ML20039E892
| ML20039E892 | |
| Person / Time | |
|---|---|
| Site: | Black Fox |
| Issue date: | 09/30/1979 |
| From: | Reuter F, Stancavage P GENERAL ELECTRIC CO. |
| To: | |
| Shared Package | |
| ML20039E857 | List: |
| References | |
| 22A4365-01, 22A4365-1, 22A4365-R03, 22A4365-R3, NUDOCS 8201110590 | |
| Download: ML20039E892 (400) | |
Text
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00'.f;E TjED 22A4365 Rev. 3
A Class I V~ Septenber 1979
- 82 JM -7 N0 :02 0Wi UDC it ;g ;
a INTERIM CONTAINMENT LOADS REPORT (ICLR)
MARK III CONTAINMENT O Approved:
- Approved:
M bC W {' -
F. Reuter, Manager P. P. Stancavage, Manager Mark III Containmen', Design Containment Engineering i
Approved: .
w '
P. W. Ianni, Manager Containment Design e
NUCLE AR ENERGY ENGINEERING DIVISION
- GENERAL ELECTRIC COMPANY SAN JOSE, calif ORNIA 95125 l
GENER AL $ ELECTRIC 8201110590 920106 !
1 090779 I
{DRADOCK 05000556 PDR
]
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( 22A43.65 11 Rev. 2 l A DISCTAIbER OF RESPONSIBILITY This document is being made available by Geneml Electric Company without consideration in the interest of promoting the apread of technical knowledge. Neither General Electric Company nor the individual authors:
A. Make any varranty or representation, expressed or implied, l
with 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 owned rights; B. Assume any responsibility for liability or damge uhich may result from the use of any information disclosed in this document; or C. Imply that a plant designed in accordance uith the recomenda-tions found in this document vill be licensed by the United States Nuclear Regulatory Comission or that it vill comply uith Federal, State or local regulations.
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22A4315 Rev. 2 111 TABLE OF CONTENTS Pagg
- 1. INTRODUCTION 1-1 l 1.1 Confirmatory Testing 1-2 1.2 Definition of LOCA 1-2 1.3 Design Margins 1-3
- 2. EEVIEW OF PHENOMENA 2-1 i
j 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 l
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- 4. DRYWELL STRUCTURE 4-1
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4.1 Drywell Loads During a Large Break Accident 4-1 l
4.1.1 Sonic Wave 4-1 4.1.2 Drywell Pressure 4-1 4.1.3 Hydrostatic Pressure 4-2 i
4.1.4 Loads on the Drywell Wall During Pool Swell 4-2 4.1.5 condensation Oscillation Loads 4-3 l 4.1.6 Fall Back Loads 4-3 4.1.7 Negative Load During ECCS Flooding 4-4 4.1.3 Chugging 4-4 J
4.1.9 Loads Due To Chugging 4-5 4.1.9.1 Chugging Loads Applied to Top Vent 4-5a 4.1.9.2 Pool Boundary Chugging Loads 4-Sa 4.2 Drywell Loads During Intermediate Break Accident 4-5 4.3 Drywell During a Small Break Accident 4-6 4.3.1 Drywell Temperryre 4-6 4.3.2 Drywell Press, < , 4-7 4.3.3 Chugging 4-8 I 4.4 Safety Relief Valve Actuation 4-8 4.5 Drywell Environmental Envelope 4-8 4.6 Top Vent Temperature (Cycling) Profile During Chugging 4-8 l
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l 22A4365 Rg. 2 , iv _
TABLE OF CONTENTS (Continued)
Page i
5 WEIR WALL 5-1 5.1 Weir Wall Loads During a Design Basis Accident 5-1
( 5.1.1 Sonic Wave 5-1 5.1. 2 Outward Load During vent clearing 5-1 5.1.3 Outward Load Due to Vent Flow 5-1 5.1.4 Chugging Loads 5-1 5.1.5 Inward Load Due to Negative Drywell Pressure 5-3 5.1.6 Suppression Pool Fallback Loads 5-4 5.1.7 Hydrostatic Pressure 5-4 5.1. 8 Safety Relief Valve Loads 5-4 5.1.9 Condensation 5-4 5.2' Weir Wall Loads During An Intermediate Break Accident 5-4 5.3 Weir Wall Loads During a Small Break Accident 5-5 5.4 Weir Wall Environment Envelope 5-5
- 6. CONTAImfENT 6-1 6.1 Containment Loads During a Large Steam Line Break (DBA) 6-1 6.1.1 Compressive Wave Loading 6-1 6.1.2 Water Jet Loads 6-1 6.1.3 Initial Bubble Pressure 6-2 6.1.4 Hydrostatic Pressure 6-2 6.1.5 Local Containwnt Loads Resulting from the Structures at or Near the Pool Surface 6-3 6.1. 6 Containment Load Due to Pool Swell at the HCU Floor 6-3 6.1.7 Fall Back Loads 6-4 6.1.8 Post Pool Swell Wave 6-4 6.1.9 condensation Loads 6-3 6.1.10 Chugging 6-5 6.1.11 Long-Tem Transient 6-5 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 6-6 6.4 Safety Relief Valve Loads 6-7 6.5 Suppression Pool Ther=al Stratification 6-7 h 042178
.22A4365 i
- Rev. 2 v TABLE OF CONTENTS (Continued)
Page 7-1
- 7. SUPPRESSION POOL BASEMAT LOADS 8-1
- 8. LOADS ON STRUCTURES IN THE SUPPRESSION POOL 8.1 Design Basis Accident 8-1 8.1.1 Vent Clearing Jet Locd 8-1 8.1. 2 Drywell Bubble Pressure and Drag Loads Due to Pool Swell 8-1 8.1.3 Fall Back Loads 8-2 8-2 8.1.4 Condensation Loads 8.1.5 Chugging 8-2 8.1. 6 Compressive Wave Loading 8-2 8.1.7 Safety Relief Valve Actuation 8-3 9-1
- 9. LOADS ON STRUCIURES AT THE POOL SURFACE 10-1
- 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 STRUCIURES AT THE HCU FLOOR ELEVATION 12-1
- 12. LOADS ON SMALL STRUCTURES AT AND ABOVE THE HCU FLOOR ELEVATION R-1 REFERENCES i
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
22A4365 R v. 3 vi i TABLE OF CONTENTS (Continued) i Page ATTACIDENT E - DRYWELL NEGATIVE PRESSURE CALCULATIO!'S E-1 J
ATTACIDENT F - WETWELL ASYMMETRIC PRESSURES F-1 ATTAGDENT G - SUBMERCED STRUCTURE LOADS DUE TO LOCA AND SRV ACUATIONS G-1 f
11 - 1 l ATTAGUENT 11 - WEIR WALL LOADS DURING DRYWELL DEPRESSURIZATION
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I-l ATTACHMENT I - POOL SWELL VELOCITY ATTAClHENT J - SCALING ANALYSES AND SMALL STRUCTURE POOL SWELL DYNAMIC LOADS J-l ATTAGIMENT K - RESPONSE TO NRC QUESTIONS K-1 ATTAGUIENT L - CONTAINMENT ASYMMETRIC LOADS L-1 ATTAGIMENT M - MULTIPLE SAFETY / RELIEF VALVE ACTUATION FORCING FUNCTION MET 110DS M-1 ATTACIIMENT N - SUPPRESSION POOL TilERMAL STRATIFICATION N-1 ATTACinENT 0 - DIGITIZATION OF FORCING FUNCTION FOR CONDENSATION OSCILLATION 0-1 0
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22A4365 Rev. 2 vii 3
(G LIST OF ILLUSTRATIONS Title Page Figure Loss-of-Coolant Accident Chronology (DBA) 2-9 2.1 2-10 2.2.1 Schematic of the Mark III Pool Swell Phenomenon 2.2-2 Typical Suppression Pool Cross Section 238 Plant 2-11 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 f t 3 in.
2.3 Idealized Quencher Bubble Pressure Oscillation in 2-16 Suppression Pool 4.1 Drywell-Loading Chart for DBA 4-9 4.2 Drywell-Loading Chart for SBA 4-10 4.3 Drywell-Loading Chart for IBA 4-11 4.4 Short Term Drywell and Containment Pressure Response to a Large Steam Line Break (DBA) 4-12 73 4-13 (s,) 4.5 PSTF Test Results - Vent Static Pressure Differential 4.5a PSTF Test Results - Vent Static Pressure Differential 4-14 4.6 Typical Drywell Pressure Traces During Condensation, Run 23, Test 5807 4-15 4.6a Condensation Oscillation Load Spatial Distribution on Drywell Wall, Containment and Basemat 4-16 4.6b Condensation Oscillation Forcing Function on the Drywell Wall 0.D. Adjacent to the Top Vent 4-17 4.7 Typical Top Vent Pressure Trace During Chugging, Run 19 4-18 Peak Pressure Pulse Train in Top Vent During Chugging 4- 19 4.7a I 4.7b Peak Force Pulse Train in Top Vent During Chugging 4-19a Average Farce Pulse Train in Top Vent During Chugging 4-19b 4.7c Average Pressure Pulse Train in Top Vent During Chugging 4-19c 4.7d 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 Chugging 4-21 4.8b Suppression Pool Chugging Normalized Peak Underpressure Attenuation 4-21a 4.8c Suppression Pool Chugging Normalized Mean Underpressure O' and Post Chug Oscillation Attenuation 4-21b 101678 i
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22A4365 Rev. 3 v111 LIST OF ILLUSTRATIONS (Continued)
Figure Title P_aje, 4.8d Suppression Pool Chugging Normalized Spike Attenuation 4-21c 4.8e Suppression Pool Chugging Spike Duration "d" as a Function 4-21d of Location in the Poo,1 4.8f Suppression Pool Chugging Normalized Paak Post Chug 4-/'a Oscillations Circumferential Underpressure Amplitude Attenuation 4-21f 4.8g Circumferential Post Chug Oscillation Amplitude Attenuation 4-21g 4.8h 4.9 Drywell - Containment Pressure Differential During Chugging 4-22 4.10 - Calculated Maximum Drywell Atmosphere Bulk Temperature and Pressure Envelope 4-24 4.11 Dryweli 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 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 l 5.4a Typical Pressure Time-History for Weir Annulus During 5-10 llg 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 t
5.6 Nor=alized Weir Annulus Pressure Pulse Attenuation 5-11 5.7 Nominal Predicted Absolute Pressure Transient in Drywell Initiated by Vessel Reflood Line Break Level 238 Standard Plant 5-12 5.8 Vent Backflow Weir Annulus Water Surge Velocity Vs. Height Above Weir Wall 5-13 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 6-11 l Test Series 5706, Run 4 0
090779
22A4365 Rev. 3 ix LIST OF ILLUSTRATIONS (Continued)
Figure Title Page
- 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 Water Level Transients in Drywell and Suppression Pool i Following DBA 6-14 6.8 Drag Loads on Protruding Structures Due to Pool Swell 6-15 6.9 Containment Loading Due to Flow AP Across HCU Floor 6-16 6.10 Typical Containment Wall and Baseast Pressure Traces During Condensation, Run 23 (Ref. Test 5807) 6-17 6.11 Containment Wall and Basemat Pressure Time Histories, Test 6-18 5807, Run 11 _
6.12 Containment Wall Chugging Pressure Time History, Test 5707, 6-19 Run 9 6.13 Basemat Chugging Pressure Time History. Test 5707, Run 9 6-20 6.14 Calculated Maximum Containment Atmosphere Bulk Temperature and Pressure Envelope for Any Rupture 6-21 6.15 Long Term Containment Pressure Following a DBA 6-22 l
6.16 HCU Floor AP vs Vent Area 6-23 1 6.17 Suppression Pool Temperature Profile for the Example j Problem 6-24 7.1 Pool Boundary Loads During Bubble Formation 7-2 8.1 Structures Within Suppression Pool-Loading Chart for DBA 8-4 i 8.2 Deleted 8-5 8.3 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-Lo'ading 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 Gurface 10-6 10.4 Drag Load on Solid Structures within 18 f t of the Pool Surface 10-7 10.5 Drag Loads for Various Geometries (slug flow) 10-8 0
090779
22A4365 Rev. 2 x LIST OF ILLUSTRATIONS (Continued)
Figure Title Page 10-6 Summary of Pool Swell 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 During DBA 11-3 12.1 Small Structures at the HCU Floor Elevation - Loading Chart During DBA 12-3
.12.2 Loads at HCU Ficor Elevation Due to Pool Swell Froth Impact and Two-Fhase Flow 12-4 O
O 101678
i 22A4365 f
D*3' xi/xti i
l LIST OF TABLES Table Title g 1.1.1 Stammary of PSTF Tests 1-4 1.3.1 Stamary of Specified and Realistic Design Values 1-6 3.1.1 S m ry of Postulated Accidents Affecting Mark III Structures 3-1 4.1 Chugging Loads 4-8b 9-1 i
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O 090779
22A4365 '
Rev. 2 xiii/xiv ,
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\ ABSTRACT This technical report provides nwnerical information for thermt i hydraulic dynamic loading conditions in GE Mzrk III Reference Plant pressure suppression containment system during a loss-of-coolant accident, safety relief valve discharge and related dynanic events.
Information and guidance has been provided to assist the convain-ment designer in evaluating the design conditions for the various structures uhich font the containment system. Observed test data, or calculations upon uhich the loads are based, are discussea. A Class III supplement to this report (22A4365AB) includes additional proprietary information in' support of the load definiti.on.
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! 1. INTRODUCTION I
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i The NRC has forwarded to each Mark III project two sets of questions; one regarding LOCA dynamic loads and one regarding safety relief valve dynamic f loads. The information in this document represents the General Electric l input for those loads where GE may have expertise from pressure suppression
! and safety relief valve test programs. General Electric will use the design load values specified herein as the basis for the 238 GESSAR license applica-tion. Thus, General Electric is committed to the defense of the calculated j design loads and corroborating test data and analyses to the NRC. Other load values or smaller margins than those provided herein may be used if the archi-tact engineer is willing to defend them through the licensing process.
1 4
! 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 l
i effact of the above loads and load combinations on the structures and equip-ment, Thus, the architect engineer is responsible for generation of a project i unique document that answers the NRC concerns.
During a loss-of-coolant accident (LOCA) 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.
4 The loading information is based on either observed test data or conservatively 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 temporal distribution of the IDCA related loads, the bar charts identify other loading conditions such as seismic accelerations, dead-weight, etc. For each bar on the chart, reference is made to the section where specific discussion of the load is presented.
i 042178 l 1
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22A5365 R;v. 2 _ l-2 ,
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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 che 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 enese 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 evnll impact forces on a variety of targets representative of small structures found in the Mark III containment annulus, and aza discussed in Attachment J.
O This document relies on a large experimental test data base from the PSTF program. See Table 1.1.1 for a summary of tests. The scaling of the large scale and 1/3 scale PSTF precludes direct application to the prototype MK III.
Conservative interpretation of these tests results, employing dimensional similitude scaling relationship, is used to arrive at specified desig:t loads for MK III. (See Attachment J.)
1.2 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 basis accident (DBA). Other small line breaks result in loss-of-coolant accidents, and although their energy release does not result in large O
042178
22A4365 Rev. 3 g O dynamic loadings, their thermal effects may control the design of structures.
The intermediate break accident (IBA) and small break accident (SBA) fall into f this category. The size of the SBA is defined as that which will not cause i
i 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 l
I Table 1.3.1 summarizes the loads due to a LOCA for the containment structures.
Reasonable design margins are clearly shown by comparing the magnitude of the values between the conservatively specified design values and the realistic expected loads. The Mark III loads presented in this document should be f
interpreted as rigid wall loads. A similar case for showing the conservatism 4
l in the loads specified for relief valve actuation is given in Attachment A.
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i It is shown in this report that the MK III dynamic loading phenomena has been conservati ely bounded and the PSTF test data is conservatively interpreted.
Parameter simulation has justified the applicatioh of the test data to MK III designs with adequate design conservatism added. Any further margin considera-tions cannot be technically envisioned. In fact, where possible, the contain-ment designer may chose to justify more realistic design values.
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090779
Table 1.1.1
SUMMARY
OF PSTF TESTS Area Number Venturi Top VentQ Initial Number Pool / Re fer-Test of Range Submergence Pressure Blowdown of Vent Primary ence Type Vents Scaling Objectivesi* Report Series Blowdowns (inch) Range (feet) (psia)
Saturated 1. Vent Clearing 4 5701 21 2 1/8-3 5/8 2.0 - 15.5 1050 1 Full Steam 2. Full Scale Condensation Demo
- 3. Drywell Pressure Saturated 2 1. Vent Clearing 4 5702 17 2 1/8-3 5/8 1.93 - 11.97 1050 Full Steam Sat trated 3 1. Vent Clearing 4 .# e$
5703 3 21/2-35/8 6.77 - 11.05 1050 Full w$
Steam
- 1. Pool Swell 7 5705 4 1 - 4 1/4 6.0 - 8.0 1065 Air 2 Full Scoping 1065 2 Full 1. Pool Swell 7 5706 7 4 1/4 6.0 - 10.0 Air
- 2. Inpact Loading 7.5 1. Chugging 16 5707 22 2 1/8 - 3 1050 Air and 3 Full Steam 5801 19 2 1/8-3 5.0 - 10.0 1050 Saturated 3 1/3 1. 1/ 3 Scale 11 Steam Demonstration
- 2. Pool Swell
- 3. Roof Density and AP
{
i h 5802 3 21/8-3 6.0 1050 Saturated 3 1/3 1. Pool Swell 11
[
l Steam l
t e
l 9 9
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O O O Table 1.1.1 (Continued)
Area -
Number Venturi Top Vent 4, Initial Number Pool / Re fe r-Tect of Range Submergence Pressure Blowdown of Vent Primary ence Series Blowdowns (inch) Range (feet) (psia) Type Vents Scaling Objectives
- Report 5803 2 2 1/8-3 5.0 - 7.5 1050 Saturated 3 1/3 1. 1/3 Scale Demo 11 Liquid
- 2. Liquid Blewdown 5804 5 2 1/8-3 5.0 1050 Saturated 3 1/3 1. Roof Density 11 Steam Density and AP Repeatability 5805 52 1-3 5.0 - 10.0 1050 Saturated 3 1/3 1. Pool Swell 12 Steam Impact {g 5806 12 2 1/2-4 1/4 5.0 - 7.5 1065 Air 3 1/3 1. Pool Swell 13 w{
5807 20 1-3 7.5 1050 Saturated 3 1/3 1. Steam 15 Steam and Condensation Liquid 6002 14 2-1/8 - 3 5 - 10 1050 Steam 9 1/9 1. Pool Svell 17 l
Multivent Effects 6003 12 2-1/2 7.5 1050 Steam 9 1/9 1. Steam 18 Condensation Multivent Effects o
- In general tests are not direct prototype simulations, but parametric studies to be used in analytic g model evaluations.
U e Y.
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Table 1.3.1
SUMMARY
OF SPECIFIED AND REALTSTIC DESIGN VALUES I SP ecified Design Basis for Engr,g Load Design Estimate Analysis Test Section Comments STRUCTURE: Drywell BREAK SIZE: Large Drywell Pressurization 30 psig 18 psig Model 4.1.2 Peak calculated 21.8 psig (Ref. 1) plus margin Ilydrostatic 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
{h
- O Wetwell Pressurization 11 paid 3-5 paid Model in 5801, 5802 12.1 Test shows pressure 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 tc small flat 5802, 5805 structureu attached to D.W. (ser. 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 paid 5702, 5703 4.1.5 See Fig. 4.6.a for
{
U Oscillation Loads Fallback Velocity for (mean) 35 ft/sec 20 ft/sec Bounding 5801, 5807 4.1.6 pressure distribution 7
Drag Loads calculation O O O
O O O ,
Table 1.3.1 (Continued) pe Design Basis fled Load - Design Es t iniat e Analysis Test Section Comments Negative Load During -21 psid -15 psid Bounding 4.1.7 Assumes no vacuum relief ECCS Flooding calculation 1
Chugging Gross structure i2 psid il psid 5801,'5802 4.1.8 Design pressures are 5803, 5804 +30 psig and -21 paid Loading within top 4.1.9.1 vent e Pre-chug under- -15.0 paid -12 psid 5707 4.1.9.2 pressure (Peak) (peak)
-9.0 paid -8 psid (mean) (mean) gu e Pulse (spike) 540 psid 500 psid 5707 Local and global pulse
- O (peak) (peak) train specified W$
214 paid 180 psid
( an) (n.?an) 250 kips 250 kips 5707 Local and global net e Net force (peak) (peak) upward vertical load 91 kips 75 kips (mean) (mean)
Loading on drywell I.D. 5.1.4 Same as weir wall specification Loading on drywell 0.D. 5707 4.1.9.2 See Table 4.1 for dura-tion and frequency e Pre-chug under- -5.8 paid -4.0 psid pressure (peak) (peak)
~2.65 paid -1.0 psid l o (mean) (mean) g 8 e Pulse (spike) 100 psid 75 psid 4
U (peak) (peak) 24 psid 20 psid (mean) (mean)
Table 1.3.1 (Cantinued)
Specified Design Basis for Engr,g Load Design Estimate Analysis Test Section Comments
- Fost-chug 16.5 psid 4.0 psid oscillation (peak) (peak) ,
12.2 psid 11.1 psid (mean) (mean)
STRUCTURE: Drywell BREAK SIZE: Intermediate ADS 4.2 See Attachment A Chugging 4.1.8- Same as large break 4.1.9.2 specification STRUCTURE: Drywell BREAK SIZE: Small Temperature 330*F/310*F 330*F/ Bounding 4.3.1 3 hr at 330*F initially, 310*F calculation next 3 hr at 310*F E'S$
Chugging 4.1.8- Same as large break . $i 4.1.9.2 specification " 8l o
b; r.
g
%J 00 03 O O O
O O O Table 1.3.1 (Continued)
Specified Design Basis Ew ' g Design Estimate Analysis Test Section Comments Load STRUCTURE: Weir Wall BREAK SIZE: Large**
Outward Load Due to Vent 10 psig 5 psig Model - 5.1.2 First 30 see of blowdown Clearing Ref. 1 5.1.3 Chugging 5707 5.1.4 Local and global loading specified e Pre-chug under- -2.15 paid -2.0 paid pressure (peak) (peak)
-0.98 psid -0.5 paid (mean) (mean) e Peak spike of pulse 43 psid 35 paid
%u (peak) (paak) train 15 psid 13 paid *gu (mean) (mean) we 12,800 lbf/ 8000 lbg Bounding 5.1.5 Attachsent H Inward Load Due to Negative Drywell vent (top vent) calculation Pressure Differential 6000 lbf (mid) 4000 lbf (bottom) pH pH Standard 5.1.7 Hydrostatic Pressure analytical techniques STRUCTURE: Weir Wall BREAK SIZE: Intermediate **
Attachment A ADS STRUCTURE: Weir Wall BREAK SIZE: Small**
330*F/310*F Bounding 5.4 330*F for 3 hr initially Temperature 310*F for next 3 hr o calculation o
} ** Chugging Loads on Weir Wall are the same for large, intermediate and small break accidents. {
Table 1.3.1 (Continued)
P" "
p sign Basis Engr'g Load Design Estimate Analysis Test Section Comments l STRUCTURE: Containment BREAK SIZE: Large Water Jet <1 psig 0 psig Attachment G 5706 6.1.2 Heasured pressure is small and is obscured by bubble 5701, 5702 Pressure Bubble Formation 10 psid 8 psid 6.1.3 5703, 5705 5706 Ilydrostatic Pressure pil pil Standard 6.1.4 analytical techniques Pool Swell Loads for 10 psid 8 paid D.W. bubble 6.1.5 Only large structures Attached Structures (bubble) pressure see bubble pressure :o w at 1 Surface 40 ft/sec 30 ft/sec Bounding 6.1.5 See Attachment I .
(drag calculation w$
- velocity)
Pool Swell at ilCU Floor 15 psi (froth 10 psi 5706 6.1.6 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 Post Pool Swell Waves 2 ft 2ft PSTF Tests 6.1.8 Negligible load Condensation 11 psid i0.6 5807, 5701 6.1.9 See Figure 4.6a 5702 l
Oscillation Loads (mean)
Chugging 5707 4.1.9.2 See Table 4.1 for dura-tion and frequency 8 e Pre-chug-under- -1.3 psid -0.8 paid 8 pressure (peak) (peak) 7
% -1.0 psid (mean)
-0.3 paid (mean) l$
O O e - - - - - - - - - -
... - . . - . . . _ - --- - - . . - , - .- . _ - .= . . . .
O O O Table 1.3.1 (Continued)
SP ecified Design Basis Engr'g Design Estimate Analysis Test Section Comments Load e Pulse (spike) 3.0 paid 2.2 psid (peak) (peak) 0.7 psid 0.6 paid (mean) (mean) .
e Post-chug 1.7 psid tl.5 psid oscillation (peak) (peak) 11.0 psid 10.5 paid (mean) (mean)
Pressurization 15 psig 5 psig Model 6.1.11 Peak calculated value (Ref. 1) is 9.8 psig plus margin 180*F* <150*F Supplement 1 6.1.11 Conservative calculated EU Temperature .# $
to Ref. I peak temperature is 173*F t.a STRUCTURE: Containment BREAK SIZE: Intermediate "$
Pressurization 5 psi 2 pei Bounding 6.2 calculation 6.2 See Attachment A ADS 4.i.8- Same as large break Chugging 4.1.9.2 specification STRUCTURE: Containment BREAK SIZE: Small Temperature 220*F 185"F Bounding 6.3 Stratification (Dome) calculation Pressure 2 psig 1 psig Bounding 6.3 Typical value calculation
'4.1.8- Same as large break Chugging 4.1.9.2 specification w ~
S *See paragraph 6.1.11
Table 1.3.1 (Continued)
Spe ified Design Basis Load Design Estimate Analysis Test Section Comments STRUCTURE: Basemat BREAK SIZE: Large
~
Hydrostatic pil pil Standard analytical techniques Bubble Formation 10 + 21.8 psid 18 psid Peak equal to 5706/4 7.0 10 psi over 1/2 pool D.W. pressure assumed to increase linearly to 21.8 psi.
See Figs. 7.1 and 6.6 Condensation 11.7 psid 1.0 psid 5807, 5702 7.0 See Figure 4.6a Oscillation Load 5701 Chugging 5707 4.1.9.2 See Table 4.1 for dura- ,y tion and frequency Qy
- V e Pre-chug under. -1.8 paid -1.5 psid See Figures 4.8b wy pressure (peak) (peak) through 4.8f for l
-1.34 psid -0.7 psid basemat attenuation (mean) (mean) e Pulse (spike) 10 psid 7.5 paid (peak) (peak) 2.4 paid 2 psid (mean) (mean) e Post-chug 12.1 psid 2.0 psid oscillation (peak) (peak) 11.3 psid 11.0 paid (mean) (mean)
STRUCTURE: Basemat BREAK SIZE: Intermediate ADS 7.0 See Attachmant A Chugging 4.1.9.2 Same as large break o specification g g h STRUCTURE: Basemat BREAK SIZE: Small Chu ging 4.1.9.2 Same as large break specification
, O o n%e Table 1.3.1 (Continued)
Specified Design Basis Analysis Test Section Comments Load Design Estimate STRUCTURE: Submerged BREAK SIZE: Large*
Structures G2.1 Load is bounded by LOCA LOCA Water Jet Loads air bubble load Load is on a sample :
4 paid
- Attachmenc G G2.2 I
LOCA Air Bubble Load structure 4 ft from the top vent axis 30 ft/sec Bounding See Attachment I Velocity for Computing 40 ft/sec Drag Loads Calculation 20 ft/sec Bounding G2.5 Fall Back Velocity for 35 ft/sec Drag Loads Calculation C2.3 Frequency 2.+3.5 Hz - ww 1.4 paid Attachment G $@
LOCA Condensation Load is on a sample
- Oscillation Loads structure 4 ft from the di vent exit "$
G2.4 Load is on a sample 1.9 psid Attachment G LOCA Chugging Loads structure 4 ft from the top vent exit G3.1 Load is neglectible outside Neglectible Attachment G a sphere circumscribed by X-Quencher Water Jet Load the quencher arms G3.2 Load is on a sample 0.5 psid Attachment G X-Quencher Air Bubble structure 9 ft from Load Quencher center STRUCTURE: Submerged BREAK SIZE: Intermediate
- 1 Structures See Attachment G ADS STRUCTURE: Submerged BREAK SIZE: Small* r Structures _
S No Additional loads generated h j
- Chugging loads are the same for large, intermediate and small break accidents.
Table 1.3.1 (Continued)
Specified Design Basis for Engr,g Load Design Estimate Analysis Test Section Comments 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 E' k!
Drag Loads calculation .*$1 w
STRUCTURE: Structures BREAK SIZE: Intermediate "[S at Pool Surface ADS See Attachment A STRUCTURE: Structures BREAK SIZE: Small at Pool Surface No addit.onal loads generated (See large break tabulation)
O S y a 2:
O O e
i a J v Table 1.3.1 (Continued)
P" "
esign Basis Engr'g Load Design Estimate Analysis Test Section Comments STRUCTURE: Structures BREAK SIZE: Large Between Pool ,
Surface and IICU Floor Slug Impact Loads l Small flat structures 115 psi 60 psi 5801, 5802 10.1 See Attachment J 5805, 5706 .
Piping 60 psi 30 psi 5801, 5802 10.1 See Attachment J 5805, 5706 Froth Impingement Loads 15 psi 15 psi 5706 10.1 See Attachment J and Figure 10.6 gu 4
40 ft/sec 30 ft/sec Bounding 10.2 See Attachment I. See -
Velocity for Computing Drag Loads calculation Figure 10.3 for grating p$
loads Fallback Velocity for 35 ft/sec 20 ft/sec Bounding 10.3 Drag Loads calculation STRUCTURE: Structures BREAK SIZE: Intermediate Between Pool Surface and 110U Floor No additional loads generated (See large break tabulation)
STRUCTURE: Structures BREAK SIZE: Small Between Pool Surface and llCU Floor (See large break g No additional loads generated tabulation) y g
5 U
Table 1.3.1 (Continued)
Spe ed
, Design Basp Load Design Estimate Analysis Test Section Comments STRUCTURE: Expansive bMEAK SIZE: 1.arge Structures at itCU Floor Eleve. tion Wetwell Pressurization 11 psig 3-5 psig Model la 5801, 5802 11.0 (3-4 see) (1-2 see) Ref. 1 5803, 5804 Froth Impingement 15 psig (100 ms) 10 psig (100 ms) 5801, 5802 5805, 5706 11.0 See Attachment J discussion l' 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
- n w Fallback and Water 1 psi 0.5 psi Bounding 12.0 Based on water flow ,
Accumulation calculation through IICU floor *$
no U' STRUCTURE: Expansive BREAK SIZE: Intermediate Structures at 11C0 Floor Elevation No additional loads generated See large break tabulation STRUCTURE: Expansive BREAK SIZE: Sumil Structures at ilCU Floor Elevation _
No additional loads generated See large break tabulation CENERAL NOTES TO TABLE 1.3.1
$; 1. Where S/R valve loads are sp'ecified in the applicable bar charts, refer to Attachment A,
$; Section AS.6 for margin discussion. 7 g 2. Not all loads for IBA and SBA are tabulated. Generally the large break load condition will govern. $l e e e' . i
1 1
Table 1.3.1 (Continued) !
4 Specified Design Basis for Engr'g Test Section Comments Load Design Estimate Analysis STRUCTURE: Small Structures at IICU Elevation 5801, 5802 12.0 See Attachment J Froth 1mpingement 15 psid 10 psid 5805, 5706 discussion ;
Model in 5801, 5802 12.0 Test shows pressure Flow Pressure 11 psid 3-5 psid Ref. 1 5803, 5804 differential of 3 to Differential 5 psi 0.5 psid Bounding 12.0 Based on water flow Fallback and Water 1 psid through HCU floor Calculation Accumulation 5"
<W 0 j ug .
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l 22A4365 i Rev. 2 _ 2-1 ,
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- 2. REVIEW OF PHENOMENA The purpose of this section of the report is to qualitatively review the sequence 1 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 safety relief valve actuation. The objective of this review is to provide an understanding of the various pool dynamic loads and their inter-relationships, and to define the dynamic loading terminology. Specific design load values are provided in subsequent sections.
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 at:nosphere. At the break exit point, the wave amplitude theoretically is reactor operating pressure (1000 psia). However, there is rapid attenuation as the wave front expands spherically outward into the drywell at sonic velocity.
As the dryvell pressure increases, the water initially standing in the vent sys-tem accelerates into the pool and the vents are cleared of water. During this 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 the suppression pool and on the containment wall opposite the vents. During the vent clearing transient, the drywell is subjected to a pressure differential and the weir wall experiences a vent clearing reaction force.
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). This bubble theoretically transmits a pressure wave through the suppression pool water and results in l loading on the suppression pool boundaries and on equipment located in the O 99te 1e= ree1- !
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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.
GE 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 riss 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 axperience 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 chase is referred to as incipient breakthrough; i.e. , ligament begins to break up.
O Ligament thickness continues to decrease until complete breakthrough is reached d
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 drywell air into the suppression pool results in a period of froth pool swell. This froth 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 loc 4ted, 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 themse'ves and on the adjacent containment and drywell. The result is a discontinuout ' essure loading at this elevation.
h.
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4 j 22A4365.
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- O l Figure 2.2-1 is a diagram that sumanarizes the various phases of pool swell and I
the nature of the dynsmic loading conditions that occur. It should be empha-4 sized that the pool swell elevation information presented on Figure 2.2-1 is l based on an assessment of the PSTF air tests. As such it is considered con-servative since Jte PSTP air test data have been interpreted and used in a consarvative manner.
i 4
l The pool swell impact and impingement target data presented in Section 10 of this document is applicable to small structures. This restriction on the app).ication of the igset 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 insediate 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 saml1 structures.
f For floors that are expansive enough to decelerate a large sector of the pool v rather than a small region of the pool in the vicinity of the target, the impul-sive loading on the floor is dependent upon the momentum of the entire slug and 1
is related to slug thickness.
l l 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.
042178
22A4365 R;v. 3 s.
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 h reduced steam flow rate leads to a reduction in the drywell/ containment pressure differential which in turn results in a sequential recovering of the horizontal vents. Suppression pool recovering of a particular vent row occurs when the vent stagaation differential pressure corresponds to the suppression pool hydro-static 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 row of vents. In addition, an oscillatory prissure loading condition can occur on the drywell and containment, but is insignificant. Since this phenomenon is steam mass flux dependent (the chugging threshold appears to be in the range of 10 lb/
sec/ft2) it is present for all break sizes. For smaller breaks, it is the '
only mode of condensation that the vent system will experience.
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 reacto. 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. Because 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 drywe)l. 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 negat$ve load condition is therefore speci-fied for drywell design.
Following vessel flooding and drywell/ containment pressure equalization, sup-pression pool water is continuously recirculated through the core by the ECCS O
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22A4365 R;v. 2 - -
25 .
ptmps. 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 RER 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 173*F and with low : term containannt spray operation the peak temperature can approach 180*F.
The increase in air and water vapor pressure at these temperatures results in a pressure loading of the contain==nt.
The post DBA contain= ant heacup and pressurization transient is terminated when the RER heat exchangers reduce the pool tagerature and containment pressure to nominal values.
2.2 INTERMEDIATE BREAK ACCIDENT (IBA)
An intermediate size break is defined as a break that is less than the DBA but is
[ of sufficient magnitude to automatically depressurize the primary system due to loss of fluid and/or automatic initiation of the ECCS systems. In practice, this O means liquid breaks greater than 0.05 f t2 and steam breaks greater than 0.4 ft2 as determined by analysis.
In general, the magnitude of dynsmic loading conditions associated with a loss i
of coolant accidents decrease with decreasing break size. However, the inter-mediate break is examined be.cause une Automatic Depressurization System (ADS) may be involved. Simultaneous actuation of the multiple safety /relier valves committed to this system introduces significant containment system loads, as discussed in Section 2.4
! 2.3 SMALL BREAK ACCIDENT (SBA) l 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 l
l l
042178
22A4365 R3v. 2 - 2-6 designed 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 h sequence of events is postulated.
With the reactor and containment operating at m-4=_= 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 pressura 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 end 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 pos2.tive pressure differential sufficient to keep the weir annulus water level depressed to the top vents and chugging can occur. Continued reactor blowdown steam is g 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 frca 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 at=os-pheric pressure, the steam-and-liquid blowdown will have a temperature 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, decompression of 1,000 psia saturated steam to atmospheric pressure results in 298'F superheated steam (86*F of superheat). h 042178
22A4365 i Rev. 2 , 2-7 -
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- . . . 1 Reactor operators are alerted to the SBA incident by the leak detection system, or high drywell-pressure signal, and reactor scram. For the purpose of eval-unting the duration of the superheat condition in the drywell, it is assumed 4
i that operator response to the small break is to shut the reactor down in an orderly ==nnar using selected relief valves and with the RRR heat exchangers controlling the suppression pool temperature. (This assumes the main condenser
! is not available and the operators must use the suppression pool for an energy I
sink. In all probability, the condanser would be available and the suppression j pool would not be involved in the shutdown.) Reactor cooldown rate is assumed "to be started 30 minutes after 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 i
- the blowdown flow rate is terminated. It should be noted that the end-of-1
- blowdown chugging phenomenon discussed in Section 2.1 will also occur during a I small break accident and will last the duration of reactor depressurization.
j .
i 2.4 SAFETY RELIEF VALVE ACTUATION 1
In addition to loads on the valves and discharge piping, actuation of the j 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 j vall, the dtywell and the containment adjacent to 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.
4 l 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 t
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 f through the holes in the quencher arms and forms a large number of small bubbles.
1 Once in the pool, the bubbles expand, coalesce and form four large bubbles.
I Each of the four bubbles expands analogous to a spring and accelerates the sur-
- rounding pool of water. The momentum of the accelerated water causes the bubble to over-expand and the bubble pressure becomes negative. This negative pressure O slows down and finally reverses the motion of the water leading to compression l
l l
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22A4365 Rev. 2... 2-8 -
of the bubble. This sequence of expansion and contraction is repeated with a maximum frequency of about 12 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).
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 associated 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; 1.e., any single component may fail to act when called upon.
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 with a single safety / relief valve actuation as an additional load for demon-strating additional capability.
Safacy relief valve loading data is discussed in Attachment A.
l l
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. -
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22A5365 Rev. 2 2
_-9 .
EVENT
- POTENTIAL LOACING CONOlTION'
(_,/ e COMPRES$3VE WAVE LOCA OCCURS ,'" LOAOING ON CONTAINMENT ORYWELL PRESSURE RISES o SONIC WAVE LOACING OF ORYWELL 1I I
o JET IMPlNGEMENT ANO eue8LE VENTS CLEAR ANO VENT m PRESSURE LOAOS ON THE AIR / STEAM PLOW STARTS CONTAINMENT e VENT CLEARING ANO VENT FLOW AP ON ORYWELL e OUTWARO FLOW AP ON WClR WALL lI e IMPACT LOAOS ON LOW POOL SWELLS IN m STRUCTUR ES A SULK MODE e OR A C LOAOS ON STRUCTURES IN AND A00VE THE POOL If BREAKTHROUGH 1f POOL SWELL CONTINUES IN A e FROTH IMPtNGEMENT ON
' FROTH' MODE AND ENCOUNTERS - HIGH STRUCTURES FLOW RESTRICTION AT HCU e FLOW AP ON HCU FLOOR A
F FLOOR , AND AOJACENT CONTA.NMENT lf OR YWE LL VENTING -
e ' FALL S ACK* LOACS ON COMPLETE STRUCTUR ES If STEAM CONOENSATION 5 e CONOCNSATION LOAOS IN POOL AT VENT ExlTS f
m o WCtR WALL ANO ORYWELL GLOWOOWN ENOS "
LOAOS DUE TO CHUGGING 1I e NEGATIVE PRESSURE ON WEIR ECCS PLOOOING OF PEACTOR WALL, ORYWELL ANO ITS VESSEL AND ORYWELL 7 PEN ETR ATIONS OEPRESSURIZATION e NEGATIVE FLOW t.P ON WCIR WALL 1I G E HEAT UP i e CONTAINMCNT PRESSURE LOAO
- ALL POTENTI AL LOCA OYN AMIC LOA 05 ARE lOENTirtEO, BUT ALL ARE NOT SIGNIFICANT ISEE TEXT FOR OETAILSI Figure 2.1. Loss-of-Coolant Accident Chronology (DBA) j i
I 042178
22A5365 -
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M40 f t 0 0 0 o 0 0
0 0 0 0 0
D 0 )
0 g o SPRAY Q
0
- NO SIGNIFICANT LOAOS g g D g 0 Q D g
L g 0 0
0
+20 l 0 Oc a o O
h h h l FROTH h h - LOW IMPINGEMENT LOAOS LOW ORAG LCA05 C
C g o Q c G 12 f t BULK POOL SWELL
- bi - HIGH SLUG VELOCITIES
- ^^-^^^^^ ^6- b~^ ~:__ bib ^ ' -'5--
^_5 5 5 --5HIGH OR 3AG 3 ^h IMPACT LOADS
^ ^ ^
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INITI AL POOL SURF ACE Figure 2.2-1. Schematic of the Mark III Pool Swell Phenomenon -
042178
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O Figure 2.2-2. Typical Suppression Pool Cross Section 238 Plant 042178
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042178
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042178 l .
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Figure 2.2-5. Containment Equipment Drain Su::p 238 Plant 042178
22A4365 2-15 Rev. 2__ .
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- 2'2A6365 Rev. 2 _ 3-1/3-2 ,
- 3. DYNAMIC LOAD TABLE 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 "bomas" which span the given 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 Drywell X X X Weir Wall X X X Containment X X X 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 Notes:
- 1. X indicates accident with significant loading conditions
- 2. For concurrent S/R valve events, see appropriate bar charts 042178
22A4365 Rev. 3 4_1 O 4. DRLELL STRUCTURE t
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 ccher 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 i
steam line break. A discussion of the loading conditions follows:
4.1.1 Sonic Wave Theoretically, a sonic cogressive wave is initiated in the drywell atmosphere i
following the postulated instantaneous rupture of a large primary system pipe.
' 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 'rywell Pressure During the vent clearing process, the drywell reaches a peak calculated differ-
' ential pressure of 21.8 psid. 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-This wetwell ment annulus restriction at the HCU floor (see Figure 4.4).
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 surface does not adversely affect the drywell pressure.
Figure 4.4 shows the drywell pressure during the DBA. It includes the HCU floor pool swell interference effe. cts. The analytical model presented in Ref. I was used to calculate these values.
090779
~ - .. -
22A4365 4-2 Rev. 2 Blockage of the weir annulus flow area by equipment located above the annulus entrance has the potential for increasing the real drywell pressure differen- h tial.
Attachment C presents data which show no potential pressure increase fo.
blockages up to 30 percent of the total area.
During the blowdown process, the drywell is subjected to differential pressures between levels because o,f flow restrictions. This value varies with the size of the restriction, but a bounding value for a 25 percent restrictica is 0.5 psi On the basis of this bounding calculation, it as discussed in Attachment D.
has been concluded that differential presaures 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 lower regions of the drywell experience an inward load due to the hydrostatic pressure associated with the pool water.
If it is asotmed that an earthquake is occurring at this time, the horizontal and vertical accelerations of the building can influence the hydro-static pressure calculations. See Attachment B.
4.1.4 Loads On The Drywell Wall During Pool Swoll During bubble formation, the outside of the dryvell wall in the pool will be A bounding range of 0 to 21.8 psid is specified subject to varying pressures. The on those sections of the drywell wall below the suppression pool surface.
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 nominal suppression pool surf ace, the pressure (see Figure 6.5).
linearly decreases from 21.8 psid to O paid over 18.0 feet Any structures in the containment annulus that are within approximately 20 feet of the initial suppression pool surface will experience upward loads during 4
i O
101678
22A4365 Rev. 3 4-2a 1' J pool swell (see Figure 12.2). If these structures are attached to the drywell wall, then the upward loads will be transmitted into drywell structure. In i
' 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.
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- 090779 i
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22A4365 Rev. 3 4-3 I
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 !
is 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 been performed in the PSTF facility.
Some pressure oscillations have been observed on the drywell wall. Figures 4.5 and 4.Sa give a su:nmary of the magnitude of the top vent exit pressure fluctuations observed during PSTF steam tests. The data has been plotted against wat submergence and is independent of this parameter.
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 decreases from approximately 110 paid to approximately 2 paid 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):
P(T) = Af' { 0.8 sin (2n x T x f(t)) !
+ 0.3 sin (4n x T x f(t))
+ 0.15 sin (6n x T x f(t))
+ 0.2 sin (8n x T x f(t)) } (psid) 0 090779
22A4365 Rev. 3 4-3a where:
P(t) = pressure amplitude (psid) for consecutive cycles beginning at time t = 3 sec.and ending at T p
n i
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) - 1.344 (log t)') Eqn (4.2) f(t) = fundamental frequency variation with time. (Hz)
= = 0.8 (2.495 - 0.225 t - 0.742 log t + 10.514 (log t)
I
- 9.271 (log t) + 3.208 (log t)4) Eqn (4.3) t = time (sec), 3 1 t 1 30, time from initiation of LOCA blowdown t = time increment for successive periods Tp <t <T p, Tp =
f 3); where n is number of cycles between 3 and 30 sec.
~
p p +T p + ... + T p n
f (3 + T 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 wetted surface of the suppression pool is shown in Figure 4.6a. The amplitudes shown a re the maximum values determined from Eqn (4.1) normalized to 1.0 at the top vent centerline.
4.1.6 Fall Back Loads In general, the data generated in the PSTF indicates that no significant loading conditions on the drywell wall occur during pool fall back. Figure 6.4 shows that suppression pool vall pressures following bubble breakthrough return to their initial pre-LOCA values during the 1.5 to 5 second period when the pool l level is subsiding. Therefore, fall back pressure loads are not specified for i Mark III drywell.
090779
22A4365 Rev. 3 4-4 Structures attached to the drywell wall expe'rience drag loads as the water level subsides to its initial level. These structures could experience drag forces associated with water flowing at 35 ft/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 to the drywell and start condensing the steam in the drywell. The rapid drywell depressurization produced by this condensation will draw non-condensable gas from the containment free space via the drywell vacuum breakers. It is during 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 llh relief system. This theoretical conservative calculation yields a drywell to containment negative pressure differential 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-ferential between the drywell and containment as shown in Figure 4.9. The maximum values of this load (+2.0, -0.7 psid) are negligible when compared to the peak positive drywell pressure used for drywell design and the negative pressure discussed in Attachment E (Peak Negative Drywell Pressure). Chugging is an oscillatory phenomenon having a period of 1 to 5 seconds.
The PSTF data shown on Figure 4.9 is from the 5801, 5802, 5803 and 5804 series of 1/3 scale PSTF tests. The data has been plotted against top vent sub-mergence with no obvious cor> 21ation. Because volumes and areas of the 1/3 scale test are correctly scaled, the tests are more appropriate as a source of chugging 090779
22A4365 Rev. 2 4-5
- () >
1 induced drywell pressure data than large scale tests 5701, 5702, and 5703 discussed in Reference 4. The large scale PSTF configuration had a drywell volume to vent area ratiu 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 races at the vent exit and water splashing into the drywell. The undersized dry-l well of the large scale PSTF would tend to exaggerate the drywell pressure 4
response.
4.1.9 Loads Due to Chugging i
In addition to the bulk drywell 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 traic 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.
l 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 oressure spike. The peol bound-ary load definition considers that the chugging loads transmitted to the dry-well wall, weir wall, basemat and containment are the result of several vents chugging simultaneously at different amplitudes. ,
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101678
22A4365 l Rev. 3 4-Sa l O1 1
1 4.1.9.1 Chugging Loads Applied To Top Vent 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 force pulse train shown in Figure 4.7b is applied vertically upward over the projected vent area concurrently with the peak For global effects, pressure pulse train to evaluate local effects at one vent.
the average force pulse train shown in Figure 4.7c is applied vertically over the projected area of all top vents 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 paid with a standard deviation of 23 psid. On this basis, the specified design value is
-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 consists of a pre-chug underpressure defined as a half sine wave, a triangular pulse (pressure spike) loading characterized by a time duration "d" and a post-chug oscillation described by a damped sinusoid.
The pulse is at its maximum magnitude and duration near the top vent on the drywell wall due to the localized nature of the phenomena. The amplitude of the pre-chug underpressure and the post-chug oscillation are also maximum at
.this location.
O 090779
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22A4365 I
Rev. 3 4-5b O For local load considerations on the pool boundary:
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 e 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 l liner buckling and vent liner stresses. Local chugging loads shall not be conbined 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:
o Pre-chug underpressure
- mean amplitude - Table 4.1 distribution - Figure 4.8c O
090779
22A4365 Rev. 3 4-6 l
e Pulse (spike) g
. mean amplitude - Table 4.1
- distribution - Figure 4.8d
- duration - Figure 4.8e e Post-chug oscillation mean amplitude - Table 4.1 distribution - Figure 4.8c 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 invcived. (See Attachment A) Figure 4.3 is a bar chart for this condition. lll 4.3 DRYWELL DURING A SMALL BREAK ACCIDENT A small steam break can lead to high atmospheric temperature conditions in the drywell. 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 normal shutdown. 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 combination of primary system pressure and drywell pressure which produces the maximum superheat temperature.
~
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 090779
22A4365 Rev. 3 4-7 O During an SBA the continuing blowdown of reactor steam will cause all the air j
initially in the drywell to be purged to the contain==nt free space. The peak l
superheat temperature is 330*F. This tagerature condition exists until the At this time, RER shutdown cooling is completed in approximately three hours.
after three hours, the pressure in the reactor pressure vessel is 150 psia j ad the corresponding superheat tegerature is 310*F. This will exist for tht e hours. These superheat tegeratures correspond to drywell atmosphere l
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 marinum normal conditicas, a small l break occurs allowing blowdown of reactor steam to the drywell. The resulting l
drywell pressure increase leads to a high drywell pressure signal that scrams lt the reactor and activates the containment isolation system. Drywell pressure l
}
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 containment 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 contain:nent. The drywell is now full of steam and a positive pressure differential sufficient to keep the weir annulus water level depressed to the top vents is maintained. Continued reactor blow-down steam is condensed in the suporession pool.
l f
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090779
22A4365 Rev. 3 4-8 4.3.3 Chugging 1 During a small break accident there will be chugging in the top vents.
l Applicable chugging loads on the drywell 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 thermal loading condition with the 21.0 psi negative pre ssure 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 envelope of drywell atmospheric pressures and tempera-tures for the spectrum of postulated loss of coolant accidents. This figure represents a conservative definition of calculated peak drywell conditions.
Figure 4.10 defines only the drywell atmospheric condition; separate analyses are required to evaluate the transient structural response to these conditions.
' These envelopes should be used judiciously, since it is not possible to have concurrent high drywell pressure and temperature.
