ML20039E899
| ML20039E899 | |
| Person / Time | |
|---|---|
| Site: | Black Fox |
| Issue date: | 01/31/1980 |
| From: | Ianni P, Reuter F, Stancavage P GENERAL ELECTRIC CO. |
| To: | |
| Shared Package | |
| ML20039E857 | List: |
| References | |
| 22A4365, 22A4365-R04, 22A4365-R4, NUDOCS 8201110597 | |
| Download: ML20039E899 (400) | |
Text
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!U i '82 JAN -7 A10:02 i 0FF:'E ~ 3E1 '- DCCKE!J31 !! P ' f!F: AliCH i CON *AIhWENT LOADS REPORT (CLR) l MARK III CONTAINMENT i i i I i 1 em Approved: Approved: M GwCC M id - f F. Reuter, Manager P. P. Stancavage, Manager j Mark III Containment Design Containment Engineering i l 1 Approved: < . % P. W. Ianni, Manager Containment Design i I. 1 NUCLE AR ENERGY ENGINEERING DIVISION e GENERAL ELECTRIC COMPANY SAN JOSE, CALIFORNI A 95125 O 8 '" ' " ^ ' O ' 8'" ' 8 1 011880 8201110597 820106 gDRADOCK 05000556 PDR
f i l ) 22A4365 g Rev. 2 i l A DISCiAIMER CF RESPONSIBILITY This document is being made available by General Electric Company uithout consideration in the interest of promoting the spread of technical knouledge, neither Ceneral Electric Company nor the individual authors: A. Make any carranty or representation, expressed or implied, uith respect to the accuracy, completeness, or usefulness of the information contained in this document, or that the use of any information disclosed in this document may not infringe privately ouned rights;
- 3. Assume any responsibility for liability or damage chich may result from the use of any information disclosed in this document; or C. Imply that a plant designed in accordance uith the recommenda-tions found in this document vill be licensed by the United States Nuclear Regulatory Comission or that it will comply uith Federal, State or local regulations.
O 042178
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22A4365 Rev. 4 111 O's - TABLE OF CONTENTS P_ age
- 1. INTRODUCTION 1-1 1.1 Confirmatory Testing 1-2 1.2 Definition of LOCA 1-2 1.3 Design Margins 1-3 2-1
- 2. REVIEW OF PHENOMENA 2.1 Design Basis Accident (DBA) 2-1 2.2 Intermediate Break Accident (lBA) 2-5 2.3 Small Break Accident (SBA) 2-5 2.4 Safety Relief Valve Actuation 2-7 2.5 Other Considerations 2-8
- 3. DYNAMIC LOAD TABLE 3-1
- 4. DRYWELL STRUCTURE 4-1 4-1
/N 4.1 Drywell Loads During a Large Break Accident V 4.1.1 Sonic Wave 4-1 4.1.2 Drywell Pressure 4-1 4.1.3 Hydrostatic Pressure 4-2 4.1.4 Loads on the Drywell Wall During Pool Swell 4-2 4.1.5 Condensation oscillation Loads 4-3 4.1.6 Fall Back Loads 4-3a Negative Load During ECCS Flooding 4-4
- 4.1.7 4.1.8 Chugging 4-4 4.1.9 Loads Due To Chugging 4-5 4.1.9.1 " hugging Loads Applied to Top Venc 4-Sa 4.1.9.2 Pool Boundary Chugging Loads 4-Sa
. 4.2 Drywell Loads During Intermediate Break Accident 4-6 4.3 Drywell During a Small Break Accident 4-6 4.3.1 Drywell Temperature 4-6 4.3.2 Drywell Pressure 4-7 4.3.3 Chugging 4-8 Safety Relief Valve Actuation 4-8 4.4 4-8 () 4.5 4.6 Drywell Environmental Envelope Top Vent Temperature (Cycling Profile During Chugging 4-8 l , 4.7 Drywell Multicell Effects 4-8a l 011880 l
22A4365 iv R v. 4 TABLE OF CONTENTS (Continued) Page 5-1
- 5. REIR WALL Weir Wall Loads During a Design Basis Accident 5-1 5.1 5-1 5.1.1 Sonic Wave Outward Load During Vent Clearing 5-1 5.1.2 5-1 5.1.3 Outward Load Due to Vent Flow 5-1 5.1.4 Chugging Loads Inward Load Due to Negative Drywell Pressure 5-3 5.1.5 5-4 5.1.6 Suppression Pool Fallback Loads 5-4 f 5.1.7 Hydrostatic Pressure 5-4 5.1. 8 Safety Relief Valve Loads 5-4 5.1.9 Condensation Weir Wall Loads During An intermediate Break Accident 5-4 5.2 Weir Wall Loads During a Small Break Accident 5-5 5.3 5-5 5.4 Weir Wall Environment Envelope 5-5 5.5 Weir Annulus Multicell Effects 6-1
- 6. CONTAINMENT Containment Loads During a Large Steam Line Break (DBA) 6-1 O 6.1 6-1 6.1.1 Compressive Wave Loading 6-1 6.1.2 Water Jet Loads 6-2 6.1.3 initial Bubble Pressure 6-2 6.1. 4 Hydrostatic Pressure 6.1. 5 Local Containment Loads Resulting from the 6-3 Structures at or Near the Pool Surface 6.1.6 Containment Load Due to Pool Swell at the HCU Floor 6-3 Wetwell Pressurization 6-4 6.1.7 Fall Back Loads 6-4 6.1.8 Post Pool Swell Wave 6-5 6.1.9 Condensation Oscillation Loads 6-5 6.1.10 Chugging 6-5 6.1.11 Long-Term Transient 6-6 6.1.12 Containment Environmental Envelope 6-6 6.2 Containment Loads During an Intermediate Break Accident 6-6 6.3 Containment Loads During a Small Break Accident 6-7 6.4 Safety Relief Valve Loads 6-7 6.5 Suppression Pool Thermal Stratification 6-7 6.6 Containment Wall Multicell Effects 011880
22A4365 R v. 2 v TABLE OF CONTENTS (Continued)
./D b Page 7-1
- 7. SUPPRESSION POOL BASEMAT LOADS 8-1
- 8. LOADS ON STRUCTURES IN THE SUPPRESSION POOL Design Basis Accident 8-1 8.1 8.1.1 Vent Clearing Jet Load 8-1 8.1.2 Drywell Bubble Pressure and Drag Loads Due to 8-1 Pool Swell 8.1. 3 Fall Back Loads 8-2 8- 2 8.1.4 Condensation Loads 8.1.5 Chugging 8-2 Compressive Wave Loading 8-3 8.1. 6 Safety Relief Valve Actuation 8- 3' .
8.1. 7 9-1
- 9. LOADS ON STRUCTURES AT THE POOL SURFACE 10-1 l 10. LOADS ON STRUCTURES BETWEEN THE POOL SURFACE AND THE HCU FLOORS 10.1 Impact Loads 10-1 10.2 Drag' Loads 10-3 10.3 Fall Back Loads 10-3 11-1
- 11. LOADS ON EXPANSIVE STRUCTURES AT THE HCU FLOOR ELEVATION i
12-1
- 12. LOADS ON SMALL STRUCTURES AT AND ABOVE THE HCU FLOOR ELEVATION 1
R-1 REFERENCES A-1 ATTACHMENT A - SAFETY RELIEF VALVE LOADS (QUENCHER) B-1 ATTACHMENT B - SUPPRESSION POOL SEISMIC INDUCED LOADS C-1 ATTACHMENT C - WEIR ANNULUS BLOCKAGE D-1 ATTACHMENT D - DRYWELL PRESSURE DISTRIBUTION 042178
I 22A4355 j Rev. 4 vi TABLE OF CONTENTS (Continued) { Page ATTACHMENT E - DRYWELL NEGATIVE PRESSURE CALCULATIONS E-1 ATTACIEENT F - WETWELL ASYMIETRIC PRESSURES F-1 ATTACICENT G - SUBMERGED STRUCTURE LOADS DUE TO LOCA AND SRV ACUATIONS G-1 ATTACIDENT H - WEIR WALL LOADS DURING DRYWELL DEPRESSURIZATION H-1 ATTAGLMENT I - POOL SWELL VELOCITY I-l ATTACIDENT J - SCALING ANALYSES AND SMALL STRUCTraE POOL SWELL DYNAMIC LOADS J-l ATTACHMENT K - DELETED K-1 ATTAGUENT L - CONTAINMENT ASY1CETRIC LOADS L-1 l ATTACHMENT M - MULTIPLE SAFETY / RELIEF VALVE ACTUATION FORCING I FUNCTION lETHODS M-1 l 1 l ATTACHMENT N - SUPPRESSION POOL THERMAL STRATIFICATION N-1 ATTACHMENT 0 - DIGITIZATION OF FORCING FUNCTION FOR CONDENSATION OSCILLATION 0-1 0 011880
22A4365 vif Rev. 4 I~ LIST OF ILLUSTRATIONS C). Title Pa_ge Figure Loss-of-Coolant Accident Chronology (DBA) 2-9 2.1 2-10 2.2 1 Schematic of the Mark III Pool Swell Phenomenon 2-11 2.2-2 Typical Suppression Pool Cross Section 238 Plant 2-12 2.2-3 Plan at Elevation 11 ft 0 in. 2-13 2.2-4 Containment Floor Drain Sump 238 Plant 2-14 2.2-5 Containment Equipment Drain Sump 238 Plant 2-15 2.2-6 Plan at Elevation (-) 5 ft 3 in. 2.3 Idealized Quencher Bubble Pressure Oscillation in 2-16 Suppression Pool 4-9 4.1 Drywell-Loading Chart for DBA 4-10 4.2 Drywell-Loading Chart for SBA 4-11 4.3 Drywell-Loading Chart for IBA 4.4 Short Term Drywell and Containment Pressure Response to 4-12 a Large Steam Line Break (DBA) 4-13 4.5 PSTF Test Results - Vent Static Pressure Differential O
\,,/ 4.5a PSTF Test Results - Vent Static Pressure Differential 4-14 4.6 Typical Drywell Pressure Traces During Condensation, Run 23, 4-15 Test 5807 4.6a Condensation Oscillation Load Spatial Distribution on ,
4-16 Drywell Wall, Containment and Basemat 4.6b Condensation Oscillation Forcing Function on the Drywell 4-17 Wall 0.D. Adjacent to the Top Vent 4-18 4.7 Typical Top Vent Pressure Trace During Chugging, Run 19 Peak Pressure Pulse Train in Top Vent During Chugging 4-19 4.7a Peak Force Pulse Train in Top Vent During Chugging 4-19a 4.7b 4-19b 4.7c Average Force Pulse Train in Top Vent During Chugging 4-19e 4.7d Average Pressure Pulse Train in Top Vent During Chugging 4.8 Typical Containment Pressure Trace During Chugging, Run 11 4-20 (Ref. Test 5707) 4.8a Typical Pressure Time-History on the Pool Boundary During 4-21 Chugging 4.8b Suppression Pool Chugging Normalized Peak Underpressure 4-21a Attenuation 4.8c Suppression Pcol Chugging Normalized Mean Underpressure 4-21b Attenuatior, [} 011880
22A4365 Rev. 4 viii LIST OF ILLUSTRATIONS (Continued) ll Figure Title Page 6.3d Suppression Pool Chugging Normalized Spike Attenuation 4-21c 4.8e Suppression Pool Chugging Spike Duration "d" as a Function of Location in the Pool 4-21d 4.8f Suppression Pool Chugging Normalized Peak Post Chug Oscillations 4-21e 4.8g Circumferential Underpressure Amplitude Attenuation 4-21f 4.8h Circumferential Post Chug Oscillation Amplitude Attenuation 4-21g 4.81 Suppression Pool Chugging Normalized Post Chug Oscillation Attenuation 4-11h l 4.8j Chugging Pressure Time-History on the Drywell Wall Adjacent to Vert 4-211 l 4.9 Drywell - Containment Pressure Differential During Chugging 4-22 4.10 Maximum Design Drywell Atmosphere Bulk Temperature and Pressure Envelope 4-24 4.11 Drywell Top Vent Cyclic Temperature Profile and Area of Application During Chugging 4-25 5.1 Weir Wall-Loading Chart for DBA 5-6 5.2 Weir Wall-Loading Chart for SBA 5-7 llh 5.3 Weir Wall-Loading Chart for IBA 5-8 5.4 Typical Weir Wall Chugging Time History - Test Series 5707, Run 1 5-9 5.4a Typical Pressure Time-History for Weir Annulus During 5-10 Chugging 5.5 Underpressure Distribution on the Weir Wall and Drywell I.D. Wall During Chugging 5-10a 5.5a Peak Pressure Pulse Train on the Weir Wall and Drywell I.D. Wall During Chugging 5-10b 5.5b Mean Pressure Pulse Train on the Weir Wall and Drywell I.D. Wall During Chugging 5-11 5.6 Normalized Weir Annulus Pressure Pulse Attenuation 5-11 5.7 Pressure Transient in Drywell Initiated by Vessel Reflood Line Break Level 238 5-12 5.8 Vent Backflow Weir Annulus Water Surge Velocity Vs. Height Above Weir Wall 5-13 5.9 Mean Pressure Pulse Train on the Weir Wall and Drywell I.D. Wall During Chugging With Superposition of Adjacent Spikes 5-14 6.1 Containment-Loading Chart for DBA 6-8 6.2 Containment-Loading Chart for IBA 6-9 6.3 Containment-Loading Chart for SBA 6-10 6.4 Observed Bubble Pressure During Pool Swell Test Series 5706, Run 4 6-11 011880
22A4365 R v. 4 ix LIST OF ILLUSTRATIONS (Continued)
&' Page Figure Title 6.5 Dynamic Loads Associated with Initial Bubble Formation in the Pool 6-12 6.6 Containment Pressure Differential During Bubble Formation 6-13 6.7 Deleted 6-14 Drag Loads on Protruding Structures Due to Pool Swell 6-15 6.8 Containment Loading Due to Flow AP Across HCU Floor 6-16 6.9 6.10 Typical Containment Wall and Basemat Pressure Traces During ,
Condensation, Run 23 (Ref. Test 5807) 6-17 Containment Wall and Basemat Pressure Time Histories, Test 6-18 6.11 5807, Run 11 Containment Wall Chugging Pressure Time History, Test 5707, 6-19 6.12 Run 9 6.13 Basemat Chugging Pressure Time History, Test 5707, Run 9 6-20 6.14 Calculated Atmosphere Bulk Temperature and Pressure Design Envelope for all Breaks 6-21 6.15 Long Term Containment Pressure Following a DBA 6-22 O' 6.16 HCU Floor AP vs HCU Floor Open Area 6-23 6.17 Suppression Pool Temperature Profile for the Large Breaks 6-24 7.1 Pool Boundary Loads During Bubble Formation 7-2 8.1 Structures Within' Suppression Pool-Loading Chart for DBA 8-4 l 8.2 Deleted 8-5 P3 Deleted 8-6 9.1 Structures at the Pool Surface-Loading Chart During DBA 9-2 10.1 Small Structures Between the Pool Surface and the HCU Floor-Loading Chart During DBA 10-4 10.2 Profile of Impact Loads on Small Structures Within 18 ft of the Pool Surface 10-5 10.3 Pressure Drop Due to Flow Across Grating Within 18 ft of the Pool Surface 10-6 10.4 Drag Load on Solid Structures within 18 ft of the Pool Surface 10-7 10.5 Drag Loads for Various Geometries (slug flow) 10-8 O O 011880 l l l l,
22A4365 Rev. 2 x LIST OF ILLUSTRATIONS (Continued) Title Page Figure 10-6 Su= mary of Pool Svell Loading Specifications for Small Structures in the Containment Annulus (Not Applicable to the Steam Tunnel or Expansive HCU Floors) 10-9 11-1 Expansive Structures at HCU Elevation - Loading 11-3 During DBA 12.1 Small Structures at the HCU Floor Elevation - Loading Chart During DBA 12-3 12.2 Loads at HCU Floor Elevation Due to Pool Swell Froth 12-4 Impact and Two-Phase Flow l I l O O 101678 1
22A4365 i Rsv. 3 xi/xii j l 4 l LIST OF TABLES t i Table Title g ! i. 1.1.1 Summary of PSTF Tests 1-4 f l 1.3.1 Summary of Specified and Realistic Design Values 1-6 l i 1 ! 3.1.1 Sumary of Postulated Accidents Affecting Mark III 1 Structures 3-1 , i '
; 4.1 Chugging Loads 4-8b i
i i 1 i 1 4 e 1 4 i 1 i i l i i l i 4 iO ! l I i i y i 1 l 5 i k i i I j. O { l u 1 t j 090779 , I 3
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J. 22A6365 Rev. 4 xiii/xiv ABSTRACT d This techn%at report provides numerical infomation for thermal hydraulic dynamic toading conditions in GE Mark III Reference Plant 1 pressure suppression containment eyetan during a loss-of-coolant } accident, safety relief valve discharge and related dynamic events. Information and guidance has been provided to assist the contain-i ment designer in evaluatir.g the design conditions for the various ' j structures which fom the containment system. Confimatory tests are completed. Observed test data or calculations upon which the toads are based, are discussed. A Class III supplement to this report (22A436SAB) includes additional proprietary informtion in support of the load definition. i\ 1 I i i i 1
- o 011880
22A4365 R v, 4 1-1 e (_ / 1. INTRODUCTION The information in this document represents the General Electric Company recommendation for containment loads. Loss of coolant accident (LOCA) and safety relief valve (SRV) dynamic loads, based upon pressure suppression and safety relief valve test programs, are included. General Electric has con-cluded the confirmatory test program for the Mark III containment configuration. These tests support and confirm the pressure suppression loads that result from the postulated LOCA and from SRV operation. The confirmatory program includes a series of scaled multivent tests that demonstrate no significant vent interaction effects for the LOCA prccess. The Caorso tests, also included, demonstrate the conservative SRV loads. General Electric will use the design load values specified herein as the basis for the 238 GESSAR license application. Other load values or smaller margins than those provided in this document may be used if the architect engineer is willing to defend them through the licensing process. O The architect engineer is responsible for the definition of load combinations, which include loads of the type described in this document, as well as more normal loads such as deadweight, seismic, wind, missile impact, jet impinge-ment, etc. The architect engineer is also responsible for determining the effect of the above loads and load combinations on the structures and equip-ment. Thus, the architect engineer is responsible for project unique con-tainment analyses.
/
During a loss-of-coolant accident and events such as safety / relief valve actuation, the structures forming the containment system and other structures within the Reactor Building experience dynamic phenomena. This report provides numerical information on the dynamic loads that these phenomena impose on the Mark III containment system structures. l The loading information is based on either observed test data or conservatively-l calculated peak values. The LOCA loading combinations are presented in the form of bar charts for each of the containment system structures. In addition [} to defining the timing of the LOCA related loads, the bar charts identify l i l 011880. _ - __ s
22A4365 Rev. 4 1-2 other loading conditions such as seismic accelerations, dead-weight, etc. For each bar on the chart, reference is made to the section where specific discus-sion of the load is presented. To provide a better understanding of the various dynamic loads and their inter-relationships, Section 2 contains a qualitative description of sequen-tial events for a wide range of postulated accidents. The air clearing loading phenomena associated with the actuation of a safety relief valve is also described. 1.1 CONFIRMATORY TESTING Impact and impingement load specifications for small structures affected by suppression pool swell, are based on the results of the PSTF air tests con-ducted in March 1974 and reported in Reference 9. The intent of these tests was to provide conservative design data. It was recognized that the data base would require extension beyond that provided by the air tests and to achieve this, additional impact tests for both small and large structures were included in the PSTF schedule. These tests involved measurement of pool swell impact forces on a variety of targets representative of small structures found in the Mark III containment annulus, and are discussed in Attachment J. This document relies on a large experimental test data base from the PSTF program. See Tabic 1.1.1 for a summary of these tests. The scaling of the large scale and 1/3 area scale PSTF precludes direct application to the proto-type Mark III. Conservative interpretation of these tests results, employing dimensional similitude scaling relationship, is used to arrive at specified design loads for Mark III. (See Attachment J.) Evaluation cf full scale Caorso SRV tests is included in Attachment A. The evaluation shows that the test result loads are significantly lower than the current design loads and the use of reduced design loads are justified. 9 011880
22A4365' 1-3 R;v. 4 l 12 DEFINITION OF LOCA A loss-of-coolant accident (LOCA) is the sudden break of a high energy pipe in the reactor coolant pressure boundary of the nuclear steam supply system. The largest postulated break could be either the break of a main steam or a recirculation line. This loss-of-coolant accident (LOCA) is the design 9 basis accident (DBA). Other small line breaks result in loss-of-coolant in large accidents, and although their energy release does not result dynamic loadings, their thermal effects may control the design of structures. The. intermediate break accident (IBA) and small break accident (SBA) fall into this category. The size of the SBA is defined as that which will not cause automatic depressurization of the reactor. The SBA is of concern because it imposes the most severe temperature condition inside the drywell. j 1.3 DESIGN MARGINS structures. l Table 1.3.1 summarizes the loads due to a LOCA for the containment 4 Reasonable design margins are clearly shown by comparing the magnitude of the values between the conservatively specified design values and the realistic should be expected loads. The Mark III loads presented in this document
. interpreted as rigid wall loads. A similar case for showing the conservatism in the loads specified for relief valve aceux* ion is given in Attachment A.
It is shown in this report that t'# V 11 dynamic loading phenomena has been t a is conservatively interpreted. conservatively bounded and the Pb ' ca. 4 Parameter simulation has justified the application of the test data to MK III f designs with adequate design conservatism added. Any further margin considera-tions cannot be technically envisioned. In fact, where possible, the contain-l , ment designer may chose to justify more realistic design v31ues. 4 O 011880
l Table 1.1.1 l SUM!!ARY OF PSTF TESTS Area Number Venturi Top Ventq, Initial Number Pool / Re fe r-Test of Range Submergence Pressure Blowdown of Vent Prima ry ence Series Blowdowns (inch) Range (feet) (psia) Type Vents Scaling Objectives
- Report 5701 21 21/8-35/8 2.0 - 15.5 1050 Saturated 1 Full 1. Vent Clearing 4 8'*""
- 2. Full Scale Condensation Demo
- 3. Drywell Pressure 5702 17 2 1/8-3 5/8 1.93 - 11.97 1050 Saturated 2 Full 1. Vent Clearing 4 Steam 5703 3 21/2-35/8 6.77 - 11.05 1050 Saturated 3 Full 1. Vent Clearing 4 .
Steam ny 5705 4 1 - 4 1/4 6.0 - 8.0 1065 Air 2 Full 1. Pool Swell 7 Scoping 5706 7 4 1/4 6.0 - 10.0 1065 Air 2 Full 1. Pool Swell 7
- 2. lapact Loading 5707 22 2 1/8 - 3 7.5 1050 Air and 3 Full 1. Chugging 16 Steam .
5801 19 21/8--3 5.0 - 10.0 1050 Saturated 3 1/3 1. 1/ 3 Scale 11 Steam Demonstration
- 2. Pool Swell
- 3. Roof Density
$ and AP 5
M 5802 3 2 1/8-3 6.0 1050 Saturated 3 1/3 1. Pool Swell 11 [ Steam
# 9 e
O - O O Table 1.1.1 (Continued) Area Number Pool / Re fe r-Number Venturi Top Venty, Initial Primary ence Submergence Pressure Blowdown of Vent Test of Range Objectives
- Report Range (feet) (psia) Type Vents Scaling Series Blowdowns (inch) 1/3 1.1/3 Scale Demo 11 ,
5803 2 21/8-3 5.0 - 7.5 1050 Saturated 3 Liquid
- 2. Liquid Blowdown 1/3 1. Roof Density 11 2 1/8-3 5.0 1050 Saturated 3 5804 5 Density and AP Steam Repeatability 5.0 - 10.0 1050 Saturated 3 1/3 1. Pool Swell 12 5805 52 1-3 Impact Fg Steam *
.n 2 1/2-4 1/4 5.0 - 7.5 1065 Air 3 1/3 1. Pool Swell 13 w" 5806 12 3 1/3 1. Steam 15 20 1-3 7. 5 1050 Saturat 5807 Staas aus Condensation Liquid 1050 Steam 9 1/9 1. Pool Swell 17 l
6002 14 2-1/8 - 3 5 - 10 Multivent l Effects 1/9 1. Steam 18 7.5 1050 Steam 9 6003 12 2-1/2 Condensation Multivent Effects
*In general tests are not direct prototype simulations, but parametric studies to be used in analytic c) g nodel evaluations. T wi O
e
- _. . . . . . . . . . . . . . . . . . . . . . .q
Table 1.3.1
SUMMARY
OF SPECIFIED AND REALISTIC DESIGN VALUES Specified "" "
""'E" for Engr'g load Design Estimate Analysis Test Section Connents STRUCTURE: Drywell BREAK SIZE: Large Drywell Pressurization 30 psig 18 psig Model 4.1.2 Peak calculated 21.8 psig (Ref. 1) plus margin !!ydrostatic Pressure pil pil Standard 4.1.3 analytical techniques Bubble Formation 0 + 21.8 psid 18 psid Max pressure 4.1.4 equal D.W.
pressure E
< >I$
Hetwell Pressurization 11 psid 3-5 psid Model in 5801, 5802 12.1 Test shows pressure *O Supplement 1 5803, 5804 differential in the "O to Ref. 1 3 to 5 psi range Pool Swell Slug impact load 115 psi 60 psi 5706, 5801 12.1 Applies to small flat 5802, 5805 structures attached to D.W. (see Fig. 10.6). See Attachment J. Froth impingement load 15 psi 15 psi 5706, 5801 Applies to small struc-5802, 5805 tures attached to D.W. (see Fig. 10.6). See Attachment J. Velocity for computing 40 ft/sec 30 ft/sec Bounding 9.0 See Attachment I drag loads (slug flow) calculation 10.2 Condensation 17 psid 14 psid 5702, 5703 4.1.5 See Fig. 4.6.a for S Oscillation Loads (mean) 5801, 5807 pressure distribution o 2j Fallback Velocity for 35 ft/sec 20 ft/sec Bounding 4.1.6 7
- calculation Drag Loads O O e - - - - -
O O O Table 1.3.1 (Continued) SeP led B is Estimate Analysis Test Section Comments Load Design
-15 psid Bounding 4.1.7 Assumes no vacuum relief Negative 1,oad During -21 psid ECCS Flooding calculation Chugging 2 psid il psid 5801, 5802 4.1.8 Design pressures are Cross structure +30 psig and -21 psid 5803, 5804 I.oading within top 4.1.9.1 vent e Pre-chug under- -15.0 psid -12 psid 5707 4.1.9.2 pressure (peak) (peak)
-9.0 psid -8 psid (mean) (mean) py e Pulse (spike) 540 psid 500 psid 5707 Local and global pulse *O (peak) (peak) train specified W$
214 psid 180 psid (mean) (mean) 250 kips 5707 Local and global net e Net force 250 kips (peak) (peak) upward vertical load 91 kips 75 kips (mean) (mean) 5.1.4 Same as weir wall I,oading on drywell I.D. specification 5707 4.1.9.2 See Table 4.1 for dura-Loading on drywell 0.D. tion and frequency e Pre-chug under- -5.8 psid -4.0 psid pressure (peak) (peak) I
-2.65 psid -1.0 psid o (mean) (mean) e g S e Pulse (spike) 100 psid 75 psid (peak) (peak) 4 3
- 24 psid 20 psid (mean) (mean)
Table 1.3.1 (Continued) E"' '" Design Basis for Engr,g 1.oad Design Estimate Analysis Test Section Comments e Post-chug 16.5 psid 14.0 psid oscillation (peak) (peak) 12.2 psid 21.1 psid (mean) (mean) STRUCTURE: Drywell BREAK SIZE: Intermediate ADS 4.2 See Attachment A Chuggirg 4.1.8- Same as large break 4.1.9.2 specification STRUCTURE: Drywell BREAK SIZE: Small 330*F/310*F 330*F/ Hounding 4.3.1 3 hr at 330*F initially, Temperature 310*F calculation n e).t 3 hr at 310*F %'I$ Chugging 4.1.8- Same as large break *$ 4.1.9.2 specification " 8l = sJ h' 03
- O e -- - - - - - -
O O r L. Table 1.3.1 (Continued) Specified Design Basis Engr'g Analysis Test Section Comments Load Design Estimate STRUCTURE: Weir Wall BREAK SIZE: Large** Outward Load Due to Vent 10 psig 5 psig Model in 5.1.2 First 30 sec of blowdown Ref. 1 5.1.3 Clearing 5707 5.1.4 Local and global loading Chugging specified e Pre-chug under- -2.15 psid -2.0 psid pressure (peak) (peak)
-0.98 psid -0.5 psid (mean) (mean) e Peak spike of pulse 43 psid 35 psid pp train (peak) (peak) @p 13 psid 15 psid (mean)
*O w as (mean)
- 12,800 lbr/ 5.1.5 Attachment 11 Inward Load Due to 8000 Negative Drywell vent (top lbg vent) Bounding calculation Pressure Differential 6000 lbr (mid) 4000 lbf (bot tom) pli Standard 5.1.7 liydrostatic Pressure pli analytical techniques STRUCTURE: Weir Wall BREAK SIZE: Intermediate **
Attachment A ADS STRUCTURE: Weir Wall BREAK SIZE: Small** Bounding 5.4 330*F for 3 hr initially Temperature 330*F/310*F 310*F for next 3 hr calculation o o
.U
** Chugging Loads on Weir Wall are the same for large, intermediate and sin 11 break accidents. {
i
Table 1.3.1 (Continued) P**I'I Design Bas's, for E.ngr,g Estimate Analysis Test Section Comments Load Design STRUCTURE: Containment BREAK SIZE: Large 0 psig Attachment G 5706 6.1.2 Measured pressure is small Water Jet <1 psig and is obscured by bubble P#""""#* 8 psid 5701, 5702 6.1.3 Bubble Formation 10 psid 5703, 5705 5706 liydrostatic Pressure pli pil Standard 6.1.4 analytical techniques 8 psid D.W. bubble 6.1.5 Only large structures Pool Swell Loads for 10 psid see bubble pressure :o w Attached Structures (bubble) pressure at Pool Surface 40 ft/sec 30 ft/sec Bounding 6.1.5 See Attachment I (drag calculation u$
- velocity) 5706 6.1.6 Pool Swell at ilCU Floor 15 psi (froth 10 psi impingement) 11 psi 3-5 psi Model in 5801, 5802 6.1.6 Test shows pressure (flow AP) Ref. 1 5803, 5804 differential in the 3 to 5 psi range Fallback Velocity for 35 ft/sec 20 ft/sec Bounding 6.1.7 Drag Loads calculation 2ft PSTF Tests 6.1.8 Negligible load Post Pool Swell Waves 2 ft il psid 10.6 5807, 5701 6.1.9 See Figure 4.6a l Condensation (meaa) 5702 Oscillation Loads Chugging 5707 4.1.9.2 See Table 4.1 for dura-tion and frequency 3 e Pre-chug-under- -1.3 psid -0.8 psid (peak) (peak) 7 pressure O
% -1.0 psid -0.3 psid lE (mean) (mean)
- O e
O E () (_) k- \m/ Table 1.3.1 (Continued) pecified Design Basis for Engr,g Load Design Estimate Analysis Test Section Comments e Pulse (spike) 3.0 psid 2.2 psid (peak) (peak) 0.7 psid 0.6 psid (mean) (mean) e Post-chug il.7 psid ii.5 psid oscillation (peak) (peak) 11.0 psid 10.5 psid (mean) (mean) Pressurization 15 psig 5 psig Model 6.1.11 Peak calculated value (Ref. 1) is 9.8 psig plus margin Temperature 185*F* <150*F Supplement 1 6.1.11 Conservative calculated f 5' S$ to Ref. 1 peak temperature is 176*F IN w STRUCTURE: Containment BREAK SIZE: Inte rmediate #'$ Pressurization 15 psig 5 psig Bounding 6.2 l calculation ADS 6.2 See Attachment A Chugging 4.1.8- Same as large break 4.1.9.2 specification STRUCTURE: Containment BREAK SIZE: Small Temperature 220*F 185*F Bounding 6.3 Local temperatures of Stratification (Dome) calculation 300/250*F are possible in the event of reactor steam /11guld blowdowns to containment. Pressure 2 psig 1 psig Bounding 6.3 Typical value calculation o U Chugging 4.1.8- Same as large break r. d. O 4.1.9.2 specification pa
*See paragraph 6.1.11 1
rable 1.3.1 (Continued) E#S' Design Basis for E,ngr,g Analysis Test Section Comments Load Design Estimate STRUCTURE: Basemat BREAK SIZE: Large liydrostatic pil pil Standard analytical techniques 10 4 21.8 psid 18 psid Peak equal to 5706/4 7.0 10 psi over 1/2 pool Bubble Formation assumed to increase D.W. pressure linearly to 21.8 psi. See Figs. 7.1 and b.6 11.7 psid 11.0 psid 5807, 5702 7.0 See Figure 4.6a Condensation 5701 Oscillation Load 5707 4.1.9.2 See Table 4.1 for dura- ,y Chugging tion and frequency Qy
. e-
-1.8 psid -1.5 psid See Figures 4.8b wy
- Pre-chug under. *
(peak) through 4.8f for pacec"ra (ncak)
. l
-1.34 psid -0.7 psid basemat attenuation (mean) (mean) e Pulse (spike) 10 psid 7.5 psid (peak) (peak) 2.4 psid 2 psid (mean) (mean) e Fost-chug 12.1 psid 12.0 psid oscillation (peak) (peak) 11.3 psid 11.0 psid (mean) (mean)
STRUCTURE: Basemat BREAK SIZE: Intermediate 7.0 See Attachment A ADS 4.1.9.2 Same as large braak c) Chugging specification g. g a u g; y STRUCTURE: Basemat BREAK SIZE: Small 4.1.9.2 Same as large break Chugg d specification
# 9 e
l pO . LJ L) Table 1.3.1 (Continued) Specified Design Basis Load Design Estimate Analysis Test Section Comments STRUCTURE: Submerged BREAK SIZE: Large* Structures LOCA Water Jet Loads G2.2 Load is bounded by LOCA air bubble load LOCA Air Bubble Load 8.2 psid* Attachment G G2.3 Load is on a sample l structure 4 ft from the top vent axis Velocity for Computing 40 ft/sec 30 ft/sec Bounding See Attachment I Drag Loads Calculation Fall Back Velocity for 35 ft/sec 20 ft/sec Bounding G2.4 Drag Loads Calculation LOCA Condensation 0.7 psid Attachment G G2.5 Frequency 2.+3.5 Itz - w Oscillation Loads Load is on a sample l w$ $? structure 4 ft from the *C vent exit *$ LOCA Chugging Loads 1.9 psid Attachment G G2.6 Load is on a sample structure 4 ft from the top vent exit Negligible Attachment G G 3.1 Load is negligible outside X-Quencher Water Jet a sphere circumscribed by Load the quencher arms 0.5 psid Attachment G G3.2 1,oad is on a sample X-Quencher Air Bubble structure 9 ft from Load Quencher center STRUCTURE: Submerged BREAK SIZE: I nt e rmed ia t e
- Structures See Attachment G ADS S STRUCTURE: Suhmerged BREAK SIZE: Small*
$ Structures b $ No Additional loads generated u
- Chugging loads are the same for large, intermediate and small break accidents.
Table 1.3.1 (Continued) Specified Design Basis for Engr,g Design Es t itaa t e Analysis Test Section Comments Load STRUCTURE: Structures BREAK SIZE: Large at Pool Surface Bubble Formation Drywell 21.8 psid 18 psi Equal to D.W. 9.0 Large structures only pressure Containment 10.0 psid Attenuated D.W. pressure Velocity for Computing 40 ft/sec 30 ft/sec Bounding 9.0 Drag Loads calculation Fallback Velocity for 35 ft/sec 20 ft/sec Bounding 4.1.6 fU Drag Loads calculation 4$u STRUCTURE: Structures BREAK SIZE: Intermediate "$ at Pool Surface ADS See Attachment A STRUCTURE: Structures BREAK SIZE: Small at Pool Surface No additional loads generated (See large break tabulation) 8 8 7 0 %
- O e
.-.~ - ..-. .- .- .. . . - . - . - . _ _ . - - . - - . - --
O O Table 1.3.1 (Continued) o pc led esign Basis Engr'g Comments Design Estimate Analysis Test Section Load STRD.?e URE: Structures BREAK SIZE: Large 4 Between Pool Surface and i ilCU Floor Slug Impact , Loads 60 psi 5801, 5802 10.1 See Attachment J Small flat structures 115 psi 5805, 5706 30 psi 5801, 5802 10.1 See Attachment J Piping 60 psi 5805, 5706 15 psi 5706 10.1 See t.ttachment J and Froth Impingement Loads 15 psi Figure'10.6 gg 10.2 See Attachment I. See .# $ 40 ft/sec 30 ft/sec Bounding Velocity for Computing calculation Figure 10.3 for grating 9$* Drag Loads loads 20 ft/sec Bounding 10.3 Fallback Velocity for 35 ft/sec calculation Drag Loads STRUCTURE: Structures BREAK SIZE: Intermediate Between Pool Surface and llCU Floor (See large break No additional loads generated tabulation) STRUCTURE: Structures BREAK SIZE: Small Between Pool Surface and itCU Floor (See large break g No additional loads generated tabulation) s g A. 03
Table 1.3.1 (Continued) SP ecified Design Basis for Engr g Load Design Estimate Analysis Test Section Comments STRUCTURE: Expansive BREAK SIZE: Large Structures at ilCU Floor Elevation Wetwell Pressurization 11 psig 3-5 psig Itodel in 5801, 5802 11.0 (3-4 see) (1-2 sec) Ref. 1 5803, 5804 Froth Impingement 15 psig 10 psig 5801, 5802 11.0 See Attachment J 5805, 5706 discussion l (100 ms) (100 ms) Flow Pressure 11 psig 3-5 psig Model in 5801, 5802 12.0 Test shows pressure Differential Ref. 1 5803, 5804 differential of 3 to 5 psi
- o "
Fallback and Water 1 psi 0.5 psi Bounding 12.0 Based on water flow ay Accumulation calculation through liCU floor
- we*
STRUCTURE: Expansive BREAK SIZE: Intermediate Structures at ilCU Floor Elevation No additional loads generated See large break tabulation STRUCTURE: Expansive BREAK SIZE: Small Structures at itCU Floor Elevation No additional loads generated See large break tabulation GENERAL NOTES TO TABLE 1.3.1 5 1. Where S/R valve loads are specified in the applicable bar charts, refer to Attachment A, s Section AS.6 for margin discussion. T' s 2. Not all loads for IBA and SBA are tabulated. Generally the large break load condition will govern. $' O O O
__ _ . . . _ _ . _ _ . _ _ _ _ _ _ . . _ . _ _ . . _ - . . _ _ . ._ _ . _ _ . _ _ . . - . _ __..._.m _ _ . . . . -_ _ . . _.
.O O O Table 1.3.1 (Continued)
Specified Engr's Design Basis for Design Estimate Analysis Test Section Comments Load STRUCTURE: Small Structures
! at IICU Elevation Froth Impingement 15 psid 10 psid 5801, 5802 12.0 See Attachment J 5805, 5706 discussion 11 psid 3-5 psid Model in 5801, 5802 12.0 Test shows pressure Flow Pressure Ref. 1 5803, 5804 differential of 3 to Differential 5 psi Fallback and Water 1 psid 0.5 psid Bounding 12.0 Based on water flow Calculation through IICU floor Accumulation
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5 4 1 22A4365
- i. Rev. 2 2-1 1
- 2. REVIEW OF PHENOMENA 3
The purpose of this section of the report is to qualitatively review the sequence of events that could occur during the course of the design basis accident (DBA), an intermediate break accident (IBA), a small break accident (SBA) and during i safety relief valve actuation. The objective of this review is to provide an i understanding of the various pool dynamic loads and their inter-relationships, lj and to define the dynamic loading terminology. Specific design load values are i provided in subsequent sections. t 2.1 DESIGN BASIS ACCIDENT (DBA) The Figure 2.1 chart shows the events occurring during a DBA and the potential loading conditions associated with these events. With the instantaneous rupture of a steam or recirculation line a theoretical . sonic wave exits the broken primary system pipe and expands into the drywell i atmosphere. At the break exit point, the wave amplitude theoretically is reactor operating pressure (1000 psia). However, there is rapid attenuation i ' as the wave front expands spherically outward into the drywell at sonic velocity. 1 1 i As the drywell pressure increases, the water initially standing in the vent sys-tem accelerates into the pool and the vents are cleared of water. During this i vent clearing process, the water leaving the horizontal vents forms jets in the suppression pool and causes water jet impingement loads on the structures within During the the suppression pool and on the containment wall opposite the vents. i vent clearing transient, the drywell is subjected to a pressure dif ferential and the weir wall experiences a vent clearing reaction force. I Immediately following vent clearing, an air and steam bubble forms at the exit of the vents. The bubble pressure initially is assumed equal to the current drywell pressure (peak calculated is 21.8 psig). '?his bubble theoretically ! transmits a pressure wave through the suppression pool water and results in 1' loading on the suppression pool boundaries and on equipment located in the i i suppression pool. " 101678
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l 22A4365 Rev. 2 2-2 As air and steam flow from the drywell becomes established in the vent system, the initial vent exit bubble expands to suppression pool hydrostatic pressure. CE Large Scale Pressure Suppression Test Facility (PSTF) Tests (Ref 4) show that the steam fraction of the flow is condensed but continued injection of drywell air and expansion of the air bubble results in a rise in the suppression pool surface. During the early stages of this process, the pool swells in a bulk mode (i.e. , a slug of solid water is accelerated upward by the air). During this phase of pool swell, structures close to the pool surface will experience loads as the rising pool surface impacts the lower surface of the structure. See Figures 2.2-4, 2.2-5, and 2.2-6. In addition to these initial impact loads, these same structures will experience drag loads as water flows past them. Equipment in the suppression pool will also experience drag loads. Data from PSTF air tests (5706) indicates that after the pool surface has risen approximately 1.6 times the initial submergence of the top vent, the water ligament thickness has decreased to two feet or less and the impact loads are significantly reduced. This phase is referred to as incipient breakthrough; i.e. , ligament begins to break up. O Ligament thickness continues to decrease until complete breakthrough is reached and the air bubble can vent to the containment free space. The breakthrough process results in formation of an air / water froth. Continued injection of dr>vell air into the suppression pool results in a period of froth pool swell. This iroth swell impinges on structures it encounters but the two phase nature of the fluid results in loads that are very much less than the impact loads associated with bulk pool swell. When the froth reaches the elevation of the floors on which the Hydraulic Control Units for the Control Rod Drives are located, approximately 20 feet above pool level, the froth encounters a flow restriction; at this elevation, there is approximately 25% open area. See Figures 2.2-2 and 2.2-3. The froth pool swell experiences a two phase pressure drop as it is forced to flow through the available open areas. This pressure differential represents a load on both the floor structures themselves and on the adjacent containment and drywell. The result is a discontinuous pressure loading at this elevation, j l 042178 l
22A4365 Rev. 2 2-3 Figure 2.2-1 is a diagram that summarizes the various phases of pool swell and (~)) the nature of the dynamic loading conditions that occur. It should be empha-sized that the pool swell elevation information presented on Figure 2.2-1 is based on an assessment of the PSTF air tests. As such it is considered con-servative since the PSTF air test data have been interpreted and used in a conservative manner. I The pool swell impact and impingement target data presented in Section 10 of this document is applicable to small structures. This restriction on the application of the impact test data is necessary since the basic tests involved targets with a width of 20 inches. For this size target, only the suppression pool water in the immediate vicinity of the target has to be re-directed by the impact impulse, thus, the impact loads are not dependent upon the pool swell water ligament thickness. Attachment J discussed application of PSTF impact data to small structures. For floors that are expansive enough to decelerate a large sector of the pool g-s) rather than a small region of the pool in the vicinity of the target, the impul-
'wd sive loading on the floor is dependent upon the momentum of the entire slug and is related to slug thickness.
As drywell air flow through the horizontal vent system decreases and the air / water suppression pool mixture experiences gravity induced phase separation, pool upward movement stops and the " fallback" process starts. During this process, floors and other flat structures experience downward loading and the containment wall theoretically can be subjected to a small pressure increase. However, this pressure increase has not been observed experimentally. The pool swell transient associated with drywell air venting to the pool typically lasts 3 to 5 seconds. Following this, there is a long period of high steam flow rate through the vent system; data indicates that this steam will be entirely condensed in a region right at the vent exits. For the DBA reactor blowdown, steam condensation lasts for a period of approximately one minute. Potential structural loadings during the steam condensation phase of the accident have been observed, are relatively small, and are included in the containment loading specification. i i 042178
22A4365 Rev. 3 2-4 As the reactor blowdown proceeds the primary system is depleted of high energy fluid inventory and the steam flow rate to the vent system decreases. This reduced steam flow rate leads to a reduction in the drywell/ containment pressure dif ferential which in turn results in a sequential recovering of the horizontal vents. Suppression pool recovering of a particular vent row occurs when the vent stagnation dif ferential pressure corresponds to the suppression pool hydro-etatic pressure at the row of vents. Toward the end of the reactor blowdown, the top row of vents is capable of con-densing the reduced blowdown flow and the two lower rows will be totally recovered. As the blowdown steam flow decreases to very low values, the water in the top row of vents starts to oscillate back and forth causing what has become known as vent " chugging." This action results in dynamic loads on the top vents and on the weir wall opposite the upper rcw of vents. In addition an oscillatory pressure loading condition can occur on the drywell and conta .. ment, but is insignificant. Since this phenomenon is steam mass flux dependent (the chugging threshold appears to be in the range of 10 lb/ sec/ft ) it is present for all break sizes. For smaller breaks, it is the only mode of condensation that the vent system will experience. h Shortly af ter a DBA, the Emergency Core Cooling System (ECCS) pumps automatically start up and pump condensate water and/or suppression pool water into the reactor pressure vessel. This water floods the reactor core and then starts to cascade into the drywell from the break; the time at which this occurs depends upon break size and location. Secause the drywell is full of steam at the time of vessel flooding, the sudden introduction of cool water causes rapid steam condensation and drywell depressurization. When the drywell pressure falls below the containment pressure, the drywell vacuum relief system is activated and air from the containment enters the drywell. Eventually sufficient air returns to equalize the drywell and containment pressures; however, during this drywell depressurization transient, there is a period of negative pressure on the drywell structure; a conservative negative load condition is therefore speci-fied for drywell de;, sign. Following vessel flooding and drywell/ containment pressure equalization, sup-pression pool water is continuously recirculated through the core by the ECCS 090779 1
22A4365 Rsv. 4 l 2-5 1 pumps. The energy associated with the core decay heat results in a slow heat up of the suppression pool. To control suppression pool temperature, operators will activate the RHR heat exchangers. After several hours, the heat exchangers control and limit the suppression pool temperature increase. The suppression pool is conservatively calculated to reach a peak temperature of 176*F and with f
' long term containment spray operation the peak temperature can approach 185'F The increase in air and water vapor pressure at these temperatures results in a ; pressure loading of the containment.
1 The post DBA containment heatup and pressurization transient is terminated when the RHR heat exchangers reduce the pool temperature and containment pressure to 1 nominal values. t 1 2.2 INTERMEDIATE BREAK ACCIDENT (IBA)
, An intermediate size break is defined as a break that is less than the DBA but is
! of suf ficient magnitude to automatically depressurize the primary system due to loss of fluid and/or automatic initiation of the ECCS systeus. In practice, this 2 means liquid breaks greater than 0.05 f t 2 and steam breaks greater than 0.4 ft () i as determined by analysis. 4 In general, the magnitude of dynamic loading conditions associated with a loss 1.' of coolant accident decrease with decreasing break size. However, the inter-mediate break is examined because the Automatic Depressurization System (ADS) may be involved. Simultaneous actuation of the multiple safety / relief valves committed to this system introduces significant containment system loads, as discussed in Section 2.4.
- 2. 3 SMALL BREAK ACCIDENT (SBA)
Small breaks are defined as breaks not large enough to automatically depressur-ize the reactor. Accident termination is dependent upon operator action and the duration of the accident is determined by operator response. The dynamic loads produced by this class of accident are small. However, there are certain con-ditions associated with smaller reactor system breaks that must be considered during the design process. 'specifically, the drywell and weir wall must be f() 011880 1
22A4365 Rev. 2 2-6 d; signed for the thermal loading conditions that can be generated by a small s team break (SBA) . For a definition of the design conditions, the following s:quence of events is postulated. With the reactor and containment operating at maximum normal conditions, a small break occurs allowing blowdown of reactor steam to the drywell. The resulting dryvell pressure increase leads to a high drywell pressure signal that scrams the reactor and activates the containment isolation system. Drywell pressure continues to increase at a rate dependent on the size of the assumed steam laak. This pressure increase to 3 psig depresses the water level in the weir annulus until the level reaches the top of the upper row of vents. At this time, air and steam enter the suppression pool. Steam is condensed and the air passes to the containment free space. The latter results in gradual pressuriza-tion of the containment at a rate dependent upon the air carryover. Eventually, entrainment of the drywell air in the steam flow through the vents results in all the drywell air being carried over to the containment. At this time, containment pressurization ceases. The drywell is now full of steam and has a positive pressure differential sufficient to keep the weir annulus water level depressed to the top vents and chugging can occur. Continued reactor blowdown steam is condensed in the suppression pool. The thermodynamic process associated with blowdown of primary system fluid is one of constant enthalpy. If the primary system break is below the RPV water level, blowdown flow consists of reactor water. Upon depressurizing from reactor pressure to drywell pressure, approximately one-third of this water flashes to steam, two-thirds remain as liquid, and both phases will be in a saturated condition at drywell pressure. Thus, if the drywell is at atmos-pheric pressure, the steam-and-liquid blowdown will have a tec:perature of 212*F. If the primary system rupture is located so that the blowdown flow consists of reactor steam, the resultant steam temperature in the drywell is significantly higher than the saturated temperature associated with liquid blowdown. This is because a constant enthalpy decompression of high pressure saturated steam results in a superheat condition. For example, deco =pression of 1,000 psia saturated steam to atmospheric pressure results in 298*F superheated steam (86*F of superhcat). 042178
22A4365 Rev. 4 2-7 Reactor operators are alerted to the SBA incident by the leak detection system, (~ N-
)
or high drywell-pressure signal, and reactor scram. For the degraded accident evaluation, rapid depressurization is assumed to be manually initiated at 10 minutes to terminate the event. For the purpose of evaluating the duration i of the superheat condition in the drywell, it is assumed that operator response to the small break is to shut the reactor down in an orderly manner using selected relief valves and with the RHR heat exchangers controlling the suppres-sion pool temperature. (This assumes the main condenser is not available and the operators must use the suppression pool for an energy sink. In all probabil-ity, the condenser would be available and the suppression pool would not be involved in the shutdown.) Reactor cooldown rate is assumed to be started 30 minutes af ter the break and at 100*F/hr. Using these procedures, leads to a reactor cool-down in approximately three to six hours. At that time, the RHR system (in the shutdown mode) maintains the reactor at 212*F or less and the blowdown flow rate is terminated. It should be noted that the end-of-blowdown chugging phenomenon discussed in Section 2.1 will also occur during a small break accident and will last the duration of reactor depressurization. 2.4 SAFETY RELIEF VALVE ACTUATION In addition to loads on the valves and discharge piping, actuation of the safety / relief (S/R) valves causes pressure disturbances in the suppression pool water which results in dynamic loads on the suppression pool floor, the weir wall, the drywell and the containment adjacent tc the pool. Structures in the pool also experience this loading. Relief valve actuation can be initiated either automatically by a reactor pressure increase to the valve setpoints or by an active system such as ADS. The phenomena which cause these loads is as follows. Prior to actuation, the S/R discharge lines contain air at atmospheric pressure and a column of water in the submerged section. Following S/R valve actuation, the pressure builds up inside the piping and expels the water column. The air follows the water through the holes in the quencher arms and forms a large number of small bubbles. Once in the pool, the bubbles expand, coalesce and form four large bubbles. Each of the four bubbles expands analogous to a spring and accelerates the sur-g rounding pool of water. The momentum of the accelerated water causes the bubble Nl to over-expand and the bubble pressure becomes negative. This negative pressure slows down and finally reverses the motion of the water leading to compression 011880
22A4365 Rev. 4 2-8 of the bubble. This sequence of expansion and contraction is repeated with a m:an frequency of about 8 Hz until the bubble reaches the pool surface. The bubble oscillation causes oscillating pressures throughout the pool. The magnitude of the pressure amplitude decreases with time and with distance from the bubble. The duration of this load is less than 1 second (See Figure 2.3). SRV steam condensation during valve discharge to the pool also occurs. This phenomenon results in low amplitude pressure fluctuations. V In evaluating the Mark III structural loads and containment /drywell capability it is necessary to properly account for the hypothetical accident related loads and their sequence of occurrence. In defining the loads for this evaluation, this report addresses the design basis accident (pipe break) and the loads a:sociated with the hypothetical concurrent earthquake, pool dynamics, and static loading. The ability of the design to accommodate these loadings, when properly sequenced, constitutes the design basis of the structure. This design basis includes the single failure criterion; i.e., any single component may lh fail to act when celled upon during loss-of-coolant accidents This report also addresses an additional consideration namely the inadvertent opening of a single S/R valve. The opening of a single valve is not a direct result of the LOCA and, furthermore, is not an expected occurrence during the accident sequence. However, the loading chart figures show the loads ae*;ciated with a single safety /rel2ef valve actuation as an additional load for demon-strating additional capability. Safety relief valve loading data is discussed in Attachment A. 2.5 OTHER CONSIDERATIONS In addition to the LOCA and S/R valve dynamic loads that have been identified in the preceding sections, other loads must be considered during the design process (deadweight, seismic accelerations, etc.) These loads are included in the loading diagrams contained in this report. lll 011880
22A4365 R;v. 2 2,9 EVENT POTENTI AL LOACING CONOf fiON
- e COMPRESSIVE W AVE
(~~s LOAOING ON CONTAINMENT (' / LOCA OCCURS m ORvWELL PRESSURE RISES
- SONIC W AVE LOAOiNG OF ORYWELL If e JET IMPINGEMENT ANO SUBBLE VENTS CLEAR ANO VENT m PRESSURE LOADS ON THE Airl/ STEAM FLOW STARTS CONT AINM E N T e VENT CLEARING ANO VENT FLOW AP ON DAvWELL e OUTYpARO F LOW AP ON WEIR WALL 1I e IMPACT LOADS ON LOW POOL SWE LLS IN m STAUCTUAES A BULK MODE e OR AG LOAOS ON STAUCTURES IN AND A80VE THE POOL If SREAKTHROUGH 1f
[e F ACTH IMPlNCEMENT ON POOL SWE LL CONTINUES IN A
' A ACTH' MOCE AND ENCOUNTE AS HIGH STAUCTUR ES e FLOW AP ON HCU FLOOR FLOW AESTRICTION AT HCU AND AOJACENT CONT AINMENT FLOOR ,
II fn) OAYWELL VENTING - e 'F ALL B ACC LOACS ON COMPLET E STR UCTU A E S If STE AM CONDENSATION 5 e CONOENSATION LOAOS IN PCOL AT VENT ExlTS f m o WEI A W ALL ANO ORYWELL BLOWOOWN ENOS " LO AOS OUE TO CHUGGING If e NEGATIVE PRESSUAE ON WEIR ECCS FLOOCING OF AE ACTDA WA LL OR YWELL AND ITS VESSEL AND CAvWELL 7 P EN ETR A T IONS DEP A ESSUAl2 ATION e NEGATIVE FLOW AP ON WEI A WALL G TEA " e CONT AINMENT PA ES$URE LOAO pO
*ALL POTENTI AL LOCA OvN AMIC LO AOS AAE tOENTirlEO. BUT ALL ARE NOT SIGNIFICANT ISEE TEXT FC A CET AILS)
Figure 2.1. Loss-of-Coolant Accident Chronology (DBA) (O) , 042178
22A4365 Rev. 2 2-10 0 0 0 o 0 0 0 0 0 0 0 D 0 3 0 a 0 Q l SPR AY 0 g 6 NO SIGNIFICANT LOAOS q Q D g L 0 0 0 20 l 0 go a o O h h h FROTH h h LOW IMPtNGEMENT LOADS LOW ORAG LOADS g ea o G g o 4 o G ( 12 f t BULK POOL SWELL - 5$$$-1$5$$$$$$$5$55&-~&^ ^ :-: --
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O -- INITI AL POOL SURF ACE Figure 2.2-1. Schematic of the Mark III Pool Swell Phenomenon 042178
22A4365 Rev. 2 2-11 I v CONTAINMENT SistCLD , SLDG '
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22A4365 2-12 Rev. 2 STE AM TUNNEL
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- O 042178
2-15 22A4365 Rev. 2 l l C) 4 FLOOR DR AIN $ UMP SELOW c ECulp OR AIN ISEE FIGURE 2 2 f syy, ggton t$EE W- '25' I
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22A4365 3-1/3-2 Rev. 2 0 3. DYNAMIC LOAD TABLE I The dynamic loading information for the Mark III containment system is presented in the subsequent sections of this report. The data is presented in bar chart form and shows the temporal distribution of loading sequences for the various structures. At any given time on a bar chart it is assumed that the particular structure being considered experiences all the loading conditions in those
" boxes" which span the J 'iven time unless a specific exception is indicated.
Each chart has applicable loading information references. Table 3.1.1 sum-marizes the accidents that influence the design of various structures. Table 3.1.1 Summary of Postulated Accidents Affecting Mark III Structures (DBA) (IBA) (SBA) Large Intermediate Small Structure Break Break Break V( g Drywell X X X Weir Wall X X X X X X Con tainment Suppression Pool Floor X X - Structures in Suppression Pool X X - Structures at the Suppression X X - Pool Surface Structures Between the Pool X Surface and the HCU Floor Structures at the HCU Floor X - Elevation l Notes: ! \ s' l. X indicates accident with significant loading conditions
- 2. For concurrent S/R valve events, see appropriate bar charts 042178
22A4365 Rev. 3 41 l l l [D 4 DRYWELL STRUCTURE V The drywell structure experiences loads during both the design basis loss-of-coolant accident and during a small steam break accident. Loads occurring during an intermediate break accident are less severe than those associated with the large and small break. The designer should consider other dynamic loads that are not included in this report. These are pipe whip, jet impinge-ment, missile, etc. 4.1 DRYWELL LOADS DURING A LARGE BREAK ACCIDENT Figure 4.1 is the loading bar chart for the drywell structure during a large steam line break. A discussion of the loading conditions follows: 4.1.1 Sonic Wave Theoretically, a sonic compressive wave is initiated in the drywell atmosphere following the postulated instantaneous rupture of a large primary system pipe. O' This phenomenon is not considered in the drywell design conditions on the basis that the finite opening time of a real break coupled with the rapid attenuation with distance and short duration does not produce any significant loading in the drywell. 4.1.2 Drywell Pressure During the vent clearing process, tne drywell reaches a peak calculated differ-ential pressure of 21.8 paid. During the subsequent vent flow phase of the blowdown, the peak pressure differential is less than 21.8 psid due to the wetwell pressurization from the two-phase pool swell flow reaching the contain-l ment annulus restriction at the HCU flocr (see Figure 4.4). This wetwell pressurization is a localized load that acts on the Drywell 0.D. below the HCU floor. Interaction between pool swell and the limited number of structures at or near the pool surf ace does not adversel,y af f ect the drywell pressure. [ 1 Figure 4.4 shows the drywell pressure during the DBA. It inclsdes the HCU () floor pool swell interference effects. The analytical model pre ented in Ref. I was used to calculate these values. 090779
22A4365 4-2 Rev. 2
- lockage of the weir annulus flow area by equipment located above the annulus entrance has the potential for increasing the real dryuell pressure dif feren-tial. Attachment C presents data which show no potential pressure increase for blockages up to 30 percent of the total area.
