ML20246B683
| ML20246B683 | |
| Person / Time | |
|---|---|
| Site: | Sequoyah  |
| Issue date: | 06/30/1989 |
| From: | Connors H, Hall J, Pitterle T WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19312B340 | List: |
| References | |
| SG-89-06-005, SG-89-6-5, WCAP-12290, NUDOCS 8907100057 | |
| Download: ML20246B683 (179) | |
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WESTINGHOUSE PROPRI'ETARY CLASS 3 i L
l WCAP-12290 SG-89-06-005 1
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1 SEQUOYAH UNIT 1 AND UNIT 2 ,
EVALUATION FOR TUBE VIBRATION INDUCED FATIGUE
~I JUNE.1989 )
\
i AUTHORS: H. J. CONNORS M. H. HU 'i 1
J. M. HALL T. S. MAGGE
. G. W. HOPKINS R. M. WILSON J. L. HOUTMAN 'H. W. YANT APPROVED: Id O/ I T. A. PITTERLE, MANAGER STEAM GENERATOR ENGINEERING WORK PERFORMED UNDER SHOP ORDER C08D42531 This. document contains information proprietary to Westinghouse Electric Corporation. It is submitted in confidence and is to be used solely for the purpose for which it is furnished and is to be returned upon request. This document and such information is not to be reproduced, transmitted, disclosed or.used otherwise in whole or in part without written authorization of Westinghouse Electric Corporation, Energy Systems Business Unit.
WESTINGHOUSE ELECTRIC CORPORATION l- NUCLEAR SERVICES DIVISION P.O. BOX 3377 l- PITTSBURGH, PENNSYLVANIA 15230 9048M:1E-0s2689
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WESTINGHOUSE PROPRIETARY CLASS 3 1
i ABSTRACT I
- t. )
i
"' On July- 15, 1987, a steam generator tube rupture e.e t occurred at the North ,
Anna Unit 1 plant. The cause of the tube rupture has been determined to be I high' cycle fatigue. The source of the loads associated with.the fatigue mechanism is a combination of a mean stress level in the tube with a superimposed alternating stress. The mean stress is the result of manufacturing residual stress, applied stress and residual stress due to denting of the tube at the top tube suppert plate, while the alternating stress is due to out of plane deflection of the tube U-bend attributed to flow induced vibration. For tubes without AVB support, local flow peaking effects )
at unsupported tubes are a significant contribution to tube vibration amplitudes. l This report documents the evaluation of steam generator tubing at Sequoyah Unit 1 & Unit 2 for susceptibility to fatigue-induced cracking of the type !
experienced at North Anna Unit 1. The evaluation utilizes operating l conditions specific to Sequoyah Unit 1 & Unit 2 to account for the plant specific nature of the tube loading and response. The evaluation also includes reviews of eddy current data for Sequoyah Unit 1 & Unit 2 to establish AVB locations. . This report provides background of the event which occurred at North Anna, a criteria for f atigue assessment, a summary of test data which support the analytical approach, field measurement results showing AVB positions, thermal hydraulic analysis results, and calculations to determine tube mean stress, stability ratio and tube stress distributions, and j accumulated fatigue usage. This evaluation concludes that one tube in Steam Generator 1 of Unit 1, one tube in Steam tenerator 1 of Unit 2, and one tube in Steam Generster 4 of Unit 2 are potentially susceptible to fatigue and require corrective action.
i 9048M 1E-Os2629
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WESTINGHOUSE PROPRIETARY CLASS 3-StlMMARY OF ABBREVIATIONS ASME.
American Society of Mechanical Engineers ATHOS -
Analysis of.the Thermal Hydraulics of Steam Generators AVB- -
Anti-Vibratic ser AVT -
All Volatile Treatment ECT: -
Electric Power Research Institute
- FFT . -
-rast Fourier Transform FLOVIB -
Flow Induced Vibrations MEVF -
Modal Effective Void Fraction OD -
Outside Diameter l- RMS -
Root Mean Square SR -
Stability Ratio TSP -
Tube Support Plate
- F- -
degrees Fahrenheit hr -
hour ksi -
measure of stress - 1000 pounds per square inch lb -
pound '
mils -
0.001 inch MW .
mega watt psi -
measure of stress pounds per square inch psia -
measure of pressure - absolute 9046M:1E-Os2689 !
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WESTINGHOUSE PROPRIETARY CLASS 3 TABLE OF CONTENTS i
L SECTION.
l 1.0 Introduction f
l 2.0 Summary and Conclusions
2.1 Background
2.2 Evaluation Criteria 2.3 Denting Evaluation 1 2.4 AVB Insertion Depths l' 2.5 Flow Peaking Factors l 2.6 Tube Vibration Evaluation l 2.7 Overall Conclusion 3.0 Background 3.1 North Anna Unit i iube Rupture Event 3.2 Tube Examination Results 3,3 Mechanism Assessment 4.0 Criteria for Fatigue Assessment I 4.1 Stability Ratio Reduction Criteria 4.2 Local Flow Paaking Considerations i 4.3 Stress Ratio Considerations 5.0 Supporting Test Data 5.1 Stability Ratio Parameters j 5.2 Tube Damping Data )
5.3 Tube Vibration Amplitudes with Single-Sided AVB Support l 5.4 Tests to Determine the Effects on Fluidelastic Instability of Columrwise Variations in AVB Insertion Depths l 5.5 References l
904BM:1E-052689 444 I
j R WESTINGHOUSE PROPRIETARY CLASS'3
' TABLE OF CONTENTS (CONTINUED) m SECTION.
6.0- Eddy Current Data and AVB Positions 6.1 AVB Assembly Design
' 6.2 Eddy Current Data for'AVB Positions 6.3 Tube Denting at Top' Tube Support Plate 6.4 AVB Map Interpretations 7.0 Thermal and Hydraulic Analysis 7.1- Sequoyah Unit 1.& Unit 2 Steam Generator Operating Conditions 7.2 : ATHOS Analysis Model 7.3 ATHOS Results 7.4 Relative Stability Ratio Over Operating History 8.0 Peaking Factor Evaluation 8.1 North Anna 1 Cotifiguration 8.2 Test Measurement Uncertainties 8.3 Test Repeatability 8.4 Cantilever vs U-Tube 8.5 Air vs Steam Water Mixture -
8.6 AVB Insertion Depth Uncertainty 8.7 Overall Peaking Factor with Uncertainty 8.8 Feaking Factors for Specific Tubes 9.0 Structural and Tube Vibration Assessments 9.1 Tube Mean Stress 9.2 Stability Ratio Distribution Based Upon ATH0S j 9.3 Stress Ratio Distribution with Peaking Factor ;
9.4 Cumulative Fatigue Usage I i
904BM.1E-0526B9 $y l
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WESTINGHOUSE PROPRIETARY CLASS 3 LIST OF FIGURES FIGURE 3-1 Approximate Mapping of Fracture Surface of Tube R9C51 S/G "C" Cold Leg, North Anna Unit 1 3-2 Schematic Representation of Features Observed During TEM Fractographic Examination of Fracture Surface of Tube R9C51, S/G "C" Cold Leg, North Anna Unit 1 3-3 Calculated and Observed Leak Rates Versus Time 4-1 Vibration Displacement vs. Stability Ratio 4-2 Fatigue Strength of Inconel 600 in AVT Water at 600*F 4-3 Fatigue Curve for Inconel 600 in AVT Water Comparison of Mean Stress Correction Models 4-4 Modified Fatigue with 10% Reduction in Stability Ratio for Maximum Stress Condition 4-5 Modified Fatigue with 5% Reduction in Stability Ratio for Minimum Stress Condition 5-1 Fluidelastic Instability Uncertainty 3ssessment 5-2 Instability Constant - B 5-3 Instabi'ity Constants, B, Obtained for Curved Tube from Wind Tunnel Tests on the 0.214 Scale U-Bend Model 5-4 Damping vs Slip Void Fraction j l i 1
904BM:1E-052689 y
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' WESTINGHOUSE PROPR!ET_ARY CLASS 3 il m LIST OF FIGURES.(Continued) d 1 FIGURE 5 Overa'11 View of Cantilever Tube Wind Tunnel Model 5-6 Top View of the Cantilever Tube Wind Tunnel Model ,
l- !
5-7 Fluidelastic Vib' ration Amp'litude with Non-Uniform Gaps 1
!5-B' Typical Vibration Amplitude and Tube /AVB Impact Force Signals for )
Fluidelastic Vibration with Unequal Tube /AVB Gaps-9 Conceptual Design of the Apparatus for Determining the Effects on Fluidelastic Instability.of Columnwise Variations in AVB Insertion Depths f
5-10 Overall View of Wind Tunnel Test Apparatus 5-11 Side View of Wind Tunnel Apparatus with Cover Plates Removed to Show Simulated AVBS and Top Flow Screen 5-12 AVB Configurations Tested 5-13 Typical Variation of.RMS Vibration Amplitude with Flow Velocity for Configuration la in Figure 5-12 6-1 AVB Insertion Depth Confirmation 6-2 Sequoyah Unit 1 Steam Generator 1 - AVB Positions 6-3 Sequoyah Unit 1 Steam Generator 2
- AVB Positions 6-4 Sequoyah Unit 1 Steam Generator 3 - AVB Positions 904BM:1E-052689 y$
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WEST 1NGHOUSE PROPRIETARY CLASS 3
- LIST OF FIGURES (Continued)
~ FIGURE-6-5 Sequoyah' Unit.1. Steam' Generator 4 - AVB Positiens
'6-6 Sequoyah Unit 2 Steam Generator 1 - AVB Positions 6-7 Sequoyah Unit 2 Steam Generator 2 -~AVB Positions
~6-8 Sequoyah Unit 2 Steam Generator 3'- AVB Positions 6-9 Sequoyah' Unit 2 Steam Generator 4 - AVB Positions 7-1 Plan View'of ATHOS Cartesian Model for Sequoyah 7-2 Elevation View of ATHOS Cartesian Model for Sequoyah 7-3 Plan View of ATHDS Cartesian Model Indicating Tube Layout 7-4 Flow Pattern on Vertical Plane of Synutry 7-5 Vertical Velocity Contours on a Horizontal Plane at the Entrance l to the U-Bend 7 Lateral Flow Pattern on Horizontal Plane at the Entrance to the U-Bend l 7-7 Void Fraction Contours on Vertical Plane of Symmetry
'7-B Tube Gap Velocity and Density Distributions for Tube at R10/C3 7-9 Tube Gap Velocity and Density Distributions for Tube at R10/C20 7-10 Tube Gap Velocity and Density Distributions for Tube at R10/C40 904BM.1E-052689 y4$
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?fa ;W WESTINGHOUSE PROPRIETARY CLASS 3-
.g , , LIST OF FIGURES (Continued)
, . , , FIGURE 7-11 Average Velocity and Density in the Plane of the U-Bends Normal to Row 10-7-12 Sequoyah Unit 1 & Unit 2 Normalized Stability Ratio Based on HighPower-(>85%) Operation 8-1 Original North Anna.AVB Configuration 8-2 Schematic of Staggered AYBs 8-3' AVB " Pair" in ECT Trace 8-4 ' North Anna 1, Steam Generator C: AVB Positions Critical Review "AVB Visible" Calls 8-5 North Anna 1, Steam Generator C R9C51 Projection Matrix B-6 North Anna R9C51 AVB Final Projected Positions
'B-7 Final Peaking Factors for Sequoyah Unit 1 & Unit 2 L 9-li Axisymmetric Tube Finite Element Model l
9-2 Dented Tube Stress Distributions - Pressure Load on Tube 9-3' Dented Tube Stress Distributions - Interference Load on Tube 9-4 Dented Tube Stress Distributions - Combined Stress Results l 9-5 Relative Stability Ratios using MEVF Dependent Damping -
Sequoyah Unit 1 904BM:1E-0526BB g$
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. LIST OF FIGURES (Continued)
FIGURE
- 9-6 ' Relative Stability Ratios using MEVF Dependent Damping -
- Sequoyah: Unit 2 9-7 ' Stress Ratio Vs. Column Number - Dented Condition -
Sequoyah Unit 1 L. '
c-8 Stress Ratio Vs.' Column' Number - Undented Condition -
Sequoyah. Unit 1 9-9 Stress Ratio Vs. Column Number - Dented Condition -
Sequoyah Unit 2 9-10 Stress Ratio Vs. Column Number - Undented Condition -
Sequoyah. Unit 2 9-11 Sequoyah Unit 1 & Unit 2.- Maximum Allowable Relative Flow Peaking i
9048M:1E-060789 iX e
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LIST 0F' TABLES
. TABLE 4-1 . Fatigue Usage per: Year Resulting From' Stability Ratio Reduction 5-1 Wind Tunnel Test on Cantilever Tube Model:
5-2 Fluidelastic Instability Peaking ~ Velocity Factors for.Columnwise
" Variations in AVB Insertion-Depths p
6-1. Summary Listing of. Unsupported Tubes - Sequoyah Unit 1 6-2 Summary Listing of' Unsupported Tubes - Sequoyah Unit 2
- 1. Sequoyah Unit 1 & Unit 2 Steam Generator Operating Conditions Used for ATHOS Analysis 7- 2 .. Sequoyah Unit 1 & Unit 2 Operating History Data 8-1 Stability Peaking Factor Due to Local Velocity Perturbation i 8-2 Comparison of Air and Steam-Water Peaking Factor Ratios. !
8-3 Effect of Local Variation of AVB. Insertion 8-4 Uncertainties in Test Data and Extrapolation 8-5 Extrapolation of Test Results to Steam Generator Conditions 1
)
l 8-6' Final Peaking Factors 8 Stability Peaking Factors for Specific Tubes - Sequoyah Unit 1 8-8 Stability Peaking Factors for Specific Tubes - Sequoyah Unit 2
-9048M:1E-05?689
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WESTINGHOUSE PROPRIETARY CLASS 3 LIST OF TABLES TABLE 9-1 100% Power Operating Parameters 9-2 Sequoyah Unit 1 - Tubes With Significant RSR's or Stress Ratios 9-3 Sequoyah Unit 2 - Tubes With Significant RSR's or Stress Ratios I
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ll WESTINGHOUSE PROPRIETARY CLASS 3
1.0 INTRODUCTION
This report documents the evaluation of steam generator tubing at Sequoyah Unit 1 & Unit 2 for susceptibility to fatigue-induced cracking of the type experienced at North Anna Unit 1 in July,1987. The evaluation includes three-dimensional flow analysis of the tube bundle, air tests performed to support the vibration analytical procedure, field measurements to establish ,
AVB locations, structural and vibration analysis of selected tubes, and fatigue usage calculations to predict cumulative usage for critical tubes.
The evaluation utilizes operating conditions specific to Sequoyah Unit 1 & ]
Unit 2 in order to account for plant specific features of the tube loading and )
response.
Section 2 of the report provides a summary of the Sequoyah Unit 1 & Unit 2 evaluation results and overall conclusions. Section 3 provides background for ,
the tube rupture event which occurred at North Anna Unit 1 including results of the examination of the ruptured tube and a discussion of the rupture mechanism. The criteria for predicting the fatigue usage for tubes having an environment conducive to this type of rupture are discussed in Section 4.
Section 5 provides a summary of test data which supports the analytical vibration evaluation of the candidate tubes. A summary of field measurements
~
used to determine AVB locations and to identify unsupported tubes is provided in Section 6. Section 7 provides the results of a thermal hydraulic analysis to establish flow field characteristics at the top support plate which are subsequently used to assist in identifying tubes which may be dynamically unstable. Section 8 presents an update of the methodology originally used to evaluate the tube rupture at North Anna Unit 1. The final section, Section 9, presents results of the structural and vibration assessment. This sectica describes tube mean stress, stability ratio and tube stress distribut'.ons, and !
accumulated fatigue usage, for the small radius U-tubes in the Seguryah Unit 1
& Unit 2 steam generators.
I 9o4BM:1E-060sB9 1-1
1 WESTINGHOUSE PROPR!ETARY CLASS 3
.2.0
SUMMARY
AND CONCLUSIONS
- The Sequoyah Unit 1 & Unit 2 steam generators have been evaluated for the susceptibility of unsupported U-bend tubing with denting at the top tube support plate to a fatigue rupture of the type experienced at Row 9 Column 51 (R9C51) of Steam Generator C at North Anna Unit 1. The evaluation of Sequoyah Unit 1 was based on Eddy Current Test (ECT) interpretation performed by TVA, while the evaluation of Sequoyah Unit 2 was based on ECT interpretation ,
performed by Westinghouse.
2.1 Background
The initiation of the circumferential crack in the tube at the top of the top.
' tube support plate at North Anna 1 has been attributed to limited displacement, fluid elastic instability. This condition is believed to have prevailed.in the R9C51 tube since the tube experienced denting at the support plate. A combination of conditions were present that led to the rupture. The tube was not supported by an anti-vibration bar (AVB), had a higher flow field due to local flow peaking as a result of non-uniform inse tion depths of AVBs, had reduced damping due to denting at the top support plate, and had reduced l fatigue properties due to the environment of the all volatile treatment (AVT) chemistry of the secondary water and the additional mean stress from the 'l I
denting.
2.2 Evaluation Criteria !
)
The criteria established to provide a fatigue usage less than 1.0 for a finite period of time (i.e., 40 years) is a 10% reduction in stability ratio that provides at least a 58 percent reduction in stress amplitude (to < 4.0 ksi) 4 for a Row 9 tube in the North Anna 1 steam generators (SG's). A reduction of this magnitude is required to produce a fatigue usage of < 0.021 per year for a Row 9 tube in North Anna and therefore a fatigue usage factor objective of greater than 40 years. This same fatigue criteria is applied as the '
principal criteria in the evaluation of Sequoyah Unit 1 & Unit 2 tubing.
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a WESTINGHOUSE PROPRIETARY CLASS 3 i y
The fluidelastic stability ratio is the ratio of the effective velocity divided by the critical velocity. A value greater than unity (1.0) indicates instability. 'The stress ratio is the expected stress amplitude in a Sequoyah' ,
. Unit 1 & Unit 2 tub'e divided by the stress amplitude for the North Anna 1, O R9C51 tube.
Displacements are computed for the unsupported U-bend tubes in Rows 11 and
. inward, .(descending row number) using relative stability ratios to R9C51 of
- North Anna 1'and an appropriate power law relationship based on instability displacement versus flow velocity. Tubes having different U-bend radii will have different stiffness and frequency and, therefore, different stress and p fatigue usage per year than the Row 9 North Anna tube.- These effects are-accounted for in a stress ratio technique. The stress ratio is formulated so that a stress ratio of 1.0_or less produces acceptable stress amplitudes and fatigue usage for the Sequoyah Unit l'& Unit 2 tubing for the reference fuel cycle analyzed. Therefore, a stress ratio less than 1.0 provides the next level of acceptance criteria for unsupported tubes for which the relative stability ratios,. including flow peaking, exceed 0.9.
The stability ratios for Sequoyah Unit 1 & Unit 2 tubing, the corresponding stress and amplitude, and the resulting cumulative fatigue usage must be evaluated relative to the ruptured tube at Row 9 Column 51, North Anna 1, Steam Generator C, for two reasons.. The local effect on the flow field due to various AVB insertion depths is not within the capability of available analysis techniques and is determined by test as a ratio between two AVB configurations. Ir. addition, an analysis and examina' tion of the ruptured tube at North Anna 1 provided a range of initiating stress amplitudes, but could only bound the possible stability ratios that correspond to these stress amplitudes. Therefore, to minimize the influence of uncertainties, the evaluation of Sequoyah Unit 1 & Unit 2 tubing has been based on relative stability ratios, relative flow peaking factors, and stress ratios.
The criteria to establish that a tube has support from an AVB and therefore eliminate it from further considerations is that it must have at least one sided AVB support present at the tube centerline. The criteria is based on test results which show that one sided AVB support is sufficient to
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1 WESTINGHOUSE PROPRIETARY CLASS 3 limit the vibration amplitude for fluidelastic excitation. AVB support is established by analysis of eddy current (EC) measurements and is a key factor in' determining the local flow peaking factors. The local flow peaking produces increased local velocities which cause an increase in stability ratio. A small percentage change in the stability ratio causes a significant change in stress amplitude. The relative flow peaking factors for Sequoyah Unit 1 & Unit 2 tubing without direct AVB support have been determined by test. These flow peaking factors normalized to the North Anna R9C51 peaking, are applied to relative stability ratios determined by 3-D tube bundle flow analysis, to obtain the combined relative stability ratio used in the stress ratio determination.
2.3 Denting Evaluation The denting evaluation for Sequoyah Unit I was based on Eddy Current Tests (ECT) performed in September of 1985 and interpreted by TVA. The evaluation of Sequoyah Unit 2 was based on the ECT performed in January and February of 1989 and interpreted by Westinghouse. For conservatism in the evaluation, all of the tubes are evaluated for two possible conditions - corroded, but not dented; and as being dented. The effect of denting on the fatigue usage of the tube has been conservatively maximized by assuming the maximum effect of mean stress in the tube fatigue usage evaluation and by incorporating reduced demping in the tube vibration evaluation.
