ML20235M911
| ML20235M911 | |
| Person / Time | |
|---|---|
| Site: | Catawba, McGuire, 05000000 |
| Issue date: | 02/20/1989 |
| From: | Conners H, Curlee N, Pitterle T WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML20011C557 | List: |
| References | |
| WCAP-12126, NUDOCS 8902280494 | |
| Download: ML20235M911 (150) | |
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L WCAP 12126 WESTINGHOUSE CLASS 3 S
> m3. [
CATAWBA UNIT 1 EVALUATION FOR TUBE VIBRATION INDUCED FATIGUE AUTHORS: H J. CONNORS J. L. HOUTMAN
? N.'J. CURLEE M. H. HU
!* T. M. FRICK A. Y. LEE b: J. M. HALL R. M. WEPFER G. W. HOPKINS APPROVED:
- 4 T. A. PITTERLE,VMANAGER STEAM GENERATOR ENGINEERING WESTINGHOUSE ELECTRIC CORPORATION NUCLEAR SERVICE DIVISION i P.O. BOX 3377 ;
PITTSBURGH, PENNSYLVANIA 15230 0371M:49/0203891
_ - - _ _ _ _ _ _ _ _ _ _ - - - - - - _ _ _ _ _ _ _ _ - - _ _ _ _ _ - _ . _ _ _ _ __--_._)
t i
l ABSTRACT l- On July 15,'1987, a steam generator tube rupture event occurred at the North Anna Unit J plant. The cause of the tube rupture has been determined to be high cycle fatigue. The source of the loads associated wiQ the fatigue mechanism is a combination of a mean stress level in the tube with a I superim' posed alternating stress. The mean stress is the result of applied loading, manufacturing-induced residual stttss and denting of the tube at the top tube support plate, while the alternating stress is due to out-of-plane I deflection of the tube U-bend attributed to flow induced vibration. For tubes without AVB support, local flow peaking effects are a significant contributor to tube vibration amplitudes.
This~ report documents the evaluation of steam generator tubing at Catawba #1 for susceptibility to fatigue-induced cracking of the type experienced at North !
Anna Unit 1. The evaluation utilizes operating conditions specific to Catawba
- 1 to account for the plant specific nature of the tube loading and response.
The evaluation also includes reviews of eddy current data for Catawba #1 to
) establish AVB locations. This report provides background of the event which occurred at North Anna, a criteria for fatigue assessment, a summary of test data which support the analytical approach, field measurement results showing AVB positions, thermal hydraulic analysis results, and calcuhtions to determine tube mean stress, stability ratio, tube stress ratic, and accumulated fatigue usage. This evaluation concludes that one of the tubes is potentially susceptible to fatigue and requires corrective action.
7 0371M:49/020 $ -2
SUMMARY
OF ABBREVIATIONS ASME. -
American Society of Mechanical Engineers ATHOS -
Analysis of the. Thermal Hydraulics of Steam Generators
.- 'AVB -
Anti Vibration Bar AVT -
All Volatile Treatment ECT -Eddy Current Test EPRI- -
Electric Power Research Institute FFT - Fast Fourier Transform FLOVIB -
Flow Induced Vibrations MEVF -
Modal Effective Void Fraction OD -
Outside Diameter RMS -
Root Mean Square SR - Stability Ratio TSP - Tube Support Plate
'F - degrees Fahrenheit hr -
hour ksi - measure of stress - 1000 pounds per square inch lb -
pound mils - 0.001 inch MW -
mega watt psi - measure of stress - pounds per square inch psia -
measure of pressure - absolute i
)
0371M:49/020389-3
1
')
i TABLE OF CONTENTS 1
-SECTION e
t-1.0 Introduction 2.0 Summary and Conclusions
2.1 Background
2.2 Evaluation Criteria 2.3 Denting Evaluation 2.4 AVB Insertion Depths 2.5 Flow Peaking Factors 2.6 Tube Vibration Evaluation 2.7 Overa11 ' Conclusion 3.0 Background 3.1 North Anna Unit 1 Tube Rupture Event
~
3.2 Tube Examination Results 3.3 Mechanism Assessment 4.0 Criteria for Fatigue Assessaient ,
4.1 Stabilit,y Ratio Reduction triteria ,
4.2 Local Flow Peaking Considerations 4.3 Stress Ratio Considerations 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 5.4 Tests to Determine the Effects on Fluidaladic Instability of Coluenwise Variations in AVB Insertion Depths 5.5 References i
O' l
l I
0371M :49/020323-4 L .-
TABLE OF CONTENTS (CONT 1 HUED)
. gfCyg3 6.0 Eddy' Current Data and AVB Positions I
6.1 Catawoa #1 AVB Assembly Design I 6.2 Eddy Current Dhta for AVB Positioni f1 6.3 Tube Denting at Top Tube Support Plate l 6.4 AVB Nap Interpretations 7.0 Thermal and Hydraulic Analysis 7.1 Catawba #1' Steam Generator Operating Conditions 7.2. ATH0S Analysis Model
}
7.3 ATHOS Results 7.4 Relative stability Ratio ~0ver Operating History i l 8.0 Peaking Factor Evaluation 8.1 Ncrth 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 8.8 Peaking factors for Specific Tubes 9.0 Structural and Tube Vibration Assessments 9.1 Tube Mean Stress 9.2 Stability Ratio Distribution Based Upon ATH0S 9.3 Stress Ratio Distribution with Peaking Factor 9.4 Cumu'lative Fetigue Usage e
0371M:49/020389-5 s
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" .AppendixkI Evaluation of the .[ Ja,b,c Cable Tube Damper i
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I 0371M:49/Ct03?9-6
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s LIST OF FIGURES T
o{, 1 FIGURE
^
3-1 Approximate Mapping of Fracture Surface of Tube R9051 S/G "C" Cold Leg,- North Anna Unit 1 3-2 Schematic Representation of Feattres Observed During TEM y p Fractographic Examination of Frtcture Surface of Tabe R9C51, S/G '
"C" Cold Leg, North Anna Unit 1 L
\
3-3 Calculated and Observed Leak Rates Versus Time .;
4-1 Vf bration Displacement vs. Stability Ratio 4-2 ' Fatigue Strength of Inconal 600 in AVT Water at 600*F
[, 4-3 Fatigue Curve for Inconel C00 in AVT Water Comparison of Mean Stress
- l. ' Correction Models 4 Modified Fatigue with 10% Reduction .in Stability Ratio for Maximum Stress Conaition l
I 4-5 Modified Fatigue with 5% Reduction in Stability Ratio for 1 Minimum Stress Condition l
l S-1 Fluidelastic Instability Uncertainty Assessment 5-2 Instability Constant - S .
i l
5-3 Instability Constants,. $, Obtained for Curved Tubes from Wind funnel t
Tests on the 0.214 Scale U-Bend Model a
j_ 5-4 Damping vs. Slip Void Fraction 0371M:49/0iC389+7
.; I 1
LIST OF. FIGURES (Continued)
FIGURE 5-S Overall View of Cantilever Tube Wird Tunnel Model 5-6 Top View of the Cantilever Tube Wind Tunnel Model 5-7 Fluidelastic Vibration Amplitude with Non-Uniform Gaps 5-8 Typical Vibration Amplitude and Tube /AVB Impact Force Signals for Fluidelastic Vibration with Unequal Tube /AVB Gaps ,
5-9 Conceptual Design of the Apparatus for Determining the Effects on '
Fluidelastic Instability of Columnwise Variations in AVB Insertion Depths 5-10 Overall View ,of Wind Tunnel Test Apparatus 5-11 Side View of Wind Tunnel Apparatus with Cover Plates Removed to Show Simulated AV8S and Top Flow Screen 1.
5-12 .AVB Configurations Tested for Catawbc #1 5-13 Typical Variation of RMS Vibration Amplitude with Flow Velocity for Configuration la in Figure 5-12 6-1 -AVB Insertion Depth Confirmation 6-2 Catawba #1- Steam Ger. orator A AVB Positions 6-3 Catawba #1- Steam Generator B AVB Positions 6-4 Cstawba #1- Steam Generator C AVB Positions 6-5 Catawba #1- Steam Generator D AVB Positions l 0371M:49/020389-8 l
L - -
L LIST OF FIGURES ~(Continued)
FIGURE
^
6-6' AVB Projection Depth = 9.00 6-7 AVB Projection Depth = 9.15 7-1 Plan View of ATH0S Cartesian Model 7-2 Elevation View of ATHOS Cartesian Model 7-3 Plan View of ATH0S Cartesian Model Indicating Tube layout 7-4' Flow Pattern on Vertical Plane of Symmetry
~
7-5 Vertical Velocity Contours on Horizontal Plane at the Entrance to
,- the U-Bend Region 7-6 Lateral Flow Pattern on Top of Top Tube Support Plate 7-7 Void Fraction Contours on Vertical Plane of Symmetry 7-8 Tube Gap Velocity and Density Distributions for Tube P.ow 10/ Column 5 7-9 Tube Gap Velocity and Density Distributions for Tube Row 10/ Column 20 i
r- 7-10 Tube Gap Velocity and Density Distributions for Tube Row 10/ Column 40 7-11 Average Velocity and Density in the Plane of the U-Bends Normal to Row 10 037Dt:49/02C389-9
1 LIST OF FIGURES (Continued)
FIGURE 7-12 Catawba #1 - Normalized Stability Ratio Based on High Power (>85%) Operation 8-1 Original North Anna AVB Configuration 8-2. Schematic of Staggered AVBs 8-3 AVB " Pair" in ECT Trace r .
. i North Anna 1,' Steam Generator C: AVB Positions Critical 8-4 Review "AVB Visible" Calls 8-5 North Anna 1, Steam Generator C, R9C51 Projection Matrix B-6 North Anna R9CSI AVB Final Projected Positions 8 Final Peaking Factors for'Cataw' oa #1 9-1 Axisymmetric Tube Finite Element Model 9-2 Dented Tube Stress Distributions - Pressure Load on Tube 9-3 Dented Tube Stress Distributions - Interference Load on Tube 9-4 Dented Tube Stress Distributions - Combined Stress Results Catawba #1 9-5 Relative Stability Ratio Using MEVF Dependent Damping - Catawba #1 9-6. Stress Ratio Vs. Column Number-Dented Condition - Catawba #1
, 9-7 Stress Ratio Vs. Column Number - Undented Condition - Catawba #1
'9-8 Maximum Allowable Relative Flow Peaking - Catawba #1 l
037Di:(9/020389-10 !
LIST OF FIGURES (Continued) fig _yEl m a,b,c A.1 A.2 A.3 A.4
.. g O
0371M:49/020389-11
LIST OF TABLES l
TABLE 4-1 Fatigue Usage per Year Resulting From Stability Ratio Reduction 5-1 Wind Tunne'l Tests on Cantilever Tube Model 5-2 Fluidelastic Instability Velocity Peaking Facter.s for Columnwise Variations in AVB Insertion Depths Catawba #1 6-1 One AVB Signals Determir.ed to be Supported 6-2 One AVB Signal Indicating Support for Flow Peaking Analysis 6-3 Summary Listing of Unsupported Tubes Catawba #1 7-1 Catawba #1 Steam Generator Operating Conditions and Comparison with
, Unit 2 Conditions Used in the 3D ATH0S Analysis 7-2 Catawba #1 Operating History Data 8-1 Stability Peaking Factor Due to Local Velocity Perturbation 8-2 Comparison of Afr and Stear, Water Peaking Factor Ratios 8-3 Effect of Local Variation of AVB Insertion l 8-4 Uncertainties in Test Data and Extrapolation !
l 8-5 Extrapolat-ion of Test Results to Steam Generator Conditions 8-6 Final Pecking Factors 87 Stability Peaking Factors for Specific Tubes j l
I 0371M:49/020389-12 1 I
l
_ __ _ _ _ _ - _ _ _ - _ . _ - _ _ _ -_ _ ___ _ A
)
LIST OF TABLES (Continued) l-~
TABLE l
9-1 100% Power Operating Parameters - Catawba #1 l 9-2 Catawba #1 Tubes with Significant RSR's or Stress Ratios i
.i
?
0371M:49/020389-13
i
]
l
1.0 INTRODUCTION
l l This report documents the evaluation of steam generator tubing at Catawba #1 for susceptibility to fatigue-induced cracking of the type experienced at North Anna Unit 1 in July,1987. The evaluation includes three-dimensional flow '
f 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 Catawba #1 in order to account for plant specific features of the tube loading and response.
Section 2 of the report provides a summary of the Catawba #1 evaluation results and overall conclusions. Section 3 provides background for the tube rupture event which occurred at North Anna Unit 1 including results of the examination of the ruptured tube and a discussion of the rupture mechanism. The criteria for predicting the fatigue usage for tubes having an environment conducive to this type of rupture are discussed in Section 4. Section 5 provides a summary of test data which supports the analytical vibration evaluation of the candidate tubes. A summary of field measurements used to determine AVB locations and to identify unsupported tubes is provided in Section 6. Section 7 provides the results of a thermal hydraulic analysis to establish flow field characteristics at the top support plate which are subsequently used to assist in identifying tubes which may be dynamically unstable. Section 8 presents an
! update of the methodology originally used to evaluate the tube rupture at North Anna Unit 1. The final section, Section 9, presents results of the structural and vibration assessment. This section describes tube mean stress, stability ratio and stress ratio distributions,-and accumulated fatigue usage, for the
. Catawba #1 steam generator small radius U-tubes.
A cable damper was developed to reduce vibration in tubes which are unsupported and subjected to relatively strong flow peaking. The cable damper and its performance are discussed in Attachment A.
0371M:49/020389-14
q i
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1
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2.0
SUMMARY
AND CONCLUSIONS i
The Catawba #1 steam generators have been evhluated for the susceptibility to a fatioue rupture of the type experienced at Row 9 Column 51 (R9C51) of Steam Gererator C at North Anna Unit 1. The evaluation used Eddy Current Test (ECT)
~
l data supplied by Duke Power Company and interpret ~ed by Duke Power and
!- Westinghouse.
2.1 Background
The initiation cf the circumferential crack in the tube at the top of the top tube support plate at North Anna 1 has been attributed to limited displacement, fluid elastic instability. This condition 1: believed to have prevailed in the R9C51 tube since the tube experienced denting at the support plate'. A combination of conditions were present that led to the rupture. The tube was not supported by an anti-vibration bar (AVB), had a higher flow field due to i 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)
. chemistry of the secondary water and the additional mean stress from the denting.
2.2 Evaluation Criteria The criteria established to provide a fatigue usage less than 1.0 for a finite period of time (i.e., 40 years) is a 10% reduction in stability ratio that provides at least a 58% reduction in stress' amplitude (to < 4.0 ksi) for a 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 l same fatigue criteria is applied as the principal criteria in the evaluation of l
Catawba #1 tubing.
l i
0371M:49/020329 15
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 Catawtia #1 tube divided by the stress amplitude for the North Anna 1, R905) tube.
Displacements are computed for unsupported U-bend tubes in Rows 12 and inward, (descending row number) using relative stability ratios to R9C51 of North Anna 1 and tn appropriate power law relationship based on instability displacement versus flow velocity. Different U-bend radius tubes and tube sizes will have different stiffness and frequency and, therefore, different stress and fatigue usage per year than the Row 9 North Anna tube. These effects are accounted for in a stress ratio technique. The stress ratio is formulated se that a stress ratio of 1.0 or less produces acceptable stress amplitudes and fatigue usage for the Catawba #1 tubing for the reference fuel cycle analyzed. Therefore, a stress ratio less than 1.0 provides the next level of acceptance criteria for unsupported tubes for which the relative stability ratio, including flow peaking, exceed 0.9.