4.6 TOP VENT TEMPERATURE (CYCLING) PROFILE DURING CHUGGING Full scale test results (Reference 16) indicate that during chugging the l water level in the weir annulus fluctuates over a 4 foot band centered at l about the top vent centerline. The weir wall and the inside drywell wall then are subjected to steam temperature (230*F) above the top vent and cold pool temperature (100*F) near the lower vents, with a transition region in-between, where the temperature fluctuates due to the chugging process.
090779 91
22A4365 w3 4-8a O
For weir annulus thermal stratification, the most severe design condition results from imposing the maximum drywell temperature (3300F) concurrent with the initial suppression pool temperature (see Section 4.3.1).
For evaluation of local effects, the cyclic temperature profile during 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 varies from 1 to 5 seconds.. The areas of application are:
e 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 vall just above each top vent, as shosn on Figure 4.11.
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 Attachment N.
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i O l 090779 i
Table 4-1 CHUGGING LOADS PRECHUG UNDERPRESSURES PULSE (SPIKEl AND P05I LMuG 05CILLAil0N AND DURATION DURATION *d'* AND fALQu[mCT PEAK (Al ME AN (A) PLAA til A,e PEAK (8) MEAN18)
-5.8 P510 -2.65 P5ID 100 P51D 24 PilD 6.50 P510 2.2 PSID DATWILL HALL 125 M5 325 M5 3 M5 8 M5 10 12 Mg 10 12 Ma
-1.3 P51D -1.0 P5ID 3 P5lD 0.7 F510 l.7 P5ID l.00 P5ID py COnTAlhM(hi 125 M5 125 M5 2 MS 2 MS . 10-12 M2 10-12.N2 $%s~
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-1.8 to -l . 3 PSID -1.34 to -1.0 P5ID 10 to 3 PSID 2.4 to 0.7 PSID 2.1 to a 1.7 PDID 1.29 to a 1.0 Pil0 8ASIMAI 125 MS 125 M5 4 to 2 M5 4 to 2 M5 10-12 N 10-12 Hz Y
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a O O O
O O .
O STRUCTURE: DRYWELL ACCIDENT: LARGE STE AM LINE BREAK (DBA)
POOL TEMPER ATURE (SECTION 4 6)
(SEC 4.5)
DRYWELL INT [HNAL PRESSURE AND TEMPER ATijRE (SEC 4.1.2 AND FIG. 4.41 LOADS DUE TO SEISMIC ACCELERATION OF THE STRUCTURES AND LOADS DUC TO SEISMIC INDUCED POOL SURF ACE WAVES lATTACalMENT 06 HYDROST ATIC PRESSURE NOT E; - THERE WILL BE NO WATER IN THE WElR ANNULUS BETWEEN 1 ANO 30 $[CONDS ISEC 4.1.31 l - POOL DUMP ST ARTS AT 5 mm ATTACHMENT A.SEC 2A SINGLE $/R VALVE ACTUATION
- THE DRYWELL HEAD LOADS DUE TO FIGURES 10.3.10.4 DESIGN CONSIDERS z POOLSWELL 10.5 AND 12.2 A HE AD SPR AY 9 SECTION 4.1.4 BREAK.
t ^
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- b F ALLB ACK LOADS 5 E CONTAINMENT FREE SPACE
<N
.u kO PRESSURIZATION DUE TO N F U M _- POOL F ALLBACK SWELL LOADSAND g@
DRYWELL AIR CARRYOVER SECTION FOR A GIVEN STRUCTURE ARE WETWELL PRESSURIZATION CHUGGING 4.1.8 AND (LOAD ON DRYWELL O.0.1 4.19 NOT COINCIDENT .
BOTH LOADS HAVE SECTION 6.1.6,12.0 NEG ATIVE LOAD DUE TO POST A DURATION OF LOCA ECCS 0.5 sec. POOL SWELL CAN OCCUR 1 TO SECTION 4.1.7 FLOODING Op ORYWELL 1.5 sec AFTER BREAK LOCA BUBBLE DEPENDING ON PRESSURE SECTION 4.l.4 HEIGHT A80VE THE LOAD POOL. FALLBACK POST LOCA LOADS OCCUR SECTION 4.1.1 SECTION 6.1.8 way WAVES 1.5 T0 5 secAFTER THE BREAK CONDENSATION OSCILLATION SECTION 41.5 1
l l l 3.0 5 10 30 100 600 0 0.1 1.0 1.5 TIME AFTER EVENT (soci o ..
g 5
e '
Figure 4.1. Drywell-Loading Chart for DBA
.T
STRUCT UR E: DRYWE LL ACCIDENT. SM ALL STE AM BRE AK (SS Al LOADS DUE TO THE SEISuic ACCELERATION OF THE STRUCTURES ( ATTACHMENT 8)
AND LOADS DUE TO SEISMIC INDUCED POOL SURF ACE WAVES HYDROSTATIC NOTE 1. THE W[tR ANNULUS WILL BE CLE ARED TO THE TOP OF THE UPPER VENTS WITHIN SEC ON PRESSURE A FEW MANUTS OF THE ACCIDENT. ITIME IS BREAK ARE A DEPENDENTI 4
- 2. POOL DUMP INCLUDED (AUTO AT 30 ernni SECTION DRYWELL ATMDSPHERE TEMPERATURE (FIGURE 4101 ,
NOTE: DLRING COOLDOWN WITH CONDENSER ISOLATED.
z SINGLE SIR VALVE ACTUATION S/R VALVES ARE OPERATED PERIODICALLY FOR (SEC 2.4 & 4.31 9 UP TO THREE HOURS lATTACHMENT Al 3 CHUGGING l
SECTION Z NOTE: CHUGGING CAN LAST UNTIL BREAK ISOLATED OR VESSEL O DEPRESSURIZED. (NOTE TWO TYPES OF LOADS) ______ _ _ _] #f g
? NOTE. NEGATIVE LOAD DUE TO FLOODING TO COOLING OF DRYWELL POST ACCIDENT IS NO MORE %. La 3 SEVERE THAN THAT FOR THE LOCA RELATED EVENT $
.a -
CONTAtNMENT PRESSURE R AISED TO 3 pu.
sRYWELL PRESSURE DIFFERENTIAL RAISED TO 3 ps.d POOL HEATUP RAISES CONTAINMENT PRESSURE TO 6 peig SECTION DRYWELL PRESSURE DIFFERENTI AL MAINTAINED AT 3 psid 432 I I i 1.0 aren 3 ha 6 hr TIME AFTER EVENT 8
o 3 Figure 4.2. Drywell-Loading Qiart for SBA /o e
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TIME AFTER INITIATION OF BLOWOOWN leec) o s~
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u u y Figure 4.6b. Condensation Oscillation Forcing Function on the Drywell Wall 0.D.
Adjacent the Top Vent
22A4365 4-18 RI,v. 2 O
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O Figure 4.7. Typical Top Vent Pressure Trace During Chugging, Run 19 0
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l (Ref. Test 5707) i l I i l l
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A PULSE (PRESSURE
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SEE TABLE 4.1 FOR VALUES FOR A,8 AND d PRECHUG UNDERPRESSURE (A SIN w ty ) FOR t g w* w/0.125 radians /sec FROM 0 TO 0.125 see a = 0.55/r,1/sec d = 2n/r, radiens/sec Figure 4.8a. Typical Pressure Time-History on the Pool Boundary During Chugging 090779 i
22A4365 Rev. 3 4-21a 1
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090779
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090779
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figure 4.8e. Suppression Pool Chugging Spike Duration "d" as a Function of Location in the Pool O
090779
22A4365 4-2.le Rev. 3 4
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, 090779 l
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Figure 4.8h. Circumferential Post Chug Oscillation Amplitude Attenuation O 1 101678 j l
22A4365 2 4-22 -
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Figure 4.9. Drywell - Containment Pressure Differential During Chugging '
042178
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I CYCLING FLutD TEMPERATURE TIME HISTORY 230*F 1
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Figure 4.11. Drywell Top Vent Cyclic Temperature Profile and Area of Application During Chugging 101678 l
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 4.1.1, this phenomenon is not included in the weir wall design conditions. A sonic compressive wave does not produce a design load condition in the drywell.
5.1.2 Outward Load During Vent clearing O 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 O 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 i
~ , -
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 verrical direc-tion is shown in Figure 5.6. For all global loads, there is no attenuation in the circumferential direction.
O O
090779
_ i I
22A4365 .
Rev. 3 5-3
) 5.1.5 Inward Load Due to Negative Drywell Pressure Due to negative drywell pressure discuseed 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 ft/ m 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 H.
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 25 feet /
sec can be calculated at the top of the weir wall. This velocity is decreased O' due to the effects of gravity with elevation and the spreading of the flow field so that the maximum elevation reached is 11 feet above the top of the weir wall as shown in Figure 5.8. Structures in the path of the water are designed for drag loads using the following equation:
o CD 0
- F =
2g c
where:
F = Drag Load Force, Ibf C = rag coefH eient D
A = Projected Area Normal to Flow, Ft p = Density of Water, 62.4 lbm/ft i gc = Newton's constant, 32.2 lbm-ft/lbf-sec2 l V = Velocity of fluid, ft/sec.
O l 090779
22A4365 R2v. 3 5-4 5.1.6 Suppression Pool Fallback Loads For the reasons presented in 4.1.6 and since the weir annulus pressure is I
controlled by vent flow during the period of interest, no supprecsion pool l fallback pressure loads are upecified for the weir wall.
5.1.7 Hydrostatic Pressure During the first second after the DBA, the water in the annulus is depressed to the bottom vent; therefore, there is no inward hydrostatic pressure load on l 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.
l 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 through the vents. AttJehment A provides loading information. Tha S/R valve load is applied to the projected vent hole area on the weir wall.
l l
5.1.9 Condensation i
There will be no loads induced on the weir during condensation, as shown by l
lack of transducer response in the tests.
5.2 WEIR WALL LOADS DURING AN INTERMEDIATE BREAK ACCIDENT l
Figure 5-3 shows the bar chart for the weir wall during the IBA. The safety relief loads associated with ADS activation are discussed in Attachment A.
The LOCA induced pressure differential across the weir wall will be small.
l O
090779 l l
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l
22A4365 5-5 !
Rev. 3 5.3 WEIR WALL LOADS DURING A SMALL BREAK ACCIDENT 4
s_/
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, i
5.4 WEIR WALL ENVIRONMENT ENVELOPE
] The temperature and pressure for the drywell envelope data (Figure 4-10) applies to the weir wall with the exception of that part of the outside face which is below the elevation of the upper vents. This region will remain sub-merged and will be maintained at suppression pool temperature. It should be
' noted that the weir wall structure is totally within the drywell and effects a
of environmental conditions should be examined on this basis, including the thermal cycling during chugging (see Section 4.6).
The first 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br /> 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.
O 090779
STRUCTURE: WEIR WALL ACCIDENT: DESIGN BASIS ACCIDENT (DBA)
WEIR WALL PRESSURE AND TEMPERATURE SECTION S.4 AND 4.6 SEISMIC - STRUCTOR AL ACCELERATION LOADS
- POOL SLOSHING LOADS (ATTACHMENTB)
HYDROSTATIC PRES $URE -(NONE BETWEEN O & 30 soci (SECTION 5.1.7)
LOADS DUE TO SINGLE S/R VALVE ACTUATION ^ ^ ^
E TlO 51 & 2A
$ NOTE: CHUGGING AND INWARD LOAD DUE TO
{ POST LOCA FLOODING ARE NOT COINCIDENT.
O 8
0
{< OUTWARD LOAD
- VENT CLE ARING SECTION 51.2 W
s g
.m OUTWARD LOAD - VENT F LOW SECTION S.1.3 CHUGGING SECTION S.l A & FIGURES 54.5.5.5.6 l w$ ,
INWARD LOAD FALLBACK SECTION DUE TO POST l l LOADS 5.16 LOCA ECCS F LOODING OF SECTION 5.1.5 l i DRYWELL SONIC SECTiON 5.1.1 WAVE f
1 I i l I I I I o, i i .5 5 30 too 600
^ '"'
- ADD Si8 DYNAMIC LOAD TO STA TIC LOAD DUE TO DRYWELL AIR PUHGED TO CONTAINMENT, VAPOR PRESSURE AT 140' F.
APPLIES TO BOTTOM 2 VENTS ONLY h Figure 5-1. Weir Wall-Loading Chart for DBA Y.
e 9 9
O O O STRUCTURE. WElR WALL ACCIDENT. SMAt L BREAK ACCIDENT IS8Al LOADS DUE TO SEISMIC ACCELERATION OF THE STRUCTURES AND LOADS DUC TO SEISMIC INDUCED POOL SURFACE AND WAVES.
IATTACHMENT Bl HYDROSTATIC NOTE: THE WEIR ANNULUS WILL BE CLEARED TO THE TOP OF THE UPPER VENTS WITHIN A F EW MINUTES OF SECTION PRESSURE THE ACCIDENT. M.7 y ATMOSPHERE TEMPERATURE 54 0 g" fD O FIGURE t
O SINGLE SIR VALVE ACTUATION NOTE: DURING COOLDOWN WITH CONDENSER 4.10 . h m
ISOLATED. PERIODICALLY FOR UP TO THREE ISEC 2 41 o HOURS lATTACHMENT Al o ----------
7 SECTION E CHUGGING NOTE: CHUGGING CAN LAST UNTIL SREAK ISOLATEDOR VESSEL DEPRESSURIZED l 5.1.4.
FIGURE 2.3 o
$ - - - - - - - - - d 5.4. 5.5 AN D 5.6 a
1mm 3 has 6 has TIME AFTER EVENT S
S 3 Figure 5-2. Weir Wall-Loading Chart for SBA 4
M
STHUCTURE. WE6R WALL ACCIDENT INTERMEDs ATE BRE AK ACCIDENT (ISA)
LOADS DUE TO SEtSMIC ACCELER ATaOtd OF THE STRUCTURES AteD LOADS DUE TO SEtSuiC aNDUCED POOL SURF ACE W AVE llIATTACHuCNT St HYGROST ATIC PRE SSUR E NOTE POOL DUMP INCLUDED AFTER ADS.WE6R ANNULUS LEVEL AT UPPER VENT LEVEL.eSEC S 176 ATTACHMENT A SINGLE S/R V ALVE ACTUATION h SECTION 2.4 ADS ACTUATED AIR RETURN TO SECTION DRYWELL 4.32 3 OUTWARD LOAD SECTaON 5.1.2 5 VENT CLEARING 5
" OUTW AR O LOAD SECTaON S.I.3
$ VENT F LOW (SMALL COMPARED TO 08 Al N
o
$" DRYWELL AIR PURGED TO POOL HEATUP RAISES CONTAINMO4T PRESSURE AND TEMPERATURE $%
CONTAINMENT 3 pe.d (SECT. 432) (SECT.5 41 ,
La vi CHUGGtNG SEC 5.1.4 AND 5.1.9 SECTIONS 514.5.1.9 VENT CLEARING i a l i e a 60 ** 500 600 1000 TtME AFTER EVENT.sec
- TIME SCALE DEPENDENT UPON BREAK SIZE. MIN! MUM VALUE OF t = 2.0 tru o
g 1 SINGLE SRV LOADS DO NOT COM8tNE WITH OTHER SP.V LO ADS d l v,
- Figure 5-3. Weir Wall-Loading Chart for IBA e CD 1
- 9 e ----- - - -
22A4365 5-9 Rev. 3 O
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i Figure 5.4. Typical Weir Wall Chugging Pressure Time History - Test Series 5707 Run 1.
090779
22A4365 Rev. 2 5-10 2
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C PRESSURE PULSE TRAIN &
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i 101678
22A4365 Rev. 3 5-10a O
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TIME
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0.080 esc n
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s 2 P = -0.98 SIN st/0.080 4 MEAN FOR 0 < t < 0.080 sec TIME
-4.98 -
0.080 esc Figure 5.5. Underpressure Distribution on the Weir Wall and Drywell I.D. Wall During Chugging O
090779
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4 PERIOD BETWEEN CHUGS = 1 TO 5 sec >
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During Chugging O O O
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4 PEHIOD BETWEEN CHUGS = 110 6 ses >-
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1 1
l 22A4365 Rev. 2 5-11 O
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THIS ATTENUATION ALSO APPLIES TO THE CIRCUMFERENTI AL D!RECTION
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22A4365 Rev. 3 6-1
- 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 sp ay activation, included pipe whip, shield building loads, jet impingement etc. that are not in this report.
6.1 CONTAINMENT LOADS DURING A LARGE STEAM LINE BREAK (DBA)
Figure 5-1 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 O' Very rapid compression of the drywell air could, theoretically, result in a compressive wave being generated in the weir annulus water. This wave could then travel down the weir annulus, through the vents and accross the pool to the containment wall. This phenomenon is not specifically included in the containment 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 /see wave were generated at 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 ioads. 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 0.8 sec. Water jet loads are negligible when compared to the subsequent air bubble pressure discussed in Section 6.1.3 and are not specifically included as a containment design load.
090779
22A4365 R v. 3 6-2 6.1.3 Initial Bubble Pressure O
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 be more representative than the large scale steam blowdown tests. The maximum 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-down. The Mark III design load which is based on these tests, is shown in Figure 6.5.
The magnitude of the containment pressure increase following vent clearing is dependent upon tha rate at which the drywell air bubble accelerates the suppres-sion pool water. Circumferential variations in the air flow rate may occur due to drywell air / steam mixture variations but it results in negligible variations in the containment bubble pressure load. (See Attachment L).
O 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 blow-down. Figure 6.4 shows a typical transient. A maximum increase of 10 psid on the containment wall was observed in the PSTF at the Mark III drywell peak cal-culated pressure of 36.5 psia; 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 presented in Attachment B is used to determine the hydrostatic pressure loads on h 090779
22A4365 6-3 Rev. 3
() the containment during an earthquake. During periods of horizontal accelerations there will be an asymmetric distribution around the circumference of the con-tainment. Also the DBA will initiate the suppression pool makeup system and the added pool water is included in the hydrostatic pressure calculations.
Figure 6-7 shows the water level transients in both the suppression pool and the drywell following the DBA.
6.1.5 Local Containment Loads Resulting from the Structures at or Near thq 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, then the upward loads are transmitted into the containment wall. Sections 9 and 10 discuss the types of loads that will be transmitted.
Localized loads on the containment wall resulting from the pressure losses 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 occurring.
6.1.6 Containment Load Due to Pool Swell at the HCU Floor (Wetwell Pressurization)
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-mately 1/2 second after top vent clearing and will generate both impingement 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 l structures attached to it and the flow pressure differential will result in an outward pressure loading on the containment wall at this location. The impingement loads will be 15 psi and the froth pressure drop across the HCU flocr has been calculated to be 11 psi; the containment wall will see an i Os 090779
22A4365 6-4 Rev. 2 11 psi discontinuous pressure loading at this elevation. Figure 6-9 shows details of the 11 psi pressure loading. The bases for both the impingement and flow pressure loading are discussed in Section 11 and 12.
When evaluating the containment response to the pressure differential at the HCU floor, any additional loads transmitted to the containment via HCU floor supports (beam seats, etc.) must be assumed to occur simultaneously. These loads 2
are based on the assumption that there is approximately 1500 ft of vent area reasonably distributed around the annulus at this elevation. For plant configu-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 Eack 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 g evidence of pressures higher than the initial static pressure.
Structures within the containment annulus below the HCU floor wi1J exoerience fall 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 ft/sec; typical drag coefficients are shown on Figure 10-5. This is the terminal 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 containment loading conditions.
O 042178
1 l
22A4365 Rev. 3 6-5
()
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 phase of the 1/3 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 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-
[')
~
ment are defined in subsection 4.1.9.2.
6.1.11 Long-Term Transient Following the blowdown, the Mark III containment system will experience a long term suppression pool temperature increase as a result of the continuing core decay heat. The operators will activate the RHR system to control the tem-4 perature increase, but there will be a period of containment pressurization before the transient is terminated. Peak calculated containment pressure is 9.8 psig (see Figure 6.14), and peak calculated suppression pool temperature l is 1730F. (With long term Containment Spray operation, the peak temperature can l
l 1
l 090779 1 l
22A4365 Rev. 2 approach 180*F.) The model used to simulate the long term post LOCA contain-ment heat up transient is described in supplement 1 to Reference 1.
1 6.1.12 Containment Environmental Envelope Figure 6.14 is a diagram showing the maximum calculated containment pressure and temperature envelope for any size of credible primary system rupture.
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, the 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 level, these loads are additive. Attachment A defines the loading magnitudes which are assumed for the S/R valve discharge.
g The seismic induced increase in suppression pool hydrostatic pressure as a result of horizontal accelerations is asymmetric. This loading sequence is discussed 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 than the loads associated with the intermediate or DBA break.
Figure 6.3 is the bar chart for this case.
There are unguarded RWCU lines in the containment that can release steam to the containment free space in the event of a rupture. The RWCU isolation 0
042178
f 22A4365 6-7 Rev. 2 O valves and flow 1Laiter for this system are designed to terminate the blow- f 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.
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.
The loads generated by S/R valve actuation are discussed in Attachment A.
)
6.5 SUPPRESSION POOL THERMAL STRATIFICATION 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 discussed in Attachment N. The momentary thermal stratification for large break accident used in containment evaluation is shown in Figure 6.17.
i
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O 101678
STRUCTURE- CONTAINMENT WALL ACCIDENT- LARGE STEAM LINE BRE AK IDB Al LOADS DUE TO SEISMIC ACCELERATION OF THE STRUCTURES AND LOADS DUE TO SEISMIC INDUCED POOL WAVES (ATTACHMENT Bn HVDROST ATIC PRESSURE NOTE POOL OUMP STARTS AT S m.n. ISECTION 61.4) gooop POOL TEMPERATURE (SECTION 6.1.11 AND 6 6) 150 F 180*F e i i
^ "^
SINGLE SIR VALVE ACTUATION SECTION 2.4 b
G 6 PRESSURIZATION OF CONTAINMENT z
WI H THE DATA i 14 CARRYOVER w o aN 4p WETWELL PRESSURIZATION O
LOAD BELOW HCU FLOOR SECTION 6.1.6,12.0 CHUGGING IkOION O
g t.a w LOADS DUE POOL SWELL AND FALL BACK LOADS SECTION 6.1.5 FOR A GIVEN STRUCTURE ARE NOT TO POOL SWELL ,
COINCIDENT, BOTH LOADS HAVE A SECTION 6.1.8 DUR ATION OF 0.5 sec. POOL SWELL CAN LOCA BUBBLE " OCWR 1 TO 1.5 sec AFTER BREAK PRESSURE LOAD POST LOCA WAVES DEPENDING ON HEIGHT ABOVE THE POOL. F ALL BACK LOADS OCCUR FALL 8ACK SECTION 6.17 1.5 TO S sec AFTER THE BREAK.
P E ENT DURING SECTION 6.1.2 VENT CLEARING.
COMPRESSIVE WAVE SECTION &1.1 CONDENSATION OSCILLATION LOADS SECT'9N 61.9 LOAD OUTWARD.
I I l i l 1.5 see 10 sec 5.0 see 10 sec 30 sec 100 sec 6 to be I see TIME AFTER EVENT 8
S Figure 6.1. Containment-Loading Chart for DBA i on
- 9 e - - - - - - -
O O O STRUCTURE : CONT AINMENT WALL ACC4 DENT. INTERMEDIATE STEAM LINE 8REAK tl8Al LOADS DUE TO SElsMIC ACCELERATION OF THE STRUCTURES AND LOADS DUE TO SEISMIC INDUCED POOL SURFACE WAVES (ATTACH NOTE; POOL DUMP STARTS AFTER ADS ISECTION 6.14)
HYDROST ATIC PR ESSUR E .
SINGLE SIR V ALVE ACTU ATION
, ATTACHMENT A.
SECTION 2.4 ADS ACTUATED a
8 5
6 AIR RETURN TO b ORYWELL (SEC 4.3.2) y ha POOL HEATUP RAISES CONTAINMENT TEMPERATURE AND PRESSURE TO E$ '
h CONTAINM NT = 3 g 5 peig. DRYWELL DIF FERENTI AL PRESSURE MAINTAINED AT 3 psid. ISECTIONS 6.2,6.5) .b 3 (SEC 43.2) i
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@ h SINGLE SRV LOADS DO NOT COMBINE WITH OTHER SRV LOADS
' TIME SCALE DEPENDENT UPON BREAK SIZE, MIN MUM VALUE OF t = 2 min O
U Figure 6.2. Containment-Loading Chart for IBA .
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STRUCTURE. CONT AINMENT WALL ACCIDENT: SMALL STEAM BRE AK LOADS DUE TO THE SEISMIC ACCELERATION OF THE STRUCTURES AND LOA.DS DUE TO SEISMIC INDUCED POOL SURFACE WAVES lATTACHMENT 88 HYDROSTATIC PRESSURE NOTE: POOL DUMP INCLti.9ED IAUTO AT 30 mini ISEC 61 di SINGLE S/H VALVE ACTUATION NOTE: DURING COOLDOWN WITH CONDENSER ISOLATED, SEC 2.4 (ATTACHMENT A)
S/R VALVES ARE OPERATED PERIODICALLY FOR UP TO THREE HOURS - -----
I CHUGGING: NOTE: CHUGGING CAN LAST UNTIL BREAK ISOLATED OR VESSEL !
DEPRESSURIZED (SEC 6.1.10,2.31 o - - - - - - _.J --
r m"
aY 4e 8 ORYWELL AIR CARRY OVER RAISES w m
f PHESSURE DIFFERENTIAL = 3 osed WU O
POOL HEATUP RAISES CONTAINMENT TEMPERATURE AND PRESSURE RISES TO 6 pse BECAUSE OF POOL HEATUP. DRYWELL PRESSURE DIFFERENTIAL MAINTAINED AT 3 psi. (SEC 4.3.2)
I i 3 he 6 he 1 8 men o TIM [ AFTER EVENT e @
Figure 6.3. Containment-Loading Chart for SBA .h O
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22A4365 Rev. 2 6-24 24 l k
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- =
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i 2 i o i 7 12 - TOP VENT CENTER LINE l
I l
8 -
a YINITIAL Y FINAL l
4 v
4 -
l l l BASEMAT l 100 120 140 160 180 I POOL TEMPE RATURE (*F) i Figure 6.17. Suppression Pool Te:::perature Profile for Large Breaks 0-101678
22A4365 7-1
._.. ._Rev. 2 .
- 7. S PPRESSION P'00L BASEMAT LOADS
]
l In addition to the normal, seismic, deadweight and hydrostatic pressure
{
loadings, that section of the basemat which forms the bottom of the suppression poo' also experiences dynamic LOCA loads ond 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 j I
21.8 psi at the dryvell wall - with the increase from 10 psi to 21.8 psi to be assumed linear and distributed over 50% of the pool width as indicated in I Figure 7.1. This specification is based on the observation that the maximum pressure that the initial bubble can ever have is the maximum drywell pressure during the accident. Data trace no. I shown on Figure 6.4 indicates that the i 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 between this point and the drywell wall will bound the actual pressure dis tribution. 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 dry. ell air carryover and the long term pressure and temperature 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.
l Safety / relief valve oscillating loads are defined in Attachment A. The net i 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.
O S
042178
+ e .
22A4365 1 R2v. 2 7-2 i
. i l
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NOTE: PRESSURES SHOWN 00 NOT INCLUDE HYDROSTATIC CONSIDERATIONS 21.8 pm 1
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BASEMAT RADIAL DIMENSION Figure 7.1. Pool Boundary Loads During Bubble Formation I
l I
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042178 l l
22A4365 Rev. 3 8-1
- 8. LOADS ON STRUCTURES IN THE SUPPRESSION POOL There are certain structures within the suppression pool which will l experience dynamic loads during both loss-of-coolant accidents and/or safety /
relief valve actuation.
- 8.1 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 creates j a dynamic load on structures submerged in the pool. However, this dynamic l 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 clearing, a single bubble is formed around each top vent. It is during the bubble growth i
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 methodology and calculation procedures for determining submerged structures drag loads are discussed in attachment G.2.3.
O 090779
22A4365 R;v. 3 8-2 1
l i
l Structures in the suppression pool should be designed conservatively for the LOCA drywell bubble pressure (see Figure 7.1) and acceleration drag gg (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 these structures will experience drag forces associated with water flowing at 35 ft/sec; that is the terminal velocity for a 20 ft free fall 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.
The methodology and calculation procedures for determining condensation loads on submerged structures are discussed in attachment G.2.5.
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.
The methodology and calculation procedures for determining chugging loads on submerged structures are discussed in attachment G.2.6.
O 090779
22A4365 j Rev. 3 8-3 O- 8.1.6 Compressive Wave Loading
- As discussed in Section 6.1.1, the very rapid compression of the drywell air theoretically generates a compressive wave. But as pointed out in Sections 6.1.1 and 6.1.2, there were no loads recorded on the containment 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.
! For structures near the containment, neither contressive or jet loads are significant.
8.1.7 Safety Relief Valve Actuation Loads on submerged structures due to safety relief valv,e actuation are discussed on Attachment G.
4 O
1 l
O 1
090779 I l
(
STRUCluRE. STRUCTURES WITHIN THE SUPPRES$10N POOL
POOL TEMPERATURE (SECTION 6.1.11 AND 65)
LOADS DUE TO SEISMIC ACCELERATION OF STPUCTURES AND LOADS DUE TO SEISMIC INDUCED POOL MOTION (ATTACHMENT Se l
HYDROST ATIC PR ESSUR E NOTE: POOL DUMP INCLUDED AT 5 men )
SINGLE S/R VALVE ACTUATION ISEC 2.48 SE TION 8 7 VENT CLEARING z WATER JET $[CTION 8st.1 h LOAD 6
DRA SECTION S.t.2 D
$ PJ 5
O FALL 8ACK SECTION S.I.3
[$
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LE CONDENSATION LOADS SECTION 8.14 PRESSURE LOAD CHUGGING SECTION 815 SECTION 8.1.2
- TYPICAL STRUCTURES ARE COMPR ESSIVE SECTION & I 6 3. S/R V ALVE LINES AND OUENCHER WAVE 2. ECCS SUCTION LINES 3 ECCS RETURN TO POOL LINES ITEST AND RELIEFI I I i l l i 1 1.5 3 6 30 100
~
- ADD S/R DYN AM;C LOAD TO STATIC LOAD DUE TO DRYWELL AIR PURGED TO CONTAINMENT VAPOR PRESSURE AT 1400F TIME AFTER EVENT. wc o
8 Figure 8-1. Structures within Suppression Pool-Loading Chart for DBA D
?
1 i
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22A4365 a Rev. 3 8-5 1 I
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090779
22A4365 Rsv. 3 8-6 0
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O 090779
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22A4365 Rev. 2 9,
( !
l
- 9. IDADS ON STRUCTURES AT THE POOL SURFACE i
Some structures have their lower surfaces either right at the suppression pool surf ace or slightly submerged. This location means that these struc-tures do not experience the high pool swell impact loads discussed in Section 10. However, they experience pool swell drag loads and LOCA induced bubble loads. Relief valve loads must also be considered. These are:
(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 exit 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
( 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 this 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 bubLle. 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.
O Pool fall back lo tds are as discussed in Section 4.1.6.
j 042178
STRUCTURE. STRUCTURFS AT THE POOL SURFACE ACCIDENT: LARGE STEAM LINE BRE AK ID8 Af POOL TEMPERATURE (SECTION 6.1.11, 6.51 LOADS DUE TO SEISMIC ACCELERATION OF THE STRUCTURES ANO DUE TO SERSMIC INDUCED POOL SURF ACE WAVES 4 ATTACHMENT 83 SINGLE S/R VALVE ACTUATION ADJACENT TO STRUCTURE (SEC 2.4) SECTION 9c AND ATTACHMENT A o LOCA BUBBLE PRES $URE LOAD SECTION 9 g
g
= ru
?C DP ' LOADS SECTaON 9 y$
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.a FALL 8ACK SECTION 4.l.6 POST LOCA SECTION 6.13 WAVES I t I t i l l5 3 6 30 TIME AFTER EVENT see o
o Figure 9-1. Structures at the Pool Surface-Loading Chart During DBA e u e e
O O -
O
l 22A4365 Rev. 2 _ lo_1 ;
- 10. LOADS ON STRUCIURES BETWEEN POOL SURFACE AND THE HCU FLOORS Equipment and platforms 1ocated in the containment annulus region, between the pool surface and the HCU platfcrm, 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 items 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.
O The remainder ef this sectien dea 1s with re1ative1F sma11 structures def1=ed 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 IMPACI 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 that 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 l
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 containment annulus, a water surface ve1ocity of 40 ft/see is assumed. Bulk pool swell would start I see after LOCA.
l l
101678 I
22A4365 Rev. 2 lo_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 f t elevation but below the HCU floors, 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 specifications 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 ef fect 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 042178
_ _ _ _ _ _ _ _ _ _ _ _ - _ _ _ A
4 22A4365 l I
Rev. 2 10-3 O
that any local increase in swell velocities will not result in loads in excess of design values.
The conservatism in these load definitions are illustrated in Attachment J.
10.2 DRAG LOADS In addition to the impact loads, structures that experience bulk pool swell are also subject to drag loads as the pool water flows past them with 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.
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042178 l
STRUCTURE SMALL STRUCTURES DETWEEN THE POOL SURF ACE AND THE HCU FLOORS ACCIDENT. LARGE STEAM LINE BRE AK LOe Al J
IMPACT LOAOS Tim 10.1 AND FIGURE 10.2 OH 10.6 F LOW (DR AG) LOADS SECTION 10.2 AND FIGURES 10.3.10.4 ANO 10.6 5
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NOTES:
e LOAO TO 8E APPLIED TO SOLIO AREA ONLY e SULK POOL SWE LL F LOW 10R AG) e IMPACT LOAOS ARE NOT SPECIPIED FOR GRATINGS 14 - o OURATION OF LOAD 0.5 SEC 5 12 -
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Q SOURCE: Chemicas Engmeer's Handbook.
g R. H. Pwry, p. 5-37 8
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O 0.5 06 0. 7 08 0.9 1.0 OPCN AREA FRACTION Figure 10-3. Pressure Drop Due to Flow Across Gracing Within 18 ft -
of the Pool Surface 042178
22A4365 Rev. 3 10-7 O '
l l
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l NOTES- 1. FOR DURATION, ASSUME STATIC LOAD
- 2. APPLIES TO FLAT SURFACES.
18 -
FOR OTHER MAMME FlWRE M -
i . 3. SOURCE: MARKS MECHANICAL ENGINEERS HANDBOOK, g 6th EDITION. PAGES 1182 J
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22A4365 Rev. 2 10-8 (REF: Flut0 MECHANICS, VICTOR L. STRE ETER, Lt3 ED. MC GRAW HILL)
^ * '" '"
oR AG COEFFICIENT
- PRESSURE PRES $URE 800Y SH APE O OlF F ERENTI AL (psil OlF F ERENTIAL losal 13 10 CIRCULAR CYLINDER 1.2 F LOW DIRECTION 0.6 2:1 7 5 E LLIPTICA L CYLINDER 0.32 4:1 4 3 ELLIPTICAL CYLINDER 0.29 8:1 3 2 E LLIPTICAL CYLINDER 2.0 22 17 SQUARE TR1 ANGLE 2.0 120* 22 17 T RI ANG LE 1.72 120* 19 14 TRIANGLE 2.15 90* 23 18 TRI ANGLE 1.60 90* 17 13 TRIANGLE 2.20 60* 24 18 TRI ANGLE I.39 60* 15 12 TAIANGLE 1.8 30* 19 15 TRIANGLE 1.0 30* 11 a SEMlTU8ULAR 2.3 h 25 19 SE Mt TUBULAR 1.12 12 9 4 (10 - 105p 4
- These drog coeffectents are conservative because ttiev are for low Reynot6s Nurnber flow conditions Use of lower values may be used of its applicabelity can be demonstrated.
Figure 10-5. Drag Loads for Various Geometries (slug flow) 042178
o o o l NOTE:
- 1. CURVE SC-D APPLifS TO HORIZONTAL RUNS OF PIPING
- 2. CURVE S A-E APPLIES TO SEAMS AND SMALL FLAT STRUCTURES FOR DURATION
- 3. SEE FIGURE 12.2 FOR HCU FLOOR SEE FIGURE 10.2 LOADS
- 4. SEE ATTACHMENT J AND FIGURE 47 FOR JUSTIFICATION E
A 115
{
l FOR DURATION OF APPLIED LOAD ,
e l SETWEEN IS AND 19 FEET, DETERMINE S SY LINEAR INTERPOLATION OF VALUES g l SHOWN ON FIGURES 10.2 AND 12.2 5 l 9 I i V D MN W s
.o . . . . . .W .
.ge "y l NOTE 1
y l[ ONLY DRAG LOADS # ME APPL.IED ABOVE Tits ilCU FLOOR l {* FROM VELOCITY DETERMINEO SY DECELERATION WITH ELEVA-I T80N. NO FROTH IMPACT. NO DRAG LOAD ABOVE 30 ft.
30 -
I FOR DURATION l -
l @SEE FIGURE 12-2 gs 16 s=1 l
l 1
, I l la le i f HEIGHT F ROM POOL SURF ACE liti e U ?
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$ Figure 10-6. Summary of Pool Swell Loading Specifications for Small Structures in the Containment Annulus (Not Applicable to the Steam Tunnel or Expansive !!CU Floors)
- 4,
22A4365 R;v. 2 11-1 O 11. toios o xx>i stvz st uciuxes ir r r acu vtoo =tzvir1o=
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 ft2vent area attthis elevation (see Sec-tion 6.1.6). Figure 11-1 shows the loa' ding 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.
The 11 psi froth flow pressure differential lasting for 3 see 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 3
mixture of the top 9 f t of the suppression pool (i.e. ,18.8 lbm/f t ). Supple-ment 1 to Reference 1 describes the analytical model used to simulate the HCU floor flow pressure differential 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 appr'oximately 10 lbm/ft 3 Reference 11 indicates the HCU floor pressure differential is in the 3 to 5 psi range.
O 042178 i
22A4365 R;v. 2 u-2 The potential for circumferential variations in the pressure transient in the wetvell ragion 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).
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I 22A4365 12-1 Rev. 3
- 12. LOADS ON SMALL STRUCTURES AT AND ABOVE THE HCU FLOOR ELEVATION i l
1 Structures at the HCU floor elevation experience " froth" pool swell which involves both impingement and drag type forces. Figure 12.1 shows the loading 4
sequences.
l I
4 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 dynamic impingement load; grating structures are not subjected to this impingement load (Reference 12).
\s 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.
4 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,/ft ). This is a conservative density assumption confirmed by the PSTF 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 pai 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 effects of gravity. The velocity of 50 f t/see is a bound of the available data (Reference 13). No pool swell is assumed for structures more than 30 ft above the suppression pool.
090779
m ..M'-"*"
22A4365 12-2 Rev. 2 The potential for circumferential variations in the pressure transient in the I 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 paid. (See Attachment F.) I 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.
Supplemant 1 to Reference 1 describes the analytical model used to simulate the HCU floor flow pressure differential and presents a comparison of model predictions with test data. The model is shown to be conservative.
O O
042178
O O O 4 STRUCTURE STRUCTURES AT THE HCU FLOOR ELEVATION ACCIDENT LARGE STE AM LINE BREAK (DBAl DOWNWARD LOADS DUE TO F ALL B ACK SECTION 12
, AND WATER ACCUMULATION (ON HCU FLOORI J
g FLOW (DR AGI TWE LOADS (UPWARD) FIGURE 12.2 g
- - g 5 (TWOPHASE FLOW THROUGH HCU FLOOR) 5 .E C g w$
5 sa F ROTH IMPlNGEMENT LO ADS (UPW AR D) l l I 1.5 1.6 5 i
TIME AFTER EVENT. sec l o
o a Small Structures at the llCU Floor Elevation - Loading Chart During DBA Figure 12.1. ,
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FROTH IMPlNGEMENT (NOTE 21 15 -
r CA.LCULATED FROTH g TWO PHASE FL OW dP c
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NOTE
- 1. DATA BASED ON HCU FLOOR LOCATED APPet0X4MATELY 20 f ABOVE POOL SURFACE 4HWLi 2 REF TEST SERIES 10ti l
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. 15 16 20 30 40 50 55
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T IM E. wc 5 Figure 12.2. Loads at ilCU Floor Elevation Due to Pool Swell Froth Impact and Two-Phase Flow g Y
0 9 O
22A4365 R-1 Rev. 2 O
REFERENCES NOT ALL THE REFERENCES APPEAR IN THE TEXT. THE FIRST 11 REFERENCES REPRESENT A COMPREHENSIVE BIBLIOGRAPHY OF REPORTS RELATED TO GE'S PSTF PROGPM.
- 1. Bilanin, W. J., The General Electric Mark III Pressure Suppression 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. NEDO-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, NEDO-20210, December 1973 (Proprietary Report).
j 6. Fourth Quarterly Progress Report: Mark III Confirmatory Test Program, i NEDO-20345, April 1974 Supplement 1 (Proprietary Report).
1
- 7. Fifth Quarterly Progress Report
- Mark III Confirmatory Test Program, s- NED0-20550, July 1974 Supplement 1 (Proprietary Report).
- 8. Sixth Quarterly Progress Report: October 1974 (Letter Transmittal to NRC Staf f.) (Proprietary Data Attached.)
l
- 9. Seventh Quarterly Progress Report: Mark III Confirmatory Test Program, NEDO-20732-P, December 1974 (Proprietary Report).
- 10. Eighth Quarterly Progress Report: Mark III Confirmatory Test Program, 4 NEDo-20853-P, April 1975 (Proprietary Report).
t
- 11. Mark III Confirmatory Test Program 1/3 Scale Three Vent Tests, NEDO-13407, April 1975 (Proprietary Report).
j 12. Mark III Confirmatory Test Program 1/3 Scale Pool Swell Impact Tests -
Test Series 5805, NEDE-13426-P, August 1975 (Proprietary Report).
- 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 Centainment 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 i (Proprietary Report).
101678
1 l
22A4365 R;v. 3 R-2
- 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)
- 9 0
090770
A-1 22A4365
]b .
Rev. 3 ATTACHMENT A SAFETY RELIEF VALVE LOADS (QUENCHER)
Pjyle A
1.0 INTRODUCTION
A-5 A2.0
SUMMARY
& CONCLUSIONS A-6 A
3.0 DESCRIPTION
OF PHENOMENA A-7 A4.0 ARRANGEMENT A-10 A4.1 Distribution in Pool (Quencher Arrangement) A-10 A4.2 SRVDL Line Routing A-ll A4.2.1 Line Lengths and Volume A-ll A4.2.2 Drywell Penetration Sleeve A-12 A4.2.3 SRVDL Vacuum Bresker A-13 AS.O QUENCHER LOADS ON POOL BOUNDARY A-28 AS.1 Pressures on Drywell, Basemat, and Containment A-28
() A5.1.1 Single S/R Valve Loads A5.1.2 Two Adjacent S/R Valve Loads A-29 A-29 AS.I.3 fen 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 A5.3 Loads on Submerged Structures A-30 A5.4 Normalized Pressure - Time History (Theoretical Raleigh Bubble) A-33 A5.5 Representatite Pressure Time History A-33 AS.6 Estimated Margins A-33 A5.6.1 Peak Bubble Pressures A-33 AS.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 Pool 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 Loads by A.E.
A-66 A-66 A6.4 Seismic Slosh Loads by A.E. A-66 090779
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22A4365 Rev. 3 1
P_jyyt 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 Ceometry 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 A8.2 SSE and OBE Considerations A-83 A8.3 LOCA Considerations A-83 A8.3.1 DBA with M.S. Line Break A-84 A8.3.2 DBA with Recirculation Line Break A-84 AB.4 Recommended Design Load Summation A-85 A9.0 FATIGUE CYCLES A-87 A10.0 RECOMMENDED CALCULATION PROCEDURES FOR MARK III USERS A-90 A10.1 Constraints A-90 A10.2 Determine SRVDL Design A-91 l 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 A10.3.2 liow to Find the Attenuated Pressure on the Drywell Wall, Basemat and Containment Wall A-94 All.0 PARAMETRIC STUDIES A-104 A12.0 BASIS AND JUSTIFICATION FOR DEVELOPMENT QUENCHER LOADS A-106 A12.1 Introduction A-106 A12.2 Test Data Application for Mark III Containment A-107 A12.2.1 Miniscale Test Observations A-107 A12.2.2 Small-Scale Test Observations A-107 A12.2.3 Large-Scale Test Observations A-107 A12.3 Physical Parameters A-108 090779
j _._.___ .
3 1
i A-3/A-4 1 22A4365
!O t
I i
i P,, age,
}
1 I A12.4 Correlation of Positive and Negative Pressure Peaks A-ll8
]
i A12.5 Development of Design Value Calculation Method A-140 A12.6 Application A-173 A12.7 References (for A12) A-183 f
i l
i 1
i l
I 4
I 090779
22A4365 A-5 Rev. 3 A
1.0 INTRODUCTION
General Electric has determined that the. quencher is a desirable alternative feature to minimize suppression pool boundary loads resulting from the air clearing phenomena in the Safety Relief Valve Discharge Lina (SRVDL) .
The quencher device will be specified for the standard 238 Mark Ill design and is recomended for BWR-6, Mark III application.
This attachment provides the following:
- a. Recomended quencher arrangement.
- b. Recommended 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 valve 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
l 22A4365 R';v . 3 A-6 l
A2.0
SUMMARY
AND CONCLUSIONS I 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 . With 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 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.
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.
Table A4.2 summarizes the SR/DL design requirements and objectives neces-sary to obtain the S/R valve pressure loads for the 238 Mark III containment identified in this attachment.
O 090779
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22A4365 A-7 -
.Rev,. 2 O .
A
3.0 DESCRIPTION
OF THE PHENOMENA Prior to the lifting of a pressure relief valve, the downstream piping between the S/RV discharge and the water surface is filled with air at drywell pressura 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 lifts, the effluent reactor steam causes a rapid pressure build up 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 pr.ak as the last of the water is expelled. The compressed cushion of air between the water slug and the effluent vapor exits the quencher and forms four clouds of small bubbles 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 surf ace.
The magnitude of the pressure disturbance in the suppression pool decreases with increasing distance from the point of discharge, resulting in a damped oscillatory load at every point on structures below the water surface.
042178
~ .. . _ . -
22A4365 Rev. 2 A-8 -
From an air-clearing standpoint, a decrease in the volume of air initially 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.
Figu l.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
- Based on back pressure specifications to which valves sre purchased l
e-l 101678 l
O O O 60 55 -
PPIPE < 625 psed n-- 50 -
NOTE. NOT TO 8E EXTRAPOLATED E OR IN TE RPOLATE D y S/R VALVE SET PRESSURE
- 1287 psui 3 PP4PE > 625 pamt S/R VALVE FLOW RATE = 317.9 as/sec .xs ta VALID ONLY FOR: 0.33 4 C 4 6.0 k 4
WHE RE C = 10 in. P PE LENGTH 1 AIR) 'y 4b - 12 en. PIPE LENGTH 1 AIR) mg WATER LEG = 181:
VALVE OPENING TIME A 0.02 sac 40 -
l l l l l l l l 1 33 0 1 2 3 4 5 6 7 8 9 to F L/D t to .n SCHED 40p o
V U
M Figure A3.1, SRVDL Air Voltune Versus fL/D with 625 paid Constraint T
- 4
22A4365 A-10 ,
Bev, 2 A4. 0 ARRANGEMENT A4.1 DISTRIBUTION IN POOL (QUENCHER ARRMiGEMENT)
Figures A4.1 and A4.2 show the elevation and plan views of the standard quenchar 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 ft.
This arrangement meets the following objectives :
- 1. Minimize drywell structural interference.
- 2. Permit water circulation through top and bottom of the drywell sleeva 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 option. An advantage to the side anchor arrangement is that it eliminates containment liner penetration for anchor requirements.
Figure A4.3 shows the recommended quencher azimuthal locations in the standard 238 pool. As shown in this figure the low, intermediate and high pressure-switch set valves are uniformly distributed around the pool to preclude concurrent adjacent valves operation.
042118 g.
22A4365 pyy Rev. 3
( ,
y 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:
l l
( Standard Plant 238 6.5 ft above basemat i
Standard Plant 218 5.5 f t above basemat
~
Standard Plant 251 5.0 ft above basemat A4.2 SRVDL ROUTING nV The SRVDL is routed by the Architectural Engineer from the first pipe i
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. The SRVDL should have a suf ficient slope in the air l leg section routing to prevent condensation accumulation in the line.
Figure A10.2 is a typical :syout 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.
O LJ 090779
__ _ J
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 A11.0.
The SRVDL from 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 at 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 vent. The sleeve may be extended 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 from the ASME 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 increase'd temperatures 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
22A4365 A 3 Rev.,2- ;
local areas are allowed to reach 650*F from steam or water jets in 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 (SRVDL1 is at the normal suppression pool level. After the S/R valve closes, the steam rammining in the line 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 wil.: occur if the SRV opens when the water is above the normal pool levd. The purpose of the discharge line vacuum breakers is to prevent the water from rising substantially above its normal level when a subse-quant 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 located in the drywell above the expected level of water rise in the line subsequent to SRV closure. This eliminates the possibility of watwell pressurization in the event of a stuck open vacuum breaker and ensures proper functioning of the vacuum breaker.
The following parameters will yield satisfactory performance for most SRVDL geometries and is recommended to satisfy the above requirements. However, plant specific analysis for vacuum breaker design should be performed by the design engineer to confirm this.
- a. The vacuum breaker effective area, (A/ d)* is equal to or greater o
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 applied across it.
I
- c. The minimum opening differential pressure to start the vacuum breaker to open is equal to or less than 0.2 PSID. 042178
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22A4365 A-14 Rev. 2 l .
5
- d. The vacuum breaker must be fully open when pressure difference is equal to or less than 0.5 PSID.
- e. The vacuum breaker should be located in the drywell at an elevation above the m.wimum water level rise in the line following a SRV closure.
- L is used to calculate flow through the vacuum breaker as follows:
A w = /aP (2pg*) (144) b R
where w = Flowrate through vacuum breaker in ibm /sec AP = Pressure differential across the vacuum breaker (PSID) 9 o = Air or steam density in ihm/ft lbm - ft.
ge"
- 2 lb f - s ee
^
= Effective area of valve in ft
'[
042178 g,
II i l 4 22A4365 A-15
! Rev. 2 i'
g
!. Table A4.1 i QUENCHER ARRANGEMENT i
Mark III Plants S/R Valve Location Quencher Elevation / Plan View
}
i 238-732 STD. Figure A4.3 Figure A4.1/ Figure A4.2
, 238-615 Figure A4.4 Figure A4.1/ *
) 218-592 Figure A4.5 Figure A4.6/
- 251-784 Figure A4.7 Figure A4.8/
- i 4
251-848 Figure A4.9 Figure A4.8/
- i i
- O
- Typical plan view similar to Figure A4.2.
4 s
l l
042178 O
22A4365 A-16 Rev. 3 i O
Table A4.2 SRVDL DESIGN REQUIREMENTS AND OBJECTIVES SRVDL DESIGN REQUIRD1ENTS (a) Maximum SRVDL Pipe Pressure 1 625 psid. (Coordinares 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 1 0.02 sec.
