ML20154R330
| ML20154R330 | |
| Person / Time | |
|---|---|
| Site: | Zion File:ZionSolutions icon.png |
| Issue date: | 08/31/1988 |
| From: | Pitterle T WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML20154R325 | List: |
| References | |
| WCAP-11949, NUDOCS 8810040226 | |
| Download: ML20154R330 (167) | |
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wCAP-11949 L
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ZION UNITS 1 AND 2 EVALUATION FOR TUBE VIBMTION INDUCED FATIGUE AUGUST 1988 i
r AUTHORS: H. J. CONNORS M. H. HU T. M. FRICK A. Y. LEE
) J. M. HALL M. R. PATEL l G. W. HOPKINS R. M. WILSON J. L. HOUTMAN H. W. YANT f
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APPROVED: Id. 8/N/h T. A. PITTERLE, MANAGER l
STEAM GENERATOR ENGINEERING This document contains information proprietary to Westinghouse Electric Corporation. It is submitted in confidence and is to be used solely for the purpose for which it is furnished and is to be returned upon request. This document and such information is not to be reproduced, transmitted, disclosed or used otherwise in whole or in part without written authorization of j
> Westinghouse Electric Corporation, Power Systems Business Unit.
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l WESTINGHOUSE ELECTRIC CORPORATION SERVICE TECHNOLOGY O! VISION 0010040226 DR 8e692 o ADOCK 0500 5 P.0, BOX 3377 PITTSBURGH, PENNSYLVANIA 15230
ABSTRACT t
j On Juiy 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
- high cycle fatigue. The source of the loads associated with the fatigue k
mechanism is a combination of a mean stress level in the tube with a superimposed alternating stress. The mean stress is the result of denting of l the tube at the top tube support plate, while the alternating stress is due to out-of plank deflection of the tube U bend attributed to flow induced l vibration. For tubes without AVB support, local flow peaking effects at unrupported tubes are a significant contribution to tube vibration amplitudes.
This report documents the evaluation of steam generator tubing at Zion Units 1 and 2 for susceptibility to fatigue induced cracking of'the type experienced at North Anna Unit 1. The evaluation utilizes operating conditions specific to Zion Units 1 and 2 to account for the plant specific nature of the tube loading l
and response. The evaluation also includes reviews of eddy current data for Zion Units 1 and 2 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 shewing AVB positions, thermal hydraulic analysis results, and calculations to determine tube mean stress, stability ratio and tube !, tress distributions, and accumulated fatigue usage. This evaluation concludes that none of the tubes potentially suscaptible to fatigue require corrective action, l
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$UMMARY 0F ASSREVIATIONS ,
ASME - American Society of Mechanical Engineers ATHOS - Analysis of the Thermal Hydraulics of Steam Generators AVB - Anti-Vibration Bar t AVT - All Volatile Treatment l ECT - Eddy Current Test I
EPRI - Electric Power Research Institute FFT - Fast Fourier Transform FLOVIB - Flow Induced Vibrations HEVF - Modal Effective Void Fraction 00 - Outside Diameter i
PMS - Root Mean Square SR - Stability Ratio TSP - Tut,e Support Plate i
'F - degrees Fahrenheit t hr -
hour j ksi - measure of stress - 1000 pounds per square inch !
lb - pound f mils - 0.001 inch MW - mega watt psi - measure of stress - pounds per square inch !
psia - measure of pressure absolute ;
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TABLE OF CONTENTS
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SECTION 1.0 Introduction l 2.0 Sumary and Conclusions j
2.1 Background
2.2 Evaluation Criteria 2.3 Denting Evaluation ;
4 2.4 AVB Insertion Depths
) 2.5 Flow Peaking Factors ,
I 2.6 Tube Vibratic:: Evaluation
- 2.7 Prior Fatigue Usage l 2.8 Overall Conclusion i
1 3.0 Background i 3.1 North Anna Unit 1 Tube Rupture Event j
, 3.2 Tube Examination Results 3.3 Mechanisn Assessment j 4.0 Criteria for Fatigue Assessment 4.1 Stability Ratio Reduction Criteria !
! 4.2 Local Flow Peaking Consideration l 4.3 Stress Ratio Considerations 1 !
5.0 Supporting Test Data 5.1 Stability Ratio Parameters 5.2 Tube Damping Data 5.3 Tube Vibration Amplitudes with Single-Sided AVB Support 1 5.4 Tests to Determine the Effects on Fluidelastic Instability of
.; Columnwise Variations in AVB Insertion Depths 1 5.5 References l
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TABLEOFCONTENTS(CONTINUED) i SECTION 6.0 Eddy Current Data and AVB Positions 6.1 AVB Assembly Design 6.2 Eddy Current Data for AVB Positions 6.3 Tube Denting at Top Tube Support Plate 6.4 AVB Map Interpretations ,
r a 7.0 Thermal Hydraulic Analysis 7.1 Zion Steam Generator Operating Conditions 7.2 ATHOS Analysis Model 7.3 ATHOS Results 7.4 Relative Stability Ratio Over Operating History ;
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8.0 Peaking Factor Evaluation 8.1 North Anna 1 Configuration 8.2 Test Measurement Uncertainties
- 8.3 Test Repeatability 8.4 Cantilever vs U-Tube ,
8.5 Air vs Steam Water Mixture 8.6 AVB Insertion Depth Uncertainty ,
8.7 Overall Peaking Factor with Uncertainty :
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- 9.0 Structural and Tube Vibration Assessments j
! 3.1 Tube Menn Stress ,
9.2 Stability Ratio Distribution Based Upon ATHOS f 9.3 Stress Ratio Distribution with Peaking Factor 9.4 Cumulative Fatigue Usage 1 l 4
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4 LIST 0F FIGURES FIGURE ,
i 3-1 Approximate Mapping of Fracture Surface of Tube R9C51 S/G "C" Cold Leg, North Anna Unit 1 3-2 Schematic Representation of Features Observed During TEM Fractographic Examination of Fracture Surface of Tube R9C51, S/G "C" Cold Leg, North Anna Unit I d
33 Calculated and Observed Leak Rates Versus Time f 4-1 Vibration Displacement vs. Stability Ratio 42 Fatigue Strength of Inconel 600 in AVT Water at 600'F 4-3 Fatigue Curve for Inconel 600 in AVT Water Comparison of Mean Stress i Correction Models 4-4 Modified Fatigue with 10% Reduction in Stability Ratio for Maximum Stress Condition 45 Modified Fatigue with 5% Reduction in Stability Ratio for Minimua Stress Condition 5-1 Fluidelastic Instability Uncertainty Assessment ,
l 52 Instability Constant - A !
5-3 Instability Constants, A, Obtained for Curved Tube from Wind Tunnel Tests on the 0.214 Scale U Bend Model ,
9 54 Damping vs. Slip Void Fraction 1 l 0221M:49/072888 7 ,
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LIST 0FFIGURES(Continued)
FIGURE 55 Overall View of Cantilever Tube Wind Tunnel Model 56 Top View of the Cantilever Tube Wind Tunnel Model 57 Fluidelastic Vibration Amplitude with Non-Uniform Gaps 58 Typical Vibration Amplitude and Tube /AVB Impact Force Signals for Fluidelastic Vibration with Unequal Tube /AVB Gaps 59 Conceptual Design of the Apptratus for Determining the Effects on Fluidelastic Instability of Columnwise Variations in AVB Insertion Depths i
1 5 10 Overall View of Wind Tunnel Test Apparatus 1 5 11 Side View of Wind Tunnel Apparatus with Cover Plates Removed to Show Simulated AVBS and Top Flow Screen i
1 5 12 AVB Configurations Tested j 5 13 Typical Variation of RMS Vibration Amplitude with Flow Velocity for
.l Configuratin la in Figure 5 12 i
61 AVB Insertion Depth Confirmation i
62 Zion Unit 1 - Steam Generator 'A' AVB Positions
, 63 Zion Unit 1 - Steam Generator 'B' - AVB Positions >
64 Zion Unit 1 - Steam Generator 'C' - AVB Positions l
65 Zion Unit 1 Steam Generater 'D' - AVB Positions 0221M:49/072888 8
LIST 0FFIGURES(Continued)
FIGURE 6-6 Zion Unit 2 - Steam Generator 'A' - AVB Positions 6-7 Zion Unit 2 Steam Generator 'B' - AVB Positions 68 Zion Unit 2 Steam Generator 'C' - AVB Positions 6-9 Zion Unit 2 Steam Generator 'O' - AVB Positions 6 10 Replacement AVB Proximity ECT Bench Test AVB Normal to Tube 6-11 Replacement AVB Proximity ECT Bench Test AVB Skewed to Tube Centerline 6 12 Theoretical Arc Lengths (AVB Normal to Tube) 6-13 R11: Actual vs Theoretical Arc Between AVB Legs Centerlines 6-14 R11: Actual vs Theoretical Arc Lengths; AVB Skewed Relative to the Tube Centerline 7-1 Plan View of ATHOS Cartesian Model for Zion 7-2 Elevation View of ATHOS Cartesian Model for Zion 7-3 Plan View of ATH0S Cartesian Model Indicating Tube layout 74 Flow Pattern on Vertical Plane of Symmetry 7-5 Lateral Flow Pattern on Horizontal Plane in the U Bend Region 7-6 Lateral Flow Pattern on Top of Tubesheet 0221M:49/072888 9
LISTOFFIGURES(Continued)
FIGURE 7-7 Tube Gap Velocity and Density Distributions for Tube at R10/C3 78 Tube Gap Velocity and Density Distributions for Tube at R10/C20 7-9 Tube Gap Velocity and Density Distributions for Tube at R10/C40 7-10 Average Velocity and Density in the Plane of the U-Bends Normal to Row 10 1
7-11 Zion Normalized Stability Ratio Based on l
- HighPower(>86%) Operation i
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.; 8-1 Original North Anna AVB Configuration l I
i 82 Schematic of Staggered AVBs l l
i 84 AVB "Patr" in ECT Trace f j 8-4 North Anna 1, Steam Generator C: AVB Positions Critical l Review "AVB Visible" Calls i
j 85 North Anna 1, Steam Generator C, R9C51 Projection Matrix l f
86 North Anna R9C51 AVB Final Projected Positions 1
87 Final Peaking Factors for Zion 1 and 2 I 9-1 Axisymetric Tube Finite Element Model f 92 Dented Tube Stress Distributions - Pressure load on Tube l
9-3 Dented Tube Stress Distributions - Interference Load on Tube I
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LIST 0FFIGURES(Continued)
FIGURE 94 Dented Tube Stress Distributions - Combined Stress Results Zion Units 1 and 2
) I j 95 Relative Stability Ratio Using MEVF Dependent Damping - Zion Unit 1 1 i 96 Relative Stability Ratio Using MEVF Dependent Damping - Zion Unit 2 97 Stress Ratio Vs. Column Number Dented Condition Zion Unit 1 a 98 Stress Ratio Vs. Column Number Undented Condtion Zion Unit 2 l
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l LIST OF TABLES l l
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, BELE 41 Fatigue Usage per Year Resulting From Stability Ratio Reduction J
5-1 Wind Tunnel Test on Cantilever Tube Model 52 Fluidelastic Instability Peaking Velocity Factors for Columnwise Variations in AVB Insertion Depths 61 Resolution of Tube Support-Tubes with Single AVB Indications l 62 Sumary Listing of Unsupported Tubes - Zion Unit 1 t
63 Sumary Listing of Unsupported Tubes - Zion Unit 2 t
71 Zion Steam Generator Operating Conditions Used for ATH0S Analysis 72 Zion Operating History Data 81 Stability Peaking Factor Due to Local Velocity Perturbation i !
i 82 Comparison of Air and Steam Water Peaking Factor Ratios 83 Effect of Local Variation of AVB Insertion f 84 Uncertainties in Test Data and Extrapolation ,
- 85 Extrapolation of Test Results to Steam Generator Conditions 86 Final Peaking Factors 8-7 Stability Peaking Factors for Specific Tubes I
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LIST 0FTABLES(Continued) l -
Ist t j 91 100% Power Operating Parameters 2
1 l 92 Sunnary of Zion Evaluation of the More Salient Unsupported j i U Bends 93 Disposition Criteria Relative to Tube U Bend Fatigue [
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1.0 INTRODUCTION
This report documents the evaluation of steam generator tubing at Zion 1 and 2 for susceptibility to fatigue-induced cracking of the type experienced at North Anna Unit 1 in July, 1987. The evaluation includes three-dimensional flow analysis of the tube bundle, air tests performed to support the vibration analytical procedure, field measurements to establish AVB locations, structural and vibration analysis of selected tubes, and fatigue usage calculations to predict cumulative usage for critical tubes. The evaluation utilizes operating conditions specific to Zion I and 2 in order to account fer plant specific 4 features of the tube loading and response.
l Section 2 of the report provides a sumary of the Zion 1 and 2 evaluation results and overall conclusions. Section 3 provides background for the tube l 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 l
provides a sumary of test data which supports the analytical vibration l
l evaluation of the candidate tubes. A sumary of field measurements used to ,
determine AVB locations and to identify unsupported tubes is provided in
! Section 6. Section 7 provides the results of a thermal hydraulic analysis to l l
l establish flow field characteristics at the top support plate which are a l subsequently used to assist in identifying tubes which may be dynamically 2
unstable. Section 8 presents an update of the methodology originally used to
) evaluate the tube rupture at North Anna Unit 1. The final section, Section 9, 4 presents results of the structural and vibration assessment. This section f describes tube mean stress, stability ratio and tube stress distributions, and j accumulated fatigue usage, for the Zion 1 and 2 steam generator small radius i U tubes. l I
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2.0 SUMARY AND CONCLUSIONS The Zion Unit 1 and 2 steam generators have been evaluated for the susceptibility of unsupported U bend tubing with denting at the top tube support plate to a fatigue rupture of the type experienced at Row 9 Column 51 The evaluation used Eddy (R9C51) of Steam Generator C at North Anna Unit 1.
Current Test (ECT) data supplied by Commonwealth Edison Company and interpreted by Westinghouse.
2.1 Background
i The initiation of the circumferential crack in the tube at the top of the top i tube support plate at North Anna 1 has been attributed to limited displacement, l
fluid elastic instability. This condition is believed to have prevailed in the R9C51 tube since the tube experienced denting at the support plate. A combination of conditions were present that led to the rupture. The tube was j not supported by an anti vibration bar (AVB), had a higher flow field due to local flow peaking as a result of non uniform insertion depths of AVBs, had
! reduced damping due to denting at the top support plate, and had reduced fatigue properties due to the environment of the all volatile treatment (AVT)
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chemistry of the secondary water and the additional mean stress from the j denting.
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l 2.2 Evaluation Criteria I The criteria established to provide a fatigue usage less than 1.0 for a finite period of time (i.e., 40 years) is a 10% reduction in stability ratio that ,
provides at least a 58% reduction in stress amplitude (to < 4.0 ksi) for a
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j Row 9 tube in the North Anna 1 steam generators (SG's). This reduction is [
required to produce a fatigue usage of < 0.021 per year for a Row 9 tube in !
North Anna and therefore greater than 40 year fatigue life objective. This ,
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same fatigue criteria is applied as the principal criteria in the evaluation of Zion tubing. i l
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l The fluidelastic stability ratio is the ratio of the effective velocity divided by the critical velocity. A value greater than unity (1.0) indicates instability. The stress ratio is the expected stress amplitude in a Zion tube l divided by the stress amplitude for the North Anna 1. R9C51 tube.
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Displacements are computed for the unsupported U bend tubes in Rows 11 and j inward, (descending row number) using relative stability ratios to RbC51 of {
North Anna 1 and an appropriate power law relationship based on instability l
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) displacement versus flow velocity. Different U bend radius tubes will have l
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different stiffness and frequency and, therefore, different stress and fatigue l
! usage per year than the Row 9 North Anna tube. These effects are accounted for in a stress ratio technique. The stress ratio is formulated so that a stress !
ratio of 1.0 or less produces acceptable stress amplitudes and fatigue usage j for the Zion tubing for the reference fuel cycle analyzed. Therefore, a stress :
ratio less than 1.0 provides the next level of acceptance criteria for f unsupported tubes for which the relative stability ratio, including flow ;
i peaking, exceed 0.9.
