ML20012D766
| ML20012D766 | |
| Person / Time | |
|---|---|
| Site: | North Anna  |
| Issue date: | 05/31/1989 |
| From: | Connors H, Frick T, Houtman J WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19293A296 | List: |
| References | |
| WCAP-12266, NUDOCS 9003280345 | |
| Download: ML20012D766 (125) | |
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WESTINGH0'USE' CLASS 3.
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NORTH ANNA UNIT 2 c'
EVALUATION FOR TUBE VIBRATION INDUCED FATIGUE i
MAY 1989 AUTHORS:
H. J. CONNORS M. H. HU T. M. FRICK A. Y. LEE J. M. HALL R. M. WILSON G. W. HOPKINS R. M. WEPFER J.~L. HOUTMAN
" * * ~
APPROVED: h J. L. HOUTMAN, ACTING MANAGER STEAM GENERATOR ENGINEERING
.1
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L WESTINGHOUSE ELECTRIC CORPORATION SERVICE TECHNOLOGY DIVISION P.O. B0X 355 PITTSBURGH, PENNSYLVANIA 15230
. 9224M:1E-050189-1 j
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k ABSTRACT
=On July 15, 1987, a steam generator tube rupture event occurred at North Anna Unit 1.
The cause of the tube rupture has 'been determined to' be high cycle fatigue. The source of the loads associated with the fatigue mechanism is a combination of a mean stress level in the tube with a superimposed alternating stress. T.he mean stress is the result of manufacturing induced residual stress, applied stress and stress due to denting of the tube at the top tube support plate, while the alternating stress is due to out-of plane deflection of the tube U-bend attributed to flow induced vibration.
For tubes without AVB support, local flow peaking effects at unsupported tubes are a significant contribution to tube vibration amplitudes.
Subsequent to the tube rupture, a tube fatigue analysis was performed for the North Anna 2 plant, and several modifications were implemented.
Downcomer flow resistance plates were installed in all steam generators, resulting in a nominal (
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Testing since this October, 1987, evaluation has determined that single-sided l
support of the tubes is sufficient to limit fluidelastic excitation of the tubes. Not-uniform AVB insertion configurations have been tested to define l
AVB. positions which, dependent upon local flow conditions, produce excessive tube vibration.
This analysis provides the justification for the removal of a majority of the sentinel plugs installed in North Anna #2.
The justification is developed from a detailed AVB insertion mapping, updated thermal / hydraulic analysis, and vibration analysis.
The fatigue analysis considered the effects reduction.
J of prior operating history on tube fatigue and of a postulated Thot The report concludes that two previously-installed sentinel plugs may be required to remain in place, depending upon future operating conditions and
-l desired service period, and the remaining 116 sentinel plugs may be removed.
9213M:1E-042789-2
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SUMMARY
0F ABBREVIATIONS l,
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American Society of Mechanical Engineers.
ASME Analysis;of the Thermal Hydraulics of Steam Generators-
-ATHOS Anti-Vibration Bar AVB AVT-All' Volatile Treatment ECT Eddy Current Test Electric Power Research Institute EPRI FFT-Fast Fourier Transform Flow Induced Vibrations FLOVIB
-Modal Effective Void Fraction MEVF Outside Diameter OD Root Mean Square RMS Stability Ratio SR Tube Support Plate TSP
'F degrees Fahrenheit hr :
hour-measure of stress - 1000 pounds per square inch-ksi l
lb pound 0.001 inch mils MW-mega watt l'
ps'i measure of stress pounds per square inch psia-measure of pressure - absolute 1
9213M:1E-042789-3
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TABLE OF CONTENTS f
F SECTION-
'1.0L Introduction 7,
2.0 Sumary and Conclusiens
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 Overall Conclusion.
3.0 Background
3.1 North Anna Unit 1 Tube Rupture Event 3.2 Tube :xamination Results 3.3 Mechanism Assessment t
- 4.0 Criteria for Fatigue Assessment 4.1 Stability Ratio Reduction Criteria 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 Fluideiastic Instability of Columnwise Variations in AVB Insertion Depths 5.5 References 9213M:1E-042789-4
.m.
TABLE OF CONTENTS (CONTINUED) j SECTION i
f 6.0 ?
_ Eddy Current Data and AVB Positions
~
3 6.1 AVB Assembly Design 7
6.2 ' Eddy Current Data for AVB Positions 6.3 AVB Projection and Mapping I
6.4' Tube Denting at Top Tube Support Plate 6.5 AVB Map Interpretations by Generator 7.0 Thermal and Hydraulic Analysis t
7.1 North Anna 2 Steam Generator Operating Conditions 7.2 ATH0S Analysis Model
'l 7.3. ATH0S Results p
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7.4 ' Relative Stability Ratio Over Operating History L
l8.0 Peaking Factor Evaluation l'
-8.1 North Anna-1 R9C51 Configuration 81 2 Test Measurement Uncertainties 8.3 Test Repeatability l
8.4 Cantilever vs U-Tube l'
8.5 Air vs Steam Water Mixture
]
.8.6 AVB Insertion Depth Uncertainty l
8.7 Overall Peaking Factor with Uncertainty 8.8 Peaking Factors for Specific Tubas
'9.0 Structural and Tube-Vibration Assessments 9.1 ' Stability Ratio Distribution Based Upon ATHOS 9.2 Stress Ratio Distribution with Peaking Factor 9.3 Cumulative Fatigue Usage t
l5213t&1E-042889-5
1 LIST OF FIGURES FIGURE 3-1 Approximate Mapping of Fracture Surface of Tube R9C51 S/G "C" Cold Leg, North Anna Unit 1 j
3-2' Schematic Representation of Features Observed During TEM i
Fractographic Examination of Fracture Surface of Tube R9C51, S/G "C" Cold Leg, North Anna Unit 1 3-3 Calculated and Observed Leak Rates Versus Time 4-1 Vibration Displacement vs. Stability Ratio 4-2 Fatigue Strength of Inconel 600 in AVT Water at 600*F 4-3 Fatigue Curve for Inconel 600 in AVT Water Comparison of Mean Stress Correction Models 4-4 Modified Fatigue with 10% Reduction in Stability Ratio for Maximum Stress Condition 4-5 Modified Fatigue with 5% Reduction in Stability Ratio for Minimum Stress Condition i
5-1 Fluidelastic Instability Uncertainty Assessment 5-2 Instability Constant - 6 5-3 Instability Constants, B, Obtained for Curved Tube from Wind Tunnel Tests on the 0.214 Scale U-Bend Model 5-4 Damping vs. Slip Void Fraction 9213M:1 E-042789-6
~
LIST OF FIGURIS (Continued) i i
i.
FIGURE i
5-5 Overall View of Cantilever Tube Wind Tunnel Model 5-6
. Top View of the Cantilever Tube Wind Tunnel Model l
i 5-7 Fluideiastic Vibration Amplitude with Non-Uniform Gaps 5-8 Typical Vibration Amplitude and Tube /AVB 1mpact Force Signals for Fluideiastic Vibration with Unequal Tube /AVB Gaps 5-9 Conceptual Design of the Apparatus for Determining the Effects on Fluideiastic Instability of Columnwise Variation; in AVB Insertion i
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 AVBs and Top Flow Screen 5-12 AVB Configurations Tested - North Anna Unit 2 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 AVB Projection Depth = 9.00 6-3 AVB Projection Depth = 9.15 6-4 North Anna #2 SG-A AVB Positions 9213M:1E-o42789-7
7, i
f LIST OF FIGURES (Continued)
FIGURE 6-5 North Anna #2 SG-B AVB' Positions 6-6 North Anna #2 SG-C AVB Positions i
7-1 Comparison of Relative Stability Ratios Calculated from 10 and 3D Methods 7-2 Plan View of ATHOS Cartesian Model for North Anna 7-3 Elevation View of ATHOS Cartesian Model for North Anna 7-4 Plan View of ATHOS Cartesian Model Indicating Tube Layout 1
7-5 Flow Pattern on Vertical Plane of Symnetry r
7-6 Lateral flow Pattern on Horizontal Plane in the U-Bend Region 7-7 Contours of Vertical Velocity Component on a Horizontal Plane in the U-Bend Region l
7-8 Tube Gap Velocity and Density Distributions for Tube Row 9/ Column 3 7-9 Tube Gap Velocity and Density Distributions for Tube Row 9/ Column 20 7-10 Tube Gap Velocity and Density Distributions for Tube Row 9/ Column 44 l
7-11 Average Velocity and Density in the Plane of the U-Bends Normal to Row 9 7-12 North Anna 2 Normalized Stability Ratio Based on High Power (>85Y,)
Operation 9213M:1E-042789 B
4 4
LIST OF FIGURES (Continued) l FIGURE 8-1 Original North Anna AVB Configuration (Configuration Ib) 8-2 Schematic of Staggered AVBs i
B-3 AVB " Pair" in ECT Trace 8-4 North Anna 1. Steam Generator C: AVB Positions Critical Review AVB Visible" Calls 8-5 North Anna 1. Steam Generator C, R9C51 Projection Matrix P
8-6 North Anna R9C51 AVB Final Projected Positions 8-7 Final Peaking Factors for North Anna Unit 2 9-1 Relative Stability Ratio and Relative Flow Peaking - North Anna 2 -
Premod 9-2 Relative Stability Ratio and Relative Flow Peaking - North Anna 2 -
Postmod Without T Reduction hot 9-3 Relative Stability Ratio and Relative Flow Peaking - North Anna 2 -
Postmod With T Reduction hot 9-4 Stress Ratio Vs. Column Nw.ber - Premod - North Anna Unit 2 9-5 Stress Ratio Vs. Column Number - Postmed - Without T Reduction -
hot North Anna Unit 2 9213M:1E-042789-9
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LIST OF FIGURES (Continued) m FIGURE
[ g + f 9-6 Stress Ratio Vs._ Column Number - Postmod With Thot.- North Anna Unit: 2 j.
- - 7-Maximum Allowable l Relative Flow Peaking - North Anna. Unit 2 I
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LIST OF TABLES I
IABLE 4-1 Fatigue Usage per Year Resulting From Stability Ratio Reduction 5-1 Wind Tunnel Test on Cantilever Tube Model 5-2 Fluidelastic Instability Velocity Peaking Ratios for Columnwise Variations in AVB Insertion Depths 6-1 North Anna #2 1 AVB Signals Determined to be Supported 2 North Anna Unit 2 Unsupported Tube Summary 7-1 North Anna Unit 2 Steam Generator Operating Conditions and Comparison with the Corresponding Conditions in Unit 1 7-2 North Anna Unit 2 Operating History Data 1
i 8-1 Stability Peaking Factor Due to Local Velocity Perturbation B-2 Comparison of Air and Steam-Water Peaking Factor Ratios l
8-3 Effect of Local Variation of AVB Insertion 84 Uncertainties in Test Data and Extrapolation 8-5 Extrapolation of Test Results to Steam Generator Conditions 8-6 Final Peaking Factor 8-7 Stability Peaking Factors for Specific Tubes - North Anna 2 9213M:1E-042789-11
l
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1
-LIST OF TABLES-(Continued) l l
TABLE l
.c 9-1 Tubes With Significant' Relative Stability Ratio and Relative Flow Peaking-t
'l 9-2 ~
Tubes With Significant Relative Stress Ratio or Flow Peaking - Stress Ratio i
i
. 9-3 Summary of North Anna 2 Fatigue Usage Factors i
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' 9213M:1E-0427a9-12
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1.0 INTRODUCTION
This report documents the re-evaluation of ste'.m generator tubing at North Anna Unit 2 for susceptibility to fatigue-induced cracking of the type experienced at North Anna Unit 1 in July,1987.
Its purpose is to identify susceptible tubes and thereby justify de-plugging tubes that were plugged in 1987 with sentinel plugs.
The evaluation includes three-dimensional flow analysis of the tube bundle, air-tests performed to support the vibration analytical procedure, field measurements to establish AYB 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 North Anna Unit 2 in order to account for plant specific features of the tube loading and response.
Section 2 of the report provides a summary of the North Anna Unit 2 evaluation results and overall conclusions. Section 3 provides background for the tube rupture event which occurred at North Anna Unit 1 including results of the examination of the ruptured tube and a discussion of the rupture mechanism.
The criteria for predicting the fatigue usage for tubes having an environment j
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 stability ratio and tube stress distributions, and accumulated fatigue usage, for the North Anna Unit 2 steam generator small radius U-tubes.
9213M:1E-o42789-13
i 2.0
SUMMARY
AND CONCLUSIONS I
i The North Anna Unit 2 steam generators have been evaluated for the
]
susceptibility of unsupported U-band tubing with denting at the top tube support plate to a fatigue rupture of the type experienced at Row 9 Column 51 (R9051) of Steam Generator C a't North Anna Unit 1.
The evaluation uses Eddy CurrentTest(ECT)datainterpretedbyWestinghouse.
)
2.1 Background
The initiation of the circumferential crack in the tube at the top of the top tube support plate at North Anna 1 has been attributed to limited displacement, fluid elastic instability. This condition is believed to have prevailed in the R9C51 tube since the tube experienced denting at the support plate. A combination of conditions were present that led to the rupture. The tube was not supported by an anti-vibration bar (AVB), had a higher flow field due to local flow peaking as a result of non-uniform insertion depths af 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 (SGs). 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 objec,tive.
This same fatigue criteria is applied as the principal criteria in the fatigue
-evaluation reported.herein, i
9213M:1E-o42789-14
j 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 North Anna tube divided by the stress an'olitude for the North Anna 1, R9C51 tube.
Displacements are computed for the unsupported U-bend tubu in Rows 11 and inward, (descending row number) using relative stability ratios to R9C51 of North Anna 1 and an appropriate power law relationship based on instability displacement versus flow velocity.
Different U-bend radius tubes will have different stiffness and frequency and, therefors, 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 so that a stress ratio of 1.0 or less produces acceptable stress amplitudes and fatigue usage for the North Anna tubing for the reference fuel cycle analyzed.
Therefore, a stress ratio less than 1.0 provides the next level of acceptance criteria for uncupported tubes for which the relative stability ratio, including flow peaking, oxceed 0.9.
The stability ratios for North Anna 2 tubing, the corresponding stress amplitud&, and the resulting cum.ulative fatigue usage must be evaluated relative to the ruptured tube at Row 9 Column 51: North Anna 1, Steam Generator C, for two reasons. The local effect ca the flow field due to various AVB insertion depths is not within the capability of available analysis techniques and is deterndned by test as a ratio between two AVB configurations, in addition, an analysis and examination of the ruptured tube at Xerth Anna 1 provided a range of initiating stress amplitudes, but could only bound the possible stability ratios that correspond to those stress amplitudes. Therefore, to minimize the influence of uncertainties, the evaluation of North Anna 2 tubing has been based on relative stability ratios, relative flow peaking factors, and stress ratios.
The criteria for establishing that a tube has support from an AVB and therefore eliminate it from further considerations is that it must have at least one sided AVB support present at the tube centerline.
The criteria is based on test results which show that one sided AVB support is sufficient to 9213M:1E-042789-15
limit the vibration amplitude for fluidelastic excitation.
AVB support is established by analysis of eddy current (EC) meesurements and is a key factor in the determining the local flow peaking factors. The local flow peaking produces increased local velocities which cause an increase in stability ratio. A small percentage change in the stability ratio causes a significant change in stress amplitude. The relative flow peaking factors for Horth Anna tubing without direct AVB-support have been determined by test. These flow peaking factors normalized to the North Anna R9C51 peaking, are applied to a
relative stability ratios determined by 3-D tube bundle flow analysis, to obtain the combined relative stability ratio used in the stress ratio l
determination.
2.3 Denting Evaluation The Eddy Current (EC) tapes were evaluated to determine the condition of the tube / tube support interface of unsupported tubes immediately below the apexes of the AVBs. Analysis of the August-September 1987 North Anna 2 eddy current
~
inspection data indicated that most of the tubes had crevice corrosion product buildups at the top tube support plate, but were not dented with deformation.
Of the 118 tubes recommended for preventive plugging,104 tubes were evaluated l
as having top tube support plate corrosion, 2 showed denting with deformation, 6 showed no detectable denting (or crevice corrosion product), and 6 were l
unreadable. Per the NRC Bulletin 88-02 definition, all but the 6 tubes showing no detectable denting are required to be considered as dented in the analysis. For conservatism in the evaluation, all of the tubes esaluated are postulated as being dented.
The effect of denting on the fatigue usage of the tube has been conservatively maximized by ascuming the maximum 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 North Anna Unit 2 SGs have two sets of Alloy 600 AVBs.
