elevation, and any type of degradation below the j
F* distance is acceptable.
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Eddy current indications (ECI's) in the top portion of the roll expansion (within the F' j
distance) may be addressed by additional plugging criteria designated as the L* criteria.
These criteria are based on the fact that leaks from throughwall tube degradation will be substantially reduced by a certain length of tube expanded into the tubesheet, shorter than the length required for structural integrity. Also, some types of potential tube degradation,
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specifically axial or near-axial cracks within the tubesheet roll expanded region, will provide l
the required structural integrity even when the cracks have propagated throughwall. Use of the L* criteria requires more definition of the tube degradation and probably more
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sophisticated eddy current inspection techniques than required for the F* criterion, but should j
address a significant number of the tubes which do not qualify for the F* criterion. Finally, a combination of the L' and F* analysis methods can be shown to address the existence of tube j
degradation in the form of axial or near-axial ECI's at a single elevation or at multiple j
elevations in the roll expansion.
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Figure 1-1.
Configuration for Tubesheet Region F* and L* Plugging Criteria for Partial Depth Roll-Expanded, Steam Generator Tubes (Single Band of l
Degradation.)
wrrmensoen 1-3 l
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i-l 2.0 DEVELOPMENT OF F* CRITERlON j
The F* tube plugging criterion is based on a semi-empirical method of quantifying the axial -
loadbearing capability of the rolled joint, resulting from the radial contact preload pressure 1
and the associated friction between the tube and TS. The presence of the tube-to-tubesheet (T/TS) radial pressure, s,, which consists of the as-manufactured pressure as changed by operating loads and temperatures, also causes significant resistance to the leakage of j
primary-to-secondary and secondary-to-primary water. It has been determined by previous
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tests, as shown in Appendix A, that sound roll expansions of F* length are essentially leak
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l tight. The use of the F* criterion obviates the necessity of determining the ECI depth,
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i number, inclination, length and circumferential spacing. Only the distance from the uppermost part of the ECI to the BRT needs to the determined. In short, the nature and l
l extent of tube degradation need not be determined. Refer to Figure 1-1.
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i Tube rupture in the conventional sense, as characterized by an axially oriented " fishmouth" opening in the side of the tube, is not possible within the tube /tubesheet roll expansion (RE).
j The reason for this is that the tubesheet material prevents the wall of the tube from expanding i
outward in response to the internal pressure forces. The forces which would normally act to cause crack extension are transmitted into the walls of the tubesheet, the same as for a l
non-degraded tube, instead of acting on the tube material. 'Ihus, axially oriented linear l
indications, e.g., cracks, cannot lead to tube failure within the RE and may be considered on f
the basis ofleakage eNects only.
P Likewise, a circumferentially oriented tube rupture is resisted because the tube is not free to f
i deform in bending within the roll expansion. When degradation has occurred such that the j
remaining tube cross sectional area does not present a uniform resistance to axial loading,.
4 bending stresses are developed which may significantly accelerate failure. When bending l
l forces are resisted by lateral support loads, provided by the tubesheet, the acceleration mechanism is mitigated and the tube separation mode is similar to that which would occur in i
a simple tensile test. Such a separation mode, however, requires the application of I
j significantly higher loads than for the unsupported case.
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In order to evaluate the applicability of any developed criterion for indications within the l
tubesheet, some postulated type of degradation must be considered. For this evaluation it was i
postulated that a circumferential severance of a tube could occur, contrary to existing plant j
operating experience. However, implicit in assuming a circumferential severance to occur is i
3 the consideration that degradation of any extent could be demonstrated to be tolerable below j
the location determined acceptable for the postulated condition.
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When the tubes have been hardiolled into the tubesheet, any axial loads developed by j
pressure and/or mechanical forces acting on the tubes are resisted by friction forces developed by the elastic preload that exists between the tube and the tubesheet. For some specific j
length of engagement of the hardroll, no significant axial forces'will be transmitted farther along the tube, and that length of tubing, i.e., F*, will be sufficient to anchor the tube in the j
tubesheet. In order to determine the value of F* for application in 51 Series steam generators,
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a testing program was conducted to measure the elastic preload of the tubes in the tubesheet.
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wemsenem 2-1
l The presence of the elastic preload also presents a significant resistance to flow of primary to secondary or secondary to primary water for degradation which has progressed fully through the thickness of the tube. In effect, no leakage would be expected if a sufficient length of J
hardroll is present. This has been demonstrated in steam generator sleeve to tube joints made by the Westinghouse hybrid expansion joint process.
2.1 Determination of Elastic Preload Between the Tube and Tubesheet Tubes were installed in the Prairie Island steam generators using a hardrolling process which i
expands the tube to bring the outside surface into intimate contact with the tubesheet hole.
The roll process and roll torque are specified to result in a metal to metal interference fit l
between the tube and the tubesheet.
l A test program was conducted by Westinghouse to quantify the degree ofinterference fit j
l between the tube and the tubesheet provided by the partial depth hardrolling operation. The data generated in these tests have been analyzed to determine the length of hardroll required to preclude axial tube forces from being transmitted farther along the tube, i.e., to establish the F* criterion. The amount of interference was determined by installing tube specimens in 1
collars specifically designed to simulate the tubesheet radial stiffness. A hardroll process representative of that used during steam generator manufacture was used in order to obtain specimens which would exhibit installed preload characteristics like the tubes in the tubesheet.
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Once the hardrolling was completed, the test collars were removed from the tube specimens and the springback of the tube was measured. The amount of springback was used in an analysis to determine the magnitude of the interference fit, which is representative of the residual tube to tubesheet radial load in Westinghouse 51 Series steam generators.
l 2.1.1 Radial Preload Test Configuration Description l
The test program was designed to simulate the interface of a tube to tubesheet partial depth hardroll for a 51 Series steam generator. The test configuration consisted of six cylindrical collars, approximately [ ]'" inches in length, [
]"' inches in outside diameter (OD), and
[
]'" inch in inside diameter (ID). A mill annealed, Alloy 600 (ASME SB-163), tubing specimen, approximately [ ]'" inches long with a nominal [
]"' outside diameter before rolling, was hard rolled into each collar using a process which simulated actual tube l
installation conditions.
The design of the collars was based on the results of performing finite element analysis of a section of the steam generator tubesheet to determine radial stiffness and flexibility. The inside diameter of the collar was chosen to match the size of holes drilled in the tubesheet.
l The outside diameter was selected to provide the same radial stiffness as the tubesheet.
l The collars were fabricated from AISI 1018 carbon steel similar in mechanical properties to the actual tubesheet material. The collar assembly was clamped in a vise during the rolling l
process and for the post roll measurements of the tube ID. Following the taking of all post l
wrmanum 2-2
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roll measurements, the collars were saw cut to within a small distance from the tube wall.
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The collars were then split for removal from the tube and tube ID'and OD measurements
- were repeated.
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Two end boundary conditions were imposed on the tube specimen during rolling. The end l
was restrained from axial motion in order to perform a tack roll at the bottom end, and was l
allowed to expand freely during the final roll.
j 2.1.2 Preload Test Results: Discussion and Analysis I
l j[L All measurements taken during the test program are tabulated in Table 2-1; 'Ihe data recorded were employed to determine the interfacial conditions of the tubes and collars.
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These consisted of the ID and OD of the tubes before and after rolling and after removal from l
the collars, as well as the inside and outside diameters of each collar before and after tube i
rolling..Two orthogonal measurements were taken at six axial locations within the collars and j
tubes. Additional data ofinterest were calculated from these specific dimensions. The '
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calculated dimensions included wall thickness, change in wall thickness for both rolling and l
removal of the tubes from the collars, and percent of spring back.
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. Using the measured and calculated physical dimensions, an analysis of the tube deflections l
was performed to determine the amount of preload radial stress present following the hardrolling. The analysis consisted of application of conventional thick tube equations to -
f account for variation of structural parameters through the wall thickness. ' However, traditional application of cylinder analysis considers the tube to be in a state of plane stress. For these j
tests, the results implied that the tubes were in a state of plane strain elastically. This is in l
agreement with historical findings that theoretical values for radial residual preload are below g
those actually measured, and that axial frictional stress between the tube and the tubesheet
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increases the residual pressure. In a plane stress analysis such stress is taken to be zero i
j (References 2 and 3). Based on this information, the classical equations relating tube deformation and stress to applied pressure were modified to reflect plane strain assumptions.
The standard analysis of thick walled cylinders results in an equation for the r:. dial deflection of the tube as:
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U=C*r+C/r (1) i 2
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- where, U = radial deflection l
r = radial position within the tube wall, and the constants, C and C are found from the boundary conditions to be funcdons of the i
2 elastic modulus of the material, Poisson's ratio for the material, the inside and outside radii, and the applied internal and external pressures. The difference between an analysis assuming j
plane stress and one assuming plane strain is manifested only in a change in the constant C -
2 The first constant is the same for both conditions. For materials having a Poisson's ratio of O.3, the following relation holds for the second constant:
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I C (Plane Strain) = 0.862
- C (Plane Stress)
(2) 2 2
The effect on the calculated residual pressure is that plane strain results are higher than plane stress results by slightly less than 10 percent. Comparing this effect with the results reported in Reference 2 indicated that better agreement with test values is achieved. It is to be noted that the residual radial pressure at the tube to tubesheet interface is the compressive radial stress at the OD of the tube.
By substituting the expressions for the constants into Equation (1), the deflection at any radial location within the tube wall as a function of the intemal and external pressure (radial stress at the ID and OD) is found. This expression was differentiated to obtain flexibility values for the tube deflection at the ID and OD respectively, e.g., dU/dP, is the ratio of the radial deflection at the ID due to an OD pressure. Thus, dU/dP, was used to find the interface pressure and radial stress between the tube and the tubesheet as:
S,, = - P, = - (ID Radial Springback) / (dU/dP )
(3)
The calculated radial residual stress for each specimen at each location is tabulated in l
Table 2-2. The mean residual radial stress and the standard deviation were found to be
[
]'" psi and [
]"' psi, respectively. In order to determine a value to be used in the analysis, a tolerance factor for [
]'" percent confidence to contain [
]'" percent of the population was calculated, considering the [
]'" useable data points, to be [
]'"
- Thus, a[
]'" lower tolerance limit (LTL) for the radial residual preload at room temperature is
[
]"' psi.
2.1.3 Residual Radial Preload During Plant Operation During plant operation the amount of preload will change depending on the pressure and temperature conditions experienced by the tube. The room temperature preload stresses, i.e.,
radial, circumferential and axial, are such that the material is nearly in the yield state if a comparison is made to ASME Code (Reference 4) minimum material properties. Since the -
coefficient of thermal expansion of the tube is greater than that of the tubesheet, heatup of the plant will result in an increase in the preload and could result in some yielding of the tube.
]
In addition, the yield strength of the tube material decreases with temperature. Both of these effects may result in the preload being reduced upon retum to ambient temperature conditions, i.e., the celd condition. However, as documented in Reference 5 for a similar investigation, tube pullout tests which were preceded by a very high thermal relaxation soak showed the analysis to be conservative.
1 The plant operating pressure influences the preload directly based on the application of the pressure load to the ID of the tube, thus increasing the amount of interface loading. The pressure also acts indirectly to increase the amount of interface loading by causing the tubesheet to bow upward, i.e., placing the roll expansions near the bottom of the tubesheet in compression for normal operating and feedline break (FLB) conditions (FLB results in a higher primary-to-secondary AP than steamline break, hence FLB is used to bound the FLB and SLB cases for this analysis). For the loss of coolant accident (LOCA) event, the tubesheet bows in the opposite direction, producing dilation of the tubesheet holes and wermenam 2-4
l
{
l l
reducing the amount of tube to tubesheet preload. Each of these effects may be quantitatively i
treated.
The maximum amount ofincrease in preload due to tubesheet bow for primary-to-secondary pressure differential will occur at the bottom, central part of the tubesheet. Since F* is measured from the bottom of the hardroll transition (BRT) and leakage is to be restricted by the F* region of the tube, the potential for the tube section within the F* region to experience a net tightening or loosening during operation is evaluated. However, the central location case is not the most stringent case for normal operation and FLB; rather, the most stringent case for normal operation and FLB involves a peripheral tube, obtained for Case 2 of normal operation, which is the case with the maximum primary-to-secondary AP. The results for this limiting case are presented in the following sections.
2.1.4 Increase in Radial Preload Due to 'Ihermal Expansion Tightening For conservatism in determining the total residual preload for normal operating conditions, tightening of the tube /tubesheet joint due to differential thermal expansion is minimized by applying the SG outlet temperature to the tubing. From the limiting case identified in Section 2.1.3 and Table 2-3, this corresponds to a cold leg temperature of $28'F. The mean coefficient of thermal expansion for the Alloy 600 tubing between ambient conditions and i
4 4
528'F is approximately 7.74*10 in/in/'F. That for the steam generator tubesheet is 7.30*10 in/in/ F. These 9 alues were reconciled as conservative with respect to the 1965 ASME Boiler and Pressure Vessel Code, which was the code of construction for the Prairie Island Units 1 and 2 SGs. Thus, there is a net difference of 0.44*104 in/in/*F between the expansion properties of the two materials. Considering a temperature difference of(528 - 70) = 458 F between ambient and operating conditions, the increase in preload between the tube (t) and the tubesheet (ts) was calculated as:
S,, = (0.44E-6)*(458)*(Collar ID) / 2 / ((dU/dPJ, - (dU/dP,))
(4)
The results indicate that the increase in preload radial stress due to thermal expansion is t
l
[
]'" psi. Note that this value applies for both normal operating and faulted conditions.
l l
2.1.5 Increase in Rsdial Preload during N.O. and FLB Due to Differential Pressure '
l l
The normal operating (N.O.) differential pressure from the primary to secondary side of the l
steam generator during the most limiting condition evaluated is 1593 psi (see Table 2-3). The intemal pressure acting on the wall of the tube will result in an increase of the radial preload on the order of the pressure value. The increase was found as:
S,, = - P, = - P (dU/dPJ / ( (dU/dP) - (dU/dP,))
(5) i i
In actuality, the increase in preload will be more dependent on the intemal pressure of the tube since water at secondary side pressure would not be expected between the tube and the i
tubesheet. However, the primary to secondary AP'is used for conservatism.
1 wrrasrmom 2-5 l
l
The increase in radial contact pressure due to differential pressure was evaluated for both normal operating (AP = 1593 psi) and faulted (AP = 2650 psi) conditions. The results indicate -
that the increase in preload radial stress is [
]* psi for normal operating conditions and
[
]* psi for faulted (FLB) conditions.
2.1.6 Change in Radial Preload due to Tubesheet Bow An analysis of the 51 Series tubesheet was performed to evaluate the change in preload stress j
that would occur as a result of tubesheet bow for interior tubes. The analysis was based on performing finite element analysis of the tubesheet and SG shell using equivalent perforated l
plate properties for the tubesheet (Reference 3). Boundary conditions from the results were l
then applied to a smaller, but more detailed model, in order to obtain results for the tubesheet holes. Basically the deflection of the tubesheet was used to find the stresses active on the i
l bottom surface and then the presence of the holes was accounted for. For the location where the increase of preload is a maximum, the radial preload stress would be' increased by
[
]* psi during normal operation and [
]* psi during faulted (FLB) conditions.
However, the interior tubes are not the limiting case for primary-to-secondary pressure -
differential. The limiting case involves peripheral tubes where tubesheet bowing has a negligible effect on tube-to-tubesheet preload. Therefore, the N.O. and FLB analyses address only tubes in the peripheral region of the tubesheet. During LOCA, the differential operating pressure is from secondary to primary. Hus, the radial preload wil! decrease by [
]* psi as the tubesheet bows downward. However, the action of the differential pressure is such that the tube is pushed toward the tube-to-tubesheet well. His case is of no consequence to the determination of F*
2.1.7 Net Preload in Roll Transition Region for N.O. and FLB Conditions l
l Combining the room temperature hardroll preload with the thermal and pressure effects results L
in a net operating preload of [
]* psi during normal operation and
[
]* psi for faulted conditions. In addition to restraining the tube in the tubesheet, this preload should effectively retard leakage from indications in the tubesheet region of the tubes.