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The stability ratios for Zion tubing, the corresponding stress and amplitude, i l
l and the resulting cumulative fatigue usage must be evaluated relative to the f
} ruptured tube at Row 9 Column 51, North Anna 1 Steam Generator C, for two i reasons. The local effect on the flow field due to various AVB insertion depths is not within the capability of available analysis techniques and is [
j determined by test as a ratio between two AVB configurations. In addition, an 1 analysis and examination of the ruptured tube at North Anna 1 provided a range
! of initiating stress amplitudes, but could only bound the possible stability i ratios that correspond to these stress amplitudes. Therefore, to minimize the influence of uncertainties, the evaluation of Zion tubing has been based on i relative stability ratios, relative flow peaking factors, and relative stress i ratios. f I f 4 The criteria for establishing that a tube has support from an AVB and therefore !
I eliminate it from further considerations is that it must have at least one I sided AVB support present at the tube centerline. The criteria is based on test results which show that one sided AVB support is sufficient to limit the l
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vibration amplitude for fluideiastic excitation. AVB support is established by analysis of eddy current (EC) measurements and is a key factor in the determining the local flow peaking factors. The local flow peaking produces increased local velocities which cause an increase in stability ratio. A small percentage change in the stability ratio causes a significant change in stress amplitude. The relative flow peaking factors for Zion tubing without direct :
AVB support have been determined by test. These flow peaking factors normalized to the North Anna R9C51 peaking, are applied to relative stability ratios determined by 3-D tube bundle flow analysis, to obtain the combined j relative stability ratio used in the stress ratio determination.
i 2.3 Denting Evaluation i
1 The Eddy Current (EC) tapes were evaluated to determine the condition of the tube / tube support interface for the first row of unsupported tubes inboard of the AVBs. Analyses of eddy current (EC) data for Zion Unit 1 shows the 4 presence of ' corrosion with magnetite' in the majority of the tube / TSP crevices. Analysis o'f EC data for Zion Unit 2 also shows the presence of j
' corrosion with magnetite' in the majority of the tube / TSP crevices. Zion Unit I has a total of 5 ' dent' (denting with deformation) signals in i unsupported tubes imediately in front of the AVBs in the 4 SGs, and Zion Unit 2 has a total of 13 Row 11 tubes in the four SGs with ' dent' signals. For i conservatism in the evaluation, all of the tubes evaluated are postulated as being dented. The effect of denting on the fatigue usage of the tube has been conservatively maximized by assuming the maximum effect of mean stress in the tube fatigue usage evaluation and by incorporating reduced damping in the tube vibration evaluation. ,
l 2.4 AVB Insertion Depths l '
l The Zion Unit 1 SGs have two sets of Alloy 600 AVBs. The ' inner' AVBs have a i rectangular cross section and extend into the tube bundle approximately as far -
as Row 11. They provide a nominal total clearance between a tube without
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) ovality and the surrounding AVBs of ( la.c inch. Including average tube j
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ovality for a Row 11 tube, the nominal total tube to AVB clearance is chout Ja.c inches.
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The outer AVBs, also have a rectangular cross section, and extend into the tube bundle approximately as far as Row 14, providing a nominal tube to AVB clearance comparable to the inner AVBs. Since the purpose of this analysis is to evaluate the potentially unsupported tubes at or near the point of maximum AVB insertion, only the dimensions and EC data pertaining to the inner AVBs is required.
The AVBs originally supplied in Zion Unit 2 have been replaced with a design which uses a pair of 405 stainless steel AVB assemblies. These AVBs also have a rectangular cross section approximately twice the width of the original AVBs. One of these AVB assemblies is located on each side of the U bend
- centerline. Except at the extreme perimeter of the tube bundle, these AVBs normally support the tubes into Row 12, providing 4 AVB signals on a given tube. These AVBs infrequently support Row 11. At the perimeter of the tube bundle, the AVBs between columns 3 and 4 and columns 91 and 92 extend inboard of Row 10, but typically not as far as Row 9. Deviations in the tube diameter l and ovality were accommodated by deflection of the flexible AVBs in some
- instancesandbyverysmallAVBtotubegaps[ ]inothers.
- The overall bundle effect is a zero or small gap condition, with the majority of the tubes lightly preloaded.
i The eddy current data supplied by Comonweslth Edison Company, were reviewed to
- identify the number of tube /AVB intersections and the location of these ,
! intersections relative to the apex of a given tube. This information was used '
in calculations to determine the deepest penetration of a given AVB into the tube bundle. For the area of interest in the Zion 1 steam generators, the AVB :
support of the tube can normally be verified if EC data shows both legs of the lower AVB, one on each side (hot leg cold leg) of the U bend. This data, indicated by a listing of two or more AVBs in the insertion depth plots, is the j method of choice for establishing tube support. For Zion 2 with the i j ' replacement' AVBs, support of a tube in Row 12 and outward can be verified !
with EC data showing all 4 AVB legs. Verification by geometric projection is !
- required with Row 11, as the presence of the 40555 AVBs cah typically be seen I in Row 11 even when the bottom of the AVB is significantly above the Row 11 !
- tube centerline.
0221M: P '082688 18 ;
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l If only the apex of a Zion 1 AVB assembly is near or touching the apex of a tube U-bend, only one AVB signal may be seen. In this case, adequate tube support cannot be assumed without supplemental input. Support can be determined if ' projection' calculations based on AVB intercepts of higher row number tubes in the same and adjacent columns verify insertion depth to a point below the tube centerline. Maps of the AV8 insertion depths for Unit I are j shown in Figures 6 2 thru 6 5. Unit 2 SGs are shown in Figures 6 6 thru 6 9. (
These AVB maps list the results of the ' projection' calculations where this [
infomation contributes to understanding of the AVB insertion depth.
2.5 Flca Peaking Factors 4
Tests were performed modeling Zion, Series 51 53 tube and AVB geometries to I determine the flow peaking factors for various AV8 configurations relative to ,
1 the North Anna R9C51 peaking factor. Tests were performed for the AVB
! geometries used in both Zion units. The test results were used to defini an <
l upper bound of the ratio relative to the R9C51 configuration. It was found i
! that the worst case flow peaking results for either Zion 1 or Zion 2 were j significantly less than for R9C51. ,
l 1 2.6 Tube Vibration Evaluation I The calculation of relative stability ratios for Zion makes use of detailed [
! tube bundle flow field information computed by the ATH0S steam generator [
} thermal / hydraulic analysis code. Code output includes three- dimensional l
) distributions of secondary side velocity, density, and void fraction, along ;
with primary fluid and tube wall temperatures. Distributions of these
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l parameters have been generated for every tube of interest in the Zion tube I 1 bundles based on recent full power operating conditions. This infomation was i factored into the tube vibration analysis leading to the relative stability f
l ratios.
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Relative stability ratios of Zion Unit I and 2 (Row 8 through Row 12) tubing [
l versus R9C51 of North Anna 1 are plotted in Figure 9-5 and 9 6, respectively.
j l These relative stability ratios include relative flow peaking factors. The [
0221M:49/082688 19 f a . - -
- . . - . -- . . _ - _ = _ - . - - _ . _ . - . - - . . - _ - _ .
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stress ratios for Zion Unit 1 are given in Figure 9-7, and the stress ratios f for Zion Unit 2 are given in Figure 9 8. These also include the relative flow l
l peaking effect, and are calculated based on clamped tube conditions with denting at the tube support plate. For all eight steam generators, the stress f
i ratios for all tubes in Rows 8 through 12 are less than 1.0, even when the I tubes are assumed to be unsupported.
, I i A summary listing of the unsupported critical tubes evaluated is given in j 1 Table 9 2. For Unit 1 tubing, the maximum relative stability ratio and stress j ratio occur in SG C:R9C60, and for Unit 2 tubing, in SG D:R11C89.
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" r For the limiting tube, SG C:R9C60 in Unit 1, the maximum cumulative fatigue usage for a 40 year operating period is calculated to be 0.04. Since this is }
1ess than 1.0, all analyzed tubing in Unit 1 is acceptable for continued [
1 service. The fatigue calculation utilized plant operating history to date and {
! assumed future operation at 100% availability with current fuel cycle l parameters. {
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A similar calculation for the limiting tube, SG 0:R11C89, in Unit 2 results in ;
a total usage of 0.014. However, because of AVB replacements in Unit 2 SGs, j the fatigue usage prior to the modification cannot be calculated precisely l since the old AVB positions and associated peaking factors were not determined. (In the calculation, they were assumed to be the same as for the f currer.t configuration). However, based on considerations relative to prior [
fatigue usage below, in conjunction with the results of vibration analyses, all j i Unit 2 tubes are judged acceptable for continued service. [
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2.7 Prior Fatigue Usage With the exception of the two end tubes in Row 12 (R12C2 and R12C93), which l
have a very small stress ratio and hence a negligible future fatigue usage.
only Row 11 and smaller tubes in Unit 2 are unsupported and subject to future fatigue usage. These tubes meet stress ratio criteria based on continued l
j operation at the same stability ratio. The predicted future usage for the
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limiting tube is .0005 per year or about 0.012 over the remaining period of operating license. The original AVB locations and tube fatigue usage, prior to installation of the AVB modification, were not determined. However, by the prior design, Row 11 tubes are expected to have had AVB support and consequently negligible prior fatigue usage. Even though AVB position evaluations for other units of similar design have indicated an occasional Row 11 tube that was not supported, the frequency of such an occurrence is less than 0.5% based on a sample of over 2800 tubes. Table 9 3 provides justification for acceptance of Row 11 tubes based on analyses which indicate that Row 11 tube stress levels in Zion Unit 2 would not lead to rapid propagation of postulated or existing cracks.
2.8 Overall Conclusion 1
1 The analysis described above indicates that the Zion tubes remaining in service are not expected to be susceptible to fatigue rupture at the top support plate l in a manner similar to the rupture which occurred at North Anna 1. Therefore.
l no modification, preventive tube plugging, or other measure to preclude such an
) event is necessary, i
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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, i
! 3.1 North Anna Unit 1 Tube Rupture Event The cause of the tube rupture has been determined to be high cycle fatigue, j
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 1 been increased to a maximum level as the result of denting of the tube at the l top tuba support plate and the alternating stress has been determined to be due 1
to out of plane deflection of the tube U bend above the top tube support caused l
! by flow induced vibration. These loads are consistent with a lower bound ,
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- fatigue curve for the tube material in an AVT water chemistry environment. The j vibration mechanism has been determined to be fluid elastic, based on the magnitude of the alternating stress.
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l A significant contributor to the occurrence of excessive vibration is the
! reduction in damping at the tube to tube support plate interface caused by the l j denting. Also, the absence of antivibration bar (AVB) support has been j concluded to be required for requisite vibration to occur. The presence of an 2
AVB support restricts tube motion and thus precludes the deflection amplitude ;
j required for fatigue. Inspection data shows that an AVB is not present fer the i
Row 9 Column 51 tube but that the actual AVB installation depth exceeded the l l minimum requirements in all cases with data for AVBs at many other Row 9 I
! tubes. Also contributing significantly to the level of vibration, and thus l I 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 j
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f properties of the tube reflect the lower range of properties expected for an I i l 0221M:49/072888 22 I
AVT environment. In sumary, the prerequisite conditions derived from the evaluations were concluded to be:
prereauisite Conditions Fatiaue Recuirements Alternating stress Tube vibration
- Dented support ,
- Flow excitation
- Absence of AVB I
Mean stress Denting in addition to applied stress Material fatigue properties AVT environment ,
- Lower range of properties r
3.2 Tube Examination Results Fatigue was found to have initiated on the cold leg outside surface of Tube R9C51 imediately above the top tube support plate. No indications of f significant accompanying intergranular corrosion was observed on the fracture ;
face or on the imediately adjacent 00 surfaces. Multiple fatigue initiation i 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 l l
system used, or approximately 90' from the geometric center of the initiation [
zone at Section D 0. High cycle fatigue striation spacings approached 1 f micro inch near the origin sites, Figure 3 2. The early crack front is believed I to have broken through wall from approximately 100' to 140', From this l 1 point on, crack growth is believed (as determined by striation spacing, striation direction, and later observations of parabolic dimples followed by I equiaxed dimples) to have accelerated and to have changed direction with the [
resulting crack front running perpendicular to the circumferential direction. ,
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1
3.3 Mechanism Assessment To address a fatigue mechanism and to identify the cause of the loading, any loadina 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 rumber of cycles involved and a corresponding low fatigue usage, even when accounting for the dented tube condition at the plate. This analysis also showed an axial tensile stress contribution at the tube 00 a short distance above the plate from operating pressure and temperature, thus providing a contribution to mean stress. Combining these effects with denting deflection on the tube demonstrated a high mean stress at the failure location. Vibration analysis for the tube developed the characteristics of 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 progrossing 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 9 ksi. The comparison to the best estimate of plant leakage (performed after the event) showed good agreement, Figure 3 3.
Based on these results, it followed that the predominant loading mechanism responsible is a flow induced, tube vibration loading mechanism. It was shown that of the two possible flow-induced vibration mechanisms, turbulence and fluidelastic instability, that fluidelastic instability was the most probable cause. Due to the range of expected initiation stress amplitudes (4 to 10 ksi), the fluidelastic instability would be limited in displacement to a range of approximately ( la,c. This is less than the distance between tubes at the apex, ( ja.c. It was further confirmed that displacement prior to the rupture was limited since no indication of tube U bend (apex region) damage was evident in the eddy-current signals for adjacent tubes.
0221M:49/072888 24
I 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. Nean stress effects were f [
1 included due to the presence of denting and applied loadings. The results i indicated that a 10% reduction in stability ratio is needed (considering the
, range of possible initiation stress amplitudes) to reduce the fatigue usage per j year to less than 0.02 for a tube similar to Row 9 Column 51 at North Anna 4
Unit 1.
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I 0221N:49/0728SS 25 !
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! Figure 3 1 Approximate Mapping of Fracture surface of Tube R9C51, 5/G 'C' Cold Leg, North Anna Unit 1 l l
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8 = 1.8/1.4 e in.
- Heavy l
Calde 3 8 s t! w in.
l 3 = 1.0/1.85 e in.
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14 Dinoles 270'- l> ) and V Intert.41 Wecking 38 3 = 1.8/4.0 m in.
)g l j 3,3 ,88#f,y fB 8B Wearly taut Aaed 8 = 4.1/6.9 v in. Otaples l
Wete: Arrows Indicate Direction of Fracture Propagation l Figure 3 2 schematic Representation of Features Observed During TEM Fractograhic Examination of Fracture Surface of Tube R9C51, $/G 'C' Cold Leg North Anna Unit 1 l
l l
l 1
- l l l l l l l l l l l Calculated and etterved leak rates versvl glas.
l 06terved values based on galeeus specialcondenter i
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~m .
3!SNA A = 5 KS!
S!sMA A = 7 KS! i
.........- g!gMA A = 3 KS! l a j
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- Figure 3 3 Calculated and observed Leak Rates Versus Time t
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I 4.0 CRITERIA FOR FATIGUE ASSESSMENT The evaluation method and acceptance criteria are based on a relative comparison with the Row 9 Column 51 tube of Steam Generator C, North Anna Unit 1. This approach is necessary because (1) methods for direct analytical prediction of actual stability ratios incorporate greater uncertainties than a relative ratto method, and (2) the stress amplitude (or dis," . cement) 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 Zion 1 and 2 tubing to Row 9 Column 51 of North Anna Unit 1, several criteria are utilized. First, stability ratios are calculated for Zion 1 and 2 steam generators based on flow fields predicted by 3 D thermal hydraulic models and ratioed to the stability ratio for Row 9 Column 51 at North Anna Unit 1 based on a flow field obtained with a 3 D thermal hydraulic model with the same degree of refinement. These ratios of stability ratio (called relative stability ratios) for each '
potentially unsupported U bend in the Zion 1 and 2 steam generators should be equivalent to 10.9 of R9C51, North Anna 1 (meeting the 10% reduction in 1
stability ratio criteria). This provides the first level of screening of 4
susceptible tubes incorporating all tube geometry and flow field differences in 1 the tube dynamic evaluation, it has the inherent assumption, however, that each tube has the same local, high flow condition present at Row 9 Column 51, North Anna Unit 1. To account for these differences, flow peaking factors can be incorporated in the relative stability ratios and the relative stress ratios.
d i
j 0221M:49/072888 29
The next step is to obtain stress ratios, the ratio of stress in the Zion I or l 2 tube of interest to the stress in Row 9 Column 51, North Anna Unit 1, and after incorporating the requirement that the relative stability ratio to Row 9 ;
Column 51 (R9C51) for the tube of interest is equivalent to s 0.9, require the j
stress ratio to be $1.0. The stress ratio incorporates the tube g n metry }
j differences with R9C51 in relation to the stress calculation and also ,
{ incorporates the ratio of flow peaking factor for the tube of interest to the
! flow peaking factor for R9C51 (flow peaking factor is defined in Section 4.2). f This should provide that all tubes meeting this criteria have stress amplitudes equivalent to i 4.0 ksi. )
l t
l j
"inally, the cumulative fatigue usage for plant operation to date and for l continued operation with the same operating parameters is evaluated. A fatigue l
usage of s 1.0 may not be satisfied by meeting the stress ratio criteria using i the reference operating cycle evaluation since the reference cycle does not ;
]
necessarily represent the exact duty cycle to date. Therefore, the time l l
history of operation is evaluated on a normalized basis and used together with I
the stress ratio to obtain a stress amplitude history. This permits the i f calculation of current and future fatigue usage for comparison to 1.0. l I
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 f a
unity (1.0) indicate instability (see Section 5.1). I L
l i 1
Motions developed by a tube in the fluidelastically unstable mode are quite r large in comparison to the other known mechanisms. The maximum modal (
displacement (at the apex of the tube) is linearly related to the bending l stress in the tube just above the cold leg top tube support plate. This [
relationship applies to any vibration in that mode. Tnus, it is possible for f j an unstable, fixed boundary condition tube to deflect an amount in the U bend j which will produce fatigue inducing stresses. l l !
f I
I 0221M:49/072*S8-30
l l
l l
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 v. triable. Taking logs of both sides of this equation leads to the slope-intercept form of a straight-line equatica 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 stchility ratio SR, and the dependent va-table y is tube (fluidelastic instability induced) displacement ret -D, ,d the slope n is renamed s.