During the blowdown process, the drywell is subjected to differential pressures between levels because of flow restrictions. This value varies with the size of the restriction, but a bounding value for a 25 percent restriction is 0.5 psi as discussed in Attachment D. On the basis of this bounding calculation, it has been concluded that dif ferential pressures within the drywell during the DBA will be small and as such, need not be included in the drywell loading specifications. 4.1.3 Hydrostatic Pressure During the period of vent flow, the water normally standing in the weir annulus is expelled into the main suppression pool and the icwer regions of the drywell experience an inward load due to the hydrostatic pressure associated with the pool water. If it is assumed that an earthquake is cccurring at this tice, the horizontal and vertical accelarations of the building can influence the hydro-static pressure calculations. See Attachment B. 4.1.4 Loads On The Drvwell Wall During Pool Swell During bubble formation, the outside of the drywell wall in ene pool will be subject to varying pressures. A bounding range of 0 to 21.8 psid is specified en those sections of the dryweil tiall below the suppression :ool surface. The basis for this specification is the knowledge that the minimum pressure increase is O psi and the maximum bubble pressure can never exceed the peak drywell pres-sure of 21.8 psid. Above the ncminal suppression pool surf ace, the pressure linearly decreases from 21.8 psid to O psid over 18.0 feet (see Figure 6.5). Any structures in the containment annulus that are within approximately 20 feet of the initial suppression pool surface will experience upward loads during O 101678
22A4365 i Rev. 3 4-2a I
/~'
pool swell (see Figure 12.2). If these structures are attached to the drywell l wall, then the upward loads will be transmitted into drywell structure. In addition, the region of the drywell below the HCU floors will experience the wetwell pressurization transient during pool swell froth at the HCU floor, as l shown in Figure 4.4 090779
22A4365 Rev. 3 4-3 Sections 9, 10, 11 and 12' discuss applied loads for equipment, floors, etc. that are located in the containment annulus. 4.1.5 condensation Oscillation Loads Following the initial pool swell transient (during a LOCA) when the drywell air l 1s vented to the containment free space, there is a period of 0.05 to 1.5 minutes (depending upon break size and location) when high steam mass flows through the top vents and condensation oscillation occurs. Vent steam mass fluxes of up to 25 lbm/sec/ft occur as a result of either a main steam or recirculation line break. Steam and liquid blowdown tests with various blowdown orifice sizes have 4 been performed in the PSTF facility. Some pressure oscillations have been observed on the drywell wall. Figures 4.5 and 4.Sa give a summary of the magnitude of the top vent exit pressure fluctuations observed during PSTF steam tests. The data has been plotted against vent submergence and is independent of this parameter. O Additional instrumentation was located on the drywell wall above the top vent in PSTF Series 5807. Typical test data traces are shown in Figure 4.6 and show the localized nature of the condensation loads. Maximum pressure amplitude f decreases from approximately !10 psid to approximately 22 psid in two feet. The condensation oscillation forcing function to be used for design is defined as a sumation of four harmonically related sine waves developed from a regres-sion analysis of the data obtained in test series 5807 (Reference 15):
^( { 0.8 sin (2n x T x f(t)) '
P(T) = 2
+ 0.3 sin ( An x T x f(t))
+ 0.15 sin (6n x T x f(t))
+ 0.2 sin (Sn x T x f(t)) ) (psid) 1
- O 090779 i
22A4365 Rav. 4 4-3s where: O P(T) = pressure amplitude (psid) for consecutive cycles beginning at time t = 3 sec.and ending at T p n A(t) = peak-to-peak pressure amplitude variation with time. (psid)
= 5.5 {3.395 - 0.106t + 1.15 log t - 7.987 (log t)2
+ 7.688 (log t)3 - 1.344 (log t)') Eqn (4.2)* l f(t)- = fundamental frequency variation with time. (liz)
= 0.8 (2.495 - 0.225 t - 0.742 log t + 10.514 (log t)2
- 9.271 (log t)3 + 3.208 (log t)') Eqn (4. 3)*
- Log terms shown in Equations 4.2 and 4.3 are log to base 10.
time (sec), 3 1 e 1 30, time from initiation of LOCA blowdown t = t = time increment for successive periods Tpg<t <T p, n Tp = g 3); where n is number of cycles between 3 and 30 sec.
~
p f 3+T p
+T p + ... + T p P(t) from Eqn (4.1) has been calculated for 4 cycles and is shown in Figure 4.6b.
Eqn (4.1) has been calculated and digitized in Attachment 0 of this report. The spatial distribution of the forcing function amplitude over the vetted surface of the suppression pool is shown in Figure 4.6a. The amplitudca shown are the maximum values determined from Eqn (4.1) norreali:ed to 1.0 at the top vont centerline. 4.1.6 Fall Back Loads In general, the data generated in the PSTF indicates that no significant loading conditions on the dryvell wall occur during pool fall back. Figure 6.4 shows that suppression pool wall pressures fo11~<ing bubble breakthrough return to their initial pre-LOCA values during the 1.5 to 5 second period whea the pool level is r.ubsiding. Therefore, fall back pressure loads are not specified for Mark III drywell. 011880
r i 22A4365 4-4 Rev. 3 Structures attached to the drywell wall experience drag loads as the water level subsides to its initial level. These structures could experience drag forces associated with water flowing at 35 f t/sec; typical drag coefficients are shown on Figure 10.5. This is the terminal velocity for a 20 f t. free fall and is a conservative bounding number. 4.1.7 Negative Load During ECCS Flooding Somewhere between 100 and 600 seconds following a LOCA (the time is dependent on break location and size) the ECCS system will refill the reactor pressure vessel. Subsequently, cool suppression pool water will cascade from the break The rapid drywell to the drywell and start condensing the steam in the drywell. depressurization produced by this condensation will draw non-condense.ble gas It is during from the containment free space via the drywell vacuum breakers. this drywell depressurization transient that the maximum drywell negative pressure occurs. However, for design purposes a conservative bounding end point calculation was performed which assumes that drywell depressurization occurs before a significant quantity of air can return to the drywell via the vacuum relief system. This theoretical conservative calculation yields a drywell to containment negative pressure dif ferential of 21 psi (see Attachment E). 4.1.8 Chugging During vent chugging, drywell pressure fluctuations result if significant quantities of suppression pool water are splashed into the drywell when the returning water impacts the weir wall. This can result in a pressure dif- f The f ferential between the drywell and containment as shown in Figure 4.9. maximum values of this load (+2.0, -0.7 paid) are negligible when compared to f the peak positive drywell pressure used for drywell design and the negt.tive e Chugging pressure discussed in Attachment E (Peak Nagative Drywell Fressure). is ar. oscillstory phenomenon having a period of 1 to 5 seconds. The PSTF data shown on Figure 4.9 f rom the 5801, 5802, 5803 and 5804 series of 1/3 scale PSTF tests. The data has been plotted against top vent sub-l' mergence with no obvious correlation. Because volumes and areas of the 1/3 scale tests are correctly scaled, the tests are more appropriate as a source of chugging 090779
\
l 22A4365 4-5 i R v. 4 5701, 5702, and 5703 (~ induced drywell pressure data than large scale tests discussed in Reference 4 The large scale PSTF configuration had a drywell volume to vent area ratio only one-third of either the full scale Mark III or the 1/3 scale PSTF configuration. Drywell pressure variations during chugging result from a combination of fluctuating steam condensation rates at the vent exit and water splashing into the drywell. The undersized dry-well of the large scale PSTF would tend to exaggerate the drywell pressure response. 4.1.9 Loads Due to Chugging In addition to the bulk dryvell pressure fluctuations, high amplitude pressure pulses are' observed when the steam bubbles collapse in the vents during chug-ging. The dominant pressure response to the top vent during chugging is of the pulse train type with the peak amplitude of the pulses varying randomly from chug to chug. The pressure pulse train associated with a chug consists of a sequence of four pulses with exponentially decreasing amplitude as shown in the typical pressure trace in Figure 4.7. The dominant pressure responses in the suppression pool during chugging is characterized by a prechug underpressure, an impulse (pressure spike), and a post chug oscillation as shown in the data trace in Figure 4.8. The chugging process as observed in PSTF tests has a random amplitude and fre-quency. Although it is expected that chugging will occur randomly among the vents, synchronous chugging in all top vents is assumed. Each vent is expected to be periodically exposed to the peak observed pressure spike. The pool bound-ary load definition considers that the chugging loads transmitted to the dry-well wall, weir wall, basemat and contair. ment are the result of several vents chugging aimultaneously at different amplitudes. The potential for asymmetric chugging loads is discussed in Attachment L. a 011880
22A4365 Rev. 3 4-Sa 4.1.9.1 Chugging Loads Applied To Top Vent, h Within the top vent, the peak pressure pulse train shown in Figure 4.7a is applied for local or independent evaluation of vents. Although some variation is observed in the pressure distribution from the top to the bottom of the vent, it is conservatively assumed that during the chugging event the entire top vent wall is simultaneously exposed to spatially uniform pressure pulses. Because some net unbalance in the pressure distribution gives rise to a vertical load, the peak f orce pulse train shown in Figure 4.7b is applied vertically upward over the projected vent area concurrently with the peak pressure pulse train to evaluate local ef f ects at one vent. For global effects, the average force pulse train shown in Figure 4.7c is applied vertically over the projected area of all top ve.nts concurrently with the average pressure pulse train within the vent shown in Figure 4.7d. As can be seen in Figure 4.7, the underpressure preceding the pressure pulse train within the top vent is very small compared to the peak (spike) over-pressure. The mean measured pressure (results from tests) was -9 psid with a standard deviation of 23 psid. On this basis, the specified design value is , l
-15 psid.
4.1.9.2 Pool Boundary Chugging Loads The chugging load applied to the pool boundary (drywell, basemat and contain-ment) is described by the typical forcing function shown in Figure 4.8a. The forcing function consista of a pre-chug underpressure defined as a half sine wave, a triangular pulse (prescu;e spike) loading ct.aracterized by a tine duration "d" and a post-chug oscillation described by a damped sinusoid. The pulse la at its maximum magnitude and duration near the tco vent on tne dryvell wall due to the locali:ed nature- of the phenomena. The amplitude of the pre-chug underpressure and the post-chug oscillation are also maximum at this location. O 090779
22A4365 Rev. 3 4-5b (T For local load' considerations on the pool boundary: LJ 4 e Pre-chug underpressure
- peak amplitude - Table 4.1 distribution - Figure 4.8b e Pulse (spike)
- peak amplitude - Table 4.1 distribution - Figure 4.8d duration-- Figure 4.8e o Post-chug oscillation
- peak amplitude - Table 4.1
- distribution - Figure 4.8f Local chugging loads should be used to evaluate local effects such as pool liner buckling and vent liner stresses. Local chugging loads shall not be Os combined with other loads.
For distribution in the horizontal (circumferential) direction, the pre-chug underpressure attenuates on the drywell, basemat and containment, as shown in Figure 4.8g. The pulse attenuation is the same as the lower portion of the vertical attenuation shown in Figure 4.8d, except that the peak is at the vent centerline, and the post-chug oscillation attenuates on the drywell, basemat and containment, as shown in Figure 4.8h. The profiles in Figures 4.8g and 4.8h represent the peak observed value at one vent, with the other vents chugging at the mean value. For global load considerations on the pool boundary: e Pre-chug underpressure
- mean amplitude - Table 4.1 .
- distribution - Figure 4.8d O
V 090779
22A4365 l Rev. 4 4-6 l l e Pulse (spike)
- menn amplitude - Table 4.1 distribution - Figure 4.8d duration - Figure 4.8e e Post-chug oscillation mean amplitude - Table 4.1 distribution - Figure 4.81 e No horizontal attenuation for this loading Global loads should be used for load combinations and for piping and equipment response calculations.
4.2 DRYWELL LOADS DURING INTERMEDIATE BREAK ACCIDENT The loading conditions caused by an intermediate break are less than those in a DBA or small break; however, they are examined because actuation of the ADS can be involved. (See Attachment A) Figure 4.3 is a bar chart for this condition.
- 4. 3 DRYWELL DURING A SMALL BREAK ACCIDENT l
l A small steam break can lead to high atmospheric temperature conditions in the l d rywell. Figure 4.2 is the bar chart for this accident. 4.3.1 Drywell Temperature For drywell design purposes, it is assumed that the operator reaction to the small break is to initiate a nor=al shutdovn. Under these circumstances, the blowdown of reactor steam can last for a 3 to 6-hour period. The corresponding design temperature is defined by finding the combtnation of primary system pressure and drywell pressure which produces the maximum superheat tenperature. Steam tables show that the maximum drywell steam temperature occurs when the primary system is at approximately 450 psia and the containment pressure is at a maximum. O Oll8PO
22A4365 Rsv. 4 4_7 During an SBA the continuing blowdown of reactor steam will cause all the air
]
v' initially in the drywell to be purged to the containment free space. The peak superheat temperature is 330*F. This temperature condition exists until the RHR shutdown cooling is completed in approximately three hours. At this time, after three hours, the pressure in the reactor pressure vessel is 150 psia and the corresponding superheat temperature is 310*F. This may exist for l three hours. These superheat temperatures correspond to drywell atmosphere only; separate analyses are required to determine transient response of the drywell wall to the elevated steam temperatures. See Section 4.5 for additional environmental information. 4.3.2 Drywell Pressure With the reactor and containment operating at maximum normal conditions, a small break occurs allowing blowdown of reactor steam to the drywell. The resulting drywell pressure increase leads to a high drywell pressure signal that scrams the reactor and activates the containment isolation system. Drywell pressure continues to increase at a rate dependent on the size of the assumed steam leak. This pressure increase to 3 psig depresses the water level in the weir annulus until the level reaches the top of the upper row of vents. At this time , air and steam enter the suppression pool. Steam is condensed and the air passes to the containment free space. The latter results in gradual pressurization of the con tainment at a rate dependent upon the air carryover rate. Eventually, entrainment of the drywell air in the steam flow through the vents results in all drywell air being carried over to the containment. The drywell is now full of steam and a positive pressure differential suf ficient to keep the weir annulus water level depressed to the top vents is maintained. Continued reactor blow-down staam is condensed in the suppression pool. v) 011880
22A4365 Rev. 4 4-8 4.3.3 Chugging 0 During a small break accident there will be chugging in the top vents. Applicable chugging loads on the dryvell and vents are discussed in Sections 4.1.8 and 4.1.9. The Mark III drywell design does not require the combination of the SBA thurmal loading condition with the 21.0 psi negative pressure load. 4.4 SAFETY RELIEF VALVE ACTUATION Relief valve operation can be initiated as a result of either a single failure, ADS operation, or by a rise in reactor pressure to the valve set points. In addition, the drywell can be exposed to S/R valve actuation loads any time the operator elects to open a valve or valves, as during an isolated cooldown. The loads generated by S/R valve actuation are discussed in Attachment A. 4.5 DRYWELL ENVIRONMENTAL ENVELOPE Figure 4.10 shows the design envelope of drywell atmospheric pressures and temperatures f or the spectrum of postulated loss-of-coolant accidents. Figure 4.10 defines only the drywell atmospheric condition; separate analyses are required to evaluate the transient structural response to these conditions. The high pressure and the high temperature conditions shown for the first 45 seconds cannot occur simultaneously and need not be con-sidered in combination. 4.6 TOP VENT TEMPERATURE (CYCLINC) PROFILE DURING CNUCCING Full scale test results (Reference 16) indicate that during chugging the water lavel in the weir 2 anulus fluctuates over a 4 f act band contered at about the top vent centerline. The weir wall and the inside drywell vall then are subjected to steam temperature (230*r) above the top vent and ccid pcM temperature (100*F) near the lower ven s, with a transition region ia-between, where the temperature fluctuates due to the chugging process. O 011880
4 22A4365 Rev. 4 4-8a l f^
~
For weir annulus thermal stratification, the most severe design condition i results from imposing the maximum drywell temperature (3300F) concurrent
- with the initial suppression pool temperature (see Section 4.3.1).
i
- For evaluation of local effects, the cyclic temperature profile during 4 chugging is shown in Figure 4.11. The cycling temperature ranges'from 1000F to 2300F; and the period is equal to the chugging period, which randomly i
varies from 1 to 5 seconds. The areas of application are:
- 4 foot horizontal band on the weir wall and inside drywell, e the upper inside vent surface, e and an area of the outside drywell wall just above each top vent.
- as shown on Figure 4.11. I The duration of the thermal cycling is identical to the duration of chugging (see bar charts, Figure 4.3). As the event proceeds, the AT reduces in amplitude due to bulk pool temperature increase. As part of the design () calculation, this bulk pool temperature should be considered and is shown in Figure 6.17. The long and short term thermal gradients are discussed in l Attachment N.
$.7 DRYWELL MULTICELL EFFECTS j
Chugging is conservatively considered to be synchronous for the Mark III load definitions. The typical pressure time history is shown on Figure 4.8a. i Superimposed chugging spikes from adjacent vents, as confirmed by multi-cell tests (ref. 18), are miner and considered to be insignificant, On Figure 4.8) these spikes are superimposed to indicate the typical $ msgnitude, i i v 011880 l
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O v STR UCTURE : DRYWELL ACCIDEN T: LARGE STE AM LINE BRE AK (DBA) PGOL TF.MPER ATURE (SECTION 4 6) DayWEt t INTERNAL PRESSURE AND TEMPER ATesHE 4 LOADS DUE TO SE'SMIC ACCELERATION (ATTACHMENT 8) H'v DROST AT LC PRf SSURE NOTE - THERE WILL BE NO WATER IN THE WEIR ANNULUS BETWEEN I AND 30 SECONDS ISEC 4.1.31
- POOL DUMP STARTS AT 5 mm SINGLE STR VALVE ACTUAT60N ATTACHMENT A,SEC 2.4
- THE DRYWELL HEAD i OADS DUE TO FIGURES 103.104 DESIGN CONSIDERS z POOL SWE L L 10 5 AND 12.2 A HE AD SPR AY 9 SECTION 41.4 BRE AK.
2 O z b F ALLBACK LOADS RE 10' ' p2 tJ O mg
$ CONTAINMENT FREE SPACE .# s-kO PRESSURIZATION DUE TO ANDF URE 4.4 F ALLBACK LOADS &$
DRYWELL AIR CARRYOVER
" SECTION FOR A GIVEN WETWE LL PRESSURIZATION CHUGGING 4I8AND STRUCTURE ARE (LOAD ON DRYWE LL O O I 419 NOT COINCIDENT .
NEGATIVE LOAD BOTH LOADS HAVE SECTION 6.1.6.12 0 DUE TO POST A DURATION OF LOC A ECCS 0 5 sec. POOL SWELL SECTION 4.l.7 FLOODING OF CAN OCCUR 1 TO ORYWELL 15 sec AFTER BRE AK LOCA BUDBLE PRESSURE SECTION 4. l.4 DEPENDING ON LOAD HEIGHT ABGVF THE POOL. FALLBACK SECTION 4.1.1 POST LOCA SECTION 6.1.8 LOADS OCCUR wy WAVES 1.5 TO S secAF1ER THE BRE AK CONDE NSA TION OSCIL L AIlON SECTION 41.5 l l l 9 o 01 1.0 3.5 J0 5 10 30 100 600 o TIME AFTER EVENT isect w b-* co co o Figure 4.1. Drywell-Loading Chart for DBA s. 8
ST H UCT UH E DH YWE L L ACCIDENT SMALL STE AM BHE AK (SSAI (AT T ACHMENT 88 LOADS DUE TO THE SEISMIC ACCELEHATION (ATT ACHMENT B1 SECTION HY DHOST A TIC NOTE I. THE WEiH ANNULUS WILL SE CLE AHED TO THE TOP OF THE UI'PER VENTS WITHIN 4,3 A Ff W MINUTS OF THE ACCIDENT. titME IS BHE AK ARE A DEPENDENTI PftESSU RE 2 POOL OUMP INCLUDED ( AUTO AT 30 ment SECTION (FIGUHE 410) , DHy WiLL ATMOSPHEHE TEMPEHATUHE NOT E . DUHING COOLDOWN WITH CONDENSEH ISOLATED. (SEC 2 4 & 4.3) S/H VALVES AHE OPEH ATED PERIODICAL LY FOH I SINGL E $ H VALVE ACTUATION UP TO THREE HOUHS (ATT ACHMENT Al _ __ ____ _ _ _q 9 SECTION
.- l CHUGGsNG O
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%5 is5~ a3 ds<
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22A6365 Rev. 1 4-20 0 This figure is PROPRIETARY and is provided under separate cover. Containment Wall Pressure Trace During Chugging, Run 11 Figure !. 8. (Ref. Test 5707) h 101678
22A4365 Rev. 3 4-21 0 1 i n PULSE (PRESSUR E SPIXE)
)s c
2 k FOST-CHUG OSCILLATION 3 ! j B SIN J 2t FORT2 4 0.25 r 3 a Be -a lt2-0 25r) 3;g , '2 FORT2 s 0.25 r E O > . 0.25 r
^ - l y TIME, t (sec) ,
g , I
+ d h i
---> t g i t,
A I SEE TABLb 4.1 FOR VALUES FCR A, B AND d PRE-CHUG UNDERPRESSURE w - e/0.125 radians /sec j (A SIN w tg) FOR tg F ROM 0 TO 0.125 sec a = 0.55tr. Itsec J = 2str, radians /sec l 1 l l r Figure 4.8a. Typical Pressure Time-History on the Pool Boundary During Chugging O 090779 t I
22A4365 Rev. 3 4-21a O V A 7.5 ft 10 0 22 3 ft I TOP VENT ,
- q. a 2 ft i .o e
O 4 d hI E E E d 5 5 5 E 8 0 0.22 0.31 f8ASEMAT 1 0.22 0.31 Figure 4.8b. Suppression Pool Chugging Normalized Peak Underpressure Attenuation O 090779
i 22A4365 Rev. 4 4-21b O I V A 7 5 ft 10 0.38 3 ft
' I TOP VENT q . ,
sk 2 ft 1.0 l o 3 J E a m 4 s E a b d 5 E v x Q
-B ASE M ATg 0.38 f 0 51 0.38 0.51 4
Figure 4.8c. Suppression Pool Chugging Normalized Mean Underpressure Attenuation 0 011880
22A4365 Rsv. 3 4-21c e v /
-- 0.03 6 _
7.5 ft 4 _ 0.6 0.8 10 0.2 04 8 L-g a g 2 e i
~
~3 SPIKE PEAK PRESSURE c' - -
TOP VENT Q
.z o- . _
0
; -2 _
o b - 8s G j 3 d _6 . d I _ ,_ i
< T z E w
-8 a-d !
g _a 1 2
-10 'O- $ '
8ASEMAT _ l 0 03 j 0.1 ,
, _ o o3 i
I I I l l O.1 i Figure 4.8d. Suppression Pool Chugging Normalized Spike Attenuation l O 090779 l ( l
22A4365 Rei. 3 4-21d (3 V J V d - 2.0 7.5 f 4 4.5 4 2 6 8 2 3 DUR ATION (msec)
& if 1f if y0 - -
TOP VENT ( y 0.5 ft b } h \ { ) e sw
+
b 5
} J 3
b U d d w ( 5 NOTE: APPLIES TO BOTH PEAK AND MEAN
-10 .
PRESSURES ' 40 p0ASEMAT 2.0 p 20 40 & Figure 4.8e. Suppression Pool Chugging Spike Duration "d" as a Function of Location in the Pool l O 090779 l
22A4365 4-21e Rev. 3 v A 7 5 ft 1.0 0.25 3 tt TOP VENT q, ' V aL 2 ft 10 9 e s i s e 5 w & E 8 8 0.32 78 AStV %T 0.25
}o.25 0.32 Figure 4.Sf. Suppression Pool Chugging Normalized Peak Post Chug Oscillations 0
090779
f f 22A4365 Rev. 3 4 21f O 4 { l I l i 0 l CONTAINMENT 10 ft ELEVATION ! l, 1 i l i !2; (DRYWELL BASEMAT INTERF ACE . ); $ 4 3 a ! 2 ORYWELL WALL i
= 4 -
10 ft ELEVATION I - l' 8 z 3 5 - l l I ! I I I ' ' ' 6 18 27 36 45 . 45 27 -18 -9 0 9 ! AZIMUTH (degrees) I i t Figure 4.8g. Circumferential Underpressure Amplitude Attenuation i, .l O 090779
22A6365 Rev. 2 4-21g O t 7 6 - l ORYWELL WALL 10 ft ELEVATION
; s -
)3 a
j4 - Y I ' 23 - 2 w 4 y2 - l i I CONTAINMENT 10 f t ELEVATION .! ORYWELL WALL BASC 1
-36 -27 -1:1 -9 0 9 18 27 36 80
-- 80 AZIMUTH (degrees) i I 4
Figure 4.Sh. Circumferential Post Chug Oscillation Amplitude
- Attenuation i
e ! 101678
22A4365 Rev. 4 4-21h O v JL 7.5 ft 1.0 0.45 3f i
'I f
TOP VENT Q, , , n 2 ft i 1.0 0 i 5
$ E E E
- ?
z
'3 OS -8 ASEM AT% 0 45
. 0 45 06
[ l l l Figure 4.81. Suppression Pool Chugging Normalized Post Chug Oscillations Attenuation l l O 011880
22A4365 4-211 Rev. 4 l l 2.4 psed AL PULSE (PRESSUR E SPIKEl 3 1 8
- POST CHUG OSCILLATION g 4 5 psid
{
\
3 3 psed f 1 ms i k a f TIME. ) i PRE CHUG UNDERPRESSURE NOTE: FORCING FUNCTION FROM FIGURE 4.8a e Figure 4.hj. Chugging Pressure Time-History on the Drivell '4all e Adjacent to Vent O 011880
22A4365 4-22 Rev. 2 O This figure is PROPRIETARY and is provided under separate cover. l l 1 l I Figure 4.9. Dryvell - Containt::ent Pressure Diff erential During Chugging l l 042178 l
i i 1 ' 22A4365 Rev. 3 4-23 i i 4 l@ l l
- \
\ ( l l
+
I i l t i I i f ( i I i I Deleted i l l l l t f 1 1 090779 1
. _ . _ - . , _ _ _ _ _ _ . ~ - _ . . _ _ _ _ _ . _ _ . _ _ _ _ . _ _ . _ _ _ . ~ _ _ _ _ _ _ _ _ _ _ _
l 1 22A6365 4-26 Rev. 4 3>2 8/ I - I I- 1 $
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! l l l -
a R R 2 * .; S S (Bisd) 3HOSS3Bd l l l
! 3 R
l 2 l X S 8 2 g (dol 3WO1YW3dW31 011880
22A4365 Rev. 2 4-25 CYCLING FLUID TEMPERATURE TIME HISTORY 2300F i i 200 - l l C l l 9
- l I 5 I i 7
e i i s' I l
$ l l C
5 'O.2P' ' ' O 2P O.2P O.4P O P 2P O TIM E (sec) 1 < P 4 5 seconds n AREA OF APPLICATION
.A ,
- 2 ft +
- c. f I
- - c _
ff!II 1 ft
- d
' _ TOP VENT l I
T _ [ . . o,
. . N. ;
i L . ,' OUTSIDE DRYWELL INSIDE ORYWELL WElR WALL a WALL
. WALL i
l T
,)
Figure 4.11. Drywell Top Vent Cyclic Temperature Profile and Area of Applicatien During Chugging 101678
22A4365 Rev. 2 5-1 !
- 5. WEIR WALL \--
The weir wall experiences loading conditions during both the design basis accident and during a small steam break accident. Figures 5.1 and 5.2 are the bar charts for these two cases. The intermediate break loads are less severe than those associated with the large and small break. Figure 5.3 is the bar chart for this case. 5.1 WEIR WALL LOADS DURING A DESIGN BASIS ACCIDENT 5.1.1 sonic wave For the reasons discussed in 4..l.1, this phenomenon is not included in the weir wall design conditions. A sonic ccmpressive wave does not produce a design load condition in the drywell. 5.1.2 Outward Load During Vent Clearing () The pressure drop at any point on the weir wall due to the acceleration of water during vent clearing is less than the local hydrostatic pressure. There-fore, there is no net outward load on the weir wall due to vent clearing. This conclusion is based on the predictions of the analytical model presented in Reference 1. 5.1.3 Outward Load Due to Vent Flow Once flow of air, steam and water droplets has been established in the vent system, there will be a static pressure reduction in the weir annulus that leads to approximately a 10 psi uniform outward pressure on the weir wall. This loading was calculated with the vent flow model described in Reference 1 and for design purposes is assumed to exist during the first 30 seconds of blowdown. 5.1.4 Chugging Loads () The pressure pulses generated inside the top vents during chugging (see Section 4.1.9) propagate toward the weir annulus. A typical trace of the 101678
22A4365 Rev. 3 5-2 pressure pulses on the. weir wall is shown in Figure 5.4. The dominant pressure response in the weir annulus during chugging is characterized by a pre-chug underpressure followed by a pressure pulse train, as shown in Figure 5.4a. The load applied to the weir annulus (weir wall, basemat and inside drywell wall) is described by a pre-chug underpressure, defined as a half sine wave as shown in Figure 5.5, followed by the pressure pulse train shoh in Figures 5.5a or 5.5b. For local load considerations the peak amplitudes are applied, and for global considerations the mean ampli-tudes are applied. Vertical attenuation of the weir underpressure is very small; for design evaluation, no attenuation should be assumed. For the pressure pulse train, the attenuation on the weir wall and drywell ID wall in the vertical direc-tion is shown in Figure 5.6. For all global loads, there is no attenuation in the circumferential direction. O O 090779
22A4365 Rsv. 4 5-3 , 5.1.5 Inward Load Due to Negative Drvwell Pressure__. v Due to negative drywell pressure discussed in Section 4.1.7, reverse water flow in the horizontal vents will lead to inward acting impingement loads on the weir wall. A simple,' steady-state flow analysis leads to flow velocities approaching 40 f t/sec if it is assumed that a 21 psi negative differential exists between the drywell and containment. This leads to a total impingement force on the weir wall of 12,800 lb. per vent applied over the projected area of the vents as shown in Attachment II. This number is based on a simple jet impingement analysis which assumes that the force on the weir wall corresponds to a change of the horizontal momentum of the water flowing through the vents. This same negative drywell condition can theoretically result in the flow of water over the weir wall into the drywell. Using the nominal predicted drywell depressurization time history shown in Figure 5.7, a peak velocity of 30 feet / l sec can be calculated at the top of the weir wall. This velocity is decreased due to the effects of gravity with elevation and the spreading of the flow field so that the maximum elevation reached is 14 feet above the top of the l weir wall as shown in Figure 5.6. Structures in the path of the water are designed for drag loads using the following equation: o C
=
D F 2g c where: F = Drag Load Force, lbf C = rag e cient D A = Area Normal to Flow, Ft c = Density of Water, 62.4 lbm/ft gc = Newton's constant, 32.2 lbm-ft/lbf-sec2 V = Velocity of fluid, ft/sec. O 011880
22A4365 5-4 Rev. 3 5.1.6 Suppression Pool Fallback Loads 1 For the reasons presented in 4.1.6 and since the weir annulus pressure is controlled by vent flow during the period of interest, no suppression pool fallback pressure loads are specified for the weir wall. 5.1.7 Hvdrostatic Pressure _ During the first second after the DBA, the water in the annulus is depressed to the Nottom vent; therefore, there is no inward hydrostatic pressure load on the weir wall. Post LOCA hydrostatic load is an outward load due to the differ-ence between the water within the weir wall and the level in the suppression pool. The influence of seismic accelerations on hydrostatic pressure distri-bution is discussed in Attachment B. 5.1.8 Safety Relief Valve Loads In the event of safety relief valve actuation, the hydrodynamic pressure oscil-lations associated with the pipe air clearing transient can reach the weir wall ll through the vents. Attachmunt A provides loading information. The S/R valve load is applied to the projected vent hole area on the weir wall. l l 5.1.9 Condensation l There will be no loads induced on the weir during condensation, as shown by l [ lack of transducer response in the tests. 5.2 WE'. R WALL LOADS DURING AN INTERMEDIATE BREAK ACCIDENT Figure 5-3 shows the bar chart for the weir wall during the IBA. The safety reitef loads associated with ADS activation are discussed in Attachment A. The LOCA induced pressure differential across the weir wall will be small.
- 0 0
090779
22A4365 5-5 Rev. 4 5.3 WEIR WLL LOADS DURING A SMALL BREAK ACCIDENT The loading sequence for the weir wall during a small steam line break is essen-tially the same as for the drywell wall with the exception that there will be no pressure differential across the weir wall other than hydrostatic pressure. Apart from that, the information in Section 4.3 applies. 5.4 WEIR WALL ENVIRONMENT ENVELOPE J The temperature and pressure for the drywell envelope data (Figure 4-10) applies to the weir wall with the exception of the temperature of that part of the outside face which is below the elevation of the upper vents. This region will remain submerged snd will be maintained at suppression pool temperature. It should be noted that the weir wall structure is totally within the drywell and effects of environmental conditions should be examined cn this basis, including the thermal cycling during chugging (see Section 4.6).
- The f1rst 6 hours of the environmental conditions defined on Figure 4.10 are based on a small steam break. Faster shutdown by operator can reduce the duration of the small break to 3 hrs. For a large break, the free volume inside the weir wall is flooded and environmental temperature conditions will correspond to the water temperature in this volume. This is less severe than the conditions of Figure 4.10.
5.5 WEIR ANNULUS MULTICELL EFFECTS Chugging is c.onservatively considered to be synchronous for the Mark III load
- definitions. The typical pressure time history is shown on Figure 5.4a.
Superimposed chugging spikes from adjacent vents, as confirmed by multicell tests (ref. 18), are minor and considered to be insignificant. On Figure 5.9 these spikes are superimposed to indicate the typical magnitude. O 011880 s
STHUCTUHL. WEtH WALL ACCIDE NT : DESIGN BASIS ACC8 DENT 406 A) SECTION b 4 AND 4 6 WFIR WALL PHESSURE AND TEMPERATUHE (ATTACHMENTB) SEtSMIC - STHUCTURAL ACCELEHATION LOADS (SECTION 5.1.78 HYDHOSTATIC PHESSURE - (NONE BETWEE N O & 30 sect --
^ ^C" " ^
ON 4 LOADS DUE TO SINGLE S/H VALVE ACTUATION NOTE: CHUGGING AND erdWARD LOAD DUE TO
$ POST LOCA FLOO0tNG ARE NOT COINCIDENT.
p 5 5o O [ OUTWARD LO AO SECT 40N 51.2
- VENT CLE AHING u s CHUGGING SECTION 51.4 & FtGUHES 5 4. S 5. 5 6 e$
OUTWAHD LOAD - VENT SECTION b.1.3 FLOW INW ARD LOAD FALLBACK SECTION A CCS ION 5.1 b LOADS b.16 F LOODING OF DHYWELL SONIC SECTION 51.1 WAVE i i I l i i i b 30 1W W Os 1 15
' ADO S/H DYNAMIC LOAD TO ST A Tsc t OAD OUE TO DHYWELL AiH PUHGEO TO CONT AINMEt4T. VAPOH PHEaSuf1E AT 140 F.
APPLIES TO BOTTOM 2 VENTS ONLY Figure 5-1. Weir Wall-Loading Chart for DBA h - Y.
s s O STHUCTU54E Wean WALL ACCIDEN T SMAt L BRE AK ACCIDENT ISBAl LOADS DUE TO SEISMIC ACCELEHATION (ATT ACHMENT Bl H YDHOST ATIC NOTE THE WElH ANNULUS WILL BE CLE AHED TO THE TOP OF THE UPPER VENTS W11HIN A FEW MINUTES OF SE CTION PHE SSUH E THL ACCIDE NT 5. I F SEC ION M ATMOSPHE RE T E MP( H ATURE $ 4 AND Pt3 g Z 9 riGuHE a> C
- 4. 0 .
U NOTF DOHING COOLDOWN Wi1 H COND1 NSE H ISOL ATE D. SlH VAL VES AHE OP( HATE O PEHIODICAL L Y F OH (SEC 2 48 pM m z SINGL E S<H VAL VE ACTUATION O UP TO THRt f HOUHS ( AT I ACHME Ni Al u _-_______~ SECTION
$ CHUGGING NOTE CHUGGING CAN LAST UNTIL BRE AK ISOLATE DOR VESSEL DEPHESSURil[D l 6FIGURE I 4+ 2 3 o
----_-_ _ - d 5 4,5 5 AN0 5 6 o
a l m.n 3 ha s 6 his TIME AF7[H [ VENT o 6-* cn Weir Wall-Loading Chart for SBA
$ Figure 5-2.
u
a ST HUCT UH r wLig watt ACC1 DENT INTERMEDIAT L BHE AK ACCIDE NI (tB A) (alt ACHMLNT Ul LOADS DUE TO SEISMtC ACCELER ATION HYDHOST ATIC PHESSURE NOTL POOL DUMP INCLUDED AFTER AOS WilH ANNULUS LEVE L AT UPPEH V[NT LEVE AT ACH ENT A SINGLE S/H VALVE ACTUATION h ADS ACTUATED AiH HETUHN TO SECTION DHYWELL 43.2 5 OUTWAHD LOAD SECTION 5.1.2 g 3 VENT CLEAFt4NG 5 o OUTWAHD LOAD SECTION 5.1.3 (SMALL COMPARED TO DBA) g n g{y VENT FLOW > 5 POOL HEATUP HAlSES CONTAINMENT PHESSUHE AND TE MPERATUHE . oJ DHYWELL AlH PUHGED TO (SECT. 5 4) CONT AINMENT 3 pud (SECT.4.3.23 *$ CHUGGING SECTf 0NS S. l A. 5.1.9 a 1 VENT CL EAHING
' e e
500 1000 e 60 TIME AFTEH EVENT.sec SINGLE SHV L.OADS 00 NOT COMBINE WIT H OTHER SHV LOADS l o TIME SCALE DEPENDENT UPON UHE AK SIZE, MINIMUM VALUE s H
$ OF t = 2 0 mm w on o Figure 5-3. Weir Wall-Loading Chart for IIIA i CD
- O e - _ - - - -
l 4 22A4365 5-9 Rev. 3 I i lO , 4 . I 1 l 4 i 1 l I I i , i 1 i j This figure is PROPRIETARY and is provided under separate cover. 4 5 0 1 4 i ii l l 1 Figure 5.4. Typical Weir Wall Chugging Pressure Time History - Test O~ Series 5707, Run 1. l r 090779 l
22A4365 Rev. 2 5-10 l 0' i n l l
; ppE3suRE PUL$E TRAW r )
9 3 a 8 A ^ TIME '
\ / f O
PRECHUG UNDERPRESSURE Figure 5.4a. Typical Pressure Time-ilistory for Weir Annulus During Chugging l l 1 l 101678 O l
. . _ . . ~ _ _ . _ _ _ . _ _ _ _ _ _ .. _.. _ ._ . . _ _ . _ _ - . _ . _ . _ _ . . _ . . _ _ _ . . _ _ _ . _ . _ . . _ . . . . _ . . _ . _ _
E ' 22A4365 i Rev. 3 5-10a i 4 f i j d 3# 3 , 8 { 3 U j a 1 i P = -2. 5 SIN wt/0.C80 I PEAK 1 w I < < w . ' " FOR 0 < t < 0.080 sec
- i. 4 TIMF 1-l
.i ' 1 4 i l i j -2.15 - *' I 0.080 sec > 1 4 a t I f i t 1 p 4 I 7, i r .a ] 9 l t
' 3 4 P MEM = --0.98 SIN et/0.ORO :'
l < 4 w FOR 0 < t < 0.000 sec , s ; e TIME 7 f : i l -0 98 - 1 i 0.080 sec j
, l l 1 t
u Figure 5.5. Underpressure Distribution on the Weir Wall and Drywell I.D. Wall ! l During Chugging 1 1 l t !O ! 090779 ll il l , i ! l } i_______.._______..,____. - _ __ <,.. _ - _ . . _ . _ . . . - . . -__ _ . . . . _ _ . . . _ _ . . -_ , - , _ _ ..,- -.
22A6365 5-10b Rev. 2 l 1 l l
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22A4365 Rev. 2 5-11 i O WElR ANNULUS l POOL SURF ACE ELEV AflON DURING CHUGGING f f 2 1 t t l I TOP VENT Q
- O I A
I
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>i I 4
1 ' l O i & E --4 5 i O 8 t U i lE
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w C
' N PE AK AMPLITUDE 43 pied - SEE FIG 5 Sa
$ ME AN AMPLITUDE 15 piso - SEE FIG 5 Sb w _g _
DUR ATION 5 miec w THIS ATTENUATION ALSO APPLIES TO THE CIRCUMFERENTI AL DIRECTION
-10 -
1 -12 - j l l l l ? Y 2 I 04 08 10 ) i 0 02 ' NORMALIZED PRESSURE O Figure 5.6. Normalized k'eir Annulus Pressure Pulse Attenuation 101678
We$*
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8 1 - o t N d O I 6 1 o I o T l A f Z e IL R A U l Q e E s E I 4 s H 1 e u V S S X y E H A b P M d c e e h t u p I 2 1 a it 5 t n
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l l 22A4365 Rev. 4 5-13 O 20 M AXIMUM WATE R HEIGHT. V = 0 to - WATER ASCENDING E
< O E
E N TOP OF WEIR WALL O b.- i WATER DESCENDING INTO O 2 9
-10 -
DRYWE LL WEIR ANNULUS
$ DOWNWARD UPWARD
' + V E LOCITY
$2 VELOCITY E
-20 -
30 - 40 40 - 20 0 20 40 VE LOCITY (f pst i Figure 5.8. Vent Backflow Weir Annulus Water Surge Velocity vs. Height l Above Weir Wall 011880
i u ev s s TE e g n i g g u h
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22A4365 Rev. 3 6-1 s 6. CONTAINMENT The containment experiences dynamic loadings during all three classes of loss-of-coolant accidents. The containment designer should consider other containment loads such as negative pressures during containment spray activation, pipe whip, shield building loads, jet impingement etc. that are not included in this report. 6.1 CONTAINMENT LOADS DURING A LARGE STEAM LINE BREAK (DBA) Figure 6-1 is the bar chart showing the loading conditions that the contain-ment structure may experience during the DBA LOCA. Design loads for the various structures in the containment annulus are presented in Sections 7 thru 12. Figures 2.2-2 through 2.2-6 show typical structures above the suppression pool in the standard plant arrangements. 6.1.1 Compressive Wave Loading () Very rapid compression of the drywell air could, theoretica'lly, result compressive wave being generated in the weir annulus water. This wave could in a then travel down the weir annulus, through the vents and accross the pool to
*he containment wall. This phenomenon is not specifically included in the contain=ent design conditions on the basis that the approximately 20 psi per second pressure rate in the drywell is not sufficiently rapid to generate a com-pressive wave in the water. In addition, even if a 20 psi /sec wave were generated de the weir annulus surface, the very significant attenuation as the wave crosses the 18.5 ft. wide suppression pool would lead to insignificant containment wall loads. This phenomena has never been observed in any GE Pressure Suppression test.
6.1.2 Water Jet Loads
. Examination of applicable PSTF data shown in Figure 6.4, indicates some evidence l of a loading of the containment wall due to the water jet associated with the vent clearing process (i.e., less than 1 psid), as indicated by the small spike at l I, w) 0.8 sec. Water jet loads are negligible when compared to the subsequent air %/-
tubble pressure discussed in Section 6.1.3 and are not specifically included as a containment design load. 090779
22A4365 6-2 Rev. 3 6.1.3 Initial Bubble Pressure The PSTF air test data for runs 3 and 4 (Ref. 7) has been examined for evidence of bubble pressure loading of the suppression pool wall opposite the vents. These tests were chosen because the drywell pressure at the time of vent clear-ing is comparable to that expected in a full scale Mark III (i.e. , approximately 20 psid and because the vent air flow rates and associated pool dynamics would The maximum be more representative than the large scale steam blowdown tests. bubble pressure load on the containment observed during PSTF testing was 10 psid as shown in Figure 6.4. Figure 6.6 is a summary of all the peak containment wall pressure observed in PSTF tests during the bubble formation phase of the blow-is shown in down. The Mark III design load which is based on these tests, Figure 6.5. . The magnitude of the containment pressure increase following vent clearing is l dependent upon the rate at which the drywell air bubble accelerates the suppres-sion pool water. Circumferential variations in the air flow rate may occur due it results in negligible. variations to drywell air / steam mixture variations but in the containment bubblo pressure load. (See Attachment L). The conservative asymmetric condition assumes that all air is vented on half of the drywell periphery and steam is vented on the other half. The large scale PSTF test data is the basis for specifying the maximum asym-metric load of 10 psi. Figure 6-6 is a summary of all the peak containment wall pressures observed in PSTF tests during the bubble tormation phase of the blow-A maximum increase of 10 psid on f down. Figure 6.4 shows a typical transient. the containment wall was obe rved in the PSTF at the Mark III drywell peak cal-culated pressure of 36., cs i Figure 6-6 shows the maximum increase close to zero. Thus, use of a 10 psid asymmetric pressure condition applied in a worst case distribution is a bounding specification will be used for containment evaluation. 6.1.4 Hydrostatic Pressure In addition to the hydrostatic load due to the suppression pool water, the data h presented in Attachment B is used to determine the hydrostatic pressure loads on 090779
l 22A4365 Rev. 4 6-3 the containment during an earthquake. During periods of horizontal accelerations-4
. V("T there will be an' asymmetric distribution around the circumference of the con-tainment. The maximum pool level above the pool bottom in the suppression pool is 22 feet and is.26 feet for the dryvell and weir. annulus.
1. 6.1.5 Local Containment Loads Resulting from the Structures at or Near the Pool Surface Any structures in the containment annulus that are at or near the suppression { pool surface experience upward loads during pool swell. If these structures
~
are attached to the. containment wall, than the upward loads are transmitted into the containment wall. Sections 9 and 10 discuss tha types of loads that will be transmitted. I Localized loads on the containment wall resulting from the pressure losses j associated with water flow past a body are depicted in Figure 6-8. The data presented in this figure is based on drag type calculations and assumes that the affected structures have design features which preclude impact type loads from occu'rring. i j 6.1.6 Contaica.~at Load Due to Pool Swell at the HCU Floor 4 (Wetwell rressurization) 1 This structure is approximately 20 ft. above the pool surface and is 8 feet above , the point where breakthrough begins. Froth will reach the HCU floor approxi-l mately 1/2 second after top vent clearing and will generate bott impingement 4 loads on the structures and a flow pressure differential as it passes through the restricted annulus area at this elevation. [ The impingement will result in vertical loads on the containment wall from any I structures attached to it and the flow pressure differential will result in an I outward pressure loading on the containment wall at this location. The impingenent loads will be 15 psi and the froth pressure drop across the ECU ficer has been calculated to be 11 psi; the containment wall will see an f 1 l O c 011880 i
~ ., _. _ _ _ , .= _ - , _ , . _ . . . . _ . - .
22A4365 6-4 Rev. 2 this elevation. Figure 6-9 shows 11 psi discontinuous pressure loading at details of the 11 psi pressure loading. The bases for both the impingement and flow pressure loading are discussed in Section 11 and 12. the When evaluating the containment response to the pressure differential at HCU floor, any additional loads transmitted to the containment via HCU floor These loads supports (beam seats, etc.) must be assumed to occur simultaneously.2 are based on the assumption that there is approximately 1500 ft of vent area this elevation. For plant configu-reasonably distributed around the annulus at rations with HCU flow vent area other than 1500 ft2 (see Figure 6-16 for the froth pressure drop). The question of circumferential variations in the pressure under-neath the HCU floor is addressed in Section 12, and Attachment F. 6.1.7 Fall Back Loads _ No significant pressure loads are indicated from the data generated by the PSTF during the period when suppression pool water is subsiding to its original level following pool swell. Figure 6-4 shows that during the 2 to 5 seconds suppression pool fall back is occurring, the pool wall pressure probes show no evidence of pressures higher than the initial static pressure. Structures within the centainment annulus below the HCU floor will experience f all back induced drag loads as the water level subsides to its initial level. For design purposes, it is assumed that these structures will experience drag forces associated with water flowing at 35 f t/sec; typical drag coef ficients are shown on Figure 10-5. This is the ter=inal velocity for a 20 ft. free fall and is a conservative, bounding number. 6.1.8 Post Pool Swell Waves Visual observations of PSTF tests indicate that following pool swell, the sur-face of the suppression pool is agitated with random wave action having peak to peak amplitudes of less than 2 ft. These waves do not generate significant l containment loading conditions. O 042178
l 4 22A4365 Rev. 4 6-5 l
. 6.1.9 Condensation Oscillation Loads During the condensation phase of the blowdown, there have been some pressure ' oscillations measured on the containment wall in PSTF tests. Figures 6.10 and 6.11 show typical traces of the containment wall pressure fluctuations '
observed during the condensation phaee of the 1/5 scale PSTF tests.
' The forcing function to be used for design is described in section 4.1.5.
The magnitude of the load on the containment wall is shown in Figures 4.6a 1'
, and 4.6b.
+
6.1.10 Chugging Examination of the PSTF data shows that attenuated vent system pressure fluc-tuations associated with the chugging phenomenon is transmitted across the suppression pool. Figures 6.12 and 6.13 show typical containment wall and basemat pressures from full scale PSTF tests. Chugging loads on the contain-J ment are defined in subsection 4.1.9.2. 6.1.11 Long-Term Transient Following the blewdown, the Mark III containment system will experience a long cers suppression pool temperature increase as a result of the continuing core decay heat. The operators will activate the RHR system to control the tem-perature increase, but there will be a period of containment pressurization before the transient is terminated. Figure 6.14 shows the envelope of contain-ment atmospheric pressure and temperature for all postulated breaks. The Figure defines only the containment atmospheric condition. Separate analyses are required to evaluate the transient structural response to these conditions. Peak design containment pressure is 15 psig and peak design containment temperature is 185*F. i J !G 1 1 l 011880 e 9ve-g - --.m-- .~r3 --$+ - + - a%cvr e.- p 5- g-- ,,%
;+, -
- -w-.- y. -- w w y
l l
)
22A6365 Rev. 4 6-6 I The model used to simulate the long term post LOCA contain-ment heat up trans_ent is described in supplement 1 to Reference 1. 6.1.12 Containment Environmental Envelope Figure 6.14 is a diagram showing the maximum design containment pressure The and temperature envelope for any size of credible primary system rupture. long term containment pressure following a DBA is shown on Figare 6.15. 6.2 CONTAINMENT LOADS DURING AN INTERMEDIATE BREAK ACCIDENT Figure 6.2 is the bar chart for the containment during an intermediate break that is of sufficient size to involve the ADS system. Since these breaks are typically quite small and because there is a two minute timer delay on the ADS system, all the drywell air will have been purged to the containment prior to the time the ADS relief valves open. Thus, thi containment will experience the loads from multiple relief valve actuation coupled with the 5 psi, pressure increase produced by the drywell air purge and pool heatup. Since the former are pressure oscillations whose magnitude is not dependent upon the datum h level, these loads are additive. Attachment A defines the loading magnitudes 1 which are assumed for the S/R valve discharge. l The seismic induced increase in suppression pool hydrostatic pressure as a result of horizontal accelerations is asymmetric. This loading sequence is discu ssed in more detail in Attachment B. 6.3 CONTAINMENT LOADS DURING A SMALL BREAK ACCIDENT No containment loads will be generated by a small break in the drywell that are any more severe th a the loads associated with the intermediate or DBA break. Figure 6.3 is the bar chart for this case. There cre vaguarded RWCU lines in the containment that can re10ase steam to the containment free space in the event of a rupture. The RWCU isolation 0 011880
22A4365 Rev 4 6-7 [ valves and flow limiter for this system are designed to terminate the blow-I down before significant containment pressurization can occur. Typically a 2 psi pressure increase may occur. Steam released by a pipe break in the containment may stratify and form a pocket of steam in the upper region of the containment. The steam temperature will be at approximately 220'F whereas the air temperature will be at approxi-mately its initial pre-break temperature. This temperature stratification should be accounted for in the design. Local temperatures of 330/250*F are possible in the event of reactor steam / liquid blowdowns to the containment. 6.4 SAFETY RELIEF VALVE LOADS Relief valve operation can be initiated as a result of either a single failure, ADS operation, or a rise in reactor pressure to the valve set points. In addition, the containment can be exposed to S/R valve actuation loads any time the operator elects to open a valve or valves as during an isolated cooldown. j The loads generated by S/R valve actuation are discussed in Attachment A. i 6.5 SUPPRESSION POOL THERMAL STRATIFICATION 1 During the period of steam condensation in the suppression pool, the pool water in the immediate vicinity of the vents is heated. For the Mark III configuration, most of the condensing steam mass and energy are released to the pool through the top vents. By natural convection the hot water rises, and the cold water,is displaced towards the bottom of the pool. The vertical temperature gradient resulting from this effect is known as thermal stratifi-cation and is ~iscussed in Attachment N. The momentary thermal stratificatien for large break accident used in containment evaluation is shown in Figure 6.17. ! 6.6 CONTAINMENT k'ALL MULTICELL EFFECTS l No multicell effects on the containment wall were observed for pool swell or condensation phenomena during multicell testing (ref 17 and 18). i 011880 i
T STHUCTURL. CONT AINMENT WALL ACC4 DENT LAHGL STE AM LINE SHE AK (DBAl LOADS DUE TO SEISMIC ACCELEHATION (ATT ACHMENT bi HV DHOST ATIC PHLSSUH E NOTE POOL DUMP ST ARTS AT 5 m.n (SECTaON 61 A) 1000F POOL TEMPERATUHL (SECTION 6.1.11 AND 6.5) tSO F is0 F I ' '
^ ^ ^^
SINGLE S/H VALVE ACTUATION SECTtO 2 5 G 5 H COM A W ME N T o TH'S LOAD IS NOT COINCIDENT H A DUE TO DHMELL A1H SEE HGUHE 4 4 u WITH THE DATA ON FIG. 61.4 CARRYOVER w o " ' E <pp O WETWELL PHESSUHlZATION
- LOAD BELOW HCU F LOOH SECTION 6.1.6.12 0 SECTIGN y O
J CHUGGING6 ll1.50 e g LOADS OUE POOL SWELL AND F ALL BACK LOADS SECTION 6.1.5 TO POOL SWEL L SE CTION 618 COINCIDENT.BOTH LOADS HAVE A DUH ATION OF 0 5 sec POOL SWEL L CAN LOCA 8088LE ' " "" PRESSUHE LOAD POST LOCA WAVES DEPENDING ON HEIGHT ABOVE THE POOL. F ALL 8 ACK LOADS OCCUH WATLHJET A I l. 1.5 TO S sec AFTER THE BHEAK. IMPINGf MENT DOHING SECTsON 6.1.2 VENT CLE AHING. COMPHES$1VE WAVE S CTsON E1.1 LOAD OUTWAHD CONDENSATION OSCILLATION LOADS SECTION b l.9 I I I I I I sec 15 sec 10 sec 5 0 set 10 src 30 set 100 sec 61O he TIME AFTER EVENT o C co
. Figure 6.1. Containment-Loading Chart for DBA
# 9 e ---_-
-. ._ - . _ _ . _ . .- . ._. . . . . _ _ _ _ .. -. . - . . - . .., ~
O O STHUCTUHE . CONT AINMENT WALL ACCIDE NI INTEHMEDI ATE STE AM LINE BHE AK llBAl IATTACHMENTBl LOADS DUE TO SElSMIC ACCELEllATION ISECTION 6141 t?YDHOST ATIC PH ESSURE NOTE POOL DUMP ST ARTS AFTEH ADS ssNGLE StH V AL VE ACTU AllON ATTACHMENT A.
" SECTION 2 4 ADS ACTUATED b
G 6 AIR RETURN TO g DRYWELL (SEC 4 3.21 u 0 m C r AINMEr T - POOL HEATUP RAISES CONTAINMENT TEMPERATURE AND PRESSURE TO 5 p9g. DflYWELL DIFFEllENTIAL PRESSURE MAINTAINED AT 3 psed (SECTIONS 6 2.6 5) =b 3 (SEC 4 3 2) >$ CONDENSATION CHUGGING SECTIONS 6.8.9 & 6 I to OSCllLATlONS l 1 1 l 30 th- 500 1000
= TIME AFTER EVENT.sec o h SINGt.E SHV L OADS DO NOT COMBINE Wliti OTHER SRV LOADS g 2 TIME SCAL E DE PENDENT UPON llHf AK SIZE. MINIMUM V ALUE OF t = 2 mm co o Figure 6.2. Containment-Loading Chart for IBA &
I e
STRUCTURE. CONT AINMENT WALL ACCIDENT. SMALL STE AM BHE AK (ATT ACHMENT 88 LOADS DUE TO THF. SESSMIC ACCELEllATION HYDHOSTATIC PHESSURE NOTE. POOL DUMP INCLUDED lAUTO AT 30 ment ISEC 61 di SEC 2.4 SINGLE S/H VALVE As'TUAllON NOTE: DURING COOLDOWN WITH CONDENSEH ISOLATED, (ATTACHMENT A) E/+4 VALVES ARE OPEF ATED PEHIODICALLY FOR UP TO THREE ilOURS - ----- CHUGG NG: NOTE: 0 HUGGING CAN LAST UNTil BHE AK ISOL ATED OH VESSEL ! DEPRESSUHl2ED ISEC 6.1.10,2.31 o --_-J - p m 5 WN z #> O DHYWELL AIR CAHH) OVEH HAISES h 2 g PHESSUHE OtFF EREN 'lAL a 3 pied
- O
( 3 POOL HEATUP HAISES CONTAINMENT TEMPEHATUHE AND PHESSURE ftlSES TO 5 ps.g DRYWELL PHESSURE DIF FEHENTIAL MAINTAINE D AT 3 pu. (SEC 4.3 21 g i i 3 he 6 hr 1 men TIME AFTEH EVENT O m m e O Figure 6.3. Containment-Loading Chart for SBA d o 8 O e
22A4365 Rev. 3 6-11 O This figure is PROPRIETARY and is provided under separate cover. O . Figure 6.4. Observed Bubble Pressure During Pool Swell - l ' Test Series 5706 Run 4 l 090779
22A4365 Rsv. 2 6-12 PRESSURE , I 4 OISTRIBUTION h 1 1 3 -
, 13f 18 f t .
+
d
+
O *
. . 2 10 pud +
- +
$ ,3 4
5 h 21 8puo
- +g w 3 .
w 2 s o-
=
. w w a Z
4
- +0 Li2 L12 ,
.O. >
. BASEMAT 9
,7 I
I I IT 10
#1 pud y
/
N 218 pud O UR ATION P MAX w FOR y ( yo, r .10 sec 3 FORy yo r = 1.0 + tv - yo)/40 (sec) WHERE
$E D
i r* DELAY OUE TO FINITE POOL SWELL VELOCIT Y w) Mo l y= HEIGHT ABOVE BASEMAT. H 4, o l vo = INITI AL POOL DEPTH. It 33 zC l (B ASED UPON 40 fos POOL SWE LL VELOCtTW
~*
g i _ yy4x = vo
- 14 4
' r+01 ..v3 0
TIME AFTER LOCA (sect Figure 6.5. Dynamic Loads Associated with Initial Bubble Formation in the Pool 9 101678
22A4365 6-13 Rev. 2 O : i l This figure is PROPRIETARY and is provided under separate cover. O . i i O Figure 6.6. Contain:nent Pressure Dif ferential During Bubble Formation 012178 l
22A4365 Rev. 4 6-14 i !. 9 i 1 I I a h a 1 1 i 4 I I I I I. t J i This page intentionally deleted 1 i l d 4 4 1 i l l 1 i i i I 1 0 011880
( 22A4365 f Rev. 2 6-15 i 1
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45 [ c 9 AR R ANGEMENT " A*" ARRANGEMENT B STR EETER, "F LUID MECHANICS,** 5th ED , P 273 SEE FIGbHE 10 4 aPMAX 3 aP Pg r b2 [ PRESSURE DISTRIBUTION ap > #RES$URE DISTRIBtJTION g _ 4 /; PO FOR ARR ANGEMENT " A" v b/2 FOR ARR ANGE MEN T ' E-
&* MAX =PD aPMAX 16I' I
* = AP
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- NOTE POOL SWE t i
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'2 VE L OCl Tv 40 's sec f
1 l
% 80 O 10 20 ek O
U Figure 6.8. Drag Loads on Protruding Structures Due to Pool Swell 042178
22A4365 Ray. 2 6-16 w e i
~
3
- 3 HCU FLOOR h
_ ____g a. t i ps w--- CONTAINMENT s Figure 6.9. Containment Loading Due to Flow aP ACROSS HCU FLOOR t l O l 042178
7 i I 22A4365 ' 2 Rev. 2 6-17 i ' t I i I I i 4 l r i . t 1 l l t !
' I; i l 1
This figure is PROPRIETARY and is provided under separate cover. 4 L LQ I f l ! I
, r l
i i ; i . j- l 1 i i } i t i l J i ! 1 J Figure 6.10. . Typical Containment Wall and Basemat Pressure Traces During ! l Condensation, Run 23'(Ref. Test 5807) l~ - i
! f
- i 4
j 101678 . i i t i i ',-n-- 3- - , , ,,,r..--y ww- .__ - - - .--.,,,m--w,*m .-v 2-.s-- w,1, - - - -- - - , - . , - - -- - - + - - .-e,m.-
22A4365 Rev. 3 6-18
)
This figure is PROPRIETARY and is provided under separate cover. O j Figure 6.11. Containment Wall and Basemat Pressure Time Histories, Test 5807, Run 11 l 090779
4 22A4365 Rev. 3 6-19 I O l I 4 i i i l l f . _ This figure is PROPRIETARY and is provided under reparate cover. s J O % s
,s t i
'e s
g O figure 6.12. Containment Wall Chugging Pressure Ti=e History _ - - Test Series 5707 Run 9 ; 090779.
