ML20064B892
| ML20064B892 | |
| Person / Time | |
|---|---|
| Site: | Pilgrim |
| Issue date: | 10/27/1982 |
| From: | TELEDYNE ENGINEERING SERVICES |
| To: | |
| Shared Package | |
| ML20064B891 | List: |
| References | |
| CON-#487-5015 2.206, TR-5310-1, NUDOCS 8301040239 | |
| Download: ML20064B892 (144) | |
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TECHNICAL REPORT O
i TECIINICAL REPORT TR-5310-1 l M ARK I CONTAINMENT PROGRAM
- O PLANT-UNIQUE ANALYSIS REPORT l OF TIIE l TORUS SUPPRESSION CIIAMBER I FOR
'O PILGRIM STATION - UNIT 1 l OCTOBER 27, 1982 l
=O 8301040239 e21230 PDR ADOCh 05000293 P PDR
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BOSTON EDISON COMPANY O 800 B0YLSTON STREET BOSTON, MA 02199 i
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l0 TECHNICAL REPORT TR-5310-1 1
i MARK 1 CONTAINMENT PROGRAM
- 8 i
l PLANT-UNIQUE ANALYSIS REPORT OF THE TORUS SUPPRESSION CHAMBER
@ FOR PILGRIM STATION - UNIT 1
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lO l OCTOBER 27, 1982 i
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- G WTELEDYNE ENGINEERING SERVICES 130 SECOND AVENUE WAllHAM, MASSACHUSETTS 022M i t317-800-3350
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O Technical Report WM TR-5310-1 -ii-ENGeEstNG SERVICES 4
1 O
l ABSTRACT l
i 1
The work summarized in this report was undertaken as a part of the Mark 1 O
i 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 j modifications that were made to the structure and internals to increase safety margins.
O In all cases, the stresses reported in this document meet the allowable levels as defined in the structural acceptance criteria (Reference 3). The g methods and assumptions used in this analysis are in accordance with USNRC NUREG 0661 (Reference 2), except as noted in the text. The modifications described in this report are also in compliance with NUREG 0661, unless i
1 otherwise noted.
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Technical Report W F W NE TR-5310-1 -iii- ENGINEERING SERVICES TABLE OF CONTENTS
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Page ABSTRACT ii D
1.0 INTRODUCTION
AND GENERAL INFORMATION 1 2.0 PLANT DESCRIPTION 4 3
2.1 General Description 4 2.2 Recent Modifications 4 2.2.1 Modifications to Reduce Loads 5 D
2.2.2 Modifications to Strengthen Structure 8 3.0 CONTAINMENT STRUCTURE ANALYSIS - 34 SHELL & EXTERNAL SUPPORT SYSTEM 3.1 Computer Modeld 34 l 3.2 Load Analysis 35
! 3.2.1 Pool Swell Loads 35 D 3.2.2 Condensation Oscillation - DBA 36 3.2.3 Chugging 37 3.2.3.1 Pre-Chug & IBA C0 37 3.2.3.2 Post Chug 37 0 3.2.4 SRV Discharge 38 3.2.5 Deadweight, Thermal & Pressure 39 3.2.6 Seismic 39
~3.2.7 Fatigue Analysis 40 0
3.3 Results and Evaluatien 41 3.3.1 Shell 41 3.3.2 Support Columns % Attachments 43 3.3.3 Support Saddles & Shell Weld 44 0 3.3.4 Earthquake Restraints & Attachments 45 3.3.5 Anchor (Tie-Down) System 46 O
Technical Report 'R TF1FrWNE TR-5310-1 -iv- ENGINEERING SERVICES TABLE OF CONTENTS (CONTINUED)
Page 4.0 VENT llEADER SYSTEM 58 4.1 Structural Elements Considered 58 4.2 Computer Models 58 i 4.3 Loads Analysis 60 l 4.3.1 Pool Swell Loads 60 I 4.3.1.1 Pool Swell Water Impact 60 4.3.1.2 Pool Swell Thrust 61 l I
4.3.1.3 Pool Swell Drag (Support Columns Only) 61 i i
4.3.2 Chugging Loads 61 4.3.2.] Downcomer Lateral Loads 61 4.3.2.2 Synchronized Lateral Loads 62 4.3.2.3 Internal Pressure 62 4.3.2.4 Submerged Drag 63 4.3.3 Condensation Oscillation - DBA 63 4.3.3.1 Downcomer Dynamic Load 63 4.3.3.2 Vent System Loads 64 4.3.3.3 Thrust Forces 65 4.3.3.4 Submerged Drag (Support Columns) 65 4.3.4 Condensation oscillation - IBA 65 4.3.5 SRV Loads 65 4.3.5.1 SRV Drag Loads 65 4.3.b Other Loads - Weight, Seismic & Thermal 66 4.4 Results and Evaluation 66 4.4.1 Vent Header-Downcomer Intersection 66 4.4.2 Vent Header-Main Vent Intersection 67 4.4.3 Vent lleader Suoport Columns & Attachments 67 4.4.4 Downtomer Tie Cars & Attachments 68 4.4.5 Vent Header Deflector & Attachments 69 4.4.6 Main Vent /Drywell Intersection 69 l 4.4./ Vent lieader, liain Vent & Downcomers - 70 Free Shell Stresses l 4 . 4 . 11 Vent lieader - Mitre Joint 70 I
4.4.9 Fatigue Evaluation 70 f
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Technical Report TE DE TR-5310-1 -v-ENGNEERNG SERVICES lAl!LL 01 LOflILNIS (CONilfiULIQ Page 5.0 RING GIRDER ANALYSIS 78
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i 5.1 Structural Elements Considered 78 5.2 Computer Models 78 I
5.3 Loads Analysis 79 5.3.1 Loads Applied to the Shell 79 i l
5.3.2 Drag Loads 80 5.3.3 Loads due to Attached Structure 81 5.4 Results & Evaluation 82
, 5.4.1 Ring Girder Web & Flange 82 5.4.2 Attachment Weld to Shell 82 6.0 TEE-QUENCHER & SUPPORT 88 l
l 6.1 Structural Elements Considered 88 6.2 Computer Models 89 6.3 Loads Analysis 89 6.3.1 SRV Loads 6.3.2 Pool Swell Loads 89 6.3.3 Chugging Loads 90 6.3.4 Condensation Oscillation Loads 90 6.3.5 Other Loads 91 6.4 Results & Evaluation 91 6.4.1 Tee-Quencher Structure 91 6.4.P Submerged SRV Line 91 6.4.3 Tee-Quencher Support 92 6.4.4 Tee-Quencher Support Brace 92 7.0 GTHER STRULTURES 95 7.1 Catwalk 95 7.1.1 Computer Models 95 l
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} Technical Report "RTF1 FfWNE TR-s310-1 -vi- ENGINEERING SERVCES TABLE OF CONTENTS (CONTINUED)
Page 7.1.2 Loads Analysis 95 7.1.2.1 Pool Swell Water Impact 95 7.1.2.2 Pool Swcll Fallback 96
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7.1.2.3 Froth Loads 96 7.1.2.4 Drag Loads (Support Columns) 96 7.1.2.5 Weight & Seismic 97 7.1.3 Results & Evaluation 97
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7.1.3.1 Main Frame 97 7.1.3.2 Support Columns & End Joints 98 7.1.3.3 Welds to Ring Girder 98 7.7 Internal Spray Header 99 7.2.1 Computer Model 99 7.2.2 Loads Analysis 99 7.2.2.1 Froth Load 99 7.2.2.2 Weight, Seismic & Ring Girder Motion 99 7.2.3 Results and Evaluation 100 7.3 Vent Pipe Bel!aws 101 7.3.1 Analysis B.ethod 101 7.3.2 Loads Considered 102 7.3.3 Results & Evaluation 102 7.4 Elictrical Penetration Boxes 102 l
l /.b Wetaell Vacuum creaker Mount 103 l
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.n :
',UPPRF%IUN P0QL TENPERATURE EVALUATION 108
- m. I N aimum Suppression Pool Temperature 108
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Pool temperature Monitoring System 109 af Lear.Ls 112
- ,i : ( 1 - USE OF SRV IEST EGTA IN ANALYSIS Al-1
/ G NDIi '
- USE OF 32 HZ CUTOFF FOR C.O. & POST CHUG ANALYSIS A2-1
/ iYl NDI X c; - ORAG VOLUMES FOR SUC"ERGED STRUCTURE ANALYSIS A3-1
l Technical Report "RTF1 FrWNE -
I TR-5310-1 -vii- ENGINEERING SERVICES FIGURES AND TABLES
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Page Figures:
2-1 Torus Plan View 12
) 2-2 Torus Composite Cross Section 13 2-3 Torus Modifications - Cross Section at Ring Girder 15 2-4 Torus Modifications - Cross Section at Mid-Bay 16 2-5 AP Pressurization System 17 2-6 Vent Header Deflector 18 2-7 Vent Header Deflector Attachment 19 2-8 SRV Tee-Quencher and Support 20 2-9 Pool Temperature Monitoring System 21 2-10 PHR Retro Line Elbow anu Support 22 2-11 Torus Support Saddles and Saddle Anchors 23 2-12 Torus Support Column and Anchors 24 2-13 Downcomer Tie Rod and Gusset Modification 25 2-14 Catwalk and Handrail Modification 26
?-15 Catwalk and Handrail Modification 27 2-16 Torus Spray Header Support Modifications 28
?-17 Thermowell Detail 29 2-18 Vacuum Breaker - Vent Intersection Modification 30 l
l 2-19 Electrical Junction Box Modification 31 t
2-20 Thermowell Locations 32
?-21 Ring Girder Weld Modification 33 3-1 Detailed Tarus Shell Model 48 l 3-? Detailed lorus Shell Model 49 3-3 Detailed Torus Shell Model 50 I
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I I . 7PTF1 pr?(NE 1 1 -yiii- ENGINEERING SERVICES FIGURES AND TABLES (CONTINUED)
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Page 3-4 Torus Beam Model (360 ) 51
> 3-5 Pool Swell - Net Vertical Load 52 3-6 Pool Swell - Average Submerged Pressure 53 3-7 Pool Swell - Torus Air Pressure 54 3 3-8 SRV Shell Pressure - Typical 55 3-9 Location of Maximum Shell Stress 56 3-10 Earthquake Restraint System 57 4-1 Detailed Vent Header Model 72 4-2 Detailed Vent Header'lodel 73 4-3 Detailed Vent Header Model 74 4-4 Vent Header Beam Model 75 4-5 Vent Header Deflector Analysis 76 4-6 Chugging Cases - Synchronized Lateral Loads 77 S-1 Ring Girder 84 5-2 Detailed Shell - Ring Girder Model G5 5-3 Detailed Shell - Ring Girder Model 86 S-4 Detailed Shell - Ring Girder Model 87 6-1 SRV Line Analytical Model 94
/-l Catwalk Computer Model (Unmodified) 104
/-2 Catwalk Computer Model (Modified) 105 7-3 Spray Header Computer Model 106 7-4 Vent Pipe Bellows Motion 107 S-1 NRC Specified Local Pool Temperature Limits 110 Al-1 SRV Test Instrumentation - Shell Al-6
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Technical Repcrt "W TEL M NE TR-5310-1 -ix-ENGINEERING SERVICES FIGURES AND TABLES (CONTINUED)
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Page Al-2 SRV Test Instrumentation - Columns Al-7 Al-3 SRV Test Instrumentation - Tee-Quencher Al-8 I
Al-4 SRV Test Instrumentation - Attached Piping Al-9 Al-5 SRV Test Instrumentation - Typical Internal Structures Al-10 Al-6 SRV Test Instrumentation - Typical Downcomers Al-ll Al-7 SRV Drag Pressures Al-12 Tables:
- 1. Structural Acceptance Criteria for 114 Class MC and Internal Structures
- 2. Plant Physical Dimensions 115
- 3. Plant Analysis Information 116 8-1 Summary of Pilgrim Pool Temperature Response 111 l
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Technical Report TR-5310-1 WF WNE ENGmEERING SERVICES
> 1.0 GENERAL INFORMATION The purpose of the Mark 1 Torus Program is to evaluate the effects of ;
hydrodynamic loads resulting from a loss of coolant accident and/or an SRV I
) discharge to the torus structure. This report summarizes the results of l extensive analysis on the Pilgrim torus structure and reports safety margins against established criteria. The content of this report deals with the torus 1 shell, external support system, vent header system and internal structures.
