ML17138B440

to Design Assessment Rept,Including Updated Pages,Figures & Tables to Vol 1,nonproprietary Version

ML17138B440
Person / Time
Site:	Susquehanna
Issue date:	08/21/1980
From:	PENNSYLVANIA POWER & LIGHT CO.
To:
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ML17138B441	List: ML17138B440 ML18030A008
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NUDOCS 8008250416
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EL 704'4" V ~ 0I j6 l SRV LINE DOWNCOMER COLUMN EL 672'-0" HIGH WATER BRACING EL 668'-0" E L 648'-0" p

IJ ~

b Rev. 3, 7/80 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT SUPPRESSION CHAMBER-SECTION VIEW s008ps'orb;  ! FIGURE 1-3

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165o 165o 180 NOTE: i INDICATES ADS-ASSOCIATED QUENCHER Rev. 2, 5/80 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT QUENCHER DISTRIBUTION FIGURE 14

SSES LICENSIHC BASIS

1. Hark II Co<<tainment - Supporting Program A. LOCA - Related Tasks

. Task , Target Used for Number ., ~eetivtt A~ettett e C~ee ieti e Documentation SSES Licensin A.l "4T" Test Program Phase I Test Report, Completed HEDO/NEDE 13442-P-01 Yes Phase I Appl Memo Completed Application Memo Yes Phase II III Test gt Rept Completed HEDO(HEBE 13468-P Yes Application Hemorandum Completed REDO(NEDE 23678-P Yes A.2 Pool Swell Model Report Hodel Report Completed HEDO/HEDE 21544"P Impact Tests PSTF I/3 Scale Tests Completed REDO/HEBE 13426-P Yes Hark I I/12 Scale Tests Completed REDO/HEDC 20989-2P Yes Impact Hodel PSTF 1/3 Scale Tests Completed REDO/HEDE 13426-P Yes Hark I 1/12 Scale Tests Completed HEDO(NEDC 20989-2P No A.5", l.oads on Submerged LOCA/Rll Air Bubble Hodel Completed NEDO/HEDE 21471-P Yes (Partial)

Structures LOCA(RH Mater Jet Model Completed REDO/HEDE 21472-P Yes fpartial)

Ring Vortex Hodel Completed Letter Report. Ho 4Q 79 Topical Report Ho Applications Methods Completed REDO/NEDE 21730-P Ho Queue. Air Bubble Hodel 3Q 79 NEDO 21471 Supplement Ho Appl. Memo. Suppleme<<t 3Q 79 NEDE 21730 Supplement Ho Quencher hir Bubble I/4 Scaling Tests Complete HEBE 23817-P No Data Eval. 4Q 79 Report Ho Steam Condensation Hethods- Plant DAR's A.6 ~ Chugging, AnalYsis and Single Cell Report Completed NEDO/NEDE 23703 P Yes Testing Hultivent Model Completed NEDO/NEDE 21669-P No 4T FSI Report Completed HEDO/HEDE 23710-P Yes A.7. Chugging Single Vent CREARE Report Completed NFDO/HEDE 2185 l-P A.9 EPRI 'fest. Evaluation EPRI - 4T Comparison Completed REDO 2166I Yes EPRI I/13 Scale .Tests 3D Tests Completed EPRI NP-441 Yes EPRI Single Cell Tests Unit. Cell Tests 3Q 79 EPRI Report Yes Rev. 2, 5/80

Task Target Used for Humber Activi~t Activi st'1~pe C~ILio Documentation SSES Licensin A.ll Nultivent Subscale Testing Preliminary NV Prog Plan Completed HEDO 23697 Yes aud Analysis NV Test. Program Plan 6 Proc. Completed NEDO 23697 Rev 1 Yes Phase I Phase I Test Report 3Q 79 Report Yes NV Test. Prog Plan 6 Proc 3Q 79 REDO 23697, Rev. 1, Supp. 1 Yes

- Phase II Phase II Test. Report. 2Q 80 Report Yes CONNAP 'fests 3Q 79 Report Yes NlH Ver'afacataon 1/10 Scale Completed NEDE 25116-P Yes A.13 Single Vent Lateral. Loads Dynamic Analysis Completed HEDO 24106-P Yes Smsaary Report Completed NEDE 23806-P Yes Suaxaary Report (Extension) 3Q 79 Report, Yes A.16 Improved Chugging Load Impulse Evaluation Completed Letter Report Ho Definition Improved Chug Load Defn. 3Q 79 Report Ho t

A.17 Steam Condensation Dscill. 4T C.O. Test. 2Q go Report Undecided Rev. 2, 5/80