From experimental iesults, it is known that the turbulence response curve (on log-log coordinates) has a slope of approximately ( Ja,b,c. Test results also show that the slupe for the fluidelastic response depends somewhat on the instability displacement (response amplitude). It has been shown by tests that a slope of ( Ja,b,c is a range of values corresponding to displacement amplitudes in the range of ( ]a,c, 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,c whee 01 and SR1 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 j displacement response for any given reduction in stability ratio. l This equrtion shows that there is benefit derived from even a very small percentage change in the stability ratio. It is this reduction in displacement i for a quite small . eduction in stability ratio that formed the basis for demonstrating that a 10% reduction in stabil'Vy ratio would be sufficient to prevent Row 9 Column 51 from rupturing by fatigue.
0221M:49/072888-31
The fatigue curve developed for the North Anna Unit I tube at R9C51 is from
(
Ja,c. Thus,
~
a,c i
~
where, o'a is the equivalent stress amplitude to a, that accounts for a maximum stress ofy a , th yield strength. Tha -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 trean stress.
The curve utilized in this evaluation is clearly well below the code curve reflecting the effect of an AVT environment on fatigue and (
Ja,c for accounting for mean stress that applies to materials in a corrosive environment.
Two other mean stress models were investigated for the appropriatoness of their use in providing a reasonable agreement with the expected range of initiating stress amplitudes. These were the l la,c shown in Figure 4 3. With a (
]a,c, the (
, la,C, i
l l
l 0221M:49/072388 32 r l
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 The slope in this equation can range from [ la,c on a log scale depending on the amplitude of displacement. Knowing the stress resulting from a change in stability ratio from SR1 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 (
ja,c. The corresponding stress level is 5.6 ksi.
The maximum stress, 9.5 ksi, would be reduced to ( Ja,c with a 10%
reduction in stability ratio and would have a future fatigue usage of i
( la,c per year at 75% availability, Figure 4-4. The minimum stress, ,
5.6 ksi, would be reduced to (. Ja,c ksi with a 5% reduction in stability ratio i would have future fatigue usage of ( Ja,c per year, Figure 4-5. iddition, if a tube were already cracked, the crack could be as large as ( neh in length and thru wall and would not propagate if the stress amplitudes sre reduced to s 4.0 ksi, i
l 0221M:49/072708 33
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, confirming the conservatism of 9.5 ksi. [
Ja c, cumulative fatigue usage may then be computed to get a magnitude of alternating stress for the last cycle that results in a cumulative 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 [
]a,c for R9C51, North Anna Unit 1, Steam Generator C, and that the major portion of the fatigue usage came in the second, third and fourth cycles. The first cycle was conservatively omitted, since denting is assumed, for purposes of this analysis, to have occurred during that first cycle. Based on this evaluation, the tube fatigue probably occurred over most of the operating history of North Anna Unit 1.
A similar calculation can be performed for the time t' story of operation assuming that [
Ja.c, On this basis, the effect of a 10% reduction in stability ratio is to reduce the stress amplitude to 4.0 ksi and results in a future fatigue 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 higher curves, 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.
022)M:49/072888-34
4.2 Local Flow Peaking Considerations Local flow peaking is a factor on stability ratio that incorporates the effect on local flow velocity, density and void fraction due to non-uniform AVB insartion 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 Zion Unit 1 and 2 tubing is relative to R9C51, North Anna Unit 1, the flow peaking factors are also applied as relative ratios, i.e., a ratio of Zion Unit 1 and 2 tubing 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 ( Ja,b,c is appropriate for R9C51 of North Anna 1. The peaking factor for a tube in Zion Unit 1 and 2 tubing is therefore divided by ( Ja,b,c and the resulting relative flow peaking is multiplied times the relative stability ratio based on ATHOS results. If the peaking factor is 1.0, the relative flow peaking is
[ Ja,b,c, As a further demonstration of the conservatism of ( ')a,b,c as the minimum flow peaking factor for R9C51, the stress amplitude of 7.0 ksi obtained from iterating on cumulative fatigue usage (and selected as the nominal value from fractography analysis) wa: used to back calculate the apparent stability ratio and then the apparent flow peaking factor. Allowing for a range of slopes of the instability curse from 10 to 30, the stability ratio is in the range of 1.1 to 1.4 and the flow peaking factor is in the range of 1.8 to 2.2. This range of flow peaking agrees with the range of flow peaking factors measured in the air tests and is considered to be the best estimate of the range of the R9C51 flow peaking factor.
! 0221M:49/072888 35 l
l
The range of stability ratios. 1.1 to 1.4, is based on a value of 0.63 obtained with ATHOS results without flow peaking and with nominal damping that is a function of modal effective void fraction (MEVF). MEVF is calculated using the formula:
_ - 44 The nominal damping reflects the nominal reduction in damping that occurs with denting at the tube support plate. Therefore, a minimum damping scenario that is independent of void fraction is not considered to be credible and is not addressed in the evaluation that follows.
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 another tube ir the same or another steam generator, the differences in (
Ja,C, ,
f l ,ca e 1 I
I i
0221M:49/082688 36
l a,c where the stability ratio (SR) includes the flow peaking effect.
1 0221M:49/072888 37
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 stress (such as for non dented tubes), different frequency tubes from the ( Ja,c.e hertz frequency of R9C51, North Anna 1, and shorter design basis service.
In the case of lower mean stress, the stress amplitude that would have caused the failure of R9051, North Anna 1, would have been higher. (
Ja,C, A lower or higher frequency tube would not reach a usage of 1.0 in the same length of time as the R9C51 tube due to the different frequency of cycling.
The usage accumulated is proportional to the frequency and, therefore, the allowable number of cycles to reach a usage of 1.0 is inversely proportional to frequency. The equivalent number of cycles to give the usage of 1.0 for a different frequency tube (
ja,c, 1
For a different time basis for fatigue usage evaluation, (
4 jac.e, l 0221M:49/072888 38
Knowing the magnituda of the stress ratio allows 1) the determination of tubes that do not meet a value of i 1, and 2) the calculation of maximum stress in the acceptable tubes, a,c Having this maximum stress permits the evaluation of the maximum fatigue usage for Zion Units 1 and 2 based on the time history expressed by normalized stability ratios for the duty cycle (see Section 7.4).
1 i
i 1
l t
0221M:49/072888 39 4
6
- - - - ,- 7.
% - . .-. . . - - , - . _ ~ , _
Table 4-1 Fatigue Usage per Year Resulting From Stability Ratio Reduction SR, % STRESS FATIGUE MEAN STRESS USAGE REDUCTION BASIS (l) CURVE (2) MODEL PER YEAR a,c
- 5. 9 yrs to fall [ ,)a,c
- 5. 9 yrs to fail [ Ja c
- 5. 9 yrs to fail ( Ja,e 10.
max.stre{g) amplit d 10.
max.stre{g) amplitdg 10.
max.stresg) amplitudet (i . ]a c 10.
max.stregg) mplitu
- 10. max. stress based on duty c lj(5)
I (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 ,= Seean + 34 '
(3) ( Ja,c, (4) Cycles to failure implied by this combination of stress and fatigue properties is notably less than implied by the operating history.
Consequently this combination is a conservative, bounding estimate.
(5)
Cyclestgfatiguecurveatthemaximumstressof(failureimpliedbytheoperatinghistory Ja, Ja, ,
3221M:49/072888 40
a,b, Figure 4-1 Vibration Displacuent vs. Stability Ratio
l l
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t
(
.t 4
.t .
1 s
l-t i e i
M t
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{
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1 1 i i
Figure 4 2 Fatigue Strength of Inconel 400 in AVT Water at 600'F i i
6 l f t
s I
r I
'I a e a,
1 I
r J
I i i i
l i
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J i Figure 4-3 Fatigue curve for Inconel 600 in AVT Water !
Comparison of Mean Stress correction Models l I
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F l I 2
l l
s,c
[r r
i e
i 1
i 1
i Figure 4 4 Modified Fatigue with 105 Reduction in Stability ,
l Ratio for Maxima Stress Condition \
l l
{
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, . , , - . _ _ -, . - - , , = - - . , . . , , - - . . . - - - ,, ,.p- -, - - - - - - , . _ . - . , _ , - - - - - - . - - -,
i l
a,e r
r 1
t P
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Figure 4 5 Modified Fatigue with 5% Reduction in Stability i
Ratio for Minimum Stress Condition l l
l l
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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 cnd 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 mode 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 r.,
where Uc (critical velocity) and Uen (effective velocity) are determined by:
2 Ue =4f D n ((m, 6 n) / I#o D))1/2 (3) and; j3 I#j/#oIUjI jn Zj u,,2 ......................
c23 N
2 jf("j/*o)4 jn 3
z j
0221M:49/072888 40
where, 0 = tube outside diameter, inches U en - effective velocity for mode n, inches /sec
= number of nodal points of the fiaite element model I N
= number of degrees of freedom in the out of plane direction mj, Uj, pj = mass per unit length, crossflow velocity and fluid density at node j, respectively
- o, mo = reference density and a reference mass per unit length, respectively (any representative values) on = logarithmic decrement (damping) l djn = 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 B = an experimentally correlated stability constant Substitution of Equations (1) and (2) into the expression which defines t
' stability ratio, and cancellation of like terms, leads to an expression in fundamental terms (without the arbitrary reference mass and density parameters). From this resulting expression, it is seen that the stability l ratio is directly related to the flow field in terms of the secondary fluid i velocity times square rocudensity distribution (over the tube mode shape), and )
inversely related to the square root of the mass distribution, square root of l l
modal damping, tube modal frequency, and the stability constant (beta).
! l 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 i discussion, and, where appropriate, the experimental bases to quantitatively l establish the uncertainty associated with each of these parameters. In 0221M:49/082688 47
addition, Section 5.3 provides the experimental basis to demonstrate that tubes with (
Ja.c. This implies that those tubes [ la,c would not have to be modified because their instability response amplitude (and stress) would be small. The very high degree of sensitivity of tube response (displacements 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.
Frecuency It has been demonstrated by investigators that analytically deterwinee frequencies are quite close to their physical counterparts obtained from measurements on real structures. Thus, the uncertainty in frequercies has been shown to be quite small. This is particularly appropriate in the case of dented (fixed boundary condition) tubes. Therefors, uncertainty 'evels introduced by the frequency parameter are expected to be insignificant (see also "Average Flow Field" subsection below).
lattability Constant (Betal The beta (stability constant) values used for stability ratio and critical velocity evaluations (see above equat bns) are based on an extensive data base i
comprised of both Westinghouse and other experimental results. In addition, previous field experiences are considered. Values have been measured for full length U bend tubes in prototypical steam / water environments. In addition, measurements in U bend air models have been made with both no AVB and variable AVB supports (Figure 5-3).
To help establish the uncertainties associated with 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 detennined. These analyses supported a beta value of ( ]a,b,c, 0221M:49/072888 48 1
A summary of the test bases and qualifications of the beta values used for these assessments is provided by figure 5-2. The lowest measured beta for tubes without AVBs was a value of ( Ja,b,c. This value is used for the beta 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 [
Ja,c, i Tube Damoina Test data are available to define tube damping for clamped (fixed) tube i supports, appropriate to dented tube conditions, in steam / water flow conditions. Prototypic U bend testing has been performed under conditions leading to pinned supports. The data of Axisa in Figure 5 4 provides the principal data for clamped tube conditions in steam / water. This data was obtained for cross flow over straight tubes. Uncertainties are not defined for the data from these tests. Detailed tube damping data used in support of the stability ratio evaluations addressed in this report are provided in Section 5.2, below.
Flow Field - Velocity Times Souare-Root Density Distribution Average and U bond local flow field uncertainties are addressed independently '
in the following. l i
J 0221M:49/072888 49 ,
i
Averaae Flow Field Uncertainties in the average flow field parameters, obtained from ATHOS 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 { ]a,e for the combined effects of average flow field, stability constant and frequency appears to be reasonable. To further minimize the impact of these uncertainties, the Zion 1 and 2 tubes are evaluated on a relative basis, so that constant error factors are essentially eliminated.
Thus, the uncertainties associated with the average velocity times square-root-density (combined) parameter are not expected to be significant.
U Bend local Flow Field Non uniform AVB insertion depths have been shown to have effects on stability ratios. Flow peaking, brought about by the "channeling" effects of non-uniform AV8s, 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 Sectier 5.2, below.
Overall Vncertaintjes Assessment Based on the above discussions, and the data provided in the following sections, 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 and stability ratios relative to the North Anna tube, R9C51, are utilized in this report to minimize uncertainties in the evaluations associated with instability constants, incal flow field effects and tube damping.
0221M:49/072888 50
5.2 Tube Damping Data The damping ratio depends on several aspects of the physical system. Two primary determinants 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-phase 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 f contributions from structural, viscous, flow dependent, and two phase damping.
The structural damping will be equal to the measured damping in air. However, in two phase flow, the damping ratio increases significantly and is dependent on the void fraction or quality. It can be shown that the damping contribution from viscous effects are very small.
Damping ratios for tubes in air and in air water flows have been measured and reported by various authors. However, the results from air water flow are poor representations of the actual conditions in a steam generator (steam water flow at high pressure). Therefore, where available, results from prototypic steam water flow conditions should be used. Fortunately, within the past few years test data on tube vibration under steam water flow has been developed for both pinned and clamped tube support conditions.
Two sources of data are particularly noteworthy and are used here. The first is a large body of recent, as yet unpublished data from high pressure steam water tests conducted by Mitsubishi Heavy Industries (MHI). These data were gathered under pinned tube support conditions. The second is comprised of the results from tests sponsored by the Eleccric Power Research Institute (EPRI) and reported in References (5 2) and (5 3).
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 determined on the basis of ( ja C 0221M:49/072888 51
- i. (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 supports at all void fractions. Denting of the tubes at the top support plate effectively clamps the tubes at that location. Therefore, the clamped tube support curve is used in the current evaluation to include the effect of denting at the top tube support plate.
The 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 tube vibration in air has shown that the l mechanical damping is only ( Ja.c rather than the .5% reported in t
Reference (5-3). Therefore the lower curve in Figure 5 4 is the Reference (5 3) data with all damping values reduced by ( .)a,c, 0221M:49/072888 52 ;
5.3 Tube Vibration Amplitudes With Single-Sided AVB Support A series of wind tunnel tests were conducted to investigate the effects of tube /AVB eccentricity on the vibration amplitudes caused by fluidelastic vibration.
['
Ja 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 centilever tube apparatus are in good agreement with corresponding characteristics observed in wind tunnel and steam flow tests using U bend tube 4 arrays. A summary of these prior results is given in Table 5-1.
An overall view of the apparatus is shown in Figure 5 5. Figure 5-6 is a top j view of the apparatus. [
I c
i i
)#.C , .
4 i
i 0221M:49/072888 53
s As shown in Figure 5 7, the tube vibration amplitude below a critical velocity is caused by (
Ja,c, Figure 5 7 shows the manner in which the zero-to peak vibration amplitude, expressed as a ratio normalized to [ Ja,c, varies when one gap remains at( Ja,c. For increasing velocities, up to that corresponding to a stability ratio of (
Ja,c. Figuro 5 8 shows typical vibration amplitude and tube /AVB impact force signals corresponding to those t l obtained from the tests which provided the results shown in Figure 5-7. As expected, impacting is only observed in the ( Ja,c, I l
It is concluded from the above test results that. (
Ja,c, 5.4 Tests to Determine the Effects on Fluidelastic Instability of l Columnwise Variations in AVB Insertion Depths l
This section sumarizes a series of wind tunnel tests that were conducted to investigate the effects of variations in AVB configurations on the initiation of fluidelasti: vibration. Each configuration is defined as a specific set of insertion depths for the individual AVBs in the vicinity of an unsupported U bend tube.
