ML20084F221
| ML20084F221 | |
| Person / Time | |
|---|---|
| Site: | Vermont Yankee  |
| Issue date: | 11/30/1983 |
| From: | TELEDYNE ENERGY SYSTEMS |
| To: | |
| Shared Package | |
| ML20084F214 | List: |
| References | |
| TR-5319-1, TR-5319-1-R02, TR-5319-1-R2, NUDOCS 8405040024 | |
| Download: ML20084F221 (408) | |
Text
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YANKEE ATOMIC ELECTRIC COMPANY 1671 WORCESTER ROAD FRAMINGHAM, MA 01701 TECHNICAL REPORT TR-5319-1 MARK 1 CONTAINMENT PROGRAM PLANT-UNIQUE ANALYSIS REPORT OF THE TORUS SUPPRESSION CHAMBER FOR VERMONT YANKEE NUCLEAR POWER STATION l l NOVEMBER 30, 1983 REVISION 2 WTELEDYNE ENGINEERING SERVICES 130 SECOND AVENUE WALTHAM, MASSACHUSETTS 02254 617-890-3360
l Technical Report l TR-5319-1 .. WTF1 Fry (E Revision 2 ENGNEERNG SERVICES RECORD OF REVISIONS REVISION PAGE DESCRIPTION 1 Cover Changed Revision 0 to Revision 1 and l date from 11/20/82 to 4/8/83 Title Changed Revision 0 to Revision 1 and date from 11/20/82 to 4/8/83 x Added Table 4 1 Changed Reference 9 to Reference 1 2 Reformatted because of change on Page 1 31 Deleted Reference 9 32 Deleted Reference 9 58 Changed Reference 9 to Reference 1 108 Deleted Reference 9 113 Added Table 4 2 Cover Changed Revision 1 to Revision 2 and date from 4/8/83 to 11/30/83 Title Changed Revision 1 to Revision 2 and dated from 4/8/83 to 11/30/83 ii Added Appendix 4 ix Added Appendix 4 xi Figure 7-1, remove (unmodified) from title 9 Complete rewrite for modified catwalk 25 Catwalk supports revised 26 Catwalk supports revised 51 Figure 3-8, change scale on ordinate axis l 66 Revised membrane stress to include seismic l and thermal from drywell 73 Figure 4-5, added dotted lines and end supports l
l R- 1 Y Revision 2 -iii- N SERVCES l RECORD OF REVISIONS l REVISION PAGE DESCRIPTION 2 90 Para. 7.1 and 7.1.1 revised for modified catwalk 92 Change Figure 7.2 to Figure 7.1 (2 places) Remove "without grating" statements (2 places) 93 Revise " actual stress" numbers Remove "without grating" (2 places) 94 Revise " actual stress" number 100 Figure 7-1, redraw to show new computer model 101 Figure 7-2, delete, replace with new sheet 107 Figure 8-1, revised top two curves t l i l l l
Technical Report TR-5319-1 WTri pr?(NE Revision 2 -iv-ENGREEMNG SERVICES 1 ABSTRACT l 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 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 ccmpliance with NUREG 0661, unless otherwise noted. Following initial submittal of this report, it was reviewed by the NRC and its consultants. The review is complete, and is documented in Appendix 4 of this report.
Technical Report
' 9p qq R vi cn 2 -V- SERVICES TABLE OF CONTENTS Page ABSTRACT ii
1.0 INTRODUCTION
& GENERAL INFORMATION 1 2.0 PLANT DESCRIPTION 4 2.1 General Description 4 2.2 Recent Modifications 4 2.2.1 Modifications to Reduce Loads 5 2.2.2 Modifications to Strengthen Structure 8 3.0 CONTAINMENT STRUCTURE ANALYSIS - 30 SHELL & EXTERNAL SUPPORT SYSTEM 3.1 Computer Models 30 3.2 Load Analysis 31 3.2.1 Pool Swell Loads 31 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 3.2.4 SRV Discharge 34 3.2.5 Deadweight, Thermal & Pressure 35 3.2.6 Seismic 35 3.2.7 Fatigue Analysis 36 3.3 Results and Evaluation 37 3.3.1 Shell 38 3.3.2 Support Columns & Attachments 39 l
3.3.3 Support Saddles & Shell Weld 41 3.3.4 Earthquake Restraints & Attachments 42 3.3.5 Anchor (Tie-Down) System 43 l
Technical Report TR-5319-1 Revision 2
-vi-WT31 WE N S 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 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 Free Shell Stresses 4.4.8 Vent Pipe - Mitre Joint 66 4.4.9 Fatigue Evaluation 67
Technical Report TR-5319-1 1%Y WE Revision 2 -vii- DGNEBMG SERVICES 1 TABLE OF CONTENTS (CONTINUED) l l Page 5.0 RING GIRDER ANALYSIS 75 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 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 5.4.2 Attachment Weld to Shell 79 6.0 TEE-QUENCHER & SUPPORT 84 6.1 Structural Elements Considered 84 6.2 Computer Models 85 6.3 Loads Analysis 85 6.3.1 SRV Loads 85 6.3.2 Pool Swell Loads 85 6.3.3 Chugging Loads 86 6.3.4 Condensation Oscilliation Loads 86 6.3.5 Other Loads 87 6.4 Results & Evaluation 87 i
- 6.4.1 Tee-Quencher Structure 87 6.4.2 Submerged SRV Line 87 6.4.3 Tee-Quencher Support 88 l
l 7.0 OTHER STRUCTURES 90 7.1 Catwalk 90 7.1.1 Computer Models 90
Technical Report TR-5319-1 W Tptprt( E Revision 2 -viii-TABLE OF CONTENTS (CONTINUED) Page 7.1.2 Loads Analysis 90 7.1.2.1 Pool Swell Water Impact 90 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 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 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 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 7.3.3 Results & Evaluation 97 7.4 Monorail 98 7.4.1 Computer Model 98 7.4.2 Loads Analysis 98 7.4.2.1 Froth loads 98 7.4.2.2 Weight & Seismic 99 7.4.3 Results & Evaluation 99 8.0 SUPPRESSION POOL TEMPERATURE EVALUATION 105 8.1 Maximum Pool Temperature Analysis 105 8.2 Pool Temperature Monitoring 106
1 TR 53 1 Y Revision 2 -ix- N SERVICES TABLE OF CONTENTS (CONTINUED) Page REFERENCES 108 APPENDIX 1 - USE OF SRV TEST DATA IN ANALYSIS Al-1 APPENDIX 2 - USE OF 32 HZ CUT 0FF FOR A2-1 C.0. & POST CHUG ANALYSIS APPENDIX 3 - SUBMERGED STRUCTURE DRAG ANALYSIS FOR A3-1 THE RING GIRDER APPENDIX 4 - NRC REVIEW COMMENTS & RESPONSES A4-1 A4.1 A4.1-1 A4.2 .., A4.2-1
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s , Technical Report - ', WTia 'FWE e on 2 - -x-
- :ENGBEERNG SERVICES i F.TGURES AND TABLES Page
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FIGURES: , 2-1 Torus Plan View 11 2-2 Torus Composite Cross Section 12 2-3 Torus Modifications - Cross Section at Ring Girder 14 i 1 2-4 Torus Modifications - Cross Sect. ion at Mid Bay 15 2-5 AP Pressurization System. 16 2-6 Vent lleader Deflector'J -- 17
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2-7 Vent Header' Deflector Attachment 18 2-8 SRV Tee-Quencher and Suppott 19 2-9 Pool Temperature Monitoring System 20 2-10 RHR Rytorn Line Elbow'drid ; Support 21 2-11 Torus S'upport Saddles and Saddle Anchors 22 n. 2-12 Torus Support Column Anchors 23 2-13 Downcomer Tie Rod and Gusset Modification 24 2-14 Catwalk and Handrail Modificd.mion~ 25 2-15 Catwalk and Handrail Modification 26 2-16 TorusSprayHeaderSuppotSodificadons. 27 2-17 Monorail '
. 28 2-18 Thermocouple Locations ,
29 3-1 Detailed Torus Shell Modei
' 44 3-2 Detailed Torus Shell:Model 45 3-3 Detailed Torus Shell'fodel _ _
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Technical Report TR-5319-1 3pg Revision 2 -xi-SERVICES ! I FIGURES AND TABLES (CONTINUED) l 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 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 5-2 Detailed Shell - Ring Girder Model 81 5-3 Detailed Shell - Ring Girder Model 82 5-4 Detailed Ring Girder - Shell Model - Ring Girder Elements 83 6-1 SRV Line Analytical Model 89 7-1 Catwalk Computer Model 100 7-2 Intentionally Omitted 101 7-3 Spray Header Computer Model 102 7-4 Vent Pipe Bellows Motion 103 7-5 Monorail Computer Model 104 8-1 Bulk Suppression Pool Temperature vs. Quencher Mass Flux 107 Al-1 SRV Test Instrumentation - Shell Al-6 Al-2 SRV Test Instrumentation - Columns Al-7 Al-3 SRV Test Instrumentation - Tee-Quencher Al-8 Al-4 SRV Test Instrumentation - Internal Structures Al-9 Al-5 SRV Orag Pressures Al-10
y- ...: - s R-5 1 1 TE E ' Revision 2 -xii- ENGrEERING SERVCES _ FIGURES AND TABLES (CONTINUED) - Page TABLES: __
- 1. Structural Acceptance Criteria for Class MC 110 Internal Structures _
- 2. Plant Physical Dimensions 111
- 3. Plant Analysis Information 112
- 4. SRV Load Cases 113 l
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. 3 - ENGNEERNG SERVICES i
Revision 2 1.0 GENERAL If60RMATION q The purpose of the Mark 1 Torus Program is to evaluate the effects of I hydrodynamic loads resulting from a loss of coolant accident and/or an SRV discharge on the torus structure. This 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 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 Section XI of the ASME Code and meet the Summer 1978 Edition of Section III for design, materials and fabrication. A great many technical reports have been written and issued as a part of this program. These reports provide detailed descriptions of the phenomena, the physics controlling the phenomena, calculational methods and detailed procedures for plant-unique load calculations. Several of these documents are listed as references in this report. The approach of this report will be to reference these documents, wherever possible, and to avoid a re-statement of the same information. A major part of this program has dealt with providing plant-unique load calculation procedures (Reference 1 is an example 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 referenced in
WTri m(NE N${U}"*P"* _2_ ENGNEERING SERVCES Revision 2 Program documentation. Examples 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 documen-tation; 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 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. 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. The loads and load combinations considered in this program required special consideration to determine the appropriate levels of ASME Code appli-cation. Reference 3 was developed to provide this standard. Table 5-1 of 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 various SRV discharge conditions. The approach used in the final evaluation of structures is to reduce this large number to a 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 TM L TR-5319-1 N NES Revision 2 f 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 g 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 Load Definition Report (Refer-
< ence 1). This has been done by identifying the applicable LDR paragraph in parenthesis immediately following the title of the load. This referencing i directs the reader to a more detailed description of the load than can be included in this report.
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Technical Report "ptTF1 FrWNE
- TR-5319-1 ENGINEERING SERVICES Revision 2 2.0 PLANT DESCRIPTION 2.1 General Arrangement The configuration of the Vermont Yankee torus structure is shown in Figures 2-1 and 2-2.
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 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). 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 opening where the vent pipe penetrates the torus shell is sealed by a bellows. The ring girder also supports the catwalk, spray header, SRV tee-quencher support, and var-ious internal piping runs. Figures 2-3 to 2-17 show several details of the I torus structure. Table 2.0 lists several of the plant specific dimensions. 2.2 Recent Modifications Many modifications have been made at Vermont Yankee over the past several years to the torus, incluaing several made dur-ing the 1983 refueling outage, both to increase its strength and also to miticate the hydrodynamic loads. The modifications are illustrated and listed in the composite sections of Figures 2-3, and 2-4, along with their installation dates. A description and illustration of each individual modification follows: f
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i 1 Report
-5 W TF1FrWNE Revision 2 ENGNEERING SERVCES l
2.2.1 Modifications to Reduce Hydrodynamic Loads Drywell Pressurization System (AP System) Installation of a system to maintain a pressure differential 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 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) is to depress the water leg in the downcomers and reduce the water slug that must be cleared, if rapid pressurization of the drywell occurs. Early generic
! testing in the Program demonstrated that this was an effective means to reduce 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 plant. As illustrated in Figure 2-5, pressure differential is maintained by 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 maintain AP.
Vent Header Deflector The vent header deflector at Vermont Yankee is illustrated 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 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 to provide accurate vent header loading for detailed analysis.
Technical Report "R TriFrT(NE TR-5319-1 ENGINEERING SERVICES Revision 2 SRV Tee-Quencher I A tee-quencher has been installed at the discharge end of each main steam relief line to replace the existing ramsheads. The quencher 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 generic testing to reduce torus shell pressure by factors 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. 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 l 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 proper operation of the SRV quencher. These modifications are illustrated in 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 which are directly affected by relief valve discharge, plus one additional bay 1 1 for complete coverage (See Figure 2-18). Five thermocouples are used to provide individual indication of temperatures in each of the five bays, while the other five thermocouples are combined to provide an average or bulk 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 have been remounted on vertical support members. The vertical support members are 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
Technical Report TN TR-5319-1 Revision 2 N SOMCES recorder is installed in the control room. The recorder is used to record the 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. 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 3 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 was 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 'sf 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 RCIC turbine exhaust. In addition, the vent drain lines were cut off and capped abov2 the pool.
. . ~ - - - 2 l- _
J WTELEINNE 1 SBMCES Revision 2 2.2.2 Modifications y Strengthen the Structure Torus Support Saddles and Anchor Bolts - Support saddles were added under each 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. The saddle is constructed of lh-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 small clearance to allow for normal radial growth of the torus due to tempera-ture changes. Torus Support Column Anchors The uplift capacity of the torus was also increased by the 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 each column location. The bolts are set eight inches into the concrete. . Downcomer Tie-Rods and Vent Header Gussets The downcomer tie-rods and vent header gussets are illus-trated in Figure 2-13.
-.i-i-
h Report yp qq Revision 2 SERVICES 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 are prevented from sliding on the downcomers by welded sto!! -lips both above w telow the clamp. 1 The gussets between the downcomers and vent header are to reduce local intersection stresses due to chugging lateral loads on the down-i comers. They are constructec fIom -inch thick SA 516 GR 70 steel plate, and are welded to the vent header and downtomers. Catwalk and Handrail The catwalk and handrail at Vermont Yankee have been removeo from 14 of the 16 bays and only remain in the two bays where the access hatches exist. The two remaining bays have been mooified as shown in Figures 2-14 and 2-15. These modifications include: i e Replacement of original support columns with four inch, schedule 80 pipe columns, e Addition of 4, four inch schedule 80 diagonal braces in each bay, e New cable handrails and support posts. e Addeo horizontal stiffeners and grating hold down clips. All of the platform grating is removed during normal plant operation except for a 3' x 3' square under the access hatches. l
5 91 YW Revision 2 EPGEERING SERVICES ( L Internal Spray Header Supports 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
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I react upward loads. These were modified as shown in Figure 2-16 to react the upward loads associated with pool swell and froth. Drywell-to-Wetwell Vacuum Breakers New aluminum discs with higher strength and better impact characteristics were installed in the drywell-to-wetwell vacuum breakers. This change was the result of information found in the Mark 1 Full Scale Test Facility during the testing.
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Technical Report TE DE TR-5319-1 ENGrEERING SERVCES Revision 2 i KEY FOR FIGURES 2-3 AND 2-4 Modification Completion Date
- 1. Column Reinforcement and Tie-Down 1977
- 2. Mitred Joint Saddle 11/80
- 3. Downcomer Ties 11/80
- 4. Vent Header Deflector 11/80
- 5. Vent Header /Downcomer Stiffening 5/83
- 6. Vent Drain Line 11/80
- 7. Monorail 10/79
- 8. Catwalk Grating and Handrail 5/83
- 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
- 13. RCIC Turbine Exhaust Truncated 11/80
- 14. Temperature Monitoring System 5/83
- 15. Modification of Drywell to Wetwell Vacuum Breakers 11/80
- 16. Saddle Anchor Bolts 11/80
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l Technical Report TN l TR-5319-1 Revision 2 N SS{\/ ICES l 3.0 CONTAINMENT 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 containment structures was accomplished using the computer models shown in Figures 3-1 to Figure 3-4. The detailed shell model shown in Figure 3-1 was used to calculate the effects of all loads on shell stress, as well as all symmetric loads on the support and anchor system. The beam model shown in Figure 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). The detailed finite element model shown in Figure 3-1 simulates one-half of the non-vent bay. It is bounded by the ring girder on one end and the mid-bay point on the other. The vent header system is assumed to be dynamically uncoupled from the shell by the support saddles and is not included in this model. This model was constructed with the assumption that the small offset that exists between the ring girder and mitre joint will not affect results; accordingly, the offset is not included in the model.
Technical Report W F W NE TR-5319-1 ENGNEERNG SERVICES Revision 2 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 as an integral part of the structural analysis. The percent of water mass used is identified in the discussion of each load calculation that follows. 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 and 10. Typical pressure-time history curves are shown in Figures 3-5 through 3-7. (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, full AP and zeroAP. Figures 3-5 through 3-7 show comparative values
Technical Report W F W NE
;gf,a'in 2 BGNEERM SERVICES !
l l and time histories for the two cases. The only difference between the analy- l ses was the input loads; the models were identical. Details of the full load l distribution can be found in References 1 and 10. 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) 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 individu .1 frequency effects to determine total stress at selected elements. The three variations in the C0 spectrum (Reference 1) were evaluated by re-scaling the results of the unit load analysis. 100% of the water mass was used for all C0 analysis. A plant-unique factor 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
1%P W NE I 1 ENGNEERING SERVICES l Revision 2 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 the SRSS 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 reported and discussed in References 6, 14 and Appendix 2 of this report. l 3.2.3 Chugging 3.2.3.1 Pre-Chugging & IBA/C0 (4.5.1.2 & 4.4.1) The pre-chug load was evaluated for both the sym-metric and asymmetric distribution described in Reference 1. Results for the symmetric pre-chug analysis were also used for IBA/C0 as described in para-graph 4.4 of Reference 1. Results for symmetric pre-chug were developed 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 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) Post chugging is defined as a spectral load across a wide band, similar in nature to the CO, 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 stress:
Technical Report TN TR-5319-1 DJGREBMG SBNCES Revision 2 Maximum Shell Stress Shell Membrane Stress Pre-Chug 1284 psi Post Chug 776 psi
- 1. Based on frequencies to 30 Hz - sum of 4 maximum +
l 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 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. No further post chug analysis was done for the shell. This position was also influenced by the fact that post chug stresses were very small . 3.2.4 SRV Discharge Calculation of stresses due to SRV line discharge pressures, 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 l guidelines given in Reference 2. In these tests, pressure amplitude and l frequency were recorded and compared to calculated values for the test condi-tions. Factors were developed that related test to calculated values for both amplitudeandfrequency(seeAppendix1). These factors were then applied to calculated load values for other SRV conditions; the structural analyses was
1 Technical Repnrt TR-5319-1 WM , I B4GrEBtNG SERVICES performed using these adjusted values. Appendix 1 discusses the in-plant test and analysis in more detail. A typical set of SRV shell pressures is shown in Figure 3-8. l The method of modeling the water mass in the SRV computer model was the subject of extensive study in this program. Initial attempts to I reproduce measured stresses by applying measured pressures to the computer models failed badly. Af ter considerable study of the nature of the SRV phenomena itself, and the differences between 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. 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. For the thermal analysis, conduction into the columns and saddles from the torus was considered. Convection from the columns and 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. 3.2.6 Seismic 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.
Technical Report TN TR-5319-1 Revision 2 N SSMCES 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 the 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 1 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: e It applies the maximum stress concentration (4.0), recognized by Section III of the ASME Code to all elements (Reference 11). and e It adds the maximum absolute stress for each load case even tlough 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 x , S y , Sxy) for the free shell, l local shell, and the supports.
Technical Report WM TR 5319 1 SENICES pev
- 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-l plete the fatigue 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 whicn actually exist. I
Technical Report 'RTri mYNE 1R-53l9-l evi, ion 2 ENGINEERING SERVCES 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: DBA.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 l Bending Element 19 l Stress Range Local Shell 28,311 psi 69,900 psi Element 147 l
Technical Report IR-5319-1 ENGNEERING SERVICES Revision 2 Compressive Buckling - Acceptable (see below) l l j Compressive Buckling - Reference 13 discusses the results of analy-i 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 f ailure; but, that a safety factor of 7 still exists for an applied SRV pressure of +29.3/-22.6 psi. The maximum l 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 (0AP) + Weight
Technical Report TN TR-5319-1 Revision 2 N SBNCES For the column-to-shell weld, load case 25 controls: Pool Swell (fulltiP) + Seismic (SSE) + SRV + Weight For these load cases, the following controlling conditions were identified: SUPPORT COLUMN - CONTROLLING AXIAL CONDITION LOAD CONTROLLING ACTUAL ALLOWABLE COLUMN DIRECTION CONDITION FACTOR FACTOR i Inner Down Axial + Bending .42 1.0 Outer Down Axial + Bending .54 1.0 COLUMN-TO-SHELL WELD LOAD CONTROLLING ACTUAL ALLOWABLE LOCATION DIRECTION STRESS STRESS STRESS Inner Down Shear 15.08 K/in 17.06 K/in Outer Down Shear 16.31 K/in 17.06 K/in r l COLUMN TIE-DOWN STRUCTURE LOAD ACTUAL ALLOWABLE LOCATION DIRECTION LOAD LOAD Inner Column Up 97 K 240 K Outer Column Up 135 K 240 K 1
Technical Report Y TR-5319-1 94GtEstNG SENICES . Revision 2 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 (OAP) + 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-die and the torus shell result from the downward loads of case 25: l Pool Swell + (full AP) + SRV + Seismic (SSE) + Weight < Controlling conditions are: SADDI.E 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) 1
Technical Report it' F W NE TR-5319-1 ENGNEERING SERVICES Revision 2 SADDLE-T0-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 l largest lateral load. This is case 15 which includes:
4 Chugging + SRV + SSE l All three of these loads have been selected to produce the highest lateral load on one earthquake restraint; contributions from the l individual 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 l STRESS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS Weld at Base Shear 2,379 psi 21,000 psi of Tie Plates 4 l l
Technical Report it' F W NE TR-5319-1 ENGINEERING SERVICES Revision 2 i L 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: Pool Swell (full AP) + Weight + SSE + SRV and load case 21 for the saddle anchors 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 f MAXIMUM MAXIMUM 0F LOCATION LOAD CAPACITY SAFETY Outside Column 76.7 K 312 K 4.07 SADDLE ANCHOR BOLTS ACTUAL FACTOR MAXIMUM LOAD OF LOAD CAPACITY SAFETY 48.78 K/ bolt 263 K/ bolt 5.41
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Revision 2 1 l Q' ( , 4.0 VENT HEADER SYSTEM -
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> The vents head'e'r system, as def irie'd in tnis section, includes the
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f 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 dynamic analysis and provided detailed stress gradient information in the downcomer/ vent header intersection region. The second vent header model is the beam model shown in Figure 4-4. This model represents a full vent bay, complete with vent pipe and downcomers; j as well as a half non-vent bay on either side. It was used to determine I 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 following elements: i e Vent Header Support Columns e Vent Pipe / Vent Header Intersection e Vent Pipe /Drywell Intersection e Vent Header Mitre Joint e Main Vent Pipe 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: 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 - Reference 16 (Bijlaard) e Mitre joint - used stress multipliers from detailed shell model (Figure 4-1) e Main vent pipe - hand analysis
O Technical Report YME TR-5319-1 ENGNEERNG SERVICES Revision 2 I L The beam model used a stiffness representation of the VP/VH inter-section taken from Reference 7. Attachment stiffness between the vent pipe and drywell was calculated using References 17 and 18. Pool swell water impact on the vent header deflector was calculated 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 I 4.3.1 Pool Swell Loads l jl 4.3.1.1 Pool Swell Water Impact J t Analysis for stresses due to pool impact and crag' was done using both computer models. Determination of shell stresses was done with the detailed model in Figure 4-1. For this analysis, force time histories ~oased on QSTF test data were used (References 4 and 10). These time histories were applied at 100 nodal points on the shell model and the dynamic response of the structure was calculated. Relative timing between loading, (Reference 1) was ; ma'intained to preserve accurate representation of longi'.udinal, and circum- l ferential wave sweep. Stresses in the vent reader /dodncomer intersection, as well as in the free shell areas, were taken diMctly from this model. Stres-ses in the downcomer tie bars were also taken from this model.~ Analysis'was done for both full and zero e.P impacts. m 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 and reducing them to nodal forces. Approximately e !< n
L Technical Report WME Revision 2 ENGBEERNG SERVICES 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. 1 4.3.1.2 Pool Swell Thrust (4.2) Pool swell thrust forces are defined as dynamic forces at each bend or mitre in the vent system, and are a consequence of the flowing internal fluids. Analysis for these loads was done using the beam l model and applying the loads statically. This is consistent with the slow s nature of the applied pressure forces. 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)
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The vent header support columns are loaded by for-ces f rom LOCA-jet and LDCA bubble drag. By inspection, it was concluded that LOCA-jet loads would not combine with water impact on the vent system due to dif ferences in timing and, therefore, would not contribute to the maximum stress calculations - LOCA jet forces were not considered further. LOCA bubble forces were calculated and the maximum normal components (radial and longitudinal) were applied simultaneously to conservatively bound the bending moments on the support column. These peak values were applied statically at the midpoint of the column. Stress calcu-lations were done by hand.
i h ( Technical Report YN TR-5319-1 ENG4EERNG SERVCES ) 1 Revision 2 L 4.3.2 Chugging Load _, 4 5 4.3.2.1 Downcomer Lateral Loads (4.5.3) Reference 1 identifies downcomer lateral loads as static equivalents with random orientation in the horizontal plane. The major consequence of this loading is to produce 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 mode 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. I 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 1. 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 pairs 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 j described in Section 4.2. i 4.3.2.3 Internal Pressure (4.5.4) Three vent system internal pressures exist during-chugging. They are:
Technical Report YN TR-5319-1 Revision 2 N SERVCES l \ l e Gross vent system pressure - a .7 Hz oscillat-s ing pressure with a maximum value of 5.0 psi. This pressure acts on the entire vent system. 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. e Acoustic downcomer pressure oscillation - a 40-50 Hz pressure at 13 psi that produces only ; hoop stress in the downcomers.
) 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-ated with pool swell pressures were used in all combined load cases involving chugging pressures; this produces conservative results. 1 l
4.3.2.4 Submerged Structure Drag (Support Columns only) Examination 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 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 bars are already included in the Downcomer Lateral Loads, which were based directly on test data.
i Technical Report TN TR-5319-1 DJGDEstNG SERVCES Revision 2 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.
The horizontal component of the C0 downcomer load i 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 will bound C0 horizontal loads. The C0 downcomer load aise produces a vertical 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 downcomer load (horizontal and vertical components) would still be bounded by CH lateral loads.
Based on this, CH lateral load results were con-servatively used in all load cases in place of C0 downcomer loads.
Technical Report SPT M NE TR-5319-1 ENGNEERNG SERVICES Revision 2 4.3.3.2 Ve_nt 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 f act that the ring modes of the structure are very high. 4.3.3.3 Thrust Forces (4.2) 3 Stresses resulting from C.0. thrust forces were conservatively taken from the pool swell thrust calculations and applied to all C0 load cases (paragraph 4.3.1.2). 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. 1
Technical Report TN TR-5.319-1 ENG4EERNG SERVCES Revision 2 4.3.5 SRV Loads 4.3.5.1 SRV Drag Loads An SRV discharge produces drag loads which act 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 Appendix 1. 4.3.6 Other Loads Deadweight and seismic stresses in the vent system were calculated using the beam model of Figure 4-4. I Seismic stresses were calculated ay statically applying the acceleration values in Table 3. Thermal stresses were determined for the steady state application of maximum vent system temperature, using hand analysis. 4.4 Results and Evaluation
.i 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 combinations listed in Table 1 have been considered.
l ) s Technical Report YME TR-5319-1 ENGNEERING SERVICES I - Revision 2 l 1 I j 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 header / vent pipe inter-section occurs as a result of pool swell water impact. The controlling load condition is case 25 in Table 1 which includes: Pool Swell (fullAP) + Thrust + Seismic (SSE) + Weight + SRV Pressure
s' Technical Report TME TR-5319-1 ENGNEERING SERVICES Revision 2 ACTUAL ALLOWABLE TYPE OF STRESS STRESS STRESS Combined Maximum Stress 28,930 psi 28,950 psi This load case was formed using a OAP 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 combination for the vent header support columns and the clevis joints at each end is case 25, Table 1. This case includes: 1 Pool Swell (full AP) + Seismic (SSE) + Weight + Thrust + SRV Controlling stress in the support column is: ACTUAL ALLOWABLE TYPE OF STRESS STRESS STRESS Axial in Column (tension) 12,028 psi 16,380 psi Controlling stress in the clevis joint at the end of the support column is: STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS Clevis Plate Shear 12,629 psi 13,840 psi 4.4.4 Downcomer Tie-Bars and Attachments The controlling load combination for stresses in the down-h 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.
s Technical Report TEE TR-5319-1 ENGNEERING SERVICES Revision 2 I The controlling case includes: Pool Swell Impact (full AP) + SSE Seismic + SRV + Weight + Pressure + Thrust The controlling stress is: ACTUAL ALLOWABLE LOCATION STRESS STRESS STRESS Tie-Bar Clamp Bending 16,800 psi 22,240 psi 4.4.5 Vent Header Deflector and Attachments i The major load on the vent header deflector occurs as a result of pool swell water impact. The controlling load condition is case 19 in Table 1 which includes: Pool Swell (fullAP) + SSE Seismic + Weight + Thrust 1 The controlling stress in the deflector is: STRESS ACTUAL ALLOWABLE LOCATION TYPE VALUE VALUE Center of Bending 10,000 psi 57,400 psi the Long Span 4.4.6 Main Vent /Drywell Intersection 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: Pool Swell (0AP) + Seismic (SSE) + Weight' + Thrust + Pressure
I 53 9 1 W Revision 2 -66_ 94GNEBWJG SBWICES The controlling stress is: ACTUAL ALLOWABLE TYPE OF STRESS STRESS STRESS Local Membrane 23,108 psi 28,950 psi The effects of all loads from the vent system, and the pres- { sure load were considered using Reference 16. Stresses due to seismic and thermal response of the drywell have been included. 4.4.7 Vent Header, Main Vent & Downcomers - Free Shell Stresses It was established by inspection of the computer results l 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.
