ML20023B302
| ML20023B302 | |
| Person / Time | |
|---|---|
| Site: | Vermont Yankee File:NorthStar Vermont Yankee icon.png |
| Issue date: | 11/30/1982 |
| From: | TELEDYNE ENGINEERING SERVICES |
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| References | |
| TR-5319-1, NUDOCS 8212270218 | |
| Download: ML20023B302 (140) | |
Text
A e
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J TELEDYNE ENGINEERING SERVICES
'A'THWE CONTROLLED ENGINEERING SERVICES DOCUMENT TES PROJ. NO._ d/
DATE_
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TECHNICAL REPORi-
.:3 TECilNICAL REPORT TR-5319-1 M ARK I CONTAINMENT PROGRAM 3
PLANT-UNIQUE ANALYSIS REPORT OF TIIE TORUS SUPPRESSION CIIAMBER FOlt
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VERMONT YANKEE NUCLEAR POWER STATION NOVEMBER 30, 1982 REVISION 0 R A OC 0
0
J O
YANKEE ATOMIC ELECTRIC COMPANY 1671 WORCESTER ROAD FRAMINGHAM, MA 01701 O
O TECHNICAL REPORT TR-5319-1
'O MARK 1 CONTAINMENT PROGRAM PLANT-UNIQUE ANA'.YSIS REPORT OF THE O
TORUS SUPPRESSION CHAMBER FOR VERMONT YANKEE NUCLEAR POWER STATION O
O NOVEMBER 30, 1982 REVISION 0
'O O
WTELEDYNE ENGINEERING SERVICES 130 SECOND AVENUE WALTHAM, MASSACHUSETTS 02254
- O 617-890-3350 l
l
Technical Report TR-5319-1
-ii-N I
ABSTRACT
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The work summarized in this report was undertaken as a part of the Mark 1 Containment Long-Term Program. It includes a summary of the analysis that was performed, the results of the analysis and a description of 19 significant modifications that were made to the structure and internals to increase safety margins.
In all cases, the stresses reported in this document meet the allowable
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levels as defined in the structural acceptance criteria (Reference 3).
The methods and assumptions used in this analysis are in accordance with USNRC NUREG 0661 (Reference 2), except as noted in the text.
The modifications described in this report are also in compliance with NUREG 0661, unless otherwise noted.
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Technical Report TN TR-5319-1
-11i-ENG4EERING SERVICES TABLE OF CONTENTS
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Page ABSTRACT ii I
1.0 INTRODUCTION
& GENERAL INFORMATION 1
2.0 PLANT DESCRIPTION 4
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2.1 General Description 4
2.2 Recent Modifications 4
2.2.1 Modifications to Reduce Loads 5
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2.2.2 Modifications to Strengthen Structure 8
3.0 CONTAINMENT STRUCTURE ANALYSIS -
30 SHELL & EXTERNAL SUPPORT SYSTEM
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3.1 Computer Models 30 3.2 Load Analysis 31 3.2.1 Pool Swell Loads 31
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3.2.2 Condensation Oscillation - DBA 32 3.2.3 Chugging 33 3.2.3.1 Pre-Chug & IBA C0 33 3.2.3.2 Post Chug 33
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3.2.4 SRV Discharge 34 3.2.5 Deadweight, Thermal & Pressure 35 3.2.6 Seismic 35 3.2.7 Fatigue Analysis 36
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3.3 Results and Evaluation 37 3.3.1 Shell 38 3.3.2 Support Columns & Attachments 39 3.3.3 Support Saddles & Shell Weld 41
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3.3.4 Earthquake Restraints & Attachments 42 3.3.5 Anchor (Tie-Down) System 43
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Technical Report YM TR-5319-1
-iv-6M TABLE OF CONTENTS (CONTINUED)
Page 4.0 VENT HEADER SYSTEM 54 4.1 Structural Elements Considered 54 4.2 Computer Models.
54 4.3 Loads Analysis 56 4.3.1 Pool Swell Loads 56 4.3.1.1 Pool Swell Water Impact 56 4.3.1.2 Pool Swell Thrust 57 4.3.1.3 Pool Swell Drag (Support Columns Only) 57 4.3.2 Chugging Loads 58 4.3.2.1 Downcomer Lateral Loads 58 4.3.2.2 Synchronized Lateral Loads 58 4.3.2.3 Internal Pressure 58 4.3.2.4 Submerged Drag 59 4.3.3 Condensation Oscillation - DBA 60 4.3.3.1 Downcomer Dynamic Load 60 4.3.3.2 Vent System Loads 61 4.3.3.3 Thrust Forces 61 4.3.3.4 Submerged Drag (Support Columns) 61
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4.3.4 Condensation Oscillation - IBA 61 4.3.5 SRV Loads 62 4.3.5.1 SRV Drag Loads 62 4.3.6 Other Loads - Weight, Seismic, & Thermal 62 4.4 Results and Evaluation 62 4.4.1 Vent Header-Downcomer Intersection 63 4.4.2 Vent Header-Main Vent Intersection 63 4.4.3 Vent Header Support Columns & Attachments 64 4.4.4 Downcomer Tie Bars & Attachments 64 4.4.5 Vent Header Deflector & Attachments 65 4.4.6 Main Vent /Drywell Intersection 65 4.4.7 Vent Header, Main Vent, & Downcomers -
66 3
Free Shell Stresses 4.4.8 Vent Pipe - Mitre Joint 66 4.4.9 Fatigue Evaluation 67 D
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Technical Report yg TR-5319-1
-v-g gg TABLE OF CONTENTS (CONTINUED)
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Page 5.0 -RING GIRDER ANALYSIS 75
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5.1 Structural Elements Considered 75 5.2 Computer Models 75 5.3 Loads Analysis 76 5.3.1 Loads Applied to the Shell 76
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5.3.2 Drag Loads 77 5.3.3 Loads Due to Attached Structure 78 5.4 Results & Evaluation 79 5.4.1 Ring Girder' Web & Flange 79
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5.4.2 Attachment Weld to Shell 79 6.0 TEE-QUENCHER & SUPPORT 84 6.1 Structural Elements Considered 84 l
6.2 Computer Models 85 6.3 Loads Analysis 85 6.3.1 SRV Loads 85
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6.3.2 Pool Swell Loads 85 6.3.3 Chugging Loads 86 j
6.3.4 Condensation Oscilliation Loads 86 6.3.5 Other Loads 87 6.4 Results & Evaluation 87 6.4.1 Tee-Quencher Structure 87 6.4.2 Submerged SRV Line 87 6.4.3 Tee-Quencher Support 88 7.0 OTHER STRUCTURES 90 7.1 Catwalk 90 7.1.1 Computer Models 90
Technical R:: port TN TR-3319-1
-vi-6 MS TABLE OF CONTENTS (CONTINUED)
O Page 7.1.2 Loads Analysis 90 7.1.2.1 Pool Swell Water Impact 90 g
7.1.2.2 Pool Swell Fallback 91 7.1.2.3 Froth Loads 91 7.1.2.4 Drag Loads (Support Columns) 91 7.1.2.5 Weight & Seismic 92 g
7.1.3 Results & Evaluation 92 7.1.3.1 Main Frame 92 7.1.3.2 Support Columns & End Joints 93 7.1.3.3 Welds to Ring Girder 93 g
7.2 Internal Spray Header 94 7.2.1 Computer Model 94 7.2.2 Loads Analysis 94 7.2.2.1 Froth Load 95 O
7.2.2.2 Weight, Seismic & Ring Girder Motion 95 7.3 Vent Pipe Bellows 96 7.3.1 Analysis Method 96 7.3.2 Loads Considered 97 0
7.3.3 Results & Evaluation 97 7.4 Monorail 98 7.4.1 Computer Model 98 7.4.2 Loads Analysis 98 O
7.4.2.1 Froth loads 98 7.4.2.2 Weight & Seismic 99 7.4.3 Results & Evaluation 99 0
8.0 SUPPRESSION POOL TEMPERATURE EVALUATION 105 8.1 Maximum Pool Temperature Analysis 105 8.2 Pool Temperature Monitoring 106 O
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Technical Report TR-5319-1
-vii-MM
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TABLE OF CONTENTS (CONTINUED)
Page
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REFERENCES 108 APPENDIX 1 - USE OF SRV TEST DATA IN ANALYSIS Al-1
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APPENDIX 2 - USE OF 32 liZ CUf0FF FOR A2-1 C.0. & POST CHUG ANALYSIS APPENDIX 3 - SUBMERGED STRUCTURE DRAG ANALYSIS FOR A3-1
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THE RING GIRDER
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Technical Report TME TR-5319-1
-viii-ENGPE954G SERVCES FIGURES AND TABLES
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Page FIGURES:
2-1 Torus Plan View 11 D
2o2 Torus Composite Cross Section 12 2-3 Torus Modifications - Cross Section at Ring Girder 14 l
2-4 Torus Modifications - Cross Section at Mid Bay 15 3'
2-5 AP Pressurization System 16 2-6 Vent Header Defiector 17 l
2-7 Vent Header Deflector Att3chment 18 2-8 SRV Tee-Quencher and Support 19 2-9 Pool Temperature Monitoring System 20 2-10 RHR Return Line Elbow and Support 21 2-11 Torus Support Saddles and Saddle Anchors 22 2-12 Torus Support Column Anchors 23 2-13 Downcomer Tie Rod "1d Gustet Modification 24 0
2-14 Catwalk and Handrail Modification 25 2-15 Catwalk and Handrail Modification 26 l
2-16 Torus Sp ay Header Support Modifications 27 lm j
2-17 Monorail 28 l
2-18 Thermocouple Locations 29 l
l 3-1 Detailed Torus Shell Model 44 lO l
3-2 Detailed Torus Shell Model 45 l
3-3 Detailed Torus Shell Model 46 3-4 Torus Beam Model (360 )
47
- O 3-5 Pool Swell - Net Vertical Load - Average Submerged 48 l
3-6 Pool Swell - Average Submerged Pressure 49 l
l 3-7 Pool Swell - Torus Air Pressure 50 lO l
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Technical R: port TN TR-5319-1
-ix-MM FIGURES AND TABLES (CONTINUED)
Page 3-8 SRV Shell Pressure - Typical 51 3-9 Location of Maximum Shell Stress 52 3-10 Earthquake Restraint System 53 4-1 Detailed Vent Header Model 69 4-2 Detailed Vent Header Model 70
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4-3 Detailed Vent Header Model 71 4-4 Vent Header Beam Model 72 4-5 Vent Header Deflector Analysis 73 4-6 Chugging Cases - Synchronized Lateral Loads 74 5-1 Ring Girder 80 52 Detailed Shell - Ring Girder Model 81 5-3 Detailed Shell - Ring Girder Model 82 l
l 5-4 Detailed Ring Girder - Shell Model - Ring Girder Elements 83 6-1 SRV Line Analytical Model 89 7-1 Catwalk Computer Model (Unmodified) 100 l
l 7-2 Catwalk Computer Model (Modified) 101 i
7-3 Spray Header Computer Model 102
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7-4 Vent Pipe Bellows Motion 103 7-5 Monorail Computer Model 104 8-1 Bulk Suppression Pool Temperature vs. Quencher Mass Flux 107 3
Al-1 SRV Test Instrumentation - Shell Al-6 Al-2 SRV Test Instrumentation - Columns Al-7
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Al-3 SRV Test Instrumentation - Tee-Quencher Al-8 Al-4 SRV Test Instrumentation - Internal Structures Al-9 Al-5 SRV Drag Pressures Al-10 D
D Technical Report
'A'TF1 Fr?(NE TR-5319-1
-x-ENGNEERING SERVICES D
FIGURES AND TABLE.S (CONTINUED)
Page TABLES:
D 1.