8 SRV Related Tasks h

Task Target Used for N bec '~cti it Act iviL~Tpe C~om lotion Documentation SSES Eicensin 8.1 Quencher Empirical Hodel DFFR tlodel Completed HEDO/HEDE 21061-P No Supporting Dal.a Completed NEDO/HEBE 21078-P Ho Ramshead Hodel OFFR Hodel Completed REDO/HEDE 21061-P Ho Supporting Dal.a Completed NEDO/HEBE 21062-P Ho Analysis Completed HEDO/NEOE 20942-P Ho

. Honticello ln-Plant e B.3 Preliminary. Test Completed HEDO/NEDC 21465-P No S/RV Tests Report Report'ydrodynamic Completed NEDO/NEOC .21581-P Ho 8.5 S/RV Quencher in-Plant Test. Plan CompleLed HEDtl 20988 Rev. 2 Ho Caorso Tests Test. Plan Addendum 1 Completed NEDH 20988 Rev. 2, Add 1 Ho

'Test. Plan Addendwa 2 Coaip le ted HEUH 209BS Rev. 2, Add 2 Ho Test SuaNaary "

CompleLed Letter Report Ho Phase I Test Report Completed HEDE-25100-P Ho Phase 11 Test Report 1Q 80 Report Ho 8.6 Thermal Hixing Hodel AnalyLical Hodel Completed NEDO/HEDC 23689-P 8.10 Konticello FSI Analysis of FS1 Completed REDO 23834 -No B.ll- DFFR Ramshead Hodel Data/tlodel Comparison Completed NSC-GEH 0394 To HonLicello Data 8.12 Ramshead SRV tlet.hodology Analytical Hethods Completed NEOO 24070 Summary 8 ..14 Queucher Empirical tlodel Hodel Confirmation 1Q 80 Report Update h

N e

~ - ~

~N

'N.

~

, Rev. 2, g/Sp

't h

C. Hiscellaneous Tasks Task Target Used for Number Activ~it A~ctiviL T n C~lv I an Documentation C.O Supporting Program Supp Prog Rpt Completed HEDO 21297 Supp Prog Rpt Rev. Completed NElS 21297 - Rev. 1 Supp Prog Rpt Rev. 4Q 78 REDO 21297 - Rev. 2 C;1 DFFR Revisions Revision 1 Completed NEOO/NEDE 21061-P Rev. 1

'i

~ Revision 2 Completed HEOO/HEDE 21061-P Nev. 2 Revision 3 . Completed HEDO/NEDE 21061-P Rev. 3 Yes (Partial)

C.3 HRC Round 1 Questions DFFR Rev. 2 Completed NEDO/HEDE 21061-P Rev. 2 Yes DFFR Rev. 2 Amendment 1 Completed REDO/NEDE 21061-P Rev. 2 Amend. 1 Yes DFFR Rev. 3. Appendix A Completed REDO/HEDE 21061-P Rev. 3 Appendix A Yes C.5 SRSS Justification interim Report 'ompleted 'HEDE 24010) Yes SRSS Report Completed HEOO/HEDE 24010-P Yes SRSS Exec. Report Completed Summary Report Yes SRSS Criteria Appl. Completed 'EDO/NEDE 24010-P Suppl. 1 Yes SRSS Bases Completed HEDO/HEDE 24010-P 2 'uppl.

Yes SRSS Justification Suppl 3Q 79 Report Yes C.6 NRC Round 2 Questions DFFR Amendment 2 Completed NElM/HEDE 21061-P Rev. 2 Amend. 2 Yes

~- DFFR Amend 2, Suppl 1 Completed HEDO/HEDE 21061-P Rev. 2 Amend. 2 Supp. 1 Yes DFFR Amend 2, Suppl 2 . Completed NEDO/NEDE 21061-P Rev. 2 Amend. 2 Supp. 2 Yes DFFR Rev. 3, Appendix A Completed NEDO/NEDE 21061-P Rev. 3 Appendix A Yes C.7 Justification of "4T" Chuggiug Loads Complete HEOO/HEBE 23617-P Yes Bounding Loads Justification Complete - REDO/NEDE 24013-P Yes Complete NEDO/NEDE 24014-P Yes Complete REDO/HEDE 24015-P Yes Complete NEDO/HEDE 24016-P Yes Complete HEDO/NEDE 24017-P Yes Complete NEDO/NEDE 23627-P Yes CIB S/RV and Chugging Prestressed Concrete FSi Reinforced Concrete Completed REDO/NEDE 21936-P Yes St,eel C.9 Honitor Morld Tests Honitor Tests End of Hone Program C.13 - Load Combinations & Criteria Justification Completed REDO 219B5 Yes Functional Capability Criteria C.14 HRC RuuiId 3 Quest.ions Letter Report Completed Letter Report. Yes DFFR, Rev. 3, Appendix A NECO/HEOE 21061-P ltev. 3 Appendix A Yes C.15 Submerged Structure Crit,eris NRC Questiou Responses 3Q 79 L<<tter Report Ycs Rev.- 2, 5/80