- 3. :linimize the SRVDL air leg volume.
- 4. Minimize length of longest SRVDL.
- 5. Minimize the contribution of fL/D to the first half of the discharge line.
- 6. Start 12" S/40 or 14" S/40 pipe just below the first anchor point to meet objective (5).
- 7. The ratio of the air legs (length of 10" S/40 pipe / length of 12" S/40 pipe = C) should be 0.33 1 C 1 5.0.
- 8. Slope lines down toward pool to avoid condensate-water accumu-lation in line (no horizontal runs) .
- 9. SRVDL vacuum breakers should be 10" size. One 1 10 ft. above the weir wall and the other just below the seismic restraint at the SRV.
090779
22A4365 Rev. 3 A-17 f) s_/
Table A.4.3 SRVDL MARK III 238 STANDARD PLANT Air Leg Length Max. fL/D Volume S/R Valve Total Length 10" S/40 12" S/40 14" S/40 (ft3) (a) (b)
V-1 79'-8" 30'-5" 49'-3" - 54.9 2.09 4.21 4
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 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 V-8 77'-2" 19'-11" 57'-3" -
55.4 2.31 4.65 V-9 77'-1" 19'-7" 57'-6" - 55.4 2.31 4.65
() V-10 V-ll 77'-5" 76'-11" 20'-1" 19'-4" 57'-4" 57'-7" 55.55 2.31 4.65 55.34 2.31 4.65 i 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 V-16 77'-4" 20'-4" 57'-0" - 55.5 2.31 4.65 V-17 72'-9" 32'-6" 3'-9" 36'-6" 55.0 2.22 4.4) l V-18 79'-5" 28'-7" 50'-10" - 55.16 2.27 4.57 V-19 81'-0" 33'-5" 47'-7" -
55.3 2.27 4.57 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 piping arrangement). (These line designs have not been optimized to
() take advantage of the maximum pipe pressure of 625 psid) .
090779
Table A.4.4 QUENC!lLit BUBBLE PRESSURE MARK III, 238 STANDARD PLANT 95-95% CONFIDENCE LEVEL Design Value-Bottom " "'"**"
Maximum Pressure (psid) Containment
,, (-) @ 1O (ps d)"
p (+) Normalized Factor Case Description B 'B @ Point 10a p, p_
Single Valve First Actuation, 13.5 -8.1 0.711 9.6 -5.8 at 100 F Pool Tenperature Single Valve Subsequent 28.2 -12.0 0.711 20.1 -8.5 Actuation, at 1200F Pool Tempe rature Two Adjacent Valves First 13.5 -8.1 0.856 11.6 -6.9 Actuation at 1000F Pool $'R$
Tempe rature 4$
ro *$
10 Valves (one Low Set and 16.7 -9.3 0.916 15.3 -8.5 Nine Next Level Low Set)
First Actuation at 1000F Pool Temperature 19 Valves (All Valve Case) 18.6 -9.9 1.0 18.6 -9.9 First Actuation, at 1000F Pool Temperature 8 ADS Valves First Actuation 17.4 -10.4 0.821 14.3 -8.5 at 1200F Pool Temperature
" Point 10 on Containments is Peak Pressure.
S M
t O O O
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22A4365 A-19 Rev. 2 -
S/R SLEEVE 3.0 f $ f CJ CJ ORYWELL WALL
/ / CJ f)
/,/- VEN LE {
8
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[\ {\
/ !((# SECTION A-A h
_- Aw-o A -PIPE SLEEVE & SUPPonT
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\ \ W .P .
y =
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- 2. :: .
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- A 11a P3 8 i V4 _ _b, &. 5-t T4 _ 4 , ,
9 " "
- 1
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+ 1.s r +
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Figure A4.1. 238 Standard Mark III Quencher Arrangenent Elevation 042178
22A4365 A-20 .-
Rev.[2 .
t 60
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s 520*
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opywett WA" / -
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Figure A4. 2. Mark III Quencher Plan view (Typ.) 238 Plant Arrangement e-Shown 042U 8
22A4365 A-21 .
Rev.' 2 00 8
V9 (346.5*) V10 (4.5 )
F047H 190) VII (22.5 8) 111311180) F041E - MPL IDENT V8 (328.5*) 1123 - PMESSURE SWITCH SET POINT fpsql F0410 I (1165) - SPRING SET POINT (psg) 112341165) 14%
t l V7 (310.5*)
F0470 / \
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1 1113 11190) 1113(1180) O A
, C
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8 V6 (292.5 ) V13 (67.5*)
11661 s! N VS Vg V10 V11 V14 \ F041A 1123(1165)
[ V4 V7 VS V12 V13 16 V3 V17 185 5*)
7 [ ,
M13(11801 - b V2 V18 -+- 1113 (1180) 270a l-. h- -- -.
1 I e
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V1 V19 8
V4 (256.5 ) RPV V 151103.5"i 4
O F041F s*% F041G Cj 1123t1165)
[ 112311165)
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/ ORYWELL WALL v. V16 (121.5")
F051G V31238.5*) \ 1113 (11901 F0518 1913 (1190) N /
x N,hk /
\
/ \ V171139.5*)
F047C 1123111801
-+ V18 (157.5*)
1
/ l F051C 1123(11651 V1 (193.5*) " '
V19(175.5*)
F0478 F041C 1113 (11801 1123(1165) 1800 LEGEN0:
AOS - O NOTES: 19 S/R VALVES Figure A4.3. S/R Valve Discharge Locations for 238-748 Standard Plant 042178
22A4365 Rsv. 2 A-22 . ..-
O' h.-
V10 (4.5*)
V9 (346.5*) F0510 V11 (22.5*)
F0470 1103(11903 p047E - MPL 10EP.T 1113 (11801 1113- PRESSURE SWITCH SET POINT (pose)
I (1180) - SPRING SET POINT (poeg) 1- 4
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g ggNNWN V12149.5*)
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V7 (310.5*) O A \ / 1123 (1165)
' F041D 8 C 'M 1123 (1195)
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o V6 (292.5 ) V13 (67.5*)
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I 1 270*
- 90*
V1 RPV l V41256.5*) J F041F t 123 (1165)
\
~ ORYWELL Wall-
- V3 (238.5*) , /g/ V16(121.5*)
F0478 \ ,g, 111341180) j
/
\- .
/ \ V17(139.5*)
11658 v21211.5*)
F0418
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1123(1165) F047C V1 t193.5*) V19 (175.5*) 1113(1180)
F05is F04tC 1113111901 1123(1166)
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10*
LEGENO: AOS
- NOTE: 16 S/R val.VES l
l Figure A4.4. S/R Valve Discharge Locations for 238-648 Standard Plant l 101678 -
1 4
I
22A4365 A-23
~
Rev.,2 08 V9 I5.3')
v4 (344.1'l F0610 3 11801 vio (26.5')
1103(11901 F041 A - WPL IDENT 1123 - PRESSURE SWITCH SET *OINT fotgl (11664 - $PR!NG SET POINT Ipsegl v7 (322.9'l F047D #p- g N 4
/
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ve (30s.s')
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1113 (19901 a y
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l vs treo.s'i po47p ,
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y, vt6 v4(259.4'l
- ~: g F041F 4 \ f v13 (100.6*)
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- NOTES. to 3/R VALVES O ris re ^' s- $'*'"t'""'"'"""" '-' ' """
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042178
22A4365 Rev. 2 A-25 .
00 V10 (366') V11 (12*l F0810 po47A )
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LEGENO:
AOS
- O NOTES:
- 1. (201 S/M VALVES i
Figure A4.7. S/R Valve Discharge Locations for 251-800 Plant O
l 101673 i
A-26 -
22A4365 -
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l 1800 l LEGENO: ADS
- O NOTES: 1. 22 S/A VALVES Figure A4.9. S/R Valve Discharge Locations for 251-848 P1 mt O
101678
A-28 32A4365 Rey. 2 O.
A5.0 QUENCHER LOAD ON POOL BOUNDARY ;
1 1
AS.1 PRESSURES ON DRYWELL, BASEMAT AND CONTAINMENT Dryws11 wall, basemat and S/R Valve Loads are calculated as discussed in Section A10.0. For the 238 Standard Plant, the m nfmum and minimum bubble pressure below the quencher just after air clearing are shown in Table A4.4. actuations (psid) .
The absolute pressure on the pool walls can be calculated by the following equation:
+ +0 (#}
(a) " containment 44 where:
P(,) = Absolute pressure at point (a) (psia) r = Distance from center of quencher to point (a) (ft)
P eontainment
= Absolute pressure of containment atmosphere (psia) h(a) = Head of water act1*.g at point (a) (ft) a = Density of pool watar % 62.4 Clb/f t ).
.1P g = Bubbla pressure attenuated by distance, r : point (a),
for rzultiple S/R vabe actuations (psid) .
The pressure decays with time and this is discussed in Section A5.4.
042178 O-
i 22A4365 A-29 Q,m Rev. 3 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 PBmax 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 Bmax = 25 psid for another SRVDL layout, the normalized values of Tables AS.1 through AS.5 would be multiplied by 25 to obtain the design pressures.
A5.1.1 Single S/R Valve Loads The normalized dynamic peak pressures AP g for a single S/R Valve Discharge valve are given in Table A5.1 and the normalized radial and circumfsrential peak values are shown in Figures AS.1, AS.2, and AS.2a.
(The valuzs given presume an air leg volume of 56.13 f t for all SRVDS's).
h (V
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 Loads The normalized dynamic peak pressures AP (r) are given in Table AS.2 and the normalized radial and circumferential peak values are plotted in Figures AS.3, AS.4, and AS.4a for the two adj acent S/R Valves V-8 and V-9.
A5.1.3 Ten S/R Valve loads
- lormalized AP (r) laods are given in Table AS.3 and the normalized values are shown in Figures AS.5, A5.6, and AS.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
22A6365 A-30 Rev. 3 O
A5.1.4 Eight S/R Valve Loads (ADS)
Normalized AP( ) loads are given in Table AS.4 and the normalized values are shown in Figures AS.7, AS.8, and AS.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 shown in Figures A5.9, A5.10, and AS.10a for all (19) valves V-1 to V-19.
A5.2 LOAD ON WEIR WALL The S/R valve loads on the weir wall are the same as those on the drywell wall except they only act on the projected area through the drywell wall vents.
AS.3 LOADS ON SUBMERGED STRUCTURE For submerged structures, the loads are specified in Section G3 of Attachment G.
O 090779
i J
t 22A4365 A-31 Rev. 3 A-32 J i6 4
1 i
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J i
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l i
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1 i
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J THESE PAGES ARE INTENTIONALLY DELETED I
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090779
I 22A4365 '
Is Rev. 2 A-33 L)
AS.4 NORMALIZED PRESSURE TIME HISTORY (Theoretical Raleigh Bubble)
The ideal pressure is normalized for the maximum AP(r) , ritive value as shown in Figure AS.11. The frequency is 5 to 12 Hz as derived from the test l 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 A5.11 is used by the designer for determining pressure amplitudes 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.
AS.5 REPRESENTATIVE PRESSURE TIME HISTORY Figure AS.12 depicts a representative pressure time history at points P1 through P4 asshown on Figure A4.1. These curves provide the designer a realistic picture of the pressure oscillations as opposed to the idealized Raleigh bubbles.
AS.6 ESTIMATED MARGINS l
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.
O V
l l 101678
22A4365 Rev. 3 A-34 O
Generalized Bottom Pressure Load Case
- A B C D e
- 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 6.3/2.2 5.9/2.5 12.1/2.9
- 4. % Margin (Based 35/23 34/22 34/24 43/24 on Predicted Maximum Bubble Pressure)
"See Section A12.5.1 for load case description.
A5.6.2 95%-95% Confidence 95%-9.i% menns that there is 95% confidence that 95% of any new data obtained will fall within the maximum levels of the current data base.
See Section A12. 5.1.2 for additional discussion.
A5.6.3 Margin The apparent margin in the specified containment design based on quencher bubble pressure is :alculated as 20 to 45%.
090779
O O o Table A5.1 MARK Ill 238-732 STANDARD Pl. ANT DYNAMIC PRESSURE FIELD FOR ONE S/R VALVE TIME - 0.15 sec (Positive Pressure psid) AP (r) 0.08 sec (Negative Pressure paid) AP (r)
V-10 S/R Valves Angle (degrees) 4.5 13.5 Reference Point 283.5 292.5 301.5 310.5 319.5 328.5 337.5 346.5 355.5 0 0 0 0 0 0 0 0 ,
1 2 0 0.274 0.334 0.423 0.566 0.805 0.984 0.805 0 0.280 0.345 0.449 0.632 1.0 1.0 1.0 3
0 0.282 0.348 0.453 0.645 1.0 1.0 1.0 4 us w 5 0 0 0.277 0.339 0.435 0.598 0.902 1.0 0.902 .} g 0 0.198 0.227 0.268 0.328 0.427 0.605 0.994 1.0 0.994 m 6
0.168 0.188 0.216 0.254 0.311 0.406 0.566 0.902 1.0 0.902 7
8 0 0.159 0.178 0.203 0.239 0.290 0.372 0.494 0.716 0.885 0.716 9 0.137 0.151 0.169 0.192 0.224 0.269 0.337 0.411 0.563 0.645 0.563 10 0.138 0.152 0.170 0.194 0.227 0.274 0.345 0.449 0.605 0.711 0.605 11 0.138 0.152 0.169 0.193 0.225 0.272 0.343 0.445 0.594 0.691 0.594 12 0.137 0.151 0.168 0.191 0.222 0.266 0.331 0.420 0.535 0.605 0.535 0 0 0 0 0 0 0 0 0 0 13 0
?
is a y 0
9
Table AS.2 MARK III 238-732 STANDARD PLANT DY!LVilC PEAK PRESSURE FIELI) FOR TWO AIU ACENT S/R VALVF.S TIME - 0.15 see (Positive Pressure) AP (r)
= 0.08 sec (Negative Pressure) AP (r)
S/R Valves V-9 V-10 Angle (Degrees)
Refe rence Point 265.5 274.5 283.5 292.5 301.5 310.5 319.5 328.5 337.5 346.5 355.5 4.5 1 0 0 0 0 0 0 0 0 2 0.274 0.334 0.504 0.657 0.910 1,0 1.0 1.0 3 0.280 0.345 0.529 0.721 1.0 1.0 1.0 1.0 4 0.282 0.348 0.533 0.733 1.0 1.0 1.0 1.0 5 0 0 0.277 0.339 0.515 0.687 1.0 1.0 1.0 1.0 6 0 0 0.198 0.227 0.333 0.399 0.504 0.688 1.0 1.0 1.0 1.0 .f (
7 0 0.168 0.188 0.274 0.316 0.379 0.479 0.646 0.989 1.0 1.0 1.0 ru h 8 0 0.159 0.178 0.258 0.298 0.354 0.442 0.573 0.807 1.0 1.0 1.0 9 0.137 0.151 0.218 0.245 0.280 0.331 0.405 0.508 0.656 0.766 0.796 0.776 10 0.138 0.152 0.219 0.247 0.283 0.335 0.413 0.526 0.697 0.841 0.856 0.841 11 0.138 0.152 0.218 0.246 0.282 0.334 0.410 0.521 0.686 0.822 0.840 0.822 12 0.137 0.151 0.216 0.244 0.278 0.328 0.399 0.497 0.629 0.736 0.757 0.736 13 0 0 0 0 0 0 0 0 0 0 0 0
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T 9 O 9
O O O Table AS.3 MARK III 238-732 STANDARD Pl. ANT DYNAMIC PEAK PRESSURE FIEI.D EDR 10 S/R VALVES TIME = 0.15 sec (Positive Pressure psid) AP (r)
= 0.08 sec (Negative Pressure psid) AP (r)
S/R Valves V-10 V-12 V-14 Angle (Degrees)
Reference Point 4.5 13.5 22.5 31.5 40.5 49.5 58.5 67.5 76.5 85.5 94.5 103.5 0 0 0 0 0 0 0 0 0 0 0 0 .
1 2 1.0 0.969 0.782 0.758 0.913 1.0 0.910 0.801 0.950 1.0 0.950 0.801 3 1.0 1.0 0.849 0.825 1.0 1.0 1.0 0.894 1.0 1.0 1.0 0.894 ,
4 1.0 1.0 0.862 0.837 1.0 1.0 1.0 0.912 1.0 1.0 1.0 0.912 yN
< Y 5 1.0 1.0 0.813 0.789 1.0 1.0 1.0 0.845 1.0 1.0 1.0 0.845 '. *;
6 1.0 1.0 0.834 0.820 1.0 1.0 1.0 0.907 1.0 1.0 1.0 0.914 "O 7 1.0 1.0 0.8 04 0.772 1.0 1.0 1.0 0.850 1.0 1.0 1.0 0.857 8 1.0 0.890 0.724 0.694 0.834 0.991 0.866 0.750 0.857 0.989 0.360 0,756 9 0.841 0.745 0.651 0.637 0.704 0.768 0.720 0.675 0.723 0.776 0.714 0.667 10 0.903 0.782 0.669 0.655 0.741 0.827 0.758 0.699 0.761 0.835 0.752 0.692 11 0.884 0.772 0.664 0.650 0.731 0.809 0.748 0.693 0.750 0.817 0.742 0.686 12 0.804 0.720 0.638 0.625 0.680 0.733 0.695 0.660 0.698 0.741 0.688 0.652 13 0 0 0 0 0 0 0 0 0 0 0 0
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O w
4 l M l
Table A5.3 (Continued)
S/R Valves V-16 V-10 V-1 Angle (Degrees)
Re fe rence Potat 112.5 121.5 130.5 139.5 148.5 157.5 166.5 175.5 184.5 193.5 202.5 211.5 1 0 0 0 0 0 0 0 0 0 0 0 0 2 0.950 1.0 0.950 0.801 0.950 1.0 0.950 0.801 0.910 1.0 0.913 0.707 3 1.0 1.0 1.0 0.894 1.0 1.0 1.0 0.894 1.0 1.0 1.0 0.776 4 1.0 1.0 1.0 0.912 1.0 1.0 1.0 0.912 1.0 1.0 1.0 0.788 5 1.0 1.0 1.0 0.845 1.0 1.0 1.0 0.845 1.0 1.0 1.0 0.739 6 1.0 1.0 1.0 0.914 1.0 1.0 1.0 0.907 1.0 1.0 1.0 0.800 7 1.0 1.0 1.0 0.857 1.0 1.0 1.0 0.850 1.0 1.0 1.0 0.753 8 0.860 1.0 0.860 0.756 0.860 0.989 0.857 0.750 0.851 0.975 0.824 0.675 _
9 0.727 0.779 0.727 0.667 0.714 0.776 0.723 0.661 0.704 0.762 0.694 0.604 'gm 10 0.764 0.838 0.764 0.692 0.752 0.835 0.761 0.685 0.743 0.821 0.731 0.622 21 0.754 0.819 0.754 0.686 0.742 0.817 0.750 0.679 0.732 0.803 0.721 0.618 12 0.702 0.744 0.702 0.652 0.688 0.741 0.698 0.646 0.679 0.726 0.669 0.592 s
13 0 0 0 0 0 0 0 0 0 0 0 0 3
0 t
u 9 .
O O O Table A5.3 (Continued)
S/R Valves Angle (Degrees) V3 V5 V7 V9 Re fe rence Point ,, 220.5 229.5 238.5 247.5 256.5 265.5 274.5 283.5 292.5 301.5 310.5 319.5 328.5 337.5 346.5 355.5 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 1
2 0.707 0.913 1.0 0.910 0.801 0.950 1.0 0.950 0.801 0.950 1.0 0.988 0.868 1.0 1.0 1.0 3 0.776 1.0 1.0 1.0 0.894 1.0 1.0 1.0 0.894 1.0 1.0 1.0 0.959 1.0 1.0 1. 0 -
4 0.778 1.0 1.0 1.0 0.912 1.0 1.0 1.0 0.912 1.0 1.0 1.0 0.976 1.0, 1.0 1.0 5 0.739 1.0 1.0 1.0 0.845 1.0 1.0 1.0 0.845 1.0 1.0 1.0 9.911 1.0 1.0 1.0 6 0.800 1.0 1.0 1.0 0.907 1.0 1.0 1.0 0.914 1.0 1.0 1.0 0.944 1.0 1.0 1.0 38N 7 0.753 1.0 1.0 1.0 0.850 1.0 1.0 1.0 0.873 1.0 1.0 1.0 0.886 1.0 1.0 1. 0 ^ *1 $
ne*
8 0.675 0.824 0.975 0.851 0.750 0.857 0.989 0.860 0.773 0.879 1.0 0.875 0.784 0.920 1.0 1.0 9 0.604 0.694 0.762 0.704 0.661 0.723 0.776 0.727 0.584 0.733 0.788 0.741 0.707 0.772 0.352 0.860 10 0.622 0.731 0.821 0.743 0.685 0.761 0.835 0.764 0.708 0.771 0.846 0.779 0.731 0.810 0.913 0.916 11 0.618 0.721 0.803 0.732 0.679 0.750 0.817 0.754 0.702 0.761 0.828 0.769 0.725 0.800 0.895 0.901 15 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0
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Table AS.4 ftARK III 138-732 STA'4DAl;D PLANT DYNAMIC PEAK PRESSURE FIELD FOR EICtfr (8) S/R VALVES TIME = 0.15 see (Positive Pressure ) AP (r)
TIME = 0.08 sec (Negative Pressure ) AP (r)
S/R Valves V-11 V-13 Angle (degrees)
Re fe rence Poin t 4.5 13.5 22.5 31.5 40.5 49.5 58.5 67.5 76.5 85.5 94.5 103.5
. 1 0 0 0 0 0 0 0 0 0 0 0 0 2 0.801 0.910 1.0 0.913 0.707 0.707 0.871 1.0 0.850 0.657 0.599 0.657 3 0.894 1.0 1.0 1.0 0.776 0.776 1.0 1.0 1.0 0.721 0.635 0.721 y, 4 0.912 1.0 1.0 1.0 0.788 0.788 1.0 1.0 1.0 0.733 0. 64 1 0.733 . . , ,
5 0.845 1.0 1.0 1.0 0.739 0.739 0.964 1.0 0.943 0.687 0.615 0.687 6 0.907 1.0 1.0 1.0 0.775 0.767 1.0 1.0 1.0 0.716 0.636 0.725 7 0.850 1.0 1.0 0.987 0.729 0.741 0.987 1.0 0.962 0.694 0.627 0.681 8 0.750 0.851 0.975 0.809 0.651 0.663 0.809 0.939 0.782 0.621 0.578 0.608 9 0.661 0.707 0.749 0.677 0.596 0.592 0.664 0.722 0.650 0.556 0.528 0.560 10 0.685 0.743 0.809 0.715 0.614 0.611 0.702 0.783 0.688 0.573 0.539 0.577 11 0.679 0.735 0.791 0.705 0.610 0.606 0.692 0.764 0.678 0.569 0.536 0.573 12 0.646 0.679 0.713 0.652 0.584 0.580 0.639 0.686 0.625 0:546 0.520 0.550 13 0 0 0 0 0 0 0 0 0 0 0 0 is a T S
e- e .e :..i
O O O Table AS.4 (Continued)
S/R Valves V-16 V-18 V-2 Angle (deg ree s)
Reference Point 112.5 l_21.5 130.5 139.5 148.5 157.5 166.5 175.5 184.5 193.5 202.5 211.5 1 0 0 0 0 0 0 0 0 0 0 0 0 2 0.893 1.0 0.910 0.801 0.910 1.0 0.893 0.657 0.S99 0.657 0.850 1.0 3 1.0 1.0 1.0 0.894 1.0 1.0 1.0 0.721 0.635 0.721 1.0 1.0 4 1.0 1.0 1.0 0.912 1.0 1.0 1.0 0.733 0.64 1 0.733 1.0 1.0 5 0.983 1.0 1.0 0.845 1.0 1.0 0.983 0.687 0.615 0.687 0.943 1.0 6 1.0 1.0 1.0 0.856 1.0 1.0 1.0 0.725 0.636 0.716 1.0 1.0 to w 7 0.971 1.0 1.0 0.835 1.0 1.0 0.971 0.681 0.627 0.694 0.962 1.0 $W 8 0.792 0.954 0.826 0.735 0.826 0.954 0.792 0.608 0.578 0.621 0.782 0.939 ro y 9 0.646 0.725 0.691 0.646 0.691 0.725 0.646 0.560 0.528 0.556 0.650 0.709 -
l 10 0.68) 0.786 0.730 0.670 0.730 0.786 0.685 0.577 0.539 0.573 0.688 0.771 11 0.674 0.767 0.720 0.665 0.720 0.767 0.674 0.573 0.536 0.569 0.678 0.752 12 0.620 0.688 0.666 0.628 0.666 0.688 0.620 0.550 0.520 0.546 0.625 0.672 13 0 0 0 0 0 0 0 0 0 0 0 0
?
Y
- $g
Talsle AS.4 (Continued)
S/R Valves V-4 V-7 Angle (degrees)
Re fe rence Point 220.5 229.5 238.5 24 7.5 256.5 265.5 274.5 283.5 292.5 301.5 310.5 319.5 1 0 0 0 0 0 0 0 0 0 0 0 0 2 0.871 0.707 0.707 0.871 1.0 0.850 0.657 0.599 0.657 0.893 1.0 0.910 3 1.0 0.776 0.776 1.0 1.0 1.0 0.721 0.635 0.721 1.0 1.0 1.0 4 1.0 0.788 0.788 1.0 1.0 1.0 0.733 0.641 0.733 1.0 1.0 1.0 5 0.964 3.739 0.739 0.964 1.0 0.943 0.687 0.615 0.687 0.983 1.0 1.0 6 1.0 0.741 0.741 1.0 1.0 1.0 0.716 0.636 0.72.5 1.0 1.0 1.0 7 0.972 0.717 0.717 0.972 1.0 0.962 0.694 0.627 0.681 0.971 1.0 1.0 :=
8 0.793 0.639 0.639 0.793 0.939 0.782 0.621 0.578 0.608 0.792 0.967 0.845 .
3 9 0.646 0.584 0.584 0.646 0.709 0.650 0.556 0.528 0.560 0.660 0.740 0.716 "$
10 0.685 0.602 0.602 0.685 0.771 0.688 0.573 0.539 0.577 0.698 0.801 0.737 11 0.6 75 0.598 0.598 0.675 0.752 0.678 0.569 0.536 0.573 0.688 0.782 0.726 12 0.621 0.572 0.598 o.621 0.672 0.625 0.546 0.520 0.550 0.635 0.704 0.673 13 0 0 0 0 0 0 0 0 0 0 0 0 S
N M .
0 9 -- -
9
22A4355 A-43 -
Rev. 2 O Table AS.4 (Continued)
V-9 S/R Valves Angle (degrees)
Refersnce Point 328.5 337.5 346.5 355.5 <
I 1 0 0 0 0 2 0.801 0.950 1.0 0.950 3 0.894 1.0 1.0 1.0 '
4 0.912 1.0 1.0 1.0 5 0.845 1.0 1.0 1.0 6 0.885 1.0 1.0 1.0 7 0.846 1.0 1.0 1.0 8 0.745 0.842 0.976 0.856 9 0.657 0.707 0.761 0.710 10 0.681 0.745 O.821 0.748 11 0.675 0.735 0.803 0.738 l 12 0.642 0.681 0.726 0.684 13 0 0 0 0 E
o l
b O 042.178
Table AS.5 MARK 111 238-732 STAllDARD Pl.AfiT DYllAMIC PEAK PRESSURE FIELD WR 19 S/R VAI.VES TIME = 0.15 sec (Positive Pressure psid)
TIME = 0.08 sec (Negative Pressure psid)
S/R Valves V-10 V-11 V-12 V-13 V-14 V-15 Angle (degree s)
Re fe rence Point 4.5 13.5 22.5 31.5 40.5 49.5 58.5 67.5 76.5 85.5 94.5 103.5 1 0 0 0 0 0 0 0 O O O O O 2 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 3 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 y 4 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1. 0 5 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 6 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 7 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 8 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 9 1.0 0.989 0.792 0.920 6.920 0.972 0.989 1.0 1.0 1.0 1.0 1.0 10 1.0 1.0 1.0 0.960 0.960 1.0 1.0 1.0 1.0 1.0 1.0 1.0 11 1.0 1.0 1.0 0.949 0.949 1.0 1.0 1.0 1.0 1.0 1.0 1.0 12 0.972 0.953 0.93/ 0.894 0.894 0.937 0.953 0.972 0.974 0.983 0.980 0.987 13 0 0 0 0 0 0 0 0 0 0 0 0 S
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- l 22A4365 A-47 i Rev. 2 '
~ '
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l Table AS.5 (continued) a j S/R valves V-8 V-9
) Angle (debrees)
- Reference Poine 328.5 337.5 346.5 355.5 i
1
! 1 0 0 0 0 f
4 2 1.0 1.0 1.0 1.0 t
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d 22A4365 Rev. 3 A-65 O
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O THIS FIGURE IS INTENTIONALLY DELETED.
l O 090779
O 22A4365 A-66 Rev. 3 O
A6.0 OTHER LOADS ON STRUCTURES IN THE POOL A6.1 LOCA AND POOL SWELL See Section 2.
A6.1.1 Forces on Pipes Due to Vent Clearing Pool Swell and Fallback The loadings are given for the quencher and reduced to effective pressure on a pipe in Table A6.1. The effective pressures of Table A6.1 can be applied nomal 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 THERMAL EXPANSION 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 expansion loads.
A6.3 SEISMIC LOADS (BY ARCHITECT-ENGINEER)
The seismic loads are to be applied by the plant designer. These are included in Quencher Anchor Loads, Section A7.0.
A6.4 SEISMIC SLOSE LOADS (BY ARCHITECr-ENGINEER).
See Attachment B.
090779 0
22A4365 - A-67
_ Rev. 2 Table A6.1 >
I TACA LOADS ON PIPES F*
p Force On Tira Quencher Water Velocity Event (sec) (ibf) (ft/sec) Ref Water Clearing 0.1 to 0.7 30 Sec. 8.1.1 and Fig. G-3 Pool Swell 0.7 to 3 40 Sec. 8.1.2 Fall Back 3 to 6 35 Sec. 8.1.3 2
CD # V
- Fp =
2g (144) A 4
l
}
042178
A-68 22A4365 Rev. 2
_ . e i
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22A4365 Rev. 2 A-69 i
A)
's.
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.
The designer should evaluate the optimum location for anchorage of the RVDL to the drywell sleeve. The analyses should consider line thermal expansion. The designer should also evaluate the drywell penetration sleeve to assure that the drywell concrete local temperature limit is not exceeded. Preliminary thennal calculations for the 238 Standard Plant drywel1 sleeve show that concrete temperatures for normal operation do not exceed 200 F and 14" Schedule 80 sleeve is acceptable. Designers should perform independent calculations to assure tHese findings.
A7.1 QUEN01ER 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 presenting a maximum loading condition.
042178
22A4365 A- 70 Rev. 3 Table A7.2 lists typical design loads for the Mark III quencher configuration.
lll 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-tion loads illustrated by F g
, yF , gM and M in Figure A7.2. These are typ-ical design loads for the quencher supporting structure.
Interface 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 and 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 o f a quencher.
A7.2.1 Codes and Standards
- a. American Society of Mechanical Engineers (ASME) Boiler and Pressure Code.
(1) ASME Section III, Nuclear Power Plant Components
- b. American National Standards Institute (NIS1)
(1) ANSI B16.25, Butt Welding Ends for Pipe, Valves , Flanges ,
and Fittings.
- c. American Institute of Steel Construction (AISC).
O 090779
l 22A4365 Rev. 3 A-71 O
A7.2.2 Desian Pressures. Temperatures .steadsi_ Configurst. inn arid Performance A7.2.2.1 Component Data
- Safety / Relief Valve, Discharge Piping and Quencher
- a. Design Freasure 570 psig
- b. Design Temperature 475 F
- c. Maximum Pressure. 625 psig
- d. Maximum Temperature. sat. steam
- e. Maximum Flow 520 metric tons /hr
- g
! O at 1190 psig j
- f. Maximum Back Pressure 40% of safety / relief valve set pressure i
- g. S/R valve Minimum 0.020 see Disc. Strok.a Time.
- h. Minimum Ambient 60 F
[ Service Temperature A7.2.2.2 SRVDL Geometry (See Section A10.)
l lO
! 090779 l
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 A5.11 A7.2.2.4 Quencher Configuration and Location.
~
- a. PROPRIETARY, Provided under separate cover
- b. PROPRIETARY, Provided under separate cover
- c. PROPRIETARY, Provided under separate cover
- d. PROPRIETARY, Provided under separate cover
- e. Quencher arm length 58.5 in, to CLQuencher
- f. Quencher pipe size / 12 in./Sched 80 (suggested).
schedule
- g. Internal Quencher 101.6 sq in.
pipe area
- h. Min clearance between >5 ft Cg Quencher and pool floor / basement i
1 l 042178 O 1
'5 i
1 22A4365 A-73 Rev . 3 i
- 1. Plane of 4 Quencher legs Horizontal
- 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)
! s tainless 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 i
i e
I
() 090779 1
i
. _ . _ _ _ , _ , . . . _ . , ,, _ . _ . . . _ . _ .,.__1 _ , .
22A4365 A-74 Rev. 2 O
Table A7.1 QUENGER AILM LOADS (Reference Figure A7.1)
Load Description 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 F , downward direction) e Earthquake load,1.25g - (lbs) at SSE 1488 (Location F , vertical direction)
Earthquaka load,1.0g - Gbsl at SSE +390 (Location F , horizontal dire.ctioni 3
- Due to single valve subsequent actuatie..
042178
22M365 A-75 l Rev. 2 ~
4 O
Table A7.2 MARK III QUE3 CHEF. ANCHOR LOADS l (Reference Figure A7.2) 1 Air Clearing Water Clearing Lateral Loads - (1bs)
F - Air and water clearing 28,510 8,553 b
LOCA vent water clearing 10,240 10,240 F - SSE Earthquake load (1.0g) quencher :nass 3,940 3,940 e
- SSE Earthquake load (1.0g), water mass - 1,680 1,680
- 10,855 10,855 F - Inlet line load t
- F g - Total base lateral reaction load 53,545 35,268 O Vertical Loads - (1bs) 111,344 14,651 F, - Air clearing Transient wave 19,000 -3,700
-LS ,000 +2,400
! Pool swell -14,742 -14,742 quencher weight +3,940 +3,940 SSE Earthquake load (1.25gl 14,925 16,425 4
F1 - Inlet line load 110,855 110,855 Water clearing +150,000/-2,000 40,064 178,271 i
i -56,866 -37,722
- F, - Total base vertical reaction load 042178
. . , - . . . - - - . - . - - . . - - . . .-w ., . e nw - - - - - - --, -- , ., --- .c - , _ . . . , , - - - - - - - , - - - ~
l 22A4365 A-76 )
Rev. 2 O
Table A7.2 (Continued)
Air Clearing Water Clearing lateral Moments Transferred to Base Plata - (ft-Ibsl Air and water clearing 37,524 11,257 M,-
Pool swell 17,751 17,751 Moments resulting from lateral loads -
2.64 x (Fb (air clearing) + LOCA vent clearing] 102,300 49,614 9,14 9,141 2.32 x Fe (eart p.de, quencher mass) 6.67 x F (inlet line) 72,402 72,402 3.00 x Fy(earthquake, water mass) 5,040
- M - Total base lateral reaction moment 239,118 165,205 Z
O Vertical Moments Transferred to Base Plate - (ft-lbs)
M - Air clearing 105,618 31,685 b
Multipla valva actuation 0 0 LOCA vent claaring 8,047 S,047 M - Inlet line moment 25,836 25,836 t
- M y - Total base vertical reaction moment 139,501 65,568
- Quancher bottom flange anchor loads. (Individual loads are time dependent and peak values are conservatively combined.)
042178 h
22A4365 g 77 Rev. 2
- O t
80' 38.4 in.
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34.3 in.
1 NOTE: '
8 LOADS APet.Y TO 90 -90' CONFIGURATION AS WELL AS 120*-40*-30*-80' v
34.3" % Fe l U f )
C ) ( J l
etevarics view ( l 4 l
l -
Figure A7.1. Quencher Arm Loads 042178 9
22A4365 g_73
-. Rev. 2 O
ORTHOGONAL INLET LINE LOAOS F; AND Mg FOLLOW THE RELATIONSHJP:
F1 Mi
- 'I Mg { K FI WHERE:9 a a 10,855 'b Mo = 25.836 ft-Ib h
C S w w 0
- l! $
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= = Mi do g E e 0
= Fa
^
Fs r Ma Fc l
l Il N "
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= = \ )
1 n n N I f 3 F
4 Me SASE REACTION LOADS Mw NOTE:
LOAOS Fb. Ec, AND Fg MAY ACT IN ANY HCRIZCNTAL DIRECTION p MCMENTS M, AND Mg MAY ACT IN ANY VERTICAL PLANE t
Figure A7.2. Quencher Load Diagram I
i 042178 f
22A4365 A-79 Rev. 2
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042178
22A4365 A-80 Rev. 2 O
l This figure is PROPRIEIARY and is provided under . separate cover.
G Figure A7.4. Sectional View of Quencher Leg (Typical Each Side) 042178 g
22A4365 A-81 Rev. 2 p
V A8.0 S/R VALVE LOAD COMBINATIONS Safety / relief valve discharge piping routed to the suppression pool is i 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 L' 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 caces where various numbers of open valves can be postulated for the 238 Standard Plant:
Case Number of Valves (1) 1 Single activa failure, normal function or operator action. (First or subsequent actuation)
(2) 2 1 nonnal plus single active f ailure of adjacant (adjacent) valve (First actuation)
(3) 10 All 11113 psi set point valves (First actuatim (4) 8 ADS Activation (71rst actuation) s (5) 19 Vessel pressure 11123 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 wf era various reactor power levels are assumed when the transient event is t
initiated. Therefore, the containment must be able to withstand any 0 042178
22A4365 A-82 Rev. 2 O
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 SYhMETRIC AlfD ASYMMETRIC LOAD CASES The following selected cases represent the asymetric 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 poi sat point valve.
B. 2 Adjacent S/R Valves - This situation can occur due to a pressure transient at low power, whic's would lif t one valve.
Concurrent with this the single active failure of an adjacent h
valve is assumed.
tea prohahility of the precisa confilnation of two adjacene valves woul.1 be very low, sinca common sat point valva discharge points are uniformly distributed around the pool. However, if the containment structural design requirements are satisfied under this asymmetric condition, subsequeist analysis need not be performed for the multitude of other
= ore probable asymmetric load cases.
The following selected cases represent the symetric cases for contain-ment loads:
C. 8 ADS Valves - This situation can occur with an intermediate break where the ADS system is activated.
042178 g
22A4365 A-83 p Rev. 2
\d 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 recommended.
From observation of Figures A5.2a, A5.4a, AS.6a, AS.8a, and AS.10a the 1 or 2 valve lo&d case is the governing case for asymmetrical considera-tions, and the 19 valve load case for maximum symmetrical consideration.
The final :tlection of valve combinatiens is the designer's (A.E.)
responsibility.
A8.2 SSE AND OBE CONSIDERATIONS Whatever 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 arc 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 accounc for the hypothetical accident related loads and their sequence o. occurrence. In defining the loads for this evaluation, this report addresses the design basis accident (pipe 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 upon.
1 O
042178 l
22A4365 A-84 Rev. 3 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 IDCA and, furthermore, is not an expected occurrence during the accident sequence. However, the loading chart figures show the loads associated with a single safety / relief valve actuation as an additional load for demonstrating additional capability.
A8.3.1 DBA With M.S. Line Break For the DBA, with M.S. line break no valves will lif t due to vessel pressure rise (Figure 4.1).
A8.3.2 DBA With Recirculation Lina Break For the DBA, with a recirculation line break no valves will lif t due to vessel pressure rise (Figure 4.1) .
O O
090779
22A4365 A-85 Rev. 2 A8.4 RECOMME:TDED DESIGN LOAD SUMMATION The design loads on MK III structures are comprised of static (dead loads ,
live loads, hydro , etc.) and alternating dynamic loads (saismic and S/R valve loads, etc.) .
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 committee for combining loads will be adopted:
1/2 R= [ (DC) i [ (AC) i=1 i=1 where O 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 compenent.
i = 1, 2, . . . number of time varying events for which the rerultant 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 then 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 j D l V 042178 l
22A4365 A-86 I 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 s1= ply, the above equation with respect to loads or stresses , etc. ,
may be represented by:
- -2 - -2 SRV Seismic SRV Seismic EF = D.L. + L.L. + Hydro + + AC DC DC AC where:
- D.L., L.L., Hydro = Static Loads
- "SRV" " Seismic' "
" Y ## "I #""~***"* #8
- DC) -
DC) - cocoonent (DC) of the dynanic response SRV' ' Seismic' = Alternating co=ponent of responsa defined
-AC)'-
- '\C) as the maximum response. value. minus tha corresponding DC componer t.
This is si:sply represented as, followsr:
g - AC CCMPCNENT OF RESPCNSE.
Ag g A [3 N (SEISMIC }
Y A NCN ALTERNATING DC COMPONENT OF OYNAMIC RESPCNSE, seismic aS oC when Cae alternating component has no DC co=nonent, tha DC ter=s drop out.
042178 9
l
22A4365 Rev. 2 O A-87 i
A9.0 FATIGUE CYCLES During the 40-year plant life, there will be safety / relief valve (SRV) dis-charge events that are anticipated to occur. Based on the many years of BWR plant operating experience, an analysis has been parformed to determine the mean frequency of occurrence of the potential events. This information is presented in Table A9.la. Some of the transients that can occur result in containment isolation; in which case, subsequent opening of a SRV will occur to remove decay heat until an alternate path such as (1) bypass of the MSIVs
, to the main condenser, or (2) RHR steam condensing mode can be established.
Table A9.lb lists the number of subsequent openings of the low setpoint valve that are determined to occur during an assumed 30-minute period for the establishment of the alternate path for decay heat removal.
The total number of valve openings recommended for use in a BWR/6 Mark III
() Containment Fatigue evaluation is conservatively set at 4200 cycles. For BWR/6 systems where " low-low set" instrumentation logic is used, the total number of valve openings is 1800 cycles. The containment designer should use this number of cycles in conjunction with the quencher pressure time-history as shown in Figure AS.ll for evaluating the containment fatigue life.
l O
101678
22A4365 A-88 Rev. 3 O
Table A9.la SAFETY / RELIEF VALVE ACTUATION Number of Valves Open for Initial Blow Isolation Mean Frequency / (All (1/2- (1/ 3 Type Events 40 Years 2/3) 2/3) -0) Event Turbine Trip (w/BP) 51 x No Load Rejection (w/EP) 29 x No Pressure Regulator Failure 26 x Yes Feedwater Controller Failure 19 x , No Trip of Both Recircu-lation Pumps 10 x No Recirculation Con-troller Failure 10 x No Loss of Feedwater Flow 29 x No Loss r f Auxiliary Power 10 x Yes Closure of all MSIV's 38 x Yes Loss of Condenser Vacuum 26 x Yes Inadvertent Relief valve Opening 4.0 x No Turbine Trip (w/o BP) 1.0 x Yes Load Rejection (w/o BP) 1.0 x Yes 090779 g
22A4365 Rev. 3 A-89 Table A9.lb l
' CYCLES OF SINGLE LOWSET SAFETY / RELIEF VALVE PER ISOLATION
' Cycles / Isolation Plant Cycles / Isolation (Low-Low Set) 218 BWR/6 - Steam Turbine Feedwater Pumps 34 15 218 BWR/6 - Motor Feedwater Pumps 39 15 238 BWR/6 - Steam Turbine Feedwater Pumps 29 15 238 BWR/6 - Motor Feedwater Pumps 34 15
't ;
O 251 BWR/6 - Steam Turbine Feedwater Pumps 24 15 251 BWR/6 - Motor Feedwater Pumps 30 15 0
090779 i
i
22A4365 A-90 Rev. 2 O
A10.0 RECOMMENDED CALCULATION PROCEDURES FOR MARK III DESIGNERS The following information provides the procedures for predicting loads on the drywell wall, basemat, and containment wall associated with the air clearing transient following the opening of a safety-relief valve for the 238 standard MARK III plant. The numbers 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 information in Section A12.0. For design purposes , a statistical evaluation of tne data was used. Design values represent a 95%-95% 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 RVDL.
(1) Peak .: .tessure 1625 psid.
(2) p'L cannot exceed those values given in Figure A3.1 at the corre-sponding pipe volume.
(3) Water Leg <17.8 ft.
Constraints ou routing the safety / relief valve discharge line are:
- 1. No more than one 90 long radius bend coming off the relief valve , and two 45 long radius bends entering the quencher in the 10" schd 80 piping. The remaining bends should be in the 042178 e
l l
I l
l 22A4365 A-91 Rev. 3 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.
- 2. The initial length of 10" schd 40 pipe be kept to a minimum.
A10.2 DETERMINE SRVDL DESIGN The following steps are reconsnended for designing the SRVDL within the above constraints and the design requirements in Table A4.2.
(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.
(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 process where a balance of 10, 12, and 14-inch SCH 40 piping is adjusted to the minimum total air volume and fL/D for the 625 psi pipe prassure constraint. It is important to insure that all the SRVDL air volume and fL/D from the SRV to the free water surface is included. Figure. A10.1 sfiows the portion of SRVDL from the SRV to the first anchor.
l (3) For the portion of the SRVDL shown in Figure A10.1, the loss coef ficients, K, for each of the three flexible joints are.
l 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 SRVDL.
090779
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 :
2 2 ~
(A10"\ +K IA10"I fL/D)Ref l0"
= K.,. +K T **1 T **1 I I + ***
10" 12" (Al2") 14" (#14") ,
whe re :
botal 10" 10" kosses 10" S/40 botaly 3,, 12" + kosses n" S/40 kotal l4,, ! 14" + kosses l4" S/40 A10" = Rydraulic area f 10" schd 40 pipe (ft )
A 12., = Hydraulic area f 12" schd 40 pipe (ft2)
A g , = Hydraulic area of 14" schd 40 pipe (f t )
D 10"
= Diameter of 10" schd 40 pipe (ft)
D g,, = Diameter of 12" schd 40 pipe (f t)
D 14,, = Diameter of 14" schd 40 pipe (ft) 042178
22A4365 Rev. 2 A-93 The friction factor "f" in the above equations should be calculated based on the pipe diameter, relative roughness of the pipe, and the Reynolds number. A Reynold's number of 0
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.
l Using the system fL/D calculated above enter Figure A3.1 with corresponding air volume. The intersection must fall on or above the 625 paid curve.
(6) Determine the quencher bubble pressure using the actual air j
volume in the RVDL, see Section A12.6.
1 i A10.3 S/R VALVE AIR CLEARING LOADS MARK III 238 STANDARD PLANT O After the quencher bubble pressure has been obtained, Section Al2.6, the next step is to calculate wall pressures based on the peak bubble value (+ and -).
A10.3.1 Absolute Pressure on-3asemat and Walls The at: soluta pressure anywhere on the dry. ell wall, basemac, and con-tainment wall in the wetwell region can be calculated by the equation:
P (al = P + + AP(r) (1) where P(a) = absolute pressure at arbitrary point "a" Gsia) r = distance from quencher center to point "a" (ft) 042178
22A4365 A-94 Rev. 2 O
P = absolute pressure of containment atmosphere (psia) h(a) = vater head acting at point "a" (ft) o = water density (approx. 62.4 lbm/ft )
AP(r) = bubble pressure attenuated by distance (r) to point "a".
The attenuated bubble pressure for one S/RV, AP(r), can be calculated from the bubble pressure, SPB , [ which is obtained from Section A12.6]
using the following equations:
SP(r) = 2 x AP 3 l (drh for r > 2r (2) o r
j AP (r) =
AP3 for r 1 2r 0 (3) g where, r, = quencher radius = 4.875 ft.
A10.3.2 How to Find the Attenuated Pressure on the Drvwell Wall.
Basemat, and Containment Wal.l.
A10.3.2.1 Develop grid to detarmine 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 llev . 2
'l
- 3. Arc distance by 360 t (vent stations) (Table A10.4) . ,
i
- 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 quencher.
- 5. The point (a) is then selected and the distance (r) to (a) is l
j 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).
i
- A10.3.2.3 Wall Pressure at Point (a) for Multiple S/R Valve.
O In the event of multiple S/RV actuation the attenuated bubble pressure, I AP B, must be calculated using the following equations:
II n
j AP(r) =
[ AP n=1 where,
[ *o \ fo r i #
- AP = 2AP g l n o (7)l n
f## n 1# o If the calculated AP(r) > AP3, set AM = APB ' '" * # = the dis-n tance from the center of the quencher to point a.
l 042178
A-96 22A4365 Rev. 3 For the cases where multiple valves are discharged due to a pressure transient, the valves in each set point group (1103,1113, and 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 difference in time of discharge due to the 10 psi dif ference in pressure setpoints of the valve groups. Using the quencher bubble model presented in Figure AS.ll, it is seen that when P
g from the 1123 psi setpoint valves occurs, the bubble pressure from the 1113 psi setpoint valves has dropped to 0.9175 Pg , and the bubble p ressure 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 preasure loading, there is significant reduc-tion in pressure at certain locations when considering the time sequenced phasing approach. The most limiting position on the containment is not a f fec ted (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 load calculation. These two factors (i.e. , time phasing and the multiple l
valve statistical consideration) have not been included in the develop-ment of the local pressure distributions on the containment wall because they do not affect the limiting local pressure. However, these factors are impo rt ant to the structural response and will be employed in the building response evaluation. Attachment M presents the method for treating these ef fects in detern.i..ing structural response used for the equipment evaluations.
, 090779
'i 22A4365 Rev. 2 A-97 O
Table A10.1 218 STANDARD PLANT P~iPE 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 4
V2 73.38 82.50 155.88 13.0 v3 73.38 82.50 155.88 13.0 i V5 73.38 72.12 145.50 12.12 I
V6 130.00 71.62 201.62 16.80' i
V7 118.38 71.50 189.88 15.82 V8 115.88 71.00 186.88 15.5 7
- O V9 115.88 71.00 186.38 15.57 y
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 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 a
V16 73.38 82.50 155.88 13.0 i
The valve numbers shown on the table above are the same valve numbers on Figure A4.5.
(}
042178 l
1
22A4365 A-98 Rev. 2 l Table A10.2 238 STANDARD PLANT PIPE SPOOL DIMEISI0IIS Dimension A Dimension B Total Di=ension (A + B)
Valve No. (in.) (in.) (in.) (ft)
V1 75.00 82.25 157.25 13.10 V2 75.00 82.25 157.25 13.10 73 75.00 82.25 157.25 13.10 V4 75.00 82.25 157.25 13.10 75 75.00 72.00 147.00 12.25 V6 75.00 71.62 146.62 12.22
/
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 V10 119.88 70.62 190.50 15.88 ggg V11 119.88 70.62 190.50 15.88 v12 120.25 70.88 191.13 15.93 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 82.25 157.25 13.10 V17 75.00 82.25 L5 7. 25 13.10 V18 75 .00 32.25 157.25 13.10 V19 75.00 82.25 157.25 13.10 The valve numbers shown on the table above are the same valve numbers on Figure A4.3.