Structural Acceptance Criteria for Class MC 110 Internal Structures 2.
Plant Physical Dimensions 111 D
3.
Plant Analysis Information 112 D
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Technical Report TF ME TR-5319-1 ENGNEBtNG WICES
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1.0 GENERAL INFORMATION The purpose of the Mark 1 Torus Program is to evaluate the effects of hydrodynamic loads resulting from a loss of coolant accident and/or an SRV 3
discharge on the torus structure.
Tr.is report summarizes the results of extensive analysis on the Vermont Yankee torus structure and reports safety margins against established criteria. The content of this report deals with the torus shell, external support system, vent header system and internal D
structures. Analysis and results for piping attached to the torus (including shell penetrations and internal piping), for the SRV line (except for the submerged portion and tee-quencher) and for the SRV line vent pipe penetration will be presented in a separate piping report, TR-5319-2.
The criteria used to evaluate the torus structure is the ASME Boiler &
Pressure Vessel Code,Section III, Division 1, with addenda through Summer 1977 (Reference 11) and Code Case N-197.
Modifications were done under O
Section XI of the ASME Code and meet the Summer 1978 Edition of Section III for design, materials and fabrication.
l A great many technical reports have been written and issued as a part of O
this program.
There reports provide detailed descriptions of the phenomena, the physics cc ' *o iling the phenomena, calculational methods and detailed procedu,
a ant-unique load calculations. Several of these documents are listed < r ;_ "
es in this report. The approach of this report will be to LO reference these documents, wherever possible, and to avoid a re-statement of the same information.
A major part of this procrxn has dealt with providing plant-unique load O
calculation procedures (References 9, Volume 1-10 are examples of this).
In most cases, the loads used to support the analysis were calculated in strict accordance with those procedures, as amended by NUREG 0661 (Reference 2).
In some cases, optional methods have been used; these methods are specifically 0
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1 Technical Report TN TR-5319-1 N NES I
referenced in Program documentation. Exemples of these are the use of plant-unique SRV test data to calibrate SRV analysis, and use of plant unique quarter scale pool swell movies to refine certain water impact and froth loads. In a few cases, analysis assumptions have been made that do not appear in Program documentation; these are identified in the text.
Extensive structural analysis was performed as a part of this evalua-tion.
The major analysis was for dynamic response to' time-varying loads.
Analysis for static and thermal conditions also form a part of this work. The computer code used to perform almost all of this analysis was the STARDYNE code, as marketed by Control Data Corporation. STARDYNE is a fully verified and accepted code in this industry; details of the code are available through
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CDC. Cases where a computer code other than STARDYNE is used will be identi-fied in the text. All dynamic analysis used damping equal to 2% of critical, unless stated otherwise.
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As an aid in processing the large amounts of calculated data, post-processors for the STARDYNE program were written and used.
These programs were limited in function to data format manipulations and simple combinations of load or stress data; no difficult computational methods were included.
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The loads and load combinations considered in this program required special cor. sideration to determine the appropriate levels of ASME Code appli-cation.
Reference 3 was developed to provide this standard.
Table 5-1 of
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Reference 3 is the basis for all the evaluation work in this report; it is reproduced in this report as Table 1.
This table shows 27 load combinations that must be considered for each structure.
The number actually becomes several times that when we consider the many different values associated with
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various SRV discharge conditions. The approach used in the final evaluation of structures is to reduce this large number to relatively small number of cases by conservative bounding.
For example, load combinations including SSE
- seismic, have a
higher allowable than the same combination
Technical Report TN TR-5319-1 MM with OBE seismic.
For these cases, our first evaluation attempt is to con-sider the SSE combination against the OBE allowables.
If this produces an acceptable result, those numbers are reported as final.
This procedure re-sults in many cases where safety margins are understated; this is the case for most of the results.
As an aid in correlating discussion of particular load analyses to detailed program documentation, most analysis described in this report has been referenced directly to a paragraph in the Lcad Definition Report (Refer-ence 1). This has been done by identifyiag the applicable'LDR paragraph in parenthesis immediately following the title of the load.
T;ais referencing directs the readcr to a acre detailed description of the load than can be included in this report.
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Technical Report TF WE TR-5319-1 ENGNEBW4G SERVICES 2.0 PLANT DESCRIPTION 2.1 General Arrangement
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The configuration of the Vermont Yankee torus structure is shown in Figures 2-1 and 2-2.
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Figure 2-1 shows a plan view of the torus.
It is made up of the sixteen (16) mitred sections, connected to the drywell by eight (8) equally
~ spaced vent pipes.
It is supported by two external columns and an inter-mediate saddle at each of sixteen places, as shown. The columns and saddles
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are connected to the basemat with anchor bolts. Four earthquake restraints, spaced equally around the torus, connect the belly of the torus to the basemat (Figure 3-10).
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Figure 2-2 shows some of the inside arrangement. Ring girders reenforce the outer shell at each of the sixteen planes defined by the external support system.
The vent header system is supported off of the ring girders and is directly connected to the drywell via the vent pipes. The oper,ing where the
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vent pipe oenetrates the torus shell is sealed by a bellows. The ring girder also supports the catwalk, spray header, SRV tee-quencher support, and var-ious interr:al piping runs.
Figures 2-3 to 2-17 show several details of the torus structure. Table 2.0 lists several of the plant specific dimensions.
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2.2 Recent Modifications Many modifications have been made at Vermont Yankee over the past
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several years and in addition several more will be made during the 1983 refueling outage to the torus, both to increase its strength and also to mitigate the hydrodynamic loads. The modifications are illustrated and listed in the composite sections of Figures 2-3 and 2-4, along with their installa-
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tion dates.
A description and illustration of each individual modification follows:
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D ENGNEERNG SEm1CES
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2.2.1 Modifications to Reduce Hydrodynamic Loads Drywell Pressurization System (AP System)
D Installation of a system to maintain a pressure differential l
between torus and drywell was the first modification of this Program.
The system is illustrated in Figure 2-5.
It is designed to maintain a minimum j
positive pressure difference of 1.7 psi between the vent system (drywell) and the airspace inside the torus. The result of this pressure difference (A P) l is to depress the water leg in the downcomers and reduce the water slug that must be cleared, if rapid p-essurization of the drywell occurs. Early generic testing in the Program demonstrated that this was an effective means to reduce j
shell pressures related to DBA pool swell.
The 1.7 psi pressure difference was selected as the basis for the Vermont Yankee plant unique quarter-scale pool swell tests and is intended to be the normal operating condition of the j
plant. As illustrated in Figure 2-5, pressure differential is maintained by l
using the nitrogen inerting system to pressurize the drywell to 1.7 psi; the torus remains at ambient pressure.
Other methods are also available to maintainAP.
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Vent Header Deflector The vent header deflector at Vermont Yankee is illustrated j
in Figures 2-6 and 2-7.
It is a 16-inch schedule 120 pipe with -inch plate welded to the sides.
The deflector extends under the belly of the vent header to j
protect the vent header from direct water impact during pool swell.
It does this by shadowing the most sensitive part of the vent header and by separating and diverting the rising pool before it can reach the vent header.
This deflector was included in the plant unique pool swell tests for Vermont Yankee O
pr vide accurate vent header loading for detailed analysis.
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Technical Report TE WE TR-5319-1 N N N ES
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SRV Tee-Quencher A tee-quencher has been installed at the dis @Sa end of each main steam relief line to replace the existing ramsheads.
The quencher
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and its support is illustrated in Figure 2-8.
The quencher serves to divide the SRV discharge bubble into hundreds of smaller bubbles and to distribute them over an entire bay. This division and distribution of SRV discharge has been shown in generi.c testing to reduce torus shell pressure by f actors of two
)
or more when compared to ramshead pressures.
The plant-unique character-istics of these devices at Vermont Yankee were determined by in-plant testing after their installation.
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The quencher support is also illustrated in Figure 2-8.
It is a 20-inch schedule 120 pipe welded to the ring girder, as shown.
Temperature Monitoring System & RHR Return Lines
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The addition of a pool temperature monitoring system and an elbow to the discharge end of the RHR return lines are both intended to assure l
proper operation of the SRV quencher. These modifications are illustrated in h
Figures 2-9, 2-10 and 2-18.
The temperature monitoring system senses pool temperature through ten thermocouples installed in the four bays of the suppression pool
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which are directly affected by relief valve discharge, plus one additional bay for complete coverage (See Figure 2-18).
Five thermocouples are used to provide individual indication of temperatures in each of the five bays, while l
the other five thermocouples are combined to provide an average or bulk h
temperature indication of the entire suppression pool. The thermocouples are the dual-element type and are protected by a stainless steel tube.
As a result of modifications to the catwalk, the thermocouple assemblies will be remounted on vertical support members. The vertical support members will be
)
off the ring girder with the end of the protection tube at the same elevation as the tee-quencher centerline.
A Westronics multi-point strip chart
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Y Technical Report TN TR-5319-1 F M $$\\/ ICES recorder is installed in the control room. The recorder is used to record the
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five individual measurements as well as the bulk temperature of the suppres-sion pool. The recorder provides a history of the temperatures as well as a means for the operator to observe temperature trends.
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The elbows on the RHR return lines were added to improve pool circulation during periods of extended SRV blowdown. Circulation of the pool with these lines assures that local-to-bulk temperature differences will be minimized and that SRV quencher performance will be maintained for the maximum possible time during extended discharge.
These two RHR return lines were l
further modified by re-routing them to the ring girders.
The ring girders react drag loads on these lines and also provide for reactions due to elbow discharge loads.
Additional SRV Vacuum Breakers Each of the four SRV discharge lines at Vermont Yankee has been fitted with a ten-inch vacuum breaker. A second ten-inch vacuum breaker will be added during the 1983 refueling outage. This modification minimizes the temporary formation of the high water leg in the SRV line which could occur after an initial actuation; and thereby prevents the high clearing loads which could occur if a second actuation occurred at that time. The location of these devices is different on each SRV line due to space limitations and is not illustrated. Analyses are on-going to optimize the location of a second set of vacuum breakers in the SRV lines.
Removal of Submerged Piping Some of the piping inside the torus extended to greater depths than was necessary for its proper functioning. This additional submer-gence resulted in drag loads on the piping that was unnecessary. In order to eliminate this unnecessary load, a piping system was cut off to provide a three foot submergence at minimum torus water level. The line affected is the
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RCIC turbine exhaust.
In addition, the vent drain lines were cut off and capped above the pool.
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SFTELEDVNE 1 1 2.2.2 Modifications to Strengthen the Structure Torus Support Saddles and Anchor Bolts Support saddles were added under e:ch ring girder as shown in Figure 2-11.
The saddles, support columns and ring girder all lie in the same plane and react all vertical loads on the torus - most of the load is reacted by the saddle.
t The saddle is constructed of 1 -inch type SA 516 GR 70 steel plates, welded to the torus shell and resting on the concrete basemat.
It is restrained from upward motion by six pairs of two-inch Williams rock bolts, set 24-inches into the basemat.
The anchor bolt restraints are set with a l
small clearance to allow for normal radial growth of the torus due to tempera-ture changes.
Torus Support Column Anchors l
The uplift capacity of the torus was also increased by the
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addition of anchors as illustrated in Figure 2-12.