Il. KNI Tests and Reports (supplied to PP&L)

Document Used for-IIumber Title Status Documentation SSES Licensin Eormation and oscillation of a spherical gas Completed AEG " Report 2241 Yes bubble Analytical model. for clarification of pressure pulsation in the wetwell after vent cleaning Completed AEG - Report 2208 Yes Tests on mixed. condensation wiI.h aIodel quenchers ~ .=. Completed KWV - Report 2593 Yes Condensation and vent cleaning tests at GKII with quenchers Completed KWV - Report 2594. Yes Concept and design of the pressure relief system with quenchers 'Completed KWV - Report 2703 Yes:

KKB vent clearing with quencher Completed KMV - Report 2796 Yes Tests on condensation with quenchers when submergence of quencher arms is shallow Completed KMV - Report 2840 Yes

8. KKB - Concept and task of pressure relief system Completed KWV - Report 2871 Yes Experimental approach to vent clearjng in a model tank Completed KWV - Report 3129 Yes

10. KKB - Specification of blowdown tests during non-nuclear hot.functional test - Rev. I dated October 4, 1974 Completed KWU/V 822 Report Yes Anticipated data for blowdown tests with pressure relief system during the non-nuclear" hot functional test at nuclear power station Brunsbuttel (KKB) Completed KNI - Report 3141 Yes

12. Results of the non-nuclear hot functional tests with thc pressure relief system in the nuclear power station Brunsbuttel Completed KNI - Report 3267 Yes

13. Analysis of the loads measured on the pressure relief system during the non-nuclear hot functional test at KKB .'Completed . . KNI - Report, 3346

14. KKB - Listing of test parameters aud important test data of the non-nuclear hot functional tests with the pressure relief system =

Completed KNI - Working Report R 521/40/77

15. KKB - Specification of additional tests for testing of the pressure relief valves during the nuclear- start-up, Rev. 1 Completed ~Q/V 822 TA Yes Rev-. 2,=5/80

Oocument Used for Humber Title Status Documentation SSES Licensin KKB - Results from nuclear start-up testing of pressure relief system Completed KMU - Morking Report Yes R 142-136/76

.17. Huclear Po~er Station Phillipsburg I- Unit 1 Hot functional Test: Specification of pressure relief valve tests as well as emergency cooling and wetwell cooling systems Completed KNI/V B22/RF 13 Yes lg. Results of the non-nuclear hot functional tests with the pressure relief system in the nuclear power station Phillipsburg Completed 'KMU - Morking Report Yes

. R . 142-3B/77 KKPI - List,ing of test parameters and im'portant test data of the non-nuclear hot funct.ional tests with the prcssure relief system Completed . KNI - Morking Report Yes R 521/41/77

20. Air oscillations during vent clearing with single and double pipes Completed AEG - Report 2327 Yes Rev. 2, 5/80

TABLE 1-2 SSES CONTAINMENT DESIGN DIMENSIONS A. Suppression Chamber Inside Diameter 88 ft 0 in Height 52 ft 6 in B. Drywell Inside Diameter of

. Base 86 ft 3 in Inside Diameter of Top 36 ft 4.5 in Height 87 ft 9 in C. Reactor Pedestal Inside Diameter Below Diaphragm Slab 19 ft 7 in Inside Diameter Above Diaphragm Slab 20 ft 3 in Wall Thickness Below Diaphragm Slab 5 ft 1 in Wall Thickness Above Diaphragm Slab 4 ft 5 in Height 81 ft 9.6 in D. Reinforced Concrete Thickness Base Foundation Slab 7 ft 9 in Containment Wall 6 ft 0 in Diaphragm Slab 3 ft 6 in

Table 1-2 (Cont'd)

E. Steel Line Plate Thickness for Base Foundation, Containment Wall; and Diaphragm Slab 0.25 in F. Suppression Chamber Columns Outside Diameter 3 ft 6 in Wall Thickness 1.25 in Height 52 ft 6 in

~ \

TIE 1-3 SSES CONTAINMENT DESIGN PARAMETERS A. Dr ell and Su ression Chamber Drr~ell Su ression Chamber

1. (a) Internal Design Pressure 53 psig 53 psig 1 (b) Internal Design Pressure in Combination 44 psig 29 psig with other Loads

2. External Design Pressure 5 psid 5 psid

3. Drywell Floor Design Differential Pressure Upward 28 psid Downward 28 psid 4~ Design Temperature 340 F 220- F 3

5. Drywell Free Volume (Minimum) 239,337 ft3 (including vents) (Normal) 239,593 ft3 (Maximum) 239,850 ft 3

6. Suppression Chamber Free (Minimum) 148,590 ft3 Volume (Normal) 1533860 ft (Maximum) 159.,130 ft 3

7. Suppression Chamber Water Volume (Minimum) 122,410 ft3 (Normal) 126,980 ft3 (Maximum) 131,550 ft

8. Pool Cross-Section Area Gross (Outside Pedestal) 5379 ft 2 Total Gross (Including Pedestal Water Area) 5679 ft Free (Outside Pedestal) 50.65 ft Total Free 5277 ft REV. 6, 4/82

Table 1-3 (Cont'd)

Drr~ell Su ression Chamber

9. Pool Depth (Minimum) 22 ft.

(Normal) 23 ft.

