ML20154A655
| ML20154A655 | |
| Person / Time | |
|---|---|
| Site: | Beaver Valley |
| Issue date: | 04/30/1988 |
| From: | Connors H, Frick T, Hall J WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19302D586 | List: |
| References | |
| SG-88-03-013, SG-88-3-13, WCAP-11800, NUDOCS 8805160078 | |
| Download: ML20154A655 (145) | |
Text
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WESTINGHOUSE NON PROPRIETARY CLASS 3 .!
l WCAP-11800 SG 88 03 013 i
i BEAVER VALLEY UNIT 1 EVALUATION FOR TUBE VIBRATION INDUCED FATIGUE I APRIL, 1988 AUTHORS H.J. CONNORS H.M. HU '
T.M. FRICK H.Y. YANT J.M. HALL A.Y. LEE' -
G.W. HOPKINS P.J. PRABHU ,
J.L. HOUTMAN R.E. SM:TH T
APPROVED:
- h. ;
T.A. PITTERLE, MANAGER +
STEAM GENERATOR LNGINEERING ;
WORX PERFORMED UNDER SHOP ORDER DL8D-01304 t
WESTINGHOUSE ELECTRIC CORPORATION !
POWER SYSTEMS BUSINESS UNIT SERVICE TECHNOLOGY DIVISION !
P. O. BOX 3377 ;
PITTSBURGH,PA 15230 0355 i 8805160078 880504 i PDR ADOCK 05000334 !
Q DCD
l l
l l
ABSTRACT l
On July 15, 1987, a steam generator tube rupture event occurred at the North Anna Unit 1 plant. The cause of the tube rupture has been determined to be y 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 seperimposed alternating stress. The mean stress is the result of denting of the tube at the top tube support 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.
This report documents the evaluation of steam generator tubing at Beaver Valley Unit I for susceptibility to fatigue induced cracking of the type experienced at North Anna Unit 1. The evaluation utilizes operating conditions specific to Beaver Valley Unit 1 to account for the plant specific nature of the tube loading and response. The evaluation also includes reviews of eddy current data for Beaver Valley Unit I to establish AVB locations. This report provides background of the event which occurred at North Anna, a criteria for fatigue 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 accumulated fatigue usage. This evaluation leads to identification of tubes potentially susceptible to fatigue, l f
I P
M448t:49/031783 !
1 i
SUMMARY
OF ABBREVIATIONS ASME - American Society of Mechanical Engineers ATH0S - Analysis of the Thermal Hydraulics of Steam Generators AV8 - Anti-Vibration Bar AVT - All Volatile Treatment ECT - Eddy Current Test EPRI - Electric Power Research Institute FFT - Fast Fourier Transform FLOVIB - Flow Induced Vibrations MEVF - Modal Effective Void Fraction 00 - Outside Diameter 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 i
Cl44M:49/031788 3 II
TABLE OF CONTENTS
$ECTION EME 1.0 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . 1-1 2.0 Summary and Conclusions . . . . . . . . . . . . . . . . . . . 2-1 2.1 Sa c kg rou nd . . . . . . . . . . . . . . . . . . . . . . . . . . 2-1 2.2 Evaluation Criteria ........... . . . . . . . . . . 2-1 2.3 Con:1usion . . . . . . . . . . . . . . . . . . . . . . . . . . 24 3.0 Background . . . . . . . . . . . . . . . . . . . . . . . . . . 3-1 3.1 North Anna Unit 1 Tube Rupture Event . . . . . . . . . . . . . 31 3.2 Tube Examination Results . . . . . . . . . . . . . . . . . . . 32 3.3 Mechanism Assessment . . . . . . . . . . . . . . . . . . . . . 33 4.0 Criteria for Fatigue Assessment . . . . . . . . . . . . . . . 4-1 4.1 Stability Ratio Reduction Criteria . . . . . . . . . . . . . . 4-2 4.2 Local Flow Peaking Considerations .......... . . . . 4-7 4.3 Stress Ratio Considerations ................. 48 5.0 Supporting Test Data . . . . . . . . . . . . . . . . . . . . . 51 5.1 Stability Ratio Parameters . . . . . . . . . . . . . . . . . . 51 5.2 Tube Damping Data . . . . . . ,
,.,;c. . . . . . . . 5-6 5.3 Tube Vibration Amplitudes with ,
,AVB Support . . . 5-8 5.4 Tests to Determine the Effects on Fluidelastic Instability of Columnwise Variations in AVB Insertion Depths . . . . . . . . 5-10 5.5 References . . . . . . . . . . . . . . . . . . . . . . . . . . 5-14 6.0 Eddy Current Data and AVB Positions . . . . . . . . . . . . . 6-1 6.1 Tube Denting at Top Support Plate .............. 61 6.2 Tube Wall Thinning at the AVB Supports . . . . . . . . . . . . 6-1 6.3 Eddy Current Data for AVB Positions. . . . . . . . . . . . . . 61 6.4 AVB Insertion Depths . . . . . . . . . . . . . . . . . . . . . 62 0144M:49/C31? M 4}
.t TABLEOFCONTEWTS(CONTINUi.D) i SECTION PAGI 7.0 Thermal Hydraulic Analysis . . . . . . . . . . . . . . . . . . 71 7.1 Beaver Valley Unit 1 Steam Generator Operating Conditions .. 71 7.2 ATHOS An alys i s Model . . . . . . . . . . . . . . . . . . . . . 7-2 7.3 ATHOS Results ........................ 74 7.4 Relative Stability Ratio Over Operating History ....... 75 i
8.0 Peaking Factor Evaluation. . . . . . . . . . . . . . . . . . . 81 8.1 North Anna 1 Configuration . . . . . . . . . . . . . . . . . . 81 8.2 Test Measurement Uncertainties . . . . . . . . . . . . . . . . 85 8.3 Test Repeatability . . . . . . . . . . . . . . . . . . . . . . 86 8.4 Cantilever vs U-Tube . . . . . . .............. 86 8.5 Air vs Steam Water Hixture . . , , . . . . . . . . . . . . . . 88 8.6 AVB Insertion Depth Uncertainty. . . . . . . . . . . . . . . . 8 12 8.7 Overall Peaking Factor with Uncertainty. . . . . ...... 8 13 4 9.0 Structural and Tube Vibration Assessments .......... 91 4 9.1 Tube Mean Stress . . . . . . . . . . . . . . . . . . . . . . . 91 9.2 Stability Ratio Distributions Based Upon ATHOS . . . . . . . . 92 9.3 Stress Ratio Distribution with Peaking Factor ........ 93 9.4 Cumulative Fatigue Usage . . . . . . . . . . . . . . . . . . . 95 014484:49/C32464 5 gy
l LIST OF F18URES 1 ffd8f. fKil 3-1 Approximate Mapping of Fracture Surface of Tube R9C51 S/G 'C' C ol d L eg , No rth Ann a Un i t 1 . . . . . . . . . . . . . . . . . . . 35 1 3-2 Schematic Representation of Features Observed During TEM ( 1 Fractographic Examination of Fracture Surface of Tube R9C51, S/G
'C' Cold Leg, North Anna Unit 1 . . . . . . . . . . . . . . . . . 36 i 3-3 Calculated and Observed Lcak Rates Versus Time ......... 37 4-1 Vibration Displacement vs. Stability Ratio . . . . . . . . . . . 4-12 i
~ 4-2 Fatigue Strength of Inconel 600 in AVT Water at 600'F . . . . . . 413 ' s ! ! 4-3 Fatigue Curve for Inconel 600 in AVT Water Compariscn of Mean Stress Correction Models . . . . . . . . . . . . . . . . . . . . . . . . 4 14 ; 4-4 Modified Fatigue with 10% Reduction in Stability Ratio for Maximum Stress Condition ....................415 9 4-5 Modified F:tigue with 5% Reduction in Stability Ratio for Minimum Stress Condition .... , . . . . . . . . . . . . . . 4-16 5-1 Fluideiastic Instability Uncertainty Assessment . . . . . . . . . 5-11 i 52 Instability Constant - # . . . . . . . . . . . . . . . . . . . . 5-16 53 Instability Constants, A, Obtained for Curved Tubs from Wind Tunnel Tests on the 0.214 Scale U. Bend Model , . . . . . . . . . . . . . 5-17 54 Damping vs. Slip Void Frtction . . . . . . . . . . . . . . . . . 5 18 81444:49/031844 6 y
LISTOFFIGURES(Continued) FIGURE Edfai i 5-5 Overall View of Cantilever Tube Wind Tunnel Model . . . . . . . . 519 56 Top View of the Cantilever Tube Wind Tunnel Model . . . . . . . . 5 20 , 5-7 Fluidelastic Vibration Amplitude with Non-Uniform Gaps . . . . . 5-21 B 58 Typical Vibration Amplitude and Tube /AVB Impact Force Signals for l t Fluidelastic Vibration with Unequal Tube /AVB Gaps . . . . . . . . 5 22 59 Conceptual Design of the Apparatus for Determining the Effects on Fluidelastic Instability of Columnwise Variations in AVB Insertion Depths . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5 23 5 10 Overall View of Wind Tunnel Test Apparatus . . . . . . . . . . . 5 24 5-1) Side View of Wind Tunnel Apparatus with Cover Plates Removed to Show Simulated AVBS and Top Flow Screen . . . . . . . . . . . . . 5 25 5-12 AVB Configurations Tested . . . . . . . . . . . . . . . . . . . . 5 26 5 13 Typical Variation of RMS Vibration Amplitude with Flow Velocity i for Configuration 1 in Figure 5-12. . . . . . . . . . . . . . . . 5 27 6-1 AVB Insertion Depth Confirmation ................ 69 62 Beaver Valley 1 - Steam Generator ' A' - AVB Positions . . . . . . 6-10 63 Beaver Valley 1 Steam Generator 'B' - AVB Positions . . . . . . 611 64 Beaver Valley 1 - Steam Generator 'C' - AVB Positions . . . . . . 612 0144m 49/c32488 7 VI f
LISTOFFIGURES(Continued) ele 8E fMi 7-1 Plan View of ATH0S Cartesian Model for Beaver Valley Unit 1. . . 7-12 7-2 Elevation View of ATHOS Cartesian Model for Beaver Valley Unit 1. 7-13 7-3 Plan View of ATH0S Cartesian Model Indicating Tube Layout . . . . 7-14 7-4 Flow Pattern on Vertical Plane of Symetry . . . . . . . . . . . 7-15 75 Lateral Flow Pattern on Horizontal Plane in the U Bend Region . . 716 7-6 Lateral Flow Pattern on Top of Tubesheet. . . . . . . . . . . . . 7-17 7-7 Tube Gap Velocity and Density Distributions for Tube at R10/C3. . 7-18 78 Tube Gap Velocity and Density Distributions for Tube at 010/C20 . 719 7-9 Tube Gap Velocity and Density Distributions for Tube at R10/C40 . 7-20 l 7-10 Average Velocity and Density in the Plane of the U Bends Normal to Row 10 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7 21 7-11 Beaver Valley Unit 1 - Normalized Stability Ratio Based on High Power (>86%) Operation . . . . . . . . . . . . . . . . . . . 7 22 0141049/C317H 8 V]I l
LISTOFFIGURES(Continued) finBE PEil 8-1 Original North Anna AVB Configuration . . . . . . . . . . . . . . 8 25 8-2 Schematic of Staggered AVBs . . . . . . . . . . . . . . . . . . . . 8-26 83 AVB "Pair" in ECT Trace . . . . . . . . . . . . . . . . . . . . . 8 27 8-4 North Anna 1, Steam Generator C: AVB Positions Critical , Review "AVB Visible" Calls . . . . . . . . . . . . . . . . . . . 8 28 a,c 85 North Anna 1, Steam Generator C, R9C51 Matrix . . . . 8 29
- - a,e 86 North Anna R9C51 AVB Final - .