042178 O
1 l
22A4365 A-99 7
Rev. 2 Table A10.3 251 STANDARD PLA?rr PIPE SF00L DI)ENSIONS Dimension A Dimension B Total Dimension (A + B)
Valve No. (in. ) (in.) (in.) (ft)
V1 73.38 82.62 156.0 13.0 V2 73.38 82.62 156.0 13.0 V3 73.38 82.62 156.0 13.0
) 156.0 13.0 i V4 73.38 82.62 V5 73.38 72.62 146.0 12.17 V6 73.38 72.25 145.63 12.14 V7 73.38 72.00 145.38 12.12 V8 144.75 71.12 216.50 18.04 V9 133.50 70.88 205.1') 17.09
~
V10 128.38 70.62 199.75 16.65 i ()i 70.38 199.13 16.60 V11 123.00 i
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 vi7 73.38 72.25 145.63 12.14 V18 73.38 72.62 146.0 12.17 vi9 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 i
Thr valve numbers shown on the table above are the same as valve numbers
(} on Figure A4.9.
042178
22A4365 Rev. 2 Table A10.4 h DRWELL AND SUPPRESSION POOL GEOMETRY VENT (V)
ANNULUS C RPV q SHIELD O ORYWELL E WALL -- *-
E LEV 0 f t4 in.O Y
PE DESTAL l
=
A
,, 8 ~ POOL SUPPRES$
- 15) ION "[ CONTAINME NT WElm WALL d m 0.50 ft 08 - ~h ~ HtGH WATER LEVEL (HWLI
_ORYWELL. g ANNULUS ,10N -
LOW WATER LEVEL ILWL3 08 (Wi
,l. +-
4;.4.50 ft P
, Y 07 .'.
~ 4.50 ft I ,
gf' 8ASE MAT -
4 ,
T1 -
27.50 ea. '
T2 --* i VENTLO TABLE 1 __
PLT SIZE /CNTMT DiA OR NO. 07 FUEL SUNOLE DESCRIPTIONS 1 218/114 I*218/120 ' *238 L 251/800 *251/864
[
t-6 8.33 t-6 8.33 (-p 5.50 t-6 7.08 (-p 7.08 A ELEV TOP OF WEIR e' ALL 455 455 482 570 570 V VENT ANNULUS AREA (ft2) 5760 6863 6382 8170 8170 v+S TOTAL ARE A (ft2)
(-) 12.58 (-) 12.58 (-) 11.16 l-l 13.00 (-) 13.00 HWL HIGH WATER LEVEL ELEV 4 25 4 25 5.67 5.92 5.92 L M!N FREEBOARD (ft)
(-) 16.92 (-) 16.92 (-) 15.53 (-l 17.33 (-) 17.33 H ORAWOOWN LEVEL ELEV 19.42 19.42 20.42 19.00 19.00 Y POOL DEPTH fft) 111.00 131.90 129.60 15330 153.90 POOL VOL (1.000 ft3 ) AT LWL 21.30 17.31 34.15 30 03 30.03 DRAWOOWN MAKEUP VOL (1000 ft 31 t-l 29.08 (-) 29.08 (-) 27.67 (-) 29.50 (-) 29.50 J t OF SOTTOM VENTS ELEV I-1 32.00 (-) 32.00 (-l 31.58 (-) 32.00 (-) 32.00 F ELEV TOP OF SASE MAT 102 102 120 135 135 NUMBER OF VENTS 34 34 40 45 45 VENT STATIONS 420 420 495 557 557 GROSS VENT ,4RE A Ift2) 2100 2100 2475 2785 2785 VENT VOLUME (ft3) 16 16 19 20 22 NUMBER OF SAFETY RELIEF VALVE 6.38 6.28 5.75 5.25 5.25 ClRCUMFERENTIAL VENT SPC'G DW IO 2223 2223 2535 2688 2688 .'. AREA (ft2) 2184 27.84 39.29 39.82 38.75 W VOLUME (1000 ft3l 267.6 267.60 301.10 351.90 351.90 P ARE A (ft2) 5525 5525 0948 7477 7477 P VOLUME (ft3)
(-l 28.98 (-l 28.98 (-) 28.58 (-l 28.33 (-l 28.33 07 RSO PT ELEV (REF1
(-) 19.35 (-l 19.35 (-) 19.46 (-) 18.71 (-l 19.21 08 RSO PT ELEV IREF)
(-) 12.35 (-) 12.35 (-) 12.46 (-l 11.71 (-) 12.21 09 RSO PT ELEV (REF)
(-) 20.85 ( 1 20.85 (-> 20 98 (-) 21.50 (-) 21.90 15 RSO PT ELEV TREF) 113.00 118.00 145.50 165.60 (LTR) 1100F STO REOO POOL VOL (1000 ft3 )
102.40 102.40 124.20 141SO (LTRI PLT 1000F VS SVCE WATER TEMP 87.60 87.60 108.50 123.20 (LTR) 900F TABLE 2 PLT SIZE 06(REF) 15(REF) A B C 0 l E l Tt T2 OlA $ 8.42 29.83 61 64 67 63 79 114 218 RAD 9.21 14 91 30.50 32.33 34.50 39.50 57 1.83 217 DIA 18.42 29 83 61 64 67 69 79 120 EI3 2A7 2II. AA0 9.21 14 91 30.50 32.33 34 50 39 50 GO
,38' OIA 19 58 31.58 65 68.67 73 83 120 1 83 217 RAD 9.79 15.79 32.50 34.33 36.50 41.50 60 DIA 21.17 32.67 67 70 75 85 130
,$y.
33.50 35 37.50 42 50 65 1 50 2%
RA0 10 58 16.33 NOTES.
- 1. PLANTS 10ENTIFIED WiTH (*)
ASTERISK ARE STANDARD PLANTS 101678
22A4365 Rev. 2 _
T 'A 0.19 &
Ff 9E 10 un. 5/s0
_ ,j-""J S
. . u -
N s, .
. d M' " puumus K
- 0.175 K a 0.077
/7 THE R MOCOUPLE '8:0.19 CONNECTION e
_ (3) FLExtBLE 8ALL JOsNTS I l K = 0.077 y
i i i l
'SEE T A8 LE A10.1. A10.2 AND A10.3 FOR VALUES OF A ANO S FOR THE 218,238. AND 251 STANDARO PLANTS Figure A10.1. Safety / Relief Discharge Piping Detail SRV to First Anchor O
042178
r.OTES;
- 1. EQUAL DIST RisuTION OF QUENCHERS IN THE SUPPRESSION POOL IS REOutRED FOR THE FOLLOWINC. FONCTION:
(Al ADS VALVES (81 SPRING SET POtNTS
- 2. NON-VERTICAL LE* ,THS OF DISCHARGE PLPE ARE LOCATED AS HIGH AlsOVE THE SUPPR ESSION POOL AS 65 PR ACTICAL
- 1. SLOPE ALL OtSCH ARGE PIPES TOWARD THE SUPPRESSION POOL
- 4. Two V ACuuM SRE AKERS ARE REQUIRED. ONE IS LOCATED AT LEAST 10 FEET ABOVE THE WElR WAl_L. THE OTHER is LOCATED AS CLOSE TO THE SAFETY RELIEF VALVE As PRACTICAL
.D4L a ee.
. ceA .s e 8
< La
.. -- ;,- , g ,- - - , F ...* ~.s r -* i.- *' Ne
- , , ' , _ -'.,,.N \ . *-
n ,,, ,
\ ,'; A \'-s 4g i " ~}
W**'*.5 *. g .
g8 g g g g l ! \ ) e-r F'*"
/s gaa. , Pla.e Wh u. - - - - . -- .' m g 'e .e. .oa au-g *
- c v- g- ' - - u n *a' , ;* '
, , .;b ... . m. .r e.w u .e . . use r. ..
, ,<- YT ti-- i e e
.. *t.a= n c=j .A g.
. .u -.a s as ..
. .; . p. .
S . .
G
' ' ' 'g** Ad
-- 3g M NR -
~
.' N 3 ,. ,,7,,, ,
,=.
., L s... a. .
~
.=-.. ...
1 : ..
... . . . . i -
= -
- g. . .
- q} , "
I. .I. 4 .'. 2, i ,.
7
, q i t
t
.e .s.e one ss m e s . ~- -
90 0;- '
l -g Oou -. l- - - . . . ............. .. . . . 4++
O*- - --)- ' *- - *- *. .L- . . . . *. .. ..... ..e. 6 . . . . . .. . . . . .. . .
" "u f ais s. g. 4 e*
. . . .n o . . o .. . t .
n N
V u >
co Figure A10.2. Safety /Reilef Valve Discttarge Piping Arrangement e' o
N G G e
O O O 2f 45*
a -
u.
4 '
- 4 COP 4784NMENT pd '
- - >p
- 0 0 '"
O n, s% '
so
- *i.
\
g
/DRYWELL O.D.
\ JU
\ 55 N
DISTANCE FROM CENTER OF OUENCHERntr IN ft.
REFERENCE POINT / ANGLE 00 90 180 270 360 450 540 630 72 0 810 13 - - -
12 16.1 1s.2 23.2 29.4 ass oss 80s es.1 64.6 71.1 11 14.1 16.4 21D 28.4 36 3 43.2 50.4 57.5 64.1 10.7 10 13.7 16.1 21.7 28.2 3o.6 43.0 50.2 67.4 64.0 70 4 9 15.1 17.3 22A ED 36.2 43 5 50.7 57.7 84.3 70 S 8 11.0 13 4 19.1 26.2 33 4 40.7 47A 54A 81.1 -
7 72 102 17.2 24.0 31.3 30.3 46.1 512 57.9 -
6 6.5 9A 16.1 22 2 29.7 38.4 423 49.2 - -
5 8.2 10A 16.3 22.4 3.7 36.2 - - - -
4 5.2 8.7 15.1 21.5 28.0 34 6 - - - -
3 6.1 9.3 15.4 21.7 3.2 34 3 - - - -
2 99 12.1 47.2 23.0 29.2 36 4 - - - -
i S
N >
5 Figure A10.3. 238 Standard Plant Distance from Center of Quencher to Pressure Point (ft) ,L O
22A4365 A-104 Rev. 3 All.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.
Recommendations for quencher location within the pool and the drywell wall penetration location minimize the flexibility in the water leg 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 bsen generated. _
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. After the bubble pressures have been determined, the procedures for determining suppression pool boundary loads identified in Section A.10.3 should be utilized.
I 090779 l
l l
i e
22A4365 A-105 Sev. 3 '
O Table All.1 QUENCHER BUBBLE PRESSURE SENSITIVITY TO SRVDL AIR VOLUME Bubble Pressure (psid)
First Actuation Subst,quent Actuation Ai V lume Maximum Allowable (ft3) f1/D at 10" SH40 Pipe P+ P- P+ P-40 1.0 9.9 -6.7 20.9 -10.4 44 1.85 10.9 -7.1 22.9 -10.9 48 2.72 ' 11.6 -7.4 24.2 -11.2 52 3.60 12.6 -7.8 26.4 -11.6 56 4.45 13.6 -8.3 28.4 -12.0 60 5.35 14.4 -8.6 29.7 -12.2 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 Opening Time, var = 20 msee.
Quencher Submergence, SUBM = 13.92 ft. (4.24 m) 090779 l
A-106 22A4365 Bev. 2 O
A12.0 BASIS AND JUSTIFICATION FOR DEVELOPED QUENCHER LOADS A12.1 INTRODUCIION To assure that the containment loads resulting from S/R valve discharge phenomena are conservatively icv 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 'nas 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-ment pressure loads for the standard 238 plant. Included in this report is a test description and a summary of test data upon which the quencher design and performance are based.
g 042178 $
O
+ - - - - - - - - . a. . a. a s , 4 4
~
22A4365 A-107 thru A-124
~
Rev. 2
, O.
\
k f
O SECTION A12.2 CONTAINS GE COMPANY PROPRI fARY INFORMATION PROVIDED UNDER SEPARATE COVER l
l l
1 0421.78
1 22A4365 A-125 l Rev. 2 1
./
A12.3 PHYSICAL PARAMETERS ;
. 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-ation data. As far as subsequent actuations are concerned, only the
[)
maximum peak bubble pressure for any series of actuations is of concern.
This maximum , was correlated with the peak of the first actuation.
A12.3.1 I=cortant Paramerers 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 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 A12.3-1 shows the interaction and interdependence of various parameters influ-encing quencher clearing. .
t, 042178 O
l 22A4365 A-126 l Rev. 2 O
A12.3.2 Overview o f the Phenomenon Since the peak pressure on the pool boundaries is the main concern, we must first look for factors that influence containment air clearing loads, i.e.:
(1) number 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 deternine 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 quencher, therefore, pool surf ace area per quencher ()7) divided by the area of the quencher (A q
= the area of the circle that circu= scribes 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 Subble Size For a given peak bubble pressure, bubble size depends en the mass and the temperature of the bubble. Since the bubble is essentially at pool temperature, one of the important parameters is pool temperature (T.g) .
042173
+
1 22A4365 A-127 Rev. 2 _
Asstsaing, constant initial conditions in the discharge line, the mass of the bubble is proportional to the initial air volume in the pipe. Since l the air is spread over the quencher area, the important dimension becomes the height of the bubble, which is proportional to the initial air
. volume (VA )/quander area (Aq); hence, V /AAQbecomes 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 poel water, the air is essentially cooled to suppression pool temperature. Therefore , for a given air mass, the peak bubble pressure depends on:
1 (1) bubble shape, and (2) mass flow race.
j A12.3.2.3.1 Bubble Shape Bubble shape refers to tha 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 che 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 bubble and is reflected in Vg .q .
042173 O
22A4365 A-128 Rev. 2 O
A12.3.2.3.2 Mass Flow Race 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. S.nce 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, therefore, lower bubbic 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 ficw 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 colu=n, which is normally equal to quencher submergence, is the main parameter affecting the peak pipe pressure, the air discharge pressure, and the velocity of water as it clears the quencher ar=s. The faster the water is expelled from the quencher arms , the faster the holes become available 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 af fects the air flow rate and, therefore, the peak bubble pressure.
The air volu=a has already been identified as a key parameter; however, it should be pointed out that the ef fect of increased air volume in this case is to reduce the mass flev rate and thereby reduce the peak bubble 042178
22A4365 A-129
~ Bev. 2
(
pressure. This is in the opposite di cection of the effect of Vg/Aq which was previously identified. In fact, as the air volume is increased, these opposing effects eventually cancel each other out.
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 (tn,) to a power of approximately
- 0. 7.
The air flow rate (or steam flow rate) is converted into a :nass flux to be suitable as a physical parameter. This is done by dividing the mass flow rate by an area such as quencher opening area (defined as the total hole area). For quenchers of similar geometry, the opening area is propor-tional to the quencher area (A ), q and therefore, bubble pressure becomes a function of the mass flux across the quencher area (Aq ). To sum =arize, air mass flux depends on37 , (m, * )/.g and the length of the water column, which is normally equal to the submergence of the quencher.
Since the maximum steam flow rate occurs only when the valve is fully open, valve opening time must also be considered as a parameter af fecting peak i 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 To sununarize, the following main parameters were identified as the ones that significantly affect the peak boundary loads :
(1) Pool area per quencher / quencher area (A g /A q );
(2) Pool temperature (Tg); i O 042178
A-130 22A4365 Rev. 2 O
(3) Initial air volume / quencher area (7 /Aq );
(4) (Steam flow race to the power 0.7)/ quencher area (5 /A );
q (5) valve opening time (var); an i (6) Lengths of water column (WCL) .
Other parameters fall in one of the following categories:
(1) Parameters which, within the range of the data, did not seem to have any effect at all on the boundary pressures, 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 (except for 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 dia=eter, st=e and arrangement of the Boles, and quencEer area.
(31 Parameters that become important only for subsequent actuations (e.g. , pipe temperature and water velocity prior to valve actuation 1 The combined effect of these parameters is accounted for by the use of a statistically determined =ulti-plier applied to the first actuation loads, i
A12.3.4 Ef fects 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 follows ,
these parameters and their effects will be discussed in more detail.
042178
1 l
22A4365 A-131 Rev. 2 4
O A12.3.4.1 Pool Area Per Quenchar/Quenchar Area (Ag /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 effect of multiple relief valve actuation.
A12.3.4.2 Pool Temperature (Ty)
Part of the energy absorbed by the air column during the compresskt 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 I low. The heat transfer rate depends on the pool temperature and vanishes when the air temperature becomes equal to pool temperature. The pool
- temperature, therefore, establishes the lower limit of the energy con-cent of the bubble at the end of bubble formation process. In other words, pool temperature af facts the so-called " bubble formation efficiency" and, thereby, the peak bubble pressure.
A12.3.4.3 Air Volume / Quencher Arna (V /A )
i The effect of air volume is rather cotepicx. On the one hand, the air column serves as a cushion to provide a low pipe-clearing pressure; on j
the other hand, more air neans a larger bubble, more energy and higher peak bettom prwsures.
l The fact that very small and very large VA both lead to negligible pressure l
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.
,042178
A-132 22A4363 Rev. 2 O
Tha strong influence of Vg/Aqon the peak bubble pressure i= plies 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 proportional to discharge pressure, affects peak bubble pressure, the latter must als;, increase with steam flow rate.
Condensation of steam on the walls of the discharge line has the effect 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 actuaticn.
Valve opening time affects the variation of steam flow rate with c1=e.
However, once the valve is fully open, the flow rate remains essentially constant for the remainder of the air clearing transient. Because air-clearing occurs af ter the valve. is fully open (in the range of practical values of valve opening timal valve opening time does not significantly affect the mean steam flow-rate or the peak bubble pressure.
A12.3.4.5 Length of Water Column (WCL)
The length of water colu=n is the submerged length of pipe to the center of the quencher. The peak pipe pressure and, therefore, discharge pressure are both affected by the length of the water column due to the longer time required to accelerate and clear a large water mass compared l l
l 042178 gl l
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22A4365 A-133 Rev. 2 O
to a scall mass. The duration of the clearing of the quencher arms i depends on the volumetric flow rate of water, which also depends on water column length. The discharge pressure affects the air flow race. '
Therefore. WCL af fects air flow rate and, thereby, the peak bubble t
pressure.
)
i Another factor which depends on WCL is the wetted pipe area available for l
cooling of the compressed air during air-clearing. This wetted area, of l
course, increases with WCL. This has the effect of reducing the energy that enters the bubble and tends to counteract the previous effect of
- WCL (Figure A12.2-8).
i 1
i e
i 9
lO i
042178 l
1
a IM 22A4365 l Rev. 2 1 1
MASS .
GF AIR
, AIR AIR PRESSURE v0LUVE d L T AND TEM.
PE R ATURE ,
BUBBLE WATER l T TEMPE R A-i b8> T E MPE R.*
TURE l TURE
._ CONTAIN.
MENT LOAOS I -
GUENCHER - QUENCHER GEOMETRY AREA 8088LE BUB 8LE PRESSURE q
< lf SIZF CONSTANf a
IME 'PE
~ -
P R ESSU R E
+
MASS OF AIR ANO
, , SUBBLE TEVPE RA- M EAR- ~
TURE CONST AN r iNG TIME VOLUME 4
NUMBER OF CUENCHERS 1 P J b AIR MASS OISCHARGE l LENGTH IW pggw gg7g
CONSTANT M A XIMUM REAC*OR STE A M P R E SSU R E ARM CLE AR.
ING TIME 8' OW 8 M ASSI FLOW AREA JL VALVE T CPENING TIME I
BUBBLE q, , % .-
~
f NITt AL I GEOMETRY AIR CONDITIONS I
OlSCH A RGE
' & LINE GECMETRY ,
OUENCHER - QUENCHER GEOMETRY AREA WATE9 SURF ACE TO
' VARY WITH SUBSECUENT ACTU ATION ; k OUENCHERAREA R ATIO et
- CONSTANT Figure A12.3-1. Relationship of Key Para =eters g 042178
A-135 22A4365 Rev. 2 O _
A12.4 CORRELATION OF POSITIVE AND NEGATIVE PRESSURE PEAKS Despite the complexity of the bubble dynamics for the quencher, a simple '
correlation exists between the peak positive and the peak negative bubble pressures. This correlation is based on the principle of conservation of energy and has been verified against the test data.
The correlation provides a convenient means for determining one of the peak pressures, provided the other peak is known. Being quite general, 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 Consider an air bubble of arbitrary geometry with a volume V, and pressure Pg (same as local absolute pressure) in thermodynamic equilibrium with the surrounding water. If the bubble is compressed to a pressure P co rre-sponding to a volume Vg and then allowed to oscillate, it will act as a spring-mass system. The pressure will oscillate between P and P and the volume will oscillata between V and V .
Conservation of energy dictates that the minimum pressure must correspond 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 sts;e as the reference state:
W = -W comp exp i
O
22A4365 A-136 Rev. 2 O
or V V min :nax PdV = - PdV (A12. 4-1)
V V o o Assuming the compression and expansion processes to be isothermal, the following relationship between P and V exists:
V P = P, V where P, = absolute surrounding pressure ;
V = initial air volume at P = =P=;
o
? = instantaneous bubble pressure; and V = instantaneous EuEEle. volu=e Rearranging:
P, V V =
p and P' V dp dV = (Al2. 4-11.
2 P
O 101678
22A4365 A-137 Rev. 1 O
~
Substituting in Equation A14.2-1 we obtain:
P P max *i" P dP dP
-P V -=- P= V --PV in
= oP oP o P, P, ,
P, P P~
I = -P V in = -P V in
= o P= = o P min which simplifies to:
P P max =
P= P
> min i
i O
- P P = P,- (Al2.4-3)
' sax min 4
- ?
'3or the case of interest, P+sos Pabs = P'= ab s g.
. ,2 W
where i
P- = minimum absoluce bubble pressure, and abs
+
P = maximum absolute bubble pressure.
abs 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 max and P That the i
min.
process is properly considered isothermal is demonstrated by comparison i to data.
(
l 042178 l
l l
22A4365 A-138 Rev. 2 In terms of gauge pressures (P+ and P") , Equation A12.4-3, by simple algebra, takes the following form:
l
~
P = P P ,/ (P + P ,) (A12.4-3A)
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-cating that the combined effects of irreversibilities result in actual bubble thermodynamics which are very well approximated by a reversible isothermal process.
Notice that large values of negative gauge pressure correspond to small
~
values of P The predictions for SWR's will be in the lower end of the abs.
45 line where the model gives conservative results.
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I 22A4365 g_139 Rev. 2 O
i 1.5 -
00 00 i
a O e5 h Ok' 00 Q o
3 1.0 -
6
% A SMALL SCALE,6 mWCL 7 SMALL SCALE 4 m WCL O LARGE SCALE, FIRST ACTUATIO*4 O LARGE SCALE,SUBSEOUENT ACTUATIONS
! I O 0.5 O$ 1.0 PREDICTED PE,(barl 1.5 2.0 Figure A12.4-1. Comparison of Eq. (A12.4-3) Predictions with Test Data O 042178 l
i A-140 22A4365 Rev. 3 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 (MPP) 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 r' 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 MPP, for both first and maximum subsequent actuations. An equation to obtain HNP values directly from MPP values is also provided.
O 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 Osmediately af ter safety /
relief valve actuation. The pressures are maximums over the oscillations.
MPP and MNP are dif ferences above and below the absolute pressure at quencher elevation, where the absolute pressure is due to atmospheric and hydrose ttic 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) ; lll 090779
22A4365 A-141 Rev. 3 (C) first actuation of an ADS valve (120 F suppression pool); and
- (D) subsequent actuation of a single valve (120 F suppression pool).
Water surface area ratio distinguishes generalized load cases 1 and 2.
Similarly, the effect of water surface as well as pool temperature dis-tinguish case 3. Generalized load case 4 must be distinguished because it was found from testing that the highest MPP and MNP occur on 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 maximum of subsequent actuations.
A12.5.1.2 Criterion 3
O The design values are to be such that there is 100 Y% confidence that at least 100(1 - a)% of actual plant MPP (or MNP) values will be less than the design values. Values for 100 % (confidence value) and 100(1 - a)%
(the percentage of the distribution of individuals) are both 95%.
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.
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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 from 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 Strategy 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 prediction.
(2) Normalize some variables for scale differences among the three test configurations and for application to the plants.
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O (3) Determine the sensitivity of MPP to each variable simultaneously, l in a prediction e.quation linear in coefficients estimated by multiple linear regression (curve fitting), thus maximizing the amount of data used to estimate each coef ficient. Retain only terms which make a statistically significant reduction in the variability of the observed values about the predictica surf ace.
(4) Predict the first actuation MPP 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 sufficiently similar to the configurations in the three tests that, after O normalization of some variables, the sensitivity of MPP in the
'J plants will be of the same magnitude as observed in each test.
Where possible, preference was given to selecting a coefficient estimated on large scale data, because that test configuration used a quenchar 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.
! (61 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 tso contributions to this variat.ce: (1) the variance of the pre-dicted value, and (2) the variance of individual values.
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22A4365 A-144 Rev. 2 O
(7) Pind 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 MPP = Abbreviation for maximum positive pressure MNP = Abbreviation for maximum negative pressure MPP1 = An observed value of MPP. on a first actuation MPPQ = An observed value of MPP on a maximum subsequent actuation MNP1 = An observed value of MNP on a first actuation MNPQ = An observed value of MNP on a maximum subsequent actuation PRD1 = A predicted value of >TP on a first actuation 4
PRDQ = A predicted value of MPP on a maximum subsequent actuation PRN1 = A predicted value of MNP on a first actuation PRNQ = A predicted value of MNP on a maximum subsequent actuation MPPDV = A design value for MPP MNPDV = A design value for MNP 041178 9
22A4365 A-145 Rev. 2 A12.5.2 Desian Value Equations for Maximum Positive Pressure and Mav4 mum Nagative Pressure Implementing the foregoing strategy, the assign value equation for MPP appears in its basic form on the first line of Table A12.5.1. That table goes on to give all subordinate equations and terms, and the design value equation for MNP, together with the succeeding sections herein where each equation is derived, or each term evaluated. Thus , Table A12.5.1 serves 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
TP has been accounted for in the term (s) used in the prediction equation.
Figure A12.5-6 shows the shell residua".s for VAAQ, with respect to that variable. These are for the s=all-scale data, from which the coef ficient for VAAQ was esti=ated. Also shown arc two straight lines, one the hort-
- ontal continuation o ? the other, which show the effect of the VAAQ term in the equation. As can be seen, the shell residuals for VAAQ reach a maxi =um and then decrease. But rather than extrapolating this decreas e , the prediction equatien was =cdified for conservatism to provide for a horisontal projection for VAAQ values exceeding 0.255.*
"That the 7AAQ values reached a maximum and began decreasing was found to be a statistically significant trend, and may be expected from physical considerations as described in Sectica al2.3. Thus , the horizontal pro-joction is appropriate and conservative. ,
t 042178
22A4365 A-151 '
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Thus, tb.e 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 permit 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.
O 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 LhTJ 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 satisfactory.
Figure A12.5-10 shows the shell residuals and effect of the water colu::m length terms (WCL 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-mated on the small-scale data, having been fcund to be not significant l
in the large-scale data, as suggested by the shell residuals' following the 0 line.
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22A4365 A-152 Re_v. 2 O
Figure A12.5-13 shows the shell residuals for AWAQ for the miniscale data, together with the ef fect of the AWAQ and AWQ2 terns. The AWA0 model is probably really asymptotic, but no use la made of it for values of AWAQ greater thsn 20.
Terms for the several variables were brought together into the single equation shown in Table A12.5.2, which predicts MPP for the first actua-tions (PRDl) . The structura of this equation is illustrated in general terms in Figure / where, in Figure A12.5-l4a, an equation linear in coefficients, in . ..adard intercept form, is illus trated. The equation is shown passing through1(x , y); and y, the predicted value at y at some x is also snown. But the standard intercept form is not convenient for 7
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 A12.5-14c. Figure A12.5-14b applies to those terms where the coefficients are estimated from the large-scale data, appropriate since f is also the average observed MPP for the large-scala data. It shows the predicted y , y , as y adjusted by a ay for x 4
Figure A12.5-14c describes the ly's for terms found as at (x7 - k) .
with coef ficients estimated on data other than large scale. Each of this s
type of ay is found as the term a3 #8 0 '
3 -*3 *3 ref is the mean value of x in the large-scale d ta, nd x3 is the ralue of x3 at which 3
andiareinthedataset from which a 3was 4
y is being calculated. x 3
estimated.
For the 238 Standard Plant, the ralue of each variable is shcun in Table A12.5.5. These values are entered in place of the variable names in the equation in Table A12.5.2, which are the x te rms in Figure A12. 5-l4.
- he actual calculation of predicted values for first actuations (PRD1) is carried out for the load cases in tonnection with calculating design values, in Section A12.5.17.
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22A4365 -A-153 Rev. 3 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 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 bwer 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 obaerved maximum subsequent actuations tend to be proportional to pre-dicted first actuations, rather than simply a fixed amount greater, for example. 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 maximun subsequent actuation from predicted first actuation was chosen passing through the origin and (x, y). Therefore, for load cases involving subsequent actuations, CMSA = 1.744.
A12.5.7 Evaluation of Tenn CONF (Confidence Coefficient)
CONF depends on the confidence statement to be made and on the number of data on which SIFV is based. The confidence statement has the form written in subsection A12.5.1.2. A value of 37 data points (the number of large-l scale data) is used for first actuations; a value of 10 data points is l
used for maximum 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."
090779
32A4365 A-154 Rev. 2 The confidence statement is valid when the distribution of individual values (in this case, of residuals about the prediction surface) is normal. That this is nearly so in the observed data is shown in Figure A12.5-16, which shows residuals 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 b oader than suggested by those data in Figure A12.5-16.
A12.5.8 Derivation of Equations for SIEV (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 = (VIF") ,
is the usual relationship between standard deviation and variance.
VIFV = VPP.D + VIND reflects the fact that VIFV is comprised of two parts: (1) VPRD, the 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
=odel in Figure A12.5-17.
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A12.5.9 Derivation of Equation 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 P?Jbl.
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 gz , is
= -
- Var y =
{' ( ay) at zg x Var x .
O 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 ~ C:MSA) :4 VVP1 and VPRM is the contribucica of the variance of the predicted maximum subsequent actuation (required by laod case c) due to the variance in C'.SA :
! VPRM = (PRD1) x VVPM.
l
!O 042178 t
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A12.5.10 Evaluation of Ters VVP1 (Variance of the Fredicted First Actuation)
VVF1 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 =ultiple 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 location of the prediction surf ace).
And in the sums of terms, sometimes called the " flaring" ter=s , the expression reflects the variance of estimate of each coefficient 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 elecent in the so-called c =atri:*
and the variance of residuals , both outputs of the multiple regression computation. The e matrix is shcwn in Table A12.5.7. The c value for pairs of coefficients 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
- he combined data set, taking advantage of the better precision of esti-mates in data sets having low residual variability. These different variances of residuals are subscripted t in Figure A12.5.13 and are tabulated in Table A12.5.7.
The $, values on Figure A12.5-18 are the values of each variable for the plant. 5.e values af x are the observed =can for variables whose coef-ficients were estimated on the large scale data, and the large-scale data value, x ref, for variables whose coefficients were esti=ated on other than large-scale data, just as distinguished in Figures A12.5-14b and A12.5-14c. For VAAQ, because the horitental portion is greater than 0.255
- Inverse matrix of coefficients in the nor=al equations.
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22A4365 A-157 Rev. 2
- O and does not involve the coefficient, x' = 0.255 was used for any cases where VAAQ 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 used for any cases where MNAQ is greater than 5.ga, i
A12.5.ll Evaluation of Term vvPM (Variance of Coefficient for Maximum Subsequent Actuation) i VVPM is the variance of CMSA, the coefficient on the predicted first Y actuation to obtain the predicted maximum subsequent actuation. Refer-ring to Figure A12.5-16, VVPM is the variance of estimate of the cos.fficient 1.744 for load cases involving the maximum subsequent actua-s tion. For load cases involving only the first actuation, where CMSA = 1.0, VVPM is not applicable and VPRM = 0.
The variance of estimate of the slope of a line through the origin .is found by the standard equation shown in Figure A12.5-19, and VVPM is evaluated in that figure as 0.01199.
A12.5.12 Derivation of Equation for VIND (Variance of Individus1 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 t
necessary to consider whether the standard deviation of residuals is apparently constant over all predicted values , or is in some way propor- l 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 daviatien of residuals should be regarded as proportional to the predicted value, both for first l actuation and for maximun subsequent actuation, according to the prediction l
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22A4365 A-158 Rev. 2 O
line shown in Figure A12.5-20. The line is based on the =ax1=um subsequent actuation data. "he average absolute residual within each 0.1 bar cell of the predicted max 1=um subsequent actuation is shown plotted versus those predicted values. The proportionality of the plotted points is clear, and is used within and beyond the range of predicted values shown. The standard deviation is obtained by the equation shown en Figura A12.5-20, the 0.798 divisor being the expected value of the upper half of the normal distribution, corresponding to the average absolute residual when the residuals are normally distributed. The p ro-cedure for calculating the expected value of the upper half of a standard normal distribution is illustrated in Figure A12.5-21.
~
A12.5.13 Evaluation of Term PROR PRCR is the coefficient to =ultiply by the predicted value to obtain the standard deviation of individual values, (VI:iD) . Its evaluation is shown on Figure A12.5-20 as 0.229. Thus, one standard deviatien of indi- g vidual values (residuals) is 22.9". of the predicted value.
A12.5.14 Derivation of Ecuation for MNPDV (Maxi =um Negative Pressure ,
Desizn Value)
- t??DV is the design value for maximum negative pressure OCTP) . It was derived in Seccion A12.4 as:
P ?_ =?;.
?' is the absolute pressure equivalent of MPP at quencher elevaticn, P' is the absolute pressure equivalent of itTP, and ?, is the absolute pressure 042178 O
I
, 22A4365 A-159 Rev. 3 O .
at quencher elevation (considering atmospheric pressure and hydrostatic pressure) . Tb obtain an appropriate pressure difference value, MNPDV = PlNF x MPPDV/(PlNF + MPPDV).
MPPDV is found in Section A12.5.3.
A12.5.15 Derivation of Equation for PINF and SUBM PINF is P, at quencher elevation, in absolute pressure units. Evaluated
- 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 meter of hydrostatic head; and SUBM is meters submergence at the centerline O. of the quencher.
J The maximum negative pressure design value correspending to any r2aximum positive pressure design value can be found using the above equations.
A12.5.16 e.atistical r 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 of fered 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. ,
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By way of further confirmation, the following two models were fitted to the maximum negative pressure (MNP) data:
MNP = Cy+C2 x MPP, and
- [ P~)
P =C + C, x l -
1
( 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 quality. Intercept C3 was n t significantly different from 0, and C was 4
not significantly different from 1.0, at even the 10% level.
In application, the predicted maximum positive pressure mtast, of course,
+
be used for P . There fore , it is of interest to fit the negative pressure data using the predicted positive pressure values in place of measured 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 C3 and C4 again not signmcantly Merent from 0 and 1, respec-tively, confirming f rom the data the appropriateness of the relation-ship P P~ =P.
+
The adequacy of P P- = P,2 using gredicted maximum positive pressures can be confirmed visually for first actuations by comparison of the residuals for large 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.
t O
042U 8
i 2214365 ;
l Rev. 2 g '
V A12.5.17 Nianerical Results for Maximum Positive and Negative Pressures 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 maximum positive pressures to show the relationship between pre-dicted and design values.
O lo 1
042178 i
22A4365 A-162 l Rev. 2 0'
Table A-12.5.1 DESIGN VALUE EQUATIONS ;71TH SUBORDUIATE EQUATIONS AND TEPliS Section Equations A-12.5.3 >TPDV = 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 A-12.5.7 CONY 1/ ~
A-12.5.8 SIFV = VIFV A-12.5.9 VIFV = VPRD + VEID A-12.5.10 VPRD = VPR1 + VPRM A-12.5.ll VPR1 = CMSA x WP1 A-12.5.12 VVP1 A-12.5.13 VPRM = PRD1 x WPM A-12,5.14 VVPM g A-12.5.15 VIND = PROR x PRED A-12.5.16 PROR A-12.5.17 'CEDV = M'iP DESIGN VALUE = PU2 x MPPDV/ QI'E
- T?3V1.
A-12.5.18 Pri? = 1.014 + G.0980.2 SU3M The ter=s are defined and tha equations are. derived in tSa Sections sHown..
3 indexes are defined in Section A12.5.3.
042178
- $2A4365
_ Rev. 2 _ A-163 and A-164 i
O l Table A12.5.2 i EQUATION FOR PREDIOTION OF PRD1, MAXIMUM POSITIVE PRESSURE 70R FIRST ACTUATION l
l a
3 4
1 1
.i _
l
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1 i (GE COMPANY PROPRIJTARY INPORMATION PROVIDED UNDER SEPARATE COVER)
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'2214365 I
- Rev. 2 A-165 thru A-174 LO Table A12.5.3 :
.~ I I, DATA AND PREDICTED VALUES f
l (10 Shaats) i l
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(CE COMPANY PROPRIETARY INFORMATION PROVIDED UNDER SEPARATE COVERl lO 3
1 d
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22A436$
Rev. 2 Table A12.5.4 VARIATION OF VARIABLES BY TEST Large-Scale Small-Scale Miniscale Dependent:
MPP Yes Yes Yes MNP Yes Yes Not Reported Independent:
VAA No SEL No MNAQ SEL* Varied ** No LNTW SEL Varied No SUBM No SEL No i
VOT Varied SEL No AWAQ No No SEL SEL = Coefficient (s) estimated on this data set was selected for j l prediction equation.
Varied = Variable was varied, but coef. not selected.
042178
A-176 22A4365 Rev. 3 O
Table A12.5.5 VALUES OF VARIABLES 10R STANDARD 238 MA10C III PLANT Ceneralized Bottom Pressure Load Case
- a. First Ac- b. First c. First d. Subsequent tuation One Actuation Actuation Actuation or Two Valves All Valves ADS Valve Single Valve Parameter (1000F Vater) (1000F Water) (1200F Water) (1200F Water) 0.23 0.23 0.23 0.23 VAAQ MNAQ 11.41 11.41 11.41 11.41 MNQ1 6.89 6.89 6.89 6.89 MNQ2 47.47 47.47 47.47 47.47 MNQJ 11.41 11.41 11.41 11.41 L;rW 3.63 3.63 3.89 3.89 l
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 7.85 20.00 20.00 g
AWAQ 20.00 3.93 AWQ2 400.00 15.44 61.62 400.00 CONF 2.15 2.15 2.15 2.91 PINF 1.43 1.43 1.77 1.43 3
Air Volume (V ) = 1.59 m 2
Quencher Area (Aq ) = 6.93 m VAAQ = VA !^Q "
Maximum Steam Flow Rate (in) = 520 metric ton /hr MNAQ = m * /Aq = 11.41 Temperature of Suppression Pool (T g ) = 37.8 C (100 F) or 48.9 C (120 F)
LNW = 3.63 or 3. 89 Length of Water Column (WCL) = 5.42 m.
WCL2 = (WCL) = 29.38 g
090779
A-177 22A4365 i Rev. 3 Table A12.5.5 (Ccatinued)
Valve Opening Time (VOT) = 20 maec.
4 Effective Water Surface Area (Ag ) = 548.05 m2 (single valve) 54.79 m (ADS valves) 27.20 m (all valves) j Water Surface Ratio (AWAQ) =gAq/A = 20.00 (single valve)*
= 7.85 (ADS valves) a '
= 3.93 (all valves) 4 MNQ1 = MNAQ if MNAQ < 6.89 MNQ1 = 6.89 if MNAQ > 6.89 MNQ2 = (MNQl) 1 MNQJ = MNAQ O Quencher Submergence to Centerline (SUBM) = 4.24 m Containment Pressure = 14.7 psia Q9.7 psia for ADS only)
= 1.0135 bar G.358 bar for ADS only)
PINF = Containment Pressure + Hydrostatic Pressure Hydrostatic Pressure = 0.098 x SUBM
= 1.0135 + 0.4158 = 1.43 bar
= 1.358 + 0.4158 = 1.77 bar (for ADS only) l O
090779
22A4365 A-178 Rev. 2 O
Table A12.5.6 VALUES FOR A STANDARD 238 MARK. III PLANT Generalized Bottom Pressure Load Case a b c d MP?DV (psid) 13.44 18.73 17.40 28,13 MPPDV (bar d) 0.927 1.29 1.20 1.94 PRED 0.603 0.851 0.790 1.11 PRD1 0.603 0.851 0.790 0.639 CMSA 1.0 1.0 1.0 1.74 CONF 2.15 2.15 2.15 2.91
, 9IFV 0.151 0.205 0.191 0.284 -
VIFV 0.0227 0.0421 0.0364 0.0805 VPRD 0.00357 0.00407 0.00363 0.0154 VPR1 0.00357 0.00407 0.00363 0.0105 VVP1 0.00357 0.00407 0.00363 0.00345 VPRM 0. O. O. 0.00490 VVPM NA* NA NA 0.0120 VIND 0.0191 0.0380 0.0327 0.0651 PROR 0.229 0.229 0.229 0.229 MNPDV (psid) 8.15 9.84 10.38 11.93 MNPDV (bar d) 0.562 0.679 0.716 0.?t3 PINF 1,43 1.43 1.77 1.43 SUBM 4.24 4.24 4.24 4.24
- NA = Not Applicable.
042178
A-179 22A4365 Rev. 2 q .
D l Table A12.5.7 c MATRIX VALUES * (Page 1 of 2) c 3.08E42EE i1 c -8.91E-04 2.75E-05 i2 c g3 0 0 2.52E-01 e g4 0 0 -6.47E-02 2.42E-01 c 0 0 3.23E-02 -2.93E-02 3.67E-03 i5 c 0 0 0 0 0 i6 c 0 0 -3. 4 8E-04 -3.30E-06 1.624E-07 i7 c 0 0 0 0 i8 e gg 0 0 0 0 0 "1j "2j "3j "4j "5j c 2.86E M i6 c 0. 4.22E-07 7
O c ia e gg
'"- 2 8.12E-04 0.
- 2 37'-
-1.953E-02 1.653E-03
- 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 VOI estimated on small-scale data 8 MNAQ estimated on large-scale data 9 MNQ2 estimated on large-scale data l ' ij "j i e for toefficients estimated on different data sets = 0
- Inverse matrix of coefficients in the normal equations. (See Figure A12.5-l'6.)
~
- " " = 10
!O 042178 i
I
A-180 22A4365 Rev. 2 _
O Table A12.5.7 (C6ntinuad)
Rnsiduals (MPP)
Mean (bar d) Variance Standard Deviation First Actuations -
37 Large-Scale Data -0.01070* 0.00938** 0.0969 70 small-Scale Data 0.00325 0.00927** 0.0938 9 Miniscale Data 0.000832 0.000493** 0.0222 Maximum Subsequent Actuations 10 Large-Scale Data 0 0.01188 0.1090 Residuals (MNP)
Mean (bar d) Variance Standard Deviation First Actuations 37 Large-Scale Data 0.00479 0.00892 0.0945 70 Small-Scale Data 0.0187 0.00214 0.0463
- In least squares fitting, the naan of residuals = 0. These are not exactly 0, due to use of an equation involvi g coefficients estLnated on sets of data other than on line shown.
- VRESID, u:. in Figure Al2.5-18.
042178 9
i 22A4365 A-1Si Rev. 2 O
(GE COMPANY PROPRIETARY INPORMATION PROVIDED UNDER SEPARATE COVER) i l
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Figure A12.5-1. MPP for First and Subsequent Actuations Positive-Large Scale -
t i
O (GE COMPMY PROPRIETARY INFORMATION PROVIDED UNDER SEPARATE COVERl i
Figure A12.5-2. MNP for First and Subsequent Actuations Negative-Large Scale i O 042178 l
l l
l
22A4365 A-182
.Rev.'2' O
\ vor AWAQ d
MNAQ 9 SUBM l
1 p _ .. _ _ _7_ _____q F1 r
~
V h
..... s _ . ,, ... .
. _ _ _ J _ _ _ _ _ _ _;
iI L _ .a
! LNTW VAAQ MNP M*P_
Figure A12.5-3. Quencher Diagram Illustrating Dependent and Independent Variables i
042178 i
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A-183 2
i 22A4365 Re_v. 2 ,
I
! Ao, m2 (OUENCHER AREA) i 0 1 2 3 4 5 6 7 8 9
. M,Ni i
i . SMALL i
l . uRc.
- STD PLANT 1
i.
VA, m3 (AIR VOLUME)
I i 0 0.4 0.8 1.2 1.6 2.0 2.4
' MINI J
- SMALL l
i iO
- -Ree
. STD PLANT VAAQ (AIR VOLUME RATIO.VA /Ac.ml 0 0.05 0.10 0.15 E20 a25 0.30 0.3
. M,Ni SMALL l . LARGE I
'
- STD PLANT 9
Figure A12.5-4. Ranges of Values / Raw and Transfor:ned Indepe.ndent Variables (Page 1 of 4) l 042178
A-184 22A4365 R ev. 2 -
O A STEAM MA$$ FLOW R ATE ltonneeths) 100 200 300 =CO 500 0
MINI i e ,
i
_ SMALL LARGE I
1
- STD PLANT I
l l
A0.7 20 3 40 60 60 70 80 l
0 10 MINI
"""" SMALL LARGE e !.TD PLANT l
i l,
MNAQ, A RATIQ [(tonnadhelo.7/m2]
o la 20 30 => so a0 70
- STD PLANT Figure A12.5-4 Page 2 of 4 i
042178 O
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22A4365 A-185 Re_v . _2_ -
VOT (VALVE OPENING TIME rm) 0 000 400 800 800 1000 1200 1400 1800 e MINI SMALL LARGE e STD PLANT l
- 'w
. (WATER TEMPERATURE,'Cl
! 0 to 20 W 40 50 N 70 80 e MINI
. SMALL LARGE i ,
e e STD PLANT a.h c.d l
LNTW 8 LOG, Twl 2.0 2.5 10 15 4.0 4.5
- MINI SMALL LARGE
- STD PLANT e.b c.d rigure A12.5-4. Page 3 of 4 0 042178 l
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~$A4365 A-186 l
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WCL, WATER COLUMN LENGTH (m)
- O I 2 3 4 5 6 e MINI
- C SMALL
- LARGE e STD PLANT Aw, m (WATER SURPACE AREAL
.0 4 8 12 16 20 24 28 MINI
! . mm a
e LARGE i e e STO Pt ANT b a.c.d AWAQ (WATER SURFACE RATIO Aw/Aq m2/m2}
O 5 10 15 20 25 30 MINI e SMALL e !ARGE
. e o e STD PLLT a d Ac Figure A12.5-4. Page 4 of 4 042178
22A4365 A-187 Rev. 2 F6 AST ACTUATIONS OBSERVED MAXIMUM POSITIVE PRESSURE O 0.1 0.2 0.3 0.4 0.5 0.6 0. 7 0.8 MINI j SMALL l
LARGE LARGE FOLLOWED BY SUBSEQUENT ACTUATIONS MAXIMUM SUBSEQUENT ACTUA110NS OBSERVED MAXtMdM POSITIVE PRESSURE O 0.1 0.2 03 0.4 0.5 0.6 0.7 0.8 LARGE O FIRST ACTUATIONS OBSERVED MAXIMUM NEGATIVE PRESSURE O 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 SM ALL LARGE l
LARGE FOLLCWED BY SUBSEQUENT ACTUATIONS MAXIMUM SUBSEQUENT ACTUATIONS OBSE RVED M AXIMUM NEGATIVi! PRESSURE 0 0.1 0.2 0.3 0. 4 0.5 0.6 0.7 0.8 l,
l LARGE Figure A12.5-5. Ranges of values; Dependent variables l
l l
l 042178 l
22A4365 A-188 Rev. 2 O
(CE COMPANY PROPRIETARY EIFORMATION PROVIDED LWDER SEPARATE COVER)
Figure A12.5-6. Shell Residuals, VAAQ Omitted, and Effect of VAAQ Term on Prediction; Small Scale Data (Coefficient Estimated on These Datal O
(GE COMPANY PROPRIETARY EIFGRMATION PROVIDED UNDER SEPARATE COVER)
Figure A12.5-7. Shell Residuals , :CIAQ Omitted, and Effect of MNAQ Ter=s on Prediction; Large Scale Data (Coef ficient Estimated
! on These Data) 1 042178 O
l l
1
i A-189 22A4365
' Rev. 2 l O ;
I (GE COMPANY ' PROPRIETARY INFORATION PROVIDED trtDER SEPARATE COVER) 1 a
Figure A12.5-8. Shell Residuals, UTJ 0.nittad, and Effect of CITJ Term on Predicticu; Large Scala Data (. Coefficient Estimated on THase Data).