~;
i R s o ;n % g \
l 22A4365 Rev. 3 6-20 l l O' h
% 4 s
S s s This fi/,ure is PROPRIETARY and will be provided under separate cover. O
\
~
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'.'s
~ , N s
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s s 2 Etgure 6.L3. Basemat Chugging Pressure Time History, s,s ' Test Series 5707 Run 9 g e - g
,q u \' s 090779 s
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22.%365 Rev. 4 6-22 i I 30 O l l 20 - 9 w' s a w I 10 - MINIMUM ECCS PUMPS OPERATING 1000F SERVICE WATER TEMPER ATURE i l i I
^
O 10 3 10 4 10 5 10 6 ,y TIME, sec
% r 20 -
1 i n 3 E ALL ECCS PUMPS OPER ATING 1000F SERVICE WATEER TEMPER ATURE 10 - 0 3 4 5 0 7 10 10 10 10 10 TIME, su O Figure 6.15. Long Term Containment Pressure Following a DBA 011880
22A4365 Rev. 4 6-23 l l J t b N N I J e Q w ! c o L C
- S
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- R R 2 (p% a wools now 011880
22A4365 Rev, 4 6-24 24 O FREE SURF ACE 20 - - - - - - - - - - - - INITI AL POOL TEMPE R ATURE 1000F 16 - TOTAL POOL MASS 8 a 106 lb POOL OtPTH 20 ft i TOTAL ENERGY RELEASE 4 a 100 Stu FINAL BULK POOL TEMPERATURE 1500F
=
=
2 9 7 12 . TOP VENT CENTER LINE
.a NOTE:
USE PLANT SPECIFIC TEMPERATURES 8 - INITI A L FINAL i 4 - l
; ! ( ) eASEMAT l 0=
100 120 140 160 180 l POOL TEMPE RATURE ("F) Figure 6.17. Suppression Pool Temperature Profile for Large Breaks O l l 011880 t
d 22A4365 R;v. 2 7-1 O 7. SUPPRESSION POOL BASEMAT LOADS V In addition to the normal, seismic, de_dweight and hydrostatic pressure loadings, that section of the basemat which forms the bottom of the suppression pool also experiences dynamic LOCA loads and oscillatory loads during safety / relief valve actuation. The safety / relief valve loads are discussed in Attachment A. The outer half of suppression pool floor will experience a 10 psi bulk pres-sure load associated with initial air bubble formation as discussed in Sec-tion 6.1.3. This pressure rise above hydrostatic is assumed to increase to 21.8 psi at the drywell wall - with the increase from 10 psi to 21.8 psi to be assumed linear and distributed over 50% of the pool vidth as indicated in Figure 7.1. This specification is based on the observaticn that the maximum pressure that the initial bubble can ever have is the maximum drywell pressure during the accident. Data trace uc. I r; town on Figure 6.4 indicates that the pressure increase is no greater than 10 psi at a point halfway across the sup-pression pool. Thus the specification that the pressure increases linearly
\s l between this point and the drywell wall will bound the actual pressure distribution. During the condensation and chugging phases of the postulated LOCA blowdown, the loading on the basemat is the same as that on the contain-ment. See Sections 6.1.9 and 6.1.10.
The containment pressure increases to 3 psi due to drywell air Larryover and the long term pressure and teeperature increases as shown on Figure 6.15. The time history of these pressure transients is as shown on Figures 6.1, 6.2 and 6.3. Safety / relief valve oscillating loads are defincd in Attachment A. The net loading on the suppression pool liner will reverse during the negative pressure phase of the oscillation, and this lifting load on the liner needs to be con-sidered during the design process. Where ground water level is a concern, this pressure should also be considered in the basemat liner design. l l l G(~T l l 042178 i i
I I 22A4365 i Rav. 2 7-2 O NOTE: PRESSURES SHOWN 00 NOT INCLUDE HYOROSTATIC CONSIDERATICNS 218 psa d
# 5 a !
a 2
,0 _______________ o o
I l i l i I I
- tJ2 : : L/2 :
BASEMAT RAOf AL OlMENSION s Figure 7.1. Pool Bounde.ry Loads During Bubble Formation O I I l 9 l 042178
l 22A4365 Rev. 3 8-1
- 8. LOADS ON STRUCTURES IN THE SUPPRESSION POOL 3
There are certain structures within the suppression pool which will experience dynamic loads during both loss-of-coolant accidents and/or safety / i relief valve actuation. 4 S.1 DESIGN BASIS ACCIDENT Figure 8.1 is the bar chart that defines the loads that structures in the suppression pool experience during the LOCA. 8.1.1 Vent Clearing Jet Load During the initial phase of the DBA, the Drywell air space is pressurized and the water in the weir annulus vents is expelled to the pool and induces a flow field in the suppression pool. This induced flow field c'reates a dynamic load on structures submerged in the pool. However, this dynamic load is less (see attachment G) than the load induced by the LOCA air bubble which forms after the water is expelled. Since the air bubble dynamic load is bounding, this load is conservatively used in place of the water jet load. The air bubble load is discussed in Section 8.1.2 and attachment G.2.2. 8.1.2 Drywell Bubble Pressure and Drag Loads Due to Pool Swell During the initial phase of the DBA, pressurized drywell air is purged into , the suppression pool through the submerged vents. After vent citaring, a single bubble is formed around each top vent. It is during the bubble growth period that unsteady fluid motion is created within the suppression pool. buring this period all submerged structures below the pool surface will be exposed to transient hydrodynamic loads. l The methodology and calculation procedures for determining submerged structures drag loads are discussed in attachment G.2.3.
- O 09')779
22A4365 R .;v . 3 8-2 Structures in the suppression pool should be designed conservatively for the LOCA drywell bubble pressure (see Figere 7.1) and acceleration drag h (attachment G). This applies to small submerged structures, e.g., pipes. 8.1.3 Fail Back Loads There is no pressure increase in the suppression pool boundary during pool fall back as discussed in Section 4.1.6. Structures within the containment suppression pool that are above the bottom vent elevation will experience drag loads as the water level subsides to its initial level. For design purposes, it is assumed that th a ne structures will experience drag forces associated with water flowing at 35 ft/sec; that is the terminal velocity for a 20 ft free f all and is a conservative, bounding number. Free fall height is limited by the HCU Floor. 8.1.4 Condensation Loads Steam condensation begins after the vent is cleared of water and the drywell air has been carried over into the wetwell. Condensation oscillation phase is vibratory in nature and induces a bulk water motion and therefore creates drag forces on structures submerged in the pool. This condensation oscil- ! lation continues until pressure in the drywell decays. l The methodology and calculation procedures for determining condensation loads on submerged structures are discussed in attachment G.2.5. l 8.1.5 Chugging Following the condensation oscillation phase of the blowdown the vent mass flux falls below a critical value and a random collapse of the steam bubbles occurs. This pressure suppression phase is called chugging and causes a high pressure wave (spike) on structures submerged in the pool, j l The methodology and calculation procedures for determining chugging loads on submerged structures are discussed in attachment G.2.6. O 090779
22A4365 Rev. 3 8-3 I i () 8.1.6 Compressive Wave Loading i j As discussed in Section 6.1.1, the very rapid compression of the drywell air i
- theoretically generates a compressive wave. But as pointed out in i Sections 6.1.1 and 6.1.2, there were no loads recorded on the containment i wall in PSTF for this phenomena. From this, it can be concluded that com-pression wave loads or structures in the suppression pool are significantly smaller than loads caused by the water jet, for structures close to drywell.
I For structures near the containment, neither compressive or jet loads are significant. 8.1.7 Safety Relief Valve Actuation Loads on submerged structures due to safety relief valve actuation are i discussed on Attachment G. i i a
- O 4
i i 1 i l i e i f O 090779
STRUCTuHE. STHuCTOHES WITHIN THE SuiVMESSION POOL
- ACCIDE N T L AHGE ST E AM t INE eHE AK tt>ea s POOL TEMPEHATURE (SECTION 61.11 AND 6 5)
LOADS DUE TO SEISMIC ACCELEHATION (ATTACHMENT Bl HYDROST ATIC PH ESSURE NOTE. POOL DUMP INCLUDEO AT S m.n SINGLE S/R VALVE ACTUA TaON ISEC 2.4)
^ ^"
SECT ON 8 VENT CLEARING z WATEH JET SECTION & l.1 LOAD 5 SECTaON 8.8 2 D Q
$ N L FALLBACK SECTION 8.13 A V O *
.a u CONDENSATION OSCILLATION LOADS SECTION 8 8 4 LE PHESSURE LOAD CHUGGING SECTION 8 I 5 SECTION 8.12
- TYPICAL STHUCTuHES ARE COMPH ESStVE SECT 60N 8 8 6 1 $/R V ALVE LINES AND OUENCHER WAVE 2. ECCS SUCTION LINES 3 ECCS RETURN TO POOL LINES ITEST AND RELIEFI l l t l l f 1 1.5 3 5 30 100
- ADD S/H DYNAMIC LOAD TO STATIC LOAD DUE TO ORYWELL AlH PUHGED TO CONTAINMENT VAPOft PHESSUHE AT 1400F TIME AFTEH E VE NT. wt o
g Figure 8-1. Structures within Suppression Pool-Loading Chart for DBA cn O O O
+
l i I l 22A4365 Rev. 3 8-5 iG 1 1 } i 1 e i 1 i l i 1 1 i i f i 1 i 1 O Deleted Sh 8-3 i i l l 1 4 l i b ' 090779 }
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i
. 22A4365 Rev. 3 8-6 i
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22A4365 Rev. 2 9_7
. ~s i
- 9. LOADS ON STRUCTURES AT THE POOL SURFACE Some structures have their lower surfaces either right at the suppression pool surface or slightly submerged. This location means that these struc-tures do not experience the high pool swell impact loads discurred in Section 10. However, they experience pool swell drag loads and LOCA induced bubble loads. Relief valve loads must also be considered. These are:
t (a) Pool swell drag loads produced by water flowing vertically past the structures at 40 ft/sec. (See Section 8.1.2 and Attachment I). , (b) Pressure loads generated by formation of the vent exi# air bubble immediately following LOCA vent clearing. This type of load will result when the structure is expansive enough to restrict pool swell and cause the bubble pressure to be transmitted through the 4 ] () pool to the under side of the structures. For the GE reference design, the TIP and drywell personnel lock platforms and the sump tanks below are the only structures in this category. All are located on the drywell wall. The maximum upward floor pressure specified for r ais design is equal to the maximum drywell pressure 21.8 psid (see Figure 4.4) . Similar structures located on the containment wall would be designed for a maximum upward floor pres-sure of 10.0 psid (see Figure 7-1) . This is conservative because the bubble pressure can never exceed the drywell pressure, and no credit is taken for the attenuation of pressure associated with the head of water above the bubble. These structures should be designed conservatively for the combined loads specified above (i.e., drag loads and bubble pressure). (c) Loads due to the safety / relief valve actuation. See Attachment A. Only structures with surfaces in the suppression pool will experi-ence the S/R valve bubble loads. Pool f all back loads are as discussed in Section 4.1.6. 042178
ST R UCT UR E STHuCTUHES AT THE POOL SuHFACE ACCsD E N T L ARGE ST E Au tine BHE AK aneAt POOL TEMPEH ATURE (SECTION 61.11,6 SI LOADS DuE TO SEBSueC ACCELEH AT60N OF THE STHUCTuHES AND DOE TO SEsSusC sNDUCED POOL SURF ACE WAVES (ATT ACHMENT 88 SINGLE SIR w ALVE ACTU ATiON ADJACENT TO STHUCTURE (SEC 2 48 SECTION 9c AND ATT ACHMENT A LOCA BUBBLE PRESSURE LOAD SECT 60N 9 g
- mm 3 m>V Z
o .A u DH AG LOADS S E CTION .' u$ 1 o s FALLBACK SECTION 4 8.6 POST LOCA SECTION 6.1 B W AV ES f f I I I 6 iS 3 5 30 TIME AF TEH EVENT ac o 4 Figure 9-1. Structures at the Pool Surf ace-Loading Chart During DBA e S 3 b G G - G
22A4365 Rev. 2 10-1 m
- 10. LOADS ON STRUCTURES BETWEEN THE POOL SURFACE AND THE HCU FLOORS Equipment and platforms located in the containment annulus region, between the pool surface and the HCU platform, experience pool swell induced dynamic loads whose magnitude is dependent upon both location and the geometry of the structure. The pool swell phenomenon can be considered as occurring in two phases, i.e., bulk pool swell followed by froth pool swell. The pool swell dynamic loading conditions on a particular structure in the containment annulus are dependent upon the type of pool swell that the structure experiences.
In addition to location, the size of the structure is also important. Large platforms or floors will completely stop the rising pool, and thus incur larger loadings whereas small pieces of equipment and structural itams will only influence the flow of a limited amount of water in the immediate vicinity of the structure. The steam tunnel and HCU floors are the only struc-tures that could be categorized as expansive. Section 11 discusses these structures. f'\.
\s-) The remainder of this section deals with relatively small structures defined as approximately 20 inches wide. Figure 10-1 is the loading bar chart for these structures. Structures at this elevation will be subjected to verti-cal loads'only. Horizontal loading mechanisms are not identified and 1/3 scale impact tests verify this conclusion.
10.1 IMPACT LOADS Figure 10-2 shows the impact loading profile that is applicable to small structures which are exposed to bulk pool swell. The PSTF air test data shows t' tat after the pool has risen approximately 1.6 times vent submergence (i.e., 12 ft.) the ligament thickness has decreased to 2 ft or less and the impact loads are then significantly reduced. However, bulk pool swell impact loading is applied uniformly to any structures within 18 ft of the pool surface as shown on Figure 10-2. For evaluating the time at which impact occurs at various elevations in the contain=ent annulus, a water surface velocity of 40 ft/see is assumed. Bulk pool swell would start (% (. l 1 sec after LOCA. 101678
+T - w - ~7 -
22A5365 R;v. 2 10-2 The basis for the loading specification is the PSTF air test impact data discussed in Reference 7. Specifically, test Series 5706 run number 4 is used. These tests involved charging the reactor simulator with 1000 psia air and blowing down through a 4.25 inch orifice. Fully instrumented targets located over the pool provided the impact data. Additional tests have been conducted which provide impact data for typical structures that experience bulk pool swell. Data from these tests (Series 5805) indicates that the specified design load is conservative. It should be noted that impact loads are not specified for gratings. The width of the grating surfaces (typically 1/4 inch) do not sustain an impact load. This has been verified in the one third scale PSTF test Series 5805. Figure 10.3 should be used for calculating grating drag loads. For structures above the 19 ft elevation but below the HCU floo.s, the froth impingement data portion shown in Figure 12.2 should be used. Again, this impingment load is applied uniformly to all small structures with the time history shown. O For structures between 18 and 19 feet above the suppression pool design loads and duration are linearly interpolated from the values shown on Figures 12.2 and 10. 2. Figure 10.6 is a summary of the loading specificationa for small structures in the containment annulus as a function of height above the pool. The influence of seismic induced submergence variations on the pool swell transient and resulting impact loads has been considered. It has been concluded that the effect on the magnitude pool swell impact load is not significant. This conclusion is based on a consideration of the influence of submergence on swell velocity and the significant load attenuation which will result from the pool surface distortions. The very significant margins between the specified loads and the expected loads (see Attachment J) provides confidence 0 042178
1 1 - l 1
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22A4365 Rev. 2 10-3 l EQ that any local increase in swell velocities will not result in loads in excess of design valties, i J.- The conservatism in these load definitions are illustrated in Attachment 1 10.2 DRAG LOADS l d
.i In addition to the impact loads, structures that experience bulk pool swell 1
are also subject to drag loads as the pool water flows past them with i velocities as high as 40 ft/sec. Figures 10-3,10-4 and 10-5 provide drag ' load information for geometrical shapes. Data is applied to all small struc- , tures in the containment annulus between the pool surface and the HCU floors. ~ 10.3 FALL BACK LOADS Fall back loads are discussed in Sections 4.1. 6, 6.1. 7, and 8.1. 3. LO I 1 4 4 4 i l i l i l O I 042178 __ - . _ . -- - , - _ _ . ~ . _ . _ . _ . - . _ , , _. . -
SiHUCTUHF SMALL STHUCTURES BETWEEN THE POOL SOHF ACE AND THE HCU FLOORS ACCIDENT. LARGE STEAM LINE BHEAK (DtlA) IMPACT LOADS F4GURE 10.2 OH 10 6 F LOW (DH AG) LOADS SECTION 10.2 AND FIGUHES 10.3.10 4 AND 10 6 2 9 e o y 2 8 F ALLBACK LOADS SECTION 10.3 gg
<p a .u 2
- 6 V La 4
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- 1 TO 15 SEC F HOM LOCA.
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V Figure 10-1. Small cructures lletween the Pool Surface and the llCU Floor-Loading Chart During DilA O O e - - - - - - - - - -
l O O O l l l -
$15 ps F OH HE AMS AND SMAL L F L A T
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22A4365 kv. 2 10-6
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18 - 16 - NOTES e LOAD TO GE APPLIED TO SOtto AHEA ONLf e BULK POOL SWE LL F LOW (OH AGI e IMPACT LOAL15 AHE NOT SPECIFIED FOH GH A TINGS 14 - e OUHATiON OF LOAD 0 S SEC 5 12 - J 2 2 w 10 - 0 u SOUHCE Chemical E ngineee's Henemouk , 3 HH Pe<<v.o 537 y I = , _ 6 - 4 -
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2 l l I l I l g OS 06 07 08 09 10 OPEN AME A FHACTION Figure 10-3. Pressure Drop Due to Flow Across Grating Within 18 ft of the Pool Surface 042178
22A4365 Rev. 3 10-7 ) O NOTES 1. FOR DURATION, ASSUME STATIC LOAD
- 2. APPLIES TO FLAT SURF ACES.
FOR OTHER SHAPES SEE FIGURE 10.5 18 -
- 3. SOURCE: MARKS MECHANICAL ENGINEERS H ANDtiOOK, 6th EDITION, PAGES 11-82
( J 5 E Y 16 - O V O 5 E a
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'0 30 a io 2o R ATIO asb Figure 10-l. Drag Load on Solid Structures within 13 ft of the Pool Surface 090779 t
22A436', Rev. 2 10-8 (REF: FLUtD MECHANICS, VICTOR L. STREETER,5th ED. MC GRAW HILL) O BASED ON V = 40 f ps 8ASED ON V
- 35 f ps ORAG COEF FICIENT* PR ESSUR E PRESSURE BODY SHAPE O OlF F ERENTI AL (pse) DIF F ER ENTI A L Ipsi) 13 10 CIRCULAR CYLINDER 1.2 , F LOW OIRECTION
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0.6 2:1 7 5 E LLIPTICAL CYLINDER 0.32 4:1 4 3 ELLIPTICAL CYLINDER 0 29 8.1 3 2 ELLIPflCAL CYLINDEtt i 2.0 22 17 SOUARE 2.0 120* 22 17 TRI ANGLE 1.72 120* 19 14 TRIANGLE TR1 A NGLE 2.15
^ 90* 23 18 k
TRl ANGt E 1.60 90* 17 13 TRIANGLE 2.20 60" 24 18 TRIANGLE 1.39 60* 15 12 TRIANGLE 1.8 30* 19 15
,RI AN.E ,.0 y. ,, 8 SEMITU80LAR 2.3 h 25 19 S E MIT U8ULAR 1.12 12 9 4 5 Rangel
- These drag coeff*Cients are conservative because tttey are for low Reynold's Numoer flow condetions (10 - 10 Use of lower values may be used if its applicability can be demonstrated.
Figure 10-5. Drag Loads for Various Geometries (slug flotE) 042178
O O O NOTE:
- 1. CURVE B C D APPLIES TO HORIZONTAL RUNS OF PIPING 2 CURVE B A E APPLIES TO BE AMS AND SMALL F LAT STRUCTURES 8 F OR DUR A TION
- 3. SEE FIGURE 12 2 FOR HCU FLOOH SE E FIGURE 10 2 LOADS
- 4. SEE ATTACHMENT J AND FIGURE J 7 FOR JUSTIFICATION 185 l
FOR DURATION OF APPLIED LOAD 3 :'c TWEEN 18 AND 19 F EET, DETERMINE S l BY Lt' f AR INTERPOLATION OF vat.UES o l SHOWNLN FIGURES 10 2 AND 12 2 9 I 9 l j D 60gm r c0 . . . . . . C h>
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l NOTE "' ONLY DRAG LOADS ARE
- l APPLIED ABOVE THE HCU FLOOR l*
g FHOM VELOCITY DETERMINED BY DECELERATION WITH ELEVA-l TION. NO F ROTH IMPACT. NO l DRAG LOAD ABOVE 30 f t M - I FOR DURATION l 1 @SE E FIGURE 12-2 l 15 ini l8 I i o I l 18 19 1f HEIGH T FHOM POOL SURF ACE (f t) e o t; 1 M rigere 10-6 Summary of l'ool Swell 1.oatling Specifications for Small Structures in tlie Cont a liunent Annnius (Not Applicable to the Steam Tunnel or Expansive llCU Floors)
22A4365 Rev. 4 11-1 p L 11. LOADS ON EXPANSIVE STRUCTURES AT THE HCU FLOOR ELEVATION At the HCU floor elevation there are portions of the floor which are com-prised of beams and grating and other portions that are solid expansive struc-tures. The bottom of the steam tunnel is at approximately the same elevation (19 ft-6 in.). The small structure portion (beams and grating) of the HCU floor is discussed in Section 12. The expansive structures at this elevation experience an impulsive loading followed by an 11 psi pressure differential. The impulsive load is due to the momentum of the froth which is decelerated by the expansive structure. The 11 psi pressure differential is based on an analysis of the transient pressure in the space between the pool surface and the HCU floor resulting from the froth flow through the 1500 ft2 vent area at this elevation (see Sec- l tion 6.1.C). Figure 11-1 shows the loading sequences and Figure 12-2 shows the loading history. PSTF test Series 5706 is the basis for the froth impingement load of 15 psi
)
lasting for 100 msec (see Reference 9). Representative tests of the expected Mark III froth conditions at the HCU floor are the 5 ft submergence tests of Series 5801, 5802, 5803 and 5804 These tests confirmed the adequacy of the 15 psi impingement load. l The 11 psi froth flow pressure differential lasting for 3 sec is based on an analysis of the transient pressure in the space between the pool surface and the HCU floor. The value of 11 psi is from an analysis that assumes that the density of the flow through the annulus restriction is the homogeneous mixture of tha top 9 f t of the suppression pool (i.e., 18.8 lbm/ft3). Supple-ment 1 to Ref erence 1 describes the analytical model used to simulate the HCU floor flow pressure dif ferential and presents a comparison of model predic-tions with test data. This is a conservative density assumption confirmed by the PSTF 1/3 scale tests which show average densities of approximately 10 lbm/ft . Ref erence 11 indicates the HCU floor presst.re differential is in the 3 to 5 psi range. !O j 011880 l
22A4365 11-2 Rev. 2 The potential for circumferential variations in the pressure transient in the wetwell region beneath the HCU floor have been examined and on the basis of bounding calculations it is concluded that the pressure variation will be less than 0.5 psid (see Attachment F). I O O i l 042178 l l
d 22A4365 11-3/11-4 Rev. 3
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22A4365 i Rev. 4 *
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a-(N V, 12. LOADS ON SMALL STRUCTURES AT AND ABOVE TEE.HCU FLOOR ELEVATION Structures at the HCU floor elevation experience " froth" pool swell which _ involves both impingement and drag type forces. Figure 12.1 shows the loading sequences. Only structures in the line of sight of the pool will experience froth pool swell loads. PSTF air tests show that the structures experience a froth impingement load of 15 psi lasting for 100 milliseconds (Reference 9). The impingement data is shown on Figure 12.2. Structures must be designed for this short term dynsmic impingement load; grating structures are not subjected to this impingement load (Reference 12). As discussed in Section 6.1.6, following the initial froth impingement there is a period of froth flow through the annulus restriction at this elevation. The froth flow pressure differential load (i.e., drag type force) specification
) of Figure 12.2 is based on an analysis of the transient pressure in the space between the pool surface and the HCU floor. The value of 11 psi is from an analysis that assumes that the density of the flow through the annulus restriction is the homogeneous mixture of the top 9 ft of the suppression pool water and the free air between the HCU floor and the pool (i.e., 18.8 lb ,/ft3). This is a conservative density assumption confirmed by the PSTF3 1/3 scale tests which show an average density of approximately 10 lb /ft .
Representative tests of the expected Mark III froth conditions at the HCU floor are the 5 ft submergence tests of Series 5801, 5802, 5803, and 5804 Reference 11 indicates the HCU floor pressure differential during these tests was in the 3 to 5 psi range (Drag load on HCU floor). Those small structures above the HCU floor that could be exposed to pool swell froth may be exposed to a drag load. The drag load is determined for the geometric shape of the structure (reference Figure 10,5) using a froth density of 18.8 lbm/ft3 as in the HCU floor AP calculation and the velocity of the froth at the elevation of the structure. The velocity used is 50 ft/sec at 19-1/2 ft above the suppression pool and is decelerated by the effacts of gravity. The velocity of 50 ft/sec is a bound of the available data (Reference 13). No pool swell is assumed for structures more than 30 ft above
-the suppression pool.
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'T s'...._' ,7 The potential fcr.circur.forential variations in the pressure transient in the wetwell region beneath' tiia-HCU floor have been examined and on the basis of bounding cale214tions it is concluded that the pressure variation will be less than 0.5 psid. (See Attachcent F.)
Since the air tests were performed, additional PSTF tests have been conducted with the specific objective of providing further data on the interaction of pool, swell with the HCU floors. The test results are in Reference 11. Supplement 1 to Reference 1 describes the analytical model used to simulate the IICU floor fic ure differential and presents a comparison of model predictions with , cata. The model is shown to be conservative. 0 e s
- 1 042178 l
l
_ _ _ _ _ . __ _ _ . _ . . _ _ _ _ __ ._ ._ __ _ __ __ _ _ . . _ _ _ . _ .m _ __m_ O i STHUCTUHE STRUCTUHES AT THE HCu rtOOR ELEVATION ACCIDE NT L AHGE STE AM LINE BRE AK (DBAl i i l DOWNWARD LOADS
. DOE TO F Att. BACK ION U AND WATER ACCUMULATION (ON HCU FLOOH) l t
en F LOW (DRAG) TYPE LOADS ,
$ (UPWAHDI tsGURE 12.2 y$ '
ITWOPttASE FLOW THROUGH HCU FLOOHi ab
< s~ t 5 .u r z Ch a u s.n i o o
Z o 4 ' 3 F ROTH IMetNGEMENT FIGURE 12 7 LOADS (UPW AH DI l 4 1 1 I 5 1.5 16 TIME AF T ER EVENT, yc O w o U
- F i p,u re 12.1. Small Structures at the llCU Floor Elevation - 1.oading Chart During Dl?A U.
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22A4365 R- I Rev. 2
~%
(O REFERENCES NOTALLTHEREFERENCESAPPEARkNTHETEXT. THE FIRST 11 REFERENCES REPRESENT A COMPREHENSIVE BIBLIOGRAPHY OF REPORTS RELATED TO GE'S PSTF PROGRAM.
- 1. Bilanin, W. J., The General Electric Mark III Pressure Supprcssion Con-tainment System Analytical Model, NEDO-20533, June 1974 and Supplement 1 August 1975.
- 2. Mark III Confirmatory Test Program Progress Report, April 1973. NEDM-10848 (Proprietary Report).
- 3. Mark III Analytical Investigation of Small-Scale Tests Progress Report, August 1973. NED0-10976.
- 4. Mark III Confirmatory Test Program Phase 1 - Large Scale Demonstration Tests, October 1974, NEDM-13377 (Proprietary Report).
- 5. Third Quarterly Progress Report: Mark III Confirmatory Test Program, NED0-20210, December 1973 (Proprietary Report).
I 6. Fourth Quarterly Progress Report: Mark III Confirmatory Test Program, NEDO-20345, April 1974 Supplement 1 (Proprietary Report).
- 7. Fifth Quarterly Progress Report: Mark III Confirmatory Test Program, NED0-20550, July 1974 Supplement 1 (Proprietary Report).
l
- 8. Sixth Quarterly Progress Report: October 1974 (Letter Transmittal to
! NRC Staff.) (Proprietary Data Attached.)
j 9. Seventh Quarterly Progress Report: Mark III Confirmatory Test Prot, ram, NED0-20732-P, December 1974 (Proprietary Re. pert).
- 10. Eighth Quarterly Progress Report: Mark III Confirmatcry Test Program, NEDO-20853-P, April 1975 (Proprietary Report).
, 11. Mark III Confirmatory Test Program 1/3 Scale Three Vent Tests, NED0-13407, April 1975 (Preprietary Report). I
- 12. Mark III Confirmatory Test Program 1/3 Scale Pool Swell Impact Tests -
> Test Series 5805, NEDE-13426-P, August 1975 (Proprietary Report). l 13. Mark III Confirmatory Test Program 1/3 Scale Three Vent Air Tests - Test Series 5806, NEDE-13435-P, November 1975 (Proprietary Report).
- 14. Test Results Employed by GE for BWR Containment and Vertical Vent Loads, NEDE-21078P, October 1975 (Proprietary Report).
- 15. Mark III Confirmatory Test Program - 1//3 Scale Condensation and Stratification Phenomena - Test Series 5807, NEDE-21596-P, March 1977 l
j (Proprietary Report). 7 l I l 101678 i l
l 22A4365 l I R v. 3 R-2 I
- 16. Mark III Confirmatory Test Program - Full Scale Condensation and Stratification Phenomena - Test Series 5707, NEDE-21853-P, August 1978 (Proprietary Report).
- 17. Mark III Confirmatory Test Program - 1/9 Area Scale Multivent Pool Swell Tests - Test Series 6002, NEDE-24648P, September 1979 (Proprietary Report)
- 18. Mark III Confirmatory Test Program, 1/9 Area Scale Condensation and Stratification Phenomena, Test Series 6003, NEDE-24720-P, November 1979 (Proprietary Report)
O O 090770
22A4365 Rev. 4 A-1 O ATTACHMENT A SAFETY RELIEF VALVE LOADS (QUENCHER) Page 1 Al.0 INTRODUCTION A-5 A2.0
SUMMARY
& CONCLUSIONS A-6 A2.1 Load Reduction A-6 l A
3.0 DESCRIPTION
OF PHENOMENA A-7 A4.0 ARRANGEMENT A-10 4 A4.1 Distribution in Pool (Quencher Arrangement) A-10 A4.2 SR7DL Line Routing A-11 A4.2.1 Line Lengths and Volume A-ll A4.2.2 Drywell Penetration Sleeve A-12 A4.2.3 SRVDL Vacuum Breaker A-13 A5.0 QUENCHER LOADS ON POOL BOUNDARY A-28 A5.1 Pressures on Drywell, Basemat, and Containment A-28 A-29 {') A5.1.1 Single S/R Valve Loads A5.1.2 Two Adjacent S/R Valve Loads A-29 I A5.1.3 Ten S/R Valve Loads A-29 A5.1.4 Eight S/R Valve Loads (ADS) A-30 A5.1.5 All (19) S/R Valve Loads A-30 AS.2 Loads on Weir Wall A-30 AS.3 Loads on Submerged Structures A-30 A5.4 Normalized Pre;sure - Time History (Theoretical Raleigh Bubble) A-33 A5.5 Representative Pressure Time History A-33 AS.6 Estimated Margins A-33 A5.6.1 Peak Bubble Pressures A-33 A5.6.2 95%-95% Confidence A-34 A5.6.3 Margin A-34 A6.0 OTHER LOADS ON STRUCTURES IN THE POOL A-66 A6.1 LOCA and Peol Swell A-66 A6.1.1 Forces on Pipes Due to Vent Clearing--Pool Swell and Fallback A-66 () A6.2 Thermal Expansion Loads A6.3 Seismic Laads by A.E. A-66 A-66 A6.4 Seismic Slosh Loads by A.E. A-66 011880
22A4365 Rev. 4 A-2 Page A7.0 QUENCHER ANCHOR LOADS A-69 A7.1 Quencher Arm Loads and Quencher Loading Application A-69 A7.2 Quencher Design Information A-70 A7.2.1 Codes and Standards A-70 A7.2.2 Design Pressures. Temperature, Loads, Configuration, and Performance A-71 A7.2.2.1 Component Data A-71 A7.2.2.2 SRVDL Geometry A-71 A7.2.2.3 Quencher Design Criteria A-72 A7.2.2.4 Quencher Configuration and Location A-72 A8.0 S/R LOAD COMBINATIONS A-81 A8.1 Symmetric and Asymmetric Load Cases A-82 A3.2 SSE and OBE Considerations A-83 A8.3 LOCA Considerations A-83 A8.3.1 DBA with M.S. Line Break A-34 A8.3.2 DBA with Recirculation Line Break A-84 A8.3.3 Other SRV Conditions A-84 l A8.4 Recommended Design Load Su=mation A-85 gg A9.0 FATIGUE CYCLES A-87 l A10.0 RECOMMENDED CALCULATION PROCEDURES FOR MARK III USERS A-90 l A10.1 Constraints A-90 A10.2 Determine SRVDL Design A-91 A10.3 S/R Valve Air Clearing Loads Mark III 238 Standard Plant A-93 A10.3.1 Absolute Pressure on Basemat and Walls A-93 l A19.3.2 How to Find the Attenuated Pressure on the Drywell Wall, Basemat and Containment Wall A-94 All.0 PARAMETRIC STUDIES A-104 A12.0 3 ASIS AND JUSTIFICATION FOR DEVELOPMENT QUENCHER LOADS A-106 A12.1 Intreduction A-106 A12.2 Test Data Application for Mark III Containment A-ll7 A12.2.1 Miniscalc Test Observations A-107 A12.2.2 5 mall-Scale Test Observations A-107 A12.2.3 Large-Scale Test Observations A-107 A12.3 Physical Parameters A-108 0 011880
^~ A~' l 22A4365 Rev. 4 i
i
- f. age, A12.4 Correlation of Positive and Negative Pressure Peaks A-118 A12.5 Development of Design Value Calculation Method A-140 A12.6 Application A-173 Al2.7 SRV Load Reduction A-209a A12.8 References (for Al2) A-210 O
l 011880 I
22A4365 A-5 Rev. 3 O A
1.0 INTRODUCTION
General Electric has determined that tha quencher is a desirable alternative feature to minimize suppression pool boundary loads resulting from the air clearing phenomena in the Safety Relief Valve Discharge Line (SRVDL) . The quencher device will be specified for the standard 238 Mark III design and is recomended for %'R-6, Mark III application. This attachment provides the following:
- a. Recomended quencher arrangement,
- b. Recomended quencher distribution in the pool.
- c. Calculation of pool boundary loads for 238 Standard Mark III application.
- d. Definition of other loads including quencher anchor loads.
- e. S/R valvc combination design load cases and estimated valve cycles.
- f. Procedures for calculating pool boundary loads for other Mark III plants,
- g. Justification and basis for quencher loads.
It should be emphasized that the specific pool boundary loads identified herein are for a particular SRVDL configuration are used for example only, and should not be used arbitrarily by other designers. Since the calculation of the quencher loads is highly sensitive to and dependent upon the SRVDL design, procedures in this attachment A are provided to obtain plant unique pool boundary loadings for other SRVDL and pool designs. 090779
1 22A4365 R7v. 4 A-6 A2.0
SUMMARY
AND CONCLUSIONS O Once the SRVDL routing is established the detailed calculation of the pool boundary loads resulting from the quencher air clearing transient is per-formed. The line air volume is the critical parameter and for the Mark III design a series combination of both 10" Schedule 40 and 12" Schedule 40 pipe is utilized in the line design. The SRVDL peak pressure is limited to 625 psid (S/R valve back pressure limit) . Table A4.3 lists the SRVDL air leg information for the 238 Standard Plant. The maximum air volume is 56.13 ft . W th this design, the maximum quencher bubble pressures are tabulated in Table A4.4. See Section A10 for clarifi-cation. This design procedure is based on single and multiple or consecu-tive actuation considerations at 95-95% confidence. To assure that the initial water leg (L < 18 feet) is not exceeded ater - following the initial actuation, vacuum breakers are used on the SRVDL. The water leg limit is a design objective for the standard 238 Mark III containment. l The design procedure requires an optimization of the SRVDL air volume to assure the 625 psid peak pressure limit is not exceeded with a minimum air volume. i Table A4.2 summarizes the SRVDL design requirements and objactives neces-sary to obtain the S/R valve pressure loads for the 238 Mark III containment identified in this attachment. A.2.1 LOAD REDUCTION The GiSSAR PDA quencher design bubble pressures sur .arized in Table A.4.4 have been shown to be conservative based on applicable plant test results. The analyses of test results provide the load reduction shown in Table A.4.5. The reduced design quencher bubble pressures for design centrolling loads are based on a 90-90 percent confidence level of the CAORSO data. Load reduc-tien justification is discussed in Section A.12.7. 011880
22A4365
^~I Rev. 2 A
3.0 DESCRIPTION
OF THE PHENOMENA Prior to the lif ting of a pressure relief valve, the downstream piping between the S/RV discharge and the water surface is filled with air at drywell pressure and temperature conditions. The discharge piping terminates at some pre-determined submerged depth in the suppression pool with the water level inside the pipe at the same level as the water level in the suppression pool. When a relief valve lif ts, the effluent reactor steam causes a rapid pressure build 4p in the discharge pipe. This rapid compression of the column of air in the pipe causes a subsequent acceleration of the water slug in the submerged portion of the pipe. During this blowout process the pressure in the pipe builds to a peak as the last of the water is expelled. The compressed cushion of air between the water slug and the ef fluent vapor exits the quencher and forms four clouds of small bubbles r~ ( ,}j that begin to expand to the lower pool pressure. This expansion leads to coalescence of the bubble cloud into four bubbles. The four bubbles continue to oscillate, displacing the water and propagating a pressure disturbance throughout the suppression pool. The dynamics of the sub-merged bubbles of air are manifested in pressure oscillations (similar to that of a spring-mass system) arising from the bubble expansion coupled with inertial effects of the moving water mass. The sequence of expansion and contraction is repeated with an identifiable frequency until the bubbles reach the pool surface. The magnitude of the pressure disturbance in the suppression pool 4 decreases with increasing distance from the point of discharge, resulting in a damped oscillatory load at every point on structures below the v ar surface. 042178 O
22A4365 Rcv. 2 A-8 From an air-clearing standpoint, a decrease in the volume of air initially h in the discharge pipe will result in a decrease in the containment loads due to relief valve discharge. Since the design limit of the safety / relief valve is 625 psid,* the discharge pipe volume must be sized so this limit will not be exceeded. There is a balance that must be reached; pool boundary loads are optimized while the safety / relief valve line pressures are not exceeded. Figure A3.1 demonstrates the effect of discharge pipe air volume on the peak pipe pressure. This figure was developed for the specific parameters listed on the figure. The pipe pressures were calculated for first actuations or opening of a safety / relief valve. O
$3ased on back pressure specifications to which valves are purchased 101678
O 60 s 55 - PPIPE =~ 625 ps=1 g- 50 - NOTE NOT TO BE EXTRAPOLATED E OR INTE RPOL ATE D y S/R VALVE SET PRESSURE = 1217 paid 3 Ppa pp ,625 psal O S/R VALVE F LOW R ATE
- 317.9 R1/nec mN
> G N l g VALID ONLY FOR 0.33 s C s 5.0 < >
4 43 _ WHE RE C - 10 m. PIPE L FNGTH ( AIRI 12 m. PIPE L ENGTH l AIR) h p WATE R LEG = 18 f t VALVE OPENING TIME s 0.02 sec 40 - I I I I I I I I I O , 1 2 3 4 5 6 7 8 9 to F t D (10 en SCHE () 401 o w H a l'igure A1.1. SRVlil. Air Volume W'rsus fl./l) witti 625 3. aid Constraint during Air Clearing
>s
22A4365 A-10 Rev. 2 A4.0 ARRXiGEMENT A4.1 DISTRIBUIION IN POOL (QUENCHER ARRANGEMENT) Figures A4.1 and A4.2 show the elevation and plan views of the standard quencher arrangement. For the 238 Standard Plant the quencher arm is located at 6.5 feet above the basemat and the inclined penetration is 45 . This results in a water leg length of N18 f t. This arrangement meets the following objectives :
- 1. Minimize drywell structural interference.
- 2. Permit water circulation through top and bottom of the drywell sleeve penetration.
- 3. Locate quencher arms at an elevation between vent holes to minimize vent discharge loads on the quencher during LOCA.
Figure A4.1 shows two support methods. The alternate position is the designer's eption. An advantage to the side anchor arrat gement is that it eliminates containment liner penetration for anchor requirements. Figure A4.3 shows the recom= ended quencher azimuthal locations in the standard 238 pool. As shew. in this figure the low, inter =ediate and high pressure-svitch set valves are uniformly distributed around the pool to preclude concurrent adjacent valves operation. 042178
22A4365 A-ll Rev. 3
\
Table A4.1 identifies the figures for S/R valve location, quencher elevation and plan view for the Mark III 238, 218 and 251 plants. As shown in Figures A4.1, A4.6 and A4.8, the elevation of the quencher arms from basemat, varies for the various Mark III plant configuration to satisfy the arrangement objectives cited above. The recommended quencher arm elevations for the three plant sizes are: Standard Plant 238 6.5 ft above basemat l Standard Plant 218 5.5 ft above basemat Standard Plant 251 5.0 ft above basemat A4.2 SRVDL ROUTING l 4
' } The SRV)L is routed by the Arc.hitectural Engineer from the first pipe s ,/
anchor point just below the S/R valve using 10",12", and 14" Schedule 40 pipe to the drywell and 10" Schedule 80 through the drywell wall to and including the' quencher. TheSRVDLshouldhaveasufficientslopeintheairl leg section routing to prevent condensation accumulation in the line. Figure A10.2 is a typical layout of the SRVDL Routing. l A4.2.1 Line Lengths and Volume Line lengths and volumes are based on the layout shown in Figure A10.2 and the S/R valve constraint of 625 psid. These lengths and volumes are shown in Table A4.3. The layout design does not represent an optimized layout with respect to pipe air volume. It is possible to reduce the air volume within the 625 psid pipe pressure constraint and thus reduce pool boundary loads. 090779
22A4365 Rev. 3 A-12 The SRVDL pipe size and line lengths shown are optimized to satisfy a S/R valve back pressure constraint of 550 psid rather than 625 psid, and at the same time minimize the air volume in the lines to obtain the pressures on the suppression pool walls. The design loads for pool boundaries and for support of the quencher device are sensitive to and dependent on the design of the Safety Relief Valve Discharge Line (SRVDL). The design requirements for SRVDL are discussed in Section A10.2 and All.0. The SR7DL f rom the 45' elbow just above the pool to the quencher is a 10" Schedule 80 pipe. (See Figure A4.1.) The increase to Schedule 80 pipe is to provide for corrosion allowance. The corrosion allowance for Carbon Steel is 0.125"/40 years / side and stainless steel is 0.002"/40 years / side. A4.2.2 Drvwell Penetration Sleeve The Drywell Penetration Sleeve is a 14" Schedule 80 pipe ac 45 which acts as a conduit for the SRVDL. The sleeve is shown in Figure A4.1 with the lower lip of the upper end just below the pool level and extending down to the top level of the top drywell ve.nt. The sleeve may be extended h as shown by dotted line, if needed for support. A.4.2.2.1 Thermal Consideration Studies indicate that the 14" Schedule 80 pipe sleeve to concrete interface does not exceed the 200*F limit for normal S/R valve operation. The design temperature criteria f rom the ASSE boiler and pressure code subsection CC-3440, concrete temperature, Section III, Division 2 is:
"a. The following temperature limitations are for normal operation or any other long term period. The temperatures shall not exceed 150*F except for local areas, such as around a penetration, which are allowed to have increased teeperatures not to exceed 200*F.
- b. The temperature limitations for accident or any other short term period shall not exceed 350*F for the interior surface. However.
O 090779
I 1 22A4365 A-13 Rev. 4 local areas are allowed to reach 650*F f rom steam or water jets in (m) the event of a pipe failure." A4.2.3 SRVDL Vacuum Breaker Vacuum breakers are provided for each of the S/R valve discharge lines to prevent excessive water rise in the SRVDL pipe above normal S/R pool level following valve actuations. At the time of initial opening of the S/R valve, the water level in the S/R Valve Discharge Line (SRVDL) is at the normal suppression pool level. Af ter the S/R valve closes, the steam remaining in the line l condenses, creating a vacuum which draws the water to a higher than normal pool water level in the line. Higher SRVDL peak pressure and _ thrust load will occur if the SRV opens when the water is above the normal pool level. The purpose of the discharge line vacuum breakers is to prevent the water from rising substantially above its normal level when a subse-quent S/R valve opening occurs, and thus, the SRVDL peak pressure is
) about the same as for the first opening.
The SRV vacuum breakers are 1ccated in the drywell above the expected level of water rise in the line subsequent to SRV closure. This eliminates the possi-bility of wetwell pressurization in the event of a stuck open vac"un breaker and ensures proper functioning of the vacuum breaker during the reflood transient. The following parameters will yield satisfactory performance for most SRVDL However, plant geometries and is recommended to satisfy the above requirements. specific analysis for vacuum breaker design should be performed by the design engineer to confirm this. j a. The vacuum breaker effective area, (A//K)* is equal to or greater 2 than 0.30 ft ,
- b. The vacuum breaker shall open (fully closed to fully open) in 0.2 second or less when an instantaneous AP of 0.5 PSID is C4 applied across it.
- c. The minimum opening differential pressure to start the vacurrt breaker to open is equal to or less than 0.2 PSID. 011880 umm - - ,r s
i l 22A4365 A-14 ! 1 R;v. 2
- d. The vacuum breaker must be fully open when pressure difference g is equal to or less than 0.5 PSID.
The vacuum breaker should be located in the drywell at an e. elevation above the maximum water level rise in the line following a SRV closure.
- L is used to calculate flow through the vacuum bresker as follows:
[K A w = /aP (2pgc) (144) - v'K whe re w = Flowrate through vacuum breaker in Ibm /sec aP = Pressure differential across the vacuum breaker (PSID)
= Air or steam density in Ibm /ft
~
t = 32.2 c 2 lbf - sec
^ = Effective area of valve in ft
',T 042178 O
I t 4 l 1 4 22A4365 ^~10 l Rev. 2 LO I i Table A4.1 i t QUENCRER ARRANGEME!.T j i i Mark III Plants S/R Valve Location Quencher Elevation / Plan View 1 ! 238-732 STD. Figure A4.3 Figure A4.1/ Figure A4.2 l i 238-615 Figure A4.4 Figure A4.1/
- i 218-592 Figure A4.5 Figure A4.6/
- I l 251-784 Figure A4.7 Figure A4.8/
- l 251-848 Figure A4.9 Figure A4.8/
- 1 4
O
- Typical plan view similar to Figure A4.2.
I 1 l 1 l I i . l 1 l 1 i i i 042178 O
- . . . , , . . . . . , . . . . - , - - . - . , - . . -- . . . - - , - - - . . - - - . . - , ~ , . ..
22A4365 A-16 Rev. 3 Table A4.2 SRVDL DESIGN REQUIREMENTS AVD OBJECTIVES SRVDL DESIGN REQUIRDtENTS (a) Maximum SRVDL Pipe Pressure 1625 psid. (Coordinates of (fl/D) and (SRVDL Air Volume) must be 1 625 psid as plotted on Figure A3.1) (b) Two vacuum breakers are required in the drywell. SRVDL DESIGN OBJECTIVES
- 1. Water leg i 18 ft.
- 2. Safety-relief valve opening time > 0.02 sec.
- 3. Minimize the SRVDL air leg volume.
l
- 4. Minimize length of longest SRVDL.
l
- 5. Minimize the contribution of fL/D to the first half of the discharge line.
! 6. Start 12" S/40 er 14" S/40 pipe just below the first anchor I point to meet objective (5). l
- 7. The ratio of the air legs (length of 10" S/40 pipe / length of 12" S/40 pipe = C) should be 0. 33 < C < 5.0.
B. Slope lines down toward pool to avoid condensate-water accumu-lation in line (no horizental runs) . I
- 9. SRVDL vacuum breakers should be 10" size. One > 10 f t. above tha weir wall and the other just below the seismic restraint at j the SRV.
090779
22A4365 A-17 Rev. 3 1 O Table A.4.3 SRVDL MARK III 238 STANDARD PLA'IT Air Leg Length Max. fL/D Volume (ft3) (a) (b) S/R Valve Total Length, 10" S/40 12" S/40 14" S/40 30'-5" 49'-3" - 54.9 2.09 4.21 V-1 7 9' -8" V-2 80'-2" 26'-11" 53'-3" - 56.13 2.46 4.95 V-3 73'-7" 33'-7" 3 ' - 9" 36'-3" 55.36 2.41 4.85 l V-4 77'-2" 20'-5" 56'-9" - 55.29 2.30 4.63 V-5 76'-11" 19'-5" 57'-6" - 55.32 2.31 4.65
' V-6 77'-1" 20'-0" 57'-1" - 55.30 2.31 4.65 ' V-7 77'-4" 20'-8" 56'-8" - 55.40 2.31 4.65 77'-2" 19'-11" 57'-3" - 55.4 2.31 4.65 V-8 77'-1" 19'-7" 57'-6" - 55.4 2.31 4.65 I
V-9 V-10 77'-5" 20'-1" 57'-4" - 55.55 2.31 4.65 76 -11" 19'-4" s7'-7" - 55 3' 2 31 ' 6s i O v-11 V-12 77'-8" 20'-11" 56'-9" - 55.56 2.31 4.65 V-13 77'-3" 20'-5" 56'-10" - 55.36 2.31 4.65 v-14 76'-5" 29'-11" 26'-9" 19'-9" 55.72 2.41 4.85 V-15 76'-11" 19'-5" 57'-6" - 55.32 2.31 4.65 i 77'-4" 20'-4" 57'-0" - 55.5 2.31 4.65 V-16 7 2 ' -9" 32'-6" 3'-9" 36'-6" 55.0 2.22 4.47 V-17 5 V-18 79'-5" 28'-7" 50'-10" - 55.16 2.27 4.57 81'-0" 33'-5" 47'-7" - 55.3 2.27 4.57 V-19 l Note:
- 1. f = 0.015
- 2. (a) is normalized to 10" schedule 40 pipe
- 3. (b) is normalized to 12" schedule 40 pipe
- 4. Design constraints are listed in Table A.4.2.
- 5. The values are based on Figure A.10.2 (Safety / relief valve discharge l
piping arrangement) . (These line designs have not been optimized to take advantage of the maximum pipe pressure of 625 psid) . O 090779
>I OD $"
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8 5 9 5 9 5 e d _ 5 8 6 8 9 8 i p - - - - - - t s n p e3( m3 i n 0 1 ap t t n n o i 6 3 C o, 6 1 6 3 pP p . . 9 0 1 5 8 4 1 1 G 2 1 1 r o t t c naa T eF0 N m 1 6 6 1 A nd 1 1 5 1 2 L i et 1 1 P azn 7 7 8 9 0 8 t ~ i . . D nl o 0 0 0 0 1 0 R oaP A Cm D r@ N o A N T S 8 l ) _ 3 E d 2 V mi)
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- c. F s v1 eeo u l o u F C e l t nvi t l i t l D eo ea t t a At a so n vo v , re nae r (O e. 1 ta r ( a r eo o e l P l e u e u e vP s a an u cn u st t p st p l 0 a VF V o t ao t exc m ec m e
aF 1 C i a ji a veA e vA V0 e0 et r dt r l N T l T 0 t l 0 l a e A a e a t at S2 n g1 gu p u p V es l V s l D1 i n nt m ot m nr o r o A o it i c e wc e 0ii NF P o 9i 1 F P o 8 a t P S a SA T TA T 1
?Ma O
O O O Table A.4.5 l REDUCED QUENCIIER 15UBBl.E PRESSURE ttARK III, 238 STANDARD PLANT Containment Design Value-Bottom P P sm i
' Maximum Pressure (psid) Containment @ Point 10 (ps..O" p (+) '(-) Normalized Factor p -p, ' Case Description B g,B @ Point 10a Single Valve First Actuation, 10.8 -6.5 0.711 7.7 -4.6 at 100*F Pool Temperature Single Valve Subsequent 18.3 -7.8 0.711 13.0 -5.5 j
Actuation, at 120*F Pool Ter.perature . Two Adjacent Valves First 10.8 -6.5 0.856 9.2 -5.5
. Actuation at 100*F Pool Temperature f.h s~
, 10 Vcives (One 1.ow Set and (Not. Used for Design)
-{
Nine Next i.evel 1.ow Set) First Actuation at 100*F j Pool Temperature i 19 Valves (All Valve Case) 12.1 - fs . 4 1.0 12.1 -6.4 First Actuatfon, at 100*F ! Pool Temperature 8 ADS Valves First Actuation 11.3 6.8 0.821 9.3 -5.5 at 120*F Pool Temperature f " Point 10 on Containments is Peak Pressure.
@Basedon90-90% con!'idencelevelofCaorsodata.
e t
=
a o
=- - ---_-___w
22A4365 3_19 Rev. 2 S/H SLE EVE \v 3.0 CJ GJ ORYWELL WALL r d' GJ VENT HOLE (TYP) f) \ 7 n ;- ;
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! SECTION A--A
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PIPE SLEEVE & SUPPORT V ,
'..\\ 14" SCHED 80
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/f W.P.
( N \ 20.4'
-4 y, VENT HOLE' ~~ Ji (Hw L)
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+ 2. t r A+
(~) % Figure A4.1. 238 Standard Mark III Quencher Arrangement Elevatien 042178 1
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22A4365 A-20 Rev. 2 O i , I l 8O* g h*
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1200 /
/ ITYP) q
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r
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/
l VENT HOLES 'h
. AIR CLEARING JET (TYPI '
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974" O N
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.ke -
'n so Figure A4.2. Mark III Quencher Plan Viev (Typ.) 238 Plant Arrangement Shown 042118
l 22A4365 A-21 Rev. 2 m V10 (4.5'l V9 (346.5 I F047H g V11122.58) 1113 (1180) F041E - MPL IDENT 1123 - PRESSURE SWITCH SET POINT ips gt V8 (328.5*) I l (11651 - SPAING SET PCINT tosegl F0410 l l 1123(1165) d 4. % , I V12 449.5*6
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APV V151103.5 6 V 41256.5*) F 041 F 4- \ % F041G 1123111651 1123111651 [ [ O s N ORYWEa WALL N'
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V2 (211.5') +-_4 o V18 f 157.5 i F0418 / l l F051C 1123111651 ' V1 M935 i V19 (175 5 i F0478 FC41C 1113411801 1123(1165i 1800 LEGEND ADS
- NOTES. 19 S!A V ALVES i
Figure A4.3. S/R Valve Discharge Locations for 238-748 Standard Plant l O 042178 1 l l
22A4365 Rev. 2 A-22 0* l l l V10 (4.5*) ) V9 4346.5*l F0510 V 11 122.5*) i F047D 1103(11901 F047E - MPL IDENT 1113(11801 1113 PRESSURE SystTCH SET POINT (psql l (11801 - SPRING SET POINT fos.gl s a
/ gg@p g- V12149.5 I
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F047C 1123(11651 V1 (193.5 I V19 (175.5 1 1113(11801 F0519 F041C 1113t1130) 1123(11651 LEGEND. ADS
- NOTE.16 S/R V ALVES Figure A4.4 S/R Valve Discharge Locations for :'38-648 Standard Plant l 101678
22A4365
^~
~
Rev, 2 O . V9 (5.3') s V81344.1*l F047A 1113(1180) V10126.5's F0510 F041 A - MPL 6 DENT 1103 (11906 1123 - PRESSURE SWITCH SET POINT est (1165) - SPRING SET POINT (pset,i r V71322.9 ) ~- j ' F0470 # N I d 1113(1180)
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/ \ D l A V11458.2 i V6 (301.8*) B C F051G F0410 f s\ 1113 111901 1123(1165)
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V5 (280.6") V6 V13 I' F047F V4 ID U O b" 1113(1180) \ v3 l y34
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v0418 l 1123(11651 V 161164.1 ) g F041C 1123(1165) 1113 (1180) 1800 i LEGENo AoS -O NOTES, 16 $/R V ALVf 5 Figure A4.5. S/R Valve Discharge Locations for 218-624 Plant 101678
_ A-24 22A4365 Rev. 2 s,R SLEEVE h .
/I s o- i ^: ,- t . /
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PIPE SLEEVE & SUPPORT I - 14" SCHED 30 4 ; ,
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()- LC- ! b 0 q y p) 7 L - "$-- Y s.s-l j pl(I ' i I I$ ,, , , , h, ,;~, 9 -- ---- ---- - Y Y 7
+ 2.1 T &
A Figure A4.6. 218 Standard Plant Mark III Quencher Arrangement Elevation 042178 h
22A4365 Rev. 2 A-25 l 1 00 V10 (356*) VII (12 i F0510 F047A V9 (340 1 1103(11901 1113(1180) V12128*) F047H F041E 1113 (11801 112311165) V8 (324 ) F0410 g# r - [ 1123(1165) / ,g 1 [sg 8 %
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- b NOTES.
- 1. (20) $/R VALVES Figure A4.7. S/R Valve Discharge Locations for 251-800 Plant O
O 101678
22A4365 A-26 Rev. 2
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0 V21 (164 1 po4tB 1123(11651 23 1180) 90) l 1800 LEGENO: ADS = O NOTE S: 1. 22 SIR VALVES l Figure A4.9. S/R Valve Discharge Locations for 251-848 P' ant i ^ 1 1 101678
A-28 22A4365 Rev. 2 O AS. O QUENCHER LOAD ON POOL BOUNDARY A5.1 PRESSURES ON DRYWELL, BASEMAT AND CONTAINMETI Drywell wall, basemat and S/R Valve Loads are calculrted as discussed in Section A10.0. For the 238 Standard Plant, the maximum and minimum bubble pressure below the quencher just after air clearing are shown in Table AL.4. actuations (psid) . The absolute pressure on the pool walls can be calculated by the following equation: (a) " containment 44 there: P g) = Absolute pressure at point (a) (psia) O r = Distance from center of quencher to point (a) (ft) P = Absolute pressure of containme .t atmosphere (psia) thent h(e) = Head of water acting at point (a) (ft)
= Density of pool water % 62.4 (lb/f t }
iP g = Subble pressure attenuated by aistance, r to point (a), for multiple 3/R valve actuations (p s id) . The pressure decays with time and this is discussed in Section A3.4. 042178 O
22A4365 A-29 Rev. 3 (3 G) The following paragraphs discuss the dynamic pressure fields , at radial and circumferential locations of the pool for the 238 standard plant (Figure A4.3 and Table A10.2) . The pressure fields are based on P Bmax C normalized to 1 psid. These dynamic peak pressure fields can be used to reflect the changes in the maximum and/or minimum bubble pressure. If for example P = 25 psid for another SRVDL layout, the normalized Bmax values of Tables AS.1 through AS.5 sould be multiplied by 25 to obtain the design pressures. A5.1.1 Single S/R Valve Loads for a single S/R Valve The normalized dynamic peak pressures AP g Discharge valve are given in Table A5.1 and the normalized radial and circumferential peak values are shown in Figures AS.1, AS.2, and AS.2a. 3 (The values given presume an air leg volume of 56.13 ft for all SRVDS's). ' This is the base case and this pressure field is used to develop any other S/R Valve combination as described in Section A10.0 A5.1.2 Two Adiacent S/R Valve Leads The normalized dynamic peak pressures AP (r) are given in Table A5.2 and the normalized radial ar ; circumferential peak values are plotted in Figures A5.3, AS.4, and A5.4a for the two adj acent S/R Valves V-8 and V-9. A5.1.3 Ten S/R Valve Loads Sormalized aP (r) laods are given in Table AS.3 and the normalized values l are shown in Figures AS.5, AS.6, and A5.6a for the ten (1103 and 1113 psi low set point) valves V-10, V-12, V-14, V-16, V-18, V-1, V-3, V-5, V-7 and V-9. 090779 l
22A4365 A-30 R:v. 3 A5.1.4 Eight S/R Valve Loads (ADS) Normalized AP( ) loads are given in Table A5.4 and the normalized values are shown in Figures A5.7, AS.8, and A5.8a for the eight S/R valves , V-ll, V-13, V-16, V-18, V-2, V-4, V-7 and V-9. A5.1.5 All (19) S/R Valve Loads Normalized AP (r) loads are given in Table A5.5 and the normalized values are shevn in Figures AS.9, A5.10, and A5.10a for all (19) valves V-1 to V-19. AS.2 LOAD ON WEIR WALL The S/R valve loads on the weir wall are the sama as those on the drywell wall except they only act on the projected area through the drywell wall vents. O A5.3 LOADS ON S13 MERGED STRUCTURE For submerged structures, the loads are specified in Section G3 of Attachment G. O 090779
22A4365 A-31 Rev. 3 A-32 9 I l l i i i THESE PAGES ARE INTENTIONALLY DELETED i l l I 090779
l 22A4365 Rev. 4 A-33 P-A5.4 ' NORMALIZED PRESSURE TIME HISTORY (Theoretical Raleigh Bubble) t The ideal pressure is normalized for the maximum 3P(r) positive value as shown in Figure AS.ll. The frequency is 5 to 12 Hz as derived from the test data shown on Figure AS.12, and the total time of oscillation is 0.75 sec. (i.e., the time for the air bubbles to rise to the surface of the pool, or attenuation has dropped the amplitude to negligible values). Figure A3.ll is used by the designer for determining pressure amplitules- with time and ' the number of pressure cycles (see Section A9.0 fatigue cycles). It should be noted that bubble pressure decays to 1/3 Pmax occur in 5 cycles for any frequency between 5 and 12 Hz. For this linear attenuation rule it is observed that the pressure amplitude is fully decayed (P = 0 psig) in 7.5 pressure cycles after the peak. The justification for this application is from examination of full scale plant data where most traces were observed to decay to a small fraction of their peak value in 2 or 3 cycles. A5.5 REPRESENTATIVE PRESSURE TIME HISTORY Figure A5.12 depicts a representative pressure time history at points F1 through P4 as shown on Figure A4.1. These curves provide the designer-a realistic picture of the pressure oscillations as opposed to the idealized Raleigh bubbles. 4 AS.6 ESTIMATED MARGINS AS . 6.1 Peak Bubble Pressures For the examples shown in this document, the maximum loads on any structure resulting from the S/R valve air clearing phenomena are governed by the peak quencher bubble load. For the Mark III Standard 238 plant these values are shown on the next page. () A5.6.2 Peak Bubble Pressure Load Reduction I See Section A12.7 for information that supports use of reduced bubble pres-sure design values. 011880
22A4365 A-34 Rev. 3 Generalized Bottom Pressure Load Case
- A B C D
- 1. Predicted Maximum +8.8/-6.2 +12.3/-7.7 +11.5/-7.9 +16.1/-9.1 Bubble Pressure, psid (+/-)
- 2. Specified for +13.5/-8.1 +18.6/-9.9 *17.4/-10.4 +28.2/-12.0 Standard 238 Design, psid
(+/-)
- 3. Pressure Margin 4.7/1.9 5.3/2.2 5.9/2.5 12.1/2.9 4 % Margin (Based 35/23 34/22 34/24 43/24 on PIedicted Maximum Bubble Pressure)
"See Section A12.5.1 for load case description.