Positions .........830 L 8 31 8-7 Final Peaking Factors . . . . . . . . . . . . . . . . . . . . . . 91 Axisymmetric Tube Finite Element Model . . . . . . . . . . . . . 9-9 i 92 Dented Tube Stress Distributions Pressure Load on Tube . . . . . 9 10 9-3 Dented Tube Stress Distributions Interference Load on Tube . . . 9-11 94 Dented Tube Stress Distributions Combined Stress Results Beaver Vall ey Unit 1. . . . . . . . . . . . . . . . . . . . . . . 9-12 Relative Stability Ratio Using MEVF Dependent Damping . . . . . . 9 13 9-5 f g6 Stress Ratio Vs. Column Number Dented Condition . . . . . . . . . 9-14 l 9-7 Stress Ratio Vs. Column Number Undented Condtion. . . . . . . . . 915 c144a:u/enus.: )
LIST OF TA8LES IAEI fKil 4-1 Fatigue Usage per Year Resulting From Stability Ratic Reduction 4 11 5-1 Wind Tunnel Test on Cantilever Tube Model . . . . . . . . . . . . 513 5-2 Fluidelastic Instability Peaking Factors for Columnwise Variations in AVB Insertion Depths . . . . . . . . . . . . . . . . . . . . . 5-14 61 ResolutionofTubeSupport-Tubeswith[- "AVB Indications . . 6-7 6-2 Summary Listing of Unsupported Tubes . . . . . . . . . . . . . . 6-8 71 Beaver Valley Unit 1 Steam Generator Operating Conditions . . . . 79 7-2 Steam Generator Operating Conditions Used for ATHOS Analysis. . . . ,7-10 7-3 Beaver Valley Unit 1 Operating History Data . . . . . . . . . . . 7-7 8-1 Stability Peaking Factor Due to Local Velocity Perturbation . . . 8 17 8-2 Comparison of Air and Steam Water Peaking Factor Values . . . . . 8 18 8-3 Effect of Local Variation of AVB Insertion. . . . . . . . . . . . 819 8-4 Uncertainties in Test Data and Extrapolation. . . . . . . . . . . 8 20 85 Extrapolation of Test Results to Steam Generator Conditions . . . . B 21 86 Final Peaking Factors . . . . . . . . . . . . . . . . . . . . , . 8 22 87 Stability Peaking Factors for Specific Tubes. . . . . . . . . . . 8 23 01444:49/03!4H 10 ]y
u LISTOFTABLES(Continued) IAALE P.AE 91 100% Power Operating Conditions for Beaver Valley Unit 1. . . . , . 9 7 92 Beaver Valley Unit 1 Evaluation of the More Salient Unsupported U Bends . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 98 I 0144N:49/0317H 11
1.0 INTRODUCTION
t This report documents the evaluation of steam generator tubing at Beaver Valley - Unit 1 for susceptibility to fatigue-induced cracking of the type experienced l 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, fleid measurements to establish AVB locations, structural and vibration analysis of selected tubes, and fatigue usage calculations to predict cemulative usage for critical tubes. The evaluation utilizes operating conditions specific to Beaver Valley Unit 1 in order to account for plant specific features of the tube loading and response. Section 2 of the report provides a sumary of the Beaver Valley Unit 1 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 conductive to this type of rupture are discussed in Section 4. Section 5 provides a sumary of test data which supports the analytical i vibration evaluation of the candidate tubes. A sumary of field measurements ' used to determine AVB locations and ultimately to identify unsupported tubes is provided in Section 6. Section 7 provides the results of a thermal hydraulic j 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 , avaluate the tube rupture at Ncrth Anna Unit 1. The final section, Section 9, presents results of the structural and vibration assessment. This section determines tube mean stress, stability ratio and tube stress distributions, and accumulated fatigue usage, forming the basis for the conclusions for Beaver Valley Unit 1. 81444:49/032868 12 1-1
2.0
SUMMARY
AND CONCLUSIONS The Beaver Valley Unit I steam generators are evaluated for the susceptibility uf 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, North Anna 1.
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. The unstable condition prevailed in the R9C51 tube when the tube experienced denting at the support plate. A combination of conditions were present that led to the rupture. The tube is not supported by an anti-vibration bar (AVB), has a higher flow field due to local flow peaking as a result of non uniform insertion depths of AVB's, has reduced damping because of denting at the top support plate, and has fatigue properties consistent with a lower bound fatigue curve for the material in an all volatile treatment (AVT) water chemistry environment. 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 in turn, would provide a 58% reduction in stress amplitude (to <4.0 ksi) for a Row 9 tube in the North Anna 1 steam generators. This reduction is required to produce a fatigue usage of <0.021 per year for the Row 9 tube. This same criteria is used as one of several criteria in the evaluation of Beaver Valley Unit I tubing. With additional effects accounted for through a stress ratio criteria that permits the calculation of fatigue usage to demonstrate tube acceptability, the ftnal criteria of cumulative fatigue usage to date plus future operation with current operating parameters can be satisfied. i The stability ratios for Beaver Valley Unit 1 tubing, the corresponding stress - and amplitude, cnd the resulting cumulative fatigue ustge must be evaluated relative to the ruptured tube at Now 9 Column 51, North Anna 1 Steam . Generator C, for two reasons. The local effect on the flow field due to - t c144M:49/C32464 13 3 2-1
various AVB insertion depths is not within the capability of available analysis techniques and must be determined by test. In addition, an analysis and examination of the ruptured tube at Nor'th Anna 1 provided a range of initiating stress amplitudes, but could only bound the possible stability ratios that correspond to these stress ampittudes. Therefore, the evaluation of Beaver Valley Unit 1 tubing has been based on relative stability ratios, relative flow a peaking factors, and relative stress ratios. The criteria for establishing that a tube has support from, an AVB and therefore eliminate it from further considerations is that at least ~'
This is established by analysis of eddy
- a,c current (EC) measurements. AVB support of the tube may be
~
established b,v,an EC indication of both legs of the AVB or may be established by th'e depth of insertion knowing the geometry of the AVB and EC ind'ications that locate it on larger radius tubes. The AVB insertion depths are a key factor in the assessment of fatigue failure susceptibility since the AVB positions determine the local flow peaking factors. The local flow peaking factor is a direct factor on the apparent stability ratio and a small percentage change causes a significant change in stress amplitude. The relative flow peaking factors for Beaver Valley Unit 1 tubing without direct AV8 support t. ave been determined by instability tests. These factors have been applied to relative stability ratios determined by 3 D tube bundle flow analysis and the combined relative stability ratio used in the stress ratio determination. Eddy current data for the tubes in the area of interest is analyzed to determine the condition of the tubes, and to identify the number and location of tube /AVB interfaces. A more complete explanation of this activity is provided in Section 6 of this report. For the purpose of tts tube.gta,bility , analyses, it was conservatively assumed that all tubes were (in structural terminology) at the seventh (top) tube support plate (TD). In fact, eddy current analyses showed that relatively few of the tubes in the area of interest were dented, or showed signs of corrosion at the TSP interface. None of the tubes in the area reviewed showed signs of thinning at AVB interfaces. Therefore, it is unlikely that any of these tubes have been subject to unstable vibration patterns. C344m:43/c3ttel 14 2-2
In addition to determining the physical condition of the tubes, eddy current data was used to determine AVB insertions depths, and the extent of AVB support for each column of tubes in the area of interest. Review of the eddy current data shows that only in the center 1/3 of the tube bundle were the tubes not supported as far in as row eight or row nine. Within the three SG's, support could not be verified for one row 12 tube, 6 row 11 tubes, and 23 row 10 tubes. Fifteen tubes are identified as the more susceptible of the unsupported tubes. All of these tubes have been evaluated as dented. These are: TUBE TUBE S/G TUBE S/G S/G A R9C35 B R9C35 C R9C60 R9C47 R9C47 R10C47 R10C43 R9C60 R10C60 R10C47 RllC47 R11C49 R11C51 R11C52 R12C51 The local flow peaking corresponding to AVB insertion depths are factored into the evaluation for each of the above tubes. This results in the relative stability ratios and stress ratios shown below. (All ratios are in comparison to R9C51, North Anna Unit 1. Steam Generator C.) Generator Tube Stability Ratio Stress A R9C35 0.73 0.31 R9C47 0.66 0.18 R10C43 0.89 0.81 R10C47 0.77 0.35 R11C49 0.88 0.64 B R9C47 0.66 0.18 R9C60 0.86 0.77 R9C35 0.86 0.77 C344N 43/032688 15 2-3
Generator Tube Stability Ratio Stress C R9C60 0.86 0.77 R10C47 0.77 0.35 R10C60 1.081 nl .0 R11C47 1.272 nl.0 R11C51 1.454 n1.0 R11C52 1.454 91.0 R12C51 >1.12 nl.0 With the exception of five tubes in '.Lm 9enerator C all tubes have relative stability ratios (including flow pen + less than 0.9 and stress ratios less than 1.0. The five tubes exceeding a ,33 criteria are R12C51 R11C52, R11C51, R11C47, and R10C60 in Steam Generator C. Six tubes have been plugged as a preventive measure to eliminate them from further consideration. Sentinel, or tell-tale, plugs were used to permit detection of tube degradd4n if it should occur in the future. One tube R10C43, was plugged in Steam Genir A, b'ased on preliminary analytical results. Final analysis indicates pl% .ng of that tube was not necessary. Of the remaining tubes that meet the criterion, the tubes at R9C35 and R9C60 have the highest stress ratio 0.77. The expected maximum stress amplitude for this tube is 3.1 ksi. Based on the operating history of Beaver Valley 1, which is presented in Figure 715 (see Section 7) in the form of normalized stability ratio versus time in days, the fatigue usage for 40 years is expected to be less than 0.346. This is less than the ASME Code limit of 1.0. The fatigue usage calculation utilizes a lower bound fatigue curve that is consistent witn the fatigue properties of the North Anna tube that ruptured. 2.3 Conclusion The Beaver Valley 1 tubes remaining in service are not expected to be susceptible to fatigue rupture at tha top support plate in a manner similar to the rupture which occurred at North Anna 1. Therefore, no modification, preventive tube plugging, or other acasure to preclude such an event is judged to be necessary other than the plugging of the five tubes in Steam Generator C. The one tube in Steam Generator A. R10C43, may be returned to service during a future outage. 01444:49/CH 46a 16 2-4
3.0 BACKGROUND
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 UrJt 1 Tube Rupture Event The cause of the tube rupture has been determined to be high cycle fatigue. The source of the loads associated with the fatigue mechanism has been determined to be a combination of a mean stress level in the tube and a 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 due to out-of-plane deflection of the tube U-bend above the top tube support caused by flow induced vibration. These loads are consistent with a lower bound fatigue curve for the tube material in an AVT water chemistry environment. The vibration mechanis:n 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 dhmping 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 fcr AVBs 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 l steam generator, i.e., AVB insertion depths. In addition, the fatigue properties of the tube reflect the lower range of properties expected for an
~
0144M:49/031788 17 3-1
AVT environment. In summary, the prerequisite conditions derived from the evaluations were concluded to be: Faticue Reouirementt Prereouisite Conoitions Alternating stress Tube vibration
- Dented support
- Flow excitation
- 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 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 was 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-bend 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 l to have broken through wall from approximately 100* to 140'. From this j point on, crack growth is believed (as determined by striation spacing, striation direction, and later observations of parabolic dimples followed by equiaxed dimples) to have accelerated and to have char.ged direction with the resulting crack front running perpendicular to the circumferential direction. 0144M:49/031788 1B i 3-2
i 3.3 Mechanism Assessment l To address a fatigue mechan 4m 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 centribution at the tube 0D a short distance abo,e 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 first mode, cantilever response of the dented tube to flow induced vibration 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. 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 9ksi. The comparison to the best estimate of plant leakage (performed after the event) showed good agreement, Figure 3-3. Based on these results, it followed that the predominant loading mechanism responsible is a flow-induced, tube vibration loading mechanism. It was shown l that of the two possible flow-induced vibration mechanisms, turbulence and fluidelastic instability, that fluidelastic instability was the most probable cause. Due to the range of expected initiation stress amplitudes (4 to 10ksi),
- the fluidelastic instability would be li,mited in displacement to a range of approximately fhis is less than the distance between
-ac -
tubes at '.he apex, Itwasfurtherconfirmedthatdisplacementprior
~
l to the ryture was limited since no indication of tube U-bend (apex region) damage was evident in the eddy-current signals for adjacent tubes. l l
~
0144M:49/0317tA 19 3-3
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 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 tdicated that a 10% reduction in stability ratio is needed (considerir the range of possible initiation stress amplitudes) to reduce the fatigue usage per year to less than 0.02 for a tube similar to Row 9 Column 51 at North Anna Unit 1. l i
@l44M:49/031788 20 3-4 i -
J l 0-0 e_ g M 1 C
- CC l, t
3 C 180*
- Region of Norringbone gPattarn ,
- so' tr-
/ N E.8 g.g i
p 7 g .