O (GE COMPANY PROPRIETARY INFOR'!ATION PROVIDED UADER SEPARAIE COVER)
Figure A12.5-9. Shell Residuals, UTJ Cmittad, and Effect of LNTW Term on Prediction; Small Scale Data G2tTW Coefficient Estimated from Large Scale)_
042178 l
A-190 22A4365 Rev. 2 O
(GE COMPANf PROPRIETARY DIPJRMATION PROVIDED WDER SEPARATE COVER)
Figure A12.5-10. Shell Residuals, WCL Omitted, and Effect o f WCL Te r=s on Prediction; Small Scale Data (Coefficients Estimated from These Data)
O (GE COMPA:lY PROPRIETARY INFORMATICN PROVIDED WDER SEPARAIE COVER)
Figure A12.5-il. Shell Residuals. VOT Onitted, and Effect of VOI Ter: on Prediction; Small Scale Data (Coefficient Esti=ated from These Data) 042178 G
i l
22A4365 A-191 Rev. 2 _
4
.l j
f l
1 (GE COMPANY PROPRIETARY INFORMATIDX PROVIDED UNDER SEPARATFJ. COVERL A
Figure A12.5-12. Shall Rasiduals, VOT Omitted, and Esffect of VOT on Pradiction., Large Scale. Data OTOI Coefficient Estimated from Small Scale; not Significant in Latge.Scalal O ,
(GF. COMPANT PROPRIETARY IN70RMATION PROVIDED UNDER SEPARATE. COVERL i
i I
( Figure A12.5-13. Shall Residuals, AWAQ Omitted, and Effact of AWAQ Tar =s on Predictf.on, Miniscale Data CCoefficients Estimated l from These Datal 042178 l
l l
12A4365 A-192 l Rev. 2
' O, FIGURE A12.514a STANDARD INTERCEPT FORM: Y Y*AO + Aq kj + A2 2, g,,
l l X1 X1 Xi FIGURE A12.514b MEAN ADJUSTED FORM:
Y y Y = Y + ra Y 4e + AY 4c 1 1 ! ^y
_ ~
- v i4e - Ai te t,- XQ ,
+A2 (X2-XI2
+
Y FROM LARGE SCALE A 6's ESTIMATED ON LARGE SCALE DATA I I
X1 X1 X1 9
Y FIGURE A12.514c REFERENCE ADJUSTED FORM:
I aY t4c
- A 3($3- X3REFI Y
+aY
+A4 (X4-X4 REF) '
+
A >
A i's ESTIM ATED ON DATA OTHER THAN LARGE SCALE I I l X3REF X3 X3 X3 Figure A12.5-14. General Foms of Prediction Equations 042178 O
. - _ . - . . . ~ . - - - _ - . _. .- - - ---.__ . .- - - - = _ _ .
h A-193 1
22A4365
_ Rev. 2 1
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l 3
{-
i f
(GE COMPANY PROPRIETARY INFOKMATION PROVIDED UNDER SEPARATE COVER) x i
a i
I f
- Figure A12.5-15. Observed MPPQ Maximum Subsequent Actuations, MPPQ Versus Predicted First Actuations, PRDl; Large Scale Data t
1 i
042178 I
O 4
I t
e
-__.____m _ _ _ _ _ _ _ _ . _ . __ . _- . - . , . . . ,_ - - - . - -,1--, ,m,,- -. _, , ...-,,.m.,. . , -r , . . _
l l
22A4365 A-194 Rev. 2 LARGE. SCALE FIRST ACTUATION 37 O j l
M M I
O I -0.2 -0.1 0 0.1 0. - 0.3
" SMALL-SCALE 70 l
5 - =
1 '
. I l I l l I l I 0
U -0. 2 -0.1 0 0. 5 3.2 0.3 5
8 i
MINISCALE 3 5
0
-0.2 -0.1 0 0.1 0.2 0.3 LARGE SCALE MAXIMUM SUBSEQUENT ACTUATION 10 5
=
0
-0. 2 -0.1 0 0.1 0.2 0.3 R ESICUAL (OBSERVED.PR E DICTE D) (bort i Figure A12.5-16. Frequency Distributing Showing Nor:nality of Residuals 0
042178
I 22A4365 A-195 Rev. 2 )
]
AN IWOtviOUAL FUTURE V ALUE'
= TRUE VALUE FOR PLANT = lRROR IN PREQlCTED VALUE
- ERROR IN INOlVIDUAL VALUE ABOUT PRGDICTED VALUE DUE TO INCEPENDENCE C9 ERRORS:
VARIANCE OF INOlvlDUAL FUTURE VALUE'
= 0 + VARIANCE OF PREDICTED VALUE + VAFI ANCE OF RESIOUALS (VlFV = VPRO + VIND)
'VALUE OF FIRST OR MAxlMUM SUSSEQUENT ACTUATION MPP.
Figure A12.5-17 Error Model tv=R RESiDi,
- LARGE SCALE VVP1 3
+ E di - El Vai= 2 % I is; Io d Ip i COV (a s aj) i e
I WHERE Vs;
- VARIANCE OF ESTIMATED OF COEFFICIENT a,
- 'C,; V RESlot, COV ta ;. a g)
= C,j (VRESlot, x VRES10t3)
WHERE ti, t) REFER TO THE SET OF TEST DATA ON WHICH THE COEFFICIENTS AND C.; REFER TO ELEMENTS IN THE G, AND a j WERE th'TIMATED. AND C,i C MATRIX. VALUES OF VRESIO t; AND C,;. ETC ARE TASULATED IN TA8LE A12.5.7
'THE EQUATION FOR VVP1 ASO /E IS FOR A PRE 0lCTION ECUATION IN THE FORM OF THE ECUATION FORIIN Fis URE A12.5146.
Figure A12.5-18 Variance of Predicted Value MPP, FirsC ACCuation Evaluated aC x0 Values VARIANCE OF SLOPE THRouGH QRIGIN:
vRESIO M54 VvPM = tittus R ATED IN FIGURE A12.5 201
!PRDI
, 0 011E4 0.998
= 0.01199 I
Figure A12.5-19. Variance of Coefficient (1.744) for Maximum Subsequent l
s Acr.uation 042178
2'2A4365 A-196 Rev. 2 O
1 (GE COMPANY PROPRIETARY IN10RMATION PROVIDED UNDER SEPARATE COVER)
Figure A12.5-20. Proportionality of Average Absolute Residuals and Predicted Values, Maximum Subsequent Actuations f" a f(z) da AVERAGE ABSOLUTE DEVIATION = = 0.S f" f tzi da P-I* UI WHERE f (2) = 5 THE STANDARD NORMAL PRORA88LITY DENSITY FUNCTION APPLIES TO NORMALLY AVG ABS DEV = 0.798 OBSTRIBUTED INOlVIOUAL VALUES. USED IN THE EQUATION FOR IVINOll i2 IN FIGURE A12.5-20.
1 I
I 8
I I
I I I ii 1
-3 -2 -1 0 1 2 3 STANDARD NORMAL DEVIATES Figure A12.5-21. Derivatica of Ratio of Average Absolute Deviation to Standard Deviatirn 042178
f
^~197 22A4365 ;
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1
. (GE COMPANY PROPRIZTARY INFORMATION PROVIDED IJNDER SEPARATE COVER)
I
~
! Figure A12.5-22. Residuals for MNP Large Scale Data iO (GE COMPANY PROPRIETARY INFORMATION PROVIDED IJNDER SEPARATE COVER) t i
Figure A12.5-23. Residuals for L'P Small Scale Data 042178
l A-198 f 22A4365 3.. 2 g
i
)
i (GE COMPANY PROPRIETARY INYORMATION PROVIDED UNDER SEPARAT?. COVER)
O i
Figure A12.5-24. Observed is Predicted Values, MPP (MPPI Versus PRDI or MPQ Versus PRDQ) 042178 h
, ~,
- f2A436F lA-199 Rev. 2 o
) $ MPPOV i
' MPPOV = MAXIMUM POSITIVE PRESSURE.
DESIGN VALUE i
PROQ = PREDICTED MAXIMUM SUBSEQUENT
! ACTUATION MP' D ~
PRO 1
- PREDICTED FIRST ACTUATION MPP a
I i
f 3
i $ 20 ~
t 2, )( upov i E b { MPPOV
- w I
{ - PROQ (PRED)
' w h 15 -
8 a
2 )b MPPOV 3
iO !
3
- PRO 1 l
j 10 -
- - PRD1
- PRO 1 1
l
- 5 1
I i
O a b C d i
GENE A ALIZED 80TTCM PREbSURE LOAD CASF Figure A12.5-25. Predicted Values and Design Values of Maximum Positive Pressure from Table A12.5.6 1
O 04217d I
= -
22A4365 A-200 Rev. 3 A12.6 APPLICATION O
The purpose of this seccion is to provide the designer with a simple and straightforward procedure for calculating the maxiuum positive and negative air-clearing pressures on the bottom of the su ppression 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.
A12.6.1 Procedure All bottom pressures otesined by these procedures have a 95-95 confidence level, and are within +1.0% of the values obtained by strict application of the techniques described in the previous chapter.
The first step in determining time bottom pressure is to calculate the predicted first actuation maximum positive pressure (PkD1). Since ehe quencher device is a fixed design (Area = 6.93 m ), the maximum flow rate is 520 metric tons per hour, and the safety / relief valve opening time is at the minimum (20 mscc). The equation for PRD1 in Table A12.5.1 can be reduced to: ,
PRD1 = 0.421
+ 2.58 (VAAQ - Q.1706)_
+ 0.1377 (LNW - 3. 83)_
+ 0.206 (WCL - 4)
- 0.0176 OvcL2 - 16)
- 0.0336 (.\WAQ - 20)
+ 0.000761 (AWQ2 - 400)
O 090779
l 1
'5I4365 A-201 Rev. 2 ,
where PRD1 = mean first actuation peak positive pressure (bars);
i VAAQ = air volume in the safety / relief valve discharge line (m ) divided by the quencher area (m ), where quencher area is defined as the area of a circle that circum-scribes the quencher. (For the standard BWR/6 Mark III 2
plant, the quencher area is 6.98 m . If VAAQ is greater than 0.255, use VAAQ = 0.255.) ;
LNTW = Natural log of Tg , where Ty 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 interf ace in the discharge pipe; wcL2 = (WcL12 ;
AWAQ = Tha effective pool surface area per quencher divided by the quencher area, where quencher area, as the VAAQ, is 6.93 m (If AWAQ is greater than 20, use 20.); and AWQ2 = (AWAQ) . I I
The above formula allows for the plant unique location of the 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 this attachment.
i
- O r
101678
A-202
~22A4365 Rev. 2 O
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 (MNPDV) is obtained from the following equation:
~
MPPDV (FINF + MPPDV) where MPPDV = Design Positive Bottom Pressure (bars);
MNPDV = Design Negative Bottom Pressure (bars);
FINF = Absolute Pressure at the Level of the centerline of the quencher ar=s (bars abs.)
To convart pressure from bars to psi, a conversion factor of 14.5 psi /bar is used.
A12.6.1.1 Development of Figuras A12.6-1 and A12.6-2 The maximum positive bottom design pressure (MPPDV1 is a function of PRDL. This fu tctional relationship was described in A12.5 and can be summarized in the following form.
042178
22A4365 A-203 i Rev. 2 ,
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I (GE COMPANY PROPRIETARY INFORMATION PR077DED UNDER SEPARATE COVER) j Al2.6.2 Exarples Given the following input, the design bottom pressures for the four cases described in Section A12.5.1.1 are calculaced:
Air Volume = 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 l
t 042178
22A4365 A-204 Rev. 2 l Water Column Length = 5.42 m Submergence to Centerline = 4.24 m 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 =6.93 1.50"
- LTN = In 37.8 = 3.63 for cases a and b In 49.0 = 3.89 for cases e and d WCL = 5.42 WCL2 = 29.38 548 Therefore, 20 is used for cases a and d.
AWAQ = 6.93 = 78.5 2.
97 2 = .9 r case b 6.93 }
5 6h3 = 7.85 for case e AWQ2 = (20) = 400 for case a and d (3. 93) = 15.44 for case b (7.85) = 61.62 for case e i
I O
101678
'~22A4365 A-205 Rev. 2 PINF = .5 + (0.098 x 4.24) = 1.43 bar for case a, b and d bar for case c f4.5 + (0.098 x 4.24) = 1.77 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.
PRD1 = 0.421
+ 2.58 (0.23 - 0.1706)
+ 0.1377 (3.63 - 3.83)
+ 0.206 (5.42 - 4 0)
- 0.0176 (29.38 - 16.0)
- 0.0336 (20 - 20)
+ 0.000761 (400 - 400) = 0.604 bars From Figura A12.6-1, for PP.D1 = 0.604, MPPDV = 0.93 bars. MNPDV is then calculatad:
= 0.56 bars MNPDV = 1.43 g,43 O.93)
Converting to psi ve get:
j MPPDV = 0.93 x 14.5 = 13.49 PSID MNPDV = 0.56 x 14.5 = 8.12 PSID 042178 O
I _ . . - - . . - _ -
l 22A4'365 A-206 Rev. 2 0!
A12.6.2.3 Case 's - First Actuation of All valves (100 F Pool Teciperature)
PRD1 is calculated from th2 equation in Section A12.6.1.
PRD1 = 0.421
+ 2.58 (0.23 - 0.1706)
+ 0.1377 (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 Figure A12.6 1, for PRD1 = 0.85, MPPDV = 1.28 bars MNPDV is then calculated:
= ars MNPDV = 1.43 1.43)
G.28 O
Converting to psi, we get:
MPPDV = 1.28 (14.51 = 18.56 psid MNPDV = 0.68 C14.51 = 9.36 psid A12.6.2.4 Case d - First Actuation of an ADS Valve. G20 ? Pool TemperatureL PRD1 is calculated from the equation in Section A12.6.1 PRD1 = 0.421
+ 2.58 C0.23 - Q.17061
+ 0.1377 0.89.- 3.83L
+ 0.206 (5.42 - 4.0L
- 0.0176 (29.38 - 16.01
- 0.0336 C7.85 - 101
+ 0.000761 C61.62 - 4001 = 0.79. bars 042178
-~22A4365 A-207 Rev. 2 .
uom Pigure A12.6-1 for PRD1 = 0.79, MPPDV = 1.2 bars. MMPDV is then en1culated:
MNPD7 = 1.77 = 0.72 bars G.2 1.77)
Converting to psi, we get:
MPPDV = 1.2 (14.5) = 17.40 psid MNPDV = 0.72 (14.5) = 10.37 psid A12.6.2.5 Casa d - Subsequent Actuation of a Single Valve (120 F Pool Temperature)
PRD1 is calculated from the equation in Section A12.6.1.
O PRD1 = 0.411
+ 2.58 (0.23 - 0.1706)
+ 0.1377 0.89 - 3.83)
+ 0.206 (5.42 - 4.01
~ 0.0176 (29.38 - 16.0)
- 0.0336 (20 - 201 i
+ 0.000761 (400 - 4001 = 0.65 bars From Figure A12.6-2 for ?RD1 = 0.64, MPPDV = 1.95 bars. MNPDV is then calculated: l l
l MNPDV = 1.43 = . ars (1.9 1.43) 042178
A-208 22A4365 Rev. 2 O.1 Converting to psi, we get:
i MPPDV = 1.95 (14.5) = 28.23 psid MNPDV = 0.825 (14.5) = 11.96 psid 1
O
)
042178 O
__ _ _ _ ___ _ _ _ _ _ _ _ _ _ m ____ - _ _ _ _ . _ _ _ ___ ___ __ _ -__ _
22A4365 A-209 RIv. 2 -
2 O NOTE: SEE EQUATION OF THIS CURVE. l PAGE A 203 3
I E ~
i >
! 2 i
l 1
I l
2 4
0 0 0.1 0.2 0.3 0.4 0.5 0.6 0;7 0.8 0.9 10 PRO 1 tbars dl i Figure A12.6-1. First Actuation Design Pressare Versus Predicted Fressere for First Actuation 3
4 I
O I
NOTE: SEE E'1UATION OF THIS CURVE.
PAGE A 203
.e, f
2 -
3 l
h i
1 -
i I
i i
l ! ! ! l
! ! l l 0,1 (,, . 2 3.3 0.4 0.5 06 0.7 08 0.9 1.0 3
PR D I ibars di l
Figure A12.6-2. Subsequent Actuation Desis;n ?ressure Versus Predicted Pressure for First Actuation 042173 i 4
.1 I
A-210 22A4365 Rev. 2 O
A
12.7 REFERENCES
- 1. Test Results Emploved by General Electric for BWR Containma.nt and vertical vent Loads,, Class III, October 1975 (NEDO-21078),
- 2. Safetv-Relief Valve Discharge Analvtical Models, May 1975, (NEDE-20942-P).
f
- 3. Comoarison of Safetv-Relief Valve Model Predictions .f.th Test Data _,
July l'.-75 (NEDE-21062-?) .
O 4
042178 O
- 9. ,
2fA4365 Rev. 2 l
\
ATTACHME:IT B 31.0 SUPPRESSION POOL SEISMIC INDUCED LOADS (To be provided by A.E.)
Bl.1 HYD%) STATIC PRESSURE During both verticsl and horizontal accelerations , the hydrostatic pres-sure distribution in the suppression pcol undergoes distortions that lead to dynamic loads on the suppression pool boundary.
Bl.2 VERTICAL ACCEL % RATION 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-n1tude of the vertical acceleration, i.e. , at any point in the suppression pool the hydrostatic pressure P , is given by H
P H" +"8 (1 4 ( ) in,2 where 3 I o = specific weight of water, lb/ft I E = depth, ft
. 9
- a = vertical acceleration, f t/sec~
j g = gravitation acceleration, ft/sec
- O 101678
22A436'S 3-2 Rev_. 2 .
O\
Bl. 3 HORIZONTAL ACCELERATION During horizontal acceleratf.ons there vill 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 normal hydrostatic pressure distribution be modified as follows.
For those structures which are providing this accelerating force, the hydrostatic pressure at any point, Pg , is given by
. loH ti a/gh lb H" 144 + 144 2 vhere p a specific weight of water lb/ft
'I = width of suppression pool water that a particular point O
on the structure is accelerating, ft a = horizontal acceleration, ft/sec g = gravitational acceleration, ft/sec On opposite surf aces, design adequacy is es tablished assuming both normal hydrostatic pressure and no hydrostatic pressure on these l surfaces.
l l
l 101678 g
C-1 22A4365 R.ev. 2
(
t ATIAC24ENT C WEIR ANNULUS BLOCKAGE The fellowing 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 NED0-20533 (Ref.1) and was generated assuming that the annulua water associated with the blocked region still has to be accelerated during the vent clearing transient but that the corresponding vent flow area is not available. During vent clearing, the water in the blocked area was conservatively assigned to the unrestricted vent stacks.
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.
l I
O l
l O
l l 042178 l
i i
22A4365 C-2 Rev. 2 8
U e '
Z
=
U u
w 2
w l
I w ~
R o
5 =0 C
C W
> u c C 2 ~
< a u
Q -I mu
< *~
2 c c
< c -
NM 0
8
' u w
"J 3
w t
U c
e a co w
\
' ! l l o
l e , " n ,
o (IS4 BSY3tf3Ns gwngg3yg 333gga 042178 l
l I
D-1 22A4365 Rev. 2 AITACHMLVI D DRYWELL PRESSURE DISTRIBUTION
\
)
INTRODUCTION 1 The purpose cf this attachment is to show the resulting pressure differ-ential across a given level with some flow restriction assuming a 25%
restriction in the drywell.
HPV s + '
x O DRYWELL 25% RESTRICTION i
f RPV
/ 25% RESTRICTION ORYWELL
- A
/
V E
The The greatest pressure differential would occur during a steam break.
flow rate is:*
D "
at t= 1 sec m g = 28,200 h g = 546 3
1D' '"
Et = 1,230 h = 1,190.9 lbm g sec g O
- Data obtained from Table D.1 101678
l 22A4365 D-2 Rev. 3 The quality of the break flow is:
gas 1,230 X= m ,
29,430 X = 0. 0418 From the quality, the enthalpy of the break is:
ho = (1 - X)h .r + Xh g h, = (1 - 0.0418)(546.0) + 0.0418(1190.9) h, = 573.0 Assuming constant enthalpy process and a final pressure of 14.7 psia, the final quality can be calculated O
h o
=h g
- (1 - X)h fg 572.95 = 1150.5 - (1 - X)970.4 X = 0.405 Using this quality the final specific volume is v= (1 - X)vg+Xv8
= (1 - 0.405)0.016715 + 0.405( 26.80) y= 10.86 ft /lbm 1
090779 9
22A4365 D-3 Rev. 2
~
l The differential pressure is then calculated using the following formulc:
i
- 2 AP = 1/2 kov but v =
2 pA Therefore
'2 ,
AF = 1/2 K ,
DA'
~
where the A is the remaining unrestricted area ,
+
A=A '
Drywell A = 0. 75(3402.0)
A = 2551
- O and K is the loss coefficient. This is maximum for an orifice'.
1 K= 2
= 2.778 0.6 Therefore we now have lbm Ft sec' * ' b*
9 ' '
2 '
29.430 (10.86) lbfsee' 1 ft' AP = 1/2 (2. 7781 32.21bmfc 144 2551 f4 ,
2
= 0.433 osid i
I
)
() 042178
,-..---,..,-.v- - , . , , ,-,~ - - . , - - - . - - . - - - - - - -
.22A4365 , D-4 ,
-Rev. 2 .
Table D.1 ,
REACTOR PRIMARY SYSTEM SLOWDOWN FLOW RATES AND FLUID ENTHALPY MAIN STEAM LINE BRZAK Liquid Liquid Steam Steam Time Flow Enthalpy Flow Enthalpy (sec) (lbs /sec) (Btu /lb) (1bs /sec) (Btu /lb) 0 0 351.6 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 20 15,400 12,270 533.2 513.2 2,220 2,435 1194.5 1198.7 g
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 1,343 300.1 460 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 l
80 1,510 226.3 56 1166.8 l 85 1,472 222.7 45 1165.6 l 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
.g 8
'~
22A4365 O aev. 2 -
Table D.1 (Continued) i i At the end of blowdown these rates are as follows :
1 1
Primary System Liquid Flow Liquid Enthalpy
- Time (sec) (ib/sec) (Btu /lb) 1 399.98 2755 177.8 400.00 2755 177.8 l
- 800 2755 162.8 4
1400 2755 156.3 1799 1720 156.8 -
1800 1720 157.8 j 2400 1720 163.7 l 3500 1720 161.8 10,000 1720 152.7 5
5 2.1 x 10 1720 148.7 5 #
1.0 x 10 1720 126.4 5
2.5 x 10 1720 114.3 5
i 5.0 x 10 1720 106.9 5
7.5 x 10 1720 104.1 1.0 x 10 1720 102.2 i
1 l
O ,
! 042178 i
j E-1 22A4365 Rev. 3 ATTACHMENT E DRYWELL 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 j
after the reflooding of the reactor vessel. It is a bounding end point
! calculation that leads to the maximum theoretically possible negative pressure.
! CALCULATION Somewhere between 100 and 600 see the ECCS system will flood the vessel l
causing instantaneous condensation of steam in the drywell. At this time 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 ) VDW
%W" g where 1
P = Pressure in D.W. initially = 16.7 psia Py = Partial pressure of vapor = $P sat R = Temperature of D.W. = 135 F l ft-lbm i R = Gas constant = 53.34 lbW R 090779 i
22A4365 E-2 Rev. 2 3
V = V lume D.W. = 274,500 ft DW
$ = Relative humidity = 0.40 1
P sat " 135 " Sat. Pressure at 135 = 2.5365 psia Therefore M [ 16.7 - 0.4(2.5365)] (274,500) 1 0 in. '
DW " 53.4(540) 2 ft g = 21,501 lbm of air Initial Mass in Contoinment (P - Pv) V cont g ,
con RI where P = Pressure in containment initially = 14.7 psia Py = Partial pressure vapor = $ P sat V = Volume of containment = 1,138,750 ft 3
-lb R = Gas constant = 53.34 gR T = Initial temperature = 80 F
$ = Relative humidity = 0.20
- Psat " Sat Pressure = 0.5067 psia 042178 g
i E-3 22A4365 Rev. 2 O __ .
[ 14.7 - 0.2 (0.50471! (1,138.750)
M C "C
= *+- -
1 in.2
, 53.34 (540) 2 ft 4
! M = 83,110 lba of air coat j
i 4
From the above air masses tne post blowdown containment pressure can be calculated.
l 4
1
=
IMRT P
l cent V + $P sat
- e j
1 <
IM =
Summation of initial air mass in cent and
, D.*J. = 102,645 lbm air t
lbm-ft i R = Gas constant = 53.34 nFOR l
t i
j T =
Final temperature = temperature of pool = 170 F i
V C
= Containment volume = 1,138,750 ft 3 4
j
) = Final relative humidity = 1.0 4
=
l P,,g Saturation pressure at 170 = 5.990 I
042178 4
1 l
i I
1 22A4365 E-4
. Rev. 2 p , 102,645 (53.34) 630 + 1. 0 (5. 990) cent 2 (1,138,750)144 "j ft
= 27.025 psia To evaluate the minimum drywell pressure at this time the following assumptions are mace:
(1) All steam in the drywell is condensed.
(2) ECCS flow out of vessel is at temperature of 170 F.
(3) Assume all the air has been purged out of drywell pressure.
(4) No vacuum breakers.
O j Using these assumptions the final drywell pressure is equal to the saturation pressure at 170 F.
l P
DW " sat l70 Therefore the negative pressure load across the drywell wall is the difference in the final pressures of the contain=ent and drywell.
l D" DW - cent
= 5.990 - 27.025
= -21 psid 0
042178
- E-5/E-6 22A4365 Rev. 2 t
The above represents a very conservative bounding calculation of the maximum theoretical negative pressure.
1 The assumptions that noncondensibles return to the drywell via the vacuum relief system and that the steam temperature in the dr>vell ,
I instantaneously drops to the suppression pool temperature are both very conservative. In addition, the real estimate of relative humidity in containment is 50% rather than the 100% assumed.
j An evaluation of the probable transient condition for this phase of the LOCA leads to the conclusion that the realistic negative pressure is less
, thsn 8 psi.
]
i i
1 l
4 i
l O
i 042178
22A4365 y_1 Rev. 2 4
ATTACHMENT F l
WETWELL ASSYMMETRIC PRESSURES INIRODUCIION 1
The purpose of this attachment is to determine the pressure gradient under the HCU floor during pool swell due to flow restriction at the HCU level.
CALCULATION 4
4 During pool swell the following conditions exist in the wetwell.
j AP gg = HCU floor pressure differential = 11 psi i
2 j Agg = Open area of HCU floor = 1500 ft 2
3 Xg = X-Section area between Pool and HCU Ficor = 400 ft
! (Vertical plane) 1
- k. = loss coefficient through floor =5 -
o = density of flow through floor = 20 ft I
From this information the flow rate through the HCU floor can be calculated.
V = A_,72 6p
- dCU g HCU ,
, \
1500 I" ( 1h*-f* I l = 2(20)(11) 144 32.2 ,
2 l 8 ft i
lb-ft-sec/
lbm i = 958,968
' SBC O 042178
22A4365 F-2 Rev. 2 l
O In ordar to calculate. a di.f ferential pressure. under the. HCU._ floor, assuma a 25% restriction at tee. ECU leyel.
25% RESTRICTION I
CRYWELL CONTAINMENT Using this 25% restriction then 1/8 of the flow will be horizontally diverted in both directions.
1/8 FLOW N e
\\ 1/8 FLOW v
This aorizontal flow rata is M
'd 8 958,968 8
042178
F-3/F-4 22A4365
_Rev. 2 l
D* 1 k=119,770 see.
Assuming the density is constant and using the above flow rate, the differential pressure under the HCU floor due to this restriction can now be calculated.
v2 ,
4'2 AP = 4 g but V' = 2 2 A=X A o *A K(p )ii' 2(g)o2(Xf) 2 1 (20) 119,770 ft 2 (32.2) 202 (400 ) 144 in 2 O = 0.483 osi t
042178 l
l
22A4365 G-1 i
Rev. 3 O
ATTACHMENT G SUBMERGED STRUCTURE LOADS DUE TO LOCA AND SRV ACTUATIONS TABLE OF CONTENTS Section Title Page Gl. INTRODUCTION G-2 i
, G2. SUBMERGED STRUCTURE LOADS DUE TO MCA G-4 G2.1 Compressive Wave Loading G-4 1
G2.2 LOCA Water Jet loads G-4 G2.3 MCA Bubble Loads G-5 I
G 2. 3.1' LOCA Bubble Loads - Sample Problem G-17 G2.4 Fall Back Loads G-23 G2.5 LOCA Condensation Oscillation Loads G-24 I
j G2.5.1 LOCA Condensation Oscillation Loads -
Sample Problems G-25 G2.6 MCA Chugging Loads G-27 G2.6.1 LOCA Chugging Loads - Sample Problem G-30 G3, SUBMERGED STRUCTURE M ADS DUE TO SRV ACTUATIONS G-33 G3.1 Quencher Water Jet Load G 33 G3.2 Quencher Bubbla Load G 33 G3.2.1 Quencher Bubble Load - Sample Problem G 37 G4. REFERENCES G 42
.l i
l 1
O 090779
22A4365 G-2 Rev. 3 Gl. INTRODUCTION h
In the following two sections, the flow induced loads on structures submerged in the suppression pool due to loss-of-Coolant (LOCA) and Safety Relief l
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 submerged structure. This compressible wave loading is negligible (discussed in subsection G2.1) . Folluving this compressive wave, the drywell is rapidly pressurized. Tha water in the weir annulus and dryvell 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 submerged structures (discussed in subsection 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 C2.3) . These air bubbles cause the pool water surface to rise until they break through the pool water surface. The pool surface slug decelerates and falls back to the original pool level (f all back, loads are discussed in subsection G2.4) . Now the stream from the break fills the drywell space and is channeled to the pool via the vent system. S team condensation oscillation starts and the vibratory nature of pool water motion causes an oscillatory load on submerged structures (discussed in subsec-tion G2.5).
O 090779
22A4365 G-3 Rev. 3 .
O 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 j i
observed which causes an acoustic load on submerged structures (discussed in the subsection G2.6) .
l 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 bubbles formed between the arms of the X-Quencher. These air bubbles are smaller in size than the U ,A 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 discuused in the subsection G3.2. -
The material in Attachment G is organized as follows:
(1) The specific analytical model is referenced, this is followed by (2) A load calculation procedure which is a sununary of the engineering process . This is followed by (3) A sample problem which demonstrates the use of the procedures.
O 090779 i
l l -- - _ _ . _ _ _ . . _ _ _ _ _
22A4365 G-4 i l
Rev. 3 l
O G2. Sl'2ERCED STRUCTURE LDADS DUE TO LOCA G2.1 Compressive Wave Loading As discussed in Section 6.1.1, the very rapid compression of the drywell air theoretically generates a compressive wave. But as pointed out in Sec-tions 6.1.1 and 6.1,2, there were no load; recorded on the containment wall 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 ca'ased 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 flo field in the suppression pool. This induced flow field creates a ll) dynami ' load on structures submerged in the pool.
However, this dynamic load 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 observation. Since the air bubble dynamic loan in bounding, this load is conservatively used in place of the water jet loaa (for air bubble load, see paragraph G2.3.) .
O 090779
22A4365 G-5
{
- ( Rev. 3 G2.3 LOCA Bubble Ioads i
During the initial phase of the DBA, pressurized drywell air is purged into i the suppression pool through the submerged vents. Af ter 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.
1 The bases of the flow model and load evaluation for the LOCA bubble-induced submerged structure load definition are derived from the model in Ref-erence G4 3. The following procedure is reconnended for calculating the loads on submerged structures.
- 1. Bubble Data
( Specific data that must be obtained are:
R: initial bubble radius, assumed to be the same as the vent 1
radius, ft.
I P,: Bubble pressure shown in Figure 4.4 (page 4-12), in psia.
p : air density corresponding to drywell conditions when the 3
drywell pressure is P,, ib,/f t .
l p: pool liquid density, Ib,/f t .
P: containment air space pressure, assumed to be constant at c
14.7 psia.
P,: initial pool pressure at the top vent centerline submergence, psia.
O f 090779
22A4365 G-6 Rev. 3 H: initial bubble submergence, same as the vent centerline h y
submergence, 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 structures, or when breakthrough occurs if the bubble does not engulf the structures.
- 3. Initial Bubble Location Initially the bubble center is assumed to be located on the vent axis at a distance equal to one vent radius from the vent exit.
- 4. Movement of Bubble Center The bubble movement consists of two components of motion:
O (a) As the bubble grows, the ~, ubble center is displaced hori-zontally away from the dryw<.11 wall. It is assumed that the bubble is always in contact with the drywell wall. Thus the he -izontal distance from the drywell wall to the bubble center at time t is equal co the instantaneous bubble radius, R(t),
at time t.
(b) The buoyance displaces the bubble center upward by Hy - Z b('}
from its initial submergence of Hy.
Ef fects (a) and (b) together determine the bubble center trajec-tory as a function of time.
O 090779
22A4365 G-7 Rev. 3 s
- 5. Bubble Dynamics The coupled bubble dynamics equations and the bubble rise equation given below can be solved for:
R(t): the bubble radius at time t 4
i(t): the bubble growth rate at time t l
- E(t)
- the rate of change of bubble growth rate at time t.
Z (t): the bubble center submergence at time t assuming no change b
j in water level.
I i Bubble Dynamics Equations:
4 R = (PB - P,) - k (G2.3-1)
'P "
1 #Cb ~
B
- B 3 p R 4wR o .
I P, ( )
PC+0k
=
- where P = bubble pressure at time t, psia B
in B
= bubble charging rate, Ib,/sec i
k = rate of specific heats for air 4
O l i
090779
22A4365 Rev. 3 @8 Bubble Rise Equation:
- 17 wp C 2- 4 D
R Z 3l(l 7 wp gR 3+gg-g = , ,
(G2.3-4) m wp R
_B+ _
where k = dg/dt = bubble rise velocity
'k = d Zb/dt = bubble rise acceleration C = drag coefficient of bubble D
g = bubble air mass e
Initial Conditions R(0) = Rg = vent radius k(0) = V g/4 where V = final water jet velocity (f rom Figure G2.3.1) 1 PB (0) = P o
Zb (0) = Hy
~
((o) := 0 mB (0) = wR p e
090779
33 *
- w o ,
. 0 2
RR A A E I 8 _
E L 1 _
R L _
A C C _
E T T -
L C N E
N E -
T N
V V E L E O M
6 V D T I 1
P 0 T y O 1 O t T MB I i l l l Il , ll i I I i
- - - c o
C C C l V V v e T M S T 4 V N 1 1
E t V e M J O
T r
- T e O t
- B 2 a 1
1 W
- l 1 l!1 II l l 1 l l I I 1 l
t n
e t
V c
E e s l a
o L (
DT 0 E t DN I E 1
1 M n MV I
T o z
l l l 1 l l l l l l I i i l l lI
_ r l
o I
_. )
1 8 I F ! I
._ E 0 I
- R T
( N k 3 E G3 V r N6 0 P a IS2 O M
- U O T DD i 6 .
EE 0 1
. TN A L 3 R E E
D 2 N
E O C
0 u I
g GT
._ Y i IF L F SAI HN TA 2
0
- - - - - O
-- - o o 0 0 0 0
7 0
6 s 4 3 m 1 j}C > Es O
oU 4
22A4365 G-10 Rev. 3
- 6. Struc.ture Data !
Location of structure, including elevacion, distance from the drywell wall, and distance from the containment wall.
Dimensions, shape, and orientation of structure. Long structures should be divided into smaller segments (with each segment approxi-mately 2 f t long) for more precise evaluation.
- 7. Distance to Structure Determine the following at time t r:
g the distance from the real bubble center to the center of structure or structure segment D,: the cross-sectional dimension of structure or structure segment in the direction of r,
- 8. Check Structure / Bubble Contact i
l l At anytime t using the value of bubble radius, R(t), from Step 5 check if R(t) 2 (rg - D,/2). If true, then loading calculations for the structure or structure segment under consideration end, j
because it is inside the bubble and the drag forces are zero.
I If not, proceed with Step 9 until the bubble breaks through the pool l
surface.
- 9. Pool Boundary Effects To account for the ef fects of pool walls, floor and free surface, use the method of images as described in Section 4.10 of Ref-erence G4.3. At any time t, first determine r t , which is the i
distance from the center of the structure or structure segment ch source or sink image (x , y , z ),
(x, y, z) to the i O
i 090779
, 22A4365 G-11
! Rev. 3 .
i (G2.3-5) 1
(* - *i) +(Y-Y) i + ( * -
- i)
Note that from Step 7, r9corresponds to the real source. Then evaluate the functions X, Y, Z as given by Equation A67 of Ref-
. erence G4.3. Using the simple notations adopted here, these functions may be written as I
N (, _ , )
X = K 3
i=0 1
- i(y - yi) -
Y = K (0
- 0) 3 i=0 1 N
i p) y Z =
K2 p t(z - z )
3 i=0 1 where the + sign is for the real source and source images, the
- sign is for sink images,1 is the total number of images con-idered, and K is a factor sed for finite bubbles to satisfy the I local pressure boundary co.idition at the real bubble surface, i.e.,
the pressure at the real bubble surf ace equals the independe.ntly calculated bubble pressure, P
- B
. 10. Number of Images ,
I A sensitivity study should be conducted to determine the number of images to be included to provide the accuracy the user desires.
As a starting point, the images shown in Figure C2.3.2 should be considered.
090779
22A4365 G-12 Rev. 3
. G l . . .
++ +
+ ++ ++ :
. -+ - -e. +
-+ - + + ++ -+
++ ++ ++ +
+ ++
++ ++ ++ l
+ + +.
- 0. e. + -e. -e- -4. -#- ++ -
_' +
++ ++ M@ ++ ++
i + ++ ,
++ ++ ++
-e. +.
e e.
-e. e.
-e. .
++ + + ++ +
+ ++ ++.
+. +
LEGEND: ++ +
h REAL SOURCE (BUBBLE) 4 IMAGINARY SOURCE p4 4
.q. IMAGINARY SINK 4 FREE SURFACE
- - POOL BOUNDARIES DRYWELL CONTAINMENT Fig * . . 3- 2 . Arrangement of Images
- G 090779 I
I 22A4365 G-13 Rev. 3 l r
- 11. To account for multiple vents effects, assume all bubbles are formed synchronously and evaluate the parameters X, Y, and Z from Equation (A80) of Reference G4.3 or by expending Equation (G2.3-6) to include all bubbles to be considered and their images. Again the neber of bubbles or vents to be considered depends on the accuracy desired.
- 12. Direction of The Flow Field f
The direction of the flow field at time t is determined by the unitvector,$,where X$+Y$+Zn
$= Y - (C2. 3-8)
I X +Y A Z
- 13. Acceleration and Velocity Using the results from Steps 5 and 11, the equivalent uniform acceleration at time t at the structure location in a finite con-tainment is 0,(t) =
R (t) *k(t) + 2R(t) k (t) X +Y +Z (G2.3-9)
The corresponding velocity U,(t) may be obtained by numerically integra ting (t). As a first approximation, U,(t) can also be evaluated from U,( t) =
R (t) k(t) X +Y +Z (C2. 3- 10) i O
090779
22A4365 G-14 i
Rev. 3 l
14 . Drag Forces h The acceleration drag is calculated from j
r b""(t)V o (C2
- 3-ll)
FA (t) =
8e where U is the acceleration component normal to the structure and V is the acceleration drag volume (from Tables C2.3.1 and A
C2.3.2 for flow normal to the structure) .
The standard drag force is calculated from
"" ( *
~
}
F,(t) = CD A C
where C is the drag coef ficient for flow normal to the structure.
An is the projected structure area normal to U (t).
at any time t to get the total load on the structures Add FA "" S segment.
The direction of total drag is normal to the submerged structures.
090779
l 22A4365 G-15 Rev. 3 C< Table G2.3.1 ACCELERATION DRAG VOLUMES FOR 'NO-DIMENSIONAL STRUCTURAL COMPONENTS (LENGTH L FOR ALL STRUCTURES)
SECTION THROUGH BODY AND UNIFORM HYDRODYNAMIC ACCELER ATION DRAG SODY FLOW DIRECTION MAgg 6 VOLUME V g R
CIRCLE % ,,g2L 2sR2L
- b Y
- wege+b)L ELLIPSE : r ase2L 4L h
b '
E LLIPSE : ; ,e2L .bt s )L 9 -.
PLATE : '
edL 2e A t
O 2b --*' 37 "
ase2L eL(4tr+ss)
RECTANGLE : r 2e , ,
h 5 1.21 ase2L eLt dar+1.21re) 2 1.36 ase2L eLietr+ 1.3ere) 1 1.51 ase2L eL(4tr+ 1.51e al 1/2 1.70 pes 2L ettatr+1,70. )
1/5 1.98 ase2L ett 4tr+ 1.90w el 1/10 2.23 pes 2L eLidtr+ 2.23e s) ,
e---2b %
'I 2 0.86 aws2L eL(2tr+0ASeal
- 2 0.76 awe 2L eLt2b+0.76es)
DIAMOND 2b 1 1/2 0.67 ave 2L eLt2tr+0.67ee) j 1/5 0.61 pes 2L eLt 2tr+ 0.61r s) c% % e/c=2.6. b/c=3 6 c
r J
O AM =
s t;
'- >" '2 " ' 2 2- - " '
4-- 2b -o-090779 l
l
22A6365 c-16 Rev. 3 Table G2. 3.2 g ACCELERATION DRAC VOLUMES FOR TliREE-DIMENSIONAL STRUCTURES HV ORODY N AMIC ACCELE R ATION DR AG 800Y AND FLOW DIRECTION gass 6 VOLUME VA DESCRIPTION
- 6/3 a R3 8/3 R3 CIRCULAR ,
DISK ,
/f ..
/ a n/6 ba2 n/6 b,2
' ^ 3 0.9 a e/6 ba2 0.9 e/6 ba2 05
[ 2 0 826 e e/6 ba2 0.826 n/6 ba2
- 1.5 0.748 a e/6 ba2 0.748 e/6 ba2 1.0 0.637 a eI6 ba2 0.637 s/6 be2
/ w.
A -
9 1 0.478 a e/4 e2b o 478 =/4 a2b b
/ 15 0 680 a n/4 a2b 0680n/4a2b HECTANGULAR / 2 0.040 a s/4 a2 b 0 840 n/4 a2b
/
PLATE
/ 2.5 0.953 a ele a2b 0.953 =/4 a2b
./ 3 e ele a26 el4 e26
- ~
a el4 a2b w/4 a2b ca3lT AN #1 2 ahTANo1 2 e e J
~
T HI ANGl'L AR E = -
PLATE [.
\
r n
3 EPH E R E * - - * - a 2/3e R 2.R3 0
090779
22A4365 G-17 Rev. 3 i
G2.3.1 LOCA Bubble Load - Sample Problem As an example the drag force on a cylindrical structure induced by the LOCA bubble from one vent will be calculated, including the boundary effects. Fig- ]
ure G2.3.3 depicts the Mark III Containment and the submerged structure in (
question. 1
{ Step 1:
i I The following data were used in the sample calculation:
l (a) Initial bubble radius: R g = 1.1458 ft t
) (b) Initial bubble velocity: vf = 15 ft/sec 3
1 i (c) Initial bubble submergence: Hy = 7.5 ft i
i bd (d) Pool water depth: H = 20 ft (e) Stagnation bubble pressure: P, = 36.5 psia (f) Containment air space pressure: P c = 14.7 psia (g) Pool liquid density: P = 62.1 lbm/f t l
(h) Air density corresponding to Po : p =
0.14 lb,/f t (i) Subble drag coefficient: Cp = 2.5 i
(j) Top vent bubble charging rate: See Figure G2.3.4.
i
- For simplicity, in this sample problem, we assume P, equal to the maximum (drywell pressure) .
'O 090779
22A4365 Rev. 3 G-18 O
I CONTAINMENT Z
I d -
DRYWELL ,
VENT 8
. .._ t
.__ .* x- t- - 2 ft -
VENT A -
VENT C ,g
)
O
\
f_1/_=___=_-l - --
/z,b
/
e.
~
p, -
I'/
/ 1
/ 7 :__ . ,
/s/uussa,/sstes d BASE MAT Figure G2.3.3. Mark III Submerged Structure for Sample Calculation 090779
22A4365 Rev. 3 G-19 O
FULL 4CALE REPRESENTATION OF MARK 131
. 1 Yj i 4
e dl di db p
- 6 IL ,d b b l' Q (n.y,al Q
- .'I . ; p" ,
=-
(no.Vo. M _
H H -
4 i
ip _
m
- L r I L T O
- D*R m :
Rm
=
TOP VENT sf BEING /\
CONSIDERED C D/2 TOP VENT BEING CONSIDERED X TOP VENT O SU8 MERGED STRUCTURE
- 1. CONSIDER ONLY TOPMOST VENT
- 2. CALCULATE FORCES DUE TO TWO NEAREST VENT EXITS O Figure G2.3.3a. Reference Dimension for Mark III Containment Submerged Structures Analysis 090779
22A4365 re20 Rev. 3 iso too -
I 3
50 - -
1 o I I l l l 0 02 *1D , 12 2n 2s TIME (sec)
Figure G2.3.4. Typical Vent Air Flow Rate for Main Steam Line Break (lbm/sec)
Steps 4, 5.
l Equations G2.3-2, G2.3-3 and the bubble rise equation G2.3-4 are solved. The results are shown in Table G2.3.3.
Step 6.
j Structure center location is X = 4.0 ft, y = 15.5 ft, Z = 0. ft. The given structure has length 2 f t and is considered as one segment.
Steps 7, 8.
i From Table G2.3.3, bubble arrives at structure at t = 0.05 sec.
I O
090779
22A4365 G-21 Rev. 3
, S teps 9, 10, 11.
1 To account for the effects of pool walls, floor and free surface the method
- of images was used. Equation (G2.3-6) was calculated for the function
[X2 + y2 + g2 which is the correction factor that accounts for boundaries .
and is applied to the velocity (Equation G2.3-10) and acceleration (Equation G2.3-9).
Table C2.3.4 presents the X, Y, Z and R = X +Y +Z . actors t, account for boundary effects on velocity and acceleraion, as well as the drag forces that these fields induce on the submerged structure. Note that the factor K (Equation A64. Reference G4.3) used to satisfy the local pressure boundary condition at the bubble surface is conservatively assumed to be equal to one. -
l I
Step 12.
- Using Table *G2.3.4 the direction of the ibduced flow field at each time' step is as follows
n n n (sec) x 2 z 4
0.0 0.655 0.756 0 0.02 0.63 0.776 0 0.050 0.588 0.811 0 Step 13.
Using Tables G2.3.3 and G2.3.4, il,(t) is calculated as:
T(sec) C,(ft/sec )
0.0 159.24 0.025 116.57 0.050 124.51 l
O 090779
.. ~ ,.
O 22A4365 G-22 Rev. 3 Table C2.3.3 LOCA BUBBLE INFORMATION l
(
Bubble Bubble Bubble Radial Bubble Radial Time Submergence Radius Growth Rate Acceleration (sec) (ft) (f t) (ft/sec) (ft./sec2) 0.0 7.500 1.146 15.000 917.416 0.025 7.480 1.451 15.370 245.153 0,050 7.422 1.880 17.897 -8.409 0.015 7.330 2.318 16.852 -64.960 0.100 7.207 2.718 15.116 -72.025 0.125 7.057 3.074 13.372 -67.092 0.150 6.882 3.388 11.801 -58.328 0.175 6.687 3.666 10.476 -47.434 0.200 6.473 3.914 9.446 -34.741 0.225 6.242 4.135 7.832 -119.975 0.250 5.998 4.284 4.012 -121.179 0.275 5.740 4.365 2.967 8.160 0.300 5.472 4.443 3.293 12.733 0.325 5.19 5 4.529 3.567 9.134 0.350 4.911 4.520 3.756 5.993 0.370 4.690 4.693 3.850 3.966 Table G2.3.4 FACTOR X, Y, Z, R AND DRAG LOAD X Y X R Time (sec) 0.0 0.06061 0.06995 0 0.09255 ,
0 .0 25 0.06112 0.07532 0 0.09700 O.050 0.06209 0.08594 0 0.1060 Acceleration Standard Acceleration Velocity Time (see) Drag (psi) Drag (psi) (ft/sec2) (ft/sec) 0.0 3.348 0.0165 159.2 1.823 0.025 2.451 0.0488 116.6 3.139 0.050 2.619 0.2229 124.5 6.706 g 090779
G-23 22A4365 Rev. 3 Step 13.
Using Tablea G2.3.3 and G2,3.4, U'=(c) is calculated as:
2 T(sac) 0=(f t/sec )
0.0 159.24 4 0.025 116.57 0.050 124.51 l Step 14.
The standard drag and acceleration drag are computed and shown in Table G2.3.4.
Also, note from Table G2.3.4 that the total drag force is primarily caused by the acceleration drag for this sample problem. In summary, the total drag I
force vs time and its direction is:
Time (sec) D (psi) 0 (degree) 0 0.0 3.1 49 0.025 2.5 51 0.050 2.8 54 where tan 0 = n /n (n , n are from Step 12).
y y G2.4 FALL 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 sup-pression 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 these structures will experience drag forces associated with
' water flowing at 35 f t/sec; this is the terrainal velocity for a 20 f t free fall and is a conservative, bounding number. Free fall height is limited by the HCU Floor. The Load computation procedura is the same as for calculating standard drag load in Step 14 of subsection G2.3 and,will not be repeated here.
Y e
?
D 090779
22A4365 G-24 Rev. 3 G2.5 LOCA CONDENSATIOE OSCILLATIONS LOADS Steam condensation begins after 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-mended for calculating the loads on submerged structures:
- 1. Note the dimension of the containment (L and H) as shown in Figure G2.3.3a.
- 2. Note the location of the submerged structure (x, y, z) .
- 3. Note the locations of the top vent exits (xgg, yd' *d)*
- 4. Determine 1/r effi f r each vent. The parameter r ef 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 g is small compared to the corresponding value for the vent nearest to the structure.
- 5. Calculate the acceleration field from N
h, = 2 (G2.5-1) i=1 eff g where 5 = 347 ft /sec is source strength determined from Mark III 1//T scale test data and N is the total number of vents i considered.
O 090779 l
22A4365 G- 25 Rev. 3 !
O
- 6. Calculate the acceleration drag force from P E= YA F " (Q2.5-2)
A R I C
l
- 7. The forcing function may be approximated by a sine wave with an amplitude equal to FA and a frequency rat.ge of 2 to 3.5 Hz.
4
- 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 analy=ed is that depicted in Figure G2.3.3.
() For simplicity, only three vents are considered here. 'The dimensions of the containment are:
L = 18.5 ft H = 20 ft Step 2.
The location of the submerged structure is X = 4.0 ft Y = 15.5 ft Z = 0ft O
090779 l
22A4365 G-26 Rev. 3 l l
l Step 3.
Ol The location of vents are:
Vent A B C 0 0 0 X,1 Y,1 12.5 12.5 12.5 Z,1 0 6.52 -6.52 Step 4.
for vents A, B and C are computed as before.
1/r- gg'i Vent A B C 31 13 13 2
Y eff' O
therefore
- 2
= 31 + 13 + 13 = 57
- e ff Step 5.
The acceleration field is (Equation G?. 5-1)
S 1 347 (f t) 1 57.8 ft/sec 2
U =
7b2r,gg L
sec 2
(18.5 f t) 2 x (57) =
i O
090779
i 22A4365 G-27 Rev. 3 1
I Step 6.
j The acceleration drag (Equation G2.5-2) per unit projected area is Fg = 1.35 psi i
Steps 7, 8.
The direction of this force is along the line joining vent A centerline I
! and the structure centerline. The forcing function is approximated by a l
sine wave with an amplitude equal to F A and a frequency range of 2 to 3.5 Hz.
i '
I G2.6 LOCA CHUGGING LOADS 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
- e on structures sdamerged in the pool. ,
1 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.
- 2. Determine location of structure (x, y, z) relative to bubble center (see Figure C2.6.1) .
- 3. Calculate distance r from chugging center to structure r = [x +y +z O
090779
~ G-28 22A4365 Rev. 3
= - s : -: --- -
- x - -_
1 ato
+ e-
~
* ===a =
4
- ' , 2 A r oro stN o V
- E .
- SIDE VIEW (Q[
f Vf Y V TOP VIEW Figure G2.6.1. Mark III Horizontal Vent Chugging O
090779
22A4365 G-29 Rev. 3 O 4.
Evaluate angle (0) between structure axis and r from i
- cos 0 =
cos a, cos a3 + cos 8, cos Sb + c s y, cos yb where (cos a,, cos 8,, cos y,) are the direction cosines of the structure axis, while (cos ab , cos Bb , e s Yb ) are the direction cosines of the vector r from the bubble center of the structure.