The lower AVBs have a rectangular cross-section and extend into the tube bundle approximately as far as Row 11.
They provide a nominal total clearance between a tube 9213M.1 E-042789-16 j
oithout ovality and the surrounding AVBs of [
Ja.c inch.
Including average tube ovality for. a Row 11 tube, the nominal total tube to AVB clearance is about (
)"'C inches.
The upper AVBs also have a rectangular cross section and extend into the tube bundle approximately as far as Row 13, 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 lower AVBs are required.
The eddy current data for North Anna 2 were reviewed to identify the number of tube /AVB intersections and the location of these intersections relative to the apex of a given tube. This information was used in calculations to determine the deepest penetration of a given AVB into the tube bundle. For the North Anna 2 steam generators, the AVB support 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 AVB support.
If only the apex of a North Anna ~2 AVB assembly is near or touching the apex l
of a tube U-bend, only one AVB signal may be seen.
In this case, adequate tube support cannot be assumed without supplemental input. Support can be determined if projection calculations based on the AVB intercepts of higher row number tubes in the same and adjacent columns verify insertion depth to a point below the tube centerline.
Maps of the~AVB insertion depths for North Anna 2 are shown in Figures 6-2 through 6-4.
These AVB maps list the results of the projection calculations from the smallest row tube for which suitable data exist to make a projection.
r 2.5 Flow Peaking Factors AVB position evaluations were used in evaluating the local flow peaking factors.
Local flow peaking produces increased local velocities which cause an increase in stability ratio. A small percentage change in the stability ratio can cause a significant change in stress amplitude. The test-based flow l
l 9213M.1 E-042789-17
+
b peaking factors are normalized to the North Anna R9C51 peaking, and are applied to relative stability ratios determined by 3-D tube bundle flow analysis to obtain the combined relative stability ratio used in the stress ratio determination.
2.6 -Tube Vibration Evaluation The calculation of relative stability ratios for North Anna 2 makes use of detailed tube bundle flow field information computed by the ATHOS steam generator thermal /hydraulis analysis code.
Code output includes three-dimensional distributions of secondary side velocity, density, and void fraction, along with primary fluid and tube wall temperatures.
Relative stability ratios of pre-DFRP, post-DFRP and post-T reduction hot North Anna Unit 2 (Row 8 through Row 12) tubing (relative to Premod R9C51 of North Anna 1) are plotted iin Figures 9-1, 9-2 and 9-3, respectively.
These relative stability ratios helude relative flow peaking factors.
Stress ratios of pre-DFRP, post-DFRP, and post-T reduction for North Anna Unit 1 hot are plotted in Figures 9-4, 9-5, and 9-6, and are calculated based on clamped tube conditions with denting at the tube support plate.
For all three steam generators, the stress ratios for all remaining tubes in Rows 8 through 11 are less than or equal to 1.0, even when the tubes are assumed to be unsupported, i
One-dimensional performance and relative stability ratio analyses of operating data for North Anna #2 have been completed for each fuel cycle since the plant became operational in 1980. These data include operation prior to the i
l installation of downcomer flow resistance plates (and prior to the tube rupture event in Unit #1), recent operation in Cycle 6 following the i
installation of the plates, and projected cperation with reduced primary water temperature and steam pressure. The latter conditions were base,d on a December 1988 test with the turbine valves wide open to obtain the lowest L
.possible steam pressure which can be obtained while still maintaining full power. Reduced steam pressures are of interest because they result in higher, potentially more limiting stability ratios.
1 9213M:1E-042789-18
Comparisons of 1-D relative stability ratios calculated for each of these
]
conditions are made with the ratios determined for the corresponding conditions in Unit #1.
In all cases, the stability ratios for the Unit #2 conditions are within 1% of the ratios calculated for the corresponding j
conditions in Unit #1.
Based on this close agreement, the results of the existing 3-0 ATH0S flow field /stabilty ratio evalution for Unit #1 are applied to Unit #2, with only small adjustment factors from the 1-D stability ratio evaluations.
Lists identifying the support conditions determined in the analysis, for use aith the AVB insertion maps, are provided in Tables 6-1 and 6-2.
Relative stability ratios and stress ratios of North Anna #2 Row 8 through 12 tubes versus R9C51 of North Anna #1 are listed by steam generator in Tables 9-1 and 9-2.
Both the relative stability ratios and the stress ratios include relative flow peaking factors. The stress ratios are calculated assuming that the tubes were dented immediately af ter initial startup.
Table 9-3 contains a summary of fatigue usage factors for tubes that have stress ratios near or greater than 1.00 (calculated using the more limiting Premod conditions and assuming the tubes became dented since the first cycle).
As can be observed in the table, all tubes currently have fatigue usage factors less than 1.00. Future usage factors have been determined for operation under current operating conditions and for conditions where T hot reduction is implemented.
Results are presented for both 40 years of total operation and for 10 more years of operation.
These results indicate that, for a total of 40 years of operation, two tubes are at potential risk if T
reduction is implemented. These two tubes (SG:A R9C60 and SG:B R9C35) het currently have usage factors equal to 0.49 but will have projected fatigue usage f actors greater than 1.00 af ter 40 years of total operation. Usage factors calculated after 10 more years of operation (with T implemented) hot have been determined to be 0.84.
2.7 Overall Conclusion The results of the fatigue evaluation indicate that currently no tubes in the North Anna Unit 2 steam generators require preventative action to preclude a 9213M;1E-o428B9-19
1 North Anna Unit 1 R9C51 type tube rupture and that any tubes currently plugged with' sentinel plugs, to detect such a rupture, can be returned to service.
Nowever, two tubes previously identified, SG:A R9C60 and SG:B R9035, will require preventive action in the future, to preclude such a rupture, after at j
least 10 more years of service, Note that in the event of a future uprating
]
or increase in general plugging level the potential for tube fatigue would need to be re-evaluated.
9213 M.1 E-042B B 9-20
1 l
I
3.0 BACKGROUND
C On July 15, 1987, a steam generater tube rupture occurred at the North Anna 1
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 circumferentici in orientation with a 360 degree extent, j
3.1 North Anna Unit 1 Tube Rupture Event e
i 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 determined to be fluid elastic, based on the magnitude of the alternating stress, t
A significant contributor to the occurrence of excessive vibration is the reduction in damping at the tube-to-tube support plate interface caused by the denting.
Also, the absence of antivibration bar (AVB) support has been concluded to be required for requisite vibration to occur.
The presence of an AVB support restricts tube motion and thus precludes the dr'fection amplitude required for fatigue.
Inspection data shows that an AVB is not present for
.the Row 9 Column 51 tube but that the actual AVB installation depth exceeded the minimum requirements in all cases with data for AVBs at many other Row 9
' tubes. Also contributing significantly to the level of vibration, and thus leading, is the local flow field associated with the detailed geometry of the steam generator, i.e., AVB insertion depths.
In addition, the fatigue properties of the tube reflect the lower range of properties expected for an 9213M,1E-042689-21
~
AVT environment.
In summary, the prerequisite conditions derived from the evaluations were concluded to be:
i Fatigue Requirements Prerequisite Conditions Alternating stress Tube vibration
- Dented support
- Flow excitation
- Absence of AVB y
I Mean stress Denting in addition to applied stress Material' fatigue properties AVT environment
- Lower range of properties 5.2 Tube Examination Results f
e Fatigue was found to have initiated on the cold leg outside surface of Tube R9C51 immediately above the top tube support plate. No indications of significant accomptnying intergranular corrosion was observed on the fracture face or on the immediately adjacent OD surfaces. Multiple fatigue initiation sites were found with major sites located at 110*,120',135' and 150', Figure 3-1.
The plane of-the U-bend is located at 45' with the orientation system used, er approximately 90' from the geometric center of the initiation zone at Section D-D.
}iigS cycle fatigue striation spacings approached 1 micro-inch near the origin sites, Figure 3-2. The early crack front is believed to have broken through-wall from approximately 100* to 140',
From this point on, crack growth is believed (as determined by striation spacing,
+
striation direction, and later; observations of parabolic dimples, followed by equiaxed dimples) to have accelerated and to have changed direction with the resulting crack front running perpendicular to the circumferential direction.
- $213M:1E-042BB9-22
'l
3.3 Mechanism Assessment To address a fatigue mechanism and to identify the cause of the loading, any loading condition that would cause cyclic stress or steady mean stress had to be considered. The analysis of Normal Upset and Test conditions indicated a relatively low total number of cycles involved and a corresponding low fatigue usage, even when accounting for the dented tube condition at the plate.
This analysis also showed an axial tensile stress contribution at the tube OD a short distance above the plate from operating pressure and temperature, thus providing a contribution to mean stress. Combining these effects with denting deflection on the tube demonstrated a high mean stress at the failure location. Vibration analysis for the tube develcped the characteristics of first mode, cantilever response of the dented tube to flow induced vibration for the uncracked tube and for the tube with an increasing critek angle, beginning at 90' to the plane of the tube and progressing around on both sides to complete separation of the tube.
Crack propagation analysis matched cyclic deformation with the stress intensities and striation spacings indicated by the fracture inspection and analysis.
Leakage data and crack opening analysis provided the relationship between leak rate and circumferential crack length.
Leakage versus time was then predicted from the crack growth analysis and the leakage analysis with initial stress amplitudes of 5, 7, and 9 ksi.
The comparison to the best estimate of plant leakage (performed after the event) showed good agreement, Figure 3-3.
Based on these results, it followed that the predominant loading mechanism responsible is a flow-induced, tube vibration loading mechanism.
It was shown that of the two possible flow-induced vibration mechanisms, turbulence and fluidelastic instability, that fluideiastic instability was the most probable l
cause. Due to the range of expected initiation stress amplitudes (4 to 10 ksi), the fluidelastic instability would be limited in displacement to a range of approximately (
Ja c.
This is less than the distance between tubes at the apex, (
Ja.c It was further I
9213M:1E 042689-23
confirmed that displacement prior to the rupture was limited since no indication of tube U-band (apex region) damage was evident in the eddy-current signals for adjacent tubes.
Given the likelihood of limited displacement, fluidelastic instability, a means of establishing the change in displacemant, and corresponding change in stress amplitude, was developed for a given reduction in stability ratio (SR). Since the rupture was a fatigue mechanism, the change in stress amplitude resulting from a reduction in stability ratio was converted to a fatigue usage benefit through the use of the fatigue curve developed.
Mean
^
- stress effects were included due to the presence of denting and applied loadings. The results 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.
f r
9213M:1 E-o42689-24
i I
i D.D t
2 M
r C
CC i
C 180*
Region of Herringbone r Pattern
/
- se -
70 -
/
N' B-8 7
g.g A
\\
/
s m
"o F.F l
Coarse Texture o
and Dimpled Rupture
( Indicates origins l
l Figure 3-1 Approximate Mapping of Fracture Surface of Tube R9C51, S/6 "C" Cold Leg, North Anna Unit I l
5 = 1.5/1.8 v in.
r Heavy Oaine Attack 3
5 s 21 y in, f
5 = 1.0/1.85 y in.
ggg.
3C O
b Parabolic
~ 90*
270' l
1-G Dieeles i
/
and V
Internal f
Necking 38 Sadly 5 = 2.8/4.0 v in.
l
/
11 Peer.ed VM M
Nearly Ecui4xed 5 a 6.1/6.9 y in.
Otmples Note: Arrows Indicata Direction of Fracture Propagation Figure 3-2 Schematic Representation of Features Observed During TEM Fractograhic Examination of Fracture Surface of Tube R9C51, S/G "C" Cold Leg, North Anna Unit 1
~ - - ' ' ' ^ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ '
~~T) i i
i i
i l
1 Calculated and obsened leak rates versus time.
Obsened values based on gaseous species condenser atr ejector seeee' SISMA A = S KS!
- SISMA A = 7 KS!
5
"""""*SISMA A = 9 KS!
I l 3
t O
Ar-dt t
9 0
Xe-135 a
ge-ey d
i 1
l w
O s
/
g r
t O
O O
e ese sees name sees sees snee asse mee mee sees som sene TIME. MINUTES Figure 3-3 Calculated and Observed Leak Rates Versus Time
WESTINGHOUSE PROPRIETARY CLASS 2 1
4.0 CRITERIA FOR FATIGUE ASSESSMENT The evaluation method and acceptance criteria are based on a relative e
corparison with the Row 9 Column 51 tube of North Anna Unit 1 Steam Generator C.
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 R9C51.
For these reasons, the North Anna Unit 2 tubing-evaluation was done on a relative basis to Row 9 Column 51 and a 10% reduction in stability ratio criteria was established to demonstrate that tubes left in
(
service would be expected to have sufficiently low vibration stress to preclude future fatigue rupture events.
i To accomplish the necessary relative assessment of North Anna Unit 2 tubing to the North Anna Unit 1 Row 9 Column 51 several criteria are utilized.
- First, stability ratios are calculated based on flow fields predicted by 3-D thermal hydraulic models and raticed to the stability ratio for Row 9 Column 51.
^
These ratios of stability ratio (called relative stability ratios) for each l
potentially unsupported U-bend in the North Anna Unit 2 steam generators should be equivalent to 1 0.9 of (the pre-modification) North Anna Unit 1 R9C31 (meeting the 10% reduction in stability ratio criteria). This provides l
the first level of screening of susceptible tubes incorporating all tube l
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. To account for these differences, flow peaking factors can be incorporated in the relative stability ratios and the relative stress ratios.
l The next step is to obtain stress ratios, the ratio of stress in the North Anna 2 tube of interest to the stress in North Anna 1 Row 9 Column 51 and
(
after incerporating the requirement that the relative stability ratio to Row 9 Column 51 (R9C51) for the tube of interest is equivalent to 1 0.9, require the stress ratio to be 5 0.
The stress ratio incorporates the tube 1
geometry differences with R9C51 in relation to the stress calculation and also 9213M:1E-042689-28
incorporates the ratio of flow peaking factor for the tube of interest to the
' flow peaking factor for _R9C51 (flow peaking factor is defined in Section 4.2).
This should provide that all tubes meeting this criteria have stress amplitudes equivalent to 1 4.0 ksi.
Finally, the cumulative fatigue usage for plant operation _to date and for continued operation with the planned operating parameters is evaluated. A fatigue usaga of 5 1.0 may not be satisfied by meeting the stress ratio i
criteria using the reference operating cycle evaluation since the reference cycle does not necesserily represent the exact duty cycle to date.
Therefore, the time history of operation is evaluated on a normalized basis and used together with the stress ratio to obtain a stress amplitude history.
This permits the calculation of current and future fatigue usage for comparison to 2.0.
4.1 Stability Ratio Reduction Criteria For fluidelastic evaluation, stability ratios are determined for specific configurations of a tube.
These stability ratios represent a measure of the potential for flow-induced tube vibration during service.
Values greater than unity (1.0) indicate instability (see Section 5.1).
Motions developed by a tube in the fluidelastically unstable mode are quite large in comparison to the other known mechanisms. The maximum modal displacement (at the apex of the tube) is linearly related to the bending stress in the tube just above the top tube support plate.
This relationship j
applies to any vibration in that mode. Thus, it is possible for an unstable, fixed boundary condition tube to deflect an amount in the U-bend which will produce fatigue inducing stresses.
_ 9213M:1 E-0426B9-29
m L
The major features of the fluidelastic mechanism are illustrated in Figure 4-1.
This~ figure shows the displacement response (LOG D) of a tube as a function of stability ratio (LOG SR). A straight-line plot displayed on 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 l
x is raised), and y is the dependent variable.
Taking logs of both sides of this equation leads to the slope-intercept form of a straight-line equation in log form, log y = c + n log x, where c = log A and represents the intercept and n is the slope.
In our case the independent variable x is the stability ratio SR, and the dependent variable y is tube (fluidelastic instability induced) displacement response D, and the slope n is renamed s.
From experimental results, it is known that the turbulence response curve (on log-log coordinates) has a slope of approximately [ la,b,c.
Test results also show that the slope for the fluidelastic response depends somewhat on the instability displacement (response amplitude).
It has been shown by tests that a slope of (
Ja,b,c is a range of values corresponding to displacement amplitudes in the range of (
Ja.c,
whereas below [
']a,e are conservative values.
l l
l The reduction in response obtained from a stability ratio reduction can be l
expressed by the following equation:
a,e where D1 and SR) are the known values at the point corresponding to point 1 of Figure 4-1 and D2 and SR2 are values corresponding to any point lower l
on this curve. Therefore, this equation can be used to determine the reduction in displacement response for any given reduction in stability ratio.
l This equation shows that there is benefit derived from even a very small percentage change in the stability ratio.