(Maximum) 24 ft.

l. Number of Downcomers 82 (Five capped: see Appendix K)

2. Downcomer Outer Diameter 2 ft.

3. Total Downcomer Vent Area 257 ft.

4. Downcomer Submergence (Minimum) 10 ft.

(Normal)

(Maximum) ll ft.

12 ft.

5. Downcomer Loss Factor 2.5 C. Safet Relief Valves

l. Opening Time

a. Delay Time (between trip and motion) 0.10 sec.

b. Response Time (close to open) 0.15 sec.

REV. 6, 4/82

Table 1-3 (Cont'd)

2. Safety and Relief Setpoints for the 16 valves.

ASME Rated Capacity at 103%

Spring Set* Pressure Switch*+ of Spring Set Valves Pressure si Set Pressure si Pressure lb./hr.

(See Figure 1-4)

B>E 1146 1076 862,400 ASCPD 1175 1086 883,950 PRRRS 1185 1096 891,380 JsL>N 1195 1106 898,800 G,K,M 1205 1116 906,250 1175 1096 883,950 1185 1086 891,380

• Will open if switch

>* Reset pressure fails 55 to 100 psi below pressure switch set point 3~ Reaction Forces (vertical, Fv, and horizontal, Fh) on valve supports during Valve Opening and Closing at 1250 psig.

a. No Flow Established Fv = 60,300 lb.

Fh = 23,600 lb.

b. At Full Flow Fv = 56,200 lb.

Fh = 24,200 lb.

Table 1-3 (Cont'd)

4. Maximum Steam Flow Rate at 70 bar (1000 psig)*

Reactor Pressure (conservative value for design calculation) 390.93 metric tons/hr (862,400 lb/hr)

• When a value is given in two sets of units, the first value is the original one; the second is an approximation provided for convenience.

D. Safet Relief Valve Dischar e Pi es

1. Outer Diameter 12 in

2. Distance of Quencher Middle Plane to Basemat 3 ft 6 in

3. Quencher Submergence (Minimum) 18.5 ft (Normal) 19.5 ft (Maximum) 20.5 ft

4. Length, Number of Bends, and Air Volume for each SRV Pipe-Pi e Len th ft- Number of Bends Quencher Inside Inside Inside Inside Air Position ~Dr ell Wetwell Total ~Dr ell Wetwell Total Volume ft (See Figure 1-4)

A 67. 67 73. 11 140.78 12 15 92.38

66. 4 73. 23 139.63 13 91.48 67.71 54. 47 122. 18 10 10 78.12 69.95 75. 16 145. 11 10 95.79 93.06 54. 47 147.53 16 16 98.03

61. 96 54.47 116.43 73.6

70. 40 75.04 145.44 12 96. 05 H 73. 09 78. 22 151.31 12 16 100.66

Table 1-3 (Cont'd)

Pi e Len th ft Number of Bends Quencher Inside Inside Inside Inside Air Position ~Dr ell Wetwell Total Dr~ell Wetwell Total Volume ft

73. 34 74. 85 148. 19 12 15 98. 2

80. 82 72. 53 153. 35 13 102.34

67. 44 54. 47 121. 91 77.91

59. 84 54.47 114. 35 71.97 N 75. 09 81.60 156.69 12 17- 105.15

71. 77 83. 91 155. 68 10 14 104.1

72. 59 54. 47 127. 06 12 12 81.95

67. 23 72. 11 139. 34 13 16 91.25

Page 1 TABLE 1&

Review of Susquehanna SES Units 1 4 2 Pool Dynamic Loadings-

-Com arison with NUREG 0487, NUREG 0487-Su lement No. 1. Lead Plant and Generic Lon Term Pro ram-NRC Acceptance Criteria Lead Plant Position Generic Long Term NUREG 0487 Su lement No. 1 (Zimmer DAR, Amendment 13) Pro ran Position Sue uehanna Position Remarks I. LOCA RELATED HYDRODYNAMIC LOADS A. Submerged-Boundary Loads 24 PSI overpressure statically applied March 20, 1979 letter. 24 Evaluating impact. Evaluation During Vent Clearing. with hydrostatic pressure to surfaces psi statically applied to indicates 24 PSI 33 psi overpressure added below vent exit (attenuate to 0 psi surfaces below vent exit overpressure is to local hydrostatic at pool surface) for period of vent (attenuate to 0 psi at conservative (see below vent exit (walls clearing for plants with (mhL)/ pool surface) for period of Subsection 4.2.1.2) and basemat)-linear at- A A VDN vent clearing. Zimmer and tenuation to pool sur- where: m mass flow in vents lb/sec LaSalle meet NUREG 0487.