The stability ratios for Catawba #1 tubing, the corresponding stress and
, amplitude, and the resulting cumulative fatigue usage must be evaluated relative to the ruptured tube at Row 9 Column 51, North Anna 1, Steam Generator C, for two reasons. The local cffect on the flow field due to various AVB insertion depths is not within the capability of available analysis techniques and is determined by test as a ratio between two AVB configurations. In addition, an 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 ratios that correspond to these stress amplitudes. Therefore, to minimize the influence of uncertainties, the evaluation of Catawba #1 tubing has been based on relative stability ratios, relative flow peaking f actors, and relative stress ratios.
The criteria for establishing that a tube has support from art AVB and therefore eliminate it from further considerations is that it must have at least one sided AVB support present at the tube centerline. The criteria is based on test results which show that one sided AVB support is sufficient to limit the I
i I
0371M:49/020389-1R
vibration amplitude for fluidelastic excitation. AVB support is established by analysis of eddy current (EC) measurements and is a key factor in determining the local flow peaking factors. The local flow peaking produces increased local velocities which cause an increase in stability ratio. A small percentage change in the stability ratio causes a significant change in stress amplitude. The relative flow peaking factor of Catawba #1 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 tuba bundle flow analysis, to obtain the combined relative stability ratio used in the stress ratio determination. '
2.3 Denting Evaluation The Eddy Current (EC) tapes were evaluated to determine the condition of the tube / tube support interface of the first row of unsupported tubes inboard of the AVBs. Analyses of eddy current (EC) data for Catawba #1 shows the presence of ' corrosion with magnetite' in roughly half of the tube / TSP crevices. One tube, SG-C Rll, C111 indicated a ' dent' (denting with deformation) signal at the top tube support plate. For conservatism in the evaluation, all of the
, tubes evaluated are postulated to be donted. The effect Of denting on the fatigue usage of the tube has been conservatively maximized by assuming the maxixin effect of mean stress in the tube fatigue usage evaluation and by incorporating reduced damping in the tube vibration evaluation.
2.4 AVB Insertion Depths The Catawba #1 SGs have two sets of Alloy 600 AVBs. The ' inner' AVBs have a 1 rectangular cross-section and extend into the tube bundle approximately as far as Row 10. They provide a nominal total clearance between a tube without ovality and the surrounding AVBs of { ]a, cinch.
The outer AVBs also have a rectangular cross section, and extend into the tubc bundle approximately as far as Row 21, 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 AiBs are required.
0371M:49/020389-17
l The eddy current data were a'nalyzed by Duke Power and Westinghouse to identify f
j the number of tube /AVB intersections and the location of these intersections l 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 Catawba #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 is the preferred method of establishing tube support.
If only the apex of a Catawba #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 tybe 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 AVB insertion depths for Catawba #1 are shown in Figures 6-2 thru 6-5. These AVB maps list the results of the
' projection' calculations from the smallest row tube for which suitable data exist to make a projection.
. 2.5 Flow Peaking Factors Tests were performed modeling Catawba #1 tube and AV8 geometries to determine the flow peaking factors for various AVB configurations relative to the North Anna R9C51 peaking factor. The test results were used to define an upper bound of the ratio relative to the R9L configuration.
2.6 lube Vibration Evaluation The calculation of relative stability ratios for Catawba #1 makes use of i detailed tube bundle flow field information computed by the ATH0S steam generator thermal / hydraulic analysis code. Code output includes three-dimensional distributions of secondary side velocity, density, and void fraction, along with primary fluid and tube wall temperatures. Distributions of these parameters have been generated for every tube of interest in the
, Catawba #1 tube bundles based on recent full power operating conditions. This information was factored into the tube vibration analysis leading to the relative stability ratios.
0371M:49/020389-18
Relative stability ratios of Catawba #1 (Row 8 through Row 12) tubing versus R9C51 of North Anna 1 are plotted in Figure 9-5. These relative stability ratios include relative flow peaking factors. The stress ratios for Catawba #1 are given in Figure 9 6 for tubes in the dented condition. These also include the relgtive flow peaking effect, and are calculated based on clamped tube conditions with denting (with deformation) at the top tube support plate.
The analysis indicates that SG-C RllClll, without modification, exceeds the stress ratio criteria. Based upon a thorough review of the operational, maintenance, and modification installation factors associated with the modification options, Duke Power decided to install cable dampers into SG-C R11Clll, SG-B R12C109 and SG-B R12C110. The cable dampers were installed in December, 1988. Although the two SG-B tubes are acceptable without cable tube dampers installed, installation of the cable dampers reduces the potential for future modifications, should plant operating parameters change.
An analysis was performed to determine the effects of installation of a four cable damper into the SG-C R11C111 tube. The cable damper extends from the hot leg tubesheet elevation, through the U-bend, and below the top of the second
, uppermost tube support plate on the cold leg. Tests performed of the cable damper indicate that a minimum additional damping of [ Ja,c is provided by the cable in a Row 11 tube. This has the effect of reducing the relative stability ratio of the SG-C RllClll tube from [ ]a,c and the relative stress ratio from [ Ja,c well below the 1.0 stress ratio criterion.
An evaluation has also been performed on the five cable tube damper installed in both row 12 tubes. This cable tube damper is similar in many ways to the four cable tube damper installed in the R110111 tube. The major exception l being that the damper is assembled with five cable tube damper similar to the tests performed on the four cable tube damper. (Appendix A of this WCAP contain a discussion of the results of these test.) Results of these tests have indicated that additional damping will be experienced by the tube with the installation of the five cable tube damper. Note that any additional amount of damping provided by the cables in the Row 12 tubes will reduce the relative stability ratio (and stress ratio) to an even lower value. The amount of 0371M:49/020389-19
i 1.
additional daniping required to produce an acceptable R12C109 or R12C110 tube is ;
0.00 because the tubes are already acceptable without any additional modification.
After the installation of cable dampers into SG-C R11C111, SG-B R12C109 and SG-B R12C110, the stress ratios of all tubes in all four steam generators in Rows 8 through 12 are less than 1.0, even when the tubes are assumed to be unsupported. A summary listing of the unsupported critical tubes and pertinent vibration parameters (prior to installation of cable dampers in the aforementioned tubes) is given in Table 9-2.
The tube remaining in service with the largest stress ratio is the R10C93 tube in Steam Generator A. This tube has a stress ratio of 0.41 without denting, and 0.49 with denting; it does not exceed the 1.0 stress ratic criterion. The calculated cumulative usage for this tube, assuming denting, and future operation at current cycle parameters with 100% availability is 0.04. Assuming continued operation at current operating conditions, there is sufficient margin
. in these tubes to allow continued operation for the remainder of the plant license period.
2.7 Overall Conclusion The Catawba Unit I tube fatigue evalutt!on identified one tube, SG-C R1]C111, requiring corrective action under dented or clamped tube / TSP interface conditions. Cable dampers were installed into Catawba #1 SG-C R11C111 SG-B R12C109 and SG-B R12C110 in December, 1988. Assuming future operation with curre.nt cycle parameters for the balarice of the plant operating license, no additional plugging or cable damper installation, further modification, or other measure is judged to be necessary to precludo, at Catawba #1, a high-cycle fatigue tube rupture similar to the North Anna Unit #1 event.
O 0371M:49/020389-20
3.0 BACKGROUND
a On July 15, 1987, a steam generator tube rupture occurred at the North Anna
. Unit 1. The ruptured tube was determined to be Row 9 Column 51 in steam generator "C". The location of the opening was found to be at the top tube support plate on the cold leg side of the tube and was circumferential in orientation with a 360 degree extent.
3.1 North Anna Unit 1 Tube Rupture Event The cause of the tube rupture has been determined to be high cycle fatigue.
The source of the loads associated with the fatigue mechanism has been determined to be a combination of a mean stress level in the tube and a superimposed alternating stress. The mean stress has been determined to have been increased to a maximum level as the result of denting of the tube at the top tube support plate and the alternating stress has been determined to be due to out-of-plane deflection of the tube U-bend above the top tube support caused by flow induced vibration. These loads are consistent with a lower bound fatigue curve for the tube material in an AVT water chemistry environment. The vibration mechanism has been determired to be fluid clastic, based on the magnitude of the alternating stress.
A significant contributor to the occurrence of excessive vibration is the reduction in damping at the tube-to-tube support plate interface caused by the denting. Also, the absence of antivibration bar (AVB) support has been concluded to be required for vibration to occur. The presence of an AVB support restricts tube motion and thus precludes the deflection. amplitude required for fatigue. Inspection data show that an AVB is not present for the Row 9 Column 51 tube but that the actual AVB installation depth exceeded the minimum requirements in all cases with data for AVBs at many other Row 9 tubes. Also contributing significantly to the level of vibration, and thus loading, is the local flow field associated with the detailed geometry of the 0371M:49/020389-21
l l
1 steam generator, i.e., AYB insertion depths. In addition, the' fatigue properties of the tube reflect the lower range of properties expected for an l
AVT environment. In summary, the prerequisite conditions derived from the l evaluations were concluded to be: I 4
0
~ f Fatioue Requirements frereauisite Conditions l Alternating stress Tube vibration j
- Dented support
- Flow excitation
- Absence of AVB t
Mean stress Denting in addition to applied stress l Material fatigue properties AVT environment .
- Lower range of properties 3.2 Tube Examination Results Fatigue was found to have initiated on the cold leg outside surface of tube R9C51 immediately above the top tube support plate. No indication of significant accompanying intergranular corrosion was observed on the fracture face or on the immediately adjacent OD surfaces. Multiple fatigue itiitiation 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' vith the orientation system used, or approximt.tely 90* from the geometric center of the irfitiation zone at Section D-D. High cycle fatigue striation spacings approached 1 micro-inch near the origin sites, Figurs 3-2. The early crack front is believed to have broken through-wall from approximately 100* to 140*. From this point on, crack growth is believed (as determined by strintion spaciing, <
striation direction, and later observaticms of parabolic dimples folloried by equiaxed dimples) to have accelerated and to have changed direction with the I resulting crack front running perpendicular to the circumferential direction.
0371M:49/020389-22
_ _ _ _ _ _ _ _ _ _ ________________________________________________________________________________d
f 3.3 Mechanism Assessment To address a fatigue mechanism and to identify the cause of the loading, any loading condition that would cause cyclic stress or steady mean stress hed to be considered. The analysis of Normal, Upset and Test canditicns indicated a
~
relatively low total number of cycles involved end a corresponding low fatigue usage, even when accounting for the dented tubc 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, thu:
providing a contribution to mean stress. Combining these effects with denting deflec^ ton on the tube demonstrated a high mean stress at the failure location. Vibration analysis for the tube developed the characteristics of first inode, cantilever response of the dented tube to flow induced vibration for the uncracked tube and for the tube with an increasing crack angle, beginning at 90* to the plane of the tube and progressing around en both sides to complete separation of the tute.
Crack crophgation analysis matched cyclic deformation with the strest intensities and striation spacings indicated by the fracture inspection and
, analysis. L?akage data and crack opening analysis provided the relationship between leak rate and circumferential crack length. Leakage yersus 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 vlbration mechanisms, turbulence and fluidelastic instability, that fividelastic instability was the most probable cause.. Due to the range of expected initiation stress amplitudes (4 to 10 ksi), the fluidelastic instability *. Ad be limited in displacement to a range of approximately [ Ja,c. This is less than the ;
distance between tubes at the apex, [ ]a,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 of adjacent tubes.
0371M:49/020389-23
___-_ ____. -_ _ - _ ____ - _ _ _ --__- _ _. ------- _0
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Given the likelihood of limited displacement, fluidelastic instability, a means of establishing the change in displacement, and corre:ponding 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 rasulting from a reduction in stability ratio was converted to a fatigue usage benefit through the use of the fatigu-a curve developed. Mean stress effects were included due to the presence of denting and applied loadings. The results indicated that a 10% reduction in stability ratio is needed (considering the range of possible initiation stress amplitudes) to reduce the fatigue usage per year to less than 0.02 for a tube similar to Row 9 Column 51 at North Anna Unit 1.
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( Indicates origins Figure 3-1 Approximate Wapping of FrMture Surface of Tube R9C51.
S/G 'C' Cold Leg, North Anna Unh 1 0281H:49/092888-21
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, Figure 3-2 St hematic Representation of Futures Obscreed During TEM Fractogr3 hic Examination of Fracture Surfr.ce of
, Tube R9051, S/ii "C" Gold Leg, North Anna Unit 1 l
0281M:49/092888-22 l
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Ot81M:49/092088-23 I
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f 4.0 CRITERIA FOR FATIGUE ASSESSMENT L
The evaluation metaod 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 ratio method, and (2) the stress amplitude (or displacement) associated with a specific value of stability ratio can only be estimated by ,
the analysis of North Anna Unit 1. For these reasons, the North Anna Unit I tubitig 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 to expected to have sufficiently low vibration stress to
{
preclude futura fatigue rupture events.
- 1. To accomplish the necessary relative assessment of Catauba #1 tubing to Row 9 Column 51 of North Anna Unit 1, several criteria are utilized. First, stability ratios are calculated for Catawba #1 tubes based on flow fields predicted by 3-D thermal hydraulic models and raticed to the stability ratio for Row 9 Colunn 51 st North Anna IJnit 1 based on a flow field obtained with a
. 3-D thermal hydraulic model with the sume degree of refinement. These ratios of stability ratio (called relative stability ratios) for each potentially ;
unsupported D-bond in the Catawba #1 steam generators should be equivalent to <_
, 0.9 of R9051, North Annb i (meeting the 10% reduction in stability ratio criteria). This provides the first level of streening of susceptible tubes i incorporating all tube geometry and flow field differences in 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 thece differences, flow peaking factors can be incorporated in the relative stability ratios and the relative stress ratios.
i C371M:49/0R389-28
_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ ________J
i The next step is to obtain stress ratios, the ratw of stress in the Catawba #1 tube of interest to the stress in Row 9 Column 51, North Anna Unit 13 and after incorporating the requiremer.t that the relative stability ratic to Row 9 Column 51-(519C51) for the tube of interest is equivalent to 5 0.9, require the stress ratio to be 51.0. The stress ratio incceporates the tube geometry differences-
~~
with R9051 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 (ficw peaking factor is defined in Section 4.2). This should provide that all tubes meeting this criteria have stress amplitudes equivalent to s 4.0 ksi.
Finally, the cumulative fatigue usage for plant operation to date and for continued operathn with the sa:ne operating parameters is evaluated, A fatigue usage 1 1.0 may not bu satist'ied by meeting the stress ratio criteria using the reference operating cycle evaluation since the reference cycle does not necessarily represent the exact duty cycl: to date. Therefore, the time Mstory of operation is evaluated on a normalized basis and used together with the s' tress ratio to obtain a stress amplitude history. This permits the
[ calculation of current and future fatigue usage for comparison to 1.0.
4.1 Stability Ratio Reduction Criteria
.For fluidelastic evaluation, stability ratios are determined for specific
[. configurations of a tube. These stability ratios represent a measure of the f' potsatial for flow-induced tube vibration during service. Yalues greater than unity -(1.0) indicate Inst &bility (see Section E,1).