) Analysis and results for piping attached to the torus (including shell pene-trations and internal piping), for the SRV line (except the submerged portion and Tee-Quencher), and for the SRV line vent pipe penetration will be pre-sented in a separate piping report, TR-5310-E.
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 phenom-ena, 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 (References 9, Volume 1-10 are examples of this). In most cases, the loads used to support the analysis were calculated in strict accordance with those procedures, as amended by NUREG 0661 (Reference 2). In some cases, optional methods have been used; these methods are specifically referenced in Program documentation. Examples of these are the use of plant-l
Technical Report YE WE TR-5310-1 N N N ES
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unique SRV test data to calibrate SRV analysis, and use of plant unique quarter scale pool swell movies to refine certain water impact and froth loads. In a few cases, analysis assumptions have been made that do not appear in Program documentation; these are identified in the text.
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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 was also done. The computer code
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used to perform almost all of this analysis was the STARDYNE code, as marketed by Control Data Corporation. STARDYNE is a fully qualified 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 identified 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-i 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 and 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 smaller number of cases by conservative bounding. For example, load combinations including SSE seismic, have a higher allowable than the same combination with OBE seismic. For these cases, our first evaluation attempt is to consider the SSE combination against the OBE allowables. If this produces an acceptable result, those numbers are reported as final. This procedure results in many cases where safety margins are understated; this is the case for most of the results.
l Technical Report YN NE TR-5310-1 N SBi\/ ICES
) 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-once 1). This has been done by identifying the applicable LDR paragraph in
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parenthesis immediately following the title of the load. This referencing directs the reader to a more detailed description of the load than can be included in this report.
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Technical Report WP WE TR-5310-1 6M
) 2.0 PLANT DESCRIPTION 2.1 General Arrangement
) The configuration of the Pilgrim torus structure is shown in Figures 2-1 and 2-2.
Figure 2-1 shows a p'.e.a 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 saddles and columns are connected to the basemat with eight anchor bolts per saddle. Four earth-
) quake 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 i
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. Openings where the vent pipes penetrate the torus shell are sealed by bellows. The ring girder also i
supports the catwalk, spray header and SRV tee-quenchers. Figures 2-3 to 2-21 show several details of the torus structure. Table 2.0 lists several of the plant specific dimensions.
2.2 Recent Modifications Over the period of the past few years, many modifications have been made to the Pilgrim torus, both to increase its strength and also to mitigate the hycirodynamic 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:
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Technical Report TE @E TR-5310-1 ENGNEERNG SERVICES l l
2.2.1 Modifications to Reduce Hydrodynamic Loads
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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 I
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system is illustrated in Figure 2-5. It is designed to maintain a minimum positive pressure difference of 1.17 psi between the vent system (drywell) and the airspace inside the torus. The result of this pressure difference ( AP) !
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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 1.17 psi pressure difference
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was selected as the basis for the Pilgrim plant unique quarter-scale pool swell tests and is intended to be the normal operating condition of the plant. l The system complies with NUREG 0661, which requires two narrow range instru- )
ment channels with less than 1 0.1 psid error. Pilgrim Station uses a narrow
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range instrument for drywell-to-wetwell AP and backup instruments to measure drywell and wetwell pressures separately. The maximum error for the above l instruments is 1 08 psid.
Downcomer Shortening l Downcomer shortening is an additional means of reducing the water slug that must be cleared from the downcomers during rapid drywell pressurization. Its primary advantage is to reduce pool swell loads during those periods of time when the A P system may be inoperative. It alsa allows for a very small water leg during normal operation, without the problems associated with higher drywell pressures. The downcomers at Pilgrim were cut
- to provide a minimum submergence of 3.0 feet (distance from bottom of down-comer to minimum torus water level). l Vent Header Deflector The vent header deflector at Pilgrim is illustrated in Fig-ores 2-6 and 2-7. It is a 16-inch schedule 120 pipe with h-inch plate welded to the sides.
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Technical Report YM TR-5310-1 6M
) The deflector extends under the belly of the vent header to protect the vent header from direct water impact during pool swell. It shields the most sensitive part of the vent header 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 Pilgrim to provide accurate vent header loadir.g for detailed analysis.
SRV Tee-Quencher
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A tee-quencher has been installed at the discharge end of each main steam relief line to replace the existing ramsheads. A typical quencher and its support are 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 f actors of two or more when compared to ramshead pressures. The plant-unique
) characteristics of these devices at Pilgrim were determined by in-plant test-ing after their installation. End cap holes in the tee-quencher were also provided to promote suppression pool mixing and minimize temperature gradients.
1 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.
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_ Temperature Monitoring System & RHR Return Lines 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, 2-17 and 2-20.
Tne temperature monitoring system senses pool temperature through a set of 26 sensors set in thermowells around the torus shell (Figures
?-17 and 2-20). Of these 26,12 are mounted in redundant pairs (6 locations)
Technical Report TN TR-5310-1 6M
) to sense local pool temperature *. The remaining 14 are mounted in seven redundant pairs and are used along with the local sensors to determine bulk )
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pool temperature. The system is hard wired to a console' in the control room. 'l/ '
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) The elbows on the RHR return lines were added to improve pool circulation during periods of extended SRV blcndown. Lirculation of the pocl "
with these lines asseras 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 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 (see Figure 2-10).
)
Additional SRV Vacuum Breakers Each of the four SRV discharge lines at Pilgrim has been
> fitted with two, ten-inch vacuum breakers, in addition to the original one-inch vacuum breakers. This modification minimizes the temporary formation of r the high water leg in the SRV line which could occur due to steam condensation oscillation after an initial actuation; and thereby prevents the hign cldar-- -
) ing loads which could occur if a second actuation occurred at that t{me.' The location of these devices is different on each SRV line due te, space con-straints, and is not illustrated. ,
Removal of Submerged Piping Some of the piping inside the torus extended to greateh ,
depths than was necessary for its proper functioning. This additional ruirer-gence resulted in drag loads on the piping that was unnecessary. In arde; tc., /
reduce the loads, these piping systems were cut off to provide a three-foot' , ,
- The si flocal paTrs are locatod in the six SRV tee-quencher bays. Only four j of these quenchers are operative; the other two are not cannected to the steam T relief lines and have been installed for possible future use. ',,
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Technical Report WTF1 Fr?(E TR-5310-1 ENGNEERNG SERVICES 5 submergence at minimum torus water level. The five lines affected were the HPCI exhaust, RCIC exhaust, HPCI drain, RCIC drain and a 18-inch diameter i
spare line.
) In addition, the eight vent drain lines (one-inch diameter) were cut off and capped above the pool. In the future, any condensation in the vent system will be manually drained, as required.
4 Removal of Monorail The monorail originally installed al Pilgrim was a low capa-city unit of relatively light design. Modification of the unit to withstand worst-case froth loads was Judged impractical. The monorail was removed, and not replaced.
2.2.2 Modifications to Strengthen the Structure i
Torus Support Saddles and Anchor Bolts Support saddles were added under each ring girder as shown
/ in Figure 2-11. The saddles, support colunns and ring girder all lie in the sene plane and react all vertical loads on the torus - most of 1.he load i.;
reacted by the saddle.
Each saddle is constructed of 1 -inch type SA 516 GR 70 steel e web plates, which are welded to the torus shell at the top, and to a clevis /
I sole plate assembly on the bottom. Anchor bolt attachments are welded to the web phtes and clevis sok plates, and eight anchor bolts per saddle restrain upwant motion. A total of 128, two-inch diameter Williams Rock Bolts were installed, with a 24-inch minimum embedment into the basemat.
}. The ancher Dalt restraints are set with a small clearance, dnd the sole plates rest Gn lubrile plates, to allow for normal radial growth
, of the torus due to temperature changes. The lubrite plates rest on a bed j- plate which is grouted to the basemat.
Technical Report WMNE TR-5310-1 ENGNEERNG SERVICES Installation of the saddles required relocation of the 18-inch core spray lines, to eliminate interference problems.
Torus Support Columns and Bases
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The torus columns were welded to the saddle webs and the bases will be held down by using the existing "J" bolts (two per column).
Figure 2-12 shows one of the two beams that bridge the base holding down the
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column. There is a gap between the beam and base that will allow for thermal growth, radial movement of the column sliding on the lubrite plate.
Downcomer Tie-Rods and Vent Header Gussets The downcomer tie-rods and vent header gussets are illus-trated in Figure 2-13.
3 The tie-rods are constructed from 21-inch i 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, caitstructed from 3/4-inch steel. The clamps
, are prevented fro,n sliding on the downcomers by welded steel clips both above and below the clamp.
The gussets between the downcomers and vent header are to 3 reduce local jiitersection stresses due to chugging lateral loads on the down-c oiners . They are constructed from 1-inch 2 thick SA 516 GR 70 steel plate, and are welded to the vent header and downcomers.
Rin1_ Girder Weld Modifcation lhe weld between the ring girder and the torus shell will be increased in the area of the tee-quencher as illustrated in Figure 2-21. The e<isting 5/16-inch fillet weld was increased to l-inch, i as shown, to help react and distribute the laterai tee-quencher reaction loads due to SRV discharge.
Technical Report "RTA WNE TR-5310-1 ENGINEERING SERVICES
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Catwalk and Handrail The catwalk and handrail at Pilgrim required substantial modification, as illustrated in Figures 2-14 and 2-15. The planned or com-pleted modifications included the following:
e Replacement of original catwalk extension support columns with four-inch diameter pipe columns.
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e Addition of bracing between catwalk channels.
e Addition of lateral supports.
e Addition of grating restraints.
e Replacment of handrails.
Electrical Junction Box Modification There are two large, barrel-shaped, electrical junction
, boxes inside the Pilgrim torus. These boxes are mounted at the top of the torus, in the path of both froth 1 and froth 2 impingement loads. They are constructed with sheet metal covers which could deform under these loads.