face. VD DN drywell volume ft3 n enthalpy of air in vent-Btu/lb L submergence ft A /A ~ pool area to vent area Por plantR wKere (mhL)/[(A /A )V >) >55, the loading increase over RydrosPatic pressure on basemat and submerged walls below vent exit is p 24 + 0.27 (mhL) /

[(A /A )>V ] -55 (attenuate to 0 psi at )os suNace).

B. Pool Swell Loads.

1. Pool Swell Analytical Model (PSAM)

a. Air bubble pres- (a) No change from NUREG 0487. (a) Accept NUREG 0487. (a) Accept NUREG (a) Accept NUREG sure-use PSAM 0487. 0487.

described in NEOE-21544-P.

b. Pool swell eleve- (b) Use PSAM with polytropic exponent (b) Accept NUREG 0487. (b) Accept NUREG (b) Accept NUREG 0487 tion-Use PSAM dcs- of 1.2 to a maximum swell height 0487 -Sup- -Supplement Ho. 1 plement Ho. 1 REV. 6, 4/82

Page 2 TABLE 1-4 NRC Acceptance Criteria Lead Plant Position Generic Long Term NUREG 0487 Su lement No. 1 Zimmer DAR Amendment 13 Pro ran Position Sus uehanna Position Remarks cribed in NEDE- which is the greater of 1.5 vent 24544-P with a submergence or the elevation cor-polytropic expo- responding to the drywell floor nent of 1.2 for uplift A P used for design assess-wetwell air com- ment per response to Question pression. 020.68 and February 16, 1979 let-ter from Shoreham provided the drywell pressure response used for the swell height is calculated according to NEDN-10320.

c. Pool swell velo- (c) No change from NUREG 0487. (c) Accept NUREG 0487 with (c) Accept NUREG (c) Following lead city-use PSAN des- velocity vs elevation 0487 with velo- plant/long term cribed in NEDE- obtained from PSAM. city vs eleva- position.

24544-P multiplied tion obtained by a factor of 1.1. from PSALM.

d. Pool swell acceler- (d) No change from NUREG 0487. (d) Accept NUREG 0487. (d) Accept NUREG (d) Accept NUREG ation-use PSAM des- 0487. 0487.

cribed in NEDE-24544-P.

e. Vetwell air com- (e) No change from NUREG 0487. (e) Accept NUREG 0487. (e) Accept NUREG (3) Accept NUREG pression-use PSALM 0487. 0487.

described in NEDE-24544-P.

f. Drywell pressure (f) No change from NUREG 0487. (f) Accept NUREG 0487. (f) Accept NUREG (f) Accept NUREG history-unique 0487. 0487.

based on NEDN-10320.

2. Loads on Submerged No change from NUREG 0487. Accept NUREG 0487. Accept NUREG 0487. Accept NUREG 0487.

Boundaries. Haximum bubble p~essure pre-dicted by PSAH is to be added uniformly to Rev. 5, 3/81

Page 3 TABLE 1-4 NRC Acceptance Criteria Lead Plant Position Generic Long Term NUREG 0487 Su lement No. 1 Zimmer DAR Amendment 13) Pro ram Position Sus uehanna Position Remarks local hydrostatic be-low vent exit (walls and basemat) and linear attenuation to pool surface. Apply to walls up to maxi-mum pool swell eleva-tion.

3. Impact Loads

a. Small structures- (a) No change from NUREG 0487. (a) Accept NUREG 0487. (a) Accept NUREG (a) Accept NUREG (For horizontal 0487. 0487.

pipes, I-beams, and other similar structures having one dimension < 20 in.). The loading function shall have the versed sine shape:

p(t)=0.5 p (1-COS

b. Large structures- (b) No change from NUREG 0487. (b) Not applicable (no (b) Not applicable (b) Not applicable not applicable, large structures). (no large (no large struc-no large struc- structures). tures).

tures are impacted by pool swell.

c. Grating-The static (c) No change from NUREG 0487. (c) Not applicable (no (c) Accept NUREG (c) Not applicable drag load, F , is grating). 0487 with velo- (no grating in to be calculated city vs eleva- pool swell zone).

by forming the tion obtained product of AP from from PSAM.