Motions developed by a tube in the fluidelastically unstable mode are quite large in comparison to the other known mech:inisms. The maximum modal l displacement (at the apex of the tube) is linearly related to the bending stress in the tube just above the cold leg top tuba support plate. This relationship applies to any vibration in that mode. Thus, it is possible for an unstable, fixed boundary condition tube to deflect an amount in the 0-bend which will produce fatigue inducing stresses.
l 4
i 0371M:40/02038b29
< s
)
I The major featureis of the fluidelastic mechanism are illustrated in Figure 4-1. This figure shows the displacement response (LOG D) of a tube as a function of stability ratio (LOG SR). A straight-line plot displayed on log-log coordinates implies a relation of the form y - A(x)", where A is a constant, x is the independent variable, n is the exponent (or power to which x is raised), and y is the dependent variable.- Taking logs of both sides of this equation leads to the slope-intercapt form of a straight-line equation in log fctm, log y - c + n log x, where c - log A and represents the intercept and n is the slope. In our case the independent variable x is the stability ratio SR, and the dependent variable y is tube (fluidelastic instability induced) displacement response D, and the slope n is renamed s.
f From experimental results, it is known that the turbulence response curve (on l log-log coordinates) has a slope of approximately [ Ja,b,c. Test results l also show that the slope for the fluidelastic response depends somewhat on the l 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:
l a,c I [ ]
where D3 and SR3 are the known values at the point corresponding to point 1 l of Figure 4-1 and D2 and SR2 are values corresponding to any point lower on this curve. Therefore, this equation can be used to determine the reduction in displacement response for any given reduction in stability ratio.
This equation shcws that there is benefit derived from even a very small l percentaac thange h tne stability ratio. It is this reduction in displacement for a quite small reduction in stability ratio that formed the basis for demonstrating that a 10% reduction in stability ratio wnuld be sufficient to
, prevent Row 9 Column 51 from rupturing by fattgue.
'0371M:49/02C309-30
__ _____ _ _ _ . _ _ _ _ _ _ _ _ _ . . - - . - - - - - - - - - - - - --]
The fatigue curve developed for the North Anna Unit I tube at R9C51 is from
[
i
. .q.
1 Ja,c. Thus,
,, . - a,c )
L a whore,a[istheequivalentstressamplitud'etoo a that accounts for a'maxitvam stress of py , the y1sid strength. The -3 sigma curve with mean stress effects ic shown in Figure 4-2 and is compared to the ASME Code Design Fatigue Curve for Incone) 6G0 with the maximum effect of mean stress.
. The curve utilized in this evaluation is clearly well below the code curve reflecting the effect of an AVT environment on fatigue and i Ja,c. for accounting for mean stress that applies to materials in a corrosive environment.
Two-other mean stress mod'els were investigated for the appropriateness of their use in providing a reasonable agreement with the expected range of initiating stress amplitudes. These were the [ Ja,c shown ir. Figure 4-3. With a [
]a,c, the [
.1 Ja,c, .
.- I l
I
)
0371M:49/020389-31 I o
4
y:
4
- f. j 1
y ^
The assessment of the benefit of a reduction in stability ratio begins nith the J
~
relationship between stability ratio and deflection. For a' specific tube-l gagietry, the displacement change is directly proportional to change in stress- f
, ~ so .that stress has the same relationship with stability rctio, .
1 f* ~ a,c a l m q y~ ,.
~
1 The slope in this equation can range frco [' ]*'C on a log scale I depending 9n the amplitude of displacement. Knowing the stress resulting from a change in stability ratio from SR3 .to SR 2
, the cycles to failure at the 'l 4 tress amplitude were obtained from the fatigue curve. A fatigue usage per )
{,
year was then determined assuming continuous cycling at the natural frequency
' f of the tube. The initial stress was datermined to be in the range of 4.0 to- i 10.0 ksi by 'the fractography analysis. j It was further developed that the maximum initiating stress amplitude was not Emore than 9.5 ksi. This was based on [
l' t
i ge , la,c. The. corresponding stress level is 5.6 ksi. 4 The maximum stress, 9,3 ksi, would be rsduced to [3.96 ksi]a,c with a 10%
reduction in ' stability ratio and would have a future fatigue usage of
[0.0'!09]a,c per year at 75% availability, Figure 4-4. The minimum stress, 5.6 ksi, would be reduce-d to [ ]a,c ksi with a 5% reduction in stability
- y. ratio and would have future fatigue usage of [ ]a,c per year. Figure
.4-5. in addition, if a tube' were already cracked, the crack could be as large
,T as [ ]a,c inch sin length and thru-wall and would not propagate if the stress abplitudes are reduced to 14.0 ksi.
l;.;
' 037tM:49/020389 32 mg , __ _ _ _ _ _ _ _
(
Subsequent to the return to poser evaluation for North Anna Unit 1, the time history. of operation wu evaluated on a normalized basis to the last cycle.
[ [- ;
Ja,c, cumulativ6 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 it that the probable stress associated with this fatigue cur've during i the last L do of operation was approximately [ ]a,c for R9C51, North I 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 omdtted, since denting is assumed, for purposes of this analysis, to have occurred during that first cycle. . Based on this evaluation, the tube fatigue probably occurred over most of the operating history of North Anna Unit 1.
, A similar calculation can be performed for the time history of operation assuming that [
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 fatigte usage of
[ ]a,c, Other combinations of alternating stress and mean stress were evaluated with
-3 sigma and -2 sigma fatigue curves te 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 ligan stress effects is conservattle. Any higher fatigue curve whether through tiesn stress, mean stress model, or probability, results in greater benefit for the same reduction in stability ratio. Further, fcr 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.
0371M:40/020389-33 j
4.2 Local Flow Peaking Considerations 1
Local flow peaking is a factor on stability' ratio that incorporates the effects
'~
of local flow velocity, density and void fraction due to non-uniform AVB insertien depths. The flow peaking factor is applied directly to the stability O ratio obtained from thermal-hydraulic analycis 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 Catawba #1 tubing is relativa to R9C51, North Anna Unit 1, the flow peaking factors are also applied as relative ratios, i.e., a ratio of Catawba #1 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 minirtum valua of ( Ja,b,c is appropriate for R9C51 of I North Anna 1. The peaking factor for a tube in Catawba #1 tubing is therefore divided by [ ]a,b,c and the resulting relative flow peaking is multiplied
. times the relative stability ratio based on ATHOS.results. If the peaking i
f actor is 1.0, the relative flow peaking is [ ]a,b,c, As a further demonstration of the conservatism of ( Ja,b,c as the minimum flow peaking factor for R9C51, the stress amplitude of 7.0 kai obtained from iterating on cumulative fatigue usage (and selected as the nominal value from fractography analysis) was used to back calculate the apparent stability ratio
! and then the apparent flow peaking factor. Allowing for a range of slopes of the instability curve from 10 to 30, the stability ratio is in the range of 1.1 to 1.4 and the flow peaking factor is in the range of 1.8 to 2.2. This range l 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.
i l
l 0371M:49/020389-34
___-_________--_____-_a
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. The nominal damping reflects' the nominal reduction in damping that occurs with denting at the tube support pl ate. 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 I to a level that would not have ruptured, 4.0 ksi. To apply this same criteria to I another tube in the same or another steam generatcr, the differences in [
l Ja,c, r E'h
[
l t
l 88 1
l 0371M:49/020389-15 I
-, y, .
,s c,qj .a by(; ,
f- ,
1
..r-s t
'h
.s.- ,,
a,c 1 , ,
I
[.-
L, i where the stability retto (SR) includes the flow peaking effect. j-
<~ i t
By establishing their equivalent'effect on the stres's amplitude that produced the tube rupture at North Anna 1, several other effects may be acco:mted for.,
, These include a lower mean stress (such as for non-dented tubes), different' !
frequency tubes from the [ ]a,c e hertz frequency of R9C51, ilorth Anna 1, o and shorter design basis service.
l a
4 l
l
)
I h371M:49/020389 , ,
fi
,j In the case of lower mean stress, the stress amplitude that would have caused the failure of R9C51, North Anna 1, would have been higher. [
. j i
~ l
-]a,c, )
l A lower or higher frequency tube would not reach a usage of 1.0 in the same )
length of time as the R9051 tube due to the different frequency of cycling.
The usage accumulated ic proportional to the frequency and, therefore, the 3 allowable number of cycles to reach a usage of 1.0 is inversely proportional to d 4
frequency. The equivalent number of cycles to give the usage of 1.0 for a t j different frequency tube [ j Ja,C;
- r for a different time basis for tatigue usage evaluation,. [
I l
l- Ja,c,e, 1
Knowing the magnitude of the stress ratio allows 1) the determination of tubes that do not meet a value of 51, and 2) the calculation of maximem stress in the acceptable tubes,
- 8,C L w I
. 1 Having this maximum stress permits.the evaluation of the maximum fatigue usage l
. for Catawba #1 based on. the time history expressed by normalized stability ratiot for the duty cycle (see Section 7.4).
0371M:49/V20389-37
il Table 4-1 1 4
Fatigue Usage per Year Resulting From Stability Ratio Reduction
. l
.SR, % STRESS FATIGUE NEAN STRESS USAGE j C
REDUCTION BASIS (l)
CURVE (2) MODEL PER YEAR
- a,c
- 5. 9 yrs to fail [ ]a,c j
- 5. 9 yrs to k fail [ Ja,c i S. 9 yrs'to fail [ ]a,c
{
- 10. '
max,stresg) amplitudet
[ ]a,c !
1 .
)* 10.
max.streg) amplitude
[ Ja c ,
10.
max. stres )
amplitude
[ Ja,c 10.
max.stres) amplitude
[ Ja,c
- 10. max, stress based on .
dutycycle(8) j_ w
[ ja,c
-(1) This gives the basis for selectiori of the initiating stress amplitude and its value in ksi. '
(2) S, is the maximum stress applied with Sm"Smean + Sa-(3) [ Ja,c, 1
(4) Cycles to failure implied by this combination of stress and fatigue properties is notably less than implied oy the operating history.
Consequently this combination is a conservative, bounding estimate.
g.* ; I l (5) Cycles to failure implied by the operating history requires [
]a,c fatigue curve at the maximurn stress of [ ]a,c, 1
0371M:49/020389-3C
_ -_ _ _ _ __ _-____-___-_a
\
I
+
~
.. ..%b e
- 2. >
Figure 4-1 Vibration Displacement vs. Stability Ratio 0281M:49/092888-36
- 7. ,
to a; <; t 3:
=.;,
(
l f
t i f, i 1 i
.I <'
.~
$c i.
f6 l .
y t')
i w
I Figure 4-2 Fatigue Strength of Inconal 600 in AVT Water at 600*F 1
0281M:49/092888-37
, .1
[,
r .
'.1 l
- ~ .1
- %c
-l E
I -l L. -
.i
, Figure 4-3 Fatigue curve for Inconal 600 in AVT Water comparison of Mean Stress Correction Models 0281M:49/092888-38
i
.4 n
- r. ,,4. c L
.e
~
i l
': 4 Figure 4-4 Nodified Fatigue with los Reduction in Stability
. Ratio for Maximum Stress Condition 0281M:49/092888-39
l L
l.
s l1 1
l W,jC I
l a
~
Figure 4-5 Modified Fatigue with 5% Reduction in Stability
, Ratio for Minianas Stress Condition i
l 0281M:49/092888-40
Ii 5.0 SUPPORTING TEST DATA i
This section provides a mathematical description of the fluidelastic mechanism, I 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 l conditions and corresponding parameters directly related to the event and associated preventative measures. The basis for estaL11shing the appropriate values and implications associated with these parameters are provided. Where appropriate, test results are presented.
5.1 Stability Ratio Parameters Fluidelastic 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. Fluidelastic 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 (eigenvector) for the linearly supported tube being considered. l The methodology is comprised of the evaluation of the following equations:
Fluidelastic stability ratio - SR = Uen/U cfor mode n, where Uc (critical velocity) and Uen (effective velocity) are determined by:
2 Uc"#f Dn Umg6n ) / (#o D)p/2 (f) and; N
- 2 g(p/p)U jo 3 4jn Z j
U "
1 en [2] I N
j1 3 d" d ;
0371M:49/020389-44
k i
( ,
M .wherei-
'" ! -D -- ' tube outside diameter, inches' K, z U en -
effective. velocity for mode n, inches /sec
.s N = number of nodal points of the. finite element model sj, Uj, pj = . ' mass per unit length, crossflow velocity and fluid l-
'densityLat node j, respectively li
.po, mo -
reference density and a reference mass per unit >
. length, respectively (any representative values) !
6 n
logarithmicdecrement-(damping)
Ijn - normalized displacement at node j in.the nth mode of vibration zj -
average of distances between code j to j-1, and j to J+1 an experimentally correlated stability constant Substitution of Equations [/] and [2) into the expression which defines stability ratio, and cancellation of like terms, leads to an expression in !
fundamental terms-(without the arbitrary reference mass'and density o parameters)'. From this resulting expression, it is seen that the stability 1 ratio is directly related to the flow field in terms of the secondary fluid velocity times square-root-density distribution (over the tube mode shape),'and inversely related to the square root of the mass distribution, square root of modal. damping, tube modal frequency, and' the stability constant (beta).
The uncertainty in each of these parameters is addressed in a conceptual manner in Figure 5-1. The remainder of this section (Section 5.0) provides a discussion, and, where appropriate, the experimental bases to quantitatively establish the uncertainty associated with each of these parameters. In 0371M:49/020389-45 l
addition, Section 5.3 provides the experiment 61 basis to demonstrate that tubes with [
]a,c. This /
implies that those tubes [ ]a,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.
Freauency It has been demonstrated by investigators that analytically determined frequencies are quite close to their physical counterparts obtained from measurements on real structures. Thus, the uncertainty in frequencies has been shown to be quite small. This is particularly appropriate in the case of I dented (fixed boundary condition) tubes. Therefore, uncertainty levels
. introduced by the frequency parameter are expected to be insignificant (see also " Average Flow Field" subsection below).
Instability Constant (Beta) 1 The beta (stability constant) values used for stability ratio and critical velocity evaluations (see above equations) are based on an extensive data base comprised of both Westinghouse and other experimental results. In addition, previous field experiences are considered. Values have been measured for full length U-bend tubes in prototypical steam / water environments. In addition, measurements in U-bend air models have been made with both no AVB and variable AVB supports (Figure 5-3).
To help establish the uncertainties associated with ATH0S flow velocity and l
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 ATH0S predicted flow conditions l and the MB-3 maasured critical velocity was determined. These analyses supported a beta value of [ ]a,b,c, 0371M:49/020389-46 l
\ - - _ - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - .
L
_ 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 i beta parameter in all stability' ratio evaluations addressed in this Report (see also " Average Flow Field" subsection below).
I Mass Distribution The mass distribution parameter is based on known information on the tube knd 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, Tube Damoina
, Test data are available to define tube damping for clamped (fixed) tube supports, appropriate to dented tube conditions, in steanVwater 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 ster.m/ water. This data was -
obtained for cross flow over straight tubes. Uncertainties are net defined for the data from these tests. Detailed tube damping data used in support of the stability ratio evaluations addressed in this report are provided in Section 5.2, below.
Flow Field - Velocity Times Souare-Root-Density Distribution Average and U-bend-local flow field uncertainties are addressed independently in the following.
e 0371M:49/020389-47
Averaae Flow Field Uncertainties in the average flow field parameters, obtained from ATH0S ;
analyses, coupled with stability constant and frequency, are essentially the same for units with dented or non-dented top support plates. If the errors associated with these uncertainties were large, similar instabilities would be expected ie 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 generatcrs without denting. Thus, an uncertainty estimate of about [ ]a,c 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 Catawba #1 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 AWs, leacts 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 it 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 Secticn 5.2, below.
Overall Uncertain _tjes 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 vibra. tion amplitude for the failed North Anna tube. Ratios of stresses and stabilfity ratios relative to the North Anna tube, R9C51, are utilized in this report to minimize uncertainties in the evaluations associated with instability constants, local flow field effects and tube damping.
0371M:49/020389-48
5.2 Tube Damping Data The damping ratio depends on several aspects of the physical system. Two l-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 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 incraases significantly and is dependent on the void fraction or quality. It can be shown that the damping contribution i from viscous effects are very .small.
+
Damping ratios for tubes ir, 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 develnped for both pinned and clamped tube support conditions.