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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 AP) + Thrust + Seismic (SSE) + Weight
+ SRV + Pressure ACTUAL ALLOWABLE TYPE OF STRESS STRESS STRESS Combined Maximum Stress 27,137 psi 28,950 psi 9
s Technical Report W M NE TR-5319-1 ENGNEERING SERVICES Revision 2 r I 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 cycles reach these maximum values. The duration of the major loads in this analysis is 900 seconds, the length of chugging associated with an SBA/IBA event. The procedure used in this analysis consists of the follow-ing steps: e For a given load and component, locate the highest I stress, e For this location, establish the stress range, o Repeat this process for all other loads in the load combination. e Add th stress ranges for all loads. e Multiply this total stress range by the appropriate stress intensification factor, e Calculate stress intensity and determine the allow-able number of stress cycles, e Determine the usage f actor, using the methods of Reference 11. The f atigue evaluation was performed for all high stress areas in the vent system. The major load, contributing to the f atigue evalua-I l l I, e
, Technical Report 7t' M NE TR-5319-1 ENGNEERING SERVCES Revision 2 tion is chugging following a DBA. The controlling load case is case 21 in l Table 1, which includes: 5 Chugging (DBA) + Seismic (SSE) + SRV + Weight The controlling usage factor for the vent system is: VENT SYSTEM FATIGUE RESULTE ACTUAL ALLOWABLE USAGE USAGE LOCATION FACTOR FACTOR At the VH Support .76 1.0 s
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Technical Report SPTA RWNE TR-5319-1 ENGNEERING SERVICES Revision 2 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 added because it is local to a specific area not affected by other stresses. 5.2 Computer Models Two computer models were used as a part of the ring girder analyses; 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 strutural 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-I
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Technical Report SPT31FTVNE TR-5319-1 ENGNEERING SERVICES Revision 2 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 girder 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 Analysis 5.3.1 Loads Applied to Shell As stated, the ring girder stresses for all symmetric loads applied to the shell were taken from the 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) i i
Technical Report W TF1FrWNE - TR-5319-1 ENGNEERING SERVCES Revision 2 (d) SRV Discharge * (e) Seismic (f) Deadweight, Thermal and Pressure ( *SRV discharge is conservatively assumed to be a symmetri-cally applied load for shell analysis. 5.3.2 Drag Loads 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 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 of this, no dynamic load factors were applied to the static analysis results (DLF = 1.0). Drag loads considered were: (a) Pool Swell Bubble (b) Pool Swell Jet (bounded by a) (c) SRV Jet (d) SRV Bubble (e) C0 including FSI (bounded by 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 based on data collected from in-plant tests. A discussion of the in-plant tests and the use of drag data from the:,e tests is given in Appendix 1. Calculation of ring girder drag loads, due to condensation 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 I
s Technical Report WTA WNE TR-5319-1 ENGINEERING SERVICES pev calculation. It produced drag volumes for the ring girder of approximately half of those that the NUREG 0661 procedure would have produced. A discussion The FSI drag calculation was based on of this is included in Appendix 3. - 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. l l
l 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, included 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. I e Vent header support column reaction loads during pool swell. e Vent header support column drag loads. l l I
s e Technical Report SeTFI FTVNE . TR-5,319-1 ENGNEERING SERVICES Revision 2 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. F 5.4 Results & Evaluations L 5.4.1 Ring Girder Web & Flange The controlling load combination for the ring girder web and f flange is load case 16 of Table 1; this includes: Pool Swell (0AP) + Pressure + Weight The controlling stress is: STRESS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS Web Membrane 15.1 ksi 19.3 ksi Flange Membrane 10.6 ksi 19.3 ksi 5.4.2 Weld to Torus Shell The controlling load combin'ation for the shell weld is load case 21 in Table 1; the controlling stress is: STRESS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS Column Region Shear 5.80 K/in 8.53 K/in (Inside) Column Region Shear 8.03 K/in 8.53 K/in (Outside) Saddle Region Shear 6.94 K/in 8.53 K/in
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Technical Report SPTF1 s ry(m ItPsit 2 ENGNEERING SERVICES s 6.0 TEE-QUENCHER AND SUPPORT The following results for the tee-quencher and supports are conservative due to the combined effect of several f actors, three of which are: e The calculational 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 l plant. l e For submerged drag loads, individual frequency components were added to produce maximum stress without regard to load direction. t l The effect of these conservatisms varies among stresses, but can be significant in some cases. 6.1 Structural Elements Considered l 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.
l The structural elements considered in this section include: e The quencher. e The submerged portion of the SRV line. e The quencher support beam and attachments. 1
Technical Report TN ( TR-5319-1 ENGNEERING SERVICES Revision 2 6.2 Computer Models The computer model used in this analysis is shown in Figure 6-1. This is a STARDYNE beam model which represents all piping and struc-ture between the drywell jet deflector and the ring girder. For these analy-ses, the ring girder was assumed rigid and the vent pipe penetration was represented by a stiffness matrix which was developed from a finite element 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. ] This model was used for both static and dynamic analysis. 6.3 Loads Analysis 6.3.1 SRV - Load The calculation of stress due to SRV blowdown was done by applying the dynamic loads to the computer model and calculating the time-history response of the system. The applied loads included both the blowdown forces on the piping and the water clearing loads at the quencher. The controlling condition was for a second, multiple valve actuation af ter an IBA/SBA break, with steam in the drywell (SRV case C3.3). This case produces a high reflood level at the time of the second actuation and produces maximum load on the support system. Loads for this analysis were developed using G.E. computer program RVFOR-04. 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 I
s Technical Report YF M jR-5319-1 ENGNESUNG SERVICES dynamic load factor of 2. It was clear from this analysis that combined pool i swell events would not control stresses - no further analysis was done. s - 6.3.3 Chugging Loads Dynamic analysis of the quencher and support system was done t t 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 all steady-state frequency excitations from 1-31 hz. Results for individual f load conditions were determined by scaling individual frequency results of the computer analysis by the appropriate pressure amplitude. The mass of the structure used in the computer analysis was adjusted to account for the "added mass" effect of the surrounding water. For FSI and post-chugging analyses, individual frequency components were combined by adding the five maximum frequency contributors to the SRSS sum of the others (see Reference 12 for discussion). The maximum value of each frequency component was used in the combination, regardless of vector direction or time of instantaneous response. FSI loads were calculated by considering the l calculated local accelerations in the pool due to the response of the entire shell. 6.3.4 Condensation Oscillation Loads The quencher and support system are subjected to conden-l sation oscillation drag and C0-FSI drag. Analysis for these loads was based on the same harmonic analysis discussed in paragraph 6.3.3, scaled to the C0 amplitudes. Each of the three C0 spectra shown in Figure 4.4.1-1 of Reference l 1 were considered. l All other discussion from paragraph 6.3.3 for chugging applies to the condensation oscillation analysis, except that the final load was determined by adding the four maximum frequency contributors to the SRSS sum of the others.
l - Technical Report TN TR-5319-1 ENGNEERING SERVICES Revision 2 6.3.5 Other Loads L Calculations of stress due to weight, thermal and seismic loads was done by using the computer model in Figure 6-1 and static analysis. Pressure stresses for the piping and quencher were calculated by hand. 6.4 Results and Evaluation The results reported in this section may be conservative depending on the effect of factors discussed in Sections 1.0 and 6.0 of this report. 6.4.1 Tee-Quencher The controlling stress in the tee-quencher itself occurs in the ramshead between the quencher arms. It is the result of a second SRV actuation after an SBA accident - load case 15 of Table 1. It includes: SRV Blowdown (case C3.3) + Chugging Drag + Weight + Seismic + Internal Pressure + Thermal The controlling stress is: I STRESS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS Bifurcated Bending 17,522 psi 24,705 psi Elbow 6.4.2 Submerged SRV Line The controlling stress for the submerged portion of the SRV line occurs in the inclined lines and is a result of load case 15 in Table 1. This case includes:
- Technical Report TE WE TR-5319-1 ENGNEERING SERVCES Revision 2 SRV Blowdown (case C3.3) + Chugging Drag + Weight + 1 Seismic + Internal Pressure + Thermal The controlling stress is: STRESS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS Vertical Bending 14,457 psi 27,000 psi Section Above First Elbow 6.4.3 Tee-Quencher Support 3 The controlling stress that was calculated for the tee-quencher support is the result of load case 15 of Table 1. This case includes: SRV Blowdown (case C3.3) + Chugging Drag + Weight + Seismic + Thermal i The controlling stress for the beam is: STRESS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS At the Brace Bending 20,465 psi 27,000 psi Connection
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Tecnnical Report yg ev n2 ENGrEERNG SERVCES 7.0 OTHER STRUCTURES 7.1 Catwalk The catwalk at Vermont Yankee has been removed except for two sec-tions; one under each access hatch. Each of these sections spans a single bay and consists of a horizontal frame structure which supports an open grating. Each section is supported from an adjacent pair of ring girders by vertical and diagonal columns as shown in Figures 2-14 and 2-15. 7.1.1 Computer Model The computer model of the catwalk is shown in Figure 7-1. It includes all of the load carrying structural members, but does not include the grating or handrails. Loads from these elements are calculated and applied to the frame as forces at the points of attachment. All catwalk analysis was performed on this linear mode?. All analysis used static application of loads, increased to account for dyna-mic amplification, where appropriate. 7.1.2 Loads Analysis Loads analysis for the catwalk was performed for the direct effects of the following loads. Indirect effects due to motion of the ring girder at the attachments points were considered, but judged to be negli-gible. 7.1.2.1 Pool Swell Drag (4.3.4) Pool swell drag loads are produced as the rising pool envelopes the main frame, grating and handrails. Loads on the frame were
Technical Report 1PTA WNE TR-5319-1, ENGNEERNG SERVICES Revision z calculated based on velocities taken from plant unique QSTF movies and the l methods in Reference 1. These were multiplied by two to account for the dynamic effect. Loads on the grating were taken from Section 4.3.4 of Refer-ence 1; these loads already include a dynamic f actor, since they are based on test data. 7.1.2.2 Pool Swell Fallback (4.3.6) Pool f allback loads were calculated and applied in accordance with Reference 1, except in unusual cases where f allback loads exceeded upward loads. In these cases, the maximum values of upward load were used for fallback also. Fallback affects the main frame and grating as well as the handrails. 7.1.2.3 Froth Load (4.3.5) Froth loads have their major effect on the catwalk handrails; and, when applied horizontally, can produce high bending stresses in the vertical handrail members. Froth loads were calculated in accordance with Reference 1, except that the froth 1 influence region was redefined using plant-unique QSTF movies. These movies show clearly that froth 1 loads do not reach the catwalk railing; the analysis was therefore performed with froth 2 loads only. Except e the handrails, the entire catwalk is submerged before froth loads reach this part of the torus; because of this, froth was only considered on the handrails. 7.1.2.4 Drag Loads (Support Columns) The submerged portion of the catwalk support col-umns are subject to loading from drag forces from the following sources:
Technical Report TR-5319-1 TN Revision 2 N SENICES (a) Pool Swell } J (b) SRV Discharge (c) Condensation Oscil11ation (d) Chugging Loads from these sources were calculated and applied to the support columns as static loads. The natural frequency of the support was calculated using hand calculations and compared to the fre-quency(s) of each source. The statically determined stress was then multi-plied by a dynamic amplification factor, developed by considering the worst 1 case frequency ratio and the fact that this is a harmonic loading. 7.1.2.5 Weight and Seismic Loads f Stress due to weight loads were analyzed using static analysis and the computer model shown in Figure 7-1. Seismic loads are small and were considered using hand analysis and scaling static stresses. 7.1.3 Results and Evaluatior. Table 1 allows stresses in the catwalk structure (exc'4uding attachments) to exceed yield; and, in certain cases, to exceed ultimate. Our analysis was based on a linear model and all stresses were maintained at yield or less. Controlling stress and load combination for various catwalk elements are listed here. 7.1.3.1 Main Frame The controlling stress in the catwalk frame occurs in the inboard supporting channel (Figure 7-1). It is a result of the combined condition that includes: Pool Swell + SRV + Seismic + Weight (case 25, Table 1)
h Report 9gg Revision 2 .g3_ ENGrEERNG SERVICES The maximum stress value is: TYPE OF ACTUAL ALLOWABLE STRESS STRESS STRESS Bending + 24,400 psi 40,600 psi Axial 7.1.3.2 Support Columns, Support Diagonal Braces & End Joints The controlling load case for the support system and end joints includes: l Pool Swell + SRV + Seismic + Weight (case 25) i Controlling stress is: TYPE OF STRESS ACTUAL ALLOWABLE STRESS LOCATION STRESS STRESS Bending Outbaord 18,445 psi 42,000 psi Diagonal Brace 7.1.3.3 Welds to Ring Girder The controlling load combination for this stress is also case 25: Pool Swell + SRV + Seismic + Weight
s ~ I 1 Y Revision 2 B4GNEstlNG SERVICES For this condition, stresses are: TYPE OF ACTUAL ALLOWABLE STRESS STRESS STRESS Tensile 18,264 psi 42,000 psi 7.2 Internal Spray Header l The internal spray header is attached to the ring girders and to a penetration on the shell. It is located at the top of the torus, above the j vent header (Figure 2-16). s 7.2.1 Computer Model The computer model used to analyze the spray header is shown in Figure 7-3. It was constructed to allow determination of stresses in a i typical multi-span area as well as at branch connections. This is a part of a piping system and piping elements were used in the model. All results were obtained through the use of static analysis, with factors applied to account for dynamic response. 7.2.2 Loads Analysis The spray header is high enough in the torus so it does not experience direct water impact-froth is the only pool swell related load that is applied. The motion of the ring girder that results from pool swell loads on the shell was considered but judged to be a negligible input to the spray header. Shell displacement at the nozzle connections was input to the computer analysis.
Technical Report YE E - TR-5319-1 ENGINEERING SERVCES Revision 2 7.2.2.1 Froth Load (4.3.5) Froth loads on the spray header were calculated as outlined in Reference 1. The worst stress condition existed for a vertically applied load. The loads were applied statically to the system (DLF = 1.25). 7.2.2.2 Weight, Seismic & Ring Girder Displacement The effects of weight, seismic and shell displace-ment were all considered by using the model shown in Figure 7-3 and applying f loads and displacments statically.
)
7.2.3 Results and Evaluation The controlling stress for the spray header piping is a result of load case 19, Table 1. This case includes: Froth, Weight, Seismic and Shell Motion The controlling stress is: SPRAY HEADER PIPING STRESS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS Branch Side Bending 25,600 psi 27,000 psi of Tee I.
Technical Report Y TR-531,9-1 ENGNEERING SERVICES Revision 2 ATTACHMENTS TO RING GIRDER STRESS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS At Ring Bearing 10,108 psi 34,200 psi Girder WELDS TO RING GIRDER STRESS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS I 3 At Ring Tension + 4,002 psi 18,000 psi Girder Shear 3 7.3 Vent Pipe Bellows The vent pipe bellows forms the pressure seal between the vent pipe and torus, and allows for relative motion between these parts. It is illus-trated in Figure 7-4. 7.3.1 Analysis Method The bellows are rated by the manuf acturer for differential motion both axially and radially. These ratings are intended to define static differences which occur over a long enough time so that dynamic response of the bellows itself can be ignored. In the present analysis, both ends of the bellows are exper-iencing dynamic motion; one end is controlled by the vent pipe-the other by the torus shell. We expect that the dynamic characteristics of the convoluted bellows should increase stresses over their static equivalents. We also expect that the convolutions will produce complex modes and stress patterns that will not couple efficiently with specific input frequencies; i.e., high dynamic reponse is not expected. Further, the "pogo" and " rolling" modes of
D'n Technical Report inP WNE TR-5319-1 ENGNEERNG SERVICES Revision 2 i the convolutions are non-linear, highly cross-coupled modes that would not be i predicted by ordinary structural codes. Our approach to the bellows evaluation is to compare the maximum calculated difference in dynamic response across the bellows to the manuf acturers allowable. We accept the bellows as adequate for all cases where a large margin exists between predicted input motion and the static capacity, as stated by the manufacturer. 7.3.2 Loads Analysis Calculation of vent pipe motion and torus shell motion was
) '
done as a part of the analysis work discussed in Sections 3.0 and 4.0 of this report. The analysis of the torus shell in Section 3.0 was based on a computer model of the non-vent bay and therefore did not account for the presence of the vent pipe hole, or the heavy shell reinforcement in that area. 4 7.3.3 Results and Evaluation The maximum differential motion across the bellows occurs as a result of case 25 in Table 1; this case includes: Pool Swell Pressure on Shell + Water Impact on the Vent System + Vent System Thrust + Pressure + Weight + SRV + Seismic For this case, the following deflections occurred: MAXIMUM MANUFACTURERS' DIFFERENTIAL STATIC MOTION ALLOWABLE-Axial Compression (in.) .038 .875 Axial Extension (in.) .038 .375 Lateral Motion (in.) .0516 .625 I i
Technical Report SPTF1FrVNE TR-5,319-1 ENGINEERING SERVCES Revision 2 All calculated values are less than 11% of the manuf ac-turer's allowables. We consider that this large difference demonstrates the acceptability of the bellows, especially if we consider that much of the load ) is either static or a single-pulse transient (maximum amplification of 2). s 7.4 Monorail I The monorail is attached to the torus ring girders at about 45 above the water level. It is a non-containment related structure and there-fore in the same category as the catwalk. It is illustrated in Figure 2-17. 7.4.1 Computer Model The computer model used to analyze the monorail is shown in f Figure 7-5. It is a beam model that represents the monorail through 180 of the torus structure. Symmetric boundary conditions were applied to the model to allow for a full 360 representation. All loads were applied to the monorail statically and a dyna mical load factor of 1.12 applied. 7.4.2 Loads Analysis The monorail is high enough over the pool so that it does not experience direct water impact. The only pool swell related load is froth. As with the catwalk, ring girder motion was considered, but judged to be negligible. 7.4.2.1 Froth Loads The monorail is located in the froth 1 region of the torus and was analyzed for these loads, oriented to produce maximum stress. This orientation was 45 to the horizontal. Froth loads were calcu-lated in accordance with the methods of Reference 1, and applied statically to the computer model.
lechnit .il lieport TEI NE TR-5319-1 ENGNEERING SERVICES Revision 2 7.4.2.2 Weight & Seismic Weight and seismic analysis was performed using the model shown in figure 7-5 and static analysis. L 7.4.3 Results and Evaluation / k The combination of froth, weight and seismic SSE (case 19, Table 1) produce the following controlling stresses: MONORAIL BEAM STRESS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS At Support Bending 26,840 psi 36,000 psi 4 MONORAIL ATTACHMENT STRESS ACTUAL ALLOWABLE TYPE STRESS STRESS Axial + 33,340 psi 38,000 psi Bending
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- Technical Report $P M T -5319-1 -105-ENGINEERING SERVCES 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 seppression pool. These requirements are stated in NUREG 0661 (Reference 2) and NUREG 0783. 8.1 Maximum Suppression Pool Temperature 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 assumes no RHR actuation. The result is a bulk temperature of: 210 - 43 = 1670 F at a mass flux rate 2 (42#m/sec-ft 200 - 43 = 157 F at a mass flux rate 2
> 94 #m/sec-f t The results of the bulk temperature analysis for the most limiting case (Figure 8-1) meet the above limit and, therefore, satisfy the require-ments of NUREG 0661. Additional analyses are being contemplated to reduce the l
r N'U!U}"*" -106-YF WE L Revision 2 ENGINEERING SERVICES r L conservatism present in the above calculation but this effort is outside the j scope of the long-term program. I 8.2 Pool Temperature Monitoring System f The NRC criteria also presents guidelines for a monitor-ing system to constantly monitor pool temperature. A monitoring system installed at vermont Yankee which uses a network of thermo-couples, hardwired to a strip chart recorder in the control room has been upgraded to meet the NRC criteria. The system is descri-bed more fully in Section 2.2.1 of this report and is illustrated in Figures 2-9 and 2-18.
L l TR-5315~1 WN Revision 2 -107- Figure 8-1 (Revised) . [ Comparison of T-Quencher Bulk Suppression f Pool Ter perature Limit to Stuck Open S/RV I From 100% Power Transient Responses f . 1-i i a _ . l -
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Technical Report W F W NE TR-5319-1 -108- ENGNEERING SERVICES Revision 2 REFERENCES
- 1. G.E. Report NED0-21888, Rev. 2, " Mark 1 Containment Program Load Defi-nition Report", dated November 1981.
- 2. NRC " Safety Evaluation Report, Mark 1 Containment Long-Term Program",
NUREG 0661, dated July 1980.
- 3. G.E. Report NED0-24583-1 " Mark 1 Containment Program Structural Accept-ance Criteria Plant Unique Analysis Application Guide" dated October 1979.
- 4. G.E Report NED0-21944 "...k Scale 2-D Plant Unique Pool Swell Test Report" dated August 1979.
- 5. G.E. Report NED0-24615 "....k Scale Suppression Pool Swell Test Pro-I gram: Supplemental Plant Unique Test", dated June 1980.
l
- 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. l 8. G.E. Report NEDE-21968 " Analysis of Vent Pipe - Ring Header Inter-section" dated April 1979.
- 9. Deleted.
- 10. G.E. Report NED0-24581, Rev. 1, " Mark 1 Containment Program - Plant Unique Load Definition - Vermont Yankee Generating Station" dated October 1981.
Technical Report '#PTF1 FrVNE I"-5319-l -l 9-Revis30n 2 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
{ Approach Based on FSTF Data for Combining Harmonic Amplitudes for Mark 1 Post Chug Response Calculations", dated May 1982.
- 13. Mark 1 Containment Program Report WE8109.31 " Buckling Evaluation of 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 Design Rules", dated March, 1982, pg. 3.
- 15. Intentionally Blank.
- 16. Welding Research Council Bulletin No. 107, " Local Stresses in Spherical & Cylindrical Shells due to External Loadings", dated August 1965 with March 1979 Revision.
- 17. Welding Research Supplement, " Local Stresses in Spherical Shells from Radial and Moment Loadings", P.P. Bijlaard, dated May 1957.
- 18. "On the Effects of Tangential Loads on Cylindrical & Spherical Shells", P.P. Bijlaard, Unpublished, Available from PVRC, Welding Research Council.
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Technical Report TE M -
-111-SERVICES TR-5319-1 Revision 2 TABLE 2 -
PLANT PHYSICAL DIMENSIONS VERMONT YANKEE TORUS Inner Diameter 27'8" 16 Number of Sections Shell Plate Thickness Vent Pipe Penetration 1.0625" 3-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 1-Beam (12.5" x 1.25" Flange & 10" x 1" Web) 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 Quantity Size Vent Pipe 8 6'9" 1.D. Vacuum Breakers 10 18" 1.D. Vent Header Support Columns 16 pairs 6" Sch. 80 96 2' 0.0., l'-ll\" 1.D. Downcomers Submergence 4.29' Min. - 4.54' Max. Water Volume @ Minimum Submergence 68,000 cu. ft.
Technical Report -112- TM Revision 2 ) L TABLE 3 f PLANT ANALYSIS INFORMATION VERMONT YANKEE Seismic Acceleration Values (G's) OBE 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 AP - 57% Zero AP - 52% Plant Unique C.0. Multiplier (Reference 1)
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lechnical Report' TR-5319-1 -113- ENGNEERNG SERVICES Revision 2 TABLE 4 SRV LOAD CASE / INITIAL CONDITIONS Any One ADS
- Multiple Design Initial Condition Valve Valves Valves 1 2 3 1 N0C* , First Act. A1.1 A3.1 A 2 SBA/IBA,* First Act. A1.2 A2.2 A3.2 3 DBA,* First Act. A1.3 1 NOC, Subsequent Act. C3.1 SBA/IBA, Sub. Act.
C 2 Air in SRV/DL C3.2 SBA/IBA, Sub. Act. 3 Steam in SRV/DL C3.3 (I) This actuation is assumed to occur coincidently with the pool swell event. Although SRV actuations can occur later in the DBA accident, the resulting air loading on the torus shell is negligible since the air and water initially in the line will be cleared as the drywell to wetwell AP increases during the DBA transient.
- ADS = Automatic Depressurization System NOC = Normal Operating Condition SBA = Small Break Accident IBA = Intermediate Break Accident DBA = Design Basis Accident
s ? Technical Report SPTF1 prt(m ' ^- ENGINEERING SERVICES IIvi3siS~n 2 APPENDIX 1 Use of SRV In-Plant Test Data for Analysis Test Data I 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 line and a second about one minute later (hot line). The test sets were about three hours apart to allow for SRV line cool down. The torus was instrumented with a combination of strain and pressure transducers as shown in Figure Al-1. Strain gages were mounted in pairs on both sides of the shell to allow separation of bending and membrane stresses. Additional gages were located on the columns (Figure Al-2), and internal structures (Figures Al-3 and Al-4). 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 collected. 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. Use of Data - Applications 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.
Technical Report TN Revision 2 Use of Data - Shell & Support System Analysis Evaluation of shell stress and support system loads due to SRV actuation 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) condition and to develop correction f actors between these measured results and values pre-dicted by generic analytical methods. The steps involved re these:
- 1. Determine maximum average shell pressure, average frequency and waveform for the four cold tests.
- 2. Calculate these same quantities for the test conditions using the generic computer programs (QBUBS 02).
- 3. Calculate calibration f actors relating predicted-to-ubial pres-sure and predicted-to-actual frequency.
- 4. Calculate predicted pressures and frequencies using the generic coniputer program, 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-sure transient, as measured in the test, is affected proportionally by the frequency correction and used as the basis for all computer model loading.
Verification of Computer Model The test data was also used to verify the accuracy of the computer model. This was done by the following method:
- 1. The computer model was loaded with the measured shell pressures.
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- 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 same points.
Correlations for shell stress were excellent - generally within 5%. Correlations to column loads were not so good - generally off by about 50%. 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-I 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 sh6wed these results: i l _ _ _ _ - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - J
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- 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.
l The data collected from Vermont Yankee was evaluated along with the data I collected by TES in three other in-plant tests. The matrix of data collected 6 is as follows: 1 Ring Catwalk Vent Girder Supports Column Downcomer (Pressure) i Millstone X X X l 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 bz.* 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 bz.
Technical Report W F W NE 1R-5319-1 Al-5 ENGNEBt!NG SERVICES Revision 2 We conclude from this that the structures involved are responding to a fairly L 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 stresses measured in the test, when applied uniformly to the submerged area. These static pressure values were plotted against distance from the quencher j 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 I pressures for the two cases, for application to structures under different loading conditions.
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Technical Report TF M TR-5319-1 A2-1 ENGNEERING SERVICES Revision 2 APPENDIX 2 Discussion of 32 Hz Frequency Cut-off for Condensation Oscillation and Post Chug Analysis TES made the decision to limit C0 and Post Chug response analysis to frequencies below 32 Hz early in the program. 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: For condensation analysis.
- 1. Load Definition - A PSD study of the C0 pressure data showed that 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 cut-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:
IethniuI ReporL NE NN TR-5,319-1 A2-2 ENGNEERING SERVICES Revision 2 / Numbe2 0f* Number 0f Max. Number of GX GXp2 Value Frequencies 1000 > 2000 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 Hz are relatively insignificant. They provide a strong indication that the 10% worst-case unconservatism discussed above will be greatly reduced by the selective 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.
s I Technical Report SPTFI FrVNE TR-5319-1 A3-1 ENGNEERING SERVICES Revision 2 i APPENDIX 3 C0/CH Drag Loads for Ring Girder Analysis 4 { TES did not follow the calculational methods of NUREG 0661 (Reference 2) for calculation of C0/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 VT times the radius of the circumscribed circle. It also specifies that the drag forces be increased by an additional factor of 2 for structures attached to the torus shell, to account for wall interference. i Application of the NUREG criteria produces a factor of 4 multiplier for drag force for flat plate structures in the fluid; and a f actor 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 SPTF1 pry ( E
^- ENGNEERING SEFNICES evision 2 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.
f Both References 1 and 2 present results as a function of the VT/D ratio where:
)
V = maximum velocity v T = period of flow oscillation D = diameter Keolegan and Carpenter show the effective hydrodynamic mass coefficient foraplatevariesfromamaximumof4ath=125to1atVT/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 at h = 15 to a minimum of 1.65 at VT/D = 0. 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 drag loads on the ring girder. For this structure, under this load, VT 7 ratios are near zero 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 loads on the ring girder. i
. W TF1FrWNE 1 A3-3 ENGNEERING SERVICES j Revision 2 The TES method to calculate these drag loads on the ring girder used the L
same references as above (A3-1 and A3-2), but accounted for calculated values of rather than the values corresponding to the maximum increases. Con-sideration of the actual f ratio for wall interference led to an interference factor of 1.65 (instead of 2). i Low values of hsuggest that the theoretical hydrodynamic mass coefficient for the ring girder is appropriate. The theoretical coefficient i 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 f actor used was related to the NUREG multiplier by: 2.0 x 1.65 , ,41 4.0 2.0 The factor used by TES was .41 x the NUREG 0661 factor.