These anchor restraints
/
were installed on each of the 32 torus support columns. The designs for inner and outer columns are slightly different and have slightly different capa-cities, but the illustration in Figure 2-12 represents both sets of locations.
The anchor into the basemat is made by 12, 1 -inch diameter anchor bolts at i
cach column location. The bolts are set eight inches into the concrete.
1 Downcomer Tie-Rods and Vent Header Gussets
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The downcomer tie-rods and vent header gussets are illus-trated in Figure 2-13.
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J
Technical Report WP WNE TR-5319-1 ENGNEERNG SERVICES
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The tie-rods are constructed from 2 -inch schedule 40 pipe and provide greatly increased capacity to downcomer lateral loads than the original tie bars.
They are attached to the downcomers with specially fabricated 24-inch pipe clamps, constructed from 3/4 inch steel. The clamps
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are prevented from sliding on the downcomers by welded steel clips both above and below the clamp.
The gussets between the downcomers and vent header are to
)
reduce local intersection stresses due to chugging lateral loads on the down-comers.
They are constructed from -inch thick SA 516 GR 70 steel plate, and will be welded to the vent header and downcomers.
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Catwalk and Handrail The catwalk and handrail at Vermont Yankee requires substan-tial modification, as illustrated in Figures 2-14 and 2-15. The modifications
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will include the following:
e Replacement of original support columns with four-inch diameter pipe columns.
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e Addition of lateral supports.
i These changes are scheduled for the 1983 outage. An alter-
)
native design is being considered which will remove 14 of the 16 catwalk sections.
Internal Spray Header Supports
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The spray header piping inside the torus is hung from the ring girder at the top of the torus, as illustrated in Figure 2-3.
The original supports for this line were a "U" shaped rest support which could not
)
react upward loads. These were modified as shown in Figure 2-16 to react the
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upward loads associated with pool swell and froth.
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D
3 Technical Report TN TR-5319-1 MM 3
Drywell-to-Wetwell Vacuum Breakers New aluminum discs with higher strength and better impact 3
characteristics were installed in the drywell-to-wetwell vacuum breakers.
This change was the result of inforrration found in the Mark 1 Full Scale Test Facility during the testing.
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WTEUIDGE Technical Report ENCBEiBMGSSMCES TR-5319-1
_11_
i EARTHQUAKETIgS (4) PLACES 90 EXTERN AL SUPPORT l
COLUMNS l
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-SADDLES l 49'
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FI G. 2-1 TORUS PLAN VIEW
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Technical Report WM J
TR-5319-1 NMRVICES 9
VACUUM BREAKER LINE
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2 3 = MODIFIC ATIONS o
FIG. 2-2 a
TORUS COMPOSITE CROSS SECTION-VERMONT YANKEE
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Technical Report TN TR-5319-1 N SBWICES KEY FOR FIGURES 2-3 AND 2-4 Modification Completion Date l.
Column Reinforcement and Tie-Down 1977 2.
Mitred Joint Saddle 11/80 3.
Downcomer Ties 11/80 4.
Vent Header Deflector 11/80 j
5.
Vent Header /Downcomer Stiffening 5/83 6.
Vent Drain Line 11/80 7.
Monorail 10/79 l
8.
Catwalk Grating and Handrail 5/83 l
9.
Drywell/Wetwell P Control 1976 10.
Safety Relief Valve Vacuum Breakers (Drywell) 5/83 11.
SRV Tee-Quencher and Supports 11/80 12.
Add RHR Return Line Elbow 11/80 l
13.
RCIC Turbine Exhaust Truncated 11/80 1
14.
Temperature Monitoring System 5/83 l
15.
Modification of Drywell to Wetwell Vacuum Breakers 11/80 16.
Saddle Anchor Bolts 11/80 D
9 Technical Report M
TR-5319-1 )
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TORUS MODIFICATIONS-CROSS SECTION AT RING GIRDER
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TR-5319-1 D D
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TORUS MODIFIC ATIONS-CROSS SECTION AT MID BAY
I NEMCES Technical Report TR-5319-1 )
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1.7 PSID - PRESSURE CURRENTLY MAINTAINED BY N INERTING SYSTEM NOMINAL 2
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TGchnical Report TM TR-5319-1 NM
)
3.0 CONIAINMENT STRUCTURE ANALYSIS - OUTER SHELL & EXTERNAL SUPPORT SYSTEM (INCLUDING ANCHORS)
The containment structure section of this report includes the analysis
)
and evaluation of the following structures:
Torus Shell Support Columns
)
Column-To-Torus Weld Support Saddles Saddle-To-Torus Weld Earthquake Restraints & Attachment
)
Anchor (tie-down) System 3.1 Computer Models
}
Analysis of the contaiiment structures was accomplished using the computer models shown in Figures 3-2 to Figure 3-4.
The detailed shell model shown in Figure 3-1 was used to calculate the effects of all loads on shell I
stress, as well as all sp metric loads on the support and anchor system. The
)
beam model shown ir. Figsre 3-4 was used to determine the effects of asymmetric loads on the support system.
Asymmetric loads on the torus structure are horizontal earthquake, SRV and chugging.
Evaluation of the support system considered the combined effect of symmetric and asymmetric loads in accord-
)
ance with the load combination table (Table 1).
l The detailed finite element model shown in Figure 3-1 simulates j
one-half of the non-vent bay.
It h bounded by the ring girder on one end and the mid-bay point on the other.
The vent header system is assumed to be i
dynTmically uncoupled from the shell by the support saddles and is not included in this model. This model wet, constructed with the assumption that the small offset that esists between the ring girder and mitre joint will not
)
affect results; accordingly, the offset is not included in the model.
)
)
Technical Report WM TR-5319-1 6M
)
This model includes 587 structural nodes, 664 plate elements, 2261 static degrees of freedom and 362 dynamic degrees of freedom.
Symmetric boundary conditions were used at both ends of the model.
)
The model was modified for various load calculations to account for differences in the percent of the water mass that is effective for that load event. In all cases., modeling of the water mass was accomplished using a 3-D virtual mass simulation os an integral part of the structural analysis. The
)
percent of water mass used is identified in the discussion of each load calculation that follows.
l U
The 360 beam model of the torus is shown in Figure 3-4.
This model
)
was used to evaluate the effects of lateral loads on the support system and earthquake restraint system.
The beam element properties were selected to simulate combined bending and shear stiffness of the sections.
Water mass was lumped with the structure weight on the wetted nodes.
)
3.2 Loads Analysis 3.2.1 Pool Swell Loads (4.3.1 & 4.3.2)
)
Analysis for pool swell loads was done using the finite element model shown in Figure 3-1.
This was a dynamic analysis performed in the time domain by applying a force-time history, to simulate the pressure-
)
time histories of the pool swell event to each node on the computer model.
Input pressure-time histories were varied in both the longitudinal and radial directions in accordance with the information in References 1, 2, 9 and 10.
Typical pressure-time history curves are shown in Figures 3-5 through 3-7.
3 (These pressure-time histories are taken directly from Reference 10, before adjustment, as required by Reference 2.
Therefore, the amplitudes shown are slightly different than the loads used in the analysis).
)
The computer analysis was run for two different pool swell conditions, fullAP and zeroAP. Figures 3-5 through 3-7 show comparative values
)
?
Technical Report TN TR-5319-1 N SBtVICES
)
and time histories for the two cases. The only difference between the analy-ses was the input loads; the models were identical. Details of the full load distribution can be found in References 1, 9 and 10.
h Plant-unique quarter scale pool swell tests showed that the effective water mass was less than 100% af ter bubble breakthrough and was slightly different for both zero and full AP conditions (Reference 4). The water mass used in the computer simulation was constant throughout the analy-sis and was set at the average of the two reduced masses identified in the quarter scale tests. The reduced and average mass values are given in Table 3.
This simplification in water mass analysis is consistent with the rela-tively slow (pseudo-static) nature of the pool swell load.
This simplifi-cation only affects the inertial (frequency) calculation; the effects of weight are accurately calculated for each load and time in the deadweight analysis.
)
3.2.2 Condensation Oscillation - DBA (4.4.1) l Analysis for condensation oscillation (CO) was also done with the structural model shown in Figure 3-1.
The condensation oscillation shell load is specified as a spectrum of pressures in 1 Hz bands (Reference 1). The analysis for this load was performed by considering the effects of unit loads at each load frequency (harmonic analysis) and then scaling and combining the individual frequency l
effects to determine total stress at selected elements. The three variations in the C0 spectrum (Reference 1) were evaluated bj re-scaling the resJlts of the unit load analysis. 100% of the water mass was used for all C0 analysis.
A plant-unique f actor was applied to the nominal condensation oscillation pressures as discussed in Reference 1; the factor is listed in Table 3.
)
The combination of individual harmonic stresses into total element stress was done by considering frequency contributions at 31 Hz and
)
)
W TELECWNE NEM MES 1 )
below.
The actual combination was done by adding the absolute value of the four highest harmonic contributors to the SRSS combination of the others for shell stress.
Loads on the support and anchor system were determined by
)
adding the absolute value of the three highest harmonic contributors to tne SR$S of the others. These combination methods and use of the 31 Hz cutoff are the result of extensive numerical evaluation of full scale test data, which is l
reported and discussed in References 6, 14 and Appendix 2 of this report.
Y 3.2.3 Chugging l
3.2.3.1 Pre-Chugging & IBA/C0 (4.5.1.2 & 4.4.1) i The pre-chug load was evaluated for both the sym-
'etric and asymmetric distribution described in Reference 1.
Results for the symmetric pre-chug analysis were also used for IBA/C0 as described in para-p graph 4.4 of Reference 1.
l l
l Results for symmetric pre-chug were developed l
directly from the unit-load harmonic analysis done for C0.
The results of that analysis were scaled to two psi (the pre-chug pressure) and all frequen-cies in the pre-chug range were scanned to determine the highest possible l
stresses.
)
Analysis for asymmetric pre-chug was performed using the beam model in Figure 3-4 by applying the unbalanced lateral load through the prescribed frequency range.
3.2.3.2 Post Chugging (4.5.1.2) l Post chugging is defined as a spectral load across a wide band, similar in nature to the C0, but much lower in amplitude.
)
Analysis done on one of the TES plants produced very low stresses and loads that were bounded by pre-chug values.
The analyses for pre-and post chug produced these results for maximum shell stres3:
Tcchnical Report TN TR-5319-1 MM Maximum Shell Stress
)
Shell Membrane Stress Pre-Chug 1284 psi l
Post Chug 776 psi 1.
Based on frequencies to 30 Hz - sum of 4 maximum +
SRSS of others.
Additional work published in Reference 12 showed that pre-chug bounded post chug (to 50 Hz) for column and saddle loads (Table
)
5-1, Ref. 12).
It also showed that PL+Pb stress due to post chug exceeded l
pre-chug by 53%.
TES analysis for post chug used the pre-chug stress values. The pre-chug stress may be increased by 53% to account for the 30 to 50 Hz contribution and they will still meet allowable stress.
l No further post chug analysis was done for the shell. This position was also influenced by the f act that post chug stresses were very small.
3.2.4 SRV Discharge l
Calculation of stresses due to SRV line d9 charge pressures, I
were also done using the finite element model in Figure 3-1.
The loading function used for this analysis was based on data collected from in-plant SRV
)
testing in this facility.
Testing was done in general accordance with the guidelines given in Reference 2.