It is this reduction in displacement for a quite small reduction in stability ratio that formed the basis for demonstrating that a 10% reduction in stability ratio would be sufficient.to prevent Row 9 Column 51 from rupturing by fatigue.
l 9213M 1E-042689-30
p ui L
i The fatigue curve developed for the North Anna Unit 1 tube at R9C51 is from i
.(
l l
I i
Ja,c. Thus,
- a,c where, a,'is the equivaler$t stress amplitude to a, that accounts for a maximum stress of a, the yield strength. The 3 sigma curve with y
mean stress effects is shown in Figure 4-2 and is compared to the ASME Code Design Fatigue Curve for Inconel 600 with the maximum effect of mean stress.
The curve utilized in this evaluation is clearly well below the code curve j
reflecting the effect of an AVT environment on fatigue and [
Ja,c for accounting for mean stress that applies to mater:als in a corrosive environment.
Two other mean stress models were investigated for the appropriateness of their use in providing a reasonable agreement with the expected range of initiating.
j stress amplitudes.,These were the [
]a,c j
shown in Figure 4-3.
With a l J
Ja,c, the [
1 I
L
]&,c, l
0295M:49/022489-31
7-1 The assessment of the benefit of a reduction in stability ratio begins with the
{
relationship between stability ratio and deflection. For a specific tube geometry, the displacement change is directly proportional to change in stress l
so that stress has the same relationship with stability ratio, a,c
}.
i 8
The slope in this equation can range from (
Ja c on a log scale depending on the amplitude of displacement. Knowing the stress resulting from a change in stability ratio from SR1 to SR, the cycles to failure at the 2
stress amplitude was obtained from the fatigue curve. A fatigue usage per year was then determined assuming continuous cycling at the natural frequency of the tube. The initial stress was determined to be in the range of 4.0 to 10.0 ksi by the fractography analysis.
It was further developed that the maximum initiating stress amplitude was not more than 9.5 ksi. This was based on [
la.c. The corresponding stress level is 5.6 ksi.
The maximum stress, 9.5 ksi, would be reduced to [
la,c with a 10%
reduction in stability ratio and would have a future fatigue usage of
(
]a,c per year at 75% availability, Figure 4-4.
The minimum stress, 5.6 ksi, would be reduced to (
Ja c ksi with a 5% reduction in stability ratio and would have future fatigue usage of [
]a,c per year, Figure 4-5.
In addition, if a tube were already tracked, the crack could be as large as [
]a,c inch in length and thru wall and would not propagate if the stress amplitudes are reduced to 5 4.0 ksi.
0295M:49/022489-32
.t Subsequent to the return to power evaluation for North Anna Unit 1, the time
'I history of operation was evaluated on a normalized basis to the last cycle, i
(
confirming the conservatism of 9.5 ksi. [
la,c, cumulative fatigue usage may then be computed to get a magnitude of alternating stress for the last cycle that results in a cumulative usage of 1.0 for the nine year duty cycle.
The result
{;
of the iterative analysis is that the probable stress associated with this fatigue curve during the last cycle of operation was approximately L
{
Ja,e for R9C51, North Anna Unit 1, Steam Generator C, and that the major portion of the fatigue usage came in the second, third and fourth cycles. The first cycle was conservatively omitted, since denting is assumed, for purposes of this analysis, to have occurred during that first cycle.
i Based on tnis evaluation, the tube fatigue probably occurred over most of the l
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 fatigue usage of (
Ja.c, Other combinations of alternating stress and mean stress were evaluated with
-3 sigma and -2 sigma fatigue curves to demonstrate the conservatism of the 10% reduction in stability ratio.
Table 4-1 presents the results of the cases analyzed clearly demonstrating that the 10% reduction in stability ratio combined with a -3 sigma fatigue curve and with maximum mean stress effects is conservative.
Any higher fatigue curve whether through mean stress, mean I
stress model, or probability, results in greater benefit for the same reduction in stability ratio.
Further, for any of these higher curves, a smaller reduction in stability ratio than 10% would result in the same 9213M;1 E-042689-33 l
benefit.
In addition, there is a large benefit in terms of fatigue usage for relatively small changes in the fatigue curve.
4.2 Local Flow Peaking Considerations Local flow peaking is a factor on stability ratio that incorporates the effect of local flow velocity, density and void fraction due to non-uniform AVB insertion depths. The flow peaking factor is applied directly to the stability ratio obtained from thermal-hydraulic analysis that does not account for these local geometry effects. Being a direct factor on stability ratio, a small percentage increase cari result in a significant change in the prediction of tube response.
Since the evaluation of North Anna Unit 2 tubing is relative to North Anna Unit 1 R9C51, the flow peaking factors are also applied as relative ratios, i.e., a ratio of North Anna Unit 2 tubing to R9C51.
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 the North Anna Unit 1 R9C51 AVB pattern compared to critical velocity for a uniform AVB pattern. As explained in Section 8.0, the minimum value of i
[
Ja,b,c is appropriate for R9C51 of North Anna 1.
The peaking factor for a tube in North Anna Unit I tubing is therefore divided by [
Ja,b,c and the resulting relative flow peaking is multiplied the relative stability l
ratio based on North Anna Unit 1 ATHOS results, appropriately scaled for North l
Anna Unit 2.
If the peaking factor is 1.0, the relative flow peaking is
[
3a,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 ksi 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 3.1 to 1.4 and the flow peaking factor is in the range of 1.8 to 2.2.
This j
range of flow peaking agrees with the range of flow peaking f actors measured 9213M:1E-042789-34
i f n thiair-tests:and is considered to. be the best estimate of the range of the -
i R9C51 flow peaking factor.
The range of stability ratios,1.1 to 1.4, is based on a value of 0.63 i
'obtained with ATH0S results without flow peaking and with nominal damping that ll is a function of' modal effective void fraction (MEVF). MEVF is calculated using the formula:
~\\ a,c l
l i
The ncminal damping reflects the nominal reduction in damping that occurs with denting at the tube support plate. Therefore, a minimum damping scenario that is independent of void fraction is not considered to be credible and is not i
addressed in the evaluation that follows.
l 4.3 Stress Ratio. Considerations In Section 4.1, a 10*4 reduction in stability ratio was established to reduce the stress amplitude on the Row 9 Column 51 tube of North Anna Unit 1 to a level.that would not have ruptured, 4.0 ksi.
To apply this same criteria to l
another tube in the same or another steam generator, the differences in [
3a,c,
a,C lW'-
i 9213M.1 E-042689-35
p i
a,e t'
m.
The quantities with subscript NA refer to R9C51, and the quantities without subscripts refer to the tube being evaluated.
Using the displacement versus stability ratio relationship defined in Section 4.1,
-i-i a,c
,s i.
l-:
l 0295M:49/022489-36
g
~
I
~
By establishing.their equivalent effect on the stress amplitude that-produced
~
the tube rupture at R9C51 several other effects may be accounted for.
These include a lower mean stress (such as for non-danted tubes) and different r
frecuency tubt s from the [ Ja,c.e hertz frecuency of R9C51, North Anna 1.
In the case 'of lower mean stress, the stress amplitude that would have caused
- the' failure of R9C51, North Anna 1, would have been higher.
[
'l
.)*54
+
A lower or higher frequency tube would not reach a usage of 1.0 in the same length of time as the RDC51 tube due to the diffsrent frequency of cycling.
The usage accumulated is proportional to the f requency and, therefore, the allowable number of cycles to reach a usage of 1.0 is inversely proportional to frequency. The equivalent number of cycles to give the usage of 1.0 for a different frequency tube [is used to obtain a stress amplitude different from 9.5 ksi that gives the equivalent result.
The ratio of these stresses becemes
-a factor times the above stress ratio expression to account for a frequency effect)a c,
' Knowing the magnitude of the stress ratio allows 1) the determination of tubes
.that' do not meet a value of 5 1,:and 2) the calculation of maximum stress in the acceptable tubes, a,c Having-this maximum stress permits the evaluation of the maximum fatigue usage
.for North Anna Unit 2 tubes based on the time history expressed by normalized stability' ratios for the duty cycle (see Section 7.4).
l l-l' 9213M:1E-042889-37
l)i.
i i
Table 4-1 Fatigue Usage per Year Resulting b
From Stability Ratio Reduction
(*
SR, %
STRESS-FATIGUE MEAN STRESS USAGE
[
REDUCTION BASIS (l)
CURVE (2)
MODEL PER YEAR l'
g
- a,c 5.
9 yrs to fail [
]a,c 5.
.9. yrs to i
fnil [
ja,c 5.
9 yrs to fail [
]a,c 10.
max.strasg) amplitudel
[
ja,c max,stregg')
10.
amplitude \\
[
]a,c 10.
max. stres )
amplitude
[
]a,c
- 10. -
max,stres;)
amplitude b
[
]a,c 10.
max. stress based on duty cycle (5)
[
ja,c (1) This gives the basis for selection of the initiating stress amplitude and its value in ksi.
(2) 'S,is the maximum stress applied with Sm"Smean + S -
a (3)
[
]a,c, (4) Cycles to failure implied by this combination of stress and fatigue properties is notably less than implied by the operating history.
Consequently this combination is a conservative, bounding estimate.
(5) Cycles to failure implied by the operating history requires [
Ja,c fatigue curve at the maximum stress of [
]a,c,
l.
)
{
9:
?
. a,b,c I
i 4
lb
.1 l
1 1,
l
-1 n
1 l!-
i h
Figure 4-1 Vibration Displacement vs. Stability Ratio
e i
g P
e-a,e
.l d
..o i
w 1
l4;.
Figure 4-2 Fatigue Strength of Inconel 600 in AVT Water at 600*F 1
n.
I t
.c
I J
f I
c.-
1 j
a,c 14'.
i i
l 4
h I
l t'
\\
I l-4 l-1 i
l i.L I
r Figure 4-3 Fatigue curve for Inconel 600 in AVT Water p
Comparison of Mean Stress Correction Models i-1 I
i.'
+
c.:
1 a,c t
I^
7 4
l:
t.
Figure 4-4 Modified Fatigue with 10% Reductlon in stability Ratio for Maximum Stress Condition l
1 -
+
-]
l l
d.
a,c i
i i
t l
i
'1
)
l l
l (I
n 1
i l
l-1 l
lr.-
l l
Figure 4-5 Modified Fatigue with 5% Reduction in Stability Ratio for Minimum Stress Condition l
i l
1 I:.,-
s 5.0 SUPPORTING TEST DATA' 9.L This section provides a mathematical description of the fluidelastic
. mechanism, which was determined to be the most likely causative mechanism for the North Anna ~ tube rupture, as discussed in Section 3.3, to highlight the physical conditions and corresponding parameters directly related to the event and associated preventative measures. The basis for establishing the appropriate values and implications associated with these parameters are provided. Where appropriate, test results are presented.
'5.1 Stability Ratio Parameters Fluideiastic ' 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 L
to' instability during service. Fluidelastic stability evaluations are performed with a computer program which provides for the generation of a I
finite element model of the tube and tube support system.
The finite element 1
model provides the vehicle to define the mass and stiffness matrices for the y
tube'and its support system.
This information is used to determine the modal
. frequencies-(eigenvalues) and mode shapes (eigenvectors) for the linearly l
supported tube being considered.
i
)
The methodology is comprised of the evaluation of the following equations:
Fluidelastic stability ratio = SR = Ven/Uc for mode n, where Uc (critical velocity) and Uen (effective velocity) are determined by:
.D )) W gy) 2 U *0I 0 II*o 6 ) / (P c
n n
o and; N
3[y IPj/Po) U k j jn *j V,
(2)
.l
= ----------------------
N 3[yI*j/*o) 'jn Zj
' 9213M.1 E-042789-44
i
- where, i
.D tube outside diameter, inches
=
effective velocity for mode n, inches /see 1
U
=
en N
. number of nodal points of the finite element model
=
number of degrees of freedom in the out-of plane direction
=
m), U, p3 =
mass per unit length, crossflow velocity and fluid 3
l' density at node j, respectively a'
reference density and a reference mass per unit p,m length, respectively (any representative values) logarithmic decrement (damping) 6
=
n l
.#jn n rmalized displacement at node j in the nth mode of vibration l.
=
average of distances between node j to j-1, and j to j+1
~
2-
=
3 an experimentally correlated stability constant 6
=
Substitution of Equations [1] and (2) into the expression which defines l
stability ratio, and' cancellation of like terms, leads to an expression in fundamental terms (without the arbitrary reference mass and density L
-parameters). From this resulting expression, it is seen that the stability ratio is directly related to the flow field in terms of the secondary fluid 1
velocity tiraes square-root-density distribution (over the tube mode shape),
l and inversely related to the square root of the mass distribution, square root l
of modal damping, tube modal frequency, and the stability constant (beta).
L.
i l
The uncertainty in each of these parameters is addressed in a conceptual manner in Figure 5-1. The remainder of this section (Section 5.0) provides a e
discussion, and, where appropriate, the experimental bases to quantitatively establish the uncertainty associated with each of these parameters.
In
' 9213M:1E-042889-45
- addition, Section 5.3 provides the experimental. basis to demonstrate that tubes with'(-
g.
]a,c.
This implies that those tubes [i
-]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 dented-(fixed boundary condition) tubes.
Therefore,uncertaint3jlevels introduced by the frequency parameter are expected to be insignnficant (see also " Average Flow Field" subsection below).
Instability Constant (Beta)
The beta (stability constant) values used for stability ratio and critical velocity evaluations (see above equations) are based on an axtensive data base comprised of both Westinghouse and other experimental resuits.
In addition, previous field experiences are considered. Values have been measured for full length U-bend tubes in prototypical steam / water environments.
In addition, measurements in U-bend air models have been made with both no AVB and variable AVB supports (Figure 5-3).
To help establish the uncertainties associated with ATH0S flow velocity and density distribution predictions on stability analyses, the Model Boiler (MB-3) tests performed at Mitsubishi Heavy Industries (MHI) in Japan were modeled using ATHOS. A beta value consistent with the ATH0S predicted flow conditions and the MB-3 measured critical velocity was determined. These analyses supported a beta value of [
la,b,c,
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 Ltubes without AVBs was a value of [
la,b,c.
This value is used for the o
beta parameter in all stability ratio evaluations addressed in this Report (see also " Average Flow Field" subsection below).
L Mass Distribution lt The mass distribution parameter is based on known information on the tube and l
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 [
3a,c, Tube DamDinQ Test data are available to define tube damping for clamped (fixed) tube supports, appropriate to dented tube conditions, in steam / water flow conditions.
Prototypic U-bend testing has been performed under conditions leading to pinned supports. The data of Axisa in Figure 5-4 provides the-principal data for clamped tube conditions in steam / water. This data was obtained for cross flow over straight tubes.
Uncertainties are not defined for the data from these tests. Detailed tube damping data used in support of the stability ratio evaluations addressed in this report are provided in Section 5.2, below.
Flow Field - Velocity Times Sauare-Root-Density Distribution Average and U-bend-local flow field uncertainties are addressed independently in the following.
-C.
i
Averaae Flow Field Uncertainties in the average flow field parameters, obtained from ATH0S
. analyses, coupled with stability constant and frequency, are essentially the same for units with dented or non-dented top support plates.
If the errors associated with these uncertainties were large, similar instabilities would be expected in the non-dented units with resulting wear at either the top support plate or inner row AVBs.
Significant tube wear has not been observed in inner row tubes in operating steam generators without denting. Thus, an uncertainty estimate of about [
la,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 North Anna Unit 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.
1 U-Bend t.ocal Flow Field l
Non-uniform AVB insertion depths have been shown to have effects on stability l
ratios.
Flow peaking, brought about by the " channeling" effects of non-uniform AVBs, leads to a local perturbation in the velocity times square-root-density parameter at the apex of the tube where it will have the largest effect
'(because the apex is where the largest vibration displacements occur).
l Detailed local flow field data used in support of the stability ratio evaluations addressed in this report are provided in Section 5.2, below.
Overall Uncertainties Assessment Based on the above discussions, and the data provided in the following sections, it is concluded that local flow peaking is likely to have contributed significantly to the instability and associated increased vibration amplitude for the failed North Anna tube.
Ratios of stresses and stability ratios relative to the North Anna tube, R9C51, are utilized in this report to minimize uncertainties in the evaluations associated with instability constants, local flow field effects and tube damping.
5'.2 ' Tube. Damping Data
(
The damping ratio depends on several aspects of the physical system. Two primary determinants of damping are the support conditions and the flow field.
It has been shown that tube support conditions (pinned vs clamped) affect the damping ratio significantly.
Further, it is affected by the flow conditions, i.e., single-phase or two-phase flow. These effects are discussed below in more detail.