2.4 AVB Insertion Depths The Sequoyah Unit 1 & Unit 2 SGs have two sets of Alloy 600 AVBs. The ' inner' AVBs have a rectangular cross-section and extend into the tube bundle l approximately as far as Row 11. Discounting tube ovality, which tends to vary with bend radius, they provide a nominal total clearance between a given tube and the surrounding AVBs of [ ]a,c inch. Considering average tube ovality for a Row 11 tube, the nominal total tube to AVB clearance is approximately [ ]a,c inches.
904BM.1E-Os2689 2-3
-WESTINGHOUSE PROPRIETARY CLASS 3 The outer AVBs have the same cross section as the inner AVBs, and extend into the tube bundle approximately as far as Row 14, providing a nominal tube'to AVB clearance comparable to the inner AVBs. Since the purpose of this analysis is to evaluate the potentially unsupported tubes at or near the point L of maximum AVB insertion, only the dimensions and EC data pertaining to the inner AVBs are required.
The eddy current data supplied by TVA were reviewed by Westinghouse to identify the number of tube /AVB intersections and the location of these
. intersections relative to the apex of a given tube. This information was used in calculations by Westinghouse to determine the deepest penetration of a l' given AVB into the tube bundle. For the area of interest in the Sequoyah Unit 1 & Unit 2 steam generators, the AVB support of the tube can normally be verified.if EC data shows both legs of the lower AVB, one on each side (hot j leg - cold leg) of the U-bend. This data, indicated by a listing of two or !
more AVBs in the insertion depth plots, is the method of choice for establishing tube support.
If only the apex of an.AVB assembly is near or touching the apex of a tube U-bend, only one AVB signal.may be seen. In this case, adequate tube support cannot be assumed without supplemental input. Support can be determined if
' projection' calculations based on the AVB intercepts of higher row number tubes for the same column-verify insertion depth to a point below the tube centerline. Maps of the AVB insertion depths for Sequoyah Unit 1 & Unit 2 are shown in Figures 6-2 thru 6-5 and Figures 6-6 thru 6-9, respectively. These AVB maps list the results of the ' projection' calculations where this information contributes to understanding of the AVB insertion depth.
2.5 Flow Peaking Factors Tests were performed modeling Sequoyah Unit 1 & Unit 2 Series 51 SG tube and AVB geometries to determine the flow peaking factors for various AVB configurations relative to the North Anna R9C51 peaking factor. The test results were used to define an upper bound of the ratio relative to the R9C51 9048%1E-Os2689 2-4
l WESTlNGH005E PROPRIETARY CLASS 3 configuration. It was found that one tube in Sequoyah Unit 1, and two tubes in Sequoyah Unit 2 had flow peaking values of the same order of magnitude as R9C51.
2.6 Tube Vibration Evaluation <
The calculation of relative stability ratios for Sequoyah Unit 1 & Unit 2 makes use of detailed tube bundle flow field information computed by the ATHOS steam generator thermal / hydraulic analysis code. Code output includes three-dimensional distributions of secondary side velocity, density, and void fraction, along with primary fluid and tube wall temperatures. Distributions of these parameters have been generated for every tube of interest in the Sequoyah Unit 1 & Unit 2 tube bundles based on recent full power operating conditions. This information was factored into the tube vibration analysis leading to the relative stability ratios.
Relative stability ratios of Sequoyah Unit 1 & Unit 2 (Row 8 through Row 12) tubing versus R9C51 of North Anna 1 are plotted in Figure 9-5 and 9-6, respectively. These relative stability ratios include relative flow peaking factors. The stress ratios for Sequoyah Unit 1 & Unit 2 are given in Figures 9-7 thru 9-10. These also include the relative flow peaking effect, and are calculated based on clamped tube conditions with denting at the tube support plate.
For Sequoyah Unit 1, examination of Table 9-2 and Figure 9-7 shows that the unsupported tube at R10C44 of SG 1 exceeds the limiting stress ratio criteria, and should be removed from service. Removal may take the form of either
' stabilizing and plugging' or ' sentinel plugging'. Of the remaining unsupported tubes in all four Unit 1 steam generators, the highest stress ratio is 0.50 and occurs at loa tion R8C59 in SG 4. The maximum fatigue usage for this tube is calculated by combining the usage for Unit 1 operating history to date plus the projected usage for future eneration. Assuming operation at 100% power with the current parameters and plugging values for '
- 100% availability, the maximum calculated fatigue usage is 0.05.
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l WEST 1NGH0VSE PROPRIETAW CLASS 3 For Sequoyah Unit 2, examination of Table 9-3 and Figure 9-9 shows that the l unsupported tubes at R10C60 of SG 1 and R9C60 of SG 4 exceed the limiting stress ratio criteria, and should be removed from service. Removal may take )
the form of either ' stabilizing and plugging' or ' sentinel plugging'. Of the {
remaining unsupported tubcs in all four Unit 2 steam generators, the highest l stress ratio is 0.95 and occurs at location R9C35 in SG 1. The maximum fatigue usage for this tube is calculated by combining the usage for Unit 2 operating history to date plus the projected usage for future operation.
Assuming operation at 100% power with the current parameters and plugging values for 100% availability, the maximum calculated fatigue usage is 0.83. ,
2.7 Overall Conclusion I
The analysis described above indicates that the Sequoyah Unit 1 & Unit 2 tubes recommended to remain in service are not expected to be susceptible to fatigue rupture at the top support plate in a manner similar to the rupture which occurred at North Anna 1. Therefore, no modification, preventive tube plugging, or other measure to preclude.such an even+ is believed necessary, other than the plugging of R10C44 of Unit 1, SG1; R10C60 of Unit 2 SG 1; and R9C60 of Unit 2 SG 4. This conclusion is based on the current power levels and flow conditions with some margin on steam pressure as identified in j Section 7 of this report. An increase in Unit power level (steam flow rate), {
a significant increase in SG plugging level, or a decrease in steam pressure below 800 psia may require supplemental reanalysis.
9048M:1E-060689 2-6
i l WESTINGHOUSE PROPRIETARY CLASS 3- I l:
3.0 BACKGROUND
L On July 15, 1987,. a steam generator tube rupture occurred at the North Anna Unit 1. The ruptured tube was determined to be Row 9 Column 51 in steam generator "C". The location of the opening was found to be at the top tube support plate on the cold leg side of the tube and was circumferential in orientation with a 360 degree extent.
3.1 North Anna Unit 1 Tube Rupture Event I
The cause of the tube rupture has been determined to be high cycle fatigue.
The source of the stresses associated with the fatigue mechanism has been determined to be a combination of a mean stress level in the tube and a L
superimposed alternating stress. The mean stress has been determined to have been increased to a maximum level as the result of denting of the tube at the top tube support plate and the alternating stress has been determined to be
- l due to out-of plane deflection of the tube U-bend above the top tube support caused-by flow induced. vibration. These. stresses are_ consistent.with a lower
' bound fatigue curve for the tube material in an AVT water chemistry l
2nvironment. The vibration mechanism has been. determined to be fluid elastic, based on the magnitude of the alternating stress.
A significant contributor to the occurrence of excessive vibration is the reduction in damping at the tube-to-tube support plate interface caused by the denting. Also, the absence of antivibration bar (AVB) support has been concluded to be required for requisite vibration to occur. The presence of an AVB support restricts tube motion and thus precludes the deflection amplitude required for fatigue. Inspection data shows that an AVB is not present for the Row 9 Column 51 tube but that the actual AVB installation depth exceeded the minimum requirements in all cases with data for AYBs at many other Row 9 tubes. Also contributing significantly to the level of vibration, and thus loading, is the local flow field associated with the detailed geometry of the steam generator, i.e., AVB insertion depths. In addition, the fatigue properties of the tube reflect the lower range of properties expected for an 1
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WESTINGHOUSE PROPRIETARY CLASS 3 1
AVT environment. In summary, the prerequisite conditions derived from the evaluations were concluded to be: j Fatigue Requirements Prerequ' site Conditions Alternating stress Tube vibration
- Dented support
- Flow excitation l
- Absence of AVB Mean stress Denting in addition to applied stress Material fatigue properties AVT environment .
- Lower range of properties !
3.2 Tube Examination Results 1 Fatigue was found to have initiated on the cold leg outside surface of Tube ,
R9C51 immediately above the top tube support plate. No indications of significant accompanying intergranular corrosion were observed on the fracture face or on the immediately adjacent OD surfaces. Multiple fatigue initiation sites were found with major sites located at 110', 120*, 135' and 150',
Figure 3-1. The plane of the U-band is located at 45' with the orientation (
system used, or approximately 90' from the geometric center of the initiation zone at Section D-D. High cycle fatigue striation spacings approached 1
]
micro-inch near the origin sites, Figure 3-2. The early crack front is !
believed to have broken through wall from approximately 100' to 140*. From I this point on, crack growth is believed (as determined by striation spacing, l striation direction, and later observations of parabolic dimples followed by I equiaxed dimples) to have accelerated and to have changed direction with the
]
resulting crack front running perpendicular to the circumferential direction, 904BM;1E-Os2689 3-2 1
1 I
WEST!NGHOUSE PR0e.. ETARY CLASS 3 3.3 Mechanism Assessment To address a fatigue mechanism and to identify the cause of the loading, any loading condition that would cause cyclic stress or steady mean stress had to be considered. The analysis of Normal, Upset and Test conditions indicated a relatively low total number of cycles involved and a corresponding low fatigue usage, even when accounting for the dented tube condition at the plate. This analysis also showed an axial tensile stress contribution at the tube 00 a 3 short distance above the plate from operating pressure and temperature, thus providing a contribution to mean stress. Combining these effects with denting deflection on the tube demonstrated a high mean stress at the failure location. Vibration analysis for the tube developed the characteristics of 1 l
first mode, cantilever response of the dented tube to flow induced vibration I for the uncracked tube and for the tube with an increasing crack angle, beginning at 90* to the plane of the tube and progressing around on both sides )
to complete separation of the tube.
1 Crack propagation analysis matched cyclic deformation with the stress intensities and striation spacings indicated by the fracture inspection and ;
analysis. Leakage data and crack opening analysis provided the relationship between leak rate and circumferential crack length. Leakage versus time was then predicted from the crack growth analysis and the leakage analysis with initial stress amplitudes of 5, 7, and 9 ksi. The comparison to the best estimate of plant leakage (performed after the event) showed good agreement, Figure 3-3.
\
Based on these results, it folicwed that the predominant loading mechanism responsible is a flow-induced, tube vibration loading mechanism. It was shown that of the two possible flow-induced vibration mechanisms, turbulence and i fluidelastic instability, that fluidelastic instability was the most probable l cause. Due to the range of expected initiation stress amplitudes (4 to l 10 ksi), the fluidelastic instability would be limited in displacement to a range of approximately [ Ja,c . This is less than the distance between tubes at the apex, [ Ja,c . It was further 904BM:1E-Os2689 3-3 l l
1
WESTINGHOUSE PROPRIETARY CLASS 3 confirmed that displacement prior to the rupture was limited since no indication of tube U-bend (apex region) damage u s evident in the eddy-current signals for adjacent tubes.
Given the likelihood of limited displacement, fluidelastic instability, a means of establishing the change in displacement, and corresponding change in stress amplitude, was developed for a given reduction in stability ratio .
(SR). Since the rupture was a fatigue mechanism, the change in stress l amplitude resulting from a reduction in stability ratio was converted to a fatigue usage benefit through the use of the fatigue curve developed. Mean stress effects were included due to the presence of denting and applied loadings. The results indicated that a 10% reduction in stability ratio is needed (considering the range of possible initiation stress amplitudes) to reduce the fatigue usage per year to less than 0.021 for a tube similar to Row 9 Column 51 at North Anna Unit 1. (
l l
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l 904BM:1E-052689 3,,4
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WEST 1NGHOUSE PROPRIETARY CLASS 3 ,
i I
I I
D-D > g id C
- C-C '
L !
C i Region of Herringbone g Pattern i
\ ,
30* - 270'- ; '
j s
B B-B , p E-E A
x TAB
%, A-A F-F
%, Cearse Texture
?. and Dimpled e Rupture l'
d Indicates Origins Figure 3-1 Approximate Mapping of Fracture Surface of Tube R9C51, S/G "C" Cold Leg, North Anna Unit 1 I
904BM 1E-052689 3-5 !
WESTINGHOUSE PROPRIETARY CLASS 3 i
5 = 1.5/2.6 w in.
r Heavy i Dxide 5 = 21 y in.
-E 3-D ,
/,
5 = 1.0/1.85 m in. ggo.
3-C l
h Parabolic
- 90* 270' - i 1-G Dic:ples k / and V
Internal Necking j 1-H !
l 3-8
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0 -
Badly 5 a 2.8/4.0 v in. ,
l j 11 Peened I
(///3/ ?.* l > !
Nearly Ecui-Axed 5 = 6.1/6.9 y in. Dir@les Note: Arrows Indicate Dirtction of Fracture Propagation Figure 3-2 Schematic Representation of Features Observed During TEM Fractograhic Examination of Fracture Surface of Tube R9C51, S/G "C" Cold Leg, North Anna Unit 1 904BM:1E-052689 3-6 !
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WESTINGHOUSE PROPRIETARY CLASS 3 l
3::::: l l l .: l l l l l l l l Calculated and observed leak rates versus time.
Observed values based on gaseous speciescondenser air ejector 1
~ secoe.- --
_e
@ SIGMA A = 5 KSI W SIGNA A = 7 MSI ,
g i SIGMA A = 9 KSI l l J O Ar-41 6 ,,,,.. D Xe-135 ~~
O. A Kt'-67 E.
i l 0- ! -
t y O 4
see.. .- g ..
tu /
a O O so-- .
l 0
l 9 900 10kp0 Sekte ee$pe gn'ee geh MM M @d 8000 TIME (Minutes)
Figure 3-3 Calculated and Observed Leak Rates Versus Time 904BM:1E-052689 g
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WESTINGHOUSE PROPR3ETARY CLASS 3 4.0 CRITERIA FOR FATIGUE ASSESSMENT l
The evaluation method and acceptance criteria are based on a relative l comparison with the Row 9 Column 51 tube of Steam Generator C, North Anna I; i
Unit 1. This approach is necessary because (1) methods for direct analytical I prediction of actual stability ratios incorporate greater uncertainties than a relative ratio method, and (2) the, stress amplitude (or displacement)
{
associated with a specific value of stability ratio can only be estimated by i the analysis of North Anna Unit 1. For these reasons, the North Anna Unit 1 l tubing evaluation was done on a relative basis to Row 9 Column 51 and a 10%
l reduction in stability ratio criteria was established to demonstrate that tubes left in service would be expected to have sufficiently low vibration i stress to preclude future fatigue rupture events.
To accomplish the necessary relative assessment of Sequoyah Unit 1 & Unit 2 i
tubing to Row 9 Column 51 of North Anna Unit 1, several criteria are )
utilized. First, stability ratios are calculated for the Sequoyah Unit 1 &
Unit 2 steam generators based on flow fields predicted by 3-D thermal hydraulic models and ratioed to the stability ratio for Row 9 Column 51 at ;
North Anna Unit 1 based on a flow field obtained with a 3-D thermal hydraulic I model with the same degree of refinement. These ratios of stability ratio (called relative stability ratios) for each potentially unsupported U-bend in the Sequoyah Unit 1 & Unit 2 steam generators should be equivalent to 5 0.9 of R9C51, North Anna 1 (meeting the 10% reduction in stability ratio criteria). This provides the first level of screening of susceptible tubes incorporating all tube geometry and flow field differences in the tube dynamic evaluation. It has the inherent assumptien, however, that each tube has the same local, high flow condition present at How 9 Column 51, North Anna Unit
- 1. To account for these differences, flow peaking factors can be incorporated in the relative stability ratios and the stress ratios.
The next step is to obtain stress ratios, the ratio of stress in the Sequoyah Unit 1 & Unit 2 tube of interest to the stress in Row 9 Column 51, North Anna Unit 1, and after incorporating the requirement that the relative stability ratio to Row 9 Column 51 (R9C51) for the tube of interest is equivalent to 5 0.9, require the stress ratio to be 5 1.0. The stress ratio j 904BM:1E-060sB9 4-1
WESTINGHOUSE PROPRIETARY CLASS 3 incorporates the tube geometry differences with R9C51 in relation to the stress calculation and also incorporates the ratio of flow peaking factor for the tube of interest to the flow peaking factor for R9C51 (flow peaking factor is defined in Section 4.2). This should provide that all tubes meeting this i criteria have stress amplitudes equivalent to < 4.0 ksi.
Finally; the cumulative fatigue usage for plant operation to date and for continued operation with the same operating parameters is evaluated. A fatigue usage of < 1.0 may not be satisfied by meeting the stress ratio criteria using the reference operating cycle evaluation since the reference i cycle does not necessarily represent the exact duty cycle to date. Therefore, the time history sf operation is evaluated on a normalized basis and used together with the stress ratio to obtain a stress amplitude history. This permits the calculatbn of curred and future fatigue usage for comparison to 1.0.
4.1 Stability Ratio Reduction Criteria For fluidelastic evaluation, stability ratios are determined for specific configurations of a tube. These stability ratios represent a measure of the potential for flow-induced tube vibration during service. Values greater than unity (1.0) indicate instability (see Section 5.1).
Motions developed by a tube in the fluidelastically unstable mode are quite large in comparison to the other known mechanisms. The maximum modal displacement (at the apex of the tube) is linearly related to the bending stress in the tube just above the cold leg top tube support plate. This relationship applies to any vibration in that mode. Thus, it is possible for an unstable, fixed boundary condition tube to deflect an amount in the U-bend which will produce fatigue inducing stresses.
i The major features of the fluidelastic mechanism are illustrated in Figure '
4-1. This figure shows the displacement response (LOG D) of a tube as a function of stability ratio (LOG SR). A straight-line plot displayed on j log-log coordinates implies a relation of the form y = A(x)", where A is a j constant, x is the independent variable, n is the exponent (or power to which 4 l
904EM:1E-052689 g i
i 1
WESTINGHOUSE PROPRIETARY CLASS 3 x is raised), and y is the dependent variable. Taking logs of both sides of this equation leads to the slope-intercept form of a straight-line equation in log form, log y = c + n log x, where c = log A and represents the intercept I and n is the slope. In our case the independent variable x is the stability l l
ratio SR, and the dependent variable y is tube (fluidelastic instability l
induced) displacement response D, and the slope n is renamed s.
From experimental results, it is known that the turbulence response curve (on log-log coordinates) has a slope of approximately [ ]a,b,c Test results
. j also show that the slope for the fluidelastic response depends somewhat on the I l
instability displacement (response amplitude). It has been shown by tests )
that a slope of [ ]a,b,c is a range of values corresponding to displacement amplitudes in the range of [ 3a,c ,
whereas below [ ]a,c are conservative values. )
l l
The reduction in response obtained from a stability ratio reduction can be j expressed by the following equation: )
- _ a,c )
_ _ l where D 1 and SR1 are the known values at the point corresponding to point 1 of Figure 4-1 and D2and SR2 are values corresponding to any point lower on this curve. Therefore, this equation can be used to determine the reduction in displacement response for any given reduction in stability ratio.
This equation shows that there is benefit derived from even a very small percentage change in the stability ratio. It is this reduction in j displacement for a quite small reduction in stability ratio that formed the basis for demonstrating that a 10% reduction in stability ratio would be sufficient to prevent Row 9 Column 51 from rupturing by fatigue.
The fatigue curve developed for the North Anna Unit I tube at R9C51 is from
[
]8,C 9048M:1E-052689 4-3 l
s
]
(
WESTINGHOUSE PROPRIETARY CLASS 3 L
'Ja,c . Thus,
- a,c ,
1
~ ~
l .:
where,oa is the equivalent stress amplitude to oa that accounts ~l for a maximum stress of o y
, the yield strength. The -3 sigma curve ~with l
mean stress effects is shown in Figure 4-2 and is compared to the ASME Code Design Fatigue Curve for.Inconel 600 with the maximum effect of mean stress.
The. curve utilized in this evaluation is clearly well below the code curve reflecting the effect of an AVT environment on fatigue and [
la,c for accounting for mean stress that applies to
' materials in a corrosive environment.
Two other mean stress models were investigated for the appropriateness of their use in providing a reasonable agreement with the expected range of initiating stress amplitudes. These were the [
la,c shown in Figure 4-3. With a [
]a,c, the'[
4
.i ya,c ,
The assessment of the benefit of a reduction in stability ratio begins with the relationship between stability ratio and deflection. For a specific tube 904BM:1E-052689 4_4 i
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WEST!NGHOUSE PROPRIETARY CLASS 3 l i
geometry, the displacement change is directly proportional to change in stress so that stress has the same relationship with stability ratio,
- - a,c q
~'
.The s ope in this equation can range from [ ]a,c on a log scale depending on the amplitude of displacement. Knowing the stress resulting from a change in stability ratio from SRy to SR2 , the cycles to failure at the stress amplitude were obtained from the fatigue curve. A-fatigue usage per year was then determined assuming continuous cycling at the natural frequency of the tube. The initial stress was determined to be in the rangc of 4.0 to 10.0 ksi by the fractography analysis.
t lt was further developed that the maximum initiat bg stress amplitude was not more than 9.5 ksi. This was based on [
la,c . The corresponding stress-level is 5.6 ksi.