A5.6.2 95%-95% Confidence 95P.-95% means that there is 95% confidence that 95% of any new data obtained will fall within the maxi =um levels of the current data base. See Section A12.5.1.2 for additional discussion. A3.6.3 Margin The apparent margin in the specified containment design based on quencher bubble pressure is calculated as 20 to 45%. O 090779
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't - , i 22A4365 A-43 Rev. 2 f Table AS.4 (Continued) V-9 f S/R Valves j i /cgle (degrees) 355,5 s Xeference Point 328.5 337.5 346.5_ 4 0 0 0 0 1 5 I 0.801 0.950 1.0 0.950 2 3 0.894 1.0 1.0 1.0 } 4 0.912- 1.0 1.0 1.0 j I 0.845 1.0 1.0 1.0 5 i 6 0.885 1.0 1.0 1.0 i i 7 0.846 1.0 1.0 1.0 0.745 0.842 0.976 0.856 ! S f 0.657 0.707 0.761 0.710 9 t 0.681 0.745 0.821 0.748 j 10 0.675 0.735 0.803 0.738 4 i N 11 0.642 0.681 0.726 0.684 i 12 I 0 0 0 0 l 13 I. 8 042178 a - - - . ~ - _ _ ~ . _ _ _ _ _ _ _ _ _ . _ ___ __,_____..,_ _ _ _ _ _ _ _ _ _
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l 22A4365 A-66 Rev. 3 A6.0 OTHER LOADS ON STRUCTURES IN T'dE POOL O A6.1 LOCA A:ID POOL SWELL See Section 2. A6.1.1 Forces on Pipes Due to Vent Clearing Pool Swell mad Fallt ack The loadings are given for the quencher and reduced to ef fective pressure on a pipe in Table A6.1. The ef fective pressures of Table A6.1 can be applied nor=al to the projected SRVDL or sleeve areas to obtain the maximum design forces. These loads are included in the quencher anchor loads in Section A7.0. A6. 2 THER'!AL EXPA1SION LOADS Figure A6.1 gives the pressure and corresponding temperature for the SRVDL as a function of fL/D. The temperature can then be applied to the SRVDL for determining thermal e.:pansion loads. A6.3 SEISMIC LOADS (BY ARCHITECT-ENGINEER) The seismic loads are to be applied by the plant designer. These are included in Quancher Anchor Loads, Section A7.0. A6.4 SEISMIC SLOSH LOADS (3Y ARCRITECI-EElNLER) See Attachment B. 090779 O
1 } 1 22A4365 A-67 i 1 ! .Rev. 2 1 !O ! Table A6.1 I LOCA LOADS ON PIPES l 1 F* i P Force On Time Quencher Water Velocity Event (sec) (lbf) (ft/sec) Ref Water Clearing 0.1 to 0.7 30 Sec. 8.1.1 and Fig. G-3 4 ) Pool Swell 0.7 to 3 40 Sec. 8.1.2 1 i Fall Back 3 to 6 35 Sec. 8.1.3 i ! 2 i CD # . *F p = 2g (144) A 4 O - i i 1 l5 4 i I 1 042178 i O i s i
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22A4365 Rev. 2 A-69 A7.0 QUENCHER ANCHOR LOADS Figures A4.1, A4.2, A4.6 and A4.8 show the general arrangement of the quencher in the pool. GE has estimated anchor loads for a bottom quencher attachment and these are defined in Tables A7.1 and A7.2 and Figures A7.1 through A7.3, for the 238 Standard Plant. Both air clearing and water clearing load cases were evaluated, as they do not occur simultaneously. As shown in Figure A7.2 the anchor loads are specified at the base of the quencher and need to be translated to the basemat for embedment design. An additional adapting pedestal is required from the quencher bottom flange to basemat. There may be advantages to side pedestal anchorage to the drywell. These decisions and investigations are left to the Architect Engineer. /) s_/ The designer should evaluate the optimum location for anchorage of the RVDL to the drywell sleeve. The analyses should consider line thermal exp ansion. The designer should also c ,aluate the drywell penetration sleeve to assure that the drywell concrete local temperature limit is not exceeded. Preliminary thermal calculations for the 238 Standard Plant drywell sleeve show that concrete temperatures for normal operation do not exceed 200 F and 14" Schedule 80 sleeve is acceptable. Des igners should perform independent calculations to assure these findings. A7.1 QUENCRER ARM LOADS AND QUENCHER LOADING APPLICATION Table A7.1 lists maximum forces exerted on the sparger arms. Corre-spending points of force application are illustrated in Figure A7.1. In design of the sparger all of these forces shall be considered as acting simultaneously in directions present1ng a maximum loading condition. 042178 {a}
22A4365 A- 70 Rev. 3 Table A7.2 lists typical design loads for the Mark III quencher configuration. These loads consist of allowable inlet line loads, typical operating loads resulting from water clearing, air clearing, LOCA, and safe shutdown earth-quake loads. The resultant of these forces, which are considered to act simultaneously in a maximum loading condition, are expressed as base reac-These are typ-tion loads illustrated by Fg , F y, M gand M in Figure A7.2. ical design loads for the quencher supporting structure. Interf ace loads for plant unique conditions must be calculated and incor-porated into the overall plant design. A7.2 QUENCHER DESIGN INFORMATION Figures A4.1, A4.2 mad A4.3 show the quencher side elevation, top elevation and elevation and angular locations in the suppression pool. The following information is given to assist the designer in the design of a quencher. A7.2.1 Codes and Standards i
- a. American Society of Mechanical Engineers (ASME) Boiler and Pressure Code.
(1) ASME Section III, Naclear Power Plant Components
- 5. American National Standards Institute (A'!SI) i
! (1) ANSI B16.25, Butt Welding Ends for Pipe, Valves , Flanges , and Fittings.
- c. Anerican Institute of Steel Construction (AISC).
O 090779
22A4365 Rev. 4 A-71 A7.2.2 Design Pressures, Temperatures, Loads, Configuration, and Performance 4 A7.2.2.1 Component Data Safety / Relief Valve, Discharge Piping and Quencher: 4
- a. Design Pressure 570 psig i
s !- b. Design Temperatura 470*F . 2
- c. Maximum Pressure. 625 psig I
i j d. Maximum Temperature. sat. steam
- e. Maximum Flow 520 retric tons /hr at 1190 psig
- f. Maximum Steady State 40% of safety / relief valve inlet pressure l Back Pressure at rated ASME flow.
l l g. S/R valve Minimum 0.020 see Disc. Stroke Time _
- h. Minimum Ambient 60 F Service Temperature A7.2.2.2 SRVDL Geometry l
l (See Section A10.) O 011880
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v*- *ry-- y -- * - - -- - -:m- a w+-g ---+u--e+ m$7g~%?m
22A4365 A-72 Rev. 2 O A7.2.2.3 Quencher Design Criteria
- a. Forces See Figures A7.1, A7.2, A7.3 and Tables A7.1 and A7.2 ,
- b. Fatigue See Section A9.0 and Figure A5.11
- c. Cycles of operation See Section A9.0 and Figure AS.11 A7.2.2.4 Quencher Configuration and Location
- a. PROPRIETARY, Provided under separate cover
- b. PROPRIETARY, Provided under separate cover
- c. PROPRIETARY, Provided rnder separate cover
- d. PROPRIETARY, Provided under separate cover
- e. Quencher arm length 58.5 in, to CgQuencher
- f. Quencher pipe size / 12 in./Sched 80 (suggested) schedule
- g. Internal Quencher 101.6 sq in, pipe area
- h. Min clearance betseen >5 ft C Quencher and pool t
floor / basement 042178
22A4365 A-73 Rev. 3 ) O
- 1. ? lane of 4 Quencher legs 'dorizontal
- j. Angle between Quencher 80 , 80 legs for greatest 80 , 120 installation flexibility
- k. Corrosion allowance:
carbon 0.240 in. (0.120 per wetted side) stainless 0.0048 in. (0.0024 per wetted side)
- 1. Min submergence to 2/3 of min water level or 6 ft C Quencher min whichever is greater
- m. Design rating 625 psig
- n. Minimum clearance between .117 inches
)
Quencher and CCCS suction 4 4 a () 090779
22A4365 A-74 Rev. 2 O Table A7.1 QUENCRER ARM LOADS (Reference Figure A7.1) Load Descriotion Mark III Air clearing - (lbs) 116,460* (Location F,, any direction normal to arm centerline} Adj acent S/R - (lbs) 1974 (Location F 3 - horizontal direction) LOCA vent - (lbs) 1,866 (Location F , horizontal direction) b Arm weight - (lbs) 390 (Location Fe, downward direction) Earthquake load,1.25g - (lbs) at SSE 1488 (Location F , ver cal direction) c Earthquake load,1.0g - (lbs) at SSE 1390 (Location Fb , horizontal directical
*Due to single valve subsequent actuation.
042178 h
. 22A4365 A-75 Rev. 2 Table A7.2 MARK III QUENCHER ANCHOR LOADS (Reference Figure A7.2) l
- Air Clearing Water Clearing Lateral Loads - (lbs)
I F - ir and water clearing 28,510 8,553 b LOCA vent water clearing 10,240 10,240 F - art qu e a .g quenc er mass 3,%0 3,%0 e
- SSE Earthquake load (1.0g), water mass 1,680 1,680 F - Inlet line load 10,855 10,855 1
*F 2 - Total base lateral reaction load 53,545 35,268 J
O Vertical Loads - (lbs) F, - Air clearing 111,344 14,651 Transient wave 19,000 -3,700
-15,000 +2,400 Pool swell -14,742 -14,742 Quencher weight +3,940 +3,940
- SSE Earthqurie load (1.25g) 14,925 16,425 j F - Inlet line load 1
110,855 110,855 Water clearing +150,000/-2,000 40,064 178,271
*F y - Total base vert .d cal reaction load -56,866 -37,722 l
l 042178
22A4365 A-76 Bev. 2 O Table A7.2 (Continued) Air Clearing Water Clearing Lateral Moments Transferred to Base Plate - (ft-lbs) M - Air and water clearing 37,524 11,257 a Pool swell 17,751 17,751 Moments resulting from lateral loads - 2.64 x { Fb (air clearing) + LOCA vent clearing] 102,300 49,614 2.32 x F (earthquake, quencher mass) 9,141 _ 9,141 72,402 72,402 6.67 x F1 (inlet line) 5,040 3.00 x F, (earthquake , water mass)
- M - Total base lateral reaction =oment 239,118 165,205 7
Vertical Moments Transferred to O. Base Plate - (ft-lbs)
'!- b - Air clearing 105,618 31,685 Multiple valve actuation 0 0 LOCA vent clearing 8,047 8,047 M - Inlet line =ocent 25,836 25,836
- M y - Total base vertical reaction =ocent 139,501 65,568 i
- Quencher bottom flange anchor loads. (Individual loads are time dependent and peak values are conservatively combined.)
042178 h
i 22A4365 A-77 i Rev. 2 l (N () 3
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22A4365 A-78 Rev. 2 O ORTHOGONAL INLET LINE LOADS F; AND M; FOLLOW THE RELATIONSHIP:
- F; M; M; 5 Mo Fi WH E RE : Fo
= 10.855 lb M O= 25,836 f t Ib h
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- tOVENTS Ma AND Mg MAY ACT IN ANY VERTICAL PLANE Figure A7.2. Quencher Load Diagram 042178
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This figure is PROPRIETARY and is provided under separate cover. O l Figure A7.4. Sectional View of Quencher Leg (Typical Each Side) 042178 O
221.4365 A-81 Rev. 4 A8.0 S/R VALVE LOAD COMBINATIONS Safety / relief valve discharge piping routed to the suppression pool is arranged so that the points of discharge within the pool are uniformly distributed. (See Figure A4.3.) The location of valve discharge around the pool is for distribution of air clearing loads as well as for con-siderations of pool thermal mixing. The number of S/R valves that can open at one time is dependent on many variables. The following table shows several discrete cases where various numbers of open valves can be postulated for the 238 Standard Plant: Case Number of Valves (1) 1 Single active f ailure , normal function or operator action. (First or subsequent n ( ,) actuation) (2) 2 1 normal plus single active f ailure of adj acent (adj acent) valve (First actuation) (3) 10 All 11113 psi set point valves (First actuation) (4) 8 ADS Activation (First actuation) (5) 19 Vessel pressure >1123 psi. (First actuation) (All) The number of S/R valves that will open during a reactor vessel pressure transient could be from 1 to 19 valves. This can be shown for situations where various reactor power levels are assumed when the transient event is initiated. Therefore, the containment mus t be able to withstand any () f- s 011880
A-82 22A4365 Rev. 2 number of valves discharging at a given moment. Since the discharge points for valves with various setpoints, or those associated with ADS , are distributed around the suppression pool, the discharge of one or two valves represents an asymmetric load on the containment. A8.1 S M TRIC AND ASYMMETRIC LOAD CASES The following selected cases represent the asymmetric cases for contain-ment loads: A. 1 S/R Valve - This situation can occur due to an operator action or a single active failure. Subsequent actuation of a SRV af ter an initial pressure transient would be limited to the single 1103 psi set point valve. B. 2 Adjacent S/R Valves - This situation can occur due to a pressure transient at low power, which would lift one valve. Concurrent with this the single active failure of an adjacent valve is assumed. The probability of the precise combination of two adjacent valves would be very low, since co==on set point valve discharge points are uniformly distributed around the pool. However, if the containment structural design requirements are satisfied under this asymmetric condition, subsequent analysis need not be performed for the multitude of other more probable asymmetric load cases. The following selected cases represent the syc=etric cases for contain-ment loads: C. 8 ADS Valves - This situation can occur with an intermediate break where the ADS system is activated. 042178 O 1
l 22A4365 A-83 Rev. 2 D. 10 Valves - This event can occur due to a low power isolation transient. E. 19 (All) Valves - This event can occur due to a high power isolation transient. For structural evaluation the 5 load cases listed above are recom= ended. From observation of Figures AS.2a, A5.4a, A5.6a, AS.8a, and AS.10a the 1 or 2 valve load case is the governing case for asymmetrical considera-tions, and the 19 valve load case for maximum symmetrical consideration. The final selection of valve combinations is the designer's (A.E.) responsibility. A8.2 SSE AD OBE CONSIDERATIONS s b'hatever asymmetric or symmetric load cases are evaluated for design,
\- these should be combined with OBE and SSE seismic levels. The seismic combination which yields the controlling stress condition, may be either (OBE or SSE) since allowables and load factors .are different for the two conditions.
A8.3 LOCA CONSIDERATIONS In evaluating the Mark III structural loads and containment /drywell capa-bility it is necessary to properly account for the hypothetical accident related loads and their sequence of occurrence. In defining the loads for this evaluation, this report addresses the design basis accident (pipe l break) and the loads associated with the hypothetical concurrent earth-quake , pool dynamics , and static loading. The ability of the design to [ accommodate these loadings , when properly sequenced, constitutes the design basis of the structure. This design basis includes the single failure criterion; i.e. , any single component may fail to act when called l fs upon. 042178 l l
22A43o5 A-84 l Rev. 4 This report also addresses an additional consideration namely the inadvertent opening of a single S/R valve. The opening of a single valve is not a direct result of the LOCA and, furthermore, is not an expected occurrence during the accident sequence. However, the loading chart figures show the loads associated wit'h a single safety / relief valve actuation as an additional load for demonstrating additional capability. A8.3.1 DBA Nith M.S. Line Break For the DBA, with M.S. line break no valve will lift due to rapid vessel depressurization (Figure 4.1). A8.3.2 DBA With Recirculation Line Break For the DBA, recirculation line break, no valves will lift due to rapid vessel depressurization (Figure 4.1) . AS.3.3 other Snv Conditions O Other SRV conditions have also been analyzed at the forcing function level and their effect (except for SRV stcam condensation) were found to be less control-ling than the base case (CESSAR PDA) SRV loads. These other conditions result from a detailed analysis of pressure traces from Caorse SRV Quencher plant tests and the postulation of an SRV discharge under LOCA conditions. These other SRV conditions are:
- 1. A water clearing spike which precedes the air bubble pressure oscillation for the SRV quencher discharge
- 2. A leaking SRV discharge
- 3. SRV steam condensation
- 4. An SR" discharge for the LOCA related conditions is postulated with a pressurized drywell and wetwell.
011880 O
22A4365 Rev. 4 A-84a O The conditions are identified since the forcing function frequency range was
\~ / } generally broader but with much lower pressure amplitude than the normal SRV base cases described in A8.0 and A8.1. Analyses showed that the effects from these additional cases were generally less than the major bubble effects from the base cases even when load reduction factors of Section A12.7 are con-sidered. Comparisons are discussed in the followins sections.
A8.3.3.1 Water Clearing Pressure Spike For One SRV, First Actuation, Normal Operating Conditiens During Caorso SRV testing, a high frequency (15 to 30 Hz) pressure spike was ' observed just prior to the air bubble oscillation as shown in the typical trace on Figure A8.3-1. This spike occurs during the water clearing portion of the SRV blowdown. Multipliers were applied to the predicted bubble pres-sure amplitude such that the Caorso data, including the spike, would be I bounded at a 90-90 one-sided statistical tolerance limit and account for i / Mark III design conditions. O A comparison was then made of Amplified Response Spectra (ARS) for the forcing function including the water spike and the GESSAR PDA currently specified waveform (Figure A5.ll). The results show that the ARS for the GESSAR specified SRV discharge waveform bounds the ARS for the waveform which includes a water spike. In summary, the water spike observed in the Caorso data is not significant due to its short duration and limited number of cycles (one to three) and its effect is bounded by the waveform of Figure AS.ll. A8.3.3.2 First Actuation of One SRV With a Pressurized Containment For the case of an SRV actuation under small break accident LOCA conditions, i when the drywell and containment are pressurized, the initial water level in the SRVDL is depressed below normal water level. This lower water level is due to pressurization of the SRV line enrough the SRV vacuum breaker. Using a simplified model, the predicted bubble pressure forcing function results in 2 water spike pressures which are lower than predicted for normal operating () conditions, and air bubble pressures which are slightly higher due to an increase in air mass from the pressurized drywell condition. This case was 011880
22A4365 Rev. 4 A-84b also found to be bounded by the waveform of Figure A5.11 for single valve, first actuation loads when a comparison was made of the amplified response spectra generated from the bubble pressure forcing functions. A8.3.3.3 Water Clearing Pressure Spike for One SRV, Second Actuation, Normal Operating Conditions Second Actuation, normal operating condition SRV bicadowns are also characterized by a high frequency water spike followed by lower magnitude, lower f requency air clearing loads as shown in Figure A8.3.2. Second actuations occur with higher initial SRV discharge line (SRVDL) temperature, higher pool temperature, and lower air mass in the SRVDL. Second actuation . forcing functions were obtained by applying multipliers to the predicted bubble pressure amplitude such that the Caorso data, including water spike, i would be bounded at the 90-90 one-sided statistical tolerance limit and account for Mark III design conditions. Comparison of ARS for this pressure load to the waveform of Figure AS.ll for one SRV, second -e.tuation, showed the CESSAR PDA specification to be bounding. AS.3.3.4 Second Ac:uation of Cne SRV With a Pressurized Containment Subsequent actuation of SRV's under accident (LOCA) conditions are not predicted to occur; thus no load specifications are required. AS.3.3.5 First Actuation of One SRV, Leaking Valve Condition During the Caorso test series one SRV was found to be leaking. Several tests were conducted with this valve to determine the effect on bubble pressure. Results showed the typical water spike followed by low amplitude high frequency (20 to 30 Hz) random oscillatory behavior which was atypical of normal air bubble response. This trace is provided in Figure A8.3.3. An evaluation of this effect was performed using a typical leaking SRV data trace from Caorso. The pressure amplitude was increased to account for design operating conditions. A comparison of amplified response spectra was made for this leaking valve trace and the Figure A3.ll waveform for one SRV, subsequent actuation. The Figure AS.11 waveform was found to be bounding. h 011880
22A4365 Rev. 4 A-84c t
,r' The probability of leaky SRV actuation in combipation with LOCA is sufficiently small such that the leaky SRV is not specified in combination with LOCA loads (C.O. and Chugging) .
A8.3.3.6 Steam Condensation. One SRV During the Caorso testing of the SRV blowdown, steam condensation effects were observed after the air bubble oscillation phase. Beundary pressure ampli-tudes of 0.5 to 3.3 psid with typical mean values about 2.0 psid, and frequency content of 40 to 110 Hz were noted. These values also apply to Mark III. An evaluation of the effect of these steam condensation loads was made by select-ing a Caorso data trace with the highest RMS pressure value. The steam con-densation trace with the highest RMS pressure value is shown in Figure A8.3.4. This case was compared to the Figure A5.11 waveform by an evaluation of amplified response spectra. The case was bounded by the GESSAR PDA specifica-tion for one SRV subsequent actuation. For load cases in which first actua-tic, SRV loads are significant to design, the SRV steam condensation trace () shown in Figure A8.3-4 should be evaluated in series with the wavefo'rm of Figure A5-ll for first actuation.
- O 011880
22A4365 Rev. 4 A-84d l O l i i l l 1 l a i i i (CE COMPANY PROPRIETARY INFORMATION PROVIDED UNDER SEPARATE COVER) 4 i I l
)
l 1 i l 4 Figure A8.3-1. SRV Quencher Bubble Pressure - First Actuation 011880
i 6 i 22A4365 Rev. 4 A-84e j l lO i l 1 1 1 ; i l h i s s i i i l i i ('GE COMPANY PROPRIETARY INFORMATION PROVIDED UNDER SEPARATE COVER) i i 4 Figure A8.3-2. SRV Quencher Bubble Pressure - Second Actuation 011880
22A4365 Rev. 4 A-84f O\ i 1 (GE C0"PANY PROPRIETARY INFORMATION PROVIDED UNDER SEPARATE COVER) Figure AB.3-3. SRV Ouencher Bubble Pressure - Leaking Valve First Actuation 011880
j i 1 1 22A4365 l ! Rev. 4 A-84g i !O ) i J l i 1 i i < i i i i I i I l i ) i. (CE COMPANY PR01'RIETARY INFORMATION PROVIDED UNDER SEPARATE CO ] e t i l 1 1 I. i a Figure A8.3-4. Pool Boundary Pressure Time-History for One Second of SRV Condensation 011880
22A4365 A-25 Bev. 2 (- ) V 5 A8.4 RECOMMENDED DESIGN LOAD SUMMATIO'! The design loads on MK III structures are comprised of static (dead loads, live loads, hydro, etc.) and alternating dynamic loads (seismic and S/R valve loads, etc.). i For postulated simultaneous occurrence of S/R valve loads and SSE, the method recommended by the Task Group on Dynamic Analysis (TGDA) of the ASME Code comnittee for combining loads will be adopted: n 1/2 n R= )[ (DC) + )[ (AC) i=1 i=1 where g-V R = resultant response of the structure, e.g., displacement, accel-eration, load or stress. DC = slowly varying or non-alternating component of the dynamic response. AC = alternating component of response defined as maximum response value minus its corresponding DC component. i = 1, 2, . . . number of time varying events for which the resultant response is calculated. The use of this method is justified by the fact that earthquake excitation is a random process with amplitude increasing to a peak and thr.n decaying and the fact that the amplitude of the S/R valve loads also rise to a peak and then decay. Therefore, considering that the dynamic responses of O 042178
22A4365 A-86 Rev. 2 O
)
such loads possess varying frequencies, amplitudes and random phase relationship with respect to each other, this method is adequate for calculating the design loads. More simply, the above equation with respect to loads or stresses , etc. , may be represented by: _?
~,
. - ~
e smic , .SRV _Seisdc~ EF = D.L. + L.L. + Hydro + + AC AC DC DC where:
- D.L., L.L., Hydro = Static Loads
- S RV' Seismic' = w y vary ng r n n-a ernating DC) -
- DC) - component (DC) of the dynamic response
~
. SRV- ' Seismic = Altemating component of response defined i AC)
- AC) as the maximum response value minus the i
corresponding DC component. l l This is simply represented as follows: l , AC COMPONENT OF RESPONSE. l
,2(SE iC)
Axn An " A - NON-ALTERNATING OC COMPONENT OF OYNAMIC RESPONSE, SEISMIC
'S oC when the alternating component has no DC component, the DC tems drop out.
042178 9
l 22A4365 Rev. 4 A-87 ( A9.0 FATIGUE CYCLES A number of safety relief valve (SRV) discharge events may occur during the 40-year plant life. An analysis, based on many years of BWR plant operations, was performed to determine the mean frequency of occurrence of these potential events. Results of the study are summarized in Table A9.1. Transients which result in containment isolation are identified in the table. During isolation events the decay heat is initially removed from the reactor vessel via the SRV's. As shown in the table, half to all of the valves are initially actuated. Subsequently, the low-low set valve cycles until the RHR steam coudensing mode is established or the main condenser becomes available. is The total actuations of the low-low set valve during an isolation event
~
15 per event. The valve nominally remains open 80-90 seconds following each opening actuation. The non-isolation transients are also listed in Table A9.1. These events typically result in a single opening actuation of several or all SRV's, but the low-low set SRV does not cycle. When all valves are actuated, the open duration is 5-10 seconds. Considering 15 actuations for each isolation event and all of the non-isolation events, the total number of SRV actuations ia approximately 1800. Each actua-tion results in certain pressure pulses in the suppression pool which are transmitted to the containment as discussed throughout attachment A. For fatigue evaluation purposes the most significant forcing function on the con-tainment due to SRV actuation is the SRV bubble pressure. The normal bubble frequency range and duration are provided in Figure A5.ll. Using this figure in conjunction with the SRV actuations provides the number of cycles affecting the containment due to the SRV air bubbles. The SRV steam condensation pressure oscillation frequency is nominally specified as 80 CPS. This value is based upon.the pressure trace shown in Figure A8.3-4. i Utilizing this frequency, the noted durations per actuatic , and the number of actuations provides the cycles in the event this load has significant stress 011880
22A4365 Rev. 4 A-87a O affect on the containment. The phenomenon creates s relatively high number of cycles but the lower pressure amplitude forcing function typically results in low stress intensities. Hence, the containment fatigue factor due to this load is minor compared to the normal bubble effect. The discussed events, actuations, frequencies and durations are applicable to all BWR6 plants as low-low set logic has been incorporated. However, specific fatigue usage factors due to the various SRV forcing functions will vary as the factors are dependent on the unique stress analyses applicable to a given plant. The SRV loading fatigue effects should be considered by the contain-ment designer. O O 011880
l 22A4365 A-88 Rev. 4 O Table A9.1 SAFETY / RELIEF VALVE ACTUATION Number of Valves
- Open for Initial Blow Isclation Mean Frequency / (All (1/2- (1/3 Type 40 Years 2/ 3)_ 2/3) -0) Event Events x No Turbine Trip (w/BP) 53 x No Load Rejection (w/BP) 30 Pressure Regulator 26 x Yes Failure Feedwater Controller x No Failure 20 Trip of Both Recircu-lation Pumps 13 x No Recirculat. ion Con- No 13 x troller Failure 30 x No Loss of Feedvater Flow 15 x Yes Loss of Auxiliary Power 40 x Yes Closure of all MS1V's Loss of Condenser 26 x Yes vacuum Inadvertent Relief x No Valve Opening 4.0 0.5 x Yes Turbine Trip (w/o BP)
Load Rejection 0.5 x Yes (w/o BP) l 1 011880 1 1 1
i i f 22A4365 ! Rev. 4 A-89 i ie l j I i i i a This page deleted intentionally. l Y l. 1 1 i !O 011880 a
22A4365 A 00 Rev. 2 O A10.0 RECOMMENDED CALCULATION PROCEI'URES MR MARK III DESIGNERS The following information provides the procedares for predicting loads on the drywell wall, basemat, and containment wall associated with the air clearing transient folicwing the opening of a safety-relief valve for the 238 standard MAPI III plant. The nu=bers are applicable for those plants having a quencher of the standard design installed on the discharge end of the pipe. The given bubble pressures are based on infor=ation in Section A12.0. For design purposes , a statistical evaluation of the data was used. Design values represent a 957.-957. tolerance statement relative to that data. The bubble pressures are predicted for the first opening and consecutive opening cases. A10.1 CONSTRAINTS The following constraints are not to be exceeded for the design of the RVJL. (1) Peak Pipe Pressure lo25 psid. (2) p'L cannot exceed those values given in Figure A3.1 at the corre-spending pipe volu=e. (3) Water Leg 117.8 ft. Constraints on routing the safety / relief valve discharge line are:
- 1. No note than one 90 long radius bend cocing off the relief vrive, and two 45 Long radius bends entering the quencher in the 10" schd 30 piping. ~he remaining bends should be in the 042178 O1
22A43,65 A-91 l Rev. 3 r a ! C 12" schd 40 piping as far down stream as possible such that no more than 50% of the total fL/D of the system is in the first half of the length of the discharge line. l 2. The initial length of 10" schd 40 pipe be kept to a minimum. A10.2 DETERMINE SRVDL DESIGN The following steps are recommended for designing the SRVDL within the l above constraints and the design requirements in Table A4.2. l (1) Layout Preparation for SRVDL Routing The designer will prepare a layout drawing similar to Figure A10.2 and later detail the SRVDL. The longest line will be evaluated first. ( ! n (2) From the longest SRVDL length the air volume and fL/D values are calculated and plotted on Figure A3.1. This is an iterative l process where a balance of 10,12, and 14-inch SCR 40 piping is adjusted to the minimum total air volume and fL/D for the 625 psi pipe pressure constraint. It is important to insure l that all the SRVDL air volume and fL/D from the SRV to the free l l water surface is included. Figure A10.1 shows the portion of SRVDL from the SRV to the first anchor. ( (3) For the portion of the SRVDL shown in Figure A10.1, the loss coefficients, K, for each of the three flexible joints are. shown on the figure. The line lengths for each. plant size is given in Tables A10.1, A10.2 and A10.3. (4) Repeat the iterative process of (2) for each of the other I SRVDL. i O 090779
F 22A4365 A-92 Rev. 2 O (5) fL/D The corresponding maximum values of fL/D are calculated in reference to the 10" pipe velocities as shown below. Pipe friction losses should be considered from the S/R valve to the surface of the water. (a) For reference to 10" pipe velocities :
/A 10" I^10"
)+
y " 10" +y*I U"
= l l + y 1 fL/D)Re f l0" .
(Al2") 14" i( ^14") ~ who re :
~
bota1 10" 10" + kosses 10" S/40 b*atal g, = ft/P12" + kosses 12" S/40 K., = fL/D L*
,,, + N osses
.o tal g,, g.,S/40 A "
7# " # **
- 10" A
g., = IIydraulic area of 12" schd 40 pipe (f t ) A 14,, = livdraulic area o f 14" schd 40 pipe (ft ) D g,, = Diame te r o f 10" s chd 40 pipe (ft) D g., = Diameter of 12" schd 40 pipe (ft) D g,, = Diameter of 14" achd 40 pipe (ft) 0 042173
l 1 l l l 1 22A4365 Rev. 2 A-93 O The friction factor "f" in the abcve equations should be calculated based on the pipe diameter, relative roughness of the pipe, and the Reynolds number. A Reynold's number of 6 approximately 3 x 10 is appropriate. Based on this Reynold's
- number and the pipe of a commercial steel a typical value of "f" is 0.015.
Using the system fL/D calculated above enter Figure A3.1 with corresponding air volume. The intersection must fall on or above the 625 psid curse. (6) Determine the quencher bubble pressure using the actual air volume in the RVDL, see Section A12.6. A10.3 S/R VALVE AIR CLEARING LOADS MARK III 238 STANDARD PLANT fm) (, After the quencher bubble pressure has been obtained, Section A12.6, the next step is to calculate wall pressures based on the peak bubble l value (+ and -). l l l A10.3.1 Absolute Pressure on Basemat and Walls 1 t The absolute pressure anywhere on the drywell wall, basemat, and con-tainment wall La the vetwell region can be calculated by the equation: (" " containment
+
44
+ AP(r) (1) where P (a) = absolute pressure at arbitrary point "a" (psia) r = distance from quencher center to point "a" (f t) 042178
22A4365 A-94 Rev. 2 O P ,g
= absolute pressure of contain=ent at::osphere (ps ia) 9 g h(a) = vater head acting at point "a" (f t) a = water density (approx. 62.4 lbm/ f t )
a?(r) = bubble pressure attenuated by distance (r) to point "a". The attenuated bubble pressure for one S/RV, aP(r), can be calculated f rom the bubble pressure , aPB, w ch is obtained f rom Sec. tion C.6] l using tne following equations: l fr) AP (r) = 2 x AP B l l fo r r > 2 r o (2) r j a? (r) = SP for r < 2r o B - (3) g uhe re , r g
= quencher radius = 4.875 ft.
A10.3.2 How to Find the Attenuated Pressure on the Drvwell Wally Basemat, and Containment "all. A10.3.2.1 Develop grid to determine values of (r)
- 1. Make a scaled layout of the pool with quencher (Figure A10.3).
- 2. Divide wall distances by four (4).
O 101678
22A4365 A-95 Rev. 2 f 3. Arc distance by 360 + (vent stations) (Table A10.4) . ., 4. Draw line (Figure A10.3) from bubble cloud extremity (i.e. , quencher radius) tangent to drywell wall and project to con-tainment. This gives the area of pressure influence for this t quencher. 4 l 5. The point (a) is then selected and the distance (r) to (a) is obtained from the layout. A10.3.2.2 Wall Pressure at Point (a) Single S/R Valve. The wall pressures are obtained from A10,3.1 equation (2) and (3) . A10.3.2.3 Wall Pressure at Point (a) for Multiple S/R Valve. l In the event of multiple S/RV actuation the attenuated bubble pressure,
; AP , must be calculated using the following equations:
B , n ! AP(r) = L. [ AP
~
n=1 f where, o 2aP # AP = B I E l # # n o ( n/ SP = AP '# # 1 # B n o If the calculated AP(r) > SP3, set AM = aP B . N te the r n = s-tance from the center of the quencher to point a. 042178
A-96 22A4365 Rev. 3 O Fct the cases where multiple valves are discharged due to a pressure transient, the valves in each set point group (1103,1113, cad 1123 psi) are assumed to discharge simultaneously. The setpoint groups, however, will discharge at different times depending on the rate of reactor pressure increase associated with the event under consideration. The most severe pressure transient is the postulated " generator load rejection with failure of the turbine bypass valve" event which results in a calculated 132 psi . per second pressure increase at the beginning of the transient. This results in a 0.075 second dif ference in time of discharge due to the 10 psi dif ferenca in pressure setpoints of the valve groups. Using the quencher bubble model presented in Figure A5.11, it is seen that when P f rom the 1123 psi setpoint valves occurs , the bubble pressure from the 1113 psi setpoint valves has dropped to 0.9175 Pg, and the bubble pressure from the 1103 psi setpoint valve is 0.835 Pg. These values are used in determining the attenuated bubble pressure at a point (a) for the multiple S/R valve cases. For local peak containment pressure loading, there is significant reduc-tion in pressure at certain locations when considering the time sequenced phasing approach. The mest limiting position on the containment is not affected (i.e., the local peak pressure is equal to the maximum positive bubble pressure, 18.6 psid). In addition, the 95-95 confidence level statistical analysis for the individual valve is conservatively applied to the multiple valve cases without consideration of the number of valves being actuated. In reality, the 95-95 confidence total load for the 19 valve case is much lower than that used in the local pool boundary lead calculation. These two f actors (i.e. , time phasing and the multiple valve statistical consideration) have not been included in the develop-mer.t of the local pressure distributions on the containment vall because they do not affect the limiting local pressure. However, these factors i are important to the structural response and will be employed in the 1 building response evaluation. Attachment M presents the method for treating these effects in determining structural response used for the equipment evaluations. 090779
22A4365 Rev. 2 ^~97 Table A10.1 218 STANDARD PLANT PIPE SPOOL DIMENSIONS Dimension A Dimension B Total Dimension (A + B) Valve No. (in.) (in.) (in.) (ft) V1 73.38 82.50 155.88 13.0 v2 73.38 82.50 155.88 13.0 V3 73.38 82.50 155.88 13.0 V5 73.38 72.12 145.50 12.12 V6 130.00 71.62 201.62 16.80 f V7 118.38 71.50 189.88 15.82 i V8 115.88 71.00 186.88 15.57 l V9 115.88 71.00 186.88 15.57 V10 130.00 71.62 201.62 16.80 V11 73.38 72.12 145.50 12.12 V12 73.38 72.38 145.76 12.15 t V13 73.38 82.50 155.88 13.0 V14 73.38 82.50 155.88 13.0 V15 73.38 82.50 155.88 13.0 l V16 73.38 82.50 155.88 13.0 l l 1 The valve numbers shown on the table above are the same valve numbers on Figure A4.5. 042178 O
22A4365 A-98 Rev. 2 O Table A10.2 238 STANDARD PLANT PIPE SPOOL DDiENSIONS Dimension A Dimension 3 Total Dimension (A + B) (ft) Valve No. (in.) (in.) (in.) 75.00 82.25 157.25 13.10 V1 75.00 82.25 157.25 13.10 V2 75.00 82.25 157.25 13.10 V3 75.00 82.25 157.25 13.10 V4 75.00 72.00 147.00 12.25 V5 75.00 71.62 146.62 12.22 V6 V7 137.62 71.38 209.00 17.42 V8 126.75 71.12 197.87 16.50 V9 120.25 70.88 191.13 15.93 119.88 70.62 190.50 15.88 V10 9 Vil 119.88 70.62 190.50 15.88 V12 120.25 70.88 191.13 15.93 i V13 137.62 71.38 209.00 17.42 V14 75.00 71.62 146.62 12.22 V15 75.00 72.00 147.00 12.25 V16 75.00 32.25 157.25 13.10 V17 75.00 S2.25 157.25 13.10 V18 75.00 82.25 157.25 13.10 V19 75.00 82.25 157.25 13.10 l l The valve numbers shown on the table above are the same valve numbers on Figure A4.3. 042178 O
t 22A4365 A-99 Rev. 2 i fi d Table A10.3 251 STANDARD PLANT PIPE SPOOL DIMENSIONS Dimension A Dimension B Total Dimension (A + B) Valve No. (in.) (in.) (in.) (ft) 73.38 82.62 156.0 13.0 V1 73.38 32.62 156.0 13.0 V2 73.38 82.62 156.0 13.0 V3 V4 73.38 82.62 156.0 13.0 73.38 72.62 146.0 12.17 V5 73.38 72.25 145.63 12.14 V6 73.38 72.00 145.38 12.12 V7 V8 144.75 71.12 216.50 18.04 V9 133.50 70.88 205.13 17.09 V10 128.38 70.62 199.75 16.65 Vil 128.00 70.38 199.13 16.60 V12 128.38 70.75 199.13 16.60 V13 128.75 71.00 199.75 16.65 V14 133.88 71.25 205.13 17.09 V15 145.00 71.50 216.50 18.04 V16 73.38 72.00 145.38 12.12 V17 73.38 72.25 145.63 12.14 V18 73.38 72.62 146.0 12.17 V19 73.38 82.62 156.0 13.0 v20 73.38 82.62 156.0 13.0 V21 73.38 82.62 156.0 13.0 V22 73.38 82.62 156.0 13.0 The valve numbers shown on the table above are the same as valve numbers on Figure A4.9. 042178
A-100 22A4365 Rev. 2 Table A10.4 g DRWELL AND SUP?RESSION ?0OL GEOMETRY VE NT (V) ANNVLUS O SHIELD O ORvWELL E RPV ( g ' WALL --=- - E LE V O ft4 'a. dk PE DEST A L A -Y'
,8 SUPPRESSION POOL (SI "[ CONTAINME NT l WEIR WALL " I 0 50 f t 4
Os - HIGH W ATER LEVE L (HWLI _anywgLt , 'O***' ' ' ANNULUS _2 00b -7 dCE
- -- H 08 (W) ~
*[* - ,
g 4.50 ft v
'K * - .'
is . ;w. ,'. ,I BASE MAT - (( #
' 50 " I
-2 it 27.50 n.
T2 - - i VENT 10 TABLE 1 PLT SIZEiCNTMT OIA OR NO. OF FUE L BU,NOLE DE SCRIPTIONS 218/114 ! *218/120 *238 251/800 *251/864 f t-6 8 J3 t-18 33 (-1550 t-17 08 t-17 08 A ELEV TOP OF WEIR WALL 455 455 482 570 570 V VENT ANNULUS ARE A (ft21 5760 6863 6382 8170 8170 V+S TOTAL ARE A (f t2) (-l 12.58 :-) 12 58 t-) 11.16 t -i 13 00 (-)13 00 HWL HIGH WATER LEVEL ELEV 4 25 4 25 5 67 5 92 592 L vin FREEBOARD (f t) t -1 16 92 (-11692 (-11550 t-117 33 1-1 17.33 H DR AWOOWN LEVE L E LEV 19 42 19.42 20 42 19 00 19 00 Y POOL DEPTH (f t) III 00 131 90 129 60 15390 153 90 POOL VOL (1.000 f 3t ) AT LWL 21 80 17 31 3415 30 03 30 03 OR AWOOWN M AKEUP VOL (1000 ft3) (-) 29 08 t-129 08 t-) 27 67 t-l 29 50 (-! 29.50 J q OF 8OTTOM VENTS ELEV
- t -l 32 00 (-) 32 00 (-l 31 58 t-l 32 00 4-) 32.00 F ELEV TOP OF BASE MAT 102 102 120 135 135 NUMBER OF VENTS VENT STATIONS 34 34 40 45 45 4I0 420 495 557 557 GROSS VENT ARE A (f t2) 2100 2100 2475 2785 2785 VENT VOLUME (ft3) 16 16 19 20 22 NUMBER OF SAFETY RELIEF VALVE 6.38 6 38 5 75 5 25 9.25 CIRCUMFE RENTI AL VENT SPC'G DW 10 2223 2223 2535 2638 2088 W ARE A (f t2 27 84 27 84 39.29 39 82 38 75 W VOLUME (1000 ft3) 267.60 301.10 351.90 351.90 P AHE A (f t2) 267 6 5525 5525 6948 7477 7477 P VOLUME (ft3D t -) 29 Er8 t-l 28 98 (-) 28 58 (-628.33 (-l 28.33 07 RSO PT ELEV (REF) t-e 19 35 t-) 19 46 6-) 18 71 t-l 19 21 08 RSO PT ELEV f REF) t -6 19.35 t -l 12.35 (-l 12.35 t-i 12.46 (-11171 6-) 12.21 09 RSO PT ELEV (REF)
(-) 20 S5 t-120 85 (-) 20 98 1-) 21 50 ( ~ ) 21.90 15 RSO PT ELEV 'REF) 118 00 118 00 145 50 165 60 (LTRI 1100F STD REGO POOL VOL 11000 f t3) 102.40 102 40 124.20 141 90 (LTR) PLT 100 F VS SVCE WATER TEMP 8760 87 60 108 50 123 20 4 LT R) 900F TABLE 2 PLT SIZE C8(REF) 15 tRF F) A B C D E T T; 29 33 61 64 67 69 79 114
,18 DIA 18 42 1 83 2.17 RAD 9.21 14 91 30 50 32 33 34 50 39 50 57 OlA 18 42 29 83 61 64 67 69 79 120 1 83 2.17
,'O.
RAD 9 21 14 91 30 50 32 33 34 50 39 50 60 19 58 31 58 65 68 67 73 83 120 217 DIA 60 1 83 230 RAD 9 79 15 79 32 50 34 33 36 50 41.50 2117 32 67 67 70 75 85 130 2 50 DIA 42 50 65 1 50 251 RAD to 58 16 33 33 50 35 37 50 NOTES 1 PLANTS IDENTIFIED WITH (*) ASTERISK ARE STAND ARD PL ANTS 101678
? l i A-101 22A4365 Rev. 2 O i i a 4 I
- A s 0.19 r PIPE 10 in 5/40 l
s-"") J'" - g
. - - --N m_ , ,
~"
,,H"
< = 0.175
=0m e
/7 t
' THE RMOCOUPLE
'820.19 CONNECT ION (3) F L E xtBLE
- B AL L JOINTS
_j l K 0 077 i ! -- 1 7
*SE E T AB LE A101. A10.2 AND A ? O 3 FO A VALUES OF A AND B FO A THE 218. 238 ANO 251 ST ANDARD PLANTS Figure A10.1. Safety / Relief Discharge Piping Detail SRV to First Anchor O
042178
NOikS
- 1. EQUAL DIST RIBUTION OF QUENCHERS IN THE SUPPRESSION POOL IS REQUIRED FOR THE FOLLOWING F UNCTION.
(Al ADS VALVES (B) SPRING SET POINTS
- 2. NON-VERTICAL LENGTHS OF DISCH AMGE PLPE ARE LOCATED AS HIGH ABOVE THE SUPPRESSION POOL AS IS PR ACTICAL
- 3. SLOPE ALL D4SCHARGE PtPES TOWARD THE SUPPRESSION POOL 4 TWO VACUUM BRE AKERS ARE REOutRED. ONE IS LOCATED AT LEAST 10 FEET ABOVE THE WE4R WALL. THE OTHER IS LOCATED AS CLOSE TO THE SAFETY RELIEF VALVE AS PRACTICAL
- 5. VgiDE NTIFIES RE LATIVE VALVE LOCATION ON MAIN STE AM LINES 32 e-a CA y WE LL iN TE RIOR wa s.L m to
.,..ww
- ~ . ~ . .. <>-
s ._ g -
,,g,,,_-.._,r...,. ..
- 74 ,_.., ,g
\ .-a .* n .. . ,
i (
, p *<.- -q v - e - s u -.
3 ] y. 4, x .v . i , -
*i**'o- r
_.. _ a
'::c. .
r ..
~
\ s--w:,r -tq +sr'
= .au .... - ,,,. ~ ,; . o ...- 7 Q - - .i
,t.., ,
.x- , y:.s j .
v ~ ~.,., p~ ,r4 c>=p N i,, is/m> <N&k rI e.>r@.o, %6#. k~.?s . s. s. Nx _..m_ s% _ s A00. i-. . y I ;,; i '. s
.g i
, _t ...
sg ! % .@ . t l fy:r .-) - ' "- Q' ) s .t Ni , ._ ._.
. . 13h.. ..... . _.
J n>,
~
7 y, c,3, d) , c c, ,, . .
.* .e. .=.
(,.,, , . ... . . ... ... .*. .. . * * . . * . . * . . * . l- g m..,.,.
- w- .s
, ,f
. o f 3 8.*C48 R& v (
- o e
tJ D-*
>1
$ Figure A10.2. Safety / Relief Valve Discharge Piping Arrangement g tJ e O O O
O O O 21* 45* ga* 54 l 63 CONT AIN ME NT 4 4 0
#p, O
gO '"/ O O a s)
- v : , j' N y,
4 x \ t
\.
, DRYWELL 0 D. ,28 y s <
.t
\ my O'ST ANCE F ROM CENTER OF QUENCHER tan t IN ft.
REFERENCE POINT / ANGLE '00 90 180 270 360 450 540 630 720 810 13 - - - - - - - - - 12 16.1 18.2 23.2 29 .4 36 6 43 9 5OS 58.1 64 6 78.1 11 14.1 16 4 21.9 28 4 35 8 43.2 50 4 57.5 64.1 70.7 to 11.7 16.1 21.7 28.2 35 6 43.0 50.2 57 4 64 0 70 6 9 15.1 17.3 22 4 28 S 36 2 43 5 50.7 57.7 64.3 70 9 8 11 0 13 6 19.7 26.2 33 6 40 7 47.9 54 6 61.1 - 7 73 103 17.2 24 0 31.3 38 3 45.1 513 57.9 - 6 65 9A 16.1 22 S 29.7 36.4 423 49.2 - - 5 8.2 '0B 16.3 22.4 28.7 35 2 - - - - 4 52 87 15.1 21.5 28.0 34 6 - - - - 3 6.1 9.3 15.4 21.7 28.2 34 8 - - - - 2 99 12 1 17.2 23 0 29 2 35 6 - - - - 1 - - -------- v~ 5 Fla,ure A10.3. 238 Stantlar(1 Plant Distance f ron Center of quenchar to Pressure Point (ft) u
22A47A; A-104 Rev. 3 O A11.0 PARAMETRIC STUDIES The containment designer may choose to lay out the SRVDL such that equipment within the drywell can be accommodated somewhat differently than the GE Standard Plant. The application of the quencher data corre-lation allows for some flexibility in the pipe routing within the previously identified constraints. Generally speaking, the greatest flexibility exists in the routing of the air leg portion of the RVDL. Recom=endations for quencher location within the pool and the drywell wall penetration location minimize the flexibility in the water leg l portion of the SRVDL. To demonstrate the sensitivity of the changes to the air leg portion of the SRVDL, with all other parameters held fixed, Table All.1 has been generated. l The basic data correlation equation shown in Section A12.6 can be used by the containment designer to determine quencher design value bottom pressures for plant unique configurations. Af ter the bubble pressures have been determined, the procedures for determining suppression pool boundary loads identified in Section A.10.3 should be utilized. L I I l 090779 O
l A-105 22A4365 Rev. 3 ( Table All.1 QUENCHER BUBBLE PRESSURE SENSITIVITY TO SRVDL AIR VOLUME Bubble Pressure (psid)
#8 ^ "* b- ^- " **""*" ^ " " ' "
Ai V lume Maximum Allowable P-P+ P- P+ (ft3) fl/D at 10" S1140 Pipe 9.9 -6.7 20.9 -10.4 40 1.0 1.85 10.9 -7.1 22.9 -10.9 44 2.72 11.6 -7.4 24.2 -11.2 48 3.60 12.6 -7.8 26.4 -11.6 52 4.45 13.6 -8.3 28.4 -12.0 56 5.35 14.4 -8.6 29.7 -12.2 60 Standard Conditions: Steam Flow Rate (in.) = 520 metric tons /hr Pool Temperature (T y ) = 100 F (first actuation) 120 F (subsequent actuation) Water Leg, WCL = 17.8 ft (5.42 m) Valve Oconing Time, VOT = 20 msec. Quencher Submergence, SUBM = 13.92 ft. (4.24 m) 090779
A-106 22A4365 Rev. 4 O A12.0 BASIS AND JUSTIFICATION FOR DEVELOPED QUENCHER LOADS A12.1 INTRODUCIION To assure that the containment loads resulting from S/R ,alve discharge phenomena are conservatively low on Mark III containment, General Electric recommends a special discharge device in the S/R valve line discharge in the suppression pool. The device selected is called a " quencher." This device has been designed for application to pressure suppression contain-ments based on a series of small and large scale tests. The quencher arrangement is shown in Figures A4.1 and A4.2 and has been scaled directly from the large scale prototype. This section describes the basis for definition of the " quencher" per-formance in Mark III Design and Section A5 presents the resulting contain-cent pressure loads for the standard 238 plant. Included in this report is a test description and a su= mary of test data upon which the quencher lh design and perfor=ance are based. Full scale tests at the Caorso plant were conducted. Data evaluation shows that the SRV bubble pressures shown in Table A4.4 are very conserva-tive and that reduced values should provide adequate design margins. See section A12.7 for further discussion. The test data is shown in References 4 and 5. l 011880 O l
22A4365 A-107 thru A-124 Rev. 2
.O r
l O SECTION A12.2 CONTAINS GE COMPANY PROPRIETARY INFORMATION PROVIDED UNDER SEPARATE COVER 042178 O
22A4365 A-125 Rev. 2 A12.3 PHYSICAL PARAMERS , Due to the complexity of the phenomena associated with quencher perform-ance, it was not feasible to conduct the quencher tests such that the effects of various parameters on maximum bubble pressure could be studied one at a time. For instance, consecutive actuation of a safety / relief valve changes the local pool temperature, pipe temperature, velocity of the water column, and mass, temperature and steam content of the air column. Each of these changes will have an effect on the peak bubble pressure, but only the combined effect can be assessed from the data. It was therefore necessary to identify the important parameters phenomeno-logically and then determine the influence of each parameter statistically. The following sections explain the reasons for choosing the physical parameters that were used in the statistical analysis of the first actu-f_ ation data. As far as subsequent actuations are concerned, only the (_, maxi =um peak bubble pressure for any series of actuations is of concern. This maximum peak,was correlated with the peak of the first actuation. A12.3.1 Imoortant Parameters A complete list of parameters which might have minor effects on peak bubble pressure would be very long. It was therefore necessary to identify the most important parameters and include the rest only if they were l found to be statistically significant. The selection of important parameters was done by phenomenological considerations as well as the qualitative observations discussed in Section A12.2. Figure Al2.3-1 shows the interaction and interdependence of various parameters influ-encing quencher clearing. 042178 l s.
22A4365 A-126 Rev. 2 O A12.3.2 Overview of the Phenomenon Since the peak pressure on the pool boundaries is the main concem, we must first look for factors that influence containment air clearing loads, i.e.: (1) num!mr of quenchers discharging air simultaneously; (2) bubble size; and (3) peak bubble pressure. Peak bubble pressure depends on the number of bubbles originally formed ~ and their distribution in the pool. The same factors also determine the final shape of the bubble. Thus , the effects of shape of the bubble are included in the bubble pressure. O A12.3.2.1 Number of Quenchers As discussed in Section A12.2.1, increasing the number of quenchers has the same effect on the peak boundary pressures as reducing the size of the pool. The size of the pool must be judged by the size of the qt .ncher; therefore, pool surf ace area per quencher (A g ) divided bv the area of the quencher (A q
= the area of the circle that circumscribes the quencher) is an important parameter. The influence of this parameter would intui-tively be expected to diminish as the value of the parameter increases.
A12.3.2.2 Bubble Size For a given peak bubble pressure, bubble size depends on the mass and the temperature of the bubble. Since the bubble is essentially at pool temperature, one of the important parameters is pool temperature (Tg) . 042173
I l 22A4365 A-127 Rev. 2 C Assuming constaat initial corditions in the discharge line, the mass of the bubble is proportional to the initial air volume in the pipe. Since the air is spread over the quencher area, the important dimension becomes the height of tne bubble, which is proportional to the initial air volume (VA )/ quencher area (Aq ); hence, VA/Aqbecomes a key parameter. A12.3.2.3 Peak Bubble Pressure Due to expansion and heat transfer to the wetted surface of the discharge pipe and contact with the suppression pool water, the air is essentially cooled to. suppression pool temperature. Therefore , for a given air mass, the peak bubble pressure depends on: (1) bubble shape, and (2) mass flow rate. O A12.3.2.3.1 Bubble Shape Bubble shape refers to the outline of the bubble at the completion of the air-clearing transient. For a given mass flow rate vs. time, the shape of the bubble is strongly influenced by the quencher area, the distribution of the holes on the quencher arms and the manner in which the holes are uncovered. For quenchers that are geometrically similar to the large-scale test device and have comparable air-clearing dynamics, the bubble forms a flat circular cylinder. The dynamics of a flat bubble depend strongly on the thickness of the babble and is reflected in Vg/Aq. 042178 0. v
22A4365 A-128 Rev. 4 O A12.3.2.3.2 Mass Flow Rate For a given quencher, the mass flow rate of air into the pool is determined by the dynamics of the air-clearing transient, and the degree of mixing of air with water and steam. Since steam condenses almost instantly upon entering the pool, mixing of air with steam or water results in more gradual introduction of air into the pool (i.e. , lower mass flow rate of air and, there fore , lower bubble pressure) . However, systems of similar geometry (viz. , simple discharge pipes of large L/D) have comparable degrees of mixing. Thus , for a given quencher and a given air temperature , the mass flow rate of air depends on the discharge pressure which is a function of: (1) the length of the water column in the discharge pipe; (2) air volume in the discharge pipe; and O (3) steam flow rate from the safety / relief valve. For a given steam flow rate, the length of the water column is the main parameter affecting the peak pipe pressure, the air discharge pressure, and the velocity of water as it clears the quencher arms. The faster the water i expelled f rom the quencher arms, the faster the holes become availaole for air flow. Since the air flow rate depends on discharge pressure and on opening area, the length of the water column is a parameter that affects the air flow rate and, therefore, the peak bubble pressure. The air volume has already been identified as a key parameter; howeve , it should be pointed our that the effect of increased air volume in this case is to reduce the mass flow rate and thereby reduce the peak bubble 011880 0
22A4365 A-129 Rev. 4 i O V pressure. This is in the opposite direction of the effect of V /Aq l which was previously identified. In fact, as the air volume is increased. l l
. these opposing ef fects eventually cancel each other out.
l l It has been determined from the numerical solution of the air-clearing problem for various steam flow rates, that the discharge pressure is pro-portional to the maximum steam flow rate (m,) to a power of approximately l 0.7. The air flow rate (or steam flow rate) is converted into a mass flux to be suitable as a physical parameter. This is done by dividing the mass flow j rate by an area such as quencher opening area (defined as the total hole j area). For quenchers of similar geometry, the opening area is propor-l tional to the quencher area (A ),q and therefore, bubble pressure becomes a function of the mass flux across the quencher area (Aq ). To summarize, air mass flux depends onAV ' ( s * )/Aq and the length of the water i V column. Since the maximum steam flow rate occurs only when the valve is fully open, i valve opening time must also be considered as a parameter af fecting peak bubble pressure. However, as long as the valve is fully open before the water column is expelled, valve opening time does not significantly affect peak bubble pressure. A12.3.3 List of Parameters l To summarize, the following main parameters were identified as the ones that significantly affect the peak boundary loads : (1) Pool area per quencher / quencher area (Ag /A q ); (2) Pool temperature (T,); g
# 011880
.- -. .~- _
A-130 22A4365 Rev. 2 O (3) Initial air volume / quencher area (V /Aq ); (4) (Steam flow rate to the power 0.7)/ quencher area (5, * /Aq ); (5) Valve opening time (VOT); and (6) Lengths of water column (VCL) . Other parameters fall in one of the following categories: (1) Parameters which, wit'ain the range of the data, did not seem to have any effect at all on the boundary pressurer , such as initial air temperature, eccentricity of the quencher relative to a circular pool, and distance of the quencher from the bottom of the pool (with constant submergence). (2) Parameters which were properly scaled for all tests (excep t fo r a few miniscale runs) and held constant in GE quencher design. These include all important geometric properties of the quencher, such as arm length and diameter, size and arrangement of the holes, and quencher area. (3) Parameters that become important only for subsequent actuations (e. g. , pipe temperature and water velocity prior to valve actuation). The combined effect of these parameters is accounted for by the use of a statistically determined multi-plier applied to the first actuation loads. A12.3.4 Effects of the Parameters Each of the parameters identified in the previous sections af fects the peak bubble pressure, sometimes in more than one way. In what fo llows , ( these parameters and their ef fects will be discussed in more detail. O 042178
22A4365 A-131 q Rev. 2 V A12.3.4.1 Pool Area Per Quencher / Quencher Area (A g /Aq ) This parameter begins to have an effect only when a large number of relief valves is actuated simultaneously. The role of this parameter is to empirically account for wall effects and for the combined ef fect of multiple relief valve actuation. 1 A12.3.4.2 Pool Temperature (Tg) Part of the energy absorbed by the air column during the compression process in the discharge line is lost by heat transfer to the surroundings , and the remainder enters the bubble. The magnitude of pressure oscilla-tions in the pool depends on the energy contained in the bubble. There-fore, for high heat transfer rates, the magnitude of the pressures will be low. The heat transfer rate depends on the pool temperature and vanishes g when the air temperature becomes equal to pool temperature. The pool V v temperature, therefore, establishes the lower limit of the energy con-tent of the bubble at the end of bubble formation process. In other words, pool temperature affects the so-called " bubble formation efficiency" and, thereby, the peak bubble pressure. A12.3.4.3 Air Volume / Quencher Area (V /A ) The ef fect of air volume is rather complex. On the one hand, the air colu=n serves as a cushion to provide a low pipe-clearing pressure; on the other hand, more air means a larger bubble, more energy and higher i peak bottom pressures. The fact that very small and very large Vg both lead to negligible pressure changes on the boundaries suggests that peak bubble pressure must increase with increasing air-volume, reach a maximum and then decrease and asymp-totically approach zero, l 042178 i
A-132 22A4365 Rev. 2 O The strong influence of Vg /A on the peak bubble pressure implies that the thickness of the flat quencher bubble is indeed its characteristic length. This indicates that the bubble expands and contracts mainly in the vertical direction as a one-dimensional spring-mass system. A12.3.4.4 Steam Flow Rate and Valve Opening Time The discharge pressure increases with steam flow rate. Since the air flow rate, which is proport!.onal to discharge pressure, affects peak bubble pressure, the latter must also increase with steam flow rate. Condensation of steam on the walls of the discharge line has the ef fect of reducing steam flux. Consecutive actuation of the valve increases the discharge pipe temperature causing a reduction in the condensation on the pipe walls. This partially explains the increase in bubble pressure with repeated actuation. O Valve opening time affects the variation of steam flow rate with time. l IIowever, once the valve is fully open, the flow rate remains essentially constant for the remaindar of the air clearing transient. Because air-clearing occurs af ter the valve is fully open (in the range of practical values of valve openin; timal valve opening time does not significantly affect the ma:.n steam flow-rate or the peak bubble pressure. A12.3.4.5 Length o f Uater Column (UCL) The length of water column is the submerged length of pipe to the center of the quencher. The peak pipe pressure and, thet . fore, discharge pressure are both affected by the length of the water column due to the j
)
longer time required to accelerate and clear a large water mass compared j 1 042178 h
22A4365 A-133 Rev. 2 O to a small mass. The duration of the clearing of the quencher arms l depends on the volumetric flow rate of water, which also depends on water column length. The discharge pressure af fects the air flow rate. Therefore, WCL af fects air flow rate and, thereby, the peak bubble pressure. , h Another factor which depends on KCL is the wetted pipe area available for ' This wetted area, of cooling of the cogressed air during air-clearing. course, increases with WCL. This has the effect of reducing the energy e that enters the bubble and tends to counteract the previous effect of WCL (Figure A12.2-8) . { i d i O 042178
22A4365 Rev. 2 MASS , GF AIR I
. AIR l A6APRESSURE VOLUME l
'L T AND TEM PE R A TURE l BUBBLE WATER s h TEMPE R A. T TE MPE R A.