In these tests, pressure amplitude and frequency were recorded and compared to calculated values for the test condi-tions. Factors were developed that related test to calculated values for both amplitude and frequency (see Appendix 1). These factors were then applied to calculated load values for other SRV conditions; the structural analyses was D
Technical Report WM TR-5319-1 g performed using these adjusted values. Appendix 1 discusses thc in-plant test and analysis in more detail. A typical set of SRV shell pressures is shown in Figure 3-8.
The method of modeling the water mass in the SRV computer model was the subject of extensive study in this program. Initial attempts to reproduce measured stresses by applying measured pressures to the computer models f ailed badly.
Af ter considerable study of the nature of the SRV phenomena itself, and the differences betweer. it, and the pool swell related loads, it appeared that a dry structure analysis should produce acceptable correlation. The method was tested and correlation of calculated-to-measured shell stress was excellent.
The dry structure analysis method was subse-quently used as a basis for all SRV analysis.
l 3.2.5 Deadweight, Thermal & Internal Pressure Deadweight, thermal and internal pressure analyses were done using the computer model shown in Figure 3-1.
Resulting stresses were calcu-lated and considered for all elements on the model.
I For the thermal analysis, conduction into the columns and saddles from the torus was considered.
Convection from the columns and l
saddles to ambient produced a calculated temperature gradient in these ele-ments.
The torus water, internals and shell were all assumed at the same
)
temperature.
l 3.2.6 Seismic 1
h i
Seismic analysis for shell stress was done by applying sta-tic G levels to the model in Figure 3-1.
Load orientation and values were adjusted for vertical and horizontal earthquakes in accordance with Table 3.
)
J
)
Technical Report TN TR-5319-1 N SBi\\/ ICES
')
The effects of lateral seismic loads on the support system were determined using the model in Figure 3-4.
The effective water mass used in this (lateral) analysis was adjusted in line with test results which showed that net dynamic reaction loads due to the water mass were substantially less
)
than those obtained from static application of the seismic acceleration.
A discussion of this fact can be found in Reference 7; the effective water mass used can be found in Table 3 of this report.
)
3.2.7 Fatigue Analysis Fatigue analysis of the torus shell and external support system is described here. Analysis of 1he shell at piping penetrations will
)
be described in TES report TR-5319-2, when the piping analysis is complete.
The f atigue analysis of the shell and support system was a conservative one which was based on applying a stress concentration factor of
)
4.0 on the most highly stressed elements for each load case.
In the case of the support system, only the column-to-torus and saddle web-to-torus welds were considered, since they have higher stresses than the rest of the support system. The process is conservative because:
)
It applies the maximum stress concentration (4.0),
4 recognized by 2ection III of the ASME Code to all elements (Reference 11).
)
and e
It adds the maximum absolute stress for each load case even though they do not usually occur at the same element.
)
The procedure used in this analysis consists of the follow-ing steps.
)
1.
For a given load, locate the maximum component-level stresses (S ' S ' 3xy) for the free shell, x
y local shell, and the supports.
)
)
Technical Report TN TR-5319-1 N SBi\\/ ICES
)
2.
For these locations, establish the stress intensity ranges and the approximate number of cycles.
3.
Repeat the process for all other loads in the load
)
combination.
4.
Add the stress ranges for all loads, independent of sign.
)
5.
Multiply these total stress ranges by 4.0 (the SIF).
)
6.
Calculate the alternating stress intensity and com-plete the f atigue analysis in compliance with Ref-erence 11.
)
Fatigue analysis resulting from chugging was done assuming that the operator would depressurize the system within 15 minutes after the chugging begins.
Plant procedures are presently under study to provide for this action.
)
3.3 Results and Evaluation Results are reported for each structural element of the containment
)
system for the controlling load combination.
Controlling load combinations are the ones that produce the smallest margins against the allowable stress -
not necessarily the highest stress.
)
All load combinations listed in Table 1 have been considered. As stated previously, most results include some level of bounding analysis and, therefore, understate the margins which actually exist.
)
t
)
)
Technical Report W F W NE TR-5319-1 ENGNEERNG SERVICES
)
3.3.1 Torus Shell Results of shell stress due to individually applied loads
)
were calculated and maintained on a component stress level until all the load combinations were formed. Stress intensities were then calculated from these total component-level values.
)
)
The controlling load combination for the shell at Vermont Yankee is case 20 in Table 1, which is:
OBA.C0 + Seismic (SSE) + Pressure + Weight
)
This load combination controls all categories of shell stress, although the location of the elements is different for the different types of stress.
The following table summarizes the controlling stresses.
)
Approximate locations of the controlling stresses are shown in Figure 3-9.
CONTROLLING SHELL STRESSES - VERMONT YANKEE
)
TYPE OF ACTUAL ALLOWABLE STRESS LOCATION STRESS STRESS Local (Pm)
Free Shell 12,957 psi 19,300 psi
)
Element 17 Local (P1) local Shell 8,952 psi 28,950 psi Element 114
)
Membrane +
Free Shell 15,542 psi 28,950 psi Bending Element 19 Stress Range Local Shell 28,311 psi 69,900 psi
)
Element 147
)
)
W P W NE
. ENGNEBUNG SERVICES
)
Compressive Buckling - Acceptable (see below)
Compressive Buckling - Reference 13 discusses the results of analy-tical studies and tests on Mark 1 torus structures to determine
)
their compressive buckling capabilities. The report concludes that SRV is the dynamic load which presents the maximum chance of com-pressive buckling failure; but, that a safety factor of 7 still exists for an applied SRV pressure of +29.3/-22.6 psi. The maximum worst-case SRV shell pressures for Vermont Yankee are +5.77 psi and
-4.81 psi, which are lower than those used in the referenced study.
Based on this, compressive buckling stresses are considered to be acceptable for the Vermont Yankee torus.
FATIGUE EVALUATION - VERMONT YANKEE CUMULATIVE USAGE FACTOR
SUMMARY
?
(Stress Intensification Factor = 4.0)
EVENT TYPE NORMAL ELEMENT OPERATION SBA/IBA DBA
)
19 0.0
.0001
.012 147
.001
.011
.078 3.3.2 Support Columns & Attachments
)
The controlling load case for the support column at Vermont Yankee is load case 16 of Table 1.
The controlling condition is the result of a downward load.
This same case controls stress for the column tie-down
)
structure, during upward loads. Load case 16 includes:
Pool Swell (OAP) + Weight
)
)
O TN Technical Report TR-5319-1 M MS O
For the column-to-shell weld, load case 25 controls:
Pool Swell (full AP) + Seismic (SSE) + SRV + Weight O
For these load cases, the following controlling conditions were identified:
SUPPORT COLUMN - CONTROLLING AXIAL CONDITION O
LOAD CONTROLLING ACTUAL ALLOWABLE COLUMN DIRECTION CONDITION FACTOR FACTOR O
Inner Down Axial + Bending
.42 1.0 Outer Down Axial + Bending
.54 1.0 O
COLUMN-TO-SHELL WELD C)
LOAD CONTROLLING ACTUAL ALLOWABLE LOCATION DIRECTION STRESS STRESS STRESS Inner Down Shear 15.08 K/in 17.06 K/in
- O Outer Down Shear 16.31 K/in 17.06 K/in
,g COLUMN TIE-DOWN STRUCTURE LOAD ACTUAL ALLOWABLE LOCATION DIRECTION LOAD LOAD O
Inner Column Up 97 K 240 K I
Outer Column Up 135 K 240 K O
)
Technical Report TN TR-5319-1 6M
)
3.3.3 Support Saddles & Shell Weld Controlling stresses saddle are associated with two dif-ferent load cases for upward and downward loads. For downward loads, case 16
)
controls:
Pool Swell (0AP) + SSE + Weight
)
Controlling saddle stresses related to upward loads are the result of case 21:
DBA.C0 + Seismic (SSE) + Weight
)
Controlling stresses in the attachment weld between the sad-dle and the torus shell result from the downward loads of case 25:
)
Pool Swell + (full AP) + SRV + Seismic (SSE) + Weight Controlling conditions are:
)
SADDLE STRESSES LOAD STRESS TYPE OF DIRECTION LOCATION STRESS ACTUAL ALLOWABLE
)
Down Sole Plate Bending 18.5 K/in 28.5 K/in Up Clamping Plate Bending 48.8 K 75 K
)
(Load Related)
D
I Technical R: port Y
TR-5319-1 ENGNEERNG SERVICES
)
SADDLE-TO-SHELL WELD LOAD STRESS TYPE OF DIRECTION LOCATION STRESS ACTUAL ALLOWABLE
)
Down Outside End Shear 11.62 K/in 13.65 K/in 3.3.4 Earthquake Restraints & Attachments
)
The earthquake restraint system is illustrated in Figure 3-10.
The controlling load case for this system is the one that produces the largest lateral load. This is case 15 which includes:
)
Chugging + SRV + SSE All three of these loads have been selected to produce the
)
highest lateral load on one earthquake restraint; contributions from the individuai loads were added directly.
The controlling stress results follow:
')
EARTHQUAKE RESTRAINT STRESS STRESS ACTUAL ALLOWABLE
- LOCATION, TYPE STRESS STRESS Tie Plate Pin Bearing 1,847 psi 34,200 psi ATTACHMENT WELD
)
STRESS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS
)
Weld at Base Shear 2,379 psi 21,000 psi of Tie Plates
)
O Technical Report "WTELEDYNE TR-5319-1 ENGNEERING SERVICES 3.3.5 Anchor (Tie-Down) System The load combination which produces the highest upload and minimum margin on the anchor bolts is case 25 for the column anchors:
O Pool Swell (full AP) + Weight + SSE + SRV and load case 21 for the saddle anchors
_J DBA.C0 + Seismic (SSE) + Weight For these cases, the anchor bolts with the smallest margins of safety (accounting for as-built conditions) are:
COLUMN ANCHOR BOLTS (Capacity of Mounting Pad - 8 Bolts)
ACTUAL FACTOR MAXIMUM MAXIMUM OF g
LOCATION LOAD CAPACITY SAFETY Outside Column 76.7 K 312 K 4.07 g
SADDLE ANCHOR BOLTS ACTUAL FACTOR PAXIMUM LOAD OF LOAD CAPACITY SAFETY O
48.78 K/ bolt 263 K/ bolt 5.41 O
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l Technical Report TN TR-5319-1 ENGrEERNG SERVICES
)
4.0 VENT HEADER SYSTEM 4.1 Structural Elements Considered
}
The vent header system, as defined in this section, includes the following structural components:
l a.
VentHeader(V.H.)
)
b.
Main Vent Pipe (V.P.)
c.
Downcomers (D.C.)
d.
Downcomer Tie Bars e.
Deflector p
f.
Vent Header Support Columns & Attachments g.
VH/DC Intersection h.
VH/VP Intersection l
1.
VP/Drywel' Intersection
)
j.
Vent Header Mitre Joint l
l The main vent bellows are considered in Section 7.0.
p 4.2 Computer Models l
Two computer models provided the means to analyze the vent header 1
system, they are shown in Figures 4-1 through 4-4.
3 The first of these is a detailed shell model, (Figures 4-1 to 4-3),
which includes a highly detailed representation of one-half of the header in a non vent bay, complete with four downcomers.
D The model also includes an approximate representation of one-half of the vent bay; this was intended to provide the proper boundary conditions and stiffness transition near the non-vent bay.
The vent bay half of the D
D
D TN I
O 1 D model was not used for stress determination. This large finite element model was used primarily to determine shell stresses in the non-vent bay; some other uses are discussed in the following text.
It was used for both static and 9
dynamic analysis and provided detailed stress gredient information in the downcomer/ vent header intersection region.