Reference (5-1) indicates that the damping ratio in two phase flow is a sum of contributions from structural, viscous, flow-dependent, and two-phase damping.
o.
The structural-damping will be equal to the measured damping in air. However, in two-phase flow, the damping ratio increases significantly and is dependent on the void fraction or quality.
It can be shown that the damping contribution from viscous effects are' very small.
Damping ratios for tubes in air and in air-water flows have been measured and reported by various authors.
However, the resu'lts from air-water flow are poor representations of the' actual conditions in a steam generator (steam-water flow at high pressure). Therefore, where available, results from prototypic steam-water flow conditions should be used.
Fortunately, within the past few years test data on tube vibration under steam water flow has been developed for both pinned and' clamped tube support conditions.
Two sources of data are particularly noteworthy and are used here. The first is a large body of recent, as yet unpublished data from high pressure steam-water tests conducted by Mitsubishi Heavy Industries (MHI). These data were gathered under pinned tube support conditions. The second is comprised of the results from tests sponsored by the Electric Power Research Institute (EPRI) and reported in References (5-2) and (5-3).
The damping ratio results from the above tests are plotted in Figure 5-4 as a function of void fraction.
It is important to note that the void fraction is determined on the basis of [
]a,c
'(Reference (5-4)).
The upper curve in the figure is for pinned support conditions. This curve represents a fit to a large number of data points not shown in the figure. The points on the curve are only plotting aids, rather than specific test results.
The lower curve pertains to the clamped support condition, obtained from Reference (5-3). Void fraction has been recalculated on the. basis of slip flow.
It may be noted that there is a significant difference in the damping iatios 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, I
the clamped tube support curve is used in the current evaluation to include the-effect of denting at the top tube support plate.
. The Reference 5-3 data as reported show a damping value of.5% at 100% void fraction. The 100% void fraction condition has no two phase damping and is considered to be affected principally by mechanical or structural damping.
Westinghouse tests of clamped tube vibration in air has shown that the 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
. (5-3) data with all damping values reduced by [
]a,c, i
l F
i L
l
5.3.- Tube' Vibration Amplitudes With Single-Sided AVB Support o,
A series of wind tunnel tests were conducted to investigate the effects of
' tube /AVB eccentricity on the vibration amplitudes caused by fluidelastic vibration.
[
la,c, Prior test results obtained during the past year using this apparatus have demonstrated that the fluidelastic vibration characteristics observed in the tests performed'with the i
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 l
. view of the apparatus.
[
1 1
1 1
I l-L Ja,C, l*
- As shown in-Figure 5-7, the tube vibration amplitude below a critical velocity j
is caused by [
6 b
Ja,c, Figure 5-7 shows the manner in which the zero-to-peak vibration amplitude, t
expressed as a ratio normalized to [
]a,c, varies when one gap remains at [
]a,c.
For increasing
-velocities, up to that corresponding to a stability ratio of [
]a,c.
Figure 5 8 shows typical l
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 [
']a,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 AVB Insertion Depths j
This section summarizes a series of wind tunnel tests that were conducted to investigate the effects of variations in AVB configurations on the initiation of fluidelastic vibration.
Each configuration is defined as a specific set of insertion depths for the individual AVBs in the vicinity of an unsupported U-bend tube.
.The tests were conducted in the wind tunnel using a modified version of the cantilever tube' apparatus described in Section 5.3.
Figure 5-9 shows the conceptual design of the apparatus. The straight cantilever tube,
s' t
b,
[
e T.
l W
W i,4,
]a,C,
.i I
]a,c.
Figure 5-11 shows the AV8s, when the side panel of the' test section is removed. Also shown is the top flow screen which 'is [
Ja,c. The AVB 3
configurations tested are shown in Figure 5-12.
Configuration la corresponds:
to tube R9C51,'the failed tube at North Anna. Configuration 2a corresponds to one of the cases in which the AVBs are inserted to a uniform depth and no local velocity: peaking effects are expected.
As shown in Figure 5-9, [
Ja,c, All the tubes except the instrumented tubed (corresponding to Row 10) are y
-[
]a,c.
As discussed in Section 5.3, prior testing indicates that this situation provides a valid model. The instrumented tube [
] ate as shown in Figure 5.10.
Its [
]a,c direction vibrational motion is measured using a non-contacting L
transducer.
[.
ja,c. The instrumented tube corresponds to a Row 10 tube as shown in Figure 5-9.
However, depending on the particular AVB configuration, it can reasonably represent a tube in Rows 8 through 11. The AVB profile in the straight tube model is the average of Rows 8 and 11. The difference in profile is quite small for these bounding rows.
[-
Ja,c using a hot-film anemometer located as shown in Figure 5-9.
]
Figure 5-13 shows the rms vibration amplitude, as determined from PSD (power spectral density) measurements made using an FFT spectrum analyzer, versus flow velocity for Configuration la (which corresponds to tube R9C51 in North Anna).
Data for three repeat tests are shown and-the critical velocity is identified.
The typical rapid increase in vibration amplitude when the critical velocity for fluidelastic vibration is exceeded is evident.
e 9
3 i
The main conclusions from the tests are-l
~
4 l '. Tube vibration below the critical velocity is relatively sma11, typical of turbulence-induced vibration, and-increases rapidly when the critical velocity for the initiation of fluidelastic vibration is exceeded..
2.
Configuration Ib (a preliminary version of R9C51 in North Anna) has the lowest critical velocity of all the configurations tested.
3.
Configuration Ib is repeatable and the configuration was rerun periodically to verify the consistency of the test apparatus.
The initial test results obtained in support of the North Anna Unit 1 evaluation are summarized in Table 5-2.
The test data is presented as a i
velocity peaking ratio; a ratio of critical velecity for North Anna tube R9C51 configuration la, to that for each North Anna Unit 1 AVB configuration evaluated.
5.5 References j
a,c
)
1 3
I a
~
- )
5
)
4 Table 5 -
Wind Tunnel Tests on Cantilever: Tube Model OBJECTIVE:.
Investigate the effects of tube /AVB fitup on flow-induced tube.
vibration.
+
APPARATUS:.
Array of cantilevered tubes with end supports [
i
]a,c, L
MEASUREMENTS: Tube vibration amplitude and tube /AVB impact forces or preload forces.
RESULTS:
.a,b,c l
i l
l l'
Y 9-0295M:49/022489-56
Table 5-2 Fluidelastic Instability Velocity Peaking Ratios for Columnwise Variation in AVB Insertion Depths (North Anna 2)
Type of Insertion Peaking Ratio Configuration U,/U 1
n a,c i
l Note: U is instability velocity at inlet for type n of AVB insertion n
3
_ configuration, b.
9213M:1 E-042689-57
1 i
h
~
a,e
?
l L
1
(
l l
1~
- s
[+
I 1
s h
m.
r Figure 5-1 Fluidelastic Instability Uncertainty Assessment s
a a
'U-Bend Test Data 1)
MB-3 Tests
)
~ * ~
$ values of [
]a,b:c 2)
MB-2 Tests
- of-[
.]a,b,c l
3)
Air Model Tests
- of [
Ja,b,c without AVBs Tendency for $ to increase in range of [
]a,b,c with inactive AVBs (gaps at AVBs)
Tendency for A 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 S values 5)
ATH0S velocity data with S of (
]a,b,c and known damping-should not significantly underestimate instability for regions of uniform U-bend flow Figure 5-2 Instability Constant - B 1
i a,b,c i
s i,
t t
,77 j.-
l I
L ii
?
lo-Figure 5-3 Instability Constants, 4, obtained for Curved Tubes from Wind Tunnel Tests on the 0.214 Scale U-Bend Model
-t
?-
P t
a,b,c' 4
t
?
4 h
6 Figure 5-4 Damping vs. Slip Void Fraction
s i
t t
q a,b,c t
p i
l' 1
i I
Figure 5 5 Overall View of Cantilever Tube Wind Tunnel Model 253681 i:
-i I
i i.
l i
i i
t, -
~
a,b,c 1
j J
1 r
a f
i f
4 k
i e
m Figure 5-6 Top View of the Cantilever Tube Wind Tunnel Model o
25368 2
.i i
i a,b.e l
i
),..
r i
L i
f i
i 6
i t
0-O Figure 5 7 Fluidelastic Vibration Amplitude with Non-Uniform Gaps i
f i
I f
a.b.c
.s-i k'
l i
l
,t I
l I
-i t
L
+
~
Figure 5 8 Typical Vibration Amplitude and Tube /AV8 Impact Force Signals for Fluideiastic Vibration with Unequal Tube /AYS Gaps m-
i I
f i
i.
)
4.b it r
1 l
l 1
j i
t i
'O Figure 5-9 Conceptual Design of the Apparatus-for Determining the f
Effects of Fluidelastic Instability of Coltunwise L
Variations in AV8 Insertion Depths i.
1 s
i L
I 1
?
r t
l#
puse a,b,c !
t t
1 r
I i
i i
1 1
?
i I
1 1
6 i
1 1
1 l
i 1
l-0 6
Figure 510 Overall View of Wind Tunnel Test Apparatus l
l 25364 3
r t
i.-
i
- a,b,c
.o i
1 E
l
?
i I
i 4
l':
l8; l
l O
Figure 511 Side View of Wind Tunnel Apparatus with Cover Plates Removed to Show Simulated AVBs for Field Modified Units f
and Top Flow Screen 25368 5
3 l
l
\\
k I
i l.
. TYPE OF AVB TYPE OF AVB TYPE OF AVB i
INSERTION INSERTION INSERTION i
E-a,e a,e a,e i
1a 2a 5a l
l.
1b 4a Sb 1
l
?
t 1e 4b i
Sc i
x h
- 10 4f Ga i
v 1w 4r 8d i
i 12
'4s 8b L
I O
I Figure 5-12 AVB Configurations Tested - North Anna Unit 2 i
i-L 9213M.1 E-042789-69 7
m,...
,,.,.y..--.
w
1 e..x,
,: n, i
P a,b,c -
i
{. t ;
i p
5 j
- v. '-
_'}
r
!a.-
- d j
.a t
?
2 i-1 j
L,
)
fv.
4 n;
q a
c (i,
i'i.
M'
[
d ll a
i 5
g j.
4.
g t
63 Figure 5-13 Typical Variation of kMS Vibration Amplitude with Flow Velocity for Configuration la in Figure 512
' i..
e-
n
' 6.0 EDDY CURRENT DATA AND AVB POSITIONS I-6.1 AVB Assembly Design L
[
1 l
la,c.e Upper AVBs which are inserted beyond the design l~
depth occasionally show on the EC traces for the Row 12 tubes. Since the purpose of this analysis is to evaluate potentially unsuoperted tubec at or near the point of maxitium AVB insertior, enly the ditmozions atd EC data pertaining to the lower AVBs is used.
0.2 Eddy Current'0ata for AVB Positions The AVB insertion depths were determined on the basis of the interpretation of the eddy current data.
To locate the AVBs, the eddy current data traces from the August-September %B7 inspection were searched ior the characteristic peaks seen in tF.e signals wMen indicate the intersection of an AVB (or a tube support plate) with the tube. A typical sigr.a1 is shown in Figure b-1.
About 1,176 tubes from a!nong the three steam generators were examined for these signals. The U-bend signals for North Anna #2 were relatively free of signal disturbances usually attributed to copper deposited on the outsides of the
- tubes, 6.3 AVB Projection and Mapping e
Since ambiguity can occur in the interpretation of the ECT data due to the inability of ECT to differentiate on which side of a tube a " visible" AVB is o
located, other information was used to assist in establishing the location of 9213M:1 E-o42689-7e
the AVBs.
[
l ja.c j
~
The projection technique is useful in determining the AVB positions where j
suspected noisy or spurious ECT signals prevent direct observation of the l
AVBs, and where data are unavailable due to plugged tubes.
(
1 l
l 1,
ja.c l
For single AVB contacts, !
Jac Table 6-1 lists the "1-AVB" signals at locations near the r,rojected apex of the AVB which have been evaluated as being supported.
1 E
i l
The AVB positian maps are rhown in Figures E-4 to 6-6.
(
l l
(
l i
l
- 0
)"'C Table 6-2 lists tubes in Rows 7 through 10 which are evaluated in the analysis as being unsupported.
9213M:1E-042789-71
p t
The observation that (
I L
)***
6.4 Tube Denting at the Top Tube Support Plate f
As previously noted, tube denting was determined to be a prerequisite for the North Anna #1-type tube fatigue mechanism, and therefore of interect in the North Anna #2 tube evaluation. Tube support plate crevice corrosion products l
and tube de.nt sizes as small as 1-2 mih (0.001" to 0.002") are Ltectable by cddy current testing. Althrsgh the difference betwenn tube denting and the presence of crevice corrosion products may be significant in terms of the tube fctigue mechanism, it is conservative to consider a tube to be dented if either tube denting with deformation or top tube support plate crevice corrosion products were detected. All North Anna #2 tubes were assumed to be dented with deformation, which produces the naxtnum effect of mean, stresses for fatigue evaluations due to yielding of the tube at the top tube support plate.
Eddy current evaluations of North Anna #2 crevices from the August-September, 1987, inspection indicated that most of the tubes had crevice corrosion product buildups at the top tube support plate, but were not dented with deformation.
Of the 118 tubes which were sentinel plugged, 104 tubes were ovaluated as having " corrosion with magnetite" at the top tube support plate intersection, 2 showed " denting with deformation", 6 showed "no detectable denting" (or crevice corrosion product), and 6 were " unreadable".
The
" corrosion with magnetite", " denting with deformation", and " unreadable" conditions are considered as meeting the NRC Bulletin 88-02 definition of
" denting". As noted above, all tubes were considered dented in this analysis.
4 l
9213M:1 E-042789-72
I 6.5.'AVB Map Interpretations by Generator A' description of the AVB position mapping'in each of the North Anna #2 i, team generators is provided below.
t SG-A 1
1 The AVB map is given in Figure 6-2.
All Row 10 Row 11 and Row 12 tubes are supported. Twenty-five (25) Row 9 tubes, sixty-eight (68) Row 8 tubes, and ninety (90) Row 7 tubes are unsupported. Sentinel plugs were installed in forty-five (45) tubes in this. steam generator.
R9060 was [
l; Ja,e and was evaluated as potentially susceptible to fatigue, but only when Thot reduction i
is implement &d This tube was previously centinel plugged, based upon the October 1987 evaluation.
p 1
i
_Of the remaining tubes, R9C35 is the highest loaded tube in this steam generator. A conservative flow peaking factor of [
l l.
w i
~
~ Jac The AVB map for SG-A (Figure 6-2) has.been corrected from the letter report (Westinghouse transmittal #
VRA-89-533, R. N. Easterling to W. R. Cartwright) to show R9C35 as acceptable l
for sentinel plug removal.
In addition the map was corrected to indicate tubes R9C77 and R9C78 as sentinel plugged; these were previously indicated as L
having been inspected, but not plugged.
?
SG-B t
b.
The AVB map is shown in Figure 6-3.
/,11 Row 10, Row 11 and Row 12 tubes are supported. Seventeen (17) Row 9 tubes, forty-nine (49) Row 8 tubes, and 9213M;1 E-042789-73
.c t
- cighty-four (84) Row 7 tubes are unsupported.
Sentinel plugs were. installed in twenty-nine-(29) tubes in this steam generator.
t R9C35 was evaluated as [
[
p f
.)a,e This tube was previously sentinel plugged, based upon' the October 1987 evaluation.
l l
b 0f the remaining tubes,'RBC60 is the highest loaded tube in this steam ponerator. A conservative flow peaking factor of (
-l
)a c r
7 l
4
-gg;C'
.The AVB map in shown in figure 6-4.
All Row 11 and Row 12 tubes are
'i
~
supportd.:-Five Row 10 tubes, twenty-seven (27) Row 9 tubes, fifty (50) Row 8 j
tubes, and ninety (90) Row 7 tubes are unsupported. Sentinel. plugs were
[
installsd in forty-four (44) tubes--in this steam generator.
l R10060 w4s evaluated as unsupported, although (
[
t
[
Ja c i
The AVB. positions [
v i
3*
y
- R9C60 and R9C35 are the highest loaded tubes in this steam generator.
[
[
ya,c 9213M:1E-042789-74
(;
r l
l TABLE 6-1 North Anna #2 1 AVB Signals Determined to be Supported t
North Anna #2 Steam Generator A i
Row 12 None Row 11 None l
t Row 10 None Row 9 Columns 10, 12-16, 61, 85 Row 8 Columns 31, 73, 74 Row 7 None North Anna #2 Steam Generator B Rev 12 None Row 11 None Row 10 None Row 9 Columns 15, 53, 60 Row 6 Columns 17, 61 Row ?