2.2 Determination of Required Engagement Distance The calculation of the value of F* recommended for application to the Prairie Island Units 1 and 2 steam generators is based on determining the length of hardroll necessary to cffset the applied loads during the maximum normal operating conditions or faulted conditions, whichever provides the largest value. Thus, the applied loads are balanced by the load carrying ability of the hardrolled tube for both of the above conditions. In performing the analysis, consideration is made of the potential for the ends of the hardroll at the hardroll transition and the assumed severed condition to have a reduced load carrying capability.
l werssenum 2-6 l
n
'~
2.2.1 Applied Loads i
1 8
The applied loads to the tubes which could result in pullout from the tubesheet during all normal and postulated accident conditions are predominantly axial and due to the internal to external pressure differences. For a tube which has not been degraded, the axial pressure' load
'l is given by the product of the pressure with the internal cross-sectional area. 'However, for a' tube with intemal degradation, e.g., cracks oriented at an angle to the axis of the tube, the intemal pressure may also act on the flanks of the degradation. Hus, for a tube which is conservatively postulated to be severed at some location within the tubesheet, the total force -
acting to remove the tube from the tubesheet is given by the product of the pressure and the' cross-sectional area of the tubesheet hole. The force resulting from the pressure and intemal area acts to pull the tube from the tubesheet and the force acting on the end of the tube tends L to push the tube from the tubesheet. For this analysis, the tubesheet hole diameter has been used to determine the magnitude of the pressure forces acting on the tube. The forces acting to remove the tube from the tubesheet are [
]'" pounds and [
]'" pounds respectively l
for normal operating and faulted conditions. Any other fe:ces such as fluid drag forces in the l
i U-bends and vertical seismic forces are negligible by comparison.
I 2.2.2 Coefficient of Friction at Tube-to-Tubesheet Interface i
In order to determine the coefficient of friction between hard-rolled tubes and the tubesheet, pull tests and hydraulic proof tests were conducted on 3/4" diameter Alloy 600 tubing, hard rolled into carbon steel collars with'an OD to simulate tubesheet rigidity (similar to the tests' l
described in Section 2.1).- After rolling, an inside circumferential cut was machined through i
i the wall of the tube at a controlled distance from the bottom of the roll transition. The i
l-samples were heat soaked at [
]'" to simulate the possible 'effect of reduced preload force due to tube yielding during manufacturing heat treatment. 'Two sets of pullout tests were conducted, on a tensile testing machine in air at room temperature, and with intemal pressure as the acting force on the tube. The pressure tests were performed at room temperature using deionized water. The pull tests performed with the tensile testing machine showed a static coefficient of friction of [
]'"
For samples which were expelled from the collars during the hydraulic proof tests, the coefficient of friction was determined to range l
from [
]'".
For tubes that leaked before tube expulsion, resulting in termination j
of the tests before expulsion, the lower bound coefficients of friction were determined to I
range from [
]'". On the basis of these results, the use of a coefficient of friction of [
]'" is considered to be conservative for the 7/8" tubes at Prairie Island Units 1 and 2 for application in determining the required engagement distance to resist tube pullout forces.
2.2.3 End Effects i
For a tube which is postulated to be severed within the tubesheet there is a material discontinuity at the location where the tube is severed. For a small distance from each assumed discontinuity the stiffness, and hence the radial preload, of the tube is reduced j
relative to that remote from the ends of the roll expansion. He analysis of end effects in thin cylinders is baseo on the analysis of a beam on an elastic foundation. For a tube with a given l
WPF/NSPn21494 2-7 i
k,
,u r,
. ~
~.
1 radial deflection at the end, the deflection of points away from the end relative to the end
- deflection is given by:
u / u,, = e*
- cosine ( A
- x )
~(6).
m
- where, A=[
]* = end effect constant.
x = distance from the end of the tube.'
Fcr the radially preloaded tube, the distance for the end effects to become negligible is the location where the cosine term becomes zero. Thus, for the roll expanded 51 Series tubes the distance corresponds to the product of "A" times "x" being equal to (x/2) or [.
]* inch.
Figure 2-1 shows a roll expansion which is postulated to be severed at the bottom of the F*
region. For a distance of [ -
]* inch above the severed end and below the bottom of the -
roll transition, the expanded joint has a reduced radial load carrying capability relative to the remainder of the F* length. He effective radial preload carried by these "end-affected" regions is calculated as follows.
i I
l The above equation can be integrated to find the average deflection over the affected length to be 0.384 of the end deflection. This means that on the average the stiffness of the material over the affected length is 0.616 of the stiffness of the material remote from the ends.
Therefore, the effective preload for the affected end lengths is 61.6 percent of the preload at regions more than [
]* inch from the ends. For example, for the normal operating net preload of [
]* psi or [
]* pounds per inch of length, the effective preload for a distance of [
]* inch from the end is [
]* pounds per inch or [
]* pounds.
2.2.4 Calculation of Engagement Distance Required, F*
The calculation of the required engagement distance is based on determining the length for l
l preload frictional forces to equilibrate the applied operating loads..The axial friction force was found as the product of the radial preload force and the coefficient of friction between -
l the tube and the tubesheet. The value assumed for the coefficient of friction was [
]*,
l from Reference 5. For normal operation the radial preload is [
]* psi or [
- ]*
I pounds per inch of engagement. Hus, the axial friction resistance force is [
-]* pounds l
per inch of engagement. It is to be noted that this value applies away from the ends of the tube. For any given engagement length, the total axial resistance is the sum of that provided by the two ends plus that provided by the length minus the two end lengths. From the preceding section the axial resistance of each end is [
]* pounds. Considering both ends i
of the presumed severed tube, i.e., the hardroll transition is considered one end, the axial I
resistance is [
]* pounds plus the resistance of the material between the ends, i.e., the total length of engagement minus [
]* inch. For example, a one inch length has an axial resistance of, l
[
]*
l-wrrmennosu 2-8
n
(
l l
' Conversely, for the maximum normal operating pressure applied load of [
]"' pounds, considered as [
]'" pounds with a safety factor of 3, the length of hardroll required is given by, F* = [
]"' = 1.07 inches.
I l
Similarly, the required engagement length for faulted conditions can be found to be 0.75 inch using a safety factor of 1.43 (corresponding to a ASME Code safety factor of 1.0/0.7 for allowable stress for faulted conditions).
The calculation of the above values is summarized in Table 2-4. The F* value thus determined for the required length of hardroll engagement below the BRT, for normal operation is sufficient to resist tube pullout during both normal and postulated accident condition loadings.
Based on the results of the testing and analysis, it is concluded that following the installation of a tube by the standard hardrolling process or the application of an additional roll expans on, a residual radial preload stress exists due to the plastic deformation of the tube and tubesheet interface. His residual stress is expected to restrain the tube in the tubesheet while providing a leak limiting seal condition, provided a sufficient length of_undegraded tubing exists below the bottom of the roll transition.
2.3 Limitation of Primary to Secondary Leakage The allowable amount of primary to secondary leakage in each Prairie Island Units I and 2 steam generator during normal plant operation is limited by plant technical specifications to a l
total of I gpm for both SGs. His limit, based on plant radiological release considerations and implicitly enveloping the leak before break consideration for a throughwall crack in the free span of a tube, is also applicable to a leak source within the tubesheet. In evaluating the primary to secondary leakage aspect of the F* criterion, the relationship between the tubesheet region leak rate at postulated FLB conditions (which bound SLB for primary-to-secondary i
pressure differential considerations) is assessed relative to that at normal plant operating.
conditions. The analysis was performed by assuming the existence of a leak path; howevsr, no actual leak path would be expected due to the hardrolling of the tubes into the tubesheet, i
2.3.1 Operating Condition Leak Considerations In actuality, the hardrolled joint would be expected to be leak tight, i.e., the plant would not be expected to experience leak sources emanating below F*.
Because of the presence of the tubesheet, tube indications are not expected to increase the likelihood that the plant would experience a significant number of leaks. It could also be expected, that if primary to l
secondary leakage is detected in a steam generator it will not be in the tube region below F*.
l Thus, no significant radiation exposure due to the need for personnel to look for I
tube /tubesheet leaks should be anticipated, i.e., the use of the F* criterion is consistent with ALARA considerations. As an additional benefit relative to ALARA considerations, l
l j
werwrnum 2-9
l precluding the need to install plugs below the F* criterion would result in a significant reduction of unnecessary radiation exposure to installing personnel.
The issue of leakage within the F* region up to the top of the roll transition (RT) includes the consideration of postulated accident conditions in which the violation of the tube wall is very extensive, i.e., that no material is required at all below F*. Based on operating plant and laboratory experience the expected configuration of any cracks, should they occur, is axial.
The existence of significant circumferential cracking is considered to be of very low probability. Thus, consideration of whether or not a plant will come off-line to search for leaks a significant number of times should be based on the type of degradation that might be expected to occur, i.e., axial cracks. Axial cracks have been found both in plant operation and in laboratory experiments to be short, about 0.5 inch in length, and tight. From field experience, once the cracks have grown so that the crack frcnt is out of the skiproll or transition areas, they arrest.
Axial cracks in the free span portion of the tube, with no superimposed thinning, would leak at rates compatible with the technical specification acceptable leak rate. For a crack within the F* region of the tubesheet, expected leakage would be significantly less. Leakage through cracks in tubes has been investigated experimentally within Westinghouse for a significant number of tube wall thicknesses and thinning lengths (Reference 6). In general, the amount of leakage through a crack for a particular size tube has been found to be approximately proportional to the fourth power of the crack length. Analyses have also been performed i
which show, on an approximate basis for both elastic and elastic-plastic crack behavior, that i
the expected dependency of the crack opening area for an unrestrained tube is on the order of the fourth power, e.g., see NUREG CR-3464. The amount of leakage through a crack will be proportional to the area of the opening, thus, the analytic results substantiate the test results.
The presence of the tubesheet will preclude deformation of the tube wall adjacent to the crack, i.e., the crack flanks, and the crack opening area may be considered to be directly proportional to the length. The additional dependency, i.e., fourth power relative to first power,is due to the dilation of the unconstrained tube in the vicinity of the crack and the bending of the side faces or flanks of the crack. For a tube crack located within the tubesheet, the dilation of the tube and bending of the side faces of the crack are suppressed.
Thus, a 0.5 inch crack located within the F* region up to the top of the roll transition would be expected to leak, without considering the flow path between the tube and tubesheet, at a rate less than a similar crack in the free span, i.e., less than the Prairie Island Units 1 and 2 technical specification limit of 1.0 gpm. Additional resistance provided by the l
tube-to-tubesheet interface would reduce this amount even further, and in the hardroll region j
the residual radial preload would be expected to eliminate it. This conclusion is supported by i
the results of the preload testing and analysis which demonstrated that a residual preload in excess of [
]* psi exists between the tube and the tubesheet at normal operating conditions.
2.3.2 Postulated Accident Condition Leak Considerations For the postulated leak source within the RE, increasing the tube differential pressure increases the driving head for the leak and increases the tube to tubesheet loading. For an wemsenem 2 - 10 s
.c
- - +
.n-,
--n-,..,,a,
i initial location of a leak source below the BRT equal to F*, the FLB pressure differential result-in an insignificant leak rate relative to that which could be associated with normal plant aperation. This small effect is reduced by the increased tube to tubesheet loading i
associated with the increased differential pressure as well as the tightening contribution of the i
tubesheet bending. Thus, for a circumferential indication within the RE which is left in service in accordance with the pullout criterion (F*), the existing technical specification limit l
is consistent with accident analysis assumptions. For postulated accident conditions, the preload testing and analysis showed that a net radial preload of about [
-]'" psi would exist between the tube and tubesheet.
For axial indications in a partial depth hardrolled tube below the BRT of the roll transition zone (which is assumed to remain in the tubesheet region), the tube end remains structurally intact and axial loads would be resisted by the remaining hardrolled region of the tube. For this case, the leak rate due to FLB differential pressure would be bounded by the leak rate for a free span leak source with the same crack length, which is the basis for the accident.
analysis assumptions.
2.3.3 Operating Plant Leakage Experience for Within-Tubesheet Tube Cracks i
A significant number of within-tubesheet tube indications have been reported for some j
non-domestic steam generator units. The present attitude toward operation with these indications present has been to tolerate them with no remedial action relative to plugging or sleeving. No significant number of shutdowns occurring due to leaks through these indications have been reported.
l 2.4 Tube Integrity Under Postulated Limiting Conditions The final aspect of the evaluation is to demonstrate tube integrity under the postulated loss of coolant accident (LOCA) condition of secondary to primary differential pressure. A review of tube collapse strength characteristics indicates that the constraint provided to the tube by the tubesheet gives a significant margin between tube collapse strength and the limiting secondary to primary differential pressure condition, even in the presence of circumferential or axial indications.
The maximum secondary to primary differential pressure during a postulated LOCA is
[
]'" psi. This value is significantly below the residual radial preload between the tubes and the tubesheet. Therefore, no significant secondary to primary leakage would be expected to occur. In addition, loading on the tubes is axially toward the tubesheet and could not contribute to pullout.
2.5 Chemistry Considerations The concern that boric acid attack of the tubesheet *due to the presence of a through wall flaw within the hardroll region of the tubesheet may result in loss of contact pressure assumed in the development of the F* criterion is addressed below. In addition, the potential for the wermsmrom 2 - 11 1;_____
-.-~-
i i
i i
l i
existence of a lubricated interface between the tube and tubesheet as a result oflocalized primary to secondary leakage and subsequent effects on the friction coefficient assumed in the development of the F* criterion is also discussed.
l 2.5.1 Tubesheet Corrosion Testing 3
4 Corrosion testing performed by Westinghouse specifically addressed the question of corrosion j
rates of tubesheet material exposed to reactor coolant. The corrosion specimens were l
assembled by bolting a steel (A336) coupon to an Alloy 600 coupon. The coupon dimensions t
were 3 inches x 3/4 inch x 1/8 inch and were bolted on both ends. A torque wrench was
]
used to tighten the bolts to a load of 3 foot-pounds. The performance of A508 in testing of 1
this nature is expected to be quite comparable to the performance of A336 (Gr. F-1) steel (a j
material used for tubesheet construction prior to A508). The arguments used in supporting j
the F' case relative to corrosion of tubesheet material due to minute quantities of primary coolant contacting the carbon steel were so conservative and had such margin that minor differences in material composition or strength would not change the conclusion.
I j
The specimens were tested under three types of conditions:
i 1
i 1.
Wet-layup conditions 2.
Wet-layup and operating conditions 3.
Operating conditions only l
The wet layup condition was used to simulate shutdown conditions at high boric acid concentrations. The specimens were exposed to a fully aerated 2000 ppm boron (as boric
)
acid) solution at 140 degrees F. Exposure periods were 2,4,6, and 8 weeks. Test solutions l
were refreshed weekly.
While lithium hydroxide is normally added to the reactor coolant as a corrosion inhibitor, it j
was not added in these tests in order to provide a more severe test environment. Previous i
testing by Westinghouse has shown that the presence of lithium hydroxide reduces corrosion of Alloy 600 and steel in a borated solution at operating temperatures.
1 t
1 Another set of specimens were used to simulate startup conditions with some operational j
exposure. The specimens were exposed to a 2000 parts per million boron (as boric acid) solution for one week in the wet-layup condition (140'F), and 4 weeks at operating conditions (600'F,2000 psi). During wet layup, the test solution was aerated but at operating conditions the solution was deaerated. The high temperature testing was performed in an Alloy 600 q
autoclave. Removal of oxygen was attained by heating the solution in the autoclave to 250'F j
and then degassing. This method of removing the oxygen results in oxygen concentrations of
]
less than 100 parts per billion.
i j
Additional specimens were exposed under operating conditions only for 4 weeks in the j
autoclave as described above.
1 High temperature exposure to reactor coolant chemistry resulted in steel corrosion rates of about I mil per year. This rate was higher than would be anticipated in a steam generator i
WPF/NSP/Imm 2 - 12
i
!n i
i l
since no attempt was made to completely remove the oxygen from the autoclave during -
l heatup. Even with this amount of corrosion, the rate was still a factor of nine less than the corrosion rate observed during the low temperature exposure. This differential corrosion rate j
' observed between high and low temperature exposure was expected because of the decreasing j
acidity of the boric acid at high temperatures and the corrosive effect of the high oxygen at j
low temperatures.
2 i
l These corrosion tests are considered to be very conservative since they were ' conducted at
}
maximum boric acid concentrations, in the absence of lithium hydroxide, with no special j
precaution to deaerate the solutions, and they were of short duration. The latter point is very j
significant since parabolic corrosion rates are expected in these types of tests, which leads one to overestimate actual corrosion rates when working with data from tests of short duration.
l l
Also note that the ratio of solution to surface area is high in these tests compared to the scenario of concern, i.e., corrosion caused by reactor coolant leakage through a tube wall into j-the region between the tube and the tubesheet.
i 2.5.2 Tubesheet Corrosion Discussion 7
l At low temperatures, e.g., less than 140 degrees F, aerated boric acid solutions comparable in strength to primary coolant concentrations can produce corrosion of carbon steels. Demerated solutions are much less aggressive and deaerated solutions at reactor coolant temperatures produce very low ce trosion rates due to the fact that boric acid is a very much weaker acid at i
high temperature, e.g.,610 degrees F, than at 70 degrees F.
i l
In the event that a crack occurred within the hardroll region of the tubesheet, as the amount j
of leakage would be expected to be insufficient to be noticed by leak detection techniques and i
is largely retained in the crevice, then a very small volume of primary fluid would be involved. Any oxygen present in this very small volume would quickly be consumed by surface reactions, i.e., any corrosion that would occur would tend to cause existing crevices to narrow due to oxide expansion and, without a mode for replenishment, would represent a very i
benign corrosion condition. In any event the high temperature corrosion rate of the carbon
[
steel in this very local region would be extremely low (significantly less than 1 mil per year).