Two sources or data are particularly noteworthy and are used here. The first is a large body of rcc.nt, as yet unpublished data from high pressure steam-water tesi.s r,onducted by Mitsubishi Heavy Industries (MHI). These data were gathered under pinned tube support conditions. The second is comprised of l the results from tests sponsored by the Electric Power Research Institute (EPRI) and reported in References (5-2) and (5-3).
l The damping rctio 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 b71M:49/c20389-49
~
(Reference (5-4)). The upper curve in thu 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 carve pertains to the clamped support condition, obtained from Reference (5-3). Void f*:.ction 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 incitde the effect of denting at the top tube support plate. ,
J The Reference 5-3 data as reported show a damping value of .5% at 100% void I
fraction. The 100% void fraction condition has no two phase damping and is considered to be affected principally by mechanical or structural damping.
l Westinghouse tests of clamped tube vibration in air has shown that the l" mechanical damping is only [ ]a,c rather than the .5% reported in Reference (5-3). Therefore the lower curve in Figure 5-4 is the Reference l
(5-3) data with all damping values reduced by [ ]a,c, i
l l
i I
s l l
i 1
0371M:49/020389.50 i
5.3 Tube Vibration Amplitudes 91th Single-Sided AVB Support i
A series of wind tunnel tests were conducted'to investigate the effects of )
tube /AVB eccentricity on the vibration amplitudes caksed by fluide'hastic j vibration. i l
)
( )
I Ja,c. Prior test results obtained during the past year using this apparatus have demonstrated that the
- f. fluidelastic vibration characteristics cbserved in the tests performed with the cantilever tube apparatus are in good agreement with corresponding characteristics observed in wind tunnel and steam flow tests using U-bend tube arrays. A summary of these prior results is given in Table 5-1.
An overall view of the apparatus is shown in Figure 5-5. Figure 5-6 is a top view of the apparatus. [
4 3
1 I
JE,C, 0??!K.491020389-51 4
---. _ _ _ _ _ _ _ _ _ _ _ _ - - - . _- _ . __ ._-.___.A
~
L
' 7 g .,
T ,/_1 1
U E g, ,.
z4 7.g e o }
q
- As shown in Figure 5-7, the tube vibration amp 1.itude below a critical velocity 3 is' caused by [
- A;4 i
.3 ,
ja,c, l
Figure 5-7 shows the nanner in which the tero-to-peak vibration amplitude, expressed as a ratio normalized to [ Ja,c, varies when one gap remains at[ .]a,c. For increasing j l
t <
. velocities, up to that corresponding to a stability ratio of [
Ja,c. Figure.5-8 shows typical vibration amplitude and tube /AVB impact force signals corresponding to those obtained from the tests which provided the results shown in Figure 5-7. As expected, impacting is only observed in the [ Ja,c, It is concluded from the above test results that, [
ja,c, 5.4 Tests to' Determine the Effects on Fluidelastic Instability of Columnwise Variations in MB Insertion Depths This section su:amarizes a series of wind' tunnel tests that were conducted to 1 investigate the effects of variations in AVB' configurations on the initiation
.of. fluidelastic' vibration. Each configuration is defined as a specific set of insertion depths for the individual AVBs in the vicinity of an unsupported U-bend tube, The tuts were toiducted in the wind tunnel using a modified version of the cantilever tube apparatus described in St4ction 5.3. Figure 5-9 shows the s
4.'
03714:49/020389-52 L
d_ ^
- -_m _ . _ _ _ _ _ _ _ _ _ _ . _ _ _ _ . - - _ _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . . _ _ _ _ .
- t
, ; conceptual ' design of the apparatus.I The straight cantilever tube, j
[
, . ~I l
L 'j 1
ja,C, c.
.[
I, ,
]a,c. Figure 5-11 shows the AVDs, when the side panel of the test sectior; is removed. Also shown is the
. top flow screen which is [.
i I
J.
{
l
]a,c. The AVB l configurations tested are shown in Fig. 5-12. Configuration la corresponds to tube.R9651, the tailed tube at North Anna. Configuration 2a corresponds to one of thc casas in which the AVBs are incerted to a uniform depth and no local velocity peaking effects are expected.
- 1) The AVBs shown in Figure 5-9 correspond to original AVBs. AVBs 3
corresponding to thoise used in field modified units, such as Catawba
)
- 1, were also made using the same precedure as for the original AVBs.
. ,. {
4 l
0371M:49/02Q89-53 L i
W. .. ,
j t ,
I'M As'shown in Figure 5-9, [.
y I
.ja,c, All the tubes except.the instrumented tubed (correspor. ding to Row 10) are
[ ]a c. As dit. cussed in Section 5.3, prior testing indicates that this situation provides a valid acdel.. The instrumented.
tube [ Ja,c as shown.in Figure 5.10.
Its [ Ja,c direction vibrational motion it measured using a non-contacting' transducer.
[
Ja,c. The instrumented tube corresponds to a Row 10 tube as shown in L, Figure 5-9. However, depending on the particular AVB configuration, it can ;
reasonably represent a tube in Rows 8 through 11. The AVG profile in the ;l
. . straight tube model is the average of Rows 8 and 11. The difference in profile I-is quite small for these bounding rows.
[ la,c using a hot-film anemometer located as shown in Figure 5-9.
Figure 5-12 shows the rms vibration amplitude, as dttermined from P50-(power spectral density) measurements made using an FFT spectrum analyzer, versus flon velocity for Configuration it.. Configuration la corresponds.to the final, <
evaluated positions of AVBs near tube R9C51 in North Anna (See Figures 8-6 and 8-7). Data for three r,epeat tests are shown and the critical velocity is
. identifieo. The typical rapid increase in vibration amplitude when the 3 i
,,- ' critical velocity for fluidelastic vibration is exceeded is evident.
o,_sm_
j i
gjoi y'
h .q k .The main conclusions from the tests. ara:
F
.1. Tube ' vibration beloy the critical ve'locity is relatively small, typical of' turbulence induced vibration, and increases rapidly when the critical velocity for tha-initiation of .fluidelastic vibration is exceeded.
.f '
]
~
h '2. Configuration Ib (which was initially thought to represent AVB positions
'.,, near R9C51 in North Anna until re-evaluation indicated f. configuration'la)'
has the lowest critical velocity of all the configurations testeds
- 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 Canwba #1 evaluation are cunnarized in Table 5-2.- The test data are prasented as velocity peaking
!! -ratios, the ratio of' critical velocity for North Anna tube R9051 configuration la, to that fur each Catawba #1-AVB configuration evaluated.
I!" '5.5 References.
i - a,, b , c '
54 0
5-3 5-4 I
J
)
0371M:49/020389-%5
~ _- . . _ . . _ _ _ _ _ - _ _ _ . _ _
l,a N,,
Table 5-1 Wind Tunnel Tests on Cantilever Tube Model
.ee p
'0BJECTIVE: Investigate the effects of tube /AVB fitup on flow-induced tube vibration.
1 APPARATUS: ' Array of cantilevered tubes with end supports [
i 3a,c, MEASURCHENTS:' Tube vibration' amplitude and tube /AVB impact forces or preload.
forces.
RESULTS:
... a,b,c
~
i sse
.; i.
2.
m j i
3.
A.
I 5.
l e
- i a
d iM:49/020389-56 !
4 1, t
- \
Table 5-2 o 9' Fluidelastic Instability Velocity Peaking Ratios
- v ,
, for Columnwise Variation.in AVB Insertion Depths ,
"*-~ '
(Catawba Unit 1) !
- [ Type of AYD' Insertion Peaking Ratio jy Configuration Uja/Un a,b,t la I 1 lb Iw L . 2a
(:<
4a 4b .
f., 4e j
~
4x 1 , 42 l l
) , Sa i i
Sb Sc f
Sg j 51 6a 6b 6C
. 6d 0 l f
ll 8a Sb 1; Sc
- l. - I NOTE: nU is instability velocity at 1alet for type n of AVB insertion configurat',on, 0371N:49/0?E389-57
- = - - - - - _ - _ - - _ _ _ _ _ - _
1
a,c l
4
'1
-(
Figure 5-1 Fluidelastic Instability Uncertainty Assessment 0224M:49/090888 56
7 ,;; p w , 1 0)f f, c ,
w c J, c U-Bgnd Test L.ata.-
- 1) Mb-3 fests S values of [- la,b,c
- 2) MB-I Tests ;
- [ p of. [ J ,b,c t
- 3) Air tiodel Tests
$ of [ ]a,b,c.without AVDs '
J
' Tendency for S to increase in range of [ ]a,b,e with inactive AVBs (gaps at AVBs)
Tendency for B to decrease toward a lower bound of
[ ]a,b,c with active AVBs Verification, of Instability Conditions
- 1) Flow conditions at critical velocity from MB-3
- 2) Measured damping for the specific tube
, 3) Calculated velocities from ATH0S 3D. analysis
- 4) $ determined from calculated critical values
, ' Good agreement with reported # values
- 5) ATH0S velocity data with B of [ ']a,b,c and known damping should not significantly underestimate instability for regions of uniform U-bend flow 1
l
', Figitre 5-2 Instability Constant - $
l I
C371M:49/070389-59
. . . . . . ~ . . .
W# .n : .
o s .- ..
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8
- f a
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Figure 5-3 Instability constants, A, obtained for Curnd Tubes frem .
3 Wind Tunnel Tests on the 0.114 Scale U-Bend leodel l
l I
i 0281M:49/092888-57 1'
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1,. , _ _ _ _ _ _ _
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,W' s Figure 5-4 Damping vs. Slip Void l'rsction 0281M:49/092888-58 t
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- . Lj.' Figure 5 5 Overall View of Cantilever Tube. Wind Tunnel Model #
i i
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i 25368-1
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i Figure 5-6 Top View of the Cantilever Tube Wind Tunnel Model e ,
25368 2 l
3
a,b,e
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, Figure 5-7 Fluidelastic Vibretton Amplitudt with Non-Uniform Gaps 0154M:49/032888-65
, a.s .
._L.
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r8;l; 5f 0)'
, a,b.c
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Figure 5-8 Typical Vibration Amplitude and Tube /AVB locact Force
. , signals for Fluidelastic Vibration with unequal Tube /AVB Gaps 0154M:49/032888-66
1
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a ,b . c l
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., Figure 5-9 Conceptual Design of the Apparatus for Determining the Effects of Fluidelastic Instability of Columnwire ;
Variations in AVB Insertion Depths 0154H:49/032888 67 !
p.-
a,b,c t
l*
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Figure 510 Overall View of Wind Tunnel Test Apparatus 25368?
'- ' =-
i:. 1 l
, - - a,b c-l 8
Figure 5-11 Side View of Wind Tunnel Apparatus with Cover Plates
, ' Removed to Show Simulated AVBs for Field Modified Units and Top Flow Screen f
- 25368-5 x - _:_-__-_--___--___-_--
1 I
l 1
'{-
1 1
1 1
1 l
TYPE OF AVB TYPE OF AVB TYPEOFAVB f INSERTION
~
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INSERTON INSERTON '
)
e #
. a h,c.
s 9 he-sbA 4
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l Figure 5-12 AVB Configurations Tested for 04awba 1 l
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Figure 5-13 Typical Variation of RMS Vibration Amplitude with Flow Velocity for Configuration la in Figus 5-12 0281M:40/092868-60 l
6.0 EDDY CURRENT DATA AND AVB POSITIONS Westinghouse performed preliminary AVB mapping utilizing the AVB-to-tube
- support plate span length data provided by Duke Power from baseline (MIZ-12) eddy current inspections performed before the plant became operational. In
~
addition, approximately 602 tubes were inspected in December,1988, to provide additional data for determining AVB positions. Analysis of the eddy current data was performed by both Duke Power and Westinghouse eddy current analysts.
6.1 Catawba #1 AVB Assembly Design i
[
.la,c.e Review of the EC data for Catawba #1 shows that AVB insertion depth is fairly uniform in the regions between Columns 31 and 84 (corresponding to the " flat" contour of the tube bundle in this region) with insertion depth in most cases to Row 9 or lower, AVB insertion depth in the remaindar of the columns is more variable, tending to a higher depth of L insertion.
l'
- i*
l l
0371M:49/020369-63 1
I 6.2 Eddy Current Data for AVB Position- l The AVB insertion depths were determined on the basis of interpretation of the 4 eddy current data. To locate the AVBs, the ECT data traces were searched for /
the characteristic peaks seen in the signals which indicate the intersection of an AVB (or a tube support plate) with the tube (Figure 6-1). Since ambiguity l can occur in the interpretation of the ECT data, due to inability of ECT to differentiate on which side of a tube a " visible" AVB is located, other information was used to assist in establishing the location of the AVBs.
[
Ja,c The number of these AVB intersections, including zero (meaning no AVB present),
was evaluated for each tube to indicate the presence or absence of AVBs.
l Figures 6-2 through 6-5 show a representation of AVB insertion distance based
~
on evaluation of the EC data. All inspected tubes in these figures, with the exception of those indicated with a "0" or a "1", have two or more " visible" AVB signals. In cases where no AVBs were indicated, a "0" is shown, and i
likewise, a "1" is shown where "one AVB" is indicated. All other tubes examined have two or more AVB signals present.
The direct observation data (the number of AVB intersections 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 observations and projections, the more conservative data are used to determine the AVB positions. Since ' direct observation' gives a 'yes - no' type of answer, the projection method is used to ' interpolate' AVB insertion depths between rows of tubes.. Greater conservatism is generally interpreted as the AVB being less inserted although consideration must also be given to the resulting flow peaking factors.
4 l
0371";O/020389-64 l
l m
1 l + , ,
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. [
f* i
.. l ja,c l
In the case where the AVB characteristic signals can not be confidently j determined due to a noisy signal or pre-existing plugged tubes, location data for the AVBs is provided for [
T f 1
i 1
1 i
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i ja,c .j s
6.3 Tube Denting at Top Tube Suppert Plate i
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 l plate. In the denting evalutiion, the EC tapes were evaluated to determine the. l I
o, I
i I
G372M:49/020389-1 a--_---___-----:.-- -
1 1
l l
condition of the tube / TSP interface for potentially unsupported tubes'in l
locations which could be susceptible to flow peaking. Ant. lysis of the data for
-Catawba #1 shows the presence of ' corrosion with magnetite' in the majority of the tube / TSP crevices examined. Of the thirteen (13) tubes evaluated in December,1988, for top tube support plate corrosion, four were found to be slightly corroded or have magnetite present and three showed denting indications (both the tubes having 13P interfaces corroded with magnetite and those dented with deformation are considered to be dented per the NRC Bulletin 88-02 definition), and six showed no detectable magnetite or corrosion.
Because the tube vibration analyses were based on the conservative assumption that all tubes in the area of interest were structurally ' fixed' in the TSP holes by denting or corrosion with associated tube deformation, the results of this phase of the examination do not influence the disposition of the tubes found to be susceptible to fatigue.
6.4 AVB Map Interpretations 6.4.1 Description by Steam Generator To review the relationship of tube stability and tube row number, it is useful to examine Figure 9-7, at the end of Section 9. This figure combines the I effects of the tube position (row and column), tube bundle flow-field (velocities, densities, and damping), and tube geometry (bend radius, etc.) to develop an allowable flow peaking ratio for each tube. The allowable flow peaking ratio is the highest flow peaking factor that a particular tube may experience before exceeding a stress ratio of 1.0 as compared to the North Anna R9C51, previously established in Section 4.0. Changes in the flow field account for some of the variation in the allowable peaking ratio, and the peripheral tubes (Column 2) can accept markedly higher peaking ratios, since their allowables are determined from an approach velocity developed from a square-root-sum-of-the-squares formulation of the gap velocity on one side of the tube, and the comparatively low bulk flow velocity on the periphery.