None North Anna #2 Steam Generator C Row 12 None Row 11 None Row 10 Columns 41-43, 46, 48, 51, 54, 55 Row 9 None Row 8 Columns 37, 58, 59 Row 7 None c
P l
l 9213M:1E M2789-75
4 t
TABLE 6-2 l
North Anna #2 Unsupported Tube Summary A
h North Anna #2 Steam Generator A Row 12 No unsupported tubes Row 11 No unsupported tubes Row 10 No unsupported tubes Row S Columns 11, 3$, 40-55, 60, 79-84 i
Row R Columns 2-16, 25, 32 35, 38-57, 60-64, 67-72, 77 93 Row 7 Colu.sns 2-18, 21-93 North Anna #2 Steam Generator 8
' Row 12 No unsupported tubes Row 11 No unsupported tubes R w 10 No unsupported tubos Rcw 9 Columns 34, 35, 40-52. 92, 93
' Row B Columns 9-16, 23-28, 31-3E, 38-56, 60, 79-E5, 91-93 Roy 7 Columns 2-57, 60-64, 67-72, 77-93 North Anna #2 Steam Generator C Row 12 No unsupported tubes Row 11 No unsupported tubes Row 10 Columns 44, 45, 49, 50, 60 Row 9 Columns 35, 40-56, 60, 61, 79-85 Row B Columns 10-16, 32-35, 38-57, 60 64, 68-70, 77-87 Row 7 Columns 2-18, 21-93 9
o I
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O Figure 6-1 AVB Insertion Depth Confirmation g
m w-w,,-
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Figure 6-2 AVB Projection Depth = 9.00
I
[
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O 7.0 THERMAL AND HYDRAUllt ANALYSIS This section presents the results of a thermal and hydraulic analysis of the flow field on the s6condary side of the steam generator using the 3-D ATHOS p computer code, Reference (7-1). The major results of the analysis are the cater /stsam 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
- results, in the following subsection, operating condition data for North Anna 2 are presented. Data for three conditions are included:
(1)operationin Cycle 5B pr'ior to the installation of downcomer resistance plates, (2) recent operation in Cycle 6 with the new downcomer resistance plates installed, and (3) operetion with reduced primary fluid temperatures. A description of en ATHOS model and some sample results previously completed for North Anna Unit 1 operation with duncomer resistance plates installed are included in the next tro sections. The final section describes an analysis of the operating history d6ta for North Anna 2. This analysis defines a parameter termed the normalized stability ratio which provides 6 relative indication of the effect of past operation on the plant's fluidelastic stability ratio. 7.1 Not th Anna 2 Steam Generator Operating Conditions Recent steam generator eperating condition data for North Anna Unit 2 were provided by Virginia Power and are summarized below. The data are representative of operation in Cycle 6 with the new downcomer resistance plates installed in all three generators. Recent Full Load Operating Parameters for North Anna 2 (SG A) a. Steam pressure - 894 psia 6 b. Steam Flowrate - 4.25 x 10 lbm/hr c. Feedwater Temperature - 434.B'F d. Primary Inlet and Outlet Temperatures - Tin = 617.9'F, Tout 552.5'F e. Thermal Load - 975.1 MWthermal (100. M of full power) 3 9213f t:1E-042789-B4
With the above data, calculations were completed using the Westinghouse SG performance computer code, GENF, to verify the plant data and to establish a g complete list of operating conditions required for the ATHOS analysis. The GENT code determines the primary side temperatures and steam flow rate required to obtain the specifie' steam pressure at the given power rating. d 3 Besides confirming these parameters, the code calculates the circulation ratio which is of primary importance to the stability ratio analysis since it, together with the steam flow, establishes the total bundle flow rate and average loading on the tubes. It also provides an overall indication of the voids within the tube bundle since the bundle exit quality is inversely proportional to the circulatite ratio (Xexit = 1/cire ratio). The calculated circulation ratio along with the other thermal / hydraulic conditions for Unit 2/ Cycle 6 are listed in Table 7-1. Note that the l cirrulation ratio includes the effect of downcomer resistance plates which were %telled in Unit 2 prior to the November 1987 restart (following the tube rupture event in Unit 1). For comparison, Table 7-1 also includes l: pirameters for operation in Oy,le SB prior to the installation of the plates. l 1he add (sd flow resistance associated with the plate has led to a significant reduction in the circulation ratio comcared to prior cperation [ la.c. The resulting decrease in bundle flow and_ loading of the tubes in the U-bend his greatly reduced the potential for fluidslaatic vibration instability. Table 7-1 also includes a set of operating conditions having reduced primary fluid temperatures which the utility is considering for future operation. These conditions were also supplied by Virginia Power and are based on an actual test with the turbine valve wide open at full power. The measured steam pressure from this test (823 psia), ther9 fore, represents the lowest l pressure which can exist, without making turbine modifications, while still L maintaining full power. Performance calculations were also completed for this set of conditions. Note that the reduced temperature condition has essentially the same steam flow rate, circulation ratio, and bundle flow rate p as exist for the recent operating condition. The steam pressure reduction from 894 to 823 psia, however, is significant since it will result in higher l 9213M.1E-o42789-85
p i I fluidelastic stability ratios as a result of Doth higher U-bend tube gap velocities and decreased damping. i Also included in Table 7-1 are the corresponding operating conditions for Unit 3 1, i.e., prior to the tube rupture event, post-rupture operation with-downcomer resistance plates installed, and proposed operation with reduced j primary fluid temperature. A comparison of the corresponding Unit 1 and 2 conditions indicates that only small differences exist in some of the parameters. The effect of differences in operating conditions on stability ratios can be determined with a one-dimensional (1D) relative stability ratio calculation method. Adjustment factors determined from the ID method also provide a means of generating simulated 3D stability ratios for an alternate set of operating conditions without having to complete a specific, detailed 3D flow field calculation. 5 The ID relative stability ratios for all three operating conditions :n both units are also listed in Table 7-1. A detailed description of the 1D relative stability ratio is provided in a later sub-section, as it applies to the analysis of historical operating data. However, for the prasent discussior, 't is sufficient to state that the ID relative stability ratio is primarily deperdent upon three operating parameters: power level / steam flow, steam pressure, and the circulation ratio, (Primary side temperatures havn only a very minor influence on stability ratios). As mentioned previously, the steam flow rate and circulation ratio influence the total bundle flow rate and tube-to-tube gap velocity in the U-bend. The steam pressure also influences the gap velocity via the void fraction and density, however, its major impact is on the tube damping. High U-bend flow along with low steam pressure results in a higher loading on the tubes with reduced damping. Both of these factors lead to higher, more limiting stability ratios. As indicated by the comparison in Table 7-1, the relative stability ratios calculated for each condition in Unit 2 are within 1% of the ratios calculated f for the corresponding conditions in Unit 1. In particular, for (' pre-modification operation, the Unit 2 value is only 1.009 x the reference Pre-Mod Unit I value. The slightly higher value is the result of decreased L 9213M.1E-o42789 86 e
damping associated with a lower steam pressure (870 vs 890 psia). For recent eperation, the ratios are 0.897 and 0.890, respectively, for Units 1 and 2. 1 g Ratios calculated for proposed operations with reduced primary temperature are higher compared to current operation. Again, however, the ratios for Units 1 s and 2 are nearly the same. 0.950 and 0.944, respectively. The fact that the operating conditions and ID relative stability ratios for Units 1 and 2 are so close is important, in that it permits the application of existing 3D stability ratios derived from ATHOS flow field calculations for Unit 1, along with small stability ratio adjustment factors derived from the ID method. In particular, a reference set of Pre-Mod 3D stability ratios is, generated for Unit 2 by applying the 1.009 adjustment factor to the existing 3D stability ratios f or Pre-Mod Unit 1. Simulated 3D stability ratios for other Unit 2 conditions can then be generated by applying the appropriate adjustment factors to this reference set of 3D ratios for Unit 2: a) for recent operation in Cycle 6, the adjustment factor is 0.890/1.009 = 0.882 and b) for reduced temperature operation it is 0,944/1.009 = 0.936. The similarity of operating conditions and ID relative stability ratios for Units 1 and 2 also means that the ATHOS 3D flow field simulation described in the next two sections based on Unit I recent operation is also applicable to current operation in Unit 2. Justification for use of a simplified, one-dimensional, relative stability ratio adjustment factor is provided by making comparisons with the results obtained from more detailed three-dimensional flow field / tube vibration calculations. The attached Figure 7-1 presents the comparison of the results of the two calculation methods for ten other 51, 44, and 27 Series generators which have been evaluated, to date. The three-dimensional results are based on use of bundle flow fields predicted with the ATHOS3 computer code (Reference 7-1). Both cylindrical and Cartesian models have been used in the ATHOS3 simulations. Note that the results plotted in Figure 7-1 do not include the effects of anti-vibration bars. The comparisons indicate that the 1D method provides a good or modestly conservative prediction of the 3D relative stability ratios for these similar l l l' 9213M1E-042789-87 l
generator models. Note, in particular, that the ID method essentially bounds the maximum 3D ratios observed for each tube row. This is so for the smaller ] g radius tubes which, based on past experience, are typically the tube rows of interest in the tube vibration / fatigue evaluations. The variation in ratios j h for the plants within each steam generator model reflects differences in the basic thermal / hydraulic operating conditions (Wsteam, Psteam, and cire ratio). Further, this plant-to-plant variation is maintained for each of the tube rows which are plotted. The' fact that the plant-to plant variation in the ID ratios follows the 3D trends, indicates that the operating condition contribution to the relative stability ratio can be adequately accounted for by the ID approach. Overall, the comparison demonstrates that the 10 calculation method can provide meaningful relative stability ratios in support of tube fluideiastic vibration / fatigue assessments. In particular, the one-dimensional technique can be used to adjust tube-specific stablity ratios determined from detailed three-dimensional calculations for the effects of differences in thermal / hydraulic operating conditions. This 1D-to-30 adjustment is justifiable as long as its applied within a group of steam generators which share a common tube bundle configurat4un, as in the case of the 27, 44, and 51 Series feed ing generators. In these situations, the overall tube bundle flow fields will be similar and the individual plant ratios will differ only as a result of the effects of variations in the basic thermal / hydraulic parameters. 7.2 ATHOS Analysis Model The calculation of relative stability ratios involves comparing the stability ratic calculated for one or more tubes to the ratio calculated for the ruptured Row 9 Column 51 tube in North Anna, it makes use of ATHOS computed i flow profiles. 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 ATH05 model. 3 The ATHOS code does not include the capability to model the prosence of the AVBs in the U-bend region. However, Westinghouse has modified the code to 9213M:1E-042689 88
U. include the capability to model the AVBs via flow cell boundary resistance factors. Practical lower limits of cell size in the ATHOS code, however, g prevent a fine grid representation of the AVB V-bar shape which, in turn, limits the accuracy of the AVB representation. ATHOS calculations have been 3 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 are essentially the same for models with or without AVBs. The ATHOS AVB modeling sensitivity studies with uniform insertion show some tendency for the AVB resistance effects to lower tube gap velocities near the central regions and to increase velocities near i the peripheral tubes. However, the magnitude of this effect is uncertain due to the limitations in ATHOS for modeling the AYBs. Further, the global flow l resistance of staggered AVB insertion would be less than that from uniform insertion. Based on the sensitivity studies using ATHOS models with and i uithout uniformly inserted AVBs, the most reliable relative stability ratios (for actual steam generators with non-uniform AVB insertion depths) are expected using ATHOS models excluding AVBs and effects of variable AVB insertion depths by using flow test results of actual AVB geometries. The North Anna analysis is based on a Cartesian coordinate system for the array of flow cells instead of the typical, and recre widely used, cylindrical coordinate system. With a Cartesian coordinate system the tube array and any AVBs are arranged in a square pitched configuration 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 North Anna steam generator consists of 13,050 flow cells having 30 divisions in the x-axis (perpendicular to the tubelane) direction,15 divisions in the y-axis (along the tubelane) direction and 29 divisions in the axial (z-axis) direction. In,the ATHOS analysis, the steam generator is considered to be symmetrical about the x-axis of the tube bundle. The model therefore, consists of one-half of the het leg and one-half of the cold leg sides of the steam generator. Figures 7-2 and 7-3 show the plan and the elevation views of the model. These two figures show the layout of the flow cells and identify locations for some of the geometric features. 9213M:1E-o42689-89
As shown in Figure 7-2, with the Cartesian coordinate system, the circular j wrapper boundary is represented by a step-wise wall as indicated by the heavy lines. All of the simulated flow cells outside the simulated wrapper boundary above the first axial slab were blocked off by specifying extremely high flow 3 resistances on the faces of the appropriate cells. Tubelane flow slots in the tube support plates are also modeled. Figure 7-4 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 fineness of the cell mesh is evident; the largest enlis contain only 20 tubes while some of the smallest cells include only three tubes. Note, in particular, that additional detail was added near the bundle periphery (!Y=12-15) to more closely model the inner radius tubes (rows 115). Five axial layers of cells were inc?uded in the U-bend near the top tube support (Figure 7-2, 12 16 to 12 21) to more closely model the flow conditions in the area of interest. 7.3 ATHOS Results The results from the ATHOS analysis consist of the thermal-hydraulic flow carameters 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 ATMOS are defined on the surfaces of a flow cell, the tube gap valocity, which is the appropriately interpolated cell velocity raticed upward to account for the minimum flow area between the tubes, and density distributions along a particular tube required for tube vibration evaluation cre determined by a post processor from the ATHOS output. The post processor generates a data file which contains the gap velocity, density, void fraction and tube metal temperature distributions for all the tubes in the model and the file serves l 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 l the ATHOS calculated velocities (defined on the cell surfaces) and density 3 (defined at the center of the cell). The post-processer performs the 9213M:1 E-o42689-90
- m necessary interpolations to determine in plane and out of plano velocity distributions at specific intervals along the length of the tubes.