Technical Report TN TR-5319-1 A3-4 ENGBEStNG SBNCES Revision 2 ) s REFERENCES A3-1 Keolegan and Carpenter, " Forces on Cylinders and Plates in a Oscillating Fluid," National Bureau of Standards, Vol. 60, No. 5, May 1959. A3-2 Sarpkaya, " Forces on Cylinders near a Plane Boundary in a Sinusoidally i Oscillating Fluid", Journal of Fluids Engineering, September 1976. 6 I am i mummiemi u - i m F'
Technical Report W F W NE TR-5319-1 ENGNEERNG SERVICES Revision 2 A4-1 APPENDIX 4 NRC REVIEW COMMENTS & RESPONSES This report was reviewed by NRC consultants during 1983. As a result of these reviews, meetings were held and additional information was exchanged. This appendix summarizes the review process and includes copies of all rele-vant information resulting from the review. Part of the review process included questions on torus attached piping systems, including main steam safety relief lines. These analyses were not reported in this 5319-1 report but the questions are included in this appen-dix, along with the other information. TES report 5319-2 documents the piping eqalysis. The event sequence that made up the review process is as follows: Review by Franklin Institute
- 1. April 25, 1983 - Questions were received from Franklin Institute.
These are included in Section A4-1 of this appendix.
- 2. June 23, 1983 - A written response was provided to the Franklin Institute questions. These are included in Section A4-2 of this appendix.
- 3. August 9,1983 - A meeting was held at the Yankee Atomic offices in Framingham, Mass. with representatives from Franklin Institute, the NCR, TES and Yankee Atomic. The meeting was to discuss the TES responses above, as well as to review the methods used for attached piping analysis. The presentation handout used at the meeting is included in Section A4-3.
=1 Revision 2 e- A4-2
*mu-E BC#EstlNG SERVICES
- 4. August 18, 1983 - At the August 9 meeting, a request was made for additional information on the calculation of usage factors for f atigue analysis. This request was answered by a TES letter on August 18, 1983; it is included in Section A4-4.
This completed the review process by Franklin Institute. \ Review by Brookhaven National Laboratories 1
- 1. June 13, 1983 - Questions were received from Brookhaven National Laboratories. These are included in Section A4-5 of this appendix.
- 2. July 26, 1983 - A meeting was held at the Yankee Atomic offices in Framingham, Mass. with representatives from Brookhaven, the NRC, TES and Yankee Atomic. The meeting was held to discuss responses to the questions above. A copy of the presentation handout from the meet-ing is included as Section A4-6.
- 3. September 2,1983 - The July 26 meeting ended with requests for addi tional information in some areas. A written response to these requests was provided on September 2, 1983. This response is included as Section A4-7 of this appendix.
This concluded the Brookhaven review, as well as the NRC review of the document. I
R 1 W P W NE Revision 2 A4.1-1 ENGeEERNG SERVICES t i
)
i REVIEW BY FRANKLIN INSTITUTE (FRC) A4.1 Review questions received on April 25, 1983
_" TF WE 94GNEstlNG SENICES Revision 2 A4.1-2 s ITEM 1 i Provide a summary of the analysis and the results for the following penetrations: e Vent pipe torus intersection e Vacuum breaker line and RCIC torus penetration ITEM 2 1 Comment on the effect of the neglected loads indicated on page 66 of Reference 4 on the stress results for the drywell-to-vent penetration. ITEM 3 Provide evidence that the fatigue criteria for the bellows, as required by Paragraph NE-3365-2, Section III of the ASME B&PV Code, are met. ITEM 4 Provide a summary of the analysis with regard to the vacuum breaker valves; indicate whether they are considered Class 2 components as required by the criteria (1). ITEM 5 Provide analyses of the piping systems not included in this report. ITEM 6 Provide details of the construction of the SRV line as it exists in the Vermont Yankee plant, specifically in the region of the elbow support, if any. l
) Technical Report WTptFrt(E ev on 2 A4.1-3 ENGNEstNG SERVICES ITEM 7 l Describe the end conditions assumed for the beam model of the vent header deflector shown in page 4-5, how these were derived, and the sensitivity of maximum calculated stresses to boundary assumptions. ITEM 8 Provide a detailed sketch of the actual diagonal brace-catwalk attach-ment, together with its stress analysis results. 1 ITEM 9 Provide the results of the buckling analysis, including the margin of safety for the catwalk components, i.e., the 4-inch diameter schedule 80 pipe supports and the 2-inch pipe brace. ITEM 10 Provide full justification for the stress values shown as representative of those that may occur in the containment shell mitre joint. Establish limits of maximum possible error. ITEM 11 Provide a list of the component materials and their corresponding metal temperatures used for the stress limit selection. ITEM 12 Indicate whether each torus attached piping and its supports have been classified as Class 2 or Class 3 piping, Class 2 or Class 3 component sup-ports, and essential or non-essential piping systems. Also, indicate whether a pump or valve associated with the piping mentioned above is an active or inactive component, and is considered operable. l I l
Technical Report
" 97 qq Rev s on 2 A4.1-3 SERVICES ITEM 13 f
With reference to Table 1 of Appendix B, indicate whether all loads base been considered in the analysis and/or provide justification, if any load has been neglected. ITEM 14 Provide a summary of the analyses for the new modifications yet to be supplied; these include items 5, 6, 10, 12 and 15 of the key for Figures 2.3 I and 2.4 of Reference 4. In addition, if the final configuration of the 1 catwalk is to be changed, update the analysis accordingly. ITEM 15
)
Provide details of fatigue analysis for piping systems. Indicate whether the f atigue usage f actors for the SRV piping and the torus attached piping are sufficiently small that a plant-unique f atigue analysis is not warranted for piping. The NRC is expected to review the conclusions of a generic presentation (6) and determine whether it is suffi-cient for each plant-unique analysis to establish that the expected usage f actors for piping are small enough to obviate a plant-unique f atigue analysis of the piping. ITEM 16 Submit a summary of the analysis for the miscellaneous internal piping. ITEM 17 The ASME Code provides an acceptance procedure for computing f atigue usage when a member is subject to cyclic loadings of random occurrence, such as might be generated by excitations from more than one type of event (SSE and SRV discharge, for example). This procedure requires correction of the
Technical Report 3p g /g n2 A4.1-4 ENGtEBWJG SEMCES stress-range amplitudes considered and of the associated number of cycles in order to account for the interspersion of stress cycles of unlike character. State whether or not the reported usages reflect use of this method. If not, indicate the effect on reported results. ITEM 18 k Justify the reason for not considering skew symmetric boundary condi-tions in the analysis of the torus shown in Figure 3.1. Evaluate the effect of the thus neglected modes. ITEM 19 Specific comments addressing the method of summation used and its com-pliance with the probability of non-exceedance (PNE) criteria of 84% stated in para. 6.3b of Reference 1 should be incorporated into the text. ITEM 20
)
Provide justification for analyzing only one SRV discharge line, as shown in Section 6.0 of Reference 4. Indicate whether all discharge lines are identical in configuration to the one modeled, and whether the model investi-gated is conservative enough to represent all lines. ITEM 21 Submit a sumary of the analysis for the vacuum breaker and its penetration. ITEM 22 Justify that the 45 model of the vent header and downcomer used in the analysis is adequate to meet the intent of the criteria which requires at least 180 .
I 1 TE M Revision 2 A4.1-5 BIGNEstNG SENICES Justify the reasons for not considering skew symmetric boundary condi-l tions to evaluate the effect of the resulting modes. L ITEM G1 Describe fully the procedures used to assess cumulative f atigue damage. l In particular, address:
- 1. Where departures from standard code procedure were introduced.
- 2. How critical points were selected and how stress (or stress inten-sity) ranges were computed.
- 3. Which cyclic loads were omitted, if any, in these computations. For example, were thermal transients given consideration?
- 4. Whether cyclic amplitudes and the associated number of cycles were adjusted to account for the interspersion of cycles of unlike char-acter.
- 5. How the cumulative usage factor was computed.
- 6. What iinpact departures from code procedures have on the margins of safety shown for each component for which cumulative usage was computed.
ITEM G2 Is the method described in Section 4.3.6 of Reference 4 for assessing thermal stress typical of all evaluations made in the report? Please discuss the tacit assumption that either:
- 1. Thermal equilibrium is achieved before other significant mechanical loads are experienced by the structure.
or
( Technical Report TR-5319-1 TME N SBMCES Revision 2 A4.1-5 r 1
- 2. Maximum transient thermal stresses are conservatively bounded by the assumptions made.
REFERENCES FOR APPENDIX B
- 1. NED0-24583-1
" Mark 1 Containment Program Structural Acceptance Criteria Plant Unique Analysis Application Guide General Electric Co., San Jose, CA October, 1979
- 2. NUREG-0661
" Safety Evaluation Report, Mark 1 Containment Long-Term Program Resolution of Generic Technical Activity A-7" Office of Nuclear Reactor Regulation July, 1980
- 3. NED0-21888, Revision 2
" Mark 1 Containment Program Load Definition Report" l General Electric Co., San Jose, CA November, 1981
- 4. Vermont Yankee Nuclear Power Station Plant Unique Analysis Report, Mark 1 Containment Program Vermont Yankee Nuclear Power Corporation November 30, 1982, TR-5319-1, Revision 0 l
- 5. NRC l " Damping values for Seismic Design of Nuclear Power Plants" October, 1973 Regulatory Guide 1.61
- 6. P. M. Kasik
" Mark 1 Piping Fatigue," Presentation at the NRC Met. ting, Bethesda, MD September 10, 1982
1 R 53 9 Y Revision 2 A4.2-1 DiGBEstNG SERVICES i REVIEW BY FRANKLIN INSTITUTE (CONTINUED) A4.2 TES Response to FRC Questions - June 1983
s P Technical Report 97 gg Rev n2 A4.2-2 N SERVICES RESPONSES TO NRC REQUEST FOR INFORMATION MARK 1 TORUS PROGRAM PLANT UNIQUE REPORT YANKEE ATOMIC ELECTRIC COMPANY l VERMONT YANKEE NUCLEAR STATION JUNE 10, 1983 l l l l
k Tcchnical Report
' ' 9g ev s on 2 A4.2-3 SENICES r
L ITEM 1 i QUESTION Provide a summary of the analysis and the results for the following { penetrations: e Vent pipe torus intersection e Vacuum breaker line and RCIC torus penetration ANSWER The vent pipe is isolated from the torus at their intersection by a large diameter bellows. Therefore, the torus shell is essentially isolated from the vent pipe Mark 1 torus loads. The bellows deflections from the original and Mark 1 loads are approximately 10 percent of the allowable design deflections. Therefore, the combined effect of loads defined by the LDR does not produce stresses greater than 10 percent of the allowable value and no further evalua-tion is required. The TES Torus Attached Piping Technical Report (TR-5319-2) is scheduled for release by Yankee Atomic Electric Company during the f all of 1983. This report will contain a sumary of the analysis and results for all torus attached piping penetrations, including the drywell/wetwell vacuum breaker line and RCIC piping penetrations. ITEM 2 QUESTION Comment on the effect of the neglected loads indicated on page 66 of Reference 4 on the stress results for the drywell-to-vent penetration.
$ 1 Y
Revision 2 A4.2-4 94GtEB54G SERVICES ANSWER The original loads on the drywell-vent pipe intersection, due to seismic and thermal response of the drywell, were not available when the PUA for the torus was issued. The effects of seismic and thermal response using original calculation methods has now been considered with torus loads without exceed-ing code allowables in that area. The next revision to the PUA report will include the Seismic and Thermal in the stress summary. The following is a summary of the local membrane stress to be reported: P1 = 23108 psi 4 28950 psi, Allowable ITEM 3 QUESTION Provide evidence that the f atigue criteria for the bellows, as required by Paragraph NE-3365-2, Section III of the ASME B&PV Code, are met. ANSWER TES has reported that the maximum calculated differential motion across the bellows is less than 10% of the rated movements for the rated cycles (^1000). Based on EJMA (*) fatigue data of unreinforced austenitic bellows, the permissible cycles for the present condition are well in excess of the endurance limit ( + (10)6 cycles). Therefore, the condition does not impact the fatigue acceptability of the bellows. (*) Standard of the Expansion Joint Manufacturers Assoc., Inc. Fifth Edition, 1980.
. _ _ _ - - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ __ b
s Y I 91 Revision 2 A4.2-5 BGEstNG SERVICES ITEM 4 QUESTION Provide a summary of the analysis with regard to the vacuum breaker valves; indicate whether they are considered Class 2 components as required by the criteria (1). ANSWER l The work recently completed for the wetwell/drywell vacuum breaker valves indicates that they do not actuate during a chugging event at Vermont Yankee. Therefore, no additional analysis beyond the original plant design scope is required at the present time. l The USNRC is in the process of reviewing the Mark I wetwell/drywell l vacuum breaker valve loading transients. Any revisions to the loading tran-sients which may result from this review will be evaluated for the Vermont Yankee vacuum breaker valves when the NRC review is completed. l ITEM 5 QUESTION Provide analyses of the piping systems not included in this report. ANSWER The analysis techniques used, piping stresses, support loads and required modifications will be summarized within TES Technical Report TR-5319-2. This report is scheduled for release during the fall of 1983. l l l
s Technical Report yp gg v 2 A4.2-6 BIGrEBtNG SERVICES k ITEM 6 l [ QUESTION Provide details of the construction of the SRV line as it exists in the ( Vermont Yankee plant, specifically in the region of the elbow support, if any. ANSWER The details of the Safety Relief Valve Discharge Line elbow support and a typical isometric of an SRVDL are included for your review. Note that all four SRVDL's are identical from the vent pipe penetration to the quencher support. ITEM 7 QUESTION Describe the end conditions assumed for the beam model of the vent header deflector shown in page 4-5, how these were derived, and the sensitivity of maximum calculated stresses to boundary assumptions. ANSWER The Vermont Yankee Vent Header Deflector is a continuous structure through the 16 torus bays. Figures 2-6 and 2-7 of the PUA Report (TR-5319-1) illustrate the end connection details. The 16-inch deflector pipe slides into a short stub which is welded to the vertical deflector support plate. This connection arrangement does not allow moment transfer; therefore, analysis was performed assuming each span was simply supported. Figure 4-5 of the same report was intended primarily to show the level of load that is applied to the V.Y. vent header model. It creates a misleading impression regarding the analysis assumptions that were used.
s Technical Report TE WE TR-5319-1 A4.2-/ Revision 2 BIGrE854G SERVICES l The analysis was actually performed for the uniform 2.9 Kip per foot load applied to a simply supported beam 19.5 feet long. This non-vent bay analysis bounds that of the vent bay and was used for both. ITEM 8 QUESTION Provide a detailed sketch of the actual diagonal brace-catwalk attach-ment, together with its stress analysis results. ANSWER We are including, as a part of this package, a set of catwalk drawings which contain the actual diagonal brace-catwalk attachment. Item 9 below discusses stress / buckling results for this structure. A summary of the stress analysis results for other major components appear in Section 7.0 of the PUAR. ITEM 9 QUESTI0; Provide the results of the buckling analysis, including the margin of safety for the catwalk components, i.e., the 4-inch diameter schedule 80 pipe supports and the 2-inch pipe brace. ANSWER The buckling analysis results for the Vermont Yankee catwalk supports are as follows:
- 1. The new vertical support leg, four-inch schedule 80 pipe, has a maximum compressive load of 9.2 K with an allowable buckling capa-city of 132.0 K. The margin of safety is equal to 13.35.
5 afR* port y Revision 2 A4.2-8 MCES
- 2. The new diagonal braces, all four per bay, are four-inch schedule 80 pipe, have a maximum compressive load of 15.3 K with and allowable buckling capacity of 77.6 K. The margin of safety is equal to 4.07.
ITEM 10 QUESTION Provide full justification for the stress values shown as representative of those that may occur in the contair. ment shell mitre joint. Establish limits of maximum possible error. ANSWER Early in the Mark 1 program it was decided that not modeling the four-inch offset strip between the ring girder and mitre joint was technically justified, and, in fact, might produce more accurate results if it was omitted. A technical concern that was avoided by omitting this four-inch strip was one related to the substantial change in grid size and pattern. The torus model responds primarily to ring and cylinder modes of the shell. We knew from early experience with this model, that the combination of the thin shell and very high water mass produced sensitive mode shapes. Our concern was that the transition from a very small grid near the ring girder to the much larger grid that would be required on the free shell might affect these sensitive modes and would have an uncertain effect on all results. It was not practical to carry the refined mesh throughout the entire model. In f act, the four-inch wide strip is closer to two inches wide. The four-inch dimension includes half the saddle thickness, the saddle-to-shell weld and the mitre joint weld. We attempted to instrument this region in one of our in-plant SRV tests, but did not have room to install the strain gages. I
s Technical Report s TR-5319-1 TM NE Revision 2 g _g SERVICES In addition to these practical limitations, we believe the assumption is technically justified based on the following information regarding shell l stresses. l The stress analysis that TES has completed confirms the following: e All major loadings on the torus shell are in the form of a uniform or hydrostatic pressure distribution. e The primary membrane stress can be calculated using basic strength of materials and will be maximum at mid-bay bottom dead center of the torus shell. e It follows that the maximum membrane stress cannot occur in the four-inch offset strip of shell in question. e All bending stresses in the region of the ring girder or mitre joint, including the four-inch offset strip may be considered to be secondary because of the gross structural discontinuity, o Since there is no primary bending, it follows that the maximum primary local plus bending stress in this region must be less than the maximum membrane stress and will therefore meet the increased allowable. e The maximum total stress (primary plus secondary) range occurs in the region of the shell adjacent to the ring girder. Since the membrane stresses are reduced in this region, the range of stress would result from the local bending produced by the increased stif-fness of the saddle and ring girder. e The bending of the shell would be symmetric about the two sides of the ring girder if the four-inch offset strip was not present.
s
*E TF WE Revision 2 A4.2-lo DJGrEstNG SERVICES e The torus structure may be considered a beam fixed at the ring girder for purposes of this discussion. The increased stiffness of l the mitre joint should, therefore, result in lower bending stresses
( in the torus shell to the mitre side of the ring girder. A review of the shell analysis results adjacent to the ring girder for the five TES plants was completed. The margin of safety on total stress for the plants ranges from .27 to 1.31. The additional margin is more than adequate to support any unexpected increase in total stress which may occur in the four-inch offset strip. e It follows that the range of total stress on the mitre side of the ring girder must be less than the range reported to the opposite side which was analyzed, o The fatigue evaluation was completed with a stress intensification factor of four (the maximum SIF required by the Code). All elements analyzed exhibited usage factors less than 10 percent of the allow-able, remote from the torus attached piping penetrations. The conclusion of this study is that it is not possible to produce a stress intensity within the four-inch offset strip between the ring girder and mitre joint wnich will exceed those allowable values reported. ITEM 11 QUESTION Provide a list of the component materials and their corresponding metal temperatures used for the stress limit selection. ANSWER The torus structure and major components were evaluated at a temperature of 200 F. This temperature conservatively bounds the maximum temperature obtained from the Plant Unique Load Definition (Reference 10 of PUA) at 1720F.
s - a R: port yp qq Revision 2 A4.2-11 ENGBEBWJG SERVCES 1 i j The major component materials are as follows: s A 516 Gr 70 Torus l Shell Support Columns Ring Girder Saddle Support Earthquake Restraints Drywell vent System Vent Pipe Vent Header Downcomers A 333 Gr 1 Vent Header Support Columns A 333 Gr 6
. Vent Header Deflector ITEM 12 QUESTION Indicate whether each torus attached piping and its supports have been classified as Class 2 or Class 3 piping, Class 2 or Class 3 component sup-ports, and essential or non-essential piping systems. Also,. indicate whether a pump or valve associated with the piping mentioned above is an active or inactive component, and is considered operable.
Tcchnical Report TR-5319-1 TF WE Revision 2 A4.2-12 94GBEstNG SERVICES 4 ANSWER All Vermont Yankee Torus Attached Piping systems have been classified as essential Class 2 piping systems and all components associated with these sys-tems are considered active, for purposes of these analyses and evaluations. ITEM 13 QUESTION With reference to Table 1 of Appendix B, indicate whether all loads have been considered in the analycis and/or provide justification, if any load has been neglected. ANSWER All loads shown on Table 1 of Appendix B in the PUA report have been considered in the analysis, except those that were specifically identified and discussed in the report. Discussion of these exceptions follows: CONTAINMENT STRUCTURE ANALYSIS All loads were analyzed on the torus shell with the exception of the post chugging load. Analysis done on one of the TES plants produced very low stresses and loads that were bounded by pre-chug values. Additional work published (Ref. 12 PUA Report) showed that pre-chug bounded post chug (to 50 Hz) for column and saddle loads. It also showed that P1 + Pb stress due to post chug exceeded pre-chug by 53%. TES analysis for post chug used the pre-chug stress values which may be increased by 53% and still meet allowable stress. (Taken from Section 3.0 of the PUAR). The attached piping reaction loads on the torus shell will be con-sidered in the Torus Attached Piping (TAP) Technical Report (TR-5319-2). These loads are a function of the final piping configuration. The local stresses will be added to the existing state of stress for the appropriate region cf the torus shell.
s Technical Rep;rt gg on 2 A4.2-13 SERVICES VENT HEADER SYSTEM (The following are taken from Section 4.0 of the } / PUAR). The following vent system loads were not analyzed: e Pool Swell Drag LOCA Jet Forces The vent header support columns are loaded by forces from f LOCA-Jet and LOCA-Bubble drag. By inspection, it was con-cluded that LOCA-Jet loads would not combine with water impact on the vent system due to differences in timing and, there-fore, would not contribute to the maximum stress calctlations, e Submerged Structure Drag (Support Columns Only) Examination 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 produce much higher stresses. For this reason, stresses in the vent header support columns were not calculated for chugging drag. e 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 dif-ferences in timing, o Condensation Oscillation - IBA Stresses and loads resulting from IBA condensation oscillation are bounded in all cases by either DBA condensation oscilla-tion or chugging.
s 51 1 W Revision 2 A4.2-14 N SBWCES 0THER STRUCTURES (The following are taken from Section 7.0 of the PUAR). All direct loads were applied to the torus catwalk. Indirect effects, due to motion of the ring girder at attachment points were considered, but judged to be negligible. Except for the handrails, the entire catwalk is submerged before froth loads reach this part of the torus; because of this, froth was only considered on the handrails. The internal spray header is attached to the ring girders and to a penetration on the shell. The motion of the ring girder that results from pool swell loads on the shell was considered but judged to be a negligible input to the spray header. Shell dispircement at the nozzle connections was input to the computer analysis. The spray header is high enough in the torus so it does not experience direct water impact; froth is the only pool swell related load that was applied. ITEM 14 QUESTION Provide a summary of the analyses for the new modifications yet to be l supplied; these include items 5, 6, 10, 12 and 15 of the key for Figures 2.3 : and 2.4 of Reference 4 In addition, if the final configuration of the catwalk is to be changed, update the analysis accordingly. ANSWER Items 6,10,12 and 15 on Figures 2.3 and 2.4 of the PVA pertain to Torus Attached Piping analyses. These items will be summarized in the TAP Technical Report TR-5319-2 scheduled to be issued in the fall of 1983.
~
"'*'" TF WE Revision 2 A4.2-15 DJGtEstNG SSWICES Item 5, the Vent Header to Downcomer Stiffener stresses are bounded by those summarized in Section 4.4.1 of the PUA. The detailed vent header model as shown in Figure 4-1 includes the stiffeners.
Since the PUAR was issued, a decision was made to remove most of the ). catwalk at Vermont Yankee. The catwalk has been removed from fourteen bays and only remains in the two bays where the access hatches exist. The non-vent bay portion of the 1/16 STARDYNE model was removed and the vent bay portion has beer. re-analyzed. A list of the catwalk modifications follows:
- 1. 4 x 4 angle support legs changed to four-inch schedule 80 pipe, pinned at both ends.
- 2. Addition of four four-inch schedule 80 pipe diagonal braces, pinned at both ends.
- 3. Additional 7 x 1/2 plate welded to the existing 4 x 3 angle for lateral stiffness.
- 4. Additional 3/4 inch steel rod or equivalent, added to increase horizontal stiffness.
- 5. New cable handrails and posts.
- 6. Additional hold-down plates for grating.
- 7. Removal of the ladders during plant operation.
The report will be revised to reflect the new stress results. A summary of these results are as follows:
s Tcchnical Report pgg o 2 A4.2-16 BIGretNGSERVICES
-e Main Frame Pool Swell + SRV + Seismic + Weight (Case 25)
Bending + Axial Stress = 24,400 psi, 40,600 psi allowable e Support Columns, Support Diagonal Braces and End Joints Pool Swell + SRV + Seismic + Weight (Case 25) Bending Stress of Outboard Diagonal Brace = 18,445 psi, 42,000 psi allowable e Welds to Ring Girder Pool Swell + SRV + Seismic + Weight (Case 25) Tensile Stress = 18,264 psi, 42,000 psi allowable ITEM 15 QUESTION Provide details of fatigue analysis for piping systems. Indicate whether the fatigue usage factors for the SRV piping and the torus attached piping are sufficiently small that a plant-unique f atigue analysis is not warranted for piping. The NRC is expected to review the conclusions of a generic presentation (6) and determine whether it is suffi-cient for each plant-unique analysis to establish that the expected usage f actors for piping are small enough to obviate a plant-unique f atigue analysis of the piping. ANSWER TES has provided typical f atigue information to the Mark 1 Owners' Group generic study for all five of the plants for which we are analyzing torus
Technical Report 96 Rv n2 A4.2-17 BIGNEstNG SERVICES attached piping. Therefore, the conclusion of the generic presentation to the NRC, which established that the fatigue usage factors are small enough to obviate a plant-unique fatigue analysis, applies. We anticipate NRC agree-ment with the generic presentation, shortly. j ITEM 16 QUESTION Submit a summary of the analysis for the miscellaneous internal piping. ANSWER The following is a sumary of the maximum stresses associated with the miscellaneous torus internal piping: Maximum Allowable Load Item Stress Type (PSI) Conditions Main Junction 12850 Bending 21600 DL + SSE I + FRTHIA Box (No. 855) Thermocouple 18080 Bending 27000 DL + SSE I + FRTHlA Junction Box Dewcell Support 2117 Bending 21600 DL + SSE I RTD Support 10690 Bending 27000 DL + SSE I Thermocouple 17874 Bending 27600 DL + SSE I + IMP + Support DRG + MH 3/4" 0 Conduit 18641 Bending 27000 DL + SSE I + FRTHlA Supports on Ring Girders Support for Main 3891 Tension 16000 DL + SSE I + FRTHIA PowerCables(from penetration to main junction box) 1h" 0 Conduit 18606 Bending 27000 DL + SSE I Supports on Monorail
l a Report yp qq Revi ~on 2 A4.2-18 ENGNEstlNG SERVICES
- Definitions DL - Deadload SSE I - Safe Shutdown Inertia FRTHIA - Froth Load (Region 1A)
IMP - Impact Load i DRG - Drag (Submerged Structure) MH - Hydrodynamic Load (Associated with Impact) ITEM 17 QUESTION The ASME Code provides an acceptance procedure for computing f atigue usage when a member is subject to cyclic loadings of random occurrence, such as might be generated by excitations from more than one type of event (SSE and SRV discharge, for example). This procedure requires correction of the stress-range amplitudes considered and of the associated number of cycles in order to account for the interspersion of stress cycles of unlike character. State whether or not the reported usages reflect use of this method. If not, indicate the effect on reported results. ANSWER The f atigue analysis of the torus shell does correct the stress-range amplitudes and associated number of cycles to account for the interspersion of stress cycles of unlike character. The reported usage f actors do reflect the use of this method. It should be pointed out, however, that the usage factors reported do not contain the fatigue usage factors at the Torus Attached Piping Penetrations. The fatigue analysis for the TAP penetrations will be discussed in detail in TES Technical Report TR-5319-2 scheduled for issue in the fall of 1983.
s ~ a _fReport yp qq Revision 2 A4.2-19 SERVICES l ITEM 18 QUESTION Justify the reason for not considering skew symmetric boundary condi-I tions in the analysis of the torus shown in Figure 3.1. Evaluate the effect , of the thus neglected modes. ANSWER It has been our position that the geometry of the torus structure, the nature of the loads imposed, and the constraints imposed by the support saddles and ring girder will force the symmetric modes to dominate shell response to the extent that asymmetric modes can be omitted; the logic follows: The nature of the loads was considered first. Most Mark 1 loads are both vertical and uniform. For these loads, asymmetric modes clearly are not excited. The loads which do not satisfy this description are SRV, asymmetric chugging and horizontal earthquake. Of these loads, earthquake is a static load, so the question of mode - shapes does not apply. (Seismic analysis of the restraint system was done on a 360 model (ref. Figure 3.4, PUAR). Chugging consists of two components, pre-chug and post chug; the post chug component of chugging is a small load and is bounded by pre-chug for all stresses controlled by gross structural response (ref. para. 3.0, PUAR). Therefore, SRV and asymmetric pre-chug are the two loads which must be addressed. Although these loads are not uniform, they always produce pressures that are in-phase in adjacent bays. Such a loading will produce response controlled primarily by symmetric modes. This is especially true if we consider the f act that both these loads can exist anywhere within a frequency band, but must be assumed to reside at the single worst fre-quency in that range. Because of the in-phase characteristic of the
Technical Report TR-5319-1 TN Revision 2 A4.2-20 BIGrEBWJG SERVICES load, that wurst single frequency will be one associated with a symetric mode, not an asymmetric one. On this basis, asymetric modes were considered to be unnecessary. I \ It is also true that the use of symetric boundary conditions implies g that the load is uniform, and because of that, introduces some conserva-tism in results. We believe this conservatism more than compensates for the small error that may be associated with neglecting asymmetric mode shapes. ITEM 19 QUESTION Specific comments addressing the method of summation used and its com-pliance with the probability of non-exceedance (PNE) criteria of 84% stated in para. 6.3b of Reference 1 should be incorporated into the text. ANSWER As we understand the question, it relates to use of the cumulative distribution function in combining dynamic load effects. The cumulative distribution function method of combining any two structural responses has not been used for any analysis. All combinations of two separate dynamic loads were done by absolute sum. ITEM 20 QUESTION Provide justification for analyzing only one SRV discharge line, as shown in Section 6.0 of Reference 4. Indicate whether all discharge lines are identical in canfiguration to the one modeled, and whether the model investi-gated is conservative enough to represent all lines.