The tests were conducted in the wind tunnel using a modified version of the cantilever tube apparatus described in Section 5.3. Figure 5 9 shows the conceptual design of the apparatus.1 The straight cantilever tube,
- 1) The AVBs shown in Figure 5 9 correspond to original AVBs. AVBs corresponding to those used in field modified units, such as Zion Unit 2, were also made using the same procedure as for the original AVBs.
0221M:49/072888 54
(
ja.c.
t
[
ja.c. Figure 5 11 shows the AVBs, when the side panel of the test section is removed. Also shown is the ;
i top flow screen which is ( ;
Ja.c. The AVB l configurations tested are shown in Figure 5-12. Configuration la corresponds j 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. i a
h 1
l 4
1 0221M:49/072888 55 ;
.I /Q As shown in Figure 5 9, [
ja,c, All the tubes except the instrumented tubed (corresponding to Row 10) are
( Ja.c. As discussed in Section 5.3, prior testing indicates that this situation provides a valid model. Tiie instrumented tube ( la.c as shown in Figure 5.10.
j Its ( Ja,c direction vibrational motion is measured using a non-contacting transducer.
l
(
Ja.c. The instrumented tube corresponds to a Row 10 tube as shown in Figure 5-9. However, depending on the particular AVB configuration, it can reasonably represent a tube in Rows 8 through 11. The AVB profile in the straight tube model is the average of Rows 8 and 11. The difference in profile is quite small for these bounding rows.
( la,c using a hot film anemometer located as shown in Figure 5-9.
Figure 5 13 shows the rms vibration amplitude, as determined from PSD (power spectral density) measurements made using an FFT spectrum analyzer, versus flow velocity for Configuration la (which corresponds to tube R9C51 in North Anna).
Data for three repeat te:;ts 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.
0221M:49/072888 56
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 (R9051 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 Zion 1 and 2 evaluation are sumarized in Table 5 2. The test data is presented as a velocity peaking ratio; a ratio of critical velocity for North Anne tuce R9C51 configuration la,
( to that for each Zion 1 and 2 AVB configuration evaluated.
l 5.5 References q l , a,c b 5-1 52 53 54 _
s -
0221M:49/072788-56 -
Table 5-1 Wind Tunnel Tests on Cantilever Tube Model OBJECTIVE: Investigate the effects of tube /AVB fitup on flow-induced tube vibration.
APPARATUS: Array of cantilevered tubes with end supports [
l Ja.c, MEASUREMENTS: Tube vibration amplitude and tube /AVB impact forces or preload forces.
RESULTS:
a,b,c
/ S 1.
2.
3.
5.
s 0221M:49/072788 57
i I
Table 5-2 Fluidelastic Instability Velocity Peaking Ratios
( for Colunnwise Variation in AVB Insertion Depths (ZION UNITS 1 & 2)
Type of Insertion Peaking Ratio Configuration Ula/Un a,b,c la lb i
2a i
3 I 4a 4b 4c 4d
! 4f
- 4n 4P
! 4q
] 5a 5b 5c Sh 6c
- 14b 14 s <
i . - . . .
l Note: Un is instai.- -
volccitj r ~=let for type n of AVB insertion l
configueat'r ,
0221M:49/0F'68 58 i
'i
v, .
=
?
l i 4
l f
i l-i i
i
.i .
r i !
t t
C v
l i i t
l w, . a '
t i
Figure 5 1 Fluidelastic Instability Uncertainty Assessment j i
I 1
i
.. i
? 0221M:49/071288 59 i 1
6 i
ga g---r- . . . - - _
l )
l U Band Test Data l l i
- 1) MB 3 Tests !
Avaluesof( Ja,b,c
- 2) MB 2 Tests
- of ( Ja,b,c
- 3) Air Model Tests
- A of ( Ja,b,c without AVBs .
Tendency for A to increase in range of ( ]a,b,c l
with inactive AVBs (gaps at AV8s)
Tendency for # to decrease toward a lower bound of ,
[ ]a,b,c with active AV8s Verification of Instability Conditions i
- 1) Flow conditions at critical velocity frca MB 3 i
! 2) Measured damping for the specific tube ;
- 3) Calculated velocities from ATHOS 30 analysis
- 4) A determined from calculated critical values i Good agreement with reported A values
- 5) ATH0S velocity data with A ef ( la,b,c and known damping should not significantly underestimate instability for regions of j l uniform U bend flow f
! i 1
i !
i j !
I ;
i
! Figure 5 2 Instability Constant - A i
I I i
0221M:49/072788 60 i
. . .= . -
D ,h C
, m i
I d
f i
i 4
I I
1, ,
1 I
l i
l i
l l
i c
l 1 l
- l
! I i !
I i
) !
l
.l 1, I 1 !
L i
l ,
l '
l Figure 5-3 Instability Constants. A. Obtained for Curved Tubes from i.
' Wind Tunnel Tests on the 0.214 Scale U Bend Model i
! i I !
I I 0221M:49/071288 4l
- i
I s,b,c i
l l
l [
l '
i i
t I
i 1
I i
i !
i ,
h
. i f
- 1 e
i i
i 1
- a f
4 i
i .
1 i l
T L
l
- 1 I
I I
l i i i ;
1 i
I
- l I i Figure 5-4 Damping vs. Slip Void Fraction ;
t 1 .
, i I
i l
I I I I, i
'l 1 i 1 __ _ , ._ _,__..___ .- _ - . .___ _ _-__ ___._ _ _ _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ . _ _ . _ . . _ _ _ _ _ _ . _ _ . , _
- '"~
a,b,c
) Figure 5 5 Overall View of Cantilever Tube Wind Tunnel Model 253ESA
~
a,b,c Figure 5 6 Top View of the Cantilever Tube Wind Tunnel Model i
25368 2
i a,b,c
'D 1
a a
t I t i
s d
J t
1 f
l 1
r L
i 1 ,
1
.i :
2 r
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- l
\
l i
i
< ?
4
. f t
1 ,
,i I j I l Figure 5 7 Fluideinstic Vibration Asp 11tude with Non Uniform Saps ,
e i
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l I
I l \
l
a,b.c '
~
[
t l
t l
J l '
f i i
1 1
a I -
1 1
1 4
]
I ;
j l i i i l i
< 1 l
i i ,
2 i Figure 5 8 Typical Vibration Amplitude and Tube /AYS !apact Force Signals for Fluidelsstic Vibration with Unequal f
l Tube /AVI taps l f I
l l I
l
l 1
a,b,c Figure 5 9 Conceptual Design of the Apparatus for Determining the Effects of Fluidelastic Instability of Celtanwise Variattens in AVR Insertion Depths l
l
i l
l
- a,b,c Figure 510 Overa'l View of Wind Tunnel Test Apparatus 25366-3
- a,b,c
)
4 1
1 i
l t
I i
Figure 511 Side View of Wind Tunnel Apparatus with Cover Plates Removed to Show Simulated AVBs for Field Modified Units and Top Flow Screen h
253(8 5 i
TYPE OF AVS TYPE OF AV8 INSERTON NSERTION
' k, bo C s boboC ,
\
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4 1
),
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4 ,
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/ (
1 Figure 5 12 AVS Confitorations Tested
' i e
i 0221M:49/072188.GS
- l a,b,c )
4
. a, i i
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t 1
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1 1
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k I
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. L 1 ,
4 ',
a :
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1
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Figure 5 13 Typical Variation of RMS Vibration A.nlitude with Flew :
! Velocity for Configuration la in Figure 512 !
a L I
t 3 i 3
i L
l I
i 6.0 ED0Y CURRENT DATA AND AVB POSITIONS i
, t
- The EC input to the Zion Unit 1 analyses is based on EC tapes generated during ;
) the 100 percent inspection of March,1988. The EC input to the Zion Unit 2 f
analyses is based on EC tapes generated during the 100 percent inspection of .
Jur.e 1987.
i i
6.1 AVB Asscobly Design 4
6.1.1 Zion Unit 1 i
j I '
6 i
i a i l i l
Ja.c.e ' Upper' AVBs which are inserted beyond the design l depth, occasionally show on the EC traces for the Row 12 tubes. Since the ,
l purpose of this analysis is to evaluate potentially unsupported tubes at or l near the point of maximum AVB insertion, only the dimensions and EC data i pertaining to the ' lower' AVBs is used. Review of the EC data for Zion Unit 1 1
~
shows that with a few exceptions in SG A, and SG C, the lower AVBs extend as ;
! far inward as Row 8 or Row 9, and tys,1cally less than 15 of the Row 9 tubes are I not supported. ;
i !
4 6.1.2 Zion Unit 2 1
The AVBs originally supplied with the Zion Unit 2 SGs were replaced in 1987. ;
{
4 The design of the replacement AVBs uses ( j i
i .
i !
! i
! i i 0221M:49/072888 66
I j :
I l Ja.c.e The nominal insertion depth of the replacement AVBs was controlled during installation by the use of temporary spacers between the outennost tube, and the attachment head on the outer end of each bar. This insertion depth was l j verified separately by inspection prior to permanently fixing the AV8s to their i retainer plate. At the ' nominal' point of the AVB assembly tolerance, ( ;
i 1
I i
) ,
ja.c.e Review of the EC data for Zion Unit 2 shows that from tube Column 3 through tube Column 92, the AVBs support all Row 12 tubes, and approximately 8 percent of the Row 11 tubes. The Il ( Jac.e always support the Row 10 tubes in Columns 3, 4, 91 !
l and 92.
I i
6.2 Eddy Current Data for AVB Positions i
i The AVB insertion depths were determined on the basis of interpretation of the !
eddy current data. To locate the AV3s, the ECT data traces were searched for the characteristic peaks seen in the signals, which indicate the intersection j of an AVB (or a tube support plate) with the tube (a typical signal for replacement AVBs is shown in Figure 6.1). Since ambiguity can occur in the !
interpretation of the ECT data, due to inability of ECT to differentiate at l i
which side of a tube a "visible' AV8 is located, other information was used to
! l l !
i i
i 0221M:49/082688 67 i i
assist in establishing the location of the AVBs. Consistency with the design of the AVB assembly, consistency of data for adjacent columns and verification by projection were utilized to determine the depth of insertion which was plotted. For the cases of single AVB contacts, the length of the contact signal, and verification by projection was used in a few instances to confirm or deny support of the tube with a single contact.
j 6.2.1 Zion Unit 1 1
- For Zion Unit 1, the number of AVB intersectiJns, including Zero (Naning no AVB present), was logged for each tube to indicate the presence or absence of l
AVBs. Where only a single intersection was indicated by the data, the length I
of this intersection was recorded to provide additional information to assess the adequacy of support for the tube. Figures 6-2 through 6 5 show the number of AVB signals found for each Zion Unit 1 tube, the condition of the tube to (
l TSP interface at the locations when it was evaluated, and a representation of .
AVB insertion distance based on evaluation of the EC data. Details of AVB projection techniques based on EC data and tests are given in Section 6.2.4.
t j 6.2.2 Zion Unit 2 i '
j The ' replacement' AVBs used in Zion Unit 2 are fabricated primarily from a l SA 405 which is a ferritic stainless steel. As a result, they have a stronger .
EC signal than the Inconel AVBs in Zion Unit 1.. This stronger signal leads to j a large number of spurious ' support' readings for tubes in Row 11. As a result, the AVB counts for the Row 11 tubes are heavily discounted, and the AVB
' insertion distance values for Zion 2 are based primarily on projectior ;
calculations using input from the spacing between the Row 12 and Row 13 AVB intercepts. Figures 6 6 through 6 9 indicate the number of AVB signals found ;
l for each Zion Unit 2 tube, the condition of the tube to TSP interface, where f evaluated, and a represent.ation of AVB insertion distance based on evaluation :
of the EC data. Details of AVB projection techniques based on EC data and tests are given in Section 6.2.4. j l
j 0221M:49/082688 68 l i
+
6.2.3 AVB Insertion Depths AVB position maps for the Zion Unit I and Unit 2 steam generators are given in Figures 6 2 through 6 9.
The direct observation data (the number of AVB intersactions seen by the eddy current probe) are the principal basis for determining the AVB positions. ,
Where the direct observations were ambiguous or there is a conflict between i observations and projections, the more conservative data are used to determine the AVB positions. Since ' direct observation' gives a 'yes - no' type of f answer, the projection method is used to ' interpolate' AVB insertion depths ,
between rows of tubes. The visual images thus produced being more easily understood when fluid flow peaking situations are evaluated.
l I ;
! Greater conservatism is generally interpreted as the AVB being less inserted although consideration must also be given to the resulting flow peaking .
]
factors.
I 1
l !
I 4
)
b
)
.I i i
! l 4
l :
i l
l 1 !
l !
l 0221M:49/082688 69 t
i ,
6.2.4 AVB Projection The projection technique is useful where noisy or spurious ECT signals prevent direct observation of the AVBs and where data is unavailable due to plugged tubes. (
Ja,c, I In the case where the AVB characteristic signa's can not be confidently determined due to a noisy signal or pre existing plugged tubes, location data !
for the AVBs is provided for (I I i 1
i' i
)
i -
r Ja,c, r
F t.
i Ja.C 6.2.5 AVB Projection Testing Four AVB indications are frequently observed at the Row 11 tubes in steam l generators with replacement AV8s. Nomally, this would be interpreted that the l
Row 11 tubes are supported by the replacement AVBs at each of the four AVB ;
legs. While this is theoretically possible if the attachment head of the AVBs 7 is in contact (or nearly in contact) with the outer row of tubes, the ,
probability of frequent occurrence is estimated to be small, since the nominal l' j
0222M:49/072888 1
d inserted position is set using spacers during assembly. The nominal inserted positionis( Ja c.
indicating that the Row 11 tubes are not supported. The maximum inserted position occurs when the attachment head contacts the outer row tubes, and is
( ia.c (Projection values are based on the location of the lower AVB tip on the AVB centerline relative to the tube centerline. The tip of the upper AVB is less inserted due to the geometry of the AVBs.) (Note Figure 6 11)
- Projections based on EC measurements for the Row 11 tubes frequently predict an AVB insertion position which is more deeply inssrted than the corresponding projections based on data for the Row 12 and Row 13 tubes in the same column.
The difference between the Row 11 and the Row 12 and 13 projections is believed to be the result of the eddy current probe seeing the tips of the upper and J lower AVBs, even though the AVB was not in contact with the tube, due to the
! strong signal generated by the magnetic properties of the Type 405S5 utilized l for the replacement AVBs.
A bench test ws performed to determine if AVBs not tr, contact with the tubes j could be seen by ECT. Figures 6 10 and 6 11 shc= the test setups used. The two tubes were separated by spacers to provide the Model 51 nominal tube pitch. The zero position of the AVB was defined as the position of the l la c coincident with the centerline of
, (
the simulated Row 12 tube, and the AVB was moved on a line normal to the tube centerline while maintaining the angle between the AVB and tube centerline, f
- Straight,1600 tubes were used for test convenience, together with prototypic ,
I AYBs. As the figures show, tests were performed, both with the upper AVB bar
! normal to the tubes to simplify testing, and at the angle of incidence for a j Row 12 tube /AVB intercept in order to verify that the ' normal' test setup rearinably represented the actual geometry, j
Figure 612 shows the expected are lengths in the Rll and R12 tubes for a range j of AVB position, relative to the R12 centerline for the AVB normal to the R12 t u bt, . The centor to center distance of the AVB hinge pins is [ ]ac ;
Thus, the expected arc length when the hinge pins are aligned with the tube I centerline is ( la,c When the AVB is moved 1.281 inches (a full l l l !
l l 0222M:49/072888 2 (
i
pitch in the test setup), the expected arc length in R11 is 1.0 inch since the hinge pins are aligned with the R11 centerline at that point. The tip of the lowerAVBis( Ja.c inch below the hinge pin centerline. (
Ja.c Test Results Figure 613 shows the measured are lengths in the simulated Row 11 tube (with normal AVB layout) together with the theoretical arc lengths for the design.