-The maximum stress, 9.5 ksi, would be reduced to [ la,c with a 10%
reduction ir stability ratio and would have a future ~ fatigue usage of
[ Ja,c per year at 75% availability, Figure 4-4. The minimum stress, 5.6 ksi, would be reduced to [ Ja,c ksi with a 5% reduction in stability ratio and would have future fatigue usage of [ Ja c per year, Figure 4-5. In addition, if a tube were already cracked, the crack could be as large as [ la c inch in length and thru-wall and would not propagate if the stress amplitudes are reduced to < 4.0 ksi.
S',osequent to the return to power evaluation for North Anna Unit 1, the time history of operation was evaluated on a normalized basis to the last cycle, confirming the conservatism of 9.5 ksi. [ la,c 904BM:1 E-052689 4,5
WESTINGHOUSE' PROPRIETARY CLASS 3
[
Ja,c, cumulative fatigue usage may then be computed to get a magnitude of alternating stress for the last cycle that results in a cumulative usag.e of 1.0 for the nine year duty cycle.. The result of the iterative analysis is that the probable stress associated with this fatigue curve during the last cycle of operation was approximately [
la,c for R9C51, North Anna Unit 1, Steam Generator C, and that the major portion of the fatigue usage came in the second, third and fourth cycles. The
' first cycle was conservatively omitted, since denting is assumed, for purposes of this analysis, to have occurred during that first cycle. Based on this evaluation, the tube fatigue probably occurred over most of the operating history of North Anna Unit 1. '
1 A similar calculation can be performed for the time history of operation '
assuming that [
Ja,c . On this basis, the effect of a 10% reduction in stability ratio is to reduce the stress amplitude to 4.0 ksi and results in a future fatigue i usage of [ }"'C .
Other combinations of alternating stress and mean stress were evaluated with
-3 sigma and -2 sigma fatigue curves to demonstrate the conservatism of the 10% reduction in stability ratio. Table 4-1 presents the results of the cases 1 analyzed clearly demonstrating that the 10% reduction in stability ratio combined with a -3 sigma fatigue curve and with maximum mean stress effects is conservative. Any higher fatigue curve whether through mean stress, mean '
stress model, or probability, results in greater benefit for the same reduction in stability ratio. Further, for any of these higher curves, a 1 smaller reduction in stability ratio than 10% would result in the same I benefit. In addition, there is a large benefit in terms of fatigue usage for relatively small changes in the fatigue curve. ]
1
}- J 904BM.1E-Os2689 -
4,g
.. - - - - - - .. -}
i 1
WEST!NGHOUSE PROPRIETARY CLASS 3 4
4.2 Local Flow Peaking Considerations Local flow peaking is a factor on stability ratio that incorporates the effect on local flow velocity, density and void fraction due to non-uniform AVB insertion depths. The flow peaking factor is applied directly to the l stability ratio obtained from thermal-hydraulic analysis that does not account I for these local geometry effects. Being a direct factor on stability ratio, a small percentage increase can result in a significant change in the prediction of tube response.
l l
Since the evaluation of Sequoyah Unit 1 & Unit 2 tubing is relative to R9C51, l North Anna Unit 1, the flow peaking factors are also applied as relative ratios, i.e., a ratio of Sequoyah Unit 1 & Unit 2 tubing to R9C51 at North Anna Unit 1. The flow peaking relative instability is obtained by testing in the air test rig der.cribed in Section 5.4, where the peaking factor is defined ,
as the critical velocity for R9C51 AVB pattern compared to critical velocity for a uniform AVB pattern. As explained in Section 8.0, the minimum value of
[ ]a,b,c is appropriate for R9C51 of North Anna 1. The peaking factor i for a tube in Sequoyah Unit 1 & Unit 2 tubing is therefore divided by
[ Ja,b,c and the resulting relative flow peaking is multiplied times the relative stability ratio based on ATHOS results. If the peaking factor is 1.0, the relative flow peaking is [ Ja,b,c ,
As a further demonstration of the conservatism of [ 3a,b,c as the minimum )
flow peaking facter for R9C51, the stress amplitude of 7.0 ksi obtained from l
iterating on cumulative fatigue usage (and selected as the nominal value from )
fractography analysis) was used to back calculate the apparent stability ratio and then the apparent flow peaking factor. Allowing for a range of slopes of the instability curve from 10 to 30, the stability ratio is in the range of 1.1 to 1.4 and the flow peaking factor is in the range of 1.8 to 2.2. This range of flow peaking agrees with the range of flow peaking factors measured )
in the air tests and is considered to be the best estimate of the range of the I R9C51 flew peakint factor, i l
l 904BLt1E-052689 g l
___ J
, . m
-WESTINGHOUSE FCOPRfETARY CLASS'3' E. ' The range of stability ratios, 1.1 to 1.4, is based on a value of 0.63 u
( ., obtained with ATHOS results without flow peaking and with nominal damping that is a function'of modal effective void fraction (MEVF). MEVF is calculated using the formulai a,c a.
'I
.i L
The nominal damping reflects the' nominal reduction in damping that occurs with
' denting.at'the tube support plate. . Therefore, a minimum damping scenario that is independent of void fraction is not considered to be credible and is not addressed in the evaluation that folicm , d
~
4.3 Stress Ratio Considerations 1
In Section 4.1, a 10% reduction in stability ratio was established to reduce ,
i the stress amplitude on the Row 9 Column 51 tube of North Anna Unit 1 to a i level that;would not have ruptured, 4.0 ksi. To apply this same criteria to ,l !
another tube in the same or another steam generator, the differences in [ !
1 j .
}a,C
_ a,c_ ,
I j
l-1 l
j 904BM.1F-Os2689 4-8
__g WESTINGHOUSE PROPRIETARY CLASS 3 a.c 4
1 W _
904BM:1E-052689 4,g
-i
).: b i
WESTINGHOUSE PROPRIETARY CLASS 3 s _By estab l ishi'ng thei r equi val ent effect on t he stress amplitude that produced 1 the tube rupture at North Anna 1, several other effects may be accounted for.
These include a lower mean stress (such as for non-dented tubes), different
-frequency tubes from the [- ]C hertz frequency of R9C51, North Anna 1, )
and shorter' design' basis service.- i In the case of lower mean~ stress, the stress amplitude that would have caused the failure of R9C51, North Anna 1, would have been higher. [
l 1
3a,c ,
'.A lower or higher frequency tube would not reach a usage of 1.0 in the same length of time as the R9C51 tube due to the different frequency of cycling.
The usage accumulated is proportional to the frequency and, therefore, the allowable number of cycles to reach a usage of 1.0 is inversely proportional to frequency. The equivalent number of cycles to give the us.ege of 1.0 for a different frequency tube [
ya.c ,
For a different time basis.for fatigue usage evaluation, [
1 ja.c.e, 904BM:1E-052689 4-10 i I .
L
WEST!NGHOUSE PROPRIETARY CLASS 3
-Knowing the magnitude of the stress ratio allows 1) the' determination of tubes that do not meet a value of 5 1, and 2) the calculation of maximum stress in the acceptable tubes,
- - a,c Having;this maximum stress permits the evaluation of the maximum fatigue usage for Sequoyah Unit 1 & Unit 2 based on the time history expressed by normalized stability ratios for the duty cycle (see Section 7.4).
l l
l l
l 1
j l
904BM:1E-052689 4,g )
1 o
WESTINGHOUSE PROPRIETARY CLASS 3 Table 4-l' Fatigue Usage per Year Resulting i From Stability Ratio Reduction
\
SR, % STRESS FATIGUE MEAN STRESS- USAGE l REDUCTION BASIS (1) . CURVE (2) MODEL- PER YEAR l
- - a,c
'5 . 9 yrs to i fail [ Ja,c
]
- 5. 9 yrs to fail [ Ja.c
- 5. 9 yrs to fail'[ Ja,c
- 10. max.stregg) amplituge c
[ ]
- 10. max,stregg) amplitge c
- 10. max.stre{g) amplitu$,c
- 10. max.stregg) amplituj$,c
- 10. max. stress based on dutycygig(5)
[ ] - _
(1) This gives the basis for~ selection of the initiating stress amplitude and its value in ksi.
.(2) Sm is the maximum stress applied with Sm = Smean + Sa.
l (3) [ Ja,c ,
(4) Cycles to failure implied by this combination of stress and fatigue properties is notably less than implied by the operating history.
Consequently this combination is a conservative, bounding estimate.
(5) Cycles"tgfailureimpliedbytheoperatinghistoryrequireg'[
fatigue curve at the maximum stress of [
] ] .
904BM:1E-Os2689 gg
WESTINGHOUSE PROPRIETARY CLASS 3 l
l l
l \
l a, b, e 1
l \
l l
1 l
l Figure 4-1 Vibration Displacement vs. Stability Ratio 9048M.1E-052689 4,73 L____ . _ _ _ . . _ _ _ _ . _ _ _
WESTINGHOUSE PROPRIETARY CLASS 3
-a,c i
l i
1 1
l i
I i
l 1
1 Figure 4-2 Fatigue Strength of Inconel 600 in AVT Water at 600*F i 9048M:1E-052689 4,p i
WESTINGHOUSE PROPRIETARY CLASS 3 I
a,c 1
I; i
I i
- 1
+ 1 i Figure 4-3 Fatigue Curve for Inconel 600 in AVT Water Comparison of Mean Stress Correction Models 9048M:1E-052689 ,
a
1 WESTINGHOUSE PROPRIETARY CLASS 3 i
i ac l
Figure 4-4 Modified Fatigue with 10% Reduction in Stability Ratio for Maximum Stress Condition I 904BM:1E-0605B9 4-16
WESTINGHOUSE PROPRIETARY CLASS 3 a,c l
l 1
1 I
i i
l l
l l
Figure 4-5 Modified Fatigue with 5% Reduction in Stability Ratio for Minimum Stress Condition i
904BM;1 E-052689 4, g o
1
o 1.
WESTINGHOUSE' PROPRIETARY CLASS 3 5.0. SUPPORTING TEST DATA: ?
, I H
j i
H This section provides a mathematical description of the fluid-elastic mechanism,'which was determined to be the most-likely causative-mechanism for -
the North. Anna tube rupture, as discussed in Section 3.3, to highlight the H physical conditions and corresponding parameters directly related to the event and associated preventative measures. The basis for establishing the appropriate values and implications associated with-these. parameters are provided. Where appropriate, test results are presented.
5.1 ' Stability Ratio Parameters Fluid-elastic stability' ratios are obtained by evaluations for specific j configurations, in terms'of active tube. supports, of a specific tube. These
. stability ratios repre :nt a measure of the potential for tube vibration due j
~to instability during service. Fluid-elastic stability evaluations are
~
performed with a computer program which provides for the.genen. tion of a finite element model of the tube and tube support. system. .ine finite element '
model provides the vehicle to' define the mass and stiffness matrices for the tube and its support system. This-information is used.to determine the modal frequencies (eigenvalues) and mode shapes (eigenvector) for the linearly l supported tube being considered.
-The methodology'is comprised of the evaluation of the following equations:
' Fluid-elastic stability ratio = SR = Uen/Uc for mode n, where Uc (critical velocity) and Uen (effective velocity) are determined by:
1
.u*0I c Dn [(m, Sn ) / I'o D2 )y (1) and; N
b h ) hjn 3[y j o Z
j U
en
- " " - " - - - - - - - - [2] )
N 2
3[3("j/*o) 'jn Z
j 9048M:1E-052689 5-1
.j
l WESTINGHOUSE PROPRIETARY CLASS 3-
. 'l
.]
1where,
-D. = tub'e outside diameter, inches J U en; = effective velocity for mode n, inches /sec N = number of nodal points of the finite element model
=- number of degrees of freedom in'the out-of plane direction i
m3 , U3 , p3 = mass per unit length, crossflow velocity and fluid density at node j, respectively pg , m, a reference density and a' reference mass per unit length, respectively (any representative values) 6 n
=- logarithmicdec* cement (damping) ojn =. normalized displacement at node j in the nth mode of vibration l
z = average.of distances between node j to j-1, and j to J+1 j 3
i B = an experimentcily correlated stability constant Substitution of Equations [1] and [2] into the expression which defines stability ratio, and cancellation of like terms, leads to an expression in fundamental terms (without the arbitrary reference mass and density parameters). From this resulting expression, it is seen that the stability ratio is directly related to the flow field in terms of the secondary fluid velocity. times square root-density distribution (over the tube mode shape),
and inversely related to the square root of the mass distribution, square root of modal damping, tube modal frequency, and the stability constant (beta). l l
The uncertainty in each of these parameters is addressed in a conceptual
- i. manner in Figure 5-1. The remainder of this section (Section 5.0) provides a I discussion, and, where appropriate, the experimental bases to quantitatively establish the uncertainty associated with each of these parameters. In
-1 '
904BM:1E-052689 5-2
l WESTINGHOUSE PROPRIETARY CLASS 3
)
. addition, Section 5.3 provides the experimental basis to demonstrate'that- l tubes'with [. l Ja c . This q implies that those_ tubes [ Ja,c would not have to be modified because their instability response amplitude (and stress) would be
-small. 'The very high degree of. sensitivity of tube response (displacements i and stresses) to changes in the velocity times square-root-density distribution is' addressed in Section 4.0. This is important in determining the degree of change that can be attained through modifications.
Frequency It has been demonstrated by investigators that analytically determined frequencies are quite close to their physical counterparts obtained from-measurements on real structures. Thus, the uncertainty in frequencies has. ,
been shown to be quite small. This is particularly appropriate in the case of dented (fixed boundary condition) tubes. Therefore, uncertainty. levels j introduced by the frequency parameter are expected'to be insignificant (see L also " Average Flow Field" subsection below).
Instability Constant (Beta)
The beta (stability constant) values used for stability ratio and critical velocity evaluations (see above equations) are based on an extensive data base comprised of both Westinghouse and other experimental results. In addition, previous field experiences are considered. Values have been measured for full length U-bend tubes in prototypical steam / water environments. In addition, measurements in U-bend air models have been made with both no AVB and variable AVB supports (Figure 5-3).
To help establish the uncertainties associated with ATH0S flow velocity and density distribution predictions on stability analyses, the Model Boiler (MB-3) tests performed at Mitsubishi Heavy Industries (MHI) in Japan were modeled using ATHOS. A beta value consistent with the ATHOS predicted flow conditions and the MB-3 measured critical velocity was determined. These ,
analyses supported a beta value of [ 3a,b,c ,
l 904BM.1E-052689 5-3
WEST!NGHOUSE PROPRIETARY CLASS 3 A summary of the test bases and qualifications of the beta values used for these assessments is provided by Figure 5-2. The lowest measured beta for tubes without AVBs was a value of [ ]a,b,c This value is used for the beta parameter in all stability ratio evaluations addressed in this Report (see also " Average Flow Field" sub.ection below).
Mass Distribution i
The mass distribution parameter is based on known information on the tube and primary and secondary fluid physical properties. The total mass per unit length is comprised of that due to the tube, the internal (primary) fluid, and the external (secondary) fluid (hydrodynamic mass). Data in Reference 5-2 suggests that at operating void fractions [
)"'C .
Tube,Damoing Test data are available to define tube damping for clamped (fixed) tube supports, appropriate to dented tube conditions, in steam / water flow conditions. Prototypic U-bend testing has been performed under conditions leading to pinned supports. The data of Axisa in Figure 5-4 provides the principal data for clamped tube conditions in steam / water, This data was obtained for cross flow over straight tubes. Uncertainties are not defined for the data from these tests. Detailed tube damping data used in support of the stability ratio evaluations addressed in this report are provided in Section 5.2, below.
Flow Field - Velocity Times Souare-Root-Density Distribution Average and U-bend-local flow field uncertainties are addressed independently in the following.
I 904BM:1E-Os26B9 5-4 l l l 1 1
WEST 1NGHOUSE PROPRIETARY.. CLASS 3 j L .
' Average Flow Field L . Uncertainties in.the average flow field parameters, obtained from ATHOS. j analyses. coupled with stability ccnstant and frequency, are essentially.the l l-same for units with dented or non-dented top support plates. If the errors 1 associated with these uncertainties were large, similar instabilities would ba )
expected in the non-dented units with resulting wear at either the top support' plate'or. inner row AVBs. .Significant tube wear has not been observed in: inner )
. row tubes in operating steam generators without denting. : Thus, an uncertainty I estimate of about [ ]a,c for the combined effects of average flow field, stability constant and frequency appears to be reasonable. 'To further j minimize the impact of these uncertainties, the Sequoyah Unit.1 & Unit 2 tubes are evaluated on a relative basis, so that constant error factors are essentially eliminated. Thus, the uncertainties associated with the average-
. velocity times square-root-density (combined) parameter are not expected to be significant.
U-Bend Local Flow Field Non-uniform AVB insertion depths have been shown to have effects on stability ratios. Flow peaking, brought about by the " channeling" effects of non-uniform AVBs, leads to a local perturbation in the velo ity times square-root-density parameter at the apex of the tube where it will have the largest effect (because the apex is where the largest vibration displacements occur). . Detailed local-flow field data used in support of the stability ratio i evaluations addressed in this' report are provided in Section 5.2, below.
1 Overall Uncertainties Assessment h - Based on the above discussions, and the data provided in the following l sections, it is concluded that local :!ow peaking is likely to have contributed significantly to the instability and associated increased vibration amplitude for the failed North Anna tube. Ratios of stresses and stability ratios relative to the North Anna tube, R9C51, are utilized in this report to minimize uncertainties in the evaluations associated with instability constants, local flow field effects tnd tube damping.
9048M.1E-052689 5-5 p
L
WESTINGHOUSE PROPRIETARY CLASS 3 5.2 Tube Damping Data The damping ratio depends on several aspects of the physical system. Two primary determinants of damping are the supporL conditions and the flow field. It has been shown that tube support conditions (pinned vs clamped) affect the damping ratio significantly. Further, it is af'fected by the flow conditions, i.e., single phase or two phase flow. These effects are discussed below in more detail.
[ la,c indicates that the damping ratio in two phase flow is a sum of contributions from structural, viscous, flow-dependent, and two phase damping. The structural damping will be equal to the measured damping in air.
However, in two phase flow, the damping ratio increases significantly and is dependent on the void fraction or quality. It can be shown that the damping contribution from viscous effects are very small.
Damping ratios for tubes in air and in air-water flows have been measured and reported by various authors. However, the results from air-water flow are poor representations of the actual conditions in a' steam generator (steam-water flow at high pressure). Therefore, where available, results from prototypic steam-water flow conditions should be used. Fortunately, within the past few years test data on tube vibration under steam-water flow has been developed for both pinned and clamped tube support conditions.
Two sources of data are particularly noteworthy and are used here. The first is a large body of recent, as yet unpublished data from high pressure steam-water tests conducted by Mitsubishi Heavy Indut,tries (MHI). These data were gathered under pinned tube support conditions. The second is comprised of the resuits from tests sponsored by the Electric Power Research Institute (EFRI) and reported in [ Ja,c ,
The damping ratio results from tha above tests are plotted in Figure 5-4 as a function of void fraction. It is important to note that the void fraction is determined on the basis of [ Ja,c 904BM:1E-060789 5-6
WEST 1NGHOUSE PROPRIETARY CLASS 3
[ :]a,c. The upper curve in the figure is for pinned support con'ditions..-This curve represents a fit to a large number of data points not shown in the figure. The points on the curve are only plotting aids, rather i than specific test results.
The lower curve pertains to the clamped support condition, obtained from
[ ]a,c Void fraction has been recalculated on the basis of slip flow. It may be noted that there is a significant difference in the damping ratios under the pinned and the. clamped support conditions. Damping is much larger for pinned supports at all void fractions. Denting of the tubes at the top support plate effectively clamps the tubes at that location.
Therefore, the clamped tube support curve is used in the current evaluation to include the effect of denting at the top tube support plate.
The [ ]a,c data as reported show a damping value of 0.5% at'100%
void fraction. The 100% void fraction. condition has no two phase damping and is considered to be affected principally by mechanical or structural damping.
Westinghouse tests of clamped tube vibration in air has shown that the mechanical damping is only [ l a.c rather than the 0.5% reported in
[ Ja,c . Therefere the lower curve in Figure 5-4 is the
[ ]a,c data with all damping values reduced by [- Ja,c ,
i I
i 904BM;1E-060789 5-7 1 l
.. .. . . 1
WESTINGHOUSE PROPRIETARY CLASS 3 5.3 Tube Vibration Amplitudes With Single-Sided AVB Support A series of wind tunnel tests were conducted to investigate the effects of tube /AVB eccentricity on the vibration amplitudes caused by fluidelastic vibration.
[
3a,c . Prior test results obtained during the past year c'.ing this apparatus have demonstrated that the fluidelastic vibratio:; characteristics observed in the tests performed with the canti' lever tube apparatus are in good agreement with corresponding characteristics observed in wind tunnel and steam flow tests using U-bend tube arrays. A summary of these prior results is given in Table 5-1.
An overall view of the apparatus is shown in Figure 5-5. Figure 5-6 is a top view of the apparatus. [
3a,c ,
904BM:1E-052689 5-8
1 n.