The second vent header model is the beam model shown in Figure 4-4.
G This model represents a full vent bay, complete with vent pipe and downcomers; as well as a half non-vent bay on either side.
It was used to determine boundary loads on the vent system components to support a more detailed stress analysis of those components.
This model was used to define loads on the 3
following elements:
e Vent Header Support Columns e
Vent Pipe / Vent Header Intersection D
e Vent Pipe /Drywell Intersection e
Vent Header Mitre Joint e
Main Vent Pipe J
The loads and moments taken from the beam model were used in further analysis to determine stresses. The calculation methods used for these stres-ses are:
O e
VH support columns - hand analysis e
VP/VH intersection - applied stress multipliers (stress intensification factors) from Reference 7 e
VP/drywell intersection - used stress multipliers from -
O Reference 16 (Bijlaard) e Mitre joint - used stress multipliers from detailed shell model (Figure 4-1) e Main vent pipe - hand analysis D
)
W F W NE Technical Report TR-5319-1 ENGNEERING SERVICES
)
The beam model used a stiffness representation of the VP/VH inter-section taken from Reference 7.
Attachment stiffness between the vent pipe and drywd1 was calculated using References 17 and 18.
Pool swell water impact on the vent header deflector was calculated i
with a hand analysis.
The impact forces were applied statically to a beam model and a dynamic load factor was applied (see Figure 4-5).
)
4.3 Loads Analysis 4.3.1 Pool Swell Loads
)
4.3.1.1 Pool Swell Water Impact Analysis for stresses due to pool impact and drag
)
was done using both computer models.
Determination of shell stresses was done with the detailed model in Figure 4-1.
For this analysis, force tine histories based
)
on QSTF test data were used (References 4 and 10). These time histories were applied at 100 nadal points on the shell model and the dynamic response of the structure was calculated. Reiative timing between loadings (Reference 1) was maintained to preserve accurate representation of longitudinal and circum-
)
ferential wave sweep. Stresses in the vent header /downcomer intersection, as well as in the free shell areas, were taken directly from this model.
Stres-ses in the downtomer tie bars were also taken from this model. Analysis was done for both full and zero AP impacts.
)
The beam model (Figure 4-4) was also used to deter-mine stress from pool swell impact and drag.
This was done with a time history dynamic analysis using loads developed by integrating the impact
)
pressures over small areas ar.d reducing them to nodal forces. Approximately
)
)
Technical R: port gg TR-5319-1 )
95 nodes along the length of the beam model were dynamically loaded in this analysis, including loads on the VP/VH intersection and vent pipe.
The results of this analysis were used to define boundary loads on VP/VH inter-
)
section, mitre joint and other elements as listed in Section 4.2.
Stress analysis for these elements was performed using the methods indicated in Section 4.2.
4.3.1.2 Pool Swell Thrust (4.2) 3 Pool swell thrust forces are defined as dynamic forces at each bend or mitre in the vent system, and are a consequence of the 3
flowing internal fluids.
Analysis for these loads was done using the beam model and applying the loads statically.
This is consistent with the slow nature of the applied pressure forces.
g The calculation was performed with the maximum value of all thrust forces applied simultaneously; this is a conservative condition.
4.3.1.3 Pool Swell Drag Loads (4.3.7 & 4.3.8) g The vent header support columns are loaded by for-ces from LOCA-jet and LOCA bubble drag. By inspection, it was concluded that g
LOCA-jet loads would not combine with water impact on the vent system due to differences in timing and, therefore, would not contribute to the maximum stress calculations - LOCA jet forces were not considered further.
g LOCA bubble forces were calculated and the maximum normal components (radial and longitudinal) were applied simultaneously to conservatively bound the benuing moments on the support column.
These peak values were applied statically at the midpoint of the column. Stress calcu-
]
lations were done by hand.
J
... ~.
f Tcchaical Report TN TR-5319-1 6M
)
4.3.2 Chugging Loads 4.3.2.1 Downcomer Lateral Loads (4.5.3)
Ref erence 1 identifies downcomer lateral loads as static equivalents with random orientation in the horizontal plane. The major consequence of this loading is to prcduce high local stress in the VH/
downcomer intersection.
The detailed shell model (Figure 4-1) was used to
)
identify stresses in the downcomer intersection due to static loads applied at the base of the downcomer. Frequencies of the first downcomer response moda were taken from a dynamic analysis on the same model (Figure 4-1) with the downcomers full of water to the operating level. This frequency was necessary
)
to determine the proper dynamic scale factor to apply to the static load.
The stress results from the statically applied load were used as a basis for a fatigue evaluation of the intersection in accord-
)
ance with Reference 9.
4.3.2.2 Chugging - Synchronized Lateral Loads
}
The random nature of the downcomer lateral chugging load provides for all combinations of alternate force orientations on adja-cent pirs of downcomers. Various load combinations were examined to deter-mine stress levels in the vent header and mitre joint as a result of these
)
loads. The cases considered are shown in Figure 4-6.
These cases were considered by applying static loads to the beam model (Figure 4-4) and determining final stresses as
)
described in Section 4.2.
4.3.2.3 Internal Pressure (4.5.4)
)
Three vent system internal pressures exist during chugging. They are:
)
D Technical Report TN TR 5319-1 N SER\\/ ICES O
e Gross vent system pressure - a.7 Hz oscillat-ing pressure with a maximum value of 5.0 psi.
This pressure acts on the entire vent system.
D e
Acoustic vent system pressure - a sinusoidal pressure varying from 6.9 to 9.5 Hz at a maxi-mum value of 3.5 psi.
This pressure affects the entire vent system.
w!
e Acoustic downcomer pressure oscillation - a 40-50 Hz pressure at 13 psi that produces only hoop stress in the downcomers.
O Responses to these pressures were estimated using hand analysis and were determined to be substantially less than those result-ing from internal vent system pressures during pool swell. The values associ-h ated with pool swell pressures were used in all combined load cases involving chugging pressures; this produces conservative results.
I 4.3.2.4 Submerged Structure Drag (Support Columns only)
Fxamination of the load combinations that include chugging makes it clear that, these cannot control maximum stress level in the support columns; combinations that include vent header water impact will O
produce much higher stresses.
For this reason, stresses in the vent header support columns were not calculated for chugging drag.
Drag forces on the downcomers and downcomer tie oV bars are already included in the Downcomer Lateral Loads, which were based directly on test data.
O O
TGchnical Report W F W NE TR-5319-1 NBtNG SBRN/ ICES
)
4.3.3 Condensation Oscillation - DBA 4.3.3.1 Downcomer Dynamic Load (4.4.3.2)
)
The downcomer dynamic load, due to condensation oscillation, is a sinusoidal pressure variation that can be equal or unequal in the two downcomers forming a pair.
)
The unequal pressure case produces a net lateral load on the downcomer much like chugging. The major considerations for this load are stresses in the downcomer intersection due to a net lateral load on one pair of downcomers and a more general stress case where loads on adjacent downcomer pairs are phased to produce gross lateral loads on the vent system or torsion in the vent header.
l l
The horizontal component of the C0 downcomer load produces the same type of loading on the vent system as the CH lateral load, in terms of the stress analysis. A comparison of the magnitudes and frequen-cies of these two loads shows that stresses resulting from CH horizontal loads j
will bound C0 horizontal loads.
The C0 downcomer load also produces a vertical l
component of load, which is not present during CH. The effects of this load were evaluated by static analysis of the detailed vent header model (Figure 4-
- 1) and consideration of dynamic amplification effects, using the beam model (Figure 4-4).
This evaluation showed that the combined effects of the C0 downcomerload(horizentalandverticalcomponents)wouldstillbeboundedby CH 1ateral loads, g
i Based on this, CH lateral load results were con-servatively used in all load cases in place of C0 downcomer loads.
O I
I b
)
Technical Report
$ 4' M TR-5319-1 EN^N SERVICES
)
4.3.3.2 Vent System Loads (4.4.4)
Vent system loads consist of a sinusoidal pressure in the vent header and downcomers superimposed on a static pressure.
The
)
dynamic pressure in the downcomers is used to calculate hoop stress only.
Stresses for all pressure loads were based on hand analysis using static analysis. The static analysis assumption is consistent
)
with the low frequency of the applied pressure and the fact that the ring modes of the structure are very high.
l 4.3.3.3 Thrust Forces (4.2)
)
Stresses resulting from C.0. thrust forces were conservatively taken from the pool swell thrust calculations and applied to l
all C0 load cases (paragraph 4.3.1.2).
l 4.3.3.4 Drag Forces on Support Columns Inspection of approximate total loads on support
)
columns due to CO, CH, and pool swell showed that condensation oscillation
- would not contribute to the maximum column load, due to differences in timing.
No detailed analysis was performed.
4.3.4 Condensation Oscillation - IBA Stresses and loads resulting from IBA condensation oscil-lation are bounded in all cases by either DBA condensation oscillation or chugging.
No detailed analysis was performed for IBA condensation oscil-lation.
Technical Report TN TR-5319-1 N SERN/ ICES
)
4.3.5 SRV Loads 4.3.5.1 SRV Drag Loads
)
An SRV discharge produces drag loads which ar.t on the vent header support columns, downcomers, and downcomer tie bars. Analysis for drag loads on these structures was based on data collected during in-plant
)
SRV tests.
Data collected during these tests was scaled to correct it for the appropriate SRV conditions and then applied to the struc-
')
tural model to determine the resulting stress. A more detailed discussion of this procedure is provided in Appencix 1.
4.3.6 Other Loads
)
Deadweight and seismic stresses in the vent system were calculated using the beam model of Figure 4-4.
)
Seismic stresses were calculated by statically applying the acceleration values in Table 3.
Thermal stresses were determined for the steady
)
state application of maximum vent systea temperature, using hand analysis.
4.4 Results and Evaluation
)
Results are reported for each structural element of the vent system for the controlling load combination. Controlling load combinations are the ones that produce the smallest margins against the allowable stress - not necessarily the highest stress. All load combinctions listed in Table 1 have
)
been considered.
h
)
Technical Report TN TR-5319-1 N MICES
)
As stated previously, most results include some level of bounding analysis and, therefore, understate the margins which actually exist.
4.4.1 Vent Header - Downcomer Intersection
)
The controlling load on the vent header-downcomer inter-section, both for maximum stress and fatigue, is the downcomer lateral load
)
due to chugging.
The worst load combination in which this load appears is case 15 of Table 1.
This cases consists of:
Chugging (IBA) + Seismic (SSE) + Weight + Pressure + Thrust
)
< SRV For this case, the following stress occurs at a point 90 from the plane of a downcomer pair.
It is primarily the result of a longi-
)
tudinal chugging force on the downcomer.
ACTUAL ALLOWABLE TYPE OF STRESS STRESS STRESS Combined Maximum Stress 36,719 psi 37,635 psi 4.4.2 Vent Header - Vent Pipe Intersection
)
The controlling load on the vent neader/ vent pipe inter-section occurs as a result of pool swell water impact. The controlling load condition is case 25 in Table 1 which in-ludes:
)
Pool Swell (fullAP) + Thrust + Seismic (SSE) + Weight + SRV Pressure
)
3
)
Technical Report it' F W NE TR-5319-1 ENGNEERING SERVICES ACTUAL ALLOWABLE TYPE OF STRESS STRESS STRESS
)
Combined Maximum Stress 28,930 psi 28,950 psi This load case was formed using a 0AP load, and was
)
evaluated to a level A allowable. This conservative evaluation was performed to eliminate the need to evaluate several other vent header load cases.