An AVB signal with an are length of (
Ja.c Figure 6 14 shows the data from the test setup illustrated in Figure 6-11 (AYB skewed relative to the tube centerlines to simulate the actual Row 12 condition). The dasa from this test are similar to the prior tests, although the arc length data from the Row 11 test are inconclusive in regard to where the AVB is first seen. Comparing the measured are lengths to the theoretical are lengths, it is found that like the case of the ' normal' test setup, the EC data continues to overpredict AVB insertion. Therefore, it is concluded that the normal test setup adequetely simulates the skewed condition.
Conclusions l
The conclusions drawn from these tests are:
- 1) The AVB is seen by ECT before the AVB provides support to the tube. The tip of an AVB can be seen as two distinct signals, when still 0.25 inches above the tube centerline.
- 2) The measured arc length, tend to be greater than the theoretical arc length.
0222M:49/072888-3
i l
l
- 3) The difference in the measured arc lengths for two adjacent rows supported by the same AVB closely match the theoretical values. .
- 4) If the AVB is seen in Row 11 by ECT, support of Row 12 is assured, if the AVB is not seen in Row 11, projection data are needed to assure Row 12 support.
- 5) Theoretical arc lengths are used for the AVB projections. This practice
! tends to overestimate the actual AVB insertion depths by up to 0.3 of a pitch. For practical applications with the modified AVBs, this practice I tends to maximize the potential for flow peaking in Row 11 and is therefore conservative. In all cases for Zion 2, the projected AVB insertion depths assure Row 12 support even if the uncertainties in the projection methods !
are incorporated.
I 6.3 Tube Denting at Top Tube Support Plate Because of the AVB geometries involved and the desire to obtain 3 rows of f
! ' projection' data where possible, the Zion Unit 1 evaluation covers Rows 8 t5 rough 12; while the Zion Unit 2 elevation covers Rows 9 through 13.
Subsequent to identifying the AVB signals, eddy current data were examined to evaluate the incidence of corrosion and/or denting at the top tube support
- plate. In the denting evaluation, the EC tapes were evaluated to determine the j condition of the tube / TSP interface for potentially unsupported tubes in j locations which could be susceptible to flow peaking. Analysis of the data for (
Zion Unit I shows the presence of ' corrosion with magnetite' in the majority of !
the tube / TSP crevices in this zone, with 7 dented tubes in SG 8. The location l of the dented tubes can be determined from the maps in Figures 6 2 through l 6 9. Analysis of the data for Zion Unit 2 also shows the presence of
' corrosion with magnetite' in the majority of the tube / TSP crevices in this l zone. Additionally, SG A had 5 dented tubes in Row 11; SG 8 had 4 dented tubes in Row 11; SG C had 3 dented tubes in Row 11; while SG 0 had 1 dented tube in
) Row 11. The location of the dented tubes can be determined from the various
'AVB' maps. Because the tube vibration analysts are based on the conservative l
assumption that all tubes in the area of interest are structurally ' fixed' in l l the TSP holes, as if by denting or corrosion, the results of this phase of the (
I i
0222M:49/072888 4
I examination, which are plotted, are informative, but do not influence the disposition of the tubes found to be susceptible to fatigue.
6.4 AVB MAP INTERPRETATIONS
, 6.4.1 Zion Unit 1 1
The Zion Unit 1 SGs retain the original AVBs which have a nominal design I insertion depth intended to support as far inward as the Row 11 tubes.
Evaluation of the EC data indicates that in the area of interest (Row 12
- through Row 8) between Columns 3 and 92, all of the Row li and Row 11 tubes.
1
- nd the majority of the Row 10, 9 and Row 8 tubes were supported in all four i SGs.
{ ' single contact' AVB signals is given in Table 6.1, and a listing of <
unsupported tubes is given in Table 6.2. All of the Row 12 and Row 11 tubes j are supported by AVBs. Three Row 10 tubes, and ) Row 9 tubes are not l supported. The highest flow peaking factors for this SG (see Sec. 8 and 9 of l i this report) were found at tube locations RSC34 and C35.
i A
l SG B j
) !
The AVB map for SG B is given in Figure 6 3. A listing of the resolutions of 1
' single contact' AV8 signals is given in Table 6.1, and a listing of ;
unsupported tubes is given in Table 6.2. All of the Row 12, 11 and Row 10 tubes are supported by AV8s. Seventeen Row 9 tubes are not supported. The >
highest flow peaking factors for this SG were found at the tube locations R9C69 l and R8C69. [
] !
)
i 0222M:49/072888 5 t
SG C The AVB map for bG C is given in Figure 6 4. A listing of unsupported tubes is given in Table 6.2. All of the Row 12 and Row 11 tubes are supported by AVBs.
Two Row 10 and fifteen Row 9 tubes are not supported. The highest flow peaking factors for this SG mre found at the tube locations R9C60 and R8C35.
SG D The AV8 map for SG-D is given in Figure 6-5. A listing of unsupported tubes is given in Table 6.2. All of ths Row 12 Row 11 and Row 10 tubes are supported by AVBs. Seven Row 9 and twenty nine Row 8 tubes are not supported. The highest flow peaking factor for this SG was found at the tube location R8C35.
6.4.2 Zion Unit 2 The Zion Unit 2 SGs are equipped with ' replacement' AVBs, which have a nominal design insertion depth intended to support as far inward as the (
la,c Evaluation of the EC data indicates that in the area of interest (Row 13 through Row 9) between Columns 3 and 92, all of the Row 13 and Row 12 tubes are supported. Inboard of Row 12, approximately 8 percent of the tubes have AVB support.
AVB maps for the SGs in Zion Unit 2 are given in Figures 6 6 through 6 9. The pattern of AVB insertion is very flat, and only minimal flow peaking is identified. The highest flow peaking factors for these SGs was found for tube locations R11C89 and C90 in SG D. (See Sec. 8 and Sec. 9 of this report.)
A listing of unsupported Unit 2 tubes is given in Table 6.3.
6.4.3 Sumary of Tube Support Conditions Resolution of tube support evaluations for single AVB indications are listed in Table 6 1. A sumary of unsupported tubes for Zion 1 is given in Table 6 2, and for Zion Unit 2 in Table 6 3.
0222M:49/072888 6
I I
l J
Table 6.1 Resolution of Tube support Tubes With Single AVB Indications i l
Zion Unit 1 SG A i
A R8C62 Projection data indicates tube is unsupported
'l
!i SG B 1
8 R9C56 Projection data indicates tube is supported - Left Side B R8C68 Projection data indicates tube is supported - Right Side 1
)
SG C l SG has no single AVB intersection indi:ations 1
SG 0 ,
l SG has no single AVB intersection indications Zion Unit 2 1
, The design and materials of the replacement AVBs do not produce the equivalent of a single AVB signal. Support, or lack of tube support is categorized in i
i Table 6.3.
l
}
i l I l
I l \
i I
0222M:49/072888 7 b
I i
Table 6.2 Zion Unit 1 l Summary Listing of Unsupported Tubes Zion Unit 1 Stemn Generator A t
Row 12 Row 12 has no unsupported tubes Row 11 Row 11 has no unsupported tubes Row 10 Columns 41, 42 and 43 are unsupported Row 9 Columns 40 through 47 and Column 51 are unsupported Row 8 Columns 27, 28, 29, 34 and 35; 39 through 56; and 60 through
! 63 are unsupported Zion Unit 1 Steam Generator B i Row 12 Row 12 has no unsupported tubes Row 11 Row 11 has no unsupported tubes Row 10 Row 10 has no unsupported tubes
, Row 9 Columns 18, 42 through 55, 60 and 60 are unsupported Row 8 Columns 11 through 14, 18, 35, 39 through 57, 60, 61 and 69 sre unsupported Zion Unit 1 Stesm Generator C Row 12 Row 12 has no unsupported tubes Row 11 Row 11 has no unsupported tubes j Row 10 Columns 49 and 53 are unsupported i Row 9 Columns 41 through 54, plus Column 60 are unsupported Row 8 Columns 35, 39 through 55, plus 60 and 61 are unsupported Zion Unit 1 Steam Generator D Row 12 Row 12 has no unsupported tubes Row 11 Row 11 has no unsupported tubes Row 10 Row 10 has no unsupported tubes
- Row 9 Columns 46, 47 and 56 through 60 are unsi'pported Row 8 Columns 3b, 40 through 61, and 69, 70, /7. 78, 83 and 84 are j unsupported l
l
(
i l 0222M:49/072888 8 l
l
Table 6.3 Zion Unit 2 Summary 1.isting of Unsupported Tubes Zion Unit 2 Steam Generator A Row 13 Row 13 has no unsupported tubes Rcw 12 Row 12 has no unsup)orted tubes Row 11 Columns 2, 5 througi 9, 12, 16 through 21, 24 through 33, 36 through 39, 42 through 48, 51, 52, 55 through 59, 62 through 65, 70, 73 through 76, 80, 81, 82, 85 through 90, and 93 are unsupported Row 10 All columns except 3, 4, 91 and 92 are unsupported Row 9 All columns are unsupported Zion Unit 2 Steam Generator B
! Row 13 Row 13 has no unsupported tubes Row 12 Row 12 has no unsupported tubes
- Row 11 Columns 2, 5, 6, 9 through 13,16 through 28, 31, 32, 33, 39 through 90, and 93 are unsupported Row 10 All columns except 3, 4, 91 and 92 are unsupported Row 9 All columns >are unsupported Zion Unit 2 Steam Generator C Row 13 Row 13 has no unsupported tubes -
3 Row 12 Row 12 has no unsupported tubes Row 11 Columns 2, 5, 6, 9 through 39, 42, 43, 44, 47 through 53, 56 J through 70, 73, 74, 75, 78 through 90 and 93 are unsupported 4
Row 10 All columns except 3, 4, 91 and 92 are unsupported Row 9 All columns are unsupported Zion Unit 2 Steam Generator D 1
Row 13 Row 13 has no unsupported tubes Row 12 Row 12 has no unsupported tubes Columns 2, 7, 10 through 38, 43, 48 through 54, 57 through i Row 11 !
69, 71 through 90 and 93 are unsupported ,
Row 10 All columns except 3, 4, 91 and 92 are unsupported Row 9 All columns are unsupported f i
i l
l 0222M;49/072888 9 ,
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Figure 6-1 AV5 Insertion Depth Confirmation
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2 F1gure 6 2: Zion Unit 1 Steam Generator A . AVB Positione
r
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8 :
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Figure 6 3: Zion Unit 1 Steam Generster B . AVB Positlene
v i
v 5
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Figurs 64: Zion Unit 1 Steam Generator C . AVS Positions 1
l .-. --. - _- -
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Figure 6 5 Zlon Unit 1 Steam Generator D . AVE Poeltlone ;
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m m me. S Figure 6-4: Zion UnN 2 Steam Generator A . AVB Posnions
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8
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f, Mgure 6 7: Zion Unit 2 Steam Generator 3. AVS Pr.,altione
1 1
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o
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s eeeeG 1 GeeGO :
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E'
" eeeee i Gemee i eeeee r i ee60e T Gemee i 88eee T ;
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mesee T eeeee i eeeee i eeeee r eeeee i; eeeee g ; i eeeee t e e' Dee eeeee s. eeeee i ll ,
s .
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- l2i: l2ie c l *= I : 12 l *i C 12 i : I*I*
m F1gure 6-4: Zlon Unit 2 Steem Generster C . AVB Poeltlene i
, - _ , - , - - - - - - . - - , . _ - , ..-,,-,r -
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=
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i eeeee l eeeee i eeeee E. yli :
l eeeee l eeeee s eeeee i :
l eeeee t eeeee t eeeee s ll eeeee t eeeee s eeeee s ,3 eeeee t _eeeee s '
eeeee s= j eeeee n eeeee , eeeee l eeeee I:: eeeee i eeeee i lh}
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1 _ aeeee t_ eeeee t _
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210 i: 12la eleI: 12la y_ l 2 1 : 15 b.'
I
, , , , , , t I 1 Figure 6 9: Zlon Unit 2 Stoem Generator D . AVE Positions l l
l
b o C,(
i I
)
Figure 6-10 Replacement AVB Proximity ECT Bench Test ,
i AVB Normal to Tube l
0222M:49/072288 19 l
1 i
i 1
< t boC,5 1 !
- 2 i
i i i 4 !
4 i !
t 1 !
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1 !
t
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Figure 6-11 Replacement AVS P N ietty ECT Sonch Test j i-AVS Skewed to Tube CarA 'line !
(
l I
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0222M:49/072280 20 i i
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i Figure 6-12 Theoretical Arc Lengths (AVE Normal to Tube) -
i l
1 i
0222M:49/072288-21 .
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i Figure 6 13 R11: Actual vs Theoretical Arc between ,
AVS Legs Centerlines l
0272M:49/072248 22 -
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4 1
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Figure 6 14 R11: Actual vs Theoretical Arc Lengths !
4 AVB Skewed Relative to the Tube Center 11ne i B
0222M:49/072288 23 l
l 7.0 THERMAL AND HYORAULIC ANALYS!!
This section presents the results of a thermal and hydraulic analysis of the flow field on the secondary side of the steam generator using the 3 0 ATH0S 1
computer code, Reference (7 1). The major results of the analysis are the ;
r water / steam velocity components, density, void fraction, and the primary and secondary fluio and tube wall temperatures. The distributions of the tube gap velocity and density along a given tube were obtained by reducing the ATH0S j results. In the following subsections, the ATH0S model and some sample results of the analysis are described. Normalized stability ratios over the operating l histories of both units were also determined and are reported in Section 7.4.
I 4
l 7.1 Zion Steam Generator Operating Conditions Recent steam generator operating condition data for the Zion units were provided by Comonwenith Edison. The data reported were:
i Data reported by Comonwealth Edison for Zion 1 and 2 1
i i
! 1. Recent Full Load Operating Parameters (Equal for Both Units) ;
I l
- a. Steam pressure - 720 psig !
' 0
- b. Feedwater Temperature and Flowrates 428'F 14x10 lbs/hr
- c. Primary inlet and Outlet Temperatures Tin 591'F Toyg 527'F j
- d. Thermal Load - - 800 MW thermal/generator {
With the above data, calculations were completed using the Westinghouse SG ,
1 performance computer code, GENF, to verify the plant hta and to establish a complete list of operating cr.nditions required for the ATHOS analysis. The [
GENF code detemines the primary side temperatures and steam flow rate required 3 I to obtain the specified steam pressure at the given power rating. Besides '
confiming these parameters, the code calculates the circulation ratio which is '
used to determine the total bundle flow rate and average loading on the tubes.
Note that the circulation ratif., calculation includes the effect of downcomer resistance plates which are present in all eight steam generatcrs of the two Zion units. The calculated circulation ratio along with the other thermal / hydraulic conditions are listed in Table 71. The ATHOS 3 0 flow field j calculation was based on these parameters.
t 0222M:49/082688 23
7.2 ATHOS Analysis Model The calculation of relative stability ratios involves comparing the stability .
ratio calculated for one or more tubes in a given plant to the ratio calculated
- for the ruptured Row 9 Column 51 tube in the North Anna Series 51 steam l generator. It makes use of ATHOS computed flow proffles for both tube bundles. Since the presence of AVBs in the U bend region of a tube bundle could influence the overall flow field and/or the local flow parameters for a r i particular tube of interest, some discussion of the treatment of AVBs is 3
necessary before presenting a description of the ATH0S model.
) The ATH0S code does not include the capability to model the presence of the
, AVBs in the U bend region. However, Westinghouse has modified the code to ,
include the capability to model the AVBs via flow cell boundary resistance factors. Practical lower limits of cell size in the ATHOS code, however, prevent a fine grid representation of the AVB V bar shape which, in turn, '
limits the accuracy of the AVB representation. ATHOS calculations have been !
performed with and without AVBs in the model. Calculations of stability ratios f relative to North Anna R9C51 show that the relative stability ratios for tubes !
! near the center of the steam generator are essentially the same for models with ;
or without AVBs. The ATHOS AVB modeling sensitivity studies with uniform
(
insertion show some tendency for the AVB resistance effects to lower tube gap [
I velocities near the central regions and to increase velocities near the !
{ peripheral tubes. However, the magnitude of this effect is uncertain due to l
the limitations in ATHOS for modeling the AiBs. Further, the global flow l resistance of staggered AVB insertion would be less than that from uniform insertion. Based on the sensitivity studies using ATHOS models with and '
) without uniformly inserted AVBs, the most reliable relative stability ratios l
(for actual steam generators with non uniform AVB insertion depths) are expected using ATHOS models excluding AVBs and effects of variable AVB insertion depths by using flow test results of actual AVB geometries. !