The junction box covers were strengthened by the addition of external reenf orc <? ment as shown in Figure 2-19. The reenforcement consists of three longitudinal struts connected to two rings which encircle the box and prevent buckling from occurring.
Drywell,-jietwell_ Vacuum _Br,eake,r, Suppor_t_
The attachment between the wetwell vacuum breaker mounting stub and the end cap of the vent pipe was strengthened by the addition of three gussets as illustrated in Figure 2-18. This modification assures
1%'N T 531 1 N
) acceptable stress levels in the intersection, during postuleted pool swe!1 impact. There are ten vacuum breakers at the Pilgrim Plant.
Spray Header Support Modification
)
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 was a "U" shaped rest support which could not
) react upward loads. These were modified as shown in Figure 2-16 to react the upward loads associated with pool swell and froth.
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Technical Report TR-5310-1 )
EXTERN AL SUPPORT CO LUM NS
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TORUS PL AN VIEW 4
) YM Technical Report TR-5310-1 )
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VACUUM BREAKER LINE qp "N ~9 VENT '
) HEADER i DRYWELL CAT WALK DEFLECTOR
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O = MODIFIC ATIONS FIG . 2-2 TORUS COMPOSITE. CROSS SECTION-PILGRIM
) Technical Report TR-5310-1 '# TF1 W E ENGNEERING SERVICES KEY FOR FIGURES 2-3 AND 2-4
) Modification Completion Date
- 1. Reenforce Supports on 4" Spray Header 12/81
- 2. RHR Elbow Return 5/80
) 3. Strengthen Electrical Penetrations 12/81
- 4. Shorten HPCI Exhaust Line 12/81
- 5. Shorten RCIC Exhaust Line 12/81
) 6. Shorten HPCI & RCIC Drain Lines 12/81
- 7. Shorten 18" Spare Line 12/81
- 8. Mitred Joint Saddle 5/80 i
- 9. Saddle Anchor Bolts 4/82
- 10. Shorten Downcomers 12/81
- 11. Downcomer Ties 5/80
- 12. Vent Header Deflector 5/80
- 13. Vent Header Downcomer Stiffening 12/81
- 14. Vent and Drain Cut and Cap 12/81
- 15. Monorail - Remove 12/81
- 15. Drywell to Wetwell P Control 1976
- 17. Add Temperature Monitoring System (Except Thermowell) Prior to fuel cycle #7
- 18. Thermowells 5/80
- 19. Add Larger Vacuum Breakers to the SRV 5/80 & 12/81 L ine in Drywell and Support Lines
- 20. Add Tee-Quenchers 5/80
- 21. Tee-Quencher Supports 5/80 & 12/81
- 22. Wetwell Vacuum Breaker Penetration Modification 12/81
- 23. Catwalk Modifications 12/81 & prior to fuel cycle #7
.?4 . Increase Ring Girder Weld Prior to fuel cycle #7 l
- 25. Column Base Holddown Prior to fuel cycle #7
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Technical Report pg TR-5310-1 g
)
3.0 CONTAINMENT STRUCTURE ANALYSIS - OUTER SHELL & EXTERNAL SUPPORT SYSTEM (INCUDING 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 Figure 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 si,cwn 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 atfect results; accordingly, the offset is not included in the model.
)
~
)
Technical Report TIO DE - '
TR-5310-1 -
NM set \/ ICES
) ,
Thismodelincludes513structuralnodes,518platediements,2242 static degrees and 364 dynamic degrees of freedom. Sy1:netric boundary condi-tions were used at both ends of the model.
)
The model was modified for various load calculations to a'ccount for differences in the percent of the water mass that is effer.tive for that load .
~
event. In all cases, modeling of the water mass was ac.omplished using a 3-D virtual mass simulation as an integral part of the ,tructural analysis. The
)
percent of water mass used is identified in the discussion uf 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 load?, on the support system and '
carthquake restraint system.
The beam element properties wxe selected,tg 7:
simulate combined bending and shear stiffness of the sectioY.,. Water 'im s/ ..
~'
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) was lumped with the structure weight on the wetted nodes. '
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3.2 Loads Analysis s
< / < -
) 3.2.1 p
_ool Swell Loads (4.3.1 & 4.3.2) '
Analysis for pool swell loads was dona .~using the firiite element model shown in Figure 3-1. Thiswasadynamicanalysispc-formad,Yn-f u. ,
a, ~
)
the time domain by applying a force time history, i.o simulate the(pressure ,
Lime Mstories of the pool swell event to each node on Abg computer modal. '
Input pressure-time histories were varied in 'both the lo(gitudinal and radial directions in accordance with the information in References 1, 2, 9 and 10.
i
) Typical pressure-time history curves are shown in Figures 3-5 through 3-7. '
( t hese pressure-time histories are taken directly from Reference 10, bet ore '-
adjustment, as required by Reference 2. Therefore, the amplitudes showrl.are '
slightly different than the loads used in the analysis).
' ', ~; -
) -
~
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l Technical Repcrt W F W NE TR-5310-1 ENGNEERING SERVICES 1
The computer analysis was run for two different pool swell l conditions, full A P and zero A P. Figures 3-5 through 3-7 show comparative l values and time histories for the two cases. The only difference between the analyses was the input loads; the models were identical. Details of the full load distribution can be found in References 1, 9 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 different for both zero and full A P conditions (Reference 4). The water mass l
used in the computer simulatio1 was constant throughout the analysis 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 relatively slow J. (pseudo-static) nature of the pool swell load. This simplification only L. 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 y 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 ther, scaling and combining the i " Nual frequency effects to determine total stress at selected elements. The w 2e 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.
y .
( The corrhination of individual harmonic stresses into total element stress was done by considerir; frequency contributions at 31 Hz and tw;
/ '<
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l l
)
Technical Report YM l TR-5310-1 N sot \/lCES '
) 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 Ref erences 6 and 14, and in Appendix 2 of this report.
)
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 asymmet.ric 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 f rom 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. l
)
Analysis for asymmetric pre-chug was performed using the beam model in Figure 3-4 by applying the unbalanced lateral load through the prescribed f requency 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 YE NE TR-5310-1 ENGNEERING SERVICES
,) Maximum Shell Stress Shell Membrane Stress
) Pre-Chug 1284 psi Post Chug 774 psi l
) 1. Based on frequencies to 30 Hz - sum of 4 maximum +
SRSS of others.
Additional work published in Reference 12 showed
) that pre-chug bounded post chug (to 50 Hz) for column and saddle loads (Table It also showed that PL+Pb stress due to post chug exceeded 5-1, Ref. 12).
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 f act 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 guidelines given in Reference 2. In these tests, pressure amplitude and f requency were recorded and compared to calculated values for the test condi-
) tions. Factors were developed that related test to calculated values for both amplitude and frequency (see Appendix 1). These f actors were then applied to calculated load values for other SRV conditions; the structural analyses were
)
)
Technical Report TN TR-5310-1 N SBt\/ ICES
)
performed using these adjusted values. Appendix 1 discusses the in-plant test and enalysis in more detail. A typical set of SRV shell pressures is shown in Figure 3-8.
)
The method of modeling the water mass in the SRV computer model was the subject of extensive study in this program. Initial atterrpts to reproduce measured stresses by applying measured pressures to the computer models f ailed badly. After 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 & In'ternal Pressure
)
Deadweight, thermal and internal pressure analyses were done using the computer model shown in Figure 3-1. Resulting stresses were calcu-lated and considered for all elements on the model.
I For the thermal analysis, conduction into the columns and saddles f rom 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 stat-ic 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.
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
)
)
Technical Report YM TR-5310-1 MM
)
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 pipicg penetrations will be described in TES report TR-5310-2, when the piping analysis is complete.
)
The fatigue analysis of the shell and support system was a conservative one, which was based on applying a stress concentration f actor of 4.0 on the most highly stressed elements for each load case. In the case of the support system, only the column-to-torus and saddle web-to-torus welds
) were considered, since they have higher stresses then 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).
nd e It adds the maximum absolute stress for each load
) case even though they do not usually occur at the same element.
The procedure used in this analysis consists of the follow-
)
ing steps.
- 1. For a given load, locate the maximum component-level stresses (S x , S , Sxy) f r the f ree shell,
, local shell, and the supports.
- 2. For these locations, establish the stress intensity ranges and the approximate number of cycles.
)
Technical Report Y TR-5310-1 N SBi\/ ICES
) 3. Repeat the process for all other loads in the load combination.
- 4. Add the stress ranges for all loads, independent of
) sign.
- 5. Multiply these total stress ranges by 4.0 (the SIF).
)
- 6. Calculate the alternating stress intensity and com-plete the 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 af ter chugg-ing begins. Plant procedures are presently under study to provide for this
) action.
3.3 Results and Evaluation Results are reported for each structural component of the containment system for the controll,79 load combination. Controlling load combinations are the ones that produce the smallest margins against the allowable stress -
not necessarily the highest stress.
)
All load combinations listed in Table 1 have been considered. As stated previously, most results include some level of bounding analysis and, therefore, understate the margins which actually exist.
)
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 f ormed. Stress intensities were then calculated from these total component-level values.
)
)
Technical Report TN TR-5310-1 gMM The controlling load combinations for the shell at Pilgrim are cases 14 and 20 in Table 1; these are: !
Case 14 - IBA.C0 + SRV + Seismic (SSE Used) + Pressure
)
+ Weight l
Case 20 - DBA.C0 + Seismic (SSE) + Pressure + Weight k
These load combinations control all categories of shell stress, although the location of the elements is different for the different types of stress. The .~ollowing table summarizes the controlling stresses.
Approximate locations of the controlling stresses are shown in Figure 3-9.
CONTROLLING SHELL STRESSES - PILGRIM ACTUAL ALLOWABLE
) TYPE OF STRESS STRESS STRESS
& LOAD CASE LOCATION (psi) (psi)
Membrane (Pm) Free Shell 13,324 19,300 (Case 14) Element 17
)
Local (PI) Local Shell 8,365 28,950 (Case 14) Element 125
)
Membrane + Free Shell 17,258 28,950 Bending Element 19 (Case 20)
Stress Range Local Shell 26,399 69,900
)
(Case 14) Element 147 Compressive Buckling - Acceptable (see below)
)
Compressive Buckling Reference 13 discusses the results of analy-tical studies and tests on Mark 1 torus structures to determine their compressive buckling capabilities. The report concludes that SRV is the dynamic load which presents the maximum chance of com-
)
)
Technical Report YN TR-5310-1 MM
) pressive buckling f ailure; but, that a safety f actor of 7 still exists for an applied SRV pressure of +29.3/-22.6 psi. The maximum worst-case SRV shell pressures for Pilgrim are +12.6 psi and -9.6 psi, which are lower than those used in the referenced study. Based
) on this, compressive buckling stresses are considered to be accept-able for the Pilgrim torus.