\~A*A 1,
/ r.F
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Coarse Texture
. and 91apled Rupture
+
g Indicatat Origins i Figure 3-1 Approximate Mapping of Fracture Surface of Tube R9C51, S/G 'C' Cold Leg, North Anna Unit 1 0144M:49/030788-21 3-5
5 = 1.5/1.6 m in. I Neuvy Caide A****" 3 5 s 21 m in. 30 f /, 5 a 1.0/1.85 e in, 130 3.C .
~
m Parabolic fl h
- 90' 270* - l 14 Dicolas
\ ) and V Internal kacking f
nr 35 \ Badly , 3 = 2.8/4.0 m in. l g 1.] Petr4d VM @ Nearly Equi.Azad 5 a 6.1/6.9 y in. Digles l i nota: Arrows Indicata Direction of Fracture Propagation Figure 3-2 Schematic Representation of Features obserytd During i TEM Fractograhic Examination of Fracture Surface of Tube R9C51, S/6 'C' Cold Leg, North Anna Unit 1 0144M:49/030788-22 3-6
^
; O l l l l l l l l Caleviated and observed leak ratas versus ties.
Observed values based on gaseous species condenser air ejector seeer- - SISMA A = 5 KSI SISMA A = 7 MSI ,
...........gIguA A . 5 KSI I l o Ar-ds 9 -
O xe-sas il ***' " A Mc-87 N
& i 1
4 --
/ g o -
w a j
/
.' o g
- .o *
. de .sim ashe asise se'e e se'ee ese abe asks amie sees TIME MINLITES
,' Figure 3 3 Calculated and Observed Leak Rates Versus Time t. 0144M:49/030788-23 3-7
4.0 CRITERIA FOR FATIGUE ASSESSMENT Evaluation against criteria to show that Beaver Valley Unit I steam generator tubing will not rupture by fatigue in the manner of North Anna Unit I can only l be done by an assessment relative to the Row 9 Column 51 tube of Steam l Generator C, North Anna Unit 1, since,1) metheds f'or direct analytical
~
prediction of actual stability ratios incorporate greater uncertainties than a relative ratio, and 2) the stress amplitude (or displacement) associated with a specific value of stability ratio can only be estimated by the analysis of North Anna Unit 1. For these reasons, the North Anna Unit 1 tubing evaluation was done on a relative basis to Row 9 Column 51 and a 10% reduction in stability ratio criteria was established to demonstrate that tubes left in service would be expected to have sufficiently low vibration stress to preclude future fatigue rupture events. To accomplish the necessary relative assessment of Beaver Valley Unit I tubing to Row 9 Column 51 of North Anna Unit 1,. several criteria are utilized. First, I stability ratios are calculated for Beaver Valley Unit I steam generators based on flow fields predicted by 3-D therm 1 hydraulic models and raticed 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 model with the same degree of refinement. These ratios of stability ratio (called relative stability ratios) for each potentially unsupported U-bend in the Beaver Valley Unit I steam generators should be equivalent to 5 0.9 of R9C51, North Anna 1 (meeting the 10% reduction in stability ratic 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 assumption, however, that each tube has .
- a ,e To account for these differences, flow l ,
peaking factors can be incorporated in the relative stability ratios or, as noted below, in the relative stress ratios. l l I l 0144M:49/C31768-24 4-1
The second criteria is to obtain stress ratios, the ratio of stress in the Beaver Valley Unit I 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 11.0. The stress ratio incorporates the )lhR9C51inrelationtothestress calculation and also incorporates the ratio of flow peaking factor for the Beaver Valley Unit 1 tube of interest to the ficw peaking factor for R9C51 (flow peaking factor is defined in Section 4.2). This should provide that all tubes meeting this criteria have stress amplitudes s 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 f 1.0 is not necessarily satisfied by meeting the stress ratio criteria since the evaluation for Beaver Valley Unit I has been performed for a reference operating cycle. The Beaver Valley Unit 1 duty cycle was more demanding in early cycles and usage would have accumulated at a more rapid rate. Therefore, the time history of operation is evaluated on a normalized ' basis and used together with the stress ratio to obtain a stress amplitude history. This permits the calculation of current 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, $dundary condition tube to deflect an amount in the U bend which will produce fatigue inducing stresses. ol44M:49/031788 25 4-2
The major features of the fluidelastic mechanism are illustrated in Figure 4-1. This figure shows the displacement response (LOG 0) of a tube as a function of stability ratio (LOG SR). A straight-line plot displayed on log-log coordinates implies a relation of the form y - A(x)D, where A is a constant, x is the independent variable, n is the exponent (or power to which 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 and n is the slope. In our case the independent variable x is the stability ratio SR, and the dependent variable y is tube (fluidelastic instability induced) displacement response D, and the slope n is renamed s. From experimental results, it is known thht the turbulence response curve (on log-log coordinates) has a slope of approximately [ ] *1'Isf results also show that the slope for the fluidelastic response depends somewhat on the instability displacement (response amplitude). It has been shown by tests that a slope of - s' a' range of values correspondin,g to displacement amplitudes in the range of , whereas below *
*a,C are conservative values.
The reduction in response obtained from a stability ratio reduction can be expressed by the following equation:
- a,e where D3 and SR3 are the known values at the point corresponding to point 1 of Figure 4-1 and D2 and 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 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. 0144M:49/031784-26 4-3
_The fatigue curve developed for the North Anna Unit 1 tube at k9051 is from
,. a , c -
l~Thus,
- a,c
~
where, a,'is the equtvalent stress amplitude to aa that accounts for a maximum stress of ya , the yield strength. The -3 sigma curve with 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 ef,fegt of an AVT environment on fatigue and
~
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
- a,c stress amplitudes. These were the ,,
sgown in Figure 4-3. Witha[ ],the
- 5,c 0144M:49/031768 27 4-4
The assessment of the benefit of a reduction in stability ratio begins with the relationship between stability ratio and deflection. For a specific tube geometry, the displacement change is directly proportional to change in stress so that stress has the same relationship with stability ratio, a,C
. .a,c The slope in this equation can range from , ,
on a log scale depending on the am,alitude of displacement. Knowing the stress resulting from a change in stability ratio from SR3 to SR2 , the cycles to failure at the stress amplitude was 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 range of 4.0 to 10.0 ksi by the fractography analysis. It was further developed that the maximum initiating stress amplitude was not more than 9.5 ksi. This was based on ,
]"Thecorrespondingstress level is 5.6 ksi.
8,C with a 10% reduction The maximum stress, 9.5 ksi, would be reduced to . , a ,c in stability ratio and would have a future fatigue usage of Jperyearat 75% availability, Figure 4-4. The minimum stress, 5.6 ksi,'would be reduced to
'wfth a 5% reduction in stability ratio and would have future fatigue
* = >
<g c pdr year, Figure 4-5. In addition, if a tube were already usage of ' '
l cracked, it cou1d be as large as ,
]id8hinlengthandthru-wallandwould not propagate if the stress amplitudes are reduced to s 4.0 ksi.
i l 01444:49/031788 28 4-5
Subsequent 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.
I'd8mulative fatigue usage
~
may then be computed to get a magnitude of alternating stress for the last cycle that results in a cumulative usage 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 [ ][o/R9C51,NorthAnnaUnit1,SteamGeneratorC,andthatthemajor 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. A similar calculation can be performed for the time history of operation assuming that ,
, )* *(in 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 usage of }a,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 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 highe- eurves, a smaller reduction in stability ratio than 10% would result in the same benefit. In addition, there is a large benefit in terms of fatigue usage for relatively small changes in the fatigue curve.
0144M:49/032468 29 4-6
4.2 Local Flow Peaking Considerations j l Local flow peaking is a factor on stability ratio that incorporates the effect l on , }8u'etonon-uniformAVB insertion depths. The flow peaking factor is applied directly to the stability ratio obtained from thermal-hydraulic analysis that does not account 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. , Since the evaluation of Beaver Valley Unit 1 is relative to R9C51, North Anna Unit 1, the flow peaking factors are also applied as relative ratios, i.e., a ratio of Beaver Valley Unit 1 to R9C51 at North Anna Unit 1. The flow peaking relative instability is obtained by testing in the air test rig described 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 [ [is'a'ppropriateforR9C51of North Anna 1. The peaking factor for a tube in Beaver Valley Unit 1 is therefore divided by' ,
)*a'nd'She resulting relative flow peaking is multiplied times the relative stability ratio based on ATH0S results. If the peaking factor is 1.0, the relative flow peaking is ]a,b,e Applying a nominal uncertainty of 15% to the ATH0S analysis result, the predicted stability ratio for R9C51 is ~*'sk'ngnominaldampingthat1sa
~
1 l function of the modal effective void fraction (slip). Applying the ~
*[c[or f
gives a corrected stability ratio of )'Ih%refore, using the minimum flow peaking factor and nominal damping that reflects reduced damping compared to undented tubes, the tube at R9C51, North Anna Unit 1, Steam Generator C is shown to be unstable by comparison to the same tube with a uniform AVB pattern. It is therefore now considered to be sufficient to address the relative susceptibility of tubes to fatigue rupture on , f a,c i 0144M:49/031788-30 4-7
4.3 Stress Ratio Considerations In Section 4.1, a 10% reduction in stability ratio was established to reduce the stress amplitude on the Row 9 Column 51 tube of North Anna Unit 1 to a level that would not have ruptured, 4.0 ksi. To apply this same criteria to anothertubeinthesameoranothersteamgenerator,thedifferencesin[
; a,c
'a,c r
l l l 0144M:49/031768 31 4-8
o,e l By establishing their equivalent effect on the stress amplitude that produced the tube rupture at North Anna 1, several other effects may be accounted for. These include a lower mean s,tres) (such as for non-dented tubes), different frequency tubes from the , , hertz frequency of R9C51, North Anna 1, and shorter design basis service (such as 30 years). 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. , a,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 a11cwable number of cycles to reach a usage of 1.0 is inversely proportional to frequency. The equivalent number of cycles to give the usage of 1.0 for a different frequency tube , j . a,c
.Y l
l 0144M:49/032483-32 4-9
For a different time basis for fatigue usage evaluation, . l a,c.e
- t Knowing the magnitude of the stress ratio allows 1) the determination of tubes that do not meet a value of 11, and 2) the calculation of maximum stress in the acceptable tubes,
'7 2 ,c
. J Having this maximum stress permits the evaluation of the maximum fatigue usage for Beaver Valley Unit 1 based on the time history expressed by normalized stability ratios for the duty cycle (see Secticn 7,4).
J 5 4 i 4 b 0144M:49/031784 33 4-10
Table 4-1 Fatigue Usage per Year Resulting From Stability Ratio Reduction FATIGUE MEAN STRESS USAGE SR, % STRESS BASIS (l) CURVE (2) MODEL PER YEAR REDUCTION
'a,e 0.0207
- 5. 9 yrs to fail (5.6)
- 5. 9 yrs to 0.0107 fail (7.0)
- 5. 9 yrs to 0.0014 fail (8.0)
- 10. 0.0209 max.stre{g) amplitude (9.5)
- 10. 0.0053 max.stre{g) amplitude (9.5)
- 10. 0.0004 max stregg) amplitude (10.3)
- 10. 0.0142 max stregg) amplitude (11.6)
- 10. max. stress 0.0020 based on duty cycle (5) ,
(9.5) (1) This gives the basis for selection of the initiating stress amplitude and its value in ksi. (2) S ,is the maximum stress applied with 5,- Smean + Sa* (3) ,, (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 to failure implied by the operating history requires 1.3 sigma fatigue curve at the maximum stress of 9.5 ksi. 0 94M:49/03178a 34 1
.- 'a,b,e
=
l l l Figure 4 1 Vibration Otsplacement vs. stability Ratio l l [ 0144M:49/030788 35 4-12
a,c Figure 4 2 Fatigue Strength of Inconel 400 in AYT Water at 600'F 0144M:49/030788 36 i 4 13
i l a,c i 1 l i i I l l l I Figure 4 3 Fatigue Curve for Inconel 400 in AVT Water ( ( Comparison of Mean Stress Correction Models l 0144M:49/030784 37 . 4 14
~ _ .. . . _ . . .
l l l l i 4 i 1 I '
- a,c !
i - f I I i ; 4 i j- ( I 3 i L i
- a. .i li e
1
- l. '
j 1 .j g l ! ' b s I i s ! i j l- t i 6 i l. Figure 4-4 Modified Fatigue with 10% Reduction in Stability ' Ratio for Maximue Stress Condition i ! ::44en :14st 11 ; ! 4-15 ! l i i !