A selection of ATHOS results for Unit 1 operation in Cycle 7B with the downcomer resistance plates installed are presented in Figures 7-5 to 7-11. g. . As discussed in a previous sect on, these resu ts are a so applicable to i l l recent operation in Unit 2. Figure 7-5 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 7-3). It is seen that in t'is U-bend region the mixture turns radially outward, normal to the curvature of the bends toward the region of least flow resistance (i.e., outside the dome formed by the U-bends). Figure 7-6 shows the resultant vectors of the radial and circumferentif + 41ocity components on the horizontal plane at Z = 16, above the top tube support plate (see Figure 7-3). The radial outward flow is more evident from this figure since it ignores the axial component. Figure 7-7 shows the contour plot of the vertical velocity component (Vz) on the same horizontal plane (2=16). The high velocity gradient around the flow slot openings in the top tube support plate is clearly shown in the figure. Figures 7-8, 7-9 and 7-10 show a sample of the individual tube gap velocity, density and void fraction distributions along three tubes at Row 9. In each figure the parameters along the length of the tube are plotted from the hot leg tubesheet end on the lef t of the figure to the cold leg end on the right. The gap velocity shown in these figures are the in plane gap velocity acting in the direction normel to the tube. The gap velocity, density and void fraction are the data needed for the tube structural calculations. Figure 7-11 shows the plot of the average in plane gap velocity normal to the tube and density profiles in the U-bend span of the tube as a function of the column number along Row 9. The average values were taken as the numerical average of the parameter over the entire 180* span of a U-bend at a given column location. The average velocity is seen to be relatively constant with f-values ranging from 9.5 to 10.4 ft/sec. The average density is also quite 3 L constant with a value about 8.8 lb/ft. The wavy shape in the curves is due to the effect of the flow slots along the tubelane in the tube support plate l on the distribution through the top tube support plate. I l l 9213M:1E-042789-91
s. 7.4 Relative Stability Ratio Over Operating History 4 One aspect of the evaluation of the North Anna 2 steam generators is to examine the operating history data and use it to determine the susceptibility ( to fatigue from fluidelastic vibration resulting from the 9 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 fluidelastic stability ratio for each period of a plant's operation (fuel cycle) to a reference stability ratio, typically 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 ratio is subsequently combined with an absolute stability ratio for the reference operating point derived'from detailed three-dimensional thermal / hydraulic and tube vibration calculations. High values for the net stability ratio, in particular, over a significant period of operation, coupled with other prerequisite conditions (e.g., absence of AVB support and denting at the top tube support plate), could indicate an increased susceptibility to fluidelastic vibration instability and fatigue. The fluideiastic stability ratio is defined as the ratio of the effective fluid velocity acting on a given tube to the critical velocity at which large amplitude fluidelastic vibration initiates: U Fluideiastic effective Stability Ratio, SR = (1) Ucritical at onset of instability In this ratio, the effective vebcity depends on the distribution of flow velocity and fluid density, and on the mode shape of vibration. The critical velocity is based on experimental data and has been shown to be dependent upon r the tube natural frequency, damping, the geometry of the tube, the tube pattern, and the fluid density, along with the appropriate correlation coefficients. 9213M:1E-042789-92
p s !? Th'e' detaileo calculation of this ratio using velocity and density Tdistributions, etc., requires three-dimensional thermal / hydraulic and tube vibration calculations which are time consuming. Alternately, a simplified, one-disnsional. version of this ratio has been used to provide a relative p assessment technique for determining the effect of past operation on the stability ratio. The normalized stability ratio is defined by the following equation: -a,c- [2] ~ In this equation "cyc x" refers to each fuel cycle and "REF" to the reference operating condition. While this simplified approach cannot account for three-dimensional tube bundle effects, it does consider the major operational parameters affecting the stability ratio. Four components make up this 2 ratio: a loading term based on the dynamic pressure (pV ), a tube . incremental mass (m) term, the natural frequency of the tube (f ), and a n damping ratio (6) term. It should be noted that the ratio is relative, in -that each component is expressed as a ratio of the value for a given fuel cycle or power level to that of the reference operating point. [ 3a,c, The particular damping correlation which is used for all normalized stability ratio calculations is based on a dented condition at the top tube support plate (a clamped condition, as discussed in Section 5.2). The clamped condition is also assumed in calculating the tube natural frequency. h' 9213M;1E-042789-93
p.- The reference three-dimensional stability ratio calculation for the North Anna 2 steam generators was based on the following operating parameters which are 4 -- representative of full power operation in Cycle 5B prior to the installation of.downcomer resistance plates: b 6 Steam Flow 4.25 x 10 lbm/hr Steam Pressure 870 psia Circulation Ratio [ Ja,c (Westinghouse calculation) In addition to the reference pre-modification stability ratio, relative stability ratios were generated for three high power levels within each of the six fuel cycles completed, to date. S.ince tube vibration and possible fatigue ~ are associated with operation at close to 100% power, only the higher power operating periods are considered important to the evaluation. The high power operating experience is summarized in Table 7-2. It lists the number of days in each fuel cycle that the unit operated within three high power intervals (85-90,90-95and95-100%). Also listed are the full load operating parameters for each cycle. Note that in using this data, it has been conservatively assumed that the total operating time within each of the three power intervals is assigned to the highest power / stability ratio condition in the interval. 4 The resulting normalized stability ratios for Unit 2 are shown in Figure 7-12. In this figure, the normalized stability ratio is plotted against cumulative' operating time above 85% power. _ The reference value (=1.00) is for the full power operating condition on which the pre-modification 3-D stability ratios are based, i.e., on operation in Cycle SB prior to the installation of the downcomer flow resistance plates. The additional flow resistance associated with the new downcomer resistance plates has resulted in a significant reduction in the total bundle flow and, in particula,r, in the flow leading on the tubes in the U-bend. This is evidenced by the 12% reduction in - stability ratio which occurred between Cycles 5B and 6. The reduced ratios at 90 and 95% power for the previous cycles are the combined result of both decreased loading on the tubes and increased damping. Higher damping is a result of lower voids in the U-bend which occurs when the steam pressure rises 9213M:1E-042789-P4
y; F at reduced power levels. The information shown in' Figure 7-12 is utilized in. .the fatigue evaluation presented in Section 9.0.. l
References:
E 7-1 L. W. Keeton, A. K. Singha1, et al. "ATHOS3: A computer Program for Thermal-Hydraulic Analysis of Steam Generators", Vol.1, 2, and 3, EPRI NP-4604-CCM, July 1986. l i I i y' i 9213M:1E-042789-95
' Table 7-1 North Anna 2 Steam Generator Operating Conditions _ -and Comparison with the Corresponding Conditions in Unit.1. ~ Recent Operation with. Proposed Operation with Pre-Modification Downcomer Resistance Reduced Primary Fluid. Operation Plates Insta11ated Temperature Unit.1 Unit 2 Unit 1 Unit 2 (Cycle 7A) (Cycle 58) (Cycle 78) (Cycle 6) Unit 1 . Unit 2-Power (Mwt) %8.3 968.3 %8.3 975.1 978.6 970.2 (100.7%) (101.1%)- (100.2%) - Steam Pressure Assigned at SG 890 870 886 894 828- '823-Outlet (psia) 6 6 6 6 6 6 Steam Flow Rate (1bm/hr) 4.24 x 10 4.25 x 10 4.25 x 10 4.25 x 10 4.29 x 10 4.25 x 10 Feedwater Temperature (*F) 440. 440. 440.5 434.8 440.5 436.4 Water level Above Tubesheet (%) 44 44 44 44 44 44 (assumed) Primary Average Temperature (*F) 586.8 586.8 585.5 585.2 580 577.7 (calc'd) Calculated Parameters a,c t Circulation Ratio Bundle Flow Rate'(lbm/hr) 10 Relative Stability Ratio Basis of Performance Analysis t -,= umns, se -- -~...
t. e y ~ x v . Table 7-2 North Anna 2 Operating History Data
FULL LOAD VALUES---------------
Distribution of Days in. Primary Steam Steam Each Power' Interval Tavg Ffow-Pressure Calculated Cycle 95-100% 90-95% 85-90% (Deg F) 10 -1bs/hr (psia) Circ Ratio Comments - a,c 'l 324 13 12 581 4.07 860 2-192 11 2 581 4.07 860 3 337 13 5 581 4.07 860 4 363 22 13 581 4.07 860 SA 114 3 0 587.8 4.08 900 Pressure Uprating 5B 266 8 6 586.8 4.25 870 Core Power Uprating-6 449* _0 _0 585.2 4.25 894 With Downcomer Resistance ~ ~ Plates' Installed ~ 2045 70 38
- EFPD's for Cycle 6 through shutdown on 02/20/89 9213M IE/042Fe9 97
i / i= RELATIVE FLUIDELASTIC STABILITY RATIOS (RETERENCED TO NORTH ANNA R9C51) I 1.a 1.7 - O 10 } 30 Cy11adrical 1.6 - 0 [ 30 Cartesian 8 } g-15-15 0 1 - now 2 aoms Below moetnal 1.4 - a - 1.3 -34 0 g{ iro} Uj g 1.2 - iso o loo y i ] 5 d 11 - ti o' g gol "4 8 ,wy",gcoruje->cyi l 3 30 { q sol al o~ og *al j o.. - ,o aal o.7 - g O.6 - 0.5 - '= u o w w a m -3 E E E E E E E E E E o.3 - g g g 2 g g g g_ g g 0.2 1 I I 27 44 51 sTEAu cENERATOR WODEL i Figure 7-1 Comparison of Relative Stability Ratios Calculated From 1D and 3D Methods l
't s 4* - ] ) SIKilATED WRAPPER BOUMMRY [ l ~ \\ l l' / N i s / N / \\ / -l l\\ / I II \\ 1 /l l II I\\ /I I i i il I\\ lI l l ll-ll l) .i ll l l Illi I vlx l .or 6 o coto tio Figure 7-2 Plan View of ATHOS Cartesian Model for North Anna l
- l
( I2= 29 -- l SEPARATOR DECKPLATE 7 9' F U-BEND i t "6-TSP 7 3 1 \\ \\ TSP 6 i 9 TSP 5 I i i TSP 4 IZ=9 i 4 TSP 3 i TSP 2 i l l TSP 1 y i 12=1 - } WRAPPER OPE!11t1G I I i-HOT LEG COLD LEG V 1 1 K 8 L. Figure 7-3 Elevation View of ATH0S Cartesian Model for North Anna
}. ' s i L .>0 3 h i FLOW SLOT IN TUBE SUPPORT PLATE l B I s N .-IY 15 ~ 1 v F ' N ns xx eesisii4444En466cis lisc444W +5E Aw> l ll l lll,[. 5 ? fili 5 2il$'!?tii'Mi' Eij1.it??9?iiti ?fUii'if th l l 8 55 I I m I I I I MiHfii!!!!!N!!Ni}i*i$545!! iisi$i!4}i!hii$i!$iii.u.pi!M4h i6iiX l l I l / 4. %..%. %. y. - ..e
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i' !. !. ! !.! ;i l i, l.l, ii. !!!i. t. !_ !. i. l,i.1 !!..!!.U..l. lil. It. i. t.:r. i. t. i t. i t.. d; ; 3 ; !..!!. !. i_ i. i.:!. !. }i.3.. g...............r i !!. U..i p;ti,i t. a, l_ i. -IY 1 3,3 1 i i .. ; i. . ii. !.!; .r 4. ' r' / HOT LEG COLD LEG i*: ^ Figure 7-4 Plan View of ATHOS Cartesian Model Indicating Tube Layout 4
w 1 1 + l ; i i ///h%wid.......,~nn, g, f,' ' M g' q"a]% 'H g'A a. vn i, w E ,,,.1 1,..uA u,, .dila&lIIililll'1 l 1l ns,v,,s., i o i e t, u i i. i ,t i l f 'i i f f;t........... ,titit" lllll\\ 1 l illl l l l ill l llll l 1 5 c i,a v s... ] l}lllll! !m i +44f 4 t ttt f f f e 9 '; llllllllll.lIllHilIlillllllllt ~ J v- , m, i.u. I lIlllllllllit t u
- 2
.....+ m,#,. lilllllllllll/H//Illlit f f f f ti s :: , s, - g i w I Illlflittfittt. h$ ,1 I 1 t i l l l l l l l l l / //.// / / / / f t i r e,,,, g,h. 4 [ t l l l i ,l 'l 'lil 'l l i g bti,,,........ l -1 ,.. + +... t \\ \\ \\ \\ \\ \\ \\ \\ \\ l l Ilit! l l t o....,,,, 1 tillIk k h \\\\\\ \\ \\ \\ \\,,,,,,,,,, ..,tf. , f f t i \\ i l l. / / /4 4 t ,,,,, n n 4 t, t t /<<//f! \\ \\, ........,srite N m... tl'~..i ,,,.....,, om HDT M G l COLD E G { Figure 7-5 Flow Pattern on Vertical Plane of Symmetry l 1.
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/ / /. m 2 1 1 I I I / / t 0 ,5 I I I I I / / I i 5 t I II I / / / r# ~ 5 2 I I II / / / / / /, u 5-I l ! I / / l' / / // r-j, I l ! I / / / / / / /} s r} f I ! / / / / / / / iiI f fm j j 't ! I I / / / / / / /s t 1 I f / / / / / / / a' kE fG ? t i ! t / / / / / /./ ' s*
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1 3.' r i N <7/ w ?N a, /./P== \\ \\ m ,,I 5 1! \\ [j' r fe Y N 'N t y' l \\ \\ l act LEC COLO LEO 'wX 2-Con #0Nthi nlarvet vtLOClfv IFT/$tC) l J I. Figure 7-7 Contours of Vertical Velocity Component on a Horizontal Plane in the U-Bend Region l l
- 1 p.
=; f f,j e 9* \\ es s-g i_ 3s - p' w u e s\\ E N esi g e / w m. I e N e o WR' d 1 / \\ u g f \\ set i /. 4 t m . esi i i i .I \\ esi i i .t => g f g T 2 g l
- m. E g g
/ E. 5l \\ l n ~ e E W i l 5 E' e 8 g. l es - v i ll l 2 e ei - i - ess y s / - ms = E i r' - ese:: = \\ w }j - ese g a s I- \\ / - esie e -i w s g a y-j esi, y E + es. -.e* m a n a e e = e.a ou--- .-%e.w e a a a = .s.-o. i,,,i,,,,,,,i,,,,,,,,,,,i,,,,,,,i,,,i,,,, e = m e e n . Figure 7-8 Tube Gap Velocity and Density Distributions for Tube Row 9/ Column 3 l l
n. 1. + t .I e P' ( -- s s. mi E-- v 5-ni! i. w \\ f e m. E o l / \\ we d-i 1> \\ y u ,1 b'
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( _E / \\ i *jE g / I as a i-a ) m -s gi \\ E s ..u. .I l a ae 3 i 1 v ll l 2 l e d-es s ms a s / i B 5 / - ese:: s = r 3 le' - esti L i s j7 me s l i s s R s / a s., s - m. n. E. r / ~ ~ 4 e e a e e e .w... e ... =... a e x i ii ,i,,ii,,,t,,ii,,,i,,,i,,,i,,ii,,,i,,,t .l i-l i i Figure 7-9 Tube Gap Velocity and Density Distributions for Tube Row 9/ Column 20 l' l-4 b
l h.. I ~ i -i 't t .e-s~~% M p / mi y s- - mil-. N, w / d I We l / 3 E o I / \\ ) m un g MS / \\ m u l i l-I I \\ I y \\ 5 E / w J-ga i 3 s
- l
{ l m'! gI 1 g. I E= .I I a 8 si h l i .g \\ ms u. = j .s, - 3 g [ s I 3 \\ \\ / ~ l \\ s .si e o s s e ie s / mi, 1 _g s s/ - es = ess=* i I $ ' ou$ N. ---.-~ .u I ...-um It 1 t !I t I !t i1 !l I f !t t I !l I t !t if !f if $f f I !I I I ! ..-u--.. Figure 7-10 Tube Gap Velocity and Density Distributions for Tube Row 9/ Column 44
y E. p t' NORTH ANNA R0W 9 AVERAGE GAP VELOCITY & DENSITY c 13 - M l E AVERAGE GAP VELOCITY a AVEllAGE DEtSITY [ 12 -- 11-10 w c) i b " aaa aaaaaaaaaanaDoanmaaana OLa .g onao y g g e i y '[ i 7 g l u _ tij: D i 0 5 10
- .5 20 25 30 35 40 45 50 COLUMN NUMBER' 1
l Figure 7-11 Average Velocity and Density in the Plane of the U-Bends Normal to Row 9
6 4 9 i. f; NORTH ANNA 2 NORMALIZED STABILITY RATIO EMSED ON HIGH POWER (>85%) OPERATION 3,g 1.02 - $B -9 g Cycles 1-4 y g,gg. C O O 0.96 - SA t 0.94 - / o. 0.92 - N ~ ERP's 1 \\ 6 19 nInstalledj x 0.88 - f a 3_ 902 Power v D 0.86 - E 0.84'- 0.82 - 0.8 -- i i i i i i i i 0 0.4 0.8 1.2 1.E 2 2.4 (Thousands) ACCUMULATED DAW AT/ABOVE INDc'D. RATIO i Figure 7-12 North Anna 2 Normalized Stability Ratio Based on High Power (> 85%) Operation 9213M:1E-042789-109
8.0 PEAKING FACTOR EVALUATION-I J 7 This section describes the overall peaking factor evaluation to defina the ' test based peaking factors for use in the tube fatigue evaluation. The j p-evaluation of the eddy current data to define the AVB configuration for North Anna 1 Tube R9C51 is described. This configuration is critical to the tube fatigue assessments as the peaking factors for all other tub 6s are utilized relative to the R9C51 peaking factor. Uncertainties associated with applying j the air model test results to the tube fatigue assessments are also included in this section. Included in the uncertainty evaluation are the following contributions: o Extrapolation of air' test results to two phase steam-water-o Cantilever tube sinvulation of U-bend tubes o Test measurements and repeatability o AVB insertion depth uncertainty S.1 North Anna 1 R9C51 Configuration 8.1.1 Background The AVB configuration of the ruptured tube in North Anna 1, R9C51, is the reference case for the tube fatigue evaluations for other tubes. In accordance with the NRC Bulletin 88-02, the acceptability of unsupported tubes is based on tube specific analysis relative to the North Anna R9C51 tube, including the relative flow peaking factors. Thus, the support conditions of the R9C51 tube are fundamental to the analyses of other tubes. Because of the importance of the R9C51 tube, the support conditions of this tube, which were originally based on "AVB Visible" interpretations of the eddy current test (ECT) data (Figure 8-1), were reevaluated using the projection technique developed since the North Anna event. The projection technique is particularly valuable for establishing AVB positions when deposits on the tubes tend to mask AVB signals such as found for the North Anna 1 tubes. The results of this evaluation are summarized below. 9213M:1E-042789-110
w i 8.1.2 - Description of the Method The~ basic method utilized was the projection technique in which the AVB 5 ~ position is determined based on measured AVB locations in larger row tubes in the same column. In this study, the projection technique was utilized in the " blind" mode, (AVBs called strictly based on the data) as well as the reverse mode (data examined on the basis of predicted AVB positions). The objective of this application was, with the greatest confidence possible, to establish the positions of the AVBs in an 8 column range.around the R9C51 tube in North Anna 1, Steam Generator C. 7 8.1.3 Data Interpretation The ECT traces for the U-bends in Rows 8-12 (in one case,13) were examined for Columns 48-55. The original AVB visible calls are shown in Figure 8-1. The data were examined by an eddy current analyst experienced in reading these traces, and by a design engineer knowledgeable in the geometry of the Model 51 V-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 ir.d1 cations where tingle AVB intersections were predicted, etc. The data evaluation method used was a critical examination of the data, which was biased toward the presence of AVBs unless a confident call of "no AVB" could be made, and then checking the consistency of the data among the tubes in a column and against the theoretical data for the predicted AVB positions. [ 9214M;1E-042789-111
o' -l e f h b 3a,c, Figure 8-4 is the "AVB visible"' map for columns 48 through 55, based on the - critical review of the data. It should be noted that the original ~ data
- interpretations and the review interpretations are consistent.