FATIGUE EVALUATION - PILGRIM
)
CUMULATIVE USAGE FACTOR
SUMMARY
STRESS INTENSIFICATION FACTOR = 4.01
) NORMAL EVENT TYPE r
_v OPERATION SBA/IBA DBA
.005 .001 .015 iv .164 .003 .026
)
3.3.2 Supgrt Co: -
& Att nhm et:
)
The controlling loac ase f or the upport column and attach-nent weld to the torus shell is load case 16 in ble 1. Controlling stresses are associated with the downward loads. Case 16 includes:
)
Pool Swell (OAP) + Pressure + Weight For this load case, the following controlling stresses were identified:
)
SUPPORT COLUMN - CONTROLLING AXIAL CONDITION LOAD CONTROLLING ACTt!AL ALLOWABLE COLUMN DIRECTION STRESS FACTOR FACTOP.
)
Inner Down Axial + Bending .307 1.0 Outer Down Axial + Bending .352 1.0
)
)
Technical Report TN TR-5310-1 MM
)
COLUMN-TO-SHELL WELD LOAD CONTROLLING ACTUAL ALLOWABLE
) LOCATION DIRECTION STRESS STRESS STRESS Inner Down Shear 12.32 K/in 28.9 K/in Outer Down Shear 10.99 K/in 28.9 K/in 3.3.3 Support Saddles & Shell Weld
(
) The controlling load case for the weld between the saddle and the torus shell, and for down loads on the saddle is load case 16 in Table
- 1. This case includes:
Pool Swell (OAP) + Weight The resulting stresses are:
SADDLE STRESSES l
LOAD TYPE ACTUAL ALLOWABLE I LOCATION DIRECTION STRESS STRESS STRESS Sole Plate Down Bending 19.65 K/in 28.5 K/in SADDLE-TO-SHELL WELD i
LOAD TYPE ACTUAL ALLOWABLE l LOCATION DIRECTION STRESS STRESS STRESS l
l Outside End Down Shear 11.93 K/in 13.64 K/in 1
)
Technical Report TM TR-5310-1 6M
)
The controlling case for the saddle associated with up load is case 21, which includes:
DBA.00 + Seismic (SSE) + Weight
)
Loads for this case are:
LOAD iYPE ACTUAL ALLOWABLE LOCATION DIRECTION STRESS LOAD LOAD I
l Clamping Up Bending 95 kips 103.2 kips Plate 3.3.4 Earthquake Restraints & Attachments The earthquake restraint system is illustrated in Figure 3-
- 10. The controlling load case for this system is the one that produces the largest lateral load. This is case 15 which includes:
Chugging + SRV + SSE All three of these loads have been selected to produce the highest lateral load on one earthquake restraint; contributions fron1
- the individual loads were added directly.
( The controlling stress results follow:
EARTHyVAKE RESTRAINT STRESS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS Tie Plate Pin 4,879 psi 27,000 psi at Slot Bearing
)
Technical Report WMg TR-5310-1 6M f ATTACHMENT WELD l
STRESS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS h
l Weld at Tie Shear 3,668 psi 21,000 psi l
Plate - Base Plate Connection i
) 3.3.5 Anchor System The torus at Pilgrim is restrained against upward loads by two-inch diameter anchor bolts in the support saddle and 1 -inch anchor bolts in the column bases that were part of the original plant design.
l The original 1k-inch column base bo?ts did not restrain up-l ward movement because they were used only to hold a base plate which was not l attached to the column. The tiedown fix will use the original bolts with l clamping plates to tie down the torus columns.
l The controlling load case for these anchor belts is cese 21 in Table 1. This case includes:
DBA.C0 + Seismic (SSE) + Weight The loads are:
i SADDLE ANCHOR BOLTS (PER SADDLE)
ACTUAL FACTOR MAXIMUM MAXIMUM OF LOAD CAPACITY SAFETY 76.6 K/ bolt 264 K/ bolt 3.45 l
l l
Technical Report TN TR-5310-1 Na NES The original column anchor bolts are a "J" bolt embedded in the concrete and constucted from structural steel. The capacity of these bolts will be determined by the stresses in the steel bolt rather than con-
) crete pull out capacity. Accordingly, they are evaluated against stress:
, COLUM;1 ANCHOR BOLTS - ORIGNAL l
) ACTUAL FACTOR I COLUMN MAXIMUM MAXIMUM OF l.0 CATION LOAD CAPACITY SAFETY l
l Inner 47.5 kips 60 kips 1.26 l
Outer 40.3 kips 60 kips 1.49 l
i i
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) Technical Report TR-5310-1 _54- EMBEEMWQM
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.g SECTION A- A FI G. 3- 10 E ARTHOUAKE RESTRAINT SYSTEM
)
Technical Report TM TR-5310-1 6M 4.0 VENT HEADER SYSTEM 4.1 Structural Elements Considered f The vent header system, as defined in this section, includes the following structural components:
- a. Vent Header (V.H.)
)
- b. Main Vent Pipe (V.P.)
- c. Downcomers (D.C.)
- d. Downcomer Tie Bars
- e. Deflector
)
- f. Vent Header Support Columns & Attachments 9 VH/DC Intersection
- h. VH/VP Intersection
- i. VP/Drywell Intersection
)
J. Vent Header Mitre Joint The main vent bellows are considered in Section 7.0.
4.2 Computer Models Two computer models provided the means to analyze the vent header system, they are shown in Figures 4-1 through 4-4.
The first of these is a detailed shell model, (Figures 4-1 to 4-3),
which includes a highly detailed representation of one-half of the header in a non vent bay, complete with four downcomers.
The model also includes an approximate representation of one-half of the vent bay; this was intended to provide the proper boundary conditions and stif fness transition near the non-vent bay. The vent bay half of the 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
)
Technical Report TN TR-5310-1 ENGNEERNG SERVICES
) 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 shawn in Figure 4-4.
This model represents a full vent bay, complete with vent pipe and downcomers; as well as a half non-vent bay on either side. It was used to determine boundary loads on the vent system components to support a more detailed stress
) analysis of those components. This model was used to define loads on the following elements:
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 l
e VP/VH intersection - applied stress multipliers (stress intensification factors) from Reference 7 e VP/drywell intersection - used stress multipliers from -
l Reference 16 (Bijlaard) l e Mitre joint - used stress multipliers from detailed shell model (Figure 4-1) e Main vent pipe - hand analysis l
l The beam model used a stiffness representation of the VP/VH inter-section taken from Reference 8. Attachment stiffness between the vent pipe and drywell was calculated using References 17 and 18.
l l
)
Technical Report TN TR-5310-1 N SBt\/ ICES
) 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 4.3.1 Pool Swell Loads
)
4.3.1.1 Pool Swell Water Impact Analysis for stresses due to pool impact and drag was done using both computer models.
)
Determination of shell stresses was done with the detailed model in Figure 4-1. For this analysis, force time histories based on QSTF test data were used (References 4, 9 and 10). These time histories
)
were applied at 100 nodal points on the shell model and the dynamic response of the structure was calculated. Relativ2 timing between loadings (Reference
- 1) was maintained to preserve accurate representation of longitudinal and circumferential wave sweep. Stresses in the vent header /downcomer inter-2 section, as well as in the free shell areas, were taken directly from this
! model. Stresses in the downcomer tie bars were also taken from this model.
Analysis was done for both full and zero A P impacts.
[
The beam model (Figure 4-4) was also used to deter-l mine stress from pool swell impact and drag. This was done with a time history dynamic analysis using leads developed by integrating the impact pressures over small areas and reducing them to nodal forces. Approximately 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.
l l
l
)
Technical Report TN TR-5310-1 6M
)
4.3.1.2 Pool Swell Thrust (4.2)
Pool swell thrust forces are definea 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 model and applying the loads statically. This is consistent with the slow 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)
The vent header support ec'umns are loaded by for-ces from LOCA-jet and LOCA bubble drag. By inspection, it was concluded that f LOCA-jet loads would not combine with water impact on the vent system due to -
differences in timing and, therefore, would not contribute to the maximum stress calculations - LOCA jet forces were not considered further.
LOCA bubble forces were calculated and the maximum normal components (radial and longitudinal) were applied simultaneously to l conservatively bound the bending moments on the support column. These peak l values were applied statically at the midpoint of the column. Stress calcu-lations were done by hand.
I 4.3.2 Chugging Loads l
l 4.3.2.1 Downcomer Lateral Loads (4.5.3)
Ref erence 1 identifies downcomer lateral loads as static equivalents with random orientation in the horizontal plane. The major consequence of this loading is to produce high local stress in the VH/
TME T 31 1 MO
) 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 recessary to determine the proper dynamic scale factor to apply to the static load.
The stress results from the statically applied load
- were used as a basis for a fatigue evaluation of the intersection in accord-ance with Reference 9.
4.3.2.2 Chugging - Synchronized Lateral Icads
)
The random nature of the downtomer lateral chugging load provides for all combinations of alternate force orientations on adja-cent pairs of downcomers. Various load combinntions were examined to deter-
) mine stress levels in the vent header and mitre joint as a result of these loads. The cases considered are shown in Figure 4-6.
These cases were considered by applying static loads to the beam model (Figure 4-4) and determining final stresses as described in Section 4.2.
l l 4.3.2.3 Internal Pressure (4.5.4)
Three vent system internal pressures exist during chugging. They are:
a Gross vent system pressure - a .7 Hz oscillat-ing pressure with a maximum value of 5.0 psi.
This pressure acts on the entire vent system.
o Acoustic vent system pressure - a sinusoidal pressure varying f rom 6.9 to 9.5 Hz at a maxi-mum value of 3.5 psi. This pressure affects the entire vent system.
l
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e Acoust'ic downcomer h,reisvie'oscIIlationIe p
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50 Hz pressure At' I3 pfi' thaE. produce's .
f only ,i hoop stress'in the downCodiers.
~ ~;; /, !i
)
Responses to these pressures wer)eJstimated using 5 a ,-
hand analysis and were determined to be substantially less::than TAose result- , .
l ing from internal vent system pressures during po'oY swell,. The' values associ- <
[/ ,,
. jv. c>
ated with pool swell pressures were used in al.L combiced< ond cases involving dt
) chugging pressures; this produces conservative,fesv,its. ' 'fS c i
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\
! 4.3.2.4 Submerged Structure Drag (Support Co umns only) M ,
A s ,
- ' % ' 'd
- Examination of the loa'd c,ombm[ations tM'y include i
4'-
chugging makes it clear that these cannot control m;aximum,atress 1) Vel in the '
support columns; combinations that inchtde'I;en '
headep water., impaIlt fl ,
produce much higher stresses. For this reaso'n, stresses in Ic vent tcadgr I '# . [
f support columns were not calculated for chugging drag. -
,'~ m ,,
f ,
3,3 Drag forces on ther dov:ncomers and downcomer tie bars are already included in the Downcomer Lateral, Loads, 'which were based j
) " '
directly on test data.
xv - - - , ,
4.3.3 Condensation Oscillation - DBA ' Q l' s 4.3.3.1 Dov.ncomer Dynamic Load (4.4.'3.2) j f
,j ' -: ..I'
'Tha dowacomer' oynamikioad, due to condensation e
oscillation, is a sinusoidal pressure ve;r'iation that can b; equel or unequal s' in the two downcomers forming a pair. '" '
/
- f .- ,'
f ,as The unequal prunfre v/yjep pr3 duces a het -1steral
~i s ~
load on the downcomer much like chugging. Themqjorconsiderafionsforthis load are stresses in the downcomer intersection due to a ne'deteral load on
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~ . Me pal: of boincomers 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 th ven't header.
y 7"
l
/' The horizontal component of the C0 downcomer load
. produces ,the same typc,0f loading on the vent system as the CH lateral load,
-in terms of the stress analysis. A comparison of the magnitudes and frequen-j cies of these two loads shows that stresses resulting from CH horizontal loads will bound C0 horizontal loads.