4.4.3 Vent Header Support Columns & Attachments
)
The controlling load combiriation for the vent header support columns and the clevis joints at each end is case 25, Table 1.
This case includes:
)
Pool Sweli (full AP) + Seismic (SSE) + Weight + Thrust + SRV j
Controlling stress in the support column is:
ACTUAL ALLOWABLE TYPE OF STRESS STRESS STRESS
)
Axial in Column (tension) 12,028 psi 16,380 psi Cor-trolling stress in the clevis joint at the end of the support column is:
LOCATION TYPE STRESS STRESS Clevis Plate Shear 12,629 psi 13,840 psi
)
4.4.4 D_owncomer Tie-Bars and Attachments The controlling load combination for stresses in the down-
)
comer tie-bar and attachments is case 25, in Table 1.
The major load is associated with pool swell impact on the crotch region of the downcomers which produces tensile loads in the tie bar.
{
Technical Report WN f
TR-5319-1 ENGNEERNG SERVICES i
i D
The controlling case includes:
Pool Swell Impact (full AP) + SSE Seismic + SRV + Weight +
(
Pressure + Thrust The controlling stress is:
t
}
ACTUAL ALLOWABLE H
LOCATION STRESS STRESS STRESS Tie-Bar Clamp Bending 16,800 psi 22,240 psi 4.4.5 Vent Header Deflector and Attachments g
The major load on the vent header deflector occurs as a result of pool swell water impact. The controlling load condition is case 19 g
in Table 1 which includes:
Pool Swell (full AP) + SSE Seismic + Weight + Thrust g
The controlling stress in the deflector is:
STRESS ACTUAL ALLOWABLE u0 CATION TYPE VALUE VALUE Center of Bending 10,000 psi 57,400 psi the Long Span 4.4.6 Main Vent /Drywell Intersection O'
The major load on the drywell penetration occurs as a result of pool swell.
The controlling load condition is case 19 in Table 1 which includes:
n Pool Swell (OAP) + Seismic (SSE) + Weight + Thrust +
Pressure O
L
Technical Report TN TR-5319-1 N SB{\\/ ICES
)
The controlling stress is:
ACTUAL ALLOWABLE TYPE OF STRESS STRESS STRESS
?
Local Membrane 22,365 psi 28,950 psi The effects of all loads from the vent system, and the pres-sure load were considered using Reference 16. Information regarding stresses
)
due to seismic and thermal response of the drywell is not available and therefore have not been included.
4.4.7 Vent Header, Main Vent & Downcomers - Free Shell Stresses It was established by inspection of the computer results that large safety margins existed in free shell regions and that minimum safety margins would be controlled by local shell stresses. No further work
)
was done for free shell stress in these structures.
4.4.8 Vent Header - Mitre Joint
)
The controlling load on the vent header mitre joint occurs as a result of pool swell water impact.
The controlling load condition is case 25 in Table 1 which includes:
)
Pool Swell (full A P) + Thrust + Seismic (SSE) + Weight
+ SRV + Pressure
)
ACTUAL ALLOWABLE TYPE OF STRESS STRESS STRESS Combined Maximum Stress 27,137 psi 28,950 psi
)
)
Technical Report W F W NE TR-5319-1 ENGNEERNG SERVICES 3
4.4.9 Fatigue Evaluation The f atigue analysis of the vent system is a conservative one which assumes that all maximum stresses occur simultaneously, and that all O
cycles reach these maximum values.
The durt. tion of the major loads in this analysis is 900 seconds, the length of chugging associated with an SBA/IBA event.
O The procedure used in this analysis consists of the follow-ing steps:
e For a given load and component, locate the highest O
stress.
For this location, establish the stress range.
e O
e Repeat this process for all other loads in the load combination, e
Add the stress ranges for all loads.
O Multioly this total stress range by the appropriate e
stress intensification factor.
O Calculate stress intensity and determine the allow-e able number of stress cycles.
e Determine the usage f actor, using the methods of iO Reference 11.
The f atigue evaluation was performed for all high Stress areas in the vent system. The major load, contributing to the fatigue evalua-
'O O
l
Technical Report TN TR-5319-1 ENGNEERNG SERVICES tion is chugging following a DBA.
The controlling load case is case 21 in Table 1, which includes:
Chugging (DBA) + Seismic (SSE) + SRV + Weight
)
The controlling usage factor for the vent system is:
VENT SYSTEM FATIGUE RESULTS
)
ACTUAL ALLOWABLE USAGE USAGE LOCATION FACTOR FACTOR At the VH Support
.76 1.0
)
)
)
)
)
5
B Technical Report T
TR-5319-1 gSsMCES 9
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5.0 RING GIRDER ANALYSIS The ring girder for Vermont Yankee is shown in Figure 5-1.
It is mounted in a vertical plane that passes through the support saddles and the support
)
columns. Because all major internal structures are supported by the ring gir-ders, the ring girders must react to the largest number of individual loads.
5.1 Structural Elements Considered
)
Elements considered in this section are:
(a)
The ring girder web and flange
)
(b)
The attachment weld to the shell Local stresses at attachments have also been considered and added; i.e., vent header support columns, etc.
The catwalk is not included in the
)
stresses reported in this section, but in all cases can be added directly to the reported stresses without exceeding allowables. It was not ddded because it is local to a specific area not affected by other stresses.
)
5.2 Computer Models Two computer models were useu as a part of the ring girder analysus; both are detailed models which also include the shell and external supports.
)
The first model is shown in Figure 5-2.
This is a detailed model, which represents one-sixteenth of the torus structure; one half bay on each side of the mitre joint.
It accurately simulates the ring girder offset
)
(four-inches from the mitre joint) as well as strucural differences between the vent and non-vent bays. Because the ring girder is not at the boundary of this model, out-of-plane motion of the ring girder can be accurately deter-mined. This model was used to evaluate all direct loads on the ring girder;
)
these include loads from attached structures such as the tee-quencher sup-ports, catwalk and vent header system, as well as all drag loads. The one-
)
a
h Technical Repsrt YF WE TR-5319-1 NNG WICES sixteenth model used for the Vermont Yankee ring girder analysis was one that had been constructed for one of the other Mark 1 plants analyzed by TES. The dimensions of this other plant are very similar to Vermont Yankee; the dia-meter of the torus is 7% larger (conservative); the shell thickness and distance between the ring girder and mitre joint are similar.
The ring airder flange in this model is slightly smaller than Vermont Yankee and, therefore, produces conservative results since lateral loads control ring girder stresses. The comparison is:
Ring Girder Flange Dimensions (inches)
Vermont Yankee:
1.5 x 8
)
Model Used:
1.5 x 6 The second model used to determine ring girder loads is the Vermont Yankee 1/32 finite element model shown in Figure 3-1.
This model was used
)
previously to evaluate shell stresses of all symmetric loads that act on the shell. These same computer analyses produce information on ring girder stress for symmetric loads. Loads evaluated with this model include weight, internal pressure, and all shell dynamic loads. The boundary conditions on this model restrict the ring girder to in-plane motion.
5.3 Loads Analysi!
5.3.1 Loads Applied to Shell As stated, the ring girder stresses for all symmetric loads applied to the shell were taken from tF appropriate analyses described in Section 3.0; these include:
(a)
Pool Swell Shell Load (Paragraph 3.2.1)
(b)
Condensation Oscillation (3.2.2)
)
(c)
Chugging (3.2.3)
)
Technical R: port TN TR-5319-1 ENGREBUNG SERVCES 3
(d)
SRV Discharge *
(e)
Seismic (f)
Deadweight, Thermal and Pressure l3
- SRV discharge is conservatively assumed to be a symmetri-cally applied load for shell analysis.
5.3.2 Drag Loads O
The ring girder is subject to drag loads from each of the dynamic shell loads as well as Fluid Structure Interaction (FSI) effects from C0 and CH.
All these loads were evaluated by using the 1/16 model and g
applying static loads out-of-plane on wetted nodes of the ring girder.
The use of static analysis was based on the assumption that the stiffening effect of the saddle, columns and column gussets would make the ring girder very stiff and would prevent frequency interaction with the dynamic loads. Because o
of this, no dynamic load factors were applied to the static analysis results (DLF = 1.0).
Drag loads considered were:
(a)
Pool Swell Bubble g
(b)
Pool Swell Jet (bounded by a)
(c)
SRV Jet (d)
SRV Bubble (e)
C0 including FSI (bounded by g) g (f)
Pre-chug including FSI (bounded by g)
(g)
Post Chug including FSI The effects of SRV jet (c) and SRV drag (d) were evaluated g
based on data collected from in-plant t :sts.
A discussion of the in-plant tests and the use of drag data from these tests is given in Appendix 1.
Calculation of ring girder drag loads, due to condensation O
oscillation and post chug FSI, was not in accordance with NUREG 0661 (Reference 2).
An alternate method of calculating drag volume was used in this load g
t
Technical Report YMM TR-5319-1 ENGNEERING SERVICES calculation.
It produced drag volumes for the ring girder of approximately half of those that the NUREG 0661 procedure would have produced. A discussion of this is included in Appendix 3.
The FSI drag calculation was based on local pool accelerations at the ring girder due to the response of the entire
)
shell. The post chug and FSI analysis considered frequencies to 31 Hz, which were combined by adding the values of the five maximum components to the SRSS sum of the others.
)
5.3.3 Loads Due to Attached Structure Loads applied to the ring girder by structures attached to
)
it were evaluated by equivalent static analysis, using the 1/16 model (Figure 5-2).
The important loads are applied in the area of the support saddle and columns which make the ring girder very stiff and minimizes dynamic inter-action.
Because of this, dynamic amplification of the static ring girder
)
stresses was not done (DLF = 1.0).
The load input to the ring girder was a result of a dynamic analysis of the attached system (or had an appropriate DLF applied) and, therefore, incluc'ed the effects of dynamic amplification on load.
)
The following loads are applied to the ring girder and were considered:
e Tee-quencher support beam thrust due to SRV dis-
)
charge.
e Tee-quencher and support drag loads.
)
Vent header s pport column reaction loads during e
pool swell.
o Vent header support column drag loads.
J J
Technical Report WW WNE i
TR-5319-1 N M N ES As stated in Section 5,1, stresses resulting from attached
)
structure have been included in the following results, except for the catwalk I
which could be added without exceeding allowables.
5.4 Results & Evaluations
)
5.4.1 Ring Girder Web & Flange The controlling load combination for the ring girder web and flange is load case 16 of Table 1; this includes:
Pool Swell (OAP) + Pressure + Weight The controlling stress is:
STRESS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS D
Web Membrane 15.1 ksi 19.3 ksi Flange Membrane 10.6 ksi 19.3 ksi J
5.4.2 Weld to Torus Shell The controlling load combination for the shell weld is load 3
case 21 in Table 1; the controlling stress is:
STRESS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS O
Column Region Shear 5.80 K/in 8.53 K/in (Inside)
Column Region Shear 8.03 K/in 8.53 K/in lV (Outside)
Saddle Region Shear 6.94 K/in 8.53 K/in D
WTERME
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7 Technical Report WTE.EME TR-5319-1 NIERVICES l
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)
FIG. 5 -4 DETAILED RING GIRDER - SHELL MODEL
)
RING GIRDER ELEMENTS
)
Tcchnical Report T TFI N E TR-5319-1 ENGNEER!NG SERVICES f
6.0 TEE-QUENCHER AND SUPPORT The following results for the tee-quencher and supports are.:or;ervative due to the combined effect of several factors, three of which are:
e The calculationai methods to determine applied loads improved af ter this analysis was complete, and would provide reduced stresses.
e Some loads were intentionally bounded by conservative values from other plants so a single calculation could be used for more than one plant.
l f
e For submerged drag loads, individual frequency components were l
added to produce maximum stress without regard to load direction.