Figure 4-40 of NED0-21060, Rev.

Rev. 5, 3/81

Page 4 TABLE 1-4 NRC Acceptance Criteria Lead Plant Position Generic Long Term NUREG 0487 Su lement No. 1 Zimaer DAR Amendment 13 Pro ram Position Sus uehanna Position Remarks 2, and the total area of the grat-ing. To account for the dynamic nature of the initial loading, the static drag load is increased by a multiplier given by:

F E/D = I+ I+(0.064Mf) fear Mf < 2000 in/sec

4. Metall Air Compres-sion

a. Mall loads-direct- (a) No change from NUREG 0487. (a) Accept 0487. (a) Accept NUREG (a) Accept NUREG ly apply the PSAN 0487. 0487.

calculated pres-sure due to wetvell compression.

Diaphragm upuard (b) No change from NUREG 0487. (b) Use A PUP = 5.5 (b) Same as lead (b) Same as lead load-calculate A PSID. plant. plant.

PUP using the cor-relation:

A PUP = 8.2 - 44F, for 0< F <0.13 A PUP = 2.5 psi, for F> 0.13 where. F =

2 VD (AV)

AB = break area AP = net pool area AV = total vent area REV. 6, 4/82

Page 5 TABLE 1 4 NRC Acceptance Criteria Lead Plant Position Generic Long Term NUREG 0487 Su lement No. 1 (Zimmer DAR, Amendment 13) Pro ran Position Sus uehanna Position Remarks VS initial wetwell air space volume VD drywell volume

5. Asymmetric Load. Use twice the IOX of maximum bubble Accept NUREC 0487-Supple- Accept NUREG 0487- Accept NUREC 4087-Apply the maximum pressure statically applied to 1/2 ment No. l. Supplement No. 1 Supplement No. l.

air bubble pressure of the submerged boundary (with calculated from PSAH hydrostatic pressure) proposed in and a minimum air March 16> 1979 letter from GE.

bubble pressure (sero increase) in a worst case distribution to the wetwell wall.

C. Steam Condensation and Chugging Loads.

.1. Downcomer Lateral Loads.

a. Single vent loads: (a) No change from NUREG 0487. (a) Accept NUREC 0487. (a) Use single vent (a) Following long See DAR,

-h static equiva- dynamic lateral term program. Subsec-lent load of 8.8 load developed Conf irmation tion 9.6.3 KIPs shall be under Task A-13 through plant for verifi-used provided: (NEDE-24106-P) . unique GKH-IIM cation of However, extra-test data on lateral tip (i) the downcomer is polate the 30 lateral bracing load.

24" in diameter. Kip and 3 msec loads.

(ii) the downcomer dom- impulse to inant natural fre- 65 Kips and 3 msec.

quency is < 7 ax, submerged.

(iii) the downcomer is unbraced or braced>

at or above approx. 8'rom the exit.

REv. 6, 4/82 .

Page 6 TABLE 1-4 NRC Acceptance Criteria Iead Plant Position Generic Long Term NUREG 0487 Su lement No. 1 Zimmer DAR Amendment 13 Pro ram Position Sus uehanna Position Remarks

-A static equiva-lent load of 8.8 Kips multiplied by the ratio of the natural fre-quency and 7 Hz for dominant na-tural frequencies between 7 and 14 Hz. Other res-trictions in (i) and (iii) apply.

-If the natural frequency of the downcomer is > 14 Hz or if bracing is closer than the exit, 8'bove a

plant specific dynamic structural calculation shall be performed using a dynamic load defined by:

F(t) "-

FO sin  %

for t<0and t ; 0 <t <r t> r where: 2 msec < r <10 msec, and the impulse I -" 2 FO (r/%) is 200 lbf-sec.

Restriction (i) also ap-plies.

Rev. 5, 3/81

0 Page 7 TABLE 1-4 NRC Acceptance Criteria Lead Plant Position Generic Iong Term NUREG 0487 Su lement No. 1 Zimmer DAR Amendment 13) Pro ram Position Sus uehanna Position Remarks

b. Multiple vent (b) No change. (b) Accept NUREG 0487. (b) Use multivent (b) Following long loads - Use the lateral load term program.

load specified in methodology do-Figure 4-10b of cumented in NEDE-21061-P, Rev. letter report 2, multiplied by "Method of Ap-a factor of 1.26 plying Hark II for downcomers Single Vent Dyna-with natural're- mic Lateral Load quencies less to Hark II Plants than 7 Hz. For with Multiple natural frequen- Vents", trans-cies greater than mitted to the 7 Hz, apply an NRC on April"',9, additional multi- 1980 under Task plier equal to A.13.

the ratio of its frequency and 7 Hz.