TURE TURE _ CONT AIN. YENT LOAOS 1 P OUENCHER - QUENCHFR GECVETAY AREA BUBBLE SUBBLE P AE SSURE f>
,f SIZE CONSTANT a L
- E
* *
- PIPE
, BUBBLE P R E SSU R E V ASS OF Ain ANO ip BUBBL.E TEYPE R A ARM CLE AR % air TURE CONST AN T ING TIME VOLUME a NUV8ER OF OsFNCHERS -
1 r
- a 6 AIR 9 ASS DISCH A R GE LENGTH
'
- FLOW RATF : : PRESS AND 2 CF w ATER T E VP COLUVN BU8BLE TEMP ' ' '
CONSTANT
- U A xtvUy REAC'OR STE AM P R E SSUR E Ad9 CLE AR -
I ING7'YE FLOW
'M ASSI Flow ARE A J L
} ' VALVE OPENING TtVE t l 8088LE g , , '* . m t NITI A L CONDITIONS GECYETRy AIR i CISCH A RGE
& LINE SECYETRY a
- QUENCHER - CUENCHER AREA GECMETRY
.* TATER SURF ACE TO C QUENCHER
'V AMV AlTH SUBSEQUE NT AC'U ATION AREA l et
- CONST ANT A ATIO Figure A12.3-1. Relationship of Key Parameters g 042178
A-135 22A4365 Rev. 2 Iv A12.4 CORRELATION OF POSITIVE AND NEGATIVE PRESSURE PEAKS Despite the complexity of the bubble dynamics for the quencher, a simple corre.'.ation exists between the peak positive and the peak negative bubble pressures. This correlation is based en the orinciple of conservation of energy and has been verified against t' .est data. The correlation provides a convenient means for determining one of the Being quite general, peak pressures, provided the other peak is known. it is applicable to bubbles of any geometry and pressure, regardless of the initial conditions in the discharge line, first or subsequent actua-tion of the relief valve. A12.4.1 Development of the Correlation and pressure Consider an air bubble of arbitrary geometry with a volume V g l
\ P, (same as local absolute pressure) in thermodynamic equilibrium with the corre-surrounding water. If the bubble is compressed to c pressure P and then allowed to oscillate , it will act as a sponding to a volume Vg and P spring-mass system. The pressure will oscillate between P and V .
and the volume will oscillate between V correspond Conservation of energy dictates that the minimum pressure must to the maximum pressure in such a way that the energy received during the compression is equal to energy transferred during the expansion, using the equilibrium state as the reference state: W = -W comp exp 042178 l0
- -~ --
22A4365 A-136 i l Rev. 2 or V V min max PdV = - PdV (Al2.4-1) v V o o Assuming the compression and expansion processes to be isothermal, the following relationship between P and V exists: VP=P V
= o where P, = absolute surrounding pressure ;
V = initial air volume at P, = P,; l P = instantaneous bubble pressure; and l V = instantaneous bubble volume. l Rearranging: P V
" O y .
P and P" V dp dV =
, (Al2.4-21 P'
O 101678
22A4365 A-137 Rev. 4 O Substituting in Equation A12.4-1 we obtain: P P
""* *I" dP max dP -PV in
- P, Vgp=- P ,V, p P, P, P P
= -P,V9 in = -P,Vg in p which simplifies to: ,
P P == mex , P, P O P P =P (A12.4-3) max tein
+ _ 9 For the case of interest, Pg, Pabs = P' abs where P
~
= minimum absolute bubble pressure, and
= maximu absolute bubble pressure.
P+bs a Notice that this relationship holds for bubbles of any shape and is not limited to spherical bubbles. Furthermore, any energy losses that will occur in the real case will tend to reduce both P,,x and P . That the process is properly considered isothermal is demonstrated by comparison to data. O 011880 j
l l 22A4365 A-138 Rev. 2
~
O In terms of gauge pressures (P+ and P ), Equation Al?.4-3, by simple algebra, takes the following form: (A12.4-3A) P~ = P P,/ (P + P ,) A12.4.3 Comoarison with Test Data Figure A12.4-1 shows a comparison of minimum absolute pressures predicted by Equation A12.4-3 and the actual measured values. Seventy data points from small-scale and large-scale tests have been plotted, covering a wide range of parameters. As can be seen, the agreement is quite good, indi-in actual cating that the combined ef fects of irreversibilities result bubble thermodynamics which are very well approximated by e reversible isothermal process. Notice that large values of negative gauge pressure correspond to small values of P~abs. The predictions for B'1R's will be in the lower end of the 45 line where the model gives conservative results. l 042178 O
22A4365 A-139 Rev. 2 l O 15 - r j I M i 0 0 .g a O
- w li e @
C O 3 10 - 00 0 5 2 4 SMALL SCALE,6 m WCL 7 SMALL SCALE,4 m WCL
- O LARGE SCALE, FIRST ACTUATION O LARGE SCALE.JUBSEQUENT ACTUATIONS 05 05 1o 1.5 20 PREDICTEO P 3 (bar)
Figure A12. 4-1. Comparison of Eq. (Al2.4-3) Predictions with Test Data 0 042178
22A4365 A-140 Rev. 3 O A12.5 DEVELOPMENT OF THE DESIGN VALUE CALCULATION METHOD A12.5.1 Introduction It is desired that design values be calculated so that, with a high confidence, a high percentage of actual values of maximum positive pressure (STP) and maximum negative pressure (MNP) will be less than the corresponding design values. The general form of such an equation, when based on test data, is to first calculate a predicted value, then add an amount which is the product of a confidence coefficient and a value which covers the uncertainty and variability in the test results. It is noted in the test data that subsequent, sequential actuations had higher MPP values than first actuations. Accordingly, equations are pro-vided for predicted values and design values for STP, for both first and maximum subsequent actuations. An equation to obtain SNP values directly from MPP values is also provided. A12.5.1.1 Objective The objective of this section is to develop the method for calculating the design value of maximum positive bottom pressure (MPP) and maximum negative pressure (MNP) at the quencher and on the floor immediately beneath, in the suppression pool of a BWR plant containment , due to oscillation of the air bubble discharged i==ediately af ter safety / relief valve actuation. The pressures are maximums ever the oscillations.
- TP and SCG are dif ferences above and below the absolute pressure at quencher elevation, where the absolute pressure is due to atmospheric and hydrostatic pressures. The generalized bottom pressure load cases of
~
interest are as follows: (A) first actuation of one or two valves (100 F suppression pool); (B) first actuation of three or more adjacent valves (100 F suppres-sion pool); 090779 I 1 I
4 22A4365 A-141 Rev. 4
/3 (C) first actuation of an ADS valve (120 F suppression pool); and i
(D) subsequent actuation of a single valve (120 F suppression pool) . Water surface area ratio distinguishes generalized load cases A and B. Similarly, the effect of water surface as well as pool temperature dis-tinguish case C. Generalized load case D must be distinguished because it was found from testing that the highest MPP and MNP occur en the second or third actuation of a valve, subsequent to the first actuation, when the valve is discharged sequentially with closure times of from 5 seconds to 1 minute. This consistent pattern for the maximum subsequent actuation is shown in Figures A12.5-1 and A12.5-2. Accordingly, design values will be found not only for the first actuation but also for the maximus of subsequent actuations. A12.5.1.2 Criterion i (
- The design values are to be such that there is 100Y% confidence that at least 100(1 - a)% of actual plant MPP (or MNP) values will be les: than the design values. Values for 100 % Y (confidence value) and 100(1 - a)%
(the percentage of the distribution of individuals) are both 95%. 4 This criterion implies that , if we should have complete knowledge of the distribution of actual MPP values, we would set the design value such that 95% of actual values are less than the design value. But the criterion further recognizes that , since we have but a finite amount of data, we must estimate that upper 95% point; but we will do so in a conservative manner such that we are 95% confident (100 %) that the true upper 95% point lies less than the one established. l 011880
22A4165 A-142 Bev. 3 O A12.5.1.3 Data Available The development of the design value calculation method for MPP is largely empirical; that is, based on the analysis of test data. Theory was drawn upon to identify variables of potential importance in prediction, and to normalize some variables for scaling dif ferences among the three sizes of test equipment. The design value calculation method for MNP, however, uses a relationship to MPP based on theory and confirmed empirically, as described in Section A12.4 and A12.5.14.2. One hundred sixteen data relevant to the plant quencher configuration were chosen f rom testing in three sizes of equipment, as follows: (1) First Actuations: 37 data from large-scale testing 70 data from small-scale testing 9 data from miniscale testing 116 (2) Maximum Subsequent Actuations: 10 data sequences from large-scale testing. A12.5.1.4 Str ategy of Statistical Analysis The design value calculation method for MPP is the result of a statistical analysis of the test data, conducted according to the following strategy: (1) Identify the measured variables of potential importance in ; 1 prediction. (2) Nor=alize some variables for scale dif ferences among the three test configurations and f;r application to the plants. l 090779
22A4365 A-143 Rev. 2 O x_/ (3) Determine the sensitivity of MPP to each variable simultaneously, in a prediction equation linear in coefficients estimated by multiple linear regression (curve fitting), thus maximizing the amount of data used to estimate each coefficient. Retain only 2 terms which make a statistically significant reduction in the variability of the observed values about the prediction surface. (4) Predict the first actuation iTP for the plant using a composite prediction equation comprised of the large scale mean MPP and a term for each variable which adjusts that mean from large-scale conditions to plant conditions. The coefficients were estimated on large, small- or miniscale data, depending on whether the variable was actually varied in that data. Accordingly, it is assumed that the plant quencher configuration is suf ficiently-similar to the configurations in the three tests that, af ter normalization of some variables, the sensitivity of MPP in the plants will be of the same magnitude as observed in each test. ("'S
Where possible, preference was given to selecting a coef ficient estimated on large scale data, because that test configuration used a quencher of physical dimensions near those of the quenchers used.
(5) When called for by the load case, predict MPP for maximum subse-quent actuations from the predicted MPP for first actuation. (6) To account for the uncertainty in estimates of coefficients and for the variability in individual values , find the variance of individual future values for plant conditions. There are two contributions to this variance: (1) the variance of the pre-
' dicted value, and (2) t e variance of individual values.
042178 n v
22A4365 A-144 Rev. 2 (7) Find the design value for each bottom pressure load case. The design value equation consists of the predicted value plus a confidence coefficient times the standard deviation future values. A12.5.1.5 Glossary
!TP = Abbreviation for maximum positive pressure
>CIP = Abbreviation for maximum negative pressure
>TP l = An observed value of MPP on a first actuation ,
MPPQ = An observed value of '.9P on a maximum subsequent actuation
>C;P 1 = An observed value of SCIP on a first actuation SciPQ = An observed value of SCIP on a maximum subsequent actuation O
PRD1 = A predicted value of MPP on a first actuation PRDQ = A predicted value of MPP on a maximum subsequent actuation PRN1 = A predicted value of MNP on a first actuation PR'iq = A predicted value of SCIP on a maximum subsequent actuation MPPDV = A design value for SIPP
- C;PDV = A design value for SC!P 042178 O
i l 22A4365 A-145 Rev. 4 ! rT l l l 4 i
%)
A12.5.2 Design Value Equatiocs for Maximum Positive Pressure and Maximum Negative Pressure Implementing the foregoing strategy, the design value equation for MPP appears in its basic form on the first line of Table A12.5.1. That table l goes on to give all subordinate equations and terms, and the design value l ! equation for MNP, together with the succeeding sections herein where each equation is derived, or each term evaluated. Thus , Table A12. 5.1 se rves as an index to the development of the design value equations for MPP and MNP. A12.5.3 Derivation of Equation for MPPDV (Maximum Positive Pressure Design Value) MPPDV = PRED + CONF x SIFV (3 where V MPPDV = the MPP design value in bars dif ference (bar d); PRED = predicted value (bars) ; CONF = confidence coefficient- and SlFV = standard deviation of individual future values (bars). This equation reflects a standard statistical relationship, when a design value is to be based on a predicted value plus an allowance for statisti-cal uncertainty and variability, implementing the design value criterion l stated in subsection A12.5.1.2. i 011880
)
22A4365 A-146 Rev. 2 Al2.5.4 Derivation of Equation for Predicted Maximum Positive Pressure (PRED) l PRED = CMSA x PRD1 where PRED = the predicted MPP: CMSA = the coefficient for maximum subsequent actuation: For first actuations, CMSA = 1.0 and PRED = PRDl. For maximum subsequent actuations, CMSA = 1.744 and PRED = PRDQ. PRD1. = the predicted >TP for first actuation. The evaluatior of PRD1 will be described first because of its fundamental rola. There are generalized load cases which involve only the first actuation, but even for the load case which involves maximum subsequent actuation, it is necessary to first predict the first actuation, for two reasons: (1) the dependence of MPP on most variables can be determined only for the first actuation because that was the only kind of test run in the small- and miniscale experiments, and most of the large scale experiments; and (2) it was found that for the 10 large-scale data having subsequent actuations, the maximum subsequent actuation could be predicted from the first in a simple, proportional manner. Evaluation of CMSA will be described in the succeeding section. 042178 O
22A4365 A-147 Bev. 2 O A12.5.5 Evaluation of Term PRD1 (Predicted First Actuation Maximum l Positive Pressure) The predicted value for first actuations for FTP is found from an equation resulting from fitting the experimental data. In Sections A12.2 and A12.3, theory was drawn upon to identify relevant variables and to normalize some variables among the three sizes of test configuration used. One-hundred sixteen data were available , 37 from large-scale testing, 70 from small-scale, and 9 from miniscale. A multiple regression (curve fitting) procedure was used on each data set separately, to estimate coefficients in an equation linear in coef ficients. Not all variables were actually varied in each experiment. For some variables, both first and second degree terms were important. No terms were retained which did not make a reduction in the variability of the data about the prediction surface, significant at the 1% level or less. Special treatment was given to twc variables. The MPP data for air volume ratio (VAAQ) were noted to
; first increase, then begin decreasing. But, for conservatism, the pre-diction surface was projected horizontally for higher VAAQ values, rather than decreasing, since some extrapolation to plant conditions is required.
For steam mass flux (MNAQ), since the three sets of data were widely separated in this variable, the principal fit was to large-scale data; but a straight line was then used to join large and mini mean oredicted values , slightly overpredicting the mean small-scale predicted value en route. Thus, in effect, MNAQ was used to provide for reprediction of all three data sets in one equation. Finally, a single prediction equation was composited from coef ficients ,
; estimated from the three sets of data. Since the large-scale configuration was closest dimensionally to that of quenchers used, coefficients estimated on that data were preferred, but use of coef ficients estimated on the I other two data sets was also necessar . The resulting prediction equatien i for MPP is shown in Table A12.5.2. Fi;u re A12.5-24 is a comparison o f ; observed to predicted MPP values for each of the three data sets.
(/>
- s. 042178 r
l l
22A4365 A-148 Rev. 2 A12.5.5.1 Raw and Transformed Variables Raw and transfor=ed variables used in tha fitting are. identified and described in Section All.5.2 and A12.5.3. Tha transfor=ed variables are. illustrated in Figura A12.54. Besides the transformed variables shown, two other variables were. fitted initially. One was air temperature. In the discharge pipe This variable. was found to not be. significant in predicting MPP. The other was pressure. upstream of the safety-relief valve. prior to actuation, which., in a plant, would be reactor pressure. In the small-scala data whera bot!i. reactor pressure and steam flow rate were measured, those two variables had a correlation coefficient of 0.6, a significant value; thus, normalized steam flow rate only was retained in. the prediction equation, The ranges of data from each. test configuration, and values for a typical plant, for the raw and transformed variables , are illustrated in Figures A12.5-4 and A12.5-5. The data, together with. predicted values, h are tabulated in Table A12.5.3. A12.5.5.2 Statistical Analysis The statistical analysis assured that the. sensitivity of MPP to each_ variable was governed by that sensitivity within a set of data known to be consistent, leaving the question of how to reconcile tha three sets of data in one prediction equation to a separate analysis. Accordingly, the multiple linear regression procedure was used on clie three sets of data separately. Least squares estimates of the coefficients were obtained. The so-called stepwise procedure was used, in backward stabilization mode, where.by variables are of fered for fitting, all are fitted at the outset, but the fitting criterion is then successively
=ade more discriminating so that in the end one. is lef t wit!t only those 042178
I i 22A4365 A-149 i Rev. 2 variables which make a significant contribution to the fit. Variables ' { j were retained if they were significant at the 1% level, or less. In 1
'this procedure, as each variable is forced out, the other variables are
- reexamined for possibly now making a significant contribution to the fit; this is a desirable feature because the inadvertant partial correlation i
between some pairs of variables means that one variable can, to some extent, play the role of a second, so that whether or not the second f variable is significant. There is, in general, no restriction en the form of each variable; a chosen variable in first degree, second degree (squared), first degree cross product with other variables, are all ! treated as separate variables in the regression procedure. Indeed, these possible forms w'ere systematically examined for their significance, ( and those forms having both statistical and physical significance were retained. Table A12.5.4 shows the data set from which a coefficient for each j variable could be estimated. There is duplication in the case of only
- three variables
- coefficients for ICAP, LNTW ano VOT could be estimated from both large and small-scale data. Per subsection A12.5.1.4(4) ,
estimates from the large-scsle data would be preferred. For !! NAP, the range of values from the small-scale data was so narrow as to not give
- a reliable estimate of the coef ficients; thus, the coefficients from the large-scale data were used. For LNTW, there was little to choose 1
between the large and small-scale coefficients ; the large-scale coefficient was slightly larger, leading to a more conservative predic-tion at higher temperatures, and so was chosen. For VOT, the coe f ficient . ' for this variable was not significant in the large-scale data and only
- barely significant in the small-scale data; for conservatism, the 1
coefficient from the small-scale data sas used. i l 042178 O
22A4365 A-150 Rev. 2 It was found that second degree terms were significant for MNAP, AWAQ
~
and WCL, reflecting curvature in the data. Such terms describe parabolas , o f course , which ma*- t suitable for extrapolation beyond the range of the data. Ae ly, special consideration was given to VAAQ and MNAQ for which preu.etion outside the range of the data would be neces-s a ry. These considerations are described in detail below, lhe second degree terms were called MNQ2, AWQ2 and WCL2, respectively. With all variables being fit simultaneously, in order to confirm that the term or terms for each variable are indeed filling the role called for by the data, it is helpful to see the pattern of the data points af ter adjustment by all terms in the prediction equation except one. These partially adjusted observed values , herein called shell residuals (in the sense that they form a shell for shewing the ef fect on prediction of those ter=s), are shown in Figures A12.5-6 through A12.5-13. Also shown, by smooth curves , are the role played by the term (s) for that variable. Conformance of these curves to the shell residuals indicates that the effect of that variable on MPP has been accounted for in the h term (s) used in the prediction equation. Figure A12.5-6 shows the shell residuals for VAAQ, with respect to that variable. These are for the small-scale data, from which the coefficient for VAAQ was estimated. Also shown are two straight lines, one the hori-
- ental continuation of the other, which show the effect of the VAAQ term in the equation. As can he seen, the shell residuals for VAAQ reach a maximum and then decrease. But rather than extrapolating this decrease, the prediction equation was aodified for conservatism to provide for a horiaontal projection for VAAQ values exceeding 0.255.*
*That the VAAQ values reached a maximum and began decreasing was found to be a statistically significant trend, and may be expected from physical Thus , the horizontal pro-considerations as described in Section A12.3.
jection is appropriate and conservative. 042178 9
22A4365 A-151 Rev. 2 r~3 b Thus, the same MPP value will be predicted for all values of VAAQ greater than 0.255, rather than decreasing values as indicated by the data. In Figure A12.5-7 the shell residuals from the large-scale data for the MNP and MNQ2 terms are shown, together with the effect of those terms on the prediction. Because the ranges of MNAQ for the three data sets differ so much, this variable was used empirically to perndt the prediction equation to predict the data in all three sets. Thus, a line tangent to the parabola at MNAQ = 6.89 was drawn so as to meet the predicted value of the mini scale data at MNAQ = 60.7. This line is shown over-predicting the mean predicted value for small scale data, an element of conservatism in the prediction equation in that the higher two mean pre-dicted values among the three data sets were permitted to govern predic-tion on MNAQ. () Figures A12.5-8 and A12.5-9 show the shell residuals for LNTW for the large-scale and small-scale data, respectively. The effect of the LNTW term is also shown. The coefficient used was estimated on the large-scale data, but it can be seen that the fit to small-scale data is also satis f acto ry. Figure A12.5-10 shows the shell residuals and ef fect of the water column length terms (UCL and WCL2) for the small-scale data. Figures A12.5-ll and A12.5-12 show the shell residuals for VOT for the small-scale and large-scale data, respectively. The effect of the VOT term in the prediction equation is also shown; its coefficient was esti-
=ated on the small-scale data, having been found to be not significant in the large-scale data, as suggested by the shell residuals' following the 0 line.
O (j 042178
22A4365 A-152 Rev. 4 Figure A12.5-13 shows the shell residuals for AWAQ for the miniscale data, together with the ef fect of the AVAQ and AWQ2 terms. The AWA0 model is l probably really asymptotic, but no use is made of it for valua of AWAQ greater than 20. Terms for the several variables were brought together into che single equation shown in Table A12.5.2, which predicts MPP for the first actua-tions (PRD1) . The structure of thh equation is illustrated in general terms in Figure A12.5-14 where, in Figure A12.5-14a, an equation linear in coefficients, in standard intercept form, is illustrated. The equation is
- -. s shown passing through (x 1
, y); and y, the predicted value of y at some x is also shown. But the standard intercept form is not convenient for a prediction equation composited from more than one set of data. Rather, the combination of mean-adjusted and reference-adjusted terms is used, as illustrated in Figures A12.5-14b and Al2.5-14c. Figure A12.5-14b applies to those terms where the coefficients' are estimated from the large-scale data, appropriate since y is also the average observed MPP for the large-s - s scale data. It shows the predicted y, 7, as y adjusted by a ay for x found as a t (k - k) . Figure A12.5-12 :: describes the ay's for terms with coef ficients estimated on data ot! er than large scale. Each of this
^
type of ay' is found as the term a 3 3 ~*3**O' *3 ref is the mean value of x3in the large-scale data, and x3 is the value of x3 at which l ys is being calculated. - xand 3 y are in the data set f rom which a 3was estimated. For the 238 Standard Plant, the value of each variable is shown in Table Al2.5.5. These values are entered in place of the variable names in the equation in Table A12.5.2, which are the x terms in Figure A12.5-14. The actual calculation of predicted values for first actuations (PRD1) is carried out for the load cases in connection with calculating design values , in Section A12.5,17. 011880 h
<f 22A4365 A-153 Rev. 3 Lm)
A12.5.6 Evaluation of Term CMSA (Coefficient for Maximum Subsequent Actuation) CMSA is the coefficient on PRD1 for first actuation for load cases involving maximum subsequent actuations. For load cases involving only 4 one actuation, CMSA = 1.0. Figure A12.5-15 shows the observed MPP values for the maximum subsequent actuations of the 10 runs versus the PRD1 values for the first actuations of those runs. The eight points without arrows are observed maximums which were, in fact, followed by lower values ; the two points with arrows are third subsequent actuations where that actuation was maximum but there were no further actuations. The important observation is that observed maximum subsequent actuations tend to be proportional to pre-dicted first actuations, rather than simply a fixed amount greater, for ex a=ple . That is, a line fitted through the points was found to have a slope significantly greater than zero. Since it would be physically reasonable for the relationship to pass through the origin, the predic-tion line for maximum subs;quent actuation from predicted first actuation was chosen passing through the origin and (x, i). Therefore , for load cases involving subsequent actuations, CMSA = 1.744 A12. 5. 7 Evaluation of Te rm CONF (Confidence Coefficient) CONF depends on the confidence statement to be made and en the number of data on which SIFV is based. The ci nfidence statement has the form written in subsection A12.5.1.2. A value o- 37 data points (the number of large-scale data) is used for first actuations ; a value of 10 data points is used for =aximum subsequent actuations , the number of those data. The corresponding CONF values for the 95-95 statement are 2.15 and 2.91. These values appear in Table A12.5.5, and are taken from standard tables for "one-sided statistical tolerance limits." g)3
\-
l 090779 f
22A4365 A-154 Rev. 4 l The confidence statement is valid when the distribution of individual lll l is values (in this case, of residuals about the prediction surface) normal. That this is nearly so in the observed data is shown in Figure A12.5-16, which shows cesiduals for large, small and miniscale first actuation predictions , and the large-scale maximum subsequent actuation predictions. The normal distribution corresponding to the histogram of maximum sub-sequent actuation residuals is considerably broader than suggested by those data in Figure A12.5-16. l A12.5.8 Der'ivaticn of Eauntions for SIFV (Standard Deviation of Individual Future Values) and VIFV (Variance of Individual Future Values) SIFV is the standard deviation of individual future values , and VIFV is the variance of individual future values: SIFV = (VIFV)1/', is the usual relationship between standard deviction and variance. VITV = VPRD + VIND (1) VPRD, the reflects the f act that VIFV is comprised of two parts: variance of the predicted value, and (2) VIND, the variance of individual values. This equation follows from the independence of the errors in predicted value and individual value as they appear in the usual error model in Figure A12.5-17. e 011880 ) (l>
22A4365 A-155 Rev. 4 O A12.5.9 Derivation of Ecuation for VPRD (Variance of the Predicted Value) VPRD, the variance of the predicted value, is found by propagation of errors on the predicted value: PRED = CMSA x PRD1. Propagation of errors is a general procedure for finding the variance of a function when the variance of each random variable in the function is known. For any function, y, of random variables , x , y = g(xg ), the propagation of errors equation for the variance of y, for errors inde-pendent among the x , is 4 n [ ay) 2 Var y = x Var x . { i i k / l at x i Application of the propagation of errors equation to the equation for PRED gives the variance of the predicted value: VPRD = VPRI + VPRM where VPRI is the contribution of the variance of the predicted first actuation: VPRI = sO12A) x VVP1 and VPP.M is the contribution of the variance of the predicted maximu= subsequent actuation (required by load case c) due to the variance in 01SA: VPRM = (PRD1) x VVTM. 011880
A-156 22A6365 Rev. 2 A12.5.10 Evaluation of Term VVP1 (Variance of the Predicted First Actuation) VVP1 is the variance of the predicted first actuation (PRD1). This variance is found by the equation shown in Figure A12.5-18, the standard expression for the variance of a predicted value from an equation found by multiple regression. In the first term, it reflects the variance in the intercept at the average of each of the independent variables (i.e. , the uncertainty in the vertical 9 cation of the prediction surf ace). And in the sums of terms, sometimes called the " flaring" tems , the expression reflects the variance of estimate of each coef ficient in the equation, and the covariances between all pairs of coef ficients which are not completely independent. Each of these variances and covariances can be computed as the product of an element in the so-called c matrix
- and the variance of residuals , both outputs of the multiple regression l computation. The c matrix is shown in Table A12.5.7. The c value for pairs of coef ficients estimated from different data sets is 0 in theory, and as confirmed by analysis. One special technique required was that the variance of residuals used to find each coefficient variance was that of the data set in which the coefficient was estimated, rather than the combined data set, taking advantage of the better precision of esti-mates in data sets having low residual variability. These dif ferent variances of residuals are subscripted t in Figure A12.5.18 and are tabulated in Table A12.5.7.
The k values on Figure A12.5-18 are the values of each variable for the plant. The values of x are the observed mean for variables whose coef-ficients were estimated on the large scale data, and the large-scale data value, x ref , for variables whose coe f ficients were estimated on other than large-scale data, just as distinguished in Figures A12.5-14b and A12.5-14c. For VAAQ, because the horizontal portion is greater than 0.255
- Inverse matrix of coefficients in the normal equations.
042178
l l l 22A4365 A-157 Rev. 2 l (O ; l and does not involve t.he coefficient, x' = 0.255 was used for any cases where VAA0 is greater than 0.255. For MNAQ, because the straight-line tangent does not involve the coefficients on MNAQ and MNQ2, x' = 6.89 was v. sed for any cases where MNAQ is greater than 6.89. A12.5.12 Evaluation of Term VVPM (Variance of Coefficient for Maximum Subsequent Actuation) ! VVPM is the variance of CMSA, the coefficient on the predicted first l actuation to obtain the predicted maximum subsequent actuation. Refer-l ring to Figure A12.5-16, VVPM is the variance of estimate of the coefficient 1.744 for load cases involving the maximum subsequent actua- ! tion. For load cases involving only the first actuation, where CMSA = 1.0, VVPM is not applicable and VPRM = 0. i The variance of estimate of the slope of a line through the origin is found by the standard equation shown in Figure Al2.5-19, and VVPM is evaluated in that figure as 0.01199. Al2.5.12 Derivation of Ecuation for VIND (Variance of Individual Values) VIND is the variance of individual values: VIND = (PROR x PRED) . Because the variance of individual values about the prediction surface is required beyond the range of measurement of some of the variables, it is necessary to consider whether the standard deviation of residuals is l apparently constant over all predicted values , or is in some way propor-I tional to predicted values. From studies of possible proportionality in all data from small and large-scale tests , both as originally fit within the data sets and as repredicted by the composite prediction equation in Table A12.5-2, it was determined that the standard deviation of residuals should be regarded as proportional to the predicted value , both for first v actuation and for maximum subsequent actuation, according to the prediction l l l l 042178 l [ i
A-158 32A4365 Rev. 2 11ia shown in Figure A12.5-20. The line is based on the maximum h subsequent actuation data. The average absolute residual within each 0.1 bar coll of the predicted maximum subsequent actuation is shown The proportionality of the plotted plotted versus those predicted values. points is clear, and is used within and beyond the range of predicted values shown. The standard deviation is obtainod by the equation shown on Figura A12.5-20, the 0.798 divisor being the expected value of the upper half of the nomal distribution, corresponding to the averago absolute residual when the residuals are normally distributed. The pro-cedure for calculating the expected value of the upper half of a standard normal distethution is illustrated in Figuro A12.5-21. A12.5.13 Evaluation of Tem PROR PROR is the coef ficient to multiply by the predicted value to obtain the standard deviation of individual values, (VIND) . Its evaluation is shown on Figuro A12.5-20 as 0.229. Thus, one standard deviation of indi-vidual values (residuals) is 22.9'." of the predicted value. g A12.5.16 Dorivation of Lluation for StNPDV Otaximum Negativo Pressure, Design Value)_ It was
- CIPDV is the design value for maximum negative pressure (>CIP ) .
dortved in Section Al2.4 as:
~
P P - Pf, .
~
tc P is the absoluto ,'ressure equivalent of StPP at quencher elevation, P the absolute pressure the absoluta pressure equivalent of BC1P, and P,,, is 042178 I O
1 22A4365 A-159 Rev. 3 (3
\_)
at quencher elevation (considering atmospheric pressure and hydrostatic
;>res sure) . To obtain an appropriate pressure difference value, MNPDV = PINF x MPPDV/(PINF + MPPDV) .
MPPDV is foand in Section A12.5.3. A12.5.15 Derivation of Ecuation for PINF and SUBM PINF is P, at quencher elevation, in absolute pressure units. Evaluated j in bars, it is: PINF = 1.014 + 0.0980 x SUBM , where 1.014 is atmospheric pressure in bars (14.7/14.5); 0.098 is bars per I meter of hydrostatic head; and SUBM is meters submergence at the centerline of the quencher. . The maximum negative pressure design value corresponding to any maximum positive pressure design value can be found using the above equations. A12.5. 6 Statistical confirmation of MNPDV The negative pressures were treated by the same statistical analysis procedure as that used for the positive pressure data discussed in this appendix. Through this analysis, it was confirmed that the predicted positive pressure can stand alone for prediction of negative pressure. The same independent variables used in positive pressure predictions were offered for fitting, together with the positive pressure, but none of these variables made a significant reduction in variability of the fit compared to the fit using positive pressure alone.
/~T 090779 'd
A-160 22A4365 Rev. 2 By way of further confirmation, the following two models were fitted to the maximum negative pressure OCIP) data:
>CIP = C7+Cy x MPP, and f P',)
P_ =C +C,* xl - j ( P+ ) MNP and MPP are the observed maximum negative and maximum positive pressure dif ferences, respectively; P" and P are observed maximum negative and maximum positive absolute pressures , and P, is the absolute pressure at quencher elevation. Both fits were highly significant and of identical Intercept C3 was not significantly different f rom 0, and C4was l l quality . not significantly different from 1.0, at even the 10% level. In application, the predicted maximum positive pressure must, of course, l be used for +P . Therefore, it is of interest to fit the negative pressure data using the predicted positive pressure values in place of measured l positive pressures. Such fitting of both of the above equations to large and small data gave fits which were significant and of identical quality, with C nd C4 again not significantly dif ferent from 0 and 1, respec-3 tively, confirming, from the data the appropriateness of the relation- , + l ship P P- = P',.
+ 9 The adequacy of ' P- = P[ using predicted maximum positive pressures can be confirmed visually for first actuations by comparison of the residuals for lar:;e and small-scale repredicted data in Figures A12.5-22 and A12.5-23.
Since there is only one term in the equation, shell residuals are not applicable. l 042178 Oi' l l
. _ _ . _ _=
.. . . . _ _ - - _ ~ . . . - - _ .
1 l 22A4365 Rev. 2
- O A12.5.17 Numerical Results for Maximum Positive and Negative Pressures e
I Numerical results for design value for each of the generalized bottom pressure load cases, using the values of Table A12.5.5 are given in Table A12.5.6. Figure A12.5-25 presents a graphical representation of ! the maximuu positive pressures to show the relationship between pre-i dicted and design values. i i . I 4 h O l 042178
. _ . , _ . _ _ _ _ _ . . _ _ _ _ __ , - _ _ _ _ _ . . . _ _ _ _ _ _ , - _ _ _ , - . . _ _ _ - _- _ . - _ - .--.1.
22A4365 A-162 Rev. 4 Table A-12.5.1 h DESIGN VALUE EQUATIONS WITH SUBORDINATE EQUATIONS AND TER'1S Section Equations A-12.5.3 MPPDV = MPP DESIGN VALUE = PRED + CONF x SIFV A-12.5.4 PRED = CMSA x PRD1 A-12.5.5 PRD1 A-12.5.6 CMSA l A-12.5.7 CONF l A-12.5.8 SIFV = VIFV A-12.5.9 VIFV = VPRD + VIND A-12.5.10 VPRD = VPR1 + VPRM A-12.5.11 VPR1 = CMSA x VVP1 A-12.5.12 WP1 A-12. 5.13 VPRM = PRD1 x VVPM A-12.5.14 VVPM A-12.5.15 VIND = (PROR x PRED) A-12. 5.16 PROR A-12.5.17 MNPDV = MNP DESIGN VALUE = PDF x MPPDV/ QDT + MPPDV)_ f PUU = 1.014 + 0.0980 2 SUBM A-12.5.18 The ter s are defined and the equations are derived in the. Sections shown. The indexes are defined in Section A12.5.3. l l 011880 0
22A4365 1 Rev. 2 A-163 and A-164 J Table A12.5.2 I EQUATION FOR PREDICTION OF PRDl, MAXIMUM POSITIVE PRESSURE FOR FIRST ACTUATION 1 1 l i 1 (GE COMPANY PROPRIETARY INFORMATION PROVIDED U'IDER SEPARATE C0\TR) . O I O 042178
22A4365 i
- Rev. 2 A-165 thru A-174 lO Table A12.5.3 DATA AND PREDICTED VALUES (10 Sheets) 1
) i 1 } } I (CE COMPANY PROPRIETARY INF0EMATION PROVIDED UNDER SEPARATE COVER) i j 1 0 042178 I
)
i < l
i i ) j A-175 22A4365 4' Rev. 4 !O i 1 I Table A12.5.4 4, l VARIATION OF VARIABLES BY TEST i Large-Scale Small-Scale Miniscale t } - )' Dependent: i 4 MPP Yes Yes Yes MNP Yes Yes Not i Reported i i 1 l Independent: f i No SEL No l VAAQ l.
>CiAQ SEL* Varied ** No I
i LNTW SEL Varied No i i SUBM No SEL No } VOT Varied SEL No , 1 AWAq No No SEL l 1 1 1
- Coefficient (s) estimated on this data set was selected for prediction equation.
** Variable was varied, but coefficient not selected.
I i i l, ' I 4 4 . 011880 i ) 4
, -a., on-- - -- ~,-w, -n-~ -e m,w , w - +s e ---wm=-e, w~ w,- - -- ----e,-m--,-----,r- , ~ . - - - - ~ ~ - ---n--- ~ - - - - - - - - - - - - - ,
A-176 22A4365 Rev. 3 Table A12.5.5 VALUES OF VARIABLES FOR STANDARD 238 MARK III PLANT Generalized Bottom Pressure Load Case
- b. First c. First d. Subsequent
- a. First Ac-tuation One Actuat ion Actuation Actuation All Valves ADS Valve Single Valve or Two Valves Parameter (100 F Water) (1000F Uater) (1200F Water) 11200F Water) 0.23 0.23 0.23 0.23 VAAQ ll.41 11.41 11.41 11.41
>CIAQ 6.89 6.89 6.89 6.89
- CiQ1 47.47 47.47 47.47 47.47
>CiQ2 11.41 11.41 11.41 11.41 MNQJ 3.63 3.63 3.89 3.89 LNTW WCL 5.42 5.42 5.42 5.42 WCL2 29.38 29.38 29.38 29.38 VOT 20.00 20.00 20.00 20.00 20.00 3.93 7.85 20.00 AWAQ 400.00 15.44 61.62 400.00 AWQ2 CONF 2.15 2.15 2.15 2.91 PINF 1.43 1.43 1.77 1.43 Air Volume (Vg ) = 1.59 m Quencher Area (A q ) = 6.93 m' VAAQ = V3/Aq = 0.23
aximum Steam Flow Rate (in) = 520 metric ton /hr
- C' Aq = m /Aq = 11.41 Temperature of Suppression Pool (T 7 ) = 37.8 C (100 F) or 48.9 C (120 F)
L:;W = 3. 63 or 3.89 Length o f Water Column ('JCL) = 5.42 m. NCL2 = (WCL) = 29.38 090779 0
I A-177 22A4365 Rev. 4 Table A12.5.5 (Continued) Valve Opening Time (VOT) = 20 msec. Effective Water Surface Area (Ag ) = 548.05 m (single valve) 54.79 m (ADS valves) 27.20 m (all valves) Water Surface Ratio (AWAQ) =gA q/A = 20.00 (single valve)
=
7.85 (ADS valves)
= 3.93 (all valves)
MNQ1 = MNAQ if MNAQ < 6.89 _ MNQ1 = 6.89 if MNAQ > 6.89 MNQ2 = (MNQl) MNQJ = MNAQ Quencher Submergence to Centerline (SUBM) = 4.24 m Containment Pressure = 14.7 psia (19.7 psia for ADS only)
= 1.0135 bar (1.358 bar for ADS only) i PINF = Containment Pressure + Hydrostatic Pressure Hydrostatic Pressure = 0.098 x SUBM
= 1.0135 + 0.4158 = 1.43 bar *
- = 1. 35 8 + 0. 415 8 = 1. 7 7 bar (for ADS only) 011S80
A-178 22A4365 Rev. 2 Table A12.5.6 g
'd. ARK. Ill PLAY. r VALUES FOR A STANDARD 238 b c d a,
Generalized Bottom Pressure Load Case 13.44 18.73 17.40 28.13 MPPDV (psid) 0.927 1.29 1.20 1.94
- @PDV (bar d) 0.603 0.851 0.790 1.11 P RED 0.603 0.851 0.790 0.639 PRD1 1.0 1.0 1.0 1.74 CMSA 2.15 2.15 2.15 2.91 CONF 0.151 0.205 0.191 0.284 SIFV 0.0227 0.0421 0.0364 0.0805 VIFV 0.00357 0.00407 0.00363 0.0154 VPRD 0.00357 0.00407 0.00363 0.0105 VPR1 0.00357 0.00407 0.00363 0.00345 VVP1
- 0. O. O. 0.00490 VPR>i NA* NA NA 0.0120 VVPM VIND 0.0191 0.0380 0.0327 0.0651 0.229 g
0.229 0.229 0.229 PROR 8.15 9.84 10.38 11.93 10;PDV (psid) 0.562 0.679 0.716 0.823 30iTDV (bar d) 1.43 1.77 1,43 PINF 1.43 4.24 4.24 4.24 4.24 SUB31
*NA = Not Applicable.
042178 t l
22A4365 A-179 Rev. 4 O Table A12.5.7 c MATRIX VALUES * (Page 1 of 2) c gy 3.08E-02 l c i2
-8.91E-04 2.75E-05 e
g3 0 0 2.52E-01 eg 0 0 -6.47E-02 2.42E-01 c 0 0 3.23E-02 -2.93E-02 3.67E-03 , 15 c 0 0 0 0 i6 c g7 0 0 -3. 4 8E-04 -5.30E-06 1.624E-07 0 0 0 0 0 "18 e gg 0 0 0 0 0 c
"lj "2j "3j "4j 5j c 2.86E-01 i6 c 0. 4.22E-07 7
c 1.09E-02 0. 2.37E-01
, c 8.12E-04 0. -1.953E-02 1.653E-03 9
*6j "7j 8j "9j 1 AWAQ estimated on miniscale data 2 AWQ2 estimated on miniscale data 3 VAAQ estimated on small-scale data 4 WCL estimated on small-scale data 5 WCL2 estimated on small-scale data 6 LNTW estimated on large-scale data 7 VOT estimated on small-scale data 8 30'AQ estimated on large-scale data 9 5c:12 estimated on large-scale data ij ji ;
e for coefficients estimated on different data sets = 0
- Inverse matrix of coefficients in the normal equations. (See Figure Al2.5-18.)
i I 011880
}
A-180 22A4365 Rev. 4 Table A12.5.7 (Continued) Residuals (!@P) Mean (bar d) Variance _ Standard Deviation First Actuations
-0.01070* 0.00938** 0.0969 37 Large-Scale Data 0.00325 0.00927** 0.0938 70 Small-Scale Data 0.000832 0.000493** 0.0222 9 Miniscale Data Maximum Subsequent Actuations 0 0.01188 0.1090 10 Large-Scale Data Residuals DC!P)
"ean (bar d) Variance Standard Deviation First Actuations 0.00479 0.00892 0.0945 37 Large-Scale Data 0.0187 0.0021.4 0.0463 70 Small-Scale Data
*In least squares fitting, the mean of residuals = 0. These are not exactly 0, due to use of an equation involring coefficients estimated on sets o f data other than on line shown.
f
- VRESID g1 in Figute A12.3-18.
011880 0
22A4365 A-181 Rev. 2 l l (GE COMPANY PROPRIETARY INFORMATION PROVIDED UNDER SEPARATE COVER) Figure A12.5-1. MPP for First and Subsequent Actuations Positive-- Large Scale O (GE COBEANY PROPRIETARY IN70RMATION PROVIDED E DER SEPARATE COVER) Figure A12.5-2. M';P for First and Subsequent Actuations Negative-- Large Scale 042178
l
/t, j A-182 22A4365 Rey, t j~
/ -
V
,l / ,
/ - vor ._,
)
f f AWAQ - VNAQ ' SUBM l
} p__f___ ,
__q - l ..... m. .. .. ,, Ix_. j . . .-: h .....
^
______J - I L_[ l LNTW VAAQ MNP MPP Figure Al2. 5-3. Quencher Diagram Illustrating Dependent and Independent Variables 042178 O
/
A-183 22A4365 Rev. 2 O . A Q. m2 (QUENCHER ARE A) 1 6 7 9 9 2 3 4 5 0 1 MINI e SMALL e e LARGE o STO PLANT VA m3 (AIR VOLUMEl 0.4 0.8 1.2 1.6 2.0 2.4 0 MINI 8 SMALL
# LARGE
# STO PLANT VAAQ (AIR VOLUME R ATlO, VA/Ag m) 0.15 0.20 0.25 0 30 0 35 0 0 C; 0.10 8 MINI SMALL LARGE e
8 STD PLANT Figure A12.5-4 Ranges of Values / Raw and Transformed Independent Variables (Page 1 of 4) O 042178
A-184 22A4365 Rev. 2 5 STE AM MASS F LOW R ATG (tonneWhr) 100 200 E q 500 I O
, VINI l
i
- SMALL LARGE j
# STO PLANT 4
507 0 10 20 30 40 SC 60 70 80
- MINT 4
- SMALL i LARGE e STO PLANT O
1 ^ MN AQ, M R ATIO [Itonnes/hr)0.7/ m2] ! 0 10 20 % 40 50 EO 70 e MINI
- SMALL j LARGE e STO PLANT i
I 4 Figure A12.5-4. Page 2 of 4 042178 i
22A4365 A-185 3 Bev. 2
- O I VOT (V ALVE OPENING TIME, ms) l I
600 800 1000 1200 1400 1600 O 200 400 MINI e GMALL i LARGE i i STO PLANT e I Tw (WATER TEMPERATURE. 'Cl 40 50 60 70 80 0 10 20 30 e MINI SMALL LARGE e e STD PLANT O a,b c.d LNTW(LOG,Twl 20 2.5 30 3.5 40 45 e MINI SMALL LARGE e e STD PLANT a.b c .d Figure A12. 5-4. Page 3 of 4 i 042178 ) O : l l
l l l 22A4365 A-186 Rev. 2 O WCL, WATER COLUMN LENGTH (ml 2 3 4 5 6 0 1 MINI e O SMALL e LARGE e STO PLANT Aw, m2 (WATER SURFACE AREA) 8 12 16 20 24 28 0 4 MINI SMALL j e e LARGE e e STO PLANT b a.c.d i AWAQ (WATER SURF ACE R ATIO Aw/Ac m2/m2) 15 20 25 30 0 5 to MINI J SMALL e e LARGE e STD PL ANT
. e a,c o d i
Figure A12.5-4. Page 4 of 4 i 042178
22A4365 A-187 Rev. 2 O FIRST ACTUATIONS OBSERVED MAXIMUM POSITIVE PRESSURE 0.3 0.4 0.5 06 0. 7 0.8 O O1 0.2 MINI SMALL LARGE LARGE. FOLLOWED BY SUBSEQUENT ACTUATIONS l MAXIMUM SUBSEQUENT ACTU ATIONS OBSERVED MAXIMUM POSITIVE PRESSURE 0.2 0.3 04 0. 5 0.6 0.7 08 O O1 LARGE FIRST ACTU ATIONS OBSERVED MAXIMUM NEGATIVE PRESSURE O O1 02 0.3 0.4 0.5 0.6 0.7 0.8 SM ALL LARGE LARGE. FOLLOWED BY SUBSEQUENT ACTU ATIONS
)
1 M AXIYUM SUBSEQUENT ACTUATIONS OBSE RVED M AxiMUM NEGATIVE PRESSURE 02 0.3 04 05 06 0.7 0.8 O 01 LARGE Figure A12. 5-5. Ranges of Values; Dependent Variables O 01.2178
22A4365 A-188 Rev. 2 O (GE CTTPA'IY PROPRIETARY I'IFOR'1ATION PROVIDED C4 DER SEPARA Figure A12.5-6. Shell Residuals, VAAQ Omitted, and Ef fect of VAAQ Term on Prediction: Small Scale Data (Coefficient Estimated on These Data) O (GE COMP A.IY PROPRIETARY INFOR'1ATION PROVIDED CIDER SEPARATE CO Shell Residuals , 'UAQ Omitted, and Ef fect of CIAQ Terms j Figure A12.5-7. on Prediction; Large Scale Data (Coef ficient Estimated r l on These Data) 042178 1 0'
22A4365 hev. I w (GE COMPriY PROPRIETARY I:4FOR5tATIO:( PROVIDED CIDER SEPARATE C9VER) Figure A12.5-8. Shell Residuals, CITti Omitted, and Effect of CIT 4 Term on Prediction; Large Scale Data (Coefficient Estimated on 3ese Data) 7x
'w (CE CO'T,CY PROPRIETARY I'iF0P.'!ATIO:I PROVIDED CIDER SEPARATE COVER)
Figure A12.5-9. Shell Residuals, CIT'i ')mitted, and Ef fect of CITil Term on Prediction; Small Scale Data (L:ITii Coef ficient Es tima te d from Large Scale) 042178 i O L) l . _ .
A-190 ' l 22A4365 Rev. I O (GE COM' ANY PROPRIETARY I'IFORMATION PROVIDED U.4 DER SEPARAT Figure A12.5-10. Shell Residuals, WCL Omitted, and Ef fect of FCL Terms on Prediction; Small Scale Data (Coef ficients Estimated from These Data) 9 (GF. c0MPA"Y PROPRIETARY I'! FORMATIO:{ PROVIDED E4 DER SEPARATE COV Figure A12.5-il. Shell Residuals, VOT Omitted, and Ef fect of VOT Term on Prediction; Small Scale Data (Coef ficient Estimated from These Data) 042178 O
- . . . .~ - - , - . . . . . _ . .. - - - - - . - - - . - . . . -
i i 22A4365 A-191 i Rev. 2 IL 1, , l l (GE CTIPANY PROPRIETARY INFORMATION PROVIDED INDER SEPARATE. COVER) 4 i i k . i 1 Figure A12.5-12. Shell Residuals, VOT Omitted, and Effect of VOT on i Prediction, Large Scale Data (y0I Coefficient Estimated i from Small Scale; not Significant in Large. Scalel. iO ' i (GE COMPATI PROPRIETARY INFORMATION PROVIDED ITJDER SEPARATF. COVER) l i Figure A12.5-13. Shell Residuals, AWAQ Omitted, and Effect of ANAQ Terms on Prediction, Miniscale Data (Coefficients Estimated ! from These Data) i f 042178 1 O I I
22A4365 A-192 Rev. 4 O e _ FIGURE A12.514a STANDARD INTERCEPT FORM: 7=AO + Ag kj + A 2 2* Ag l l X1 X1 X1 FIGURE A12.514b Y YE AN ADJUSTED FORM: ~ 37, ,
~ y
~T _= +%AY140 Y ' %d 714C 1 I aY 14b = A t ($ 1 X 11 7
+A2 (X3-X2)
+
7 FRCM LARGE SCALE A,'s ESTIMATED ON LARGE SCALE DATA I i X1 X1 X1 O Y FIGURE A12 514c AEFERENCE ADJUSTED . FORM: , 5 A ' 14c
- A3(X3 - X3 REF) +aY Y
. A4 tX4 x4 REF) -
A, s ESTiM ATED ON DAT A OTHER T- AN LARGE SCALE I l i X3 X3 X3REF X3 s Figure Al2.5-14. General Forms of Prediction Equations 011880
4 22A4365 A-193 Rev. 2 O (GE COMPNTY PROPRIETARY INFCMfATIO:7 PROVIDED LWDER SEPARATE COVER) O Figure A12.5-15. Observed )!PPQ Maximum Subsequent Actuations , MPPQ Versus Predicted First Actuations , PRD1; Large Scale Data i 042178 . O
22A4365 A-194 Rev. 2 l LARGE-SCALE FIRST ACTUATION 37 i 5
=
s===
~
i 1 0{ lO t i 0 t 0.1 i 02 03 1
-02 -0.1
- SMALL SCALE 70 I -
5 .
'l 1 I I I II I I
i g 0. 3 0 0.1 02 $ -02 -01 z w MINISCALE 9 5 I I I l I I I g 0 0.1 0.2 03
-02 -0.1 LARGE SCALE M AXIMUM SUBSEQUENT ACTU ATION 10 5
=
I l 1 I 0.1 02 03
-02 -01 0 R E SIOU A L (O BSE R VE D.P R E DICTE Oi (barl Figure A12.5-16. Frequency Distributing Showing Normality of Residuals 042178
22A4365 A-195 Rev. 2 1 AN INQ1Vf0UAL FUTURE V ALUE'
= TRUE VALUE FOR PLANT + ERROR IN PREQlCTEQ VALUE + ERROR I IN INQiVIOUAL VALUE A80UT PREQlCTEQ VALUE OUE TO INDEPENCENCE OF ERRORS.
VARIANCE OF INOlvlOUAL FUTURE VALUE'
= 0 + VARI ANCE OF PREDICTED VALUE + VARI ANCE OF RESIQUALS i
(VIFV = VPRO + VINO)
'VALUE OF FIRST OR MAXIMUM SUBSEQUENT ACTUATION MPP Fig'Jre A12.5-17 Error Model (VAR RESIOi, = LARGE SCALE VVP1 =
+ 1 d,- D Va , + 2 I E li,- 7,8 tij 7) COV (c a )
3 g
, , 3
'<8 WHERE Va, = VARI ANCE OF ESTIMATED CF COEFFICIENT a,
= C,, V RESici, CO V ia ,.a g)
= C,3 (VRESlot, X VRESIQg31"2 WHERE t,. tj REFER TO THE SET CF TEST D ATA ON WMcCH THE COEFFICIENTS c , AND C g WERE ESTIMATEO. ANO CjiANDC4 REFER TO ELEMENTS IN THE C MATRIX V ALUES OF VRESIO t, AND C.,. ETC.. ARE TABULATED IN T ABLE A12.5.7 THE ECUATION FOR VVP1 ABOVE IS FOR A PREDICTION EQUATION IN THE 80RM OF THE ECUATION FOR YIN FIGURE A12.514b Figure Al2.5-18 Variance of Predicted Value MPP, First Actuation Evaluated at x 0 Values VARI ANCE OF SLOPE THROUGH ORIGIN-VRESiO MSA VVPM
- llLLUSTR ATED IN SIGURE A12 E.201 IPRot 2 0 01188 0 991
= 001199 Figure A12. 5-19. Variance of Coef ficient (1.744) for Maximum Subsequent Actuatiun g
( 042178
22A4365 A-196 Rev. 2 O (GE CO' IPA'iY PROPRIETARY INFOR!!ATION PROVIDED IT.4 DER SEPARATE COVER) Figure A12.5-20. Proportionality of Average Absolute Residuals and Predicted Values , 'taximum Subsequent Actuations f" f(z) d: AVER AGE ABSOLUTE DEVI ATION = =05 f" f(zldr I WHERE f (z) = P~ II y 2r THE STANDARD NORMAL PROBABILITY DENSITY FUNCTION APPLIES TO NORMALLY AVG ABS DEV = 0 798 DISTRIBUTED INDIVIOUAL VALUES USED IN THE EQUATsON FOR IVIND11/2 IN FIGURE A12 5 20. I i I I I I 1 !I I
-3 -2 -1 0 1 2 3 STANDARD NORMAL DEVIATES rigure A12.5-21. Derivation of Ratio of Average Absolute Deviation to Standard Deviation l
l 042178 I O 1
A-197 22A4365 Rev. 2 l l O ! (GE COMPANY PROPRIETARY INFOR'!ATION PROVIDED UNDER SEPARATE COVER) e Figure A12.5-22. Residuals for MNP Large Scale Data O V (GE COMPANY PROPRIETARY INFORMATION PROVIDED U'iDER SEPARATE COVER) Figure A12.5-23. Residuals for .EiP Sc:all Scale Data 042178
A-198 22A4365 Rev. 2 i O\ f i 4 i (GE COMPANY PROPRIETARY I'iFORMATION PROVIDED UNDER SEPARATE COVER) l i e i a I I Figure A12.5-24. Observed is Predicted Values, MPP (MPPI Versus PRDI or "PQ Versus PRDQ) 042178 O
22A4365 A-199 . Rev. 2 30 lb MPPOV MPPOV = MA xiMUM POSITIVE PRESSURE. DESIGN VALUE PROQ
- PRE 01CTED MAXIMUM SU8 SEQUENT ACTUATION MPP
\
25 - PRD1 - PREDICTED FIRST ACTUATION MPP j i 1 l l , $ 20 - a b ,
)b MPPOV 3 bb MPPOV $ - PROQ (PR EO)
[ E i5 - 8
\/
2 = MPPOV 3 2 E
- - PRO 1 4
E -
- PRO 1 10 -
- PROl
- PRO 1 5 -
C a n e d GENER ALIZED BOTTOM PRESSURE LOAC CAS. Figure A12. 5-25. Predicted Values and Design Values of Maximum Positive Pressure from Table Al2.5.6 042178 I
22A4365 A-200 Rev. 3 Al2.6 APPLICATI03 The purpose of this section is to provide the designer with a simple and straightforward procedure for calculating the maximum positive and negative air-clearing pressures on the bottom of the suppression pool beneath the quencher. These pressures are to be used in the development of suppression pool boundary loads for the design of the containment. The development of boundary loads is discussed in Section A10. Al2.6.1 Procedure All bottom pressures obtained by these procedures have a 95-95 confidence level, and are within +1.0% of the values obtained by strict applicatica of the techniques described in the previous chapter. The first step in determining the bottom pressure is to calculate the predicted first actuation maximum positive pressure (PRD1) . Since the l quencher device is a fixed design (Area = 6.93 m ) , the maximun flow rate llh is 520 metric tons per hour, and the safety / relief valve opening time is at the minimum (20 msec). The equation for PRD1 in Table Al2.5.1 can be i reduced to: i PRD1 = 0.421
+ 2.58 (yAAQ - 0.1706).