Observing the variation in allowable flow peaking versus row number, about a 10% decrease in the allowable flow peaking factor per row can be seen. Since a flew peaking ratio of 1.0 indicates that the flow peaking factor for the tube being examined is identical to that of North Anna R9051, it can be seen from Figure 9-7 that a Catawba #1 Row 10 tube would require a peaking factor 0371M:49/020389-2
_L
i i
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equivalent or greater than Nerth Anna R9C51 to exceed the stability ratio criteria.
SG-A The AVB map is given in Figure 6-2. All Row 12 tubes are supported. Support 3 could not be established for sixteen (16) Row 11 tubes, thirty-five (35) Row 10 l tubes, fifty-seven (57) Row 9 tubes, one hundred and nine (109) Row 8 tubes, and one hundred snd ten (110) Row 7 tubes. AVB. Insertion depths vary gradually ,
from Column 2 through Column 30, slightly from Colcan 31 to Column 81, and vary to.a greater extent between Columns 82 and 113.
R10C92 is the highest loaded tube in this steam generator, having a relative stability ratio, including the effects of local flow peaking, of 0.78. The flow peaking factor selected for this tube is that of [
]a,c Eddy current' analysis of R10C92 indicated
" corrosion plus magnetite" at the top tube support plate crevices; as with all tubes, the analysis assumed it to be dented with deformation. The stress ratio.
relative to North Anna R9C51 is 0.49, which is the highest stress ratio of any of the tubes in the four Catawba #1 steam generators (after taking into account the installation of cable dampers in the SG-B and SG-C tubes discussed below).
The AVB positions for RICC92 are conservatively shown [
]a,c pjow pealing factor selections are discussed in .Section 8.8.
The relative flow peaking factor for R11C98 has been revised since the letter report (STD-7.2.2.1-8471) to [
Ja,c &nd the relative stress
'0372M:49/020389-3
_ a
I
-o j ratio (assundng denting) to 0.22. R11C98 is the next highest-loaded tube in SG-A, after R11C105 and R110106, with relative stability and stress ratios of j 0.76 and 0.35, respectively. {
1 l SG-B I
The AVB map is given in Figuro 6 3. Support could not be established for three Row 12 tut;es, eight Row 11 tubes, twelve (12) Row 10 tubes, twenty-four (24)
Row 9 tubes, eighty (80) Row 8 tubes, and one hundred and five (105) Row 7 s g tubes. AVB insertion depths vary gradually from Column 3 to Column 29 and from Column 84 tn Column 102, slightly from Column 30 to Column 83, and vary to a greater extent between Columns 103 and 113.
The highest tube stress ratios in this steam generator were found in tubes j 4
R12C109 and R12C110.
The stress ratios for these tubes was 0.89, which is below 1.0 criteria for continued operation. However, based upon a thorough review of operational maintenance, and modification factors associated with the modification options,
, Duke Power decided to install a ( Ja,c in each of tubes R32C109 and R12C110. For a discussion of the effect of the cable darnpers on the tube stability ratio and stress ratio, refer to Section 9.0 and Appendix A. The denting evaluation of R12C109 and R12C110 indicated neither corrosion or magrietite present in top tube support plate crevices.
R12C104,[
l J
ja,c q l
0372M:49/020389-4 )
I _ _ _ _ _ _ _ _ _ _______. _ _ _ __
- ls i
Row 8 data were not available for Columns 44 through 60, and AVBs were i positioned using Row 9 and Row 11 ' projections. A bounding analysis for these
- tubes is discussed in the portion of this section under "SCcD", however, it is
.. sufficient to note at this point that [
, jo,c j L .
R9C92 is the highest loaded tuba in this steam generator (after taking into account the installation cf cable dampers into the aforementioned tubes),
having a relative stability ratio, including the effects of local flow peaking, of 0.67. The flow peaking factor selected for this tube is also that of
[ Ja,c The stability and stress ratios relative to North Anna R9C51 for R9C92 are 0.67.and 0.25, respectively. The flow peaking factor selection for r19C9F is dise.ussed in Section 8.8. A denting evaluation was not performed for th'is tube.
SG-C i
l The AVB map is given in Figure 6-4. All Row 12 tubes are supported. One Row
~
11 tube, six Row 10 tubes, sixteen (16) Row 9 tubes, twerity-two (22) Row a tubes, and one hundred and three (103) Row 7 tubes are unsupported. AVB insertion depths vary gradually from Column 11 ta Column 31 and from Column 78 to Column 100, slightly from Column 32 to Column 77, and vary to a greater ,
extent from Columns 2 to 10 and Columns 101 to 113.
A[ ]a,c was installed in tube R110111. The cable extends from the hot leg tubesheet elevation, through the U-bend, and below the ,
tup of the second uppermost tube support plate on the cold leg. Analysis and testing were performed to determine the effects of the cable damper on tube vibration. Tests indicate that a minimum additional damping of [ la,c is provided by the cable damper in a Row 11 tube. This has the effect of reducing the relative stability ratio of SG-C R110111 from [ ]a,c and the relative stress ratio from [ ]a.c For a further discussion of the cable dampers, refer to Section 9.0 and Appendix A. Eddy r.urrent analysis
- l of R11C111 indicated dent signals at the hot leg tube support plate to the cold I leg tube support plate and at positions along the U-bend; a standard 0.610" )
~
i
)
i 0372M:49/0203B9-B l
I m
~
diameter probe could not be passed through, the tube, although a 0,590" diameter
. probe was able to traverse the U-bend. It was not possible to determine if corrosion product's had clamped the tube in the support plate, j
~ ;
The AVB positions for R10C92 are conservatively shown [
l l
la,c Flow peaking factor selections are discusseo in i Section 8.8.
f AVB positions near R10C4 were evaluated [
Ja,c R10C4 is the highest loaded tube in this staam generator (since a cable damper was installed R110111), having a relative !
, stability ratios including the effects of local flow peaking of 0.68, and a i relative stress ratio of 0.24. The flow peaking factor selections are j
. discitssed in Section 8.8.
l- SG-D The AVB map is given in Figure 6-5. All F.ow 12 and Row 11 tubes are supportec. Eight Row 10 tubes, forty-three (43) Row 9 tubes, ninety-eight (98)
Row 8 tubes, and all (112) Row 7 tubes are unsupported. AVB insertion depths vary gradually throughout tLe columns.
R8C27 is the highest loaded tube in this steam generator, having a relative I stability ratio, including the effects of local flow peaking, of 0.62, and a t relative stress ratio of S.21. No cable dampers were installed in SG D.
l -
i Row 9 data were not available for Columns 37 through 39, and Columns 58 through
- 61. Projection values from Row 10 and Row 11 tubes in Columns 50 through 62,
,, where support of Row 9 tubes could not be confirmed, [
l 0372M:49/020389 5
y .e ,
.ta l (, '
0 .' /
1 i
1 1
)a,c
[
b Flow: peaking factor. selections are discussed in Section 0.8. A denting evaluation was not performed for ti,is tube, f
- 6 A.2 Sunmary of Support Conditions
.f Tubes which were evaluated as having "1" AVB signal, and determined to be supported based upon projection data and adjacent tube diecks are listed in Table'61, . k")th the exception of tubes listed in Table 6-2, "I AVB Signals' p
Indicating Support for Flow Peaking Analysis",- all tubes which are drawn with sn AVB inserted to a depth even with the tube centerline are supporteda The y 1
tubes in Table 6-2, altimugh unsupported, are less limiting than neighboring !
I ' tubes in which flew peakiro is produced, and therefore are infrequently' evaluated as " salient" tubes in .%ction 9, Tabic 9-2. Table 6-3, which l includes those tubes listed in Table 6-2, provides a sumary listing of all unsupported tubes. The AVB projections shown are in nearly all instances the-projected values from the tube in the next-highest rw in a particular colutan.
If the projection from a particular next-highest row tube varies significantly l with those of adjacent and higher row tubes, possibly due to a s5gnal from a 4 deposit, all surrounding data are analyzed, and a revised projection value is determined.'
.o p, . <
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l 037tM:49/020380 7
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TABLE 6-1 j Catawba Unit #1 ,
4 One AVB Sigr,als Determihad to be Supported Catawba #1 Steam Generatar_A Row 12 None Row 111 Colu:r,ns 78, 107,.108 Row 10 Columns 13, 15, 28, 83, 91
- , Row 9 Columns 28, 36, 45-47, 52, 53, 55, SS*66, 68, 70-72, 74, 75, 78 Row 8 None Ro1 7 None ie Catawba #1 Steam Generator B Row 12- Columns 104, 105 1 Row 11 Column 113 Row 10 Column 112 Row 9 None
. Row 8 Columns 7-9, 16, 30, 34, B2-84 Row 7' - Column 26 l CatawbiLf] Steam Generator Q A
Row 12 Column 93 i Row 11 Columns 12, 93 I Row 10 Columns 103, 108 Rew 9 Column.28 l
.9cw S None !
Row 7 ' Column 26
.Cidulwbl.1LEttuLAtatrator o )
i i
Row 12 Row 11 None idone l
-l Row 10' Columns 4, 107, 106 (
Row 9- Columns 24, 46-40, 64, 72, 97, 99 l Row a Cclumns 14, 30, 39, 89 j
-Row 7 None j l ... 1 1
j
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1 0372M:49/020389-8 e.
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, "12 ,
o ., a fjls 4 TABLE 6-2 j, "
. One AVB Signal Jndicatf ng1upport' for Flow Peaking Analitis i
j w
Calavba #1' Steam Generator f l(.
Row 12 None. l Rcw 11 Columns 6, 10 103 .l o = Row,10 Columns 17, 92 )
E, Row 9. Columns 77,89 0 Row 8 Columns 2.8, 29
' i Row 7- None y
Catawba #1 Steam Generatorfl Row 12- Column 108 1 Row li None .
-Row 10 Columri 113 P.ow 9' None' Row 8' Co',umns 7li, 76, 78, 79 '
Row 7 Columns'4, 5, 13, 85, 86
[AtJLWba #LSteam Gegrator C
- Row 12 None Row 11 Nonen .
Raw 10 Column ~s' 109, 112, 113 (
Lt Row 9 f(one .i Row 8; None Row 7 Coluinns 14, 15, 27 i
datawba #1 Steam Generator.il Row 12 None- !,
Row 12 None Row 10 None Row-9 ' Columns 20, 21, 43, 44, 52-55, 57-61, 103, 104 Row 8. Column 40
~Rew 7 None i
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0372M:49/0Zb319-S y.
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-l TABLE 6-3 Catawba Unit #1 Unsunoorted Tube Summary I
Catawba #1 Steam feltrat.o_r._f Row II No unsupported tubes Row 11 Columns 2-6, 9, 10,96-100, 103-106 Row 10 Columns b12, 16, 17,92-113 Row 9 Columns 2-23,42-44, 48, 50, 69, 73, 76, 77, 83,89-113 Row 8 Columns 2-85,89-113 ,
Row 7 Columns 2-06,89-113 '
Chtawba #1 Steam Generator j .
1 Row 12 Columns 108-110 Row 11 Columns 2, 104-110 Row 10 Columns 2, 101-110,, 113 {
Sow 9 Columns 2,91-113 Row 8 Columns 2, 18-22, $2, 33, 36-41, 43-81,87-113 Row 7 Columns 2-11, 13-24,29-113 !
. Catawba #1 Steam Generator _(
Row 12 No unsupported tubes '
. Row 11 Column 111 Row.10 Columns 4, 109-113 <
1 Row 9 Columns 2-9, 106-113 L l Row 8 Columns 2-9, 92, 93, 102-113 Row 7 Columns 2-25, 27, 28, 32-67, 69-76, 78-84,88-113
,qafta_w.ka #1 Stean Generator D Row 12 No unsupported tubes Ruw 11 No unsupported tubes l Row 10 Columnr. 2, 5, 8, 109-113 Rou 9 Columns 2-6, 17-23, 43-45, 50-62, 73, 76, 103 113 Row 8 Columns 2-12, 15-24, 27, 31-37, 40-84, 88,91-113 ] '
Row 7 Columns 2-113 1 i
! l
! I l
l 037Dt. 49/020389-10
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Figure 6-I AVB Insertion Depth Confirmation 0238H 49/063088-5 w
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i 7.0 THERMAL AND HYDRAULIC ANALYSIS 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-D ATH0S computer code, Reference (7-1). The major results of the analysis are the water / steam velocity components, density, void fraction, and the primary and secondary fluid and tube wall temperatures. The distributions of the tube gap velocity and density along a given tube were obtained by reducing the ATHOS ,
I resul ts. In the following subsections, the ATH0S model and some sample results j of the analysis are described. Normalized stability ratios over the operating history of the unit were also determined and are reported in Section 7.4. )
I 7.1 Catawba #1 Steam Generator Operating Conditions Recent steam generator operating condition data for the Catawba #1 unit were provided by Duke Power. The data are representative of full power operation during May 1988, midway through the most recently completed fuel cycle, Cycle
- 3. With these data, calculations were comp 1sted using the Westinghouse SG performance computer code, GEND3, to verify the plant data and to establish a
, complete list of operating conditions required for the ATH0S analysis. The GEND3 code determines the primary side teinperatures and steam flow rate required to obtain the specified steam pressure at the given power rating.
Besides confirming these parameters, the code calculates the circulation ratio which is used to determine the total bundle flow rate and average loading on the tebes. 11 also provides an overall indication of the voids within the tube bundle since the bundle exit quality is inversely proportional to the circ ratio (Xexit = 1/ circ ratio).
The calculated circulation ratio along with the other thermst/ hydraulic conditions are listed in Table 7-1. These parameters are representative of the average conditions of the four Catawba generators. Also listed are a set of recert operating conditions from Cycle 4 of McGuire Unit 2. These conditions from McGuire 2 form (d the basis for the input to the ATH0S flow field calculation which has been used in the statility ratio analyses for both the
. Catawba and McGuire units. With respect to their effect on tube stability ratios, three of the operating parameters listed in Table 7-1 are of prime importance: power level / steam flow, steam pressure, and the circulation 0372M:49/CC389-18
l
{
ratio. (Primary side temperatures have only a very minor influence on stability ratios). The steam flow rate and circulation ratio influence the total bundle flow rate and tube-to-tube gap velocity in the U49nd. The steam pressure also influences the gap velocity via the void fraction and density, however, its major impact is on the tube damping. High U-bend flow along with l
~
low steam pressure results in a higher loading on the tubes with reduced damping. Both of these factors lead to higher, more limiting stability ratios. l As indicated by the comparison shown in Table 7-1, the conditions for McGuire Unit 2 are nearly identical with those of Catawbu. The only different.es to
{
note are that Catawba is operating with a slightly higher steam pressure and j slightly lower circ ratio and flow in the tube bundle. An overall measure of i i
the effect of these small differences in operating conditions on the stability j ratio can be obtained from a parameter i.ermed the 10 relative fluidelastic stability ratio which is also listed in Table 7-1. A detailed description of .
this parameter is provided in later sub-sections. However, for the present discussion it is sufficient to state that a higher value for this ratio indicates less margin to fluidelastic vibration instability and tube fatigue.
As indicated in Table 7-1, any effects of the small differences in operating ;
conditions are offsetting, as both plants are calculated to have essentially identical relative stability ratios. The beneficial effects of slightly higher pressure / lower flow for Catawba are offset l,y its lower circ ratio which produces higher bundle quality and'less damping. With identical 10 relative stability ratios, the McGuire 2 ATHOS U-bend distributions can also be applied i te the Catawba 1 stability ratio evaluation.