2. Submerged Boundary Loads

a. High Steam Flux (a) No change from NUREG 0487. (a) Accept NUREG 0487 with (a)+Use Condensation (a) Use IWEGS/MARS (a) Application Loads additional plant Oscillation acoustic model procedure unique empirical load load specifica- documented in documented Sinusoidal pres- specification. tion based on NEDE-24822-P in SSES sure fluctuation NEDE-24288-P. with sources DAR, Sec-added to local derived from GEM tion 9.5.

hydrostatic. II-H steam con-Amplitude uniform densation tests.

below vent exit, linear attenuation to pool surface.

4.4 psi peak-to-peak amplitude.

2-7 Hz frequencies.

NEDE-21061-P, Rev 2.

Rev. 5, 3/81

Page 8 TABLE 1-4 NRC Acceptance Criteria Lead Plant Position Generic Long Term NUREG 0487 Su lament No. 1 Zimmer DAR Amendment 13 Pro ram Position Sus uehanna Position Remarks

b. Medium Steam Flux (b) No change from NUREG 0487. (b) Accept NUREG 0487 with (b) Use Condensa- (b) Same as (a).

Loads. additional plant unique tion Oscilla-empirical load specifi- tion load Sinusoidal pres- cation. specification sure fluctuatioa based on NEDE-added to local 24288-P.

hydrostatic. Amp-litude uniform be-low vent exit, linear attenuation to pool surface.

7.5 psi peak-to-peak amplitude.

2-7 Hz fr~qu~ncies.

NEDE-21061-P, Rev. 2

c. Chugging. (c) No change from NUREG 0487. (c) Accept NUREG 0487 with (c) Use IMEGS/MARS (c) Same as (a).

additional plant acoustic model

-Uniform loading unique empirical load presented in condition- specificatioa. NEDE-24822-P with Maximum amplitude sources derived uniform below vent from 4T-CO. Ap-exit, liaear at- plication metho-tenuation to pool dology documented surface. +4.8 in NEDE-24302-P.

psi max overpres sure, -4.0 psi max underpressure.

(Peading resolu-tion of FSI con-cerns)

NEDE-21061-P )

Rev. 2.

-Asysaetric loading condition - Maxi-REV. 6, 4/82

Page 9 TABLE 1-4 NRC Acceptance Criteria Lead Plant Position Generic Long Term NUREG 0487 Su lement No. 1 (Zimmer DAR Amendment 13 Pro ram Position Sus uehanna Position Remarks mum amplitude uni-form below vent exit - linear at-tenuation to pool surface. +20 psi max overpressure,

-14 psi max under-pressure.~ 20-30 Hz frequency, peripheral varia-tion of amplitude follows observed >

statistical dis-ribution with maximum and mini-mum diametrically opposed. 'HEDE-21061-P, Rev. 2 Rev. 5, 3/81

Page 10 TABLE 1-4 NRC Acceptance Criteria Lead Plant Position Generic Long Term NUREG 0487 Su lement No. 1 Zimmer DAR Amendment 13 Pro ram Position Sus uehanna Position Remarks II. SRV-RELATED HYDRODYNAMIC LOADS A. Pool Temperature Limits All Hark II facilities No change from NUREG 0487. Accept NUREG 0487. Accept NUREG 0487. Accept NUREG 0487.

shall use quencher type Mass & Energy analy-devices. The suppres- sis documented in sion pool local temp- SSES DAR Appendix I.

erature shall not ex-ceed 200 F for all plant transients involving SRV operations. Heasure-ments from temperature sensors located on the containment wall in the sector containing the discharge device at the same elevation as the device can be used as local indication.

B. Air Clearing Loads.

a. Hethodology for bub- (a) Accept "Interim T-Quencher load (a) T-Quencher load speci- (a) T-Quencher load- (a) Same as lead ble load prediction Definition" with the following fication presented in Same as lead plant.

T-quencher - use modifications: Susquehanna DAR, Subsec- plant.

ramshead methodology described in Sec.

-Bubble frequenncy-3 to ll

-Peak Pressure Multiplier for Hz tion 4.1.3.

NUREG 0487 -

Accept Supplement X-Quencher load-3 i of NED0-21061-P, Subsequent Actuation - 1.5 No. 1 modifications Plant unique Rev. 2. -Vertical Pressure Profile- except use bubble fre- load definition.

maximum amplitude from basemat quency in SSES DAR and x-quencher - Use Sec to 2.5'bove quencher center a peak pressure multiplier 3.3 of NED0-21061-P, line, linear attenuation to of 1.5 for all actuations.

Rev. 2. zero at pool surface.

-Multiple SRV Actuations-

1) linear ABSS superposition of peak single values with all bubbles in phase.