The effect of these conservatisms varies among stresses, but can be
)
significant in some cases.
6.1 Structural Elements Considered The configuration of the quencher and support is shown in Figure 2-8.
Vermont Yankee has four discharge lines, each enters the pool at a 30 angle.
f The structural elements considered in this section include:
l e
The quencher, e
The submerged portion of the SRV line.
e The quencher support beam and attachments.
)
)
Technical Report TN TR-5319-1 ENGBEBtNG SERVICES U
6.2 Computer Models The computer model used in this analysis is shown in Figure 6-1.
O This is a STAR 0YNE be*m model which represenu all piping and struc-ture between the drywell jet deflector and the ring girder. For these ruly-ses, the ring girder was assumed rigid and the vent pine penetration was represented by a stiffness matrix which was developed from a finite element O
model of the penetration.
Releases were modeled between the quencher and support plates to allow for free rotation of the quencher arms in the sup-ports.
O This model was used for both static and dynamic analysis.
6.3 Loads Analysis O
6.3.1 SRV - Load The calculation of stress due to SRV bicwdown was done by applying the dynamic loads to the computer model and calculating the time-O history response of the system. The applied loads included both the blowdown forces on the piping and the water cleai :
9ds at the quencher.
The i
controlling condition was for a second, mult. pie valve actuation af ter an IBA/'BA break, with steam in the drywell (SRV case C3.3). This case produces O
a high reficod level at the time of the second actuation and produces maximum load en the support system. Loads for this analysis were developed using G.E.
computer program RVFOR-04.
O 6.3.2 Pool Swell Loads The effects of pool swell jet and bubble loads on the quen-cher and support system were conservatively estimated by static analysis and a O
O l
)
Technical Report "WTF1FrVNE TR-5319-1 ENGINEERING SERVICES
)
dynamic load factor of 2.
It a s clear from this analysis that combined pool swell events would not control stresses - no further analysis was done.
6.3.3 Chugging Loads
)
Dynamic analysis of the quencher and support system was done for drag loads due to pre-chug, post chug and chugging FSI.
All of these analyses were based on a set of harmonic analysis which provided results for D
all steady-state frequency excitations from 1-31 hz. Results for individual load conditions were determined by scaling individual frequency results of the computer analysis by the appropriate pressure amplitude.
9 ihe mass of the structure used in the computer analysis was adjuste E
. TORUS
.hu D
D FIG. 7-5 3
TORUS MONORAIL 180 COMPUTER MODEL-VERMONT YANKE E
)
TGchnical Repsrt Wgg TR-5319-1
-105-g 8.0 SUPPRESSION P00L TEMPERATURE EVALUATION The Mark 1 modification which added tee-quenchers at the discharge end of the SRV lines required that we consider the high temperature performance characteristics of these devices. Several meetings took place where the high
)
temperature effectiveness and condensation stability of the devices was dis-cussed.
An important consideration in high temperature performance, is the mixing characteristics of the device and the attendent local-to-bulk tempera-ture difference (d t).
In response to these concerns and to assure reliable operation of these devices, the NRC has set limits on maximum pool temperatures for tee-quencher operation, as well as guidelines for a temperature monitoring system for the
)
suppression pool. These requirements are stated in NUREG 0661 (Reference 2) and NUREG 0783.
t 8.1 Maximum Suppression Pool Temperature
)
1 Analysis for maximum bulk pool temperature was performed by Yankee Atomic Electric Company.
The bulk pool temperature was conservatively determined by subtracting the 43 local-to-bulk temperature difference iden-
)
tified in the Monticello in-plant test from the local temperature limits defined in NUREG-0783. This is conservative since the 43 A T assu.nes no RHR actuation. The result is a bulk temperature of:
)
U 210 - 43 = 167 F at a mass flux rate 2
(42#m/sec-ft 200 - 43 = 157 F at a mass flu.v rate 2
'> 94 #m/sec-ft The results of the bulk temperature analysis for the most limiting
)
case (Figure 8-1) meet the above limit and, therefore, satisfy the req: ire-ments of NUREG 0661. Additional analyses are being contemplated to reduce the
)
I i
sJ Technical Report 9pgg TR-5319-1
-106-O conservatism present in the above calculation but this effort is outside the scope of the long-term program.
8.2 Pool Temperature Monitoring System O
The NRC criteria also presents guidelines for a monitoring system to constantly monitor pool temperature.
A monitoring system installed at Vermont Yankee which uses a network cf thermocouples, hardwired to a strip O
chart recorder in the control room will be upgraded to meet the NRC criteria.
The system is described more fully in Section 2.2.1 of this report and is illustrated in Figures 2-9 and 2-18.
O O
O O
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O
FIGURE.,8-1 TR -5319-1 Comparison of T-Qtamuher Bulk Suppression Pool Temperature Limit to Stuck Open S/RV From 100% Power Transient Responses
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)
Technical Report
-108-W TA mYNE TR-5319-1 ENGINEERING SERVICES I
REFERENCES 1.
G.E. Report NED0-21888, Rev. 2, " Mark 1 Containment Program Load Defi-nition Report", dated Novanber 1981.
)
2.
NRC " Safety Evaluation Report, Mark 1 Containment Long-Term Program",
NUREG 0661, dated July 1980.
3.
)
G.E. Report NED0-24583-1 "Marh 1 Containment Program Structural Accept-ance Criteria Plant Unique Analysis Application Guide" dated October 1919.
4.
)
G.E Report NE00-21944 "...h Scale 2-D Plant Unique Pool Sweli Test Report" dated August 1979.
5.
G.E. Report NED0-24615 "....
Scale Suppression Pool Swell Test Pro-
)
gram: Suppemental Plant Unique Test", dated June 1980.
6.
G.E. Report NEDE-24840 " Mark 1 Containment Program - Evaluation of Har-monic Phasing for Mark 1 Torus Shell Condensation Oscillation Loads" October 1980.
)
7.
G.E. Report NEDE-24519-P " Mark 1 Torus Program Seismic Slosh Evaluation" dated March 1978.
)
8.
G.E. Report NEDE-21968 " Analysis of Vent Pipe - Ring Header Inter-section" dated April 1979.
9.
)
G.E. Report NEDE-24555P " Mark 1 Containment Program - Application Guides 1-6, 9, 10, Volume 10 and 6, Rev. 3, others are Rev. 2."
10.
G.E. Report NE00-24581, Rev. l, " Mark 1 Containment Program - Plant
)
Unique Load Definition - Vermont Yankee Generating Stetion" dated October 1981.
3
O Tcchnical Report
'#PTF1 AT/NE TR-5319-1
-109-ENGINEERING SERVICES REFERENCES (CONTINUED) 11.
ASME B&PV Code,Section III, Division 1, through Summer 1977.
12.
Structural Mechanics Assoc. Rept. SMA 12101.05-R001, " Design g
Approach Based on FSTF Data for ComL'ining Harmonic Amplitudes for Mark 1 Post Chug Response Calculations", dated May 1982.
13.
Mark 1 Containment Program Report WE8109.31 " Buckling Evaluation of g
a Mark 1 Torus", dated January, 1982.
14.
Structural Mechanics Assoc. Report SMA-12101.04-R003D, " Response Factors Appropriate for Use with C0 Harmonic Response Combination g
Design Rules", dated March, 1982, pg. 3.
15.
Intentionally Blank.
O 16.
Welding Research Council Bulletin No. 107, " Local Stresses in Spheric;! & Cylindrical Shells dua to External Loadings", dated August 1965 with March 1979 Revision.
O 17.
Welding Research Supplement, " Local Stresses in Spherical Shelis from Radial and Moment Loadings", P.P. Bijlaard, dated May 1957.
18.
"On the Effects of Tangential Loads on Cylindrical & Spherical g
Shells", P.P. Bijlaard, Unpublished, Available from PVRC, Welding Research Council.
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-53
-111-NN TABLE 2 PLANT PHYSICAL DIMENSIONS
)
VERMONT YANKEE T0f,ilS Inner Diameter 27'8" Number of Sections 16 Shell Plate Thickness Vent Pipe Penetration 1.0625"
)
Top Half
.533" Bottom Half
.584" SUPPORT COLUMNS
)
Quantity Size Outer 16 I-Beam (12.5" x 1.25" Flange & 10" x 1" Web)
Inner 16 I-Beam (12.5" x 1.25" Flange & 10" x 1" Web)
)
I Base Assembly Sliding, Anchored by Modification RING GIRDER Quantity 16 Size T-Beam (8" x 1.5" Flange, 1.5" x 20" (Average) Web)
EARTHQUAKE RESTRAINT SYSTEM
)
Quantity 4
Type Support Saddles (Pin Jointed)
DRYWELL VENT SYSTEM
)
Quancity Size Vent Pipe 8
6'9" I.D.
Vacuum Breakers 10 18" I.D.
)
Vent Header Support Columns 16 pairs 6" Sch. 80 Downcomers 96 2' O.D.,
l'-11 " I.D.
Submergence 4.29' Min. - 4.54' Max.
)
Water Volume 0 Minimum Submergence 68,000 cu. ft.
l
Technical Report
-112-WN TR-5319-1 MM
)
TABLE 3 PLANT ANALYSIS INFORMATION
)
HRMONTYANKEE
)
Seismic Acceleration Values (G's)
_0BE SSE (.5% damping)
Vertical
.05
.09 Horizontal
.07
.14
)
Effective Water Mass for Horizontal Seismic Load (Reference 7) 34.9%
)
Effective Water Mass during Pool Swell Uplift (Reference 4)
)
Full A P - 57%
Zero AP - 52%
Plant Unique C.0. Multiplier (Reference 1)
)
.929 P
\\
Technical Report PTNE TR-5319-1 Al-1 N SME APPENDIX 1 Use of SRV In-Plent Test Data for Analysis Test Data The in-plant SRV tests used to support structural analysis were run at Vermont Yankee in March, 1981.
The data was, collected in a series of four
')
tests, each consisting of one actuation with a cold lir.e and a second about l
one minute later (hot line).
The test sets were about three hours apart to allow for SRV line tool down.
- )
The torus was instrumented with a combination of strain and pressure transducers as shown in Figure Al-1.
Strain gages were mounted ir pairs on both sides of the shell to allow separation of bending and membrane stresses.
l i
Additional gages were located on the columns (Figure fi-2), and internal
)
st'"Jctures (Figures Al-3 and Al-4).
l Two independent data collection systems were used to provide a check on system accuracy. The major system was a multiplexed FM tape system on which all data was cellected.
The second system was a hard wired oscillograph to produce direct, quick-look readout on several channels.
In all, 78 transducers were used during the testing. Some difficulty was experienced with the shell pressure gages and some gages did not work prop-erly; however, the remaining gages provided sufficient data to fulfill test objectives.
D Use of Data - Applications i
The SRV test data was used to calibrate computer analysis of the shell and support systems and also to establish actual numbers for SRV drag loads on submerged structures.
D
Technical Report TME TR-5319-1 Al-2 N NES e
Use of Data - Shell & Support System Analysis Evaluation of shell stress and support system loads due to SRV actuation e
was done with a large detailed computer model as discussed in para. 3.2.4 of the report.