- 5. Calculate chugging load from
=
F C
" (0 o# o) sin 0 f 2(2.53) sin 0
- where A is the projected area of the structure normal to its own axis, AP r o9
= 2.53 psi-f t as the pulse strength.
- 6. Include the effect of another vent by repeating Steps 1 through 5.
( 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.
- 8. Ob tain time history as follows:
load duration is 2 maec period between individual chugs is 1 to 5 seconds For long structures, break the structure into separate ,ections 4
9.
and calculate the load on each section as above.
O 090779
22A4365 G-30 Rev. 3 G2.6.1 IDCA Chugging Loads - Sample Problem Step 1.
The submerged structure to be analyzed is depicted in Figure G2.3.3.
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.
l I
Distar.ce r is calculated as r = (4-0) + (15.5 - 14.5) = 4.12 f t for Vent A i
2 r = 4 +1 + 6.52 = 7.71 ft for Vent B or C Time for signal to arrive at submerged structure:
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 090779 1
22A4365 C-31 Rev. 3 4
O Step 4.
The angle G between the structure axis and the centerline of the bubble for the Vents A, B and C (see Figure C2.3.3) are:
I i Vent A B C 9 90* 32.3' 32.3*
! Step 5.
i i The chugging load from each vent is calculated from
)
Fe ,
2.53 Sin 0
! A r i
thus for each vent Vent A B C r U 1.23 0.35 0.35 A
Step 7.
i j Add the three forces linearly, to obtain the total force per unit area
= 1.9 psi L
i l i I
090779
22A4365 G-32 Rev. 3 Step 8.
The load duration is 2 msec, at 1.9 psi n
1.9 pel y .
r o.co2 ses :
The period between chugs is 1 to 5 sec. Hence the time history is e
1.9 pas -
--- 1.TO 6 sec e
090779
22A4365 G-33 Rev. 3 i
/"}
(_/
G3. SUBMERGED STRUCTURE LOADS DUE TO SRV ACTUATIONS i 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.
- 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.
- 2. Determine the initial location of the four bubbles. Each bubble will be assumed to form at the intersection of hole pattern canter-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 the bisector between adjacent arms and is tangent to the boundaries I
or structures.
O 090779 l
22A4365 G-34 R:v. 3
- 3. Obtain values of the following parameters from Table A4.4 and the specific plant documents:
P ,g: maximum bubble pressure, psia P ,1 ,- minimum bubble pressure, psia T ggy: initial pool temperature, *R H: quencher arm submergence, ft q
V: initial air volume in the safety relief valve discharge g
3 line (SRVDL), f t P:g initial air pressure in SRVDL, psia T:g initial air temperature in SRVDL, *R P: containment air space pressure, psia k: specific heat ratio of air p: water density at T ggy, lb,ft
- 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 = Pool i
, ft 3
(G3. 2-1) i min 0
090779
- - - - - - - - - - - - - - - - - - - - - - - - - _ . _ - - - - - - - - - - - - - - - - - - - - - - - - - - - . - - - ----------- u
22A4365 G-35 j Rev. 3 O and A
4 R,, = V , ft (G3.2-2)
- 5. To account for the vertical motion of the bubbles, the bubble rise equation given below must be solved simultaneously with the bubble dynamics equations for R(t), R(t), I(t) and Z3 (t), where i
R(t) = bubble radius at time t R(t) = bubble growth rate at time t 4
E(t) = rate of change of the bubble growth rate at time t j Zb (t) = submergence of bubble center at time t i
Bubble Dynamics Equations E(t) =
f (PB - =
}- (C3' ~ )
P " -
B t/R (G3.2-4)
B P, = (G3.2-5)
] PC+ b Bubble Rise Equation t
wp C -
"# E D b + "B E z
3
=
3 (G3.2-6) m up R B+23
'O i
090779 ,
1 I
22A4365 G-36 Rev. 3 Y
1 i i = u e a r mass (G3.2-D m
B 4R air Ti R
ir
= gas constant of air Initial Conditions:
=
R(0) R 5L(0) = 0 Zb (0) = H Z3 (0) = 0 P(}
B min
- 6. Determine the location of images of the four source bubbles to account for the effects of pool walls, floor and free surface.
Then calculate the parameters X, Y, and Z, which are defined by Equation (A21) of Reference G4.3 (in actual computation, it is more convenient to use Equation (2.2-6)) .
- 7. For multiple quenchers use Equation (A21) of Reference G4.3 to evaluate the parameters X, Y and Z. Note the Heaviside step functions H(t-t )g and H(s -t) g are introduced to account for phasing relations among the quenchers of interest.
- 8. Using the results from Stepa 5 through 7 calculate the equivalent uniform acceleration, U,(t), at time t at the structure location f rom U,( t) =
R (t) N(t) + 2R(t) $l (t) X +Y + (G3.2-8) o 090779
i 22A4365 G-37
- Rev. 3 O
The correspondir.g velocity, U, (t), may be obtained by numerically integrating O(t) . As a first approximation, U,(t) can also be l evaluated from U,( t) =
R (t) d(t) X +Y +Z (G3.2-9)
- 9. The acceleration drag is calculated from b" (t) VA P FACE) " (G3.2-10) g where 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 b
V pU t Fs (t) = CD A, [g(c) (G3.2-ll) where C is the drag coefficient for flow normal to the structure.
D A is the projected structure area normal to U (t). Add F and n A Fg at any time to t to get the total force on the structure or structure segment. The direction of the total force is normal to the submerged structure, i
G3.2.1 Quencher Bubble Load - Sample Problem Steps 1, 2, 3.
The following geometrical and bubble data were used in the sample calcula-tion of the loads from one quencher to the structure sho rn in Figure G3.1.
(a) Maximum Bubble Pressure: Pmax = 39.3 psia (b) Minimum Bubble Pressure: P min. = 10.1 psia 090779
22A4365 G-38 Rev. 3
= 560*R e
(c) Initial Pool Temperature: T pool (d) Quencher Arm Submergence: Hg
= 13.9 ft (e) Initial Air Volume in the Relief Valve Discharge Line (SAVDL)
Vi = 56.13 ft 3 (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.
h The air leaving the quencher forms four independent and identical spherical bubbles which oscillate in phase while rising. The Bubble Dynamics equa-tions (G3.2-3), (G3.2-4) and (G3.2-5) and Bubble Rise Equation (G3.2-6) were solved and shown below.
Bubble Bubble Radial Bubble Radia)
Bubble Submergence Radius Growth Rate Acceleration Time (ft/sec 2)
(sec) (f t) (f t) (ft/sec) 1.695 0 -467.021
- 0. 13.900 1.691 -0.249 692.474 0.050 13.822 1.695 0.359 -441.970 0.100 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 e
090779
22A4365 G-39 Rev. 3 0 E-NDE vlew c--- _________
5.2. 6A - 4.11 y V
Z
- I s u - .
l .
DUSSLE #3 augSLE #1 l l
- j rX j L _. _ __ __ _l 19.2. SA,0.0)
(23,83,041 \ SUPPRESSION j POOL TOP VIEW 800sLE #4 m
j W s.2.as -4.n t 04-4.44 -dos PLAN VIEW W ATER SUMP ACE 8 a Y
\
\ \ n
\ \ ,
N
\
\ H qft
= sn hi lr \
L_
m 4
O c--
Q L __q g, _ _, _g f
=
p _
J 5.6 ft ir V BASEMAT Figure G3.1. Four-Bubble Model for Quencher Air Discharge 090779 l
L_-____---_________ __ _ ___
22A4365 G-40 Rev. 3 Steps 6, 7. l To account for the effects of pool walls, floor and fill surface, the method of images given in (A21) of Reference G4.3 was used. For simplicity, the correction factor K of (A75) of Reference G4.3 is assumed to be one. The resulting parameters of X, Y and Z are shown as below.
t X Y Z
- 0. 0.0364 0.0170 -0.0311 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 -0.0289 0'.25 0.0320 0.0072 -0.0279 Step 8.
Using Equations G3.2-8 and G3.2-9. U,and 0,are calculated as follows:
t U. U.
O. O. -68.162 0.05 -0.0339 94.264 0.10 0.0483 -59.420 0.15 -0.1064 52.240 0.20 0.1077 -35.642 0.25 -0.1172 -21.182 O
090779
22A4365 G-41 Rev. 3 O Step 9.
The normal acceleration and velocity components to the subaierged structure are calculated and their associate acceleration and standard drag are also coornted and shown as below.
Standard Acceleration Drag Drgg 2 t U U' (x 10 ~ psi) (x 10 psi) 1 - _2L 5L
- 0. O. -53.780 0. -22.74 0.05 -0.0282 78.522 -0.642 ,
33.20 0.10 0.0396 -48.724 1.266 -20.60 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 l
The direction of the force is normal to the submerged structure and is given as Xn y+Yh 2
X +Y +Z (Refer to Steps 6, 7.)
090779
~
22At365 G-42 Rev. 3 l 9
G4. REFERENCES
- 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, NECE-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 Condensation 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 LOCA and Safety Relief Valve Ramshead Air Discharges, NEDE-21471, supplement 1; June, 1978.
O 090779
22A4365 H-1 Rev. 3 ATTACHMENT H
SUBJECT:
WEIR WALL LOADS DURING DRYWELL 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
+ (6.35a +1.5a-0.75)(1}.0
'21.0 = -3.25 + 2(32.2 44) _
0.5(4]2
+
1}.0 + 5.71 21.0 = -5.20 + 2(32 ) 144) 0.5 2
+ 6.35(1, I +1.5(1,I - 0.75
( + (1.55a - a ) .0 1.0
+ 5 . 71 + 1. 95 2 2 62.4 (1-a-O m (1-a-O m 0.5 + 7.10 12.0 21.0 = -7.15 + 2(32.2)(144) 4.12
+ 1.55 -
- ) + (1.55a - a )
, 0 (1}.0 l
2" l
+ + 5.71 + 1.95 + 1.95 l (1}.0 . i 090779
22A4365 Rev. 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
=
V 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 momentem loss;
=
oV A -
F ge Results:
E' V = 37.91 = 40 see top V = 32.09 = 35 see mid f
V = 26.65 =30 ,,"
bot F = 12800 lbf top F "
mid F
bot f
e 042178 i
9
1 22A4365 H-3 Rev. 2 O Tab 1e a-1 4 1 Pressure due to top vent submargence 2 Pressure due to mid vent submergence 3 Pressure due to bottom vant submergence 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 botton and mid vents 14 Friction loss 15 Loss in Iee junction
>O 16 Pressure loss due to elevation difference between nid and top vents 17 Friction loss s 18 Loss in Tee junction ,
i 19 Pressure loss due to elevation difference between top vent and top of vier wall
, 20 Friction loss 3
a 21 Loss at Exit i
i O
O 042178
22A4365 R,v. 2 g-4 Table H-2 AP A Q K 1 -3.25 - aM -
2 -5.20 - bM -
3 -7.15 -
(1-a-b)M -
4 -
4.12 an 0.5 5 - - aM 0.0 6 -
12.00 M 6.35a + 1.Sa - 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 I -hD_, )
1,55 (1,)
16 1.95 -
(1-a)M -
17 - -
(1-a)M 0.0 2
18 -
12.00 M 1.55a - a 19 5.70556 -
M s 20 - -
M 0.0 21 -
12.00 M 1.0 AP - Pressure Drop due to static head A - Flow Are2 Q - Volu=e Flow Rate K - Loss Coefficient (from Idel Chik) 2 / 2 AP = = ! !K 2qc 2qc
(
a - 2 of volume flow rate through tpp vent b - % of volume flow rate through middle vent M - Total volume flow race through the entire vent system 042178
i l
22A4365 R;v. 2 H-5/H-6 O
SURFACE AREA WlER ANN.
NUMSER OF VENTS
= 400 ft2
= 120
,h \% ;
SURFACE AREA PER 3 VENTS = 12 f t2 # - ~
^
~
DIAMETER OF VENTS = 27.5" AREA OF VENT CROSS SECTION = 4.12 ft2
@ N e e 3
WlER WALL HEIGHT = 26*1" k8) bh HWL HEIGHT = 20'5" '
(TOP VENT HEIGHT = 12*11*
(MlO VENT HEIGHT = 8*5" (SOTTOM VENT HEIGHT = 3*11**
P
\\\
C ~ DW
- 2I "
hh h h h l
- 0W d
4e
. n
( ke 'c
$e @ @
,W e fe Mr N/
4e 4e 4e te e e e AV M^ W e
4e 4e 4e e e e
.W W W Figure H-1. Mark III Vent System Network =
042178 s
22A4365 Rev. 3 I-1/I-2 (v~)
ATTACHMENT I POOL SWELL VELOCITY Early in the Mark III program it became necessary to establish 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.
The following is a summary of the calculations:
Using the 238 reference design, the total venting area between the drywell and containment is 481 ft2, 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 pressure) then the pressure ratio across the vent system is 0.52.
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 surface area of 5900 ft gives a pool surface velocity of 34 ft/sec. For design purposes, this was rounded off to 40 ft/see to cover such uncertainties as bubble temperature and pressure.
j 090779
.. . . - . . .. . - .....=_-_ - . .- . -. ... - - . . . . - . - . . - - -
! l I
1
! 22A4365 :
e Rev. 3 J-1 through J-52
- O j' ATTACHMENT J
! SCALING ANALYSES AND SMALL STRUCTURE i POOL SWELL DYNAMIC IDADS i
1 t i '
I P
i i
j i ,
i y .
l l
l i
!O
- ATTACHMENT J is PROPRIETARY and is provided under separate cover.
b t
1 I
f I
i i
i I
i 090779 l s
/'
4 22A4365 K-1
_ Rev. 2 _.
< ATTACHMENT K I RESPONSES TO NRC QUESTIONS The following are responses to questions received in March 18, 1975 letter from J. F. Stolz:
QUESTION Kl.1 i
Provide large size plan and section drawings of the containment which illus-l trate the structures, equipment, and piping in and above the st:ppression pool.
- All equipment and structural surfaces which could be subjected to suppression pool hydrodynamic loading should be specifically identified and described on these drawings.
l
, ANSWER TO QUESTION Kl.1 This input is the responsibility of the containment designer and is provided
() in Part II, Section 14.
QUESTION Kl.2 Provide a graphical chronology of all potential pool dynamic loads which identifies the source of the load (i.e., pool swell froth impingement), the time interval over which the load is active, and the structures which are affected. (Reference GESSAR Response 3.82.)
ANSWER TO QUESTION Kl.2 This information is presanted in Sections 2, 4, 5, 6, 7, 8, 9, 10, 11, and 12 f this report. The discussion consists of (1) a description of the phenomena associated with the EOCA and S/R valve dischargc, (2) a time history depiction of the phenomena as related to weirwall, drywell, pecl bottom, containment shell, and structures at or above the pool surface. This information covers the small break accident intermediate break accident, and the DBA LOCA as well as potential S/R valve combinations. For each phenomena a load profile is
/)
i specified for the various structures and an identification of the basis for the loads as well as justification for the value is indicated.
042178 ,
A
22A4365 R v. 2 K-2 QUESTION Kl.3 <
O For each structure er group of structures, provide the anticipated load as a j function of time due to each of the pool dynamic loads which could be imparted to the structure.
ANSWER TO QUESTION Kl.3 This information is included in this report as discussed under item Kl.2 above.
QUESTION Kl.4 For each structure or group of structures, provide the total load as a function of time due to the sum of anticipated pool dynamic loads.
ANSWER TO QUESTION Kl.4 This information is included in this report as discussed under item Kl.2 above.
The loading charts in each section identify all the individual loads as a func- g tion of time such that for any time period the applicable loads can be deter-mined by the designer.
QUESTION Kl.5 Describe the manner in which the pool dynamic load characteristic shown in 1.4 above is integrated into the structural design of each structure. Specify the relative magnitude of the pool dynamic load compared to other design loads for the structure.
ANSWER TC QUESTION K1.5 This input is the responsibility of the containment designer and is provided in Part II, Sections 15 through 19.
042178 O
\
22A4365 Rev. 2 g_3 QUESTION Kl.6 Describe tha manner by which potential asymmetric loads were considered in the containment design. Characterize the type and magnitude of possible assym-metric loads and the capabilities of the affected structures to withstand such a loading profile. Include consideration of seismically induced pool motion which could lead to 1ccally deeper sub:r.ergences for certain horizontal vent I stacks.
ANSWER TO QUESTION Kl.6 The potentially significant asymmetric load situations are discussed in Attach-ment L. The containment designer must identify what his design is capable of withstanding relative to the situations identified in Attachment L or other arbitrary situations he may elect to postulate for the demonstration of overall capability (refer to Part II, Sections 15 through 19).
i QUESTION Kl.7 Provide justification for each of the load histories given in Kl.3 above by the use of appropriate experimental data and/or analysis. Reference to test data should indicate the specific test runs and data points and the manner by which they were converted to loads.
ANSWER TO QUESTION Kl.7 This information is included in the discussion of loadn identified in various
! sections of this report as indicated in item K1.2 above.
QUESTION Kl.8 i
For those structures subject to pool dynamic loads provide your anticipated schedule for completion of the structural design, procurement of materials and actual construction.
O 042178
- l l
22A4365 Rev. 2 g.4 l
ANSWER TO QUESTION Kl.8 9l The containment designer or the utility must provide this information and is provided in Part II, Section 19.
QUESTION Kl.9 Discuss your specific plans to be responsive to the concerns of ACRS. As noted in their letters on the Perry plant, the committee believes that a more basic understanding of certain phenomena such as oscillations, vent interaction, pool swell, and dynamic and asymmetric loads on suppression pool and other con-tainment structures is required. "The Committee emphasizes the importance of directing the test and analytical programs toward providing not only empirical lesign correlations but also toward more detailed evaluations of the relevant two-phase phenomena in order to enable the better application of a specific set -
of scaled tests to a range of actual reactor conditions." If reference is to be made to GE analytical methods development, a finalized breakdown of areas of investigation performed by GE and a time schedule as to their availability is requested. We require that those areas for which analytical results are avail- g able, but not yet submitted to the Staff, be documented for our review as soon as possible.
ANSWER TO QUESTION Kl.9 In a March 26, 1975 letter to R. L. Tedesco from G. L. Gyorey, a brief discussion of the GE Mark III investigation program was given together with the presenta-tion of a matrix which represented a concise picture of what has been accomp-lished to date and is scheduled during the remainder of the Mark III Verification Program. Additional discussion is included throughout this report on phenomena description and the significance of the associated loads relative to the design.
I 0
042178
22A4365 K-5 Rev. 2 _
~
The following are responses to questions received in April 22, 1975 letter from J. F. Stolz:
QUESTION K2.la Specify the nun.ber of safety relief valves, their design flow rate, and discharge line size. Provide a listing of the operating conditions under which these valves would be operated either manually or automatically.
ANSWER TO QUESTION K2.la There are 19 safety relief valves. Each valve can operate automatically or by l power actuation. Automatic operation refers to the condition when line pres-sure exceeds the set point pressure and the line pressure itself opens the valve. Power actuation uses an aux 5.liary feature to open the valve when the line pressure is less 'than the automatic set point pressure.
2 Line size is 10 inch schedule 40 and the design flow rate is 925,000 lb/hr.
Identically designed valves, line size and design flow rate is used for all operating modes.
Operating Modes:
Operating Mode Actuation Method Description .
Overpressure relief Automatic Valves function to limic line operation pressure rise and to avoid abnormal operation and pressure loading.
, Overpressure safety Automatic or Valves function as safety valves operation power actuation and open to prevent nuclear sys-tem overpressurization and cvoid failure of the reactor coolant 1
- pressure boundary.
O 042178 i
22A4365 K-6 Rev. 2 .
l ANSWER TO QUESTION K2.la (Cont) 9' Ocerating Mode Actuation Method Description Depressurization Automatic Selected valves (8) are part of operation the automatic depressurization system (ADS). These valves open automatically to depressurize the reactor during events involving small breaks in the nuclear system process barrier. This function is required for the emergency core cooling system (ECCS).
QUESTION K2.lb Describe, with the aid of drawings, th'e routing of the discharge line to, and orientation in, the suppression pool and the design of the discharge line exit.
ANSWER TO QUESTION K2.lb This part of the first item is the responsibility of the containment designer (Part II, Section 14). See also Attachment A to Part I.
QUESTION K2.2 Provide the load specification for the suppression chamber structure to accommodate actuation of one or more safety relief valves.
ANSWER TO QUESTION K2.2 -
Information is included in Section AS.0, Attachment A.
QUESTION K2.3 Provide the design load capability for the suppression chamber structure.
042178
1 22A4365 Rev. 2 < E-7 ft d ANSWER TO QUESTION K2.3 i This is the responsibility of the containment designer and is provided in Part II, Sections 15 through 19.
, QUESTION K2.4 Provide justification for the load specification given in K2.2 above by the
! use of appropriate experimental data and analysis. If the General Electric Company is responsible for specifying these loads, a statement to that effect is sufficient.
ANSWER TO QUESTION K2.4 d
This information is included in Section A12, Attachment A.
QUESTION K2.5 Identify, with the aid of drawings, any components or structures in the suppres-sion peol region, other than the bounding walls of the suppression chamber, and the location of such components relative to the relief valve discharge line exits. Discuss the structural capability of these components to accommodate loads due to relief valve actuation.
ANSWER TO QUESTION K2.5 This is the responsibility of the containment designer and is provided in
, Part II, Section 14.
i QUESTION K2.6 Estimate the maximum number of single and multiple relief valve openings over the life of the plant.
042178
22A4365 Rev. 2 K-8 l ANSWER TO QUESTION K2.6 ,
The maximu:n number of single and multiple re. lief valve openings over the life of the plant.
See section 9.0 of Attachment A for answer.
QUESTION K2.7 Identify the maximum temperature limits of the st.ppression pool with the reactor at power. This temperature limit should include provisions for the testing requirements of relief valves.
ANSWER TO QUESTION K2.7 See Appendix K-A to response for Question K2.9.
QUESTION K2.8 O
Specify the operator actions that are planned when the specified temperature limits are exceeded.
ANSWER TO QUESTION K2.8 See Appendix K-A to response for Question K2.9.
QUESTION K2.9 Present the temperature transient of the suppression pool starting from the specified temperature limits for the following transients:
- a. Main steam line isolation
- b. Semi-automatic blowdown
- c. Stuck open relief valve For purposes of the above analysis, the minimum water level should be assumed g in the suppression pool.
042178
22A4365 K-9
_ Rev. 2 _
ANSWER TO QUESTION K2.9 i I. INTRODUCTION 4
3WR plants take advantage of tha large thermal capacitance of the suppres-sion pool durinr, plant transients requiring relief valve actuation. The discharge of each relief valve is piped to the suppression pool, where the steam is condensed, resulting in a temperature increase of the pool water, but a negligible increase in the bulk containment pressure. Most transi-f eats that result in relief valve actuations are of very short duration and have a small effect on the suppression pool temperature. However, there are some events which present the potential for substantial energy releases to the suppression pool that could result La undesirable high pool temperatures if timely corrective action is not taken.
1 Elevated suppression pool temperatures during extended relief valve opera-tion at high reactor pressures have become a major concern recently in
() light of occurrences at two European BWR plants. At local suppression pool temperatures in excess of approximately 160*F and at moderate to high relief valve flow races, severe, continuous structural vibrations were encountered.
4 l
The possibility of encountering the abov'e condition is unlikely due to rigid
^
technical specifications on the pool temperature during power operation and the large capacity for heat absorption. This is supported by the fact that such an occurrence has never happened at a domestic BWR site. However, since the possibility does exist when assuming limiting situations of peak service water temperature, technical specification pool temperatures, etc.,
it is important that potential situations leading to this phenomenon be 4
recognized and procedural controls, temperature limits and instrumentation be utilized to avoid it.
II. PLANT TRANSIENT EVENTS t
i Since the discharge of the safety relief valves are piped directly to the
() suppression pool, any S/R valve actuation will result in some temperature rise in the pool. Most S/R valve actuations are of short duration (seconds) 042178
22A4365 R;v. 2 K-10 and result in negligible temperature increase. Three events however present the potential for substantial energy addition to the suppression pool via h the relief valves.
These events are (1) stuck open S/R valve, (2) system isolation, and (3) automatic depressurization. Appendix K-A proposes temperature limits and procedures for each of these events. A brief description of each of these events is given in the following paragraphs.
- 1. Stuck Open S/R Valve -
In the event of a stuck open S/R valve, the suppression pool temperature will increase at a rate dependent on the S/R valve flow rate, pool size, and heat removal system capability. The only means of terminating the energy input to the pool is to scram the reactor and depressurize the RPV (assuming that the S/R valve cannot be closed). To avoid the vibration zone, the RPV must be depressurized such that steam velocity at the discharge end of the S/R valve piping in the suppression pool is subsonic before the pool temperature in the vicinity of the discharge g reaches' 160*F. The action recommended for a stuck open S/R valve is given in Appendix K-A.
- 2. Primary System Isolation Whenever the primary system is isolated from the main condenser for whatever the reason, the reactor is scrammed automatically and the fuel relaxation and decay heat energy is initially removed automatically from the RPV via vessel pressure actuation of the S/R valves. The water level in the RPV is controlled by the feedwater system, RCIC, or RPCS. Mark III plants employ condensing type heat exchangers that are utilized to remove energy directly from the RPV or from the suppression pool after primary system isolation. The proposed actions to be taken during a postulated isolation event are given in Appendix K-A.
O 042178
22A4365 Rev. 2 K-ll
- 3. Auto Depressurization Systems (ADS)
(])
i Activation of ADS results in rapid depressurization of the primary system due to simultaneous opening of pre-selected relief valves.
This system is automatic but can also be activated manually. Typi-cally the RPV is depressurized to 150 psi or less in approximately 10 minutes. During this transient, the bulk suppression pool tempera-ture will rise on the order of 40*F. Uniformly spaced ADS relief valve discharges in the suppression pool will result in near uniform mixing of the pool during this depressurization transient. While no specific procedures are required, the pool temperature limits proposed in Appendix K-A are demonstrated to be adequate such that the vibration zone is not entered during this event.
l III. RESULTS OF TRANSIENT t.NALYSES ,
Seven specific transients were analyzed to demonstrate conformance to the specified pool temperature ILnits of Appendix K-A. The specific cases analyzed, their initial conditions, and the results are summarized in Table K-1. Figures K-1 through K-7 show the transient response of the sup-pression pool and reactor pressure for the seven cases. These results show the pool temperature is below the allowable limit (160*F), when the RPV is at or above 200 psia. This pressure represents the minimum reactor pressure at which sonic discharge of steam into the suppression pool can occur.
QUESTION K2.10 The temperature instrumentation that will be installed in the pool and the l
sampling or averaging technique that will be applied to arrive at a definitive pool temperature.
i +
ANSWER TO QUESTION K2.10 The following is a discussion of the Suppression Pool Temperature Monitoring System that is recommended for the GE Standard Plants.
i O
6 042178
22A4365 R;v. 2 K-12 General Description 9
Suppression pool temperature instrumentation is included to provide an alarm due to high suppression pool temperatura in order that the plant operator will have adequate information regarding operating status of the suppression pool.
System Description
Sensors: Commercial grade thermocouples (T/C) or resistance temperature device (RTD) compatible with existing plant equipment.
Quantity: Two sensors shall be provided at each monitoring location.
Location: Sensors shall be installed no more than one foot below the normal water level, but no less than 3 inches below the minimum tech. spec. water level.
Sensors shall be located within 30 ft (line of sight) of each relief valve ais-charge location. Sensor groups may be shared (i.e., one sensor grcup may pro-vide coverage for more than one relief valve discharge location). The relief valve discharge location shall be defined by the intersection of the centerlines of the downcomer pipe with the horizontal discharge pipes.
NOTE: The 30-ft specification is based on the results of relief valve tests performed at the Q w. Cities site. Extended blowdown tests showed uniform water temperature in a region within 2 ft of the pool surface and 43 ft on either side of the discharge point. Therefore, 30 ft is con-sidered conservative for the maximum distance between source and sensor, and thus the indicated temperature will be representative of the tempera-ture at the point of discharge.
Recording: Pool temperature shall be monitored on recorders in the control rocm.
Two sensors from each sensor group will be recorded.
Time Constant: The time constant of the final T/C or RTD installation shall be no greater than 15 sec. The time from signal output of sensor to initiation of function shall be no greater than 0.5 sec. The difference between measure =ent reading and actual lo.al temperature shall be within 2*F.
O 042178
22A4365 g,13
. _ _ _ . ,Re v 2 _ ,_,
Set Points: 1 Instrument set points for alarm shall be established so that the reactor plant can be shut down and depressuri:ed to less than 200 psia.before the suppression pool temperature reaches the threshold temperature for steam
~
quenching vibration for the S/R valve discharge device employed.
'The following are responses to questions received.in November 10, 1975 letter-from V.-A. Moore:
QUESTION K3.1 Page 5-2 of Appendix 3B to GESSAR states that 1/3 scale tests have shown chugging loads with amplitudes as high as 60 psi. Discuss the significance of these loads with respect to the weir wall load specification of 15 psi.
ANSWER TO QUESTION K3.1 This information has been incorporated in Section 5.1.4, page 5-2.
O. QUESTION K3.2 Page 6-2 of Appendix 3B to GESSAR has deleted a load specification which was present in NEDO-11314-08 (Preliminary). This load was due to a postulated asym-metric air bubble distribution in the suppression pool resulting in a 10 psi' bubble load on one-half the containment periphery and 0 psi load on the other half. We require that this lord specification be included as a design basis for Mark III containments. Pages 6-2, L-2 and L-4 should be revised accordingly.
ANSWER TO QUESTION K3.2 There may be small circumferential pressure variations at vent clearing due to asymmetric considerations; however, the General Electric Company does not believe that there is a mechanism for having the containment exposed to a LOCA bubble pres-sure of 10 psi on one-half of the suppression pool while the other half has O psi load. However, pages 6-2, L-2 and L-4 have been revised as required by the Staff to include this arbitrary asymmetric load case.
042178
22A4365
_ _Riv. 2 _ K-14 QUESTION K3.3 O
Pages 2-8 and 2-9 of Appendix 3B to GESSAR propose that loads due to the inad-vertent actuation of a safety / relief valve (single-active failure) concurrent with a loss-of-coolant accident (LOCA) be considered only for evaluating addi-tional containment capability. We require that this load combination be included as a design basis for the containment.
ANSWER TO QUESTION K3.3 The General Electric Company does not believe that the peak dynamic loads from a safety / relief (S/R) valve discharge can occur at the same instant the peak dynamic loads of the LOCA event are present in the suppression pool. This includes the consideration of the effects of imposing the " single active compo-nent failure criterion." Dynamic loads from S/R valve di,scharge are considered together with appropriate LCCA related dynamic loads as a design basis; however, consideration of the time history is given as shown on the loading condition bar charts. This is consistent with the treatment of LOCA + S/R events when evaluating reactor plant perfor=ance.
QUESTION K3.4 Figure J-7 in Appendix 3B to GESSAR specifies design pool dynamic loads for small structures located up to 20 ft above the sappression pool surface. Provide similar specifications, with appropraite justification, for both small and expansive structures above the 20-ft elevation up to the elevation at which pool dynamic loads become negligible.
l ANSWER TO QUESTION K3.4 The dynamic loads for expansive structures at the ECU floor clavation, including the bottom of the steam tunnel, are presented in Section 11. The dynamic loads for small structures at and above the HCU floor elevation are presented in Section 12.
l 042178
22A4365
_ Rev. 2 ._ K-15 QUESTION K3.5 With respect to Section J of Appendix 3B to GESSAR, provide the following:
- a. The referenced dimensional analysis based on the Buckingham x Theorem.
Discuss how parameters such as liquid density viscosity, buoyancy and gravity were considered.
- b. Justify scaling drywell pressure at vent clearing rather than the integrated drywell pressure over the pool swell transient.
- c. Discuss the basis for the " predicted" froth velocity of 50 ft/sec for Mark III.
- d. Discuss the basis for assuming that impulse is directly proportional to froth velocity.
- e. Discuss the basis for the assumed triangular load profile ror the steam tunnel load (50 psi over 45 msec) due to froth impingement.
ANSWER TO QUESTION K3.5
- a. Section J 6 has been added to present the Buckingham e analysis,
- b. A discussion of the importance of the drywell pressure profile has been incorporated into Section J.4.
- c. The basis for the predicted froth velocity has been incorporated into Section J.4.
- d. The basis for assuming that froth impulse is directly proportional to velocity is a correlation of impulse versus velocity generated from the PSTF roof input data of test Series 5801. Refer to Figures 4-65 and l 4-66 of NEDM-13407P (Reference 11). These plots clearly demonstrate that impulse does increase linearly with velocity.
042178
22A4365 Rev. 2 K-16
- e. The Steam Tunnel has been relocated to the HCU floor elevation and the l dynamic loading conditions are presented in Section 11.
1 l
QUESTION K3.6 Provide a more detailed scaling analysis than that in Section J which addresses the following:
- a. Specify the portions of the pool dynamics transient to which the scaling analysis is applicable (i.e., is it valid for bulk pool swell only .
or can it be applied to breakthrough, froth characteristics, loads imparted to structures),
- b. Depending on your response to (a) above, justify that the selected characteristic equations are adequate to account for effects such as bubble penetration into the pool, migration to the surface, growth, thermodynamics, and dynamics of breakthrough.
- c. Justify the selection of characteristic length (s) as being appropriate for the principal phenomena under consideration.
- d. Discuss the applicability of the scaling analysis to a test facility which is not uniformly geometrically scaled.
- e. For those phenomena to which the scaling study is to apply (as speci-fled in (a) above), provide additional comparisons between test data of different scale to verify the scaling relationships.
ANS'n'ER TO QUESTION K3.6 Additional scaling analysis discussion is presented in new Section J.7.
O 042178
22A4365 K-17
_.?*v .2 QUESTION K3.7 Section A12.3 of Appendix 3B has identified quencher exhaust area, amount of gas in the discharge pipe, free water surface area, and pool temperature as important parameters for the quencher design. Discuss and justify that vent submergence and air discharge rate are not important parameters for the quencher.
ANSWER TO QUESTION K3.7 Results of tests showing the effect of submergence of the quencher have been presented in Section 2.3.12 of Reference 14. As the test results show, the effect of quencher submergence is negligible when submergence is in the range of 4 to 6 meters (12 to 18 f t) . For smaller submergences the containment loads decrease; however, GE quencher loads are all for submergences of > 4 meters.
i Air discharge rate per unit quencher area is a significant parameter affecting air clearing loads. This parameter is proportional to:
/~h
%/ A outlet A
quencher This ratio is approximately the same for the small and large-scale tests discussed in Reference 14; however, it is smaller for the GE quencher design. Therefore, had this parameter been considered in the loads from the GE quencher, slightly lower load predictions would have resulted.
QUESTION K3.8 Section A12.3 and Figure A12.1 identify three important scaling parameters:
- 1. air volume in pipe / quencher cross section;
- 2. free water surface area / quencher cross section; and
- 3. total air volume at time of blowout / total quencher opening area. l C) l 042178
22A4365 K-18 R;v. 2 However, the statistical approach, which was used to determine the scaling factors for the small-scale test results and then used to predict the loads for the Mark III, has been developed using three different scaling parameters (i.e., air volume, water-surface area and water temperature). Discuss and justify using the three scaling factors in the statistical approach to predict the pressure for the Mark III quencher response instead of using those three key scaling parameters identified in Section A12.3.
ANSWER TO QUESTION K3.8 The titles of Sections A12.7.2.1 and A12.7.2.2 have been changed to be more representative of the information covered. The titles now read:
A12.7.2.1 Scaling for Air Volume / Quencher Area A12.7.2.2 Scaling .for Water Surface Area / Quencher Area Thus, the scaling factors used in the statistical analysis are indeed:
- 1. air volume / quencher area; g
- 2. water surface area / quencher area; and
- 3. water temperature.
In addition, the title of ites three in Figure A12.1 has been corrected to read
" Quencher Cross Section + Total Quencher Opening Area." The reason this param-eter was not considered is explained in the response to Question K3.7.
, QUESTION K3.9 Clarify the following important terms used in the report:
- a. Air volume in pipe and total air volume at time of blowout. Distin-guish the physical meaning between these two terms.
- b. Free water surface area. Does the free water surface area include the ,
entire water surface area or the effective water surface area? The difference between these two surface areas could be significant. g1 ;
l 042178
22A4365 K-19
. Rev. 2
() c. Water temperature. Does it mean initial pool water temperature, or local maximum water temperature during transient, or average bulk water temperature during transient?
ANSWER TO QUESTION K3.9 I a. " Air volume in pipe" refers to air volume prior to valve actuation.
" Air volume at time of blowout" refers to the volume of compressed air in the discharge pipe at the time when the air begins to enter the pool.
Air volume at time of blowout is always smaller than the initial air volume.
(
- b. " Free water surface area" refers to that portion of the pool surface which is exposed to the containment atmosphere.
- c. " Water temperature" refers to initial pool temperature. The effect of local pool temperature is implicit in the overall increase in loads due to subsequent actuation.
(O-]
QUESTION K3.10 Figure A12.12 shows that the Mark III maximum load prediction is calculated by 7
adding the pressure difference between first and subsequent R/V actuation from the large-scale test data to the scaled Mark III predicted first actuation load.
Justify that the scaling factors are not needed for the pressure difference between first and subsequent actuation from the large-scale test data.
ANSWER TO QUESTION K3.10 As indicated in Figure A12.2, the pressure, difference between first and subsequent actuation for Mark III is obtained by scaling the pressure difference obtained 1 from large-scale test data.
i O
042178
22A4365 K-20 R;v. 2 QUESTION K3.ll O
Figure A12.14 shows the large-scale test data and the analytical fit for the first and subsequent actuation. Provide the following:
- a. a description of the analytical fit (The description should include the analytical model and all assumptions used in the model.);
- b. justification that the differential pressure between the subsequent actuation and the first actuation for the analytical fit remains constant for varied reactor pressure;
- c. a figure showing the maximum and minimum pressure.s resulting from the first and subsequent actuation for each test result of the large-scale test; and discuss how the maximum and minimum pressure will be deter-mined from the large-scale test result; and,
- d. a best fit curve for the data in the figure requested on c. above.
O ANSWER TO QUESTION K3.11
- a. The maximum positive and negative floor pressures were estimated with means and variances using a stepwise multiple linear regression pro-cedure, operated in backward stabilization mode. It was assumed that the variances about the predicted means were constant and that the action of each parameter on the mean '.rss independent of all the other parameters.
- b. Subsequent valve actuations are assumed to occur at the set point of the valve; thus, reactor pressure is not a variable,
- c. Figure A12.14 provides the requested information. The maxima and minima were determined by inspection of the pressure traces.
- d. Figure A12.14 provides best fit curves for first and subsequent actions.
O 042178
22A4365 K-21 Rev. 2 __
QUESTION K3.12 .
Regarding the statistical analysis approach, provide the following:
2
- a. justification fer using mean value of the test data calculated by the statistical analysis as the basis of scaling factor instead of bounding value;
- b. an analysis of the . significance of regression (F-test);
- c. an operating characteristic curve or power curve for each scaling parameter; 4
- d. there are two curves for the_ test data of each scaling parameter.
One is for the upper bound and the other for the lower bound. Verify which curve was used in the regression analysis; and 4
- e. in Section A12.7.4, a factor of 1.3/1.5 is used for the posicive as well as the negative pressure predir. tion'for the large-scale test.
Describe the basis for the selection of this scaling factor.
ANSWER TO QUESTION K3.12
- a. Both means and variances were used in t!,e scaling procedure. This results in improved stability of the scr:. ling factors to number of I observations and thus the scale factors are more precisely known. The large variability of the bounding values renders that approach j impracticable.
2 2
- b. In testing the hypothesis that op p cP/X, r that the variance of the i y values is not equal to the variance o the residuals, the usual 2 2 notations a and S exchange meanings - (S = P (not detecting op p cP/X when it is true} and a = P (detecting o p p c /X when it is not. S be-P comes the significance level and 1-n the power of the test. Testing at the levels 1-a = 0.5 and 8 = 0.05, which is the maximum power level pos-sible given the few data points, the values of the F statistics and of the critical values are given below.
042178
22A5365 K-22 R;v. 2 F F av air volume 1.956 1.04 water area 5.54 1.03 pool T(+) 2.588 1.02 pool T(-) 2.015 1.02 2 2 Since the hypothesis is that a p a , the acceptance region is satis-y x fled in all cases, at a significance level - 0.05 and a power level of 0.5.
- c. Graphs of operating characteristic curves for 1-s = 0.5 and B = 0.05 are not readily available, but the degrees of freedom v = d.o.f.
1 and v = d.o.f. residual, are given below: 2 "1 "2 air volume 7 5 water area 8 6 pool T(+) 9 7 pool T(-) 9 7 These determine the operating characteristic curve.
- d. Neither bounding curve was used. See the answer to K3.12a.
- e. The factor 1.3/1.5 arises from the fact that, as the air volume increases, positive loads increase faster than the negative loads.
Typically for a 50% increase in positive loads, the negative load increase is approximately 30%. Therefore, the scaling factor for negative loads is taken to be smaller than that for positive loads by a factor of 1.3/1.5. Small-scale test data to justify this scaling factor is provided in Referec e 14. O 042178
,s .
22A4365 ' K-23, Rev. 2___ QUESTION K3.13 - Figure K-8 shows that the bubble pressure is sensitive to fL/D of the S/R valve discharge line. Provide the following:
- a. experimental result to verify this sensitivity; and b, discuss why this parameter has not been identified in the small-scale tests and has not been included in the scaling parameters.
1 ANSWER TO QUESTION K3.13 The curve plotted in Figure K-8 shows the relationship between S/R Valve Dis-charge Line air volume and parameter fL/D for the air leg in pipe for a peak pipe pressure of 550 psid. The area above and to the left of the curve depicts region of pipe pressure less chan 550 psid and the area below and to the right of the curve is the region of pipe pressure greater than 550 psid. O It should be noted that bubble pressure depends only on the air volume and not on parameter fL/D. The values of bubble pressure shown on the right-hand side ordinate in Figure K-8 correspond uniquely to the air volume on the left-side ordinate and have been taken from Figure All.4, which shows the relationship between Bubble Pressure and Pipe Air Volume, independently of parameter fL/D. QUESTION K3.14 The maximum and minimum loadings from nultiple S/R valve lifts are calculated to result in the same loadings as from a single S/R valve actuation. Justify this calculated result analytically and experimentally. ANSWER TO QUESTION K3.14 ) The =ethod for calculation of attenuated pressure at some location on the pool boundary due to single and multiple S/R valve actuation is described in l 042178 . J
22A5365 K-24 Rev. 2 , Sections A10.3.2.2 and A10.3.2.3. In the event of multiple S/R valve actuation, the attenuated pressure at a point, r, is calculated using the following equation: h i I n 2 1/2 AP(r) = I AP
,n=1 ",
where AP = attenuated pressure at r due to n S/R valve n = number of S/R valves slowing down simultaneously. If the calculated pressure is ap(r) > AP , the B maximum bubble pressure, set ap(r) = AP . This directly follows from the assumption of incompressible B Potential Flow, for which the continuity equation gives the following: 7 4=0 (1) where $ is the Velocity Potential. Also, the Bernoulli equation for incompressible potential flow reduces to: 2 N+2 at
+ E + gy o
=
F(t) (2) where F(t) is a function of time only. Equation 2 could also be written as: 7 (E+2 at + E + gy)I p j
= 0 l 2h or, bat (7 $) + 7 l b l
+ 7P
= 0 (3)
(2/ p For one or more free oscillating (in-phase) bubbles inside the pool, V = 0 at the extremes of oscillations and, since v $ = 0 from Equation 1, Equation 3 reduces to
~~
7P=0 (4)
- 042178
1 22A4365 K-25 Rev. 2 _ t (i e., the pressure distribution incide the pool satisfies Laplace's equation. 1 Hence, for a single or multiple bubbles oscillating (in-phase) in the pool, at the extreme of oscillations,. pressure at any point in the pool cannot exceed the maximum pressure; neither can it be less than the minimum pressure inside the bubbles. QUESTION K3.15 - Figure AS.11 shows the prediction of idealized quencher bubble pressure oscilla-1 tion in suppression pool based on the Raleigh bubble model. It is noted that the Raleigh bubble model does not result in good prediction as reported in the Topical Report NEDO-10859. This model has been revised by the consideration of energy dissipation and results in better agreement as presented in the Topical Report NEDE-20942-P. Regarding this modified model, provide the following:
- a. reanalyze the bubble pressure oscillation by using the modified model;
- b. assumptions used in the analysis;
- c. comparison between the predicted bubble pressure oscillation and the small-scale as well as the large-scale test results; and,
- d. discussion of the effect of boundarf conditions on the bubble pres-sure oscillation.
ANSWER TO QUESTION K3.15 . Mark III containment Air Clearing Loads for quencher were predicted from the small and large-scale test data and using the methodology described in Section A12.7. The design frequency range for these loads is based on the observed frequency data in the large-scale tests, as shown in Figure AS.12. It should be noted that ] I Raleigh's bubble model was not used to predict the load and frequency values for Mark III quencher air clearing loads; however, this model was used to define an approximate idealized wave shape based on predicted load and frequency values, and which was simple enough to be used for structural analysis. 042178 j
l 1 22A4365 R;v. 2 K-26 1 QUESTION K3.16 O Section AS.1 presents a formula for calculating the absolute pressure on the pool walls. In this formula, hydrostatic head is included as one of the param-eters. Because of this consideration, the negative pressure on the wall will be reduced substantially. Provide justification for including the hydrostatic head in the calculation of absolute pressure on the pool walls. Include the following:
- a. Verify that the pressure transducers used in the small scale and large scale had been adjusted so that a zero reading on the transducers represents containment air pressure plus the static head of water due to submergence of that transducer;
- b. It is noted that the pressure transducers were located in various eleva-tions of the pool in the small-scale and large-scale test. Verify which pressure transducers gave the maximum negative prassure;
- c. Discuss how the static head of water will be incorporated in the statistical techniques to scaling the small-scala test for prediction g of Mark III loads.
ANSWER TO QUESTION K3.16 The air clearing loads for Mark III Quencher, listed in Section A5.6.1, represent the fluctuation (or dynamic component) of absolute load on the pool boundaries. Hence, to calculite the absolute load, the hydrostatic head (or static component) must be superirosed over the fluctuation load (or dynamic component) . Sec-tion AS.1 presents the equation for calculation of the absolute load.
- a. The air clearing loads on the floor and valls of the pool (as determined from small and large-scale tests) represent the fluctuation of loads over and above the initial static head. Hence, a zero reading on the pressure transducers represents wetvell air pressure plus the static head of water due to submergence of that transducer.
O 042178
l 1 1 22A4365 Rev. 2 K-27 m The maximum positive and negative pressures were, in most cases, read b. by the pressure transducers closest to the quencher. In the case of small-scale test (Figure A12.2), the pressure transducers PS , P 6, P7 and P read the maximum positive and negative pressures. 3 In the case of large-scale test, the maximum values of both positive and negative loads were measured, in most cases, by the pressure . transducer DA3 (Figure A12.7) located in the plane of the quencher and on the containment wall, and, in some cases, by transducer DA4 located right below the quencher. 5
- c. Test data indicate that for submergences between 4m and 6m. pressures are not affected by submergence. Since all the data used in the analysis fell within that range (as does Mark III quencher submergence),
the static head of water was not incorporated in the statistical analysis. QUESTION K3.17 Section AS.6.1 specifies that the design negative pressure for Standard 238 plant is (-) 11.0 psid, while the calculated peak negative pressure is also (-) 11.0 psid. No margin is allowed for the design. Provide justification for this specified design pressure. ANS'4ER TO QUESTION K3.17 The mean expected negative pressure for the Standard 238 plant with the S/R dis-charge pipe routing used in the analysis is only (-) 6.7 psi. The design load of (-) 11 psi represents a 39% margin. A design load of 90-90 confidence level is a conservative estimate of the maximum expected load. The recommended design, positive or negative, for any Mark III containment are the 90-90 confidence level loads determined by the S/R discharge piping arrangement using the methods identi-fied in Section AS. Section A5 has been revised accordingly. I O
)
1 042178
22A4365 g 28 Br.v. 2 QUESTION K3.18
- O In your respcuse to Question K2.9, it is stated that: "At local suppression pool temperatures in excess of approximately 160*F and at moderate-to-high relief valve flow rates, severe, continuous structural vibrations were encountered."
And the 160*F is specified as the allowable limiting temperature for the sup-pression pool water. Subsequently, you responded to Question K2.10 that pool temperature monitoring locations shall be within 30 ft of each safety / relief discharge location. Provide the following:
- a. Justify the adequacy of the specified location of pool temperature sensor, and
- b. Provide more specific recommendation of the temperature sensor location .
such as the relative location of the sensor 'o the relief valve discharge. ANSk'ER TO QUESTION K3.18 The requested information has been provided in the revised response to h Question K2.10. QUESTION K3.19 The dynamic pressures for the base mat, dryvell wall and containment wall shown on Tables AS.1 through A5.4 appear not following the method of dynamic pressure calculation as shown on Section A10.3. Justify this inconsistency. ANS*4ER TO QUESTION K3.19 A re-analysis of the Section A10.3 procedure and the data shown in Tables A5.1 through A5.4 does not reveal any inconsistencies. O' l ( l 042178 ) l l
22A4365 Rev. 2 K-29 The following are responses to questions received in May 20, 1976 letter from v J. F. Stolz. QUESTION K4.1 Update topical repor, 11314-08 (nonproprietary) and NEDE-11314-08 (proprietary) entitled: "Information Report Mark III Containment Dynamic Loading Conditions" to include the modified method of quencher load prediction which was presented on April 2, 1976 at a meeting in Bethesda, Maryland. Based upon the discussion held at that meeting, we found the methodology used to predict the loads acceptable to us. Thus, to confirm our understanding, include the following:
- a. description of the modified method of quencher load prediction;
- b. detailed calculation of the maximum positive pressure and negative pressure for the condition of one SRV discharge;
- c. reanalysis of the pressure field of the surrounding structures of the suppression pool based on the loads calculated by the modified method; and
- d. reanalysis of the reaction loads for the quencher supports.
ANSWER TO QUESTION K4.1
- a. The modified method for establishing the design value quencher bottom pressures is provided in revised Section A12.
- b. The calculation of the maximum positive and negative bottom pressures are provided for the single SRV discharge case, as well as for the ADS and all SRV discharge cases in revised Section A12. l
- c. The bottom pressures determined from the methods presented in Section A12 are used as described in Section A10 to develop the sup- l pression pool pressure fields for the various load cases considered e 11 1d e ste nd ris r O ter de isn- the vre ret the 1 a c -
042178
22A4365 K-30 R;v. 2, . have been normalized so that boundary loads can readily be determined for variations in design value bottom pressures due to TP.V discharge llg line configuration changes. Table A4.4 summarizes the bottom pressures due to SRV discharge line configuration changes. Table A4.4 summarizes the bottom pressures and containment wall pressure (at Pt. 10) for the present Std. 238 Mark III SRV discharge line routing.
- d. The reaction loads for the quencher supports are shown in Tables A7.1 and A7.2, as well as Figure A7.3.