~
i
{
Contrast the proposed concern for corrosion relative to F* with the fact that Westinghouse has j
qualified boric acid for use on the secondary side of steam generators where it is in contact i
with the full surface of the tubesheet and other structural components made of steel. The i
latter usage involves concentrations of 5 - 10 ppm boron, but, crevice flushing procedures j
have been conducted using concentrations of 1000 to 2000 ppm boron on the secondary side j
. (at approximately 275 degrees F where boric acid is more aggressive than at 610 degrees F).
4 Relative to the lubricating effects of boron, the presence of boric acid in water may change the wetting characteristics (surface tension) of the water but Westinghouse is not aware of any i
significant lubricating effect. In fact, any corrosion that would occur would result in oxides j
that would occupy more space than the parent metals, thus reducing crevice volume or possibly even merging the respective oxides.
R
. werssemom 2 - 13
2.6 Summary of F* Evaluation l
On the basis of this evaluation, it is determined that tubes with eddy current indications in the tubesheet region below the F* pullout criterion of 1.07 inches can be left in service. Tubes with circumferentially oriented eddy current indications of pluggable magnitude and located a distance less than F* below the bottom of the hardroll transition should be removed from service by plugging or repaired in accordance with the plant technical specification pluggmg
)
limit. The conservatism of the F* criterion was demonstrated by preload testing and analysis commensurate with the requirements of RG 1.121 for indications in the free span of the tubes.
For tubes with axial indications, the criterion which should be used to determine whether tube plugging or repairing is necessary should be based on leakage since the axial strength of a tube is not reduced by axial cracks. Under these circumstances it has been demonstrated that significant leakage would not be expected to occur for through wall indications greater than
[
]* inch below the bottom of the hardroll transition. Therefore, an F* distance of 1.07 inches achieves a leak limiting condition.
The application of F* in this report addresses existing roll expansions, performed in the l
factory. The approach may also be applied to roll expansions which are added to the SG T/TS joints of the partial depth roll expansion design, such as the design of the steam generators of Prairie Island Units 1 and 2. In this case, the reroll qualification program must
~ l demonstrate that the criteria of this report leading to an F* length of 1.07 inches (without NDE uncertainty) are satisfied for the rerolled condition. The F* distance for the additional roll expansion case should provide both the resistance to pullout and resistance to significant leakage. The reroll must then result in 21.07 inches of undegraded tube length above the tip of any crack left in service. Testing and analysis of the Prairie Island sludge conditions should be performed to validate the F* distance for the additional roll expansion. The F*
distance for the additional roll expansion qualification should address the potential conditions of hard or soft sludge, partially or completely filling the tube-to-tubesheet hole crevice before the additional rolling process. Based on Westinghouse test and analytical evaluations for the retrofit-expansion case, the F* length of additional roll expansion is expected to be essentially the same as the F* length of factory roll expansion determined in this report. The permissible number of successive applications of additional roll expansion depends on the tube conditions above the additional roll expansion. If the tubes are not effectively clamped at the tube support plates or top of the tubesheet (such as due to denting), the reroll process, in principal, can be repeated to ultimately approach a full depth tubesheet expansion; however, the process qualification program must address the effect of tubesheet bow on tube-tubesheet contact l
pressure as a function of height in the tubesheet. With clamped tube conditions, the process l
qualification program should address the stress levels developed in the tube joint by the l
rolling process. Westinghouse tests indicate that a few reroll applications can be applied for tubes dented at the top of the tubesheet.
l l
l wrrmsen m u 2 - 14 l
i l
l l
i Table 2-1 Model 51 SG Tube Roll Preload Test - Test Data Test Lecesen Cener ID PreJtes CeNet 00 Prden Tube ID 8efore Rett Tube CD sofore Rail No.
No.
8 Dog. M Deg.
Avg.
O Deg. M Deg. Avg.
8 Dog. M Des.
Avg.
8 Dec. M Dog.
Avg.
a,c.e 1
1
~
2 4
l 5
6 Average 2
1 2
3 4
l 5
6 Average 3
1 3
l 3
4 5
6 Average l
6 1
i 2
3 4
5 6
Average J
l 7
1 2
3 4
l 5
l 6
Average l
8 1
2 3
l i
4 I
5 l
6 i
Average Cd Avgs:
l l
l t
2-15 TABLES.WPF/112894
)
.l Table 2-1 (Continued)
Model 51 SG Tube Roll Preload Test - Test Data Tobe ID Peetten Test Lecasen Coller OD Poet 4e8 Tebe ID Post-Reg Coller Removed g
Pre 4ted Couer
-TebeID No.
No.
Thichsees e Deg. 90 Dog.
Ave.
DeNe O Deg. 90 Deg.
Avg.
G#euth 0 Dag. 90 Deg.
Avg.
e,c.e i
1 1
2 3
4 5
6 Average 2
1 2
3 4
5 6
f Average 3
1 2
3 4
5 6
Average 6
1 2
3 4
5 6
Average i
I 7
1 2
3 4
5 6
Average 8
1 2
3
)
4 6
Average Col Avgs.
l TABl.ES.WPF/ll2894
i l
l l
Table 2-1 (Continued)
Model 51 SG Tube Roll Preload Test - Test Data l
Test Locesen Tube oO PeeIJtes Post.
TWck.
Celier RadH TubeID Collar Resnowed Roll mese Fles.
Race.
Spelag.
No.
No.
O Dog.
to Deg.
Aug.
TWck.
Red.
du/dP, M
Seck a,c,e 1
1 2
3 4
5 6
Average 2
1 1
2 I
3 4
5 6
Average l
3 1
2 l
3 l
4 l
5 1
6 Average 6
1 l
2 3
4 5
6 Average l
7 1
2 l
3 4
5 i
6
)
Average l
l 8
1 2
3 l
4 5
i 6
Average Col. Avgs:
Notes:
1.
Al measured dmensens are r) inches.
- 2.. The OD stress is cakmlated usmg the measuredID spnngback 3
The radi rate is a term that amcars frecpntly in the analyse and a found as (OD*2+!D*2)/(OD*2-ID*2).
TABLES.WPF/112294
4 l
Table 2-2 Model 51 SG Tube Roll Preload Test - Stress Analysis Results Thorniel Oper.
Test Lecesos TubeID Tube Tube Flex.
CD OD OD Exp.
Tube Preneure Tetel Tetel Spdag-Flex.
dU)dP, Redel Heep Axlel Radial Flat Radiel Radial voeinees i
No.
No.
Back dU/dP, Strees Strees Strees Strees dU/dP, '
Strese Strees Strees a,c.e 2
1 3
4 5
6 Average l
2 1
2 3
)
4 5
6 l
Average i
I 3
1 2
- l 3
i 4
5 l
6 i
l Average l
l 6
1 2
3 4
5 6
Acerage 7
1 2
l 3
4 1
5 1
6 Average I
8 1
l 2
3 4
5 6
kerage Col Avgs:
Notes:
1.
The 00 stress a calculated tsing the measured 10 springback TABLES.WPF/ll2894 i
i i
Table 2 3
}
Calculaten of F,* t ength for TITS Herdron Interface Prairie Island tJnits 1 and 2 Maxwnum RCS Pressures for all PCWG Cases -
% Tube ~
RCS Steem Pri Sec Limstmg (tl Oe4te P N.O. -
Herdron Tetet N.O.
N.O..
N.O.
~
Ret.-Case Fluggeg Pressure Pressure DelteP Thot Tcold Diff Temp Ratio (2) T retio (3)
Sr== (4)
Srt(5)
Pretoed Sr(6)
Pr (7) Pond 18) -
Pa (9) 3Pa F' (10)
(psiel (psial (psial (dog. F)
Idog. F1 (deg. F1 i
g
~
1*
0%
2250 750 1500 599.1 535.5 2*
15 %
2250 657 1693 592.1 527.9 3**
3.87%
2250 700 1550 590 530
~
~
3,C Meu F* -
1.07
~
t
References:
Ceses 1 and 2 represent des 6gn beeis parametere for 0% and 15% plugging levels beoed on Westinghouse thermelthydroute performance codes.
- Case 3 represents the current operating parameters et Prairie telend Units 1 and 2 es suppeled by Northem States Power.
i l
4 4
2-19 r
-.n
---,,---.r,--n~
---,,.-a
.-n-.
..,,.---e-w.
v-n.-
-.-n,
. ~ -.
.., ~
l I
Table 2-4 Preload Analysis Summary
[
Material Propertnes:
Tube /fubesheet Dimensions (Testedy a,c l
Elastic Modulus 28.7E+06 Init. Avg. Tube OD 1
Poisson's Ratio 0.30
. Init. Avg. Tube Thickness l
1600 Coeff. of Therm. Exp.
7.74E-06 in/in/*F Init. Avg. Tubesheet ID j
TS Coeff. of Therm. Exp.
7.30E4 in/in/*F Actual Tiunnmg j
Operatag AT.
460*F Apparent Thmmng N.0, AP
!$93 psi Faulted AP 2650 psi i
' Additional Analysis Input:
i l
a,c Coefficient of Friction I
Tubesheet Bow Stress Reduction y
N.O.
FLB
- End Effects:
~
Mean Radius (Rolled)
Lower Tolerance Limit Factor Tluckness (Rolled) 95/95 LTL 2.16 (N=36) 3.
'l End Effect Length Load Factor EVALUATION OF REQUIRED ENGAGEMENT LENGTH, F*
Elastic Analysis:
N.O.
FLB l
a,c RT Preload (LTL)
Thmna! Expansion Preload 1
l Pressure Preload i
j l
Tubesheet Bow Loss Net Preload Net Radial Force I
Net Axial Resistance i
Applied lead I
i Analysis Load End Effect Resistance Net Analysis Load I
End Effect Length l
Add. Length Required l
Tctal Length Required, F*
1.07in.
0.75 in.
h 2-20
- 1 l
1 Table 2-4 (continued) t Preload Analysis Summary l
NOTES:
- 1) 95/95 Lower Tolerance Limit rolled preload used.
l
- 2) For Normal Operation, r. safety factor of 3.0 is used.
- 3) For Faulted Conditions, a safety factor of 1.43 is used (correspondag to ASME Code use of 0.7 on ultimate strength.
- 4) 1he required length does not include eddy current inspection uncertainty for the location of the bottom of the hard solt '
. 5) Preload stresses used were for the most senogent case, i.e., cold leg, peripheral lar=han 1his miehizes the thermal expansma preload and eliminates the preload due to a W bowmg l
t 9
i I
l 1
1 l
l m
2-21
)
wermanitaa4 l
1
Figure 2-1 "End-Affected" Regions in a Tube Postulated to be Severed at the Bottom of F* Length Top of Tubesheet
)
ij
+
1 1
.m t
1 l
Top'of Roll Transitionk
{,BottomofRollTransition 1
7*l j
Bottom of Roll Transition a
4
(
l,c.e j
F*
a.c.e g
j l
u 1
Postulated Severed'End i
/
i i
i
/
)
?
i 5
r Bottom of Tubesheet I.
N 2-22 4
..~
3.0 DEVELOPMENT OF La CRITERIA 3.1 General The proposed L* roll expansion region pluggina criteria are evaluated for the four tube modes recommended to be considered by NRC Regulaury Guide 1.121, for three steam generator conditions (i.e., normal operation and two transient conditions). This evaluation includes a combination of analyses and adjustment of test data from an L* program for 3/4 inch diameter tubes to support the L* criteria for the 7/8 inch diameter tubes at Prairie Island Units I and 2, outlined in Table 3-1. Tests conducted for 3/4 inch tubing are documented in Appendix A. This evaluation is also based on the results of the F* analysis for Prairie Island ~
. Units 1 and 2, previously discussed in Section 2.
The potential eddy current indications in the partial depth roll expanded region of the Prairie Island SGs can be conservatively characterized as one of two limiting types of degradation and can be addressed with the F* and L* tube plugging criteria. The F* criterion was addressed in Section 2 of this report; it addresses all types of degradation, independent of -
crack orientation or size, provided the degradation does not extend within the F* distance of -
l.07 inches from the bottom of the roll transition. The L* criteria address axial or near-axial degradation which occurs below a distance, L*, from the bottom of the roll transition. Since the L* distance is shorter than F*, certain types of degradation within the F* region can be left in service if the degradation satisfies the L* criteria. The following sections support the development and incorporation of L* criteria into the plant Technical Specification.
i l
3.2 Leakage Considerations The selection of an L* length has as a primary consideration the goal of minimizing significant leakage from tubes which have been accepted using the L* criteria. Significant leakage would be an aggregate leakage from the L* tubes which would cause a shutdown or operation at reduced power due to safety, regulatory, or operational reasons. Determination of the leak rate through the interface between the expanded tube and the tubesheet is not conducive to calculational methods. For some tube joint ECT data, more than one degradation array, or degraded roll expansion (DRE) perjoint may be found. Refer to Figure 3-1. For various degradation arrays, the resistance-to-leakage (RTL) of the T/TS interface extending over a portion of the sound roll expansion portion (SRE) immediately below the BRT has been determined by testing for 3/4 inch tubing. These test resuts are extrapolated for use in 7/8 inch tubing in order to develop the L* criteria in the following sections.
In the case of the Prairie Island Units I and 2 partial depth roll expansions, degradation to which the L* criteria may be applied occurs within the approximately [
]'" inch long, hardrolled portion of the first approximately [
j inches of the tube from the tube end at the primary side of the tubesheet. The hardroll begins approximately [
]'" inch above the tube end. The uppermost DRE can be considered to be the most limiting location, i.e., with i
the highest leakage because the uppermost DRE fc,rms a manifold which intercepts all leakage flow, if any exists, in the T/TS interface.
wermsnmm 3-1
The application of L* in this report addresses existing roll expansions. The approach may also be applied to roll expansions or hydraulic expansions which are added to steam generator 1
tube joints of the partial depth roll expansion design, such as the steam generators of Prairie Island Units 1 and 2. In the retrofit-expansion case, the L* portion would perform the leakage resistance function of the new joint just as it would for the existing joint in the present case. The strength, or anchorage, function can be performed by the existing joint.
However, for this to apply, there is a limit on the type and extent of degradation in the original roll expansion which caused the tube to require plugging in the firet place. This degradation is limited to the type which has a minimal effect on the axial loadbearing capaciy of the tube, i.e., axial or near-axial, linear degradation.
3.3 Strength Considerations Determination of the loadbearing capability of DRE's was based on the limiting assumption that the L* region recommended length of 0.50 inch, plus ECT uncertainty, provides no axial fixity. Therefore, the most stringent axial loads, viz., pullout loads, were assamed to be applied directly to the uppermost DRE. The criteria also considers the degradation to be axial or te have an inclination angle,4, of up to [
]*, and to be throughwall. In this case, the minimum axial length of sound, undegraded roll expansions between the BRT and the top of the highest-elevation degradation, designated as L*, for operation with acceptable leakage was determined. In this criteria, the tube axial loads are borne by the sound portion of the tube i
above L*, the uppermost degraded portion, the sound portions, other axial or near-axial degradation portions, if any, below the uppermost degraded portion and extending to the T/TS l
weld. Determination of the characteristics, e.g., ECI number, inclination angle, etc., of all of l
the degraded regions in a given tube joint would be necessary if the tube pullout load were to be applied to the T/TS weld. However, this load need not be assigned to be reacted by the SRE within the L* distance and the weld. Instead, it can be reacted by the appropriate aggregate length of SRE's of the tube below the uppermost degradation array if this aggregate length is sufficient. Below this pullout load reaction length (PLRL), the degradation need not be quantified to the same extent as within the length; it need be quantified only to the same extent that degradation is quantified below the F* distance as in Section 2. In the most stringent case, these SRE's are interspersed with DRE's and several interspersed SRE's may be assigned to react the pullout load. Part of the pullout load is reacted by the SRE within the L* distance. Each DRE must transmit part of the pullout load to the SRE(s) and DRE(s) below it; the pullout load to be transmitted decreases with decreasing elevation. Therefore, the required DRE strength could also reduce with elevation. However, for the sake of simolicity, all DRE's were treated the same. Calculation of the holding force from each SRE was performed in the same way that the holding force was calculated for the F* distance in Section 2. 'Ihis involved calculation of the axial frictional force, preventing pullout, as the product of the T/TS radial contact pressure (s,) of the area of the SRE and the T/rS static coefficient of friction. The reduced holding force of the two ends of each of the SRE's was also considered. Therefore, it was concluded that the configuration involving multiple bands of DRE's could be determined by calculation and did not need to be tested.