. . _ . - _ _ _ _ _ . . . _ , _ , - _ . - . , _ . . . - - - - _ . . . - . . . . _ - , . . _ . _ ~ , _ _ _ . . _ . _ _ . _ . _ - -
i a a,c I j i t i; 5 5 iw l r s i l l
' l t
I Figure 4-5 Modified Fatigue with 5% Reduction in Stability ; Ratio im Minimum Stress Condition [ L l
.:ss u ::::u.n 4-16 !
l l 5.0 SUPPORTING TEST DATA 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 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 configurations, in terms of active tube supports, of a specific tube. These stability ratios represent a measure of the potential for tube vibration due to instability during service. Fluid-elastic stability evaluations are performed with a computer program which provides for the generation of a finite element model of the tube and tube support system. The 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 modo shapes (eigenvectors) for the linearly supported tube being considered. The methodology is comprised of the evaluation of the following equations: Fluid-elastic stability ratio - SR = Ven/Uc for mode n, whers Uc (critical velocity) and Uen (effective velocity) are determined by: 2 Uc -Bf D D))@ [1] n Um e6 n) / (#o and; (#j/#c) Uj kjn z) U . ...................... [2] , N 2 jf3 ("j/*o) Ijn z j i 0144M:49/C31788 40 5-1
I l i l where, l D = tube outside diameter, inches U en - effective velocity for mode n, inches /see N = number of nodal points of the finite element model mj, Uj, pj = mass per unit length, crossflow velocity and fluid density at node j, respectively po, mo = reference density and a reference mass per unit length, respectively (any representative values) S n
= logarithmic decrement (damping) r (jn = normalized displacement at node j in the nth mode of vibration zj - average of distances between node j to j-1, and j to J+1 A
= an experimentally 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 arbitr'ary 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).
The uncertainty in each of these parameters is addressed in a conceptual manner in Figure 5-1. The remainder of this section (Section 5.0) provides a discussion, and, where appropriate, the experimental bases to quantitatively establish the uncertainty associated with each of these parameters. In l l 0144M:49/031768 41 5-2
addition, Section 5.3 provides the experimental basis to demonstrate that tubes with - - -.a,e
~
, This implies thatthosetubeswith[ Nuldnothavetobemodified because their instability response amplitude (and stress) would be small. The very high degree of sensitivity of tube response (displacements 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.
Freauency 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 gagttcularly appropriate in the case of dented , tubes. Therefore, uncertainty levels introduced by the frequency parameter are expected to be insignificant (see also "Average Flow Field" subsection below). Instability Constant (Bet 3_). 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 experierices 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 ATHOS 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 ,
)a,b,e 0144M:49/031784 42 5-3
i A sumaary of the tcst bases and qualifications of the beta values used for these assessments is provided by Figurg,-). The lowest measured beta for tubes without AVBs was a value of , ) This value is used for the beta l parameter in all stability ratio evaluations addressed in this Report (see also "Average Flow Field" subsection below). Mass Distribution 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 the t a,e Tube Damoinq 4
. a,c Test data are available to define tube damping for ,
, tube supports, appropriate to dented tube conditions, in steam / water flow
- conditions. Prototypic U-bend testing has been performed under conoitions
~
leading to' '
's0pports. The data of
[nFigure54providesthe principal data for } 'W6econditionsinsteam/ 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 Sauare Root Density Distribution l 3 Average and U-bend-local flow field uncertainties are addressed independently in the following. 0144M:4 /031784 43 5-4
l l Averaae Flow Field Uncertainties in the average flow field parameters, obtained from ATH0S analyses, coupled with stability constant and frequency, are essentially the same for units with dented or non-dented top support plates. If the errors associated with these uncertainties were large, similar instabilities would be 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 estimate of about , jfor the combined effects of average flow field, stability constant and frequency appears to be reasonable. To further minimize the impact of these uncertainties, the Beaver Valley Unit 1 tubes are evaluated on
- 'a,e ~
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 at,out by the "channeling" effects of non uniform AVBs, leads to a local perturbation in the velocity 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 evaluations addressed in this report are provided in Section 5.2, below. Overall Uncertainties Assessment Based on the above discussions, and the data provided in the following
- ections, it is concluded that local flow peaking is likely to have contributed significantly to the instability and associated increased vibration amplitude for the failed North Anna tube. Ratios of stresses ar.d 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 and tube damping.
01440 49/031768 44 5-5
5.2 Tube Damping Data The damping ratio depends on several aspects of the physical system. Two primary detenninants of damping are the support conditions and the flow field. It has been shown that tube support conditions (pinned vs clamped). affect the damping ratio significantly. Further, it is affected by the flow conditions, i.e., single-ohase or two-phase flow. These effects are' discussed below in more detail. Reference (5-1) 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 ef fects are very small. Dartping 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, wl.ere 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 b Mitsubishi Heavy Industries (MHI). These data ae weregatheredunderf, , tu e support conditions. The second is comprised of the results from tests sponsored by the Electric Power Research Institute (EPRI) and reported in References (5-f:) and (5-3). \ 1 The damping ratio results from the above tests are plotted in Figure 5-4 as a function of void fraction. It is.important to note that the void fraction is
~
~
8'C ! determined on the basis of . l I 1 l o!44M:49/031766 45 5-6
(Reference (5-4)). The upper curve in the figure is for pinned support conditions. 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 than specific test results. The lower curve pertains to the clamped support condition, obtained from
^
Reference (5-3). 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 supporti 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 Reference 5-3 data as reported show a damping value of .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 tub,e vibration in air has shown that the mechanical damping is only( , rather than the .55; reported in Reference (53). Therefore the lower curve in Figu,re,5-4 is the Reference (5-3) data with all damping values reduced by , , i 1 i t 0144M:49/031788 46 5-7
l l 5.3 TubeVibrationAmplitudesWith( [ANSupport i A series of wind tunnel tests were conducted to investigate the effects of tube /AVB eccentricity on the vibration amplitudes caused by fluidelastic vibration.
.a
-L,C Prior test results obtained during the past year using this apparatus have demonstrated that the fluidelastic vibration characteristics observed in the tests performed with the cantilever tube apparatus are in good agreement with corresponding characteristics observed in wind tunnel and steam flow tests using U-bend tube arrays. A sumary 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. ,
,8 , C l ousn:4s/emas u L
5-8
As shown in Figure 5-7; the tube vibration amplitude below a critical velocity is caused by , a.c Figure 5-7 shows the manner in which the zero-to-peak vibration amplitude, ~ expressed as a ratio normalized to( ',*[arieswhenonegapremainsat -
. e,e
, For increasing velocities, up to that corresponding to a stability ratio of
'.** 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. As expected, impacting is only
. ",a , e observedinthe[ ,
~
It is concluded from the above test results that, for a ,
..a , e
.f i
! 5.4 Tests to Determine the Effects on Fluidelastic Instability of ! Columnwise Variations in AVB Insertion Depths l This section summariz'as 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 configuration is defined as a specific set of { insertion depths for the individual AVBs in the vicinity of an unsupported l U-bend tube. l The tests were conducted in the wind tunnel using a modified version of the l l cantilever tube apparatus described in Section 5.3. Figure 5-9 shows the , conceptual design of the epparatus.* The straight cantilever tube, ( - l 0144M:49/031788 4 5-9
C
'a,e 9
The ' can be individually positioned. The test array has 8 columns and 13 rows with 7 movable AVBs and fixed half width AVBs at the . a,e sides of the test section.
. Figure 5-11 shows the AVBs, when the side panel of the test section is removed. Also shown is the top flow 3 screenwhichis[ {
.a.
~ '
. Ihe AVB configurations tested are shown in Figure 5-12. Configuration l'a 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.
J l 1 0144M:49/0324H 49 5-10
1 l As shown in Figure 5 9, -
)
a a,C Allthetubesexceptthe,jngtrumentedtubed(correspondingtoRow10)are[ rigid
, As discussed in Section 5.3, prior testing indicates ,
that this situation provides a valid model. The instrumented tube
'nh shown in Figure 5.10. Its(~[Nction vibrational motion is measured using a non-contacting transducer.
7'#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 ll. The difference in profile is quite small for these bounding rows. difng a hot-film ( , anemometer located as shown in Figure 5-9. Figure 5-13 shows the ras vibration amplitude, as determined from PSD (power spectral density) measurements made using an FFT spectrum analyzer, versus flow velocity for Configuraiton Ib (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. l 01444:49/032468 50
The main conclusions from the tests are:
- 1. Tube vibration below the critical velocity is relatively small, typical of turbulence-induced vibration, and increases rapidly when the critical velocity for the initiation of fluidelastic vibration is exceeded.
, 2. Configuration Ib (R9C51 in North Anna) has the lowest critical velocity of all the configurations tested.
- 3. Configuration Ib is repeatable and the configuration was rerun periodically to verify the consistency of the test apparatus.
The initial test results obtained in support of the Beaver Valley Unit 1 evaluation are sumarized in Table 5-2. The test data is presented as a velocity peaking ratio; a ratio of critical velocity for North Anna tube R9C51 configuration la, to that for each Beaver Valley Unit 1 AVB configuration evaluated. 5.5 References
'a,c 5-1 5-2 53 i !
5-4 i l l l l l C144M:49/032468 51
v Table 5-1
>ind Tunnel Tests on Cantilever Tube Model DBJECTIVE: Investigate the affects of tube /AVB fitup on flow induced tube vibration.
~
APPARATUS: Array of centilevered tubes with and supports
'.ar c MEASUREMENTS: Tube vibration amplitude and tube / AVE impact forces or preload forces.
RESUI.TS:
.,, a,b,c l
l
}.
I f 2. 3. 4. l 5. 6. 1 8144m:49/03:484 52 5-13
Table 5-2 Fluidelastic Instability Velocity Peaking Ratios for Columnwise Variation in AVB Insertion Depths (Beaver Valley 1) Type of Insertion Peaking Ratio Configuration Ula/Un
a,b,e la Ib 2a 3
4a 4b 4e 4d 4f Sa ! Sb Sc 6c , , J ' Note: Un is instability velocity at inlet for type n of AVB insertion configuration. einn:4s/esins 53 5-14
1 i i ! 1 . F i 1 i ! a,c s I' , I r. 1 i f I j ' i i i L i. ll 1 ! I ! i ; 1 , t i t i l 6
- l i
l . . l l I l { , I i i Figure 5-1 Fluidelastic Instability uncertainty Assessment l t
?
1 1 0144M:49/030788 47 : 5 15 ' i 8
4 U0 Bend Test Data i
- 1) MS 3 Tests a
- a,b,e A values of , ,
- 2) M8-2 Tests - a,b,c A of , ,
- 3) Air Model Tests A of ~
~$lkoutAVBs
~
w Tendency for A to increase in range of , w iik'
- inactive AVBs (gaps at AVES)
Tendency for A to decrease toward a lower bound of
,,lk wi 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 30 analysis
- 4) 'A determined from calculated critica) values GoodagreementwithreportedA,valu,e(knowndampingshouldnot ATHOS velocity data with $ of ~ and l 5)
I significantly underestimate instability for regions of uniform U bend flow l l Figure 5 2 Instability constant - A cim;onstres ss 5-16 L
,j l I i I 1 v 1 )
"' a,b.c .a .
) + i. l ) I l l l l l
'h w Figure 5-3 Instability Constants, A. 00tained for turved Tubes fre.