7.2 ATH0S Analysis Model The calculation of relative stability ratios involves comparing the stability I 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 generator. It makes use of ATH0S computed flow profiles 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 particular tube of interest, some discussion of the treatment of AVBs is necessary before presenting a description of the ATH0S model. )
0372M:49/020389-19
The ATHOS c de does not include the capability to model the presence of the AVBs in the U-bend region. However, Westinghouse ha:: modified the code to include the capability to model the AVBs via ficw cell boundary resistance
, factors. Practical lower limits of i: ell size in the ATH0S code, however, prevent a fine grid representation of the AVB V-bar share which, in turn, limits the accuracy of the AVB representation. ATH0S calculations have been performed with and without AVBs in the model. Calculations of stability ratios relative to North Anna R9C51 show that the relative stability ratios for tubes near the center of the steam generator ard essentially the same for models utth or without AVBs. The ATH0S AVB modeling sensitivity studies with uniform insertior, show some tendency for the AVB resistance effects to lower tube gap velocities r, ear the central regions and to increase velocities near the peripheral tubes. However, the magnitude of this effect is uncertain due to the limitations in ATH0S for modeling the AVBs. Further, the global flow resistance of staggered AVB insertion would be less than that from uniform insertion. Based on the sensitivity studies using ATH0S models with and without uniformly inserted AVBs, the most reliable relative st;bility ratios (for actual steam generators with non-uniform AVB insertion depths) are
~
expected using ATH0S models excluding AVBs and effects of variable AVB insertion depths by using flow test results of actual AVB geometries.
The Catawba (McGuire) analysis is based on a Cartesian coordinate system for the array of flow cells instead of the typical, and more widely used, cylindrical coordinate system. With a Cartesian coordinate system the tube array and any AVBs are arranged in a square pitched cor, figuration which is in-line with the coordinate axes. This alignment provides an improved representation of the tube region of interest in the bundle.
The ATHOS Cartesian coordinate system model for the Catawba (McGuire) steam generator consists of 18,720 flow cells having 30 divisions in the x-axis (perpendicular tc the tubelane) direction, 16 divisions in the y-axis (along the tubelane) direction and 39 divisions in the axial (I-axis) direction. In the ATHOS analysis, the steam generator is cor.sidered to be symmetrical about l' the x-axis of the tube bundle. The model therefore, consists of one-half of the hot leg and one-half of the cold leg sides of the steam generator.
1 0372M:49/02038340
I L
Figures 7-1 and 7-2 show the plan and th'e 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 i
As shown _in Figure 7-1, with the Cartesian coordinate system, the circular !
wrapper boundary is represented by a step-wise wal'1 as indicated by the heavy lines. All of the simulated flow cells outside the simulated wrapper boundary i above the first axial slab were blocked off by specifying extremely high flow resistances on the' faces of the appropriate cells. Tubelane flow slots.in the '
tube support plates are modeled also.
Hgure 7-2 shows the elevation view of the model on the vertical plane of symmetry of the steam generator. The feedwater nozzle is located at axial indices 12-11 and 12. Ten axial layers of cells were included in the U-bend near the top tube support (Figure 7-2, IZ-27 to IZ-36) to more closely model the flow conditions in the area of interest.
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 in the various flow cells. The grid lines in the Cartesian model are in-line with the tube array, providing for all of the tubes to be within the boundary of the flow cells. The fineness of the cell mesh is evident; the largest cells contain only 251.ubes while some of the smallest cells include only three tubes. Note, in particular, that additional detail was added near the bundle periphery (IY=12-16) to more closely model the inner radius tubes.
7.3 ATH0S Results The results from the ATH0S analysis consist of the thermal-hydraulic flow.
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 ATH0S are defined on the surfaces of a flow cell, the tube gap velocity and density distributions 0372M:49/020383-21 1
1 along a particular tube required for tube vibration evaluation are determined by a post-processor from the ATH0S output. The post-processor generates a data file which contains this information for all the tubes in the model and the. l
. file serves as part of the input data required for tube vibration analyses.
Because the majority of the flow cells contain more than one tube inside a cell, the tube gap velocity and density surrounding a tube are obtained by interpolation of the ATH0S calculated velocities (defined on the cell surfaces) and density (defined at the center of the cell). The post-processor performs the necessary interpolations to determine in-plane and out-of-plane velocities
]
at specific intervals along the length of the tubes. {
l Figure 7-4 shows a vector plot of the flow pattern on the vertical plane of symmetry of the steam generator (the vectors are located at the center of the flow cells shown in Figure 7-2). The zig-zag flow pattern through the split flow preheater is clearly shown in the figure. On the hot leg side the vertical flow upward through the half-moon cut-out at the center of flow distribution Plate A is also clearly shown. The vertical velocity (VZ )
component entering the U-bend region on the hot leg side is about twice that of the cold leg side as seen in Figure 7-5 (at model vertical layer index IZ-27).
The figure also shows the high VZ -component of the flow leaving the three flow slots on the top tube support plate (PLATE T) at the middle of the figure. The lateral velocity components, VR - /Vx 2 + yy 2, on the same horizontal plane (IZ-27) are shown in Figure 7-6. Viewing Figures 7-5 and 7-6 it is seen that at the entrance to the U-bend region the vertical velocity component is about twice that of the lateral velocity resultant on the hot leg side, but is about three times of that on the cold leg side. Figure 7-7 shows the plot of the void fraction contours on the vertical plane of symmetry of the steam generator. In the preheater the void fraction is essentially zero. By comparison, the hot leg side void fraction develops rapidly from the lower bundle region. In the U-bend region the void fraction is about 0.9-0.95 on the hot leg side, decreasing to about 0.60 at the bundle periphery on the cold leg side.
Figures 7-8, 7-9 and 7-10 show a sample of the individual tube gap velocity and density distributions along three tubes at Row 10. In each figure the gap velocity and density along the length of the tube are plotted from the hot leg 0372M:49/020389-22 1
l
l l
tubesheet end 9n the left of the figure to the cold leg end on the right, fhe mixture gap velocity and density distributions are required as part of the input for tube vibration analysis to determine the tube sMility ratios.
These data were generated by the ATH0S post-processor for each tube in the j model and stored in a data file. The data file was then utilized in the subsequent stability ratio calculations. Figure 7-11 shows the plot of the j average in-plane gap velocity normal to the tube and density profiles as a l 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 I U-bend at a given column location. The average velocity values are between 7.3 and 8.4 ft/sec. The velocity variations seen in the figure at Columns 22, 38 and 50 are related to the effects of the flow slots along the tubelane of the top tube support plate.
t 7.4 Relative Stability Ratio Over Operating History One aspect of the evalestion of the Catawba #1 steam generators is to examine the operating history data and use it to detardne the susceptibility to fatigue from fluidelastic vibration resulting from the 3-1/2 years of
, operation. This assessment has been completed through the use of a parameter termed the normalized stability ratio. The normalized stability ratio compares the fluideiastic stability ratio fc. each period of a plant's operation (fuel :
cycle) to a reference stability ratio based on a recent operating condition.
A plot of this ratio against operating time, therefore, provides a relative indication of the effect of past operation on the plant's fluidelastic stability ratio. This normalized time-dependent ratlo is subsequently combined with an absolute stability ratio for the reference operating point derived from j detailed three-dimensional thermal / hydraulic and tube vibration calculations.
l High values for the net stability ratio, in particular, over a significant j period'of operation, coupled with other prerequisite conditions (e.g., absence l of AVB support and presence of denting at the top tube support plate), could indicate an increased susceptibility to fluideiastic vibration instability and fatigue.
4 I
0372M:49/020389-23 I
7 h,
- j The fluidelastic stability ratio is defined as the ratio of the effective fluid velocity acting on a given tube to the critical velocity at which large i
-amplitude fluidelastic vibration initiates J
Fluidelastic Ueffective Stability Ratio, SR - [1] !
l U j critical at onset of instability 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 experiment 31 data and has been shown to be dependent upon the tube natural frequency, damping, the geometry of the tube, the tube pattern, and the fluid density, along with the appropriate correlation coefficients.
The detailed calculation of this ratio using velocity and density distributions, etc., requires three-dimensional thermal / hydraulic and tube
~
vibration calculations which are time consuming. Alternately, a simplified, one-o.mensional version of this ratio has been used to provido 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: l a,c
[2]
I Ir. this equation "cyc x" refers to each fuel cycle and " ROP" to the recent operating condition. While this simplified approach cannot account for three-dimensional tube bundle effectc, it does consider the major operational parameters affecting the stability ratio. Four components make up this ratio:
- l. a loading term based on the dynamic pressure (pV 2 ), a tube incremental mass (m) term, the natural frequency of the tube (fn), and a damping ratio l- (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 1evel to that of the recent operating point.
0372M:49/020389-24 i
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particular damping correlation which is used for all normalized stability ratio j calculations is based on a dented condition at the top tubeisupport plate (a' i clamped condition, as discussed in Section 5.2). The clamped condition is also assured in calculating the tube natural frequency. ,
As discussed previously in Section 7.1, the reference three-dimensional stability ratio calculation for the Catawba #1 steam' generators was based on-the following operating parameters which are representative of recent full power operation in McGuire'#2:
Steam Flow 3.77 x 106 lbnyt hr Steam Pressure 1007 psis 4 Circulation Ratio [ ]a,c (Westinghouse calculation)
'A series of calcul'ations were completed to generate a normalized stability ratio for each of tt.e three fuel cycles since Catawba 1 became operational in
' June, 1985. Data for this evaluation are summarized in Table 7-2. Included are typical values for full load steam pressure and primary fluid average temperature in each cycle. The number 'of days that.the plant has operated withi~n three power intervals (85-90%, 90-95%, and 95-100%) above 85% of full j power are also listed. -Since tube vibration and possible fative are '
associated with higher power operation, only the higher power operating periods are' considered important to the evaluation. Further, because the reported
' steam pressures for the first three fuel cycles differed by less than 10 psi, I the conditions for recent operation during Cycle 3 (May,1988) were also applied to the previous fuel cycles. I l
0372M:49/0203M-25 l l
1
The resulting normalized stability ratios are shown in Figure 7-12. In this figure, the normalized stability ratio is plotted against cumulative operating time above 85% power. Note that. the ratio assigned to each of the high power intervals listed in Table 7-2 (85-90%, 90-95%, and 95-100%) and' plotted in Figure 7-12 hcs been conservatively based on the highest power level in each
^
interval. Figure 7-12 indicates that the full power stabilty ratio for Cycles 1-3 is essentially the same as for the reference McGuire 2 condition.
Further, with only three cycles of operation, the time spent at high power /high relative stability ratio has been short (<700 days). Figure 7-12 also shows that the stability ratios and operating periods at the lower' power intervals are negligible compared to the full power results.
References:
7-1 L. W. Keeton,. A. K. Singha1, et al . "ATHOS3: A Computer Program for Thermal-Hydraulic Analysis of Steam Generators", Vols.1, 2, and 3, EPRI NP-4604-CCM, July, 1986.
e I
r 0372M:49/020389-26
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l
, ' Table 7-1 g Catawba #1 St'eam Gsnerator Operating Conditions and Comparison Y 4,' 'iith McGuire UnitL2 Conditions Used in the 3D ATH0$ Analysis !
l c McGuire 2 Catawba 1 Corditiens
+
". o !
6perating Used in 3D i l
.C9Dditions ATHOS Analysis j Thermal Power (% of. Full Potver) 100. 99,8 Steam Flow' Rate (lbm/hr)- 3.79 x 106 3.77 x 10 6 e :Feedwater_ Inlet Temperature (*F) 439 437.7 i
Steam Pressure (psia) 1025 1007 ,'
^
Water Level (% of span) 66 67.2- ;
~
. Primary Inlet / Outlet Temperatures 618/561 616/559
_ (*F).
_ Calculated Parameters i
' Circulation Ratio 2.26 2.35 f Bundle' Flow Rate (lbrn/hr) 8.57 x 106 B.86 x 106 l
ID Relative Stability Ratio 0.732 0.736 i
i I
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0372M:49/E20389-27
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Figure 7-1 Plan View of ATHOS Cartesian IkMiel a :
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Figure 7-3 Plan View of ATN0$ Cartesian Ikniel Ind1cating Tube Layout i-I' l
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FigureJ-4 Flow Pattern on Veritcal Plane of symmetry
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Fipwre 7-8 Latt,rel Flow Pattern on Top of Top Th W Plate 1
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. Figure 7-8 Tube Gap Velocity and Density Disstibutions for Tube Row 10/ Column 5 4
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Figure 7-10 Tube Gap Velocity and Density Distrit,ations for Tube Row 10/Colutto 40
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vigure 7-11 Average Velocity and Density in the Plane of the l U-Bends Norm 61 to Row 10 j l
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i CATAWBA 1 NORMA'LIZED STABILITY RATIO
- f 3 BASED ON HIGH POWER (>B5%) OPEM10N 1.04 :--- - - - -
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Figure 7-12 Catawba #1 Norinalized Stability Ratio Based
, on High Prmer (>85%) Operation l a
- 0361H:49/012389 40 l- +
i 8.0 PEAKING FACTOR EVALUATION This section describes the overall pecking 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 al? 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 tabe sinrolation of U-bend tubes o Test measurements and repeatability e f?!B i~sertion n depth uncertainty 8.3 North Anna 1 Configurctfun
. 8.1.1 Background i
The AV8 configuration of the ruptured tube in North Anna, R9C51, is the reference case for the tube fatigue evaluations for other plants. In accordince with the NRC Bulletin 88-02, the acceptability of unsupported tvSes I in steam generators at other plants is based on tuba specific analysis relative to the North Anna R9CGI tube, including the relative flow 96aking facters.
Thus, the support conditions _ ci the R9C51 tube are fundamental to the analyses of other tubes. 'ec3use B of the importa ce of the North Anna tube, the support ccnditions of this tube, which were originally based on "AVB t'isible" interpretationsoftheaddycurrenttest(ECT) data (Figure 8-1),were reevaluated using the projection technique developed since tile North Anna event. The projection technique is particularly vsluac19 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 summrized below.
t
' 372M:49/020389-41 t __ A
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 incations in larger row tubes in the same column. In this study, the projection technique was utilized in the
~
" blind" mode, (kVBs called strictly based on the data) as well as the reverse mode (data exNnined on t% basis of predicted AVB positions). The objective of this application was, with the greatest confidence possible, to establish the positions of the AVBs in an 8 column range around the R9C51 tube in North Anna 1, Steam Generator C.
8.1.3 Data Interpretation The ECT traces for the U-bends in Rows 8-12 (in one case, 13) were examined for Columns 4B-55. The original AVB visible calls are shown in Figura 81.
The data were examined by an eddy current analyst experienced in reading these traces, and by a design engineer knowledgeable in the geometry of the Model 51 U-bend region.
The intent of this review was to determine if the presence or absence of AVBs as shown in Figure 8-1 could be confirmed using the 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 evidence of data " spikes" where AVBs were predicted, offset indications (multiple spikes) where offset AVBs were predicted, single indications where single AVB intersections were predicted, etc. The data evaluation method used was a critical examination of the data, which was biased toward the pesence of AVDs unless a confidant 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. [
e 0372M:49/0203B9-42
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, : Figu'r e 8-4.is the "AVB visible" nip for solumns ,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.
e 8.1.4 Proje'etions The [ ]a,c ECT traces were utilized for projecting the position of the AVBs'according to the standard formatlaf the projection method.
, iThe 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 es three, projection values 0372M:49/020%19-43 l!:
I
P L.