Thus far in this discussion, the strength aspects of the uppermost portion of the T/TS joint has involved strictly axial pullout loads on the tube. Other loads, such as axial compression, wemseimm 3-2
~
- = -
i i
l bending and torsion about the tube vertical axis were considered but were judged to be negligible. For example, the axial compression, downward, load occurring during large break i
LOCA, hereinafter referred to in this report as LOCA, was judged to affect the tube at the j
upper SRE to an insignificant extent.. If the T/TS interface within the L* distance were not leaktight and permitted the LOCA pressure to penetrate to the DRE(s), radial compression i
would act on the DRE(s). The within-TS tube collapse strength characteristics in the presence
{
of axial and circumferential indications are discussed in Section 2. It was pointed out in j
Section 2 that significant margin exists between tube collapse strength and the limiting j
secondary-to-primary pressure differential (LOCA).
1 The effects on the tube of other applied loads such tube static and dynamic bending in the vicinity of the top of the roll expansion were also considered.- However, owing to the.
j significant distance of the intended L* region from the TTS, no tube free span static or -
}
dynamic bending effects are expected at the L* region.
/
i i
3.4 Determination of Resistance to Leakage of Degraded Roll Expansions --
}
Prairie Island Units 1 and 2 i
3.4.1
Background
i There is generally a correlation, for a given tube-to-tubesheet joint design, between the tube.
i l
pullout strength, i.e., force, and tube roll expansion thinning; there is a desired thinning range j
for maximum strength. The desired thinning is achieved by the use of a certain roll expansion tool operated in a given stalling torque range. This facet of power plant heat exchanger design decreased in importance approximately two to three decades ago, with the l
widespread use of welded-and-rolled joints because of the addition of the weld to the j
structure. The resistance to leakage (RTL) of a given roll-only design was generally j
consistent with the maximum strength requirement for the design. However, the RTL of the i
expanded portion of a welded-and-rolled design typical of nuclear steam generators has not j
been extensively studied because of the very good sealing and strength properties of the weld.
i Therefore, the RTL of short lengths of the rolled Prairie Island Units 1 and 2 design is not available. In order to determine the RTL of the Prairie Island SG (7/8 inch) tubing, results from tests conducted for an L* program for 3/4 inch tubing were utilized and extrapolated to the 7/8 inch tubing of Prairie Island Units 1 and 2.
i
~
3.4.2 Applicability of 3/4 Inch Tube Tests i
i Resistance to leakage through the interface between a rolled tube and the tubraheet is j
determined by the axial length (or simply length), width (tube circumferenc e) and other geometrical characteristics of the interface. The length was a parameter in the 3/4 inch tube tests of Appendix A. The leakage flow is one dimensional, along the length of the interface, parallel to the tube axis. The total leakage will be proportional to the circumference of the i
tube. If the test conditions, length and other interface geometric characteristics are the same, leakage for a 7/8 inch tube test article (and in a steam generator using 7/8 inch tubing) would be greater than a 3/4 inch tube test article by the ratio of the tube expanded outside diameters, j
l.17. This 1.17 ratio on 7/8 inch to 3/4 inch tube diameter leakage conservatively ignores i
l j
WPF/NSP/121494 3-3 l
i I
f i
that the crack leak rate for a given crack length would be lower for the 7/8 inch tubing, due L
to the thicker tube wall (0.050 vs. 0.043 inch).
l j
The key parameter which determines interface flow resistance is the contact pressure between j
the rolled tube and the tubesheet. Higher contact pressure results in higher flow resistance.
. Other geometric factors which affect the interface flow characteristics, such as the tube and i
tubesheet surface finishes, are expected to be approximately the same for 3/4 inch and 7/8 j
inch tubes. The interface flow resistance is therefore a function only of the contact pressure, j
interface length and tube OD circumference.
i The contact pressure is determined by the tube and tubesheet characteristics and the roll l
l process. Test programs have been performed to simulate the interface of a tube-to-tubesheet-j roll expansion for both 3/4 inch tubes (shown in the appendix) and 7/8 inch tubes l
(Section 2.1). Evaluations of the results of these two test programs have shown that the 2
contact pressure is more than 20 percent higher for the 7/8 inch tubes than for the 3/4 inch tubes. If the 3/4 inch test leakage rates are therefore multiplied by the ratio of the expanded i
OD's of the tubes,1.17, the resulting numbers represent conservatively high leakage rates for j
equivalent 7/8 inch tube leakage tests (and steam generators using 7/8 inch tubing). It would i
be reasonable to consider the effects of the diameter increase on leak rate and the additional resistance afforded by the higher contact pressure for the 7/8 inch tube, in comparison with l
that of the 3/4 inch tubes, to cancel each other. Since test data are unavailable to assess the
[
beneficial impact of the 7/8 inch contact pressure on leakage, for conservatism, the benefit of increased contact pressure will be neglected and the leak rate data for the 3/4 inch tubes will j
be adjusted upward by 17 percent. These data are contained in Table 3-3. The 3/4 inch j
tests and test results are described in the appendix to this report. The adjustment of the leak j
rates, to account for the increased leakage cross-sectional flow area of the 7/8 inch tube j
configuration, in comparison with that of the 3/4 inch tube configuration, is shown in Table 3-3.
i j
3.4.3 Resistance to Leakage i
The purpose of this section is to determine tube-to-tubesheet joint resistance to leakage 4
j for application of L* plugging criteria for Prairie Island Units 1 and 2. Testing conducted for j
3/4 inch OD x 0.043 inch wall thickness tubes is extrapolated to determine RTL for the t
7/8 inch OD x 0.050 inch wall thickness tubes of Prairie Island Units 1 and 2. The tests for i
i 3/4 inch tubing, described in Appendix A of this report, were performed with T/TS test i
specimens fabricated of short sections of prototypic tubes rolled into collars which provided the same structural compliance as a unit cell of the TS. The tests were performed at prototypical pressures, temperatures and tube axial loads for normal operation (N.O.), the most stringent faultmi primary-to-secondary condition, Feed Line Break (FLB), and the most j
stringent faulted secondary-to-primary condition, Loss of Coolant Accident (LOCA).
l
~
i The effect of prototypical loads on 3/4 inch tube joint leakage was addressed in the leak test.
l For example, a load such as primary-to-secondary pressure differential during N.O., causes direct pressure effects on the tube in the L* region to increase the T/TS radial contact pressure, s,.
It also causes tubesheet bending. The differential pressure increases the TfrS
]
radial contact pressure for all tubes in the TS, peripheral as well as interior, i.e.,
}
wemsem!494 3-4
i away-from-periphery, tubes. However, the effect of upward bending of the tubesheet due to differential pressure is different for roll transitions at the top of the tubesheet (i.e., full depth expanded) than for roll transitions near the bottom of the tubesheet (i.e., partial depth roll expansions). For full depth expansions, the tube region near the roll transition (TTS) j experiences a decrease in radial contact pressure between tube and tubesheet for interior tubes i
and therefore a decreased RTL. This bending effect, which is detrimental for interior full depth expanded tubes, was accounted for in the laboratory tests conducted for 3/4 inch tubing.
For the Prairie Island Units I and 2 partial depth roll expansions, the tubing below the roll transition experiences an increase in T/TS radial contact pressure for interior tubes, since the roll transitions are well below the neutral axis of the tubesheet. Peripheral tubes would experience much less of a related increase. However, it is clear that in all of the Prairie Island Units 1 and 2 roll expansions the RTL would either be unaffected or enhanced by tubesheet bending. Therefore, the RTL tests for 3/4 inch tubing, which simulated a loss of l
RTL due to tubesheet bending, provide conservative results for Prairie Island Units 1 and 2.
The basis for applying the 3/4 inch test results to the Prairie Island Units 1 and 2 steam generator tubing is furnished in Table 3-2.
The acceptable leakage for a reasonable number of 3/4 inch OD tube joints was established for normal operating and faulted conditions. h was obvious that for all acceptable-leakage cases, the 7/8 inch joint must also exhibit acceptable strength. It was concluded that the 7/8 l
inch tube (L*) strength determination could be decoupled from the leakage determination, as l
it was for 3/4 inch tubes. However, prototypical radial contraction or cxpansion of the tube, causing a reduction or increase in the T/rS radial contact pressure (s,), was a part of the leakage determination. Leakage testing of the 3/4 inch tubes was performed at prototypical temperatures and pressure differentials. The primary-to-secondary pressure differential effect and a secondary-to-primary pressure differential effect were expected to have negligible impact on the tube joint strength and strength testing was performed at room temperature and
)
without pressure differentials.
i 3.5 Leakage Criterion and Calculated Leakage from Degraded Roll Expansions --
1 Prairie Island Units I and 2 3.5.1 Leakage Criterion The leakage acceptance criterion was developed for evaluation of the projected leak rate for Prairie Island Units 1 and 2. Individual acceptance values were developed for differential pressure conditions for normal operation, feedline break and LOCA.
Normal ()pration -
The total normal operation primary-to-secondary leakage allowed from both steam generators is specified in the Prairie Island Units 1 and 2 Technical Specitication as I gpm (1440 gpd).
This is further interpreted to allow leakage of up to I gpm in either SG, not to exceed a total of I gpm in both SGs. Both units have imposad administrative limits of 0.3 gpm during normal operation. It is reasonable to allocate only a fraction of the permissible leakage to tubes with degradation left in service using the L* criteria. Therefore, the permissible leakage.
1 wrrm an w 3-5
i 1
j for L* considerations is selected as 150 gpd (0.1 gpm); the leakage contribution from L* tubes
{
must not exceed this value. The number of tubes to which the L* criteria can be applied has arbitrarily been set at [
]* tubes per steam generator. For the specified number of 1
applicable tube ends, an allowable normal operating L* tube leakage limit of
[
]* per L* tube was established (based on the i
equivalence of 75,000 drops per gallon).
Faulted Condition: Feedline/Steamline Break Postulated feedline break conditions provide the maximum primary-to-secondary differential pressure across the tube, and a pressure differential value of 2650 psi is used based on generic
{
equipment specifications for the 51 Series steam generator. During a postulated steam line break (SLB), RCS pressure will be limited to the pressurizer safety valve liftoff setting of i
approximately 2575 psia, resulting in a pressure difference across the SG tubes of 2560 psi, Previous L* evaluations have limited the maximum L* contribution to primary-to-secondary leakage during a SLB to 0.333 gpm in the faulted loop, based on limiting the maximum L*
faulted condition leakage to 1/3 of the Technical Specification primary-to-secondary leakage limit. For Prairie Island, this value will be maintained. For the purposes of this report, the maximum number of tubes to which the L* criteria can be applied is [6 i
jm, The NRC staff has determined that, provided RCS and secondary side radioactivity levels are maintained within current Technical Specification limits, offsite doses will remain well within the accepted guidelines of 10 CFR 100. Since realistic L* leakage from up to [
]* tuba ends will result in faulted condition leakage far below the current Technical Specification -
limits for normal operation, offsite doses will remain within the current analysis results.
Further conservatism is provided by assuming the faulted condition leakage is constant over the course of the event. Upon assuming the SLB, a rapid RCS cooldown event occurs, thereby dropping RCS pressure to approximately 1000 psi. RCS pressure is eventually reestablished by safety injection and residual heat effects withir, the core. This pressure reestablishment of near normal RCS pressures would be projected to occur between 10 and 30 minutes after the event, based on comparison with similar plants. Hence, actual faulted condition leak rates due to reduced RCS pressure would be lower than those obtained by assuming constant leak rates throughout the event.
i wermSPnna 3-6
i L
)
l Faulted Condition: LOCA i
f Typically, the small leakage from steam generator tubes into the primary system is not a significant consideration in the analysis of a postulated LOCA, provided the postulated LOCA i
i in-leakage is less than the Technical Specification steam generator primary-to-secondary tube leakage limit. The test data of Table 3-3 indicate that the LOCA leakage is two orders of '
}
magnitude less than the normal operating condition test leakage. - Additionally, it is generally :
l j
accepted that for cases where the secondary system pressure is greater than the primary l
system pressure, the potential ID initiated cracks of the roll expansion would tend to close i
due to pressure effects. These conditions provide a sufficient basis to consider LOCA.
l in-leakage insignificant upon implementation of the criteria.
3.5.2 Leakage Projected for Prairie Island Units 1 and 2 i
j.
Normal Operation j
l The test data contained in Table 3-3 indicate that of the sixteen 3/4 inch tube samples tested at non-faulted conditions, only three exhibited leakage, or approximately 19 percent of the test -
l tubes leaked. In actuality, it is reasonable to assume that this distribution would bound the j
L* tubes which exhibit leakage in the field. This assumption is validated by the conservatisms included in the leakage testing. These conservatisms include the fact that the l
test primary-to-secondary pressure differential (AP) of 2250 psi is 657 psi greater than the i
limiting Prairie Island Units 1 and 2 normal operating AP of 1593 psi. His AP was
)
necessary in the test because the secondary side pressure was ambient. Subcooling of the j
primary side liquid, i.e., use of a higher pressure at the prototypic temperature, was necessary j
to insure that the fluid state at the entrance of the T/TS interface was prototypic, thereby -
I l
causing a conservative pressure differential for the test. The data of Table 3-3 indicate that j
the percentage of tubes that leaked and the leak rates are directly affected by the test pressure.
Based on this fact, the percentage of field tubes to which L' is applied for Prairie Island j
j Units 1 and 2 will be bounded by 19 percent is reasonable. Therefore, L* tube leakage will be bounded by 19 percent of 150 gpd, or approximately 29 gpd. Further conservatism is l
provided by the fact that the allowable leakage limit is more than twice the maximum j
adjusted (for 7/8 inch tubing) normal operating condition test data point. The bounding j
leakage value of 29 gpd uses the allowable limit of 0.00017 gpm (12.75 dpm) per indication.
1 From Table 3-3, the maximum measured leak rate at normal operating conditions was 0.00006
{
gpm (4.45 dpm) and the other two leakers ranged from 0.12 to 0.6 dpm. He average leak.
rate for the 16 tests at normal operating conditions is oiily 0.25 dpm. Consequently, the j
expected operating leakage for 600 L* tubes left in service would be negligible and much less j
than the 29 gpd bounding limit.
l'
[
Alternately stated, the average leak rate per L* tube times the number of L* tubes must be g
less than the operating leak limit goal of 150 gpd as developed in Section 3.5.1. For j
operational considerations, a fraction ~ of the 150 gpd limit is desirable such as the bounding limit of < 30 gpd (0.02 gpm) developed above. For 600 L* tube ends, the desirable average
]
leak rate at normal operating conditions would then be < 0.000033 gpm (2.5 dpm). As noted j
above, this average leak rate would exceed the required leak rates for as-manufactured J
i p
4 l
WPFMSPnnos 3-7
_J
l j
hardroll expansions. Thus, normal operating leakage for an L* of 0.5 inch with
]
as-manufactured hardroll expansions is expected to be negligible.
For applications of field reroll expansions (extension of hard roll expansion above the as-manufactured length), the reroll qualification program must demonstrate that the normal i
operating leakage limit of 0.1 gpm from up to 600 L* tubes is satisfied for an L* length of 0.5 inch. However, the average leak rate of 0.000033 gpm for up to 600 L* tubes left in i
service, while not required for Tech. Spec. limits, is a reasonable guideline value for reroll expansions. Westinghouse tests for rerolls in crevice sludge, including hard sludge tests, were i
performed for reroll lengths of about [
]"' inch and resulted in average leak rates of
[
]'", with the larger leak rates j
typically occurring in crevices fully packed with very hard sludge. For a 0.5 inch reroll 1
length, the average leak rate would be about [
]'"
Thus, the allowable leak rate limit of 0.1 gpm for L' application to 600 tubes can be achieved with reroll expansions having L* = 0.5 inch. To achieve the desirable goal (in contrast to the requirement of < 0.1 gpm) of < 0.02 gpm normal operating leakage with reroll expansions, a j
number of options can be applied in the implementation of L*
These options include I
arbitrarily increasing the applied L* 1ength to about 1.0 inch or limiting the number of 0.5 inch L* lengths applied to about 300 tubes for crevice conditions with hard sludge. The Westinghouse reroll qualification tests indicate that the diameter difference between the reroll and as-manufactured expansions can be applied as a conservative indicator of potentially hard crevice conditions, and the limited number of tube ends or larger L* lengths could be applied only for the rerolls indicative of hard sludge conditions.
)
Feedline Break
{
The 3/4 inch tube FLB leakage test results are shown in Table 3-3, along with the projected results for 7/8 inch tubes. Although the FLB values are an order of magnitude greater than the N.O. values at respective X's, the FLB leakages are small relative to the initial FLB leak
]
acceptance criteria of [
]"' dpm, per tube end. The fact that the (3/4 inch) FLB samples were made with smaller torques than the N.O. samples correlated with the increased flow for the FLB samples. He slightly higher AP for the FLB test had a negligible effect on the flow i
because, based on Henry's Law for critical flow through small passages, the FLB flow would i
only have been expected to be larger than the N.O. flow by approximately the ratio of [
]"' percent.