+ 0.1377 (LNW - 3. 83) l
+ 0. 206 (WCL - 4)
- 0.0176 (TTCL2 - 16) l
- 0.0336 (AWAQ - 20)
+ 0.000761 (AWQ2 - 400) 0 090779 l
22A4365 A-201 Rev. 4 whe re PRD1 = mean first actuation peak positive pressure (bars); VAAQ = air volume in the safety / relief valve discharge line 3 ? (m ) divided by the quencher area (m'), where quencher area is defined as the area of a circle chat circum-scribes the quencher. (For the standard BWR/6 Mark III plant, the quencher area is 6.93 m . If VAAQ is greater than 0.255, use VAAQ = 0.255. ) ; LNTV = Natural log of Tg , where T g is the suppression pool temperature ( C). WCL = The actual length of the water leg from the centerline of the quencher arm to the air-water interface in the discharge pipe; WCL2 = (WCL)~ ; { AUAQ = The effective pool surface area per quencher divided by the quencher area, where quencher area, as the VAAQ, is l 6.93 m (If AWAQ is greater than 20, use 20.); and i AVQ2 = (AWAQ) . The above formula allows for the plant unique location of tne quencher in the suppression pool and for plant unique routing of safety / relief valve discharge piping within the constraints identified in Section A.10 and, !- as stated above, is only applicable to the quencher design described in j this attachment, r\ V 011550
A-202 22A4365
~
Rev. 2 Using the value determined for PRDl, the corresponding maximum positive design pressure (MPPDV) is obtained from Figure A12.6-1 or A12.6-2. Using MPPDV, the negative design pressure (SCIPDV) 1 = obtained from the following equation: MPPDV (PINF + MPPDV) where MPPDV = Design Positive Bottom Pressure (bars); SC4PDV = Design Negative Bottom Pressure (bars); PINF = Absolute Pressure at the Level of the centerline of the quencher arms (bars abs.) To convert pressure from bars to psi, a conversion factor of 14.5 psi /bar g is used. A12.6.1.1 Development of Figures A12. 6-1 and A12.6-2 The maximum positive bottom design pressure (MPPDV) is a function of PRDl. This functional relationship was described in A12.5 and can be summarized in the following form. 042178 g
l 22A4365 A-203 Rev. 2 O . 1 I i (GE COMPA'iY PROPRIETARY INFOR'!ATION PROVIDED UNDER SEPARATE COVER) Al2.6.2 Examples Given the following input, the design bottom pressures for the four cases described in Section A12.5.1.1 are calculated: Air Volice = 1.59 m Pool temperature = 37.8 C (100 F) for cases a and b 49.0 C (120 F) for cases e and d 042178 l
22A4365 A-204 Rev. 4 Water Colu:::n Length = 5.42 m h Submergence to Centerline = 4.24 n 2 Ef fective Pool Area per Quencher = 548 m for case a and c 27.2 m for case b 54.8 m for case d A12.6.2.1 Calculation of Variables VAAQ = = 0.23 93 l LNT'1 = Ln 37.8 = 3.63 'for cases a and b La 49.0 = 3.89 fer cases c and d l WCL = 5.42 NCL2 = 29.38 g 548
= 79.08. Therefore, 20 is used for cases a and d. {
NJAQ = 6.93
'7 j,]'=3.93forcaseb 34.8 '00 ?*' '
- 6.93 "
AWQ2 = (20) = 400 for case a and d ( 3. 93) = 15.44 for case b (7.85) = 61.62 for case c 011S30
22A4365 A-205
~
i" Rev. 2
- -O 14'7 l PINF = 74.5 + (0.098 x '4.24) = 1,43 bar for case a, b and d 19 7 * *' * * " 1' *# # "*** "
14.5 j A12.6.2.2 Case a - First Actuation of One or Two Valves (100 F Pool Temperature) i . PRD1 is calculated from the equation in Section A12.6.1. i j PRD1 = 0.421 I + 2.58 (0.23 - 0.1706)
+ 0.1377 (3.63 - 3.83)
+ 0.206 (5.42 - 4.0)
! - 0.0176 (29.38 - 16.0) 1
- 0.0336 (20 - 20)
+ 0.000761 (400 - 400) = 0.604 bars
]() .i From Figure A12.6-1, for PRD1 = 0.604, MPPDV = 0.93 bars. MNPDV is then i
' calculated:
1 0.93 MNPDV = 1.43 = 0. 56 bars (1.43 +-0.93) ! Converting to psi we get: J MPPDV = 0.93 x 14.5 = 13.49 PSlD MNPDV = 0.56 x 14.5 = 8.12 PSID i 042178
- O 4
___. . . _ . _ _ . _ . . _ _ _ . . _ _ _ _ . --_. _ .. _ . . ~ _ . _ _
A-206 22A4365 Rev. 's Case b - First Actuation of All Valves (100 F Pool Temperature) A12.6.2.3 PRD1 is calculated f rom the equation in Section A12.6.1. P RD1 = 0.4 21
+ 2.58 (0.23 - 0.1706)
+ U.1377 (3.63 - 3.83)
+ 0.206 (5.42 - 4.00)
- 0.0176 (29.38 - 16.00)
- 0.0336 (3.93 - 20)
+ 0.000761 (15.44 - 400) = 0.85 bars From equation pg. 203 (top) when PRD1 = 0.85, MPPDV = 1.28 bars MNPDV is then calculated:
= .6 ars MNPDV = 1.4 3 (1.28 1.43)
O Converting to psi, we get:
- TPDV = 1.28 (14.5) = 18.56 psid XNPDV = 0.68 (14. 5) = 9.86 psid Case c - First Actuaticn of an ADS Valve (120 F Pool Iemperature)_ '
A12. 6. 2. 4 PRD1 is calculated from the equation in Section A12.6.1 PRD1 = 0.421
+ 2.58 (0.23 - 0.17061
+ 0.1377 (3.89 - 3.83) l
+ 0. 206 (5.42 - 4.0).
l
- 0.0176 (29.38 - 16.0) l t
- 0.0336 (7.85 - 20) l
+ 0.000761 (61.62 - 4001 = 0.79 bars 011880
i 22A4365 A-207 Rev. 4
.3 From Figure A12.6-1 for PRD1 = 0.79, MPPDV = 1.2 bars. MMFDV is then calculated:
1.2 MNPDV = 1.77 = 0.72 bars 2 + 1.77) Converting to psi, we get: MPPDV = 1.2 (14.5) = 17.40 psid l MNPDV = 0.72 (14.5) = 10.37 psid A12.6.2.5 Case d - Subsequent Actuation of a Single Valve (120 F Pool-Temperature) PRD1 is calculated from the equation in Section A12.6.1.
~
4 O PRD1 = 0.421
+ 2.58 (0.23 - 0.1706)
+ 0.1377 (3.89 - 3.83)
+ 0. 206 (5.42 - 4.0)
- 0.0176 (29.38 - 16.0)
I - 0.0336 (20 - 20) ! + 0.000761 (400 - 400) = 0.639 bars l Froc equation pg. 203 (top) when PRD1 = 0.64, MPPDV = 1.95 bars. MNPDV is then calculated:
- MNPDV = 1.43 7,43)
= 0.825 b ars (7,9 l
O 011880
4 A-208 22A4365 l Rev. 2 O i Converting to psi, we get: MPPDV = 1.95 (14.5) = 28.23 psid
'CPDV = 0.825 (14.5) = 11.96 psid 1
i e 0 i 042178 1 \ O
22A4365 A-209 Bev. 2 l 2 i !O l NOTE: SEE EQUATION OF THIS CURVE, , PAGE A 203 l 4 1 4 i $ 1 - 2 s 2 .i i i 1 i i i ! i ! ! ! I 0 08 09 10 0 0.1 02 03 04 05 06 07 PRD1 dbars di s Figure Al2.6-1. First Actuation Design Pressure Versus Predicted Pressure for First Actuation f 3 i NOTE: SEE EQUATION OF THIS CURVE. PAGE A-203 2 l l 5 4 l >
?
!i i -
l t I I I I I i 1 9 07 08 09 to Ot 02 03 04 05 06 0 PR O1 tbars dl Figure A12.6-2. Subsequent Actuation Design Pressure Versus Predicted Pressure for First Actuation 042178
22A4365 R;v. 4 A-209 2 A12.7 SRV LOAD REDUCTION MARK III STANDARD 238 PLANT Evaluations of in-plant SRV discharge test results indicate that a significant conservatism exists in the Empirical Load Model upon which Mark III Load predictions are based. Test results-are applicable to Mark III analysis ' since the same quencher and flow ratas and comparable discharge lines are used in both designs. Statistical evaluations of the tesc data were performed typically at two levels of confidence and were compared to ccrresponding Empirical Load Model test predictions. Model predictions were based on test conditions of mass flow, pool temperature and valve opening times. The impact of the use of test conditions in the predictions is that the model data comparisons performed for this evaluation are on a comparable bas'is. For des,ign application, bounding values of plant operating conditions are used to give design margin. The margin in the model itself is demonstrated in the Caorso Plant Model/ data comparisons summary in Table 12.7-1. The Caorso data, documented in ref' ence 4 and 5, support load reductions of 50%
~
for single valve first actuat.vas and subsequent actuations. Preliminary data from the second phase of Caorso testing suggests a comparable margin exists for multiple valve actuations. O For the Mark III Standard 238 Plant Design a smaller load reduction is recom-mended to cover differences in the Caorso anu Mark III Standard 238 Plant configurations and to maintain an adequate conservative margin. Specifically, if load reduction is desirad, a load reduction factor of 35% can be applied to design controlling loads for the SRV air bubble pressures as shown in Table A.4.5. The Attachment A methods for calculating loads are applicable. Load reduction factor is applied subsequent to the load calculation I precedure. O 011880
i
'^
- 22A4365 Rev. 4 A-209b l
s . Table Al2.7-1 CAORSO MODEL/MEAN DATA COMPARISONS 4~
! Potential Load f: - Prediction Data Reduction Factor
-(psid) (psid) (Peak Pos./ Peak Neg.)
lA i (SINGLE VALVE, FIRST ACTUATION) 1 8.5/-6.2' 4.3/-2.8 0.52 1 ! (SINGLE VALVE, SUBSEQUENT ACTUATION) t i t' . j 14.9/-9.0 5.8/-3.1 0.63 1 I I (MULTIPLE VALVE ACTUATION - FOUR VALVES) I j 11.2/-7.5 5.0/-3.4 0.55 !O e (MULTIPLE VALVE ACTUATION - EIGHT VALVES)
- 11.2/-7.5 5.6/-4.6 0.45 1
L I + 4 s i 4 O j 011880
.-._...,,..,._e.. . . . , - , . , . -.,- . .,m. ,,-,-,,,.v.,.._._.-,_,,.,,,_.m. ..,,-,-..,._,...-,_my +~--
A-310 i 22A4365 Rev. 4 A
12.8 REFERENCES
- 1. Test Results E=ploved by General Electric for BWR Containment and Vertical Vent Loads, Class III, October 1975 (NEDO-21078),
- 2. Saf etv-Relief Valve Discharge Analvtical Models_, May 1975,
(:ZDE-20942-P).
- 3. Cocoarison of Safetv-Relief Valve Model Predictions With Test Data, July 1975 (NEDE-21062-P) .
- 4. Caorso Phase I Test Data, NEDE-25100-P, dated May 1979.
- 5. Caorso Phase Il Test Data, NEDE-24757-P, dated March 1980.
O 011880 0
- - .. - - --- . - . . ~ ~ ._ . - . - - - - . - -
B-1 2iA4365 Rev. 2 ATTACD'ENT B Bl.0 SUPPRESSION POOL SEISMIC INDUCED LOADS (To be provided by A.E.) Bl.1 HYDROSTATIC PRESSURE During both vertical and horizontal accelerations, the hydrostatic pres-sure distribution in the suppression pool undergoes distortions that lead to dynamic loads on the suppression pool boundary. Bl.2 VERTICAL ACCELERATION ; The following methods can be used for design evaluation. For evaluating vertical accelerations, it may be assumed that the normal hydrostatic pressure increases by an amount that is directly dependent upon the mag-a () nitude of the vertical acceleration, i.e. , at any point in the suppression pool the hydrostatic pressure PH , s given by IcH lb P H"I 2
; (l44 (1 + "f 8) in ,
where c = specific weight of water, lb/ft ?
- R = depth, ft j a = vertical acceleration, ft/sec e
i g = gravitation acceleration, ft/sec 101678 i
--.-.-,--r.-,-_-.
r,w - .- s --. v-. - - -. , .---e--- --- - - - , - -.
+m, -
B-2 22A4365
~
Rev. 2 O Bl. 3 H0RIZONTAL ACCELERATION During horizontal accelerations there will be a non-symmetric modification of the hydrostatic pressure depending upon whether a particular surface is accelerating the mass adjacent to it or not. It is suggested that the I normal hydrostatic pressure distribution be modified as follows. For those structures which are providing this accelerating force, the hydrostatic pressure at any point, PH, s ghen by OH W a/ lb p , , H 144 ," 144 in.2 where
- = specific weight of water lb/ft W = width of suppression pool water that a particular point on the structure is accelerating, f t a = horizontal acceleration, ft/sec g = gravitational acceleration, f t/sec On opposite surf aces, design adequacy is es tablished assuming both normal hydrostatic pressure and no hydrostatic pressure on these surfaces.
101678 O
C-1 ! 22A4365 Bev. 2 'O ATTACECT C WEIR A2.'ULUS BLOCKAGE The following figure (C-1) indicates the effect of 0,10, 20 and 30*.' blockage of the weir annulus on the drywell pressure. This figure was obtained using the analytical model presented in NEDO-20533 (Ref.1) and i was generated assuming that the annulus water associated with the blocked region still has to be accelerated during the vent clearing transient but that the correspending vent flow area is not available. During vtat clearing, the water in the blocked area was conservatively assigned to the unrestricted vent stacks. 1 Due to weir annulus flow blockage considerations horizontal pipe routings should be avoided in the region of 5 ft. above the top of the weir wall. l0 f i f i 1 1 L i P i I 042178
, . ..- - , .-- ,y-y . , , , . - - - - , -
,,,.,m. e,,- -
,,ew m- - - , -
o 22A4365 C-2 Rev. 2 i
- G w
i e
=
w a u i < z w i 4 I I i I w R w u j C to w e 1 > a 0 - u
> c l
5 m
; a o -7 2
- $ w -
I < e = 3 < c x = i U e, o a p i a E W l g l 2 i i u l o u
- o o I
,e, i
w I 1 i-i i, I ! l 1 o i' s < r, n a osd> ssvawoN 3anssawd 173AAwa 042178 l 1
i D-1 22A4365 Rev. 2 p V ATTACIDENT D DRYWELL PRESSURE DISTRIBUTION INTRODUCTION j The purpose of this attachment is to show the resulting pressure dif fer-ential across a given level with some flow restriction assuming a 253 res triction in the drywell.
\
\
\ + t y i
\
ORYWELL 25% RESTRICTION
, Rev
/ 25% RESTRICTION ORYWELL .4
/
V The ! a The greatest pressure dif ferential would occur during a steam break. l flow rate is:* lb ~ Btu 1 see mg h f = 546 at t= = 28,200 sec
$ = 1,230 sec h = 1,190.9 g lbm g
O
- Data obtained from Table D.1 101678
22A4365 D-2
~
Rev. 3 The quality of the break flow is: y , gas , 1,230 m total X = 0.0418 From the quality, the enthalpy of the break is: ho = (1 - X)hf + Xhg hg = (1 - 0.0418)(546.0) + 0.0418(1190.9) B h, = 573.0 b Assuming constant enthalpy process and a final pressure of 14.7 psia, the final quality can be calculated O h =h - (1 - X)h fg o g 572.95 = 1150.5 - (1 - X)970.4 X = 0.405 Using this quality the final specific volume is y = (1 - X)v g +Xv g
= (1 - 0.405)0.016715 + 0.405( 26.80) y= 10.86 ft /lbe 090779 O
f 22A4365 p_3 i Rev. 2
- O The differential pressure is then calculated using the following formula
*2 2
I SP = 1/2 kov but v = cA i i
- Therefore
*2 AP = 1/2 K 2 cA l.
i i
- where the A is the remaining unrestricted area A=A 0.75 Den I
. A = 0. 75(3402.0) ! A = 2551 O and K is the loss coefficient. This is maximum for an orifice. I 1 K= 2
= 2.778 0.6
. Therefore we now have t 2 pt 3 lbm 2 sec
'
- lba
~
lb fs e e' 1 ft' a? = 1/2(2. 778) 29,430 (10.86) 2 4 32.21bmft 144
.551 ft in.2
- 0.433 psid i
l l 042178
22A4365 D-4 Rev. 2 Table D.1 REACTOR PRIMARY SYSTDI BLOWDOWN FLOW RATES AND FLUID ENTHALPY MAIN STEAM LINE BREAK Liquid Liquid Steam Steam Time Flow Enthalpy Flow Enthalpy (sec) (1bs /sec) (Btu /lb) (lbs/sec) (Stu/lb) 0 0 551.6 11,540 1190.0 0.203 0 549.2 10,650 1190.7 0.204 0 549.2 9,960 1190.7 0.99 0 546.2 8,840 1191.4 1.0 28,200 546.2 1,230 1191.4 2.0 27,800 548.3 1.231 1190.9 3.0 27,450 549.8 1,390 1190.5 4.0 27,000 550.5 1,560 1190.3 5.0 22,660 550.5 1.454 1190.3 10 18,000 546.2 1,800 1191.4 15 15,400 533.2 2,220 1194.5 20 12,270 513.2 2,435 1198.7 25 9,030 485.7 2,387 1202.7 30 6.060 450.7 2,110 1205.3 35 4,150 410.0 1,590 1204.9 40 2,750 370.8 1,128 1201.4 45 2,082 333.0 750 1195.5 50 1,843 300.1 460 1188.2 55 1,736 274.7 280 1181.6 60 1,665 256.5 180 1176.3 65 1,635 246,3 126 1173.2 70 1,585 237.7 93 1170.5 75 1,545 231.4 70 1168.5 80 1,510 226.3 56 1166.8 85 1,472 222.7 45 1165.6 90 1,430 220.9 37 1165.0 95 1,390 217.0 30 1163.6 100 1,355 215.0 25 1163.0 105 1,330 212.9 21 1162.2 110 1,300 210.7 18 1161.5 042178
~ . _ . _ . . . . _ . _ _ . .
_ .. =_ __ .._ . _ . . . . . . _ _ _ . _ _. i ' D-5/D-6
- ~ 22A4365 i Rev. 2 5
1 Table D.1 (Continued) ! At the end of blowdown these rates are as follows: I ! Primary System Liquid Flow Liquid Enthalpy Time (sec) (1b/sec) (Btu /lb) , i 1 1; 4 399.98 2755 177.8 1 j 400.00 2755 177.8 800 2755 162.8 l 1400 2755 156.3 l 1720 156,8 j 1799 ! 1800 1720 157.8 2400 1720 163.7 ; 4 3500 1720 161.8 10,000 1720 152.7 2.1 x 10 1720 148.7 ) 1.0 x 10 1720 126.4 5 114.3 ! 2.5 x 10 1720 5 106.9 2 5.0 x 10 1720 7.5 x 10 1720 104.1 I 6 102.2 1.0 x 10 1720
~
?
J l O 042178
E-1 22A4365 Rev. 3 ATTAC'rD!EhT E DRT,TELL NEGATIVE PRESSURE CALCULATION INTRODUCTION The purpose of this attachment is to document the very conservative methods used to calculate the negative drywell pressure that could occur af ter the reflooding of the reactor vessel, it is a bounding end point calculation that leads to the maximum theoretically possible negative pressure. CALCULATION 4 . Somewhere between 100 and 600 see the ECCS system will flood the vessel At this time 4 causing instantaneous condensation of steam in the drywell. all the air initially in the drywell will have been purged into the con-
' tainment. To evaluate the containment pressure at this time, the initial quantity of air in both the drywell and containment is needed.
Initial mass in D.W. (P - Py) V DW M e "DW g where F = Pressure in D.W. initially = 16.7 psia l Py = Partial pressure of vapor = 4P g R = Temperature of D.W. = 135 F i 1 ft-lbn R = Gas constant - 53.34 lbFOR
\
090779 l
- j. l
22A4365 E-2 Rev. 2 V = V lume D.W. - 274,500 ft DW
)= Relative humidity = 0.40 P,, =P = Sat- pressure at 135 = 2.5365 psia 135 Therefore in. l M " [ 16. 7 - 0.4(2.5365)]
53.4(540) (274,500) 2 ft 1 Si DW
= 21,501 lbs of air l l
l Initial >fass in Containment (P - Py) V g con " RT whe re P = Pressure in containment initially = 14.7 psia Py = Partial pressure vapor = $ Pg ,g V = Volume of containment = 1,138,750 ft R = Gas cons tant = 53.34 T = Initial te=perature = 80 F
$ = Relative humidity = 0.20
*P = Sat pressure = 0.5067 psia sat 042178 g
i i 4 E-3
! 22A4365 l Rev. 2 1 ;
1
' I
[ 14.7 - 0.2 (0.5067)] (1,138,750) M = 144 in.2 J cont 53.34 (540) ft 2 i i . i 1 M = 83,110 lbm of air g From the above air masses the post blowdown containment pressure can be calculated. l p = I'tRT 1 cont V c
+ $P sat .
} l EM = Summation of initial air mass in cont and l D.W. = 102,645 lbm air i s 1bm-ft R = Gas constant = 53.34 lbFOR L i T = Final temperaturc = temperature of pool = 170 F 3 4 i V = Containment volume = 1,138,750 ft c
? = Final relative humidity = 1.0 P,,g = Saturation pressure at 170 = 5.990 l
i 4 042178 l0 T i f i i 1
22A4365 E-4 i Bev. 2 l l 102,645 (53.34) 630 1 P = + 1.0 (5.990) cont i"*2 ~ (1,138,750)l44 ft 1
= 27.025 psia l
To evaluate the minimum drywell pressure at this time the following assumptions are made: (1) All steam in the drywell is condensed. (2) ECCS flow out of vessel is at temperature of 170 F. l l l (3) Assume all the air has been purged out of drywell pressure ~. l l (4) No vacuum breakers, t i Using these assumptions the final dryvell pressure is equal to the lh saturation pressure at 170 F. I i I P ~ = 5.990 psia DW sat 170 l l j Therefore the negative pressure load across the drywell wall is the di f f e rence in the final pressures of the containment and drywell. I l PD = PD'4 -P cent i
= 5.990 - 27.025
= -21 psid 042178 h
~ .
E-5/E-6 22A4365
,s~ Rev. 2 V,.
SLTiARY i. The above represents a very conservative bounding calculation of the f maximum theoretical negative pressure. The assumptions that noncondensibles return to the drywell via the vacuum relief system and that the steam temperature in the drywell instantaneously drops to the suppression pool temperature are both very j conse rvative. In addition, the real estimate of relative humidity in containment is 50% rather than the 100% assumed. An evaluation of the probable transient condition for this phase of the LOCA leads to the conclusiot. that the realistic negative pressure is less than 8 psi. O O l 042178 j
4 1
- 22A4365 y_1
' Rev. 2 ATTACICENT F j
; WETWELL ASSY!OiETRIC PRESSURES l
INTRODUCTION f4 i The purpose of this attachment is to ' determine the pressure gradient under i the HCU floor during pool swell due to flow restriction at the HCU level. i
- i. CALCULATION i
I j During pool swell the following conditions exist in the wetwell.
= 11 psi AP = HCU floor pressure differential j HCU
=
2 A = Open area f HCU floor 1500 ft HCU !O j XA = X-Section area between Pool and HCU Floor = 400 f t 2 l (Vertical plane) 4 k = loss coefficient through floor = 5 o = density of flow through floor = 20 ft y i j From this infor=ation the flow rate through the HCU floor can be i calculated. ! A bCU = g /2 P HCU . m r lbe-ft
= 2(20)(11) 144 I"., 32.2
/5 ft' lb-ft-sec'Y bm
= 958,968 sec I
O 042178 4
,~ .-..- -. - - , . . - . . . . - - , , .. . - , . . . . , - - - . - . , . - - - .
~
F-2 22A4365 Rev. 2 In order to cal:ulate. a dif ferential pressure under the. HLU_ floor, assume a 25% restriction at the ECU le. vel. 25% RESTRICTION
/ /
/'
ORYWELL _ CONT AINMENT Using this 25% restriction then 1/8 of the flow will be horizontally diverted in both directions. 1/8 FLOW N 1/8 FLOW This horizontal flow rate is e, ,
'dCU
' 'd 8
, 958,968 8
042178 O
e j i4 F-3/F-4 22A4365 i' Rev. 2 i f
- lbm i M.W = 119,770 sec i
! Assuming the density is constant and using the above flow rate, the
- differential pressure under the HCU floor due to this restriction can i now be calculated.
1 i i 2 { AP = Ko V;- but V 2 = '*1' ,, A=X 3 4
-8 a 2*A~
1 i
- } 2 K(p )M li 2(g):2 (XA2) i ' 2 1
1 (20) 119,770' ft 2 { 2 (32.2) 20 (400 ) 144 in 4
!O = 0.483 psi i
i 1 i i 1 i i i b t i b i l 1 042178 l i l - . _ . . . ..- - - - - _ . . . - . - - . . . - - . - . - - , . - - . . _ - , . . - - . - . _ . , . _ - . - - - - . - . . - . , ,
22A4365 G-1 i Rev. 3 ATTACHMENT C SUBMERGED STRUCTURE LOADS DUE TO LOCA AND SRV ACTUATIONS TABLE OF CONTENTS Section Title Page i Gl. INTRODUCTION G- 2 G2. SUBMERGED STRUCTURE LOADS DUE TO LOCA G-4 G2.1 Compressive Wave Loading G-4 G2.2 LOCA Water Jet Loads G-4 G2.3 LOCA Bubble Loads G-5 G2.3.1 LOCA Bubble Loads - Sample Problem G-17 G2.4. Fall Back Loads C-23 l G2.5 LOCA Condensation Oscillation Loads G-24 f"%j G2.5.1 LOCA Condensation Oscillation Loads - . v Sample Problems G-25 ) G2.6 LOCA Chugging Loads G-27 i ! G2.6.1 LOCA Chugging Loads - Sample Problem G-30 i i G3. SUBMERGED STRUCTURE LOADS DUE TO SRV ACTUATIONS G 33 ! G3.1 Quencher Water Jet Load CH33 l G3.2 Quencher Bubble Load G 33 l 1 G3.2.1 Quencher Bubble Load - Sample Problem (L37 : I C4. REFERENCES (642 i i k } } 4 iI O 4 0 090779
- . . - . - . .- - _ - . _ - .. .=- -
22A4365 c-2 Rev. 4 O Gl. INTRODUCTION In the following two sections, the flow induced loads on structures submerged l l in the suppression pool duc to Loss-of-Coolant (LOCA) and Safety Relief Valve (SRV) actuations are discussed. During LOCA, steam rapidly escapes from the break and creates a compressive wave in the drywell air space. This wave is transmitted from the weir wall water surf ace to the suppression pool and finally to the subr.arged structure. This compressible wave loading is negligible (discussed in subsection G2.1) . Following this compressive wave, the drywell is rapidly pressurized. The water in the weir annulus and drywell vents is expelled to the suppression pool. A highly localized induced flow field is created in the pool and a dynamic loading is then induced on subuerged structures (discussed in su5a..ction G2.2) . After the water is expelled from the vent system, the air in the dryvell air space, prior to the LOCA event, is forced from the top vents and forms expanding bubbles which create moderate dynamic loads on submerged structures (dis-cussed in the subsection G2.3) . These air bubbles cause the pool water surface to rise until they break through the pool water surface. The pool surf ace slug decelerates and falls back to the original pool level (f all back loads are discussed in subsection G2.4). At this point steam from the break fills the drywell space and is channeled to the pool via the vent system. Steam condensation oscillation starts and the vibratory nature of pool water motion causes an oscillatory load on submerged structures (discussed in subsection G2.5). O 011880
1 22A4365 G-3 Rev. 4 4 O V This condensation oscillation continues until pressure in the drywell decays. This is followed by a somewhat regular but less persistent vibration called chugging. During this chugging period, a high wave propagation spike is observed which causes an acoustic load on submerged structures (discussed in the subsection G2.6) . During SRV actuations, the dynamic process of the steam blowdown is quite similar to LOCA but the load is mitigated by the X-Quencher device attached at the end of each SRV discharge line. Two types of loads are important. One is due to the water jet formed at the confluence of the X-Quencher arm discharges (discussed in subsection G3.1) and another is due to the four air l bubbles formed between the arms of the X-Quencher. These air bubbles are smaller in size than the LOCA air bubbles, reside longer in the, pool, and oscillate as they rise to the surface of the pool. The load created by these bubbles are discussed in the subsection G3.2. () The material in Attachment G is organized as follows: (1) The specific analyticel model is referenced, this is followed by (2) A load calculation procedure which is a summary of the engineering process. This is followed by (3) A sample problem which demonstrates the use of the procedures. O 011880
22A4365 C- 4 Rev. 4 G2. SUBMERGED STRUCTURE LOADS DUE TO LOCA G2.1 Compressive Wave Loading As discussed in Section 6.1.1, the very rapid compression of the dryvell air theoretically generates a compressive wave. But as pointed out in Sec-tions 6.1.1 and 6.1.2, there were no loads recorded on the containment wall l in PSTF for this phenomena. From this, it can be concluded that compression wave loads on structures in the suppression pool are significantly smaller than loads caused by the water jet, for structures close to drywell. For structures near the containment, neither compressive or jet loads are signifi-cant. - G2.2 LOCA Water Jet Load During the initial phase of the DBA, the Drywell air space is pressurized and the water in the weir annulus vents is expelled to the pool and induces a flow field throughout the suppression pool. This induced flow field is not lk limited to direct jet contact and creates a dynamic load on structures submerged in the pool. However, since none of the submerged structures in the Standard Plant are in the direct path of these jets, the dynamic load on these structures is less than the load induced by the LOCA air bubble which forms af ter the water is expelled. Examination of Reference G1 and G2 test data confirms this observa-tion. Since the air bubble dynamic load is bounding, this load is conserva-tively used in place of the water jet load (for air bubble load, see paragraph G2.3.). O 011880
I 22A4365 G-5 Rev. 4 O G2.3 LOCA Bubble Ioads i During the initial phase of the DBA, pressurized drywell air is purged into the suppression pool through the submerged vents. After vent clearing, a single bubble is formed around each top vent. It is during the bubble growth period that unsteady fluid motion is created within the suppression pool. During this period all submerged structures below the pool surface will be exposed to transient hydrodynamic loads. The bases of the flow model and load evaluation for the LOCA bubble-induced submerged structure load definition are derived f rom the model in Ref-erence G4.3. The following procedure is reconsnended for calculating the loads on submerged structures. l
- 1. Pool Dimensions and Bubble Data J
Specific data that must be obtained are: R: initial bubble radius, assumed to be the same as the vent radius, 1.146 ft. P: Drywell transient pressure obtained from Figure 4.4 (page 4-12), in psia. p : air density corresponding to drywell conditions when the drywell pressure is P , Ib /ft . o m p: pool liquid density, 62.4 lb /ft . I m l P: containment air space pressure, assumed to be constant at 14.7 psia. P,: initial pool pressure at the top vent centerline submergence, Psia. H: pool depth, 20 ft. _ 011880
22A4365 G-6 Rev. 4 l l L: pool length, 18.5 ft gl D: unit cell pool width (Figure G2.3.6c), 7.97 f t. y: initial bubble location from bottom of pool, same as the top g vent centerli:.e,12.5 f t.
- 2. Duration of Loads Loads on submerged structures due to LOCA air clearing begin when the air bubble forms at the vent exit immediately following air clearing and end when the bubble engulfs the structure, or when breakthrough occurs if the bubble does not engulf the structures.
- 3. Initial Bubble Location Initially the bubble center (x , y , z ) is assumed to be located l on the vent axis at a distance equal to one vent rciius from the vent exit.
- 4. Movement of Bubble Center The bubble center movement is used only in the calculation of the bubble engulfment time. The effects of bubble movement on the pre-dicted load are conservatively accounted for by a factor of 2 (see step 14).
- 5. Bubble Dynamics The bubble dynamics equations given below can be solved for:
R(t): the bubble radius at time t k(t): the bubble growth rate at time t k(t): the rate of change of bubble growth rate at ti=e t. 011880
d 22A4365 G-7/8 i Rev. 4 !O 4 Bubble Dynamics Equations:
- I
-f (0
- 1)
R = (PB - =) ~ 1 P s P R l P = 3k - /G2. 3-2) 3 l _4nR o , a
#( ~Y) ( '
P, = P C o where P = bubble pressure at time t, psia B i T 6 B
= bubble charging rate, lb,/sec (see eq. A46, Reference G4.3) l k = ratio of specific heats for air, 1.4 4
Initial Conditions R(0) = R = vent radius = 1.146 ft l f i I k(0) = V /4 = 13 ft/sec f where V g = top vent water jet velocity at vent clearing (from Figure G2.3.1) = 52 ft/sec 4 1 4 P(
- a t P ven c ear ng " Psia Urom H gure W
-f B o
. L mB (0) = 1 nR o 3 io ,
3 i l
' A plot of bubble radius vs. time -has been obtained from the bubble l l- i
? dynamics equations and is presented in Figure G2.3.2. l ! l i, 011880
!l l ND $, p*
E - d _ .O. 0 - 2 RR A A E I 8 E L 1 H L A C C T E T L N C E N E T N E V V E M L O 6 V D T I 1 P D T y O I O T M B i i l l 1 I I i l l l i ' t i
- - - c o . C C C l e
V V V T . T M B N I 4 V 1 E t V e _ M l. O T r T e O t l l B l l Il A il l I I l l I I I I 2 1 W a t _ n e _ t V _ c E e s l L DT ( a _ o E t DN I E I t M n
- o
_O I MV T z l I i l l l 1 I [ l l i ' i l l 1 I r l I o - I I 8 I
- T 0 I N
3 E k G3 V r NS I 0 P a M S2 UO O T DD I 6 . EE 0 1 . TN A L . R! t 3
- E N O D 2 E C
. GM L L e HA 4 r U I C I 0 u GT g i I Y F ' L l SA I
. HN TA 2
_ 0
- - - - - O 0 0 0 0
- 0 0 0 2 1
' 6 5 4 3 jh C 5 }>
O 9;58
22A4365 G-9a Rev. 4 a i O 10 1 I i l 1 8 - I I 6 - i 3
~
3. 5 O $ e I m 4 - l 2 - O_ ! I l l l l 0 0.1 0.2 0.3 0.4 0.5 06 0.7 TIME AFTER VENT CLEARING, t (sec) O I V Figure G2.3.2. Bubble Radius, R(t), vs. Time 011880
22A43o5 G-10 R:v. 4 l
- 6. Structure Data O'
Location of structure, including elevation, distance from the drywell wall, and distance from the containment wall. Dimensions, shape, and orientation of structure. Long structures f should be divided into smaller segments (with each segment approxi-mately 2 ft long) for more precise evaluation.
- 7. Distance to Structure Determine the following at time t D:
s the cross-sectional dimension of structure or structure segment in the direction of the bubble center. (For this calculation, the bubble center is assumed to be displaced horizontally a distance equal to R(t)).
- 8. Check Structure / Bubble Contact At anytime t using the value of bubble radius, R(t), from Step 5 check if R(t) > v/[x - R(t)]2 + (y - y )2 + (z - z )2 - Ds /2.
If true, then loading calculations for the structure or structure segment under consideration end, because it is inside the bubble and the drag forces are zero. If not, proceed with Step 9 until the bubble breaks through the pool surface.
- 9. Pool Boundary Effects To account for the effects of pool walls, floor and free surface, use the method of images as described in Section 4.10 of Ref-erence G4.3. First determine r y k, which is the distance from the center of the structure or structure segment (x, y, z) to the source or sink image at (x , y , z,),
t O 011880 m
G-11 22A4365 j Rev. 4 I[v \ (G2. 3-5) r ijk (* ~'*i) +(Y-Y) j + (* ~ *k) l-Note that r corresponds to the real source. Then evaluate the functions X, Y, 2 as given by Equation A67 of Reference G4.3. Using t the notations adopted here, these functions may be written as N N N (,1)j (x_x) I X = K 3 k=-N j--N i=-N 1jk (-1) (y - y ) , l (G 2. 3-6) , Y = K 3 l k=-N j=-N i=-N ijk (-1)) (z - z ) 3 k=-N j=-N i=-N ijk I j v b where N is the total number of images considered, and K is a factor used for finite bubbles to satisfy the local pressure boundary condition at the real bub 5de itsrf ace, i.e. , the pressure at the real bubble surface ec'; is he ?ndependently calculated bubble pressure, PB . The K t- is not a function of the structure's location in the pool. It it- a tu.. .un of bubble radius and the bubble image function. Values of K have been computed for the unit cell shown in Figure G2.3.6c and are provided as a function of time in Table G2.3.1. t I
- 10. Number of Images The results of a sensitivity study show that 11 sets of images will provide adequate convergence. A typical arrangement of image sets in l
i the vertical plane is shown in Figure G2.3.3. The K factors shown in Table G2.3.1 are based on 11 sets of images. 011880 l
22A4365 G-11a Rev. 4 r3
\
v) i W G u e w C OOceCOCeceeOOOCeC: COCCCCCCOOCOOC CCC CCCCCC CCe::: C-6 LWWWWW 0 C WWWWWWW@WWWUUWWWWWTONC-ec NNCONo vocN@NcCvT E Q
=-
NC4-NM@cM@cN@cMCNT-CCMcN N---CO@@O,CCCNN NNCCCCCCC- aU NNNNNN-~~-e--===-------- g ---~~~------~~-. . . -------- . . . .. g a
. .............. - C COOCCCCCOOOOOOCCCCCCCCCC' W >
0 COCOCCCOOOOOOOCOCCCCCCCC OOOOCOCCOOOOOOCCCCCCCCCC g W
, C U wwyyWyWWyyWyyyWW CCCCCO:
_WL;;_y
- C Z G COCCCCOOOOCOOO 000C M COCOCOCCCCCOOCCCCCC C C C C C C C C C C C C C C C C C C C C C L*
W CC@@CC--NNMMvvCCCCNNCCO@
- v. v. i. v. C. C. C C. C. C. C. C. C. C. C. C. C. C. C. C. C C. C C OCOCOOOOOOOOOOOOCCCCCOOC OCCCCCCCCCCOOCCCCCCOOCCCCCCCCCCOOCOCOOOOCCCCCCCC OCCCOCOCOCCCCOCOOOOOOOOCCOOCOCCCCOOOOCCOOCCCCCCO WWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWW MovNNN--CMcNNcovNitc-CCCCNo-vCvNNMcOCM NTN cMOC MN-c-NT-ecveN----NMicN@NvbCMc@MN-C@vCMCMCMC4@C-c NO@NcvMNOSCNcCvnN-O@@NccCvvMN--Co@CCNNccCCvvMMMN
- 0 M ccCCCCCCCivvvvvvvvvMMMMMMMMMMMMMMNNNNNNNNNNNNNNN
. E .............~.~....... ...........................
M - OCOOCCOOCCCCCOOOOOOOOOOOOOCCCCCCCCOOCCOOOOCOOOOO n/ P t N N C; . C c > OCCCCCCCCOOCOCOOOOCOOOOOOOOOCOOCC 00000000C0C0000 OOOCCOOOCOOCCCOOOOOOOOOOOOCCCCCCCCOOOOCOOOOOOOOC i 9g WWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWW OCOCCCCCOCCOOOOOOOOOOOOOOCOCCCCCCCCCOOCCOOOOOOOC Q g m OOCCOCCOOOOOOOOOOOOOOOOCOCOCCCCCCCOOOCOOCOOCOCCC OCCCCCCCCCOCOCOCOCOCOCOCOCOCOCCCCCOnOCOCOCOCOCCC w g vvCCCCNNCC@@CO--NNMMvsnccQNSCC@@CC--NNMcivCCccNN NNNNNNNNNNNNMMMMMMMMMMMMMMMMMMMMvivvvvvvyvvvvvv4 bbbbb0 b Obb bb bbbbb bbb bb bb bbbbbb CCCOOOOOOOCOCOCOOCCCCCOOOCCCCOOOCCCCCCCCCCOOCCCC OOOOOOOOOCCCCCOOOOCCOOOCOOCCCCOOOOCCCCCCOOCCOOCC WWWWWWWWWWWWWWWWWWWWWWWWWWWWW@WWWWWWWWWWW@WWWWWWWW Nocc@C--CS-McN-M-ONCC-CC ctocCCCCvc MOONMNcM Tct
@MCC-M-@cMCCCCOM -T-M@CNCCCONNvCMC@O-iCMcNCCCCN C CT v@-McNOMCv-CCCCCN@-vbCv7MCv@C-NMONMcCONCCCM-@*
~@CcvN-@NCCNN-C@CNNCCCveMNN---COO @@@CCCCNNANcccc g CivvivinMMMMMMMNNNNNNNNNNNNNNNNNN-------~~~~----
bbb b bbb bb0 bb bb bbb bbb bbb bbbbbbbb0 bb
-------CCCCCCCCOOOOOOOCCCCCCCCCCCCC OOOOCCCCOOCCCOO
^
U N----------C.000C0000C0C000000000 00000000000 e s e e o e e e e e s e e e e e e WWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWWW6LWLWWWW G M COOOOOOOCCOOCCCCOOCCCCCCOCCCCCOOOOCCCCCCCCCCCCC COOCOOCOCOOCCCCCCCCCCCCCCCCCCCCOCOCCCCCCCCCCCCC w w OCOOOOOOOOOOCCCCCCOOCOCCCCCCCCCCCCCCCCCCCCCCCCC OOCOCOCOCOCCCCCCCCCCC--NNMMvvCCccNSCC@@CC--NNMM C--NNOMvvCoccNSCC@@--~~~~~~~~~---~~----NNNNNNNN bbb 0bb bbbbbbbb bbb0bbbbbb bbbbbbbb bbb bbb c) 011880
22A4365 G-12 Rev. 4 4
+
O
- ++ +
++ t 1
+ ++ ++ :
1
. + ++ ++
l 4 + ++ ++ + 4
++ ++ ++ +
+ ++
++ ++ ++
! + + + ++ ++ ++
++ ++ ++ ++ ++ +
1
++ ++ +!@l '+ ~ ++
+
- + ++ ++ ++ ++
e I )
++ ++ ++ t l
I
+ ++ ++.
.e. ++ 44 J
LEGEND. ++ + j h REAL SOURCE (BUBBLE) l 4 IMAGINARY SOURCE 94 4 ! IMAGINARY SINK h FREE SURF ACE _ /_ POOL BOUNDARIES i N
! CONTAlWENT l ORYWELL [g Figure C2. 3. 3. Arrangement of Images i O 011880
. . . _ . - - _ . _ - ._. . __ - _ ... - ~. ._. .
22A4365 G-13 Rev. 4 I 0 11. Since the bubbles are. symmetrically' located in the circumferential
' direction, and are formed synchronously, the ef fects of multiple bubbles can be evaluated through image methods applied to a unit cell 4 -
as shown in Figure G2.3.6c, A single pool' segment is therefore modeled by means of a rectangular unit cell of equal size.
- 12. Direction of The Flow Field I
t
- The direction of the flow field at time t is determined by the unit vector, n, where I
Xn +Yn +Zn g n = (G2. 3- 8) X +Y A Z
- 13. Acceleration and Velocity O
i Using the results from Steps 5 and 11, the equivalent uniform acceleration at time t at the structure location in a finite con-j tainment is i
-2 .. .9 '
X' + Y2 + 2' (C2.3-9)
,(t) =
R (t) R(t) + 2R(t) R'(t) . The corresponding velocity U,(t) may be obtained by numerically i integrating h(t). 9 9 9 9 X' + Y' + Z' U,(t) = R'(t) R(t) (G2. 3- 10) i Characteristic U,(t) and 0 (t) plots [ U ,(t) K ; b,(t)K
}
fX#+Y2,72 X+Y+2) O 4 are presented in Figures G2.3.4 and G2.3.5 cespectively, i 011880 < l 1
. . - . -- - . - . . _ . - . - . . . . , . - . n,- , , , . . . . - ~ , . , . , . - . . . - , , , , , , , . - , . - . . -
4 l l 22A4365 G-13a Rev. 4 O 18 16 - 14 - I e x 12 - N W N e- t,*> 10 _ j i i i' i 8 a s - e w j N G E i e - i O
=
s 4 - i 1 . '/ o i I I I I l O 0.1 0.2 0.3 0.4 0.5 0.6 0.7 1 TIME AFTER VENT CLEARING, t (sec) \ Figure C2.3.4. Characteristic Velocity vs. Time 011880 t
. I G-13b 1 22A4365 l I Rev. 4 O
I 10 - 4 I i l i g. n X
! N i
g n l' _ > n 6 i
.jx 4
l ' i l 9 i Z l U
- w U
U 4 -
!' 5 D
4 Z s 1 0 2 J i i 1 0 i i l I I I I I 0.5 06 07 O 01 02 03 04 TIME AFTER VENT CLEARING, t (wcl
- O Characteristic Acceleration vs. Time i
Figure C2.3.5. 011880
22A4365 G-14 Rev. 4
- 14. Drag Forces The acceleration drag is calculated from b (t) VA (G2. 3-ll)
FA (t) ~ g C where d, is the acceleration component normal to the structure and V is the acceleration drag volume (from Tables G2.3.2 and A G2.3.3 for flow normal to the structure) . The standard drag force is calculated from 2 Fs (t) = C A D n 2 c where C is the drag coefficient for flow normal to the structure. D A is the projected structure area normal to U (t). n Add F and F at any time t to get the total load on the structure 3 l segment. The loads predicted by this procedure agree with the Mark III sub-merged structures test data (Reference G4.1). For additional conservatism, the final load should be multipled by a factor of 2 to cover the effects of a moving source. The direction of total drag is normal to the submerged structures. O 011880
i G-15 22A4365 Rev. 4 Table G2.3.2 l ACCELERATION DRAG VOLUMES FOR TWO-DIMENSIONAL STRUCTURAL COMPONENTS
' LENGTH L FOR ALL STRUCTURES)
SECTION THROUGH HYORODYNAMIC ACCELER ATION ORAG BODY AND UNIFORM VOLUME V, FLOW OIRECTION MAgg BODY R CIRCLE : awR2 L 2nR2L
?
-> b 7
- wea, bit ELLIPSE : r awa2L il b
C ? awb2 L abia+blL j ha
- 5
+ wA PLATE 2a mt 2 t
O H --- , 9 Aw a2L aL(4b+wa)
+e 28 .t(4e.1,,4w.)
RECTANRE to 1.14 ana2L Jh b 1.21 owa2 L a L(4b+ 1.21 r al 2 1.36 0ws2L a L(4b+ 1.36w el 1 1.51 owa2L a L(4b+ 1.51w al
.t( 4e.1,70..
II2 1.70 ona2L 1/5 1.98 ona2t a t( 4e.1,gg. .l 1/10 2,23 pw a2L alt eo+ 2.23w a) s/b
~
&2b %
7 2 0 85 aws2L eL(2b+0.85*si 2 1 0.76 ona2L eL(2b+0.76,al OIAMONO 2b 1/2 0.67 ona2L a L(2b+ 0.67* s) d 1/5 0 61 9w s2L a L(2b+ 0.61* si c --- , % a/c=2.6. b/c=3 6 c
~ ~
lO
,,,4. =
Lt
!a 2. , , .. 2L ,,., ,. a2 2c,2 - b. c, , L
.-u-011880
22A4365 c-iti Rev. 4 Tablo G2.3.3 ACCELERATIO!! DRAC VOI.UttES FOR T!!REE-DittE.tStotlAI. STRUCTURES H v Dft ODY N AMIC ACCELE ft AtlON OH AO DODY AND FLOW OlHECilON 9 455 VOLUME V4 DESCniPflON
/
- - 8/3pR3 8/3 nl C t f1CU L A H j-DISK
! t>/a u e/0 ba2 h,0 ba2
/
C
- I 0 9 e ein ba,-
E LLIP flC AL 3 0 9 =/6 t a2 g,$g f 0H26oelpb42 0 H16 wi6 be2 [ 2
- 15 0 748 e w/6 bs2 0.14u e/6 ba2 1.0 0 637 o e/6 ba2 0.637 *l6 ba2 ble A, -
0 4 rn , .i4.2b 04;o.,4.2b 9 i b tD 0 6140 a */4 42b 06140*t4a 2h
/
/
o n 40, j a .2b 0 340 ,14 42b HCCTANGULAn ) ,V 7 FLAff
/ / 25 O us ),, .e 4 .2b 0 933.i4 .2b
/ 3 ,> = t 4 42b .i4 .'b p
~
* ,, e t s .In e ,4 .2b
,4
.,3n A N ,,i 3<2
. w N od2 db f p't AP4 GUL AH hh -
^/1 ^
PLATE j/ 0 i%
\
'r - .
H
* * / e 2/3. n 3 2. n a sr.Ent 0
011880
22A4365 G-17 Rev. 4 (J G2.3.1 LOCA Bubble Load - Sample Problem As an example the drag force on a cylindrical structure induced by the LOCA j bubble f rom one vent will be calculated, including the boundary ef fects. Fig- ! ure G2.3.6 depicts the Mark III Containment and the submerged structure in - l- que stion. Step 1: The following data were used in the sample calculation: (a) Initial bubble radius: R = 1.146 ft (b) Initial bubble velocity: v = 13 ft/sec * (c) Initial bubble submergence: H-y = 7.5 ft 20 ft, L 18.5 ft, D = 7.97 ft
(d) Unit cell dimensions: H = =
(e) Containment air space pressure: P = 14.7 psia (f) Pool liquid density: p = 62.4 lbm/ft 3 1 (g) Obtain drywell transient pressure from Figure 4.4 (Page 4-12) I J (h) Initial drywell temperature: 100*F (1) Vent friction factor is zero l Step 3 1.146 f:, Y = = The initial bubble location is X, = 12.5 ft and 2 9 9 3.986 ft. O V 4 011880 g
22A4365 G-18 Rev. 4 e I CONTAINMENT ORYWELL
- V" '"'
VENTB + I f t- -2n
~
VENT A , VENT C Y z (a) PLAN VIFW OF FOOL SEGMENT e f-'/l _=_ -
= _ = . . . . _ . . , .."
t.. .. .., f// _.. g: E
/l//H1/HHHH/HH
/ 7 lbl ELEVATION VIEW OF UNIT CELL
////
BASE MAT Figure G2.3.6. Mark III Submerged Structure for Sample Calculation 011880
22A4365 G-19 Rev. 4 5.921t POOL SEGMENT UNIT CELL 1
-+x n h
h
- 7.97 f t 9 42 f t 6.52 ft i
4 l ' II } V l
- ~ _ .
y j 4 m L 0 8.5 f t) V z (C) EQUIVALEN'.' UNIT CELL FOR IMAGE COMPUTATION Figure G2.3.6. Mark III Submerged Structure for Sample Calculation (Continued) O 011880
I~ l 1 22A4365 G-20 Rev. 4 Step 5. The bubble radius vs. time, R(t), is obtained from Figure G2.3.2. Step 6. Structure center location is x = 4.0 ft, y = 15.5 ft, z = 3.986 ft. The structure length = 2.0 ft, diameter = 1.0 ft, projected area = 2.0 ft2, acceleration drag volume = '3.14 ft3 (Table G2.3.2), and the standard drag coefficient = 1.2. The structure is assumed parallel to the Z axis (see Figure G2.3.6), and is considered as one segment. Steps 7, 8. D - j = 0.5 ft; R> [4 - R] + (15.5 - 12.5)2 + (0)2 _ 1.0 and R(t) for bubble engulfment = 2.4 ft. From Figure C2.3.2 for R(t) = 2.4 ft, t = 55 ms = bubble engulfment time. Steps 9, 10, 11. To account for the effects of pool walls, floor and free surface the method of images was'used. A single pool segment is modeled by means of a rectangular unit cell of equal area (see Figure G2.3.6c) . Equation (G2.3-6) is calcu-lated for the function /X2 + y2 + Z2 which is the correction factor that accounts for boundaries and la applied to the velocity (Equation G2.3-10) and acceleration (Equation G2.3-9). K values are obtained from Table G2.3.1 as a function of time. Values for X, Y, Z are computed and the results, for this sample problem, are X/K = 0.07576, Y/K = 0.13101, and Z/K = 0.00. O 011880
1 G-21 1
-22A4365 Rev. 4
.i
- f. 7'g Steps 12, 13.
! At time t.obtain the characteristic velocity and characteristic acceleration from Figures G2.3.4 and'G2.3.5 respectively. Combine these results with those
'from steps 9, 10, 11 as follows:
1 j '
=
[ U (t) K )1fg [h I U= (t)x lgx,2,7272 j g gjK 4 $ / U,(t) K )[y)
= jl -
l l U (t)I
\ YX +Y +Z/h/
f [ U,(t) K [Z i U.(t), = 2+Y2 # 72 7 [j X
/ U,(t) K p
=IE)x (g2 #72,72 )
I
.j i
' / ,(t) K ) h
= l l l l 0* (t)Y \ YX2,y2,7j 2 j ,
l 4' b (*t) K ' /z h ) !I
.(t)
~
=
- YX" + Y2+Z/( j )
2 i Results of these calculations, for this sample problem, are presented in . Table G2. 3.4. l < l t l Step 14. ! h The acceleration and standard drags are added together to give the total drag. The total x, y, and z components of drag, and the total resultant i
. drag, calculated for this sample problem, are shown in Table C2.3.5. These results should be multiplied by a factor of 2 as previously stated.
i 011880 i _,_ -. . , - . _ . , . , . . . . , . - . _ , - . . . _ . _ . _ . _ _ . _ . , - _ - . ~ . , . _ . _ . _ - _ . . , . . . .
22A4365 G-22 Rev. 4 Table G2.3.4 FLOW FIELD AT SUBMERCED STRUCTURE CENTER FOR SAMPLE PROBLEM t(sec) Acceleration (ft/sec ) Velocity (ft/sec) x y z x y z 0.000 53.1 91.9 0.0 0.7 1.2 0.0 0.005 57.2 98.9 0.0 0.9 1.6 0.0 0.010 61.4 106.3 0.0 1.2 2.1 0.0 0.015 63.5 113.3 0.0 1.6 2.7 0.0 0.020 69.2 119.7 0.0 1.9 3.3 0.0 0.025 72.5 125.4 0.0 2.2 3.9 0.0 0.030 75.3 130.2 0.0 2.6 4.5 0.0 0.035 77.5 134.1 0.0 3.0 5.2 0.0 0.040 79.3 137.1 0.0 3.4 5.9 0.0 0.045 80.5 139.2 0.0 3.8 6.6 0.0 0.050 81.2 140.5 0.0 4.2 7.3 0.0 0.055 81.5 140.9 0.0 4.6 8.0 0.0 0 0 011880
l
. 22A4365 G-23 i Rev. 4-Table G2.3.5 TOTAL FORCES ON SUBMERGED STRUCTURE FOR SAMPLE PROBLD1 t(sec) x-Force (1b f) y-Force (lb f) - z-Force (lb f) Resultant Force (lbf) 0.000 325.6 563.1 0.0 650.5 0.005 352.3 609.2 0.0 703.7 3
0.010 381.2 659.2 0.0 761.5 0.015 410.1 709.3 0.0 819.3 438.1 757.7 0.0 ~875.2 i 0.020 0.025 464.8 803.9 0.0 928.6 0.030 '490.0 847.4 0.0 978.9 4 0.035 513.7 888.4 0.0 1026.2 1 0.040 535.9 926.9 0.0 1070.7 i 0.045 556.8 962.8 0.0 1112.2 0.050 576.2 996.5_ 0.0' 1151.1 i 0.055 594.4 1027.9 0.0 1187.4 Note: Multiply these results by a factor of 2 to obtain the final answers. 1 ! G2.4 FALL BACK LOADS i
-s g These is no pressure increase in the suppression pool boundary during pool l
4
'd. fall back as discussed in Section 4.1.6. Structures within the containment i suppression pool that are above the bottom vent elevation will experience 1
4 drag loads as the water level subsides to its initial level. For design purposes, it is assumed that these structures will experience drag forces associated with water flowing at 35 ft/sec; this is the terminal velocity for a 20 ft free fall and is a conservative, bounding number. Free fall J height is limited by the HCU Floor. The Load computation procedure is the sate as for calculating standard drag load in Step 14 of subsection C2.3 and will not be repeated here. 4 1 O 011880
22A4365 G- 24 Rev. 4 G2.5 LOCA CONPENSATION OSCILLATIONS LOADS Steam condensation begins af ter the vent is cleared of water and the drywell air has been carried over into the wetwell. This condensation oscillation phase induces bulk water motion and therefore creates drag loads on struc-tures submerged in the pool. The basis of the flow model for condensation oscillation load definition is derived from the work in Reference G4.4. The following procedure is recom-r,anded for calculating the loads on submerged structures:
- 1. Note the dimension of the containment (L and H) as shown in Figure G2.3.6.
- 2. Note the location of the submerged structure (x, y, z) .
- 3. Note the locations of the top vent exits (xgt, yg, zg t) .
- 4. Determine 1/r 9
effi 2 f r each vent. The parameter r gg is defined in h Appendix A of Reference G4.4 to account for the ef fects of pool boundaries and free surface by the method of images. Exclude those vents for which 1/r is small compared to the corresponding value for th a vent nearest to the structure.
- 5. Calculate the acceleration field from N
- 5 1 U, (G2.5-1)
=7 _
2 i=1 eff he re S = 188 ft /sec is source strength determined f rom Mark III 1/>7 scale test data and N is the total number of vents considered. 011880
22A4365 G-25 Rev. 4 O 6. Calculate the acceleration drag force from pU V,
=
" ^ (Q2.5-2)
F g A c
- 7. The forcing function may be approximated by a sine wave with an amplitude equal to Fg and a frequency range of 2 to 3.5 Hz.
- 8. The direction of the resultant force is approximately along the line joining the structure and its nearest vent.
G2.5.1 LOCA Condensation Oscillations Loads - Sample Problem Step 1. The submerged structure to be analyzed is that depicted in Figure G2.3.6. l () For simplicity, only three vents are considered here. The dimensions of the containment are: L = 18.5 ft H = 20 ft Step 2. The 1ccation of the submerged structure is y = 4.0 ft Y = 15.5 ft Z = 0 ft O 011880
22A4365 G-26 Rev. 4 Step 3. The location of vents are: Vent A B C X g1 0 0 0 Y 12.5 12.5 12.5 g1 Z 0 6.52 -6.52 gt Step 4. 2 1/r gg for vents A, B and C are computed as before, i Vent A B C 31 13 13 2 r eff' G theref ore
- 2
= 31 + 13 + 13 = 57
'eff S_tep 5.
The acceleration field is (Equation G2.5-1) 188 (f t) 1 2 U" = ,S . 1
= ., x (5 7) = 31.3 ft/sec 2
L' r ggg sec (18.5 f t)' i O 011880
22A4365 G- 27
, Rev. 4 s.
1 1 Step 6. The acceleration drag (Equation G2.5-2) per unit projected area is F = 0.73 psi A Steps 7, 8. The direction of this force is along the line joining vent A centerline and the structure centerline. The forcing function is approximated by a sine wave with an amplitude equal to FA and a frequency range of 2 to 3.5 Hz. G2.6 LOCA CHUGGING LOADS . 4 Chugging occurs af ter drywell air has been purged, and the vent mass flux falls below a critical value. Chugging then induces acoustic pressure loads () on structures sdamerged in the pool. The basis of the flow model for chugging load definition is derived f rom the work in Reference G4.4. The loads on submerged structures due to chugging are calculated from the procedure described below.
- 1. Locate the bubble center at 2.0 f t above the top vent centerline. l
- 2. Determine location of structure (x, y, z) relative to bubble i
center (see Figure G2.6.1) .
- 3. Calculate distance r from chugging center to structure r = [x +y +z J--
011880
22A4365 G-28 Rev. 4 _. _---_ :_ ~ ~- _:__ __ -_ l l l a sto v## 0/' STRUCTURE i
-o-- s to
/
/y ' '
FC /A * , t i SIDE VIEW STRUCTURE
'M TOP VIEW l
f Figure G2.6.1. Mark III Horizontal Vent Chugging 011880
22A4365 G-29 Rev. 3 O 4. Evaluate angle (0) between structure axis and r from cos e = cos a, cos ab + cos 8, cos 8b + cos y, cos yb where (cos a , cos S,, cos yg ) are the direction cosines of the s structure axis, while (cos a , cos B ' 8 Y ) .tre the direction b b b cosines of the vecter r from the bubble center of the structure. 4 4 . 5. Calculate chugging load from F C ( o# o) sin 0
= 2(2.53) sin 0
. where A is the projected area of the structure normal to its own axis, t.,P r o
= 2.53 psi-f t as the pulse strength, i.
- 6. Include the effect of another vent by repeating Steps 1 through 5.
< s) The pulse width is 0.002 seconds. Include those vents for which the signal arrives at the submerged structure within 0.002 second of each other. Use 4000 f t/sec for the acoustic velocity in water.
- 7. Add the two forces linearly.
l 8. Obtain time history as follows: l load duration is 2 msee i I period between individual chugs is 1 to 5 seconds l 1 I
- 9. For long structures, break the structure into separate sections and calculate the load on each section as abcve.
l 090779
i 22A4'365 G-30 Rev. 4 G2.6.1 IDCA Chugging Loads - Sample Problem Step 1. The submerged structure to be analyzed is depicted in Figure G2.3.6. l Step 2. Bubble location is X=0ft Y = 12.5 + 2.0 = 14.5 ft Z = 0 ft Structure location is X = 4.0 ft - Y = 15.5 f t Z = 0. f t Step 3. O Distance r is calculated as r = (4-0) 2 + (15.5 - 14.5) 2 = 4.12 ft for Vent A r = 4 +1 + 6.52 = 7.71 ft for Vent B or C Time for signal to arrive at submerged strticture: t = 4.12/4000 = 0.001 second, Vent A t = 7.71/4000 = 0.002 second, Vents B or C. Note that vents located further than B or C will have a signal that travels to the structure in more than 0.002 second. O 011880
22A4365 G-31 Rev. 4 O- Step 4. The angle 9 between the structure axis and the centerline of the bubble for the Vents A, B and C (see Figure G2.3.6) are:
- Vent A B C 0 90' 32.3* 32.3' Step 5.