1 WESTINGHOUSE PROPRIETARY CLASS 3
. .i As'shown in Figure 5-7,'the tube vibration amplitude below a critical velocity-
, 1s.causedbyL[ {
3a,c ,
i figure'5-7 shows the manner in which the zero-to peak vibration amplitude, l
expressed as a ratio normalized to [ ]a,c, varies when one gap remains at [' ]a,c For increasing velocities, up to that corresponding to a stability ratio of (
i Ja,c . Figure 5-8 shows
. typical vibration amplitude and tube /AVB impact force signals corresponding to those obtained from the tests which provided the results shown in Figure 5-7.
l As expected, impacting is only observed in the [ ]a,c ,
i l It is concluded from the above test results that, [
3a,c ,
5.4- Tests to Determine the Effects on Fluidelastic Instability of Columnwise' Variations in AVB Insertion Depths This section summarizes a series of wind tunnel tests that were conducted to investigate the effects of variations in AVB configurations on the initiation of fluidelastic vibration. Each cenn guration is defined as a specific set of insertion depths for the individual AVBs in the vicinity of an unsupported U-bend tube.
The tests were conducted in the wind ~annel using a modified version of the cantilever tube apparatus descr Wed in Section 5.3. Figure 5-9 shows the conceptual design of the apparatus. The straight cantilever tube, [
1 ja,c 904 BM.1 E-052689 5-9
w ,
, ~ WESTINGHOUSE PROPRIETARY CLASS 3-a
? ?[ , ,
L.
i ..
l.
$i ja,c,
['
'I
]a,c Figure 5-11 shows the-
-AVBs, when the side panel of the test section is removed. Also shown is the.
top flow screen:which is [
l Ja,c . The
=AVB configurations tested are shown in Figure 5-12. Configuration la corresponds to tube.R9C51, the failed tube at' North Anna.- Configuration 2a corresponds to one of the cases in which the-AVBs are ' inserted to a uniform depth'and'no local velocity peaking effects are expected.
As1shown in Figure 5-9, [
3a,c ,
.All the tubes except the instrumented tubed (corresponding to Row 10) are ;
[ ]C As discussed in Section 5.3,. prior i
testin'g indicates that this situation provides a valid model. The I
904BM;1E-Os2689 5-10
.A
WESTINGHOUSE PROPRIETARY CLASS 3 instrumented tube [ ]a,e as shown in Figure 5-10. Its [ Ja,c direction vibrational motion is measured using a non-contacting transducer.
[ t
]a,c The instrumented tube corresponds to a Row 10 tube as shown in Figure 5-9.- However, depending on the particular AVB configuration, it can reasonably represent a tube in Rows 8 through 11. The AVB profile in the straight tube model is the average of Rows 8 and 11. The difference in profile is quite small for these bounding rows. ,
[ la,c using a hot-film anemometer located as shown in Figure 5-9.
Figure 5-13 shows the rms vibration amplitude, as determined from PSD (power spectial density) measurements made using an FFT spectrum analyzer, versus
' flow velocity for Configuration la (which corresponds to tube R9C51 in North Anna). Data for three repeat tests are shown and the critical velocity is identified. The typical rapid increase in vibration amplitude when the critical velocity for fluidelastic vibration is exceeded is evident.
1 The main conclusions from the tests are:
y
- 1. Tube vibration below the critical velocity is relatively small, typical of turbulence-induced vibration, and increases rapidly when the critical ;
vel'oeity for the initiation of fluidelastic vibration is exceeded.
l
- 2. Configuration la (R9C51 in North Anna') has among the lowest critical velocity of all the configurations tested.
- 3. Configuration Ib, with a similar geometry, but slightly higher peaking l factor than la, has been run periodical: - .o verify the consistency of the test apparatus and its calibration. l j
i 904BM:1E-0605B9 5-11 l i
,. s WESTINGHOUSE PROPRIETARY CLASS-3'
.p , ,
, ine initial test results obtained.in sLpport of.the Sequoyah Unit 1 & Unit'2- .j Javaluation are summarized in Table,5-2. The test data is presented as a" j velocity peaking ratio;'a ratio of critical velocity for North Anna tube R9C51 L configuration la, to that for each Sequoyah Unit 1 & Unit 2 AVB:
- configuration evaluated.. <
1 1
i
< 5.5 References ,
j
_ac-4 l
- i. '.
i,. g i
i i
1 l
4 l
l:
- 904BM:1E-0526B'a 1
5-12
WESTINGHOUSE PROPRIETARY' CLASS 3 Table 5-1 ,
m Wind Tunnel Tests on Cantilever Tube Model-DBJECTIVE: , Investigate the ' effects.of tube /AVB fitup on flow-induced tube vibration.
APPARATUS: Array of cantilevered' tubes with end supports [
ya,c ,
MEASUREMENTS: Tube vibration amplitude and tube /AVB impact forces or preload forces.
RESULTS:
a,b,e l
l l l l
4 L
1.
9048M;1E-052689 5-13
WESTINGHOUSE PROPRIETARY CLASS 3' Table 5-2 Fluidelastic Instability Velocity. Peaking Ratios for-Columnwise Variation in AVB Insertion Depths (Sequoyah 1 and 2)
. Type of Insertion Peaking Ratio Configuration U3 ,/U n
- , a,b,c la lb ig im 14 it
.1z 2a 4a
'4b 4d 4e 4n 45 4v Sa 5b Sc i 5d Sg-Sh Si 6a .
Bd >
l Note: 0 is instability velocity at inlet f6r type n '
ofAVBinsertionconfiguration. .
1 I
l 9048M:1 E-0526BB 5-14
l WESTINGHOUSE PROPRIETARY CLASS 3 I
u- u 4I I 1
1 Figure 5-1 Fluidelastic Instability Uncertainty Assessment j 904BM 1E-052EB9 5-15
__]
WEST 1NGHOUSE PROPRIETARY CLASS 3 U-Bend Test Data 1)- MB-3 Tests 6 values of [ Ja,b,c
- 2) MB'2-Tests 6 of ( 'Ja,b,:
- 3) Air Model Tests S of [ -Ja,b,c without AVBs Tendency for S to increase in range of [ Ja,b,c.with inactive AVBs (9aps at AVBs)
Tendency for 6 to decrease toward a lower bound of [ la,b,c with active AYBs Verification of Instability Conditions
- 1) Flow conditions at critical velocity from MB-3
- 2) Measured damping for the specific tube
- 3) Calculated velocities from ATH0S 3D analysis
- 4) B' determined from calculated critical values Good agreement with reported B values
- 5) ATHOS velocity data with B of [ Ja,b,c and known damping should .i not significantly underestimate instability for regions of uniform .
U-bend flow i
I I
s Figure 5-2 Instability Constant - B l I
L 904BM:1E-052689 5-16
--_---_----_--____-z-_--___--_---__-----
T WESTINGHOUSE PROPRIETARY CLASS 3 i
a, b, c i
l i
)
i l
J j
l -
Figure 5-3 Instability Constants, s, Obtained for Curved Tubes from Wind Tunnel Tests on the 0.214 Scale U-Bend Model 904BM:1E-0526E9 5-17
1 WESTINGHOUSE PROPRIETARY CLASS 3 a, b, c l
i
)
I Figure 5-4 Damping vs. Slip Void Fraction 904BM:1 E-052689 5-18 i
_.-------------------------..____--.--------------.-------.j
WESTINGHOUSE PROPRIETARY CLASS 3
_ a, b, c Figure 5-5 Overall View of Cantilever Tube Wind Tunnel Model l
9048M:1E-052689 5-19 1
WESTINGHOUSE PROPRIETARY CLASS 3 l
l a, b, c l
I l
l
)
Figure 5-6 Top View of the Cantilever Tube Wind Tunnel Model 1 1
j 9048M:1E-052689 1
WESTINGHOUSE PROPRIETARY CLASS 3 a,b,c
~
l
\
l Figure 5-7 Fluidalastic Vibration Amplitude with Non-Uniform Gaps 9048M:1E-052689
- l. 5-21
-.---.------_------------_----------------_--------.-a
1 l
WESTINGHOUSE PROPRIETARY CLASS 3 l
a, b c 1
1 (
l l
l l
)
(
l
{
l i
l i
I t
figure 5-8 Typical Vibration Amplitude and Tube /AVB Impact force l Signals for Fluideiastic Vibration with Unequal !
Tube /AVB Gaps so48u.1t-oszscs 5-22
WESTINGHOUSE PROPRIETARY CLASS 3 a, b, c l
l 1
l 1
l J
I l
Figure 5-9 Conceptual Design of the Apparatus for Determining the Effects on Fluidelastic Instability of Columnwise Variations in AVB Insertion Depths ;
l l
9048M:1E-052689 5-23
1-WEST!NGHOUSE PROPRIETARY CLASS 3 a,b,c 1
l l
l l
l-l l
Figure 5-10 Overall View of Wind Tunnel Test Apparatus 9048M:1E-052BB9 5-24
WESTINGHOUSE PROPRIETARY CLASS 3 1.
a,b,c I l i
Figure 5-11 Sido View of Wind Tunnel Apparatus with Cover Plates Removed to Show Simulated AVBs for Field Modified Units and Top Flow Screen 904BM.1E-05268C 5-25 u __j
WESTINGHOUSE PROPRIETARY CLASS 3 a, b, c.
l i
-a l
1 l
Figure 5-12 AVB configurations Tested for Sequoyah 1 and 2 1
1 904BM:1E-052689 5-26
WESTINGHOUSE PROPRIETARY CLASS 3 a,b,c l
l Figure 5-13 Typical Variation of RMS Vibration Amplitude with Flow Velocity for Configuration la in Figure 5-12 9048M:1E-052689 5-27
._-______J
g% , ' ,
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tapes generated during the inspection performed in September of'1985:
. Thd EC inputL to the Sequoyah Unit 2 analysis is based on EC tapes genereted lduring the inspection performed' in the January and February of.1989.
> 6.1-AVBAssembikDesign
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Ja,c.e , Upper' AVBs which are inserted beyond the design depth,1 occasionally show on the EC traces for the Row 12 tubes. Since the
. purpose of this analysis is to evaluate potentially unsupported tubes at or near the point of maximum AVB insertion, only the dimensions and EC data pertaining to the ' lower' AVBs are normally used.
.6.2 Eddy Current Data for AVB Positions The AVB insertion depths were oetermined on the basis of interpretation of the eddy current data. To locate the AVBs, the ECT data traces were searched for
'the characteristic peaks seen in the signals, which indicate the intersection of= an AVB (or a tube support plate) with the tube (a typical signal for AVBs is shown 'in Figure 6.1). Since ambiguity can occur in the interpretation _of the ECT data, due to inability of ECT to differentiate at which side of a tube l' a " visible" AVB is located..other information must often be used to assist in establishing the location of the AVBs. Consistency with the design of the AVB assembly,; consistency of data for adjacent columns, and verification by projection were utilized to determine the depth of insertion which was plotted, i i
9048M;1E-0526B9 6-1
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'l WESTINGHOUSE PROPRIETARY CLASS 3 The flow peaking and tube support evaluation for Sequoyah Unit I was based on a Westinghouse. review of AVB insertion plots developed and provided by the TVA staff at.Sequoyah.. These plots for SG's 1 through 4 of Unit 1 are shown in
' Figure 6-2 through 6-5.
Westinghouse developed similar plots.for the SG's in Unit 2, using the EC tapes from the nost recent outage. In these plots the number of AVB intersections, including zero (meaning no AVB present), was logged for each I tube to indicate the presence or absence'of AVBs. Figures 6-6 through 6-9 ishow the number'of AVB signals found for each tube, and a representation of AVB insertion distance based on evaluation of the'EC data. Details of AVB
. projection . techniques based on EC data and tests are provided in Section 6.2.2.
6.2.1 AVB Insertion Depths AVB position maps for the Sequoyah Unit 1 & Unit 2 steam generators are given
.in Figures 6-2 through 6-9.
The direct observation data (the' number of AVB intersections seen by the eddy current probe) are the principal basis for determining the AVB positions. Two or more AVB indications indicate that a tube is supported. The presence of a single AVB indication may signify that: the tube is just tangent with the apex of. the AVB, the AVB signal is obscured by the presence of other detectable features (such as. deposits), or the AVB is close enough to be detected by the eddy current, but is not deep enough to provide support for !
the tube. A' typical ambiguity is the case where two'AVB indications are found i at a tube on one side of an AVB, but only one AVB indication is detected at the next tube in the same row. Geometrical projection calculations based on data from higher numbered rows may be needed to determine whether the tube with the single indication is supported . Where the direct observations are ambiguous or there is a conflict between observations and projections, the more conservative interpretation is used to determine the AVB positions.
Since ' direct observation' gives a 'yes - no' type of answer, the projection method is used to ' interpolate' AVB' insertion depths between rows of tubes. f The visual images thus produced are more easily understood when fluid flow 4
l SO48M:1E-052689 6-2 q l
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i 6.2.2 AVB Projection
.The. projection technique is useful where noisy or spurious.ECT signals prevent' positive location of an AVB or where data'is unavailable due to'a tube having been plugged. .. [
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6.3 . Tube Denting at Top Tube Support Plate B:cause of the AVB geometries involved and the desire to obtain 3 rows of
' projection' data where possible, the Sequoyah Unit 1 & Unit 2 evaluation covers Rows 8 crough 12 as a minimum. Subsequent to identifying the AVB l signals, eddy current data were examined to evaluate the incidence of corrosion and/or denting at the top tube support plate. In this evaluation, the EC tapes were examined primarily to determine the condition of the tube / TSP interface for potentially unsupported tubes in locations which could be susceptible to flow peaking. The examination of potentially unsupported tubes in Sequoyah Unit 1 by TVA found 48 dented tubes between row 1 and row
- 15. (Ref. Fig. 6-2 through 6-5). The remainder of the tubes examined had indications of magnetite to varying degrees at the tube to tube support interfaces. The examination of potentially unsupported tubes in Sequoyah Unit 2 by Westinghouse was limited to potentially unsupported tubes in areas of potentially high flow peaking immediately in front of the AVB's. None of the tubes were dented; however, minimal indications of magnetite were detected to varying degrees at some of the tube to tube support plate interfaces. (Ref.
Fig. 6-6 through 6-9). Because the tube vibration analyses are based on the conservative assumption that all tubes in the area of interest are structurally ' fixed' in the TSP holes, as if by denting or corrosion, the results of this phase of the examination are of interet,t but do not influence the disposition of the tubes found to be susceptible to fatigue.
6.4 AVB Map Interpretations The Sequoyah Unit 1 & Unit 2 SG AVBs, have & nominal design insertion depth )
intended to support as far inward as the Row 11 tubes. Evaluation of the EC j data indicates that in Unit 1, in the area of interest (12 through 8) between eclumns 2 and 93, all of the Row 12 tubes, all of the Row 11 tubes, all but I three Row 10 tubes, and all but nine Row 9 tubes and approximately half of the Row 8 tubes are supported. In Unit 2, between columns 2 and 93, all of the Row 12 tubes, all of the Row 11 tubes, all but two Row 10 tubes, and all but thirty-three Row 9 tubes, and approximately half of the Row 8 tubes are supported.
Bo48M:1E-060589 6-4 j
WESTINGHOUSE PROPRIETARY CLASS 3 Sequoyah Unit 1.SG 1 The AVB map for SG 1 is shown in Figure 6-2. A listing of unsupported tubes is given in Table 6-1. All of the Row 12 and Row 11 tubes are supported. One Row 10 and three Row 9 tubes are unsupported. The highest significant flow peaking factors in this SG (see Sec. B and Sec. 9 of this report) were found at tube locations R10C44, R9C4, R9C43 .and R9C44. The tube at R10C44 exceeds the limiting stress ratio criteria; leading to the recommendation that R10C44 should be removed from service. Peaking ratio, stability ratio, and stress ratio modifications reduce the stress ratios for the remaining tubes to acceptable values of less than unity.
Sequoyah Unit 1 SG 2 The AVB map for SG 2 is shown in Figure 6 3. A listing of unsupported tubes is given in Table 6-1. This SG has no unsupported tubes outside Row 8. The highest flow peaking factors in this SG (see Sec. 8 and Sec. 9 of this report) were found at tube locations R8C35 and R7C60. Peaking ratio, stability ratio, and stress ratio modifications reduce the stress ratios for all tubes in SG 2 to acceptable values of less than unity. ,
Sequoyah Unit 1 SG 3 The AVB map for SG 3 is shown in Figure 6-4. A listing of unsupported tubes is given in Table 6-1. All of the Row 12 and Row 11 tubes are supported. Two Row 10 and two Row 9 tubes are unsupported. The highest flow peaking factors in this SG (see Sec. 8 and Sec. 9 of this report) were found at tube location R8C61. Peaking ratio, stability ratio, and stress ratio modifications reduce the stress ratios for all tubes in SG 3 to acceptable values of less than unity.
Sequoyah Unit 1 SG 4 l The AVB map for SG 4 is shown in Figure 6-5. A listing of unsupported tubes l
is given in Table 6-1. All of the Row 12, Row 11, and Row 10 tubes are 9048M:1 E-052689 6-5
j
WEST 1NGHOUSE PROPRIETARY CLASS 3 supported. Four Row 9 tubes are unsupported. The highest flow peaking factors in this SG (see Sec. 8 and Sec.'9 of this report) were found at tube locations R9C9, R9C10, R9C11,.and R9C91. Peaking ratio, stability ratio, and
~
stress ratio modifications reduce the stress ratios for all tubes in SG 4 to H' aeceptable values of less than~ unity.
P Sequoych Unit 2 SG 1 J
The AVB map for SG 1 is'shown in Figure 6-6. A listing of unsupported tubes is given in Table 6-2. All of the Row 12, and Row 11 tubes are. supported.
Two Row 10 and 24 Row 9 tubes are unsupported. The highest flow peaking ifactors'in this 3G (see Sec. 8 and Sec. 9 of this report) were found at tube locations R10C60, R9C14, R9C35,.R9C40, and R9C56. The tube at R10C60 exceeds the limiting stress ratio' criteria; leading to the recommendation'that R10C60
.should be removed,from service. Peaking ratio, stability ratio, and stress
. ratio modifications reduce the stress ratios for the remaining tubes in SG 1 to acceptable values of less than unity.
1 Sequoyah Unit 2 SG 2
.The.AVB map for SG-2 is shown in Figure 6-7. A listing of unsupported tubes is given in Table 6-2. All of the Row 12, Row 11, and Rrw D tubes are suppcrted. Two Row 9 tubes are unsupported. The highest " low peaking factors in this SG (see Sec. 8 and Sec. 9 of this repert) were found at tube locations R9C35, R8C35, kSC40, and R8C60. Peaking ratio, stability ratio, and stress ratio modifications reduce the stress ratios for all tubes in SG 2 to acceptable values of less than unity.
Sequoyah' Unit'2 SG 3 l The AVB map for SG 3 is shown in Figure 6-8. A listing of unsupported tubes j is given in Table.6-2. All of the Row 12, Row 11, and Row 10 tubes are supported. Five Row 9 tubes are unsupported. The highest flow peaking j factors in this SG (see Sec. S and Sec. 9 o' this report) were found at tube 904BM 1E*0s2689 6-6 x i
WESTINGHOUSE PROPRIETARY CLASS 3 locations R8C9, RSC10, R8C35, RBC60, and R8C61. Peaking ratio, stability ratio, and stress ratio modifications reduce the stress ratios for all tubes in SG 3 to acceptable values of less than unity. j Sequoyah Unit 2 SG 4 The AVB map for SG 4 is shown in Figure 6-9. A listing of unsupported tubes is given in Table 6-2. All of the Row 12, and Row 11, and Row 10 tubes are supported. Two Row 9 tubes are unsupported. The highest flow peaking factors in this SG (see Sec. 8 and Sec. 9 of this report) were found at tube locations ,
R9C60 and R8C9, R8C10, R8C35, and R8C60. The tube at R9C60 exceeds the limiting stress ratio criteria; leading to the recommendation that.R9C60 should be removed from service. Peaking ratio, stability ratio, and stress ratio modifications reduce the stress ratios for the remaining tubes in SG 4 to acceptable values of less than unity.
l 9048M 1E-052689 6-7 l
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WESTINGHOUSE PROPRIETARY CLASS 3 Table 6-1 Sequoyah Unit 1 Summary Listing of Unsupported Tubes SG #1 Row 12' Row 12 has no unsupported tubes Row 11- Row 11 has no unsupported tubes Row 10- l Columns.44 is unsupported Row 9 Columns 3, 43, & 44 are unsupported Row'8 Columns 2 thru 6, 9, 10, 34, 43 thru 48, & 59 are unsupported
'SG #2 Row.12 Row 12 has no unsupported tubes Row 11- Row 11 has no unsupported tubes Row 10-. Row 10 has no unsupported tubes Row 9. Row 9 has no unsupported tubes Row 8 Columns 2'thru 16, 35, 49, 51, & 79 thru 84 are unsupported SG #3 Row 12 Row 12 has no unsupported tubes Row 11 . Row 11 has no unsupported tubes Row 10 Columns 2 & 3 are unsupported Row 9 Columns 2 & 3 are unsupported Row 3 Columns, 2 thru 5, 10 thru 16, 34, 35, 40 thru 55, 60 & 61 are unsupported I
SG #4 Row 12 Row 12 has no unsupported tubes Row 11 Row 11 has no unsupported tubes Row 10 Row 10 has no unsupported tubes Rcw 9 Columns 9, 10, 11, & 91 are unsupported Row 8 Columns 2 thru 15, 34, 35, 39 thru 54, 59, 82 thru 85, & 88 thru 93 are unsupported b
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Table 6-2 D Sequoyah Unit 2 Summary Listing of Unsupported Tubes SG #1 Row 12 Row.12 has no unsupported tubes Row 11 Row 11 has no unsupported tubes Row 10 Columns 52 & 60 are unsupported
, ' Row 9 ' Columns 11 thru 14, 35, 40 thru 56, 60, & 61 are unsupported L Row 8 Columns 2 thru'16, 23 thru 28, 31 thru 35, 38 thru 57, 60 I
thru 64, 67 thru 70, 79 thru 86 are unsupported SG #2 Row 12- Row 12 has no unsupported tubes Row.11 Row 11 has no unsupported tubes Row 10 Row 10 has no unsupported tubes Row 9 Columns 35 & 83 are unsupported Row 8 Columns 2 thru.16, 25, 26.-34, 35, 39 thru 56, 60, 61,-& 79 thru 93 are unsupported SG #3 Row 12 Row 12 has no unsupported tubes Row 11 Row 11 has no unsupported tebes Row 10 Row.10 has no unsupported tubes s Row 9 Columns 2, 34 45, 49, & 60 are unsupported Row 0 Columns 2 thru 6, 9, 10, 13 thru 16, 35, 39 thru 56, 60 thru 63, 83 thru 86, & 91 are unsupported i SG #4 Row 12 Row 12 has no unsupported tubes l Row 11 Row 11 has no unsupported tu'oes Rcw 10 Row 10 has no unsupported tubes Row 9 Columns 35, & 60 are unsupported Row 8 Coiumns 2 thru 7, 10 thru 16, 34, 35, 39 thru 57, 60, 61, 79 thru 89, 92, & 93 are unsupported 1
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Tube / Tube support interf aces at TSP #7 have been examined.