4 The conservatism of the 0.333 gpm allowance for all L* tubes is shown when the leakage j
data is analyzed. The maximum adjusted leakage data point of [
]"'
Considering the conservative nature of the leak test setup
[
]"', postulated leakage will be bounded by the 0.333 gpm value by a large margin.
1 i
[
e 4
werensritmisa4 3-8
l l
i i
[
l-
~ ).'"
For reroll applications, the above acceptance criteria of [
]'" dpm should apply and also -
L result in negligible accident condition leakage. Westinghouse tests on reroll expansions into-sludge filled crevices indicate that accident condition leak rates are only moderately increased 3
over that at normal operating conditions. For the as-manufactured conditions described above, the smaller torques applied for the FLB samples, to simulate relaxation of the expansion,
)
result in significant leak rate increases and percent leakers compared to the normal operating condition tests. Westinghouse reroll tests [
]'" showed an average leak rate of j
l
[
]'" with [
]'" of the specimens leaking. Although these tests I
l were performed for about a [
]"' expansion, scaling the leak rates to a 0.5 inch I
expansion would result in large margins against the leak rate acceptance criteria. Thus, the Westinghouse tests show that a reroll L* length of 0.5 inch would be adequate to limit accident condition leakage to acceptable values.
3.6 General Application of the 3/4 Inch Tube Axial Locdbearing Test Results to
)
Prairie Island Units 1 and 2 The strength tests of the 3/4 inch tube degraded roll expansions, described in Appendix A, i
Section A.3, were applied to Prairie Island Units 1 and 2 by means of a fracture mechanics evaluation. This evaluation is provided in Section 3.7. The 3/4 inch data applied to the 7/8 inch case because of the commonality of tube joint materials, respective material properties and fabrication processes between the two designs. These features are shown in Table 3-4.
3.7 Application of 3/4 Inch Tube Test Results for the Axial Loadbearing Capability of l
Degraded Roll Expansions to Prairie Island Units 1 and 2 i
l 3.7.1 Maximum Axial Load Criterion l
Sinele Band Deeradation The maximum axial tensile load which a degraded roll expansion (DRE) must bear will be the.
most stringent of the N.O. and FLB axial loads. The LOCA axial load is axially compressive and cannot cause tube pullout.
l Multiole llanJ_Deeradation If degradation occurs in discrete circumferential bands of one or more linear indications, i.e.,
as single band degradation (SBD) or arrays in a roll expansion above the F* elevation, thereby preventing use of F*, it may also occur in multiple discrete bands below the F*
elevation. This is termed multiple band degradation (MBD). As discussed earlier, it's desired to avoid quantifying each of the possible several degradation bands to apply the pullout load -
to the T/TS weld. Instead, it's desired to apply the load only to the uppermost sound roll portions (expansions) which are interspersed with the uppermost degradation bands. For'the wermsmum 3-9
l 1
1 1
purpose of simplicity, each degradation band requires the same strength. This strength is the same as that fc a SBD Therefore, for the MBD case, the number of individual bands which must be quantified is one less than the number of SRE's (N) needed to react the pullout load. If N is 3, N-1 is 2. Conservatism may be added to this calculation in the form of a larger N than is needed.
Normal Operation The tube ultimata k.m required for normal operation at the Prairie Island Units 1 and 2 condition resulting in the largest N.O. AP (15 percent steam generator tube plugging, Tw of 592*F) is approximately [
].* A safety factor of three is applied to this load. Therefore, the tube axial loadbearing requirement is
[
]* lb. This load is the highest for the three conditions considered.
Feedline Break The Prairie Island Units 1 and 2 tube ultimate load required for FLB is approximately [
]* A safety factor of [
]* for allowable stress for faulted conditions) is proposed for this load.
Therefore, the tube axial loadbearing requirement for FLB is [-
]* lb. As discussed in Appendix A, this load need not be further augmented for dynamic loading effects during FLB i
because of the sequence of events during an FLB.
Loss of. Coolant Accident l
An axial loadbearing requirement for a LOCA event acts to move a tube downward, the
. opposite of pullout. Therefore, this condition does not apply to pullout. (However, the DRE l
must not collapse under LOCA conditions. This condition was addressed in Section 2 and collapse for the 7/8 inch tube case is not expected.)
3.7.2 Evaluations Application of the L* approach requires that the degraded region has sufficient strength to-resist axial pullout forces with the appropriate safety margins. The limiting case is the axial force generated by a pressure load equal to 3 times the operating pressure differential. For Prairie Island Units I and 2, this force is [
]* lbs., generated by a 3AP value of l
4779 psi. Ifligaments of sound material in the degraded region are not subject to significant l
local bending loads, only 28 percent of the mean circumference of the expanded region need I
be present to meet the axial strength requirements. Typically, degradation found in roll expansions occurs as multiple parallel cracks. If these cracks are inclined at an angle to the l
tube axis, as shown in Figure 3-4, then an axial force leads to bending of the ligaments between cracks and rotation of the tube and crack array as deformation proceeds. The axial strength of a tube with an array of" slanted" cracks depends on the slant, a.k.a., inclination l
crack angle and, to a lesser extent, on the number of cracks present. The following paragraphs describe a pull strength model for slanted crack arrays. This model was benchmarked by pull strength tests on 3/4 inch diameter tubing. A comparison of test data and a pull strength design curve for 3/4 inch diameter tubing is shown and a pull strength wemsenum 3 - 10
=.
1 l
- i-design curve for 7/8 inch diameter tubing is presented. This latter curve is appropriate for Prairie Island Units 1 and 2.
i As a tube with an array of parallel slanted cracks is loaded axially, deformation proceeds as follows. Initial clastic elongation is followed by plastic yielding. Depending on the number of cracks and the crack inclination angle, yielding may first occur in the virgin unrolled tube i
section or in the crack array. Large numbers of cracks and high crack angles favor yielding of the crack array. Yielding of the crack array proceeds by plastic bending of the ligaments between cracks and subsequent rotation of the cracks toward the longitudinal axis of the tube.
This is easily visible in pull strength tests. As rotation occurs, the applied axial force must j
increase substantially since the effective moment arm for ligament bending is continually i
decreasing. Strain hardening of the ligaments adds to the geometric hardening produced by crack rotation. As the load is increased, general yielding of the. virgin unrolled tube section may occur, followed by strain hardening and then yielding of the rolled but uncracked tube section. If deformation proceeds to this point, the rolled tube peels away from the tubesheet and the axial pull strength is limited by the tube-to-tubesheet weld. At some point, depending l
on the crack morphology, fracture will terminate the process of plastic deformation.
Using the notation of Figure 3-4, the axial load required to yield a tube with an array of l
slanted cracks is
[
]*
This expression only considers ligament bending and assumes deformation is confined to the initial minimum ligament crossection. The axial plastic displacement resulting for yielding and then rotation of the crack array is given by l
[
jw If the crack array plastic displacement is added to the baseline load displacement record of an uncracked tube, then the computed load displacement records closely approximate actual measured load displacement records. The load displacement record for the uncracked tube essentially provides elastic tube and test fixture displacement and an indication of yielding m the uncracked section at high loads. Figure 3-5 shows a test of a 3/4 inch diameter tube terminated by ligament fracture. Figure 3-6 shows the test record for another 3/4 inch diameter tube where the test load has become high enough to yield the uncracked as well as cracked sections. (Note: Figures 3-4 through 3-7 are repeated in Appendix A.)
l 1
The prediction of ultimate pull strength load as opposed to yield load is an elastic-plastic fracture mechanics problem rather than a plastic collapse problem because maximum loads are caused by crack tearing (beyond some minimum crack angle). A J-integral approach has been applied to the problem of estimating the onset of crack tearing with good success. The equation describing the plastic load-displacement behavior of the crack array was integrated to related plastic work, crack length and displacement, The compliance defm' ition of J was then used to compute J, Since the value of J for the onset of crack tearing in Alloy 600 is very high, the neglect of eltstic contributions to J is appropriate.
werssnmm 3-11 J
~.
i l
In pull tests of 3/4 inch diameter Alloy 600 tubes containing slanted EDM notches to simulate cracks, maximum loads were consistent with an applied J of about 3600 in.-lbslin.' This is true when the crack _ array limits the pull strength.- For low numben of cracks and small crack angles, yielding of the uncracked sections developed and pull strengths were limited by the tube to tubesheet welds. The use of narrow EDM slits to simulate cracks in tests of the 3/4 t
inch diameter tubing is justified by the fact that the initial notch root radius was much smaller than the final blunted crack tip crack opening displacement (COD) which led to crack tearing.
For the 3/4 inch diameter tubing, calculated pull strengths as a function of crack angle were adjusted to high temperature, lower tolerance limit (LTL) material properties to construct a pull strength design curve. (Lower bound statistical tolerance limits, LTL, for pull strength-values were computed in accordance with the accepted industry practice such that there is a
[
]" than LTL values.) At small crack angles, actual test data were adjusted to LTL properties to arrive at design curve values..The crack length selected for the design curve was 0.5 inch. This is a reasonable bounding value for throughwall crack lengths in standard roll transitions.
Figure 3-7 illustrates the pull strength design curve for 3/4 inch diameter tubing along with actual test data. The good consistency between test results and calculations for 3/4 inch diameter tubing provides the basis to extend such calculations to construct a pull strength design curve for 7/8 inch diameter tubia; I
l A pull strength design curve for 7/8 inch diameter was constructed by making appropriate changes in the input to the slant crack pull strength model. The tube diameter, wall thickness and LTL tensile properties were adjusted. The J level at the onset of tearing was kept at the same value. This was somewhat conservative since at the very low material thicknesses of l
interest, a 20 percent increase in the toughness might be expected for the larger wall thickness tubing. The crack length was maintained at a value of 0.5 inch. He baseline pull strength at l
low crack angles was obtained by multiplying the experimentally based value for 3/4 inch diameter tubing by the ratio of the circumferences of the 7/8 and 3/4 inch diameter tubing and -
then adjusting for differences in LTL flow strengths. The 7/8 inch diameter pull strength design curve is shown in Figure 3-8. It is about [
l
]*
3.7.3 Results Using the normal operating plant conditions for 15 percent steam generator tube plugging and a Tso, of 592'F indicates that the most stringent axial load developed during normal operating conditions for Prairie Island Units 1 and 2, [
]* degree inclination angle.
wermemim 3 - 12
~
L l
l l
The curve of Figure 3-7 should be applied to tubes with relatively well characterized ECI's,
[
e.g., tubes for which substantiated rotating pancake coil (RPC) ECT data are available. The
(
strength criterion per this curve, must be met for tubes with eddy current indications.
3.7.4 Conclusion
\\
(
With sufficiently quantified information about degradation in the anchor regions of the Prairie i
Island Units 1 and 2 S/G roll expansions, use of the Design Curve is expected to provide I
relicble operation. (The leakage criterion for this case is discussed elsewhere in this report and involves an upper limit on the number of tubes per steam generator which may be dispositioned.)
j 3.8 Pullout Load Reaction and RPC Inspection Lengths - Prairie Island Units 1 and 2 l
3.8.1 General j
The methods used to inspect, evaluate, and define the tube degradation for the application of the L* altemate plugging criteria require an RPC eddy current probe or other advanced l
inspection method. If sufficient length of sound expanded tube exists in the upper portion of the roll expansion, the lower portions of the expansion contribute relatively little to the structural integrity and resistance to leakage of the tube. Criteria have been developed which limit evaluation to the more important upper portion of the tube expansion.
~
1 The analysis to support limited evaluation of the degradation in a tube is based on a -
combination of the methods used to qualify the F* and basic L* criteria.' The analysis demonstrates that sound portions of expanded tube between degraded sections can provide for structural integrity of the tube. In the same manner that the friction force between the tube and the tubesheet in an F* tube will resist pullout, the sound portions of a tube with degradation meeting the L* criteria will provide a frictional force to resist pullout. The calculation of the length with degradation within the portion of the expanded tube required to l
resist pullout includes the possible reduction in friction force next to the degraded portions due to end effects. The length of tube required to resist pullout plus the lengdi of the -
l degraded tube portions establishes the length of tube which must be inspected and evaluated j
by rotating pancake coil or other advanced inspection methods. Therefore, rather than survey l
the entire length of the RE, to the T/fS weld, only that portion near the RT which will l
provide adequate anchoring for all axial loads need be surveyed. This length, in addition to l
the SBD or MBD within it, is the minimum length, beginning at the BRT and extending I
downward, and is referred to as the pullout load reaction length (PLRL).
3.8.2 Single Band Degradation The PLRL for this case consists of four parts, the SRE within L* consisting of the central and end-effect portions (the recommended 0.5 inch), pips the required SRE immediately below the SBD. Calculations for N. O. and FLB are shown in Table 3-5. The limiting case was N. O.,
for which a PLRL'of [
]'" inch was determined. Addition of the SBD length,0.50 inch, and the ECT uncertainty of 0.10 inch to the PLRL of [
]'" inch resulted in a length of i
l wermonum 3 - 13
nu.
=,._.a_
e a
i~
.m..
.m.
..mm I
i f
[
]'" inch of RE to be RPC inspected. The RPC inspection length is the length below the bottom of the roll transition.
i j
3.8.3 Multiple Band Degradation l
j The PLRL for this case consists of six parts, the SRE within L*, i.e., the recommended 0.5 inch consisting of the central and end-effect portions, the uppermost two degradation bands l
and the two SRE's below these two respective bands.. Calculations for the N.O. and FLB are i
shown in Table 3-5. The PLRL for N.O. conditions was 1.34 inch. An RPC inspection j
length of [
]'" inch for this N.O. condition was determined by addition of the' two i
degradation band lengths and the ECT uncertainty of [
]'" inch to the PLRL. For the 4
FLB conditions, the PLRL was determined to be 0.88 inch. Addition of the postulated two j
degradation bands and the ETC uncertainty of [
]'" to the PLRL resulted in a length of 1
l
[
]'" inch of RE to be RPC inspected.
l 3.8.4 Technical Specifications Requirements i
- In specifying the length of degraded tube required to be inspected by an advanced inspection j
technique, for the plant Technical Specifications, the information provided above can be combined into a single set of criteria. The calculated length of tube required to be inspected is 1.81 inches for the SBD case. Inspection of the top 1.81 inches rounded off to 2.0 inches, 4
)
of the tube should determine if there is at least [
]'" inch, plus ECT uncenainty, of sound
[
tube with no more than [
]'" band of axial degradation separating the sound portions and 1
with an L* of 0.50 inch plus ECT uncertainty. A similar calculation for the MBD case, for j
the postulated two degradation bands of 0.25 inch axial extent, axially separated by 0.50 inch, discussed above, results in a length of 1.98 inches rounded off to 2.0 inches to be inspected.
1 3.9 Summary of L' Evaluation 1
l 3.9.1 General j
The semi-empirical evaluation uncoupled the strength and projected leakage effects. The --
]
issues were determined separately.
1 l
3.9.2 Leakage.
.i Normal Operation i
I i
Average, projected per-joint leak rates for the N.O. condition for all X's from 0.25 through t
2.00 inches, were low relative to the permissible, per-joint leak rate of [
]'" for 600 joints. This was based on the arbitrary selection of 150 gpd as a fraction of the 1 gpm I
(1440 gpd) Tech Spec. allowable leak rate for both S/Gs during N.O. for L* tube ends.
1 Using the 150 gpd leakage limit,600 of the tube ends in the S/G could be dispositioned at all i
of these X's (plus a typical ECT uncertainty). The average per-joint, projected leak rates in drops per minute are [
f
]'" for the X's of 0.25, 0.50,1.00 and 2.00 inches,
^
j wemsenmu 3 - 14
\\
4 d
respectively. Thus, the expected leak rates at normal operating conditions for the as-i i
manufactured expansions are negligible and well below acceptance limits.
i For field reroll applications, the reroll qualification program must demonstrate that the normal operating leakage limit of 0.1 gpm is satisfied for an L* length of 0.5 inch and up to 600 L' i
. tubes. Westinghouse tests for rerolls in sludge packed crevices show that this normal j
operating leakage limit can be satisfied for reroll conditions. However, if it is desirable to l
reduce operating leakage to well below this limit when reroll applications in hard sludge l
packed crevices are encountered, increased leakage margins can be obtained by increasing the
{
undegraded length below the reroll BRT to about 1.0 inch or by reducing the number of.
j rerolls in hard sludge to the order of 300 tube ends as discussd in Section 3.5.2.
i Eau. lted Conditions i
}
Based on the FLB/SLB permissible per-joint leakage criteria of [
]'^* and the small i
average projected leak rates, as many as 600 joints could be dispositioned at X's of 0.25 to
]
2.00 inches with resultant primary to-secondary leakage of only a fraction of the Prairie Island i
FSAR value of 1.0 gpm. The average per-joint, projected leak rates were [
)*^* for X's of 0.25,0.50,1.00 and 2.00 inches, i
respectively. For field reroll applications, the reroll qualification program must demonstrate i
that the accident condition leakage limit of 1,0 gpm is satisfied for an L* length of 0.5 inch i
applied to 600 tubes. Westinghouse tests for rerolls in sludge-packed crevices show leak rates j
well below the acceptance limits, and L* lengths for as-manufactured conditions would be i
satisfied for reroll conditions.
i l
LOCA The average, projected LOCA leak rates for all X's, were nearly~two orders of magnitude less i
l j
than the N.O. average, projected leak rates averaged for all X's and two orders of magnitude less than the comparable FLB values. As stated in Section 4, projected LOCA leakage (in-i leakage) related to L' is insignificant. No criteria needed to be developed for it.
l
~
Limitine Case i
]
Of the two important conditions at an X of 0.50 inch, N.O. and FLB, neither condition is close to being limiting. At this X, the average, per-joint projected leakage is an order of magnitude less than the permissible value, for the as-manufactured expansion. For field reroll i
i applications, Westinghouse tests indicate that the acceptance limits on leakage are satisfied j
with normal operation as the more limiting condition. Additional leakage margin for hard packed crevices can be obtained by increasing the undegraded tube length above 0.5 inch or i
by limiting the number of tubes to which L* is applied for hard packed crevices.
l 3
3.9.3 Strength The calculated axial strength of DRE's far exceeds the required axial load for the most j
stringent postulated plant condition. The design curve of Figure 3-7 on crack angles for up to.