. Wind Tunnel Tests on the 0.214 Scale U Bend Model M3;M:49/0303BB 39 5-17 1: ,
a,b,e i I f i i l l l l l l l
=
Figure 5-4 Desping vs. Slip Void Fraction 0144M:49/030788 50
' 5-18
. ._. ._._- _.._ _ ___ _ _ _ . _ _ _ _ . _ = . _ _ _ _ . . _ . . _ . . . _ _ . _ _ _ __ _.
i t 1 t 6 i ;
- i. - 9 a,b c ,
i a f
~I 4
i i, ! ,6 > L I I 2 0 [ i 1 , i ! i ! I i t i 5
- i i
4 . j i f i l 1 4 , 1 l i t k t 4 l t i I i i i i i e L s
- l t
i I L t t I I i Figure S-5 Overall View of Cantilever Tube Wind Tunnel Model j - t i I L t
- I t
, 5 19 !
F r - I~ l
*a,b,c a
I. A k t i 4 F i l r I I i 4 'l i i I f i i i i
, l' l I 1
4 h i s : i il 4 . i , L 1 c 2 I i l 4 Figure 5-6 Top view of the Castilsvar Taha Wind Tunnel Model l l t 5-20 I
i
. 1 i
i i 4 1 1 i 1 l t 4 i i '
- a,b.c l' f l ;
1 : t r
- i i
1 t } . ! 1 i a
- t i
I i, t l f I t ! i i l l I i i i I i i f , L l, ! ! 4 i ! I i W l I i : i t i Figure S-7 Fluidelastic Vibration Amplitude with Non-Uniform Gaps i t 5-21 .l - . _
._._-_1
4 l
' I
) a.b.c l I 1 t i L 1 f i 5 i
)
i ' . - . .. -. s ' Figure 5-8 Typical Vibration Amplitude and Tube /AVB Impact Force ] signale for Fluidelastic Vibration with Unequal , 1 Tube /AVB says I i f I i* ill 3 l 5 22
" Y a,b,c l J FiTure 5-9 Conceptual Design of taa Apparatus for Determining the Effects of Fluideiastic Instability of columnwise
. Variations in AVB Insertion Depthe 5 23
. *a,b c l . s Figure 5-10 overall view of Wind Tunnel Test Apparatus 5-24
j l
.l ,
, 1
)
l v l j i I
,o 9 a.b.c i l
i , i i 1 1
> r 4
l : 8 . + i j t i 4 I I i i i 1 1 i . I l-i
'N i
t 1 .
. t i
Figure 5-11 side View of Wind Tunnel Apparatus with cover Plates l Removed to show Simulated AY5s and Top Flow Screen . l t, t i l 5-25 i
- k. -
j TWE OF AVB TYPE OF AVB wasRTON MSElm0N a,b.c a,b c 1a 4d ~ 1b 4f ta Se i 3 5b da 5c 4b 6c L 4e Fgure 512 AVB CONFIGURATIONS TESTED FOR BEAVER VALLEY 1 5-26
i t l t t i !
. < r 1
i 9 k
,. a,b.c :,
l i I. I 6 J I t t 4 i i I ! i i i i
. r d
f
- i i !
i l i i I' I , t t j l i l ! i I i i 1 : t
- [
! i , t J li I r ! Figure 5 13 Typical Variation of RMS Vibration Amplitude with Flow l l Velocity for Configuration la in Figure 512 r l P
- - - - - - -------~---,-n, -w. -
i 6.0 ED0Y CURRENT DATA AND AVB POSITIONS 6.1 Tube Denting at Top Tube Support Plate Eddy current data from the 1986 and 1987 outages were examined to evaluate the incidence of corrosion and/or denting at the top tube support plate for the ! tubes in rows 8 through 12. Because the tube vibration analyses were based on the conservative assumption that all tubes in the area of interest were structurally [ Mthe TSP holes, as if by denting or corrosion, the
]
results of this phase of the examination, which are plotted, are informative, but do not have an impact on the final decision as to whether to plug a tube or not. 6.2 Tube Wall Thinning at the AVB Supports No tube wall thinning was observed at the tube /AYB intersections in rows 8 through 12 of the bundle. 6.3 Eddy Current Data for AVB Positions The AVB insertion depths were determined on the basis of interpretation of the eddy current data. To locate the AVBs, the ECT data tracos were searched for the characteristic peaks seen in the signals, which indicate the intersection of an AVB (or a true support plate) with the tube (Figure 6.1). The number of these intersection, including zero, were logged for each tube to indicate the presence or absence of AVBs. Where only a single intersection was indicated by i the data, the length of this intersection was recorded to provide additional information to assess the adequacy of support for the tube. Figures 6.2, 6.3,
.and 6.4 show the number of AVB signals found for each tube, as well as the l
i condition of the tube to TSP interface at the locations wlyen it was evaluated, i Since ambiguity can occur in the interpretation of the ECT data, due to i inability of ECT to differentiate at which side of a tube a "visible" AVB is located, other information was used to assist in establishing the location cf the AVBs. Consistency with the design of the AVB assembly, consistency of data for adjacent columns and verification by [ %eutilizedtodetermine i 0144M:49/03246849
the depth of insertion which was plotted. For the cases of [
! i a,c Data from the 1986 outagc were the primary data source used to determine the positions of the AVBs. In a few casts, the 'new' tapes from the 1987 outage, which was underway at the time, were reviewed to confirm laterpretation of some signals from the '86 tapes. The Eddy Current data analysis was performed principally by the SG Inspection and Analysis group of Westinghouse STD.
6.4 AVB Insertion Depths AVB position maps for steam generators A through C are given in figures 6.2 through 6.4. The number of visible AVB indications per tube, [ ] o'l AVB insertion, and the condition of the tube to TSP interface are shown on these figures. The direct observation data (the number of AVB intersections seen by the eddy current probe) are the principal basis for determining the AVB positions. Wherethedirectobservations,wgreambiguousorthereisaconflictbetween the more conservative data are used to determine observationsand{ , i the AVB positions. Since ' direct observation' gives a 'yes - no' type of j
)**$he visual images thus produced being more easily understood when fluid flow peaking situations are evaluated. !
l Greater conservatism is generally interpreted as the AVB being less inserted although consideration must also be given to the resulting flow peaking factors. No attempt was made to inflate the flow peaking factors through , conceptually possible, but unreasonable, interpetation of the data. L l 0144M:49/C32486-60 6-2
6.4.1 AVB Assembly Design The design of the AVB assembly for the Model 51 Steam Generator includes ,
)*'h'o* AVBs are found outside
)**fteiewoftheeddycurrentdatafor Beaver Valley Unit I shows that and only in the middle third of the tube bundle, do the lower AVB's not extend as far inward as row 8 or row 9, and typically less than 20 row 10 tubes are not supported.
Eddy current data,sometimes indicates the presence of AVBs in column 2 and/or
- 91. Since . .
e a,C 6.4.2 AVB , [ The' '$echnique is useful where noisy ECT signals prevent direct I obse'rvation of the AVBs, where testing is impossible due to plugged tubes, and l in some instances to resolve ambiguities in the ECT data. Since , a,e In the case where the AVB characteristic signals can not be confidently determinedduetoanoisysignalorpreexistingpluggedtubes,[ a,c.e 1 . c14w:u/mna 61 6-3
[ t a a.c l 6.4.3[ ]8$tactArc The adequacy of support provided by'~ ~$ontactindicationsmustberesolved
~
in the cases where a potentially susceptib1e tube is concerned, since plugging j the tube may be required if adequate support cannot be shown. Preliminary 1 - analysis indicated that row 10 tubes with peaking factors greater than the failed tube at North Anna, and any tube in rows 11 and 12 were potentially j susceptible. Consequently,theprimaryfocusregarding[ [rh"contactswas for the tubes in rows 10 through 12. Arc lengths can be utilized as a basis for determining the support conditions ofthetubeswhenclear[ },ior the tubes in rows 8 through 12, the theoretical contact length of an AVB whose m
$$hes. In this analysis, a , ] *'*
i 01444:49/C317 M 62
- 6-4
r ,
/rclengthdatawasutilizedinparttoverifytheAVB positions in S/G A at R10C60 and R9C60.
i ~ 6.4.4 AVB Map Interpretations. ,,i $/G - A The AVB map for SG A is given in Figure 6.2. In regions where the potential for high flow peaking factors exists, the tubes in R11C49, 50, and 51, are ,,
, conservatively assumed to be unsupported based on ~
to
~
locate the AVB's. The fact that each tube has a ~ ][d contact signal suggests that they are in fact, most likely supported on one side or the ' other. AVB's near tube R10C43 are conservatively [ ] *' contact signals. The resulting flow peaking and tube vibration analyses during the plant outage led to the recomendation that this tube be plugged. Final analysis indicated that plugging of this tube was not necessary. AVB locations aroundtubeR9C60arebasedinparton[ [a'r'ccontactreadings. (seesec. 6.4.3) The remainder of the AYB plots are based on , ,
- S/G B The AVB map for SG B is given in Figure 6.3. The SG has no unsupported tubes
) in rows 12 or 11, and one unsupported tube in row 10. The most likely sources of flow peaking factors were the AVB geometries near the tubes in locations R10C47, R9C35, and R9C60. All of these locations were conservatively assumed to be u: rupported. None had flow peaking factors high enough to require plugging. S/G - C ; The AVB map for SG C is given in Figure 6.4. In regions where the potential for high flow peaking factors exists, the tubes in R11C51 and C52; R10C42, 43, 1 65 C144M;49/C324E4-E3
l l l I 4 44, 45, 46, 47, 50, 51, 52, and 53; and R9C60 are clearly unsupported. l Additionally the tubes in locations R12C51; RllC47; and R10C40, 41, and 60 were conservativelyassumedtobeunsupportedbasedon( t'o' !
AYhcontact I locate the AVB's. The fact that each of thesc tubes has a ~ ; signal suggests that they are in fact, most likely supporte'd on one side of the other. The resulting flow peaking and tube vibration analyses led to the recommendation that the tubes at locations R12C51, R11C47, 51, and 52, and R10060 be plugged.
6.4.5 Summary of Tube Support Conditions Resolution of tube support evaluations for single AVB indications are listed in Table 6-1. A summary of unsupported tubes is given in Table 6 2 .l J
}
i l I 6-6 0144*:49/C31768 64
Table 6.1 . . a,e Resolution of Tube ?.apport - Tubes With', , ,AVB Indications A R11C46 .
' a,c A R11C49 A R11C50 A R11C51 A R10C41 A R10C42 A R10C44 A R10C45 A R10C60 A R9C39 A R9C40 A R9C41 , A R9C44 A R9C60 B R10C46 l B R10C47 B R10C46 1
B R9C34 8 R9C35 B R9C40 B R9C55 B R9C60 l i C R12C51 1 1 C R11C42 C R11C46 C R11C47
~C R11C53 l C R10C40 i C R10C41 C R10C48 C R10C54 C R10C60 C R9C35 C R9C38 C R9C39
- C R9C55 l C R9C56 ,
a Tube locations inboard of R9 do not develop sufficiently high flow peaking factors to be of analytical interest. 0144M:49/031788 65 g,7
Table 6.2 Sumary Listing or Unst:pported Tubes l l Eram Generator 'A' Row 12 No Unsupported Tubes Row 11 Columns 49, 50, 51 Row 10 Columns 43, 46; and columns 47 through 54, Row 9 Column 35; and columns 39 through 56 Row 8 Columns 12 through 16; 22 through 27; 29 through 31; 34 through i 57; and 60 and 61
,1 team Generator 'B'
! Row 12 No Unsupported Tubes l Row 11 No Unsupported Tubes Row 10 No Unsupported Tubes Row 9 Column 35; columns 42 through 55; and 60 l Row 8 Columns 10, 11, 12, 34, 35, 38 through 57; 60, 61 79, 82, 83, 92,
~
l and 93 Steam Generator 'C' Row 12 Column H
' low 11 Columns E. H. E l Row 10 Columns 40 through 47; 50 through 54; and fQ Row 9 Columns 39 through 54; and 19 Row 8 Columns 35; columns 38 through 56; 60 and 61 l
I Analy;is indicates that tubes shown in boldface ty2st require plugging for susceptibility to North Anna rupture event. l l 0144M:49/C12488 66 6-8
4
- a,e ;
N l l i s w Figure 6-1 AVB Insertion Depth Confirmation 0144M:49/030788-59 6-9
- - . _ _ , - - -~---e---n-, --- w -m - , ,- - - , - - - , - - , - , - - - - ,-,---,----+nn.