Alltube interfaces have have some indication of corrosion and/or magnetite Tubes highlighted have been examined and found to be dented with deformation at TSP # 7 Figure 6-2 Sequoyah Unit 1 Steam Generator 1 - AVB Positions
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Tube / Tube support interfaces at TSP #7 have been examined.
All tube interfaces have have some indication of corrosion and/or magnetito Tubes highlighted have been examined and found to be dented with deformation at TSP # 7 Figure 6-3 Sequoyah Unit 1 Steam Generator 2 - AVB Positions
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Tube / Tube support interfaces at TSP #7 have been examined.
All tube interfaces have have some indication of corrosion and/or magnetite Tubes highlighted have been examined and found to be dented with deformation at TSP # 7 figure 6-4 Sequoyah Unit 1 Steam Generator 3 - AVB Positions 9167M:1E-051789 6-13
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Tube / Tube support interfaces at TSP #7 have been examined.
All tube interfaces have have some indication of corrosion and/or magnetite Tubes highlighted have been examined and found to be dented with deformation at TSP # 7 Figure 6-5 Sequoyah Unit 1 Steam Generator 4 - AVB Positions
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figure 6-7 Sequoyah Unit 2 Steam Generator 2 - AVB Positions 904BM 1E-052689 6-16 l
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17.0 TL;RMAL AND HYDRAULIC ANALYSIS This section' presents the results of a thermal and hydraulic analysis of the flow field on the secondary side of the si.eam generator using the 3D ATH0S computer code, Reference (7-1). The major results of the analysis are the water / steam velocity components, density, void fraction, and the primary and s:condary fluid and tube wall temperatures. The distributions of the tube gap .
volocity and density along a given tube were obtained by reducing the ATHOS rssults. The ATH0S distributions used in the Sequoyah snalysis are based on an enveloping set of operating conditions which were defined to provide some added margin to the evaluation. Application of these enveloping conditions to the Sequoyah Unit 1 and 2 generators provides conservative U-bend parameters for the calculation of fluidelastic stability ratios.
In the following subsections, the operating condition data for both Sequoyah units is presented along with a comparison with the enveloping conditions used in the 3D tube bundle study. A description of the ATHOS model and some sample results are included in the next two sections. The final section describes an analysis of the operating history data for Sequoyah 1 and 2. This analysis defines a parameter termed the normalized stability ratio which provides a relative indication of the effect of past operation on the plant's {
fluidelastic stability ratio.
7.1 Sequoyah Steam Generator Operating Conditions Full power steam generator operating condition data for the Sequoyah units was provided by the Tennessee Valley Authority. These data are listed in Table 7-1. A single set of data was supplied for Unit I which is applicable to both the most recent full power operation in Cycle 3 (through August 1985) and all previous operation since the plant started up in July 1980. Similarly, one
. set of data applies to all full power operation in Unit 2 from startup in j l
November 1981 through June 1985. Note that both units experienced three year !
l shutdowns which extended from 1985 to 1988. Unit I returned to full power in December 1988; Unit 2 returned to full power in April 1988.
4 I
9049M;1E-060sE9 7-1 I
______________j
n l' H WESTINGHOUSE PROPRIETARY CLASS 3
) -With' the'above data, calculations were completed using the Westinghouse SG performance computer code, GENF, to verify the plant data and to establish a g Lcomplete. list of. operating conditions required.for an ATH0S analysis. The
'GENF code determines the primary side temperatures and steam flow rate 14 . required to obtain the specified steam pressure at'the given power rating.
Besides confirming these parameters, the code calculates the circulation ratio
~
fwhich is of primary importance.to the stability ratio analysis.since it, together with the steam flow, establishes the total bundle flow rate and average loading'on the tubes. It also provides an-overall indication of the voids within the tube bundle since the bundle exit quality is inversely proportional to the circ' ratio (xexit = 1/ circ ratio).
The calculated circulation' ratio along with the other thermal / hydraulic conditions are listed in Table 7-1. Also listed are the enveloping conditions which were defined for use in the ATHOS flow field calculation. With respect l to their effect on tube stability ratios, three of the operating parameters l
are of prime importance: power level / steam flow, steam pressure, and the circulation ratio. (Primary side temperatures have only a.very minor
' influence on stability ratios). As mentioned previously, the steam flow rate-and circulation ratio influence the total bundle flow rate and tube-to-tube gap velocity in the U-bend. The steam pressure also influences the gap velocity via the void fraction and density, however, its major impact is on the tube damping'. High U-bend flow along with low steam pressure results in a higher. loading on the tubes with reduced damping. Both of these factors lead to higher, more limiting stability ratios.
As indicated by the comparison shown in Table 7-1, tho enveloping conditions essentially match Sequoyah's steam flow rate, circulatten rstio, and bundle flow rate. However, the lower steam pressure (800 psia) detined for the enveloping conditions adds conservatism to the Sequoyah analysis since tube damping is reduced. An overall measure of the effect of differences in operating conditions on the stability ratio can be of tained ' rom a parameter termed the 10 relative fluidelastic stability ratio whbh is also listed in Table 7-1. A detailed description of this parameter is provided in later sub-sections. However, for the present discussion it is sufficient to state that a higher value for this ratio indicates less margin to fluidelastic 9049tt1E-052SB9 7-2 l
l
WESTINGHOUSE PROPRIETARY CLASS 3 vibration instability and tube fatigue. The higher relative stability ratio calculated for the ATH0S conditions implies that the enveloping conditions aru more limiting and will provide added conservatism when the resulting U-bend distributions are applied to the Sequoyah stability ratio evaluation.
7.2 ATH05 Analysis Model The calculation of relative stability ratios involves comparing the stability ratio calculated for one or more tubes in a given plant to the ratio calculated for the ruptured Row S Column 51 tube in the North Anna Series 51 steam generator. It makes use of ATHOS computed flow profiles for both tube bundles. Since the presence of AVBs in the U-bend region of a tube bundle l could influence the overall flow field and/or the local flow parameters for a particular tube of interest, some discussion of the treatment of AVBs is j necessary before presenting a description of the ATHOS model.
The ATH0S code does not include the capability to model the presence of the ;
AVBs in the U-bend region. However, Westinghouse has modified the code to include the capability to model the AVBs via flow cell boundary resistance factors. Practical lower limits of cell size in the ATHOS code, however, ,
prevent a fine grid representation of the AVB V-bar shape which, in turn, limits the accuracy of the AVB representation. ATHOS calculations have been performed with and without AVBs in the model. Calculations of stability ratios relative to North Ar.na R9C51 show that the relative stability ratios i for tubes near the center of the steam generator are essentially the same for models with or without AVBs. The ATH0S AVB modeling' sensitivity studies with !
uniform insertion show some tendency for the AVB resistance effects to lower )
tube gap velocities near the central regions and to increase velocities near I I
the peripheral tubes. However, the magnitude of this effect is uncertain due 1 to the limitations in ATHOS for modeling the AVBs. Further, the global flow I) resistance of staggered AVB insertion would be less than that from uniform ]
insertion. Based on the sensitivity studies using ATH0S models with and I without uniformly inserted AVBs. the most reliable relative stability ratios (for actual steam generators with non-uniform AVB insertion depths) are expected using ATHOS models excluding AVBs and effects of variable AVB l
9049MdE-052689 7-3 I l
l.
WESTINGHOUSE PROPRIETARY CLASS 3 J
l insertion depths. Those AVB effects are accounted for by using flow test results of actual AVB' geometries. This approach has been utilized in the Sequoyah analysis, i
The Sequoyah analysis is based on a Cartesian coordinate system for the array 4 of flow cells instead of the typical, and more widely used, cylindrical coordinate system. With a Cartesian coordinate system, the tube array and any. j AVBs are arranged in a square pitched configuration which is in-line with the i coordinate axes. This alignment provides an improved representation of the i tube bundle.
The ATHDS Cartesian coordinate system model for the Sequoyah steam generator consists of 13,050 flow cells having 30 divisions in the x-axis (perpendicular to tM tubelane) direction,15 divisions in the y-axis (along the tubelane) {
direction and 29 divisions in the axial (z-axis) direction. In the ATHOS
. analysis, the steam generator is considered to be symmetrical about the x-axis of the tube bundle. The model therefore, consists of one-half of the hot leg and one-half of the cold leg sides of the steam generator. Figures 7-1 and ,
1 7-2 show the plan and the elevation views of the model. These two figures l show the layout of the flow cells and identify locations for some of the geometric features. ;
i 1
As shown in Figure 7-1, with the Cartesian coordinate system, the circular J wrapper boundary is represented by a step-wise wall as indicated by the heavy lines. Downcomer inlet ports are located on the extreme left and right sides )
of the model with a solid boundary at the top (IY=15') and the plane of symmetry at the bottom. All of the flow cells outside the simulated wrapper 3 boundary above the first axial slab were blocked off by specifying extremely {
high flow resistances on the faces of the appropriate cells. Tubelane flow slots in the tube support plates are also modeled. {
Figure 7-3 reproduces the plan view of the model but with the tube layout arrangement superimposed. This ficare illustrates the locations of the tubes in the various flow cells. The fineness of the cell mesh is evident; the largest cells contain only 20 tubes while some of the smallest cells include ,
only three tubes. Note, in particular, that additional detail was added near l
l 0049M.1E-Os2689 7-4 1
y 1 WESTINGHOUSE PROPRIETARY CLASS 3 the bundle periphery (IY=12-15) to more closely model the inner radius tubes
_ (rows <15). For this same reason, five thin' axial layers of cells were.
j included in the U-bend near the top tube support (Figure 7-2, 12=16 to IZ521) to more closely model the flow conditions in the area of interest.
- 7.3 ATH0S Results l
The.results'from the ATHOS' analysis consist of the thermal-hydraulic flow ,j parameters necessary to describe the 3-D flow field on the secondary side of j the steam generator (velocity, densciy, and void fraction) plus the 'i distributions.of'the primary fluid and mean tube wall temperatures. The ;
secondary side mixture velocity is :omposed of three components (Vx, Vy, and l Vz).which ATHOS computes on the surfaces of the flow cell. Since the local gap ve.locity surrounding a tube is required in the vibration analysis, a <
post processor is used which: a) interpolates among the velocity components for the cells located nearest to the tube of interest and, b) accounts for the minimum flow area between tubes to calculate the tube-to-tube gap velocity.
The post processor performs the necessary interpolations to determine both' in plane and out-of plane gap velocities at specific intervals along the
' length of a tube. It also interpolates on the ATHOS cell-centered density and void fraction to determine the required local parameters along the tube length. . The output of the post processing is a data file which contains these parameter distributions for all the tubes in the generator and which provides a portion of the input data required for tube vibration analyses.
. Figure 7-4 shows a vector plot of the flow pattern on the vertical plane of symmetry of the steam generator (the vectors are located at the center of the flow cells shown in Figure 7-2). It is seen that in the U-bend region the mixture turns radially outward, normal to the curvature of the bends toward the region of least flow resistance (i.e., outside the dome formed by the U-bends). Also, because of higher heat flux and void generation, the velocities in the hot leg are higher than in the cold. This difference persists up to the entrance to the U-bend as indicated by the velocity contours shown in Figure 7-5. Here, the axial component of velocity in the hot leg is about 50% higher than in the cold leg. This figure also indicates the high axial flow component which has just exited the three tubelane flow 9049M:1E-052689 7-5
s ,
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-WESTINGHOUSE. PROPRIETARY. CLASS 3 ilots in the top tube support plate. The. lateral velocity component, .l
'Vr.=/Vx2 , yy ;2in the same_ horizcntal plane (12=16) is shown in Figure 7-6.- Comparing Figures 7-5' and.7-6, it is seen that at the ' entrance to the LU-bend the vertical velocity. component is several times higher than the lateral velocity component-in both the' hot and cold legs.
Figure 7-7 presents a plot of the.. void fraction contours on the vertical plane of symmetry of the-steam generator. It is evident that the void fraction develops rapidly on'the bot leg side of the lower tube bundle. The higher voids'in the hot leg continue'up into the U-bend where the void fraction varies from about 0.8-0.9 in the hot leg and from 0.6 to 0.85 in the cold leg.
Figures 7-8,~7-9 and 7-10 provide a sample of the individual tube gap velocity.
and density distributions as computed by the ATHOS post processor. Results for three Row 10 tubes are plotted. In each figure the gap velocity and density along the length of the tube are plotted from the hot leg tubesheet on the left side of the figure to the cold leg end on the right.
Figure 7-11 presents a plot'of the average in plane gap velocity normal to the tube and density profiles as a function of the column number along Row 10.
The average values were taken as the numerical average of the parameter over the entire 180* span of'a U-bend at a given column location. The average velocity is seen to be relatively constant with values ranging from 10.5 to 10.9 ft/sec. However, the average density shows more variation across the bundle with lower values present in the interior of the bundle and higher values on the periphery.
7.4 Relative Stability Ratio Over Operating History One aspect of the evaluation of the Sequoyah steam generators is to examine the operating history data and use it to determine the susceptibility to
. fatigue from fluidelastic vibration resulting from the three cycles of operation in each unit. This assessment has been completed through the use of a parameter termed the normalized stability ratio. The normalized stability ratio compares the fluidelastic stability ratio for each period of a plant's operation (fuel cycle) to a reference stability ratio typically based on a 9049M;1E-Os2689 7-6
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r J WESTINGHOUSE PROPRIETARY CLASS.'3
'recent operating condition. i A' plot'of this ratio ~against operating time, i 1therefore, provides a relative. indication of,the effect of past operation on the: plant's fluidelastic stability ratio. This normalized time-dependent
~
ratio'.is subsequently combined with an absolute stability ratio for the.
reference operating conditions' derived from detailed three-dimensional-
' thermal / hydraulic and. tube vibration calculations. High values for the. net.
< st$bility, ratio,inparticular,overasignificantperiodofoperation, coupled with other pr@equisite conditions (e.g., absence of- AVB support and I denting at the top tube support plate), could indicate an increased i susceptibility to fluidelastic vibration instability and fatigue.
The fluidelastic stability ratio is defined as the ratio of the effective fluid velocity acting on a given tube'to the critical velocity at which large amplitude fluidelastic vibration initiates:
U Fluidelastic effective Stability Ratio, SR = _ [1]
U critica! at onset of instability In this ratio, the effective velocity deoends on the distribution of flow velocity and fluid density, and on the mode shape of vibration. The critical velocity is based on experimental data and has been shown to be dependent upon the tube natural frequency, damping, the geometry of the tube, the tube pattern, and the fluid den:ity, along with the appropriate correlation
. coefficients.
The detailed calculation of this ratio using velocity and density distributions, etc., requires three-dimensional thermal / hydraulic and tube vibration calculations which are time consuming. Alternately, a simplified, i one'-dimensional version of this ratio has been used to provide a relative assessment technique for determining the effect of past operation on the stability ratio. The normalized stability ratio is defined by the following equation:
9049M;1E-052689 7-7
WESTINGHOUSE PROPRIETARY CLASS 3-w ,
~
a,c
'[2]
3 inthisequation"cycx"'referstoeachfuelcycleand" ROP"to.'the.recent operating condition. While this simplified approach cannot account for:
three-dimensional tube bundle effects, it does consider the major operational parameters affecting the' stability ratio. Four. components make up this ratio: a 1oading term based.on'the dynamic pressure (pV
~
2 .
), , tube incremental mass (m) term,' the natural frequency of the tube'(fn ), and a.
dampingratio-(6)' term. It should be noted.that the ratio is relative, in that each component is expressed as a ratio of the value for a given fuel cycle or power level to that of the recent operating point.
[
3a,c ,
The particular damping correlation which is used for all normalized stabilit.y ratio calculations is based on a dented condition at the top tube support
-plate (a clamped condition, as discussed in Section 5.2). The clamped condition is also assumed in calculating the tube natural frequency.
As discussed previously in Section 7.1, the reference three-dimensional stability ratio calculation for the Sequoyah steam generators was based on a set'of enveloping operating parameters which provided some added margin to the analysis. These same conditions were the basis for the reference components in the ID normalized stability ratio calculation, labeled " ROP" in equation 2.
Relative stability ratio calculations were completed for the Sequoyah 1 and 2 full power conditions and for two lower power levels, 90 and 95% of full power in Unit 1. Since tube vibration and possible fatigue are associated with 1
1 9049M:1E-Os2689 7-8 m
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'l WESTINGHOUSE PROPRIETARY CLASS 3- j operation at close'to 100% power, only the higher power operating periods ~are considered'important to the evaluation'. The high power operating experience for both Sequoyah units is summarized in Table 7-2. It lists the number of- q L days in:each fuel cycle that the unit operated within three high power e intervals (85-90, 90-95 and 95-100%). Since the basic operating conditions for each plant's three fuel cycles are identical, the days within'all the i cycles can be combined and the resulting totals used in prep'aration of.the stability ratio curves. Further, it has been conservatively assumed that the total operating time within each of the three power intervals is assigned to the highest power / stability ratio condition in tha interval. ;
1 The.resulting normalized stability ratios for both Sequoyah units are shown .in Figure 7-12. In this figure, the normalized stability ratio is plotted l b
- j. against cumulative operating time above 85% power. The reference value !
l
(=1.00) is for the full power operating condition on which the 3D stability ratios are based. As shown, the full power Sequoyah 1 value is about 7% lower than the reference value which is indicative of the conservatism which was !
built into the 3D analysis. The ratio calculated for Unit 2 is about 5% below-L the reference value. The slightly smaller margin for Unit 2 is the result of I
the lower steam pressure and higher steam flow in Unit 2 compared to Unit 1.
The reduced ratios for Unit 1 at 90 and 95% power are the combined result of- 1 L both decreased 1 riding on the tubes and increased damping. Higher damping is a result of lower voids in the U-bend which occurs when the steam pressure rises at reduced power levels. 4 Figure 7-12 also includes an additional ratio which is appropriate to i calculating the fatigue usage for projected operation in Unit 2. At the time f
of the last full power operation in 1985, no tubes were plugged in any of the Unit 2 generators. During 1988, row 1 tubes were plugged such that the current plugging levels are 94, 94, 95 and 94 tubes in generators 1-4, <
respectively. If the plant does return to full power at some future time, this plugging is calculated to reduce the steam pressure slightly. This ,
results in a slightly higher normalized stability ratio (0.955 versus 0.951).
The calculation of future fatigue usage for Unit 2 (Section 9.0) will make use j of this higher value.
Bo49M:1E-052689 7~9
WESTINGHOUSE PROPRIETARY CLASS 3 f
References:
7-1 L. W, Keeton, A. K. Singha1, et al. "ATHOS3: A computer Program for Thermal-Hydraulic Analysis of Steam Generators", Vol. 1, 2, and 3, EPRI NP-4604-CCM, July 1986.