~ r a Report yg 5 Revision 2 A4.2-21 DiGDEstNG SERVICES f ANSWER Analysis of the SRV discharge line has been done and will be reported as f two separate analyses. Analysis of the quencher, quencher supports and piping in the torus is reported in TES Technical Report TR-5319-1. Analysis of the vent pipe penetration and all upstream piping and supports will be reported in TR-5319-2, scheduled for release later this year. This separation is possible because stresses in the piping and structure in the torus are controlled by water clearing and pool drag loads alone. Stresses in the penetration and the drywell are affected by all loads, includ-g ing gas clearing. The separation of analysis was made to provide early results for torus wetwell piping, which previously had been identified by the NRC as an area of concern. The portion of the SRVDL shown in Figure 6-1 of the PUAR is identical for all Vermont Yankee discharge lines. ITEM 21 QUESTION Submit a summary of the analysis for the vacuum breaker and its penetration. ANSWER The vacuum breaker piping and penetration analysis for the torus and vent pipe penetrations will be contained in the Torus Attached Piping Technical Report TR-5319-2 scheduled for release by Yankee Atomic Power Company in the fall of 1983.
"ha 3
Report yg ENGBEERNG SERVICES Revision 2 A4.2-22 f ITEM 22 I QUESTION Justify that the 45 model of the vent header and downcomer used in the analysis is adequate to meet the intent of the criteria which requires at least 180 . 1 Justify the reasons for not considering skew synmetric boundary condi-tions to evaluate the effect of the resulting modes. ANSWER A generic analysis was performed using a 180 0segment vent system beam model with symmetric boundary conditions for the appropriate asymmetric loading cases. The two loading cases considered are synchronized chugging and static seismic. The static seismic values of 0.179 horizontal and 0.19 vertical used envelop the original plant design seismic spectra for the five TES plants analyzed (Nine Mile Point, Millstone, Vermont Yankee, Fitzpatrick and Pilgrim). The combined seismic and chugging stresses of the 180 0segment model are less than the combined stresses of the 450 segment model because of the conservative assumptions used to apply the anti-symmetric chugging load on the 450 model. The ratios of the combined seismic and chugging stress of the 180 /45U models are: 970 psi /7851 psi = 0.13 for the downcomers 3630 psi /6020 psi = 0.6 for the vent headers
( Technical Report r TR-5319-1 YN Revision 2 A4.2-23 NSEMCES Theref ore, the combined stress analysis reported in the PUAR using the results from the 450model is conservative. s* , ITEM G1 QUESTION L Describe fully the procedures used to assess cumulative fatigue damage. In particular, address:
- 1. Where departures from standard code procedure were introduced.
- 2. How critical points were selected and how stress (or stress inten-sity) ranges were computed.
- 3. Which cyclic loads were omitted, if any, in these computations. For example, were thermal transients given consideration?
- 4. Whether cyclic amplitudes and the associated number of cycles were adjusted to account for the interspersion of cycles of unlike char-acter.
- 5. How the cumulative usage factor was computed.
- 6. What impact departures from code procedures have on the margins of safety shown for each component for which cumulative usage was computed.
ANSWER The following items highlight the major considerations used to assess the cumulative fatigue damage for the torus structure. A description of the actual procedure used is described in Section 3.2.7 Fatigue Analysis of the
Technical R:ptrt gg so 2 A4.2-24 S PUAR. The Fatigue Analy.is of the torus was completed using the procedures set forth in Section NE-3221.5 " Analysis for Cyclic Operation" of the ASME BPVC. e The cumulative f atigue usage f actors were conservatively calculated using the maximum stress intensification f actor recognized by the ASME BPVC of 4.0. e The maximum alternating stress intensity for a particular loading cvent is calculated independently of other loading events (Sa = Sr/2). The alternating stress intensities are then conservatively J combined by absolute summation. We are, therefore, assuming that each loading case will increase the magnitude of the stress range for the number of cycles over which it acts, e Critical points were chosen based on the stress analysis. Those elements in the region of the torus shell which exhibited the maxi. mum membrane, bending and total stress intensity as shown in Figure 3-9 of the PVA were analyzed for f atigue. e Section 3.2.3.2 on Post Chugging indicated that pre chug stress values for the torus bounded post chug. Therefore, the f atigue analysis was completed using the pre-chugging stresses. The dura-tion of loading for both chugging events is identical. e Thermal transients were not given consideration. Item G2 addresses this subject. e The torus attached piping penetrations will be addressed for fatigue in TES Technical Report TR 5319-2. e As discursed in Item 17, adjustments to the cyclic amplitudes and the associated number of cycles were made.
f Tcchnical Report 96 v o 2 A4.2-25 I ITEM G2 , QUESTION 4 f 15 the method described in Section 4.3.6 of Reference 4 for assessing thermal stress typical of all evaluations made in the report? Please discuss the tacit assumption that either: 1
- 1. Thermal equilibrium is achieved before other significant mechanical loads are experienced by the structure.
or
- 2. Maximum transient thermal stresses are conservatively bounded by the assumptions made.
ANSWER _ The resultant alternating stress intensity from one cycle of LOCA ther-mal transient event will not significantly affect the magnitude of the cumula-tive fatigue usage factor. The following discussion is provided to support our decision not to consider the thermal transient events and complete our fatigue analysis with steady state thermal results where required. The ASME BPVC NE 1221.5 Analysis for Cyclic Operation, Section (d) vessels not Requiring Analysis for Cyclic Service. Number (4) Temperature Olfference - Similar Material states: 4 A temperature difference fluctuation shall be considered to be signifi-cant if its total algebraic range exceeds the quantity 5/t
Technical R: port gg i i so 2 A4.2-26 S where S is the value of Sa obtained from the applicable design f atigue curvefor(10)6 cygg,,, for carbon steel, this quantity is approximately 700 f. The PULD temperature transients for the five plants which were analyzed by TES were reviewed with the following results: e All wetwell and drywell temperature transients for the 50A and IBA events were less than 70'F. e All wetwell temperature transients for the 00A events were less than 700 f. e The maximum 00A drywell transient of the five plants considered was 0 217 F. Therefore, the only transient of concern 15 the 00A drywell temperature. The major portion of the 00A transient occurs very early in the event (witnin the f trat 1.6 seconds) while pool swell 15 still in progress. Since the PUAAS does not require that the 00A pool swell events be considered for the fatigue analysis or primary plus secondary stress intensity range, the temperature transient may be excluded fro.n further consideration. ft.e of fects of the Transient Thermal Conditions assor,tated with the LOCA related events can, therefore, be escluded from further consideration.
l II!N!f;'l"U#* WN Revisten ? A4.3 1 N REVIEW BY FRANKl.!N INSTITUTE (C0f4TINUED). t A4.3 Presentation Handout at FRC/f4RC/TES/YAEC Mocting on August 9, 1983 t w
Il$d*lR@ ort SPTri m(NE 5 Revision 2 A4.3-2 ENGNEERNG SERVICES ) MARK 1 CONTAINMENT PROGRAM REVIEW 0F THE I PLANT UNIQUE ANALYSES REPORTS 0F. THE 1
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icchnical Report. TR-5319-1 SPTA WNE Revision 2 A4.3-4 ENGNEERNG SERVICES i i GENERAL ., o TORUS ATTACHED PIPING SYSTEMS (TAP) e MAINSTEAMSAFETYRELIEFVALVEDISCHARGEPIPING l e PIPING e PE5ETRATIONS e SUPPORTS i i e ACTIVE COMP 0NENTS e MODIFICATIONS I l
Technical Report SPTF1 prT(NE ) n2 A4.3-5 h SRV PIPING ANALYSIS e SRVDL'S PER PLANT (4-6) e DISCHARGE END (PUAR-1) TORUS SRV PIPING TEE-QUENCHERS QUENCHER SUPPORT BEAM e MAIN STEAM TO MAIN VENT PENETRATION PIPING SUPPORTS ACTIVE COMPONENTS MAIN VENT PENETRATION l l 1
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Technical Report SPTF1 W NE ev n2 A4.3-8 6 fPPLICABLE CODES AND CRITERIA I e ANALYSIS OF PIPING AND SUPPORTS SECTION III ASME BOILER AND PRESSURE VESSEL CODE,1977 EDITION, INCLUDING SUMMER 1977 ADDENDA e DESIGN AND FABRICATION OF SUPPORTS ASME BPVC SECTION XI e LOAD COMBINATIONS AND STRESS LEVELS i MARK 1 CONTAINMENT PROGRAM STRUCTURAL ACCEPTANCE CRITERIA PLANT UNIQUE ANALYSIS APPLICATION GUIDE (PUAAG)
Technical Report SPTF1 FD(E v n2 A4.3-9 SRV PIPING LOADS e GAS CLEARING SUDDEN PRESSURIZATION DUE TO RAPID SRV OPENING UNBALANCED DYNAMIC FORCES G.E. COMPUTER CODE RVFOR-04 A1.2 B0UNDING ANALYSIS OF EACH LINE o WATER CLEARING DISCHARGE WATER ACCELERATION UNDER SRV DISCHARGE LINE PRESSURE C3.3 B0UNDING ANALYSIS, MAXIMUM REFLOOD (G.E. RVRIZ) e SUBMERGED STRUCTURE DRAG (P00L MOTION) ON TEE-QUENCHER, SRV PIPING, SUPPORT BEAM, AND MAIN VENT PENETRATION P0OL SWELL - JET LOADS l
- BUBBLE LOADS CONDENSATION OSCILLATION - SOURCE INDUCED DRAG
- FLUID STRUCTURE INTERACTION (FSI) DRAG CHUGGING - SOURCE INDUCED DRAG
- FSI DRAG l SRV DISCHARGE - DRAG FROM ADJACENT QUENCHERS LOADS REPORTED IN PUAR-1
Technical Report SPTri Fry (E ev on 2 A4.3-10 SRV PIPING LOADS (CONTINUED) e WEIGHT, THERMAL AND SEISMIC ANALYSES ARE BASED ON EXISTING FSAR AND ESTABLISHED SYSTEMATIC EVALUATION PROGRAM DATA e THERMAL EXPANSION SRVDL CONTAlhMENT
Technical Report TR-5319-1 1%P W NE Revision 2 A4.3-11 ENGNEBBIG SERVCES l SRV PIPING ANALYSIS COMPUTER MODEL e STARDYNE COMPUTER CODE e REPRESENTATION OF THE MAIN STEAM LINE BY A FULL 6 X 6 STIFFNESS MATRIX DEVELOPED FROM A STATIC ANALYSIS OF THE MAIN STEAM LINE OR MODELING OF THE MAIN STEAM LINE WITH EACH SRVDL e REPRESENTATION OF THE STIFFNESS OF THE MAIN VENT PENETRATION BY A SET OF SIX ATTACHMENT SPRINGS, DEVELOPED BY COMPUTER ANALYSIS OF THE PENETRATION AREA e FULL REPRESENTATION OF THE TEE-QUENCHER AND QUENCHER SUPPORT BEAM IN THE PIPING MODEL e FULL REPRESENTATION OF THE BRACKETS BETWEEN THE QUENCHER AND SUPPORT BEAM WHICH ALLOWS FREE TORSIONAL ROTATION TO THE QUENCHER ARMS e DAMPING AT 2% OF CRITICAL FOR ALL TIME HISTORY ANALYSIS
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Technical Report 97 qq Rh n2 A4.3-14 ENGNEN SERVICES SRV PIPING ANALYSIS METHOD l ! e INDIVIDUAL TIME HISTORIES FOR WATER AND GAS CLEARING LOADS APPLIED AT EACH BEND AND ELB0W e DYNAMIC ANALYSIS e LINE UNIQUE LOADS e BOUNDING ANALYSIS WITH WORST CASE GAS AND WATER CLEARING LOADS e DAMPING 2% OF CRITICAL e CALCULATIONAL TIME INCREMENTS FOR SOLUTION MONITORING AT .0025 SEC e RESPONSE FREQUENCIES TO 50 HZ e INTERNAL PRESSURE BY HAND e THERMAL, WEIGHT AND SEISMIC ANALYSIS USING THE SAME MODEL l
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S 91 $ OgP WNE Revision 2 A4.3-16 ENGNEERNG SERVCES SRV PIPE SUPPORT ANALYSIS e ANALYSIS BY HAND AND COMPUTER l e STAAD COMPUTER PROGRAM USED FOR ANALYSIS OF COMPLEX SUPPORTS e ANALYSIS EXTENDED TO INCLUDE THE ATTACHMENT WELD TO SUPPORTING STEEL e SUPPORTING STEEL WAS INCLUDED IF LOAD CAPACITY WAS SUSPECT e MAIN VENT PIPE SUPPORTS: DETAILED EVALUATION OF STRESSES IN THE MAIN VENT WALL BIJLAARD ANALYSIS IN COMBINATION WITH INTENSIFIED FREE SHELL STRESSES DUE TO VENT HEADER LOADING SRV MAIN VENT PENETRATION ANALYSIS e BIJLAARD ANALYSIS TO DETERMINE LOCAL PENETRATION STRESSES DUE TO SRV DISCHARGE LINE LOADS e LOCAL STRESSES ADDED TO INTENSIFIED FREE SHELL STRESSES e VERTICAL ENTRY LINES HAVE A COMPLEX PENETRATION REINFORCEMENT REQUIRING A DETAILED FINITE ELEMENT ANALYSIS (ANSYS COMPUTER CODE)
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Technical Report TR-5319-1 TTF1 FTWNE Revision 2 A4.3-20 ENGNEERNG SERVCES / L SRVDL EVALUATION e 27 LOAD COMBINATIONS (PVAAG) e 13 COMBINATIONS INCLUDE SRV e CONSERVATIVE (BOUNDING) LOAD CASE EVALUATED' AGAINST LEVEL B ALLOWABLES DW + (SSE)2 + (BLOWDOWN) = 1.2 Sh e LESS CONSERVATIVE COMBINATIONS EVALUATED WHERE REQUIRED DW + (SSE)2 + (BLOWDOWN)2 = 1.8 Sh DW + OBE = 1.2 Sh DW + BLOWDOWN + 1.2 Sh (THESE CASES REPRESENT LOAD COMBINATIONS 15, 1 AND 2) l e FATIGUE EVALUATION OF SRV LINES WAS UNDERTAKEN AS A GENERIC MARK 1 PROGRAM EFFORT USING B0UNDING ASSUMPTIONS THE EFFORT CONCLUDES THAT FATIGUE WILL NOT BE A PROBLEM SRV DISCHARGE PIPING DATA PROVIDED TO GENERIC EFFORT
s Technical Report TR-5319-1 1%P W NE Revision 2 A4o3-21 ENGNEERNG SERVICES l CLASS 2 AND 3 PIPING
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Technical Report
' 1PTF1 m(NE eh n2 A4.3-22 ENGNEERNG SERVICES SRV SUPPORTS AND STRUCTURAL STEEL e PIPE SUPPORTS AND STRUCTURAL STEEL WERE EVALUATED USING A WORST CASE LOAD CONDITION CONSERVATIVE BLOWDOWN CASE SSE SEISMIC (SRSS'D WITH BLOWDOWN)
WORST CASE THERMAL LOAD DEADWEIGHT SRV MAIN VENT PENETRATION e LOADS FROM THE TORUS AND DRYWELL PORTIONS OF THE PIPING WERE COMBINED TO PERFORM THE ANALYSIS IN ACCORDANCE WITH ASME BPVC SUBSECTION NE e ANALYSIS INDICATES THAT THE MAXIMUM LOAD CAN BE CYCLED FOR A MINIMUM 0F 7500 CYCLES MAJOR LOAD 50 CYCLES MAXIMUM l NORMAL SRV ACTUATIONS PRODUCE LESS THAN 4500 CYCLES AT SIGNIFICANTLY LESS LOAD SAFETY RELIEF VALVES AND VACUUM BREAKERS e STRESSES IN ADJACENT PIPING MEET LEVEL B CRITERIA TO INSURE PROPER OPERATION OF VALVES
Technical Report TR-5319-1 7PTF1 Frh'NE Revision 2 A4.3-23 ENGNEERING SERVCES
SUMMARY
OF SRV LINE MODIFICATIONS l e INSTALLATION OF TEE-QUENCHER DISCHARGE DEVICES AND QUENCHER SUPPORTS I e INSTALLATION OF TWO TEN-INCH VACUUM BREAKERS ON EACH SRVDL e MODIFICATION OF MAIN VENT PIPE PENETRATIONS AS REQUIRED e MODIFICATION TO SRVDL SUPPORTS AS REQUIRED e MODIFICATION TO SUPPORT STEEL IN THE DRYWELL AS REQUIRED l
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' Technical Report TR-5319-1 TE NE Revision 2 A4.3-25 ENGNEstlNG SERVICES l i l TORUS ATTACHED PIPING (TAP) i e APPR0XIMATELY 15 LARGE BORE PIPING SYSTEMS (* 4 IN. DIA.) PER PLANT e ANALYSIS OF LARGE BORE PIPING SYSTEMS IS FROM THE TORUS TO THE FIRST ANCHOR e SMALL BORE PIPING $E 4 IN. DIA.) WAS ANALYZED TO THE FIRST ANCHOR OR A DISTANCE WHERE THE TORUS LOADS C0'LD J BE CONSIDERED NEGLIGIBLE e BRANCH PIPING CONNECTED TO TAP SYSTEMS e TORUS PENETRATION STRESSES e PIPING INSIDE THE TORUS ATTACHED TO TAP SYSTEMS e PUMP AND VALVE LOADS e ALL PIPE SUPPORT AND ANCHOR LOADS
Technical Report TR-5319-1 TF WE Revision 2 A4.3-26 ENGNEERNG SERVCES i k APPLICABLE CODES AND CRITERIA e TAP PIPING ANALYSIS ASME BPVC SECTION III, 1977 e BRANCH PIPING ANALYSIS ASME BPVC SECTION III, 1977 ORIGINAL DESIGN CODE FROM FSAR e SUPPORT ANALYSIS ALL TAP AND BRANCH SUPPORTS WERE ANALYZED IN ACCORDANCE WITH THE AISC CODE AND INCLUDING NRC BULLETIN 79-02 REQUIREMENTS ALL STRUCTURAL SUPPORT STEEL -- AISC e LOAD COMBINATIONS AND STRESS LEVELS MARK 1 CONTAINMENT STRUCTURAL ACCEPTANCE CRITERIA PLANT UNIQUE ANALYSIS APPLICATION GUIDE
1 l m;;!;:1 a*" c amunE Revision 2 A4.3-27 ENG4EstNG SERVICES TAP LOADS e THE ORIGINAL DESIGN LOADS WHICH INCLUDE WEIGHT, THERMAL AND SEISMIC BASED ON EXISTING FSAR AND ESTABLISHED SYSTEMATIC EVALUATION PROGRAM DATA e SHELL MOTION LOADS WERE OBTAINED FROM PLANT UNIQUE SHELL RESPONSE DATA DEVELOPED AND REPORTED IN PUAR-1 e SHELL ANALYSIS PROVIDES A TIME HISTORY RESPONSE IN 5 000F e POOL SWELL e SRV DISCHARGE e CONDENSATION OSCILLATION e POOL DRAG AND IMPACT LOADS ON TORUS INTERNAL PIPING WERE CALCULATED IN ACCORDANCE WITH THE LOAD DEFINITION REPORT e SUBMERGED PIPING LOADS: C0 SOURCE AND FSI ORAG POST CHUG SOURCE AND FSI DRAG PRE-CHUG DRAG SRV BUBBLE AND JET LOADS POOL SWELL BUBBLE DRAG AND FALLBACK l l
s R 53 1 TME Revision 2 A4.3-28 N SBh/ ICES TAP LOADS (CONTINUED) 4 e STRUCTURES AB0VE THE P00L: POOL SWELL WATER IMPACT AND DRAG FROTH FALLBACK e SUBMERGED STRUCTURE LOAD SPECTRUM WAS DYNAMICALLY ANALYZED FOR PIPING SYSTEMS, THE SPECTRUM INCLUDES CO AND CH SOURCE AND FS! DRAG e REMAINING SUBMERGED STRUCTURE LOADS WERE APPLIED A:t SEPARATE STATIC CASES WITH WORST CASE ORIENTATIONS
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If9 TF WE Revision 2 A4.3-37 ENGPEBUNG SERVICES REPRESENTATION OF TORUS SHELL FOR PIPING ANALYSIS l e DYNAMIC INTERACTION BETWEEN THE PIPING AND THE TORUS e GROUND SPRINGS IN THE PIPING MODEL REPRESENT THE TORUS CONNECTION e STIFFNESS VALUES CALCULATED BY APPLYING UNIT LOADS AND MOMENTS ND TO THE 1/32 FEM. THE IN-PLANE (TORSION) MOTION OF THE SHELL IS CONSIDERED NEGLIGIBLE LARGE BORE TAP (>4 IN. DI A.) e FINITE ELEMENT MODEL OF EACH TAP SYSTEM INCLUDING PIPING INSIDE THE TORUS, AND GROUND SPRINGS TO REPRESENT THE TORUS SHELL e STARDYNE COMPUTER CODE DYNAMIC TIME HISTORY ANALYSIS e DAMPING VALUES OF 2% IS CRITICAL e PRE-CHUG LOAD WAS NOT RUN FOR EACH TAP SYSTEM AS IT IS ALWAYS BOUNDED BY DBA C0 e DRAG LOADS DUE TO INTERNAL PIPING WERE CALCULATED BY HARMONIC ANALYSIS FOR THE SPECTRAL LOADS ( e PIPE STRESSES DUE TO WELDED ATTCHMENTS WERE ANALYZED
TTELEDGE ENQNEERNGSEMCES lecimii.al Report TR-5319-1 Revision 2 A4.3-38 l l TORUS EXTER!iAL PIPE FORCE-TIME HISTORIES 4 p PEflETRATION APPLIED AT THIS POINT
- ~
b TORUS INTERflAL PIPE TORSI 0ft PIPE RIGID 5 DEGREES OF FREEDOM Ifl OTHER DIRECT 10fis l TAP pef!ETRATI0ff REPRESENTATION (TYPICAL) l
l I g$ ihT WE Revision 2 A4.3-39 ENGNEstlNG SERVCES SMALL BORE TAP (14 IN. DIA) e COMPLEX SYSTEMS TREATED IDENTICALLY TO LARGE BORE SYSTEMS e MARK 1 DYNAMIC LOADS LIMITED TO DBA C0 e SIMPLE SMALL BORE SYSTEMS COULD BE REDUCED TO SINGLE MASS APPR0XI-MATIONS AND ANALYZED BY HAND BRANCH PIPING e GENERALLY MODELLED WITH TAP SYSTEMS (APPR0XIMATELY 1/6 DIAMETER RATIO) e BRANCH PIPING CONSIDERED FLEXIBLE IN COMPARIS0N TO THE MAIN RUN PIPE WAS GENERALLY DECOUPLED AND EVALUATED BY STATIC COMPUTER ANALYSIS UTILIZING TOTAL THERMAL AND DYNAMIC DISPLACEMENTS AT THE BRANCH CONNECTION e SOME SMALL BORE BRANCH PIPING WAS REVIEWED AND JUDGED ACCEPTABLE WITHOUT RIG 0ROUS ANALYSIS
l Technical Report 9%P WNE TR-5319-1 , Revision 2 A4.3-40 ENGrEBilNG SERVICES i TAP PENETRATION ANALYSIS e LOADS FROM PIPING RESPONSE DUE TO SHELL MOTION e LOADS DUE TO SUBMERGED DRAG AND/0R P00L IMPACT ON INTERNAL TAP e WEIGHT, THERMAL AND SEISMIC PIPING LOADS e MARK 1 EVENT COMBINATION TORUS SHELL LOADING e BIJLAARD ANALYSIS TO ACCOUNT FOR LOCAL PENETRATION STRESS DUE TO PIPING LOADS e FREE SHELL STRESSES WERE INTENSIFIED TO ACCOUNT FOR THE DISCONTINUITY
lethnical Report TR-5319-1 1e 71 W NE Revision 2 A4.3-41 ENGNEBtNG SERVICES ANALYSIS METHOD FOR PIPING SUPPORTS e ALL PIPING SUPPORTS FOR THE TAP AND BRANCH PIPING ANALYZED e STAAD AND STRUDL COMPUTER PROGRAMS USED IN COMBINATION WITH HAND ANALYSIS e BASEPLATES AND ANCHOR BOLTS ANALYZED USING NRC BULLETIN 79-02 e LOCAL TORUS SHELL STRESSES WERE EVALUATED FOR LOAD FROM TAP SUPPORTS CONNECTED TO THE TORUS SHELL VACUUM BREAKER ANALYSIS e WETWELL-TO-DRYWELL VACUUM BREAKERS NOT A PART OF THE MARK 1 PROGRAM e SEPARATE TRANSMITTAL TO NRC e MILLSTONE VALVE DISC IMPACT VELOCITIES FROM MARK 1 LOADS ARE LESS THAN NORMAL CLOSURE DUE TO WEIGHT ALONE o VERMONT YANKEE VALVES 00 NOT ACTUATE DUE TO MARK 1 LOADS e PILGRIM VALVE ANALYSIS IS INCOMPLETE PENDING NRC REVIEW 0F CONTINUUM DYNAMICS DATA
I Technical Rc; port
' 97 qq ev s on 2 A4.3-42 N NES EVALUATION e FOR PURPOSES OF THIS EVALUATION, ALL TAP SYSTEMS ARE CLASSIFIED AS ESSENTIAL e CONSERVATIVE BOUNDING 0F 27 (PUAAG) EVENT COMBINATIONS:
CASE NO. MAJORLOAD(S) ALLOWABLE 3 SRV + SSE LEVEL B (1.2 Sh) 16 ZERO DELTA P LEVEL B AS MODIFIED BY FOOTNOTE 21 DBA C0/CH + SSE SAME AS 16 25 P0OL SWELL + SRV SAME AS 16 IS SSE + SRV + CH SAME AS 16 e SEISMIC STRESSES WERE COMBINED WITH THE MARK 1 LOADS BY SRSS METHOD e EVALUATION OF VALVES WAS BASED ON STRESSES IN THE ADJACENT PIPING. PIPE STRESSES MEETING LEVEL B CRITERIA WERE CONSIDERED ADEQUATE l TO ASSURE PROPER OPERATION OF THE VALVE e PIPING SUPPORTS WERE EVALUATED FOR THE SAME LOAD COMBINATIONS AS THE PIPING
TR 19 1 YME Revision 2 A4.3-43 N MES l 1 l I FATIGUE EVALUATION e MARK 1 GENERIC STUDY CONCLUDES THAT THE STRESS LEVELS AND CYCLES INVOLVED IN THESE SYSTEMS WILL NOT PRODUCE A FATIGUE PROBLEM e ALL TES PLANTS CONTRIBUTED TO THE MARK 1 FATIGUE DATA BASE e THE FATIGUE EVALUATION OF THE PENETRATION SHOWED THAT THE MAXIMUM LOAD COULD BE CYCLED ON EACH PENETRATION FOR AT LEAST 10,000 CYCLES e MAJOR LOADS PRODUCE LESS THAN 1,000 CYCLES I
Technical Report gg ev on 2 A4.4-1 SEMCES l i i I REVIEW BY FRANKLIN INSTITUTE (CONTINUED) A4.4 Response to FRC Request for Additional Data August 18, 1983 i l l l
- l
Technical Report y'TELEDYNE TR-5319-1 ENGNEERING SERVICES Revision 2 A4.4-2 130 $l( ONO Avf Ntil Wat tway MA%WHilst195 0.7 54 e61 h en0 6150 tWs 1710, h4 f*J 86 August 18, 1983 5960-6 Mr. Nat Subramonian Franklin Research Center 20th and Race Streets Philadelphia, PA 19103
Subject:
SRVDL/ Vent Pipe Penetration Fatigue Usage Factors
Dear Nat:
I have enclosed the subject data for your review per our August 9, 1983 meeting at Yankee Atomic Electric Company. Please note that the fatigue usage factors presented are bounding cases which were conservatively calculated assuming 10,000 cycles of loading at the maximum alternating stress intensity amplitude. If you have any questions, please call Mr. Nicholas Celia, Manager, Engineering Projects or me. Very truly yours, TE ~ ENG NEER " ::""'rES _ m --- _ ond M. . .E. MInager, Projects RMP: alt enclosure cc: N. S. Celia R. H. Berks R. A. Enos R. White (YEAC) M.Franceschina(NUSCO) G. Mlleris M. Lenhart BEC0)1 BECO
R 53 9 1 Y Revision 2 A4.4-3 MN SRVOL/ VENT PIPE PENETRATION FATIGUE USAGE FACTOR
SUMMARY
e Maximum Cycles Alternating Stress Assumed Allowable Usage Plant (KSI) n N Factor Vermont Yankee 29.5 10,000 20,500 .49 Pilgrim 21.1 10,000 65,000 .15 Millstone - Inclined 37.6 10,000 10,000 1.0 Vertical 36.5 10,000 10,400 .96 l l
Technical Report WMg TR-5319-1 Revision 2 A4.5-1 6 gg l l t l \ l l l REVIEW BY BROOKHAVEN NATIONAL LABORATORIES (BNL) A4.5 Review Questions Received on June 13, 1983
Technical Report TR-5319-1 A4.5-2 W TF1FT?( E Revision 2 DJGREtNG SERVICES ITEM 1 PUAR Section 2.2.1, AC Sections 2.13.8.2 and 2.13.8.3 The temperature monitoring system described in the PUAR using a total of ten thermocouples placed at five different torus locations l differs from the Acceptance Criteria in several important respects. For local temperature the criteria state that "For practical pur-poses, the average water temperature observed in the sector con-taining the discharge device at shell locations on the reactor side of the torus downstream of the quencher centerline at the same elevation as the quencher device and at the quencher support may be considered as the " local temperature". In Vermont Yankee, the thermocouples are on the " outboard" side of the torus (side away from the reactor) and for two of the four SRV discharge bays are upstream of the quencher centerline. There are no thermocouples at or near the quencher supports. Therefore, measuring local tem-perature in the sense of the Acceptance Criteria cannot be accom-plished by this system. For bulk temperature the criteria state that "Each licensee shall demonstrate that there is a sufficient number and distribution of pool temperature sensors to provide a reasonable means of bulk temperature". The brief description and illustration in the PUAR don't demonstrate that a reasonable meas-ure of the bulk pool temperature can be obtained from the Vermont Yankee system. The PUAR does not make clear whether the intention at Vermont Yankee is to measure bulk pool temperature or to measure local temperature directly, or both. Explain how the local tem-perature limit is to be determined: from bulk pool measurement or direct local measurement, and justify the adequacy of the corres-ponding measurement in light of the above comments regarding dif-ferences from the Acceptance Criteria. ITEM 2: PUAR Section 3.2.1, AC Section 2.4 Regarding the pool swell loads on the torus shell, describe how the longitudinal and azimuthal multipliers (LOR Table 4.3.2-1) were
I Y 1 1 Revision 2 A4.5-3 SSWICES used in conjunction with the submerged pressure histories to per-form the torus shell evaluations. Provide an example of a time history at a particular location (e.g., 0 = 180 at Z/1 = 0.0) to f illustrate their use. l ITEM 3: PUAR Section 3.2.3, AC Section 2.12.1 Regarding the pre-chugging and IBA/C0 load analysis, the PUAR states that results for the symmetric pre-chug load were developed directly from the unit-load harmonic analysis done for C0. Does this mean that the water mass was accounted for as in C0 (i.e.,100% water mass), and was this loading applied for the cycle duration stated in the LDR? 1 TEM 4: PUAR Section 3.2.4, AC Section 2.13 The PUAR states that the modeling of the water mass in the SRV load computer model was fraught with difficulty. When the water mass was included in the model, measured outputs could not be reproduced by applying measured input to the computer model. A dry structure analysis produced acceptable results, however, and therefore, the dry structure analysis method was subsequently used as a basis for all SRV analysis. This is a very troublesome point. Since there is no physical reason cited in the PUAR for using a dry containment in the SRV analysis, one is lef t with the impresssion that there is an error somewhere in the modeling which is fortuitously compensated for by introducing a second modelling error, i.e, non-inclusion of the water in the torus. A further difficulty is the implication these modeling results have for other loads such as C0 and chugging for which a fluid-structure computer model is also used and wiere the water was included in the analysis. Since no verifying measure-ments for these loads could be made, the possibility exists that these calculations are badly off the mark. Justify the exclusion of the torus water from the SRV analysis on physical grounds and explain why these physical reasons differ for the C0 and chugging loads.