~. ,
J s*
f
~,1, ..
gThe C0 downcomer load also produces a vertical qcmponent 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 (horizoidal and vertical components) would still be bounded by b, CH lateral loads.
Based on this, CH lateral load results were con-servatively used in all load cases in place of C0 downcomer loads.
f ..
1 l 4.3.3.2, Vent System Loads (4.4.4) 1
- Vent system loads consist of a sinusoidal pressure
)
r- in the vent header and downcomers superimposed on a static pressure. The dynanic pressura in the downcomers is used to calculate hoop stress only.
Stresses for all pressure loads were based on hand analysis using static analysis. The static analysis assumption is consistent with the low frequency of the applied pressure and the fact that the ring modes of the structure are very high.
)
k .
W TF1A WNE 5 l 4.3.3.3 Thrust Forces (4.2)
Stresses resulting from C0 thrust fcrces were conservativel, taken from the pool swell thrust calculations and applied to
)
all C0 load cases (para. 4.3.1.2).
4.3.3.4 Drao 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-I lation are bounded in all cases by either DBA condensation oscillation or chugging. No detailed analysis was performed for IBA condensation oscil-lation.
4.3.5 SRV Loads 4.3.5.1 SRV Drag Loads
)
An SRV discharge produces drag loads which act on the vent neader support columns, downcomers, and downcomer tie bars. Analy' sis 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.
)
)
Technical Report W F W NE TR-5310-1 ENGNEERING SERVICES 4.3.6 Other Loads
)
Deadweight and seismic stresses in the vent system were cal-culated using the beam model cf Figure 4-4.
)
Seismic stresses were calculated by statically applying the acceleration values in Table 3.
)
Thermal stresses were determined for the steady state appli-cation of maximum vent system temperature, using hand analysis.
4.4 Results and Evaluation
)
Results are reported for each structural element of the vent system for the controlling load combination. Controlling load combinations are the ones that produce the smallest margins against the allowable stress - not
)
necessarily the highest stress. All load combinations listed in Table 1 have been considered.
As stated previously, most results include some level of bounding analysis and, therefore, understate the margins which actually exist.
4.4.1 Vent Header - Downcomer Intersection The controlling load on the vent header-downcomer inter-section, bcr.h for maximum stress and fatigue, is the downcomer lateral load due to chugging. The worst load combination in which this load appears is Cdse 27 of Table 1. This cases consists of:
Chugging (DBA) + Seismic (SSE) + Weight + Presssure + 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.
)
Technical Report WP WNE TR-5310-1 NN NES
-) 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
)
ACTUAL ALLOWABLE TYPE OF STRESS STRESS STRESS
}
Combined Maximum Stress 28,930 psi 28,950 psi This load case was formed using a 0 P load, and was eval-uated to a level A allowable. This conservative evaluation was performed to f eliminate the need to evaluate several other vent header load cases.
4.4.3 Vent Header Support Columns and Attachments f 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
Pool Swell (fullAP) + Seismic (SSE) + Weight + Thrust + SRV Pressure 1
As betore, the evaluaticn was conservatively performed using OAP loads and a level A allowable.
Controlling stress in the support column is:
l l
l l
)
Technical Report 7%' F W NE TR-5310-1 ENGNEERNG SERVICES
) ACTUAL ALLOWABLE H PE OF STRESS STRESS STRESS Axial in Column (tension) 17,420 psi 18,000 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 13,015 psi 15,200 psi 4.4.4 Downcomer Tie-Bars and Attachments
)
The controlling load combination for stresses in the down-comer tie bar and attachments is case 25, in Table 1. The major load is associated with pool swell impact on the crotch region of the downcomers which
)
produces tensile loads in the tie bar.
The controlling case includes:
)
Pool Swell Impact (full AP) + SSE Seismic + SRV + Weight +
Pressure + Thrust
)
Zero A P pool swell loads and service level allowables were conservatively used in this analysis.
The controlling stress is:
I STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS Tie Bar Clamp Bending 16,800 psi 22,240 psi
)
)
)
Technical Report TE DE TR-5310-1 ENGNEERING SERVICES 4.4.5 Vent Header Deflector and Attachments
)
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 (full AP) + SSE Seismic + Weight + Thrust The controlling stress in the deflector is:
)
STRESS ACTUAL ALLOWABLE LOCATION TYPE VALUE VALUE Center of Bending 21,000 psi 23,700 psi the Long Span The controlling stress for the attachments is:
)
STRESS ACTUAL ALLOWABLE LOCATION TYPE VALUE VALUE
)
Fillet Weld Shear 15,800 psi 17,100 psi 4.4.6 Main Vent /Drywell Intersection The major load on the drywell penetration occurs as a result of chugging. The controlling load condition is case 21 in Table 1 which includes:
Chugging + Seismic (SSE) + Weight + Pressure (Drywell)
The controlling stress is:
ACTUAL ALLOWABLE TYPE OF STRESS STRESS STRESS Primary and Secondary 39,970 psi 69,900 psi
)
Technical Report W TFIFrT(NE l TR-5310-1 ENGINEERING SERVICES The effects of all loads from the vent system, and the pres-
)
sure load were considered using Reference 16. Information regarding stresses due to seismic and thermal response of the drywell is not available and therefore have not been included.
)
4.4.7 Vent Header, Main Vent & Downcomers - Free Shell Stresses It was established by inspection of the computer results that large safety margins existed in free shell regions and that minimum
)
safety margins would be controlled by local shell stresses. No further work was done for free shell stress in these structures.
4.4.8 Vent Header - Mitre Joint
)
The controlling load on the vent header mitre joint occurs as c 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
)
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 lcogth of chugging associated with an SBA/IBA event.
)
The procedure used in this analysis corisists of the follow-ing steps:
)
1 b
Technical Report 7%' F W NE TR-5310-1 ENGINEERING SERVICES e For a given load and component, locate the highest stress.
e For this location, establish the stress range.
)
e Repeat this pr_ cess for all other loads in the load combination.
. Add the 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.
A The f atigue evaluation was performed for all high stress
)
areas in the vent system. The major load, contributing to the f atigue evalua-tion, is chugging following a DBA. The controlling loaa case is case 21 in Table 1, which includes:
Chugging (DBA) + Seismic (SSE) + Weight The controlling usage factor for the vent system is:
3 VENT SYSTEM FATIGUE RESULTS ACTUAL ALLOWABLE USAGE USAGE LOCATION FACTOR FACTOR VH Support .76 1.0
_ o
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DETAILED VENT HEADER MODEL
) Technical Report WM TR-5310-1 ' gm
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FIG . 4-5
, VEN T HE ADER DEFLECTOR AN A LYSIS - PILGRIM
Technical Report WTELEDGE TR-5310-1 NSERVICES
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ANALYSIS CASES FOR SYNCHRONIZED LATERAL CHUGGING FIG 4-6
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'WTFI FTWNE b1 -78 . ENGNEBUNG SERVICES
)
5.0 RING GIRDER ANALYSIS The ring girder for Pilgrim is shown in Figure 5-1. It is mounted in a
)
vertical plane that passes through the support saddles and the support col-umns. Because all major internal structures are supported by the ring gir- j ders, the ring girders that must react to the largest number of individual j 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, catwalk, etc.
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 structural 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-t - - -
)
Technical Report TE WE TR-5310-1 ENGINEERING SERVICES
}
sixteenth model used for the Pilgrim 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 Pilgrim; the diameter of the torus, shell thickness and distance between the ring girder and mitre joint are all similar. The ring girder fla.nge in this model is slightly smaller than Pilgrim and, therefore, produces conservative results since lateral loads control ring girder stresses. The comparison is:
)
Ring Girder Flange Dimensions (inches)
Pilgrim: 1.5 x 7 Model Used: 1.5 x 6
)
The second model used to determine ring girder loads is the. Pilgrim 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)
)
)
Technical Report SPTF1 FrT(NE TR-5310-1 ENGNEERNG SERVICES
)
(d) SRV Discharge *
(e) Seismic (f) Deadweight, Thermal and Pressure
- SRV discharge is conservatively assumed to be a symmetrically 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 on the 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 amplification was 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 But ble (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 1
based on data collected from in-plant tests. A discussion of the in-plant tests and the use of drag data from these tests is given in Appendix 1.
Calculation of ring girder drag loads due to condendation 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
)
)
Technical Report TR-5310-1 W F W NE ENGNEERING SERVICES calculation. It produced drag volumes for the ring girder of about half of
)
those that the NUREG 0661 procedure would have produced. A discussion of this is included in Appendix 3. The FSI drag calculation was based on local pool accelerations at the ring girder due to the response of the entire shell. The post chug and FSI analysis considered frequencies to 31 Hz, which were com-
)
bined by adding the values of the five maximum components to the SRSS sum of the others.
5.3.3 Loads Due to Attached Structure
)
Loads applied to the ring girder by structures attached to it were evaluated by equivalent static analysis, using the 1/16 model (Figure 5-2). The important loads are applied in the area of the support saddle and
)
columns which make the ring girder very stiff and minimizes dynamic inter-action. Because of this, dynamic amplification of the static ring girder stresses was not done (DLF = 1.0). The load input to the ring girder was a result of a dynamic analysis of the attached system (or had an appropriate DLF
)
applied) and, therefore, 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.
e Vent header support column reaction loads during pool swell.
e Vent header support column drag loads.
)
e Catwalk support column reactions and drag.
)
)
Technical Report TR-5310-1 SPTFI WE ENGNEERING SERVCES
)
As stated in Section 5.1, stresses resulting from attached structure have been included in the following results.