4 wemsenum 3 - 15 t
i j
.I 30 indications must be satisfied for L* application to either as-manufactured or field reroll joints. Based on the limiting assumption that the L* region, of length 0.50 inch provides no axial fixity, the strength of the uppermost degradation band was determined. For the purpose ofillustration, a plant condition was postulated wherein the uppermost degradation band consisted of 10 ECI's, approximately 0.50 inch long and inclined approximately 15 degrees from the axis. The strength of this configuration, based on a fracture mechanics analysis which was developed for the 7/8 inch tube configuration from ultimate pull force test data in the 3/4 inch tube program, was greater than approximately [
]'" lb. The required strength was 2973 lb. (including the safety factor of 3), and occurs at the N.O. condition.
The margin exceeded a fa:::cr of approximately 3.0. If future degradation were to increase to 14 ECI's per DRE and to continue to appear off-axial, with a $ of 15 degrees, this allowable load would reduce to approximately [
j.u With the axial fixity afforded by the SRE within the L* region, approximately [
]'"
This maximum axial reaction force must be transmitted through the first degradation band. The fracture mechanics analysis has determined that for the 10 ECI case, with a degradation inclination angle of 4 = 15 degrees, a decollared tube with degradation of this extent will support an axial load of [
]'". Therefore, a more realistic comparison of DRE axial strength at $ = 15 degrees, of [
]'".
3.9.4 Recommended L* Criteria l
With sufficiently quantified information about degradation in the upper approximately 2.00 inch length of degraded roll expansions, use of the recommended L* value of 0.50 inch, plus ECT uncertainty ([
]'" inch was used for an example), is expected to provide continued l
reliable operation of the Prairie Island Units I and 2 S/G's.
Inspection of a 2.0 inch or longer distance should determine if there is at least [
]'" inch, plus ECT uncertainty, of sound tube with no more than two bands of axial degradation separating the sound portions. The largest & permissible for the ECI's in a degradation band i
is 30 degrees for up to 30 axial indications in the band. (The 4 = 45 degrees curve is presented only for trending purposes.) A reasonable number of tube ends in a S/G which may be dispositioned using L* criteria is 600. When there is no degradation in the uppermost 1.07 inches (plus NDE uncertainty) below the bottom of the as-manufactured or field reroll transition, the F* criterion applies and the tube end would not be counted in the 600 tube guideline.
werusen:1494 3 - 16
+
Table 3-1 Technical Approach to Meet Regulatory Requirements Postulated Mode Condition Tube Rupture Tube Leak Tube Collapse Leak Before Break Normal Preload Preload N/A Operation Analysis Analysis Preload Pull Tests Proof Tests Analysis a
Proof Tests Leak Tests
- FLB + SSE Preload Preload N/A a
Analysis Analysis Leak Tests Pull Tests Proof Tests Proof Tests Leak Tests
- *LOCA +
N/A Preload Limit N/A SSE Analysis Analysis Leak Tests Existing i
Test Data
- FLB: Feedline Break (Note: FLB is most stringent Faulted Condition)
SSE: Safe Shutdown Earthquake I
- LOCA: Loss of Coolant Accident P
WPF/NSP/120W4 3 - 17
l Table 3-2 Basis For Use of 3/4 Inch Tube Joint Leakage Data for Projected Leakage for Prairie Island Units 1 and 2 Degraded Roll Expansions Prairie Island Units 1 and 2 PARAMETER 3/4 Inch Tube 2/_8 Inch Tube Tube material
. Alloy 600 MA Alloy 600 MA Tubesheet Mat'l A508 Cl. 2 A508 Cl. 2 Tube-to-TS Joint Sample
- Design Structural Evaluation Structural Evaluation
- Mat'Is Tube - prototypic Tube - prototypic (used in analysis)
Unit Cell (Collar) -
Unit Cell - prototypic Approximate yield (used in analysis) strength of S/G TS Fluid Conditions Prototypic or Prototypic e.g., primary side pressure "overtest" (used in analysis)
Tubesheet deformation Tested and Structurally analyzed (due to applied loads) structurally analyzed l
+
wprssenxeu 3 - 18
1 Table 3-3 3/4 Inch Tube Roll Expansion Leakage Test Resuits and Projected Leakage for Prairie Island Units 1 and 2 (7.8 inch tubes)
Roll Sample Expansion Plant Differential Temp., *F Leak Rate, dpm Number Length. in Condition Pressure. psi Planned Actual 3/4" tube 7/8" tube a,c.e 9
8 man m
werssennes4 3 - 19
Table 3-3 (continued) 3/4 Inch Tube Roll Expansion Leakage Test Results and Projected Leakage for Prairie Island Units 1 and 2 (7/8 Inch Tubes)
Roll Sample Expansion Plant Differential Temp.,'F Leak Rate, dpm Number Length. in Condition Prassure. pJi Planned Actual 3/4" tube 7/8" tubs r
m
_,a,c.e I
a l
l i
l NOTES:
\\
Limiting Prairie Island AP at N.O. is 1593 psi, use of higher AP during test is required for subcooling primary side to ensure liquid phase at potential leak path. Higher AP is conservative.
l-Gross Leak; Collar ID had Excessive Roughness.
For all tests, Collar ID surface roughness was "as drilled and reamed", with 250 RMS typical and with actual maximmn of approximately 300 RMS.
l wemSPnH894 3 - 20 I
Table 3-4 Basis for Use of 3/4 Inch Tube Joint Loadbearing Test Data for Projected Loadbearing Capability for Prairic Island Units 1 and 2 Degraded Tube Joints Steam Generator Parameter 3/4 Inch Tube Prairie Island Units 1 and 2 Tube Material Alloy 600 MA Alloy 600 MA Tube Mechanical Properties Prototypical used in test &
Prototypical used in analysis analysis Tube-to-TS Joint Interference Determined by test Determined by test & was l
Fit (Contact pressure) conservative relative to 3/4 mch tube test results Applied Loads Prototypical or "overtest" Prototypical used in analysis l
i l
l wrrmsenxes4 3 - 21 l
i i
. J Table 3-5 Calculation of Pullout Load Reaction Length and Tube Length to be RPC Inspected 1
R,C,0 I
1 l
0 1
l O
t wrrusennsu 3 - 22
.. ~. =
Table 3-5 (continued)
Calculation of Pullout Load Reaction length and Tube Length to be RPC Inspected -
a,c,e i
4 l
a 1
i l.
l l
1 l
i i
l
[
I I
r i
l werssennsu 3 - 23
9 Table 3-5 (continued)
Calculation of Pullout Load Reaction length and Tube Length to be RPC Inspected a,c,e l
1 I
l i
e em 8
wPren us94 3 - 24
i l
i l
l
- l l
Table 3-5 (continued) l Calculation of Pullout Load Reaction Length and j
Tube Length to be RPC Inspected a,c.e i
i I
?
l I
I
(*
I i
l l
i l
l b
i l
i 1
l werennsw 3 - 25 w
c n
- I
~.
I a,c,e '
'l J
i y
e I
l t
l l
l k
.I i
i l
l 1
I i
l 1
Figure 3-1.
Configuration for Tubesheet Region L* Alternate Plugging Criteria
)
l for Partial Depth Roll-Expanded Steam Generator Tubes, Multiple l
Band Degradation.
i l
l wrrmanneu 3 - 26
I i
1 i'
f Figure 3-2 s
i Postulated Leakage Configuration for Prairie Island Units 1 and 2 Steam Generator i
Normal Operation and FLB Conditions
}
l a
i a,c.e l
l 1
i e
1 1
l wrrasen 2*
3 - 27
.... - - -..=...
i Figure 3-3 Postulated I4akage Configuration for Steama Generator LOCA Conditions-r d
1 l
i a,C,c f
i I
i i
l l
I i
1 o
i m
+
l l
i i
1 l
O wrrasenns94 3 - 28
.. ~. _
i J
3
?
=
1, Figure 3-4 i
3 j:
Model for Plastic Collapse j
7/8 Inch Diameter, Mill Annealed, Alloy 600 Tubing l
I a,C,e i
i 4
i
)
-I t
4 5
s-r i
1 1
e i
l 1
i i
J I
l l
l l
1 J
0 i
wrrasenum 3 - 29
)
Figure 3-5 Comparison of Load-Displacement Records, Computed vs. Measured for 30 Slots at & = 45*,3/4 Inch Tubes 8,C,e l
I l
l t
l e
WPFNSP411*
3 - 30
~
! ~ ~ - -
i 1
j Figure 3-6 Comparison of Load-Displacement Records, Computed vs. Measured for 30 Slots at $ = 30",3/4 Inch Tubes 8,C,e I
.1 1
l I
l WPF/NSP/111894 3 -.
.l i
1-4 1
i Figure 3-7 l
Ultiniste Pull Force as a Function of Slot Angle i
for Collared and Decollared Tubes,3/4 Inch Tubes
' 8,C,C i
i 4
1 5
i
{
1-1 i
I, I
3 i
j i
]
1 3
l
}
a I.
I 4
i
{
a 4
),
I 4,
i, 1
i i.
l 4
3 i
t i
i i
l 4
I I
1 I
i i
1 1
s s
1 a
I wPrmspnnsu 3 - 32 j
t w.
-ww w
-r y.
i l
l t
l Figure 3-8.
l Pull Force Design Curve For--
7/8 Inch Tubes at Prairie Island Units 1 and 2 a,c.e 9
0 I
h f
I i
5
?
h I
l l
t i
i i
i 6
werssennaw 3 - 33 I
. ~ - -
i
4.0 REFERENCES
.?
1.
United States Nuclear Regulatory Commission, Regulatory Guide 1.121, " Bases for i
Plugging Degraded PWR Steam Generator Tubes," August,1976.
l L 2.
Goodier, J.
N., and Schoessow, G. J., "The Holding Power and Hydraulic Tightness of -
l l
Eacpanded Tube Joints: Analysis of the Stress and Deformation," Transactions of the i
A.S.M.E., July,1943, pp. 489-496.
'3.
Grimison, E. D., and Lee, G. H., " Experimental Investigation of Tube Expanding,",-
Transactions of the A.S.M.E., July,1943, pp. 497-505).
l c
4.
ASME Boiler and Pressure Vessel Code,Section III, " Rules for Construction of Nuclear i
- Vessels," The American Society of Mechanical Engineers, New York, New York,'
1%5 Edition.
5.
WCAP-11241, "Tubesheet Region Plugging Criterion for the Central Nuclear de ASCO, ASCO Nuclear Station Units 1 and 2 Steam Generators", October,1986. (Proprietary) l 1
L l
6.
WCAP-10949, "Tubesheet Region Plugging Criterion for Full Depth Hardroll Expanded
)
Tubes," Westinghouse Electric Corporation, September,1985. (Proprietary) l l
1:
l i
i I.
l
)
y l
wrrasenax 4-1
l l
i APPENDIX A j
L* TESTS FOR 3/4-INCH DIAMETER TUBE STEAM GENERATORS A.1 TEST FOR THE DETERMINATION OF RESISTANCE TO LEAKAGE OF -
DEGRADED ROLL EXPANSIONS -- 3/4 INCH TUBE S/G j
I i
- A.l.1 Background i
The resistance to leakage (RTL) of short lengths of this tube joint design was not available j
and a test was designed to determine it.
A.I.2 Objective i
The purpose of this test was to determine tube-to-tubesheet joint resistance to leakage for i
short axial lengths of sound, i.e., undegraded, rolled joints. He tests were performed with l
T/TS test specimens fabricated of short sections of prototypical tubes rolled into collars which j
provided the same structural compliance as a unit cell of the TS. The tests were performed at j
prototypical pressures, temperatures and tube axial loads for normal operation (N.O.), the i
most stringent faulted primary-to-secondary condition, i.e., feedline break (FLB), and the most j
stringent faulted secondary-to-primary condition, Loss of Coolant Accident (LOCA).
l q
The effect of prototypical loads on 3/4 inch tube joint leakage was addressed in the leak test.
For example, a load such as primary-to-secondary pressure differential during N.O., causes direct pressure effects on the tube in the L* region to increase the T/TS radial contact pressure, s,. It also causes tubesheet bending. The differential pressure increases the T/TS 1
radial contact pressure for all tubes in the TS, peripheral as well as interior, i.e.,
away-from-periphery, tubes if the L* region is above the neutral bending axis of the TS.
However, the upward bending also causs a decrease in radial con'act pressure between tube J
t and tubesheet for interior tubes and therefore a decreased RTL. This bending effect, which is detrimental for interior tubes, was accounted for in the laboratory.
The acceptable leakage for a reasonable number of 3/4 inch OD tube joints was established for normal operating and faulted conditions. It was obvious that for all acceptable-leakage i
cases, the joint was also required to exhibit acceptable strength. It was concluded that the '
l (L*) strength testing could be decoupled from the leakage testing and that the strength limits could be determined by a separate test. De primary-to-secondary pressure differential effect and a secondary-to-primary pressure differential effect were expected to have negligible impact on the tube joint strength. However, prototypical radial contraction or expansion of the tube, causing a reduction or increase in the T/TS radial contact pressure (s,), was achieved L
in the leakage test.
l l
i f
WPF/NSPAppdxA/lll894 i A-1
~
. -.., ~
i
' A.I.3 Test Equipment
- 1) Tubes, Alloy 600 (1600) mill annealed (MA), [-
]"' inch wall l
(nominal).
'l
i
[
}' ".
- 3) Airetool Roll Expander, No. [
]'"; rolls per original (shop) fabrication specification.
i
- 5) Laboratory leak test equipment such as pressurizing systems, furnaces, etc.
A.I.4 Test Major Steps
- 1) Determine roll torque reduction to simulate the effect of tubesheet bending on reduction of tube-to-tubesheet s, for interior tubes. (Note: As discussed in Section 3.2, the resistance to leakage for the 3/4 inch tube configuration was conservatively used for the 7/8 inch tubes of Prairie Island Units 1 and 2, based on a higher interference fit contact pressure.
for the 7/8 inch tubes. Part of this conservatism was based on the increase'in contact p
pressure at the tubesheet bottom, during the most important operatirig conditions, i.e., N.O.
l and FLB, involving tubesheet upward bowing, for the 7/8 inch tubes. Under the same conditions, the upward bowing of the tubesheet causes a reduction in contact pressure for the full-depth expanded,3/4 inch tubes, at the top of the tubesheet.
t
- 2) Fabricate leak test samples with [
] in. of sound roll between the BRT and the DRE. This distance is referred to as "X".
Each sample has a separate "X".
Refer to_
Figure A.1-1. Concentrate the most samples for X = 0.5 in., the desired length of sound RE above the degradation. An X of [
] in. is important for trending purposes only; X values of [
]'" in, were expected to be leaktight or to have negligible leakage 2a)
Estimate collar ID surface finish in area ofleak path. During the roll expansion process, the tube becomes fully plastic and flows into very small depressions occurring on the S/G tubesheet hole surface, or on the tubesheet simulant (collar) hole surface in this case. Conceptually, the amount of tube flow, determined by hole surface roughness, may be related to joint RTL. Ascertain that the collar finish is on the order of the prototypical finish, i.e., [
]"' RMS.
2b)
Fabricate T/TS samples per Figures A.1-1 and A.1-2. Set roll expansion motor l
torque to the middle of the torque range specified during S/G manufacturing, as reduced to account for TS upward bending., The torque reduction shall correspond to the s, caused by TS bending.
l-wrrmSPAppdxA/120894 2 A-2
2 1
j 2b-1) Partial roll expansion (non-interference fit). Remove tube from collar.
l 2b-2) Drill holes in tube to simulate DRE crack top tips. Reinsert tube in collar.
i i
2b-3) Finish (hardroll) RE.