8.1.4 Projections The.( .Ja,c ECT traces were utilized for projecting the position of the AVBs according to the standard fo' mat of the projection method. r The results of'the projections are presented in Figure 8-5, which.shows a matrix of projections for tube rows 8 through 13 in columns 48 through 55. -For many of the tubes, more than one, and as many as three, projection values 9214M:1E-042789-112
l P i are shown.- Multiple projections 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 ~ support that tube. As many as four different projections 'are possible if it. B is assumed that the tube is supported by the upper and lower AVBs, and both -upper and lower bars are staggered in elevation as shown in Figure 8-2. The logic in arranging the projection data is based on the following two rules: Rule' 1. The projections of the same AVB based on different tubes in the same column ( Ja.c, [ Ja,c, Rule 2. Two adjacent tubes in the same row ( Ja,c Consequently, the difference in the [ 3a,c The implementation of this is that if the position (either lef t or right) =of a projected AVB is assumed for a column, then the projections in the adjacent columns are also (- Ja,C 9214M:1E-042789-113
-t, The arrangement of-the AVBs as shown in Figure 8-5 satisfies the rules above and is consistent with the rupture _of R9C51. The resulting AVB arrangements, s based on the projection matrix-of Figure 8-5 is shown in Figure 8-6. = 8;1.5 Conclusions -The general AVB arrangement surrounding the ruptured tube in North Anna 1 Steam Generator C, which was the basis for the analysis, is c r firmed by a detailed critical review of the ECT data. Differences exist in the AVB pattern between tube columns 48-49, in which the AVBs appear to be less inserted than previously indicated. The pattern of Figure 8-6 is the best fit to the rules which were adopted for determining the position of the AVBs, as well as consistent with explanation of the tube failure. The basis of the review was a projection technique which utilizes data from tubes one or more rows removed from the actual inserted position of the AVB to -determine the position of the AVB. The intent of the review was to establish the. positions of the AVBs by confirming or eliminating features of AVB . alignments such as side to side offsets, etc. of the AVBs adjacent to the tubes. Overall, the conclusions regarding the positions of the AVBs around R9C51 in North Anna 1, Steam Generator C are based on consistency among all the available data. 8.2 Test Measurement Uncertainties The descriptions of the peaking factor tests and apparatus were provided in Section 5.4. All practical measures were taken to reduce uncertainties. l -Nevertheless, some still remain and should be properly accounted for. The important parameter measured during testing that has a significant impact on peaking factor is the air velocity. The air velocity at test section inlet was measured using a ( 3a,c Based on considerable experience with the use of such instruments, it is known that the magnitude of uncertainty is very small. A[ ]a,c measurement uncertainty is used in this analysis based on past experience. S 214M:1 E-042789-114
9 -) i l . 8.3 Test Repeatability 'During the peaking factor testing of AVB configuration, each test was ] _ performed at least two times to confirm repeatability. It has been ] ' demonstrated that the tests are quite repeatable with the results often falling within 2 or 3% of one another for the repeat tests. An upper bound .value of 5% was used in the current uncertainty analysis. 8.4 Lantilever 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 intortion depth on the initia' tion of fluidelastic vibration. The following assumptions are used: a,e =m. 9214M:1E-042789.115
1 T - For.'the purposes'of this estimate.-the geometry of the cantilever measuring . tube in=the' air test model is compared with the geometry of a. prototypical Row' - 10 tube. (- b l t D:
- l Ja,c,
1 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 '[ t i ? I 9214M;1E-042789-116
ja,".- Using these values, the ratio of the effective velocity for J the cantilever measuring tube to that for the U-bend tube is about [ -Ja,c for the case treated. A similar evaluation can be made for a Row 10 tube that lies in the projection or shadow of'an AVB that is inserted to a d6pth required to support a Row ~9 tube. ( 3a c, The net result is that the ratio of the effective velocity for the cantilever -tube to that for the U-bend tube is about [ Ja,c, These results indicate that, for the particular assumptions used, the canti_ lever tube model appear: to be a reasonable representation of the U-bend with respect to determining relative peaking factors for different AVB . configurations. This evaluation also shows that, on the average, the magnitude.of the systematic uncertainty associated with tho use of cantilever-tube to simulate the U-bend is about [ la,c, 8.5 Air vs Steam-Water Mixture The local peaking factors from the air tests can be applied to the steam generator steam / water conditions either as a direct factor on the mixture velocity and thus a direct factor on a stability ratio, or as a factor on the steam velocity only with associated impacts on density, void fraction and damping. :This method leads to a reduction in tube damping which enhances the peaking factor compared to the direct air test value. For estimating an absolute stability ratio, this application of the peaking factor is a best estimate approach. However, for the evaluation of tubes relative to stability ratio criteria, it is more conservative to minimize the peaking factor for the North Anna Unit 1 tube R9C51 through direct application of the air test 9214M:1E-042789-117
i' i.t peaking factor. This conservative approach is therefore used for evaluating 1 tube acceptability, l 1 a Under uniform AVB insertion (or. aligned AVB insertion), there are no local open channels for flow to escape preferentially. Therefore, air flow is g q approximately the same as steam / water flow relative to velocity -i perturbations. Under non uniform AVB insertion the steam / water flow may l differ from air, as the steam and water may separate from each other when an l obstruction, such as an AVB, appears downstream. The water would continue j 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 l dater phase, i i I Based on the above discussion, the Fg are considered to more appropriately { apply to the steam phase. Thus, it follows that mixture mass velocity for the l . tube subject-to flow perturbation can be written as follows: -i a,e i -i 'where D _is the vapor density, D g f the water density, F, the velocity peaking factor determined from air tests, j
- the nominal superficial vapor g
velocity, and jf* the superficial water velocity. Steam quality can then be determined as follows: a,e t The Lellouche-Zolotar correlation (algebraic slip model), as used in the ATH0S code, is applied to determine void fraction. Subsequently, mixture density, velocity and damping coefficients for the tube which is not supported and subject to flow perturbation is evaluated. Therefore, similar to the air 9214M:1E-042789-118 ._. b
a velocity peaking factor, local scaling factors of mixture density and velocity .and damping coefficient can be readily determined. Finally, a local stability l peaking factor for fluidelastic vibration can be calculated as follows: a,c hore Fs is the stability peaking factor, Fd the density scaling factor, Fv the velocity scaling factor, and Fdp th$ damping coefficient scaling factor. If we use the air velocity peaking factor without translating to steam / water conditions, then a,e + As shown in Table 8-1 stability peaking factors for the steam / water mixture are slightly higher than air velocity peaking factors. The difference between the steam / water and air peaking factors increases as the air peaking factor increases. For application to tube fatigue evaluations, the ratio of the peaking factor for a specific tube to that for North Anna R9C51 is the quantity of interest. Larger values for this ratio are co.iservative for the tube fatigue assessment. The North Anna R9C51 peaking factor is one of the highest peaking factors. As discussed in Section 8.7, a peaking factor of nearly [ la,c is determined for the R9C51 tube. The differences between [ la.c. Typical values are shown in Table 8-2. These results show that the direct application of-the air test data yields the higher relative peaking factor compared to R9051.. To obtain conservatism in the peaking f actor 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 i 9214M.1E-o42789 119
associated with the conservative use of air velocity peaking factor is [ ja.C, The conclusion that peaking factor for steam water flow would be higher due to the dependency of damping ratio on void fraction was supported by an alternate p study. In this study, a section of steam generator tubes were simulated using the ATHOS code under prototypic flow conditions. The objective of this study e:as to examine the magnitude of the changes in void fraction and thus stability ratio as a consequence of non uniform AVB insertion patterns. The current version of ATHOS has modeling limitations that prevent accurate modeling of local geometry effects. In addition, it is believed that an analysis using two-fluid 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 ATHOS model, the ATHOS results were processed by the TLOVIB 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 [ la.c. This alternate calculation provides independent corroboration of the prior discussion regarding the stability peaking factors under steam-water conditions vs in air. 8.6 AVB Insertion Depth Uncertainty The most significant uncertainty for the low peaking configurations is not in the test results, but in the determination of actual AVB insertion patterns adjacent to specific tubes. The methodology used for obtaining the AVB insertion patterns from eddy current data can ascertain the AVB location only approximately. The effect on peaking factor resulting frein this uncertainty is addressed using test results of AVB configurations that varied from one another by up to [ Ja.c, 9214M:1E-042*89-120
.c Based on maps of AVB insertion depth of various plants, several configurations have been tested for determining fluidelastic instability flow rate by an air ( cantilever model. Stability peaking factors were then determined from the ratio of critical flow rate for a uniform AVB insertion configuration to a e specific configuration. Figure 8-7 summarizes the AVB configurations tested. Position of AVB insertion depth is determined from Eddy Current Test (ECT) data. Positioning of AVB from ECT data reading is subject to uncertainty; its accuracyisprobablyabout[ Ja.c. A change of an AVB insertion depth in a given configuration leads to a different configuration, and thus a different peaking factor, A review of the tested AVB type has been made and results 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 ( )*'C. Such a trend can be explained; a decrease in a specific AVB depth aill open up more channels for incoming fluid to distribute and thus less flow perturbation. However, this applies only to those changes without inducing the reinforcement of flow perturbation from upstream to downstream. On the average, the uncertainty in peaking factor resulting from small variations in AVB insertion (of the order of 1/2 tube pitch) is found to be [ la,c, l ) S214M:1 E-042789-121
f 8.7 Overall Peaking Factor with Uncertainty As discussed in the previous subsections, there are several aspects to be considered in applying the laboratory test data to steam generator conditions. These considerations were reviewed one at a time in those subsections. This section will integrate the pieces into one set of stability peaking factors. Looking forward to how these peaking factors are used in the analysis (Section 9), the relative stability ratio calculated for a given tube without the consideration of flow peaking is corrected using the ratio of the peaking factor of tne specific tube to that of the North Anna Unit 1, SG-C R9C51 tuce (Configuration la). It is to be noted that the test results would be applied as ratios of a specific tube peaking factor to the R9C51 peaking factor. This will reduce the influence of some uncertainties since the systematic uncertainties would affect both the numerator and the denominator in the ratio of peaking factors. The major difference will be in those configurations whose peaking factors are significantly lower than that of R9C51. The approach employed here is intended to provide that conservative peaking factors are employed for such apparently low peaking configurations. The uniform AVB configuration (2a) is selected as a reference configuration, and the peaking factors of all configurations tested are recomputed on the basis of this reference. As discussed below, some of the test uncertainties are applied to the reference case to account for its significantly low peaking relative to the R9C51 configuration. The uncertainties in the test results and their extrapolation are those due to test measurements, test repeatability, cantilever tubes in the test vs U-tubes in the steam generator, and air tests vs steam water mixture. These were discussed in more detail in the previous subsections. The magnitude of these uncertainties are listed in Table 8-4. 1 9214M 1t-042789-122 i
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 [ 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( Ja,c instead of the [ Ja,c appearing in Table 5-2. The next three uncertainties in Table 8-4 are systematic uncertainties, it could be argued that these appear in the peaking factors of both the specific tube under consideration and the R9C51 tube and are therefore counter balanced. However, the relative magnitude of these may be different, particularly for configurations with much lower peaking than R9C51. Therefore it was judged that the ( la c. Similarly, as noted above, the effect on peaking factor due to the uncertainty in the field AVB configuration is also included in this reference case. Thus,[ ]a,c. The peaking factor of the reference configuration 2a (Table 8-5) is raised by this amount to a value of ( ]a,c, The change in peaking factors of configurations la and 2a resulting from the application of uncertainties as described above are shown in Column 3 of Table 8-5. The peaking factors of all configurations are recomputed on the basis of this reference configuration (2a). These values are displayed in Column 4 of Table 8-5. Some of the uncertainties were applied to the reference configuration (2a) in order to apply them to all low peaking configurations conservatively. Thus, no configuration should have a lower peaking factor than this reference configuration. Therefore, when a peaking factor value less than [ ]a,c is calculated for any configuration, (in Column 4 of Table 8-5), it should be altered to [ la,c. Further, for some of the configurations that are
conceptually similar, the more limiting (higher) value is used. For example, a peaking factor of ( Ja,e is used for configurations 5a and Sb based on their similarity to configuration Sc. The final stability ratio peaking factors calculated on this basis (with configuration 2a as the reference) are shown in Table 8-6. 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 fs.ctors. The uncertainty estimates were developed from both test and analysis results as described in Sections 8.2 to 8.6. The largest single uncertainty of ( Ja,c is attributable to uncertainties of up to [ ]a,c or. determination of AVB insertion depths from field eddy current data. This relatively large uncertainty is applicable only to low peaking conditions where the AVB uncertainties can contribute to small peaking factors. The definition of "no flow peaking" was increased to encompass the small peaking effects from AVB insertion uncertainties. For the AVB patterns leading to significant peaking factors, AVBs were positioned within uncertainties to maximize the peaking factor. For these configurations, variations of AVB insertion within these uncertainties are expected to reduce the peaking factor compared to the final values of Table 8-6 and Figure 8-7. 2. Including uncertainties directed toward conservatively decreasing the peaking factor for the North Anna tube R9C51, the final R9051 peaking 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 The AVB positions on each insertion pattern of Figure 8-7 should be carefully noted.[!
l* Ja,e i-p- , [' la,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 North Anna 2, [ )"'C Based on the R9C51 tube vibration analysis, flow peaking factors on the order of ( ) a,c for-Row 8 tubes and above [ ] a,e for Row 9 tubes would be required for tube fatigue to be a concern. [ 3a,e Determination of peaking factors for identified tubes shown in Table 8-7 are described in detail.- Table 8-7 is divided into small tables for, ease in following the description. 9214M:1E-042789-125
8.8.1 Steam Generator A The following table gives the peaking factors for Steam Generator A tubes with I, unique configurations of AVB insertion depths. c P-Steam Type of AVB Peaking Generator. Row No Column No insertion Depth Factor .A' 8 64 a,c 9 60 35 All of the Remaining ( 3a,e 8.8.2 Steam Generator B The following are a list of 5 tubes with unique AVB configurations.. Steam Type of AVB Peaking Generator Row No Column No insertion Depth Factor B 8 81 a,c 60 31 9 44 35 All of the Remaining For RBC81 and R8C31, ( 3a,e i 9214M.1E-042789-126 i I ___j
-8.8.3 Steam Generator C Tubes with unique AVB configurations were evaluated. The following table lists ~ b their peaking factors and types of AVB configurations, e Steam Type of AVB Peaking -Generator. ' Row No Column No Insertion Depth Factor ~ 'C 9 83 a,e 60 55 -40 35 10 60 R9C83 belonged to ( 3a c 4 9214M:1E-042789-127
f Table 8 1 I Stability Peaking Factor Due to Local Velocity Perturbation Sr.aling Factors for Steam / Water Air Velocity Void Stability Peaking Fraction Density Velocity Damping Peaking
- Factor, Scaling,
- Scaling, Scaling,
- Scaling, Factor, F
F Fd f F F a y y dp 3 . a,c NOTE: 1. Stability peaking factor for steam / water mixture is calculated as follows: a,c 2. Damping scaling factor is calculated using modal effective void fraction of [ ]a,e for R9C51 tube, i j
- y..
4 i t Table 8-2 ' Comparison of Air and Steam-water Peaking Factor _ Ratios 1. Air-Air Steam Steam- . Peaking - Peaking Peaking Peaking-Factor Ratio
- Factor
. Ratio a,c 9 1
- n'.