QUESTION K4.2 Provide justification for the following assumptions, which are made as part of the GE-proposed criteria (letter from I.F. Stuart to R.S. Boyd on April 7,1976):
- a. no pressure transient of the primary system would cause more than one SRV operating in multiple actuations;
- b. for determining the sequence for a group of SRV's discharge, the reactor pressure transient of 140 psi per second is considered the most severe pressure transient for SRV actuation; and
- c. the basis for selection of the groups in (b) .
ANS'JER TO 0.UESTION K4.2
- a. As shown in Figures A4.3, A4.4, A4.5 and A4.7, all groupings of SRV's by setpoint pressure limit the number of low setpoint valves (1103 psi) to one. The pressure transient events listed in Table A9.la are con-sidered in design. For isolation events, once the initial transient is over, the cycling of this low setpoint valve will remove decay heat until such time as, another normal path of heat removal is established.
It is conservatively assumed that this will take 30 minutes to perform. O 042178
l 22A4365 K~31 Rev. 2 i
- b. The valve sequence and rate of pressure increase for the most severe transient is discussed in the additional information provided in Section A10.3.2.3.
- c. The basis for selecting the SRV setpoints, setpoint groupings, and discharge location in the suppression pool include consideration for:
(1) minimizing the SRV cycling following an isolation event (1 low setpoint valve); (2) symr.otrical load distribution (valves in setpoint group uniformally distributed around the pool); and (3) reactor vessel overpressure protection requirements (see Section 5.2.2). O lO 042178
Table K-1 BWR/6 S/R VALVE TRANSIENT
SUMMARY
Service Poul luta141 Reactor Temp at Scram RPV kPV ilX Water Poul Pallow Fallow Dcpream Depress initial kitR Temp. Came kcactor Temp. Initial (*F) Notes (*k) Time kale Time Time flE__ (*F) (paia) Event No. CunJittun Scram IG min 2 HX 100 200 137 AJJ1tional valves Stutk open i 105% katcJ !!O e IDO*F Uncunt r olled must be opened to after kellei Valve Power prevent RPV scram repressurization 10 min ! HX 100 200 137 AJJitional valves Stuck open 2 105% kated 110 e it0*F Uncontrolled Scram must be opened to after kellet Valve Power prevent RPV scram repressurization 20 bJ Uncontroll=J e Time 1/2hr 2 ltX 100 200 130 AJJitional valves ep Stuck Open 3 ImolateJ 120 W Time must be opened to of valve after lC 73 kellet Valve sticks scram prevent RPV La luolattun repressurisation DJCQ 0 120*F When 1/2 hr 2 HK 100 200 140 HPCS and RCIC laulation 4 thusation 120 4 Time IOO*F/hr T poog after available to main-trum 105% of and reaches scram tain RPV water level pepressurization power !=olation 120*F 1 XX 100 200 148 tiPCS and RCic avall-ipulation 5 imolattuu 100 & Time of 100*F/hr 10 min 1/2 hr able tu maintain imulation after alter out and trum 105% RPV water level Power scram scram (partial) twpreneurization Unconttulled Auto. N/A N/A 100 200 123 ADS 6 105% 100 Autu. un kated High on low-Power Drywell luw Premmure water level At time Uncontrolled N/A 10 min 2 HX 100 200 136 ADS 7 laulated 120 of alter isolation ADS O C' N H M CD
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22A4365 K-35 J Rev. 2 o n i l0 G J 3 i E w J a I z o
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1 l 22A4365 g,41 Rev. 2 APPENDIX K-A INTERIM OPERATING PROCEDURES AND POOL TEMPERATURE LIMITS A. POOL TEMPERATURE LIMITS
- 1. Continuous Power Operation - Maximum allowable suppression pool tem-
! perature during continuous power operation shall not exceed the I technical specification limit for power operation. l
- 2. Testing During Power Operation - Maximum allowable suppression pool temperature during RCIC, HPCI, or S/R valve or other testing at power which adds heat to the suppression pool shall not exceed +10*F above the technical specification limit for continuous power opera-tion. The pool temperature must be returned to the continuous 4 power limit as specified in the Technical Specifications.
- 3. Reactor Operation - The maximum allowable suppression pool tempera-ture during reactor power operation at greater than 1% of rated power shall not exceed 110*F.
- 4. Reactor Isolation - Maximum allowable suppression pool temperature resulting from isolation of the RPV and concurrent scram from A.1 above shall not exceed 120*F when operating pressure and temperature are being maintained.
e a
- 3. GENERAL OPERATING PROCEDURES
- 1. Continuous Power Operation - Should the suppression pool temperature exceed the limits of A.l. above, all available suppression pool cooling heat exchangers should be initiated.
- 2. Testing During Power Operation - Should the suppression pool tempera-ture exceed the limits of A.2. above, testing should be immediately, i
terminated. t O 042178
22A4365 R;v. 2 K-42
- 3. Reactor Operation - Should the suppression pool temperature exceed the limits of A.3, above, the reactor should immediately be scrammed and depressurized at a 100*F/hr or as required, not to exceed pool temperatures of 160*F before the RPV reaches 200 psi.
4 Reactor Isolation - Should the suppression pool temperature exceed the limits of A.4. above, the reactor should immediately be scrammed and depressurized at the rate of 100*F/hr or as. required, not to exceed a pool temperature of 160*F before r% RPV reaches 200 psi. NOTE FOR GENERN. OPERATING PROCEDURES. Events Adding Energy to Suppression Pool - Dr. ring testing or other events which could result in energy addition to the suppression pool, a man should be posted at the control panef with the specific purpose of observing the suppression pool temperature response. C. OPERATION PROCEDURE FOR SPECIFIC EVENTS RESULTING IN POTENTIALLY HIGH SUPPRESSION POOL TEMPERATURE DURING S/R VALVE DISCHARGE
- 1. Stuck Open S/R Valve at Power
- a. Scram the plant, preferably by placing mode switch in shutdown, as soon as it is recognized that valve will not close or if pool temperature reaches 110*F.
- b. Initiate all available pool cooling heat exchangers imediately.
- c. Continue attempts to close stuck-open valve.
- d. Following scram, the stuck-open safety / relief valve will immedi-ately depressurize the RPV to some lower value. Continue depres-surization through the main condenser, isolation condenser, or by opening additional relief valves. If relief valves are used, they should be separated from the stuck-open relief valve to assure uniformity of energy insertion to the pool in order to g 042178
4 22A4365 K-43 Rev. 2 C)\
's_ reduce vessel pressure below 200 psi prior to reaching pool temperatures of 160*F.
- e. Proceed with plant shutdown, i
- 2. Stuck-Open S/R Valve with Reactor Isolation
- a. A stuck-open S/R valve will result in immediate depressurization l of RPV to some low pressure depending on decay heat addition.
1 i Continue as in C.c. and d. above. t l 1
- b. All available heat removal systems are already in operation and i
should continue heat removal at maximum capacity. 4
- c. Proceed with plant shutdown.
l
- 3. Loss of Main Condenser (i.e., MSIV Isolation, Scram Automatic)
O I Plants with Isolation Condenser
- a. Initiate isolation condenser immediately to minimize energy dump to pool.
- b. Initiate all available pool cooling heat exchangers immediately if pool camperature exceeds service water temperature.
- c. Continue trying to reestablish main heat sink.
l
- d. If suppression pool temperature exceeds 120*F, initiate shutdown at 100*F/hr or as required not to exceed pool temperatures of 160*F before the RPV reaches 200 psi.
Plants with RHR Non-Condensing Type Systems () a. Activate all available RHR pool systems i= mediately after pool temperature exceeds service water temperature. 042178
22A5365 g_ g Rev. 2
- b. If full complement of RHR systems are available, plant may be maintained at hot-pressurized condition providing maximum pool temperature does not exceed 120*F.
- c. Hold RPV pressure between 800 and 1000 psi by manually actuating opposing S/R valves around the pool.
- d. Restore main heat sink as soon as possible to terminate dump to pool.
- e. If pool temperature exceeds 120*F with full complement of systems operating, depressurize plant at 100*F/hr, or as required not to exceed pool temperatures of 160*F before RPV reaches 200 psi.
- f. If a full complement of RHR systems are not available, initiate plant depressurization immediately and depressurize plant at 100*F/hr or as required not to exceed pool temperatures of 160*F before RPV reaches 200 psi.
O Plants with RER Condensing Type Heat Systems
- a. If both systems are available, place in steam condensing mode within 30 minutes,
- b. Full condensing capacity should be used to maintain pool below 120*F.
- c. Hold pressure between 1000 and 800 psi by manually actuating dif-ferent S/R valves around the pool.
l
- d. If pool temperature exceeds 120*F initiate plant depressurization l to the condensing Ex at a rate of 100*F/hr or as required not to exceed a pool temperature of 160*F before RPV reaches 200 psi.
O 042178
l l 22A4365 K-45/K-46 Rev. 2
- e. If only one of two RHR systems is available, place in condensing i mode within 30 minutes and begin plant depressurization at 100*F/ l hr using maximum condansing capacity and S/R valves if necessary.
- f. Proceed with shutdown.
D. GENERAL CONSIDERATIONS Experimental data indicates a 160*F suppression pool temperature limit will prevent excessive steam condenstaion loads during relief valve discharge with sonic conditions at the discharge exit. The following are some general recommendations to even further reduce the probability of the occurrence of excessive steam condensation loading.
- 1. Suppression pool water temperature should be maintained as low as practical below the technical specification limit by operation of the pool cooling heat exchangers during periods of low service water O c ver c=re -
- 2. Testing should be performed as follows:
- a. Before testing, cool, if possible, the suppression pool to lower temperatures.
- b. All testing should be of as short duration as possible. Also consider having RER system in operation during test.
- c. Testing should be done during low power levels if possible and preferab1v during the return to power at the beginning of each operating cycle because decay heat levels are lower.
O 042178 1 I
22A4365 Rev.,2 ._ L-1 O
%d ATTACHMENT L Containment Asymmetric Loads This attachment discusses the potential for circumferential variations in the LOCA dynamic loads and relief valve loads. The asymmetric loads are identified and the data being used for containment design evaluation is present.
Table L-1 is a tabulation of the postulated phenomena which could cause loading asymmetries. With the exception of items 6 and 12, the above table either provides a reference for the asymmetric loads that are significant and should be considered, or pro-vides a reference that justifies the assumpti.on that a particular phenomenon does not lead to asymmetric loads of significance. The following is discussion of items 6 and 12. O As discussed in section 6, the maximum containment pressure increase associated with the bubble formation that follows 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-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 could lead to an increase in the load. However, PSTF data shows a very weak relationship between submergence and the containment pressure increase caused by bubble formation. 3 The survey of the PSTF data shown in Figure 6.6 shows that for tests having the i 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 that varia-tions in suppression pool depth due to seismically induced waves will not lead to significant asymmetric containment bubble loads. i The 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 terns of vent flow l composition. Any steam in the vent flow would be condensed and this would lead 042178
Table L-1 Is there the potential ior significant Asymmetric Loads asymmetric Being Used for containment loads Design Evaluation Comments Phenomena
- 1. Seismic induced pool No See Attachment B surface waves
- 2. Seismic induced changes in Yes See Attachment B the pool hydrostatic pressure Yes See Attachment A g'
- 3. Relief valve actuation *
. s~
ro *$ No 0 Loads are of negligible
- 4. Jet Loads during vent magnitude (see 6.1.2) cleaning
- 5. Sonic and compressive waves No Both 0 Loads are of negligibic magnitude (see 4.1.1 and 6.1.1)
Yes 0-10 psi See following discussion.
- 6. Bubble pressure load
- 7. IICU floor flow pressure No O See Attachment F differential S
U No 0 Loads on the containment ra M 8. Fall back are of negligible magnitude h (see 6.1.7) .
- O e
O O O Table L-1 (Continued) Is there the potential for significant Asynenetric Loads asymmetric Being Used for containment loads Design Evaluation Comunents Phenomena
- 9. Post LOCA waves No 0 Loads on the containment are of negligible magnitude (see 6.1.8)
O This is a relatively slow
- 10. Containment pressurization No charging process. See i
Figure 4.4. IU Loads are small (see
- a-
- 11. Condensation Oscillations No 0 6.1.9) N[e No O See following discussion.
- 12. Chugging
- 13. Pool Swell loads with No O See Section 10.1 and following discussion.
seismic induced waves present 1 s~ r
" b
32A4365 Rev. 3 L-4 to a less rapid pool acceleration and thus a reduced pressure load on the con- h tainment wall. It should be noted that PSTF data shows that the high degree 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 significant circumferential variations in the containment bubble formation loeds. In addition, Attachment D shows no significant circumferential variations in drywell preseure that could lead to variations in vent flow rates and thus pool swell. Despite strong evidence that circumferential variation in the containment bubble load . will not occur, arbitrary loading combination of 0 paid on one side of the containment with a simultaneous 10 paid load on the other side should be considered to account for any uncertainties about asymmetric loading conditions. Data from General Electrics ongoing Mark III test program has shown that the containment wall does not experience any loading during vent chugging (Eighth Quarterly Progress Report, Mark III Confirmatory Test Program, NEDO-20853, April 1975 contains a discussion of this data). Thus it is concluded that chugging does not represent a soucce of asymmetric containment loads. I O 090779 j
l 22A4365 I Rev. 2 - M-1 Ci V ATTACHMENT M MULTIPLE SAFETY / RELIEF VALVE ACTUATION l FORCING FUNCTION METHODS 1 ! TABLE OF CONTENTS 1 SECTION TITLE PAGE 4 M
1.0 INTRODUCTION
M-5 M2.0 RANDOM PARAMETERS M-5 . M2.1 Reactor Vessel Pressure Rise Race M-6 , M2.2 Valve Setpoint M-6 I M2.3 Valve Opening Time M-7 4 M2.4 Quencher Bubble Frequency Distribution 'M-7 l M3.0 MONTE CARLO TRIAL SIMULATIONS M-13 M3.1 Approach M-13 . M3.2 Bubbic Arrival Time M-14 M3.2.1 Calculation of Reference Arrival Time M-14 M3.2.2 Adjusement of Bubble Arrival Time for Valve Setpoint Variations M-14 1 M3.2.3 Adjustment of Bubble Arrival Time for Valvs Opening Time Variations M-14 M3.3 Quencher Bubble Frequency variations M-15 M3.3.1 Adjustment of Qusncher Bubble Frequency for Discharge Line volume M-15 M3.3.2 Adjustment of Quencher Bubble Time History for Selected Frequency M-15 0 042178
/
22A4365 . Riv. 2 M-2 TABLE OF CONTENTS (Continued) g t s SECTION TITLE PAGE M4.0 FACTORS AFFECTING PRESSURE DISTRIBUTION ON THE SUPPRESSION POOL BOUNDARY M-15 M4.1 Bubble Pressure Attenuation M-16 M4.2 Line-of-Sight Influence M-16 M4.3 Combination of Multible SRV Pressure Time Histories M-16 M5.0 FORCING FUNCTIONS FOR NSSS EQUIPMENT EVALUATIONS M-16 M5.1 Time Sequencicg M-16 M5.2 Pressure Time Histories M-17 M5.3 Vertical Basemat Force and Overturning Moment M-17 M5.4 Fourier Spectra - M-17 M6.0 STRUCTURAL RESPONSE ANALYSIS M-18 l \ 0 l I 042178 g
U 22A4365 M-3/M-4 i Rev. 2 ,, LIST OF ILLUSTRATIONS FIGURE TITLE PAGE 1 M2-1 Probability Density Function vs Pressure Rise Rate M-8 j j M2-2 Probability Density Function vs Valve Group 4 Setpoint Variation M-9 M2-3 Probability Density Function vs Valve Opening Time < Variatira M-10 l l l M2-4 Probability Density Function vs Bubble Frequency M-11 l' M2-5 Quencher Bubble Pressure Time History M-12 MA-1 Basemat Load vs Time MA-5 MA-2 Fourier Spectrum of Basemat Force MA-6 i MA-3 Fourier Spectra of Vertical Basemat Force MA-7 i O 042178
22A4365 v-5
~ Rev. 2_
O 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 suppression pool during multiple 3RV" discharge events. The random varicbles 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 future 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 Attachment A. It should be noted that test data indicated randomness in the peak
'd pressure amplitude which could also be used for determining structural response.
This is also being studied for future application. Of the SRV cases identified for consideration in containment structural design (Table A4.4 of Attachment A), the expected bounding vertical rc:.ponse at equip-ment locations is based on the all valve case. The expected bounding horizontal 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 these four cases, the Fourier Spectra of the forcing functions for 59 Monte Carlo simulations of the event are plotted. A bounding forcing function is i then selected in each of the frequency ranges of interest for use in developing the dynamic responses at a selected location on the containment structure (i.e., basemat, drywell, and containment). These dynamic responses are then employed for NSSS and B0P equipment evaluations. A dynamic time history analysis is performed to determine the acceleration time histories, respon'se spectra, and displacements needed. Dynamic responses for equipment evaluations are made by l p . 101678
22A6365 M-6 Rev. 2 enveloping the results from tl's selected trial cases with the largest Fourier g 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 probability 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 setpoints; thus, the valves within the group 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.11 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 over 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 applies, i.e., two-third decay over 5 cycles, or 0.133 decay per cycle, i The quencher bubble pressure time history in Figure AS.ll of Attachment A is an j 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 j and negative portions as shown in Figure M2-5. The P and P g ratios 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 frequency nf the air bubble is a function of the SRV discharge line air volume. The distribution of bubble frequencies for a discharge line air volume of 50 ft. is shown in Figure M2-4 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 data was 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 i bounds of 5 and 12 Hz. 1 101678
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22A4365 M-13 Rev. 2 o C M3.0 MONTE CARLO TRIAL SIMULATIONS M3.1 Approach l l There are four SRV cases that are considered to get bounding forcing functions l for the equipment evaluations. They are: i i _ Single Valve subsequent actuation Two adjacent valves ADS valves 4 - All valves In each of these cases, 59 Monte Carlo crials are performed in which appropriate l 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 two adjacent valves cases, the valve set point tolerance and pressure rise rate considerations are not incorporated for obtaining the forcing function because the entire group of ADS valves is simultaneously acti-i vated by a single signal. For all valve case all variables are considered. The all valve trials each consist of selecting a random pressure rise race 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 ti=e differ-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. Once the bubbles are in the suppression pool, each bubble frequency is randomly varied by selecting a frequency from a unique distribution for the discharge line volume involved. See Figure M2-4 for typical distribution for discharge line 3 with an air volume of 50ft . The bubble time history for each valve location O 101678
22A4365 g_14 Rev. 2 , is then used to determine the forcing function on the suppression pool boundary ggg by utilizing the methods described in Section 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-domly selected as for the multiple valve trials. For the single valve case only the bubble frequency is varied. M3.2 Bubble Arrival Time y0.2.1 Calculatioa of Reference Arrival Time The arrival time for each air bubble in the suppression pool relative to the icwest set SRV is a function of the SRV setpoint arrangement and the reactor pressure rise race. Assuming no telerance on setpoints, no variation in valve cpening time (VOT), and randomly selecting a pressure rise rate (PRR), the arrival times of the bubbles in the suppression pool are computed by dividing 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. s 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 I as calculated in Section M3.2.2 by slightly increasing or decreasing the VOT O 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 sho.in in Figure M2-3. 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 f t to the actual air volume in the SRV discharge line. For example, the adjustment of frequency for a 100 ft line volume is: , 3 8.1 Hz xg/gg50= 8.1 x 0.79 = 6.4 Hz O Examples for the other characteristics: Volume Mean Std. Dev. Lower Bound Upper Bound (f t ) (Hz) (Hz) (Hz) (Hz) 50 8.1 1.7 5 12 100 6.4 1.3 4 9.5 l M3.3.2 Adjustment of Quencher Bubble Time Historv for Selected Frequency In each Monte Carlo trial, a random number generator code is used to select a frequency from each of the frequency distribution curves generated in Section y 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 positive pulse period constant. The pressure cycle period, positive pressure pulse time and negative pressure pulse time are adjusted by multiplying each by the ratio of the reference frequency (8 Hz) to the selected frequency. For example, for 6 Hz: l l 101678
22A4365 M-16 I Rev. 2
'f ;
Pressure cycle period = 0.125 sec. = 0.167 sec. 6 Hz 8 Positive pressure pulse time = 0.05 sec. = 0.067 sec. 6 Negative pressure pulse time = 0.075 sec. = 0.100 sec. Number 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 from the quencher is 2rg/r where r - radius of the quencher - (4.87 f t) and r 12r, (see Section . A.10.3.1 of Appendix 3B) . r = true spatial distance from the quencher center to the node. M4.2 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 frem that quencher (see Section A.10.3.2.1 of Appendix 3B). M4.3 COMBINATION OF MULTIPLE SRV PRESSURE TIME HISTORIES The time sequencing application provides a given phase relationship between h quencher bubbles. The pressure at each node point and time acep is calculated by combining the contribution from each valve (in the line of sight) using algebraic sumation. At each node where the 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 equal to the maximum bubble pressure at the same instant in time. O 101678
22A4365 M-17 Rev. 2 M5.0 FORCING FUNCTIONS FOR NSSS EQUIPMENT EVALUATION a i I MS.1 TIME SEQUENCING I 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 equipment evaluations. The bounding forcing function from 59 trials will result in a 95% confidence level that 95% of the time the actual forcing I function will be less than the forcing function determined by the Monte Carlo l .
- technique.
MS.2 PRESSURE TIME HISTORIES Fif ty-nine (59) cases of pressure distribution on the pool boundary are calcu-lated using the random parameters delineated in Section M2.0. MS.3 VERTICAL BASEMAT FORCE AND OVERIURNING MOMENT 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). MS.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 into three frequency intervals as determined below:
l
- Reference 1: Cooley, J.W., & Tukey, J.W., (1965), "An Algorithm for the Machine Calculation of Coaplex Fourier Series," Mathematics of Computation, Vol.19, No. 90, pp 297-301.
- 2. Shingleton, Richard C., "On Computing the Fast Fourier O' Transform," Communication of Applied Computation Mathematics, (10(10) 1967, pp 647-654, 101678
l i 22A4365 M-17A ~ Ref. 2 i O Step 1. Adjust the mean frequency of each safety / relief valve discharge line for air volume differences, see Sub-section NG.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 fa, and 2.5 fm to 3.5 fm. I where fm = N r Ii; i = 1,...,N N = total no. of valves actuated The basemat loading cases with the largest spectral value within each , frequency interval (from the 59 cases) are selected for determination of equipment responses. O i l [ h I (:J' l i 101678
28A4365 g,gg . Rev. 2 rt6.0 STRUCTURAL RESPONSE ANALYSIS g Forcing functions corresponding to the case selected in each frequency range (solacted in Section M5.,) are used as input to the structural .snalysis. Structural dynamic analysis in then performed for these selected cases. The resulting dynamic responses are then envoloped for NSSS and B0P equipment evaluations. O t 0 042178
22A4365 MA-1 ~ Rev. 2 . O APPENDIX MA I I 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 basemat 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: (1) Pressure rise rate distribution per Subsection M2.1. n
%J (2) Pressure setpoints variation per Subsection M2.2.
Mean Standard
, Setpoint Deviation Valves (psi) (esi) 1 1103 2 9 1113 2 9 1123 2 (3) Valve opening ctme variations per Subsection M2.3 ,
Standard deviation = 0.009 sec. Step 1 An 80 psi /sec vessel pressure rise rate was randomly selected from Figure M2-1.
- Note that this example is for the 238 BWR/6 Mark III standard plant with a ganged valve arrangement. ,
i I' l 101678
- ,-r-
22A4365 Rev. 2 MA-2 - Step 2 , l The valve pressure setpoints are randomly selected from a random number generator code using the' distribution given in Figure M2-2. The valve pressure set-points from a typical random selection are 1104.5 psi,1114.3 psi, and 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 setpointg) (psi) i sec) = Pressure rise rate (psi /sec) where i = 2, 3 (the number of subsequent valve groups), and 1 = the reference valve. Hence, for i = 2, the valve opening time for the first group of 9 valves is: T , 1114.3 - 1104.5 = 0.1225 sec 2 80 Step 4 The bubble arrival time is calculated by adding the group valve opening time and a randomly selected delta time for each valve using the valve opening time distribution shown in Figure M2-3. Therefore, for each quencher the bubble arrival time = T p) + individual valve opening time (IVOT). , 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 1 g- Valve No. IVOT (sec) Valve No. IVOT (sec) Valve No. IVOT (sec) (_g) 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 4, Note that a mean value of 0.057 see is included in the above numbers. Adding these values to the group Tg calculated in Step 3 and normalizing to have the
] ; first bubble arrive at zero time results in the following bubble arrival times:
i Arrival Time Arrival Time Arrival Time Valve No. (sec) Valve No. (sec) Valve No. (sec) i 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
)
M.3 BUBBLE FREQUENCIES 1 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) j i 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 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 O. e = 1.80 Hz. With nonuniform line lengths, Subsection M3.2.1 is used to develop unique QBF distribution curves from which a frequency is rariomly selected for each line. 101678
\
22At365 MA-6 j Rev. 2 - M.C The forcing function is calculated by computing the pressure distribution h around the pool boundary using the criteria defined in Section M4.0 which are: (1) 2r,/r attentuation, r = quencher radius and r 2 2r,. (2) Line-of-sight influence. (3) Algebraic summation at each time step of the individual pressure wt es . . (4) Truncation of the total calculated pressure to the maximum bubble pressure of any of the pressure waves in the pool at eacn 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 gg 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 intervals. Out of thesa 59 trials, the maximum crial 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 dynamic analynes are performed for these potential critical cases. The resulting dynamic responses are then enveloped for NSSS and BOP equipment evaluation. 101678 O 1 l l
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22A4365 _77g_g Rev. 2 - l O n EXAMPt.E l RUN 12 RUN 48 - - - == R UN 26 RUN 12 M RUN 37
. [-- M RUN 48 3
$ RUN 26 l
i 5 / \
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% s e l d-RUN 37 k" o---
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. . , FREQUENCY U 0
- 1 2 [3 fi = PREQUENCY INTERVAL.
O Figure MA-3. Fourier Spectra of Forcing Functions
; NOTES:
- 1. 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 4
'l
- Run 12 is max. for frequency interval, f;
'2
- Run 26 is max. for frequency interval, f 4
' ~3 l
- 3. Run 37 is a typical non-maximum case. l
- 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
042178
22A4365 N-1 Rev. 2 l ATTACHMENT N SUPPRESSION POOL THERMAL STRATIFICATION N
1.0 INTRODUCTION
1 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 f because of the energy release. For the Mark III suppression pool configurar. ion, 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 , O' and N-2 present the typical temperature proilles 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 1 25 sec), indicating concentrated energy discharge through the top vent. As blowdown proceeds (t 1 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 f rom 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 i blowdowns of the same break size. This is attributed to the smaller total energy release associated with the steam 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 () stratification than small breaks because energy deposition in the pool is more 042178
22A4365 N-2 Rev. 3 rapid. Since the specific heat of water is essentially constant within the temperature range from 70*F to 200*F, tne temperature rise of the pool is independent of the initial pool temperature for a given amount of energy input. As a result, the initial pool temperature has little effect on thermal stratification. N1-3 APPLICATION To determine the maximum temperature 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 Ceries 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 Pool % of Total Energy From Pool Top Depth (H1/H) Deposition (E ) 1 20 23 2 20 23 g 3 20 22 4 20 20 5 20 12 1 To obtain the temperature profile for a prescribed initial pool temperature (Tg ) and total blowdown energy, the bulk pool temperature (T) from energy balance at the end of the blowdown was calculated, then the mean temperature l l (T ) for each pool segment was determined from: 1 T = E (T - T ) +T th where H is the total pool depth, 1H is the height of the i segment, and E is the fraction of total energy deposited in the 1 E segmen*. Assuming the mean temperature of each segment occurs in the middle of the segment, the temperature profile is readily plotted. Note the above table is valid only for a top vent initial submergence of 7.5 ft. O 090779 l i
I 22A4365 f Rev,. 2 N-3 q , O N1-3.1 Stratifir ation 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 lbm of water was assumed. The mean pool temperature after energy release is: 4 x 108 = 100 + 50 = 1500F T = 100 + 8 x 106 x1 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 = 1300F s rigure N-3 shows the resulting pool temperature profile. Note that, although the temperature difference from top to bottom is almost 300F, the peak tem-perature is only 7.50F above the mean. N1-3.2 Stratification During Intermediate and Small Break Accidents Figure N-4 shows the localized nature of the energy addition as observed in the PSTF Full Scale Tests (Reference 16) . The localized energy addition (through the top vent) from the full 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-sentially unchanged and the upper pool region was uniformly heated. This thermal stratification profile will not persist in actual conditions, since ECCS suction and return will promote pool mixing. The long term profile will essentially be as shown in Figure N-3. 101678
22A4365 N-3a Rev. 2 . \ iO Since Figure N-3 from 1//5 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 recommended for design evaluations that the maximum temperature gradient shown in Figure N-3 be applied from the lower pool region up to the top vant centerline. For the upper parts of the pool (above top vent centerline) the temperature profile, from full scale and 1//3 scale tests , shows uniform heating (Reference Figure N-3) . ) I 1 1 () 1 i, 4 i i N 4 i 1
)
101678 i
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22A4365 - Par. 2 N'4
~1 i <
O This figure is PROPRIETARY and is provided under separate cover. O l Figure N-1. Typical Transient Temperature Profiles Near Drywell Wall, Kun 22 101678
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. 1
-22A4365 -
Rev. 2 N-5 . l o . ~r*- ; .49 . l t ,- g Ay - f ;:~ . k' i 4 1 { t 4 i 1 This figure is PROPRIETARY and is provided under seperate cover. } i, O f I 3 3 i s I 1 l
/
1 l I t- , j - Figure N-2. Typical Transient Temperature Profiles Near Containment Wall,
- r -- Run 22 b
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l 22A6365 N ~ Rev. 2 24 g _ _ _ _ _ _ _g REE SURF ACE INITI AL POOL TEMPERATURE 100'F a 106 g L 16 POOL DEPTH 20 ft 4x10 B8 n, TOTAL ENERGY RELEASE FINAL BULK POOL TEMPER ATURE 150*F 2 . . . TCP VENT CENTERLINE b 12 - G e O S - T_FINAL T-INITIAL 4 - I BASEMAT j I t 140 160 180 100 120 POOL TEMPER ATURE (*F) Suppression Pool Temperature Profile for Large Breaks O, Figure N-3. 101678
22A4365
- Rev. 2 N-7 O
1
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(S 2) I 170 5 no 7*C) INITIAL POOL TEMPERATUR E 15 120D (48 9'C) INITIAL POOL TEMPER ATURE I
" 8 705 (21 C) !
p., INITIAL POOL TEMPERATURE 13 \ 1 - efO:?@)
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d LEVEL PROBE j (2.l) - A A O ! I I I I I I l_ 1.0 2.0 3.0 4.0 5.0 6.0 ' (0.3) (0.6) (0.9) (1.2) (1.5) (1.3) DISTANCE FROM ORYWELL WALL - ft (mi O Figure N-4 Postulated Maximum Steam Bubble Travel As a Function of Pool Temperature (Reference Test 5707) 101678
w 22A4365 ~ Rey. 2 1 ATIACHMENI O l DIGITIZATION OF FORCING FUNCTION FOR CONDENSATION OSCILLATION J O 101678
O O ATTACILMEll ,L O Mark III Condensation Oscillation Forcing Function for t - 3.0 to 30.0 Second TIME PRESSURE TIME PRESstmE TIME PRESSURE TINE ' PRE S91RE Tl4E PRE SSURE TIME PRESSuGE ( SEcl (PSIO) (SEC) (Psini (SEC) (PSin) (SEC) (PSIDI (SEC) (PSIO) ( SEcl IPSIO 3.000 0 3.513 4.3476 1.907 0.0000 4.462 -3.9782 4.918 0 5 . 34 7 3.6317 3.010 2.8550 3.522 4.5573 4.007 0.1085 4.471 -3.8659 4.920 2.3142 5.355 3.9089 3.021 5.2012 3.532 4.5921 4.016 -0.0129 4 . 48 0 -4.0328 4.9 28 4.2160 5. 36 4 3.8390 3.031 6.6810 3.542 4.2596 4.026 -0.5001 4.489 -4.5403 4.9 37 5.4855 5.373 3.5601 3.0 41 7.1869 3.552 3.5169 4.0 35 -1,3191 4.498 -5.2716 4.946 5.8254 5 . 381 2.9394 3.052 6.8755 3.562 2.4865 4.044 -2.3701 4.507 -5.9475 4.955 5.5732 5.390 2.0782 3.062 6.0942 3.572 1.4047 4.054 -3.3525 4.517 -6.2168 4.943 4.9399 5. 398 1.1740 3.072 5.2487 3.581 0.5246 4.063 -4.0605 4.526 -5.7791 4.972 4.2546 5.407 0.4304 3.003 4.6683 3.591 0.0135 4.073 -4.3774 4.535 -4.4990 4.988 3.7784 5.416 0.0113 3.093 4.4692 3.601 -0.1138 4.082 -4.3442 4.544 -2.4694 4.990 3.6226 5.424 -0.0958 3.104 4.5990 3.618 0.0000 4.092 -4.1443 4.553 0 4.998 3.7279 5.433 0.0000 3.814 4.8208 3.628 0.11 38 4.101 -4.0273 4.562 2.3846 5.007 3.9077 5.448 0.0951 3.124 4.8576 3.631 -0.01 35 4.118 .-4.2005 4.571 4.3443 5.016 3.9375 5.450 -0. 0l l 1 3.1 35 4.5059 3.641 -0.5246 4.120 -4.7299 4.580 5.5803 5.025 3.6524 5.459 -0.4385 3.14i 3.7203 3.650 -l.4048 4.830 -5.4918 4.589 6.0028 5.033 3.0156 5.467 -1.l741 - 3.155 2.6102 3.660 -2.4866 4.139 -6.1958 4.598 5.7427 5.042 2.1320 5.476 -2.0782 3.166 1.4859 3.670 -3.5170 4.149 -6.4764 4.607 5.0902 5.058 1.2044 5.484 -2.9394 3.876 0.5549 3.680 -4.2597 4.158 -6.0205 4.616 4.3840 5.060 0.4498 5.493 -3.5602 3.1 86 0.0143 3.690 -4.5921 4.167 -4.6869 4.625 3.8913 5.068 0.0186 5.502 -3.8380 3.197 -0.1204 3.700 -4.5573 4.177 -2.5725 4.634 3.7328 5.077 -0.0976 5.500 -3.8089 po N 3.207 0.0000 3.709 -4.3476 4.1 86 0 4.643 3.8413 5.086 0.0000 5.519 -3.6316 E% 3.217 0.1204 3.789 -4.2249 4.196 2.4696 4.652 4.0266 5.095 0.0976 5.527 -3.5311 *
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PRE SSURE TIME PRESSURE T I 'tE PRE SSURE TIME IPSIDI PHESSURE 7 (4E PRE SSORE ISEC) (PSIDI ( SEC) TIME PRE SSURE TIME ISEcl I PS10) 7.H28 - 3.1085 (PSIO) (SEcl (PSIDI I SEC) IPSini 3.3379 7.425 0.0000 (SEcl 6.61 0 0 7.019 0.0872 7.837 - 3.2151 5.774 0.0000 6.895 -1.4905 7.027 3.51 99 7.43) 7.845 -3.3533 6.203 - 3. 3919 6.618 2.1064 3.5467 7.441 -0.0304 5.783 0.0931 6.626 3.8315 7. 0 35 -0.4018 7.853 -3.7760 5.798 -0.0lll 6.211 -3.5378 4.9294 7.044 3.2899 7.450 7.061 -4.3842 6.220 -3.9817 6.634 2.7163 7.458 -l.0759 5.800 -0.4290 5.3026 7. 052 7.869 -4.9462
-l.14H8 6.228 -4.6251 6.642 l'9204 7.4 66 -l.9045 5.308 6.237 -5.2181 6.658 5.0729 7.060 .
7.474 -2.6938 7.877 -5.1702 5.811 -2.0335 6.659 4.4964 7.068 1.0849 -3.2626 7.885 -4.8062 5.82i -2.876I 6.245 -5.4546 3.8726 7.076 0.4051 7.482 7.093 -3.7416
-3.4H15 6.253 -5.0704 6.667 0.0104 7.490 -3.5872 -2.0537 5.834 6.675 3.4392 7. 0H 4 7.498 -3.4906 7.901 5.842 -3.7553 6.262 -3.9474 3.2974 7.092 -0.0879 7.909 0
{ 5.850 -3.7269 6.270 -2.1666 6.68 1 3.3932 7.101 0.0000 7.506 -3.3300 7.917 2.044l
-3.5554 6.278 0 6.692 0.0879 7.514 -3.2359 5.859 6.700 3.5569 7.109 7.925 3.7239 S.867 -3.4550 6.287 2.133? 3.5840 7.117 -0.0I04 7.522 -3.3758 7.933 4.78 31 6.295 3.8870 6.706 -0.4052 7.5 30 -3.8005 5.876 -3.6035 4.9929 6.716 3.3245 7.125 7.519 -4.4126 7.948 5.1455 5.884 -4.0577 6.303 2.7449 7.133 -1.0850 7.949 4.9226
-4.7883 6. 3f l 5.3710 6.724 -1.9205 7.547 -4.9783 5.893 6.733 l.9406 7.841 7.957 4.3632
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-5.l649 6.345 3.4835 6.757 0.0I05 7.579 -2.0670 7.981 3.1997 5.927 -4.0208 3.3399 6.765 -0.0888 7.174 -3.5199 7.587 0 7.989 3.2927 5 .9 35 -2.2069 6.353 6.774 0.0000 7.182 - 3.35 79 7.997 3.4515 5.944 0. 6.368 3.4370 7.190 - 3.2611 7.595 2.0539 3.4779 E "*
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6.403 1.9657 6.815 -l.9407 7.635 4.3841 8.037 1.0638 5.9 86 5.2183 6.823 -2.7449 7.231 -5.2474 8.045 0.3973 5.994 4.6253 6.411 f.1104 7.239 -4.3780 7.643 3.7759 0.0102 6.419 0.4847 6.8 31 -3.3246 7.651 3.3533 8.053 6.002 3.9836 0.0107 6.839 -3.5840 7.247 -3.7975 7.659 3.2858 8.068 - 0.08 62 6.011 3.5 377 6.427 6.847 -3.5569 7.255 -2.0843 8.069 0.0000 { 6.019 3. 3919 6.436 -0.09 00 6.856 -3.3932 7.263 0 7.6 68 3.3085 8.077 0.0862 6.027 3.4905 6.444 0.0000 7.272 2.0672 7.676 3.4681 3.6588 6.452 0.0900 6.864 -3.2974 3.7660 7.684 3.4945 8.085 -0.0102 ' 6 .0 36 6.872 -3.4392 7.280 3.2415 8.093 -0.3973 6.044 3.6868 6.460 -0.0107 6.800 -3.8727 7.288 4.8175 7.692 8.101 -1.0639 3.4198 6.469 -0.4147 7.296 5.2038 7.700 2.6743 6.052 -l.11 05 6.888 -4.4965 7.708 1.8922 8.809 -1.8832 6.06l 2.8235 6.477 6.897 -5.0729 7.304 4.9783 8.887 -2.66S6 6.069 I.9963 6.485 -1.9657 7 . 31 2 4.4126 7.786 1.0689' l.1277 6.494 -2.7003 6.905 -5.3026 3.8004 7.724 0.3992 8.125 -3.2261 6.078 6.913 -4.9293 7.320 0.0103 8.133 -3.4779 6.086 0.4211 6.502 -3.3674 6.921 -3.8374 7.328 3.3751 7.732 8.141 -3.4515 6.094 0.0108 6.510 -3.6302 6.930 -2,1063 7.336 3.2359 7.740 -0.0866 8.149 -3.2927 6.103 -0.0914 6.588 -3.6027 7.344 3.3100 1.748 0.0000 l 6.527 -3.4369 6.938 0 3.4906 7.756 0.0866 8.157 -3.1997 6.111 0.0000 6.946 2.0845 7 . 35 2
-0.0103 8.165 - 3.3373 6.189 0.0914 6.5 35 - 3. 3399 7.361 3.5872 7.764
-0.0109 6.543 -3.4835 6.954 3.7976 3.2626 7.772 -0.3992 8.573 -3.7579 f 6.128 6.962 4.8781 7.369 8.181 -4.1633 6 .1 36 -0.4212 6.552 -3.9226 5.2474 7. 317 2.69 37 7.700 -1.0690 8.889 -4.9226 f
-1.8278 6.560 -4.5544 6.970 8.9045 7.7 88 -1.8923 6.144 6.978 5.0201 7. 3d5 8.197 -5.1455 6.153 -I.9963 6.5 68 -5.8383 4.4496 7. 39 3 1.0759 7.796 -2.6164 8.205 -4.78 33 6.576 -5.3710 6.987 u.4usu 7.804 -3.2486 6.161 -2.8236 4.995 3.8323 7.401
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---i O O O TIME PRESSURE TIME PRE SSURE TIME PRESSURE TIME PRESSURE TIME PE3SSURE TIME PRESSURE (SEC) (PSIDI ( SEC) IPSID) (SEC) (PSIDI (SEC) (PSID) (SEC) '(PSIDI ( SEC) (PSIDI 8.229 O. 8.628 3.2757 9.024 0.0000 9.422 -3.2748 9.818 0 10.213 3.2972 8.237 2.0374 8.6 36 3.4337 9.034 0.0857 9.4 30 -3.1824 9.826 2.0405 10.228 3.4562 l
8.245 3.7887 8.644 3.4599 9.042 -0.0102 9.438 -3.3192 9.8 34 3.7874 10.229 3.4826 l 8.253 4.7677 8.652 3.2094 9.049 -0.3950 9.446 -3.7376 9.842 4.7758 10.237 3.2305 ! 8.261 5.1287 8.660 2.6498 9.057 -1.0577 9.454 -4.3396 9.849 5.1366 10.245 2.6672 , 8.269 4.9065 8.668 1.8734 9.065 -1.8722 9.462 -4.8960 9.857 4.9441 10.253 1.8857 ! 8.277 4.3490 8.676 1.0503 9.073 -2.6481 9.470 -5.1976 9.865 4.3557 10.260 1.0653 0.285 3.7456 8.684 0.3952 9.081 -3.2073 9.478 -4. 75 74 9.873 3.7584 10.268 0.3978 8.293 3.3264 8.692 0.0802 9.089 -3.4576 9.485 - 3.70 34 9.881 3.3315 10.276 0.0802
- 8. 301 3.1893 8.700 -0.0858 9.097 -3.4314 9.493 -2.0328 9.889 3.1942 10.284 -0.0843 8.309 3.2820 8.708 0.0000 9.105 -3.2735 9.501 0 9.897 3.2870 10.292 0.0000 8.317 3.4403 8.716 0.0858 9.113 -3.1888 9.509 2.0358 9.905 3.4456 10.300 0.0863 8.325 3.4665 8.724 -0.0102 9.121 -3.3I78 9.517 3.7089 ' 9.913 3.4719 10.308 -0.0803 8.333 3.2155 8.732 -0.3953 9.129 -3.7360 9.525 4.7641 9.928 3.2205 10.316 -0.3979 8.341 2.6549 8.740 -l.0584 9.137 -4.3378 9.533 5.1249 9.928 2.6589 10.324 -1.0654 8.349 I.8770 8.748 -I.8735 9.145 -4.8939 9.548 4.9028 9 .9 36 1.8799 10.332 -1.8858 8.357 1.0604 8.755 -2.6499 9.153 -5.1855 9.549 4.3457 9.944 1.0620 10.339 -2.6673 -
- 8. 365 0.3960 8.763 -3.2094 9.161 -4.7554 9.557 3.7428 9.952 0.3966 10.347 -3.2305 8.373 0.0102 8.778 -3.4599 9.169 -3.7020 9.565 3.3239 9.960 0.0102 10.355 -2 1824 8 . 381 -0.0859 8.779 -3. 4 3 3 F 9.174 -2.0319 9.573 3.1869 9.948 -0.0860 10.363 -3.4562 8.389 0.0000 8.787 -3.2757 9.884 0 9.580 3.2795 9.9 76 0.0000 10.378 -3.2972 8.397 0.0859 8.795 -3.1832 9.192 2.0330 9.588 3.4377 9.984 0.0860 10.379 -3.204l ps n 8.405 -0.0102 8.803 -3.3201 9.200 3.7037 9.596 3.4639 9.992 -0,0102 10.387 -3.3419 8 4p 8.483 -0.3960 8.811 -3.7386 9.208 4.7574 9.604 3.2831 10.000 -0.3966 10.395 -3.7638 e 8.421 -1.0604 8.819 -4.3407 9.216 5.1876 9.612 2.6528 10.008 -l.0621 10.403 -4.3692 ~
'a 8.429 -1.8778 8.827 -4.8972 0.224 4.8959 9.620 1.a756 10.085 -1.8800 10.411 -4.9294 **ES 8.437 -2.654 8.835 -5.1890 9.232 4.3396 9.628 8.0596 10.023 -2.6590 10.418 -5.8526 8.445 -3.2 3 r 4 8.843 -4.7566 9.240 3.7375 9. 6 36 0.3957 10.031 -3.2205 10.426 -4.7898 8.453 -3.4665 8.851 -3.7045 9.248 3.3892 9.644 0.0102 10.039 -3.4719 10.434 -3.7288 8 .4 68 -3.4402 8.859 -2.03 J 3 9.256 3.1824 9.652 -0.0859 10.047 -3.4455 10.442 -2.0467 8.469 -3.2819 e.867 0 9.264 3.2749 9.660 0.0000 10.055 -3.2870 10.450 0 8.477 -3.1893 8.875 2.0321 9.272 3.4328 9.667 0.0859 10.063 -3.1942 10.458 2.0547 8.485 -3.3264 8.883 3.7021 9.279 3.4590 9.675 -3.0102 10.078 -3.1315 10.466 3.7432 8.493 -3.7457 8.891 4.7554 9.287 3.2084 9.683 -0.3957 10.079 -3.7515 10.474 4.8082 8.5 01 -4.3490 8.899 5.1155 9.295 2.6491 9.691 -1.0596 10.087 -4.3557 10.482 5.8722 8.5 09 -4.9066 8.907 4.8939 9.303 8.8729 9.699 -1.8757 10.094 -4.9841 10.490 4.9481 8.517 -5.1287 8.915 4.3377 9.311 1.0581 9.707 -2.6529 10.;02 -5.1366 10.497 4.3858 8.525 -4.7677 8.922 3.7360 9.319 0.3951 9.715 -3.2131 10.810 -4.7750 10.505 3.7774 8.5 33 -3.7136 8.9 30 3.3178 9.327 0.0802 9.723 -3.4639 10.818 -3.7173 10.583 3.3546 8.5 41 ~2.0372 8.938 3.1881 9.335 -0.0857 9.731 -3.4376 10.126 -2.0403 10.521 3.2143 8.549 0 8.946 3.2735 9.343 0.0000 9.739 -3.2795 10.134 0 10.529 3.3098 8.557 2.0335 8.954 3.4314 9.351 0.0857 9.747 -3.1869 10.842 2.0468 10.537 3.4494 8.565 3.7046 8.962 3.4576 9. 359 -0.0102 9.755 -3. 3239 10.150 3.7290 10.545 3.4959 8.5 72 4.7587 8.970 3.2072 9.347 -0.3952 9.762 -3.1428 10.859 4.7899 10.553 3.2428 H.580 5.1190 8.978 2.6480 9.375 -1.0581 9.770 -4.3457 10.866 5.1526 10.568 2.6774 8.5&B 4.8972 8.986 l.0722 9.382 -l.8730 9.778 -4.9029 10.174 4.9293 20.569 1.8929 8.596 4.3407 8.994 1.0576 9.190 -2.6492 9.786 -5.8249 10.181 4.3692 10.576 1.0694 r 8.604 3.1185 9.002 0.3950 9. 39 8 -3.2006 9.794 -4.764l 10.189 3.7630 10.584 0.3993 c3 c) 8.612 3,32o3 0.0102 -3.7088 10.197 9.010 9.406 -3.4590 9.802 3.3418 10.592 0.0103 a w
8.620 3.1832 9.018 -0.0857 9.484 -3.4328 9.810 -2.0357 10.205 3.2041 10.400 -0.0866 03 e
TlWE PRESStmE 7tNE PRESSURE TlWE PRESSURE TIwE PRE SSURE (PSIDI TIME PRE SSURE TIME PRESSU4E ISEC) (PSlo) (SEC) (PSIDI I SEC) (PSIDI ( SFCI (PSIDI (SEcl (PSIDI 12.191 0.0000 12.588 -3.4249 ( SEC) 11.794 3. 3003 10.60H 0.0000 11.003 -1.3245 18.199 0. 11.802 3.5438 12.199 0.0891 12.596 -3.3232 0.0866 l1.011 ~3.2307 11.406 2.0858 12.207 -0.0104 12.604 -3.4711 10.616 11.414 3.7998 11.880 3.5702 82.612 10.624 -0.0103 11.019 -3.3696 11.818 3.3117 12.215 -0.4105 -3.9088
-3.1943 11.422 4.8809 -l.0992 12.620 -4.5384 10.632 -0.3994 II.027 5.2505 13.826 2.7342 12.223 10.640 -1.0694 11.035 -4.4055 II.430 18.834 8.9311 12.230 -l.9456 12.428 -5.1203
-1.8930 11.043 -4.9701 11.4 38 5.C230 12.238 -2.7519 12.636 -5. 3521 10.646 18.446 4.4522 II 842 1.0921 -4.975J 10.656 -2.6774 18.051 -5.1953 11.850 0.4078 12.246 -3.3330 12.644
-3.2428 11.059 -4.8206 11.454 3.8344 12.254 -3.5931 12.652 -3.8712 10.661 11.462 3.4054 11.858 0.0105 12.660 -2.1259 10.678 -3.4959 11.066 -3.7598 11.866 -0.0845 12.262 -3.5659 f II.074 -2.0636 18.470 3.2650 12.270 -3.4018 12.667 O.