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o goose 8 eeGGG J s OseeS i _ GGGGG 11 GGGGG il EoG669 : G9999 s : OG999 : : di GOGGO i: _ 8G999 : : 99990
- : GOGGG a :
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. 8 e l i c l :121 a l = b f 21 :121 a l a f I 21 :1 S l a l =
w ' f % v % 1-
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l Fgure 6 2: Beaver Vaney Unit 1 Steam Generator A AVB Positions 6 10
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~
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l ei=iei.i.
~ l }i g33 f'((
l
@ LJ Fgure 6 3: Beaver Valley Unt 1 Steam Generator B AVB Poshions 6 11
a eGeee Ei } d l eeeee s eeeee : ::
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- : a a
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gj 99999 : _eeeGe t : : 88999 } t 8 : e eeeee t eeeee n eeeee ls , , j .) m eeeee l_ _00009 al eeees e.: i 1 i 3 eeeee : : 2000G s1 eeeee : II} : 1 eeeee eesse si eGeee a : eeeee l_ eeeGO : eeeee : jiq : esees : eeeee : : A m e e e_ : : f
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l _eeeee ?: esees il eeeee : : iij eeeee eeeee : : eeeee : : eeeee i_ eeeee ll eeeee : : O00 : 0 eeeee li eeeee si eeeee : 4 i eeeOO li eeOGO : s eeeee : :
,I 9e999 : : eeGee si eeeee : : 6 eeeee 8 I3eGee si eeeee : !
i_ '980e9 11 eeeee : a ' l A series of calculations were completed to generate a normalized stability l ratio for each of the six fuel cycles since the plant became operational in June 1976. Data for this evaluation is summarized in Table 7-3. Included are
- ycle average values for full load steam pressure and primary fluid average temperature. The number of days that the plant has operated above 85*. of full power are also listed. Since tube vibration and possible fatigue are associated with operation at close to 100% power, only these higher power
- operating periods are considered important to the evaluation. The operating l parameters listed in Table 7-3 were then input to the Westinghouse "GENF" l
I 0145M:49/031188 7 7-7
computer code to determine the overall performance of the steam generator, in particular, the circulation ratio for each fuel cycle. These calculated values are also listed in the table. l The resulting normalized stability ratios are shown in Figure 7-11. In this figure, the normalized stability ratio for each fuel cycle is plotted against cumulative operatir.g time above 85% powir. Note that the ratio assigned to each of the high power intervals listed in Table 7-3 (85-90%, 90-95%, and 95-100%) and plotted in Figure 7-11 has been conservatively based on the highest power level in each interval. Figure 7-11 indicates that the full power normalized ratio has been a nearly constant throughout operation. However, small cycle-to-cycle variation which is predicted is related to changes in steam pressure; higher pressures yield lower. stability ratios and vice versa. Higher steam pressures result in higher U-bend density, lower U-bend velocity, and increased damping as a result of lower voids in the U-bend. The higher damping, together with decreased loading on the tubes, result in the lower normalized stability ratio which is indicated, for example, i in cycle 6. Figure 7-11 also shows that the stability ratios and operating periods at the lower power intervals are negligible compared to the full power results and could have t'een disregarded in the analysis, i ! References
~
s a,c 7-1 m b l l I l l I 0145M:49/011188 8 _ -_7 - 8
Table 7-1 Beaver Valley 1 Steam Generator Operating conditions Power B87 of 825 MWT Steam Pressure 837 psia Feedwater Flow Rate 3.90 x 106 lbm/hr Feedwater Inlet Temperature 435.5'F Water Level 44% of Harrow Range Span (Assumed) Primary Inlet Temperature 604.5'F . Primary Outlet Temperature 546'F i l l l l l 0145M:49/031188 9 7-9
1 Table 7-2 Steam Generator Operating Conditions Used for
- ATHOS Analysis Power 887 MWT Primary Flow Rate 3.60 x 107 lb/hr Primary Inlet Temperature 604.5'F Primary Outlet Temperature 546 'F 6
Feedwater Flow Rate 3.90 x 10 lb/hr Feedwater Inlet 435.5'F f Temperature Water Level from Tubesheet 505.9 inches Steam Pressure 83'/ psia Circulation Ratio 4.99 I l l 0145M:49/03118810 7-10
L l r> Table 7-3 8eaver Valley thtt 1 Ope.atins History Data l 4 Full Lond* I Days at Each Pouer Level Steam
- Pressure TAVG* Calculated Cycle seeinnine End Gir-901 M-955 M (estal (Det F1 Cfrc. Ratio
. .a,c 30-Nov-79 29 39 223 829 573
, 1 14-JUN-76 1 25-0EC-8 24 37 159 835 573 2 20-N0v-86 10-Jun-83 8 22 241 834 573 3 09-JUL-t2 j 11-0CT-84 8 8 297 824 574 4 24-5EP-83 . d 20 331 422 574 j 5 04-JAll-85 17-leW-86 13 i
^
16 7 364 837 576
- s 25-AUS-86 11-DEC-87 . .
I _ _ 98 133 1615 i J P l
- For 56 8 1
i 0145ft:49/031188 w - - , , . - . - , ,- ,- -,
w M
, 'a,c.e s
I . l l l ! Figure 7-1 Plan View of ATH05 Cartesian Model for Beaver Valley 1 1 l 0145M 49/030888-74 7-12 t
l a,c,e l l l l l 4 . l . l l i 1. l Figure 7-2 Elevation View of ATH03 Cartesian Model for Beaver Valley 1 0145H:49/030888-75 7-13
I l a,c.e l _ I wb l l i i l t 1 l Figure 7-3 Plan View of ATH05 Cartesian Model for Beaver Valley 1 Indicating Tube Layout l 0145M:49/030888-76 7-14 ; l l
a,e f l r l l Figure 7-4 Flow Pattern on Vertical Plane of Smtry 0145M:49/030ES8 77 7-15
O i i
, a,e f
l a 1 Figure 7-5 Lateral Flow Pattern on a Horizontal Plane in the U Bend Region 0145H:49/030888-78 7 16
i
, a,c 1
P i Figure 7-6 Lateral Flow Pattern on Top of Tubesheet 0145M:49/030888 79 7,37
. a,c 5
l l l 6 l l l l Figure 7-7 Tube tap Velocity.and Density Distributions . for Tube Row 10/Coluri 3 l l 1 0145M:49/03088880 7-18
Sa,e e l l l l l l l Figure 7-8 Tube tap Velocity and Density Distributions for Tube Row 10/ Column 20 i i 0145M:49/030888 81 7-19 i
- a,e e
i i
\
l Figure 7-9 Tube 8ap Velocity and Density Distributions for Tube Row 10/Colism 40 t i i 0145M:49/030888 82 7-20
i I d o l i 4
- a,e J
l 3 i . i 1 i l l . f Figure 7 10 Average Velocity and Density in the Plane of the U-8 ends Normal to Row 10 l 0145M:49/030888 83 7-21
- ----,,,,-,------,-----.---,-,,---,,_----,-n,, , c .c n n-- , , - - - - - , - , ,
l l l N 3 .
- # ~
c /\- \ -
;.. 1 n
2 i
- i S-2 l .
r 1 r, - u
- 7 l
, . . I.w l 3
n . r
-E kt 1 .
= k .
o 1 N - o
.s ri 3 .
.No e
r i
, , e ,,,,a e a e e s , e a e
- 95-*32I2222s;ooeeoe 3828G8 l
I ss eoeeoeeee e ende ws / x mm ws l l Figure 7-11 Seaver Valley 1 Normalind Stability Ratic BasedonHighPower(>86%) Operation t l 0145M:49/030888 84 7-22
8.0 PEAXING FACTOR EVALUATION This section describes the overall peaking factor evaluation to define the test based peaking factors for use in the tube fatique evaluation. The evaluation 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 fatigue assessments as the peaking factors for all other tubes are utilized relative to the R9C51 peaking factor. Uncertainties associated with applying the air model test results to the tube fatigue assessments are also included 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 8.1.1 Ba.ckground 1 The AVB configuration of the ruptured tube in North Anna, R9C51, is the reference case for the tube fatigue evaluations for other plants. In accordance with the NRC Bulletin 88 02, the acceptability of unsupported tubes in steam generators at other plants is based on tube specific analysis relative to the North Anna R9C51 tube, including the relative flow peaking factors. Thus, the support canditions 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" l interpretationsoftheeddycurrenttest(ECT[ data (Figure 8-1),were reevaluated using the } }de'velopedsincetheNorthAnna event. The , [.**The results of this evaluation are sumarized below. l 0145M:49/031158 23 8-1
~
)
1 8.1.2 Description of the Method i
, a,c The basic method utilized was the ,
in which the AVB position ,is determined based on ' I*'[nthisstudy,the ffe'chniquewasutilizedinthe j'Theobjective 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. l
- The data were examined by an eddy current analyst experienced in reading -
these traces, and by a design engineer knowledgeable in tha geometry of the l Model 51 U-bend region. The intent.of this review was to determine if the presence or absence of AVBs asshowninFigure8-1couldbeconfirmedusingtheAVB(
}
technique. Preliminary [ l i l l l l I i 8,C Q149I:49/031184 24
l i k b 4 a,e Figure 8-4 is the "AVB' visible" map for columns 48 through 55, based on the i critical review of the data. It should M noted that thet original data 1 interpretations and the review interpre'tations are consistent. a,e 8.1.4 , } I l P
'a,c !
6 t 0345M:49/031184 25 8-3 ,
^
b
,] a , e
,a,e The logic in arranging the( ,
data is based on the following two rules:
.a.c Rule 1. The , ,
of the same AYB based on different tubes in the same column must be consistent. t a,e Rule 2. Two' adjacent tubes in the same row '
,]tonsequently, the difference in the {
a,e i i c14 m 4s/o31tes rs 8-4
I The implementation of this is that if the position (either left=BsC or right) l
=3 . ,
,A95is assumed for a column, then the in the )
of a[ , , adjacent columns are also , _, , The arrangement of the AVBs as shown in Figure 8 5 satisfies the rules above 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 The general AVB arrangement surrounding the ruptured tube in North Anna-1, Steam Generator C, which was the basis for the analysis, is confirm.d 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.
Thebasiso,fthereviewwasa[ ]te'chniquewhichutilizes[ ! ~*'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. l 8.2 Test Measurement Uncertainties i The descriptions of the peaking factor tests and apparatus were provided in l Section 5.4. All practical measures were taken te reduce uncertainties, j Nevertheless, some still remain and should be properly accounted for. The important parameter measured during testing that has a significant impact on , l ! 0145M:4$f431188-27 8-5
1 l peaking factor is the air velocity. The air velocity at test section inlet wasmeasuredusinga] ) Based on considerable experience with the use of such instruments, it is known that the magnitude of uncertainty is very small. A((measurementuncertaintyisusedinthis l analysis basad on past experience. 8.3 Test Repeatability During the peaking factor testing of AVB configuration, each test was performed at least two times to confirm repeatability. It has been demonstrated 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 A first order estimate can be made of the validity of modeling a U bend tube by a cantilever tube in tests to determine the effects of AVB insertion depth on the initiation of fluidelastic vibration. The following assumptions are used: . a,e
- 2.