) 5.4 Results & Evaluations 5.4.1 Ring Girder Web & Flange
) The controlling load combination for the ring girder web and flange is load case 16 of Table 1; this includes:
Pool Swell (OAP) + Internal Pressure + Weight
)
The controlling stress is:
LOAD STRESS ACTUAL ALLOWABLE LOCATION DIRECTION TYPE STRESS STRESS
)
Web Down Membrane 16.6 ksi 28.95 ksi
+ Bending
}
Flange Down Membrane 15.0 ksi 19.3 ksi 5.4.2 Weld to Torus Shell I
The controlling load combinations for the shell weld are load cases 21 and 25 as indicated below. These cases include:
Lead Case 21
)
DBA.C0 + Seismic (SSE) + Pressure + Weight Load Case 25 Pool Swell (full AP) + Saismic (SSE) + Pressure +
I Weight
)
)
Technical Report 1%P W NE TR-5310-1 ENGNEERING SERVCES
)
For these cases, the controlling loads are:
RING GIRDER /SHELL WELD LOADS
')
LOAD STRESS ACTUAL ALLOWABLE LOCATION CASE TYPE LOAD LOAD Inner Column 21 Shear 7.46 K/in 8.53 K/in
)
Region Outer Column 21 Shear 8.34 K/in 8.53 K/in Region
)
Saddle Region 25 Shear 8.29 K/in 8.53 K/in
)
)
)
)
)
)
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) WTELEDGE Technical Report TR-5310-1 mm
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) WTERME ENGDEERNGSERVICES Technical Report TR-5310-1 )
MITRE JOINT OFFSET -~ --
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)
RING GIRDER ELEMENTS
Technical Report TF WE TR-5310-1 ENGNEERNG SERVICES 6.0 TEE-QUENCHER AND SUPPORT The following results for the tee-quencher and support are conservative due to the combined effect of several factors, three of which are:
)
e The calculational methods to determine applied loads improved af ter this analysis was complete, and would provide reduced stresses.
Computer program RVFOR-04 was used to calculate SRV blowdown loads.
)
RVFOR-05 is now available and gives less conservative loads.
e Some loads were intentionally bounded by conservative values from other plants so a single calculation could be used for more than one
}
plant.
e For submerged drag loads, individual frequency components were added to produce maximum stress without regard to load direction.
The effect of these conservatisms vary among stresses, but can be significant in some cases.
)
6.1 Structural Elements Considered The configuration of the quer.cher and support is shown in Figure 2-
- 8. Pilgrim has four discharge lines, each enters the pool at a 20 angle.
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.
)
M Technical Report YM M ^
TR-5310-1 ENGNEERNG SERVICES .(
6.2 Computer Models ' 'N "
) '
\
ThecomputermodelusedinthisanalysisisshowninFigure61N' 7
. \+
This is a STARDYNE beam model which represents all piping and stru6-
)
ture between the drywell jet deflector and the ring girder. For these analy-ses, the ring girder was assumed rioid 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 supports. .
l- l ,
This model was used for both static and dynamic analysis.
) ,
6.3 Loads Analysis s 6.3.1 SRV - Load
)
The calculation of stress due to SRV blowuov;n'was dole by applying the dynamic loads to the computer model and calc $lating the time- >
history response of the system. The applied loads inclused both the bbwdown
)
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-05. '
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 dynamic load factor of 2. It was clear from this analysis that combined pool
) swell events would not control stresses - no further analysis was done.
) -
)
Technical Report $pg TR-5310-1 NBtNG set \/ ICES h
6.3.3 Chugging Loads Dynamic analysis of the quencher and support system was done T
for drag loads due to pre-chug, post chug and chugging FSI. All these analyses were based on a set of harmonic analysis which provided results for k'c( - all steady-state frequency excitations from 1-31 Hz. Results for individual T load conditions were determined by scaling individual frequency results of 3 the computer analysis by the appropriate pressure amplitude.
s ,
s
, s The mass of the structure used in the computer analysis was l' ] q 'adju;;ted to account for the "added mass" effect of the surrounding water. For
) ESi and postdugging analyses, individual frequency components were combined ng the five maximum frequency contributors to the SRSS sum of the otnnrs !sec keference l?.for discussion). The maximum value of each frequency
' component was used in the combination, regardless of vector direction or time
~'
or- if.s tantaneous .respon:d;. FSI loads were calculated by considering the c a,itpl ate <1 local accelt: rations in the poo.1 due to response of the entire
, shell. '
N.
) ,
6.3.4 Condensation Oscillation Loads
, . .n -
., .The quencher -a91 support systergre subjected to conden-sation oscillailc r drag and C0 3S'l drag. Analysis fo'r't%se-ioads was based
) onthesameharmonicanaly..isdiscus:edinparar[aph6.'3.3,xsr.aledtotheC0 amplitudes. Each of the thice C0 spectra shown i'rfigure 4.4.1-1 of Reference s
1 were <.onsidered. +
)_ '
All other discu% ion from paragraph 6.3.3 for chuggirg
.applij.;s to the condensation oscillation, analnis 3 except 4 hat the final load
' A1(det .riained i by adding thc f onc maximum frequency contributors to the SPSS
.a wm of'the oi.hers.
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and seismic
) loads was done by using the computer model in Figke 6-1 aitd static analysis.. vj
,'7 -n ,,w
,P Pressure stresses for the piping and quencher were calculated'by hand.
s 6.4 Results and Evaluatica l'/y*y' '
)
r\
The results renorted in this section may tv, d!uservative (epinding ,
ontheeffectoffactorsdiscussedinSections1.0(and6.0ofthisQp, -
ort.
, 1
) 6.4.1 Tee-Quencher f, ' , , '}. '
j -.
r e
, y nN
- l The controlling stress ,inithe te'e-qecacher itself occurs in r
the ramshead between the quencher ar,ms. It is the r,esult of a second SRV
) actuation after an SBA accident lohd case 15 of Table _1. It includes:
' 1."
, , )
SRV blowdown (case C3.2,) + Chugging Drag + WaigIt'+ SeismiS +
Internal Pressure + Thermal'
,s
, / ', I '
)
5 The controlling stress is:
j u
'~
STRESS STRESS ACTUAL ALLOWABLE -
LOCATION TYPE STRESS STRESS ** '
)
s Bifurcated Bending 18,671 psi 24,705 psi Elbow y i 6.4.2 Submerged SRV Line -
The controlling stress for the submerged portion of the SRV ,
e' >
line occurs in the inclined lines and is a result of load case 15 in Table 1,' '
)
This case includes:
) '
j : 'W M W ENGiNEN NES
' ' h)53
~
1 '
) -
SRV Blowdown (case C3.3) + Chugging Drag + Weight + Seismic +
Int.ernal Pressure + Thermal The controlling stress is:
) . \
STREdS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS f
) Vertical Bendir.g 9,094 psi 27,000 psi Section Above First Elbow
) 6.4.3 Tee-Quencher Support The controlling stress that was calculatec :or 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 +
i Thermal
)
The controlling stress for the bean is:
STRESS STRESS ACTUAL ALLOWABLE LOCU :0N TYPE STRESS STRESS
)
At the Brace Bending 17,361 psi 27,000 psi Connection f ., 6.4.4 Tee-Quencher Support Brace The controlling stress that was calculated for the tee-quencher support diagonal brace is the result of load case 15 of Table 1.
) This case includes:
)
Technical Report T3 NE TR-5310-1 ENG#EBUNG SERVICES SRV Blowdown (case C3.3) + Chugging Drag + Weight + Seismic +
Thermal The controlling stress is:
)
l STRESS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS In Brace Bending 12,719 psi 27,000 psi l
f
?
)
)
(
)
l . - - _
Technical Report WTERWIE TR-5310-1 _g4 MM i
?
l
)
\
l
)
VENT PIPE PENETRATION
)
l l
o h
o o
l
) QUENCHER SUPPORT TEE QUENCHER o m o- wo
) 0---o-o -o--o- - o-- e o
)
FIG.6-I PILGRIM ANALYTIC MODEL
)
i 1
) i Technical Report TR-5310-1 TE DE ENGNEERNG SEFMCES l 7.0 OTHER STRUCTURES
)
7.1 Catwalk l
The catwalk structure is attached to the ring girders and provides
{
) 360 access to the inside of the torus. It consists of a horizontal frame structure which supports sections of open grating. At each ring girder, one side of the frame is attached directly to the ring girder by a short hori-zontal member; the far side of the catwalk is supported by vertical columns
) which connect to the ring girder below the water line. Additional pipe column supports were added and the hand rail is wire rope as shown on Figures 2-14 and 2-15.
) 7.1.1 Computer Model The computer models of the catwalk are shown in Figures 7-1 and 7-2 for the original and modified catwalk. It represents the structure
) for one full bay, beginning at mid-bay. They include all of the load carrying structural members, but do 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 these linear models.
All analysis used static application of loads, increased to account for dyna-mic amplification, where appropriate.
)
7.1.2 Loads Analysis l 1
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 envelops the main frame, gr ating and handrails. Loads on the frame were
)
Technical Report TE WE TR-5310-1 NBRNG SBT\/ ICES 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 l dynamic effect. Loads on the grating were taken from Section 4.3.4 of Refer-ence 1; these loads already include a dynamic factor, since they are based on test data.
7.1.2.2 Pool Swell Fallback (4.3.6)
Pool fallback 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 traximum values of upward load were l used for fallback also. Fallback affects the main frame and grating as well as the handrails.
)
l l 7.1.2.3 Froth Load (4.3.5)
Froth loads have their major effect on the catwalk han& ails; and, when applied horizontally, can produce high bending stresses in the vertical handreil members. Froth loads were calculated in accordance
- with Reference 1, except that the froth 1 influe:.ce 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 for the handrails, the entire catwalk is i submerged before froth loads reach this part of the torus; because of this, froth was only considered on the handrails.
l 7.1.2.4 Drag Loads (Support Columns)
The submerged portion of the cotwalk support col-umns are subject to loading from drag forces from the following sources:
) (a) Pool Swell (b) SRV Discharge
)
l l
I Technical Report W P W NE TR-5310-1 N9tlNG NES 1
(c) Condensation Oscillation I
)
(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 considerir.g the worst case frequency ratio and the fact that this is a harmonic loading.
7.1.2.5 Weight and Seismic Loads Stresses due to weight loads were analyzed using f static analysis and the computer model shown in Figure 7-2. Seismic loads are small and were considered using hand analysis and scaling static stresses.
7.1.3 Results and Evaluation Table 1 allows stresses in the catwalk structure (excluding attachments) to exceed yield; and, in certain cases, to exceed ultimate. Our l
analysis was based on a linear model and all stresses were maintained below
)
the stress at which a plastic hinge would form. Controlling stress and load i
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, Point A in Figure 7-2. It is a result of the combined condition that includes:
Pool Swell + SRV + Seismic + Weight (case 25)
The raximum stress value is:
?
I
)
Technical Report TE NE TR-5310-1 ENGNEERNG SERVICES
) TYPE OF ACTUAL ALLOWABLE STRESS STRESS STRESS Bending + Axial 24,700 psi 40,600 psi
)
7.1.3.2 Support Columns, Support Diagonal Braces & End Joints The controlling load case for the support system
)
and end joints includes:
SRV + Seismic + Weight (case 3)
) Resulting stresses are:
TYPE OF STRESS ACTUAL ALLOWABLE STRESS LOCATION STRESS STRESS Axial + Column to 8,415 psi 12,600 psi Bending Ring Girder 7.1.3.3 Welds to Ring Girder l The controlling load combination for this stress is l also case 25:
I Pool Swell + SRV + Seismic + Weight For this condition, stresses are:
f TYPE OF ACTUAL ALLOWABLE l
l STRESS STRESS STRESS l
l l Shear 8,007 psi 42,000 psi
)
)
Technical Report WTA FrVNE TR-5310-1 ENGNEBRNG SBR \/ ICES 7.2 Internal Spray Header
)
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 vent header (Figure 2-3).