10.679 -3.4694 0.0000 11.082 11.478 3.3599 l1.874 12.675 2.1482 10.687 -3.3098 0 11.484 3.5219 18.882 0.0815 12.278 -3.3058 3.9009 10.695 -3.2363 II.090 2.0742 -0.0105 12.286 -3.4479 82.683 11.098 3.7788 11.494 3.5488 13.889 12.691 5.0108 j 10.703 -3.3546 3.2919 11.897 -0.4079 12.294 -3.8825 10.711 -3.7774 11.104 4.8540 l1.500 12.302 -4.5078 12.699 5.3902 1 5.2215 11.509 2.7179 81.005 -1.0928 12.707 5.1567 10.719 -4.3859 11.114 11.913 -1.9332 12.310 -5.0857 10.727 -4.9482 II.122 4.9953 11.517 f.9216 12.318 -5.3160 12.715 4.5707 4276 18.525 1.0855 11.928 -2.7343 82.723 3.9364 l 10.735 -5.1722 11.1 30 11.929 -3.3187 12.326 -4.94l8
-4.8088 11.138 4'81 3 34 11.531 0.4054 12.334 -3.8411 12.731 3.4960 10.742 11.146 3.3865 11.541 0.0104 II.937 -3.5702 12.739 3.3519 10.750 -3.7431 11.549 -0.0880 18.945 -3.5431 12.342 -2.1116 3.4493 10.758 -2.0545 11.153 3.2470 II.953 -3.3801 12.350 0 12.747 WN 11.161 3.3413 11.557 0.0000 2.8261 12.755 3.6156 j 10.766 0 11.565 0.O n80 11.961 -3.2846 12.357 E Sf 10.774 2.0638 II.169 3.5025 -3.4259 12.365 3.8734 12.763 3.6432 10.782 3.7599 18.177 3.5292 18.573 11.581
-0.0104
-0.4054 II.969 II.977 - 3.85 77 12.373 4.9754 12.771 3.3795 *da llon 10.790 4.8296 11.185 3.2717 -4.4791 12.381 5.3521 12.779 2.7902 La 10.798 5.1953 II.193 2.7029 11.589 -1.0H56 II.985 12.389 5.1202 12.787 8.9727 88.596 -1.9214 II.992 -5.05 33 1.1144 30.806 4.9702 11.201 8.9109 12.000 -5.2824 12.397 4.5384 12.795 1.0795 ll.604 -2.7180 12.803 0.4162 10.814 4.4054 II.209
-3.2919 12.008 -4.9102 12.4DS 3.9088 10.821 3.7942 18.217 0.4011 11.612 12.413 3.4713 12.811 0.0107 3.3696 11.225 0.0804 II.620 -3.5488 12.016 -3.8226 3.3282 12.819 -0.0003 10.829 11.628 -3.5219 12.024 -2.0981 12.421 10.837 3.2307 11.232 -0.0875 12.032 12.429 3.4249 12.827 0.0003 3.3246 11.240 0.0000 II.636 -3. 3599 0.
3,5901 12.835 0.0901 10.845 -3.2650 12.040 2.1118 12.437 10.853 3.4849 18.248 0.0875 II.644 12.048 3.8473 12.445 3.6175 12.843 -0.0107 3.5815 11.256 -0.0104 18.652 -3.4054 12.453 3.3556 12.851 -0.4162 10.861 11.660 -3.8346 12.066 4.9458 10.869 3.2573 11.264 -0.4012 5. 31 60 12.461 2.7705 12.859 -1.1145 II.272 -l.0796 18. 668 -4.4523 12.064 12.867 -1.9728 10.877 2.6893 12.072 5.0857 12.469 1.95Ba 10.885 1.9014 l1.260 -l.9180 11.676 -5.0230 4.5078 12.477 1.8065 12.875 -2.7901
-2.7029 11.481 -5.2505 12.080 12.883 -3.3793 10.893 8.0741 II.288
-4.8809 82.088 3.8824 12.485 0.4132 0.4011 11.296 -3.2737 11.691 12.493 0.0104 12.890 -3.6432 10.900 II 499 -3.7997 12.096 3.4479 10.908 0.0103 11.304 -1.5292 12.104 3. 3058 12.501 -0.0897 12.898 -3.6856 11.312 -1.5025 II.707 -2.0854 0.0000 12.906 -3.4493 10.916 -0.0870 11.715 0, 12.lil 3.4018 12.509 12.914 -3.3589 10.924 0.0000 l1.389 -3.3413 2.0983 12.119 3.5459 12.516 0.0897 10.932 0.0870 11. 327 -3.2470 11.723 12.524 -0.0l07 12.922 -3.4960
- 11. 731 3.8227 12.827 3.5931 12.930 -3.9366 10.940 -0.0103 11.335 - 3. 18 66 4.9403 12.1 Mi 3.3329 12.532 -0.4133 10.948 -0.4012 11.343 -3.Hl34 18.739 12.540 -l.1066 12.938 -4.5707 11.747 5.2821 12.143 2.7518 12.944 -5.8567 10.956 -1.0742 I I . 351 -4.4277 12.151 1.9455 12.548 -l.0588 I I . 359 -4.9951 l1.755 5.0532 12.556 -2.7706 82.954 -5.3902 10.964 -1.9014 88.763 4.4790 12.359 1.0991 -5.0107 o 10.972 -2.6194 11.367 -5.2215 0.4104 12.564 -3.3556 12.942 r 10.979 -3.2573 11.375 -4.85 W 11.778 3.8576 12.167 12.175 0.0806 12.572 -3.6175 12.970 -3.9008 1 11.383 - 3. 7787 11.779 3.4258 S 10.987 -3.5115 3.2846 i 2. i B 3 -0.0895 i2.5 80 -3.5901 io.995 -3.4849 ii.39: -2.0740 ii.784 g;
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O O O TIME PRESSURE TIME PRESSURE TlWE PRES SURE TIME PRESSURE TIME PRE SSURE TI4E PRESSURE (SEC) (PSID) ( SEcl (PSIDI (SFC) (PSIDI (SEC) (PSin) (SEC) (PSID) ( SEC) (PSID) 12.986 0 13.385 3.5014 13.784 0.0000 14.187 - 3. 55 74 14.590 0 14.995 3.4474 12.994 2.1571 13.393 3.6701 11.794 0.0924 14.195 -3.4569 14.599 2.2452 15.003 3.8231 13.002 3.9298 13.401 3.6983 13.802 -0.0110 14.203 -3.6056 14.607 4.0903 15.011 3.8525 13.010 5.0479 13.409 3.4305 11.810 -0.4258 14.211 -4.0600 14.685 5.2540 15.020 3.5734 13.018 5.4301 13.417 2.8324 13.888 -1.1402 14.220 -4.7840 14.623 5.6568 15.028 2.9505 13.026 5.1948 13.425 2.0025 13.824 -2.0183 14.228 -5.3183 14.631 5.4069 15.036 2.0860 13.034 4.6045 13.433 1.1313 13.834 -2.8547 14.236 -5.5592 14.639 4.7925 15.044 1.1784 3 13.042 3.9457 13.441 0.4225 13.842 -3.4576 14.244 -5.1678 14.647 4.1277 15.052 0.4401 1 13.050 3.5218 13.449 0.0109 13.850 -3.7274 14.252 -4.02 38 14.655 3.4657 15.060 0.0113 13.058 3.3767 13.457 -0.0917 13.858 -3.6991 14.260 -2.2082 14.663 3.5146 15.068 -0.0955 13.066 3.4748 13.465 0.00 00 13.864 -3.5289 14.268 0 14.671 3.6167 15.076 0.0000 13.074 3.6424 13.473 0.0917 13.874 -3.4293 14.276 2.2265 14.679 3.1981 15.084 0.0955 13.082 3.6702 13.481 -0.0109 13.882 -3.5767 14.284 4.0563 14.687 3.8201 15.093 -0.0113 13.090 3.4045 13.489 -C.4225 13.890 -4.0276 14.292 5.2104 14.696 3.5435 15.101 -0.440s 13.096 2.8108 13.497 -1.1383 13.898 -4.4763 14.300 5.6049 14.704 2.9256 15.109 -l . l 785 13.106 1.9873 13.505 -2.0026 13.904 -5.2758 14.308 5.3521 14.712 2.0684 15.187 -2.0861 13.114 1.1227 13.513 -2.8324 13.914 -5.5147 14.314 4.7527 14.720 1.1685 15.125 -2.9505 13.122 0.4192 13.521 -3.4305 13.922 -5.1265 14.324 4.09 34 14.728 0.4364 15.133 -3.5736 13.130 0.0108 13.529 -3.6983 11.930 -3.9909 14.332 3.6152 14.736 0.0112 15.848 2.8525 13.138 -0.0910 13.537 -3.6703 13.938 -2.1905 14.340 3.4154 14.744 -0.0947 15.849 -3.8233 13.146 0.0000 13.545 -3.5014 13.946 0. 14.348 3.5967 14.752 0.0000 15.I58 -3.6474 E h* 13.154 0.0980 13.553 -3.4025 11.954 2.2084 14.157 3.7597 14.760 0.0947 15.864 -3.5444 4> , 13.162 -0.0108 13.561 -3.5488 13.962 4.0232 14.365 3.7884 14.768 -0.0112 15.874 -3.696d - La 13.170 -0.4193 13.569 -3.9961 11.970 5.1679 14.373 3.5141 14.776 -0.4364 15.182 -g.1628 N $; 13.178 -1.1227 13.577 -4.4398 13.978 5.5592 14.381 2.9013 14.785 -l.1684 15.190 -4.8331
, 13.186 -l.9874 13.585 -5.2346 13.984 5.3183 14.389 2.0583 14.793 -2.0685 15.198 -5.4529 13.194 -2.8109 13.593 -5.4716 13 994 4.7139 14.397 1.15 88 14.808 -2.9257 15.206 -5.6998 13.202 -3.4045 13.601 -5.0864 1:.002 4.0600 14.405 0.4327 14.809 -3.5435 15.214 -5.2985 13.210 -3.6702 13.609 -3.9597 14.010 3.6056 14.413 0.019! 14.887 -3.8208 15.222 -4.1249 13.218 -3.6424 13.417 -2.1734 14.018 3.4569 14.421 - 0.09 39 14.825 -3.7911 15.231 -2.2640 13.225 -3.4748 13.625 0, 14.027 3.5574 14.429 0.0000 14.833 -3.6167 15.'239 0 13.233 -3.3767 13.633 2.1907 14.035 3.7290 14.437 0.09 39 14.841 -3.5146 15.247 2.2837 13.241 -3.5219 13.641 3.9910 14.041 3.7574 14.445 -0.0112 14.849 -3.6657 15.255 4.1605 13.249 -3.9658 13.649 5.1265 14.051 3.4854 14.453 -0.4128 14.857 -4.1277 15.263 5.1442 13.257 -4.6046 13.457 5.5147 14.059 2.8777 14.468 -1.1589 14.845 -4.7926 15.271 5.7488 13.266 -5.1949 13.665 5.2758 14.067 2.0345 14.469 -2.0514 14.874 -5.4070 15.279 5.4997 13.273 -5.4301 13.673 4.6762 14.075 1.1493 14.477 -2.9044 14.882 -5.6518 15.288 4.8748 ,
13.281 -5.0478 13.681 4.0275 14.083 0.4292 14.486 - 3.5141 14.890 -5.2539 15.296 4.1985 l 13.289 -3.9297 13.689 3.5767 14.091 0.0110 14.494 -3.7884 14.898 -4.0901 15.304 3.7285 13.297 -2.1569 13.697 3.4291 14.099 -0.0931 14.502 -3.7596 14.906 -2.2450 15.382 3.5749 13.305 O. 13.706 3.5289 14.107 0.0000 14.510 -3.5866 14.914 0. 15.320 3.d787 13.383 2.1736 13.714 3.6992 14.185 0.0931 14.518 -3.4454 14.922 2.2642 15.328 3.8542 13.321 3.9599 13.722 3.7274 14.123 -0.01ll 14.524 -3.6353 14.930 4.1250 15.336 3.8856 13.329 5.0865 13.730 3.4575 14.131 -0.4293 14.534 -4.C934 14.938 5.2986 15.345 3.6041 13.337 5.4716 13.738 2.8544 14.839 -1.8494 14.542 -4.7528 14.946 5.6998 15.353 2.9758 13.345 5.2346 13.746 2.0183 14.147 -2.0346 14.550 -5.3421 14.955 5.4529 15.361 2.1039
$ 13.351 4.6397 13.754 l.1402 14.155 -2.8777 14.558 -5.6049 14.963 4.8332 15.369 1.1886 ?
H 13. 361 3.9961 13.762 0.4258 14.163 -3.4854 14.566 -5.2103 84.978 4.8627 15.377 0.4439 La m 1 3. 169 3.5408 13.770 0.0810 14.171 -3.7574 14.574 -4.0562 14.979 3.6968 15.385 0.0114 11.377 3.4025 13.778 -0.0924 14.179 -1.7290 14.582 -2.2263 14.987 3.5444 15.393 -0.0963
TiuE PRE SSURE TluE PRE SSU9E PRE SSURE TiuE PRE S5tlRE IPSIDI ( SECl (PSIDI TlWE PRE SSURE TIME ISEcl IPsini ISEcl 17.462 -3.8773 IlWE PRES 98RE (PSIDI ISECl IPSIDI 3.8094 17.044 0.0000 (SEC) IPSID) ( SEcl 16.220 0. 16.632 17.054 0.l006 17.470 -3.7679 15.402 0.0000 15.850 -3.7106 16.228 2.3441 16.640 3.93 32 17.062 -0.0820 17.479 -3.9799 15.480 0.0963 15.018 -3.6059 16.236 4.2705 16.648 4.0236 17.071 -0.4638 17.487 -4.4252
-0.08:4 15.826 -3.7609 5.4855 16.656 3.7323 17.495 -5.1380 15.418 -4.2350 16.244 3.0815 17.079 -1.2418 15.426 -0.4439 15.834 5.9009 16.645 -2,1981 17.504 -5.7967 15.842 -4.9178 16.252 2.1787 17.087 -6.0591 15.434 -1.1886 -5.5475 16.261 5.6452 16.673 17.095 -3.1089 17.512 15.442 -2.1040 15.858 5.0037 16.681 1.2308 17.520 -5.6326 15.859 -5.7987 16.249 0.4596 17.104 -3.7655 15.450 -2.9759 -5.3904 14.277 4.3095 16.689 17.182 -4.0593 17.529 -4.3849 15.450 -3.6043 15.867 16.285 3.8272 16.698 0.0118 17.120 -4.0286 17.537 -2.4068 15.467 -3.8856 15.875 -4.1964 14.294 3.6694 16.7o6 -0.0097 17.129 -3.8432 17.545 O.
15.475 -3.8562 15.883 -2.3033 16.302 3.7760 16.714 0.0000 17.554 2.4284 15.898 0 16.723 0.0997 17.137 -3.7347 17.562 4.4240 15.483 -3.6787 15.900 2.3237 16.310 3.9582
-0.0118 17.845 -3.8953 5.6827 15.491 -3.5749 4.2333 16.318 3.9884 16.731 17.154 -4.3862 17.570 15.499 -3.72tl6 15.908 16.327 3.6996 16.739 -0.4597 -5.0927 17.579 6.8830
-4.1945 15.916 5.4377 16.747 -l.2309 17.862 17.587 5.8482 15.507 5.8494 16.335 3.0545 17.170 -5.7456 15.516 -4.P,143 15.924 16.343 2.1596 16.756 -2,17 B8 17.179 -6.0058 17.595 5.1834 .
-5.4990 15.932 5.5960 1.2200 16.764 -3,0956 17.604 4.4645 15.524
-5.7488 15.941 4.9608 16.351 16.772 -3.7324 17.187 -5.5830 17.612 3.9648 15.5 32 4.2719 16.360 0.4556 17.195 -4.3463 15.540 -5.3441 15.949 16.368 0.0117 16.780 -4.02 37 11.621 3.8014
-4.8603 15.957 3.7938 -0.0988 16.789 - 3.99 31 17.203 -2.3856 17.429 3.9118 15.544 3.6374 16.376 17.282 0.
15.556 -2.2835 15.965 16.384 0.0000 16.797 -3.8094 17.220 2.4070 17.637 4.l005 p 32 0 15.973 3.7431 16.805 -3.7019 4.1388 m to 15.564 3.9231 16.392 0.0989 17.228 4.3851 17.646 4$LJ 15.573 2.3035 15.982 -0.0817 16.883 -3.8610 17.654 3.8326 15.990 3.9536 16.408 16.822 -4.3477 17.237 5.6327 3.1644 15.581 4.1966 3.6673 16.409 -0.4557 17.245 6.0591 17.662 "#ES 15.589 5.3905 15.998 16.830 -5.0480 17.678 2.2372 5.7987 16.006 3.0279 16.417 -I.2201 16.838 -5.6951 17.253 5.7966 I.2639 15.597 16,014 2.1407 16.425 -2.159' -5.95 30 17.262 5.1379 17.479 15.605 5.5474 16.434 -3.0540 16.847 17.270 4.4258 17.688 0.4720 15.683 4.9178 16.023 1.2094 16.855 -5.5339 17.696 0.012I 0.4516 16.442 -3.6997 17.278 3.9298 15.622 4.2349 16.038 16.450 - 3.9 Pd 4 16.863 -4.30 81 3.7679 17.704 -0.1024 3.7609 16.039 0.0816 16.878 -2.3646 17.287 0.0000 15.630 -0.0980 16.458 -3.fo82 17.295 3.8773 17.713 15.633 3.6059 16.047 -3.7760 16.880 0 17.721 0.1024 16.055 0.0000 16.467 16.888 2. 3d58 17.303 4.0644 15.646 3.7107 16.064 0.0980 16.475 -3.6694 4.3464 17.312 4.0054 17.729 -0.0122 15.654 3.8896 16.483 -3.8272 16.896 3.7989 17.738 -0.4720 16.072 -0.0116 16.905 5.58 38 17.320 15.663 3.9193 16.000 -0.4517 16.491 -4.3096 6.0058 17.328 3.8365 17.746 -l.2639 15.678 3.0356 16.088 -l.2094 16.500 -5.0038 16,913 5.7456 17.337 2.1175 17.755 -2.2373 15.679 3.0056 14.508 -5.6452 16.928 1.2527 17.763 -3.1644 15.687 2.8222 16.096 -2.1408 16.929 5.0927 17.345 -3.8327 16.516 -5.9009 17.353 'O.4678 17.771 15.695 1.1989 16.105 -3.0280 -5.4854 I 6.9 38 4.3862 17.780 -4.131S 16.813 -3.6674 16.524 3.8952 17.362 0.0120 15.703 0.4477 16.532 -4.2703 16.946 87.788 -4.100% 15.782 0.0815 16.121 -3.9536 -2.34 39 16.954 3.7347 17.370 -0.8085 17.796 -3.9118 16.129 -3.9236 16.541 3.8432 17.378 0.0000 15.720 -0.0971 14.549 0. 16.963 0.8085 17.805 -3.8014 15.728 0.0000 16.137 -3.7438 16.978 4.0286 17.387 -3.9648 16.146 -3.6 374 16.557 2.3648 17.395 -6.0128 17.813 15.736 0.0971 4.3082 16.979 4.0593 17.822 -4.4645
-0.0185 16.154 -3.79 38 16.565 16.988 3.7654 17.403 -0.4679 15.744 16.862 -4.2720 16.574 5.5340 3.1089 17.412 -l.2528 17.830 -5.1837 15.752 -0.4478 16.582 5.9530 16.996 17.838 -5.8482 15.761 -l.1990 16.170 -4.9601 17.004 2.1960 17.420 -2.2176 17.847 -6. l l 30 16.179 -5.5960 16.590 5.6951 1.2417 17.423 -3.1366 15.769 -2.8223 16.599 5.0479 17.012 -3.79n9 17.855 -5.6826 c3 15.777 -3.0017 16.187 -5.8494 4.3476 17.02l 0.4637 17.437 17.863 -4.4239 :
-3.6356 16.195 -5.4 176 15.607 17.029 0.0819 17.445 -4.0954 0'
$ 15.785 16.615 3.8610 17.454 -4.0643 r 15.793 -3.9193 16.203 -4.2331 16.623 3.7089 17.037 -0.1006 15.801 -3.B896 16.211 -2.3235 m O O e [t
O O O TIME PRESSURE TIME PRES SURE TIME PRESSURE TIME P RE SSURE TIME PRESSURE TluE PRESSURE ( SEC) (PSIO) (SEC) (PSID) (SEC) (PSID) (SEC) (PSID) ( SEC) (PSID) ( SEC) (PSIDI 17.880 0 18.301 3.9886 18.725 0.0000 19.150 -4.0524 19.579 0 20.010 4.1599 17.889 2.4500 18.310 4.1737 18.733 0.1052 19.159 -3.9380 19.597 2.5601 20.089 4.3606 17.897 4.4634 18.388 4.2055 18.742 -0.0125 19.167 -4.1073 19.596 4.6639 20.027 4.3939 17.905 5.7332 18.327 3.9010 18.750 -0.4847 19.876 -4.6250 19.605 5.9909 20.036 4.0757 17.914 6.1673 18.335 3.2208 18.758 -1.2979 19.184 -5. 3499 19.613 6.4445 20.045 3.3658 17.922 5.9001 18.343 2.2778 18.767 -2.2974 19.193 -6.0584 19.622 6.1653 20.053 2.3791 17.931 5.2297 , 18.352 1.2864 18.775 -3.2494 19.202 -6.3327 19.630 5.4647 20.062 1.3440 17.939 4.5041 18.360 0.4804 18.784 -3.9356 19.210 -5.8869 19.639 4.7066 20.074 0.5019 17.947 4.0000 18.369 0.0824 88.792 -4.2428 19.219 -4.5829 19.648 4.8798 20.079 0.0129 17.956 3.8351 18.377 -0.1042 18.801 -4.2104 19.227 -2.5154 19.656 4.0075 20.088 -0.1089 17.964 3.9466 18.386 0.0000 18.809 -4.0169 19.236 0 19.665 4.1239 20.097 0.0000 17.973 4.8369 18.394 0.1042 18.888 -3.9035 19.244 2.5378 19.674 4.3228 20.805 0.1089 17.981 4.1685 18.403 -0.0124 18.824 -4.0713 19.253 4.6234 19.682 4.3558 20.814 -0.0129 17.990 3.8667 18.411 -0.4805 18.835 -4.5845 19.261 5.9388 - 19.698 4.0405 20.123 -0.5020 17.998 3.1925 18.419 -1.2865 18.843 -5.3229 19.270 6.3985 19.699 3.3359 20.131 -1.3441 18.006 2.2571 18.428 -2.2772 18.852 -6.0053 19.279 6.8117 19.708 2.3585 20.140 -2.3792 18.015 1.2751 18.436 -3.2209 18.860 -6.2772 19.287 5.4172 19.717 1.3324 20.l49 -3.3652 18.023 0.4762 18.445 -3.9011 18.869 -5.8353 19.296 4.6656 19.725 0.4976 20.157 -4.0758 18.0 32 0.0122 18.453 -4.2055 18.877 -4.5427 19.304 4.1434 19.734 0.0128 20.166 -4.3939 18.040 -0.1033 18.462 -4.1736 18.884 -2.4934 19,313 3.9726 19.742 -0.1080 20.175 -4.3606
, 18.048 0.0000 18.470 -3.9816 18.894 0 19.321 4.0981 19.751 0.0000 20.183 -4.8599 18.057 0.1033 18.479 -3.8692 18.903 2.5857 19.330 4.2853 19.760 0.1080 20.892 -4.0425 ,n 18.065 -0.0123 18.487 -4.0356 18.911 4.5830 19.339 4.3180 19.768 -0.0128 20.201 -4.2463 mN 18.074 -0.4762 18.496 -4.5442 18.920 5.8870 19.347 4.0053 19.777 -0.4976 20.209 -4.7477 47, 18.082 -l.2752 18.504 -5.2762 18.928 6.3327 19.356 3.3049 19.786 -1.3325 20.218 -5.5125 u 18.090 -2.2572 18.512 -5.9526 18.937 6.0583 19.364 2.3380 19.794 -2.3586 20.227 -6.2198 b3((
18.099 -3.1926 18.521 -6.2221 18.946 5.3699 19.373 I.3208 19.803 -3.3360 20.235 .-6.5007 18.107 -3.8667 18.529 -5.7841 18.954 4.6249 19.388 0.4932 19.811 -4.0405 20.244 -6.0431 18.116 -4.1685 18.538 -4.5028 18.963 4.1073 19.390 0.0127 19.820 -4.3558 20.253 -4.7045 18.124 -4.1369 18.546 -2.4715 18.971 3.9380 19.399 -0.1070 19.829 -4.3228 20.261 -2.5822 18.133 -3.9466 18.555 O. 18.980 4.0524 19.407 0.0000 19.837 -4.1239 20.270 0 18.148 -3.8351 18.563 2.4936 18.988 4.2479 19.486 0.8070 19.846 -4.0075 20.279 2.6048 18.149 -4.0000 18.572 4.5429 18.997 4.2803 19.424 -0.0127 19,855 -4.8798 20.288 4.7455 18.158 -4.5042 18.580 5.8354 19.005 3.9704 19.433 -0.49 33 19.863 -4.7066 20.296 6.0957 18.166 -5.2297 18.509 6.2772 19.084 3.2781 19.448 -1. 3209 19.872 -5.4647 20.305 6.5572 18.175 -5.9002 18.597 6.0053 19.022 2.3176 19.450 -2.3381 19.880 -6.1653 20.314 6.2731 18.183 -6.1473 18.606 5.3228 19.031 1.3093 19.459 -3.3070 19.889 -6.4445 20.322 5.5602 18.198 -5.7332 18.614 4.5844 19.039 0.4889 19.467 -4.0054 19.898 -5.9908 20.331 4.7889 18.200 -4.4632 18.623 4.0713 19.048 0.0126 19.476 -4.3180 19.906 -4.6638 20.340 4.2529 18.208 -2.4498 18.631 3.9035 19.054 -0.106l 19.484 -4.2852 19.915 -2.5598 20.349 4.0776 18.217 0 18.640 4.0869 19.065 0.0000 19.493 -4.0481 19.924 0 20.357 4.1968 18.225 2.4717 18.648 4.2106 19.074 0.1061 19.502 -3.9726 19.932 2.5824 20. 366 4.3984 16.234 4.5030 18.657 4.2428 19.082 -0.0126 19.510 -4.1435 19.941 4.7046 20.375 4.4320 18.242 5.7841 18.665 3.9356 19.091 -0.4890 19.519 -4.6657 19.950 6.0432 20.383 4.1111 18.250 6.2221 18.674 3.2494 19.099 -1.3094 19.527 -5.4172 19.958 6.5 008 20.392 3.3943 18.259 5.9525 18.682 2.2973 19.108 -2.3177 19.536 -6.1187 19.947 6.2891 20.40I 2.3998 18.267 5.2768 18.691 1.2978 19.116 -3.2782 19.544 -6. 3985 19.976 5.5124 20.410 1.3557 s 18.276 4.5441 18.699 0.4847 19.125 -3.9704 19.553 -5.9387 19.984 4.7476 20.418 0.5063 l $ 18.284 18.293 4.0355 18.708 0.0125 19.l33 -4.2603 19.562 -4.62 32 19.993 4.2862 20.427 0.01 1) ? os 3.0692 18.786 -0.I052 19.142 -4.2478 19.570 -2.5376 20.001 4.0425 20.436 -0.1098 'd 5 i l l
TIME PRESSURE TINE PRE SSURE TlWE PHESSURE TIME PRESSHRE Tl4E PRESSURE (SEC) IPSID) (SEC) (PSIO) TIME PRESSURE (SEC) (PSIDI (SEcl (PSin) 22.662 -4.4844 (SEC) (PSID) ( SEC) (PSID) 4.3484 22.212 0.0000 20.881 -4.2323 21.321 0 28.765 0.1146 22.671 -4.2898 20.444 0.0000 2.6725 21.774 4.5508 22.223 20.453 0.1098 20.890 ~4.1128 21.330 28.703 4.5856 22.230 -0.0134 22.480 -4.4741
-0.0838 20.899 -4.2896 28.339 4.8687 22.239 -0.5283 22.689 -5. 0 3H2 20.462 21.348 6.2519 28.792 4.25 16 22.698 -5.8497 20.479 -0.5063 20.908 -4.8303 6.7275 21.801 3.5119 22.248 -l.4146 20.479 -1.3558 20.916 -5.6081 21.357 2.4829 22.256 -2.5039 22.707 -6.5997 20.925 -6.3271 21.366 6.4360 28.810 22.716 -6.8905 20.468 -2.3999 21.375 5.7046 21.818 1.4027 22.265 -3.5415 22.725 -6.412H 20.497 -3.3944 20.934 - 6.68 38 4.9832 28.827 0.5238 22.274 ~4.2804
-4.9923 20.505 -4.1112 20.943 -6.1482 21.381 0.08 35 22.283 -4.6241 22.734 20.952 -4.7H63 21.392 4.3633 23.836 -4.5898 22.743 -2.7402 20.514 -4.4320 4.1034 21.845 -0.1837 22.292 0
-4.3984 20.960 -2.6278 21.401 22.301 -4.3779 22.752 20.523 21.410 4.3050 21.854 0.0 000 22.761 2.7638 1 20.5 31 -4.1960 20.969 0.
28.410 4.5826 21.863 0.1837 22.310 -4.2543 5.0314 ! 20.540 -4.0776 20.978 2.6499 -0.01 35 22.319 -4.4372 22.770 20.987 4.8276 28.428 4.5471 28.872 22.779 6.4660 20.549 -4.2529 21.437 4.2179 21.881 -0.5239 22.328 -4.9965 6.9556
-4.7889 20.995 6.2018 22.337 -5.8013 22.788 20.558 21.004 6.6705 21.445 3.4824 21.890 -l.4028 22.346 -6.5451 22.797 6.6542 20.566 -5.5603 28.454 2.4628 28.899 -2.4830 5.8988 20.575 -6.2731 21.013 6.3816 21.908 -3.5120 22.355 -6.8414 22.806
-6.5572 21.022 5.6564 21.463 1.3909 22.364 -6.3598 22.885 5.0798 20.584 4.87I7 21.472 0.5894 21.916 -4.2536 22.824 4.5112
* '20.592 -6.0956 21.'0 38 21.481 0.0134 28.925 -4.5856 22.373 -4.9510 4.3253
-4.7453 21.040 4.3264 22.382 -2.7175 22.834 20.601 21.048 4.1481 21.490 -0.1127 21.934 -4.5508 22.843 4.4510 20.610 -2.6046 21.499 0.0000 28.943 -4. 3414 22.391 0.
4.6657 20.619 0. 21.057 4.2686 28.952 -4.2189 22.400 2.7404 22.852 ,u 2.6273 21.066 4.4745 28.507 0.8127 22.409 4.9925 22.861 4.7013 a 20.627 4.5086 21.516 -0.0834 21.968 -4.4003 22.870 4.3609 4 p 20.636 4.7865 21.075 21.970 -4.9549 22.418 6.4129 20.645 6.1483 21.084 4.1822 28.525 -0.5195 22.427 6.8985 22.879 3.6005 La 21.092 3.4530 21.534 -1.3930 21.979 - 5. 75 30 22.888 2.5456 " 8l 20.654 6.6138 21.988 -6.4905 22.436 6.5996 20.662 6.3273 21.101 2.4413 28.543 -2.4622 -6.7844 22.445 5.8496 22.897 f.4381 28.110 8.3791 28.552 -3.4825 21.997 22.906 0.5370 20.671 5,6083
-4.2179 22.006 -6.3068 22.454 5.0388 20.680 4.8302 21.189 0.5850 28.561 22.463 4.4742 22.915 0.01 36 21.128 0.0132 21.570 -4.5478 22.015 -4.9098 22.924 -0.1865 20.689 4.2896 22.023 -2.6949 22.472 4.2898 20.697 4.8128 28.136 -0.8817 28.578 -4.5826 0 22.481 4.4145 22.933 0.0000 0.0000 21.587 -4.3050 22.032 22.942 0,1165 20.706 4.2323 21.145 22.04l 2.7877 22.490 4.6274 20.715 4.4364 21.154 0.1887 28.596 -4.1834 4.9512 22.499 4.6627 22.952 -0.013a 28.163 -0.0133 21.605 -4.3633 22.050 22.961 -0.5378 20.724 4.4703 22.059 6.35 99 22.506 4.3251 20.733 4.1466 28.172 -0.5151 21.614 -4.9833 6.8414 22.587 3.5710 22.970 -l.4382 21.180 -I.3792 21.623 -5.7047 22.048 22.979 -2.5457 20.748 3.4236 22.077 6.5450 22.526 2.5247 20,750 2.4205 28.189 -2,4414 21.632 -6.4360 5.8013 22.535 1.4263 22.988 -3.6006 28.198 -3.4538 21.640 -6.7274 22.086 22.997 -4.3609 l 20.759 f.3674 21.649 -6.2538 22.095 4.9964 22.54A a 0.5326 20.768 0.5106 28.207 -4.1823 22.104 4.4372 22.553 0.0137 23.006 -4.7013 20.276 0.0131 21.216 -4.5087 21.658 -4.8685 22.562 -0.1856 23.015 -4.6657 28.667 -2.6722 22.183 4.2543 0.0000 , 23.024 -4.4510 21.225 -4.4745 20.785 -0.8108 21.676 0 22.322 4. 3779 22.578 23.033 -4.3253 l 20.794 0.0000 21.233 -4.2686 2.6958 22.131 4.5894 22.580 0.l156 20.803 0.1108 21.242 -4.1481 28.685 22.589 -0.0137 23.042 -4.581J l
-4.3264 21.694 4.9099 22.140 4.624l 23.058 -5.0799 20.811 -0.0132 21.251 22.149 4.2893 22.598 -0.5327 21.260 -4.8787 21.703 6.3069 23.060 -5.8981 20.820 -0.5107 6.7844 22.858 3.5484 22.607 -1.4264 20.829 -l.3675 28.269 -5.6565 28.718 22.867 2.5038 22.687 -2.5248 23.070 -6.6543 l
20,838 -2.4206 21.277 -6.3816 28.720 6.4905 22.626 -3.5750 23.079 -6.9556 21.286 -6;6706 28.729 5.7529 22.176 8.4145 23.088 -6.4659 20.846 -3.4237 21.738 4.9548 22.185 0.5282 22.635 -4.3252 5.0336 c3 20.855 -4.1467 21.295 -6.2010 0.0136 22.644 -4.6627 23.097 " 28.747 4.4002 22.194
$s 20.864'-4.4703 28.304 -4.8274 21.756 4.2809 22.203 -0.1146 22.653 -4.6274 20.873 -4.4364 28.383 -2.6496 g
0 O e-.,
O O .O TIME PRESSURE TIME PRESSURE TIME PRESSURE TIME PRESSURE TIME PRESSURE TIME PRESSURE (SEC) (PSID) ( SEcl IPSIO) (SEC) (PSID) (SEC) (PSIO) ( SEC) (PSIDI ( SEC) (PSIO) 23.185 O. 23.572 4.5240 24.033 0.0000 24.498 -4.5970 24.967 0. 25.440 4.7060 23.124 2.7858 23.582 4.7422 24.043 0.1194 24.507 -4.4672 24.976 2.8989 25.449 4.9330 23.133 5.0758 23.598 4.7784 24.052 -0.0142 24.517 -4.6592 24.985 5.2812 25.459 4.9707 23.142 6.5191 23.600 4.4 325 24.061 -0.5503 24.526 -5.2465 24.995 6.7838 25.468 4.6108 23.151 7.0127 23.609 3.6596 24.070 -1.4735 24.535 -6.0916 25.004 7.2974 25.478 3.8068 23.161 6.7088 23.618 2.5874 24.080 -2.6083 24.545 -6.8725 25.014 6.9883 25.487 2.6915 23.870 5.9465 23.627 1.4617 24.089 -3.6892 24.554 -7.1837 25.023 6.1879 25.497 1.5205
. 23.879 5.1215 23.437 0.5458 24.098 -4.4683 24.563 -6.6780 25.033 5.3295 25.506 0.5678 23.188 4.5483 23.646 0.0140 24.107 -4.8170 24.573 -5.1987 25.042 4.7330 25.516 0.0146 23.197 4.3608 23.655 -0.1184 24.117 -4.7805 24.582 -2.8515 25.052 4.5379 25,525 -0.8232 23.206 4.4875 23.664 0.0000 24.126 -4.5605 24.591 0. 25.061 4.6697 25.535 0.0000 23.215 4.7040 23.673 0.ll84 24.135 -4.4317 24.601 2.8763 25.070 4.8950 25.544 0.1232 23.225 4.7399 23.683 -0.0141 24.844 -4.6223 24.610 5.2408 25.000 4.9323 25.554 -0.0146 23.234 4.3967 23.692 -0.5459 24.154 -5.2049 24.619 6.7380 25.089 4.5752 25.563 -0.5679 23.243 3.6301 23.701 -l.4418 24.163 -6.0433 24.629 7.2406 25.099 3.7775 25.573 -1.5204 23.252 2.5665 23.780 -2.5875 24.172 -6.8180 24.638 6.9269 25.108 2.6707 25.582 -2.6914 23.261 1.4499 23.719 -3.6597 24.181 -7.1267 24.648 6.8398 25.118 8.5087 25.592 -3.8069
- 23.270 0.5414 23.729 -4.432S 24.191 -6.6250 24.657 5.2800 25.127 0.5634 25.602 -4.5108 23.279 0.0139 23.738 -4.7784 24.200 -5.1575 24.666 4.6961 25.137 0.0145 25.618 -4.9707 23.289 -0.1175 23.747 -4.7422 24.209 -2.8308 24.676 4.5026 25.146 -0.1222 25.628 -4.9330 23.293 0.0000 23.756 -4.5240 24.218 0 24.685 4.6334 25.856 0.0000 25.630 -4.7060 23.307 0.1175 23.765 -4.3963 24.228 2.8537 24.694 4.8569 25.865 0.1222 25.640 -4.5732 E" 23.316 -0.0140 23.775 -4.5853 24.237 5.1989 24.704 4.89 39 25.874 -0.0145 25.649 -4.7698 23.325 -0.5415 23.784 -5.1632 24.246 6.6781 24.713 4.5396 25.884 -0.5635 25.659 -5.3710 -
4hk c2 23.334 -1.4500 23,793 -5.9949 24.256 7.1837 24.723 3.7481 25.193 -l.3088 25.668 -6.2368 23.343 -2.5666 23.802 -6.7635 24.265 6.8725 24.732 2.6499 N 8l 23.352 25.203 -2.6708 25.678 -7.0356 ! -3.6301 23.811 -7.0697 24.274 6.0915 24.741 1.4970 25.212 -3.7776 25.687 -7.354I { 23. 362 -4.3967 23.821 -6.5720 24.284 5.2464 24.758 0.5590 25.222 -4.5753 25.697 -6.8364 l 23.371 -4.7399 23.830 -5.1162 24.293 4.6592 24.760 0.0144 25.233 -4.9323 25.706 -5.3221 l 23.380 -4.7039 23.839 -2.8082 24.302 4.4672 24.769 41213 25.241 -4.8950 25.716 -2.9212 23.389 -4.4875 23.848 0 24.312 4.5970 24.779 0.0000 25.250 -4.6697 25.725 O. 23.398 -4.3608 23. H57 2.8 311 24.321 4.8187 24.788 0.1283 25.259 -4.5379 25.735 2.9439 23.407 -4.5403 23.867 5.1577 24.330 4.8555 24.798 -0.0144 25.269 -4.7330 25.744 5.3632 23.416 -5.1216 23.876 6.6251 24.340 4.5039 24.807 -0.5598 25.278 -5.3295 25.754 6.8891 23.426 -5.9465 23.885 7.1268 24.349 3.7186 24.836 -1.4971 25.288 -6.1880 25.764 7.4107 23.435 -6.7089 23.894 6.8100 24.358 2.6298 24.826 -2.6500 25.297 -6.9813 25.773 7.0897 23.444 -7.0127 23.904 6.0432 24.368 8.4852 24.835 -3.7482 25. 307 -7.2974 25.783 6.2840 [ 23.453 -6.5190 23.983 5.2048 24.377 0.5546 24.845 -4.5397 25.316 -6.7837 25.792 5.4122 l 23.462 -5.0749 23.922 4.6223 24.386 0.0143 24.854 -4.89 39 25.326 -5.2810 25.802 4.8064 l 23.471 -2.7855 23.931 4.4347 24.395 -0.l203 24.863 -4.8569 25.335 -2.8986 25.812 4.6083 23.480 0 23.941 4.5605 24.405 0.0000 24.873 -4.6334 25.345 0 25.828 4.7422 23.490 2.8084 23.950 4.7805 24.414 0.1203 24.882 -4.5026 25.354 2.9214 25.831 4.9710 23.499 5.1164 23.959 4.8170 24.423 -0.0143 24.891 -4.6962 25.364 5.3223 - 25.840 5.0089 23.508 6.5721 23.969 4.4682 24.433 -0.5547 24.901 -5.2881 25.373 6.8365 25.850 4.6463 23.517 7.0697 23.978 3.6891 24.442 -l.4853 24.910 -6.1398 25.383 7.3541 26.859 3.8 361 23.526 6.7634 23.987 2.6082 24.451 -2.6292 24.920 -6.9270 25.392 7.0355 25.869 2.7122
- p. 23.536 5.9948 23.996 8.4734 24.461 -3.7187 24.929 -7.2406 25.402 6.2360 25.879 , 1.5322 o 2J.545 5.1632 24.006 0.5502 24.470 24.938 -6.7309 25.488 o
-4.5040 5.3709 25.888 0.5722
(( 23.554 4.5853 24.015 0.0142 24.479 -4.8555 24.948 -5.2399 25.421 4.7697 25.898 0.0147 4) gj 23.563 4.3963 24.024 -0.1194 24.489 -4.8187 24.957 -2.8761 25.430 4.5732 25.907 -0.1241
PRESSURE p TIME PRESSURE TIME PRESSURE TIME PRE SSURE (SEC) (PSIOl TIME PRESSURE TIME (PSID) (SEC) (PSID) TIME PRE SSURE (SEC) (PSII)) (SEC) 0.0000 28.373 -4.9575 (SEC) (PSIDI ISEC) (PSIDI 27.376 4. fH 62 27.872 25.917 0.0000 26.399 -4.7784 26.884 0. 5.8289 27.882 0.1288 28.383 -4.8115
-4.6435 26.894 3.0110 27.386 -0.0153 28.393 -5.0247 25.927 0.l241 26.408 27.396- 5.1610 27.892
-0.0147 26.418 -4.8431 26.904 5.4855 27.902 -0.5939 28.403 -5.6580 25.936 26.914 7.0462 27.405 4.7873 28.413 -6.5693 25.946 -0.5722 26.428 -5.4535 27.485 3.9526 27.912 -3.5903 25.955 .-l.5323 26.437 -6.3320 26.924 7.5797 2.7045 27.922 -2.8150 28.423 -7.4115
-7.1437 26.934 7.2583 27.425 -3.9816 - 29.433 -7.747I 25.965 -2.7323 26.447 27.435 1.5787 27.9 32 25.975 -3.8362 26.457 -7.4672 26.943 6.4273 0.5895 27.941 -4.8223 28.443 -7.2017 26.953 5.5356 27.445 -5.198F* 28.453 -5.6065 25.984 -4.6463 24.466 -6.9415 27.455 0.0152 27.951 25.994 -5.0089 26.476 -5.4039 26.961 4.9160
.27.465 -0.1279 27.961 -5.1593 28.463 -3.0773 0
26.486 -2.9661 26.973 4.7134 27.971 -4.9219 28.473 26.003 -4.9710 26.983 4.8504 27.475 0.0000 28.483 3.0995 26.013 -4.7422 26.495 O. 27.485 0.1279 27.981 -4.7829 26.505 2.9881 26.992 5.0843 27.991 -4.9886 28.493 5.6468 26.022 -4.6083 27.002 5.1238 27.494 -0.0152 28.503 7.2533l 26.032 -4.8065 26.515 5.4448 27.504 -0.5896 28.001 -5.6873 26.524 6.9940 27.012 4.7522 28.011 -6.5222 28.514 7.8025 26.042 -5.4123 24.534 7.5235 27.022 3.9236 27.514 -l.5788 28.021 -7.3583 28.524 7.4645
- 26.051 -6.2848 27.032 2.7740 27.524 -2.7946 28.534 6.6162 26.061 -7.0897 26.544 7.1976 27.534 -3.9527 28.031 -1.6915
-7.4107 26.554 6.3796 27.041 1.5671 28.048 ~7.8500 28.544 5.6984 26.070 26.563 5.4944 27.051 0.5852 27.544 -4.1373 -5.5662 28.554 5.0606 26.080 -6.8890 0.0158 27.554 -5.1410 28.051 26.573 4.8796 27.061 28.061 -3.0552- 28.564 4.8520 26.090 -5.3630 24.583 4.6785 27.071 -0.l270 27.564 -5.l289 28.574 4.9930 g4 "N 26.099 -2.9436 0.0000 27.573 -4.8862 '28.078 0 26.592 4.8144 27.081 28.081 3.0775 28.584 5.2338 26.809 0, 26.602 5.0466 27.090 0.1270 27.583 -4.7492 5.6067 28.595 5.2737 -
24.118 2.9463 27.593 -4.9524 28.091 wm 26.128 5.404) 26.682 5.0852 27.100 -0.0158 27.603 -5.5766 28.108 7.2018 28.605 4.8919
- 26.13tl 6.9486 26.622 4.7170 - 27.l10 -0.5853 27.683 -4.4748 28.lli 7.7478 28.615 4.0389 26.147 7.4672 26.631 3.8945 27.120 -1.5672 27.423 -7.3049 28.125 7.4l15 28.625 2.8556 7.8437 26.641 2.7534 27.130 -2.7741 28.l 38 6.5693 28.635 1.6132 26.157 26.651 1.5555 27.140 -3.9237 27.433 -7.6357 28.141 5.6579 28.645 0.6024 26.167 6.3319 27.149 -4.7522 27.643 -7.0986 26.176 5.4535 26.666 0.5809 23.151 5.0246 28,655 0.0155 26.470 0.0149 27.159 -5.8231 27.653 -5.5258 4.8875 28,664 -0.1307 26.886 4.8431 27.662 -3.0330 28.162 26.196 4.6435 26.680 -0.1260 27.169 -5.0843 28.872 4.9575 28.676 0.0000 0.0000 27.179 -4.8503 27.472 0.
28.686 0.1307 26.205' 4.7784 26.490 27.682 3.0554 28.182 5.8966 26.215 5.0089 26.700 0.1260 27.189 -4.7134 5.5664 28.192 5.2363 28.696 -0.0155 26.709 -0.0150 27.198 -4.9161 27.692 28.706 -0.6025 26.225 5.0475 27.702 7.1508 28.202 4.8572 26.234 4.6887 26.719 -0.5810 27.208 -5.5357 7.6915 28.212 4.0802 28.786 -l.6133 26.729 -1.5556 27.218 -6.4274 27.712 t28.726 -2.8557 26.244 3.8653 27.722 7.3582 28.222 2.8353 26.254 2.7328 26.738 -2.7536 27.228 -7.2513 28.232 1.6017 '28.737 -4.0390 27.238 -7.5797 27.732 6.5221 26.263 1.5438 26.748 -3.8946 27.742 5.4173 28.242 0.5988 28.747 -4.8920'* 26.273 0.5765 26.758 v4.7170 27.247 -7.0461 4.9885 28.252 0.0154 28.757 -5.2737 26.768 -5.0852 27.257 -5.4853 27.752 28.767 -5.2338 26.283 0.0148 27.267 -3.0108 27.762 4.7829 28.262 -0.1298 26.292 -0.1251 26.777 -5.0466 27.772 4.9219 28.272 0.0000 28.777 -4.9929. 26.302 0.0000 26.787 -4.8844 27.277 0 5 l593 28.282 0.1298 28.787 -4.8520 26.797 -4.6785 27.287 3.0333 27.782 28.797 -5.0606 26.312 0.125l 27.297 5.5260 27.792 5.1987 28.292 -0.0154 26.321 -0.0849 26.807 -4.8796 27.802 4.8223 28.302 -0.5982 29.807 -5.6984 26.331 -0.5766 26.886 -5.4947 27.307 7.0982 27.812 3.9815 28.382 -l.6018 28.818 -6.6863, 24.826 -6.3797 27.316 7.6357 28.322 -2.8354 28.828 -7.4645 26.34l -l.5439 27.326 7.3049 27.822 2.8149 0 26.350 -2.7329 26.836 -7.1976 28.332 -4.0103 28.838 -7.8025 26.846 -7.5235 27.336 6.4748 27.832 1.5902 28.342 -4.8572 28.848 -7.2532 /O y 26.360 -3.8654 26.370 -4.6887 26.855 -6.9939 27.346 5.5765 27.842 0.5938 28.352 -5.2363 28.858 ~5.6466 r.
-5.4446 27.356 4.9523 27.852 0.0353 -3.0993 26.379 -5.0478 26.865 27.862 -0.l288 28.363 -5.1966 28.868
']o 26.389 -5.0008 24.875 -2.9884 27.366 4.7482
., g
o O O TIME PRESSURE TIME PRESSURE TI4E
- PRESSURE ISECl IPSID) (SEC) IPSID) (SEcl (PSID)
!' 28.878 0 29.390 5.0615 29.908 0.0000 28.889 3.1215 29.401 5. 30 77 29.918 0.8335 l 28.899 5.6867 29.411 5.3482 29.929 -0.0159 ! 28.909 7.3047 29.421 4.9610 29.939 -0.6152 28.919 7.8577 29.432 4.0960 29.950 -1.6474 28.930 7.5173 29.442 -2.8959 29.960 -2.9860 28.940 d.6638 29.452 3.6359 29.970 -4.1244 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*8863 20.483 -0.8126 28.981 5.0283 29.494 0.0000 28.993 L.2708 29.504 0.1326 29.001 5.3111 29.514 -0.0157 29.018 -4.9265 29.525 -0.6180 29.022 4.0675 29.535 -1. 3 36l 29.032 2.8758 29.545 -2.8960 29.042 . 1.6246 29.555 -4.0968 29.052 0.6067 29.566 -4.9610 29.062 0.0156 29.574 -5.3482 29.073 -0.1316' 29.58d -5.3077 29.083 0.0000 ,'29.597 -5.0635 F ** 29.093 0.1316 29.607 -4.9205 <$ 29.103 -0.0156 29.617 -5.1321
- f; 29.114 -0. 00 #J 29.o/u -3.7tuv b2 cn 29.124 -1,6247 '"
29.638 -d.7098 29.134 -2.8759 '29.648 ~7.5 700 29.144 ~460676 29.659 -7.9827 29.154 -4.9266 29.469 -7.3557 29.165 -5.3831 29.679 -d.7263 29.175 -5.2708 29.690 -3.8430 29.185 -5.0283 29.700 0
'29.195 ~4.8863 29.18Q 3.1653 29.206 -5.0964 .29.72r 5.7668 29.216 -5.7388 29.731 ~7.4067 29.226 -6.6638 29.748 '7.9675 29.236 -7.5174 29.752 746223 29.246 -7.8577 29.762 6.756l 29.257 -7.3046 '29.773 5.8188 29.267 -S.4865 29.783 5.1675 29.277 -3. l yl2 29.793 ~4.9546' 29.287 0 29.804 5.0985 29.298 '3.1433 29.814 5.3444 29.308 5.7265 29.825 5.3852 29.318 7.3558 29.835 4.9953 29.329 7.9127 29.844 "4.1243
(; 29.339 ~ 7.5699 29.856 '2.9359 e 29.349 667097 29.866 ~l.6473 29.360 5.7788 29.877 '0.6852
- 0) 29.370 5.8 320 oo 29.887 0.0158 29.380 4.9205 29.898 -0.1335 f g;
4 9 _. _.-_ _____ -}}