1 l l
0222M:49/072884 25 I
i
! The Zion analysis is based on a Cartesian coordinate system for the array of flow cells instead of the typical, and more widely used, cylindrical coordinate i system. With a Cartesian coordinate system the tube array.and any AV8s are l arranged in a square pitched configuration which is in line with the coordinate i axes. This alignment provides an improved representation of the tube region of I interest in the bundle.
l The ATH0S Cartesian coordinate system model for the Zion steam generator ;
) consists of 13.050 flow cells !.aving 30 divisions in the x axis (perpendicular i l to the tubelane) direction,15 divisions in the y axis (along the tubelane)
I direction and 29 divisioas in the axial (2 axis) direction. In the ATH0S
~
analysis, the steam generator is considered to be symmetrical about the x axis of the tube bundle. The model therefore, consists of one half of the hot leg f' and one half of the cold leg sides of the steam generator. Figures 7 1 and 7 2 show the plan and the elevation views of the model. These two figures show the layout of the flow cells and identify locations for some of the geometric
) features.
I !
l As shown in Figure 7 1, with the Cartesian coordinate system, the circular (
l wrapper boundary is represented by a step wise wall as indicated by the heavy t
) lines. All of the simulated flow cells outside the simulated wrapper boundary 1 above the first axial slab were biccked off by specifying extremely high flow i i
i resistances on the faces of the appropriate cells. Tubelane flow slots in the tube support plates are modeled also. (
l i
[
] Figure 7 3 reproduces the plan view of the model but with the tube layout arrangement superimposed. This figure illustrates the locations of the tubes g l
l in the various flow cells. The fineness of the cell mesh is evidents the l largest cells contain only to tubes while some of the smallest cells include
- only three tubes. Nota, in particular, that additional detail was added near l the bundle periphery (!Y=12 ll) to more closely model the inner radius tubes j
) (rows $15). Five axial layers of cells were included in the U bend near the j i
top tube support (Figure 7 2, !Z 16 to !! 21) to more closely model the flow i i
conditions in the area of interest. f l !
J i
r
! 0222M:49/072848 26 i
- . - _ - _ _ _ - . , . . _ - _ _ - _ - - - . -. _---J
J 7.3 ATH0S Results The results from the ATHOS analysis consist of the thermal-hydraulic flow l parameters necessary to describe the 3 D flow field on the secondary side of the steam generator plus the distributions of the primary' fluid and mean tube wall temperatures. Since the velocity components computed by ATHOS are defined
- on the surfaces of a flow cell, the tube gap velocity, which is the [
i appropriately interpolated cell velocity ratioed upward to account for the flow area between the tubes, and density distributions along a particular tube ;
I required for tube vibration evaluation are determined by a post processor from ,
i the ATHOS output. The post processor generates a data file which contains this
! information for all the tubes in the model and the file serves as psrt of the :
I input data required for tube vibration analyses. Because the majority of the l flow cells contain more than one tube inside a cell, the tube gap velocity and I density surrounding a tube are obtained by interpolation of the ATHOS j calculated velocities (defined on the cell surfaces) and density (defined at j j
the center of the cell). The post processor performs the necessary {
interpolations to determine in plane and out of-plane velocities at specific i intervals along the length of the tubes. i Figure 7 4 shows a vector plot of the flow pattern on the vertical plane of !
symmetry of the steam generator (the vectors are located at the center of the j l flow cells shown in Figure 7 2). It is seen that in the U bend region the j mixture turns radially outward, normal to the curvature of the bends toward the !
l region of least flow resistance (i.e., outside the done formed by the !
l U bends). Figure 7 5 shows the resultant vectors of the radial and j circumferential velocity components on the horizontal plane at Z = 21, the i sixth plane above the top tube support plate (see Figure 7 2). The radial f ,
) outward flow is more evident from this figure since it ignores the axial !
j component. It may be noted that the radial velocity at this axial location is
] low at the center c,' the bundle and increases with radius. Figure 7 4 shows that the extal component is about four times greater than the radial ;
]
component. Figure 7 6 shows the lateral velocities on the top of the l I tubesheet. l I I
{ Figures 7 7, 7 8 and 7 9 show a sample of the individual tube gap velocity and !
j density distributions along three tubes at Row 10. In each figure the gap j j velocity and density along the length of the tube are plotted from the hot leg i
- tubesheet end on the left of the figure to the cold leg end on the right. l
- 0222M
- 49/072848 27 j i .
I Figure 710 shows the plot of the average in plane gap velocity normal to the tube and density profiles as a function of the column number along Row 10. The ,
average values were taken as the numerical average of the parameter over the !
entire 180' span of a U bend at a given column location. The average velocity is seen to be relatively constant with values ranging from 8.5 to g.1 ft/sec. However, the average density varies across the bundle with lower values present in the interior of the bundle and higher values on the l periphery.
) 7.4 Relative Stability Ratio Over Operating History j One aspect of the evaluation of the Zion steam generators is to examine the ,
l operating history data and use it to determine the susceptibility to fatigue j from fluidelastic vibration resulting from the 15 years of operation. This i assessment has been completed through the use of a parameter termed the l 1 normalized stability ratio. The normalized stability ratio compares the -
l fluidelastic stability ratio for each period of a plant's uperation ,(fuel l l cycle) to a reference stability ratio based on a recent operating condition. A l l plot of this ratio against operating time, therefore, provides a relative I
! indication of the effect of past operation on the plant's fluidelastic stability ratio. This normalized time dependent ratio is subsequently combined .
with an absolute stability ratio for the reference operating point derived from ;
- i. detailed three dimensional thermal / hydraulic and tube vibration calculations, i High values for the net stability ratio, in particular, over a significant l
) period of operation, coupled with other prerequisite conditions (e.g., absence !
! of AVB support and denting at the top tube support plate), could indicate an !
l increased susceptibility to fluidelassic vibraticn instability and fatigue. I i !
The fluidelastic stability ratio is defined as the ratio of the effective fluid f velocity acting on a given tube to the critical velocity at which large
- amplitude fluidelastic vibration initiates
l
~
Fluidelastic Ueffective (
5tability Ratio. SR = (1) l I. !
U critical at onset of instability l l I
1 :
0222M:49/072888 28 f
l 1
In this ratio, the effective velocity depends on the distribution of flow velocity and fluid density, and on the mode shape of vibration. The critical velocity is based on experimental data and has been shown to be dependent upon the tube natural frecuency, damping, the geometry of the tube, the tube pattern, and the fluid density, along with the appropriate correlation coefficients.
The detailed calculation of this ratio using velocity and density distributions, etc., requit es three dimensional thermal / hydraulic and tube vibration calculations which are time consuming. Alternately, a simplified, one dimensional version of this ratio has been used to provide a relative assessment technique for determining the effect of past operation on the stability ratio. The normalized stability ratio is defined by the following equation:
- 6,5 i
(2)
In this equation "cyc x' refers to each fuel cycle and ' ROP' to the recent operating condition. While this simplified approach cannot account for l three dimensional tube bundle effects, it does consider the major operational parameters affecting the stability ratio. Four components make up this ratio:
2 a loading +.erm based on the dynamic pressure (pV ), a tube incremental mass (m) term, the natural frequency of the tube (fn ), and a damping ratio (6) term. It should be noted that the ratio is relative, in that each component is expressed as a ratio of the value for a given fuel cycle or power level to that of the recent operating point.
(
0222M:49/0728SS 29
l
)
- Ja.C. The l particular damping correlation which is used for all normalized stability ratio !
- calculations is based on a dented condition at the top tube support plate (a f E
clamped condition, as discussed in Section 5.2). The clasped condition is slso j assumed in calculating the tube natural frequency.
! l As discussed previously in Section 7.1, the reference three dimensional l j
stability ratio calculation for the Zion steam generators was based on the (
- following operating parameters which are representative of both recent and all
) previous full power operation in both units:
1 Steam Flow 3.4g x 100 lbsVhr
{ Steam Pressure 73ti psia t i circulation Ratio [ lac (Westinghousecalculation) 1 J
l In addition to the reference stability ratio for this full power condition, i
) relative stability ratios were generated for two lower power levels, 90 and 95% (
of full power. Since tube vibration and pos'ible fatigue are associated with s f
! operation at close to 100% power, only the higher power operating periods are l d considered important to the evaluation. The high power operating experience J for both Zion units is summarized in Table 7 2. It lists the number of days in j each fuel cycle that the unit operated within three high power intervals ;
y (85 90, 90 95 and 95 1005). Since the basic operating conditions for all ten f j fuel cycles are identical, the days within all the cycles can be combined and j j the resulting totals used in preparation of the stability ratio curves. !
Further, it has been conservatively assumed that the total operating time j within each of the three power intervals is assigned to the highest j power / stability ratie condittin in the interval. j i !
( The resulting normalized stability rettes for Units 1 and ! are shown in Figure l l 7 11. In this figure, the normalized stability ratio is plotted against l
! cumulative operating time above 865 power. The reference value ( l.00) is for j the full power operating condition on which the 3 0 stability rettes are (
based. Figure 7 11 indicates that Unit 2 has experienced more time at higher i L
i I 4
f r
\
l I
! 0222M:49/072848 30 !
t 4
I (
L___.______.____ _ - _ . _ _ - _ . . . _ . _
l I
stability ratios, compared to Unit 1. The reduced ratios at 90 and 954 power i are the combined result of both decreased loading on the tubes and increased ;
. damping. Higher damping is a result of lower voids in the U bind which occurs i when the steam pressure rises at reduced power levels.
i
)
References:
71 L. W. Keeton, A. X. Singha1, et al. "ATH0$3: A computer Program for i l
Thermal Hydraulic Analysis of Steam Generators', Vol. 1, 2, and 3 l URI NP 4604 CCM, July 1984, '
i t
i 1 I t
'l !
r
] l 4 o 4
I i 1
) !
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! (
4 !
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i {
l i 0222M:49/072884 31 t I
_ . , - - , - - - . - , , , - - , _ . , , ~ , . - , . - -- . , - , - . - - - . ,- - - .
.e .- . . . , _. - _ - . . _ , ., - . _, . -
L 4 Table 7 1 l t'
Zion Steam Generator Operating Conditions Used for ATH0$ Analysis f
! r I
Power 812.5 MWT 7
Primary Flow Rate 3.40 x 10 lb/hr l
{. !
- Primary Inlet Temperature 591'F i
I Primary Outtet Temperature 527 'F [
Steam Flow Rate 3.49 x 100 lb/hr l Feedwater Inlet 428'F
' Temperature i
) Water Level from Tubesheet 506.4 inches Steau Pressure at SG outlet 735' psia t i
Circulation Ratio 3.2 ;
l r I
I A '
i i
j ;
1 i 1 ,
1 I i l
'! l
- c i
I I
1 f
l I i !
i I
022;M:49/072444 32 i
. __.. . . - . . . - - - - - - . - . . __ - _ , _ . - . - - _ . - - - _ _ , _ _ . - _ , _ - _ _ _ - - , ._f
', T4ble 7 2 '
Zion Operating History Data l 1 l r
3 i
e
............... li. . n ! ................
1
- ............... ii . n i ................
tail at incu poeta ttytt I (Art at (4:n W tt ll ,
l l I l ti INI ';*1:1 l! 4 ** '
1 '
Cet;I ! HI (H 911001 90 *l 11 01 1 MI (4 i ......... Il li......... ......... ......... ...... ....... l ......... ......... ......... ...... .......
i3 it it) .
l hs 13 0!.e.r.;6 3 3 11 1 h !w 73 07 ha 71 '.
!!! 61 I Il ..r 77 $. FH 10 1.I !! il I I t bJee ?6 Obin *? ..
]
j I (.*:.t 11 16 59 *1 !?! 13 l$ ! 1. ast.'l ths.r ft I:
3 let H 6 I Il n er.79 ci=.,Is 1 .! ;
. I f1 % . ?1 96Oti?9 I l.8 !$ ;
5 I lb8H lJ !! ha ll !!? 13 to t !$ 2s1 90 ll les fl ? . E i
1 !! At e Il 13 FH l? h6 1 10 i el ter Il I. FH 83 til '
6
? I " '.1 il 01les13 !!! 9 16 I #7 n4, 13 37a.r96 !!) 1 18
. 1 09 hl 94 01 in l: !?6 !) to ,
I ! tb.et l. 31 hn 15 186 1
!.5 I f4 FM 16 21PrIf Hf e 1) !
1 1 18 .'.. l: e..In V6 Il 1 19 I l .'s. 17 f. f t il 11 111 16 I (I 8 49 11 ci 8., N 1 .
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. . . - _ _ _ . _ _ , _ _ _ _ . . - - - . . --_..-r_
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Figure 7 1 Plan View of ATH0$ Cartesian Nedel for tien i i !
i I I i i l r
r i r 0222M:49/071200 4 l t
d
. - _ . , - - , . , . - - - - - - . . - - . - - - . . , . . . . . , - - - . - - - - - . . , - - , - , - . . - - - , - . ~ - . - - - - - - - - - - - - - - - - - , - - - -, - - - . - . - - - - - .
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! Figure 7 t Elevetten View of AT M S Cartesian Itede) for Zion !
i ,
I i
-i l
0!!!M:49/0712M SS j t I i
i boCot t
4 I
t i
4 i
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s Figure 7-3 Plan View of ATHos cartesian Nedel for Zion i
- Indicating Tube Layout 0222M
- 49/071288 ~54 l
l sh,C ot, i
i i
l d
1 4
/
t l
Figure 7-4 Flow Pattern on Vertical Plane of Symmetry 1
\
l f
i 0222M:49/071288 37 1
d.C ,R s
r I
i i
/
s Figure 7-5 Lateral Flow Pattern on a Horizontal Plane in the U Bend Region 0222M:49/071288 36
6,C, e
?
/
Figure 7-6 Lateral Flow Pattern on Top of Tubesheet 0222M:49/071288 M
I l
l i
h ,C , 6.
/ w
/
N Figure 7 7 Tube Gap Velocity and Density Distributions I
for Tube Row 10/ Column 3 i
0222M:49/071288 40
i l
h C o t.
S f
1 a
t s '
/
l l
i
.i Figure 7-8 Tube Gap Velocity and Density Distributions for Tube Row 10/ Column to i
l 0222M:49/071288 41
l .
I b ,C o t s /
l l
l Figure 7-9 Tube tap Velocity and Density Distributions l for Tube Row 10/ Column 40 0222M:49/071288-4L 1
.1
1 l
/ O A,Q I 1
N I
/ l l
l l
l Figure 7 10 Average Velocity and Density in the Plane of the U Bends Normal to Row 10 0222M:49/071288 45
/
1.t 1.00
1.06 -
g i.o4 -
k 1.02 -
te -
r
=
j a94 - lpLMIT2 8
ass -
1 s au- m ' d. _9
) 092 -
8
'g 0.9 - fg
& cas - * ----+
3 ass- l l
au - '
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i o.s , , , , , , , , , , , , , ,
0 0.4 0.8 1.2 1.s 2 2.4 la I (thousande) e ="Tm om a nacs mno l
l I
Figure 7-11 Zion ihmnalized Stability Ratio Based on Hi @ Power (>45%) Operation 0222M:49/071288 &L.
l l _ . . _ . _ _ _ _ _ _ . _ _ _ _ . _ _ _ _ _ . _ _ . _ _ _ _ _ _ , _ _. ___
8.0 PEAKING FACTOR EVALUATION This section describes the overall peaking factor evaluation to define the test based peaking factors for use in the tube fatigue 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 reeatability '
o AVB insertion depth uncertainty 3.1 Nortn Anna 1 Configuration 8.1.1 Background The AVB configuration of the ruptured tube in North Anna, R9C51, is the reference case for the tube fatigue evaluations for other plants. In 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 R9051 tube, including the relative flow peaking factors.
Thus, the support conditions of the ROC 51 tube are fundamental to the analyses of other tubes. Because of the importance of the North Anna tube, the support conditions of this tube, which were originally based on "AVB Visible" interpretations of the eddy current test (ECT) data (Figure 8 1), were reevaluated using the projection technique developed since the North Anna event. The projection technique is particularly valuable for establishing AVB positions when deposits on the tubes tend to mask AVB signals such as found for the North Anna 1 tubes. The results of this evaluation are summarized below.
0222M:49/072888 44
l l
8.1.2 Description of the Method The basic method utilized was the projection technique in which the AVB position is determined based on measured AVB locations in larger row tubes in the same column. In this study, the projection technique was utilized in the "blind" mode, (AVBs called strit.ly based on the data) as well as the reverse mode (data examined on the basis of predicted AVB positions). The objective j
of this application was, with the greatest confidence possible, to establish the positions of the AVBs in an 8 column range around the R9C51 tube in North Anna 1, Steam Generator C.