The chugging load from each vent is calculated from Fc , 2.53 Sin e A r thus for each vent O ve t i e c 1.23 0.35 0.35 Step 7. Add the three forces linearly, to obtain the total force per unit area
= 1.9 psi O
011880
22A4365 G-32 Rev. 3 Step 8. The load duration is 2 msec, at 1.9 psi (h 4 j ,_ r 0.002 sec : Hence the time history is i The period between chugs is 1 to 5 sec. 4 j l 19 psi - I
-+-- 1 To 5 sec -
O 090779 l i
)
l l 22A4365 G-33 Rev. 3 (3 V G3. SUBMERGED STRUCTURE LOADS DUE TO SRV ACTUATIONS G3.1 Quencher Water Jet Load Following the actuation of a safety relief valve (SRV), water is rapidly discharged through the X-Quencher device attached at the end of the SRV line. A highly localized water jet is formed around the X-Quencher arms. The load induced outside a sphere circumscribed around the quencher arms by the Quencher water jet is small. There are no submerged structures located within the sphere mentioned above in the standard Mark III arrange-ment. The induced load for submerged structures located outside a sphere circumscribed by the quencher arm is negligible and is ignored.. G3.2 Quencher Bubble Load The analytical model for quencher air bubble loads on submerged structures is presented in Reference G4.3 and G4.5. The following procedure is recom-() mended to apply the analytical model for calculating loads on submerged structures due to quencher air bubbles. 4
- 1. Determine the location, dimensions, shape and orientation of the submerged structure. For more precise evaluation long structures should be divided into smaller segments with each segment being approximately 2 ft long.
I
- 2. Determine the initial location of the four bubbles. Each bubble l
will be assumed to form at the intersection of hole pattern center- l lines from adjacent arms (see Figure G3.1) . If the presence of pool boundaries or other structures prevent bubble formation at the location thus determined, assume the bubble is located along j the bisector between adjacent arms and is tangent to the boundaries or structures. l l r l 090779
22A4365 G-34 R';v . 3
- 3. Obtain values of the following parameters f rom Table A4.4 and the specific plant documents:
P : maximum bubble pressure, psia g minimum bubble pressure, psia Pmin: T g 7: initial pool temperature. R 11 : quencher arm submergence, ft 9 V:g initial air volume in the safety relief valve discharge 3 line (SRVDL), ft
~
P: initial air pressure in SRVDL, psia f T: initial air temperature in SRVDL, *R g P: containment air space pressure, psia c k: specific heat ratio of air a: water density at Tp9gy, Ib fe
- 4. Assume that the maximum volume of each bubble occurs when the pressure is at its minimum and the air in the bubble attains the surrounding pool water temperature and calculate the maximum bubble radius from V T P V = E ,ft (C3.2-1)
T P e 090779
l l /
).' -
~,} (
# 22A4365 G-35 Rev. 3
'y and l
1- > 1/3 R = V , ft (G3.2-2)
- 5. To account for the vertical motion of. the bubbles, the bubble rise I equation given below must be solved simultaneously with the 4
bubble dynamics equations for R(t), R(t), E(t) and Z3 (t), where 3 R(t) = bubble radius at time t 1 ~ l R(t) = bubble growth rate at time t i N( t) ' = rate of change of the bubble growth rate at time .t J 1 Z3 (t) = submergence of bubble center at time t Bubble Dynamics Equations
~
l 5(t) = (PB - )- (G3.2-3) I t i P " ~ (O * ~ ) l B B P, = Z (G3.2-5) ! PC+ b { . i' ., Bubble Rise Equation l i rp C b
~
## 8 * *B 8 D ( ~
j Z 3 b m B+23 "#
% e lO
)! v i-V 090779 i
- . _ . - __,. , ~ , - . . . . _ , - _ . - . , . . , , , _ . _ . . . ~ , - -
22A4365 G-36 R;v. 4
~ " * ***
- m =
fRair i R,g = gas constant of air Initial Conditions:
=
R(0) R R(0) = 0 Zb (0) = H . Zb (0) = 0 PB (0) = P min
- 6. Determine the location of images of the four source bubbles to account for the ef fects of pool walls, floor and free surface.
Then calculate the parameters X, Y, and Z, which are defined by Equation (A67) of Reference G4.3
- 7. For multiple quenchers use Equation (A79) of Reference G4.3 to j evaluate the parameters X, Y and Z. Note the Heaviside step functions H(t-t 7) and H(s -t) are introduced to account for phasing relations accng the quenchers of interest.
S. Using the results f rom Steps 5 through 7 calculate the equivalent uniform acceleration, ,(t), at time t at the structure location f rom
- 7 ..
., 7 /' '
U,(t) = R'(t) R(t) + 2R(t) R~(t) g X' + Y' + 2 (G3.2-8) 0 011880
I 22A4365 G-37 Rev. 3 The corresponding velocity, U, (t), may be obtained by numerically integrating N(t) . As a first approximation, U,(t) can also be ! evaluated from U,(t) = R (t) R(t) X +Y +Z (G3.2-9)
- 9. The acceleration drag is calculated from U (t) VA #
F A g c where 0 -n is the acceleration component normal to the structure and V is the acceleration drag volume for flow normal to the structure. A The standard drag force is calculated from O , PU' (t) F,(t) = C A (G3.2-11) D n g c where C is the drag coef ficient for flow normal to the structure. t D An is the projected structure area normal to U,9(t) . Add FAand l F5 at any time to t to get the total force on the structure or structure regnent. The direction of the total force is normal to the submerged structure. G3.2.1 Quencher Bubble Load - Sample Problem Steps 1, 2, 3. I i l The following geometrical and bubble data were used in the sample calcula-tion of the loads f rom one quencher to the structure shown in Figure G3.1. (a) Maximum Bubble Pressure: Psax
= 39.3 psia
= 10.1 psia (b) Minieum Bubble Pressure: P min.
h 090779
G-38 22A4365 Rev. 3 (c) Initial Fool Temperature: T pool = 560*R
= 13.9 ft (d) Quencher Arm Submergence: Hg (e) Initial Air Volume in the Relief Valve Discharge Line (SRVDL) 3 Vi = 56.13 ft (f) Initial Air Temperature in SRVDL: Ti = 560*R (g) Containment Air Space Pressure: Pc = 14.7 psia
= 14.7 psia (h) Initial Air Pressure in SRVDL: Pa (i) Figure G3.1 shows the geometrical locations of the bubbles and the structure for which quencher bubble load will be calculated.
The coordinate system is also shown in Figure G3.1. Steps 4, 5. lll The air leaving the quencher forms four independent and identical spherical bubbles which oscillate in phase while rising. The Bubble Dyaamics equa-tions (C3.2-3), (G3.2-4) and (G3.2-5) and Bubble Rise Equation (G3.2-6) were solved and shown below. Bubble Bubble Bubble P.sdial Bubble Radial Submergence Radius Growth Rate Acceleration Time (ft/sec 2) (sec) (ft) (f t) (ft/sec) i 1.695 0 -467.021 !
- 0. 13.900 !
1.691 -0.249 692.474 0.050 13.822 1.695 0.359 -441.970 O.110 13.611 1.692 -0.814 398.944 0.150 13.314 1.694 0.851 -282.412 0.200 12.972 1.694 -0.962 -175.039 0.250 12.608
. . 1 1
090779 ;
22A4365 G-39 Rev. 3 SUPPRESSION POOL l
' ^ ^ ^ - ^ ^ -
^
BURBLE m2 l (5.7,6.5 - 4.11 v Z
- = i s o l
BUBBLE 83 8088LEs1 f RPV O, - s -x l O> l ' W _J 'l . L __ _ _ (9.2,65.00) 12.5.65.00) \ SUPPR ESSION j POOL TOP VIEW BUBBLE.4
\
(144.45-40) d (57,65-4.11 E PLAN VIEW h WATER SURF ACE E a Y N N \ n
\ \,
\ 4
\ H qit l l 4 5 ,i lr \
! t
[ f - ~ ~l _, l-7~ l
^
V '--1 \ J ' --- W i i
Q v " .
5 6 ft } l i 1r if 8ASEMAT Figure G3.1. Four-Bubble Model for Quencher Air Discharge 090779
22A4365 G-40 Rev. 4 Steps 6 7. To account for the ef f ects of pool walla, floor and fill surf aco, the method of inagen given in (A62) of Ref erence G4. 3 was used. For simplicit y, the corroetion f actor K of (A84) of Reference G4. 3 is annumed to be one. The resulting parameters of X, Y and Z are shown a below. l l I t X Y 7. O. 0.0164 0.0170 -0.0111 ) 0.05 0.0361 0.0163 -0.0310 0.10 0.0356 0.0146 -0.0305 0.15 0.0346 0.0134 -0.0298 0.20 0.0334 0.0097 -0.02H9 0.25 0.0 120 0.0072 -0.0279 O 3 top _A. l'rs ing I qua t ionn G 3.2-8 and Gl.2-9. U, and 0, are calculated as follow : g l'. II ., H. O. -6H.162 0.05 . 0.0 ) J9 94.204 0,10 0 . 0 '. M ) -54.420 0,i3 -0 ,10 6 '. 52.2!.0 0 . .'O 0.1077 - 1 *, . 6 4 2 0.25 -0.1172 -21.182 O 0118A0
22A4365 G-41 l Rev. 3 .o Step 9. The normal acceleration and velocity components to the submerged structure are calculated and their associate acceleration and standard drag are also computed and shown as below. r Standard Acceleration Drag Drgg t U U (x 10 psi) (x 10 2 psi) _ _an. =m
-22.74
- O. O. -53.780 0.
0.05 -0.0282 78.522 -0.642 33.20 { 0.10 0.0396 -48.724 1.266 -20.60 f 0.15 -0.0865 42.471 -6.042 17.96 0.20 -0.0849 -28.086 5.821 -11.88 0.25 -0.0906 -16.374 -6.68 -6.92 a O . . . . . i l I The direction of the force is normal to the submerged structure and is given } as ! . + Xn~ y + Y rg i 2+Y2 ,, 72 i , I ! (Refer to Steps 6, 7.) i i L
- O 090779 i
_ . _ - . , . _ _ . . ~ . _ . ,_.. _ _ . . _ . . , _ . . , . . .. . . . _ . _ _ _ . . _ _ . .
~
G-43 22A4365 Rev. 4 G4. REFERENCES h
- 1. Mark III Confirmatory Test Program - Full Scale Condensation and Stratification Phenomena - Test Series 5707, NEDE-21853-P, August, 1978 (Proprietary Report) .
- 2. T. H. Chuang, Mark III One-Third Areas Scale Submerged Structure Tests, NEDE-21606P, October, 1977.
- 3. F. J. Moody, Analytical Model for Estimating Drag Forces on Rigid Submerged Structures Caused by LOCA and Safety Relief Valve Ramshead Air Discharges, NEDE-21471; revised by L. C. Chow and L. E. Lasher, September, 1977.
4 L. E. Lasher, Analytical Model for Estimating Drag Forces on Rigid Submerged Structures Caused by Steam Condansation and Chugging, NEDo-25153, July,1979.
- 5. T. H. Chuang, L. C. Chow,'and L. E. Lasher, Analytical Model for Estimating Drag Forces on Rigid Submerged Structures Caused by j LOCA and Safety Relief Valve Ramahead Air Discharges, NED0-21471, supplement 1; June, 1978.
O 011980
22A4365 H-1 Rev. 3 4 ATTACHMENT H
SUBJECT:
WEIR k'ALL LOADS DURING DRWELL DEPRESSURIZATION METHOD The calculations of the velocity of the water in the vent system during the negative drywell containment differential pressure are conservatively calcu-lated using the network shown in Figure H-1. The explanation of this network is given in Table H-1. The particular values used to determine the velocity as well as the unknowns are given in Table H-2. Neglecting inertial terms the equation for each loop is obtained by using the formula, The equations for the three paths are
+ (o.35a +1.5a-0.75)fl.0 21.0 = -3.25 + 2(32. 144) 0.5(4{2 9-
* + 5.71 1 .0 ,
/ '
)
+ + 1.5 21.0=-5.20+2(322$144) 0.5 (4 6.35(1 )-0.75I
* + 5.71 + 1.95
( + (1.55a - a ) .C;
+
i , n.0--7.15+2<32 Sit 1cc>o . 5 N1M + 7 2o Pif0T, y 2
+ (1.55a - a )
*(1.55(y_ - h_ ) .,
1 .0
+ + 5.71 + 1.95 + 1.95 (1}0 090779 s
22A4365 R,v. 2 H-2 These three equations are solved simultaneously for the quantites a, b and m. The velocity in the vents can be calculated using the equation am y , top 4.12 V = mi 4.12 y , (1-a-b)m bot 4.12 The impingement force on the weir wall (behind the individual vents) can be calculated using the equation for momentum loss; F = v^ gc - Results: V top
= 37.91 = 40 see
'" h V
mid
= 32.09 =35see
f* V = 26.65 = 30 sec bot F = 12800 lbf F = 9800 lbf g F = 7200 lbf 042178 9
22A4365 H-3 Rev. 2 V Table H-1 1 Pressure due to top vent submergence 2 Pressure due to mid vent submergence 3 Prsssure due to bottom vent submerget 4 Loss at entrance 5 Friction loss 6 Loss at Tee junction 7 Loss at entrance 8 Friction loss 9 Loss at Tee junction 10 Loss at entrance 11 Friction loss 12 Loss at Tee junction 13 Pressure loss due to elevation difference between bottom and mid vents 14 Friction loss 15 Loss in Tee junction ( }) 16 Pressure loss due to elevation dif ference between mid and top vents 17 Friction loss 18 Loss in Tee junction 19 Pressure loss due to elevation difference between top vent and top of wier wall 20 Friction loss 21 Loss at Exit b) v 042178
22A4365 H-4 Rev. 4 Table H-2 A Q K LP 1 -3.25 - aM 2 -5.20 - bM 3 -7.15 - (1-a-b)M - 4 - 4.12 aM 0.5 5 - - aM 0.0 6 - 12.00 M 6.35a + 1.5a - 0.75 7 - 4.12 bM 0.5 8 - - bM 0.0 9 - 12.00 (1-a)M 6.35 + 1.5 - 0.75 10 - 4.12 ( 1-a-b) M 0.5 11 - - (1-a-b)M 0.0 12 - 12.00 (1-a-b)M 7.10 13 1.95 - (1-a-b)M - 14 (1-a-b)M 0.0 15 - 12.00 (1-a)M 1.55 , 16 1.95 - (1-a)M - 17 - - (1-a)M 0.0 12.00 M 1.55a - a 18 - 19 5.70556 - M - 20 - - M 0.0 21 - 12.00 M 1.0
'P - Pressure Drop due to static head A - Flow Area Q - lcluce Flow Rate K - Loss Coefficient (fro: Idel Chik)
.1P = K ~q - =1 EK lgc 2gc a-
- of volume flow rate through top vent b '. of volume ficw rate through middle vent M - Total volume flow rate through the entire vent syste:
O 011550
22A4365 Rev. 2 H-5/H-6 (- N SURFACE ARE A WlER ANN.
- 480 f t2 , . h ,
NUM8EP OF VENTS = 120
^
SURFACE AREA PER 3 VENTS
- 12 f t2 DIAMETER OF VENTS = 2 7.5" @ '-
ARE A OF VENT CROSS SECTION = 4.12 f t2 WlER WALL HEIGHT
- 26*1" e e @e HWL HEIGHT
( TOP VENT HEIGHT
- 20'5" 12 11"
\\ kk q MID VENT HEIGHT = 8'5" 5 (BOTTOM VENT HEIGHT . 311a kh k PC-POW
- 21 pse Qg g g @
DW
%e O ke 'c 4e e e e )e AV .W AN 4e 4e 4e
}e e e e
,W .W AN b
4e 4e 4e @ e e
.W .W .W Figure H-1. Mark III Vent System Network 0'.2178
22A4365 Rev. 3 I-1/I-2 G'\ ATTACHMENT I POOL S'n' ELL VELOCITY Early in the Mark III program it became necessary to esteblish an upper bound pool swell velocity. This was accomplished by conservatively assuming that during the pool swell transient the top two rows of vents are open with air only flowing in each (air test data shows that breakthrough occurs just as the second vent opens - See NEDO-20550 Ref. 7). The pool surface velocity was then calculated using a simple volumetric flow rate calculation. l The following is a summary of the calculations: Using the 238 reference design, the total venting area between the drywell 2 2 atd containment is 481 ft , thus the area for two rows of vents is 320 ft , Assuming that during the majority of the pool swell transient the drywell pressure is typically 35 psia and the pressure of the air in the submerged bubble is typically 18 psia (atmosphere pressure plus 3 psi hydrostatic A pressure) then the pressure ratio across the vent system is 0.52. - y) Under these circumstances, classical compressible flow theory for flow in ducts with friction will give an inlet Mach number of 0.35 for a duct with a total loss coefficient of 3.5 (this is the Mark III value used in SAR calculations) with drywell stagnation conditions of 35 psia and 300*F (adiabatically compressed from 135'F initial conditions), this gives an air mass flux of 54 (lb/sec)/ft 2 or a total flow rate of 17,300 lbm/sec. Assum-ing that the temperature of the air in the bubble is equal to the pool temperature and using a pool eurface area of 5900 ft gives a pool surface velocity of 34 ft/sec. For design purposes, this was rounded off to 40 ft/sec I to cover such uncertainties as bubble temperature and pressure. 4 4 0 090779
l l 22A4365 ! l I Rev. 3 J-1 through J-52 q
- v
' ATTACHMENT J I SCALING ANALYSES AND SMALL STRUCTURE POOL SWELL DYNAMIC 10 ADS i l 1
l l l 1 l l l 4 O ATTACHMENT J is PROPRIETARY and is provided under separate cover, t i 1 1 O 090779 i i
i a 1 4 22A4365 K-1 ! Rev. 4 ATTACINENT K f RESPONSES TO NRC QUESTIONS i J l' i
@ This attachment has been deleted Pages K-2 through K-46 removed l i t
L l l 011880
. ..~
22A4365 L-1 Rev. 4 L1.0 Containment Asymmetric Loads This attachment discusses the potential for circumferential variations in' the ;
.LOCA dynamic loads and relief valve loads. The asymmetric loads are identi-fied and the data being used for containment design evaluation is presented.
Table L-1 is a tabulation of the postulated phenomena which could cause-
-loading asymmetries. The table either provides a reference for the asymmetric loads that are significant and should be considered, or provides a reference that justifies the assumption that a particular phenomenon does not lead to asymmetric loads of significance.
L2.0 Asymmetric Pool Swell As discussed in section 6, the maximum containment pressure increase associated with the bubble formation that f ollows vent clearing is specified as 10 psi. The basis for this specification is data from the large scale air blowdown tests that were conducted as part of the Mark III test program. Circumferential varia-Os tions in this relatively small pressure increase could result from either seis-mically induced submergence variations or variations in the vent flow composition (i.e. , air / steam mixture variations) . Increased submergence cculd lead to an increase in the load, liowever, PSTF data shows a very weak relationship between submergence and the containment pressure increase caused by bubble formation. The survey of the PSTF data shown in Figure 6.6 shows that for tests having the same drywell pressure at vent clearing, variations of up to 6 ft in submergence lead to variations in the bubble load of 2 to 3 psi; it is concluded tha; varia-tions in suppression pool depth due to seismically induced waves will not lead i to significant asymmetric containment bubble loads. , i 1he bubble loading specification of 10 psi being used for Mark III design was derived from an air test and is thus the most conservative in terms of vent flow composition. Any steem in the vent flow would be condensed and this would lead to a less rapid pool acceleration and thus a reduced pressure load on the con-I It should be noted that PSTF data shcws that the high degree tainment wall. I of turbulent mixing in the drywell during a LOCA leads to a uniform mixture of air and steam in the vent flow. This condition will also exist in the full
- scale Mark III and this uniform vent flow composition will preclude any 4
011880
~-.
Table L-1 is there the potential for significant Asymmet ric leads asymmetric 15eing I' sed for Design Evaluation comunents Phenomena con t ainme nt loads Seismic induced pool No See Attachment B 1. surf;ce waves Selsn:ic induced changes in Yes See Attachment B 2. the pool hydrostatic pressure Yes See Attachment A y<: >lj
- 3. Relief valve actuation .O u
NO 0 Loads are of negligible No
- 4. Jet 1.oads duting vent magnitude (see 6.1.2) cleaning lioth 0 Loads are of negligible Sonic and compressive waves No
- 5. .nagn i t ude (see 4.1.1 and 6.1.1) 0-10 psi See following discussion.
liubble pressure load Yes
- 6. ,
No O See Attachment F
- 7. IICU floor flow pcessute differential S
es No 0 1.oads on the containment t-* h 8. Fall back are of negligible magnitude 4 (see 6.1.7).
- O e
- f. J i
Table I.-l (Continued) i i Is there the potential for l Asymmetric 1.oads significant asymmetric Being Used for , Design Evaluation Comments t l i Phen.mena containment loads l No 0 Loads on the containment
- 9. Post LOCA waves are of negligible magnitude (see 6.1.8)
No O This is a relatively slow
- 10. Containment pressurizatinn charging process. See Figure 4.4. ,
awN J No 0 Loads are small (see - 3
- 11. Condensation oscillations 6.1.9) mg 1 See following discussion.
No O
- 12. Chugging :
i I. No O See Section 10.1 and
- 13. Pool Swell loads with following discussion.
selspic induced waves , i present ! i l i l i 4 o I i b F U b : m 1 4 t
22A4365 L-4 Rev. 4 significant circumferential variations in the containment bubble formation loads. In addition Attachment D shows no significant circumferential varia-tions in drywell pressure that could lead to variations in vent flow rates and thus pool swell. Despite strong evidence that circumferential variation l in the containment bubble load will not occur, an arbitrary loading combination of 0 psid on one side of the containment with a simultaneous 10 psid load on the other side should be considered to account for any uncertainties about asymmetric loading conditions. The conservative asymmetric condition assumes that all air is vented on half of the drywell periphery and steam is vented on the other half. The large scale PSTF test data is the basis for specifying the maximum asym-metric load of 10 psi. Figure 6-6 is a summary of all the peak containment wall pressures observed in PSTF tests during the bubble formation phase of the A maximum increase of 10 psi blowdown. Figure 6-4 shows a typical transient. on the containment wall was observed in the PSTF at the Mark III drywell peak calculated pressure of 36.5 psia; Figure 6-6 shows the maximum increase close to zero. Thus, use of a 10 psi asymmetric pressure condition applied in a worst case distribution as a bounding specification will be used for contain-ment evaluation. L3.0 Ass = metric Chugging An analysis was performed to determine the possible asymmetric chugging It was assumed that all of the vents chugged simultaneously, but all effects. vents on one half of the dryvell were at a maximum 90-90; tolerance limit pressure while the other half of the vents in the opposite 180* sector were I at a ninimum 90-90*. tolerance limit pressure. The resulting differential ' f orces were then applied to the pcol boundaries and weir annulus. Overturn-M4 maner.ts wre calculated and compared to the currenti f specified asymmetric pool well load (Sections 4.1.9 and 6.1.3). The current specification loads (from asymmetric pool swell) result in moments t. ice as large as the asym- ,
'.H ric chugging moments. Since the current asymmetric pool swell specifica- f, tion bounds the conservatively calculated asv :netric chugging results by a large margin, asy : metric chugging is not a design basis lord.
O 011880
22A4365 M-1 Rev. 2 ATTACHMENT M MULTIPLE SAFETY / RELIEF VALVE ACTUATION FORCING FUNCTION METHODS TABLE OF CONTENTS TITLE PAGE SECTION M
1.0 INTRODUCTION
M-5 M-6 M2.0 RANDOM PARAMETERS M2.1 Reactor Vessel Pressure Rise Rate M-6 M2.2 Valve Setpoint M-6 M2.3 Valve Opening Time M-7 M2.4 Quencher Bubble Frequency Distribution M-7 M3.0 MONTE CARLO TRIAL SIMULATIONS M-13 M3.1 Approach M-13 M3.2 Bubble Arrival Time M-14 M3.2.1 Calculation of Reference Arrival Time !!-14 M3.2.2 Adjustment of Bubble Arrival Time for Valve Setpoint Variations M-14 M3.2.3 . Adjustment of Bubble Arrival l Time for Valve Opening Time Variations M-14 l M3.3 Quencher Bubble Frequency Variations M-15 l M3.3.1 Adjus: ment of Quencher Bubble Frequency for Discharge Line Volume M-15 j M3.3.2 Adjustment of Quencher Bubble Time History for Selected Frequency M-15 042178 i
i 22A4365 M-2 Rev. 2 TABLE OF CONTENTS (Continued) h PAGE SECTION TITLE M4.0 FACTORS AFFECTING PRESSURE DISTRIBUTION ON Tile M-15 SUPPRESSION POOL BOUNDARY M-16 M4.1 Bubble Pressure Attenuation Line-of-Sight Influence M-16 M4.2 M4.3 Combination of Multible SRV Pressure Time M-16 Histories M-16 M5.0 FORCING FUNCTIONS FOR NSSS EQUIPMENT EVALUATIONS M-16 50 .1 Time Sequencing M-17 MS.2 Pressure Time Histories M5.3 Vertical Basemat Force and Overturning M-17 Moment Fourier Spectra M-17 M5.4 M-18 M6.0 STRUCTURAL RESPONSE ANALYSIS 0 042178
22A4365 M-3/ M-4 Rev. 2 .O LIST OF ILLUSTRATIONS TITLE PAGE FIGURE M-8 M2-1 Probability Density Function vs Pressure Rise Rate M2-2 Probability Density Function vs Valve Group M-9 Setpoint Variation M2-3 Probability Density Function vs Valve Opening Time M-10 Variation Probability Density Function vs Bubble Frequency M-ll M2-4 M-12 M2-5 Quencher Bubble Pressure Time History MA-5 MA-1 Basemat Load vs' Time Fourier Spectrum of Basemat Force MA-6 MA-2 Fourier Spectra of Vertical Basemat Force MA-7 MA-3 O O 042178
22A4365 M-5 Rev. 2 M
1.0 INTRODUCTION
This attachment describes the procedure for determining the safety / relief valve (SRV) discharge 95-95 percent confidence level forcing functions that are imposed on the containment structure to obtain structural responses which are used as input for the evaluation of equipment located within the containment. The pro-cedure utilizes the random nature of several parameters that significantly in-fluence the phase relationship of the individual air bubbles formed in the
~
~
l suppression pool during multiple SRV discharge events. The random variables that are utilized in this procedure are 1) SRV Setpoint Tolerance, 2) Valve Opening Time, 3) Reactor Vessel Pressure Rise Rate, and 4) Quencher Bubble Frequency. Other parameters that. influence the phase relationship are being studied for f'uture application.
' The maximum positive and negative bubble pressures for each individual discharge location are determined by using the method described in Section A12.6 of O)
(, Attachment A. It should be noted that test data indicated randomness in the peak pressure amplitude which could also c used for determining structural response. This is also being studied for future coplication. Of the SRV cases identified for consideration in containment structural design (Table A4.4 of Attachment A), the expected bounding vertical response at equip-ment locations is based on the all valve case. The expected bounding horizontal f response is based on either the single valve subsequent actuation, two adjacent { valves, or the all valve case. The ADS case is also evaluated. From each of i these four cases, the Fourier Spectra of the forcing functions fer 59 Monte Carlo simulations of the event are pio ted. A bounding forcing function is then selected in each of the frequency ranges of interest for use in developing the dyr.amic responses at a selected location on the containment structure (i.e., basecat, drywell, and centainment). These dynamic responses are then e= ployed for NSSS and BOP equipment. evaluations. A dynamic time history analysis is performed to determine the acceleration time histeries, response spectra, and displacements needed. Dynamic responses for equipment evaluations are made by O . ! 101678 l
- . - - , ~ .,n
1 22A4365 M-6 Rev. 2 enveloping the results from the selected trial cases with the largest Fourier spectra magnitude in each frequency interval. For clarification, an example is presented in Appendix M.A to this attachment. M2.0 RANDOM PARAMETERS M2.1 Reactor Vessel Pressure Rise Rate (PRR) The pressure rise rate distribution for BWR/6 plants is shown in Figure M2-1. The distribution is determined from an evaluation of BWR/6 transient events. The figure represents the probabl.lity density function for pressure rise rates for events opening > 2/3 of the SRV's, weighted by the relative occurrence of the events and averaged over all reactor conditions anticipated during the last 40% of an operating cycle. The lower limit of 40 psi /sec is the minimum pres-sure rise rate expected to open 2/3 of the SRV's. The upper limit of 140 psi /see has a high probability of not being exceeded for any operating condition. It should be noted that the PRR variable is only used in the all valve case Monte Carlo event simulations. M2.2 Valve Setpoint The setpoints for SRV's on BWR/6 are arranged in three groups with redundant logic trains consisting of a pressure transducer and three pressure switches. The logic for the 238 BWR/6 design consists of one pressure switch set at 1103 psi, nine on a pressure switch set at 1113 psi, and the remaining nine on a pres-sure switch at 1123 psi. A testability feature is also included which utilizes pressure trip instrumentation. The tolerance on the pressure switch setpoints with this testability feature is based on a normal (Gaussian) distribution with a standard deviation of 2 psi as shown in Figure M2-2. For the grouped arrange-ment, the standard deviation is applied to the group seccoints; thus, the valves within the ,;roup will have the same adjustment. O 101678
22A4362 Rev. 2 M-7 The SRV arrangement and pressure setpoints for the Mark III standard plants are identified in Figures A4-3 through A4-9 of Attachment A. The actual location of the quenchers in the suppression pool is defined by the purchaser. M2.3 Valve Opening Time (VOT) Test data indicates that there is a normal distribution for the VOT with a standard deviation of 0.009 seconds as shown in Figure M2-3. M2.4 Quencher Bubble Frequency Distribution (QBF) A . typical forcing function for a quencher SRV bubble with a frequency of 8 Hz is shown in Figure AS.ll of Attachment A. The bubble lasts effectively 0.75 seconds in the suppression pool. In the 8 Hz bubble, the pressure decays to one-third of the peak value ever 5 cycles and a complete pressure cycle oscilla-tion period lasts 0.125 seconds, 0.05 seconds for the positive pulse and 0.075 secons for the negative pulse. For other frequencies, the same damping definition frg applies, i.e., two-third decay over 5 cycles, or 0.133 decay per cycle. V The quencher bubble pressure time history in Figure AS.ll of Attachment A is an idealized bubble model . For the purposes of this procedure a pressure time history curve is constructed by assigning half sine waves to both the positive and negative portions as shown in Figure M2-5. The P ,, and P min rari s and the positive and negative pulse duration periods are maintained. This provides a time history that is more representative of the test observations and allows for computer simulation. Quencher test data shows that the f requency of the air bubble is a function of the SRV disenarge line air volume. The distribution of bubble frequencies for a discharge line air voluma of 50 ft. is shown in Figure }C-4 and is used as the reference for this procedure. This reference value is the SRV line volume from the operating plants from which the Quencher bubble frequency ?sta was i obtained. The normal distribution for' the curve has a mean frequency of 8.1 Hz with a standard deviation of 1.7 Hz. It is truncated at the minimum and maximum bounds of 5 and 12 Hz. O 101678
. - . . _ ,_, _ , . _ . - ,.___m ..-..-y . . - _ _ . , _ . . . . , , , _ , . _ . .
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l 22A4365 g_y3 r Rev. 4 f~ C M3.0 MONTE CARLO TRIAL SIMULATIONS M3.1 Approach There are four SRV cases that are considered to get bounding forcing functions far the equipment evaluations. They are: _ Single Valve subsequent actuation
- Two adjacent valves
- ADS valves All valves j
In each of these cases, 59 Monte Carlo trials are performed in which appropriate random variable adjustments are selected for the parameters listed in Section M2.0. For the single valve subsequent actuation case only the quencher bubble frequency is varied. For the ADS and two adjacent valves cases, the valve set point l
! tolerance and pressure rise rate considerations are not incorporated for obtain-ing the forcing function because the entire group of ADS valves is simultane-ously activated by a single signal. For all valve case all variables are considered.
I The all valve trials each consist of selecting a random pressure rise rate from Figure M2-1 and a random pressure switch setpoint for each group of SRVs using Figure M2-2. This information is used to compute the bubble arrival time differ-4 ence or separation between each group of valves. These bubble arrival times are adjusted for each individual valve by randomly selecting a time variation due to valve opening time (VOT) using Figure M2-3. 1 i Once the bubbles are in the suppression pool, each bubble frequency is randomly varied by selecting a f requency f rom a unique distribution for the discharge line g volume involved. See Figure M2-4 for typical distribution for discharge line fJ
\- with an air volume of 50ft 3
-The bubble time history for each' valve location 011880
l 22A4365 M-14 Rev. 2 is then used to determine the forcing function on the suppression pool boundary by utilizing the methods described in fection A10.3.1 of Attachment A. For the ADS and two adjacent valve cases, each trial assumed that all valves are actuated together and then bubble phasing is adjusted by randomly selecting a time variation due to VOT for each valve. Each bubble frequency is then ran-dcmly selected as for the multiple valve trials. For the single valve case only the bubble frequency is varied. M3.2 Bubble Arrival Time M3.2.1 Calculation of Reference Arrival Time The arrival time for each air bubble in the suppression pool relative to the lowest set SRV is a function of the SRV setpoint arrangement and the reactor pressure rise rate. Assuming no tolerance on setpoints, no variation in valve opening time (VOT), and randomly selecting a pressure rise rate (PRR), the arrival times of the bubbles in the suppression pool are computed by dividing g the nominal setpoint differences (i.e., ap = 10 and 20 psi for BWR-6) by the PRR. It should be noted that SRV discharge line lengths are not considered. For BWR-6 with nominal setpoints at 1103, 1113, and 1123 psi the time separation is 0.077 and 0.154 seconds, based upon PRR = 130 psi /sec. M3.2.2 Adjustment of Bubble Arrival Time for Pressure Setpoint Variations Each all valve Monte Carlo trial will include an adjustment of the bubble arrival times as calculated in Section M3.2.1 by slightly increasing or de-creasing the valve setpoint for each group of valves. This is done by using a random number generator code to select valve setpoint variation from the distribution shown in Figure M2-2. M3.2.3 Adjustment of Bubble Arrival Time for Valve Opening Time Variations Each Monte Carlo trial will include an adjustment of the bubble arrival time as calculated in Section M3.2.4 by slightly increasing or decreasing the VOT l 101678
22A4365 M-15 Rev. 2
- for each valve. This is done by using a random number generator code to select VOT variation from the distribution shown in Figure M2-3.
1 M3.3 Quencher Bubble Frequency Variation M3.3.1 Adjustment of Bubble Frequency for Discharge Line Air Volume As indicated in Section M2.4 the frequency of the quencher bubble is a function of the SRV discharge line air volume. A reference line air volume of 50 ft has been selected to generate the bubble pressure time history shown in Figure M2-5. For each SRV discharge line volume a unique frequency distribution is generated by adjusting all of the characteristics (mean, standard deviation, lower bound, upper bound) of the reference distribution curve by multiplying by the cube root of the ratio of 50 ft to the actual air volume in the SRV discharge line. For example, the adjustment of frequency for a 100 ft line volume is: 8.1 Hz x = 8.1 x 0.79 - 6.4 Hz o Examples for the other characteristics: Y Volume Mean Std. Dev. Lower Bound Upper Bound (ft ) (Hz) (Hz) (Hz) (Hz) 50 8.1 1.7 5 12 100 6.4 1.3 4 9.5 i I M3.3.2 Adjustment of Quencher Bubble Time History for Selected Frecuency In each Monte Carlo trial, a random number generator code is used to select a j
- i
! frequency f rom each of the f requency distribution curves generated in Section M3.3.1. For each frequency selected, a time history of the Quencher bubble pressure oscillation is generated by adjusting the reference time history (8.0 Hz). This is accomplished by maintaining the ratio of negative to i positive pulse period constant. The pressure cycle period, positive pressure pulse time and negative pressure pulse time are adjusted by multiplying each D by the ratio of the reference f requency (8 Hz) to the selected frequency. d l For example, for 6 Hz: l 101678 l
22A6365 M-16 Rev. 2 8 Hz Pressure cycle period = 0.125 sec. 6 Hz
= .1 s i'c .
8" Positive pressure pulse time = 0.05 sec. 6H
= 0.067 sec.
Negative pressure pulse time = 0.075 sec. 86 Hz Hz
= 0.100 sec.
Nu=ber of cycles per , Bubble duration ,0.75 sec. = 4.5 cycles 0.75 sec. duration Pressure cycle period 0.167 sec/ cycle M4.0 FACTORS AFFECTING PRESSURE DISTRIBUTION ON THE SUPPRESSION POOL BOUNDARY M4.1 BUBBLE PRESSURE ATTENUATION The attenuation of the bubble pressure with distance r f rom the quencher is 2r /r where r - radius of the quencher - (4.87 f t) and r 3 2r 9(see Section A.10.3.1 of Appendix 33) . r = true spatial distance from the quencher center to the node. Mk.2 LINE-OF-SIGHT INFLUENCE . The line-of-sight criterion for the bubble pressure states that points which cannot be seen through a direct line from the outer radius of the quencher arms to the location in question will not be af fected by the pressure f rom that quencher (see Section A.10.3.2.1 of Appendix 3B). M4.3 COM3INATION OF MULTIPLE SRV PRESSURE TIME HISTORIES The time sequencing application providea a given phase relationship between quencher bubbies. Ihe pressure at each node point and time step is calculated by combining the centribution f rom each valve (in the line of sight) using al;ebraic sue ation. At each node where th: total calculated pressure at any time step exceeds the maximum pressure (positive or negative) from any of the contributing valves, the calculated pressure at the specific time step is set I equal to the maximum bubble pressure at the same instant in time. O 101678
l 1 22A4365 M-17 Rev. 2 MS.O FORCING FUNCTIONS FOR NSSS EQUIPMENT EVALUATION , l MS.1 TIME SEQUENCING Time sequencing with random parameters is used to arrive: at the forcing function for the multiple SRV air-clearing events referenced in Section M3.1. A Monte Carlo technique is used to generate the building forcing function for j equipment evaluations. The bounding forcing function from 59 trials will result in a 95% confidence level that 95% of the time the actual forcing
' function will be less than the forcing function determined by the Monte Carlo technique.
l M5.2 PRESSURE TIME HISTORIES i 1 Fif ty-nine (59) cases of pressure distribution on the pool boundary are calcu-lated using the random parameters delineated in Section M2.0. . I M5.3 VERTICAL BASEMAT FORCE AND OVERTURNING MOMENT J O The total basemat force is calculated as a function of time by integrating the node pressures over the suppression pool basemat incremental areas. The over-turning moments (about two perpendicular horizontal axes through the basemat center upper surface) are calculated, as a function of time, by integrating the product (node pressure x the incremental area moment arm x the incremental area) over the suppression pool boundary (containment, basemat, and drywell wall). M5.4 FOURIER SPECTRA Fourier spectra
- of the vertical basemat force and overturning moment for the 59 cases are developed for selecting the cases used to determine dynamic res-ponses for equipment evaluations. The significant frequency range is divided 4
into three frequency intervals as determined below: l r J i
- Reference 1: Cooley, J.W., & Tukey, J.W., (1965), "An Algorithm for the
! Machine Calculation of Complex Fourier Series," Mathematics of Co=putation, Vol.19, No. 90, pp 297-301. () 2. Shingleton, Richard C., "On Computing the Fast Fourier Transform," Communication of Applied Computation Mathematics, (10(10) 1967, pp 647-654 101678
4 22A4365 M-17A Rev. 2 l i Step 1. Adjust the mean frequency of each safety / relief valve ) discharge line for air volume differences, see Sub-section M3.3.1. Step.2. Calculate the mean frequency (fm) for all applicable safety relief valve discharge lines. Step 3. Establish the frequency intervals based on 0.5 fm to 1.5 fm, 1.5 fm to 2.5 fm, and 2.5 fm to 3.5 fm. 2
- 1 where fm = N fi; i = 1,...,N N = total no. of valves actuated
- The basemat loading cases with the largest spectral value within each frequency interval (f rom the 59 cases) are selected for determination of equipment responses.
O I i i i i i !:O l J 101o,3 l i- . . . - - -
I l 22A4365 M-18 Rev. 2
- 16 . 0 STRUCTURAL RESPO!!SE 4ALYSIS Forcing functions corresponding to the case selected in each frequency range (selected in Section M5.4) are used as input to the structural analysis.
The Structural dynamic analysis is then performed for these selected cases. resulting dynamic responses are then enveloped for NSSS and BOP equipment evaluations. O O 042178
I 22A4365 MA-1 l Rev. 2-APPENDIX MA-EXAMPLE OF TYPICAL TIME SEQUENCING APPLICATION This example is provided to clarify the time sequencing procedures provided in this attachment. Typical random parameter values are used to outline the steps required to determine the bounding vertical base =at force. Examination of the Fourier spectra for the vertical basemat force and overturning moments permits calculation of bounding equipment responses. Guidelines for selecting the bounding responses for equipment evaluations are included. MA. RANDOM PARANETERS The following random parameters are used: pressure setpoints, valve opening time, and vessel pressure rise race. The random parameter values used in this example problem are: J (1) Pressure rise rate distribution per Subsection M2.1. (2) Pressure setpoints variation per Subsection M2.2. 4 i
- l
. Mean Standard 2
, S et point Deviation Valves (psi) (psi) 1 1103 2
- 9 1113 2 j 9 1123 2 (3) Valve opening time variations per Subsection M2.3 ,
i l Standard deviation = 0.009 sec. i Step 1 An 80 psi /sec vessel pressure rise rate was randomly selected from Figure M2-1. 7 I
- Note that this example is for the 238 B'a'R/6 Mark III standard plant with a ganged valve arrangement.
- 101678 4
22A4365 MA-2 Rev. 2 Step 2_ g The valve pressure setpoints are randomly selected f rom a random number generator code using the distribution given in Figure M2-2. The valve pressure set-1114.3 psi, and points from a typical random selection are 1104.5 psi, 1124.6 psi. Step 3 The relative valve opening time for each of the two groups of 9 valves is calculated: Valve setpoint g (psi) - Valve setpoint(g (psi) i sec) = Pressure rise rate (psi /sec) where i = 2, 3 (the number of subsequent valve groups), and 1 = the reference valve. llence, for i = 2, the valve opening time for the first group of 9 valves is: T = 1114.3 - 1104.5 = 0.1225 sec 80 Step 4 Ihe bubble arrival time is calculated by addi.4g the group valve opeting time and a randemly selected delta time for each valve using the valve openinh time distribution shown in Figure M2-3. Therefore, for each quencher the (IVOT). bubble arrival time - T(group) + individual valve opening time For this sample problem, the typical set of randomly selected IVOT's for the distribution values stated above are: 101678
22A4365 Rev. 2 MA-3 Valve No. IVOT (sec) Valve No. IVOT (sec) Valve No. IVOT (sec) 1 0.067 7 0.067 13 0.056 2 0.069 8 0.051 14 0.061 3 0.065- 9 0.062 15 0.056 4 0.059 10 0.065 16 0.065 5 0.063 11 0.058 17 0.057 6 0.038 12 0.057 18 0.071 19 0.069 Note that a mean value of 0.057 see is included in the above numbers. Adding these values to the group T 1 calculated in Step 3 and nor.alizing to have the first bubble arrive at zero time results in the following bubble arrival times: Arrival Time Arrival Time Arrival Time Valve No. (sec) Valve No. (sec) Valve No. (sec) 1 0.125 7 0.125 13 0.243 2 0.256 8 0.238 14 0.127 3 0.123 9 0.120 15 0.243 4 0.247 10 0.0 16 0.124 5 0.122 11 0.246 17 0.245 6 0.225 12 0.116 18 0.129 19 0.256 4 O M.3 BUBBLE FREQUENCIES Bubble frequencies for individual quenchers are randomly selected from a random number generator code using the distribution shown in Figure M2-4. Typical random bubble frequency values for the 19 quenchers are: Valve No. Frequency (Hz) Valve No. Frequency (Hz) 1 6.56 11 7.22 2 9.77 12 5.39 3 9.15 13 5.68 4 5.01 14 8.60 5 9.33 15 9.86 6 6.88 16 7.04 7 9.41 17 11.08 8 9.10 18 8.68 9 7.92 19 8.52 10 11.14 J NOTE: For this example, all lines are considered as uniform in length and fre-quencies are randomly selected from one Quencher Bubble Frequency (QBF) distribution curve (Figure M2-4). In this example, mean = 8.23 Hz and p c = 1.80 Hz. With nonuniform line lengths, Subsection M3.2.1 is used to d develop unique QBF distribution curves from which a frequency is randomly selected for each line. 101678 l
22A4365 MA-4 Rev. 2 M.C The forcing function is calculated by ccmputing the pressure distribution around the pool boundary using the criteria defined in Section M4.0 which are: (1) 2rg/r attentuation, r =g quencher radius and r 2r g. (2) Line-of-sight influence. (3) Algebraic summation at each time step of the individual pressure waves. (4) Truncation of the total calculated pressure to the maximum bubble pressure of any of the pressure waves in the pool at each time step. The basemat force vs time shown in Figure MA-1 is computed for a typical trial case. The Fourier spectrum of this basemat force (Section M5.0) is calculated g in Figure MA-2. M.D A Monte Carlo technique is used to generate 59 forcing functions. This gives 95% confidence and 95% probability that these loads will not be exceeded. The significant frequency range for building and equipment evaluation is then divided into several frequency interv11s. Out of these 59 trials, the maximum trial case is selected for each frequency interval based on the peak Fourier amplitudes of the integrated vertical basemat forces or overturning moment, in that frequency interval. Figure MA-3 shows an example of this selection procedure. Structural The dynamic analyses are performed for these potential critical cases. resulting dynamic responses are then enveloped for NSSS and BOP equip =ent evaluation. O 101678
l 22A4365 MA-5 Rev. 2
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22A4365 MA-6 i Rev. 2 4 6 O i 9 0 O i W t o i Q u 2 y
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22A4365 MA-7/MA-8 Rev. 2 O b I EXAMPLE l RUN 12 RUN 48 - - == R U N 26 RUN12 H RUN37 w l M RUN 48 b h RUN 26
/ \
/ \
"' \
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~2 3 ,, -
RUN 37
. FREQUENCY fq fe f fr a FREQUENCY INTERVAL Figure MA-3. Fourier Spectra of Forcing Functions
_ NOTES:
- l. Fourier Spectra of forcing function for all 59 Monte Carlo runs are plotted.
- 2. The above example shows maximum forcing functions in the three selected frequency intervals.
Run 48 is max. for frequency interval, f. 1 1 Run 12 is max. for frequency interval, f,
^2 Run 26 is max. for frequency interval, f.
1 3
- 3. Run 37 is a typical ncn-maximum case.
- 4. The time histories for Runs 48, 12, and 26 are used in developing dynamic responses.
- 5. The dynamic responses that result from these forcing functions are then enveloped for NSSS & BOP equipment evaluations.
O l l I l 042178
22A4365 N-1 Rev. 2
~ /%
ATTACHMENT N SUPPRESSION POOL THERMAL STRATIFICATION is) N
1.0 INTRODUCTION
During the period of steam condensation in the suppression pool, from a postu-lated LOCA, the pool water in the immediate vicinity of the vents is heated because of the energy release. For the Mark III suppression pool configuration, 't most of the mass and energy is released to the pool through the top vents. As a result, the top portion of the pool is heated more than the lower portion. By natural convection the hot water rises and the cold water is displaced toward the bottom portion of the pool. The vertical temperature gradient resulting from these effects is known as thermal stratification. N1.2 REVIEW OF TEST DATA During the LOCA blowdown, the pool vertical temperature profile varies not only with time but also with the distance from the vent exit. Figures N-1
/"'s and N-2 present the typtcal temperature profiles for a large break liquid \',/
blowdown. In Figure N-1, which shows the profiles measured for the half pool near the drywell wall, the temperature peaks at the elevation of the top vent during the initial stages of the blowdown (t E 25 sec), indicating concentrated energy discharge through the top vent. As blowdown proceeds (t 2 25 sec), the temperature profile smooths out due to thermal mixing, turbulence, and pool agitation by chugging. In the other half of the pool away from the drywell wall, the temperature profile, as shown in Figure N-2 is not as steep as that of Figure N-1 at the early stages of the blowdown. However, toward the end of the blowdown the temperature profiles are nearly the same throughout the entire pool. In general, the steam blowdowns in PSTF give less stratification than liquid blowdowns of the same break size. This is attributed to the smaller total energy release associated with the secam blowdowns. For the full scale plant the energy from either break is equal. Thermal stratification is also dependent on the break size for the same blowdown fluid type. Large breaks create more t /'"g stratification than small breaks because energy de,osition in the pool is more l \_) 042178
22A4365 N-2 Rev. 3 rapid. Since the specific hea- of water is essentially constant within the temperature range from 70*F to 7.00*F, the temperature rise of the pool is independent of the initial pool temperature for a given amount of energy input. As a result, the initO 1 pool temperature has little ef fect on thermal stratification. Ni- 3 APPLICATION To determine the maximum temperat tre profile for structural evaluation, it is assumed that the energy deposition distribution as a function of submergence is the same for the 1/3 area scale (Test Series 5807, Ref.15) as for the full scale plant. Dividing the pool depth into five equal segments, the percentage energy deposition distribution for the maximum stratification expected is established as follows: Height of Segment i Segment No. (1) in % of Total Fool % of Total Energy From Pool Top Depth (H /H) Deposition (E ) 1 20 25 2 20 23 22 g 3 20 4 20 20 5 20 12 To obtain the temperature profile for a prescribed initial pool temperature (T ) and total blowdown energy, the bulk pool temperature (T) from energy o balance at the end of the blowdown was calculated, then the mean temperature (T ) for each pool segment was determined from: Tg = E (i - I g) g
+T g where H is the total pool depth, H is the height of the i segment, and E t ch is the fraction of total energy deposited in the i seg=ent. Assuming the the mean temperature of each segment occurs in the middle of the segment, temperature profile is readily plotted. Note the above table is valid only for a top vent initial submergence of 7.5 ft.
O 090779
22A4365 Rev. 2 N-3 r s
.] N1-3.1 Stratification During Large Break Accident For design evaluation of the large break accident, a total energy discharge of 4 x 108 Btu into a 1000F pool with 8 x 10 6 lba of water was assumed. The ,
mean pool temperature after energy release is: x 108 100 + 50 = 150 F T = 100 + 8 x 106 xi T1 = T2 = 0.23(150-100) (5) + 100 = 157.50F T3 = 0.22(150-100) (5) + 100 = 1550F T4 = 0.2(150-100) (5).+ 100 = 1500F T5 = 0.12(150-100) (5) + 100 = 130oF O) ( , Figure N-3 shows the resulting pool temperature profile. Note that, although the temperature dif f erence f rom top to bottom is almost 300F, the peak tem-perature is only 7.5 F above the mean. N1-3.2 Stratification Durina Intermediate and Small Break Accidents Figure N-4 shows the localized nature of the energy addition as observed in i the PSTF Full Scale Tests (Reference 16). The localized energy addition f i (through the top vent) f rem the f ull scale tests is more representative of the smaller accident breaks. Test results show that, for a very limited blowdown (about 2 minutes) with much less energy added to the pool than prototypical, the temperature in the lower pool region (%6 feet) was es- . i sentiall*, unchanged and the upper pool region was uniform 1v heated. This - thermal stratification profile will not persist in actual conditions, since ; i ECCS suction and return will promote pool mixing. The long term profile will essentially be as shown in Figure N-3. /~N O 101678
22A4365 N-33 Rev. 2 Since Figure N-3 from 1//3 PSTF results shows a thermal gradient near the bottom of the pool, and full scale tests (Reference 16) show the gradient at higher locations, it is conservatively recoc: mended for design evaluations that the maximum te=perature gradient shown in Figure N-3 be applied from the lower pool region up to the top vent centerline. For the upper parts of the pool (above top vent centerline) the te=perature profile, from full scale and 1//3 scale tests, shows uniform heating (Reference Figure N-3). O } \ O i 101678
22A4365 Rev. 2 N-4 O; l I l d l 1 I This figure is PROPRIETARY and is provided under separate cover.
- G i
4 i c 1 I l Figure N-1. Typical Transient Temperature Profiles Near Dryvell Wall. l l Run 22 101678 i l _ . - - - _ . _ . -_ , . . . . _ _ _ _ _ _
l l 4 22A4365 l N-5 Rev. 2 ! l O This figure is PROPRIETARY and is provided under seperate cover. O l i Figure N-2. Typical Transient Temperature Profiles Near Containment '4all, Run 22 O 101678
22A4365 N-6 Rev. 2 24 0 20 - - -- - N AEE SU AF ACE INITI AL POOL TEMPERATURE 100 F 16 - POOL DEPTH 20 f t 8 TOTAL ENERGY RELEASE 4=10 8tu FINAL BULK POOL TEMPE R ATURE 150 F O . . . . . TOP VENT CENTERLINE 12 i 1 a e 8 - T_ FINAL INITI A L 4 - i I J B ASEMAT l g 120 140 160 180 100 POOL TEVPER ATURE I F1 rigure N-3. Suppression Pool Temperature Profile for Large Breaks l 101678
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! 22A4365 I Rev. 2 O ] i i i i l 1 l 1 4 i ATTACINENT 0 I l i i t at DIGITIZATION OF FORCING FUNCTION FOR CONDENSATION OSCILLATION i j i i l . i i i. 4 O 101678
ATTACllMENT O Mark 111 Condensation oscillation Forcing Function for t - 3.0 to 30.0 Second P RE S 91RF TIuF PursstIDE TIME P RE SSunF TIME PRE SSURE TIME P RESSilflE rimf PRFSSHnE TIuf (PSIO (PSID) (SFCI (PSID) (SFC) (PSID) (SFC) (PSIO) (SEC) (PSIDI ( SEC l (SEC) 4.911 5.347 3.4331
- 1. m s) 0, 1.51 3 4. 1476 1.907 0.0000 4.462 -1.9792 0.
1.522 4.5511 4.007 0.1085 4.471 -3.8659 4.920 2.3142 5.355 3.80H9 3.010 2.H550 4.928 4.2160 5.344 3.8390 3.021 5.2012 3.5 12 4.5921 4.086 -0.0120 4.480 -4.0131 4.4H9 -4.5401 4.91T 5.4155 5.373 3.5608
- 1. 0 18 6. 6H I O 1.542 4.2596 4.026 -0.5001 2.9394 3.041 7. l H 69 3.552 3.5169 4.0 35 -1.1191 4.498 -5.2Flo 4.946 5.8256 5 . 388 4.044 4.507 -5.9415 4.955 5.5732 5.390 2.0782 3.052 6.8155 3.562 2.4865 -2.1701 4.9399 5.398 1.1740 1.062 6.0942 3.572 1.404F 4.054 -3.1525 4.517 -6.2168 4.941 4.063 4.526 -5.7791 4.9 72 4.2546 5.407 0.4394 3.012 5.2487 1.5HI 0.5246 -4.0405 3.7784 5.486 0.0111 3.OHI 4.6613 3.591 0.0815 4.071 -4.3774 4.535 -4.4990 4.981
-0.18 3H -4.3442 4.544 -2.4694 4.990 3.6226 5.424 -0.0951 3.091 4.4692 3.601 4.082 5.433 0.0000 1.104 4.5990 3.611 0.0000 4.092 -4.1443 4.553 0 4.998 3.7279 4.101 -4.0271 4.562 2.3846 5.007 3.C07 7 5.441 0.0951 1.114 4.0206 3.621 0.113H 5.450 -0.0181 1.824 4.H576 3.631 -0.01 h 4.118 -4.2005 4.571 4.3443 5.016 1.9375
-4.7290 4.580 5.5803 5.025 3.6524 5.459 -0. 4 3HS 3.1 71 4.5059 3.641 -0.5246 4.120 5.467 3.145 3.7203 3.650 -l.404H 4.1 30 -5.491H 4.589 6.0028 5.031 1.0156 -l.1741 3.155 2.6102 3.660 -2.4866 4.139 -6.1958 4.59H 5.7427 5.042 2.1120 5.476 -2.0782 1.4H59 3.670 -1.5170 4.149 -6.4764 4.607 5.0002 5.051 1.2044 5.4H4 -2.9394 3.166 5.OM) 0.449H 5.491 -3.5602 3.5 16 0.5549 3.6HO -4.2597 4.158 -6.0205 4.616 4.3840 0.014 1 3.690 -4.5921 4.161 -4.6H69 4.625 3.69 13 5 .0 68 0.0ll6 5.502 - 3.R 140
- 3. l H6 5.071 -0.0976 5.510 -3.8019 M tJ 3.191 -0.1204 1.700 -4.5571 4.177 -2.5725 4.614 3.7328
-4.1476 4,tRS 0. 4.64 3 3.8411 5.086 0.0000 5.519 -3.6316 @ $f 3.207 0.0000 3.709 4.0266 0.0976 5.527 - 3.5318 *e 1.211 0.1204 1.789 -4.2249 4.194 2.4496 4.652 5.095 4.205 4.4991 4.660 4.0573 5.103 -0.0116 5.516 -3.6829 w I$
1.22H -0.0141 3.729 -4.4065 5.545 -4.1471
- 3.2 3H -0.5550 3. 7 39 -4.9619 4.214 5.7792 4.669 3.7615 5.112 -0.4498
-l.4060 3.749 -5.7612 4.221 6.216H 4.678 3.1073 5.128 -1.2045 5.553 -4.8151 3.24H 5.1 30 -2.1121 5.562 -5.4324 3.259 -2.6104 3.759 -6.4901 4.232 5.9474 4. 68 7 2.1969 3.768 -6.7948 4.243 5.2716 4.696 1.2411 5 .1 38 -3,0157 5.570 -5.6784 1.269 -3.7203 0.4615 5.147 -3.6525 5.579 -5.2786 1.280 -4.5060 3.77H -6.1157 4.251 4.5402 4.705 4.260 4.0121 4.714 0.0110 5.156 -3.9375 5.588 -4.1093 3.2W) -4.H574 3.THH -4.9167 5.165 5.596 -2.2554 3.300 -4.H20H 1.798 -2.6981 4.269 3.8659 4.723 -0.1006 -3.0077
- 3. 00H 0. 4.2 7H 3.9 7H2 4.712 0.0000 5.173 -3.7279 5.605 O.