3. i 0149t:48/C31168 28 8-6
4. J 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 prototvoical Row 10 tube.
.a,C The compari, son 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. ,
1 i l 4 l
.ac l
i i i 0145M:49/031184 29 8-7 - - -
e i
.n.c
, Using these values, the ratio of the effective velocity,fgr the cantilever measuring tube to that for the U-bend tube is about ( ,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 dept required to support a Row 9 tube. , l I l
]a,eThe net ~ result is that the ratio of the effective velocity for the cantilever
*'C tube to that for the U bend tube is about [ [
These results indicate that, for the particular assumptions used, the cantilever tube model appears to be a reasonable representation of the U-bend l 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 tubetosimulatetheU-bendisabout((a,e 8.5 Air vs Steam Water Mixture The local peaking factors from the air tests can be applied te the steam ( generator steam / water conditions either as a direct factor on the mixture
; C145M:43/0311 H 30
! 8-8
l i velocity and thus a direct factor on a stability ratio, or as a factor on the steam velocity only with associated impacts on density, void fraction and damping. This method leads to a reduction in tube damping which enhances the
- peaking factor compared to the direct air test value. For estimating an absolute stability ratio, this application of the peaking factor is a best i
estimate approach. However, for the evaluation of tubes relative to stability ratio criteria, it is more conservative to minimize the peaking l 4 factor for the North Anna Unit I tube R9CSI through direct application of the air test peaking factor. This conservative approach is therefore used for evaluating tube acceptability. 1 Under 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 relative to velocity perturbations. Under non uniform AVB insertion the steam / water flow may i 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 u water phase. f Based on the above discussion, the Fj are considered to more appropriately ; apply to the steam phase. Thus, it follows that mixture mass velocity for ; j the tube subject to flow perturbation cari be written as follows:
~
- a . C.
i 1
~
is the vapor density, Df the water density, Fa the velocity L where D g j peaking factor determined from air tests, j g* the nominal superficial vapor i velocity, and jf* the superficial water velocity. Steam quality can then be determined as follows: l l l I l 0145M:49/0311aa 31 8-9
. a,c The[ , ,
[,*asusedinthe ATHOS code, is applied to determine void fraction. Subsequently, mixture density, velocity and damping coefficients for the tube which is not supported and subject to flow perturbation is evaluated. Therefore, similar to the 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,e ,
where sF is the stability peaking factor, Fd the density scaling factor, Fy the velocity scaling factor, and Fdp the damping coefficient scaling factor. If we use the air velocity peaking factor without translating to steam /waterconditions,then
- - a,e As shown in Table 8-1 stability peaking factors for the steam / water mixture are slightly higher than air velocity peaking factors. The difference between the steam / water and air peaking facters 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[ ~ l is determined for the R9C51 tube. The differences between j
'**kypical values are shown in Table 8 2. These results show that the direct 1
clawas/entes n 8-10
application of the air test data yields the higher relative peaking factor compared to R9C51. To obtain conservatism in the peaking factor evaluation, the' - . a , c' 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 , kher than the air velocity peaking factor. On the average, the uncertainty ~ associated with the conservative use of air velocity peaking factor is , The conclusion that 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 protypic flow conditions. The objective of this study was to examine the magnitude of the changes in void fraction 4 and thus stability ratio as a consequence of non-uniform AVB insertion patterns. The current version of ATHOS 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. Consequently, 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 AVB's in the ATHOS model. The ATHOS results were processed by the FLOVIB code to determine stability ratios for the specific tubes of interest. The calculation was
- repeated using a non-unifom AVB 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 [ [.Thisalternate calculation provides independent corroboration of the prior discussion regarding the stability peaking factors under steam water conditions vs in air.
014H:49/C31184 33 8-11
l l I 8.6 AVB Insertion Depth Uncertainty The most significant uncertainty for the low peaking configurations is not in the test results, but in the determination of actual AVB insertion patterns adjacent to specific tubes. The methodology used for obtaining the AVB insertion patterns from eddy current data can ascertain the AVB location only to within approximately[ , [ The effect on peaking factor resulting from this uncertainty is addressed using test results of AVB configurations that varied from one another by up l '" to . Based on maps of AVB insertion depth of various plants, several configurations have been tested for determining fluideiastic instability flow rate by an air cantilever model. Stability peaking factors were then determined from the ratio of critical flow rate for a uniform AVB insertion configuration to a specific configuration. Figure 8-7 sumarizes the AVB - configurations tested. 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 change of an AVB 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 sumarized in Table 8 3. As can be seen, a decrease in depth of an appropriate AVB tends to decrease the peaking factor, for l
instance, a I,*'%uch 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.
On the average, the uncertainty in peaking factor resulting from small variations in AYB insertion (of the order of 1/2 tube pitch) is found to be a a a,c i 0145M:49/031168 34
i 8.7 Overall Peaking Factor with Uncertainty As discussed in the previous subsections, there are several aspects to be considered in applying the laboratory test data to steam generator f conditions. These considerations were reviewed one at a time in thdse l subsections. This section will integrate the pieces into one set of l stability peaking factors. i 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 i factor of the specific tube to that of the North Anna R9C51 tube (Configurationla). It is to be noted that, of all the configurations tested, configuration Ib produced the highest peaking factor, followed very l l closely by 4c, la and 5e. This is encouraging in the sense that it tends to explain why, of all the tubes in service, the R9C51 tube was the one to experience the fatigue rupture. It is to be noted that the test results would be applied as ratios of a l specific tttbe peakinc factor to the R9C51 peaking factor. This Will reduce ! 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 R9C51 configuration. The uncertainties in the test results and their extrapolation 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 ; i I i ; 8145M:49/C311st 35 ; ! 8-13 ;
l 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 are random errors and can occur in any test. Therefore, these are treated i together. The total random uncertainties are calculated by (
)* ' CThe RSS value of these is[ ,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 the denominator of the final pe'aking factor ratio. Thus the peaking factor for configuration la (R9g) is reduced by this amount to yield a value of( [
instead of the ,
) appearing in Table 5-2.
The next three uncertainties in Table 8-3 are systematic uncertainties. It could be argued that these appear in the peaking factors of both the spec.ific tube under consideration and the R9C51 tube and are therefore counter - 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, similarly, as noted above, the effect on peaking factor due to the uncertainty in the field AVB configuration is also included in this reference case. Thus, 7'Shepeakingfactor
~
of the reference configur,ation 2a (Table 8 5) is raised by his amount to a value of , , The change in peaking factors of configurations la and 2a resulting from the application of uncertainties as described above are shown in Column 3 of i Table 8-4. 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 4. j 014!M:49/031168 36 8 14
Someoftheuncertaintieswereappliedtothereferenceconfiguration(2a)in order to apply them to all low peaking configurations conservatively. Thus, i no configuration should have a lower peaking factor than this reference l configuration. Therefore, when a peaking factor val e less than( [i!* calculated forjgcconfiguration, (in Column 4 of Table 8-5), it should be altered to , , further, for some of the configurations that are conceptually similar, the,mgre limiting (higher) value is used. For example, a peaking factor of ~ is used for configurations 5a and 5b based on their similarity to configuration Sc. The final stability ratio peaking factors cair.ulated on this basis (with configuration 2a as the reference) are shown in Table 8-6. It may be noted that the peaking fa tors vary in the range [ ['theR9C51 peaking e factor being ,
,, Figure 8-7 shows the final peaking factors with the pictorial representation of the AVB insertion patterns.
Table 8 7 shows the result of applying the peaking factors to specific tubes in the Beaver Valley Unit 1 steam generators.
- The overall, conclusions from the peaking factor assessment are
l l
- 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 in sections 8.2 to 8.6. The largest single uncertainty of ]is ,
attributable to uncertainties of up to ,
*d8 determination of AVB insertion depths from field addy current data. Thi.t relatively large uncertainty is applicable only to low peaking conditions where the AVB uncertainties can contribute to small peaking factors. The l
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, AVB's were positioned within uncertainties to maximize the peaking factor. For these configurations, l
i variations of AVB insertion within these uncertainties are expected to reduce the peaking factor compared to the final values of Table 8-6 and t Figure 8 7. ci m :o /esti u 37 8-15
- 2. Including uncertainties directed toward conservatively decreasing the peaking factor fjog,the North Anna tube R9C51, the final R9C51 peaking
. factor is , , relative to a no flow peaking condition such as with uniform AVB insertion depths.
- 3. The final peaking factors include peaking effects greater than the R9C51 tube (such as configuration 4c) although this is believed to be a consequence of the conservative uncertainty analysis and is not likely to be representative of actual peaking effects.
4 I 1 l l l c14 m as/e m sa as 8-16
Table 8 1 Stability Peaking Factor Due to Local Velocity Perturbation l Scaling Factors for Steam / Water Air Velocity Void Stability Peaking Fraction Density Velocity Damping Peaking Factor, Scaling, Scaling, Scaling, Scaling, Factor, Fy Fy F F Fg Fd dp s T a,e li
~
! NOTE: 1. Stability peaking factor for steam / water mixture is I calculated as follows: , a,e
- 2. Damping scaling factor is calculated
. . a ,e using modal effective void fraction of for R9C51 tube.
i i 1 k t i i D1494:49/031168 39 8-17 :
1 1 Table 8 2 I Comparison of Air and Stets water Peeking Factor Ratios i t
- Air Air Staus Steam i Peaking Peaking peaking Peaking Factor Ratio Factor Ratio
' a,c d
I i 1 [ l l i , i i i 0845m:49/C311 M 40 gg
Table B-3 Effect of Local Variation of AVB Insertion A to B AVB Peaking Peaking Ratio Type A Type B Variation Factor A Factor B (8/A) 5b 5a 4a 5c 5c Sa 5b a,e Sa , 5c 4a , 5a 5c , l ' l 4 i I 6 i i i j l 1 1 ! i I i , I F l , I i 014M:49/031118-41 1 8-19
i Table 8 4 Uncertainties in Test Data and Extrapolation Source of Uncertainty lygg Macaitude. % a,e
- 1. Velocity measurement Random
- 2. Test repeatability Random
- 3. Cantilever vs U-tube Systematic
- 4. Air vs steam-water mixture Systematic 1 5. Field AVB configuration i , .
l l I
- This is not an uncertainty associated with the test data.
It results f' rom the inaccuracy in determining the true AVB , position in the field using eddy current data. l t
?
4 c14 % 49/c31184 42 g,g
Table 8-5 Extrapolation of Test Results to Steam Generator Conditions Peaking Factor Test Data with Referenced to Confiauration QAla Uncertainties Confia. 2a a ,b c 3, lb l 2a 3 4a i I 4b i s ! 4C . i I 4d i 4f i i i 5a ! Sb 1 i ! Sc ! 6c 1 l ! I P j l l 1 ) J C14W:49/0311M-43 8 8-21
Table 8 6 FINAL PEAXING FACTORS Confiauration Peakina Factor a,b.c lb i 2a 3 i l l 4a 4b 4c
, 4d .
l Af , i , l I Sa l 5b 5e [ l .
; 6c
! L - l 1 l i i i
\
l i l l . 1 l 1 l 1 a 0145M:49/0311 H 44
! 8-22 i r
t i Table 8 7 g Stability Peaking Factors for Specific Tubes BEAVER VALLEY UNIT 1 STEAM GENERATOR R0W NO. COLUMN NO. PEAKING FACTOR
- t I A 8 60 a,b,e 8 61 9 35 10 43 ALL OF THE REMAINING B 9 35 9 60 i ALL OF THE REMAINING C 9 60 i 10 60 11 47 11 51 11 52 ,
! 12 51 ALL OF THE REMAINING h
- The peaking factor is divided by 1.47 to obtain the relative flow peaking factor to R9C51 of North Anna 1.
i I l i i 014 9 :49/03!684-45
.___._ ... 8-23. _ _ _ ,
1 1
~
l 0000 OO u 0000 OO l 'o OOOO @@ 00000@O O@@@@
- 0000@@@@O0000 50 49 48 47 46 45 44 56 55 54 53 52 51 n AVB AVB @ FAILED TUBE V VISIBLE b INVISIBLE O PLUGGED Figure 8-1 Original North Anna AVB Configuration 0142M:49/030788 49 I 8-24
i
. l l
l i 1
%a, s en ts ' '
1 ,e "M " ' -
, , 5 nj I*tli.>, .
ap,,.' -E w y c. 3 <> ~ i
~ )(
- fs --a '
ps V 7.?
.313,
_ fi.u.PA r i l t I Figure 8-2 Schernatic of Staggered AVBs 0142M:49/030788 50 B-25
i i i
.,,. i t
. .- o
< l t
*
- 5,C l
t I-
) i t
I t I i i t h 6 l-i i
> . . ,4 1
l l i i . I I . i, I' i i ) 4 : ! l i 4 i t i j
. i s !
l i : l t
- i l,
i i Figure 8-3 AVB 'Patr' in ECT Trace l : i i i ! op;M:49/030788 51 8-26 t
l l l Row h 22 'ogeoegeggoooo . it OOGOOOGOOOOOO '
" 0000000000000 .
00000@O000000 8 0000000000000 Column 56 55 54 53 52 51 50 49 48 47 46 45 44 g Plugged Tube g Faded Tube l L",Ta a*5a'*2'.'2125i"C'f4"s2 ^.7,'.7o'.12"e'fai avai'.ab's . Figure 8-4 North Anna 1. Steam Generator C. AVB Positions , critical Review 'AVB Visible' Calls 0142M:49/0307SS 52 l
i l l
. m.. .... . M.