Data collected from the in-plant tests was used to define the actual shell pressures and decay time for a benchmark (test) coadition and to develop correction factors between these measured results and values pre-
~
J dicted by generic analytical methods. The steps involved are these:
1.
Determine maximum average shell pressure, average frequency and waveform for the four cold tests.
J 2.
Calculate these same quantities for the test conditions using the generic compJter programs (QBUBS 02).
D 3.
Calculate calibration factors relating predicted-to-actual pres-sure and predicted-to-actual frequency.
4.
Calculate predicted pressures and frequencies using the generic D
computer prm um, for other SRV conditions.
5.
Apply the calibration factors calculated in step (3) to all other predictions for pressure and frequency. The duration of the pres-O sure transient, as measured in the test, is affected proportionally by the frequency correction and used as the basis for all computer model loading.
O Verification of Computer Model The test data was also used to verify the accurar.y of the computer model.
This was done by the following method:
O 1.
The computer model was loaded with the ineasured shell pressures.
O
Technical Report YM TR-5319-1 Al-3 6M 2.
The model was run and stresses at all strain gage locations were calculated.
3.
Comparisons were made between computer predicted shell stress and measured shell stress at the sama points.
Correlations for shell stress were excellent - generally within 5%.
Correlations to column loads were not so good - generally off by about 50%.
I This difference in computer results for test conditions was handled by devel-oping a second calibration factor for supports only, and combining it with the previous pressure calibration factor.
The results were two different cali-bration factors to be applied to final analysis - one for the shell and one for the columns. The factors developed and used are:
Shell pressure =.21 x predicted
)
Support load =.4 x predicted Multiple Valve Contributions
)
For cases where more than one valve actuates, the contributions from other valves were added directly (same signs).
The maximum value used was 1.65 x the pressure from a single valve (Reference 2).
)
SRV Test Data for Drag Loads The data collected during the Vermont Yankee in-plant test included strains measured on submerged structures.
Strain gages, positioned to show
)
bending stress due to drag loads, were installed on the catwalk support column and vent header support column.
Figure Al-4 shows the locations of these gages, relative to the quencher. The test data showed these results:
)
)
)
Technical Report TN TR-5319-1 Al-4 6M
?
1.
Structural response occurred at the natural frequency of the struc-ture only.
)
2.
Responses were much less than would be predicted by Program analy-sis methods - as much as an order of magnitude lower.
The data collected from Vermont Yankee was evaluated along with the data
)
collected by TES in three other in-plant tests. The matrix of data collected is as follows:
Ring Catwalk Vent Girder
)
Supports Column Downcomer (Pressure)
Milistone X
X X
Nine Mile Point X
X Vermont Yankee X
X
)
Fitzpatrick X
X X
An important consideration in the application of this data was the pos-sibility that resonant structural response might occur at some other SRV
)
condition. This was considered and dismissed based on two separate arguments; they are:
1.
If a major frequency component existed in the drag force, it would
)
be detectable on each of the structural responses for a given test.
This did not occur.
2.
The response frequencies of the structures tested (structural natu-
[
ral frequencies) ranged from 8.1 to 38 hz.*
If any single strong frequency existed in the drag load, one of the structural responses should have demonstrated some degree of resonant response - none did.
)
- Actual values were 8.1, 8.2,14.5,15, 21, 23, 24, 25, 29, 30, 34 and 38 hz.
J
)
Technical Report WP WNE TR-5319-1 Al-5 6 M9
)
We conclude from this that the structures involved are responding to a fairly uniform random field and that the test data represents useable data for all SRV conditions.
)
The next step in the process was to calculate an equivalent static load for each structure. This is the static load that produces the same bending stresser measured in the test, when applied uniformly to the submerged area.
These static pressure values were plotted against distance from the quencher
)
and Figure Al-5 was developed.
This curve represents the equivalent static drag pressures, including quencher jet loads.
It is scaled upward from test conditions to more severe SRV cases by the ratio of the calculated shell pressures for the two cases, for application to structures under different loading conditions.
)
)
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Technical Report TN TR-5319-1 A1-6 mgifMCE!S
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SRV TEST INSTRUMENTATION-VERMONT YANKEE INTERN AL-SHELL
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t Technical Report WTELEDGE TR-5319-1 Ab8 ENCBEERNGSEMCES 0
)
^
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--- S R V Li N E DC 30 FROM HORIZONTAL BUBBLE c'
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4 GAGEi AT 90
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FIG. Al-3 SRV TEST INSTRUMENTATION -VERMONT YANKEE
)
TEST QUENCHER 8 SUPPORT
D Technical Report WTELEDGE TR-5319-1 Al-9 NAGSEMCES e
9 A
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C VENT HEADER SUPPORT COLUMN 4 GAGES AT 90
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o FIG. Al-4 lO SRV TEST INSTRUMENT ATION - VERMONT YANKEE INTERNAL STRUCTURES
)
WTELEDGE Technical Report NSBMCES TR-5319-1 Al-10
)
f 4
)
33 SRV CASE Al.1 3 "NM
)
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)
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HORIZONTAL DISTANCE FROM BUBBLE (f t)
D
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EQUIVALENT SRV DRAG FROM IN PLANT TESTING i
FIG A l-5
)
Tcchnical Report WM TR-5319-1 A2-1 ENGNEERNG SERVICES l
APPENDIX 2 Discussion of 32 Hz Frequency Cut-of# for f
Condensation Oscillation and Post Chug Analysis TES made the decision to limit C0 and Post Chug response analysis to f
frequencies below 32 Hz early in the pregrara. The decision was the result of several considerations that led to the conclusion that the 32 Hz cut-off would produce realistic results.
The basis for use of a 32 Hz cut-off involved strong fundamental argu-ments, both in the loads used for the analysis, and in the stress analysis itself.
The primary arguments are different for CO, and for Post Chug, and are given here:
l For condensation analysis.
(
1.
Load Definition - A PSD study of the C0 pressure data showed that f
frequencies above 25 Hz accounted for only 10% of total power (Ref-erence 1, page 4.4.1-10).
This means that a system with flat frequency response to 50 Hz would suffer a 10% unconservative stress error if a 25 Hz cui.-off was used. Since we are using a 32 Hz cut-off and our system is highly responsive at low frequencies (not flat), we should expect a much smaller error.
2.
Structural Response Analysis - The relative importance of loads
)
below and above 32 Hz can be judged based on examination of the modal frequencies and generalized coordinates of the structure in both frequency ranges.
If we consider the characteristics of a typical torus model in these ranges, we find:
)
l
)
)
Technical Report WNNE TR-5319-1 A2-2 ENGNEERNG SERVICES Numbe 0f*
Numbe 0f Max.
2 2
Number of GX GX Valu Frequencies 1000
> 000 GX
)
Below 32 Hz 44 25 14 167,858 32-50 Hz 34 5
1 4,594
)
- Product of generalized weight and the square of the participation factor - used as an indicator of modal response strength.
These figures show that for condensation oscillation, frequencies below
)
32 Hz clearly dominate the response and frequencies above 32 liz are relatively insignificant.
They provide a strong indication that the 10% worst-case unconservatism discussed above will be greatly reduced by the sel.ective nat-ure of the structural response.
We should logically expect the structural
)
response characteristics, and the fact that we are using a 32 Hz cut-off, instead of 25, to reduce the 10% maximum error to less than 5%. An error of this magnitude is consistent with other assumptions which must be made in the analysis and is considered acceptable.
A further statement regarding the validity of this approach may be found in References 11 and 14.
)
For the Post Chug load, the second consideration of structural response is also valid, but the load definition is not as heavily skewed toward the low frequency end as is C0.
The decision for handling post chug was heavily influenced by the fact that it produced very low stress and, in fact, that shell membrane stresses would be bounded by pre-chug.
This is discussed further in Section 3.2.3.2 of this report.
)
D
)
Technical Report SPTE M E TR-5319-1 A3-1 MM $8{\\/ ICES
)
APPENDIX 3 C0/CH Drag Loads for Ring Girder Analysis TES did not follow the calculational methods of NUREG 0661 (Reference 2) for calculation of CO/CH drag loads on the ring girder.
This appendix describes the method that was used; the differences with the NUREG method; and the basis for the change.
The NUREG analysis method specifies that acceleration drag forces (and effective hydrodynamic mass) for flat plates be based on an equivalent cylin-der with radius equal to TE times the radius of the circumscribed circle. It also specifies that the d:ag forces be increased by an additional factor of 2 for structures attached to the torus shell, to account for wall interference.
Application of the NUREG criteria produces a factor of 4 multiplier for drag force for flat plate structures in the fluid; and a factor of 8 multi-plier for flat plate structures in the fluid and attached to the shell. These
)
values are referenced to a drag force equal to 1.0 for flat plate calculations based on potential flow theory and neglecting interference effects.
)
These increases in loads are supported by data available in Reference A3-1 and A3-2.
Keolegan and Carpenter show in Reference A3-1 that the drag
~
forces on a plate in an oscillating flow may be a factor of 4 higher than the
)
)
)
Technical Report
$ 7 K FTY( g TR-5319-1 A3-2 ENGNEERING SERVICES i
i
)
theoretical force based on potential flow. Sarpkaya shows in Reference A3-2 that forces on a cylinder near a boundary, may be twice as high as forces away from the boundary.
Both References 1 and 2 present results as a function of the VT/D ratio where:
l l
V = maximum velocity T = period of flow oscillation D = diameter l
l Keolegan and Carpenter show the effective hydrodynamic mass coefficient foraplatevariesfromamaximumof4atp"T = 125 to 1 at VT/D = 0. (pure potential flow).
Sarpkaya shows an increase in the hydrodynamic mass coef-
)
ficient for a cylinder near a boundary that varies from a maximum f actor of 2 l
atf=15toaminimumof1.65atVT/D=0.
l l
NUREG 0661 appears to use the bounding values from both of these refer-ences to formulate its' analysis method. It implies by this that large values of h will exist in the torus.
In fact, this is not true for C0 and CH j
drc3 loads on the ring girder. For this structure, under this load, VT 11 l
raitz are near zerc and the use of maximum multipliers should not be neces-sary. It is on this basis that we have used an alternate method to calculate
)
C0 and CH drag leads on the ring girder.
b l
WTFI STh'NE 1
A3-3 ENGNEERING SERVICES n
The TES method to calculate these drag loads on the ring girder used the j
same references as above (A3-1 and A3-2), but accounted for calculated values offratherthanthevaluescorrespondingtothemaximumincreases.
Con-g sideration of the actual f ratio for wall interference led to an interference factor of 1.65 (instead of 2).
f suggest Low values of that the theoretical hydrodynamic mass coefficient for the ring girde-is appropriate. The theoretical coefficient
- )
for this structure is estimated by an equivalent cylinder with a radius equal to the circumscribing radius. Use of this cylinder results in a hydrodynamic mass coefficient equal to two. The total factor used was related to the NUREG 3
multiplier by:
2.0 1.65 4.0 2.0
_,47 x
D Tne factor used by TES was.41 x the NUREG 0661 fac'or.
O O
O
/
)
Technical Report TM TR-5319-1 A3-4 6 ME
)
REFERENCES
)
A3-1 Keolegan and Carpenter, " Forces on Cylinders and Plates in a Oscillating Fluid," National Bureau of Standards, Vol. 60, aa. 5, May 3959.
)
A3-2 Sarpkaya, " Forces on Cylinders near a Plane Boundary in a Sinusoidally Oscillating Fluid", Journal of Fluids Engineering, September 1976.
)
)
J D
D D
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.