1 i
4 2c)
Heat TffS samples in furnace at [-
]*". hours to duplicate the most
]
stringent conditions caused by post-weld heat treatment of the channelhead-to-TS weld during shop fabrication.
l l
2d)
Leak test the FLB primary-to-secondary configuration with [
']*" in. of sound roll between the BRT and the. degraded portion of the joint. Refer to~
q l
Figure A.1-1. The limiting case for the degradation is a 360* circumferential, throughwall " crack", machined in the tube. However, because this machining may.
j loosen the tube for small X values and therefore possibly bias the testing, a better method is to use the discrete-hole approach. 'Ihe number of holes [
]'" is chosen to be large, to approximate or exceed the number of linear ECI's observed j
j in these S/Gs. 'After the FLB samples are tested, convert the samples to the N.O.
configuration by rerolling to the higher, N.O., torque; [
]*"
j hours and leak test at the N.O. conditions.
l
)
2e)
Leak test the secondary-to-primary (LOCA) configuration using the same sequence F
as used in primary-to-secondary configuration above. The'LOCA samples are l
j completely separate from the FLB/N.O. samples. Refer to Figure A.1-2.
j i
i i
j A. I.5 Test Facility i
i The tests were performed at the Westinghouse Research and Development Laboratory. A j
four-cell manifold was used to apply pressurized, deionized water simultaneously to the tube ID's of four specimens for primary to-secondary tests; i.e., the normal operating and FLB l
conditions, and to the tube OD's for secondary-to-primary tests, i.e., the LOCA condition.
j System pressure was measured by a transducer system with a 5 psi accuracy. The electrical j
resistance heaters and insulation were positioned on the T/fS specimen such that the j
temperature acting over the leak path and ending at the BRT, the top end of the leak path in j
the plant, was [
]'" for the N.O. and FLB conditions. Above the BRT, the temperature j
dropped as designed by placement of the insulation so that any leakage was captured and condensed as droplets. The droplets were manually counted during the test period, usually 60 minutes. For the LOCA tests, the temperature acting over the leak path and beginning at the BRT was [
]*". In this case, all leakage was captured from the tube ID and condensed as droplets and counted manually during the 60 minute test period.
1 r
)
1 1
3 WPF/NSPAppdxA/l11894.3 A-3
)
)
i
1 A.I.6 Test Procedure The test was performed in accordance with a written procedure, a referenceable document, under Quality Assurance (QA) surveillance. The tube specimens were fabricated from QA-controlled stock; the collars were fabricated of known-strength AISI 1018 cold rolled carbon steel and were also fabricated under QA surveillance.
9
'1 i
WPFHSPAppdxAllll894 4 f.4
^
i l
Figure A.1-1
)
Roll Expansion Leakage Test Sample for 3/4 Inch Tube Steam Generator J
' Normal Operation and FLB Conditions l
I a,c,e 1
l 1
i i
i
')
lI, 1
i i
l l-X = 0.25, 0.50,1.00, or 2.00 In, Tolerance to be 0.030 In.
l NOTES:
- 1. Primary Side source, dionized water, subcooled
- 2. Step roll 4 passes, include overlaps, partial roll and hardroll
- 3. Roll 1st Pass (outside of test region) at high torque to ensure leaktightness.
- 4. Plug this end of tube.
l
- 5. Test FLB condition first, then roll top 1 or 2 passes only, at appropriate higher torque, to L
mclude X, to achieve N.O. T/TS radial pressure, then leak test N.O. condition.
WPF/NSPAppdxAltli894 s A-5 l
1 i
i Figure A.1-2 j
Roll Expansion Leakage Test Sample for 3/4 Inch Tube Steam Generator LOCA Conditions j
i i
a,c e s
i i
J J
I l
1 l
f X = 0.25, 0.50,1.00, or 2.00 In. Tolerance to be 10.030 In.
NOTES:
- 1. Secondary side source, deionized water, subcooled l
- 2. Step roll 4 passes, include overlaps, partial roll and hardroll
- 3. Plug this end of tube.
WPFNSPAppdxA/Ill394 6 A-6 l
L l.
A.2 TEST RESULTS FOR LEAKAGE FROM DEGRADED ROLL EXPANSIONS --
3/4 INCH SG TUBES l
A.2.1
_ Leakage' Criterion The leakage acceptance criterion was developed for evaluation of the leak rate test results.
Individual acceptance values were developed for test differential pressure conditions -
simulating normal operation,- feedline break and LOCA.
l Normal Operation The total normal operation primary-to-secondary leakage allowed in any one steam generator is specified in the Standard Technical Specification as 500 gpd, or 0.35 gpm for steam generator tubes. For conservatism, it was decided that the leakage contribution from L* tubes must not exceed this approximatley one-third of this value, or 150 gpd (0.1 gpm). The _
i number of tubes to which the L* criteria can be applied was arbitrarily set at [
]'" tubes.
For the specified number of applicable tube ends, an allowable normal operation L* leakage limit of [
]'" per L* tube was established. (This was based on the equivalence of 75,000 drops per gallon.)
a i
The use of [
]'". of the limit in the Technical Specification, i.e., m allocation factor i
of[
]'" to determine the leakage rate criterion was an arbitrary decision for
~
operational flexibility. The initial leak rate acceptance criterion may have been reevaluated in i
the light of actual test results and the allowance for potential degradation in the steam generator. Use of a larger fraction of the Technical Specification leakage limit and/or use of a smaller leak test acceptance criterion closer to actual results also could have been used to support application of the L* criteria to a number of tubes Isrger than the number on which p
the initial leak acceptance was based.
Faulted Condition: Feedline Break Postulated feedline break conditions provide the maximum primary-to-secondary differential pressure across the tube. The steamline break (SLB) provides the most stringent radiological l
conditions for postulated accidents involving a loss of pressure or fluid in the secondary system. For establishment of the leak rate acceptance criteria for faulted conditions, the assumed SLB leakage rate is used with the feedline break pressure differential and was referred to as FLB in the test. This was the most stringent case.
l The primary-to-secondary pressure differential for this condition was [
]*" psi. It _should l
be noted that dynamic loads on the tube joint, viz., secondary side fluid drag on the tube l
during the postulated accident, need not be added to this value. This is because the safety valve relieving event and the maximum fluid drag event are not concurrent. Therefore, the l
[
]'" psi was the most stringent condition for test and an additional pressure differential, corresponding to the drag load need not be considered. This sequence of events also eliminates the need for augmenting the axial load during the test to determine the axial l
WPF!NSPAppdrNill89u A-7
t L
L i
loadbdaring strength of DRE's. The axial load during the leak test is prototypical because it is -
i p
the end cap load caused by the prototypical differential pressure.
Faulted Conditice LOCA i
Typically the small leakage from steam generator tubes into the primary system is not a -
l significant consideration in the analysis of a postulated LOCA.
The test conditions used to simulate the LOCA'condi ions of [
_ ]'" psi secondary-t to-primary differential pressure and a temperature equal to normal operating temperature were-conservative for analyzed LOCA conditions.
A.2.2 Roll Torque Reduction Test Results j
This test related s, to rolling torque (T). ' Simulation of the TS-bending effect for the -
subsequent leak tests in the laboratory was accomplished by reducing the T corresponding to the s, reduction in the plant. This was necessary because the L* length was near the top of -
the tubesheet and therefore detrimentally affected by tubesheet bending upward during j
primary-to-secondary side pressure differentials.
i 1
J Leakane Test R_gsits
)
- I The N.O., FLB and LOCA leakage test results are shown in Table A.2-1, l
i i
l 1.
wrrMSPAppdxA/Il1894 :
A-8 i
TABLE A.2-1 ROLL EXPANSION LEAKAGE TEST RESULTS - 3/4 INCH TUBE S/G Roll Sample Expansion Plant Differential Temp.,'F Leak Length. in.
Condition Pressure. pi Planned Actual Rate. dom
_a,c e i
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l t
D l
wn NSPAppdd/111894 9 A-9
?
t TABLE A.2-1 (Continued)
ROLL EXPANSION LEAKAGE TEST RESULTS - 3/4 INCH TUBE S/G Roll Sample
- Expansion Plant
- Differential Temp., F Leak Number Leneth. in.
Condition Pressure. pli Planned Afinal Rate. dom a,c e.
NOTES:
I'rairie Island AP at N.O. is 1593 psi; use of higher AP during test is required for subcooling primary side to ensure liquid phase at potential leak path. Higher AP is conservative.
Gross Leak; Collar ID had excessive roughness.
- For all tests, Collar ID surface roughness was "as drilled and reamed", with 250 RMS
. typical and with actual maximum of approximately 300 RMS.
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i wPF/NSPAppdxA/Il!89410 A-10
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J A1 TEST FOR THE DETERMINATION OF AXIAL LOADBEARING CAPABILITY f
OF DEGRADED ROLL EXPANSIONS - 3/4 INCH TUBE S/G 1
A.3.1 Introduction i
4 j
Tests were conducted to determine the axial loadbearing capability of the 3/4 inch diameter.
j l
tube steam generator degraded roll expanded T/TS joints. The simulated degradation J
consisted of a single band of axial or near-axial slots. The results of the single band -
1 degradation (SBD) tests were also applictble to the most stringent, i.e., uppermost band of l
multiple band degradation (MBD) in roll expanded tube joints. The MBD configuration did l
not require testing. A series of tests was performed to determine the ultimate pull force (i.e.,
l tensile strength) for non-degraded and artificially degraded, prototypical, [
l j
]"' in. (nominal) wall Alloy 600MA tubes rolled into cold rolled carbon steel AISI j_
1018 collars. - The limitir.g test condition involved application of the pullout load directly to
}
the top of the SBD. The roll expansion was performed utilizing a prototypical 3/4 inch tube i
steam generator T/TS roller with regulated rolling torque to provide proper levels of tube j
thinning and joint preloads. Slots were electric discharge machined (EDM) into the wall of j
tube specimens to simulate throughwall indications. All aspects of the test were designed to i
bound typical plant respective conditions. Refer to Table A.3-1.. For example, tubes were.
i
}
machined to have [
]"' slots inclined at angles of [
]"' degrees from 2
the axial centerline of the tubes. The length of the slots was [
]"' in. which was i
considered to bound typical plant lengths by a factor of approximately 2, based on eddy
)
)-
current test assessments. Based on tube pull destructive examination results that indicate a
]
throughwall distance less than the measured eddy current distance, using a 2 times multiplier-l to the eddy current crack length data as a representative throughwall crack length is quite i
l conservative. The test axial distance "X" between the' tops of the slots and the BRT was
[
]"' in.; plant "X" values for the intended tubes was approximately 0.5 in. The smaller I
(test) X was judged to bound plant potential X values; lower strength generally results from a j.
smaller X. The collars were geometrically sized to simulate the equivalent structural reaction of a unit cell of the steam generator tubesheet. Figures A.3-1 and A.3-2 illustrate the basic i
shape and geometries of the collared and decollared specimens, respectivelyc The
]
never collared tube specimens are shown in Figure A.3-3.
j i
i All tests were performed at ambient atmospheric conditions employing a tension-compression i
machine shown in Figure A.3-4. Because the plant tubes are stressed at elevated temperature, j
the room temperature test results obtained with tubes of [
]"' ksi UTS were analytically
]
adjusted (decreased) to account for the small effects of temperature and UTS, for the Pull i
Strength Model discussed later in this report.
1 i
A.3.2 Objective i
l The objectives of this test were: (1) Determine the ultimate pull force necessary to i
structurally fail non-degraded and artificially degraded [
]"' inch wall j
Alloy 600 tubes rolled into collars that simulated a' unit cell of the steam generator tubesheet.
(2) Measure the ultimate pull force necessary to structurally fait degraded non-rolled alloy wnmSPApphAUllt94 II A-11 1
l tubes free of collars. (3) Evaluate the type of failure mechanism that induced structural
- i failure in the tubes.
A.3.3 Test Equipment l
- 1) Tubes, prototypical Alloy 600MA, [
]* in. wall (nominal thickness).
)
Refer to Table A.3-1 for specimen features, and to Table A.3-2 for specific specimen configurations tested.
- 2) Tubesheet simulants (collars), cold rolled carbon steel, AISI 1018, 4.5 In. long.' Thir material is in general use for simulants because the yield strength and modulus ofit approximate the respective properties of the tubesheet material. Refer to Figure A.3-1.
- 3) Airetool Roll Expander, No. [
]* rolls per original (shop) fabrication
- specification.
-]
1
-l I
- 5) Zetac Tension-Compression Testing Machine. Refer to Figure A-3-4.
l
- 6) Other laboratory strength test equipment such as fumaces, measuring instruments, etc.
A.3.4 Test Major Steps Collared Tubes j
The collared tubes were prepared by rolling to approximately two per cent thinning, then removed from the collar for EDM of the slots, if slots were required; the tubes were then replaced in the respective collars, hardrolled to the prototypical midrange torque and welded.
The specimens were then " pulled" in a configuration similar to that shown in Figure A.3-4.
Decollared Tubes l
The decollared samples were prepared by rolling the tubes to a small amount of wall thinning.
l a non-interference fit, then removed from the collar for EDM of the slots; the tubes were then l
replaced in the respective collars, and (hard) rolled to the specified midrange torque.
After hardrolling, the specimens were decollared, i.e., the collars were carefully removed by machining. This step prevented the collar from providing the known beneficial effects, primarily friction, during the strength test. The slot geometry and axial location are shown in Figure A.3-2. Table A.3-2 identifies the specimen configuration in terms of the number of slots, angle, and slot top location with reference to the BRT.
WPF/NSPApphAllll894.12 A-12
Never-Collared Tube.s -
These tubes were prepared for pulling by EDM of the slots. The reason for including this configuration was to show that it involves higher, i.e., non limiting ultimate strengths. The specimens were pulled in a configuration similar to that shown in Figure A.3-4.
Test Facility The tests were performed at the Westinghouse Research and Deveopment Laboratory. A Zetac tension - compression test machine was used to apply a continuously increasing axial pull force to the tube specimens. Figure A.3-4 is a sketch of the test machine. The tube top l
ends were permitted to rotate as dictated by the slot angle,4. A certain amount of rotation about the tube vertical axis could occur in the plant in keeping with the conservative assumption that the SRE within L* provides no fixity to the TS, owing to the tube straight leg acting as a torsional spring and being anchored at the U-bend. It was judged that this provided lower ultimate values than if axial rotation were prevented.
A.3.5 Data Acquisition System Axial pull forces were recorded on Cartesian coordinates as a function of tube top end i
displacement by employing an X-Y analog plotter.
l A.3.6 Procedure l
The specimens were vertically mounted in the Zetac testing machine as shown in Figure A.3-4. A steadily increasing axial pull was applied to the specimens until ultimate load failed the tube. A curve was simultaneously drawn by an X-Y analog plotter of the pull force versus tube top deflection. Curves were produced for non-degraded and degraded tubes; l
non-degraded tubes were tested to establish reference data in order to assess the strength of l
the degraded tubes.
l The pull specimens were tested per Table A.3-2.
t' j
wprinsPAppdxAll1189415 A-13
i 5
Table A.3-1 Selection of Bounding Features for Axial Loadbearing Test of Degraded Roll Expansions a,C,e sup b
I i
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l A-14 i
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.- e i.
r-
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t Table A.3-2 Axial Loadbearing Strength Test i
of Degraded Roll Expansions Array Top Distance Slot Specimen Specimen Below Bottom No.of
- Angle, Ultimate Number Confieur.
of Transition. im Slots Degrees Load.ItL a,c.e
)
)
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'l f
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t l
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i A-15 i
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\\
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Figure A.3-1 i
1 Geometry of Slots in Collared Tube Specimen -
l Ultimate Strength Test l
A-16 l
a c.e l
l i
I i
I l
l 1
\\
i j
1 i
9 Figure A.3-2 Geometry of Slots in Decollared Tube Specimen -
Ultimate Strength Test t
A-17
d 1-EC4-f 4
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1 i
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1 4
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Figure A.3-3 j
Geometry of Slots in Unexpanded, Never-Collared Tube Specimen - Ultimate Strength Test f
e 4
A-18
.~-
a.c.e a
i Figure A.3-4 Sketch of Zetac Tension and Compression Testing Machine A-19 f
4 l
a
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A.4.0 CRITERION AND TEST RESULTS FOR THE AXIAL LOADBEARING CAPABILITY OF DEGRADED ROLL EXPANSIONS - 3/4 INCH TUBE S/G A.4.1.