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9049M:1E-052689 7-1Q
1 M- WESTINGHOUSE PROPRIETARY CLASS 3-Table 7-1 Sequoyah Steam Generator Operating Conditions .and Comparison
-with Enveloping Conditions Used in the 3D ATH0S Analysis Enveloping Condition's Sequoyah Sequoyah Used in 3D' Unit 1 Unit 2 ATHOS Analysis
.SG Thermal Power. 856.0 857.1' 864 (=1.01*PS6)'
(MWT)
' Steam Flow Rate 3.71 x 10 6 3.75 x 10 6 3.73 x 106 (lbm/hr) l l
Feedwater Inlet 431.1 438 430-
' Temperature (*F)
Steam Pressure 890 874 800
-(psia) l Water Level 44 44 44
(% of span)
Primary Inlet / 611.1/544.5 607/547 600/538 Outlet Temperatures-(*F)-
Calculated Parameters Circulation Ratio a,c Bundle _ Flow Rate (lbm/hr)-
10 Relative Stability Ratio _ _
l l 1 9049M:1E-052689 7-11 u
___________________-_-__-_______-_____-______________-____J
3 e WESTINGHOUSE PROPRIETARY CLASS 3 Table 7-2 Sequoyah Operating' History Data n.
DISTRIBUTION OF DAYS IN TOTAL EACH POWER INTERVAL CYCLE BEG END DAYS 95-100% 90-95% 85-90%
Unit 1 1 05-Jul-80 10-Sep-82 797 341 8 18' 2 18-Jan-84 20-Feb-84 398 284 7 5 3 15-Apr-84 23-Aug-85 495 311 4 3' 1690 936 19 26 (941 EFPD's)
Unit 2 1 06-Nov-81 18-Jul-83 619 .389 -- --
2 12-Nov-83 28-Sep-84 321 297 -- --
3 14-Dec-84 18-Jan-89 1496 377 -- --
2436 1063 (EFPD's) i 1
904BM:1E-0526es 7-12 i
WESTINGHOUSE PROPRIETARY CLASS 3 a,b,c r
i Figure 7-1 Plan View of ATH0S Cartesian Model for Sequoyah 9049M 1E-052689 7-13
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! Figure 7-3 Plan View of ATHOS Cartesian Model for 1
l Sequoyah Indicating Tube Layout 9049M:1E-052689 7-15 ;
o WESTINGHOUSE PROPRIETARY CLASS 3 u, .. ;
a, b, c l- _ m
% l I
t i- ',
\.
f5 r n
/
\!
Figure 7-4 Flow Pattern on Vertical Plane of Symmetry 7~16 9049M 1E-052689 _ ~[
s
WESTINGHOUSE PROPRIETARY CLASS 3 a, b, c Figure 7-5 Vertical Velocity Contours on a Horizontal Plane at the Entrance to the U-Bend 9049M:1E-052689 7-17
WESTINGHOUSE PROPRIETARY. CLASS 3 a, b, c I
l i
1 Figure 7-6 Lateral Flow Pattern on a Horinntal Plane ';
at the Entrance to the U-Bend 1
9049M:1E-052689 7-18
_ . _ _ _ _ d
WESTINGHOUSE PROPRIETARY CLASS 3
-- - a, b, c Figure 7-7 loid Fraction Contours on Vertical Plane of Symmetry 9049M:1E-052689 7-19
WEST 1NGliOUSE' PROPRIETARY CLASS 3 a, b, c l
l l
l Figure 7-8 Tube Gap Velocity and Density Distributions J for Tube Row 10/ Column 3 l l
l 9049M;1E-052689 7-20 !
WESTINGHOUSE PROPRIETARY CLASS 3 a,b,c Figure 7-9 Tube Gap Velocity and Density Distributions for Tube Row 10/ Column 20 9049M 1E-052E89 7-21
}-
- o. ,4 WESTINGHOUSE PROPRIETARY CLASS 3 a,b,c Figure 7-10 Tube Gap Velocity and Density Distributions for Tube Row 10/ Column 40 9049M 1E-052689 7-22
WESTINGHOUSE PROPRIETARY CLASS 3 a,b,c Figure 7-11 Average Velocity and Density in the Plane of the U-Bends Normal to Row 10 904SM:1E-052689 7-23
i WESTINGHOUSE PROPRIETARY CLASS 3 l
J l
4 a,b,c .
1 1
l Figure 7-12 Sequoyah Normalized Stability Ratio Based {
on High Power (>85%) Operation 9049M:1E-052689 7-24
1 WESTINGHOUSE PROPRIETARY CLASS 3 l
)
8.0 PEAKING FACTOR EVALUATION !
l This section describes the overall peaking factor evaluation to define the f test based peaking factors for use in the tube fatigue evaluation. The l
- valuation of the eddy current data to define the AVB configuration for North Anna-1 Tube R9C51 is described. This configuration is critical to the tube j fatigue assessments as the peaking factors for all other tubes are utilized J relative to the R9C51 peaking factor. Uncertainties associated with applying !
the air model test results to the tube fatigue assessments are also included l in this section. Included in the uncertainty evaluation are the following contributions:
o Extrapolation of air test results to two phase steam-water o Cantilever tube simulation of U-bend tubes o Test measurements and repeatability o AVB insertion depth uncertainty 8.1 North Anna-1 Configuration !
1 8.1.1 Background {
The AVB configuration of the ruptured tube in North Anna, R9C51, is the reference case for the tube fatigue evaluations for other plants. In i accordance with the NRC Bulletin 88-02, the acceptability of unsupported tubes in steam generators at other plants is based on tube specific analysis j relative to the North Anna R9C51 tube, including the'relctive flow peaking l factors. Thus, the support conditions of the R9C51 tube are fundamental to the analyses of other tubes. Because of the importance of the North Anna tube, the support conditions of this tube, which were originally based on "AVB Visible" interpretations of the eddy current test (ECT) data (Figure 8-1), j were reevaluated using the projection technique developed since the North Anna event. The projection technique is particularly valuable for establishing AVB positions when deposits on the tubes tend to mask AVB signals such as found for the North Anna 1 tubes. The results of this evaluation are summarized below.
i 9049M:1E-052689 8-1
WESTINGHO SE PROPRIETARY CLASS 3 801.2' Description of.the_ Method The' basic method utilized was the projection technique in which the AVB-position.is determined base'd'on measured AVB locations'in. larger row tubes in the same column. In this study, the projection technique was utilized in.the
" blind" mode, (AVBs called' strictly based on the data) as well as the: reverse mode'(data examined on the basis of predicted AVB positions). The objective.
of this application was, with the greatest confidence possible, to establish the positions of the AVBs in an 8 column range around the_R9C51 tube in North
' Anna 1, Steam Generator C..
8.1.3 Data Interpretation The ECT traces for the U-bends in Rows 8-12 (in one case, 13) were examined for Columns 48-55. The original AVB visible calls are shown in Figure 8-1.
The data were examined by an addy current analyst experienced in reading these traces, and by~a design engineer knowledgeable in the geometry of the Model 51 U-bend region.
The intent of this review was to determine if the presence or absence of AVBs as shown in Figure 8-1 could be confirmed using the AYB projection technique.
Preliminary projected AVB positions were based on geometric data provided for a few of the tubes near R9C51. The features which were sought were evidence of data " spikes" where AVBs were predicted, offset indications (multiple spikes) where offset AVBs were predicted, single indications where single AVB intersections were predicted, etc. The data evaluat' ion method used was a critical examination of the data, which was biased toward the presence of AVBs j unless a confident call of "no AVB" could be made, and then checking the ;
consistency of the data among the tubes in a column and against the j theoretical data for the predicted AVB positions. [
l
)a,c i
9049M:1E-052689 8-2 J
,, + ,
')._
ch WESTINGHOUSE PROPRIETARY CLASS 3 T .
'i
=
i;
, )'.
Figure 8-4 is the "AVB visible" map for columns 48 through 55, based on the L critical review of the' data. It should be noted that the original data interpretations and the review interpretations are consistent.
8.1.4- Projections The[~ )s,c ECT traces were
^'
utilized for projecting the position of the AVBs according to the standard format of the projection method.
The.results of the projections are presented in Figure 8-5, which shows a 1 ' matrix' of projections for tube rows 8 through 13 in columns 48 through 55.
For many of the tubes, more than one, and as many as three projection values are shown.. Multiple projections are expected for a tube if the AVBs on either b side of the tube are not at the same elevation, or if the upper and lower AVB 9049M:1E-Os2689 8-3
'1 1
, k" \
r
- WESTINGHOUSE PROPRIETARY CLASS 3 support that' tube. As many'as four'different projections are possible if it is assumed that the tube is supported by the. upper and lower AVBs, and both upper and lower bars are staggered in elevation'as'shown in Figure 8-2.
The logic'in arranging the projection data is based on the following two rules:
Rule l'. The projections of the same AVB based on different tubes in the -
same column [ Ja,c ,
[-
l 3a,c ,
a u.
Rule 2. Two adjacent tubes in the same row I 8 'C 1 . Consequently, the difference in the
{
3a,c ,
The implication of this is that if the position (either left or right) of a projected AVB is assumed for a column, then the projections in the adjacent columns are also [
ja.c ,
9049M:1E-052689 8-4 '
.i
~
,4 h ', (
l' t (SESTINGHOUSEPROPRIETARYCLASS3
,. ..f
'The
- arrangement of the AVBs as shown in, Figure 8-5 satisfies the rules above 1 'q and is consistent;with the rupture'of R9C51. The resulting AVB arrangements, based on the projection matrix of Figure 8-5 is shown in Figure.8-6.
~
8.1' . 5 Conclusions l .
.The general-AVB arrangement surrounding the ruptured tube in North Anna-1, Steam Generator C which was the basis'for the analysis, is confirmed by a detailed critical-review of the ECT data. Differences exist in the AVB pattern between tube columns 48-49, in which the AVBs appear-to be less inserted than previously indicated. The pattern of~ Figure 8-6 is the best fit-to'the rules which were adopted for determining the position of the AVBs, as well'as consistent with explanation.of the-tube failure.
The basis.of the review wasLa projection technique which utilizes data from tubes one'or more rows removed'from the actual inserted position of the AVB to determine the position'of the AVB. The intent of the review was to establish the' positions of the AVBs by confirming or eliminating features of AVB alignments such as side to side offsets, etc. of the AVBs adjacent to the tubes. Overall, the conclusions regarding the positions of the AVBs around R9C51 in North Anna-1, Steam Generator C are based on consistency among all the available data.
B.2 Test Measurement Uncertainties The descriptions of the peaking factor tests and apparatus were provided in
.Section 5.4. All practical measures were taken to reduce uncertainties.
Nevertheless, some still remain and should be properly accounted for. The important parameter measured during testing that has a significant impact on peaking factor is the air velocity. The air velocity at test section inlet was measured using a [ la,c Based on considerable experience with the use of such instruments, it is known that the magnitude of uncertainty is very small. A[ ]C measurement uncertainty is used in L this analysis' based on past experience.
9049M:1E-0s2689 8-5 t
WESTINGHOUSE PROPRIETARY' CLASS 3 I8.3' Test Repeatability I
During the peaking factor testing of AVB configuration, each test was performed at least two. times to confirm repeatability. It has been
/d:monstrated that the tests are quite repeatable with the results often
- , . falling within 2 or.3% of'one another for the repeat tests. An upper bound; value of 5% was used in'the current uncertainty analysis.
8.4 Cantilever vs U-Tube e :
1:
A first order estimate can-be made of.the validity of modeling a U-bend tube
~
L by a cantilever tube in tests.to determine the effects of AVB insertion depth l' on the initiation of fluidelastic vibration. .The following assumptions are used:
a,C i
em 9049M:1 E-052689 . 8-6
WESTINGHOUSE PROPRIETARY. CLASS 3 For the purposes of this estimate,'the' geometry of the cantilever measuring -
tube in the air test model is~ compared with the geometry of a prototypical Row.
10 tube. [
ja,c ,
The comparison between a U-bend tube and the model tube involve the consideration of an effective velocity associated with the flow perturbation caused by the AVBs. [:
I
)a,c
- 9049M;1E-052689 8-7
l f
W WESTINGHOUSE.PROPRIETARYCL^SS3 I ~
i Ja,c . Using these values, the ratio of the effective velocity for the cantilever measuring t'ube to that for the U-bend tube is about.[ Ja,c for the case treated.
A similar evaluation can be made for a Row 10 tube that. lies in the projection or shadow of an AVB that is inserted to a depth required to support a Row 9.
tube. [
3a,c ,
Thenetresultisthattheratiooftheeffectivevelocityforthecantilever tube to that for'the U-bend tube is about [ Ja.c ,
These results indicate that, for-the particular assumptions.used, the cantilever tube model appears to be a reasonable representation of the U-bend with respect to determining relative peaking factors for different AVB configurations. .This evaluation also shows that, on the average, the magnitude of the systematic uncertainty associated with the use of cantilever tube'to simulate the U-bend is about [ Ja,c ,
8.5 Air vs Steam-Water Mixture
.The local peaking factors from the air tests can be applied to the steam generator steam / water conditions either as a direct facto.- on the mixture velocity and thus a direct factor on a stabilit, ratio, or as a factor on the j I
steam velocity only with associated impacts on density, void fraction and damping. This method leads to a reduction in tube damping which enhances the a I
peaking factor compared to the direct air test value. For estimating an
- p. . absolute stability ratio, this application of the peaking factor is a best estimate approach. However, for the evaluation of tubes relative to stability 1 ratio criteria, it is more conservative to minimize the peaking factor for the j
9049M:1E-052689 8-8 l
j i
i di WESTINGHOUSE PROPRIETARY CLASS 3 North Anna Unit 1. tube R9C51 through direct application of the air test
)
peaking factor. This conservative approach is therefore used for evaluating i
. tube' acceptability.
K Urder uniform AVB' insertion (or aligned AVB' insertion),' there are no local open channels for flow'to escape preferentially. Therefore, air flow is approximately the same'as steam / water flow relat,ive to velocity perturbations. Under non uniform AVB insertion the steam / water flow may
. differ from air, as the steam and water may separate from each other when an obstruction, such as an AVB, appears downstream. The water would continue along the same channel'while steam readily seeks a low resistance passage and thus turns into adjacent open channels. Two phase tests indicate a tendency for. steam to preferentially follow the low pressure drop path compared to the water phase.
' Based on the above discussion, the F$ are considered to more appropriately i apply to the steam phase. Thus, it follows that mixture mass velocity for the i
L tube subject to flow perturbation can be written as follows:
1 a,c l
l where D is the vapcr density, D f the water density, F, the velocity peaking factor determined from air tests, jg* the nominal superficial vapor velocity, and jf* the superficial water velocity. Steam quality can then be determined as follows:
a,c l
.=
9049M:1E-052689 8-9 J
m a ,
WESTINGHOUSE PROPRIETARY. CLASS 3 l
LThe Le11ouche-Zolotar correlation (algebraic slip model), as used in.the.ATHOS code', is applied to determine' void fraction. Subsequently, mixture density, .I v31ocity and damping. coefficients for the tube which is not supported and.
subject to flow perturbation is evaluated. Therefore, similar.to th's air
~
velocity peaking factor, local scaling' factors of mixture density'and velocity and damping coefficient can be readily determined. Finally, a. local stability
-peaking factor for fluidelastic vibration can be calculated as follows:
._ _. a,c where Fs is the stability peaking factor, Fd the density scaling factor, Fv the velocity scaling. factor, and Fdp the damping coefficient scaling factor.
If we use the air velocity peaking factor without translating to steam / water conditions, then a,c As shown in Table 8-1 stability peaking factors for the steam / water mi.xture are slightly higher than air velocity peaking factors. The difference between
.the steam / water and air peaking factors increases as the air peaking factor increases.
'For application to tube fatigue evaluations, the ratio of the peaking factor for a specific tube to that for North Anna R9C51 is the quantity of interest.
Larger values for this ratio are conservative for the tube fatigue assessment. The North Anna R9C51 peaking factor is one of the highest peaking factors. 'As discussed in Section 8-7, a peaking factor of nearly [ 3a,c is determined for the R9C51 tube. The differences between [
]a,c Typical values are shown in Table 8-2. These results show that the direct application of the air test data yields the higher relative peaking factor compared to R9C51. To obtain conservatism in the peaking factor evaluation, [
3a,c ,
9049M.1E-052689 8-10 a
WESTINGHOUSE PROPRIETARY CLASS 3 Comparing the values in the first and last columns of Table 8-1, it may be noted that the stability peaking factor for steam water is [ ]a,c higher than the air velocity peaking factor. On the average, the uncertainty associated with the conservative use of air velocity peaking factor is ,
[ Ja,c , .
The conclusion that the peaking factor for steam water flow would be higher due to the dependency of damping ratio on void fraction was supported by an alternate study. In this study, a section of steam generator tubes were simulated using the ATHOS code under protoypic flow conditions. The objective of this study was to examine the magnitude of the changes in void fraction and I thus stability ratio as a consequence of non-uniform AVB insertion patterns.
The current version of ATH0S has modeling limitations that prevent accurate modeling of local geometry effects. In addition, it is believed that an analysis using two-fluid modeling procedure is mandatory to a calculation of the peaking factors for a steam generator to account for the preferential steam flow along the low resistance path. Conseque'ntly, the intent of this analysis is only to help bound the uncertainty on void fraction effects from extrapolating the air tests to steam-water.
First the analysis was conducted with uniformly inserted AVBs in the ATH0S model. The ATH0S results were processed by the FLOVIB code to determine stability ratios for the specific tubes of interest. The calculation was repeated using a non-uniform AVE insertion pattern in the model. The results show that the void fraction distribution changes as a result of flow perturbation. Further, the impact on stability ratio resulting from the changes in void fraction profiles was about [ ]a,c This alternate calculation provides independent corroboration of the prior discussion regarding the stability peaking factors under steam-water conditions vs. in air.
B.6 AVB Insertion Depth Uncertainty The most significant uncertainty for the low peaking configurations is not in j j the test results, but in the determination of actual AVB insertion patterns j adjacent to specific tubes. The methodology used for obtaining the AVB I
i 9049M:1E-052689 B-ll 1
.- ..a
1 WESTINGHOUSE PROPRIETARY CLASS 3 insertion patterns from eddy current data'can ascertain the-AVB location only to within approximately [ ]
Ja c . The effect'on peaking. factor resulting from this uncertainty is addressed using test'results of AVB configurations that varied from ons )
ansther by up to [ ]a,c ,
Based on maps of AVB insertion depth of various plants, several. configurations q have been tested for determining fluidelastic instability flow rate by an air !
cantilever model. Stability peaking factors were then determined from the y ratio of critical flow rate for a uniform AVB insertion configuration to a y specific configuration. Figure 8-7 summarizes the AVB configurations tested.
1 Position of AVB' insertion depth is determined from Eddy Current Test (ECT) data. Positioning of AVB from ECT data reading is subject to uncertainty; its accuracy is probably about [ ]a,c A change of an AVB L insertion depth in'a given configuration leads to a different configuration, and thus a different peaking factor. A review of the tested AVB type has been made and results summarized in Table 8-3. As.can be seen, a decrease in depth of an appropriate AVB tends to decrease the peaking factor, for instance, a
[
la,c Such a trend can be explained; a decrease in a specific AVB depth will open up more channels for incoming fluid to. distribute and thus less flow perturbation. However, this applies only to those changes without inducing the reinforcement of flow perturbation from upstream to downstream, i
On the average, the uncertainty in peaking factor resulting from small variations in AVB insertion (of the order of 1/2 tube pitch) is found to be
[ Ja,c , i l l l 8.7 Overall Peaking Factor with Uncertainty l
As discussed in the previous subsections, there are several aspects to be considered in applying the laboratory test data to steam generator conditions. These considerations were reviewed one at a time in those subsections. This section will integrate the pieces into one set of stability peaking factors.
9049M.1E-Os2689 8-12
)
i WESTINGHOUSE PROPRIETARY CLASS 3 I 1
Looking forward to how these peaking factors are used in the analysis (Section 9), the relative stability ratio calculated for a given tube without the ;
consideration of flow peaking is corrected using the ratio of the peaking factor of the specific tube to that of the North Anna R9C51 tu'oe (Configuration la).
1 1
It is to be noted that the test results would be applied as ratios of a ,
l
-specific tube peaking factor to the R9C51 peaking factor. This will reduce J the influence of some uncertainties since the systematic uncertainties would affect both the numerator and the denominator in the ratio of peaking l factors. The major difference will be in those configurations whose peaking factors are significantly lower than that of R9C51. The approach employed !
here is intended to provide that conservative peaking factors are employed for such apparently low peaking configurations.
The uniform AVB configuration (2a) is selected as a reference configuration, and the peaking factors of all configurations tested are recomputed on the ,
basis of this reference. As discussed below, some of the test uncertainties are applied to the reference case to account for its significantly low peaking relative to the RSC51 configuration.
4 The uncertainties in the test results and their extrapolatim are those due to test measurements, test repeatability, cantilever tubes in the test vs U-tubes in the steam generator, and air tests vs steam-water mixture. These were discussed in more detail in the previous subsections. The magnitude of these uncertainties are listed in Table 8-4.
Of these uncertainties, those due to measurement and repeatability of tests r e randem errors and can occur in any test. Therefore, these are treated together. The total random uncertainties are calculated by [ I
]a,c The RSS value of these is
[ ]a,c . Since these can occur in any test, these are to be applied to all tests. One way of doing this is to apply it to the R9C51 value, that being in l the denominator of the final peaking factor ratio. Thus the peaking factor for configuration la (R9C51) is reduced by this amount to yield a value of j [ ]a,c instead of the [ ]a,c appearing in Table 5-2.