H E R m H ... H
.... .... H. ..H. ..!. ...
n
~
ab \s m M h Ri E M.rs.. E W E re M e
\ \s m e sc s
\ s
\ e s nh
\ aa an - - .... ... .r. '.. ..,. :l;,, ... .., ,.., . . , l , , , . . ..
m x. u a. a 7:< z a-e e
~~
m 6 sa .* u u x . ._ I
~
m- - E
.. . k'A$ $E i$ 552 ER
... . . , . . .e r . , , =, .. v P IN.$.
u u 4v 4v 43 9 8 M E E H EE H EB
. .. a. :e. , .v., 2. 2. 2. 2. 2. 2. .. .
=. s C55 C54 C53 C52 C51 C50 C49 C48
- Preseasy upper An f,',Unreemwe U
. . SeC SeC
'Lme sW.' .ggg sa.. {
}
l Figure 8 5 North Anna 1, Steam Generator
, . a.c C,
R9C51 AVB , , Hatrix 0142M:49/030785 53 8-28
' 'N N N l4 4
0 00 0 0 0 00 0 0 0 0 1 12 it O 00 0 0 0 00 0 0 0 0 to O OO O O O OO O.,0 O..O.. .. 9 0 00 0,0 @ O 0 0 0 0 0 0
. O'000000000000 o
i i t i Figure 8-6 North Anna R9C51 AVB Final p'ositions 0142M:49/030788 54 8-29
i l PEAKING TYPE OF AVB PEAKING TYPE OF AVB INSERTON FACTOR INSERM FACTOR
- a,b,e '- - a ,b ,c 1a 4d 1b 4f 2a 5a 3 Sb 4a Sc 4b Sc
~
4c Figure 8 7 FINAL PEAKING FACTOR FOR BEAVER VALLEY 1 8-30
O 9.0 STRUCTURAL AND TUBE VIBRATION ASSESSMENTS 9.1 Tube Mean Stress This section summarizes an analysis to determine stresses in a dented 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[ 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 , [,
- The analysis is performed using the finite element model shown in Figure 9-1. The model prescribes [
- a,c Two reference cases were run using the finite element model, the first for aprimary-to-secondarysidepressuregradient,andthesecondfora[ [
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. Tube stresses due to the thru-wall thermal gradient are calculated to be 9.8 ksi using conventional analysis techniques. A plot showing the combined stress distribution along the tube length, incorporating appropriate scale factors for the Beaver Valley Unit 1 operation conditions, is provided in Figure 9-4. The maximum axial tensile stress is 23.1 ksi and occurs approximately 0.125 inch above the top surface of the support plate. Adding, for conservatism, the surface stress 0145M:49/031788 46 9-1
f due to pressure, 0.75 ksi, gives an applied mean stress of 23.3 ksi. In addition to the applied stress, residual stresses exist in the tube as a result of the manufacturing process. For mill annealed tubes with subsequent straightening and polishing, residual stresses exist in the tube. The stresses are compressive at the tube surface, but 5-10 mils below the surface, the stress levels change to 10-15 ksi tensile, Reference 9-1. Combining the applied and residual stresses results in a cumulative mean stress of 38.9 ksi. 9.2 Stability Ratio Distribution Based Upon ATH0S An assessment of the potential for tubes to experience fluid elastic ~ instability in the U-bend region was performed for each of the tubes in rows eight through twelve. This was performed using FASTVIB, a Westinghouse proprietary finite element based computer code and PLOTVIB, a post processor to FASTVIB. These codes were written to predict the individual responses of an entire row of steam generator tubing exposed to a tube location dependent fluid velocity and density profile. The program calculates tube natureal frequencies and mode shapes using a linear finite element model of the tube. The fluid elastic stability ratio V,/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 conditin. The velocity, density and void fraction distributions are determined using the ATH0S computer code, as described in Section 7.3. Also input to the code are the WECAN generated mass and stiffness matrices used to represent the tube. (WECAN is also a Westinghouse proprietary computer code.) Additional input to FASTVIB/PLOTVIB consists of tube support conditions, fluid elastic stability constant and turbulence constants. This process was performed for the Beaver Valley Unit 1 steam generator tubes and also for the North Anna Row 9 Column 51 tube (R9C51) using similarly appropriate ATHOS models. Ratios of the Beaver Valley Unit I 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 Beaver Valley Unit 1 tubes relative to the ruptured tube at North Anna Unit 1. 0145M:49/031788 47 9-2
l Figure 9-5 contains the results of this process for each of the rows under investigation. This figure was generated using the following conditions for both Beaver Valley Unit 1 and Horth Anna Unit 1:
- 1) Tube is fixed at top tube support plate.
- 2) Void fraction dependent damping used.
- 3) No AVB supports active.
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 R9C51. (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. This figure indicates that all tubes in Row 8 and 9 can be considered acceptable with some tubes being unacceptable in Row 10. Essentially all inner tubes in Rows 11 and 12 lay above this line. Tubes above this line, that are not supported by AVBs, require further evaluation to determine the acceptability of the tubes. Section 9.3 contains the results of the further evaluation for these tubes. 9.3 Stress Ratio Distribution With Flow Peaking An evaluation was performed to determine the ratio of the Beaver Valley Unit 1 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 1.47) and bending moment factors. Sections 4.2 and 4.3 contain additional information and describe the calculational procedure used to obtain the results presented in this section. The results contained in this section are based upon the following conditions:
- 1) Tube is fixed in top tube support plate.
- 2) Damping is void fraction dependent.
c14sm49/en4 s 4s 9-3
- 3) AVBs do not provide support.
- 4) 10% criteria with frequency effects used.
- 5) The tubes are assumed to be dented or undented.
- 6) Flow peaking factors are used.
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. Using this criteria indicates that the stress acting on a given tube will not produce a fatigue event in a manner similar to the rupture that occurred in the R9C51 tube at North Anna Unit 1. Figure 9-6 contains the stress ratio results for each of the Beaver Valley Unit 1 tubes in Rows 8 through 12. This figure is applicable for tubes that are dented (tube deformation) at the top tube supp. ort plate. As can be observed in the figure, the tubes in Rows 8 and 9 fall below the acceptance line indicating that the tubes are acceptable with respect to U-bend fatigue. All of the tubes in Row 10, except for Column 35, and all of the tubes in Row 11, except for the Column 43, 44, and 47 also fall below the acceptance line. The tubes in these two rows (10 and 11) that do not lay below the acceptance line have large stress ratios due to the large peaking factors associated with these tubes. Most of the tubes in row 12, except for tubes in column 3, 4 and 34 through 47, also fall below the acceptance line. This indicates that the tubes are not acceptable, with respect to U-bend fatigue, if the tubes are not supported by AVBs and are dented at the top tube support plate. Figure 9-7 contains results similar to the information presented in the previous figure except that the tube is assumed to be undented, but clamped due to support plate corrosion, at the top tube support plate. The conclusions that can be drawn from this figure are almost identical to the conclusions that can be drawn from the previous figure. This figure indicates that all of the tubes in rows 8 and 9 fall below the acceptance line indicating that the tubes are acceptable with respect to U-bend fatigue. All of the tubes in row 10, except for column 35, and all of the tubes in row 11, except for column 43, 44, and 47 also fall below the acceptance line. The tubes in these two row (10 and
- 11) that do not lay below the acceptance line have large stress ratios due to 0145M:49/032886-49
the large peaking factors associated with these tubes. Most of the tubes in row 12, except for tubes in column 3, 4 and 36 through 47, also fall below the acceptance line. This indicates that the tubes are not considered acceptable, with respect to U-bend f atigue, if the tubes are not supported by AVB and are firmly fixed, but not dented, at the top tube support plate. 9.4 Cumulative Fatigue Usage All tubes that are unsupported and have a stress ratio 1 1.0 have a maximum stress amplitude that is < 4.0 ksi (from 9.5 ksi) since a 10% reduction in the stability ratio for the North Anna Row 9 Column 51 tube was the criteria basis. The stability ratios for the Beaver Valley Unit 1 tubing are based on the current operating parameters and with future operation on the same basis, the tubes will not rupture as a result of fatigue if 1) they meet the stress ratio criteria of 11.0 and 2) their current and future fatigue usage will total less than 1.0. All tubes in the evaluation have been considered to be dented. Table 9-2 contains a summary of the combined relative stability ratios and the stress ratios for the most critical tubes in each of the steam generators. With the ex:eption of five tubes in Steam Generator C, all tubes have relative stability ratios less than 0.9 and stress ratios less than 1.0. The five tubes exceeding these criteria are R12C51, R11C52, R11C51, R11C47 and R10C60 in Steam Generator C. Six tubes have been plugged as a preventive measure to eliminate them from further consideration. Sentinel, r ' ell-tale, plugs were used to One permit detection of tube degradation if it sho': 4 occur in the future. tube in Steam Generator A R10C43, was plugged based on preliminary results. Acceptability of the Beaver Valley Unit 1 tubing for fatigue is accomplished by demonstrating the acceptability of the tubes with the highest stress ratio, 0.77, at Row 9 Column 35 and Column 60 of Steam Generator B and Row 9 Column 60 of Steam Generator C. The maximum stress with the current operating conditions is , aa = 0.77 (4.0) = 3,08 ksi 0145M:49/032888 50 9-5 _
Based on the relative stability ratio over the operating history presented in Section 7.5, the alternating stress for each operating cycle can be determined. Computing stress ratios for each cycle establishes the maximum alterneting stress for each operating cycle. The number of cycles of vibration is obtained for each fuel cycle by multiplying the number of days tjmes the number of cycles per day at the frequency of the Row 9 tube, , ) hertz. Conservatively assuming that all fuel cycles have been at 3.08 ksi, the cumulative fatigue usage for the operating license period (with future operational years at the same operating conditions as cycle 6) would be 0.346. All of the unplugged Beaver Valley Unit 1 tubes, therefore, meet the fatigue 5 usage requirement of 1.0 with denting assumed to exist from initial startup. 0145M:49/032488 51 g,g
Table 9-1 100% Power Operating Parameters Beaver Valley Unit 1 Primary Pressure = 2235 psi Secondary Pressure = 799 psi Pressure Gradient = 4360 psi Primary Side Temperature - 576*F Secondary Side Temperature = 518'F. Tube Temperature = 547'F c14sm:4s/e m sa sz g,7
Table 9-2 Beaver Valley Unit 1 Evaluation of the Salient Unsupported U-bends RELATIVE STRESS STEAM GENERATOR IMai STABILITY RATIO (1) B&I1Q(I) A R9C35 0.73 0.31 R9C47 0.66 0.18 R10C43 0.89 0.81 R10C47 0.77 0.35 R11C49 0.88 0.64 B R9C47 0.66 0.18 R9060 0.86 0.77 R9C35 0.86 0.77 C R9C60 0.86 0.77 R10C47 0.77 0.35 R10C60 1.081 >>1.0 R11C47 1.272 >>1.0 R11C51 1.454 >>1.0 R11C52 1.454 >>1.0 R12C51 >1.12 ,
>>1.0 (1) All ratios are in comparison to R9C51, North Anna 1, Steam Generator C.
c14sm:sslosuse 53 98
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a,c Figure 9-1 Axisyneetric Tube Finite Element Model 0145M:49/030888-116 g,9
a a,c l Figure 9-2 Dented Tube Stress Distributions l Pressure Load on Tube 0145M:49/030484 117 9-10
'a,c a
b ) Figure 9 3 Dented Tube Stress Distributions Interferenes Load on Tube 0145M:49/030888 ll8 9-11
- ~
a,c Figure 9-4 Dented Tube stress Distributions combined Stress Results leaver Valley Unit 1 0145M:49/030488 119 9-12
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Figure 9 5 Relative Stability Ratios Using MEYF Dependent Damping 0145M:49/030888 120 9,)3
a,e Figure 9 6 Stress Ratio vs. Column Number - Dented Condition
- 0145M:49/030888 121 9-14
,, a,C
.I Figure 9 7 Stress Ratio vs. Colan 10mber - Undented Condition 0145M:49/030448 Itt 9-15
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