9-1 TM l Revision 2 A4.5-4 DiGrERNG SERVICES l ITEM 5: PUAR Appendix 1, AC Section 2.13.9 The Acceptance Criteria call for the torus shell to be instrumented with strain gages, accelerometers, and pressure transducers during SRV in-plant tests. Since no accelerometers were used in the Ver-mont Yankee torus, explain how data from the other instrumentation was u.;ed to compansate for the lack of accelerometers. ITEM 6: PUAR Appendix 1, AC Section 2.13.9 Appendix 1 of the PUAR mentions that calibration factors relating predicted to actual pressures and predicted to actual frequencies were obtained by comparing QBUBS02 calculated values with the same quantities measured in the four in-plant tests. This appendix further states that verification of the computer model led to a further calibration factor for the column loads. Provide more details on how the calibration factors relating QBUBS and the in-plant tests were obtained, especially how the Vermont Yankee method conforms to the model calibration guidelines of the Acceptance Cri-teria (AC Section 2.13.9.2). Provide information on the actual forcing function used including amplitudes, frequency content and pressure wave forms. Were separate calibration f actors obtained for subsequent actuations? Also, provide more detail on the calibra-tion factor for column loads and explain why it is invariant over the frequency range of the loading. ITEM 7: PUAR Section 4.2, AC Section 2.10 The static load magnitude imposed on the vent header deflector in the analysis described in the PUAR seems appropriate if Figure 4.3.9-1 in the PULD accurately shows the initial impact pressure spike. Does this figure show the correct impact magnitude or should it be modified as per paragraph 1 of Section 2.10.1 of the Accep-tance Criteria? 1
a Report yp gg Revision 2 A4.5-5 ENGeEBWG SERVICES ITEM 8: PUAR Section 4.3.2, AC Section 2.12.2 ! 1 Regarding the downcomer lateral chugging loads: What is the funda-mental tied downcomer frequency? What was the corresponding I dynamic load f actor? What was the resultant static equivalent load l used in the stress analysis of the downcomer? ITEM 9: PUAR Section 4.3.2, AC Section 2.12.1 For synchronized multiple downcomer lateral chugging loads, the Acceptance Criteria Specification is based on an exceedance proba-bility for 10-4 per LOCA. The PUAR shows that two load cases were considered for multiple lateral loads. Why were only two load cases necessary and what static loads were applied? To what exceedance probability did these load magnitudes correspond? TTEM 10: PVAR Section 4.3.1.1, AC Section 2.6.2 In the calculation of stresses in the downcomers resulting from pool swell water impact, was the virtual mass of water near the downcomer accounted for, and was the eight psid pressure called for in the Acceptance Criteria applied over the bottom 50 of the angled portion of the downcomer? ITEM 11: PVAR Section 4.3, AC Section 2.14 Were the LOCA bubble drag loads calculated according to Acceptance Criteria specifications as given in Section 2.14.2 of the AC? ITEM 12: PUAR Section 5.2 Many of the ring girder loads for Vermont Yankee were analyzed using a computer model constructed for another Mark l plant of "similar" dimensions. What is the other plant? What are the dimensions of the
Report yp qq Revision 2 A4.5-6 BJGREstNG SBtVICES ring girder and surrounding shell structure of the other plant? Are attachments to the ring girder similar? Were the loads used on this attachments to the ring girder similiar? Were the loads used on this model the Vermont Yankee loads or the loads from the other plant? ITEM 13: PUAR Section 5.3.2, AC Section 2.14 Calculation of ring girder drag loads were not in accordance with the Acceptance Criteria. Therefore, provide the details of a sub-merged structure load calculation for a given segment of the ring girder. Include numerical values of a VT/D calculation, as well as source strength, as a function of frequency. In addition, provide the acceleration volume, drag coefficient, interference effect mul-tiplier and pertinent geometric parameters and configurations used in the calculation. ATEM 14: PUAR Section 5.3.2, AC Section 2.14.5 The PUAR states that FSI effects are accounted for in the submerged structure loadings. Additional detail is needed on how this was done. Is the criteria for including FSI effects the same as that stated in the AC? How were the FSI loadings obtained? Is the boundary acceleration added to the local fluid acceleration as sug-gested in the AC or has another method been used? 1 TEM 15: PUAR Section 7.1.3 The PUAR states that the catwalk structure stresses were computed without the catwalk grating. Does this mean that the grating is normally absent and will only be put in place when the catwalk is used? If the grating is always in place, by what amount will it raise catwalk stresses?
=-
Revision 2 A4.5-7
.mm-N SENICES ITEM 18: PUAR Section 4.3.1, AC Section 2.6 Provide pool swell impact and drag transient histories used in the calculation of pool swell loads on the main vent, vent header and downcomers. Provide enough detail to show how the load histories applied at the nodal points of the shell and beam models comply with the Acceptance Criteria.
ITEM 19: PUAR Sections Al, AC Sections 2.14.3 and 2.14.4 Use of SRV test data for submerged structure drag loads represents an exception to the Acceptance Criteria. The method described in Appendix 1 of the PUAR needs to be reviewed further, however, sev-eral problems which arise immediately are listed here:
- 1. The frequency content of different SRV load cases have been shown by experience to be different - multiple valve actua-tions show a lower frequency content than single valve tests.
The PUAR method does not address this problem. Arguments in the PUAR that " structures involved are responding to a fairly uniform random field" are unconvincing.
- 2. Using a uniformly distributed pressure as a way to obtain static loads giving strains equivalent to those measured can lead to nonconservatisms when Figure Al-5 is used to predict static drag pressures on structures whose geometry is dif-ferent from those on which the strains were measured.
- 3. Scaling the static drag pressure upward from test conditions to more severe SRV cases by the ratio of calculated shell pressures is an oversimplification which uses a global parame-ter to scale local effects. The local pressure on a submerged object due to simultaneous multiple SRV actuation can ratio very differently from the torus shell pressures, depending on the phasing and location of the quencher relative to the object.
T 1 t I 5319 1 W Revision 2 A4.5-8 DiGDEstNG SERVICES 1 ITEM 20: PUAR Section 6.0, AC Sections 2.14.3 and 2.14.4 \ The PUAR analysis of the tee-quencher, its support and the sub-merged portion of the SRV line does not mention quencher water jet or bubble drag loads on these structures. Where have these loads been included or why have they been ignored? ITEM 21: Provide the loads that were used in the torus attached piping.
l i Technical Report TR-5319-1 YM l Revision 2 A4.6-1 N SEMCES ! l REVIEW BY BROOKHAVEN NATIONAL LABORATORIES (CONTINUED) A4.6 Presentation Handout for BNL/NRC/TES/YAEC Meeting on July 26, 1983 l
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l Revision 2 A4.6-2 M 1 i , r, .I l I VERMONT YANKEE i I MARK 1 TORUS PROGRAM l
. RESPONSES TO PUAR QUESTIONS l
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i c Report WTri mm Revision 2 A4.6-3 DiGDEstNG SERVICES ITEM 1: PUAR Section 2.2.1. AC Section 2.13.8.2 and 2.13.8.3 QUESTION The temperature monitoring system described in the PUAR using a total of 10 thermocouples placed at five different torus locations differs from the Acceptance Criteria in several important respects. For local temperature the criteria state that "For practical purposes, the average water temperature observed in the sector containing the discharge device at shell locations on the reactor side of the torus downstream of the quencher centerline at the same elevation as the quencher device and at the quencher support may be considered as the local temperature". In Vermont Yankee the thermocouples are onthe" outboard"sideofthetorus(sideawayfromthereactor)andfortwo of the f our SRV discharge boys are upstream of the quencher centerline. There are no thermocouples at or near the quencher supports. Therefore, measuring local temperature in the sense of the Acceptance Criteria cannot be accomp. lished by this system. For bulk temperature the criteria states that "Each licensee shall demonstrate that there is a sufficient number and distribution of pool tempcrature sensors to provide a reasonable means of bulk tempera-ture". The brief description and illustration in the PUAR do not demonstrate that a reasonable measure of the bulk pool temperature can be obtained from the Vermont Yankee system. The PVAR does not make clear whether the intention at Vermont Yankee is to measure bulk pool temperature or to measure local temperature directly, or both. Explain how the local temperature limit is to be determined. from bulk pool measurement or direct local measurement, and justify the adequacy of the corresponding measurement in light of the above corrnents regarding differences from the Acceptance Criteria.
. Technical Report s.
TR-5319-1 , ., t Revision 2 'A4.6-4 Item 1 W Verndn't Yankee Suppression Pool Temperature Monitoring System s Originally installed in 1976, upgraded in 1983 Provides Bulk Pool temperature Information Only (Tech Spec in terms of Bulk pool Temperature) 910 T/C's in 5 locations (similar to Monticello)
-4' T/C's ~in S/RV Bays + 1 other
-1 T/C per Bay provides. bay temperature
-1 T/C per Bay averaged to provide bulk pool temperature
-Shallow submergence (2 feet) of T/C's will provide a conservative measurement of bulk pool. temperature
-Operator actions prompted earlier i .
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Technical Report 9mmm TR-5319-1 Revision 2 A4.6-6
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R 1 Y Revision 2 A4.6-7 DGPEstNG SERVICES ITEM 2: PUAR Section 3.2.1, AC Section 2.4 QUESTION Regarding the pool swell loads on the torus shell, describe how the longitudinal and azimuthal multipliers (LDR Table 4.3.2-1) were used in con-junction with the submerged pressure histories to perform the torus shell evaluations. Provide an example of a time history at a particular location (e.g., 0 = 180 degrees at Z/j = 0.0) to illustrate their use. l l
1 Report 9p q I 1 MN Revision 2 A4.6-8 l e LOR Table 4.3.2-1 longitudinal and azimuthal multipliers, e Linear interpolation for time and location. nd e Resultant multiplier applied to each node in 1/32 torus model. l e Pressure corrected as per Section 2.3 of Appendix A (NUREG 0661). l l e Pressure applied as a nodal force. 1 e Detailed calculation is included for reference. l l l l l
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Technical Report TM NM TR-5319-1 Revision 2 A4.6-10
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Technical Report TR-5319-1 WN NN Revision 2 A4.6-11 VT. YANKEE, 5 R I hG - 1/32 tiOr.EL, 1.7 DELTA P. WET N00Et ONLY T AdLE A-1 M RING IN T ERPOL AT ION 2 R ING = 1 2 3 4 5 Z /L : 1.JCJ .875 .750 .625 .517 . EVENT - TIME ( M tEC) ST ART C 1.30. 1 0C0 1.000 1 00C 1 000 PEAK DOWNLOAD 230 1 139 1.129 1.~94 1.060 '. 019 ZERO DOWNLOAD 340 .999 .998 .995 .992 .987 , PEAK UPLOAD 420 .94 4 .948 .967 .985 .995 REDUCED UPLOAD 568 .966 .964 .982 .997 1 003 ZERO UPLOAD 1110 1.30' 1.0CC 1. C0 1.000 1 00C I l 1
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Technical Report l TR-5319-1 YN NN ' Revision 2 A4.6-12
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VT. V ANKEE. 5 R l NG - t/32 MODEL. 1.7 DCLTA P.. WET NODES ONLY TABLE A-2 M TIME INTERPOLATION Z RING : 1 2 3 4 5 Z/L : 1.000 .875 .750 .625 .517 tit <C ~ (MSEC) 0 1.:3; 1.300-1.000 1.0:0 1.000 ; 30 1.320 1.019 1.114 1. 0 ". 9 1.003 90 1.163 1.053 1.042 1.027 1.009 1"C 1.067 1.054 1.347 1.030 1.009 1*5 1.277 1.074 1.054 1.C35 1.011 . 122 1. 0 'd 2 1.0/9 1.358 1.037 1.212 13C 1.087 1.^3* 1 261 1.C39 1.012 143 1.214 1.37C 1.~66 1.042 1.213 165 1.111 1 1?6 1.073 1.05C 1.~16 19C 1 121 1.115 1.085 1.C54 1.317 20C 1.131 1.129 1 094 1.06C 1 019 22C 1 115 1 1.0 1.0d0 1. 0 '2 C 1.014 240 1. ': M 1 111 1.066 1.091 1.:10 233 1.357 1.354 1.038 1.021 1.000 290 1.147 1.045 1 031 1.016 .998 3;C 1.033 1.C3G 1 024 1.C 11 996
1 - l Technical Report WTELEDGEENWEERNGSEWICES TR-5319-1 Revision 2 A4.6-13 VT. YANKEE, 5 RIhG - 1/32 MODEL, 1 7 DZLTA P. WET NODES 0NLY TAdLE B-1 M . THETA INTERP0i.ATION THETA 1HETA : 180 170 ' 160 150 140 EVENT T IME ( M SE C ) START : 1.003 1.;CO 1.0J0 1.200 1.003 PEA 4 00MNLOAD 223 1.205 1.200 1 159 1.083 .954 ZERO DOWNLOAD 340 .947 .950 .948 . 9 '* 0 .990 PEA 4 UPLOAD 420 . 9. 8 .926 .93T 9 t+ 0 .9'35 HECUCEO UPLOAO 568 .953 .976 .979 .372 .997 ZERO UPLUA0 111 C 1.003 1.C30 1.000 1.303 1.003 e
Technical Report WM NSGMCES TR-5319-1 Revision 2 A4.6-14 VT. YANKEE, 5 RIAG - 1/32 M0',EL, 1.7 DELTA P, WCT 400ES ONLY TAILLE r-2 M TIME INTERPOLATION THETA THE T A = *80
. 173 160 150 140 f .
TIME (MSEC) G 1.000 1 200 1.100 1. 3 ;3 1.0C0 1.024 .993 - 33 1.~31
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Technical Report ymmm TR-5319-1 Revision 2 A4.6-15 - VT. YANKEE, 5 R I AG - 1/52 *i]#EL, 17 DELTA P, WET N3 DES ONLY RING NO. 2 RIAG N3. 3 RING No. 4 RITJG lJ0. 1 Z/L: 1.000 2/L: .875 t/L: .753 Z/L: .625 NODE THETA AREA NJ)E TE T A AREA NGDE TH: TA AREA *19 3 E THETA AREA 11 100 154 47 132 3.~ 8 . 83 1C0 3 *. S . 12 1:0 308. 12 lit 312, 48 11C 623. 84 112 e23. 122 11G 603. 13 12" 32:. 49 122 641. 85 120 641. *24
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23 14 *. 439. 59 *4 99 9 . 95 14 999. V.144 140' 199. 24 13: 514 5: 133 1029. 46 1 30 1 ;' 2 8 . 1*5 1 30 1*29 25 12; 52;. 41 120 1051. 97 120 1:51. *4A 12* 1:51. 26 11C 535 12 110 1069. 98 11C 1:59. 15 1 1 ", 1:59 , 27 LO: 269. 53 100 538. 99 1:0 533. 152 1:" L33. m
Technical Report WTELEDGEEMBEEM4GN TR-5319-1 Revision 2 A4.6-16 VT. Y A N K E E, 5 R I NG .- 1/32 HODEL, 1. 7 DELT A P . WET NODE S ONLY TABLE C - SUBM;8GED LOADS TIME : 333 MS AREA
- PRESSURE FORCE NO3E RIf4G Tri!TA MZ M-THETA (30.IN.) (P3I) (LdS) 11 1 1 00 1.?24 1.127 154.1C .2 -28.
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fh!! [j$ E TE NE Revision 2 A4.6-17 ENGrEB54G SERVICES ITEM 3: PUAR Section 3.2.3, AC Section 2.12.1 QUESTION Regarding the pre-chugging and IBA/C0 load analysis, the PUAR states that results for the symmetric pre-chug load were developed directly from the unit-load harmonic analysis done for C0. Does this mean that the water mass was accounted for as in C0 (i.e.,100 percent water mass) and was this loading applied for the cycle duration stated in the LDR?
TR 5 9 Y Revision 2 A4.6-18 N ES e "STARDYNE" Lanczos modal extraction on the 1/32"dtorus model, e 100 percent water mass dimensional fluid. e Unit load harmonic analysis. e Frequency range from 6 to 10 Hz. e Maximum response to steady state, e Fatigue usage calculated with a cycle duration of 900 seconds.
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l 4.5.1.2 Lord Definitien A4.6-20 , Rigid wall torus chugging load definitions to be applied as a wall load to a structural model of the Mark I Containment are defined below. Because the load definitions are derived as rigid wall forcing functions it is intended It also is that the torus fluid be considered _in the structural evaluations. intended that the pre-chug and post-chug analyses be steady state analyses. I Pre-Chug Load Amplitude and Circumferential Two cases shall be evaluated Distribution independently: Symmetric Distribution
. 2.0 psi uniform axially along the torus centerline at bottom dead center.
Asymmetric Distribution Values shown in Figure 4.5.1-3. l 1 Vertical Cross Section Linear Attenuation with submergence Distribution along the vetted perimeter as shown in Frequency The frequency producing the maximum response in the range from 6.9 to
. 9.5 Hz.
Pre-Chug cycle Duration 0.5 seconds every 1.4 seconds fo:.the appropriate total duration defined in Table 4.5.1-1. These loads are to be applied about the local static pressure at the appropri-ate times in the blowdown (see Table 4.5.1-1). Table 4.5.1-1 CllUCCING ONSET AND DURATIO!!S Onset Time Duration Break Size After Break After Onset l DBA 35 seconds 30 seconds IBA 5 sceends 900 seconds SBA 300 serends 900 se: ends
" *P TE WE Revision 2 i
A4.6-21 ENG4EstNG SERVICES ITEM 4: PUAR Section 3.2.4, AC Section 2.13 QUESTION The PUAR states that the modelling of the water mass in the SRV load computer model was fraught with difficulty. When the water mass was included in the model, measured outputs could not be reproduced by applying measured input to the computer model. A dry structure analysis produced acceptable results however, and therefore the dry structure analysis method was subse-quently used as a basis for all SRV analysis. This is a very troublesome point. Since there is no physical reason cited in the PUAR for using a dry containment in the SRV analysis, one is lef t with the impression that there is an error somewhere in the modelling which is fortunately compensated for by introducing a second modelling error, i.e., non-inclusion of the water in the torus. A further difficulty is the implication these modelling results have for other loads such as C0 and chugging for which a fluid-structure computer model is also used and where the water was included in the analysis. Since no verifying measurements for these losds could be made, the possiblity exists that these calculaticnt ere badly off the mark. Justify the exclusion of the torus water from the SRV analysis on physical grounds and explain why these l physical reasons differ for the C0 and chugging loads.
((_: y 1%'E WE Revision 2 A4.6-22 N SERVICES e Actual shell pressure readings from in-plant tests, o Transducers were placed to measure pressure at the water-shell interf ace. e Torus response frequency is greater than SRV forcing function. e Fluid Structure Interaction (FSI) is not significant. e Transducer measurement is the pressure exerted by the water on the shell, i.e., water mass is associated with the load. e By contrast, CO, CH loads were obtained from the Full Scale Test Facility (FSTF). e FSTF data was computer modified considering the affects of FSI and water mass.
- e. Final loading was developed for use in a structural model which includes l water mass to obtain plant unique FSI effects.
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hI ga f ep rt 9pp qq Revision 2 A4.6-24 ENGrEBtNG SERVICES ITEM 5: PUAR Appendix 1, AC Section 2.13.9 i QUESTION The Acceptance Criteria calls for the torus shell to be instrumented with strain gauges, accelerometers, and pressure transducers during SRV in-plant tests. Since no accelerometers were used in the Vermont Yankee torus, explain how data from the other instrumentation was used to compensate for the lack of accelerometers. l l
I g$ Y Revision 2 A4.6-25 DGeEEWG SERVICES e Preliminary evaluations of accelerometer data showed response to be influenced by high frequency modes. I e Accelerometer data was consistent for all TES plants tested. e Low frequency strain gauge data excellent. e The shell frequency response to the test discharge was obtained from shell strain ga rp. cata, o Mode shapes were ar. /tically developed using the STARDYNE 1/32"d torus model. e Accelerometer data was not used. l
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Revision 2 A4.6-26 ENGrEstNG SERVICES ITEM 6: PUAR Appendix 1, AC Section 2.13.9 QUESTION Appendix 1 of the PUAR mentions that calibration factors relating pre-dicted to actual pressures and predicted to actual frequencies were obtained by comparing QBUBB502 calculated values with the same quantities measured in the four in-plant tests. This appendix further states that verification of the computer model led to a further calibration f actor for the column loads. Provide more details on how the calibration f actors relating QBUBBS and the in-plant tests were obtained, especially how the Vermont Yankee method con-forms to the model calibration guidelines of the Acceptance Criteria (AC Section 2.13.9.2). Provide information on the actual forcing function used including amplitudes, frequency content and pressure wave forms. Were separate calibration factors obtained for subsequent actuations? Also pro-vide more . detail on the calibration f actor for column loads and explain why it is invariant over the frequency range of the loading.
l 1 1 Report 9p gg Revision 2 A4.6-27 ENG4EERNG SERVICES e SRV dry structure techniques were extensively studied by Mark 1 Program, e Mark I study compared data with Monticello. o Excellent correlation of calculated-to-measured shell stress e Low frequency dynamic load application produced results which were veri-fied by static analysis. o Calibration factors relating QBUBBS and in-plant tests:
- 1. Pressure time history data was obtained from the four SRV tests at Vermont Yankee for several locations on the torus shell. Both cold and hot line conditions wert tested, but all structural evaluation was based on cold line data.
- 2. The data collected for each of the four tests was best fit to the longitudinal and circumferential oistribution (profile) that is credicted by QEU6BSL2. This step is required to provice cncugn data points to derive the structural model.
- 3. The maximum best fit curve (for the four ccid tests) was useo to form a calibration f actor between measured results and results pre-dicted by QBUBB502 for the test condition. Tnis factor bounced all cold tests for the and all hot tests for the plant. It was used to aojust all predictions from QBUBBS for all other conditions.
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$ _ "'" 7%T WE l Revision 2 A4.6-28 N SENICES 1
l j 4. Pressure waveform and frequency were also based on the maximum best l fit case; the specific data was taken from the maximum response point in this test, which was bottom center of the shell.
- 5. The actual SRV analysis was done using a load that combined maximum shell pressure and maximum frequency into one bounding case. Case A1.2 provides the highest shell pressures; case C3.2 provides tne highest frequency.
- 6. The C3.2 frequency, as calculated by QBUBBS, was increased by 40 percent to account for possible frequency shifts.
- 7. As a further conservatism, the A1.2 and C3.2 cases that were com-bined were for the worst lines, not necessarily the same line.
S. The final conservatism was introduced by acconting for multiple line actuation by direct addition of pressures assuming all lines produced the worst A1.2/C3.2 combined load. e Column Load Calibration Factor - generic study performed. e Three TES plants . vere initially tested without saddles fully installed. e Torus columns yield total vertical reaction force resulting from an SRV discharge, e Column Strain Gauge data used to determine axial load.
#1 Report yp gg I 1 Revision 2 A4.6-29 SERVICES 1
e Dry structure analysis at test conditions performed for comparison to test data. l l e Calibration factors relating test and analysis: Average = .19 Minimum = .40 e Analyzed column loads are high because of assumed QBUBBS uniform distribution. { e Millstone was tested with and without saddles. The addition of saddles had a negligible effect on the column calibration factor, l e Conservative facto. developed. l l l
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Technic'al Report TR-5319-1 TN Revision 2 A4.6-32 ENG4EERNG SERVICES . ( ITEM 7: PUAR Section 4.2, AC Section 2.10 QUESTION The static load magnitude imposed on the vent header deflector in the analysis described in the PUAR seems appropriate if Figure 4.3.9-1 in the PULD accurately shows the initial impact pressure spike. Does this figure show the ccrrect impact magnitude or should it be modified as per paragraph 1 of Section 2.10.1 of the Acceptance Criteria? 1 I m _____ _ _ _ _ __ _ _ _ _ _ _
l l NS$f}'P WF WNE Revision 2 A4.6-33 SERVICES e Acceptance Criteria Section 2.10.1 deals with instrumentation response time during quarter scale testing. e The V.Y. Vent Header was tested in the QSTF without a deflector. ) i e The Vent Header Load was analytically derived from Acurex analytical criteria. o A semi-emperical approach as described by the Acceptance Criteria Sec-tion 2.10.2 was used. e AC Section 2.10.1 of Appendix A does not apply.
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R 53 9 Revision 2 A4.6-37 O ITEM 8: PUAR Section 4.3.2, AC Section 2.12.2 l QUESTION Regarding the downcomer lateral chuggir.g loads: what is the fundamental tied downcomer frequency? What was the corresponding dynamic load f actor? What was the resultant static equivalent load used in the stress analysis of the downcomer? I l
Technical Report pgg ev n2 A4.6-38 O e The Fundamental downcomer frequency is 8.63 Hz. e Vent header /downcomer intersection stiffened by gusset plates. O In the plane of the downcomers: DLF = 0.08134 Maximum design load (RSEL) = 9,176 pounds Maximum RSEL range for fatigue evaluations = 11,857 pounds e Out of plane loading: DLF = 0.02785 Maximum design load (RSEL) = 3,142 pounds Maximum ~RSEL range for fatigue evaluation = 4,060 pounds o The FSTF DLF is 0.027.
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I g1 Y Revision 2 A4.6-41 DJGNEstlNG SERVICES ITEM 9: PUAR Section 4.3.2, AC Section 2.12.1 QUESTION For synchronized multiple downcomer lateral chugging loads, the Accep-tance Criteria Specification is based on an exceedance probability for 10-4 per LOCA. The PUAR shows that two load cases were considered for multiple lateral loads. Why were only two load cases necessary and what static loads were applied? To what exceedance probability did these load magnitudes cor-respond? 4 l
Report 9p gg Revision 2 A4.6-42 DJGBEBWJG SERVICES o Pool chug synchronization. e Probability of exceeding a given force magnitude during multi-downcomer chugging is determined from FSTF data. e Twelve downcomers were considered to be chugging synchronously. e Resultant load applied 9699 pounds e Probability of non-exceedance (PNE) is less than 1 in 100,000 chugs (10-5), o Assuming all downcomers chugging synchronously with a PNE of 10-4, the resultant FSTF load is 600 pounds. e A Generic Analysis was performed using a 180 0beam model, e Static applications of seismic and anti-symmetric chugging were applied. e The ratios of the combined seismic and chugging stress of the 180 degree /45 degree models are: 1 970 psi /7851 psi = 0.13 for the downcomers 3630 psi /6020 psi = 0.6 for the vent headers e The combined stress analysis reported in the PUAR using the results from the two cases analyzed with the 45 degree model is conservative and eliminates the need for additional cases on the 45 model.