I
}. are shown. Multiple projectiofis are expected for a tube if the AVBs on either side of the tube are not at the same elevation, or if the upper. and lower AVB l support that tube. As many as four different projections are possible if it is l*
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:
R91e 1. The projections of the same AVB based on different tubes in the same column [ ]a,c,
[
1
.a ja,c, i
Rule 2. Two adjacent tubes in the same row [
]a,c. Consequently, the difference in the [
Ja,C, A
0372M:49/020389-44
____________________________________________________________________.__________a
c -
1 ;
I l L ,
1 f
The implementation 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' adjac9nt columns are also [ j
'~
]a,c, The arrangement'of the AVBs as shown in Figure 8-5 sntisfies the rules above J l
and ils consistent with the rupture of R9C51. The resulting AVB arrangements, '
based on the projection matr u of Figure 8-5 is shown in Figure B-6.
8.1.5 Conclusions The general AVB arrangement surrounding the ruptured tube in North Anna-2, Steam Generator C, which was the tmsis for the ant. lysis, 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 l
, 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 AV3 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 j R9C51 in North Anna-1, Steam Generator C are based on consistency among all the available data.
B.2 Test Measurement Uncertainties The descriptions of the peaking factor tests and apparatus teere provided in Section 5.4. All practical measures were taken to reduce uncertainties.
, Nevertheless, some still reDa4 and should be properly accounted for. The important parameter measured during testing that has a significant impact on
- i 1
0372M:49/020389-45 l
- - ___ -_ _a
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i
. perking factor is the. air velocity. The air velocity at test section inlet'
~
l was measured using a [ ]a,c Based cn considerable l' '
experience with the 'use of such instruments, it'is known that the magnitude
, of uncertainty is very tmall. A{ ]a,c measurement uncertainty is used in
.this analysis based on past experience.
~8.3 Test Repeatability
)
During the psaking factor testing.of AVB configuration, each test was i 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 tha 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 1
i 037tM:49/020389-46
)
m a,c 1
g For the purposes of +:lis estimate, the geometry of the cantilever measuring tube in the air test model is compared with the geometry of a prototypical Row 10 tube. [
ja,c, The comparison between a U-bend tube and the model tube involve the consideration of an effective velocity associated with the flow perturbation caused by the AVBs. [
)a,c 0372M:49/020389-47
4
.[
l
. I
]a,c. Using these values, the ratic of the effective velocity for the cantilever measuring tube to that for the U-bend tube is about
[ ]a,c for the case treated.
A similar evaluation can be made for a Row 10 tube that lies in the projection or shadow of an AVB that is inserted to a depth required to support a Row 9 tube. [
ja,c, The net result is that the r;tio of the effective velocity for the cantilever tube to that for the U-band tube is about [ ]a,c, These results indicate that, for the particular assumptions used, the cantilever tube model appears to be a reasonable representation of the U-bend with respect to determining relative peaking factors for different AVB configurations. This evaluation also shcus that, on the average, the magnitude of the systematic uncertainty associated with the use of cantilever tube to simulate the U-bend is about [ ]a,c, 8.5 Air vs Steam-Water Mixture The local peaking factors from the &ir tests can be applied to the steam generator steam / water conditions either as a direct factor on the mixture 0372M:49/020389-48
s, velocity and thus'a direct factor on a stability ratio, or as a factor on the Jteam velocity only with associated impacts on density, void fraction and damping. This method leads to a reduction in tube damping which enhances the l peaking factor compared to the direct air test value. For estimating an l
7 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 l factor for the North Anna Unit I tube R9051 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 A98 insertion), there are no local open channels for flow to escape preferentially. Ther6 fore, air flow is approximately the same as steam / water flow relative to velocity perturbations. Under non-uniform AVB insertion the steam / water fled may differ from air, as the steant and water may separate from each other when an obstruction, such as an AVB, appears downstrr .1, The water would continue
, along the same channel while steam readily seeks a low resistance passage and thus turns into adjacent open channels. Two phase tests indicate a tendency
, for steam to preferentially follow the low pressure drop path compared to the ;
water phase.
Based on 9.he above discussion, the Fj are considered to more appropriately i apply to the steam phase. Thus, it follows that mixture mass velocity for the tube subject to flow perturbation can be written as follows:
r - a,c
. where gD is the vapor density, Dr the water density, aF the velocity peaking factor determined from air tests, .jg* the nominel superficial vapor velocity, and jf* the superficial water velocity. Steam quality can then be determined as follows:
l 0372049/020389-49
______2
<- ' a,c The Le11ouche-Zolotar correlation (algebraic slip model), as used in the
( ATH0S code, is applied to determine void- fraction. Subsequently, mixture i 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 l and velocity and damping coefficient, can be readily determined. Finally, a local stability peaking factor for fluidelastic vibration can be calculated '
as follows:
~
e a,c 4
where sF is the stability peaking factor, Fd the density scaling factor, Fy the velocity scaling factor, and F dp the damping coefficient scaling factor. If we use the air velocity peaking factor without translating to
. . steam / water conditions, then
- a,c 3
As shown in Table 8-1 stability peaking factors for the steam / water mixture are slightly higher than air velocity peaking factors. The difference between-the steam / water and air peaking factors increases as the air peaking factor increases.
For application to tube fatigue evaluations, the ratio of the peaking factor for a specific tube to that for North Anna R9C51 is the auantity of interest. Larger values for this ratio are conservative for the tube fatigue assessment. The North Anna R9C51 peaking factor is one of the highest peaking factor.s. As discussed in Section 8.7, a peaking factor of nearly }
(. Ja,c is determined for. the R9051 tube. The differences between [
- Ja,c. Typical values are shewn in Table 8-2. These results show 0372M 49/020089-50 a
c a that the direct application of the air test data yields the higher relative peaking-factor compared to R9C51. To obtain conservatism in the peaking factor evaluation, [
]a,c, )
Comparing the values in the first and last columns of Table 8-1, it may be noted that the stability peaking factor for steam water is [ Ja c higher than the air velocity peaking factor. On the average, the uncertainty )
associated with the conservative use of air velocity peaking factor is
[ ]a,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 ATH0S code under prototypic 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-uniforoi AVB
., insertion patterns. The current version of ATH0S has modeling limitations that prevent accurate modeling of local geometry effects. In addition, it is
- believed that 'an analysis using two-fluid modeling procedure is mandatory to a calculation of the peaking factors for a steam generator to account for the preferential steam flow along the low resistance path. 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 ATH0S model. The ATH0S results were processed by the FLOVIB code to determine stability ratios for the specific tubes of interest. The calculation was repeated using a non-uniform 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 [ ]a,c. This alternate
-._ calculation provides indy,endent corroboration of the prior discussion regarding the stability peaking factors under steam-water conditions vs in
- air.
0372H:49/020389-51
\
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 y adjacent to specific tubes. The methodology used for obtaining the AVB insertion patterns from eddy current data can ascertain the AVB location only approximately. Tne effect on peaking factor resulting from this uncertainty is tddressed using test results of AVB configurations that varied from one another by up to [ ]a,c, !
3ased 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 i configuration to a specific configuration. Figure 8-7 summarizes the AVB configurations tested.
g Position of AVB insertion depth is determined from Eddy Current Test (ECT) data. Positioning of AVB from ECT data reading is subject to uncertainty; its accuracy is probably about { ]a,c. A change of an AVB insertion depth in a given configuration leads to a different configuration, and thus a different peaking factor. A review of the tested AVB type has been made and results summarized in Table 8-3. As can be seen, a decrease in depth of an appropriate AVB tends to decrease the peaking factor, for instance, a {
]a,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 tiow perturbation. However, this applies only to those changes without inducing the reinforcement .of flow perturbation from upstream to downstream.
On the average, the uncertainty in peaking factor resulting from small 3 variations in AVB insertion (of the order of 1/2 tube pitch) is found to be l
[ ja,c, 0372M:49/020389-52
8.7 Overall Peaking Factor with Uncertainty i
As discussed in the previous subsections, there arn several aspects to be i considered in applying the laboratory test data to steam generator conditions. These considerations were reviewed one at a titie in those subsections. This section will integrate the pieces into one set of !
stability peaking factors. j
\
{
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 tubr:
(Configuration la).
It is to be noted that the test results would be applied as ratios of a specific tube peaking factor to the roc 51 peaking factor. This will redt.ce the influence of some uncertainties since the systematic uncertainties would g affect both the numerator and the denominator in the ratio of peaking factors. The major difference will be in those configurations whose peaking
. factors are significantly lower than that of R9C51. The approach employed here is intended to provide that conservative peaking factors are employed for such apparently low peaking configurations.
The uniforr., AVB configuration (2a) is selected as a reference configuration, and the peaking factors of all configurations tested are recomputed on the basis of this reference. As discussed below, some of the test uncertainties are applied to the reference case to account for its significantly low peaking relative to the R9C51 configuration.
i The uncertainties in the test results and their extrapolation are those due to test measurements, test repeatability, cantilever tubes in the test vs il-ttbes in the steam generator, and air tests vs steam-water mixture. These
, were discussed in more detail in the previous subsections. The n,agnitude of these uncertainties are listed in Table 8-4.
i 0372H:49/020389-53
o ,
Y l 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 [ l Ja,c. The RSS value of these is- )
{-Ja,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 denominator of'the final peaking factor ratio. Thus the peaking factor for configuration la (R9C51) is reduced by this amount to yield a valueof[ ]a,c instead of the [ la,c appearing in Table 5-2.
1 l
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 r. consideration and the R9C51 tube and are therefore counter L ,
balanced. However, the relative magnitude of these may be different, i particularly for configurations with much lower peaking than R9C51' t
Therefore it was judged that the [
]a,c. Similarly, as noted above, the effect on 3 peaking factor due to the uncertainty in the field AV8 configuration is also included in this reference case. Thus, [
.c Ja,c. The peaking factor of the reference' configuration 2a (Table 8-5) is raised by this amount to a value of [ Ja,c, ,
The cbange in peaking factors of configurations la and 2a resulting from the applicatio6 of uncertainties as descr1 bed above are shov'1 in Column 3 of l Table 8-5. The peaking factors of all configurations are recomputed on the basis of this reference configuration (2a).. These values are displayed in Column '4 of Table 8-5. 3 Some of the uncertainties were applied to the reference configuration (2a) in order tc 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 [ la,c is calculated for any configuration, (in Column 4 of Table 8-5), it should be
. altered to [ ]a,c. Further, for some of the configurations that are 0372M:49/020389 54
_ - _ - - - _ - - ~
i 1
conceptually similhr, the more limiting (higher) value is used. For example, a peaking factor of [ ]a,c is used for configurations 5a and 5b based on
,' their similarity to configuration Sc.
i
..j The final stability ratio peaking factors calculated on this basis (with i
configuration 2a as thh ' reference) are shown in Table 8-6.
The overall conclusions from the peaking factor assessment are: )
- 1. As noted in Table 8-4, five elements have been included in the uncertainty evaluation for the peaking factors. The uncertainty estimates were developed from both test and analysis results as descrit ed in Sections 8.2 to 8.6. The largest single uncertainty of [ ]a,cis attributable to uncertainties of up to [ ]a,c on determination of AVB insertion depths from fie~ld eddy currer.t data. This relatively large uncertainty is applicable only to low peaking conditions where the AVB uncertainties can contribate to small peaking factors. The 3 definition of "no flow peaking" was increased to encompass the small peaking effects from AVB insertion uncertainties. For the AVB patterns ,
.7eading to significant peaking factors, AVBs were positioned within uncertainties to maximize the pesking factor. For these configurations, I variations of AVB insertion within these uncertainties are expected to redyce the peaking factor compared to the final values of Table 8-6 and Figure 8-7.
- 2. Including uncertainties directed toward conservatively decreasing tne I
peaking factor for the North Anna tube R9C51, the final R9C51 pening factor is [ ]a,c relative to a no flow peaking condition such as with uniform AVB insertion depths.
8.8 Peaking Factors for Specific Tubes ;
l i I '
l4 'The A)!B positions on each insertion pattern of Figure 8-1 should be carefully noted. Where the AVBs are shown at the top of the test tube (configuration 4b for example), the AVBs at least partially block the flow part the test ,
tube and low ficw peaking factors are typically obtained. Where the AVBs are l l \
l 0372M:49/0t0389-55 !
)
Tz shown at the centerline of- the tube row above. the test tube, the- flow past the test tube is'not restricted and significant flow peaking can be obtained.ja.c
-[
Ja,c Table 8-7' summarizes the results of peaking factors.
Figure 8-7 shows the peaking factors with the pictorial representation of the AVB insertion configurations.
In applying the methodology to Catawba 1, [
]a,c Based on the Catawba tube vibration' analysis, flow peaking factors on the order of [ ] a,c for row 8 tubes and above [ -] a,c' for ruw 9 tubes would be required.for tube fatigue to be-a concern.
[
_. I L
L_ ..
Ja,C Determination of peaking factors for ideritified tubes shown in Table 8-7 are '
described in detail. Table 8-7 is broken into small tables for ease in following the description.
2 1.
0372M:49/020389-56 J
p s, , (, 1
-m, ,
i ,
.8.8.1 Steam Gencrator A I 5: The following table gives the peaking factors for steam generator A tubes
-t4 , with unique configurations of AVB insertion depths.
- c .;
.i: Steam Type of AVB Peaking Generator Row No Column No Insertion Depth Factor i.
3 .-
,e 9 R,C 4.-.
A :8 89 i
85, 84 9 90, 89 83 76, 73,-69,-50 10 -93 11:' '106, 105 98 I-
. Row 8 tubes shown belong to [ ,
.]a,C -!
-For R9C90 and R9C89 tubes, [
I.
1r 1
s
(',
.j; ja,c t172M.49/020J89 l
1
)
I 8.8.2 Etp s Generator R l
The following are row 8 tubes with unique AVB configurations.
. Steam Type of AVB Peaking Generator Row No Column No Insertion Depth Factor a,c B 8 87 80, 60, 39, 20 2
For RBC87 and R8C2 tubes, [
ja,c q The following table shows peaking fr.ctors for the remainina tubes.
Steam Type of AVB Peaking Generator Row No Column No Insertion Depth Factor B 9 92
' ""A 2
10 102 2
11 110 104 2
12 110 109 "
~ j For R9C92 tube, a conservative call was made using [ j l
ja,c I
0372M:49/020389-58 l
, - -_ . _= - _- - __
- -l .
b' For R12C110 tube, !
Ja,c
.. 8.8.3 3,tpam Generator C Tubes with unique AVB configurations were evalutted. The following table !
4 lists their peaking factors and ' types of AVB configurations used.
Steam . Type of AVB Peaking Generator Row No Column No Insertion Depth Factor
). , a,c
'C 8 .93 -
9 9 -
10 111, 110 4
3 11 111 j, ,
For R8C93 tube,-[
Ja,C o
0372M:49/020389-59 i
i
8.8.4 Stum. Generator D I
The following table presents results of peaking factors determined for row ]
8 tubes with unique configurations of the AVB insertion depths. l Steam Type of AVB Peaking Generatcr Row No Column No Insertion Oepth Factor m a,e i D 8 92 88 37 33, 32, 31 27 s a w.
R8C92 tube was of (
A ja,c Row 9 and row 10 tubes with unigte AVB configurations are listed below together with their pe& ng factors.
Steam Type of AVB Peaking Generator Row No Column No Insertion Depth Factor
~
, ,a,c D 9 76, 73, 62 56 51 l
45
, 23, 22 10 111, 110 8
5
- a 0372M:49/020389-60
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Table 8-1
.o Ja - ' Stability. Peaking Factor Dye to Local Velocity Perturbation Scaling Factors.for Steam / Water Air Velocity Void Stability Peaking Fraction. Density Velocity Damping Peaking Factor,' 'Scoling, Scaling, Scaling, Scaling, Factor,
! F a- I.v Fd Fy F dp I s l g a,c -
.c j
L l{ !
fiOTE: 1. 5tability ppaking factor for steam / water miviture is esiculated as follows: ,
r- , a,c
,, J l 12 . Damping scaling factor is calculated using modal effective veid fraction of [ ]a,c for R9051 tube. !