)
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 typical multi-span area as well as at branch connections. This is 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 in the only poci swell related load that is applied.
) The motion of the ring girder that results from pool swell l
loads on the shell was considered but judged to be a negligible input to the spray header. Shel! displacement at the nozzle connections was input to the computer analysis.
l 7.2.2.1 Froth Load (4.3.5)
Froth loads on the spray header were calculated as j outlined in Reference 1. The worst si. cess condition existed for a vertically applied load. The loads were applieu statically to the system (DLF = 1.3).
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 loads and displacments statically.
)
Technical Report W F W NE TR-5310-1 -100- ENGNEN SBR \/ ICES 7.2.3 Results and Evaluation l
The controlling stress for the spray header piping is a result of load case 19 in Table 1.
)
This case includes:
Froth, Weight, Seismic and Shell Motion
)
The controlli1g stress is:
SPIAY HEADER PIPING
)
STRESS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS Mid-bay Bending 22,784 psi 32,880 psi ATTACHMENTS TO RING GIRDER STRESS STRESS ACTUAL ALLOWABLE
) LOCATION TYPE STRESS STRESS Support Hold Bearing 17,298 psi 34,200 psi Down Plate WELDS TO RING GIRDER STRESS STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS
.)
At Ring Tension + 1,868 psi 18,000 psi Girder Shear
)
)
)
Technical Report TF WE TR-5310-1 -101- ENGNEERING SERVICES 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 the convolutions are non-linear, highly cross-coupled modes that would not be
) predicted by ordinary structural codes.
Our approach to the bellows evaluation is to compare the maximum calculated difference in dynamic response (displacement) 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
)
I
)
Technical Report SPTA WNE l TR-5310-1 -102- ENGNEERNG SERVICES i
) computer model of the non-vent bay and therefore did not account for the l presence of the vent pipe hole, or the heavy shell reenforcement in that area.
) 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.) .062 .625 All calculated values are less than 11% of the manu-f acturer'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).
)
7.4 Vacuum Breaker Penetration Reinforcement The vent header penetration for the drywell-to-wetwell vacuum breakers required modification as a result of load case 19 of Table 1.
)
l
)
Technical Report YE NE TR-5310-1 -103- ENGNEERING SERVICES h
This case includes:
Pool Swell + Seismic + Weight PENETRATION STRESS l
l STRESS ACTUAL ALLOWABLE TYPE STRESS STRESS h
Primary + Secondary 46,013 psi 57,900 psi The modification was to add three stiffener plates to each pene-tration as shown in Figure 2-18.
7.5 Electrical Juntion Canister
) The electrical canisters (2) near the top of the wetwell required modification as a result of load case 19 of Table 1. The canisters are subject to Level D service limits which allows for yielding of the material.
I The stress reported is compared to yield.
This case includes:
Pool Swell (froth) + Seismic + Weight l
MAXIMUM CANISTER STRESS 1
STRESS ACTUAL ALLOWABLE LOCATION TYPE STRESS STRESS k
Canister Support Arm Bending 46,000 psi 46,000 psi (yield) l The modification support arm and ring to each penetration is shown in Figure 2-19.
)
)
Technical Report WM TR-5310-1 -104- M
)
1
)
l RING GIRDER VENT BAY ATTACHMENT k
) POINT QNON-VENT BAY l
?
)
RING GIRDER l ATTACHMENT
)
POINT SHELL ATTACHMENT
) POINT
)
l
) FIG.7-1 UNMODIFIE D C ATWALK COMPUTER MODE L-PILGRIM
IEU!8*l "* ^ -los- "
)
l l
)
VENT BAY l RING GIRDER l ATTACHMENT l POINT l
)
QNON-VENT BAY
)
)
RING GIRDER ATTACHMENT
) POINTS SHELL ATTACHMENT POINT
)
F I G . 7- 2
)
MODIFIED CATWA LK COMPUTER MODEL - PILGRIM
)
R- 310-1 -106-i
?
sp*'\ON A TR ::x 7:l g REACTOR T +I 8 ,
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SUPPORT (TYP. at 16 PL ACES)
)
FIG.7-3 h
INTERNAL SPRAY HE ADER 360* COMPUTER MODEL-PILGRIM
5 WTELEDGE Technical Report TR-5310-1 -107- ENCBEERNGSSMCES
)
.062 CALCULATED BELLOWS - =
\, # .03 8
, MOTION
.038 k
r .o62 f
i
% sN 1 sx
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F IG. 7- 4
, M AXIMU M VENT PIPE - TORUS SHE LL R EL ATIVE MO TIO N ZERO-6P POOL SWELL- PILGRIM
)
Technical Report WF WhE TR-5310-1 -108- N 'G N ES
)
8.0 SUPPRESSION P0OL 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
) nixing characteristics of the device and the attendent local-to-bulk tempera-ture difference (6t).
In response to these concerns and to assure reliable operation of these
) devices, the NRC has set limits on maximum pool temperatures for tee-quencher operation, as well as guidelines for a temperature monitoring system for the suppression pool. These requirements are stated in NUREG 0661 (Reference 2) and NUREG 0783.
)
8.1 Maximum Suppression Pool Temperature NUREG 0783 presents maximum pool temperature limits for tee-
) quencher operation at different flow rates, and for several different plant conditions. Evaluation of the Pilgrim Plant for these conditions was done by General Electric Company under contract to Boston Edison. The results of that work are reported here.
)
The local pool temperature limits for the Pilgrim Plant and in accordance with NUREG 0783 are given in Figure 8-1. These correspond to a minimum quencher submergence of 7.0 feet.
)
General Electric evaluated the Pilgrim Plant for the following seven conditions that bound those specified in NUREG 0783. This analysis was performed using al initial pool temperature of 90 F.
)
)
)
Technical Report WP WNE TR-5310-1 -109-N N N ES
) 1A Stuck-open SRV during power operation with one RHR loop available.
l IB Stuck-open SRV during power operation assuming reactor isolation l
due to MSIV closure.
)
1
! 2A Isolation / scram and manual depressurization with one RHR loop available.
h 2B Isolation / scram and manual depressurization with the failure of an SRV to reclose (SORV).
2C Isolation / scram and manual depressurization with two RHR loops
) available. This case demonstrates the pool temperature responses when an isolation / scram event occurs under normal power operation, i.e., when all systems are operating in normal mode.
)
3A Small-break accident (SBA) with manual depressurization; accident mode with one RHR loop available.
3B Small-break accident (SBA) with manual depressurization and f ailure of the shutdown cooling system.
The result of the G.E. analysis showed that the pool temperature remained below the limiting values shown in Figure 8-1 for all seven cases and is therefore acceptable. The results of the analysis for the seven cases are listed in Table 8-1.
8.2 Pool Temperature Monitoring System The NRC criteria also presents guidelines for a monitoring system to constantly monitor pool temperature. A monitoring system will be installed at Pilgrim which uses a network of RTDs set in thermowells in the torus wall, hard wired to a display console in the control room. The system is described l more fully in Section 2.2.1 of this report and is illustrated in Figures 2-9, l 2-17 and 2-20.
l l
l
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, , . - - - -v TABLE 8-1 NSir RESULT
SUMMARY
OF PILGRIM POOL TEMPERATURE RESPONSES g;;
.E Number Maximum Maximum Maximum Local [
of SRVs Cooldown Bulk Pool Pool Q Case Manually Rate Tempe5 ture Tempe5ture No. Event Opened ( F/hr) ( F) ( F)
I 1A SORV at Power, 1 RHR Loop 0 880* 148 182 18 SORV at Power, Spurious 0 710 152 195 I Isolation, 2 RHR Loops l 2A Rapid Depressurization at 4 1000 157 196 ,,
Isolated Hot Shutdown, p 1 RHR Loop l 28 SORV at Isolated Hot 0 710 140 182 Shutdown, 2 RHR Loops 2C Normal Depressurization 4 100 147 176 at Isolated Hot Shutdown, 2 RHR Loops
- 3A SBA-Accident Mode, 4 (ADS)** 4400 156 190 g l 1 RHR Loop 3B SBA-Failure of Shutdown 4 100 148 174 Cooling Mode, 2 RHR Loops 1
- When the main condenser becomes available.
l
Technical Report "RT I RVNE TR-5310-1 -112- ENGINEERING SERVICES
> REFERENCES
- 1. G.E. Report NED0-21888, Rev. 2, " Mark 1 Containment Program Load Defi-nition Report", dated November 1981.
D
- 2. NRC " Safety Evaluation Report, Mark 1 Containment Long-Term Program",
NUREG 0661, dated July 1980.
> 3. G.E. Report NEDO-24583-1 " Mark 1 Containment Program Structural Accept-ance Criteria Plant Unique Analysis Application Guide" dated October 1979.
> 4. G.E Report NED0-21944 " . . . Scale 2-D Plant Unique Pool Swell Test Report" dated August 1979.
- 5. G.E. Report NEDO-24615 "....k Scale Suppression Pool Swell Test Program:
> Supplemental Plant unique Test", dated June 1980.
- 6. G.E. Report NEDE-24840 " Mark 1 Containment Program - Evaluation of Har-monic Phasing for Mark 1 Torus Shell Condensation Oscillation Loads"
> October 1980.
- 7. G.E. Report NEDE-24519-P " Mark 1 Torus Program Seismic Slosh Evaluation" dated March 1978.
- 8. G.E. Report NEDE-21968 " Analysis of Vent Pipe - Ring Header Inter-section" dated April 1979.
> 9. G.E. Report NEDE-24555P " Mark 1 Containment Program - Application Guides 1-6, r, 10 - Vcl. 6 and 10, Rev. 3, others are Rev. 2".
- 10. G.E. Report NED0-24565, Rev. 2, " Mark 1 Containment Program - Plant Unique Load Definition - Pilgrim Station: Unit 1," dated May 1982.
r
Technical Report YE WE TR-5310-1 -113- . NN NES REFERENCES (CONTINUED)
)
- 11. ASME B&PV Code,Section III, Division 1, through Summer 1977.
- 12. Structural Mechanics Report SMA-12101.05-R001, "Dasign Approach for
)
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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_ i
Technical Report "M TF1FrT(NE TR-5310-1 -115- ENGINEERING SERVICES TABLE 2 f PLANT PHYSICAL DIMENSIONS PILGRIM TORUS b
Inner Diameter 29'6" Number of Sections 16 Shell Plate Thickness I
Vent Pipe Penetration 1.125" Top Half .568" Bottom Half .629" i
SUPPORT COLUMNS Quantity Size Outer 16 I-Beam (12.5" x 1.25" Flange & 10" x 1.125 Web)
Inner 16 I-Beam (12.5" x 1.25" Flange & 10" x 1.125 Web)
Base Assembly Slidir.;
RING GIRDER P
Quantity 16 Size T-Beam (7" x 1.5" Flange, 1.5" x 20" (Average) Web) g EARTHQUAKE RESTRAINT SYSTEM Quantity 4 Type Support Saddles
> DRYWELL VENT SYSTEM Quantity Size Vent Pipe 8 6'9" I.D.
p Vacuum Breakers (Internal) 10 18" I.D.