8.1.3 Data Interpretation The ECT traces for the U bends in Rows 8 12 (in one case, 13) were examined for Columns 48 55. The original AVB visible calls are shown in Figure 8-1.
The data were examined by an eddy current analyst experienced in reading these traces, and by a des'gn engineer knowledgeable in the geometry of the Model 51 U-bend region.
The intent of this review was to determine if the presence or absence of AVBs as shown in Figure 8 1 could be confirmed using the AVB projection technique. Preliminary projected AVB positions were based on geometric data provided for a few of the tubes near R9C51. The features which were sought were avidence of data "spikes' where AVBs were predicted, offset indications (multiple spikes) where offset AVBs were predicted, single indications where single AVB intersections were predicted, etc. The data evaluation method used was a critical examination of the data, which was biased toward the presence of AVBs unless a confident call of "no AVB" could be made, and then checking the consistency of the data among the tubes in a column and against the theoretical data for the predicted AVB positions. (
0222M:49/072888 45
I l
I l
l l
b Ja,C,
. Figure 8 4 is the "AVB visible" map for columns 48 through 55, based on the critical review of the data. It should be noted that the original data interpretations and the review interpretations are consistent.
8.1.4 Projections i The [ Ja,cECTtraceswere utilized for projecting the position of the AVBs according to the standard ,
format of the projection method. i The results of the projections are presented in Figure 8-5, which shows a matrix of projections for tube rows 8 through 13 in columns 48 through 55.
For many of the tubes, more than one, and as many as three, projection values 0222M:49/072888 46
l l
l i
are shown. Multiple projections are expected for a tube if the AV8s on either !
side of Ds tube are not at the same elevation, or if the upper and lower AVB l supp:.. that tube. As many as four different projections are possible if it is assumed that the tube is supported by the upper and lower AVBs, and both upper and lower bars are staggered in elevation as shown in Figure 8 2.
The logic in arranging the projection data is based on the following two rules:
Rule 1. The projections of the same AVB based on different tubes in the same column [ Ja c,
[
ja,c, Rule 2. Two adjacent tubes in the same row [
-]ac. Consequently, the difference in the (
Ja.C, 0222M:49/072888 47
l l
l The implication of this is that if the position (either left or right) of a projected AVB is assumed for a column, then the projections in the l adjacent columns are also (
ja.c, 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 matr'x of Figure 8 5 is shown in Figure S 6.
8.1.5 Conclusions The general AV8 arrangement surrounding the ruptured tube in North Anna-1, Steam Generator C, which was the basis for the analysis, is confirmed by a detailed critical review of the ECT data. Differences exist in the AVB pattern between tube columns 48-49, in which the AVBs appear to be less inserted than previously indicated. The pattern of Figure 8 6 is the best fit to the rules which were adopted for determining the position of the AVBs, as well as consistent with explanation of the tube failure.
The basis of the review was a projection technique which utilizes data from tubes one or more rows removed from the actual inserted position of the AVB to determine the position of the AVB. The intent of the review was to establish the positions of the AVBs by confirming or eliminating features of AVB alignments such as side to side offsets, etc. of the AVBs adjacent to the tubes. Overall, the conclusions regarding the positions of the AVBs around ROC 51 in Nortn Anna-1, Steam Generator C are based on consistency among all the available data. ,
8.2 Test Measurement Uncertainties The descriptions of the peaking factor tests and apparatus were provided in Section 5.4. All practical measures were taken to reduce uncertainties.
Nevertheless, some still remain and should be properly accounted for. The important parameter measured during testing that has a significant impact on f
0222M:49/072888 48
peaking factor is the air velocity. The air velocity at test section inlet was measured using a (: Ja,c. Based on considerable experience with the use of such instruments, it is known that the magnitude of uncertainty is very small. A[ la,c measurement uncertainty is used in this analysis based 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,c s l 0222M:49/072288 49 l l
. _ - ._ - -- --. -_ I
. , a,c b
Fcr the purposes of this estimate, the geometry of the cantilever measuring tube in the air test model is compared with the geometry of a prototypical Row 10 tube. (
i i
I Ja,c, f
I The comparisen 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. (
c l
f i
l [
i I
\
t L
Ja.c j 0222M:49/072288 50 i
[
Ja c. Using these values, the ratio of the effective velocity for the cantilever measuring tube to that for the U bend tube is about !
( ja,c for the case treated.
A similar evaluation can be made for a Row 10 tube that lies in the projection or shadow of an AVB that is inserted to a depth reautred to support a Row 9 tube. [
ja,C, The net result is that the ratio of the effective velocity for the cantilever tube to that for the U bend tube is about [ la,c, These results indicate that, for the particular assumptions used, the cantilever tube model appears to be a reasonable represent & tion of the U bend
- with respect to detennining relative peaking factors for different AVB l configurations. This evaluation also shows that, on the average, the magnitude of the systeteatic uncertainty associated with the use of cantilever l tube to simulate the U bend is about [ Ja.c,
[
8.5 Air vs Steam Water Mixture [
l
.he local peaking factors from the air tests can bs applied to the steam ;
generator steam / water conditions either as a direct factor on the mixture i !
j 0222M:49/072888 51
- _ _ .~ , , _ _ _ _. _
l 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 perking factor compared to the direct air test value. For estimating an absolute stability ratio, this application of the peaking factor is a best estimate approach. However, for the evaluation of tubes relative to stability ratio criteria, it is more conservative to minimize the peaking factor for the North Anna Unit 1 tube R9C51 through direct application of the air test peaking factor. This conservative approach is therefore used for evaluating tube acceptability.
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 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 pa'ssage and thus turns into adjacent open channels. Two phase tests indicate a tendency )
for steam to preferentially follow the low pressure drop path compared to the water phase.
Based on the above discussion, the Fj are considered to more appropriately ;
i apply to the steam phase. Thus, it follows that mixture mass velocity for l the tube subject to flow perturbation can be written as follows- l a <
a,c l
l l
l where 90 is the vapor density, Of the water density, F, the velocity l i
peaking factor determined from air tests, jg
- the nominal superficial vapor ,
velocity, and jf* the superficial water velocity. Steam quality can then l be determined as follows*
l i 0222M:49/072288 52 !
l I
~
a,c The Le11ouche Zolotar correlation (algebraic slip model), as used in the :
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: l
~
- a,c
[
~
where 3F is the stability peaking factor, Fq the density scaling factor, ,
Fy the velocity scaling factor, and Fdp the damping coefficient scaling l factor. If we use the air velocity peaking factor without translating to steam / water conditions, then
- - a,c As shown in Table 81 stability peaking factort for the steaa/ water mixture are slightly higher than air velocity peaking factors. The difference '
betwean the steam / water and air peaking factors increases as the air peaking factor increases, l
- For application to tube fstigue 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 f assessment. The North Anna R9C51 peaking factor is one of the highest j I
] peaking factors. As discussed in Section 8.7, a peaking factor of nearly ja c is determined for the R9C51 tube. The differences between ( l Ja.c. Typical values are shown in Table 8 2. These results show !
d 0222M:49/072288 53 l
that the direct application of the air test data yields the higher relative peaking factor compared to R9C51. To obtain conservatism in the peaking factorevaluation,[
Ja,C, Comparing the values in the first and last columns of Table 81, it may be noted that the stability peaking factor for steam water is [ Ja,c higher than the air velocity 5,eaking factor. On the average, the uncertainty associated with the conservative use of air velocity peaking factor is
( Ja.c, 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 protoypic flow conditions. The objective of this study was to examine the magnitude of the changes in void fraction 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 model ngi 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 AVBs in the ATHOS model. The ATHOS results were processed by the FLOV!B code to determine stability ratios for the specific tubes of interest. The calculation was ,
repeated using a non uniform 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 [ ja.c. This alternate !
calculation provides independent corroboration of the prior discussion regarding the stability peaking factors under steam water conditions vs in air.
l 0222M:49/072888 54
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 tebes. The methodology used for obtaining the AVB insertion patterns from eddy current data can ascertrin the AVB location only to within approximately (.
Ja,c. The effect on peaking factor resulting from this uncertainty is addressed using test results of AVW configurations that varied from one another by up to ( Ja,c, I
Based on maps of AVB insertion depth of various plants, several configurations have been tested for determining fluidelastic 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 A?B ,
configurations tested.
l 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 [ Ja c. 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 se3n, a decrease in depth of an appropriate AVB tends to decrease the peaking factor, for instance,a(
Ja.c. Such a trend can be explained; a decrease in a specific AVB depth will open up more channels for incoming fluid to distribute and thus less flow perturbation. However, this applies only to those changes without inducing the reinforcement of flow perturbation from upstream to downstream.
On the ave; age, the uncertainty in peaking factor resulting from small ,
variations in AVB insertion (of the order of 1/2 tube pitch) is found to be
[ ja.C, t
0222M:49/072888 55
8.7 Overall Peaking Factor with Uncertainty
. As discussed in the previous subsections, there are several aspects to be considered in applying the laboratcry test data to steam generator conditions. These considerations were reviewed one at a time in those subsections. This section will integrate the pieces into one set of stability peaking factors.
. 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 j closely by 4c, la and 5e. This is encouraging in the sense that it tends to explain why, of all the tubes in servics, the R9C51 tut,e was the one to experience the fatigue, rupture.
i.
It is to be noted that the test results would be applied as ratios of a specific tube peaking factor to the R9C51 peaking factor. This will reduce a l the influence of some uncertainties since the systematic uncertainties would affect both the numerator and the denominator in the ratio of peaking !
factors. The major difference will be in those configurations whose peaking t factors are significantly lower than that of R9C51. The approach employed .-
here is intended to provide that conservative peaking facters are employed for (c h apparently low peaking configurations.
4 The uniform AV8 configuration (24) is selected as a reference configuration, and the peaking factors of all configurations tested er? 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 j peaking relative to the R9C51 configuration.
j The uncertainties in the test results and their extrapolation are these due j j to test measurements, test repeatability, cantilever tubes in the test vs i
! U tubes in the steam e nerator, and air tests vs steam. water mixture. These i i
t
, 0222M:49/072848 56
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 together. The total random uncertainties are calculated by (
t)a,c. The RSS value of these is
[ la.c. Since these can occur in any test, these are to be applied to all tests. One way of doing this is to apply it to the R9C51 value, that being ;
in the der eminator of the final peaking factor ratic. Thus the peaking factor for configuration la (R9C51) is reduced by this amount to yield a valueof( Ja.c instead of the ( la,c appearing in Table 5 2.
i The next three uncertainties in Table 8 4 are systematic uncertainties. It could be argued that these appear in the peaking factors of both the specific tube under consideration and the R9C51 tube and are therefore counter 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 (
Ja.c. 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,(
Ja.c. The peaking factor of the reference configuration 2a (Table 8 5) is raised by l thisamounttoavalueof( Ja,c, i
The change in peaking factors of configurations 1A and 2a resulting from the ;
I application of uncertaintiis as described above are shown in Column 3 of Table 8 5. The peaking factors of all configurations are recomputed on the basisofthisreferenceconfiguration(2a). These values are displayed in ;
Column 4 of Table 8 5. h I
i i
l 0222M:49/072888 57 ,
l Some of the uncertainties were applied to the reference configuration (2a) in i order to apply them to all low peaking configurations conservatively. Thus, no configuration should have a lower peaking factor than this reference configuration. Therefore, when a peaking factor value less than ! Ja,c is calculated for any configuration, (in Column 4 of Table 8-5), it should be l alteredto( Jac. Further, for some of the configurations that are ;
conceptually similar, the more limiting (higher) value is used. For example, a peaking factor of ( Jacisusedforconfigurations5aandSbbasedon their similarity to configuration Sc.
The final stability ratio peaking factors calculated on this basis (with configuration 2a as the reference) are shown in Table 8 6. It may be noted that the peaking factors vary in the range ( la,c, the R9C51 peaking factor being ( Ja.c. Figure 8 7 shows the final peaking factors I with the pictorial representation of the AVB insertion patterns.
i Table 8-7 shows the result of applying the peaking factors to specific tubes in the Zion 1 and 2 steam generators.
The overall conclusions from the peaking factor assessment are: '
- 1. As noted in Tab'e 8 4, five elements have been included in the uncertainty eva uation for the peaking factors. The uncertainty estimates were developed from both test and analysis results as described in Sections 8.2 to 8.6. Thelargestsingleuncertaintyof( Ja c is attributable to uncertainties of up to ( ]a,c on :
determination of AVB insertion depths from field addy current data. This relatively large uncertainty is applicable only to low peaking conditions where the AVB uncertainties can contribute to small peaking factors. The definition of 'no flow peaking' was increased to encompass the small peaking effects from AVB insertion uncertainties. For the AVB patterns l
leading to significant peaking factors AVBs were positioned within i
uncertainties to maximize the peaking factor. For these configurations, j
variations of AVB insertion within these uncertainties are expected to j reduce the peaking factor compared to the final values of Table 8 6 and Figure 8 7.
l I
l 0222M:49/072848 58
- 2. Including uncertainties directed toward conservatively decreasing the i peaking factor for the North Anna tube R9C51, the final R9C51 peaking ;
factoris[ ]a,c 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 I consequence of the conservative uncertainty analysis and is not likely to !
be representative of actual peaking effects.
i i
i j
j l
j
\ ,
, l l
l i
l l ,
1 '
l l l l 1
l ;
i !
j i 0222M:49/072888 59 l P
Table 8 1 Stability Peaking Factor Due to Local Velocity Perturbation Scaling Factors for Steam / Water ;
Air ,
Velocity void Stability Peaking Fraction Density Velocity Damping Peaking ,
Factor, Scaling, Scaling, Scaling, Scaling, Factor, 1 F, Fy Fd Fy Fq, F s
( 5 a,c ,
i i
1 j
NOTE: 1. Stability peaking factor for steam / water mixture is i calculated as follows: l 1 ' % a,c 4
i i ,
- 2. Damping scaling factor is calculated using modal
) effective void fraction of [0.839)a,c for R9C51 tube. l I
1 !
3 i
4 i
I t
l 0222M:49/072288 60
- Table 8 2 Comparison of Air and Steam water Peaking Factor Ratios I Air Air Steam Steam a
Peaking Peaking Peaking Peaking Factor Ratio Factor Ratio P e $ a.C ,
i f
6 s 4
t I
l l
i l
l a
l i 1 J
l i i i
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I l
l l l 1 i l 0222M:49/0722H 61 l a r 1 l
r i
Table 8 3 i Effect of Local Variation of AVB Insertion [
i l
'l A to 8 AVB Peaking Peaking Ratio l.
Type A Type B Variation Factor A Factor B (8/A) i
' .s a,c j i !
I J ;
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b l <
a.C J
L I
J t
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i i i r l !
4 i
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L I i I
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l l I l
i L
[
i l
! i l r 0222M:49/072288 42
{
i
Table 8 4 1
Uncertainties in Test Data and Extrapolation Source of Uncartainty Iygg Maanitude. % I
- $ a,c 1.
2.
3.
4.
5.
I i
i l
i i i l I i
l I' 1
i !
l 1
l
- This is not an uncertainty associated with the test data.
1 It results from the inaccuracy in determining the true AVB ;
position in the field using eddy current data.
i i
l l
0222M:49/072288 63
i i l 1
Table 8 5 i
Extrapolation of Test Results to Steam Generator Conditions l
^
Peaking Factor Test Data with Referenced to (onfiauration M Uncertainties Confia, 2a
- g a,C
)
L I
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i i 0222M:49/072288 64 r
1 Table 8 6 l
f FINAL PEAKING FACTORS
! t I I Confiauration Peakina Factor b
t
,
- a,c 1
o
! \
4 l
e r l
1 i
l 4
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2 h P l
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l 0222M:49/072288 65
(
i
Table 8 7 Stability Pecking Factors for Specific Tubes Zion Unit 1 Steam Peaking Generator Row No Column No Factor
- a,C A 8 35 34 10 43 42 41 All of the Remaining B 8 69 61 60 35 18 9 69 60 18 i All of the Remaining !
l C 8 61 60 35 9 60 10 53 49 1 All of the Remaining
, D 8 70 69 >
35 9 46 All of the Remaining s <
- The peaking factor is divided by [ Ja.c to obtain the relative flow i i peaking factor to R9C51 of North Anna 1. !
i l
1 0222M:49/072288 66 l
Table 8-7
- Stability Peaking Factors for Specific Tubes ,
Zion Unit 2 -
i Steam Peaking Generator Row No Column No Factor * ;
1 , , a.c l ,
)
A 9-11 All '
i i B 9 11 All C 9 11 All r
0 11 90 89 l
43 l
l 7 i
All of the Remaining 4 dl j 1
- The peaking factor is divided by ( Ja.c to obtain the relative flow 1 peaking factor to R9C51 of North Anna 1.