1 0373M:49f020389-1 f I
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- w. ,
Comparison of' Air and Steam-Water Peaking Factor' Ratios- .i
- , , i I
c Air Air Steam Steam Peaking- Peaking Peaking Peaking Faci.or ' Ratic Factor Ratio !
1-4
-% a a
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7 Effect.of Local Yariation of AVB Insertion L,,
-A to D AVB PeaMng Peaking Ratio ,
Type A Type B- Variation Factor A_ Factor B (B/A) a,C l
j, J J
+ a,C
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Uncertainties in Test Data and Extrapolation 4
Source of Uncertainly 12 3. Maanitu W
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!, It results from the inaccuracy in determining the true AVB ;
position in the field using eddy current data. ,
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> ;l Extrapolation of Test Results to Steam Generator Conditions- -
i
, i Peaking Factor j Test Data with Referenced to Cor. figuration ' O.tta Uncertainties? Configuration 2a
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Final Peaking Factor .
c " -l
- ,1 , ' y Configuration Peaking Factor l
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p Tabl'e B-7
- h. Stability Peaking Factors for Specific Tubes P'
L, Catawba Unit 1-
.:' Steam .
Type of #.VB Peaking Generator Rots No Column No Insertion Factor p -- -;-
A' 8 P9 I '
,J 85, 84
'y-s 9 90, 89 i 83 1
'76, 73, 69, 50
'It 93 11' 106, 1.05 98 All of the Remaining ,
'B 8 87 ;
80,'60, 39, 20 2
9 92 2
10 102 -;
2'
, 11 110 104
,. 2 12 110 109 All of the Romaining C 8 93 '
9 9 10 111, 110 4
11 111 All of the Remaining 'l D 8 92 88 37 33, 32, 31 27 9 76, 73, 62 56 51 45 23, 22 +
10 111, 110 8 L 5
All of the Remaining i, ,
Note: Row 7 tubes were not listed because the maximum allowable peaking factor for Catawba Unit 1 is at least greater than 1.90 for row 7 tubes.. Configurations for the row 7 tubes yield peaking factors much less than 1.90.
t 0373M:49/020389-7
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, Figurs 8 5 North Anna 1, Steam Generator C R9C51 AVB
[ e. ]a,e Katrix
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. TYPE OF AVB PEAKING TYPE OF AVB PEAKING TYPE OF AVB PEAKING INSERTION FACTOR INSERTION FACTOR INSERTION FACTOR
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Figure 87 Final Peaking Factors for Catawba 1 4
9.0 STRUCTURAL AND TUBE VIBRAlION ASSESSMENTS 9.1 Tube Mean Stress This section summarizes the analysis to determine stresses in a dented, but undeformed, tube at 100% power. Loads imposed on the tube torrespond to steady-state pressure, differential thermal expansion between the tube and the suppert plate, and a thru-wall thermal gradient. The analysis assumes the tube {
to be [ Ja,c at cold shutdown.
A summary of the temperature and pressure paramaters, used for mean stress calculations, at 200% power in the vicinity of the top support plate is provided in Table 9-1. The tube temperature corresponds to the average of the 5
primary-side water temperature and the plate temperature. The resulting tube / plate radial interference is [ ]a,c, Stresses due to differential pressure and interference loads are calculated i using finite element analysir with the model shown in Figure 9-1. The taodel prescribes [
]R.c Two reference cases were run using the finite element model, the first for a primary-to-secondary side pressure gradient of 1000 psi, and the second for a
[ ]a,c inch rntial interference between the tube and plate. The pressure case incorporates the axial load on the tube by applying a pressure loading along the top face of the model. Plots showing the distribution of stress for the tube outer surface for the two reference cases are provided in Figures 9-2 and 9-3. Thermal bending stresses due to the thru-will thermal gradient are calculated to be 7.3 ksi using conventional analysts techniques. The rembined stress distribution along the tube length, in Figure 9-4, was obtained by combining the thermal bending stresses and the reference solutions with
- appropriate multipliers based on 100% power operating parameters.
C373M:49/020389-15
The maximum hxial tensile stress is 19.0 ksi and occurs approximately 0.133 l
inch above the top surface of the support plate. Adding, for conservatism, the surface stress due to pressure,1.0 ksi, gives an applied mean stress of 20.0 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 compres3ive at the tube surface, but 5-10 pils below the surface, the stress leyels change to 10-15 ksi tensile. Combining the arplied and residual stresses results in a cumulative )
mean stress of approximately 3b ksi, assuming tube denting without deformation.
l If a tube is dented with deformation, the mean stress is limited by tube ]
yielding. For the case of dented tubcs with deformation, the maximum j effect of mean stress was incorporated by using omax " Oy in determining i stability ratios and fatigue usage. )
9.2 Stability Ratio Distribution Based Upon ATH0S
. An assessment of the potential for tubes to experience fluid elastic instability in the U-bend region has been perforn'ed for each of the tubes in )
rows eight through twelve. This analysis utilizes FASTVIB, a Westinghouse proprietary finits element based computer code, and PLOTVIB, a post processor to FASTVIB. These codes predict the individual responses of an entire row of stcain generator tubing exposed to a location dependent fluid velocity and density profile. The program calculates tube natural frequencies and niode j shapes using a linear finite element model of the tube. The fluid elastic 5 stability ratio V e/Uc (the ratio of the effective velocity to the critical velocity) and the vibration amplitudes caused by turbulent'e are calculated for a given velocity / density / void fraction profile and tube support condition. The !
velocit: , density and void fraction distributions are determined using the ATH0S computer code as described in Section 7.3. The WECAN generated mass and stiffnest matrices used to represent the tube are also input to the code.
(WECAN is also a Westinghouse proprietary co'nputer code.) Additional input to
. FASTVIB/PLOTVIB consists of tube support conditions, fluid elastic stabiHty constant, turbulence constants, and location dependent flow peaking factors.
i 0373M:49/020389-16 j
h l
T -
This process was performed for the Catawba #1 steam generator tubes and also for the North Anna Row 9 Column 51 tube (R9C51) using similarly appropriate ATH0S models. Ratios of the Catawba #1 results to those for North Anna Unit 1 R9C51 were generated to produce a quantity that could be used to provide an initial assessinent cF the Catawba #1 tubes relative to the ruptured tube at North Anna Unit 1.
Figure 9-5 shows the results of this process for each of the rows under investigation. The relative ratios are obtained using the following conditions for Catawba #1 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.
. It is to be noted that the stability ratios plotted in Figure 9-5 are s
composites of all steam generators. That is, any peaking effect for a given tube location indicated on the plot represents the maximum value of ;
the peaking factor in all steam generators at that location.
A horizontal line is drawn at the relative stability ratio value of 0.90.
This identifies the point where a ten percent reduction in stability ratio exists relative to North Anna 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. 1 Figure 9-5 indicates that most tubes in Rows 8 through 12 of the Catawba #1
, steam generators, with the exception of R11C111 in SG-C and R12C109, R12C110 in SG-B, lie below the 90% line.
0373M:49/020389-17
l 9.3 Stress Ratio Distribution with Peaking Factor I
An evaluation was performed to determine the ratio of the Catawba #1 tube i
^
stress over the North Anna R9C51 tube stress. This ratio is determined using
, relative stability ratios discussed in the previous section, relative flow I peaking factors (Table 8-7 factors divided by [ ]a,c) tube size, and ,
bending moment factors. Sections 4.2 and 4.3 contain additional information I 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) Location-dependent flow peaking effects
- 6) Tubes are assumed to be dented with deformation (labeled with denting) or clamped at the top support plate due to crevice corrosion (labeled without denting).
A tube can be considered acceptable if the stress ratio is less than 1.0 when calculated using the procedure described in Sections 4.2 and 4.3 and including the conditions listed above and subject to confirmation of fatigue usage acceptability. Conformance to these requirements implies that the stress acting on a given tube is expected to be insufficient to produce a fatigue event in a manner similar to the rupture that occurred in the R9C51 tube at North Anna Unit 1.
l 0373M:49/020389-18 l
. - - . . ~ . . . _ . . _. _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _ _ . _ _ . _ _ _ - _ _ _ _ - - _ _ _ _ . _ . _ _ . ~ _ -__-_-__U
Figure 9-6 shows the results of the stress ratio calculations for each of the Catawba #1 tubes-in Rows 8 through 12. As in Figure 9-5 for the stability ratio, the stress ratios in Figure 9-6 represent the composite ratios for all Catawba #1 steam generators. (Refer to Table 9-2 for critical tubes in
[,
individual steam generators). 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. Figure 9-7 contains the results for the case where tube deformation is not present. These figures )
demonstrate the effects of varying the applied mean stress on the tube. Using l the mean stress, present in the undented results, produces stress ratio values i that are lower than stress ratios calculated for tubes in the dented condition.
As can be observed in Figures 9-6 and 9-7, one tube R11C111, located in SG-C lay above the 1.00 stress ratio line for t>oth the dented and undented condition l at Catawba Unit 1. Final evaluations have indicated that denting was present at the top TSP durir,g the latest inspection period.
l As noted in Section 9.5, it is conservatively recommended that all tubes with stress ratios exceeding acceptance criteria of 1.0 be removed from service.
An evaluation has also been performed to determine the required relative flow peaking that will produce a stress ratio not greater than 1.0. Figure 9-8 contains the results of this process for all tubes in Rows 8 through 12. The j figure was generated using the conditions outlined previously with the additional constraint that the tubes are dented. Note that this figure reads :
opposite of the previous figures, i.e., the top curve in the figure corresponds to Row 8 and the bottom curve corresponds to Rew 12. Maximum Allowable Relative Flow Peaking is the required relative flow peaking (0.68 corresponds to no flow peaking) that, if used on the given tube, will produce a stress ratio (with denting) not to exceed 1.0. This curve can be used to identify the relative flow peaking required Defore preventive action would be recommended and, when used in conjunction with the actual flow peaking associated with each 0373M:49/020389-19
-__-___-_-_-_-_--_---_-___---_-__-_=__L---_A
i i
tube, to determine the margin (if any) present. This has also been performed in Table 9-2.. The colurm with heading " Max Allow Rel FPeak" identifies the relativc flow peaking facture that would be permitted, on a tube by tube basis, before the stress ratio criteria wod d be excceded. As can be observed in the table and figure, the inner row tubes have larger values of allowable relative flow peaking when compared to the outer rows. ]
l 9.4 Cumulative Fatigue Usage !
l l
All tubes that are unsupported and have a stress ratio f 1.0 have a maximum {
stress amplitude that is < 4.0 ksi (from 9.5 ksi) duce a 10% reduction in the j stability ratio for the North Anna Row 9 Column 51 tube was the criteria basis. The stability ratios for the Catawba #1 tubing are based on the current operatirg parameters and with f)ture operation on the Sams basis, tha tubes are not expected to rupture as a result of faticjue if 1) they meet the stress ratio criteria of 51.0 and 2) their current and futur.a fatig"e usage will totsl less than 1.0.
Based on the above analyses, most Catawba Unit I tubes meet the relative stress
_ retio criteria. Preventative action has been recommended fer those tubes that do not meet the stress ratio criteria. Taun ?-2 provides a sunaary of the combined relative stability ratios and the stress retios for the more salient unsupported tubes in Rows 8 through 12.
t Acceptability of the Catawba Unit I tubing for fatigue is acco-911shed by demonstrating the acceptability of the tube remaining in seuice with the j highest stress ratio that does not exceed the 1.0 stress ratio criteria. In the Catawba Unit 1 plant this tube is iocated at R10C93 in SG-A a'd n has a stress ratio of 0.49. Note that Table 9-2 indicates that three other tube have larger stress ratios than this tube: RllC111 located in steam generator C, along with R12C109 and RlRC110 located in steam generator B. These tubes !
have had cable stabilizers installed that effecthely reduce the stress ratio .
. to values lower than 0.49, therefore these tubes are not the most limiting :
tubes remair.ing in service. Appendix A contains a discussion of the effects of
. installing cable tube dampers in the tubes of interett.
0373M:49/020389-20
- - - - _ _ _ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - _ - _ _ _ _ _ _ - _ _ _ _ _ . . _ _ _ _ _ _ _ ._]
, i 1
l, Assuming the worst case tube (R10C93) has been dented since the first cycle and l L continues to operate ulider conditions, the total usage including the remaining i term of the operating license would be 0.04.
Reference:
l- 91 Westinghouse Research & Development Report-77-ID2-TUCOR-R2, " Residual Stresses in Inconel 600 Stean Generator Tubes - Part II: Straight Tubes", ;
Westinghouse Research Laboratories, Proprietary Class 2, D. L. Harrod, October 21, 1977.. l
'1 f
S 0373M 49/020363-21 3
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., Table 9-1 1
100% Power Operating Parameters Cetawba #1 Bounding Values for Mean Stress Calculations o !
t Primary Pressure - 2250 psia Secondary Pressure = 1000 psia Pressure Gradient = 1250 psi L
l
.3 Primary Side Temperature * = 589.8'F t
. Secondary Side Temperature - 544.6*F f Tuoe Temperature - 567.2'F i
Average of Thot = 618'F and Tcold = 562*F. i i
t..
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. TABLE 9 Catawbi Unit 1 Tebes With Significant.RSR's or Stress RatiosL j
- - Steam: Row s ol . : Rel. Flow ' Max Allow RSR* Stress Ratio i Gen. No . -- .No. Peaking Rel. FPEAK FPEAK W Dent' W/0 Dent l
?A. 8 84P35,89_ 0.68 '1.32 0.43 0.03 0.02
+ J 69,89,90 0.68' 1.19 0.49 -0.05 -0.04 83 -0.78 l'.18 0.57- 0.10 -0.08 .j 50,76,73- 0.68 1.21' O.48 0.04 -0.04 -
10 93: 0.95 1.08 -0.78 0.49 -0.41 -
11 .105,106 , 0.7G 0.91' O.76 0.35 0.29 98 0.79 0.94 0.76 -0.37 0.31 B 8 87 0.68 -1.32 0.43 0.03 0.02 20,39,60,80 0.72 1.36 0.43' O.03~ 0.02 I.
9 92 .~ 0.95 - 1.21 0.67 0.25 0.21 2 0.68- 1.73 0.34 0.01 0.01 ,
i 10 102 0.68 1.03 0.58 0.10 0.08 !
, 2 0.68 1.45 0.41 0.02 0.01 11 110 0.68- 0.90 0.69 0.21 0.17 c h' 104- 0.68 0.89 0.67 -0.18 0.16 2 0.68 1.21, 0.51 0.04 0.03 12- 110 0.79 0.80 0.92 0.89 0.75 109 0.80 0 81 0.92 0.89 0.75 C 8 93 0.80 1.38' O.48- 0.05 0.04 94 9 0.68 1;14- 0.51 0.06 0.05
- l. 10 110,111 -0.76 1.01 0.66 0.20- 0.17 4 0.79 1.02 0.68 0.24 0.20 11 111. 0.97 0.89 0.99 1.96. 1.66 ;
.D 8 88,92 0.68 1.33 0.42 0.03 0.02 37- 0.79 1.48 0.44 0.03 0.03
- - 31,32,33 0.76 1.32 0.47 0.05 0.04 27 1.00 1.33 0.62 0.21 0.17
- l. 9 62,72,76,45 0.68 1.19 0.49 0.05 0.04 L 56 0.79 1.17 0.58 0.11 0.09 51 0.72 1.27 0.48 0.04 0.04
-- 22,23 0.76 1.21 0.54 0.07 0.06 10 110,111 0.76 1.01 0.66 0.20 0.17 5,8 0.68 1.01 0.59 0.11 0.09 I
0373M:49/020389-23 1
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Interference Load on Tube l
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