Table 8-3 Effect of Local Variation of AVB Insertion o. A to B AVB Peaking Peaking Rotio i Type A Type B Variation Factor A Factor B (B/A) . a,c Sb Sa 4a 5c Sc Sa a,c Sa 5b Sc 4a Sa Sc P e 1
Table 8-4 Uncertainties in Test Data and Extrapolation Source of Uncertainty lyng Maanitude. % a,c 1. 2. 3. 4. 5. This is not an uncertainty associated with the test data. It results from the inaccuracy in determining the true AVB position in the field using eddy current data.
Table 8-5 Extrapolation of Test Results to Steam Generator Conditions Peaking Factor Test Data with Referenced to Configuration Data Uncertainties Configuration 2a a,c 9214M:1 E-042789-132
Table 8-6 Final Peaking Factor c Configuration Peaking Factor 8,C 9214M:1E-042789-133
Table 8-7 Stability Peaking Factors for Specific Tubes North Anna 2 Steam Type of AVB Peaking Generator Row No Column No Insertion Factor s A 8 64 a,e 9 60 35 l All of the Remaining B 8 81 60 31 9 44 .35 r All of the Remaining C 9 83 60 55 40 ~ 35 10 60 All of the Remaining +- 9214M:1E-042789-134 - - + {
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- 1. 23
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j-e s .....s.- w-p ey, +c +:. go.3, ......,,,L........ ~~~... 1. 1 l .l. l l. l l. l l. l l NTS I. I Figure 8-2 Schematic of Staggered AVBs
4-a.c i i Figure 8-3 AVB ' Pair" in ECT Trace
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-__5_REFEHWERREE ,8.. .. c _ EM E 3 E2 4 EG R... E' H EN...... 1 2R 4 A EN. IE EN...H B BE RE IR ~~~ 15 61:5 8888-390 8 ~ as. '... 'M n hW..nA..pEd e,s.sa h as' it it L L' R8 = = Ivs Ivs vah ! ',7vs Ive! Ive Ivs IvoIve! I Ivs Eve 1 C58 C54 CRB C82 C81 C80 C40 C48
- P'emmasy upsmans h hWW 444gh Side' Pqemton.
Figure 8-5 North Wma 3, Steam Generator C, R9C51 AVB [ %C Matrix
O O O O 0 0 0 0 O O.O.O. O@ O'OOO 0000000000 $6 34 53 52 51 30 49 48 47 46 45 44 ( n g i )
TYPE OF AVB PEAKING TYPE OF AVB PEAKING TYPE OF AVB PEAKING INSERTION FACTOP INSERTION FACTOR INSERTION FACTOR ~ a,c a,e 1a 2a 5a 1b 4a Sb le 4b SC 10 di 6a Sw 4r 8d 12 4s 8b Figure 8-7 Final Peaking Factors for North Anna Unit 2 9214M:1E-042789-141
9.0 STRUCTURAL AND TUBE VIBRATION ASSESSMENTS ~ 9.1 Stability Ratio Distribution Based Upon ATHOS An assessment of the potential for tubes to experience fluideiastic instability in the U-bend region has been performed for each of the tubes in rows eight through twelve. This analysis utilizes FASTVIB, a Westinghouse proprietary q 4 finite element based computer code, and PLOTVIB, a post processor to FASTVIB. These codes predict the individual responses of an entire row of steam generator tubing exposed to a location dependent fluid velocity and density profile. The program calculates tube natural freque'cies and mode shapes using a linear finite element model of the tube. The fluitelastic stability ratio q U,/U (the ratio of the effective velocity to the cri'.ical velocity) and t the vibration amplitudes caused by turbulence are calculated for a given 4 velocity / density / void fraction profile and tube support condition. The velocity, density and void fraction distributions are determined using the ATHOS computer code as described in Section 7.3. The WECAN generated mass and stiffness matrices used to represent the tube are also input to the code. (WECAN is also a Westinghouse proprietary computer code.) Additional input to FASTVIB/PLOTVIB consists of tube support conditions, fluidelastic stability constant and turbulence constants, l This process was performed for the North Anna Unit 2 steam generator tubes under consideration and also for the premod North Anna Unit 1 Row 9 Column 51 tube (R9C51). Ratios of the North Anna Unit 2 results to those for the premod l R9C51 Unit 1 tube were generated to produce a quantity that could be used to provide an initial assessment of the North Anna Unit 2 tubes relative to the ruptured R9C51 Unit 1 tube. Note that three separate steam generator conditions are to be considered in this evaluation: j 1) Premod - The conditions that existed in the steam generator before the I downcomer resistance plate was installed. 9214M 1E-0427B9-142
h
- 2) Postmod without T Reduction - Conditions that exist af ter hot installation of the downcomer resistance plate (not including any T
reduction effects). hot
- 3) Postmod with T Reduction - Conditions that exist af ter
[ hot installation of the downcomer resistance plate (including the effects of T reduction). hot q Relative stability ratios (and stress ratios) have been generated for each of the conditions listed above. Section 7 contains details of the calculations used to define the fluid conditions used in the evaluation. However, it can be noted in this section that results for both the Postmod With T reduction hot case and Post' od Without T reduction case were obtained through the use of m hot e ratios applied to the Premod results. Figure 9-1, 9-2 and 9-3 contain values of relative stability ratio (including flow peaking) for each of the 3 cases described above. The relative ratios contained in the figures were obtained using the following conditions for the Premod Unit 1 R9C51 tube and also for the Unit 2 tubes under consideration:
- 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 9-1, 9-2 and 9-3 are composites of all steam generators using mirror image tubes., That is, any peaking effect for a given tube located on the plot represents the maximum value of the peaking factor in all steam generators at that location.
i 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 the premod Unit 1 R9C51 tube. (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 the premod R9C51 value. These figures indicate that several tubes in Row 9 have relative stability rstics that lay above the 0.90 line. These tubes (SG: A R9C60, SG:B R9C35, SG:C R9C35 and R9C60) are enveloped by a single value and appear in the figure as a single point. (Note that all tubes in Rows 11 and 12 are supported and therefore can be removed from consideration. These tubes were included in the figures for completeness and comparison purposes.) Table 9-1 contains a summary of values of relative stability ratio, including flow peaking, for the tubes with significant relative flow peaking or relative stability ratio for the cases described above. As can be observed in the table, all tubes that have a relative stability ratio greater than 0.90 for the Premod conditions have relative stability ratios less than 0.90 both the Postmod and Postmod with T reduction cases. This indicates that hot additional analysis is recuired to determine acceptability of these tubes. 9.2 Stress Ratio distribution with Peaking Factor An evaluation was performed to determine the ratio of the North Anna Unit 2 tube stress over the premod North Anna Unit 1 R9C51 tube stress. This retic is determined using relative stability ratios discussed in the previous section, relative flow peaking f actors (Table 8-7 f actors divided by [ Ja,c) anc bencing moment factors. Sections 4.2 and 4.3 contain additional 'nformation and cescribe the calculational procedure used to obtain tne resul<s presented in this section. The results presenteo below are based upon the #ollowing conditions: 1) Tube is fixed at the top tube support plate, 2) Damping is void fraction dependent, 9214M 1 E-042789-144
- 3) Tubes have no AVB support,
- 4) 10% criteria with frequency effects,
- 5) Tubes are assumed to be dented (emax' y)
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 t'ube is expected to be insufficient to produce a fatigue event in a manner similar to the rupture that occurred in the R9C51 tube at North Anna Unit 1. Figures 9-4, 9-5 and 9-6 show the results of the stress ratio calculations for each of the North Anna Unit 2 tubes in Rows 8 through 12 for the three cases described earlier. 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 tcp tube support plate corrosion plus magnetite in the crevice without tube deformation. As can be observed in Figures 9-4, 9-5 and 9-6, several tubes have stress ratios that lay above the 1.0 acceptance line for the pre-mod case. These tubes (SG: A R9C60, SG:B R9C35, SG:C R9C35 and R9C60) are enveloped by a single value and appear in the figure as a single point. (Note that all tubes in Rows 11 and 12 are supported and therefore can be removed from consideration. These tubes were included in the figures for completeness and comparison purposes.) As with the relative stability ratio figures, the stress ratio figures are also composites of all three steam generators using mirror image tubes. Specifically, 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. Table 9-2 contains a summary of values of stress ratio for the tubes with significant relative flow peaking or relative stability ratio for the cases 9214M;1E-o42789-145
l described previously. As can be observed in the table, all the tubes that have stress ratios greater than 1.00 for the Premod conditions have strtss ratios less than or equal to 1.00 for both the Postmod and Postmod with Tg reduction cases. Note that acceptance, based upon 40 years of operation, is determined in part on tubes having stress ratios less than or e::ual to 1.00. The tubes having stress ratios greater than 1.00 for the Premod condition but having stress ratios less than or equal to 1.00 for the two Postmod cases must be evaluated in detail to determine the actual and projected fatigue usage associated with each tube. Final acceptance will be determined using this method. 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 9-7 = j contains the results of this process for all the tubes in Rows 8 through 12. This figure was generated using the conditions described earlier for the Premod case. The Premod case was selected to because it is the most limiting of all the conditions currently under consideration. Note that this figure reads opposite of the previous figures, i.e., the top curve in the figure eorresponds tc Row 8 and the bottom curve corresponds to Row 12. Maximum Allowable kelative 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 not to exceed 1,0. 3 This curve can be used to help identify the relative flow peaking required before preventative action would be recommended and, when used in conjunction h with the actual flow peaking associated with each tube, to determine the l margin (if any) present. This has also been performed in Table 9-2. The column with heading " Max Allow Flow Peak" identifies the relative flow peaking factor that would be permitted, on a tube by tube basis, before the stress ratio criteria would be exceedad. As can be observed in the tables anci figures, the inner row tubes have larger values of allowable relative flow peaking when compared to the outer rows. Il 9214M 1E-042789-146
'.9.3 ' Cumulative Fatigue Usage All tubes that are unsupported and have a stress ratio 1 1.0 have a maximum stress amplitude that is < 4.0 ksi (from 9.5 ksi) since a 10% reduction in the stability ratio for the North Anna Unit 1 R9C51 tube was the criteria basis. Stability ratios have been niculated for all the North Anna Unit 2 tubes using the three cases described earlier; Premod, Postmod Without Thot Reduction and Postmod With T Reduction. The tubes are not expected to hot rupture as a result of fatigue if:
- 1) they meet the stress ratio criteria of g 1.0 and 2) their current and future fatigue usage will total less than-1.0.
Determining acceptability of the North Anna Unit 2 tubes is complicated by the fee trat several tubes have stress ratios greater than 1.00 for the Dremod cea L ions, but have values loss than 1.00 for the Postmod conditions and inas j than or equal to 1.00 for the Postmed with T reduction conditions. hot . Acceptance of these-tubes must be determined by calculating the actual fatigue usage factors-for each tube on a case by case basis. Tubes with current and projected fatigus usage-factors less than 1.00 will be acceptable (with l respect to U-bend fatigue) and will not require preventive action or can he returned to service'if currently sentinel plugged. 1 Table 9-3 contains'a summary of fatigue usage factors for tubes that have - stress ratios near or greater than 1.00 (calculated using the more limiting Premod conditions and assuming the tubes became dented since the first i cycle). As can be observed in-the table, all tubes currently have fatigue usage factors less than 1.00. Future usage factors have been determined for operation under current operating conditions and for conditions where T hot reduction is implemented. Results are presented for both 40 years of total operation and for 10 more years of operation. These results indicate that, for a total of 40 years of operation, two tubes are at potential risk if T reduction is implemented. These two tubes (SG:A R9C60 and SG:B R9C35) hot currently have usage factors equal to 0.49 but will have projected fatigue usage factors greater than 1.00 after 40 years of total operation. Usage factors calculated after 10 more years of operation with T implemented hot have been determined to be 0.84. 9214M:1E-042789-147 =
-- l The results of the fatigue evaluation indicate that currently no tubes in the North Anna Unit 2 steam. generators require preventative action to preclude a North Anna Unit 1 R9C51 type tube rupture and that any tubes currently plugged with sentinel l plugs, to detect such a rupture, can be returned to service. However, two tubes previously identified, SG:A R9C60 and SG:B R9C35, will require preventive action in the future, to preclude such a rupture, after approximately 10 more years of service. Note that in the event of a future uprating.or increase in general plugging level the potential for tube fatigue l would need to be re-evaluated. 4 l i -l 9214M:1E-042789-148
Table 9 North Anna'f2 Tubes with Significant Flow Peaking or Relative Stability Ratio RELATIVE STABILITY RATIO
- RELATIVE FLOW PEAKING (Assumes all tubes are dented with deformation)
Relative Stability Ratio
- Re1 Flow Peak Postmod Postmod S.G.
Row Column Premod w/o T With'T hot hot A 8-64 0.810 0.720 0.755 1 9 11 0.648 0.577 0.607 35 0.851 0.757 0.796 40;55 0.687 0.611 0.643 60 0.961 0.854 0.899 79-84 0.654 0.581 0.612 8 8 31 0.608 0.540 0.568 60 0.817 0.726 0.764 81 0.583 0.518 0.546 9 34 0.680 0.605 0.636 35 0.960 0.854 0.899 40-52 0.687 0.611-0.643 92 0.637 0.566 0.596 93 0.462 0.411 0.433 C 9 35 0.930 0.828 0.870 '40-56 0.687 0.611 0.643-60 0.930 0.828 0.870 61-0.680 0.605 0.636 79-85 0.654 0.581 0.612 10 44 '0.797 0.709 0.746 45 0.798 0.709 0.747 49 0.798 0.709 0.747 50 0.798 0.709 0.747 60 0.798 0.709 0.747
- Tubes which are currently sentinel plugged which are recommended to remain
. sentinel plugged. l l 9214M:1E-0427ff 3.149 - ~
L Table 9-2 North Anna #2 Tubes with Significant Flow Peaking or Relative Stability Ratio STRESS RATIO 1 (Assumes all tubes are dented with deformation) Rel. Stress Ratio i Flow Max Allow Postmod Postmod S.G. Row Column Peaking Flow Peak Premod W/0 T With T hot hot ~ ~ A 8 64 0.68 0.36 0.47 a,e 9 11 0.16 0.09 0.11 35 0.73 0.38 0.50 40-55 0.22 0.12 0.16 60 1.61 0.75 1.00 79-84 0.17 0.09 0.11 8 8 31 0.14 0.07 0.10 60 0.72 0.37 0.49 81 0.11 0.06 0.08 i 9 34 0.21 0.11 0.15 35 1.61 0.75 1.00 i 40-52 0.22 0.12 0.16 92 0.15 0.08 0.10 93 0.03 0.01 0.02 C 9 35 1.21 0.63 0.83 l 40-56 0.22 0.12 0.16 I 60 1.21 0.63 0.83 61 0.21 0.11 0.15 79-85 0.17 0.09 0.12 10 44 0.43 0.22 0.30-45 0.43 0.22 0.30 49 0.43 0.23 0.30 50 0.43 0.23 0.30 60 0.43 0.23 0.30
- Tubes which are currently sentinel plugged which are recommended to remain sentinel plugged.
t - 9214M:1E-042789-150
i Table 9-3 Summary of North Anna Unit 2 Fatigue Usage Factors Usage Factor Usage Factor Current 40 Year Total Life 10 More Years Tube S.G. Usage w/o T With T w/ T With T hot hot hot hot R9C35 A 0.02 0.03 0.08 0.02 0.04 R9C60 A 0.49 0.75 1.68 0.57 0.84 i R9C35 B 0.49 0.75 1.68 0.57 0.84 R9C35 -C 0.17 0.29 0.69 0.21 0.32 R9C60 C 0.17 0.29-0.69 0.21 0.32 4 l l ' 9214M:1 E-042789-151
w ~ .NA12 - PREMODD RSR*REL : FLOW PEAKING: 8,C Figure 9-1 Relative Stability Ratto and Relative Flow Peaking - North Arma (Mit 2 Promod (Composite of all Steam Generators with Umbrella Flow Peaking) 9214M:1E-042789 I' L_ __ ___ _
- NA 2 - iPOSTMOD -ER'SR*REL4 FLOW : PEAKING
.8,C-' j 3 '~ Relative Stability Ratio and Relative Flow Peaktrg - Postumod Without T t Figure 9-2 Reduction (Cosmposite of alS SGs with unerella Flow Peaking) 9214M:1E-042789 I
r r N C, a 9, GN IKA t E oh T P h t ir t W do m O tso L P F 2 t i L nU E a m R A h tr R o N g S n i R kaeP w o l F ev T i ta O le R H !w fa T o i ta R y t i l 2 i ba t S A ev N i ta le R 3 9 eru 9 g 8 i 7 ~ F 2 4 0 E 1 M: 4 1 2 9
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- iNAR2i: - PREMODi JSTRESS : RATIO WITH1 DENTING
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