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R 5319 1 W Revision 2 A4.6-45 SERVICES ITEM 10: PUAR Section 4.3.1.1, AC Section 2.6.2 QUESTION In the calculation of stresses in the downcomers resulting from pool stell water impact, was the virtual mass of water near the downcomer accounted for, and was the 8 psid pressure called for in the Acceptance Criteria applied over the bottom 50 degrees of the angled portion of the downcomer? I
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h!, g'.'j "E WME Revistor. 2 A4.6-47 SENICES l l l e The mass of the volume of water defined by the submerged portion of the downcomer was included. I e Nodal mass was equal to the tributary volume of water. I e The eight psi pressure was applied over the bottom 50 degrees of the angled portion of the downcomer. l
1 i h 1 YME Revision 2 A4.6-48 ENGBEstNG SERVICES l e Acceptance Criteria Specifications were followed for LOCA bubble drag loads, e Includes all constraints and modifications of the flow field. e The Drag Load Assessment as required by AC 2.14.2.2 was completed, e GE Program LOCAFOR was used for the calculation up to bubble contact. e Drag loads on structures above the downcomer exit elevations were calcu-lated using the local maximum pool surface velocity. l l l
Technical Report WT51 WE R on 2 A4.6-49 SERVICES ITEM 11: PUAR Section 4.3, AC Section 2.14 QUESTION Were the LOCA bubble drag loads calculated according to Acceptance Cri-l teria specifications as given in Section 2.14.2 of the AC? ANSWER Yes, all loads were considered. l l l l l
l a Report 9p gg Revision 2 A4.6-50 ENGNEBUNG SERVCES ITEM 12: PUAR Section 5.2 l QUESTION l l Many of the ring girder loads for Vermont Yankee were analyzed using a computer model constructed for another Mark I plant of similar dimensions. Mhat is the other plant? What are the dimensions of the ring girder and surrounding shell structure of this other plant? Are attachments to the ring girder similar? Were the loads used on this model the Vermont Yankee loads or the loads from the other plant? l 1
Ig Y Revision 2 A4.6-51 SERVICES e The Vermont Yankee ring girder was analyzed using the Millstone model, e The plant unique loads for Vermont Yankee were used. e Critical Dimensions: Ring Girder Depth Shell Thickness Web Thickness i e Critical location ring girder to torus shell weld. e Controlling load case for the ring girder is a lateral pressure load. e Larger ring girder depth for Millstone resul'ts in a larger load at the weld due to a longer moment arm, e Thicker shell at Millstone results in higher weld load due to the increased shell stiffness. l e Reduced saddle at Millstone results in additional load concentrations at welds, i
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al Report 97 gg Revision 2 A4.6-54 ENGNEBMNG SERVICES ITEM 13: PUAR Section 5.3.2, AC Section 2.14 QUESTION Calculation of ring girder drag loads were not in accordance with the Acceptance Criteria. Therefore provide the details of a submerged structure load calculation for a given segment of the ring girder. Include numerical values of a VT/D calculation, as well as source strength, as a function of frequency. In addition, provide the acceleration volume, drag coefficient, interference effect multiplier and pertinent geometric parameters and con-figuration used in the calculation. l l
c a Report 9p gg Revision 2 A4.6-55 BIGNEstNG SERVICES e NRC increased factors for acceleration drag volume are shown in Slides 1, 2, 3 and 4 and are justified by data for high value of VT T i l e The theoretical acceleration drag volume for cylinders is twice that of a l l flat plate. The NRC increase (f actor of 4 on acceleration drag volume) for flat surfaced structures by using the cylinder with diameter equal to 2 DC is justified by results from Keulegan & Carpenter shown in Slide 5, if VT 125. V If VT 0, then Slide 5 indicates the NRC method is very conservative l D I e The NRC increase (f actor of 2 on acceleration drag volume) for wall interference is supported by Slide 6 obtained from Sarpkaya, if VT is greater than approximately 5. V If VT is 0, then the hydrodyamic mass coefficient increases from the the(reticalvalueof2to3.3foracylindernearaboundary. l e TES's method effectively uses a f actor of (3.3/2) = 1.65 for the increase of I drag volume for wall interference instead of the factor 2.0 l And, instead of using the Deq = 2 DC increase in acceleration drag volume (f actor of 4). We considered the ring girder a cylinder with a diameter of DC, resulting in an increase by a factor of two over ! the value obtained if the girder was a flat plate. l
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Technical Report 1 TR-5319-1 WTELEDGE ENCNEEIW4GN Revision 2 A4.6-59 1 1 { h Deff' 20C J 87 Effective Cylinder using NRC factors for Flat Surfaced Structures and Wall Interference l' i l l l Wall interference increases the acceleration drag volume by a factor of two
Technical Report TR-5319-1 WM MM Revision 2 A4.6-60 i t 5 - [ Y's factor
, / of 4 bl using D =rt Dc
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pure potential l flow on flat plate I - 0 2S 50 FS 60 0 125 Uni/ D Fucunc 12. Variation of inertin coe.ficient of plates. Diarneter (inthril: ,3 2. 5 2 1. 5 1.25 1 0.75 0.5 Conespan.itnt symbol: OA.O O A El + + , 4 Increases in Cm, Hydrodynamic Mass Ccefficient on a Flat Plate i
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1 1 1 Technical Report WTELEDGEEN(BEEfW4GSERVICES TR-5319-1 Revision 2 M.5-6)
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53 9- WW ' Revision 2 A4.6-62 Em SERVICES ITEM 14: PUAR Section 5.3.2, AC Section 2.14.5 QUESTION The PUAR states that FSI effects are accounted for in the submerged structure loadings. Additional detail is needed on how this was done. Is the criteria for including FSI effects the same as that stated in the AC? How were the FSI loadings obtained? Is the boundary acceleration added to the local fluid acceleration as suggested in the AC or has another method been used? 1
l l Report yp qq l Revision 2 A4.6-63 SENICES l l e FSI effects must be included when local fluid accelerations are less than twice the torus boundary acceleration. e Continuum Dynamics was contracted by GE owners' group to " map" the pool accelerations, e Plant Unique C0/CH shell response was used for mapping from the 1/32ND FEM analysis, e Phasing and direction of the fluid accelerations at a given frequency in the region of the structure were available to calculate the FSI contri-bution of fluid forces on a segment. e Loads are calculated in 1 Hz intervals to 31 Hz. l l l
l Technical Report l TR-5319-1 SP F W NE ! Revision 2 , A4.6-64 ENGNEERING SERVICES ITEM 15: PUAR Section 7.1.3 Question I The PUAR states that the catwalk structure stresses were computed with-out the catwalk grating. Does this mean that the grating is normally absent and will only be put in place when the catwalk is used? If the grating is always in place, by what amount will it raise catwalk stresses? l l e
T 9 P SP F W NE Revision 2 A4.6-65 ENGNEERNG SERVCES 1 e Catwalk was removed from 14 bays. l e Catwalk exists in two access hatch bays only. e The grating is removed during normal plant operation. l l l e Analysis is complete. e Report to be revised in future. i I l I l l l
Technical Report N SE:HVN ' TR-5319-1 Revisinn 2 A4.6-66 a) STEEL CABLE NQ
.l: ) .
N RE T A NT BARS
\ '
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during operation 9[
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[' 4" SCH 80 Pl? f D V ' is ,.
/ ' .: , i 4 "DI A STEEL PIPE hd 9 .
RING GIRDERS g) FIG.2-15 C t,TWALK S H A N D R AIL MOD;FIC AT!CN (2 Bus onY)
Technical Report TR-5319-1 SPTA WNE
'wision 2 A4.6-67 ENGNEERING SERVICES ITEM 16: PUAR Section 8.1, AC Section 2.13.8, NUREG-0783 Section 5.1 QUESTION The use of a local temperature of 210 in the equation for mass flux rate 42 #m/sec-f t2 on page 105 of the PUAR seems to be based on a misinterpretation of the guidelines in NUREG-0783. In order to get to 210 , the quencher submergence must be at least 14 feet (14 feet of water corresponds to a total pressure of about 20.8 psi, so the saturation temperature is 230U . Subtract the 20 subcooling and one gets 210 ). Although no exact submergence of the quencher for Vermont Yankee can be found in the PUAR, it cannot be much more than seven feet. Therefore, the saturation temperature minus the 20 subcool-ing at that submergence will be not much above 200 F. Also, Figure 8-1 does not clearly answer the question of maximum bulk pool temperature. What is the maximum bulk pool temperature reached during any of the transients required for consideration and does it conform to the Acceptance Criteria in light of the above comments?
i
Technical Report TR-5319-1 Revision 2 A4.6-68 i ITEM 16 l T-Quencher Suppression Pool Temperature Limits Vermont Yankee NUREG-0783 2 Mass Flux, Ibm /sec-ft <42 <42 T-Quencher Submergence, f t 7* 14 Total Pressure, psia 17.6 20.8 Saturation Temperature, F 221 230 T minus 20UF, UF 201 210 sat Local Pool Temperature Limit, F 200 (nominal) 210
'.3- Bulk Pool Temperature Limit, F a) RHR Off 157 --
b) RHR On 185 -- Maximum Bolk Pool Temperature, UF al75 -- (Stuck Open S/RV From Full Power)
- Actual Submergence - 7'9"
1 Technical Report WTELEDGE EN(MEERNG SGMCES TR-5319-1 Revision 2 A4.6-69 Comparison of T-Otamcher Bulk Suppression Pool Temperature Limit to Stuck Open S/RV From 100% Power Transient Responses 220 -
- -! I
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Technical Report ymmm TR-5319-1 Revision 2 Figure 8-1 (Revised) Comparison of T-Quencher Bulk Suppression Pool Temperature Limit to Stuck Open S/RV From 100% Power Transient Responses 220
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Technical Report WMMM TR-5319-1 Revision 2 A4.6-71 e ~ O C _0 . 5 o
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a Technical Report TN N A4*6-72 TR-5319-1 Revision 2 NEDE-25542-P GE COMPA.W PROPRIETARY ass III M c u l i C_C L_L_ O T/t t out Tic rJ $ I
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9 Technical Report A4.6-73 WTERMEmm i TR-5319-1 Revision 2 3 tnt.2:s42-P GE COMPh<r IKOf RIE!ARY } Class III I ficn'io c ELL O TlC L v ' n i ' 4 W' J SOUTH 0*
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Report ygg Revision 2 A4.6-75 N ES i ITEM 17: AC Section 2.1 QUESTION Section 2.1 of the Acceptance Criteria states that "as part of the PUAR, each licensee shall specify procedures (including the primary system parame-ters monitored) by which the operator will identify the SBA, to assure manual operation of the ADS within the specified time period. Longer time periods may be assumed for the SBA in any specific PUA, provided: (1) the chugging load duration is correspondingly increased, (2) the procedures to assure man-ual operation within the assumed time period are specified, and (3) the poten-tial for thermal stratification and asymmetry effects are addressed in the PUA." The PUAR does not specifically address the above requirement. Clarifi-cation is needed.
i Technical Report TR-5319-1 Y Revision 2 A4.6-76 1 l l e The USNRC in conjunction with the TMI BWR Owners' Group is in process of reviewing the Plant Emergency Procedure Guides (EPG's), e V.Y. will review and implement the EPG's as necessary for compliance. l l l
l TR 9 iD l Revision 2 A4.6-77 ENGretlNG SERVICES l
\
l 1 ITEM 18: PUAR Section 4.3.1, AC Section 2.6 l l QUESTION Provide pool swell impact and drag transient histories used in the calcu-lation of pool swell loads on the main vent, vent header and downcomers. Provide enough detail to show how the load histories applied at the nodal points of the shell and beam models comply with the Acceptance Criteria. I i I l l l
Technical Report
' 9 ev s on 2 A4.6-78 N N ES i
l e Pool Swell Impact and Drag Transient Histories for Vent System - General. e Pool Surface Displacement and Velocity Curves taken from PULD. e Main Vent was analyzed in accordance with AC Section 2.6.3. e Effective volume for acceleration drag and buoyancy reduced to include only the submerged portion of the vent (SSE response from GE). e The impact was distributed to correspond to the subdivision of the vent into an infinite number of segments, e Pool transient history for main vent at a particular location provided. e Downcomers (addressed in Item 10) were analyzed per AC Section 2.6.2. e Vent Header was analyzed in accordance with AC Section 2.6.1. e Method of analysis is outlined in the response with a typical set of FEM loading functions. i
Technical Report WTELEDGE ENCNEBWW2SBMCS l TR-5319-1 l Revision 2 A4.6-79 . e
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Technical Report TR-5319-1 WTELEDGEENCNEEIWGSGMCES Revision ? A4.6-81
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Technical Report WTELEDGE EN(MEEfW4G SERVCES TR-5319-1 Revision 2 A4.6-82 g SURRARf 0F FROTH LOAOS WATER LEVEL .10 623 VERMONT YANKEE DELTA Psi.T EXTENDEO CURVES VT. VAN 4EC MAIN VENie POOL SWELL ' IMPACT AND ORAto OPER. DELTA P L FRO 1H I TIME FROTM II TIME OF PRESSURE OF IMPACT PRCSSURE IMPACT
.T895 0.- 6.
2.369' O. - 6. 3 948 6. ,0.; 5.527 0. 6. T.106 ~ 0. O. 8 685 C. O. 1 . SURNARf 0F FROTH LOA 03 WATER LEVEL 10 623 WERMONT YANKEE DELTA Pa17 EXTERDED CURVES MAIN. VENT ABOVE VENT / VENT HOR. INTER $EC. L FROTH I TINC FROTH It TIME OF PRESSURE OF IMPACT PRESSURE IMPACT
; 1 494 1 337 .3018 .
4.482 1 312 .3010 f.470 1 28T .3010 j 1
- ~
SUMMARY
OF POOL FALLBACK LOADS WATER LEVEL ' 10 623 .- .- VERMONT YANKEE . . DELTA Ps'I.T EXTEh0C0 CURVES ~ , . . j . VT. TANKEC MAIN VENT, POOL SWELL IMPACT AND ORA 8e OPER. DELTA P
' ' '1 SUMMART OF POOL FALLEACK LOADS WATER LEVEL 10 623 VERMONT YANKEE . DELTA Ps't.T EXTENDE0 CURVCS MAIN VENT ABOVE VENT / VENT HOR. INTERSEC.
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Technical Report TN N TR-5319-1 Revision 2 A4.6-83 WATER LEVEL = 10.623 ELEMENT 1 SUS ELEMENT 1 . POOL SURF ACE AT x -LOCATION CF ELERERT - TIME Y VELOCITY CW RW CY RY -
.569 3 930 .7 450 .745 13.000 .725 9 848
.512 3.185 11 503 .745 15 440 .F25 8 820
.455 2.042' 18.229- 745 24.468 .725 7.244 398 1 221 28.983 .745 38 903 .725 6.112
.341 . .135 23.131 .745 31 049 .725 4.241 l
.2A5 -1 467 18.206 .745 24 437 .725 2.404 I
.228 -2.445 12.862 .745 17 265 725 1 355
.171 -2.933 7.361 .745 9 881 .725 .343
.114 -3.025 4.908' e745 6 587 .725 .255
.057 -3 118 2 454 .745 3 294 .725 .128
.000 -3.210 .000 .745 .000 .725 .000 e FORCE-TIME LOAD DATA's FORCE OF EMPACT s 2 24865E*04 TINC OF IMPACT s 2.89287E-01 FORCE OF ORAG s 1.04634E+04 TIME OF MAX IMPACT a 2 89287E~C1 VELOCETY AT IMPACT s 1 86200E+01 TIME or ORAG a 3.35706E-G1 MAX TIME a 5.690 00 E -01 ORA 0 COEF. 1 344E.00 UNIFORM IMP ACT s' .14241E+0* -
UNIFORM DR AG a .66266E.09
** POOL F ALLB ACK LOADS ARE NOT APPLICARLF AT CLEMENT FROTH REGION 1 ELEMENT IS HELOW VENT HEADER FROTH I DOES NOT APPLY FROTH REGION It
- i
** ELEME4T !$ NOT APPLICARLC TO FROTH IMPINGEMENT LOADS REGION !! POOL PROFtli FROTH F ALLB ACK LO ADS FROTH FALLBACK OENSITY mAft0 s .2a0 FROTH FALLSACK PRESSURE (PSI) 1 662 l
i i Technical Report WTELEDGE ENCMEBW4GN TR-5319-1 Revision 2 A4.6-84 l l
SUMMARY
OF IMP ACT AND ORAG LOADS UATER LEVEL 10.623.. VERMONT YANKEE DELTA Ps't.T LXTCh0E0 CURVES WT, YANKEE MAIN VENT #00L' SWELL IRPACT AND ORAS, OPER. OCLTA P L FORCE OF TIME OF FORCE OF TIME UNIFORM- UNIFORM
- FRON IMPACT IMPACT. ORAS ORAS IMPACT ORAS REF (LOS) (SEC) (L833 (SEC) (L8 /FTD (L8 /FT)
.790 .225E*05 .2893 .10SE*05 .3357 .142E+05 .663E+04 2 37 .250E+05 .3007 .105E+05 .3447. .158E*05 .663E.0 4 3.95 .269E*05 .3089 .105E*05 .3514 170E*05 .663E.04 5.53 .302E*05 .3227 .105E*05 .3627 .191E*05 .66 3E+0 4 7.11 .396E+05 .3561 .10 5 E
- 0 5 .3910 .251E+05 .663r+04 8.68 .524E*05 .3928 .105C+05 .4232 .332E+05 .663E+04 1
SUMMARY
OF IMPACT AND ORAG LOADS U4TER LEVEL 10.623 l VERMONT YANKEE DELTA P=1 7 EXTENDED CURVES MAIN VENT ABOVE VENT / VENT HOR. INTERSEC. ' L FORCC 0F TIME OF FORCE OF TIME UNIFORR UNIFORK FROM IMPACT 1RPACT ORAS ORAS IMPACT OR44 RC8 (L85) (SEC3 (L83) (SEC3 (L8/FT) (L8/Fil 1 49 .283E+05 .4697 .132E*05 .5266 .946C+04 . 443r.04 l , I
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l Technical Report WTELEDGE mm TR-5319-1 Revision 2 aa. A M J VT. YAMMEC MAIN VENT, POOL SWLLL IRP ACT AND OR AG, OPER. DELTA'P, ,-
' 1 ELEMENT NUMBER SUS ELEMENT NUMBER . . 1 AREA t. < 10 723- '
J ,, 708.020 HYDRO. MASS i *
- 3.396 CIRCUMSCRIBING RADIUS LENGTH 6.F92 TOP THETA .336 80TT05 THETA .336
- N*
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' INCREMENT IN N, 1 4910 l INCREME47 IN Y
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1 Report 97 gg Revision 2 A4.6-87 94GNERING SERVICES ITEM 19: PUAR Section Al, AC Section 2.14.3 and 2.14.4 QUESTION Use of SRV test data for submerged structure drag loads represents an exception to the Acceptance Criteria. The method described in Appendix 1 of the PUAR needs to be reviewed further, however, several problems which arise immediately are listed here:
- 1. The frequency content of different SRV load cases have been shown by experience to be different - multiple valve actuations show a lower frequency content that single valve tests. The PUAR method does not address this problem. Arguments in the PUAR that " structures involved are responding to a farily uniform random field" are unconvincing.
- 2. Using a uniformly distributed pressure as a way to obtain static loads giving strains equivalent to those measured can lead to non-conservatisms when Figure Al-5 is used to predict static drag pres-sures on structures whose geometry is different from those on which I the strains were measured.
- 3. Scaling the static drag pressure upward from test conditions to more severe SRV cases by the ratio of calculated shell pressures is an oversimplification which uses a global parameter to scale local effects. The local pressure on a submerged object due to simul-taneous multiple SRV actuation can ratio very differently from the torus shell pressures, depending on the phasing and location of the quencher relative to the object.
j
}[fj'!cd{"Prt itTm m(NE l Revision'2 A4.6-88 ENGREstNG SERVICES l l
SRV DRAG LOADS e How were the effects of frequency differences between SRV conditions considered? e How will the application of uniform pressure be affected by different structure geometries? e Why was-drag force scaled to shell pressure? l
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i Technical Report TR-5319-1 TN Revision 2 A4.6-89 BIGSEB51G SSWICES FREQUENCY SCALING e Evaluation of Test Data showed:
- 1. Submerged structures respond at their natural frequencies regard-les: of bubble frequency.
- 2. No resonant-type response occurred in any submerged structure regardless of natural frequency.
e Test results do not support the idea of harmonic excitation.
Technical Report SPT51 WE s on 2 A4.6-90 ENGNEERNG SERVCES I Response Frequencies of Instrumented Submerged Structures 8.1 hz 8.2
- 12. (approx.)
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n aca.cuum Technical Report m gg. TR-5319-1 A4.b-91 Revi,lon 7 In-Plant SRV Test Results Comparative Frequency Content l Millstone - Test 4-Cold 1 l
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Technical Report TR-5319-1 SPF M Revision 2 A4.6-93 N SENICES
\
AFFECTS OF DIFrERENT GE0METRIES e Test data was collected for t Small Cylinders (V.H. supports) Large Cylinders (downcomers) Angle Section (catwalk supports) No other geometries are involved. e Most structures with significant drag load were instrumented during one of the SRV tests. l l
l Technical Report SPTA RVNE TR-5319-Revision 2 A4.6-94 ENGNEERNG SERVCES Submerged Structures at Vermont Yankee Total 16 In SRV Bays 8 Not Affected by SRV Drag -4 4 Vent Header Support Columns Downcomers Downcomer Tie Bars Catwalk Support Columns l l l l l
Technical Report $PM Re s on 2 A4.6-95 DRAG FORCE SCALED TO SHELL PRESSURE e For hot and cold conditions, drag loads were compared to: Bubble Pressure and Shell Pressure e Correlation with shell pressure was consistently good. l
I g$1 Y Revision 2 A4.6-96 DJGrEBWJG SERVICES ITEM 20: PUAR Section 6.0, AC Sections 2.14.3 and 2.14.4 l l OUESTION The PUAR analysis of the t-quencha'., its support and the submerged por-tion of the SRV line does not menti sn quencher water jet or bubble drag loads on these structures. Where b e these loads been included or why have they been ignored? l
Tet.hnical Repori 1PTF1 WE ev s n2 A4.6-97 SENICES l l l l l e LOCA jet and bubble drag loads from downcomers were considered in the tee-quencher, support and pipe stress analysis, e The maximum pool swell event combination (Event 25) which includes these loads is bounded by a C0/CH event combination (Event 15). e Water jet drag loads do not act on these structures, e Equivalent adjacent structures have a two bay spacing. l l l l
Technical Report 9p qq Rev s n2 A4.6-98 SERVCES ITEM 21: _UESTION Q l Provide the loads that were used in the torus attached piping.
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Technical Report yg Rev n2 A4.7-1 94G#EstNG SERVICES l REVIEW BY BROOKAVEN NATIONAL LABORATORIES (CONTINUED) A4.7 Response to Additional BNL Questions September 2, 1983 i I l l l ! I
P WP WNE Revision 2 A4.7-2 ENGPEERNG SERVICES LOAD FACTOR FOR C0 SHELL DISPLACEMENTS Q. Provide further justification for the use of the 1.15 f actor that was applied to shell response, to define inputs for torus attached piping (question by Prof. G. Bienkowski). A. Since this question was asked, Professor Bienkowski has received a copy of SMA report number SMA 12101.04-R0020, " Response Factors Appropriate for Use with C0 Harmonic Response Combination Design Rules". This report provides additional information on the basis and development of the 1.15 factor. In a telephone conversation on September 9,1983, between N. Celia (TES) and Prof. Bienkowski, it was concluded that this additional document provided the necessary justification to answer the question and additional information was not required, i l 1
Technical Report WTFI PTT(NE R S on 2 A4.7-3 VERMONT YANKEE SUBMERGED STRUCTURE SRV DRAG LOADS Q. Provide the margin of safety on all submerged structures which are affected by SRV drag loads. A. Those structures that are affected by SRV drag loads are shown on the following table with their resultant margin of safety on SRV drag. l l l l l
i Technical Report TN !
.TR-5319-1 Revision 2 A4.7-4 DIGSEstNGSERVCES !
l ) l SRV DRAG MARGIN
SUMMARY
Controlling Condition SRV Allow. Load A1.2 Total Allow SRV Drag Item Location Type Case Load Load Load Mult. ) t l Downcomer VH/DC Primary 14 16547 psi 26729 psi 37635 psi 1.83 ! Intersection Local & i Bending
- Downcomer Clamp Bending 25 1000 psi 16800 psi 22240 psi 6.44 a Tie Bar fCataalk Inboard Axial & 25 1712 psi 20157 psi 42000 psi 13.76 t
Support Diagonal Bending Brace Ring Web Membrane & 16 1387 psi 19733 psi 28950 psi 7.65 Girder Bending i Ring Node 522 Shear 15 .587 5.916 8.53 5.45 Girder lb/in lb/in Ib/in Vent Header Clevis Shear 25 50 psi 5698 psi 13840 psi 163.84 Support Plate 4 4 Vent Header Column Tension & 14 .094 f/F .595 f/F 1 f/F 5.31 Column Bending l i i
Technical Report Y DE TR-5319-1 Revision 2 ENGNEERNG SERVCES A4.7-5 The lowest safety margin for SRV drag is for the downcomer intersection. This is accounted for, in part, by the fact that the SRV bubble frequency during the test is the same as the natural frequency of the downcomer (8.6 Hz). This comparison was made based on Fourier analysis of the test data for bubble pressure and downcomer strain for Vermont Yankee, and Millstone, which has an identical downcomer. Evaluation of the downcomer intersection is also known to be conservat-ive because the lateral loads due to chugging and SRV drag were both taken in a direction about the downcomer to produce a maximum intersecticn stress. The unlikelihood that both these levels will occur simultaneously in the critical lateral direction, in phase, further contributes to the conservatism in the evaluation, l f, l 4
1 Y Revision 2 A4.7-6 VERMONT YANKEE SRV TEST PRESSURES , l Q. Describe physical differences between Vermont Yankee and other plants that would account for the low shell pressures measured during the in-plant SRV test. A. We have examined physical dimensions and test data from the four similar plants tested by TES, and found several contributing factors. A compari-son of the major physical dimensions shows V.Y at the bottom end of the range; these dimensions correlate with measured test pressures, as follows: Air Water Quencher Maximum Voluge Volgme Depth Test Pressure (ft ) (ft ) (feet) (psi) Pilgrim 58.9 17.8 7.3 7.8 Millstone 46.5 12.3 8.9 7.8 Fitzpatrick 58.2 11.8 9.0 4.7 Vermont Yankee 41.9 11.25 5.9 2.5 This comparison places V.Y. at the low end of these important parameters in every category. As a part of this review, an additional check was made to correlate gross shell motion to measured pressures. Tests at Millstone and Vermont Yankee each had a displacement transducer mounted on the outside of the shell at bottom center. A ratio of maximum shell displacement and maximum shell pressure was developed for all tests. The results are:
h[ a Report WTF1 WE Revision 2 A4.7-7 BGEERNG SERVCES Millstone Max. Displacements Average of_hax. Pressures = .017 Average of l Vermont Yankee l Average of Max. Displace.nents = .014 Averr10 of Max. Fressures l This ccmparison demonstrdtes the consistency of the displacement and pressure readings at V.Y. and helps confirm the accuracy of the pressure readings from the test.
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Technical Report WM o 2 A4.7-8 BIGrEStNG SERVICES METHOD OF CONDENSATION OSCILLATION PHASING Q. Describe the method used to phase the harmonic components of load on the torus structure resulting from Condensation Oscillation (CO). A. The analysis of the torus structure for C0 consists of the shell stres-ses, torus reaction forces and torus attached piping penetration dis-placements. The most current document available on harmonic phasing for C0 loads is Reference 4 which evaluates the additional M-12 and M-11B test data in detail. Excerpts of Reference 4 are included for your review with the references. SHELL STRESSES A unit load harmonic analysis of the torus structure (Reference 5 - DYNRE2) was performed in accordance with the Load Definition Report (LOR) (Reference 1 - Section 4.4) for the Condensation Oscillation load-ing (CO). The analysis of the torus shell was completed to 31 Hz as suggested by Section 6, page 11 of Reference 4. Component stresses were extracted from the computer analysis as a function of frequency for those element locations chosen for further stress analysis. The procedure for the torus shell C0 stress analysis is initiated by factoring the individual element component stresses by the three sets of C0 baseline rigid wall amplitudes defined by the LDR (Reference 1 - Table 4.4.1-2). The factored component stress harmonics were then combined by the 84% NEP harmonic combination rule as defined by Reference 4. That is, the four highest contributors are assumed to be in-phase and are combined absolutely prior to being summed with the SRSS of the remaining contributions to 31 Hz. The resultant component stresses from each of the three sets of C0 amplitudes are compared and the maximum resultant is obtained. Principal stress intensities are calculated from the resultant component i stressesinaccordancewiththeASMEBPVC(Reference 2).