- 3. 318 -4.5990 5.61 3 2.2071 1.321 -4.4492 3.H I 1 2.5727 4.2H7 4.1701 4.741 0.l006 5.182 -3.6226 4.6H70 4.2019 4. 75 0 -0.0119 5.191 -1.7704 5.622 4.0209 3.338 -4.6613 1.H27 4.?94 5.2 00 -4.2546 5 . 6 10 5.1649 1.342 -5.2488 3.H 36 6.0205 4.306 3.H977 4.759 -0.4635 6.4164 4. 31 5 1.2001 4.768 -l.2412 5.208 -4.9399 5.619 5.5560 3.'52 -6.0943 3.046 5.287 -5.5732 5.647 5.3151 1.362 -6.8156 3.H 55 6.1958 4.324 2.2752 4.777 -2.1970 4.311 1.2851 4.786 -3.1074 5.226 -5.8256 5.656 4.7111 3.313 -7.1H69 3.H65 5.4911 4.0571 4.7209 4.142 0.4H00 4.795 - 3. 76 36 5 .2 35 -5.4155 5. 66 4 3.381 -6.6H09 3.H74 5.672 3.6015 J.391 -5.2010 3.HH 4 4.2005 4. 15 1 0.0823 4.004 -4.05 13 5.241 -4.2859 4.361 -0.1041 4.H12 -4.0266 5.252 -2.3140 5.681 3.4550 1.404 -2.8547 3.H91 4.0271 5.690 3.5554 O. 3.902 4.1443 4.170 0.0000 4.H28 - 1.8413 5.261 0 3.414 -3.7328 5.269 2.2657 5.698 3.7269 1.424 2.69H9 3.912 4.3442 4.370 0.1041 4.R10 4.1774 4.188 -0.0124 4. H 19 -3.8911 5.2 TH 4.1095 5.706 3.7551 1.434 4.9169 1.921 3.444 6. 3 t SH 1 . 9 11 4.0605 4.397 -0.4800 4.H48 -4. 1948 5.2HF 5.2167 5. il5 3. 4 H 14 4.H57 -5.0902 5.295 5. 6 7H 4 5. ?? 3 2.87^O 3.451 6.1941 1.940 3.1525 4.407 -1.2H54 5.732 2.0114 6.4991 2.1702 4.486 -2.2751 4.866 -5.7423 5.304 5.4321 H 3.461 1.950 4.875 5 . 312 4.8150 5.740 1.14Hi S 1.471 5.7618 3.959 1.1190 4.425 -1.2181 -4.002H 5.749 0.4290 o 4.9610 4.4 14 -1.8977 4.684 -5.5902 5.321 4.1410 os y
1.4H1 1.493 4.4065 1.969 1.9 /8 0.5000 0.0129 44 1 -4.2019 4.891 -4. 1441 5 . 3 30 1.6829 5.757 0.0110 $ 4.902 -2.3844 5. 3 3H 1.5 li s 5.766 -0.0931 1.501 4.2249 3.9HH -0.1005 4.452 -4.1701
Tl*tF PRFSsHRE TIME PHFSSURE TIME Pin:SSHOE PUF SSilRF T I ul' Put SSWF TIur ear ssurni fluF ( SI CI (PSIDI (SFel (PSlin (SIC) (PS101 tsFCI (PSlin (MC) (PSitu (SI C) IPSins 7.425 0NW l.H28 -1.10M5 6.19S 4.610 0 7.019 1.1519
- 5. / F4 0.0000 -1.4905 3.58 99 7.4 i l 0.06/2 7.637 -1.2151 S.7H1 0.0911 6.203 - 1. 1989 6.616 2.1064 7.027 7.645 -1.3533 1.8115 7.039 1.5167 7.441 -0.0804 S.7>l -0.0818 6.211 -1.5IFH 6.626 7. H5 3 - 3.7160
-0.4290 6.220 -1.9811 6.634 4.9294 7.044 3.28 99 7.450 -0.40lH S.Ha) 7.052 2.7163 7.456 -l.0759 F.Hol -4.1842
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-5.3153 6. 120 5. 41H1 6.711 1.9406 7.141 -1.9205 7.547 5.908 6.741 1.0961 7.449 -2.7164 7.555 -S.20 lH F.957 4.3632 5.980 -5.5560 6.12H 4.5544 7.965 3.75 79 S.918 -5.1649 6. 1 14 3.9225 6.749 0.4094 7.I58 -3.2900 7.561 -4.8174 6.757 0.0805 7.166 -3.5467 7.5 71 -3.7659 7.973 3. 33 F 1 S.921 -4.0208 6. 145 3.4815 7.9H1 3.1997 5.9 h -2.2049 o.35 1 3.3399 6.765 -0.01186 7.174 -3.5899 7.5 79 -2.0670 3.4170 6.774 0.0000 7.162 - 3. 15 19 7.561 0. 7.989 3.2921 5.944 0 6.361 2.0539 7.997 3.4515 3.6021 6.782 0.0HHH 7.100 - 3.2611 7.595 5.952 S.96) 2.1668 3.9475 6.169 6.17H 3.6 302 6.790 -0.0106 7.19H - 3.40 14 7.601 3.74I7 H.005 3.4719 EU<>
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$TIMF PRESSHQE TIME ( SEC) (P5 tot I Si C) jPSID) (SFC) (Pslin (SFCI (PSIH) (SFCI (PSIO) (SEC) (PSIDI 0.0000 9.422 - 3. 2 T 4 B 9.818 0 10.213 3.2972 8.229 0. d.628 1.2757 9.026 10.228 3.4562 2.0374 8 . 6 16 3.4117 9.0 14 0.0H57 9 .4 10 -3.1824 9.826 2.0405 H.231 9.8 34 3.7174 10.229 3.4826 0.245 3.7117 8.644. 3.4599 0.042 -0.0102 9.418 -1.3102 10.237 3.2305 4.7677 8.652 3.2094 9.049 -0.1950 9.446 -3.7176 9.842 4.7758
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H.271 4.3490 a.676 1.0583 9.0 11 -2.6488 9.470 -5.1876 0.1952 0.081 9.478 - 4. 75 14 9.873 3.7584 10.268 0.3978 8.215 '3.1456' h.684 -3.2073 9.888 3.3315 10.276 0.0102 8.29 1 3.3264 8.692 0.0102 0.089 -3.4576 9.485 - 3.7016
-0.085H 9.097 -3.4314 9.493 -2.0128 9.889 3.1942 10.284 -0.0863
- 8. 301 3.id93 8.700 10.292 0.0000 3.2H2O 0.0000 9.805 -3.2735 9.508 O. 9.897 3.2870 H.3UJ 8.108 1.4456 10.100 0.0863 H. 31 1 3,4403 8.716 0. 065tl 9.11 1 -3.1811 9.509 2.0158 9.905
-0.0102 9.128 - 3.3178 9.517 3.7089 9.983 3.4719 10.308 -0.0103 H.325 3.4665 8.724 3.2205 10.186 -0.3979 H. 313 3.2155 8.132 -0.3951 9.120 -3.7360 9.525 4.7641 9.921 9.5 17 -4.3378 9.513 5.1249 9.928 2.6589 10.324 -1.0654 H.34l 2.6549 8.740 -1.0564 1.8709 10.332 -1.8859 H.J49 1.8770 H.748 -0.H735 0.145 -4.H919 9.541 4.9028 9 .9 36 9.151 -5.1155 9.549 4.3457 9.944 1.0620 10.319 -2.6673 H.351 1.0604 8.755 -2.6499 9.952 0.3966 10.347 -3.2305 6.365 0.3940 H.763 -3.2004 0.161 -4.7554 9.557 3.7428 8.771 -3.4599 -3.7020 9.565 3.3219 9.960 0.0102 10.355 -3.4826 H.313 0.0802 9.169 9.968 -0.0860 10.363 -3.4562 H. 3H1 -0.0059 8.779 -3.4331 9.174 -2.0119 9.573 3.1849 H.7H7 -3.2751 0.184 O. 9.580 3.2795 9.976 0.0000 10.371 -3.2972 H.3HO 0.0000 10.379 -3.2041 m sa
- 8. 39 7 0.0H59 8.795 -3.18 12 9.192 2.0330 9.5H8 3.4377 9.984 0.0860 m bJ H.405 -0.0102 8.803 -3.3208 9.200 3.7017 9.596 3.4619 9.992 -0.0102 10.387 -3.3489 8.813 -3.7386 9.208 4.7574 9.604 3. 21 ll 10.000 -0.3966 10.395 -3.7611 .# $
tl. 4 8 I -0.3960 10.403 -4.3692 w 8.421 -1.0604 8.H10 -4.1407 9.236 5.1176 9.612 2.6528 10.003 -l.0621 9.620 1.8756 10.015 -1.8H00 10.411 -4.9294 0 H.429 -l.8 711 8.H27 -4.8972 9.224 4.8959 8.835 9.212 4.3306 9.628 1.0596 10.023 -2.6590 10.418 -5.1526 H.43T -2.6549 -5.1100 10.426 -4.7898 H.445 -3.2856 8.H43 -4.7586 9.240 3.7375 9. 6 36 0.3957 10.011 -3.2205 H.451 -3.4665 8.fl5 8 -3 7045 9.248 3.1192 9.644 0.0102 10.039 -3.4719 10.414 -3.7288 H.H59 -2.0333 9.256 3.1824 9.652 -0.0659 10.047 -3.4455 10.442 -2.0467 0.461 -3.4402 10.055 -3.2810 H 460 -3.2819 8.H61 0 9.264 3.2749 9. 660 0.0000 10.450 0. 8.8 15 2.0321 9.2 72 3.4 32tl 9.667 0.0359 10.063 -3.1942 10.45H 2.0547 11.4 17 -3. I t19 3 10.466 3.7412 H.4HS -3.3264 8. 8tt 3 3.7021 9.279 3.4590 9.675 -0.0102 10.078 -1.3115 8.HVI 4.7554 9.287 3.2r'H6 0.683 -0.3957 10.070 -1.7515 10.474 4.80H2 tl. 491 -3.7457 10.067 -4.1557 10.4H2 5.1722 fl .501 -4.3490 8.899 5.1155 9.295 2.6491 9.691 -l.0596 8.907 4.H919 9.10 3 9.699 -l.8757 10.004 -4.9141 10.490 4.9481 8.509 -4.9066 I.8729 10.102 -5.1366 10.491 4.3858 H.517 -5.12 tl i 8.915 4.3171 9.111 1.0581 9.707 -2.6529 3.7160 9.389 0.3951 9 . 71 5 - 3. 21 31 10.110 -4.7750 10.505 3.7774 ti.525 -4.767F 8.922 10.118 10.513 3.354i 11.5 31 -3.7136 8.9 30 3.3178 9. 127 0.0102 9.723 -3.4639 -3.7173
- 9. 3 35 -0.0857 9.731 - 3.4 176 10.826 -2.0403 10.528 3.2861 H.541 -2.0372 8.938 3.18ll 10.529 3.3098 8.549 0, 8.046 3.2735 9.141 0.0000 9.739 -3.2795 10.814 0.
H.954 1.4114 0.0857 9.747 - 3.1869 10.l42 2.0468 10.537 3.4694 H.551 2.0135 9.35 3.7290 10.545 3.4959 U.565 3.7046 H.962 3.4576 9.359 -0.0102 9.755 - 3. 32 39 10.850 3.2072 0 . 16 7 -0.3952 0.762 - 3. 7 4 2ft 10.l58 4.7899 10.553 3.2428 H.512 4.7567 8.970 5.1526 10.561 2.6774 it.5HO 5.1190 8.9 18 2.6400 9.175 -1.05HI 9.770 -4.3457 10.166 1.fil22 9.382 -1.8730 9. 778 -4.9029 10.174 4.9293 10.569 1.8929 H.5H*l 4.H972 ft.9H6 4.3692 10.576 1.0694 fl.596 4.3407 8.994 1.0516 9.190 -2.6492 9.786 -5.8249 10.131 0.3950 -1.20H6 9.794 -4.7641 10.189 3.7630 10.584 0.1991 o
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puissunE TIME PRE SSORE T*uE P0FSSURE T l *4E PRESSH9E flME P4F S914E TiuF P4L ssuitt flur (PSinn (SEC) (PSIO ( SI C ) (PSID (SECS (PSI 0l (SICS ( P SI O) (SEC) ( SE C ) (PSIDI 0.0000 12.586 -3.4249 I I .(N) 3 -1.1245 ll.199 0. 18.794 3. 1801 12.891
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-1.0694 11.0 35 -4.4055 18.4 10 5.2505 11.626 10.640 5.0210 72.230 -l.9456 12.628 -5.1201 10.646 -l.8910 II.043 -4.9101 11.438 II.814 f .O lli 12.o lo -5.3528
-2.6714 11.051 -5.1951 18.446 4.4522 II.H42 1.0921 12.211 -2.7519 10.454 12.244 -3.3330 12.644 -4.9758
-1.2428 11.059 -4.8296 11.454 3.H146 II.450 0. 40 TH 10.661 11.462 1.4054 11.H58 0.0105 12.254 -3.5931 12.652 -3.8712 10.671 -3.4959 11.066 -3.7598 12.262 -1.5659 12.660 -2.1259 11.074 -2.0636 11.470 1.2650 ll.noo -0.0615 10.679 -3.4694 18.874 0.0000 12.270 - 3. 40l H 12.667 0.
10.6HT -3. 3 09tl ll.0H2 0 II.478 3.1599 12.675 2.1482 10.695 -3.2163 18.090 2.0742 II.4H4 1.5219 ti.862 0.0315 12.278 -1.3058 3.7183 11.494 3.5 4 Hft 11.889 -0.0105 12.286 -3.4479 12.683 3.0009 10.703 -3.3546 11.09H 5.0804 10.711 -3.7774 II.106 4.H540 11.501 3.2989 ti.H97 -0.4019 12.294 -1.8825 12.698 11.500 2.7179 11.905 -1.0921 32.302 -4.507H 12.699 5.3902 10.719 -4.3859 11.114 5.2215 -5.0857 12.707 5.3567
-4.94H2 11.122 4.9961 18.587 1.9286 II.913 -1.9332 12.310 '
10.727 12.118 12.715 4.570F 10.735 -5.1722 11.1 30 4 4276 11.525 1.0855 11.928 -2.7143 -5.3160 3.9346 60.742 1.#HI 14 11.531 0.4054 11.029 -3.3817 12.326 -4.9448 P2.723 l
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-3.54H8 12.016 -3.8226 12.413 3.4713 12. Hit 0.0107 l o. H29 3.3696 11.225 0.0104 l1.620 3.3282 12.619 -0.0003 Id.HJ1 3.2307 18.212 -0.0875 11.62H -3.5219 12.024 -2.0988 12.421 11.240 0.0000 11.616 -3.3599 12.012 0 12.429 3.4249 12.827 0.0000 lo.N45 3.3246 12.815 0.09c1 10.853 3.4849 11.248 0.0875 11.644 -3.2650 12.040 2.1918 12.437 3.5901 11.254 -3.4054 12.048 3.H4F3 12.445 3.6875 12.H43 -0.010F 10.861 3.5815 -0.0104 II.652 3.3556 12.H51 -0.4162 f o.0 69 3.2573 II.264 -0.4012 II.660 -3.8346 12.056 4.9488 12.453 II.272 -1.0796 11.668 -4.4523 12.064 5. 31 60 12.461 2.7705 12.859 -1.1945 I O. fl 71 2.6893 12.867 -l.9721 1.9014 I I . 2fl0 -1.9810 11.676 -5.0230 12.072 5.0857 12.469 1.95H8 lo.HHS 12.477 1.1065 12.875 -2.7901 10 H91 1.0741 II.2H3 -2.7029 11.681 -5.2505 12.080 4.5018 18.294 -3.2731 11.691 -4.8H09 12.088 3.&l24 12.485 0.48 32 12.883 -3.3794 10.900 0.40ll 12.491 0.0106 12.890 -3.6412 50.906 0.0103 18.104 -3.5202 11.699 -3.7997 12.096, 3.44 19 10.986 -0.0670 11.312 -1.5025 11.707 -2.0856 12.104 3. 3158 12.508 -0.0897 12.898 -3.6856 10.924 0.0 000 11.319 -1.3411 11.715 O. 12.111 3.4018 12.509 0.0000 12.906 -3.4491
- 88. 321 11.723 2.0981 12.110 3.5659 12.514 0.0807 12.914 -3.3589 10.932 0.0810 -3.2470 12.524 -0.0107 12.922 -3.4960 10.940 -0.0I03 11.315 -3.3466 11.711 3.822i 12.127 3.5911 11.719 4.9103 12.115 3.3129 12.512 -0.4133 12.930 -3.9166 10.944 -0.4012 II.343 - 3. H I 34 2.7588 12.540 -l.looo 12.918 -4.5707 10.956 -1.0742 I I . 351 -4.4277 11.747 5.2H21 12.141 11.755 5.0512 12.158 1.0455 12.543 -l.9583 12.944 -5.1567 10.964 -l.9014 11. 159 -4.9953 12.954 10.972 -2.6894 11.367 -5.2215 11.761 4.4790 12.159 1.0991 12.556 -2.7706 -5.3902 o
- 88. 771 1.H516 12.167 0.4104 12.544 -3.3556 12.942 -5.080F r 10.979 -3.2571 18.375 -4.85 N 12.970 -3.0008 L 10.9H1 11. 3H 3 -1.7787 11.779 3.4258 12.875 0.0806 12.572 -3.6175 os 10.995
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( v / g) TIuF PRFsstlRE TluF PRFSSilpF TIME PRE SSURE TIME PRE SSURE TIME PDEssuRE TIME PRESSURE (SEC) (PSID) ( SEc l (PSID) (SFCI (PSID) (SFCI (PSID) (SFCI (PSID) ( SEC ) (PSID)
- 12. 9 H6 0 83.135 3.5014 13.1A4 0.0000 14.187 - 3. 55 14 14.590 O. 14.995 3.6474 12.994 2.1571 11.391 3.6101 11.794 0.0924 14.195 -3.4569 14.599 2.2452 15.001 3.8231 11.002 3.929H 13.401 1.6981 11.002 -0.0110 14.201 -3.6056 14.601 4.0003 15.011 3.8525 41,080 5.0419 11,409 3.4105 11.HIO -0.425H 14.218 -4.0400 14.615 5.2540 15.020 3.5714 11.01H 5.4301 13.417 2.H124 I1.819 -l.1402 14.220 -4.7840 14.621 5.6518 15.028 2.9505 11.026 5.194H 13.425 2.0025 13,H26 -2.01 H 1 14.228 -5.3183 14.631 5.4069 15.016 2.0860 11.034 4.6045 13.411 1. l li l 1 1.H 34 -2.8547 14.236 -5.5592 14.639 4.7925 15.044 1.17H4 13.042 3.9457 I1.441 0.4?25 l 1.H 42 -3.4576 14.244 -5.16FH 14.647 4.1217 15.052 0.4408 11.050 3.5218 11.449 0.0109 11.H50 -3.7214 14.252 -4.0211 I4.655 3.6657 15.060 0.0113 13.058 3.3767 13.457 -0.0917 13.H58 -3.6991 14.260 -2.20H2 14.661 3.5146 15.060 -0.0955 11.064 3.4748 13.465 0.0000 13.H 64 -3.5289 14.26H 0 14.678 1.6867 15.076 0.0000 11.014 3.6424 13.411 0.0917 11.H 74 -3.4291 14.216 2.2265 14.679 3.7918 15.084 0.0955 13.0H2 3.6102 13.4HI -0.0809 13.882 -3.5767 14.2H4 4.0563 14.687 3.8201 15.093 -0.0111 13.000 3.4045 11.409 -0.4225 11.800 -4.0276 14.292 5.2804 14.696 1.5435 15.801 -0.4401 13.091 2.Hl0H 11.497 -1. l il l 13. tl98 -4.6763 14.300 5.6049 14.704 2.9256 15.109 -1.1185 11.104 l.9873 13.505 -2.0024 11.906 -5.2758 14.308 5.3521 14.712 2.0604 15.117 -2.OH61 13.114 8.1227 11.553 -2.H124 113.914 -5.5147 14.316 4.7527 14.720 1.1685 15.125 -2.0505 11.122 0.4192 13.521 -3.4305 11.922 -5.8265 14.124 4.00 14 14.728 0.4364 15.131 -3.5714 11.130 0.0108 13.529 - 1. 60 H 1 11.930 -3.9909 14. 112 3.6152 14.7 3% 0.0112 15.141 - 3. H5 25 11.135 -0.0910 13.537 -3.6101 11.918 -2.1905 14.340 3.4354 14.744 -0.0947 15.149 -3.8233 13.846 0.0000 11.545 -3.5014 13.944 0 14.148 3.5167 14.752 0.000G 15.158 -3.6414 kU 13.154 0.0910 13.5%3 -3.4025 11.954 2.2084 14.157 3.7597 14.760 0.0947 15.166 -3.5444 ,4 g 11.162 -0.0108 13.561 - 3.5 4HH 11.962 4.0212 14.165 3.7884 14.768 -0.0812 15.l74 - 3.6968 un 13.170 -0.4193 43.569 -3,0961 11.970 5.1679 14.373 3.5141 14.776 -0.4364 15.182 -4.162H N ((
13.1 1H -1.1221 11.577 - 4. 619H 13.973 5.5592 14.381 2.9083 14.785 -I.16H6 15.190 -4.8331 11.186 -1. 9r5 7 4 13.5H5 -5.2144 11.9n4 5.3181 14.189 2.0513 14.793 -2.06H5 15.198 -5.4529 , 13.194 -2.8109 13.593 ~ 5.4 716 13.994 4.71 19 14. J)7 1.1588 14.801 -2.9257 15.206 -5.6998 13.202 -3.4045 13.601 -5.0864 14.002 4.0400 14.405 0.4127 14.809 -3.5435 15.214 -5.29H5 13.210 -3.6702 13.609 -3.9597 14.040 3.6056 14.413 0.0111 14.All -3.8201 15.222 -4.1249 13.210 -3.6424 11.617 -2.8734 14.018 3.4569 14.421 - 0.09 39 14.825 -3.7911 15.231 -2.2640 13.222 -3.4148 13.625 O. 14.027 3.5574 14.429 0.0000 14.831 -3.6867 15.239 O. 11.233 -3.3761 11.633 2.1907 14.0 15 3.7290 14.417 0.09 39 14.841 -1.5146 15.247 2.2817 11.248 -3.5219 13.641 1.9910 14.041 3.7574 14.445 -0.0182 14.849 -3.6657 15.255 4,1605 1.1.249 -3.9650 13.649 5.l265 84.058 3.4H54 14.453 -0.412H I4.857 -4.8277 15.263 5.3442 11.251 -4.6046 11.657 5.5141 14.059 2.8777 I4.468 -1.1589 14.865 -4.7926 15.278 5.7488 1 1. 2 N> -5.1949 13.665 5.2758 14.067 2.0145 14.469 -2.0514 14.874 -5.4010 15.279 5.4907 13.273 -5.4101 13.671 4.6162 14.015 1.1493 14.477 -2.9014 14.882 -5.6518 15.288 4.8748 13.2H1 -5.0478 11.681 4.0275 14.001 0.4292 14.4H6 - 3.5148 14.890 -5.2539 15.296 4.19HS 11 ?H9 -3.9297 13.689 3.5767 14.091 0.0110 14.494 - 3. 7HH 4 14.896 -4.0901 15.304 3.72HS
-2.1569 13.697 3.4291 14.099 -0.0911 14.502 -1.7596 14.906 -2.2450 15.312 3.5749 11.291 15.320 3.6781
- 13. l(h 0. 11.706 1.5289 14.101 0.0000 14.510 -1.5H66 14.914 0 11.385 2.1736 13.784 1.6092 14.115 0.09 il 14.51H -3.4454 14.922 2.2642 15.128 3.R$62 1.9599 13.722 3. 12 14 14.12? -0.0111 14.526 - 3. o 15 3 14.930 4.1250 15.316 3.8856 I I. J21 3.6041 11.129 5.0865 11.710 3.4575 14.1 9 -0.4291 14.534 -4.0334 14.938 5.2906 15.345 1 1. l lH 2.8546 14. ' -1.I494 14.542 -4.7528 14.946 5.6998 15.353 2.975H 13.111 5.4116 5,4529 15. 168 2.80 N 13.145 5.2146 11.746 2.01 H 1 14.4 -2.0146 14.550 -5. 3528 14.955 4.6 N 1 11.154 1.1402 14 14.55H -5.6i)49 14.961 4 H 312 15.349 1.1686 f
$ -2. Hill 13.351 . m e-* 1 1. 16l 3.9961 13.162 0.425H I4.161 -J.4H54 84.566 -5.2101 14.971 4.1627 15.371 0.4439
-1.7514 14.574 -- 4 . m 62 14.979 3. 696H 15.1% 0.0884
$ 1 5. 169 1.54dH 11.110 0.0110 14.171 14.179 14.5H2 -2.2263 14.981 3.5444 15.391 -0.0963 m I l. 311 1.4025 13.718 -0.0924 -l.1/90
T!WE i%E ssupE TlWE PRE SSINE TIuF 90F ssfinF TIuF PREsiOHE (SEC) IluE PDF SSU OF TluE PlW 'XORE (SFC) (PSint (5FCI :PSIDI (PSIO) ('d C l IPSIDI ( SI C I (PSIDI 0.0000 17.462 -3.8711 (SEC) (PSIOl 16.220 0 36.612 3. 8W 4 17.046 15.402 0.0000 15.Hl0 -1.7806 17.054 0.1006 17.470 -3.7679 14.228 2.1448 16.640 3.9112 15.410 0.0961 15.Hl3 -1.tO59 4.2105 lo.64H 4. 02 16 17.062 -0.0820 17.479 -1.9/97 15.41H -0.0114 15.H26 - 1. 7 M N 16.236 16.656 3.7321 37.071 -0.4618 17.487 -4.4252 15.H 14 -4.2 19) 16.244 5.4H55 !7.495 -5.13H0 15.426 -0.4439 16.252 5.0000 16.665 1.08:5 17.079 -l.2418 15.434 - 1. I nd a 15,842 -4.9871 2.1787 17.087 -2.1988 17.504 -5.796T
-5.5475 16.261 5.6452 16.673 -3.1089 17.512 - 6.0591 35.442 -2.1040 15.H51 5.0011 16.4H8 l.2108 17.095 15.450 -2.9759 15.H59 -5.7987 16.269 17.104 -3.7655 17.520 -5.4326 16.277 4.3095 16.689 0.4596 17.529 -4.3849 15.459 -3.604) 15.H67 -5.1904
- 16. 2 R5 3.8272 16.698 0.0818 17.112 -4.059 1 15.467 -3.HH56 15.875 -4.1964 16.706 -0.0997 17.120 -4.028t. 17.537 -2.4068
-1.8562 85.881 -2.3033 16.204 3.6604 17.545 O.
15.475 16.302 3.7760 16.784 0.0000 17.129 -3.8432 15.4H1 -3.6787 15.89: 0, 0.0397 17.117 -3.7347 17.554 2.4284 15.900 2.3231 16.380 3.95H2 16.123 17.562 4.4240 15.498 -3.5749 16.738 -0.0818 17.145 -3.8953
-3.7286 15.908 4.2 131 16.118 3. 9 NH 4
).154 -4.3862 17.570 5.6827 15.499 5.4377 16.327 3.6996 16.739 -0.4597 17.579 6.1810 15.501 -4.1985 15.916 3.0545 16.747 -1.2109 17.142 -5.0927 15.516 -4.8748 15.924 5.H494 16.135 17.170 -5.7456 17.587 5.84H2 15.932 5.5960 16.143 2.8594 16.756 -2.1788 17.595 5. l u is 15.524 -5.4998 1.2200 16.764 -3.0946 17.179 -6.0058 15.532 -5.7488 15.948 4.9601 16.151 17.187 -5.5810 17.404 4.4645 15.949 4.2719 16.160 0.4556 16. 772 -3.7324 17.682 3.9648 15.540 -5.344l
- 14. 168 0.0817 16. 7H0 -4.0237 17.195 -4.3463 15.548 -4.1603 15.957 3.7938 17.203 -2.3856 17.421 3.8084
-2.2835 15.965 3.6174 16.376 -0.0988 14.789 - 3.99 31 17.629 3.9118 15.556 0.0000 16.797 -3.8094 17.212 0, 15.564 O. 15.973 3.7431 16.384 17.220 2.4070 iF.637 4,1005 gn 15.982 3.9231 16.102 0.0989 16.805 -3.7019 17.646 4.1318 8 N 15.57 ) 2.3015 16.813 - 3. H 610 17.224 4.3851 15.581 4.1966 15.590 3.9516 16.401 -0.0817 16.822 ~4.3477 17.231 5.6327 17.654 1.8326 4$
5.3905 15.998 3.6411 lo.40a -0.4557 6.0591 17.462 3.1644 '*
- 15. 5 H9 16.417 -l.2203 l o.H 30 -5.0480 47.245 " [S 15.597 5.7961 14.006 3.0279 -5.6951 17.253 5.79ei 17.671 2.2372 16.014 2.1407 16.425 -2.1597 16.838 15.605 5.5474 lo.847 -5.95 10 17.262 5.1379 17.679 1.2639 15.oli 4.9171 14.023 1.2094 16.434 -3.0546 17.270 4.425l 17.688 0.4720 16.031 0.4516 14.442 -3.6997 16.855 -5.5 319 0.0121 15.622 4.2349 16.863 -4. 1393 17.278 3.9298 17.696 15.630 3.7600 16.039 0.0116 16.450 -3.9884 17.287 3.7679 17.704 -0.1024 16.047 -0.09 a0 14.458 -3.9582 16,878 -2.3646 0.0000 15.6 }l 3.6059 16.880 O. 17.295 3.8773 17.711 3.7107 16.055 0.0000 16.467 -3.7760 17.721 0.1024 15.646 0.0980 16.475 -3.6694 16.8H8 2. 3358 17.30) 4.0644 15.654 3.H896 16.064 14.896 4.1464 87.312 4.0954 17.729 -0.0122 15.66) 3.9193 16.072 -0.0186 16.481 -3.8272 17.320 3.7989 17.738 -0.4720 16.0HO -0.4517 16.491 -4.3096 16.005 5.58 11 15.671 3.6156 -5,00 38 16,913 6.0058 17.128 3.8365 17.744 -1.2619 15.679 3.0016 16.0H8 -1.2094 16.500 17.331 2.2175 17.755 -2.2371 16.096 -2.1408 16.508 -5.6452 16.921 5.7456 15.687 2.1222 16.929 5.0727 17.345 1.2527 17.763 -3.1444 15.695 1.1989 16.105 -1.0280 16.584 -5.0009 0.4678 17.778 -3.8327 16.524 -5.4H54 16,938 4.3962 17.351 15.701 0.4477 16.113 -3.6674 3.8952 17.362 0.0120 17.780 -4.1389 0.0115 16.128 -3.9516 14.532 -4.2703 16.946 -4.1005 15.il2 14.954 3.7 347 17.370 -0.1015 17.788 15.720 -0.0971 16.129 -1.9236 16.548 -2.3439 0.0000 17.796 -3.9118
-3.7411 14.549 O. 16.943 3.8412 17.378 15.728 0.0000 16.137 4.0286 17.387 0.1015 17.805 -3.8014 16.146 - 3.6 374 16.557 2.1648 16.978 -3.9641 15.736 0.0978 16.979 4.0593 17.395 -0.0121 17.883 15.744 -0.0115 16.154 - 3.7918 16.565 4.3082 -0.4679 17.822 -4.4645 16.574 5.5340 16.088 3.7654 17.403 15.752 -0.4478 16.162 -4.2720 J. lod 9 87.412 -l.2528 17.830 -5.1817 86.170 -4.9601 16.582 5.9530 16.996 17.838 -5.8482 15.768 -1.1990 16.590 5.6958 17.004 2.10 10 17.420 -2.2876 15.769 -2.1221 16.179 -5.5960 17.012 8.2417 17.428 -3.1166 17.847 -6.11 M 15.771 -3.0017 16.187 -5.8494 16.599 5.0479 -1.7919 17.855 -5.6826 c) 16.607 4.3476 17.028 0.4637 17.431
$ 15. 7H5 -3.6156 16.195 -5.4176 17.029 0.0119 17.445 -4.0954 17.861 -4.4239 i
-1.9891 16.203 -4.2131 16.635 3.8680 -4.0643 H 15.793 16.623 3.7019 17.037 -0.8006 17.454
$ 15.808 -3.H896 16.211 -2.3235 w
# 9 e
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[ \ v) TIME P RE SSURE TIME PRESSURE TIME PAESSURE FIME PRESSURE TIMF PDFSSURE TIME PRESSURE (SFCI (PSIDI (SFCI (PSID) (SEC) (PSIDI I SEC) IPSID) ( SEC) (PSini (SEcl (PSIDI 19.579 O. 20.010 4.1599 11.889 O. 18.301 3.9816 18.725 0.0000 19.850 -4.0524 18.731 0.1052 39.159 - 3.9 3HO 19.517 2.5608 20.019 4.360o l i. Htt9 2.4500 I H . 310 4.1737 4.6639 20.027 4.3939 4.4634 18.318 4.2055 13.742 -0.0125 19.167 -4.1073 19.596 II.H91 19.176 -4.6250 19.605 5.9909 20.016 4.0757 17.905 5.7312 18.327 3.9010 18.750 -0.4847 20.045 3.3658 1H.335 3.2208 18.758 -l.2979 19.184 -5.169C 19.613 6.4445 87.984 6.1673 19.622 6.1653 20.053 2.3798 11.922 5.0001 18.343 2.2778 18.767 -2.2914 19.193 -6. 0S H 4 1.1440 1.2864 18.775 -3.2494 19.202 -6.3127 19.630 5.4647 20.062 17.931 5.2297 18.352 19.679 4.7064 20.011 0.5089 17.939 4.5048 18.360 0.4804 18.784 -3.9156 19.210 -5.8669 0.0129 18.792 -4.2428 19.219 -4.5429 19.648 4.1798 20.079 17.947 4.0000 18.169 0.0124 4.0075 20.088 -0.8099 3.8151 IH.371 -0.8042 18.808 -4.2104 19.227 -2.5154 19.656 17.956 19.665 4.1239 20.097 0.0000 11.964 3.9464 I H. 3H 6 0.0000 18.809 -4.0169 19.236 0. 0.1089 0.1042 18.818 -3.9035 19.244 2.5178 19.674 4.3228 20.105 17.973 4.1169 I H. 39 4 4.6234 19.682 4.3558 20.114 -0.0129 17.988 4.36HS 10.403 -0.0824 18.826 -4.0711 19.253 18.835 -4.5845 19.26l 5.9188 19.691 4.0405 20.123 -0.5020 11.990 3.6667 IH 4ll -0.4805 19.699 3.3359 20.831 -4.3441 11.998 3.1925 18.449 -l.2865 18.841 -5.3229 19.270 6.3985 18.42H -2.2 772 18.852 -6.0051 19.279 6.1187 19.708 2.3585 20.140 -2.3792 16.006 2.2571 19.787 1.3124 20.149 -3. 1652 18.015 1.2751 13.436 -3.2209 18.860 -6.2712 19.287 5.4172
-3.9011 18.H69 -5.8353 10.296 4.6656 19.725 0.4976 20.157 -4.0758 16.023 0.4162 IH.445 19.734 0.012H 20.166 -4.3939 I H .0 32 0.0122 18.453 -4.2055 18.877 -4.5427 19.104 4.1414 18.462 -4.1736 18.H86 -2.4934 19.313 3.9726 19.742 -0.1080 20.175 -4.3606 I H.0 40 -0.1013 0.0000 20.183 -4.8599 18.048 0.0000 16.47( -3.9816 18.H94 O. 19.328 4.08HI 19,751 18.901 2.5157 19.330 4.2853 19.760 0.1080 20.192 -4.0425 ,g 16.051 0.1033 IH.474 -3.8692 19.768 20.201 -4.2141 mN 15.065 -0.0123 IH.4H7 -4.0356 18.911 4.5830 19.319 4.3180 -0.0128 4g 18.496 18.920 5.8870 10.347 4.0053 19.777 -0.4976 20.209 -4.7477 16.074 -0.4762 -4.5442
- 3. 3169 19.786 -1.3125 20.288 -5.5825 La IS.0H2 -1.2752 18.504 -t.2762 18.928 6.3327 10.156 IH.512 -5.9526 18.917 6.05H3 19.364 2.3180 19.794 -2.3506 20.227 -6.2891 N ((
18.090 -2.2572
~6.2221 18.946 5.3699 19.373 1.3208 19.801 -1.3360 20.235 -6.5007 16.099 -3.1926 18.521 -4.0405 20.244 -6.0438 18.529 -5. 7H
- l 18.954 4.6249 19.181 0.4932 19.811 1H.107 -3.8667 19.300 0.0827 10.820 -4.155H 20.253 -4.7045 IH.114 -4.8685 16.516 -4.5028 88.963 4.3073 20.268 -2.5822 16.544 -2.4FIS 18.971 1.9380 19.399 -0.1070 l9.829 -4.322H 1H.824 -4.1369 19.407 0.0000 19.8 17 -4.1239 20.210 0.
16.138 -1.9466 18.555 O. 18.980 4.0524 2.6048 2.4936 18.988 4.2419 19.416 0.1070 19.846 -4.0075 20.279 Ill .148 -3.8351 18.563 -4.179H 20.283 4.7455
!B.149 111.572 4.5429 18.997 4.2001 19.424 -0.0127 19.855
-4.0000 3.9704 19.433 -0.4933 19,863 -4.7066 20.296 6.0957 IH.15H -4,5042 la.5HO 5. H 354 19.005 20.305 6.5572 18.5H9 6.2212 19.014 3.2781 19.441 -1.3209 89.872 ~5.4647 lH. loo -5.2297 -2.3 131 19.880 -6.1653 20.314 6.2731
-5.9002 18.597 6.0053 19.022 2.3176 89.450 1 H .175 19.031 1.1093 19.459 -3. 3370 10.889 -6.4445 20.122 5.5602 16.1H1 -6.8473 18.606 5.1228 -5.9908 20.331 4.7899 18.198 -5.7332 38.414 4.5844 19.039 0.4HH9 19.467 -4.0054 10.898 19.048 0.0126 19.476 -4.3180 19.906 -4.6638 20.340 4.2529 16.200 -4,4632 18.623 4.0713 19.915 -2.5598 20.349 4.0776 18.208 -2.4498 I H . 6 31 3.9015 89.056 -0.1061 19.484 -4.2352 4.1961 lH.640 4.0169 19.065 0.0000 19.493 -4.0988 19.924 0, 20.357 16.287 O.
19.502 -1.9724 19.932 2.5824 20.346 4.39H4 18.225 2.4717 18.648 4.2106 69.074 0.1061 4.4320 4.24?8 19.082 -0.0126 10.510 -4.1435 19.941 4.7046 20.375 l *l. 214 4.5010 18.651 19.950 6.0432 20.383 4.1818 16.242 5.7641 IH.665 1.9156 19.091 -0.4890 19.519 -4.6657 19.527 -5.41 72 19.958 6.5001 20.392 3.3945 I II . 29) 6.2221 IH.414 1.2494 19.099 -I.1094 6.2198 20.40I 2.399H I tl. 259 5.9525 IH.6H2 2.?971 89.80H -2.3877 19.516 -4.1317 19.961 19.544 -6.3985 19.976 5.5124 20.410 1.3557 1H.261 5.2 161 18.691 1.297H 89.186 -1.27H2 O.5061 0.4441 19.125 -3.9704 89.551 -5.9 bl7 19.9H4 4.7476 20.41H c) s I H . 216 4.5448 IH.699
-4.6212 19.993 4.2162 20.427 0.01 3) e 16.284 4.0155 IH.iOH 0.0125 19.8 11 -4. 2 H01 19.562 -0.lO9H "
S 19.142 -4.241H 19.570 -2.5116 20.001 4.0425 20.436 os IH.293 3.H692 18.786 -0.1052 u OD
TlWF PRE SSURE TlHE P RE S"lR E TlWE PRE SSURE TlWE P RESSil4E l TIME PRESSURE TIME PRESSH4E (PSID) (SEC) (PSID) (50C1 s (PSID) ( SEC) (PSin) ( SEC) (PSID) (SEC) (PSIDI ( SFC ) 0.0000 22.662 -4.4144 20.444 0.0000 20.HMI -4.2321 28. 121 0 21.765 4. 34I4 22.212 20.453 0.1098 20.H90 -4.182H 21.310 2.6725 21.714 4.5508 22.228 0.4146 22.671 -4.2893 20.462 -0.0111 20.H99 -4.2804 28.310 4. H N17 21.7H3 4.5854 22.2D) -0.0134 22.3H0 -4.4746 20.478 -0.5043 20.90H -4.8101 21. 148 6.25 19 21.792 4. 25 16 22.239 -0.5283 22.689 -5. 0 3 H2 20.479 -1.3558 20.914 -5. 60H 1 21.357 6.7275 21.801 3.5119 22.248 -l.4146 22.698 -5.8497 20.4bd -2.3999 20.925 -o.3273 21.364 6.4160 21.H10 2.4829 22.256 -2.5039 22.707 -6.5907 20.491 -3.3944 20.934 -o.ol Di 28.375 5.7046 21.Hla I.4027 22.265 -3.5415 22.716 -o.8965 20.505 -4.1112 20.941 -6.1482 21.381 4.91 32 21.827 0.5238 22.274 -4.2H94 22.725 -o.4128 20.514 -4.4120 20.952 -4.7863 21.192 4.3633 28.8 36 0.0135 22.283 -4.6243 22.734 -4.9923 20.523 -4.3964 20.960 -2.6278 21.401 4.1834 21.a45 -0.11,37 22.292 -4.5891 22.743 -2.7402 20.5 38 -4.1960 20.969 0 21.410 4.1050 28.854 0.0000 22. 301 -4.3779 22.752 0 20.540 -4.0776 20.97H 2.6499 21.410 4.5124 21.863 0.1837 22.310 -4.2543 22.761 2.7618 20.549 -4.2529 20.9H 7 4. H2 76 21.428 4.5471 21.872 -0.0135 22.319 -4.4372 22.770 5.0311 20.558 -4.7889 20.995 6.2011 21.437 4.2179 21.881 -0.5239 22.328 -4.9965 22.719 6.4660 20.564 -5.5603 21.004 6.6706 21.445 3.4824 28,890 -1.4028 22.337 -5.8013 22.788 6.9556 20.572 -6.2711 28.01 3 6.3816 21.454 2.4621 28.899 -2.4830 22.344 -o.545l 22.797 6.6542 20.584 -o.5572 21.022 5.6564 21.463 1.3909 21.908 -3.5120 22.355 -o.8414 22.806 5. 89 H 8
- 20.592 -o.0956 21.'031 4.8787 21.472 0.5894 21.016 -4.25 16 22.364 -o.3598 22.885 5.0798 20.601 -4.745 3 21.040 4.3264 21.481 0.0134 21.925 -4.5856 22.373 -4.9510 22.824 4.5112 20.610 -2.6046 21.048 4.1481 21.490 -0.1127 21.934 -4.5508 22.382 -2.7175 22.814 4.3253 20.619 0 21.057 4.2686 21.499 0.0000 21.943 -4. 3414 22.391 0. 22.843 4.4510 20.627 2.6271 28.0 64 4.4745 21.507 0.1827 21.052 -4.2iGv 22.400 2.7404 22.852 4.6657 l x ,3 20.634 4.7865 21.075 4.5086 21.586 -0.0134 21.068 -4.4003 ?2.400 4.9925 22.868 4.7313 mu 20.645 6.1483 21.084 4.1822 21.525 -0.5195 28.970 -4.9549 22.418 6.4829 22.870 4.3609 4 $;
20.654 6.6138 21.092 3.45 3) 21.5 34 -1.1910 21.979 -5. 75 10 22.427 6.8985 22.879 3.6005 ca 20.662 6.3273 21.108 2.4413 21.541 -2.4622 21.988 -6.4905 22.4 35 6.5996 22.888 2.5456 '
- 8l 20.671 5.6083 21.110 1.3798 21.552 -3.4825 21.997 -o.7844 22.445 5.8494 22.897 1.4381 20.6H0 4.8302 21.189 0.5150 21.561 -4.2879 22.006 -o.3068 22.454 5.0381 22.906 0.5370 1 20.689 4.2H94 21.128 0.0832 21.570 -4.5471 22.015 -4.9098 22.463 4.4742 22.915 0.01 38 I 20.o97 4.1828 21.136 -0.1117 21.5 78 -4.5126 22.023 -2.6949 22.472 4. 898 22.924 -0.1865 20.706 4.2323 21.145 0.0000 21.587 -4.3050 22.032 0 22.481 4.4145 22.933 0.0000 20.715 4.4364 21.154 0.1387 21.594 -4.1834 22.041 2.7877 22.490 4.6274 22.942 0.1865 20.724 4.4703 21.161 -0.0133 21.605 -4.3633 22.050 4.9582 22.499 4.6627 22.952 -0.0138 20.733 4.1466 21.172 -0.5851 21.684 -4.9133 22.059 6.35 99 22.5 01 4.3251 22.961 -0.5375 20.741 3.4236 21.180 -l.3792 28.623 -5.7047 22.048 6.84I4 22.517 3.5710 22.970 -l.4382 20.750 2.4205 21.189 -2.4414 28.632 -6.4160 22.071 o.5450 22.526 2.5247 22.979 -2.5457 20.759 1.3674 21.198 -3.4531 28.640 -o.7274 22.086 5.8013 22.535 1.4263 22.988 -3.6006 20.766 0.5106 21.207 -4.1823 21.649 -o.2538 22.095 4.9964 22.54A 0.5326 22.997 -4.3609 20.776 0.0831 21.216 -4.5087 21.658 -4.8685 22.104 4.4372 22.553 0.0137 23.006 -4.7013 20.705 -0.1108 21.225 -4.4745 21.667 -2.6722 22.113 4.2543 22.562 -0.1156 23.015 -4.6657 20.794 0.0000 21.233 -4.2686 21.676 0 22.122 4. 3779 22.571 0.0000 23.024 -4.4510 g 20.803 0.1108 21.242 -4.1481 21.685 2.6951 23.131 4.589: 22.580 0.1156 23.033 -4.3253 20.811 -0.0132 21.251 -4.3264 21.694 4.9090 22.140 4.6241 22.589 -0.0137 23.042 -4.5113 20.H2O -0.5107 21.260 -4.H787 28.703 6.3069 22.349 4.2893 22.598 -0.5327 23.051 -5.0799 20.829 -1.3675 21.269 -5.6565 28.711 6.7844 22.858 3.5484 22.607 -I.4264 23.060 -5.8988 20.8 31 -2.4206 21.277 -o.3816 21.720 6.4905 22.167 2. 50 3C 22.617 -2.5248 23.070 -o.6543 20.846 -3.4237 21.286 -o.6706 21.729 5.7529 22.176 1.4145 22.626 -3.5710 23.079 -6.9556 20.H55 -4.1467 21.295 -o.20I0 21.738 4.9548 22.185 0.5282 22.635 -4.3252 23.088 -6.4650 g 20.864, -4.4703 21.304 -4.8274 28.747 4.4002 22.194 0.0116 22.644 -4.6427 23.097 -5.031s a Q
e 20.873 -4.4364 21.31 3 -2.6496 21.756 4.2889 22.203 -0.1146 22.653 -4.6274
- C
= 9 9 e
J O - O TIME PRESSURE TIME PRESSURE TIME PRESSURE TlWE PRE SSURE TIMF PRESSURE TiuF PRF SStJRE ISEC) (PSIDI (SEC1 (PSIDI IPSID (SFCI (PSID) (SEC) (PSID) (SFCI (PSID) ( SF( L 24.498 -4.5970 24.967 0. 25.440 4.7060 21.115 O. 23.572 4.5240 24.033 0. CXXX) 2.8989 25.449 4.9330 4.7422 24.043 0.l194 24.507 -4.4672 24.976 23.124 2.7858 21.582 24.985 5.2812 25.459 4.9707 21.133 5.0751 23.598 4.7784 24.052 -0.0842 24.517 -4.6592 25.468 4.6108 24.068 -0.5501 24.526 -5.2465 24.995 6.7838 23.142 6.5191 23.600 4.4 325 24.515 -6.0946 25.004 7.2974 25.478 3.8068 21.151 7.0127 21.609 3.6596 24.070 -1.4735 6.9813 25.487 2.6915 2.5874 24.080 -2.6063 24.545 -6.8725 25.014 21.161 6.7088 23.618 -7.1817 ?5.023 6.1879 25.497 1.5205 21.627 I.4617 24.089 -3.6892 24.554 25.506 0.5678 21.170 5.9465 24.563 -6.6780 25.033 5.3295 21.179 5.1215 21.637 0.5458 24.098 -4.4681 4.7330 25.516 0.0146 24.107 -4.8870 24.573 -5.1987 25.042 21.189 4.5483 21.646 0.0140 24.582 -2.8535 25.052 4.5379 25.525 -0.1232 23.197 4.3608 21.655 -0.1184 24.117 -4.7805 25.063 4.6697 25.535 0.0000 21.664 0.0000 24.126 -4.5605 24.591 0. 25.544 0.1232 21.206 4.4875 -4.4317 24.608 2.8763 25.070 4.8950 23.215 4.7040 21.673 0.1184 24.I35 5.2401 25.0 80 4.9123 25.554 -0.0146 4.7199 23.681 -0.0148 24.144 -4.6223 24.610 25.563 -0.5679 23.225 24.'689 6.7110 25.089 4.5752 23.234 4.3967 21.692 -0.5459 24.154 -5.2049 25.099 3.7775 25.573 -1.5206 23.708 24.163 -4.0411 24.629 7.2406 21.241 3.6301 -1.4618 24.638 6.9269 25.108 2.6707 25.582 -2.6916 23.252 2.5665 23.710 -2.5875 24.172 -6.8100 25.l18 1.5087 25.592 -3.8069 23.719 -3.6597 24.188 -7.1267 24.648 6.1398 23.261 1.4499 -6.6250 24.657 5.2880 25.127 0.5634 25.602 -4.6l08 23.270 0.5414 23.729 -4.4325 24.191 4.6968 25.137 0.0145 25.611 -4.9707 23.279 0.0139 23.738 -4.7784 24.200 -5.1575 24.666 25.146 -0.1222 25.621 -4.9330 21.747 -4.7422 24.209 -2.8308 24.676 4.5026 2 3. 2e> -0.1175 24.685 4.6134 25.156 0.0000 25.630 -g.7060 21.298 0.0000 23.756 -4.5240 24.248 24.228 0 2.8537 24.694 4.8569 25.165 0.1222 25.640 -4.5732 E" 23.307 0.1875 21.765 -4.3963 5.1989 24.704 4.89 39 25.174 -0.0145 25.649 -4.7698 4g 23.316 -0.0140 23.775 -4.5853 24.237 -0.5635 25.659 -5.3710 us 24.246 6.6781 24.713 4.5396 25.184 23.325 -0.5415 2.1. 18 4 -5.1632 25.193 -l.5088 25.668 -6.2361 N 8; 23.793 -5.9949 24.256 7.1837 24.723 3.748l -7.0356 23.334 -l.4500 6.8725 24.732 2.6499 25.203 -2.6700 25.678 23.343 -2.5666 23.802 - 6. 16 35 24.265 25.212 -3.7776 25.687 ~7.354l 21.818 -7.0697 24.274 6.0985 24.741 1.4970 23.352 -3.63)1 5.2464 24.751 0.5590 25.222 -4.5753 25.697 -6.8364
- 23. 362 -4.3967 21.821 ~6.5720 24.284 -4.9323 25.706 -5.3221 23.830 -5.1162 24.291 4.6592 24.760 0.0144 25 .2 38 23.371 -4.7 399 24.302 4.4672 24.769 -0.1213 25.241 -4.8950 25.716 -2.9212 23.380 -4.7039 23.819 -2.8082 0.0000 25.250 -4.6697 25.725 0,
-4.4815 23.848 0 24.382 4.5970 24.779 21.389 4.8887 24.788 0.1213 25,259 -4.5379 25.735 2.9439 21.398 -4.3608 23.857 2.8 311 24.121 25.744 5.3632 23.867 5.1577 24.310 4.8555 24.798 -0.0144 25.269 -4.7330 21.407 -4.5483 4.5039 24.807 -0.5591 '25.278 -5.3295 25.754 6.8891 23.416 -5.1216 23.876 6.6251 24.340 25.764 7.4107 7.8248 24.349 3.7186 24.816 -l.4978 25.288 -6.1880 21.426 -5.9465 23.885 -2.65 00 25.297 -6.9813 25.773 7.0897
-6.7089 23.894 6.8880 24.158 2.6293 24.826 21.473 1.4852 24.835 -3.7482 25. 307 -7.2974 25.783 C.2840 23.444 -7.0127 21.904 6.0432 24.368 25.792 5.4122 23.913 5.2048 24.377 0.5546 24.845 -4.5397 25.316 -6.7837 23.453 -6.5190 24.854 -4.89 39 25.324 -5.2810 25.802 4.8064 23.462 -5.0749 23.922 4.6223 24.186 0.0341 25.812 4.6083 4.4317 24.195 -0.1201 24.863 -4.8569 25.335 -2.8986 23.478 -2.7855 23.931 25.345 O. 25.828 4.7422 23.480 0 21.941 4.5605 24.405 0.0000 24.873 -4.6334 25.831 4.9780 24.414 0.3203 24.882 -4.5026 25.354 2.9284 23.490 2.8084 e 3.950 4.7805 25.364 5.3223 -
25.840 5.0089 21.959 4.8170 24.421 -0.0141 24.891 -4.6962 23.499 5.8864 24.001 -5.2a88 25.371 6.8365 25.850 4.6463 21.508 6.5721 2).969 4.4682 24.431 -0.5547 25.381 7.3541 25.859 3.836l 23.978 1.6891 24.442 -1.4853 24.910 -6.1398 23.517 7.0697 24.920 -6.9270 25.392 7.0155 25.869 2.7822 23.524 6.7634 23.987 2.6082 24.451 -2.6292 6.2360 25.879 1.5322 1.4714 24.468 -3.7187 24.929 -7.2406 25.402 r c) 23.516 5.9948 21.996 24.938 -6.7109 25.411 5.3700 25.888 0.5722 e g o 23.545 S . I N12 24.006 0.5502 24.470 -4.5040 4.7607 25.898 0.0147
")
24.479 -4.8555 24.948 -5.2399 25.421
$ 21.554 4.5351 24.015 0.0142 25.430 4.5732 25.907 -0.1241 4.3961 24.024 -0.1194 24.489 -4.8137 24.957 -2.8761 N 23.561 m
TIWF PHFSSURE TIWF PRESSURE TIME PRESSURE TIME PRE SSURE IINE PHE SSilRE TiuF PRE SSilR E (PSID) (SEC) (PSIDI (SEC) (PSID) (SECl (PSID) (SEC) (PSID) ( SEC) (PSIDI (SLC) 27.872 0.0000 28.373 -4.9575 0.0000 26.399 -4.7784 26.884 0 27.376 4.3462 2S.vi7 25.927 0.1241 26.408 -4.6435 26.894 1.0110 27.186 5.8210 27.882 0.1248 28.383 J4.Hl75 26.418 -4.8431 26.004 5.4855 27.196 5.1610 27.892 -0.0153 28.393 -5.0247 25.9 36 -0.0147 27.902 -0.5939 28.403 -5.6560 25.946 -0.5722 26.428 -5.4515 24.914 7.0462 27.405 4.7813 26.437 -6.3120 26.924 7.5797 27.415 1.0526 27.912 -1.5903 28.413 -o.5693 25.955 -l.5323 2.7945 27.922 -2.8150 28.423 -7.4115 25.965 -2.7123 26.447 -7.1 37 26.934 7.2583 27.425 25.975 -1.8362 26.457 -7.4472 26.941 6.4271 27.435 1.5787 27.912 -3.9816
- 28.433 -7.7471 25.984 -4.6463 26.466 -6.9415 26.951 5.5156 27.445 0.5895 27.945 -4.8223 28.443 -7.2017 26.476 -5.4039 26.961 4.9160 27.455 0.0852 27.951 -5.1987- 28.453 -5.6065 25.994 -5.0089 27.941 -5.1593 28.463 -3.0773 26.003 -4.9710 24.486 -2.9661 26.973 4.7114 27.465 -0.1279 26.981 4.8504 27.475 0.0000 27.978 -4.9289 28.473 0 26.013 -4.7422 26.495 C. 28.483 3.0995 26.022 -4.6083 26.505 2.9887 26.992 5.0843 27.485 0.1279 27.988 -4.7829 5.4448 27.002 5.1211 27.494 -0.0152 27.991 -4.9886 28.493 5.6468 26.0 32 -4.8065 26.515 -5.6873 28.503 7.2533-26.042 -5.4123 26.524 6.9940 27.012 4.7522 27.504 -0.5896 28.001 7.5235 27.022 3.9236 27.514 -1.5788 28.011 -o.5222 28.514 7.8025 26.051 -6.2848 26.534 -7.1583 28.524 7.4645 26.061 -7.0897 26.544 7.1976 27.032 2.7740 27.524 -2.7946 28.021 6.3796 27.041 1.5671 27.534 -3.9527 28.038 -7.6915 28.5 34 6.olo2 26.070 -7.4107 26.554
-7.1500 28.544 5.6984 26.000 -6.8990 26.563 5.4946 27.051 0.5852 27.544 -4.7873 28.041 26.090 -5.3630 26.573 4.8796 27.061 0.0151 27.554 -5.1610 28.051 -5.5662 28.554 5.0606 26.583 4.6785 27.071 -0.1270 27.544 -5.1219 28.061 -3.0552 28.564 4.8520 26.091 26.109
-2.9436 0 26.592 4.8144 27.081 0.0000 27.573 -4.8842 ' 28.071 0. 28.574 4.9930 gy 26.118 2.9663 26.602 5.0466 27.090 0.1270 27.583 -4.7492 26.081 3.0775 28.584 5.2338 < k-26.128 5.4041 26.412 5.0852 27.800 -0.0151 27.593 -4.0524 28.091 5.6067 28.57i 5.2737 C 26.13s 6.9416 26.622 4.7170 27.810 -0.5853 27.603 -5.5766 28,108 7.2088 28.605 4.8989 N&
24.147 7.4672 26.631 3.8945 27.120 -1.5672 27.613 -6.4748 28. lit 7.7478 28.615 4.0389 24.157 7.1437 26.441 2.7514 27.130 -2.7741 27.623 -7.3049 28.121 7.4115 28.625 2.8556 26.167 6.3119 26.651 1.5555 27.140 -3.9237 27.633 -7.6357 28.131 6.5693 28.635 1.6832 26.176 5.4535 26.643 0.5809 27.149 -4.7522 27.643 - 1. 0a8 5 28.141 5.6579 28.645 0.6024 26.186 4.8431 2a.670 0.0149 27.159 -5.1231 27.653 -5.5258 28.151 5.0246 28.655 0.0855 26.196 4.6435 26.680 -0.1260 27.160 -5.0843 27.662 -3.0330 28.162 4.8175 28.666 -0.1307 26.205' 4.7784 26.690 0.0000 27.179 -4.8503 27.672 0. 28.172 4.9575 28.676 0.0000 26.215 5.0089 26.700 0.1260 27.189 -4.7134 27.682 3.0554 28.182 5.1966 28.686 0.1307 26.225 5.0471 26.709 -0.0150 27.198 -4.9168 27.692 5.5664 28.192 5.2363 28.696 -0.0155 26.234 4.6817 26.719 -0.5810 27.208 -5.5357 27.702 7.1501 28.202 4.8572 28.706 -0.6025 26.244 3.8653 26.729 -1.5556 27.218 -6.4274 27.782 7.6915 28.212 4.0102 28.716 -l.6833 ' 24.254 2.7328 26.738 -2.75 36 27.228 -7.2513 27.722 7.3S82 28.222 2.8353 28.726 -2.8557 26.261 1.5438 24.748 -3.8946 27.238 -7.5797 27.732 6.5228 28.2 32 1.6017 28.737 -4.0300 26.273 0.5765 26.758 v4.7170 27.247 -7.046l 27.742 5.6173 28.242 0.5988 20.747 -4.8920 26.283 0.0148 26.768 -5.0852 27.257 -5.4853 27.752 4.9885 28.252 0.0154 28.757 -5.2737! 26.292 -0.8251 26.777 -5.0466 27.267 -3.0108 27.762 4.7120 28.262 -0.8298 28.767 -5.23381
- 26. 302 0.0000 26.787 -4.8144 27.277 0 27.772 4.9289 28.272 0.0000 28.777 -4.9929; 26.312 0.1251 26.797 -4.6785 27.287 3.0333 27.782 5.1593 28.282 0.1298 28.787 -4.8520 26.328 -0.0149 26.807 -4.8796 27.297 5.5260 27.792 5.1987 28.292 -0.0154 28.797 -5.0606 26.331 -0.5766 26.816 -5.4947 27.307 7.0V82 27.802 4.8223 28.302 -0.5982 28.807 -5.6984 26.341 -1.5419 26.826 -o.3797 27.116 7.6157 27.812 3.9815 28.312 -1.6018 28.818 -o.6163 26.350 -2.7129 26.836 -7.1976 27.326 7.1049 27.822 2.8149 28.322 -2.8354 28,828 -7.4645 27.336 6.4748 27.832 1.5902 28.332 -4.0103 28.838 -7.8025 ?
g 26.160 -3.8654 26.846 -7.5235
-6.9939 27.146 5.5765 27.842 0.59 38 28.342 -4.8572 28.848 -7.2532 s r 26.370 -4.6817 26.855 O
- 26. 379 -5.0478 24.865 -5.4446 27.356 4.9521 27.852 0.0151 28.352 -5.2363 28.858 -5.6466 os 26.389 -5.0088 24.875 -2.9884 27.366 4.7482 27.862 -0.1268 28.363 -5.1966 28.868 -3.0993 4 e . -- -
e --
A O kJ TIME PRESSURE TIME PRESSURE TIME PRESSURE (SEC) (PSIDI ( SEC) (PSID) (SFCI (PSIDI 28.878 0. 29,370 5.0635 29.908 0.0000 28.889 3.1215 20.401 5.3077 29.018 0.1335 28.899 5.6867 29.481 5.3482 29.929 -0.0150 ' 28.909 7.3047 29.428 4.9610 29.9 39 -0.6152 28.919 7.8577 29.43? 4.0960 29.950 -1.6474 28.930 7.5171 29.442 2.8959 29.960 -2.9160 28.940 6.663L 29.452 a . 6 359 29.970 -4.8244 28.950 5.7387 29.463 0.6809 29.981 -4.9954 28.960 5.0964 29.473 0.0157 29.991 -5.3852 28.970 4.8463 20.483 -0.1326 28.981 5.0283 29.494 0.0000 28.991 5.2708 29.504 0.1326 29.001 5.3111 29.514 -0.0157 29.011 -4.9265 29.525 -0.6110 29.022 4.0675 29.5 35 -1.6368 29.032 2.8758 29.545 -2.8960 29.042 8.6246 29.555 -4.0961 29.052 0.6067 29.566 -4.9610 29.062 0.0156 29.576 -5.3482 29.073 -0.1316 ' 29.586 -5. 3077 g (2 29.083 0.0000 29.597 -5.0635 < > 29.093 0.1316 29.607 -4.9205 $ 29.103 -0.0156 29.617 -5.8321 ka os 29.114 -0.0002 29.ots -3.fluv 29.124 -146247 29.638 -6.7098 29.134 -2.8759 '29.648 -7.5700 29.144 -4.0674 29.659 -7.9127 29.154 -4.9266 29.669 -7.3557 29.165 -5.3811 29.679 -$.7263 29.175 -5.2708 29.690 -3.1430 29.185 -5.0283 29.700 0. 29.195 -4.8863 29.710 3.1651 29.206 -5.0964 29.72r 5.7661 29.236 -5.7388 29.731 '7.4067 29.226 -6.6631 29.741 7.9675 29.234 -7.5174 29.752 7.6223 29.246 -7.8577 29.762 6.7561 29.257 -7.3046 29.773 5.8188 29.267 -5.6865 29.783 5.1675 29.277 -3.1212 29.793 ^4.9546' 29.287 0 29.804 5.0985 29.298 3.1433 29.884 5.3444 29.308 5.7265 20.825 5.3852 29.318 7.3558 29.835 4.9953 29.329 7.9127 29.846 '4.1243 29.339 7.5699 29.856 '2.9159
>" 29.349 657097 29.866 ~l.d473 0 29.360 5.7788 29.877 0.6152
$ 29.370 5.1 320 29.887 0.0158 j) os 29.380 4.9205 29.898 -0.1335 e 6*}}