I l
9049M;1E-052689 8-13
]
m f{
3' WESTINGHOUSE PROPRIETARY CLASS 3 j 1
The next three uncertainties in Table 8-4 are systematic uncertainties. It could be argued that these appear in the peaking factors of both the specific tube under consideration and the R9C51 tube and are'therefore counter j
- balanced. However, the relative magnitude of these may be'different, particularly for configurations with much' lower peaking than R9C51. Therefore -
it was' judged that the [
]a,c Similarly, as.noted above, the effect on peaking factor
. j
-due to the uncertainty in the field AVB configuration is also included in this 1 reference case. Thus,[
la c The peaking factor of the reference configuration 2a (Table B-5) is raised by this amount to a value of [ la.c ,
The change in peaking factors of configurations la and 2a resulting from the j application of uncertainties as described above are shown in Column 3 of Table 8-5. The peaking factors of all configurations are recomputed on the basis of ~
this reference configuration (2a). These values are displayed in Column 4 of Table 8-5.
Some of the uncertainties were applied to the reference configuration (2a) in order to apply them to all low peaking configurations conservatively. Thus, no configuration should have a lower peaking factor than this reference configuration. Therefore, when a peaking factor value less than [ )"'C .
is calculated for any configuration, (in Column 4 of Table B-5), it should be altered to [ 3a,c . Further, for some of the configurations that are conceptually similar, the more limiting (higher) value is used. For example, a peaking factor of [ Ja.c is used for configurations 5a and 5b based on
.their similarity to configuration Sc.
The final stability ratio peaking factors calculated on this basis (with configuration 2a as the reference) are shown in Table 8-6.
l l
l- l 9049M.1 E-0!i2689 8-14 l
...j
1 WESTINGHOUSE PROPRIETARY CLASS 3 The overall conclusions from the peaking factor assessment are:
- 1. As noted in Table 8-4, five elements have been included in the' uncertainty evaluation for the peaking factors. The uncertainty estimates were developed from both test and analysis results as described. I in Sections 8.2 to 8.6. The largest single uncertainty.of [ ]a,c $3 i attributable to uncertainties of up to [ ]a,c on determination of AVB insertion depths from field eddy current data. This-relatively large uncertainty it, applicable only to low peaking conditions .;
where the AVB uncertainties can contribute to small peaking factors. The i definition of "no flow peaking" was increased to encompass the small ,
peaking effects from AVB insertion uncertainties. For the AVB patterns leading to significant peaking factors, AVBs were positioned within uncertainties to maximize the peaking' factor. For these configurations, variations of AVB insertion within these uncertainties are expected to reduce the peaking factor compared to the final values of Table 8-6 and Figure 8-7.
- 2. Including uncertainties directed toward conservatively decreasing the peaking factor for the North Anna tube R9C51, the final R9C51 peaking factor is [ Ja,c relative to a no flow peaking condition such as with uniform AVB insertion depths.
8.8 Peaking Factors for Specific Tubes Peaking factors for Sequoyah Unit 1 & Unit 2 were determined using the methodology described above. Table 8-7 and Table 8-8 summarize the results of peaking factors for Unit 1 and Unit 2. The AVB positions on each insertion pattern of Figure 8-7 should be carefully noted. Where the AVBs are shown at '
the top of the test tube, (configurations it, 4n for example),.the AVBs at least partially bloc' the flow past the test tube and low flow peaking factors t
are typically cbtained. Where the AVBs are shown at the centerline of the tube row above the test tube, the flow past the test tube is not restricted and significant flow peaking can be obtained.
9049M:1E-052689 8-15 I
+.
" WESTINGHOUSE PROPRIETARY CLASS 3-In' applying the' methodology to.Sequoyah 1 and 2, maps of the AVB insertion
'd:pths shown in Figures 6-2 through 6-9 were first reviewed. The second step waItoidentifythoseuniqueandmeaningfulconfigurationsofAVBinsertion
.dtpths in locality. In'doing so, maximum allowable; flow. peaking ~ factors were also reviewed column by column for rows 8 through 12. Based on the Sequoyah 1 and 2 tube vibration analysis, flow peaking factors greater than [ la c for row 8 tubes and (- -]a,c;for row 9 tubes would be required for tube fatigue to be a concern.
After conservative estimates of peaking factors were made for specific tubes, those.having peaking factors near,the maximum allowable value were identified and AVB insertion depth accuracy was reviewed for the tube involved and its neighboring tubes. If needed, stability velocity for the tube with the identified configuration'n of the AVB insertion depth ~was determined.using the
. Westinghouse R&D Cantilever, Air'Model. The peaking factor was then calculated using the stability velocity.
Determinations of peaking factors for identified tubes shown in Table 8-7 and Table 8-8 are described in detail below. Table 8-7 and Table 8-8 are broken into small tables for ease in following the description.
)
O i
9049M:1E-052689 8-16 l
+k WESTINGHOUSE PROPRIETARY CLASS 3 8.8.1' Steam Generator 1 of Sequoyah 1 The following' table gives the peaking factors for tubes with' unique configurations of.AVB insertion' depths.
Steam Type of AVB Peaking' G:,nerator Row No. Column No. ' Insertion Depth Factor a,c:
1 .7. 33, 34 <5d 59 ~5c 60 ~5b 8 9, 10 <5b 34, 59 ~4n 9 4 4b,lj 43, 44 <5h, ~5a 10 44 ~5d _ _
For R7C33 and R7C34 tubes, type [ la,e was applicable and a peaking factor of [ 3a c resulted. Tubes R7C59 and R7C60 belonged to types [ and Ja,c, respectively, and thus a peaking factor of.[ la,c was obtained for them.' Type [ 3a,c was used to represented for both R8C9 and R8C10 tubes and a peaking factor of [ la,c was'given. Type [ la,c was a
. good choice for tubes R8C34 and R8C59 and thus there were no flow peakings.
For tube R9C4, [
la,c was selected and a peaking factor of [ 3a,c resulted. Type [ la,c was selected for both R9C43 and R9C44 tubes and a peaking factor of [ ]a,c was obtained. For R10C44 tube, [
ja.c ,
L !
9049M:1E-052689 8-17
WESTINGHOUSE PROPRIETARY CLASS 3
]
8.8.2 Steam Generator 2 of Sequoyah 1 1
Tubes with unique AVB configurations are listed below together with their peaking factors.
Steam Type of AVB Peaking G:nerator Row No. Column No. Insertion Depth Factor !
a,c I
2 7 34, 35 ~5a 60 <4e, ~5a 89 4a 8 35 ~4f 46, 51 4b _ _ . .
For tubes R7C34 and R7C35, type [ Ja,e was selected to provide a peaking factor of [ Ja,c, Type [ ]a,c was used for R7C60 tube and a peaking factor of [ la,c resulted. Tube R7C89 belonged to type [ Ja,c and a peaking factor of [ Ja.c was obtained. Type [ Ja,c was a conservative choice for RBC35 tube and a peaking factor of [ Ja,c was given. Both R8C46 and R8C51 tubes were considered to be type [ ]a,c and a peaking factor of [ Ja,c resulted.
8.B.3 Steam Generator 3 of Secuoyah 1 The following table presents results of peaking factors. )
Steam Type of AVB Peaking Generator Row No. Column No. Insertion Depth Factor
_.a , e 3 8 34, 35 ~5a, Sh 60 ~5h i 61 ~5i _ _
1 9049M.1 E-052689 8-18
kl WESTINGHOUSE PROPRIETARY CLASS 3 T Type [ la,c was selected for tubes R7C34 and R7C35 and a peaking factor of
-[- .]a,c resulted._ Type'[L Ja,c was considere'd for RBC60 tube and a
'pGating factor,of [~ _3a,c was obtained. ' Tube R8C61 belonged to type J
[ . Ja,c and a peaking' factor of'[' la,clwas given.
8.8!4 Steam Generator 4 of Sequoyah 1-The following table presents results of' peaking factors determined for row 7 and row 8' tubes.
Steam Type of AVB Peaking Generator-- Row No. Column No.- Insertion. Depth Factor l
.a , c -
4 7 30, 31 ~5d 59,.60 ~5i 68, 69 ~5a 8 24, ~5a
'34, 35 ~5d !
59 4f-82, 83 Sa 9 9, 10, 11 ~5a 91 4b, <4a Type [ la,c was selected for tubes R7C30 and R7C31 and a peaking factor of
[ ]a,c resulted. Type [ la,c was considered for both R7C59 and R7C60 tubes and a peaking. factor of [ Ja,c was obtained. Both R7C68 and R7C69
-tubes were represented by type.[ Ja,c and a peaking factor of'[ Ja,c was obtained.
For' tubes R8C24 and R8C25, type [ la,e was selected to yield a peaking factor of [ la,c Both R8C34 and R8C35 were considered to be type l [ -Ja,c and a peaking factor of [ Ja,c resulted. Tube R8C59 belonged to type [ l .ca and a peaking factor of [ ]a,c was obtained. Type [ la,c 9049M:1E-052689 8-19 h
)
i I
WESTINGHOUSE PROPRIETARY CLASS 3.-
l 1
- was a good choice f'or tubes R8C82 and R8C83 and a peaking factor of
[. Ja,c was given. For' tubes R9C9, R9C10 and R9C11, type [ la,c was )
sblected to yield a peaking' factor of [ Ja,c; this was a conservative I
- 9. choice. Type-[ la,c was a conservative selection to result in a peaking factor of [~ la,c for tube R9C91.
'{
8.8.5' Steam Generator 1 of Sequoyah 2 i
~The following table presents results of peaking factors for tubes with unique configurations of.AVB insertion depths.
Steam Type.of AVB Peaking Generator Row No. Column No. Insertion Depth Factor
._ _a c 1 9 56, 40 Sa 35 4f 14 4b 10 60 ~1m _ ._
For R9C56 and R9C40 tubes, type [ la,c was applicable and thus a peaking factor of [ Ja.c resulted. Type [ Ja,c was a good choice for R9C35 tube and a peaking factor of [ la,c was obtained. For R9C14 tube, the
[
]s,c was applicable to give a peaking factor of [ Ja,c . As for R10C60 tube, type
[ Ja,e was selected and a peaking factor of [ Ja,c was obtained.
l l
9049M.iE-052689 8-20
WESTINGHOUSE PROPRIETARY CLASS 3 1
8.8.6 Steam Generator 2 of Sequoyah 2 Tubes with unique AVB configurations are listed below together with their p;aking factors.
Steam Type of AVB Peaking G:nerator Row No. Column No. Insertion Depth Factor
_a , c 2 8 60 ~4s 40 6a 35 ~1g, ~4d 9 35 ~1q, ~1z _ _,
For R8C60 tube, type [ la,c was selected and thus a peaking factor of
[ la,c resulted. Type [ ]a,c was applicable for R8C40 tube and a peaking factor of [ la,c was obtained. For R8C35 tube, type [ Ja,c was considered and a peaking factor of [ la,c was given. As for R9C35 tube, type [ la,c was applicable and a peaking factor of [ la.c was obtained.
8.8.7 Steam Generator 3 of Seauoyah 2 The following table presents peaking factors for R8C91 ar.d R8C61 tubes.
Steam Type of AVB Peaking Generator Row No. Column No. Insertion Depth Factor
_ a,c 3 8 91 ~4b 61 ~5a _ _
l 9049M:1E-052BB9 8-21
WESTINGHOUSE PROPRIETARY CLASS 3
< Fo'r R8C91 tube, the [-
Ja.c was applicable and a peaking factor of [ la,c was obtained. As for R8C61 tube, type
[ .]a,c was selected to yield a peaking factor of [ la,c ,
-The remaining tubes with unique AVB configurations are listed below with their peaking factors.
~ Steam Type of'AVB . Peaking G2rierator . . Row No. Column No. Insertion Depth Factor
. a,c 3 8 60 ~5g 35 ~4f, ~4s 10 <5e ,
9 -<5a 9 60 ~1t _ _
Type [ ]a,c was considered for R8c60 tube and a peaking factor of
[ Ja,c resulted.. For R8C35 tube, type [ la,c was used to yield a peaking factor of [ ]"'C Tubes R8C10 and R8C9 belonged to types [ and la,c, respectively, they have a peaking factor of [ ]C Type
[ ]a,c was a good choice for tube R9C60 and a peaking factor of
[ ]a,c resulted.
l
\ l i
L i
9049M:1 E-060589 8-22
.j i
n.
WESTINGHOUSE PROPRIETARY CLASS 3
.8.8.8 Steam Generator 4 of Sequoyah 2 Tubes having unique configuration of. AVB insertion depths are tabulated with their peaking factors.
Steam Type of AVE Peaking Gtnerator Row No.. Column No. Insertion Depth Factor
_ _a c q
4 8 80' ~5e 79 ~5b 60 ~4s, <4v 35 ~5g 34 <5g 9 60 ~1m 35 <8d , _
b Types [ la,c were applicable for R8C80 and R8C79 tubes, respectively and thus a peaking factor of [- Ja,c resulted. For RBC60 tube, type
[ Ja,c was a conservative choice tube and a peaking factor of [ Ja,c was obtained. Type [ Ja,c was considered for both R8C35 and R8C34 tubes and a peaking factor of ( Ja c was then used. Type [ Ja,c was selected to yield a peaking factor of ( Ja,c for R9C60 tube. Type
[ la,c was selected for R9C35 tube and a peaking factor of [
Ja,c resulted.
~
\
l l
9049M:1E-052689 8-23 I
'A . WESTINGHOUSE PROPRIETARY CLASS 3 Table 8 Stability Peaking Factor Due to Local Velocity Perturbation Scaling Factors for Steam / Water Air Velocity. Void Stability Peaking Fraction Density = Velocity Damping Peaking Factor, ' Scaling, Scaling, Scaling, Scaling, Factor, Fa Fv Fd Fv .Fdp Fs
, __a,c l
NOTE: 1. Stability peaking factor for steam / water mixture is calculated as follows:
,_ a,c
- 2. Damping scaling factor is calculated using modal effective void fraction of [ 3a,c for R9C51 tube.
i 9049M;1E-052089 8-24 !
i WESTINGHOUSE PROPRIETARY CLASS 3 I J
i Table 8-2 Comparison of Air and Steam-water Peaking Factor Ratios Air Air Steam Steam Peaking Peaking Peaking Peaking Factor Ratio Factor Ratio
_ _aC s 1
(
9049M.1E-052689 8-25
_ _ - -- __A
.w.
WESTINGHOUSE PROPR!ETARY CLASS 3 "
..J.l ,
t Table 8 -
Effect of Local Variation of AVB Insertion l
l 1
-A to B AVB Peaking Peaking Ratio
. Type A Type B Variation Factor A Factor B~ (B/A)
-- a ,c
' a,C 9049M:1E-052689 8-26 i
i WESTINGHOUSE PROPRIETARY CLASS 3 Table 8-4 i l
Uncertainties in Test Data and Extrapolation Source of Uncertainty Type Magnitude, %
a,c
- 1. Velocity measurement
- 2. Test repeatability
- 3. Cantilever vs U-tube
- 4. Air vs steam-water mixture
- 5. Field AVB configuration
- This is not an uncertainty associated with the test data.
It results from the inaccuracy in determining the true AVB position in the field using eddy current data.
1 WESTINGHOUSE PROPRIETARY CLASS 3 ,
1 Table 8-5 ;J 1
' Extrapolation of Test Results to Steam Generator Conditions Peaking Factor ]
Referenced to Test Data with )
Configuration- Data Uncertainties Configuration 2a
._ a , c la lb .
lg I 1m 1q lt Iz- l 2a j 4a 4b Ad 4e-4n 4s 4v Sa Sb Sc 5d Sg Sh Si 6a 8d 9049M 1 E-052689 " 8*28 J
WESTINGHOUSE PROPRIETARY: CLASS 3
' Table.8-6 Final Peaking Factors Configuration Peaking Factor
./
'la lb '.
ig im Iq' It Iz.
2a 4a
. 4b 4d
.: 4e l
l 4n y
4s l.
4v Sa Sb l
Sc 5d' 5g Sh Si-6a 8d _ _,
9049M:1E-052689. 8-29
- ,,. o i
r ,
l WESTINGHOUSE PROPRIETARY CLASS 3 l
Table 8-7 Stability Velocity Peaking Factors for Specific Tubes Sequoyah 1 Steam .
Type of AVB Peaking. 1
- Generator Row No. ' Column No. Insertion Depth- Factor J,c
- 1. 7 33, 34 . '<5d i 59 ~5c l 60 ~5b 8 9, 10 <5b 34, 59 ~4n 9 '4 4b 43,44 Sa 10- 44 ~5d
'All of the Remaining-2 7 34, 35 ~5a 60 <4e, ~5a 89 ~4a 8 35 ~4f 46, 51 4b All of the Remaining 3 8 34, 35 ~5a, Sh 60 ~5h 61 ~5i All of the Remaining 4 7 30, 31 ~5d 59, 60 ~51 68, 69 ~5a 8 24, 25 ~5a 34, 35 ~5d 59 ' 4f 82, 83 Sa 9 9, 10, 11 <5a 91 4b, <4a All of the Remaining - -
9049M.1E-052689 8-30
.a
.e , t, elk ]
1 ,
j t.
i WESTINGHOUSE PROPRIETARY CLASS 3
.~.
Table 8-8 Stability Velocity Peaking Factors'for Specific Tubes Sequoyah 2 I l
. Steam , Type of AVB Peaking
. G6nerator Row No. Column No. Insertion Depth Factor
~
1 9 56 6a~ .
40 6a 35 4f 14 4b 10' 60 ' . ~1m All of the Remaining 2 8 60 .~4 s 40 6a 35 ~19, ~4d 9 35 ~1q, ~1z All of the Remaining 3 8 91 4b 61 ~5a 60 ~5g 35 ~4f, ~4s 10 <5e 9 <5a
~1t 9 .60 All of the Remaining 4' 8 80 ~5c 79 ~5b 60 .~4s, <4v 35 ~5g 34 <5g 9 60 .~1m 35 <Bd-All of the Remaining - -
l b
8-31 9049M.1E-052689
WESTINGHOUSE PROPRIETARY CLASS 3 i
I i
n o , o e o o u l
" . OOOO O 'O 0 O 000 5' O 'O 'OLO OOLO0 00 0 0 O l0 l0 0 0 0 OL e 'O LG ,$;G.O ,
e ,O '0 00000 00000
- .0,0 0lO 000000000 .
85' ff 84l 53' 52 51 50' 48' 48 47' 4E 45' 44 i , , , l i e r^itse runs O*V' VISIBLE @ avsINVIS!BLE
$ PLUGGED f
Figure 8-1. Original North Anna AVB Configuration (Configuration Ib) 9049M:1E-052689 8-32
2 I
WESTINGHOUSE PROPRIETARY CLASS 3
- l. .
I
! i I
i F.n Ng
,,... ...<. w -
r
~~ - - l~. , p.. . . . . A. . . r. ' .
s
. i .; 7,, ,,
2.i F.b*4 , . W4+- -
(('^ih.5';'l Uh4 f;a l'
l.
l.
l.
1 l.
l.
l.
l.
l.
l Figure 8-2. Schematic of Staggered AVBs 9049M 1E-052089 8-33
WESTINGHOUSE PROPRIETARY CLASS 3 l
a, c l
1 l
~
- t. ,
l Figure 8-3. AVB " Pair" in ECT Trace 1
l 9049M:1E-052SB9 8-34 1 1
\
\
' WEST!NGHOUSE PROPRIETARY CLASS 3 j i
l o o G u e I
=,'O,0 22
' m m 00000 00 00 0 0
" 'O O G O O O Ge DO OOO
" O D D D DDDDpO,000 4G ,@ OOOOOOOOOO
- O@GO000000000 column 56 ,55 ,54 53 ;52 ,51 ,50 ,49 48 ,47 ,46 ,45 44
$ Ngged Tube g FaledTube Embers h cire!es in cebmn rance 48 55 reptsort rsacahte' AVE traersocilon s*gna:s.
I based on artcaltwiew of the ECT traces Open circle h this range means te gaa k availatie.
i l
Figure 8-4. North Anna 1, Steam Generator C, AVB Positions Critical Review "AVB Visible" Calls 9049M:1E-052689 8-35 h
i WESTINGHOUSE PROPRIETARY CLASS 3 l
m,. . . n ,,_ _
l BRB c I._.,M ,M -_ ER E -M ,E .... ,..
B ER .... . . . . ..
E H !!...lH G..
R $... l...Ed..k..
l . ...
BR B..
E H il.,.lN...
I G E W...
... .., ... . . . m..
E...
W..,
R E. .. , k. R R . ..
B R.ME
- i. . .
IR Bag:. m._ .
as>A.. 4 s g EE M,FK.
a._ . m ... :. BEE.,
H,EG....
z_ ..
HE 4.. =.
CSS C84 CS3 C82 CE, CSC Cat C44
- [
U theenmo.
Y *Pumnuer WuurWB
% .e Sid. W Sids' Figure 8-5. North Anna 1, Steam Generator C, l
R9C51 AVB [' Ja,c Matrix 9049M.1E-052689 8-36
WESTINGHOUSE PROPRIETARY CLASS 3
< u m u u o r ,, ,, .
n 00 d 0 00 ,0 0 0 0 0 0 u' O'O 0 0 O0 0 0 '0 0 0 0 a OO ,0 0 0 0:0 0 0. 0. 0..