Maximum Axial Load Criterion i
j Sinnie Rand Deeradation i<-
]
The maximum axial tensile load which a DRE must bear will be the most stringent of the -
1 N.O. md FLB axial loads. The LOCA axial load is axially compressive and cannot cause tube pullout.
5 Multiole Band DearadatiSD
-l b
j
- If degradation occurs in discrete cirmmbential bands i.e., as single band degradation (SBD),
j
[
or arrays, in an RE above the F* ele.n.n thereby preventing use of F*, it may also occur in J
multiple discrete bands below the F* eiewtion. This is termed multiple band degradation
}
(MBD). ' As discussed earlier, it's desired to avoid quantifying each of the possible several j-degradation bands to apply the pullout load to the T/TS weld. Instead, it's desired to apply
)
j the load only to the uppermost sound roll portions (expansions) which are interspersed with i
the uppermost degradation bands. For the purpose e f Jenplicity, each degradation band.
l requires the same strength. This strength i:: the sara o aat for the SBD. Therefore, for the
-l l
MBD case, the number ofindividual bands which naa. oc quantified is one less than the j
number of SRE's (N) needed to react the pullout load. If N is 3, N-1, is 2. Conservatism j
may be added to this calculation in the form of a larger N than is needed.
l t
l Normal Operation i
{
j The tube ultimate load required for normal operatian is the "end cap" load resulting from the l
i N.O. primary-to-secondary side pressure differentid oultiplied by the cross-sectional area of I
the roll expanded tube OD. A safety factor of three is applied to this load. 'Ihis load is the highest for the three conditions considered and the data for the 15 slot case is plotted in Figure A.4-3.
q j
Feedline Break i
j The tube ultimate load required for FLB for these 3/4 inch tube SGs is the endcap load j
resulting from the FLB primary-to-tecondary pressure differential multiplied by the crocs-i sectional area of the roll expanded tube OD. A safety factor of [
]'" for allowable stress for faulted conditions) was used for this i
load. This load did not need to be further augmented for dynamic loading effects during FLB because of the sequence of events during an FLB. This load was less than the N. Op. load and therefore was not plotted on Figure A.4-l I
j l
WPWSPAppdxA/111894.20 A-20
i
)
Loss of. Coolant Accident An axial loadbearing requirement for a LOCA event ' acts to move a tube downward, the opposite of pullout.
Therefore, this condition does not apply to pullout.' (Howeve., the DRE.
must not collapse under LOCA conditions. This condition was tested with prototypical, i.e.,
compressive, end cap loads, h the leak test section.)
A.4.2 Strength Test Results A sample of the strength test results for non-slotted and slotted tubes is presented in.
- Figures A.4-1 and A.4-2. The results for the entire test series were previously summarized in Table A.3-2. [
t
]'" Structural failure of the tube was attributed to the axial and shear forces acting through the tube wall as a function of the vectorial relationship between the axial pull force and the angle of the respective slot to the tube axial centerline.
The strength test results discussed above dealt with pullout loads. No deleterious structural effects are expected from LOCA i.e., compressive loads on DRE's. No collapse of the tubes l
in the LOCA leakage resistance test occurred.
(
The basic objective of the L* program was to demonstrate the acceptability, for leakage and strength, of degradation arrays of known or bounding dimensions and configurations. The leakage issue was addressed elsewhere in this report. The strength design curve will be developed here and is based on the strength test results.
l A.4.3 Evaluations and Conclusions I
Pull Strength Modej As a tube with an array of parallel slanted cracks is loaded axially, deformation proceeds as follows. Initial clastic elongation is followed by plastic yielding. Depending on the number -
j of cracks and the crack angle, yielding may first occur in the virgin unrolled tube section or in the crack array. Large numbers of cracks and high crack angles (4) favor yielding of the crack array. Yielding of the crack array proceeds by platic bending of the ligarnents between cracks and subsequent rotation of the cracks toward the longitudinal axis of the tube. This is easily visible in pull strength tests. As rotation occtes, the applied axial force must increase substantially since the effective moment arm for ligament bending is contmually decreasing.
WPF/NSPAppdxAll1189(21 A-21 i
<~'
l Strain hardening of the ligaments adds to the geometric hardening produced by crack rotation.
As the load is increased, general yielding of the of the virgin unrolled tube section may occur, followed by strain hardening and then yielding of the rolled but uncracked tube section. If deformation proceeds to this point, the rolled tube peels away from the tubesheet and the axial pull strength is limited by the tube-to-tubesheet weld. At some point, depending on the crack morphology, fracture will terminate the process of plastic deformation.
l Using the notation of Figure A.4-4, the axial load required to yield a tube with an array of slanted cracks is a,c,e,
[
]
This expression only considers ligament bending and assumes deformation is confined the j
initial minimum ligament crossection. The axial plastic displacement resulting for yielding I
and then rotation of the crack array is given by l
i a,c.e
[
}
If the crack array plastic displacement is added to the baseline load displacement record of an J
uncracked tube, then the computed load-displacement records closely approximate actual measured load displacement records. The load displacement record for the uncracked tube essentially provides elastic tube displacement and an indication of yielding in the uncracked section at high loads. Figure A.4-5 shows a test of a 3/4 inch diameter tube terminated by ligament fracture. Figure A.4-6 shows the test record for another 3/4 inch diameter tube l
where the test load has become high enough to yield the uncracked as well as cracked sections These figures show that there is good agreement between the calculated and test measured load displacement curves and that the mathematical model can be used to accurately predict the integrity limits of a tube with inclined cracking within the L* region.
t The 3/4 inch tube design curve (Figure A.4-7) was developed to be used for comparison with tube axial loads, multiplied by suitable factors of safety caused by the most stringent normal and faulted operation conditions.
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Figure A.41 Pull Force as a Fuction of Pull Displacement for a Collared, Nondegraded Tube (Specimen No. I1)
A-23
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Figure L4-2 Pull Force as a Function of Pull Displacement l-for a Collared Tube (Specimen No. 3), with (15)-30 Degree Slots, l
Tops of Slots 0.25 Inch Below the Bottom'of Roll Transition A-24
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i Figure A.4-3 Ultimate Pull Force as a Function of Slot Angle for Expanded, Collared and Decollared Tubes,15 Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition - 3/4 Inch Tubes A-25
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i Figure A.4-4 Model for Plastic Collapse l
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'I CRDSSHEAD DISPLACEMENT Figure A.4-5 Comparison of Load Displacement Records, Computed Vs Measured for 30 Slots at & = 45 3/4 Inch Tubes
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for 30 Slots at & = 30' - 3/4 Inch Tubes i
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Figure A.4-7 3/4 Inch Tube Steam Generator Degraded Tube Pull Strength Design Curve for L*
A-29 1
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1 A.5 CONCLUSION With sufficiently quantified information about degradation in the anchor regions of the applicable 3/4 inch tube S/G roll er pansions, use of the Design Curve is expected to provide reliable operation. (The leakage criterian for this case is discussed elsewhere in this appendix and involves an upper limit on the number of tubes per steam generator which may be dispositioned thusly.)
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WPF/N5PAppdxAll12994 30 A-30
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l Exhibit F Prairie Island Nuclear Generating Plant License Amendment Request Dated January 9,1995 Westinghouse Authorization Letter,.
CAW-94-764,.and Accompanying Affidavit Proprietary Information Notice Copyright Notice J
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i Westinghouse Energy Systems Nudu Tenobgy Wsbn Electric Corporation sox 355 Pittsburgh Pennsylvania 15230 0355 December 16,1994 3
CAW-94-764 Document Control Desk U.S. Nuclear Regulatory Commission Washington, DC 20555
]
Attention: Mr. William T. Russell APPLICATION FOR WITHHOLDING PROPRIETARY
^
INFORMATION FROM PUBLIC DISCLOSURE
}
Subject:
WCAP-14225, "F* and L* Abe Plugging Criteria For Tubes With Degradation In The hbeshed Roll Expansion Region Of he Prairie Island Units 1 And 2 Steam Generators" (Proprietary)
Dear Mr. Russell:
I De proprietary information for which withholding is being requested in the above-referenced report is further idpotified in Affidavit CAW-94-764 signed by the owner of the proprietary information, Westinghouse Electric Corporation. He affidavit, which accompanies this letter, sets forth the basis on which the information may be withheld from public disclosure by the Commission and addresses with specificity the considerations listed in paragraph (b)(4) of 10 CFR Section 2.790 of the Commission's regulations.
1 Accordingly, this letter authorizes the utilization of the accompanying Affidavit by Northern States Power Company.
Correspondence with respect to the proprietary aspects of the application for withholding or the Westinghouse affidavit should reference this letter, CAW-94-764, and should be addressed to the undersigned.
Very truly yours, N. J. Li 1, Manager RHO /bbp Nuclear Safety Regulatory & Licensing Activities Enclosures Kevin Bohrer/NRC(12H5)
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usa w sev 1
4 CAW-94-764 i
j COMMONWEALTH OF PENNSYLVANIA:
1 ss 2
COUNTY OF ALLEGHENY:
i Before me, the undersigned authority, personally appeared Henry A. Sepp, who, being by me duly sworn according to law, deposes and says that he is authorized to execute this Affidavit on l
behalf of Westinghouse Electric Corporation (" Westinghouse") and that the averments of fact set forth in this Affidavit are true and correct to the best of his knowledge, information, and belief:
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j Henry A. Sepp, Manager Regulatory and Licensing Initiatives Sworn to and subscribed tb before me this
/lo day l
of aA
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Nowd Seaf i
Denise K HeMctson,Nowy Pubic Mom *C Iunus D/e.nd1A/D9 m$
oEco,AReym *lm i
Notary Public l
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i l~ CAW-9'4-764 (1)
I am Manager, Regulatory and Licensing Initiatives,'In the Nuclear Technology Division, of l
the Westinghouse Electric Corporation and as such, I have been specifically delegated the -
function of reviewing the proprietary information sought to be withheld from public disclosure in connection with nuclear power plant licensing and rulemaking proceedings, and am authorized to apply for its withholding on behalf of the Westinghouse Energy Systems Business Unit.
(2)
I am making this Affidavit in conformance with the provisions of 10CFR Section 2.790 of the Commission's regulations and in conjunction with the Westinghouse application for '
withholding accompanying this Affidavit.
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l (3)
I have personal knowledge of the criteria and procedures utilized by the Westinghouse Energy Systems Business Unit in designating information as a trade secret, privileged or as.
confidential commercial or financial information.
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(4)
Pursuant to the provisions of paragraph (b)(4) of Section 2.790 of the Commission's regulations, the following is furnished for consideration by the Commission in determining whether the information sought to be withheld from public disclosure should be withheld.
(i) ne information sought to be withheld from public disclosure is owned and has been held in confidence by Westinghouse.
i l
l (ii)
The information is of a type customarily held in confidence by Westinghouse and not customarily disclosed to the public. Westinghouse has a rational basis for determining I
the types of information customarily held in confidence by it and, in that connection, utilizes a system to determine when and whether to hold certain types of information in confidence. De application of that system and the substance of that system constitutes Westinghouse policy and provides the rational basis required.
6
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, CAW-94-764 Under that system, information is held in confidence if it falls in one or more of several types, the release of which might result in the loss of an existing or potential competitive advantage, as follows:
(a)_
ne information reveals the distinguishing aspects of a process (or component, -
structure, tool, method, etc.) where prevention of its use by any of Westingiv=2se's competitors without license from Westinghouse constitutes a competitive economic advantage over other companies.
(b)
It consists of supporting data, including test data, relative to a process (or component, structure, tool, method, etc.), the application of which' data secures a compaitive economic advantage, e.g., by optimization or improved marketability.
(c)
Its use by a competitor would reduce his expenditure of resources or improve his competitive position in the design, manufacture, shipment, installation, assurance of quality, or licensing a similar product.
(d)
It reveals cost or price information, production capacities, budget levels, or commercial strategies of Westinghouse, its customers or suppliers.
(e)
It reveals aspects of past, present, or future Westinghouse or customer funded development plans and programs of potential commercial value to Westinghouse.
(f)
It contains patentable ideas, for which patent protection may be desirable.
Here are sound policy reasons behind the Westinghouse system which include the following:
j
! CAW-9'4-764
'(a)
He use of such information by Wesiinghouse gives Westinghouse a competitive advantage over its competitors. It is, therefore, withheld from disclosure to protect the Westinghouse competitive position.
4 (b)
It is information which is marketable in many wrys. He extent to which such information is available to competitors dimbishes the Westinghouse i
ability to sell products and services involyMg the u.s of the information.
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-(c)
Use by our competitor would put Westinghouse at a competitive disalivantage by reducing his expenditure of resources at our expense.
(d)
Each component of proprietary information pertinent to a particular competitive advantage is potentially as valuable as the total competitive I
advantage. If competitors acquire components of proprietary information, any-one component may be the key to the entire puzzle, 'thereby depriving Westinghouse of a competitive advantage.
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(e)
Unrestricted disclosure would jeopardize the position of prominence of I
Westinghouse in the world market, and thereby give a market advantage to the competition of those countries.
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(f) ne Westinghouse capacity to invest corporate assets in research and development depends upon the success in obtaining and maintaining a competitive advantage.
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(iii)
He information is being transmitted to the Commission in confidence and, under the j
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provisions of 10CFR Section 2.790, it is to be received in confidence by the Commission.
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1 CAW-94-764 l
(iv)
He information sought to be protected is n6t available in public sources or available information has not been previously employed in the same original manner or method i
to the best of our knowledge and belief.
i (v)
The proprietary information sought to be withheld in this submittal is that which is f
. appropriately marked in "F* and L* Tube Plugging Criteria for Tubes with Degradation in the Tubesheet Roll Expansion Region of the Prairie Island Units 1 and -
)
2 Steam Generators", WCAP-14225 (Proprietary), December,1994, being -
transmitted by Northern States Power Company letter and Applicaton for Withholding.
1 Proprietary Information from Public Disclosure, to NRC Document Control Desk, i
Attention: Mr. William T. Russell. He proprietary information as submitted for use by Northern States Power Company for the Alternate Tube Plugging Criteria is i
expected to be applicable in other licensee submittals in response to certain NRC requirements for justification for similar changes in Structural Design Basis.
t His information is part of that which will enable Westinghouse to:
(a)
Provide documentation of the analyses methods and testing for determining plugging margin.
4 (b)
Establish the minimum wall thickness in compliance with Regulatory Guide 1.121.
(c)
Establish the stress limits versus thinning of the remaining tube wall.
(d)
Establish the maximum allowable leakage in support of the leak-before-break criteria.
(e)
Assist customers in obtaining NRC approval.
Further this information has substantial commercial value as follows:
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l 4-CAW-94-764 (a)
. estinghouse plans to sell the use of'similar information to its customers for W
purposes of meeting NRC requirements for licensing documentation.
(b)
Westinghouse can sell support and defense of the technology to its customers in the licensing process.
Public disclosure of this proprietary information'is likely to cause substantial harm to the competitive position of Westingivese because it would enhance the ability of mmpetitors to provide similar documents and licensing defense services for "
i commercial power reactors without commensurate expenses. Also, public disclosure -
of the information would enable others to use the information to meet NRC-i j
requirements for licensing documentation without purchasing the 'right to use the information.
y The development of the technology described in part by the information is the result of applying the results of many years of experience in an intensive Westinghouse effort and the expenditure of a considerable sum of money.
In order for competitors of Westinghouse to duplicate this information, similar technical programs would have to be performed and a significant manpower effort, having the requisite talent and experience, would have to be expended for developing, i
testing and analytical methods.
Further the deponent sayeth not.
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Proprietary Information Notice Transmitted herewith are proprietary and/or non-proprietary versions of documents furnished to the NRC in connection with requests for generic and/or plant-specific review and approval.
In order to conform to the requirements of 10 CFR 2.790 of the Commission's regulations concerning the protection of proprietary information so submitted to the NRC, the information which is proprietary in the proprietary versions is contained within brackets, and where the proprietarji information has been deleted in the non-proprietary versions, only the brackets remain (the information that was contained within the brackets in the proprietary versions having been deleted).
He justification for claiming the information so designat,ed as proprietary is indicated in both versions by means of lower case letters (a) through (f) contained within parentheses located as a superscript immediately following the brackets enclosing each item of information_being identified as proprietary or in the margin opposite such information. Rese lower case letters refer to the types of information Westinghouse ctistomarily holds in confidence identified in Sections (4)(li)(a) through (4)(ii)(f) of the affidavit accompanying this transmittal pursuant to 10 CFR 2.790(b)(1).
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Copyright Notice l
The reports transmitted herewith each bear a Westinghouse copyright notice. The NRC is permitted to make the number of copies of the information contained in these reports which are necessary for its internal use in connection with generic and plant-specific reviews and approvals as well as the issuance, denial, amendment, transfer, renewal, modification, suspension, revocation, or violation of a license, permit, order, or regulation subject to the requirements of 10 CFR 2.790 regarding restrictions on public ' disclosure to the extent such information has been identified as proprielary by l
Westinghouse, copyright protection notwithstanding. With respect to the non-proprietary versions of these reports, the NRC is permitted to make the number of copies beyond those necessary for its internal use which are weazy in order to have one copy available for public viewing in the appropriate docket files in the public document room in Washington, DC and in local public document rooms as may be required by NRC regulations if the number of copies submitted is insufficient for this purpose. Copies made by the NRC must include the copyright notice in all instances and the proprietary notice if the original was identified as proprietary.
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