1 1 Technical Report WM l 7 M .7-9 l Justification for use of the 84% NEP rule is discussed in Reference 4 (excerpted). This reference specifically deals with the evaluation of the M-12 test data which is the bounding test for all major FSTF responses as shown in Section 6, page 8. Further, this data indicates that the 84% NEP is an excellent approximation for shell stresses, while the 50% NEP closely approximates displacements and reaction forces. The 84% NEP harmonic combinatic.n rule was also chosen for analysis of the torus shell stresses based on the recommendations of Reference 4 which indicates (Section 6, page 10) that the 50% NEP rule will underestimate stresses by 10 to 15 percent. Note that this report judges that this under estimate will not occur for the design analysis and suggests that additional conservatism may be added into the design approach by use of the 84% NEP harmonic combination rule. REACTION LOADS (BASEMAT LOADS) Reaction forces were obtained using the procedure previously outlined for shell stresses. However, the three highest contributors were assumed in-phase and combined absolutely prior to being summed with the SRSS of the remaining harmonic contributions. This 50% NEP rule is used in accordance with Reference 4. DISPLACEMENTS FOR PIPING Displacement data obtained from the torus structural analysis at the piping penetrations was randomly phased to develop a single artifical
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response time history in accordance with page 2 of Ref erencA 6. The time history amplitude was increased by 15% before application to a specific torus attached piping system. 6 j y ,' (
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TR 5319 W Revision 2 A4.7-10 NO l EXCERPTS FROM REFERENCE 4 TO SECTION A4.7 (SMA REPT. 12101.04-R0010) PROVIDED AS PART OF RESPONSE TO LAST QUESTION IN SECTION A4.7 1 l
Technical Report TR-5319-1 Revision 2 A4.7-ll
- 1. INTROD'J TION 1.1 PURPOSE AND SCOPE Or WOM The loads on the Mark I containment system resulting from conden-sation oscillation (CO) phenmena are based on the Full Scale Test Fa:ility (FSTF) tests, in particular, the CO load definition is based on FSTF tests M-7 and M-8 which simulate various conditions for the Mark I torus and suppression pool. Recently, additional FSTF tests, M-IIB and M-12, have been conducted for the purposes of assessing test repeatability and the effects of increased suppression pool temperature, respectively.
A previous investigation (Reference 1) of the procedure for computing structural response to CO loadings has been conducted in which potential sources of conservatism in CO load amplitudes, harmonic phasing combination and structural response computation have been evaluated relative to FSTF M-8 test measurements. In this report, the results of a similar investigation based on the FSTF M-llB and M-12 tests are presented. In the previous investigation (Reference 1), it was concluded
- that the design LOR (Load Definition Report) CO load amplitudes and struc-tural response computation method did not introduce an overly conservative bias relative to FSTF measurements (M-8 test). The LOR amplitudes seemed to be slightly conservative and the structural response calculation method was slightly unconservative. Slightly unconservative structural response calculation f or computation of FSTF response was not believed to be of concern. There are a number of f actors which might influence the compari-son between calculated and measured response and the conclusion that structural response computation is slightly unconservative. It has been
- pointed out in Reference 1 and reiterated herein, that a'lthough some of these f actors would lead to underestimation of FSTF behavior, they would be conservative for the purposes of design of structures to CO loadings.
Consequently, it has been concluded that the structural response calcula-tion approach w:uld not be unconservative for design analyses. 1-1
Technical Report TR-5319-1 Revision 2 A4.7-12 , The C0 load amplitudes to be used for design have been defined ; in the frequency domain for 50 harmonic frequencies ranging from 0 to 50 < Hz by representing the measured FSTF data as a Fourier series. In Reference 1, it has been demonstrated that harmonic response amplitudes can be assumed to be randomly phased with respect to one another, and simple design rules for harmonic combination accounting for random phasing behavior have been developed. For example, it has been shown that taking the absolute summation of the three highest harmonic ampli-tudes combined with the square root of the sum of the squares (SRSS) of the remaining harmonic amplitudes gives response which is approximately at the 50 percent non-exceedance probability (NEP) level and the absolute sum of the four highest harmonic amplitudes combined with the SRSS of the remaining harmonic amplitudes gives response which is approximately at the 84 percent NEP level. Note that the above NEP levels are defined in accordance with conditional re,lative phasing only cumulative distribution function (CDF) curves such tha,t the 50 percent NEP rule will maintain the level of conservatism associated with the load amplitudes and structural l response calculation method and not add additional conservatism due to phasing. In this report, the following steps have been undertaken in order to evaluate FSTF M-11B and M-12 behavior in terms of; 1) load
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amplitude conservatism; 2) harmonic phasing combination and; 3) structural response computation.
- 1. Represent the measured FSTF integrated downward loading on the torus as a Fourier series ranging from 0 to 50 Hz at intervals of 0.914 liz. The Fourier series is given by Fourier amplitudes and Fourier phase angles. Note that the loading is investigated in four, one second segments, and Fourier series are derived for each segment. Also, all work described below has been conducted in these four one second segments for each FSTF test.
- 2. Compute conditional relative phasing only CDF curves using the measured Fourier amplitudes and assumed random phasing.
These CDF curves are then compared with measured integrated downward loading and loading computed by the harmonic pnasing combination rules as an indication of whether or not the rules achieve the desired NEP level for these tests as well as wnetner or not the design rules realistically simulate measured FSTF phasing behavior, l
- i-2 1
Technical Report TR-5319-1 - Revisjon 2 A4
- 3. Develop rigid wall Fourie.7-13r amplitudes and phase angles from i
the measured FSTF flexible wall integrated downward loading I by removing the effects of fluid-structure interaction I (FSI), Rigid wall loading independent of fluid-structure interaction effects can be applied as loading to an analytical model of the fluid-structure system of any Mark I plant to obtain response incorporating the plant unique j effects of FS!; thus, the LDR CO load is expressed as a l rigid wall loading. Convert the rigid wall Fourier amplitudes for integrated downward loading to equivalent i bottom dead center (BDC) pressure by using the LDR spatial l distribution of pressure on the torus. FSTF rigid wall BDC l pressure amplitudes may then be compared directly with LDR C0 loading.
- 4. Investigate the FSTF Fourier phase angles corresponding to rigid wall loadings for deterministic phase relationships.
Conduct a brief search of the FSTF Fourier phase angles for deterministic phase relationships with the dominant frequency and other frequencies by comaring phase angles from each of the four, one second segments considered.
- 5. Compute the structural response of the FSTF torus and support columns at various locations using FSTF rigid wall l Fourier amplitudes and phase angles and the standard compu-tation approach used for design analyses. Comparison of response computed in this manner with measured FSTF response gives an indication of conservatism in the response compu-tation technique.
- 6. Generate CDF curves for structural response using FSTF rigid wall amplitudes and assumed random phasing. Compute structural response from FSTF rigid wall amplitudes and harmonic combination design rules and compare to the CDF curves to check applicability of the design rules to M-llB and M-12 tests. Also, compare structural response using design rules and LDR amplitudes in order to achieve a comparison of the FSTF and LDR load amplitudes at the response level. .
- 7. Develop factors by which the response computed from a single time history analysis based on random harmonic phasing can be scaled to produce response for which there is 90 percent confidence of being at 50 and 84 percent non-exceedance probabilities.
;-3
Technical Report TR-5319-1 Revision 2 A4.7-14 l
- 8. Determine the effe-t of only including harmonics up to 30 H: instead of 50 H2 as given by the CO load definition report on response computed from load amplitudes for these FSTF tests.
In the remainder of this chapter the C0 loading phenomena will be briefly described, the FSTF M-11B and M-12 tests will be discussed and the LDR C0 load definition will be outlined. In Chapters 2 and 3, the M-12 test evaluation is presented and in Chapters 4 and 5, the M-11B test evaluation is presented. Chapters 2 and 4 pertain to FSTF results examined at the loading level and Chapters 3 and 5 pertain to FSTF results examined at the structural response level. In Chapter 6, the test results f or FSTF M-8, M-11B and M-12 tests are compared and conclusions from this study are presented. I
Technical Report TR-5319-1 Revision 2 A4.7-lb
- 6.
SUMMARY
AND CONCLUSIONS 6.1
SUMMARY
Recently conducted condensation oscillation (CO) loading tests at the Full Scale Test f acility (FSTF) designated M-Ils and M-12 have been investigated for the purposes of evaluating the design approach for CO loadings in terms cf the most recent test data. The approach recommended in Reference 1 for design of the Mark I torus shell and supports subjected to CD loading is to employ the Mark I LDR CO load amplitudes for 50 harmonics from 0 to 50 Hz along with structure response amplification f actors based on a 2 percent damped analytical model to define harmonic response amplitudes from 0 to 50 Hz. It is then recommended that the harmonic response amplitudes be combined in a manner to achieve 50 percent non-exceedance probability (NEP) by taking the absolute summation of the 3 highest harmonic response amplitudes plus the square-root-of-the-sum-of-the-squares (SRSS) combination of the remaining harmonic response amplitudes. This design approach has potential sources of conservatism from:
- 1. LDR C0 load amplitudes
- 2. Relative phasing of harmonic load amplitudes
- 3. Response computation method The major purpose of the study presented in this report is to evaluate these potential sources of conservatism by comparison with FSTF M-llB and M-12 test data.
The Mark I LDR Design Basis Accident CO load amplitudes are primarily based on the results of the FSTF M-8 test, simalating a large liquid line break. The FSTF M-llB initial test conditions were identical to these for the M-S test. The FSTF M-12 test is identical to M-8 and M-115 with one exception. For M-12, the initial wetwell pool temperature was aco;; 40 percent higher than that for M-S and M-llB. In the 0-1
TR-5319-1 Revision 2 A4.7-16 following section, 2h@ loading and structural response from the M-8, M-11B and M-12 tests are compared on a consistent basis. The M-8 loading and response values are taken from Reference 1. - L 1 6.2 COMPARISON OF MARK I CO LOADING AND RESPONSE BASED ON M-8, l M-11B AND M-12 TEST DATA AND LDR LOAD AMPLITUDES I 6.2.1 Loadino Level Comparisons < In this section, the load amplitude from the FSTF M-8, M-11B and M-12 tests are compared on a consistent basis with LOR load amplitudes. In addition, the design rule for harmonics amplitude combination is evalu-ated against a:tual phasing of the FSTF loading by comparing peak load
. calculated using measured load amplitudes and design rule amplitude combi-nation against measured peak loading.
Rigid wall loads calculat'ed from FSTF rigid wall load amplitudes and LDR load amplitudes both using the design rule for harmonic amplitude combination are presented below: l Peak Calculated Case Rigid Wall Load (osi) LDR 6.56 M-8 5.22 M-11B 4.04 M-12 6.72 The FSTF values shown are the largest values from the various time periods considered. The data shown above demonstrates that the LDR amplitudes lead to conservative peak loadings for M-8 and M-11B and essentially the same peak loading as measured for M-12. The LDR load is a factor of 1.26, 1.62 and 0.93 times the largest M-8, M-llB and M-12 loadings, respectively.
~ ~
6-2 I l
TR-5319-1 Revision 2 A4.7-17 Peak flexible wall loads calculated from measured FSTF flexible wall amplitudes and the 50 percent NEP design rule for harmonic amplitude l combination are compared with measured peak flexible wall loads below: Peak Calculated Peak Measured Flexible Wall Load (k) Flexible Wall Case Usina Phasing Design Rule Load (kl M-8 359 332 M-11B 284 260 M-12 437 414 The values presented in the above table are the largest values from the various time periods considered for each FSTF test. The data in this table indicates that at the loading level, the design rule for harmonic combination introduces a small amount of conservatism, on the order of 5 to 10 percent ' relative to the actual phasing of the measured FSTF flexible wall loading. 6.2.2 Response Level Comparisons In this section, calculated structural responses based upon measured FSTF load amplitudes and measured phasing are compared to measured FSTF responses. In addition, calculated responses using LOR amplitudes and design rule phasing are compared to calculated responses using measured FSTF load amplitudes and design rule phasing. Finally, calculated responses using measured FSTF amplitudes and measured phasing are compared with calculated responses using measured FSTF amplitudes and design rule phasing, in this manner, the three potential sources of conservatism listed in Section 6.1 for the design approach recomended in Reference 1, can be evaluated relative to FSTF test data. l l l L-3
TR-5319-1 Revision 2 A4.7-18 Peak calculated torus shell and torus support response deter-mined from measured FSTF amplitudes and phasing are compared with peak measured FSTF response at comon locations below: Peak Calculated Peak Measured Response Resoonse Quantity M-8 M-11B M-12 M-8 M-11B M-12 BDC Axial Stress (ksi) 2.0 1.5 2.2 2.3 1.6 2.7 BDC Hoop Stress (ksi) 1.8 1.4 2.4 2.6 1.4 2.9 BDC Displacement (in.) 0.12 0.08 0.12 0.11 0.08 0.14 Inside Coltnn Force (k) 98 80 120 93 68 109 Outside Colinn Force (k) 107 85 130 110 81 141 All values presented in the above table are the largest peak values from the various time periods considered for each test. From the data presented above, it may be seen that the calculational technique under-estimates the measured torus shell stresses by between 0 and 30 percent and matches the measured torus displacements and coltnn response relatively closely. In Chapter 3, the follo.ing potential reasons for differences between calculated and measured FSTF response were outlined.
- 1. The analytical model is an approximate representation of the FSTF and inaccuracy of unknown bias may be introduced.
- 2. The assumption of 2 percent damping, which is appropriate for design responses from combined loads near one-half
, yield stress, may be too high for the low level FSTF . response. if the damping were only about 1.5 percent, resonant harmonic response amplitudes would be computed to be significantly larger such that the combined calculated response would be conservative compared to the measured response. In our judgement, damping on the order of 1.5 1 6-4 ) i l
TR-5319-1 Revision 2 A4.7-19 . percent woulc be more appropriate f or FSTF response levels than 2 percent. It must be emphasized that damping of 2 percent is believed to be approcriate f or design analyses in which stress levels of one-half yield are allowable.
- 3. The response a .plification f actors were computed at a pre-sele:ted f requency interval rather than at f requen:ies asso iated with modal properties of the structure. This procedure intrcdu:es slight uncertainty of unknown bias f or the evaluation of F5TF behavior as the a:tu'al relationship between the load amplitude and response amplification over each .914 bz interval is approximated by the values at the pre-selected frequencies. For design analyses, it is our understanding (and our recorrnendation) that a conservative procedure will be utilized in which the LDR amplitudes are multiplied by the maximurn response amplification f actor in each of the corresponding one hz frequency intervals. In this manner, some added conservatism is introduced for design analyses which does not exist for this evaluation of F5TF resoonse.
4 FSTF loadings a4e represented by the integrated vertical load as determined f rom the combination of individual pressure transducers. This integrated loading would be expected to produce reasonable estimates of colunn f orces which depend mostly on the overall loading condition. On the other hand, calculation of torus shell stresses from the integrated loading could easily result in the calcu-lated stress values being lower than the measured stress values if local variations in pressure over the surf ace of , the shell occur and strongly influence shell stresses. Since the LDR load amplitudes are also based on integrated vertical load, this potential source of unconservatism would exist for design analyses as well as for this evalu-ation of FSTF behavior. Summa-izing these four potential reasons for differences between calculated and measured response; 1) the analytical model is a source of uncertainty f or this FSTF evaluation and would also be a source of uncertainty f or design and analyses; 2) 2 percent damping is a source of I unconservatism for this FSTF evaluation but would be conservative for design analyses; 3) the response amplification f actor procedure for the FSTF evaluation is a source of uncertainty but would be a source of conservatism by the approach to be used f or design analyses; and 4) the usage of integrated vertical load to represent the F5TF and LDR C0 load 6-5
I TR-5319-1 Reeision 2 A4.7-20 conditions is a source of unconservatism in the computation of to*us shell response for both this FSTF evaluation and design analyses. Considering all of the above f actors, it is concluded that the response calculation method would not be unconservative for design analyses. Calculated maximum responses using design rule phasing and measured load amplitudes from the three FSTF tests versus the LOR 4 de inition are compared below: Calculated Response Using Design Rule Phasing Resconse Ouantity LOR M-8 M-118 M-12 2.4 2.4 1.4 2.1 BDC Axial Stress (ksi) 2.5 2.2 1.5 2.3 BDC Hoop Stress (ksi) BOC Displacement (in.) 0.14 0.13 0.08 0.13 129 113 80 126 Inside Column Force (k) 145 126 89 14 0 Outside Column Force (k) The above responses are from the largest values over the time periods considered. The tabulated values presented above provide a comparison of the LOR amplitude relative to FSTF load suplitudes at the response level. This data demonstrate that the LOR anplitudes lead to conservative responses f or all FSTF tests. The LOR response exceeds the M-E resoonse by 0 to 15 pe-cent, the M-118 response by 60 to 75 percent and the M-12 response by 2 to 14 percent.
- 6-6 m
TR-5319-1 Revision 2 A4.7-21 Calculated structural response using FSTT measured amplitudes and measured phasing are compared to calculated response using f 5Tf measured amplitudes and design rule phasing in the following table: , I Peak Calculated Response FSTF Phasina Desian Rule Phasino M-12 M-8 M-118 M-12 Resosse Ouantity M-S M-113 . 2.2 2.4 1.4 2.1 BD; Axial Stress (ksi) 2.0 1.5 2.4 2.2 1.5 2.3 BD; Hoop Stress (ksi) 1.8 1.4 0.12 0.13 0.08 0.13 BD: Displacement (in.) 0.12 0.08 120 113 80 126 Inside Coltinn Force (k) 93 80 130 126 89 140 Outside Coltnn Force (k) 107 85 F or this table, the values are also from th': largest responses for the time periods considered From the above table, it may be seen that the design rule for hamonic response amplitude combination (50 percent NEP) closely matches , the maximum responses obtained using the FSTF measured phasing relations (7 percent less to 20 percent greater). It may be concluded that, on the average, design rule responses are between 0 and 8 percent grester than corresponding response calculated from measured FSTF phase angles. A similar table to that presented above can be developed to illustrate the conservatism introduced by absolute sumation of hamonic amplitudes as f ollows: 6-7
in-aarv-i Revision 2 A4.7-22 l Peak Calculated Response FSTF Phasino Absolute Summation Resoonse Quantity M-S M-11B M-12 M-8 M-11S M-12 BDC Axial Stress (ksi) 2.0 1.5 2.2 3.6 2.5 4.0 BDC ..oop Stress (ksi) 1.8 1.4 2.4 4.0 3.2 5.4 BDC Displacement (in.) 0.12 0.03 0.12 0.21 0.15 0.26 Inside Column Force (k) 93 80 120 214 173 311 Outs de Column Force (k) 107 85 130 239 1 90 342 The calculated responses presented above denonstrate that absolute sumnation of hannonic amplitudes generally overestimates the measured phasing relations from the FSTF tests by a f actor in excess of two. Finally, it is of interest to compare response obtained from the reconnended design approach with the highest measured FSTF response. Thus, the response calculated from LOR suplitudes and both the 50 and 84 percent NEP design rules (absolute sum of 3 highest harmonics plus SRSS of remaining for 50 NEP; absolute sum of 4 highest plus SRSS of remaining f or 84 percent NEP) are presented below along with the highest maximum , measured response ,f rom the FSTF tests. Design Rule Maximum Measured Response FSTF Response Resoonse Ouantity 50% NEP 84% NEP M-8 M-11B M-12 BDC Axial Stress (ksi) 2.4 2.6 2.3 1.6 2.7-BDC Hoop Stress (ksi) 2.5 2.8 2.6 1.4 2.9 BDC Displacement (in.) 0.14 0.15 .11 0.08 0.14 Inside Column Fo-ce(k) 129 143 93 68 109 Outside Colunn f orce(k) 145 160 110 81 141 6-8
i TR-5319-1 i Revision 2 A4.7-23 From the above data, it may be seen that the 53 percent NEP design rule and LDR amplitudes give torus shell stresses which are; 0.96 to 1.04 times maxims measured M-8 response; 1.50 to 1.79 times maxime measured M-118 response; a}nd,0.86 to"0.89 times maximum measured M-12 response._ l Design rule torus displacements are ?.27,1.75 and 1.00 times maximtrn , l l measured M-8, M-11B . 3d M-12 response, respectively. Design rule coltrin f orces are; 1.32 to 1.39 times maximtin measured M-8 response; 1.79 to 1.93 times maximum measured M-118 response; and 1.03 to 1.18 times maximtra m<.asured M-12 response. Thus, the design approach for response calculation using LDR amplitudes and the 50 percent NEP hamonic combination rule generally gives response values which exceed the maximtin measured responses from the FSTF tests. ',The' ~only~ exception being the Csus~dhii t str'e'is'es"from the M-12' test.which' are underestimat'e'dTby]1[t'o'
~
CIS[pefcinT. 7!Tiiay be"noted that 'using ths 84 percent NEP harmonic' ~ ~ ' '
*implitude combination rule reduces this margin for. M-12 to'rus'ihell "str' esses'to bet' ween:3 and *4 : percent.
The results presented in the above table and described above were achieved because:
- 1. The LDR load amplitudes lead to calculated responses which are O to 15 percent greater than those obtained from the measured M-8 and M-12 load amplitudes and which are 60 to 75 percent greater than those obtained from the measured M-11B load amplitudes.
- 2. The 50 percent NEP response combination rule leads to peak responses which are between 7 percent lower and 20 percent greater than responses calculated from measured FSTF phasing. This combination rule introduces slight conservative bias as, on the average, responses obtained from the 50 percent NEP combination rule are between 0 and 8 percent greater than responses calculated from measured phase angles.
?
6-9
M.7-24
- 3. The structural response analysis method (2 percent damped -
dynamic analysis) leads to calculated torus shell stresses which are between 0 and 33 percent lower than measured FSTF response; calculated torus displacements which are from 9 percent greater to 14 percent lower than measured FSTF response and calculated column fo-ces which are between 18 percent greater and 8 percent lower than measured FSTF response values. For the M-12 and M-8 tests, the conservatism introduced by Items 1 and 2 is not sufficient to compensate for the unconservatism introduced by Item 3 in the case of torus shell stres,es. In our judgment, unconservatism of Iten 3 may be predominantly attributed to: 1) use of an assuned t damping f actor that is too high for the low level FSTF response; and
- 2) use of integrated downward load and ignoring local pressure variations on the shell to calculate shell stresses. Since the intent of design analyses is to calculate the response to CO loads plus other concurrent loads at stress levels near one-half yield stress, the assunption of 2 percent damping is appropriate.
- Conservative damping and conservative application of maximum response amplification f actors as discussed earlier for design analyses should lead to conservatively calculated response.
The design approach of using LOR amplitudes, the 50 percent NEP hannonic combination rule and 2 percent damped structure response provides a reasonably accurate estimation of FSTF~ response.(Under'-
, estimation of' torus'shell stres'ses by 10 to 15 percent is a relatively'. '
~ ~
'.small Anount 'and, in ouf judiyInent', 'will' not" occur [ for ' design' analyses for[.
. . . . . ~ . . - - ... .. .
the reasons discussed---- . earlier in this sectio,n. As a result, the purpose of this work has been fulfilled as the sources of conservatism or unconservatism have been evaluated and reasonable rules for treating harmonic phasing have been developed. Note that if add _itj on.a],. conservatism were to be added into this design approach, a small load
, f actor or use of the 84 percent NEP harmonic combination rule migh_t, be util,i z ed. Providing additional conservatism requires consideration of the conservatism of the FSTF tests and in the specification of allowable stresses in addition to the amplitude, phasing and structural response combination method addressed in this report.
6-10 l
TR-5319-1 ' Revision 2 A4.7-25 6.3 CONCLUS IONS Considering the data f rom the FSTF M-12 and M-llB tests, the design approach recomended in Reference 1 remains valid for the reasons' described above. This app.oach con:;ists of using'LDR anplitudes and s' uttural response amplification f. -tors associated with 2 percent damping to determine narmonic response coefficients. Individual harmonic responses are then combined by a design rule which accounts for the
- r. idomness of phasing of the harmor'cs. The design response should be computed as the abs. .ute sum of the three highest individual hamonic responses plus the SRSS combination of the remaining harmonic response amplitudes.
In addition, the alternate time history approach design rule based on M-8 test data as recomended in Reference 1 has also been validated by this evaluation of M-llB and M-12 test data. The alternate design rule recomended consists of a single time or frequency domain analysis in which the response time history is constructed from the LOR Fourier amplitudes and an assuned set of random Fourier phase coefficients for each harmonic. The results are scaled by 1.15 to account for uncertainty associated with the single set of Fourier phase coefficients used. It has been demonstrated in this report that _ harmonic response amplitudes greater than 30 Hz do not significantly contribute to structural response. As a result, it is also recomended that the hamonics f rom 30 to 50 Hz may be ignored in response computations f or either the harmonic amplitude combination rule or time history scale f actor design approaches. Note that this conclusion is based solely on consideration of torus shell stresses and support colunn forces and additional evaluation wauld be necessary to determine if a 30 Hz cutoff w]uld be justifiable for other structures, piping and equipment. 6-11
i A4.7-26 Finally, it has been demonstrated in this report that the 1 1 I recomended design rule f or harmonic response amplitude combination intro-du:es conservative bias on the order of 0 to 8 percent and that absolute spation of harmonic amplitudes overestimates the effect of measured phasing relations by a f actor of 2 or more. Thus, it is recomnended that l abso'ute smation not be considered for harmonic amplitude combination. e I 4
)
6-12 l l
Technical Report TR-5319-1 Revision 2 A4.7-27 9 g ENGNEERING SERVICES REFERENCES TO SECTION A4.7
- 1. Mark 1 Containment Program Load Definition Report NED0-21888, Rev 2 November, 1981
- 2. ASME Boiler and Pressure Vessel Code Section III, NE-3000 through Summer 1977 Addenda
- 3. Mark 1 Containment Program Structural Acceptance Criteria Plant Unique Analysis Application Guide October, 1979
'4 . Evaluation of FSTF Tests M-12 and M-11B Condensation Oscillation Loads and Response SMA 12101.04-R001D August, 1981 P.1-2, 6-8 to 6-12
- 5. STARDYNE-3 System 1977 System Development Corp.
2500 Colorado Ave. Santa Monica, CA 90406
- 6. Response Factors Appropriate for Use with C0 Harmonic Response Combination Design Rules March, 19E2, SMA 12101.04-R002D i
l l
Appendix 5 - Teledyne Additional Information in Response to NRC Review of the Vermont Yankee PUA Report. Appendix 5.1 Information presented at February 16, 1984 meeting between Teledyne, NRC, Brookhaven, Northeast Utilities, and Yankee Atomic Electric Company (YAEC) at YAEC offices in Framingham, MA. Appendix 5.2 Responses to open items resulting from February 16, 1984 meeting at Framingham, MA. Appendix 5.3 Responses to questions discussed in March 9 and 12, 1984 conference calls between Teledyne, NRC, Brookhaven, and YAEC. i
Appendix 5.1 SPTELEDYNE ENGNEERING SERVCES l l l l MARK 1 TORUS PROGRAM RESP 0flSES TO REVIEW QUESTIONS FOR PUA REPORTS FOR MILLSTONE Afl0 VERMONT YAiiKEE February 16, 1984
i SPTELEDYNE ENGINEERING SERVICES 4 a i MAJOR ISSUES RELATED TO SRV TEST AND ANALYSIS f I e SRV tests with drywell A P. e SRV test shell pressure magnitudes. I e SRV drag test results and analysis method. I r i a 6 e i 1
I WM ENGNEERING SERVICES I USE OF SRV TEST DATA FOR ANALYSIS i i !} Effect of AP on SRV Analysis r 4 e Drywell-wetwell AP was in place during SRV tests. I* e A P was in place to meet plant operating requirements, i and was not related to the SRV test. l 1 l e AP was one of several test conditions that had to be adjusted before determining the correct load j I for the controlling SRV case. Others are reactor pressure, J ]. SRV line temperature, etc. l js l e A P affected the correlation factor relating test and j. 4 design conditions - but SHOULD HAVE NO EFFECT O'l THE I STRESSES AT THE DESIGN CONDITION. I L t e . , . , , - - E . .. r., , .
W TELEDYNE ENGNEERING SERVICES l TYPICAL SRV TEST AND LOAD CASE PARAMETERS I MILLSTONE LINE E . Analysis Test Test Condition A1.1 A1.2 AP 1.08 psi 1.08 psi 1.2 1.5 i Drywell Pressure 15.58 psi 15.58 psi 15.7 41.6 l SRV Line Temp. 186 F 186 F 115 260 Water Volume in 3 SRV Line 12.16 Ft 12.16 Ft 3 12.34 11.S4 j t L f n Factor relating Factor relating 1 test to analysis test conditions (not affected to desian conditions
- by AP) (affected by A P) i I
i I
i-l i TN ENGNEERNG SERVCES-i i SRV TEST - SHELL PRESSURE MAGNITUDES !4 Maximum Test Maximum Pressure
!1 Pressure for Analysis (PSI) (PSI) l Monticello 6.7 N/A ,
k ( Pilgrim 7.8 12.6 4 1 Millstone 7.8 13.0 l Fitzpatrick 4.7 7.8 i f Vermont Yankee 2.5 5.8 l_ Brown's Ferry 6.8 N/A 1 Peach Bottom 9.3 .N/A I Fermi No Test 14.4(1) ) i i J i } ( . l
'[ N/A = Not Available.
! <- II) Maximum Pressure at Fermi was based on Monticello data. { ! f .t i ). l , i' t. 1 { l ,l L -
WTELEDYNE
! ENGNEERING SERVCES i
i S SRV TEST - SHELL PRESSURE MAGNITUDES i e Test pressures on TES plants were comparable to other plants (except for Vermont Yankee). e Data to compare maximum pressures for analysis are not readily available; however, these are calculated directly from test pressures using generic computer codes, so should show similar differences. l I I l 1 , a l
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! SRV* DRAG TEST RESULTS AND ANALYSIS METHOD l a-i > t i s J I a* , .\ e Bickground t and development' of method-x I - 7-l , I e Comparison of cdiculated-to-mea?ured pressures and reasons
- I l for differences- -
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i W TELEDYNE t ENGNEERING SERVICES t SRV DRAG ANALYSIS METHOD - BACKGROUND i I e Predicted SRV shell pressures and load profiles were very conservative - tests showed they exceeded test data by factors of 2 to 5. e SPV shell pressure predictions benefited from extensive in-plant test data. I e SRV drag loads were based on trany of the same assumptions as SRV shell pressure, but without test data for calibration. e Initial comparisons of SRV drag results between worst case predictions and actual stresses showed consistent overpredictions between 500% and 1000%.' e SRV shell pressure overpredictions were discovered early i enough to get NRC agreement for use of test data - SRV drag I problems were not. t 4 6
6 TN ENGeEERING SEFMCES
}
4 S.R.,V , DRAG , ANALJTl_C AL__PR_EDICTIOJi_{ GEN,ERIC) i e Assumes four large bubbles w ~ill fccm in the pool. '
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f c ( . . _ . . . . . . . . . . . . . . _...._..... . - - - . .