- OOO .O090000000
.000000.0000000 M S3 M S3 52 51 ! 2 49 '48 47 46 di da l l l i i
j Figure 8-6. North Anna R9C51 AVB Final [ la,c Positions (Configuration la) 9049M 1E-052689 8-37
WESTINGHOUSE PROPRIETARY CLASS 3 1
TYPE OF AVB PEAKING TYPE OF AVB PEAKING TYPEOFAVB PEAKING .
FACTOR WSERTON FACTOR INSERTION FACTOR- INSERTION
- e, w, c -- m, s, c s, %.c-h
[-
L Figure 8-7. Final Peaking Factors for Sequoyah 1 and 2 i
9049M;1E-052689 8-38
WESTINGHOUSE PROPRIETARY CLASS 3 9.0 STRUCTURAL AND TUBE VIBRATION ASSESSMENTS 9.1 Tube Mean Stress This section summarizes the analysis to determine stresses in a dented but undeformed tube at 100% power. Loads imposed on the tube correspond to steady-state pressure, differential thermal expansion between the tube and the support plate, and a thru-wall thermal. gradient. The analysis assumes the tube to be [ ]a,c at cold shutdown.
A summary of the temperature and pressure parameters at 100% power in the-vicinity of the top support plate are provided in Table 9-1. The tube temperature corresponds to the average of the primary-side water temperature and the plate temperature. The resulting tube / plate radial interference is
[ ]a,c ,
Stresses due to differential pressure and interference loads are calculated using finite element analysis with the model shown in Figure 9-1. The model prescribes [
3a,c Two reference cases were run using the finite element model, the first for a primary-to-secondary side pressure gradient of 1000 psi, and the second for a
[ ]a,c inch radial interference between the tube and plate. The pressure case incorporates the axial load on the tube by applying a pressure loading along the top face of the model. Plots showing the distribution of stress for the tube outer surface for the two reference cases are provided in Figures 9-2 and 9-3. Thermal bending stresses due to the thru-wall thermal gradient are calculated to be 7.6 ksi using conventional analysis techniques.
The combined stress distribution along the tube length, in Figure 9-4, was obtained by combining the thermal bending stresses and the reference solutions with appropriate multipliers based on 100% power operating parameters.
9049M.1E-052689 9-1
WESTINGHOUSE PROPRIETARY CLASS 3
.Tha maximum axial tensile stress is 20.1 ksi and occurs approximately 0.134 inch above the top surface of the support plate. Adding, for conservatism, the surface stress due to pressure, 0.9 ksi, gives an applied mean stress of ;
21.0 ksi.- In addition to the applied stress, residual stresses exist in the '
tube as a result of the inanufacturing process. For mill annealed tubes with subt equent straightening and polishin.), residual stresses are compressive at the tube surface, but 5-10 mils below the surface, the stress levels change to ]
10-15 ksi tensile. Combining the applied and residual stresses results in a {
cumulative mean stress of approximately 36.0 ksi, assuming tube denting without deformation.
If a tube is dented with deformation, the mean stress is limited by tube yielding. For the case of dented tubes with deformation, the maximum f.
effect of mean stress was incorporated by using emax = oy in determining stability ratios and fatigue usage.
9.2 Stability Ratio Distribution Lased Upon ATH0S I An assessment of the potential for tubes to experience fluid elastic instability in the U-bend region has been performed for each of the tubes in rows eight through twelve. This analysis utilizes FASTVIB, a Westinghouse proprietary finite element based computer code, and PLOTVIB, a post processor to FASTVIB. These codes predict the individual responses of an entire row of steam generator tubing exposed to a location dependent fluid velocity and density profile. The program calculates tube natural frequencies and mode shapes using a linear finite element model of the tube. lne fluid elastic stability ratio Ue/Uc (the ratio of the effective velocity to the critical velocity) and the vibration amplitudes caused by turbulence are calculated for a given velocity / density / void fraction profile and tube support condition. The velocity, density and void fraction distributions are determined using the ATHOS computer code as described in Section 7.3. The WECAN generated mass and stiffness matrices used to represent the tube are also input to the code.
l (WECAN is also a Westinghouse proprietary computer code.) Additional input to FASTVIB/PLOTVIB consists of tube support conditions, fluid elastic stability l
constant, turbulence constants, and location dependent flow peaking factors.
9049M.1E-Os2689 9-2
.)
WESTINGHOUSE PROPRIETARY CLASS 3 This process was performed for the Sequoyah Unit 1 & Unit 2 steam generator tubes and also for the North Anna Row 9 Column 51 tube (R9C51) using similarly appropriate ATH0S models. Ratios of the Sequoyah Unit 1 & Unit 2 results to those for North Anna Unit 1 R9C51 were generated to produce a quantity that could be used to provide an initial assessment of the Sequoyah Unit 1 & Unit 2 i tubes relative to the ruptured tube at North Anna Unit 1. l Figures 9-5 and 9-6 contains the results of this process for each of the rows under investigation for Units 1 and 2. The relative ratios are obtained using the following conditions for the Sequoyah units and North Anna Unit 1: 1
- 1) Tube is fixed at the top tube support plate,
- 2) Void fraction dependent damping,
- 3) No AVB supports are active,
- 4) Location dependent flow peaking factors.
lt is to be noted that the stability ratios plotted are composites of all steam generators. That is, any peaking effect for a given tube location on the plot represents the maximum value of the peaking factors in all steam generators at that location.
A horizontal line is drawn at the relative stability ratio value of 0.90.
This identifies the point where a ten percent reduction in stability ratio exists relative to North Anna R9CSI. (See Section 4.1 for a discussion of the stability ratio reduction criteria.) All the tubes with ratios above this line would be considered to have stability ratios larger than ninety percent of North Anna R9C51.
These figures indicate that most tubes in Rows 8 thru 11 of the Sequoyah Unit 1 & Unit 2 steam generators lie below the 90% line. Note that all of Row 12 is supported and therefore the relative stability ratios presented in these figures for this row can be disregarded. <
9049E1E-052689 9-3
+ ,
WESTINGHOUSE PROPRIETARY' CLASS 3
-'All unsupported tubes, with the exception of'R10C44 SG 1 in Unit'1,
. R10C60 SG 1 and R9C60 SG 4 in Unit 2 have-RSR values (including flow peaking).~
- Icss.than 0.90.
4
- 9.3 Stress Ratio Distribution with Peaking' Factor e
An evaluation was performed'to determine the ratio of the Sequoyah Unit--I &
Unit 2-tube stress over the North Anna R9C51 tube stress. This ratio is determined using relative stability ratios discussed in the previous section,:
relative flow peaking factors (Table 8-7 factors divided by-[. .]a,c) and bending moment factors. Sections 4.2 and'4.3 contain additional information.
and describe the calculational procedure used to obtain'the results'pr'esented-
~
-in this section. The results presented below are based upon the following conditions:
1 1) Tube is' fixed at'the top tube support plate,
- 2) . Damping is void fraction dependent,
~
- 3) . Tubes have no AVB support,
- 4) 10%_ criteria with frequency effects, l
- 5) Tubes are assumed to be dented or undented (fixed at top support plate but without tube deformation).
A tube can be considered acceptable if the stress ratio is less than 1.0 when calculated using the procedure described in Sections 4.2 and 4.3 and including the conditions listed above and subject to confirmation of fatigue usage acceptability. Conformance to these requirements implies that the stress acting.on a given tube is expected to be insufficient to produce a fatigue
' event in a manner similar to the rupture that occurred in the R9C51 tube at l North Anna Unit 1.
p 9049M.1E-Os26B9 9-4
WESTI'NGHOUSE' PROPRIETARY CLASS 3 9.3.1:.Sequoyah Unit 1 Tubing .
I Figures.9-7 and 9-8 show the results'of the stress ratio calculations for the j
-d nted:(magnetite. clamping with tube deformation) and undented (magnetite .!
clamping without deformation) cases respectively for each of the Sequoyah- !
, Unit 1 tubes in Rows 8 through 12.
As-in the case of stability ratios, the plotted stress ratios represent a ,
- " composite set for.all four' steam generators in this unit. The critical tube !
summary list in Table 9-2 is prepared from these figures in conjunction with
-flow peaking and AVB support condition.information for individual tubes-in !
sach steam generator. l
' As discussed in Section 6, for the critical . tubes listed in Table 9-2, it is l probable that. support plate corrosion may.not have progressed to the point of !
rigidly clamping the tube (no denting, case in Table 9-2). However, in the absence of, the knowledge of actual support plate crevice conditions, the tubes i are conservatively assumed to be clamped in this evaluation. On that basis,
- examination of Table 9-2 shows that all the unsupported tubes in all steam -l gtnerators do not exceed the limiting stress ratio criteria for tubes in the undented configuration. :
l All Sequoyah Unit 1 tubes, under the currently assumed conditions of crevice i clamping without distortion, meet the stress ratio criteria. If denting is considered, the most limiting tube is in SG 1 at location R10C44 and has a stress ratio greater than 1.00. This tube is a candidate for removal from service using sentinel plugs as tube stabilizers to restrain motion.
9.3.2 Sequoyah Unit 2 Tubing Results of.the stress ratio calculations (composite of all three steam generators) for Sequoyah Unit 2 tubes in Rows 8 through 12 are shown in l
Figures 9-9 and 9-10, for the denting and no-denting case, respectively. A summary listing significant results for the critical tubes in each of the three steam generators is given in Table 9-3. j l-9049M:1E-052689 9-5 i
_.j
9
@4 f 1 - ,
'O WESTINGHOUSE PROPRIETARY CLASS 3 examination of' Figure 9-9 and Table 9-3 shows that there are two unsupported tubes that exc'eeds'a stress ratio of 1.0, when assumed clamped'at the' TSP.
These tubes are located in SG 1 R10C60 and SG 4 R9C60 and'are candidates for removal from service using sentinel plugs or tube stabilizers to restrain tube motion..-All remaining Sequoyah Unit 2 tub'es, under the currently assumed dinted conditions meet the stress rstio criteria.
9.3.3 Maximum Allowable Relative Flow Peaking
.An evaluation has also been performed to determine the required relative flow peaking that will produce a stress ratio not greater than 1.0. Figure 9-11 contains the results of-this process for all the tubes in' Rows 8 through 12.
The figure was generated using the conditions outlines previously with the additional constraint that the tubes are dented. Note that this figure reads-opposite of the previous-figures, i.e., the top curve in the figure
-corresponds to Row 8 and the bottom curve corresponds to Row 12. Maximum Allowable Relative Flow Peaking is the required relative flow peaking-(0.68-corresponds to no flow perking) that, if used on the given tubo, will produce-
'a stress ratio (with denting) not to exceed 1.0.
This curve can be used to identify the relative flow peaking required before preventive action would be recommended and, when used in conjunction with the actual _ flow peaking associated with each tube, to determine the margin
'present. This has also been performed in Table 9-2 and Table 9-3. The column with heading " Max Allow Flow Peak" identifies the' relative flow peaking factor that would be permitted, on a tube by tube basis, before the stress ratio criteria would be exceeded. As can be observed in the table and figure, the inner row tubes have larger values of allowable relative flow peaking when i compared to the outer rows.
-9.4 Cumulative Fatigue Usage
- All tubes that are unsupported and have a stress ratio < 1.0 have a maximum I
j' stress amplitude that is < 4.0 ksi (from 9.5 ksi) since a 10% reduction in j the stability ratio for the North Anna Row 9 Column 51 tube was the criteria l
9049M:1E-052689 9-6
)
l WESTINGHOUSE PROPRIETARY CLASS 3 basis. The stability ratios for the Sequryah Unit I and 2 tubing are based on the current operating parameters and with future operation on the same basis, the tubes are not expected to rupture as a result of fatigue if 1) they meet the stress ratio criteria of 51 0 and 2) their current and future fatigue ;
usage will total less than 1.0.
l Based on the above analyses, most Sequoyah Unit 1 and Unit 2 tubes meet the .
ralative stress ratio criteria. Preventative action nas been recommended for those tubes that do not meet the stress ratio criteria. Tables 9-2 and 9-3 provide a summary of the combined relative stability ratios and the stress ratios for the more salient unsupported tubes in Rows 8 through 12 for Units 1 and 2 respectively.
9.4.1 Unit 1 Cumulative Fatigue Usage Acceptability of the Sequoyah Unit 1 tubing for fatigue is accomplished by >
demonstrating the acceptability of the tube with the highest stress ratio, that does not exceed the 1.0 stress ratio, 0.50 in R8C59 SG 4. Assuming the tubes have been dented since the first cycle and continue to operate under current conditions, the total usage including the remaining term of the operating license would be 0.05.
9.4.2 Unit 2 Cumulative Fatigue Usage Using a similar approach, the maximum calculated total usage is 0.83 for the highest stressed, unplugged, unstabilized and non-supported Unit 2 tubing at location SG 1 R9C35 with a stress ratio (with denting) of 0.95. This calculation assumes that the tube became dented during the first cycle of operation and the unit will continue to be operated at the current conditions. !
i l
9049M1E-052689 9-7 i
. _ - _ - _ _ _ _ _ _ _ - - ._ _ U
' 1:.
WESTINGHOUSE PROPRIETARY CLASS 3
. Table 9-1 100%: Power Operating Parameters - Sequoyah Unit 1 &' Unit 2 Bounding. Values'for.Hean' Stress Calculation.
Primary' Pressure = 2250 psia Secondary Pressure ~= 890 psia Pressure Gradient = 1360 psi Primary Side Temperature * = 578'F Secondary Side Temperature = 531'F Tube Temperature = 554*F s,
W Average of Thot = 611'F and Tcold = 545'F.
1 l
9049M:1E-052689 9-8
3 I '
WESTINGHOUSE PROPRIETARY CLASS 3-0' . .
TABLE 9-2 SEQUOYAH UNIT 1:-LTUBES WITH SIGNIFICANT RSR's OR STRESS RATIOS S/G Row- Col. ' Flow Max Allow RSR* Stress Ratio No Peak Flow Peak FPEAK w Dent w/o Dent-
~ #
0.53 0.08 0.07
'l 7 59- 0.76 60 0.76 0.52 0.07 0.07 33 0.82 0.55 0.10 0.09 34 O.82 . 0.56 0.10 0.09 8 34 0.68 0.55 0.08 0.07 1 9 0.76 0.60 0.13 0.12 59 0.68 0.57- 0.09 0.08
-10 1.76 0 0.60 0.13- 0.11 9- 4 0.68 0.62 0.12 0.11 44' O.76 0.77- 0.43 0.38 43 0.76 0.77 0.43 0.39 10 44 0.82 0.93 1.03 0.92
- 2. 7. 34 0.76' O.52 0.07 0.06 89 0.79 0.53 0.08 0.07 60 0.95 0.65 0.25 0.22
.35 0.76 0.52 0.07 0.07 8 35 0.93 0.76- 0.48 0.42 51 0.68 0.56 0.09 0.08 46 0.68 0.56 '0.09 0.08 3 8~ 60 0.68 0.56 0.09 0.08 35 0.76 0.62 0.16 0.14 61 0 . 18 0 0.65 0.20 0.18 34 0.76 0 . 16 2 0.15 0.13-4 7 69 0.76 0.52 0.08 0.07 59 0.80 0.55 0.10 0.09 68 0.76 0.52 0.07 0.07 31 0.82 0.54 0.09 0.08 60 0.80 0.55 0.10 0.09 30 0.82 0.54 0.09 0.08 l
8 24 0.76 0.61 0.15 0.13 1 35 0.82 0.67 0.24 0.21 i 83 0.76 0.60 0.13 0.11 25 0.76 0.62 0.15 0.14 59 0.93 0.76 0.50 0.44 82 0.76 0.60 0.13 0.11 I' 0.20 34 0.82 0.66 0.23 9 11 0.76 0.73 0.30 0.27 10 0.76 0.72 0.30 0.27 9 0.76 0.73 0.30 0.27 91 0.79 0.72 0.28 0.25 4049M:1E-052689 9-9 j
WESTINGHOUSE PROPRIETARY CLASS 3 TABLE 9-3 SEQUOYAH UNIT 2 - TUBES WITH SIGNIFICANT RSR's OR STRESS RATIOS i
S/G- Row -Col. Flow Max Allow RSR* Stress Ratio
-Flow. Peak FPEAK w Dent w/o Dent ;
No Peak. ;
r- - a,c 1- 9 11-14 0.68 0.624 0.13 0.12
- 34. 0.68. 0.648 .0.16 0.15 ,
35 0.93 0.892' 0.95 0.85 40-56 0.68 0.666 -0.19 0.17 60 0.68 0.652 0.17 0.15 i 61 0.68 0.648 0.16 0.15 l 0.772 0.36 0.32 ;
J 10 51 0.68 '
52 0.68 0.773 0.36 0.32 60 1.02 1.137 > 2.00 > 2.00 l
2- 8' 35 0.98 0.800 0.64 0.57 1 60 0.96 0.783 0.57 0.51 9 35 0.68 0.652 0.17 0.15 83 0.68 0.626 0.13 0.12 1
l 3 8 9 0.76 0.599 0.13 0.12 10 0.76 0.598 0.13 0.11 35 0.96 0.783 0.57 0.51 j SO 0.76 0.620 0.16 0.14 )
61 0.76 0.615 0.15 0.13 1 9 2 0.68 0.437. 0.02 0.02 4 1
3 0.68 0.609 0.12 0.10 45 0.68 0.662 0.18 0.16 49 0.68 0.659 0.18 0.16 i
60 0.68 0.652 0.17 0.15 61 0.68 0.649 0.16 0.15 4 8 34 0.76 0.614 0.15 0.13 35 0.76 0.620 0.16 0.14 60 1.05 0.856 0.93 0.83 ,
79 0.76 0.572 0.10 0.09 i 80 0.76 0.579 0.11 0.10 9 35 0.85 0.815 0.58 0.52 )
60 1.02 0.978 1.92 1.72 1 1
l 9049M:t E-060589 9-10 l
WESTINGHOUSE PROPRIETARY CLASS 3 a,c 1
1 I
l Figure 9-1 Axisymmetric Tube Finite Element Model 9049M:1E-052689 9-11
-WESTINGHOUSE PROPRIETARY CLASS 3
-a,e e N 9
l
~l a
O b*
e gee-e e5oi o
mO"ag@
MO E ' w*8 $'1
$*585
- $$8 EgE B
E 8
M Figure 9-2 Dented Tube Stress Distributions Pressure Load on Tube I
9049M:1E-052689 9-12 b_______ _ _ . .
WESTINGHOUSE PROPRIETARY CLASS 3 ,
a,c 1
r h i
ao W
w 585 I
59ls E9 =
g ms mN4E
- E$$
$e1 o
W 5
0 J
Figure 9-3 Dented Tube Stress Distributions Interference Load on Tube 9049M:1E-052689 9-13
WESTINGHOUSE PROPRIETARY CLASS 3 a,c I
l i,
1 i
Figure 9-4 Dented Tube Stress Distributions i
Combined Stress Results l
l l 9049M;1E-052689 9-14
WESTINGHOUSE PROPRIETARY CLASS 3 a,c 1
Figure 9-5 Relative Stability Ratios Using MEVF Dependent Damping - Sequoyah Unit 1 (Composite of all Steam j Generators with Umbrella Flow Peaking) l 9049M;1E-052689 9-15
.j
- i WEST!NGHOUSE PROPRIETARY CLASS 3 :
1 I
1 a,c l
-1 I
l I
1 l
I Figure 9-6 Relative Stability Ratios Using MEVF Dependent f
Damping - Sequoyah Unit 2 (Composite of all Steam Generators with Umbrella Flow Peaking) r 9049M:1E-052689 9-16
l WESTINGHOUSE PROPRIETARY CLASS 3 a,c Figure 9-7 Stress Ratio vs. Column Number - Dented Condition - Sequoyah Unit 1 (Composite of all SGs with Umbrella Flow Peaking) 9049M:1E-052689 9-17
WESTINGHOUSE PROPRIETARY CLASS 3 a,c Figure 9-8 Stress Ratio vs. Column Number - Undented Condition - Sequoyah Unit 1 (Composite cf All SGs with Umbrella Flow Peaking) 9049M.1E-052689 9-18
WESTINGHOUSE PROPRIETARY CLASS 3 a,c Figure 9-9 Stress Ratio vs. Column Number - Dented Condition - Sequoyah Unit 2 (Composite of All SGs with Umbrella-Flow Peaking)
WESTINGHOUSE PROPRIETARY CLASS 3 a,c r.
Figure 9-10 Stress Ratio vs. Column Number - Undented Condition - Sequoyah Uni 2 (Composite of All SGs with Umbrella Flow Peaking) 9049M.1E-052689 9-20
i WESTINGHOUSE PROPRIETARY CLASS 3
.l f
I a,c {
i
- l i
i l
i j
4 i
L .
i t
Figure 9-11 Sequoyah Unit 1 & Unit 2 - Maximum Allowable Relative Flow Peak'ing 9049M;1E-052689 9-21 i