J i
l
[
t I
i 0222M:49/072288 67 !
)
- '= 00 00 0 OO OO 00 00 0 00 00 l
'o 00 00 0 G 00 00.0000 00000 l 00.00@@0000000 88 ' 85 ' 54 ' 53 52' 51 50 45 ' 48 47 45 45 ' 44
/ ;
O ^ISIBLE V' ve e Fritmo rust l V O ^ INVISIBLE
$ PLUGGED l
! l
! l 1 .
l l
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j !
Figure 8 1 Original North Anna AYS Configuration j (Configuration ib) l l :
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Tigure 8-2 Schematic of Staggered AVBs l 1
1 1
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i Figure 8 3 AV8 'Patr* in ICT Trace r l
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l ,-.. . . _ . . __ - _ _ _ - , , ,
. . - - . - - - . . . .,_,--r. . . - , - . v, - - - - _ _ , - . . . - . - - , .- , , , , , . , _ . . _ , . ., _ - - . , , ,, . _ _ . , _ , . - . _ . - - - - . - - - _ . . - , ,
! % u G u l 22 O,O@OO OGO OOOOO l u0 0000 0G0 00000 ;
1 "0 000000@ O0000 SO@@@@O000000 l . '
OGGOGGOOOOOOO
\ C4kmnl 56 SS 54 53 52 51 80 49 48 41 46 45 44 i :
g PWgged Tube g FaAedTube
' l wo m m a e e w l w mnr.n . a.u r.i m .r r. 4e m z w.m an s e . :
based on artcalreview of the ICT recas Open airWe in Ws rings meare ne das k l i Waid,W.
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1 Figure 8 4 North Anna 1, steam Generator C. AVI Positions '
Critical Review ?AVB Visible' Calls l
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HM WB E .... M.....
E R !! M H... E E... R a
_ _ ... ... ... ... .... .... .,. ,... M<... ..
! nn - - 8.M. ..5..E....
- u. l 5 5 5....
E I 5 3 E 13 28
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iiiih LL .
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, .%(i iR11 J". L AL L Ll = =
LL C55 C54 .C53 C32 C50 C51 Cat C44
=:,t g e
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- u. - -
a,e a,c !
l 5 w a.-l ; g w u.; -
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Figure 8 5 North Anna,1. Steam Go,ne, rater C, RSC51 AVB Matrix '
t
0 00 0 0 0 0 22 O OO
! it O OO O C0000 .
to O OO O 0 0,,0.0.O.
- O 00 0 0 0 0 0<O'0 0 0
.000000000d000 49 48 47 46 45 44 I
$6 55 54 53 32 51 50 k
1 i
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1 Figure 8 6 North Anna RfC51 AVB Final esitiens
TYPE OF AVS PEAXNG FACTOR TYPE OF AVS PEAXNG FACTOR msERTx:w usa 3 TION
'8,b,c '
h , b , c, i
i f
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4 i
, s s n ,
Figure 8 7 Final Peaking Factors for Zion 1 and 2 i
I 9.0 STRUCTURAL AND TUBE VIBRATION ASSESSMENTS ,
9.1 Tube Mean Stress i
This section sumarizes the analysis to determine stresses in a dented tube but undeformed at 100% power. Loads impored 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
( la,c at cold shutdown.
A sumary of the temperature and pressure parameters at 100% power in the vicinity of the top support plate are provided in Table 91. The tube temperature corresponds to the average of the primary side water temperature and the plate temperature. The resulting tube / plate radial interference is j
( Ja.c,
, Stresses due to differential pressure and interference loads are calculated j using finite element analysis with the model shown in Figure 91. The model J prescribes (
l Ja.c Two reference cases were run using the finito element model, the first for a primary to stcondary side pressure gradient of 1000 psi, and the second for a ,
( Ja.c inch radial interference between the tube and plate. The pressure ;
case incorporates the axial load on the tube by applying a pressure loading i along the top face of the model. Plots showing the distribution of stress for i
, the tube outer surface for the two reference cases are provided in Figures 9 2 and 9 3. Thermal bending stresses due to the thru wall thermal gradient are -
calculated to be 8.05 ksi using conventional analysis techniques. The combined l
] stress distribution along the tube length, in Figure 9-4, was obtained by i combining the thermal bending stresses and the reference solutions with :
appropriate multipliers based on 100% power operating parameters.
0222M:49/072888 68
i The maximum axial tensile stress is 22.5 ksi and occurs approximately 0.125 inch above the top surface of the support plate. Adding, for conservatism, the surface stress due to pressure, 0.74 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 are compressive at the tube surface, but 5 10 mils below the surface, the stress levels change to 10-15 ksi tensile. Combining the applied and residual stresses results in a cumulative !
mean stress of approximately 38 ksi, assuming tube denting without deformation.
If a tube is dented with deformation, the mean stress is limited by tube yielding. For the case of dented tubes with deformation, the maximum effect of mean stress was incorporated by using amax " 'y in determining stability ratios and fatigue usage.
9.2 Stability Ratio Distribution Based Upon ATHOS An assessment of the potential for tubes to experience fluid elastic instab'lity in the U bend region has been cerfomed for each of the tubes in 1 rows eight through twelve. This analysis utilizes FASTVIB, a Westinghouse proprietary finite element based computer code, and PLOTVIB, a post processor l to FASTV!B. These codes predict the individual responses of an entire row of steam generator tuing exposed to a location dependent fluid velocity and density profile. The program calculates tube natural frequencies and mode I shapes using a linear finite element model of the tube. The fluid elastic stability ratic U ,/Uc (the ratio of the effective velocity to the critical i
velocity) and the vibration amplitudes caused by turbulence are calculated for
)
I a given velocity / density / void fraction profile and tube support condition. The velocity, density and void fraction distributions are determined using the f ATHOS computer code as described in Section 7.3. The WECAN generated mass and l stiffness matrices used to represent the tube are also input to the code, i
(WECAN is also a Westinghouse proprietary computer code.) Additional input to l FASTVIB/PLCTV!B consists of tube support conditions, fluid elastic stability l
constant, turbulence constants, and location dependent flow peaking factors.
l l
l I
0222M:49/072888 69
This proc 6ss was performed for the Zion Units 1 and 2 steam generator tubes and also for the North Anna Row 9 Column 51 tube (R9C51) using similarly appropriate ATHOS models. Ratios of the Zion Units 1 and 2 results to those for North Anna Unit 1 R9C51 were generated to produce a quantity that could be used to provide an initial assessment of the Zion Units 1 and 2 tubes relative to the ruptured tube at North Anna Unit 1.
Figures 9 5 and 9 6 contain the results of this process for each of the rows under investigation for Units 1 and 2, respectively. The relative ratios are obtained using the following conditions for both Zion units and North Anna Unit 1:
- 1) Tube is fixed at the top tube support plate,
- 2) Void fraction dependent damping,
- 3) No AVB supports are active,
- 4) location dependent flow peaking factors.
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. . .
These figures indicate that with the exception of the two tubes, R12C3 and R12C92 (both are supported in all SGs), all tubes in Rows 8 thru 12 of the two units lay below the 90% line.
9.3 Stress Ratio Distribution with Peaking Factor '
i An evaluation was performed to determine the ratio of the Zion Unit I and 2
) tube stress over the North Anna R9C51 tube stress. This ratio :
1 i
0222M:49/072888-70 i /
is determined using relative stability ratios discussed in the previous section, relative flow peaking factors (Table 8 7 factors divided by
( Ja.c) 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 presented below are based upon the following conditions:
- 1) Tube is fixed at the top tube support plate,
- 2) Damping is void fraction dependent.
- 3) Tubes have no AVB support,
- 4) 10% criteria with frequency effects,
- 5) Tubes are assumed to be dented or undented (both situations were considered, but the evaluation is based on the more limiting, dented case).
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 Itsted above and subject to confirmation of fatigue usage acceptability. Conformance to these requirements implies that the stress acting on a given tube is expected to be insufficient to produce a fatigue event in a manner similar to the rupture that occurred in the R9C51 tube at North Anna Unit 1.
Figures 9 7 and 9 8 show the results of the stress ratio calculations for each of the Zion Unit I and 2 tubes in Rows 8 through 12. These ratios are applicable for tubes that are dented (tube deformation) at the top tube support plate. This case bounds the clamped tube condition with no tube deformation, i.e., the case corresponding to the NRC definition of denting with top tube support plate corrosion plus magnetite in the crevice without tube deformation. The current tube conditions at Zion correspond to this latter definition of denting.
0222M:49/072888 71
/
As can be observed in Figures 9 7 and 9 8, all tubes in Rows 8 through 12 of both Zion units fall in the acceptance region with respect to U bend fatigue, even when assumed to be unsupported. For Unit 1 (original AVB design), all tubes in Row 11 and above are supported. For Unit 2 (replacement AVB design),
with the exception of the two and tubes in Columns 2 and 93 of Row 12 of each SG, all tubes in Row 12 and above are currently supported.
9.4 Cumulative Fatigue Usage All tubes that are unsupported and have a stress ratio s 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 Zion Units 1 and 2 tubing are based on the current operating parameters and with future operation on the same basis, the tubes are not expected to rupture as a result of fatigue if 1) they meet the stress ratio criteria of 11.0 and 2) their current and future fatigue usage will total less than 1.0.
All tubes in the evaluation have conservatively been considered to be dented with deformation. Based on the above analyses, all Zion 1 and 2 tubes meet the relative stress ratio criteria under the current AVB conditions. Table 9 2 provides a sununary of the combined relative stability ratios and the stress ratios for the more salient unsupported tubes in Rows 8 through 12.
9.4.1 Unit 1 Tubing Acceptability of the Zion Unit 1 tubing for fatigue is accomplished by demonstrating the acceptability of the tubes with the highest stress ratio, 0.49 in $GC:R9C60. Assuming the tubes have been dented since the first cycle and continue to operate under current conditions, the total usage including the remaining ters of the operating license would be 0.04. In the event of a future uprating of the plant, the potential for tube fatigue must be re evaluated.
0222M:49/072EM ?1 2
f' t
i 9.4.2 Unit 2 Tubing j 1 ,
Using a similar approach, the maximum calculated total usage is 0.014 for the !
]'
highest stressed Unit 2 tubing at location $G D:R11C89 with stress ratio of !
0.41. This calculation assumed that tube loadings including local flow effects !
l prior to the AVB modifications were same as those for the current !
j configuration. The fatigue usage prior to the modification was not calculated j precisely since the old AVB positions and associated peaking factors were not l j determined. However, the following considerations justify tube acceptability l
{ with unknown prior fatigue usage, j l !
9.4.3 Tube Acceptability with Prior Fatigue Usage
?
l Currently, with the exception of both end tubes in Row 12 (i.e., R12C2 and [
]
R12C93), only Row 11 and shialler tubes in the lion Unit 2 are unsupported and !
j subject to future fatigue usage. As discussed above, these tubes meet stress [
8 ratio criteria based on continued operation at the same stability ratio, and l the predicted future usage for the limiting tube ($G D:R11C89) is about 0.0005 [
per year. !
I By the prior AVB design Row 11 tubes are expected to have had AVB support and j l
consequently negligible prior fatigue usage. However, AVB position evaluations l l
l for other units have indicated an occasional Row 11 tube that was not f l supported. Based on a sample of 2852 Row 11 tubes tn similar model steam {
generator designs, less than 0.5% of the tubes have been determined to be )
unsupported. The percentage of unsupported tubes in Row 10 is somewhat higher l l on the order of 3.35; however, the stress ratio for a smaller tube is !
! progressively lower as shown in Figure g 6. Also, based on prior analyses, only a small fraction of unsupported tubes have been judged to have flow j peaking factors large enough to result in relative stability ratios of greater l
! than 0.9. Therefore, the probability of Row 11 and smaller tubes in Zion l Unit 2 steam generators having a significant prior fatigue usage and/or a large l
! crack is very small. !
l !
l !
l l l
l l.
0222M:49/0728M 73 m
Table 9 3 susuarizes the acceptance basis of the limiting Row 11 tubes for Zion Unit 2 based on the fact that Row 11 tube stress levels, expected under the 4
current conditions, would not lead to rapid propagation of existing but
, undetected cracks. >
If a crack does not exist now, the fatigue usage in the tubes is probably low.
j Sut if a crack were to initiate during subsequent operation, the crack growth i
) rate would be insignificant since even a 30' (0.22 in.) thru wall,
! circumferential crack does not result in crack tip stresses that are above the l l threshold for crack propagation, 4.5 ksi in. For reference, a 30' f thru wall crack is estimated to allow a leakage rate of 35 to 40 gallons per l l day (gpd). l r
, If it is postulated that a crack does exist now in some tube, the crack size ;
) would be bounded by the crack that was in the tube at R9C51, North Anna 1. ;
j before it began to leak. That crack is estimated to have taken about 12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> l j to propagate from its initial thru wall length to a length that allowed a -
1eakage rate of 500 gpd. The most highly loaded Row 11 tube has a stress [
amplitude of 11.7 ksi. Based on a comparison to the stress amplitude that !
- caused the rupture of R9C51 North Anna tube, the time to reach the same level i of Technical Specification leakage rate in a Row 11 tube of the Zion Unit 2 ;
steam generators would be greater than 780 days. In other words, even with the
- bounding crack size of North Anna R9C51 tube, the crack tip stress intensity in the limiting Zion tube is too small to cause (rapid) crack propagation, f f
1 <
e :
3
) -
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! i 1 1 0222M:49/072888 74 l
t Table 9 ; :
t 100% Power Operating Parameters l l Zion Units 1 and 2 l Bounding Values for all Cycles i
< i l Primary Pressure = 2250 psia
! Secondary Pressure - 735 psia l Pressure Gradient = 1515 psi
- 1 2
t l
I !
i Primary Side Temperature * = 559'F ;
- Secondary Side Temperature 508.6'F !
t i
Tube Temperature = 533.8'F ;
l a i a :
Average of Thot = 591'F and Tcold = $27'F. !
I i l :
1 i
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0222M:49/072888 75
Table 9 2 Summary of Zion Evaluation of the Saltent
- Unsupported U bends i
i UNIT / RELATIVE STRESS L 13 E STABILITY RATIO (I) E ll) a
- Zion Unit 1 l
SG.A 8 34 .546 .08 I l [
SG B 4 18 .611 .20
, 60 .549 .08 61 .546 .08 l l.
69 .651 .20 ;
9 18 .751 .37
! 69 .788 .44 SG C 8 60 .549 .08 <
61 .549 .08
- 9 60 .792 .49 <
SG D 8 35 .709 0.33 l
70 .554 .09 ;
I All SGs 10 All 1 44 s .16 l ;
i
.I Zion Unit 2 ,
l i
SG 0 11 89 0.81 0.41 !
90 0.81 0.40 l All SGs 12 2 & 93 0.70 0.15 J
l I l !
! (1) All ratios are in comparisor.
- R9C51, North Anna 1, Steam Generator C. i i !
I a !
I j 0222M:49/0/2848 76 ;
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I i Figure 9 3 Dehted Tube Stress Distributions t i
Interforence Load on Tube i i
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1 Table 9-3 i
Disposition Criteria Relative to tube U Bend Fatigue Zion Unit 2 Present Condition Pre Mod Condition Diseosition tull !
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AV8 Support AVB Support Accept Stable Tube
! No AVB Support Accept 1, 2, 3 No AVB Support .
j Stress Ratio <1 AVB Support Accept 4
! No AVB Support Accept 4 !
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) Stress Ratio >l ... Sentinel- l l Plug (
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- 1. Tube is currently stable and has no indication of cracking. !
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- 2. Fatigue usage due to normal operating loads, excluding tube vibrations, is less than .001/ year. !
l l 3. If it is assumed that prior to AVB modifications the tube accumulated l fatigue usage such that potential to develop a crack exists, the rate of i
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- propagation will be very low. Leak monitoring and EC inspection during l
?! scheduled outages will provide adequate control for detection and/or an !
l orderly shutdown. :
4 l
4 Based on the current methodology / criteria, alternating stress is less than t 4.0 kst due to fluidelastic vibrations. A crack, should one develop during
(
operation, will not propagate at a rapid rate, thus allowing detection !
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and/or an orderly shutdown. For a currently unsupported tube, the maximum
! future usage assuming denting and full power operation is .0005/ year (at (
j location SG 0:RllC89). [
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! i 1 i j 0222M:49/072888 77 i __-_____-__t
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0238M:49/063088 63
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i 4 j Figure 9-2 Dented Tube Stress Distributions Pressure Load on Tube i
0238M:49/063088 64 i
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