Vent Header Support Columns 16 pairs 6 Sch. 80 Downcomers 96 24" Minimum Submergence 3'0"
)
Technical Report TF WE TR-5310-1 -116- M M]CES TABLE 3 PLANT ANALYSIS INFORMATION PILGRIM Seismic Acceleration Values (G's) b OBE SSE Vertical .06 .10 Horizontal .08 .15 Effective Water Mass for Horizontal Seismic Load (Reference 7) 25.3%
Effective Water Mass during Pool Swell Uplift (Reference 4)
Full o P - 50%
p Zero esp - 30%
Plant Unique C.0. Multiplier (Reference 1) p .917 b
x .
, n, W 4 Technical Report YEDE\-
TR-5310-1 -Al-1 6 MM APPENDIX 1 *
) : . ,
Use of SRV In-Plant Test Data for Analysis
. . s Y
Test Data "'
) -
The in-plant SRV tests used to support structural analysis were run it-Pilgrim in August, 1980. The data was collected in a series of four tests, each consisting of one actuation with a cold line and a second abcut cne
) minute later (hot line). The test sets were about three hours apart'to allow for SRV line cool down.
The torus shell was instrumented with a combination of strain and pres ~
) sure 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), the SRV quencher (Figure Al-3) and an attached piping system (Figure 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, 84 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 cbjectives.
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.
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Technir.al Report TB-5 0-1 II EE i Al-2 JNG4EBW4G SERVICES 4*
Use of Data - Shell & Support System Analysis Ev'aluallon of shell stress and support system loads due to SRV actuat%nl
, was done with a large detailed computer model as discussed in para. 3.2.4 of '
t5e report. Data collected from the in-plant tests was used to define thf
) -
actual shell pressures and decay time for a benchmark (test) condition and to develop. correction factors between these measured resu,1ts and values pre-dfctedbygenericanalyticalmethods. The steps involved are these:
)
- 1. Determine maximum average '; hell pressure, averuge frequency and a
waveform for the four cold tests.
- 2. Calculate these same quantities for the test conditions using the
)
generic computer progr.mb (Q9UBS 02). '
s ,.
- 3. CalcJlate calibration f actors relating predicted-to-actual pres-
)
sure and predicted-to-actual frequer.cy.
- 4. Calculate predicted pressures and frequencies using the generic l.Gmputer program, for other SRV conditions.
)
- 5. Apply the calibration f actors calculated in step (3) to all other predictipns for pressure and frequency. The duration of the pres-sure tro.^.sient, as measured in the test, is affected proportionally
) by the frequency correction and used as the basis for all computer model lo,tding.
t i'arification of Co,IpJter Model
)
The test data was also used to verif y the accu acy of the computer model.
Tnis was done by tho follcwing method:
%)
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) Technical Report TN TR-5310-1 Al-3 66
)
- 1. The computer model was loaded with the measured shell pressures.
- 2. The model was run and stresses at all strain gage locations were calculated.
- 3. Comparisons were made between computer predicted shell stress and I 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 f actor for supports only, and combining it with the previous pressure calibration factor. The results were two different cali-bration f actors to be applied to final analysis - one for the shell and one for the columns. The factors developed and used are:
Shell pressure = .62 x predicted l Support load = .4 x predicted
)
Multiple Valve Contributions 1
For c3ses where more than one valve actuates, the contributions from j 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
)
In the period after the Pilgrim SRV test and May, 1982, TES ran in-plant l SRV tests in four plants, and collected SRV drag information on submerged structures in accordance with the following table:
1 r
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Technical Report TN TR-5310-1 Al-4 NMES
)
Catwalk Vent Ring Supports Column Downcomer Girder j Millstone X X X Nine Mile Point X X Vermont Yankee X X Fitzpatrick X X X X
)
Data collected in these tests was evaluated, analyzed and used to develop loads for SRV drag on submerged structures for these four plants, as well as for Pilgrim. (The Pilgrim in-plant SRV test pre-dated this period, at a time
) when the very conservative nature of calculated SRV drag loads had not been established).
During these tests, strains were measured on the structures indicated
) (except for the ring girder, which was a pressure measurement). The strain gages were positioned to show bending stress due to the combined effect of SRV jet and bubble drag. Figures Al-5 and Al-6 show typical test instrumentation on these structures.
)
Evaluation of the test data for the four plants showed these important results:
) 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 - generally less than one-tenth of predicted loads.
An important consideration in the application of this data was the possi-bility that resonant structural response might occur at some other SRV condi-
) tion. This was considered and dismissed based on two separate arguments; they are:
)
)
Technical Report "RTF1 FrWNE TR-5310-1 Al-5 ENGINEERING SERVICES
- 1. If a major frequency component existed in the drag force, it would be detectable on each of the structural responses for a given test.
This did not occur.
> 2. The response frequencies of the structures tested (structural natu-ral frequencies) ranged from 8.1 to 38 Hz.* If any single strong frequency existed in the drag load, one of the structural responses should have demonstrated some degree of resonant response - none I did.
We conclude from this that the structures involved are responding to a f airly 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 I
stresses measured in the test, when applied uniformly to the submerged area.
These static pressure values were plotted against distance from the quencher and Figure Al-7 was developed. This curve represents the equivalent static drag pressures, including quencher jet loads. It is scaled upward from test I conditions to more severe SRV cases by the ratio of the calculated shell pressures for the two cases, for application to structures under different loading conditions.
> *ActuaT values were E1, 8.2,14.5,15, 21, 23, .1, 25, 29, 30, 34 and 38 Hz.
b
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) Technical Report WTERME TR-5310-1 Al-6 N
-*44"
- O Rl8 e PRESSURE P8 P7 O STRAIN ROSETTE-F-
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- BOTH SIDES OF
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E I SHELL w
3 14 8'
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--- 24'4 -- 20b
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qs Rl3 O
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f LOWER HALF OF BAY D
)
FIG. Al-l
) SRV TEST INSTRUMENTATION- PILGRIM SHELL GAGES
Technical Re ort Y TR-5310-1 Al-7
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)
RING GIRDER RING GIRDER h r
- \ /
l \
/
l \ '
b <
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t /
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t
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IB SADDLES NOT COMPLETE u .
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FIG. A l-2
)
SRV TEST INSTRU MENTATION - PILGRIM COLUMN GAGES '
Tk530 Al-8 l
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t l
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SRV LINE o
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t 4 G AGES 2 AT 9 0 * - c /
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) QUENCHER SUPPORT 4 G AGES AT 90*
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l STR AIN G AGES
)
FIG . Al-3
) SRV TEST INSTRUMENTATION - PILGRIM SRV QUENCHER 8 SUPPORT
__________u__________. .______u____ _ _ _ _ _ _ _
,__ u .___
SHELL h m' '
CBA ]
ACCE LEROMET"ER TORUS SHELL WITHIN I"OF WEL
.A BAY BETWEEN Da B.
h A
_. e a
C' NB
.C LINE WlO.202 83 A = VERTICAL B = LONGITUDIN AL C = TRANSVERSE 1
i FIG. Al- 4 SRV TEST INSTRUMENTATION - PILGRIM ATTACHED PIPING
T Ias.auffWi
) Technical Report MM TR-5310-1 Al-10 i
o o
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N}+ i q >
l N n:. - -:.- ~
l B d 1 -
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i STRAIN GAGE b
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)
)
FIG. A l-5 SRV TEST INSTRUMENTATION-TYPIC AL
)
INTERNAL STRUCTURES
) Technical Report 1P TR-5310-1 Al-11 M 4 STRAIN GAGES AT 90" f
A
)
A
)
'N /
)
+-+
Df)^ l C\ B
) SECTION A-A
)
SRV TEST INSTRUMENTATION-TYPICAL DOWNCOMER FIG. Al-6
)
Technical Report WTELEME TR-5310-1 Al-12 EMBEBW4GN
)
)
4 l
)
c SRV CASE Al .1
$" NM
) $
D
$2 VY w
9 f $
) U NM o 4 g
W MS O 2 4 6 8
) HORIZONTAL DISTANCE FROM BUBBLE (f t)
)
)
EQUIVALENT SRV DRAG FROM IN PLANT TESTING FIG A l-7
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) Technical Report "WTF1 FrT(NE TR-5310-1 A2-1 ENGINEERING SERVICES
) APPENDIX 2 l
Discuss _ ion of 32 Hz Frequency Cut-off for Condensation Oscillation and Post Chug Analysis l
l TES made the decision to limit C0 and Post Chug response analysis to b
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 - l PSD study of the C0 pressure data showed that frequencies above M Hz accounted f 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 D
cut-off and our system is highly responsive at low frequencies (not flat), we should expect a much smaller error.
- 2. Structural Response _A_nalysis - The relative importance of loads below and above 32 Hz car 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:
D
Technical Report 1% P W NE TR-5310-1 A2-2 ENGNEERING SERVICES
) Max.
Numbe 20f* Numbe 20f Number of GX CX Value Frequencies 1000 > 000 GXJ2 Below 32 Hz 44 25 14 167,858
)
32-50 Hz 34 5 1 4,594 1
- 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 f requency 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.
)
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WP WNE 1 A3-1 9JGNEstlNG SBNICES APPENDIX 3
)
C0/CH Drag Loads for Ring Girder Analysis
)
TES did not follow the calculational methods of NUREG 0661 (Reference 2) for calculation of 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 \/ 2 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.
Application of the NUREG criteria produces a factor of 4 multiplier for
)
drag force for flat plate structures in the fluid; and a factor of 8 multi-plier for flat plate structures in tne 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 f actor of 4 higher than the 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.
)
l
)
wTm mVNE h [1 1 A3-2 NBilNG SBR \/ ICES Both References 1 and 2 present results as a function of the VT/D ratio I where:
V = maximum velocity
)
T = period of flow oscillation D = diameter
)
Keolegan and Carpenter show the effective hydrodynamic mass coefficient foraplatevariesfromamaximumof4atf=125to1atVT/0=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 will exist in the torus. In fact, this is not true for CO and CH drag loads on the ring girder. For this structure, under this load, VT D
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 CO and CH drag loads on the ring girder.
) The TES method to calculate these drag loads on the ring girder used the sam >_ references as above (A3-1 and A3-2), but accounted for calculated values of f rather than the values corresponding to the maximum increases. Consid-
) eration of the actual DVI ratio for wall interference led to an interference factor of 1.65 (instead of 2).
)
l Technical Report '#PF WE TR-5310-1 A3-3 ENGNEERING SERVICES Low values of hsuggest that the theoretical hydrodynamic mass coefficient for the ring girder is appropriate. The theoretical coefficient '
~
l for this structure is estimated by an equivalent cylinder with a radius equal i to the circumscribing radius. Use of this cylinder results in a hydrodynamic l
mass coefficient equal to two. The total f actor used was related to the NUREG multiplier by:
2.0
- 1.65 * '41 4.0 2.0 The factor ed by TES was .41 x the NUREG 0661 factor.
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Technical Report i%' E D E TR-5310-1 A3-4 N MCES l
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, REFERENCES 1
l h 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 Oscillating Fluid", Journal of Fluids Engineering, September 1976.
)
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