/
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..ssumes bubt.le will rise in the ;,.00,1 and converoe on submerc.ed structures sith position and, phasing to.Jmaximize applied-pressure. '
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( ,.# e Assumes a discrete, sinusoidal pressure at th'e- bubble g i perimeter.
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' e Prsdictions t-are heavily dependent on bubble , position and Program methods: use worst case conditions.
i phase. Other, t less severe. conditions produce: lower pr'essures and can dupli-1 cate,testjpressuresineverycase. c \ -
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e Program (TQFOR 02)' attenuates" pressure amplitude over approd-
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-5
-10 0.18 0.24 0.30 0.36 0 0.06 0.12 TIME AFTER 8088LE ENTERS POOL (sect I
i t i a s 1 Figure 5.2.5-1. Sample Predicted Time History of Total X-Forces g on Downcomer 5.2.5-5 Revision : i ( I t
TME ENG#EERING SE3NICES 1 I SRV DRAG - IN-PLANT TEST PROGRAM e A program was begun imediately to collect and evaluate as much data as possible. e Eight tests (4 hot and 4 cold) were run at four different plants. e Response strain was measured for the following structures: 1 Catwalk Vent Supports Column Downcomer Millstone X X X Nine Mile Point X X Vermont Yankee X X Fitzpatrick X X e The calculated natural frequencies of tested structures ranged from 8.1 Hz to 38 Hz. f i 1 l l
W TELEDYNE ' ENGNEERNG SERVCES I J t SRV DRAG - IN-PLANT TEST RESULTS e All 72 sets of data showed consistent results. l t e All measured responses were complex (random) with varying degrees of clean response at the component natural frequency. e There was never an indication of resonant behavior or the presence of any coherent force at the bubble frequency. 1 e Consistent absence of the bubble frequency in all response is proof that the type of load described in the LDR never occurred. e Structures tested had natural frequencies evenly distributed between 8 and 38 Hz. e Bubble frequencies (measured from shell pressure) varied from 5-13 Hz during the tests, t l t
TME ENGEERING SERVICES
! SRV DRAG - POSSIBLE REASONS FOR DIFFERENCES I
BETWEEN PREDICTED AND MEASURED RESULTS 1
- 1. The predicted drag pressure is very dependent on the position and phasing of the bubbles. The bubbles may not have the physical form and coherent phasing characteristics necessary to provide worst-case conditions.
- 2. The program does not realistically attenuate the bubbles as a function of time. This exaggerates the strength of the bubble as it approaches the structures. Application of TQFOR-02 produces practically no attenuation of bubble pressure at pool breakthrough (up to 2 seconds). By contrast, shell pressure decays in about .5 second.
- 3. T-quencher bubbles are long and thin, not spherical. It may be that these bubbles do not maintain their oscillatory behavior as the bubble rises.
- 4. Photographs from the QSTF indicate that the bubbles may start breaking up before they rise in the pool.
- 5. Rising bubbles may pass submerged structures too quickly to excite resonant behavior.
l 1
- 6. Structural response may appear random because of non-linear l l
effects. The moving bubble may change the "added mass" effect i of the surrounding water. i
i TME ENGNEERING SERVICES f SRV DRAG - EFFECTS OF CHANGING BUBBLE FREQUENCIES e Frequency changes are important only to the extent they they influence structural response levels i through resonance. e Review of test data provided strong evidence that SRV drag is not frequency dependent. e Test bubbles had strong frequency components from 5-13 Hz. I e Structures tested included natural frequencies at 8.1, l 8.2, 12.2 and 14.5 Hz - no resonant type behavior was found in any response (most low frequency structures were subsequently removed), e In some cases, cold test frequencies were very near a structural resonance for a particular component and hot test frequencies were not. No difference in response was noted, e Bubble frequencies were not apparent in any response data, i Therefore, the sinusoidal force does not exist at least to 13 Hz. I 1
WNNE ENGNEERING SERVCES i
, SRV DRAG - EFFECTS OF CHANGING BUBBLE FREQUENCY i }
e Test data gives actual measurements over approximately 70% of the bubble frequency range, e In order for a resonance problem to develop. we must assume that there will be a fundamental change in the character of the bubble load in the last 30% of the range. 1 e The entire range in question (13.1-17.8 Hz) is the result of the +40% uncertainty band placed on hot actuations. t
! e The +40% band was based on the fact that a few of the original Monticello tests showed a si_ngle cycle. at an elevated fre-quency. No test ever showed any continuous frequency forces in this 40% band; i.e., in the band where we do not have plant s
unique test data, e Based on the above. we conclude that the plant unique test data is sufficient to answer questions on frequency changes. I
t Technical Report TME TR-5319-1 Revision . . DiGeEstNG SERVICES A4.7 4 i SRV DRAG MARGIN
SUMMARY
. Centrolling Condition SRV Allow.
i Load A1.2 Total Allow SRV Drag i Item Location Type Case Load Load Load Mult. i jDowncomer VH/DC Primary 14 16547 psi 26729 psi 37635 psi 1.83 Intersection Local &
) Bending i
1
) Downcomer Clamp Bending 25 1000 psi 16600 psi 22240 psi 6.44 Tie Bar i
Catwalk Inboard Axial & 25 1712 psi 20157 psi 42000 psi 13.76 i Support Diagonal Bending Brace s Ring Web Membrane & 16 1387 psi 19733 psi 28950 psi 7.65 Girder Bending 4 Ring Node 522 Shear 15 .587 5.916 8.53 5.45 Girder Ib/in Ib/in Ib/in
!I Vent Header Cleiis Shear 25 50 psi 5698 psi 13840 ps1 163.84
.lSuoport P! ate Vent lleader Column Tension F. 14 .094 f/F .595 f/F 1 f/F 5.31 Cuiumn Bending I
i l
l W N NE ENGNEERING SERVICES I i SRV DRAG - EFFECT OF C_ HANGING FREQUENCIES (CONTINUED) e The preceeding table shows the safety margins for all sub-merged structures related to SRV drag load. e The downcomers (V.H. Intersection) have the lowest margin at 1.83. e The natural frequency (1st mode) of the downcomers is 8.6 Pz (based on the large finite element model used in the PUAR). 1 o Fourier analysis of test pressures shows bubble frequencies ranging from approximately 5 to 13 Hz (hot and cold tests). but centered in the 8-9 Hz range for cold pop tests. e Tests provide measured data at the worst bubble / structural frequency for the downcomers. Other bubble frequencies cannot
; induce resonance and cannot be worse conditions; i.e., the I stated safety margin is the minimum value for all frequencies.
N'-U -eeQ a ee O WTELIDrNE ENGMERNGSERVICES TEST 3 COLD P-!S NODE = 2 a
]- 00F. 2 8
sn u 8 d- .~ d2 VERMONT YANKEE c. d-W r Do (D v Cd-e C. d-O N d ' , 12.00 20.00 o.00 4. N a'. 00 1C.00 24.00 28.00 32.00 36.00 4 0. ' FREQUENCY (HZ)
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i TME ENG#EERING SERVICES I
} } SRV DRAG - COMPARISON TO OTHER DATA 1
l COMPARIS0N OF DRAG PRESSURES VSED BY TES WITH PVAR FOR FERMI TES FERMI
- I Downcomer Radial 3.0 2.13 Downcomer Later al 3.0 3.25 f
V.H. Support Column 2.1 .60 l I d
- Tables 3-2, 2-22 and 3-2.2 23 in Fermi PVAR.
l l 4
?
a SeTELEDYNE ENGNEERING SERVCES 1 I i t RESPONSE TO QUESTIONS FROM THE i VERMONT YANKEE ', PUAR REVIEW l l
} \
J t
TM ENGNEERNG SERVICES I J R_ESPONSE TO PUAR Q'_!ESTIO_NS F_0R. VERM_0NT _Y_ANK,EE_ i 1 e A concern has been expressed because the pressures recorded for the Vermont Yankee in-plant SRV test are substantially lower than other plants. 4 Maximum Measured Pressures V.Y - 2.5 psi i Most Other Plants 1 6.7 to 7.8 psi i e The entire process, including data measurement, processing and final analysis, has been checked and no errors have been found. e To assure that Code violations would not occur for small changes in the SRV pressures, an evaluation of safety margins was done for all SRV cases; a table of results follows. 1 l t o The tabulated results show that measured SRV pressures at V.Y. could increase to 8.37 before any Code allowable is exceeded. This exceeds pressures measured and used at similar plants. l,
- . . . . . . - - . . - .. . _.- . .. . .._ . . = _
l 4 YME
.i l
! ENGSEERNGSEMCES !
1 i i j .
- VERMONT YANKEE SRV TEST RESULTS I ALLOWABLE MULTIPLIERS ON SRV i
] i TEST PRESSURES TO MEET CODE ALLOWABLES r t I i Structural Element Multiplier
- Shell Stress 3.35 l l '
- I Support Column 15.26 l Column to Shell Weld 5.24 l
j Column Tie-Down Structure 8.00 l Saddle Stresses 3.48- , 1 I a Saddle to Shell Weld 5.40 i -l . Earthquake Restraint. > 10 ! Attachment Weld > 10 i l 1 Column Anchor Bolts 4.07 i Saddle Anchor Bolts '4.72
\ Ring Girder Web & Flange 16.7:
Weld to Torus Shell -3.94. s
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i YME ENGNEERNG SERVCES
' la. Provide a detailed quantitative description of the procedures used to l develop a design torus pressure signature for any zero drywell-to-wetwell dif ferential pressure (AP) SRV load cases that have been ana-o lyzed. This should include a listing of all input values to the QBVBBS code and the basis for their selection (e.g., if a single discharge line volume was used, does it represent a worst-case choice?) Also, describe the results given by QBVBBS calculation in terms of the pressure ampli-tude and frequency and how these results are adjusted by the calibration e
factor derived from the in-plant tests. i
; lb. Supply the same information for any non-zero 6 P load cases which are I considered to bound the zero AP load cases.
e The general procedure used to process test data was presented at the BNL/NRC/TES/YAEC meeting on July 26,1983 (Item 6 on the agenda) and is included here along with a sample listing of QBVBBS input data. Data presented is for the worst (longest) line, o All test data was processed for use in AP analysis cases. There are no zero A P cases required for analysis. e Calibration for pressure amplitude was done by multiplying QBVBBS results by the factors reported in the PUAR; i.e., Shell pressure = .21 x predicted Support loads = .4 x predicted 1 l e Calibration for frequency was done by adjusting the time axis of the test data. uniformly.
l i W TELEDYNE ENGNEERING SERVICES l
SUMMARY
OF PROCEDURE TO PREPARE IN-PLANT SRV TEST DATA FOR ANALYSIS
- 1. Pressure time history data was obtained from the four SRV tests at Verment Yankee for several locations on the torus shell. Both cold and hot line conditions were tested, but all structural evaluation was based on cold line data.
I 2. The data collected for each of the four tests was best fit to the longitudinal and circumferential distribution (profile) that is pre-dicted by QCUBBS02. This step is required to provide enough data points to drive the structural model.
- 3. The maximum best fit curve (for the four cold tests) was used to form a calibration f actor between measured results and results predicted by l
QBVBBS02 for the test condition. This factor bounded all cold tests and all hot tests f or the plant. It was used to adjust all predictions from QBVBBS for all other conditions. I 4. Pressure waveform and frequency were also based on the maximum best fit case; the specific data was taken from the maximum response point in this test, which was bottom center of the shell.
- 5. The actual SRV analysis was done using a load that combined maximum shell pressure and maximum frequency into one bounding base. Case A1.2 pro-vides the highest shell pressures; case C3.2 provides the highest fre-quency.
- 6. The C3.2 frequency, as calculated by QBVBBS, was increased by 40 percent to account for possible frequency shif ts, l
i l k l
i ; WTELEDGE l ENGINEERING SERVICES l I j INPUT LISTING FOR VERMONT YANKEE DESIGN CASE i 1 YANKEE LINE C CASE Al-2 IBA DATA AN=2.0,BN=4.0 BN2=4.0 RI=0.0163,EL1=2.3,EL2=3.74,EL3=5.05.EL4=7.83,EL5=7.83 HAl=0.075,HA2=0.093.HA3=0.147.HA4=0.317,HA5=0.0,
. AARM=0.706,XLQ=9.39, i EH=10.87,EHC=5.,RMAJ=49.0.AP=730.,
l P l =42.1.T I =260. ,WL= 3.6, V I = 42.60, EMA= 5.51, P0 =311.9, TWC = .12546. , PWW=40.3.TR=143. . VG=115. , JM420,FFAC=1.0,BFAC=1.0, PF AC =1.65,JPLOT=1, ALPHA ( 1 ) =0. ,LPO INT =1 l BETA (1)=0.0.0.175.0.349,0.524.0.698,0.8731.047,1.222.1.396.KP0lNT=9,(SEND) I LAST i
! l QBUBS = Input Terms AN = Number of Quencher Arms I BN = No. of Bubbles BN2 = No. of Bubbles af ter Coalescense q
l R1 to AARM = Quencher Geometry Factors 3 EH to AP = Pool and Torus Geometry Factors PI = Initial Pressure in SRV Line (psia) ' TI = Initial Temp. in SRV Line (F) , l WL = Water Leg (ft) 3 ! VI = Initial Gas Vol. in Line (f t ) , EMA = Initial Air. Mass in Line (1bm)' i P0 = Max. Pressure at Air-Water Interf ace (psia) (from RVFOR) TWC = Time 'of Water Clearing (sec) (from RVFOR) f
- PWW = Wetwell Pressure (psia) f TP = Pool Temperature (UF )
I VG = Specific' Volume of Saturated Steam at TP (f t3 /lbm) ! Ji = No. Points to be Plotted , FFAC = Frequency Factor (to stretch or compress waveform) 4 BFAC = Bubble Pressure Factor j . PiAC = Shell Pressure Factor .i ! . ALPHA, BETA = Define Locations of Which Pressure Calc. ' _ _ , , , , , , - , , - - - , , - - - - . - , - - - , . - -- .--w- ,
' I I
i I I
- 2. Identify the load cases above using the notation of Section 4.3 of NUREG-0661 (Event Combination Numbers). Then, describe how the bounding of one load case by another is established.
e Bounding method considered two important factors, e For shells, level A & B allowables are the same, o Earthquake loads are small, so we don't use the higher stress allowables which are listed, e Most evaluations reduced to the same 5 or 6 bounding cases shown in the figure; however, each structure was considered separately and some had other bounding cases. l 1
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- 3. Outline the steps used in the calculation of the SRV discharge line clearing loads and the thrust loads on the T-Quencher arms. In parti-cular, provide the values of the various parameters listed in 5.2.1.3 of the LOR which influence SRV discharge line clearing loads and which are used as input of the SRVDL line clearing analytical model. Justify these values on the basis of providing worst case loads f or V.Y.
e All methods and procedures to calculate SRV clearing loads and thrust loads on the arms were as described in the LDR. e RVRIZ and RVFOR conputer codes were used for all computations i for Vermont Yankee as well as for Millstone (Question $16). e Water clearing thrust and thrust loads on the arms were calcu-lated for the worst (longest) line and these values were applied to all other lines, e Maximum water clearing thrust is associated with second actua-tion case C3.3. Second actuation was assumed to occur at the point of maximum linc reflood (worst case). e Gas clearing loads are calculated for each line for controll-ing case A1.2. e Maximum shell pressures are associated with case C3.2 which is calculated for the worst (longest) line.
1 l
, i StTELEDYNE ENGNEERING SERVCES l t
I Item 3 (Continued) Following are the input parameters f or the worst V.Y. line for case C3.3. Worst line was determined by running the longest and shortest lines for reflood and discharge. These are listed as they appear in para. 5.2.1.3 of the LDR. I a. Initial water leg - 14.09 ft. l
. b. Air volume - 39.89 f t.3 l
- c. Line dian eter - .797 f t. to .835 f t. (varies)
- d. Steam flow rate 299 i
I U e, initial line temperature 350 F (143 F in torus)
- f. Line configuration and hydrodynamic losses will be provided sep-arately, if required.
9 SRV main disc stroke time - 02 sec. 1 1 A t.__._ _ .
i 4 I i 1 I 4 4 j 4a. The stress at the VH/DC' intersection due to SRV drag loads calculated by , VY PUAR methods was previously given as 16547 psi. What would this stress be if drag loads were calculated according to AC approved methods? , How do maximun stresses due to SRV drag on other submerged structures 3 compare when drag loads are calculated by AC approved methods? i e l i j 4b. Which data points in Figure Al-5 of the VY PUAR correspond to downcomers I and what are the natural frequencies of these downconers? i i j e - 4a covered by opening discussion. 4 ! e 4b also covered by opening discussion. .!l Pressure on VY downcomers was taken at 3 psi. ll l - Natural frequency calculated at 8.6 Hz.
- I i
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- 5. Provide a more precise and quantified basis for the SRV submerged struc-ture drag methodology outlined in Appendix 1 of the VY PUAR. Justify in a rigorous manner the elimination of frequency effects from the calcu-lations.
? e e Covered in opening discussion. I I I I
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1 N 3 Al - SRV 1 and mitigattore via AP control' is a new approach that has not been , t s j evaluated or 4pproved by tV: staff. ) A2 - 10 perform such an evaluation, the staf f will require detailed presen-. tation of all pertinent data test data. An acceptable format for the presentation of such t)ata is submittal of a test report patterned af ter l others submitted to the staf f (e.g.. GE's Monticello T-quencher Test Report, NEDE-21864-F, 'GE's Caorso Cross ' Quencher Test Report - NEDE- { , ! 23100 D or NYTEC!i't Kupshang Cross Quencher Test Report).
, i
)+ - A3 Since thq Q'.'BBS methodology will be used to extrapolate to design condi-f, tions, we also require presentation of the relevant Q'.'BBS trends with { ,
, plant parameters, initial condjttons, etc. with AP in place. The sens-itivity of these trends"wi'.h variations inAP should also be supplied.
1 I . . A4 - A clear description of the load cases hsed for design should be supplied. s Each load c33e shou'd be character' ired by' stating the value of all variabi.:s whict! detennine torus shell loads. Thiy is especially impor-tant for any load cases involving a subsequent actuation. , t
\ - -
s3 t , e Tyegeitemshavebeencoveredinthegeeningpresentation.
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- j' ENGBEERNG SEMCES j t A5 - Plant design should also be capable of withstanding the SRV loads that
) . would be experienced with zero AP in the same sense that the staff
- requires structures to withstand DBA loads with zero AP (see Note (1)
} for event combination number 14 in Section 4.3.3 of NVREG-0661). The I applicants will be required to perform such an additional evaluation. I Use of the Monticello plant design torus pressures would be acceptable for this purpose i e The opening presentation illustrated that there are no Mark 1 3 load cases at O'AP (for a plant normally operating at AP). e Normal SRV actuation without AP in operation is bounded by normal actuations before AP was installed (with ramshead dis-charge). Based on this, OdP SRV actuation'is assumed accept-able, without further analysis.
; I I
I i s 1, .
Appendix 5.2
"# TELEDYNE l
ENGINEERING SERVICES o m . . . . . u~ n. , t " @ 6 l February 28, 1984 l 5960-10 U.S. huclear Regulatory (,ognission j 7900 Nor'd Avenue ) Bethesda, MD 20814 Attn: B. Siegel, Program Manager Room 416, Old Phillips Builtiing I
Subject:
Fark 1 Containment Prograst-PUAR Review for Vermont Yankee and M1)lstone Gertlemen: The following are responses to the three open items from tre: NRC/BNL/ YAEC/NUSC0/TES PUAR review meeting c; february 16, 1984. l I
- 1. Use of Previous Operating Experience to De Onstrate the Accept-I ability of 0 A P, SRV Actuation At the referenced meeting, TES' position was that previcus operat-t ing experience with zero A P SRV actuations was sufficient to eliminate the need for a specific analysis at the 0 A P cmdition.
This position is based on the fact that both Millstone and Vermont Yankee's normal operating condition was at .:ero 6 P for several years, beginning with init 441 plant operation & to 1976. During this period, the SRV lines discharged through ramsheads; dozens of actuations were experienced at both plants, at zerc A P, without adverse effects. Both plar.ts have since been modified so the SRV lines discharge through tee-quenchers, rather than tae previous i amsheads. Instal-
- h. tion 0 # the tee-quenchers has reduced pressures and ioads substantially ' rom those already expcrienced with the ramsheads.
} Tests run at the Monticello plant for both ramsheads and tee-mnchers shows that tee-quenchers reduced shell pressure by roximately 68%. Due to similarities in the plants and the dis-1 03rge devices, we believe these results are generally applicable l to both Vermont Yankee and Millstone.
Based on the above, we believe that 0 AP SRV actuations at Vermont Yankee ed Millstone will produce loads and stresses that are well within the capacity of the plant struttures and that no further analys >s is necessary to confirm this position. l
- 2. Additional Inf,ormation to Evaluate the Effects of A P on Test Results The inforfration requested is included in Tatle 1, attached to this letter.
, ENGINEERS AND METALLURGISTS l
N clear Regulatory Commission
'PPTELEDYNE February 28, 1984 ENGINEERING SERVICES Page Two
- 3. Additional Information on Basis of .4 Calibration Factor for Column Loads The PUAR's for Vermont Yankee and Millstone discuss the use of calibration factors, developed from SRV test data and applied to analysis models. The PUAR's identify two calibration factors that were used; one for shell stress and one for column (support) loads.
The calibration factor for columns is .4, which is a multiplier applied directly to the calculated column (support) loads. This same factor was used for all plants analyzed by TES and was the bounding (smallest) calibration adjustment resulting from four sets of in-plant tests. Results for the four sets of tests follow: Ratio of sis Data from In-plant Tests Up Loads Down Loads Cold Tests Hot Tests Cold Tests Hot Tests Millstone Test #1 .37 .26 .4 .12 (Oct.79) Millstone Test #2 .35 .20 .31 .14 (June 81) Nine Mile Point .38 .29 .26 .24 Pilgrim .27 .1 .19 .1 Average Correction .33 .21 .29 .15 Factor Factor for - -
.4 -
Minimum Amount of Correction The table illustrates that the .4 calibration factor represents the least amount of calibration that was measured in any of the tests and that up and down loads as'well as hot and cold actuations were bounded. This single generic calibration was used for all TES plants beera addition of support saddles made later plant unique instrumentattur, impractical. The tabulated data was assembled before support sad-dies were installed.
'PPTELEDYNE ENGNEERING SERVICES U.S. Nuclear Regulatory domission
- 5960-10 February 28, 1984 Page Three The question of how the addition of saddles might affect calibra-tion was answered by the second Millstone test. In this test, the ;
saddles were complete except that they were not yet shimed and attached to the basemat. In this form, they provided nearly their full effect on structural modes and response, but still allowed instrumentation of the original support columns. The table shows that addition of the saddles reduced support loads in all cases except one hot case, which was very lightly loaded. -Based on this data, we conclude that addition of the saddles makes use of the .4 factor conservative. Calibration for shell stress was done on a plant unique basis based on measured shell response from that plant. These numbers are reported in the PUAR's. Sincerely yours, TELEDYNE ENGINEERING SERVICES Nicholas S. Celia Manager, Engineering Projects NSC:sig cc: R. White (YAEC) R. Smart (NUSCO) M. Francischina (NUSCO) DFL (TES)
1 U.S. Nuclear Regulatory Comission 5960-10 '#PTELEDYNE * - February 28, 1984 ENGNEERING SERVICES Page Four TABLE'1 SRV TEST DATA FOR V.Y. AND MILLSTONE VERMONT YANKEE - PARCH 14, 1981 1 2 3 4 Initial Reactor Power (%) 42.4 41.2 42.8 45 Steam Press (PSIG) 957 958 960 961 Torus Water Temp. ( F) 69.8 65.2 63.0 59.8 Drywell Press. (PSI) 16.5 16.59 16.54 16.52 Torus Press. (PSI) 14.65 14.67 14.67 14.66 . P (PSID) 1.85 1.92 1.87 1.86 Line Temp. @ Nozzle (UF ) 199 202 207 217 Max. Shell Press. (psi) 2.3 N/A 2.5 2.1 Min. Shell Press. (psi) -2.5 N/A -2.1 -2.1 _ Initial Water Leg (f t)* 4.12 3.79 4.02 4.07 MILLSTONE - JUNE 17, 1981 Initial Reactor Power (%) 37 38 37 33 Steam Press. (PSIG) 1004 1002 1000 1001 Torus Water Temp. (UF ) 69 70 70 72 DrywellPress.(PSIA) 15.6 15.7 15.7 15.78 Torus Press. (PSIA) 14.45 14.55 14.55 14.6 P (PSID) 'k.15 1.15 1.15 1.18 = Line Temp. 0 Nozzle ( F) 145 152 162 186 Max. Shell Press. (psig) 6.5 7.3 7.8 4.5 Min. Shell Press. (psig) -5.1 -5.4 -5.1 -3.8 i Initial Water Leg (ft)* 6.67 6.67 6.67 6.60 ' N/A - Not readily available, can be supplied if required. Pressures less than test 3.
- Length of water slug, measured from center line of tee-quencher arms.
Appendix 5.3 W TELEDYNE ENGNEERING SERVICES VERMONT YANKEE MARK 1 PLANT UNIQUE ANALYSIS REVIEW TELEPHONE CONFERENCE CALLS ON mRCH 9 and MARCH 12, 1984 : Following the February 16 meeting and presentation, there were two con- a ference calls to resolve final questions. A summary of these calls follows, as they affect the Vermont Yankee Plant (these calls also covered some plant-unique Millstone questions that are not sunrnarized here). The March 9 call - included NRC, YAEC, BNL, and TES personnel. The March 12 call was between ENL - and TES. . Question #1 - How was the SRV calibration f actor developed for the V.Y. support columns? Documentation describing the method of determining the calibration fac- 2 tors was listed, as follows: 3 i
- a. PVAR, Appendix 1, page Al-2 outlines the general method of calcu-lating and applying all calibration f actors. --
- b. Item 6 from NRC/BNL/TES meeting on July 26, 1983. Documented in PVAR, Rev. 2, Appendix 4.6, page 4.6-26. This gives a more detailed description than Item a and outlines the method used for column 4 calibration.
=
- c. February 16, 1984 meeting, Item VY-1A, gives further detail and an input listing of a typical QBUBBS computer run,
- d. TES letter of February 28,1984, Item 3, discusses column multiplier in detail.
Af ter this summary, there was a specific question from BNL regarding the exact figure used to calibrate the column loads for V.Y. TES' response to this question was covered in the March 12 call. The calibration f actor was .4 _ (the smallest column factor) X .21 (the QBUBBS factor). This provides' an ' overall factor of .084. 5 m.
-i imensi - -
W TELEDYNE ENGINEERING SERVICES Question #2 - Provide further information on the " windowing" technique used to reduce SRV test data. (Ref. Item #15 at the Feb. 16 meeting). The use of a time window for measurement of maximum pressure peaks was adopted to assure that amplitude readings in regions of steep pressure grad-ients would not produce unconservative results. Measurement of the maximum pressure in a specified time window for all pressure readings assures that a non-conservative set of pressures will not be applied to the shell for the stress analysis. There was agreement that this was correct and conservative for shell stress; but, that it might not be conservative for the columns. The concern was that the conservatism in column loads due to windowing is backed out when we correct by the .4 f actor discussed in Question #1, above. If a plant has an .. actual correction less than .4 (.4 is a bounding generic number), then column loads can be over-corrected and unconservative. The response to this concern was made on the March 12 call and included the following points:
- a. The .4 calibration f actor is not entirely due to windowing. Another major contributor is the f act that we used QBUBBS to define the overall pressure profiles on the shell, (with overall amplitude scaling based on test data). We know this procedure is conservative and judge that it contributes nearly as much to the .4 f actor as windowing. This reduces the effect of any unconservatism that may exist due to windowing.
- b. The .4 calibration factor was measured for a plant without support saddles. We know that a greater calibration f actor can be justified for a plant with saddles. Data indicates that a 22% conservatism exists in the .4 f actor because of this; this offsets any equivalent non-conservatism that may exist due to windowing. (Ref. Feb. 28 letter, Item 3).
WNNE ENGNEERNG SERVICES
- c. The February 28 letter lists the test data collected for support columns. The cold test, down loads show a range of about 2 (.4 max., .19 min.); this represents the maximum possible non-conservatism. If we reduce this by 50% of the difference to account for the use of a QBVBBS pressure distribution and an additional 22%
for the addition of saddles, the maximum factor becomes 1.39, or a maximum possible 39% non-conservatism,
- d. The February 16 presentation included an allowable SRV multiplier which could be applied to calculated results and still satisfy code allowables. The smallest allowable multiplier for the support sys-tem is 3.48 (an allowable 248% increase) for saddle stress. By comparing this to the maximum estimated 39% non-conservatism above, it is clear that no code violations can occur.
Response to this question was based on an assumption that a worst case situation could occur at V.Y. There is no indication that it actually did occur. Question #3 - What is the water leg in the SRV line at minimum P (1.7 psi)? The water leg is 4.81 feet, measured from the center of the quencher arms. The balance of the discussion in these telephone calls did not apply to the Vermont Yankee plant. I _ _ . .}}