Rev. 5, 3/81

2) if the combined peak pressure exceeds, local single value peak use the lower value

Page ll TABLE 1-4 NRC Acceptance Criteria Lead Plant Position Generic Long Term NUREG 0487 Su lement No. 1 Zimmer DAR Amendment 13 Pro ram Position Sus uehanna Position Remarks

b. SRV Discharge Load (b) Same as NUREG 0487 but load case (b) Accept NUREG 0487-Sup- (b) Accept NUREG (b) Accept NUREG Cases. The follow- 4 is not included. plement No. l. 0487-Supplement 0487-Supplement ing load cases shall No. l. No. 1.

'be considered for design evaluation of containment struc-tures and equipment inside the contain-ment:

1. Single valve, first and subse-quent actuation.

2. ADS valve actua-tion.

3. Two adjacent valve first actuation.

4. All valves dis-charged sequential-ly by setpoint.

5. All valves dis-charged simulta-neously by assum-ing all bubbles are oscillating in phase.

Rev. 5, 3/81

Page 12 TABLE 1-4 NRC Acceptance Criteria Lead Plant Position Generic Long Term NUREG 0487 Su lement No. 1 Ziaxaer DAR Amendment 13 Pro ram Position Sus uehanna Position Remarks

c. Bubble Frequency. (c) 3-11 Hz. (c) Plant unique frequency (c) Same as lead (c) Following frequen- Additional T-quencher - a range range based on Susque- plant. cy range document- study per-of bubble frequency hanna DAR. ed in Susquehanna formed con-of 4-12 Hz is the DAR. firming con-minimum range that servation of shall be increased required to include if frequency range in Sus-the frequency pre- quehanna DAR dicted by the. rams- Subsec-

-'see head methodology 10.2.3). 'ion together with i'0X margin.

X-quencher - a range X-quencher bubble of bubble frequency frequency being of 4-12 Hz shall be developed by Burns evaluated. & Roe based largely on Caorso test data.,

c. Quencher Arm and Tie Down Loads.

l. Quencher Are No change from NUREG 0487. Accept NUREG 0487. Load T-quencher arm Following long term Loads. Vertical Specification in SSES DAR loads are presen- programs and lateral are Subsection 4.1.2.5 used ted in Susquehanna loads are to be to verify the conserva- DAR, Section 4.1.2.5.

developed on the tism of this approach.

basis of bound- X-quencher-Accept ing assumptions NUREG 0487.

for air/water dis-charge from the quencher and con-servative combi-nations of maxi" mum/minimum bubble pressures acting on the quencher per NEDE-21061-P, Rev. 2.

REV. 6, 4/82

1 Page 13 TABLE 1-4 NRC Acceptance Criteria Lead Plant Position Generic Long Term NUREG 0487 Su lement No. 1 (Zimmer DAR Amendment 13 Pro ram Position Sus uehanna Posi.tion Remarks

2. Quencher Tie-down Loads. No change from NUREG 0487. Accept NUREG 0487. Load T-quencher tie-down Following long term The vertical and specification in SSES DAR loads are defined program.

lateral arm load Subsection 4.1.2.6 used in Susquehanna DAR, transmitted to to verify conservatism. Subsection 4.1.2.6.

the basemat via the tie-down plus vertical transient X-Quencher-Accept wave and thrust NUREG 0487.

loads calculated from a standard momentum balance are to be calcu-lated based on con-servative clearing assumptions per NEDE-21061-P, Rev.

2.

Rev. S, 3/81

Page 14 TABLE 1-4 NRC Acceptance Criteria Lead Plant Position Generic Iong Term NUREG 0487 Su lement No. 1 Zimmer DAR Amendment 13 Pro ram Position Sus uehanna Position Remarks III. LOCA/SRV SUBMERGED STRUCTURE LOADS A. LOCA/SRV Jet Loads.

l. LOCA Downcomer Jet Accepts alternative methodology pre- The LOCA downcomer jet Ring matex model de- Following lead plant Load sented in Zimmer DAR dealing with load is calculated by veloped by Burns 6 position.

LOCA jet load. the methodology presented Roe used for WPPSS Calculate based on in the Zimmer DAR, Sub- Unit g2. Remaining methods described section 5.3.2.1. plants following lead in NEDE-21730 and plant methodology.

the following cons-traints and modifi-cations:

(a) Standard drag at the time the jet first encounters the structure must be multiplied by the factor:

6-V CD X'R where:

V =acceleration volume a

as defined in NEDE-21730.

D==drag coefficient as defined in NEDE-21730.

A ~rojected area as defined in NEDE-21730.

R.ment 1

exit radius.

Rev. 5, 3/81

Page 15 TABLE 1-4 NRC Acceptance Criteria Lead Plant Position Generic Long Term NUREG 0487 Su lement No. 1 (Zimmer DAR Amendment 13 Pro ram Position Sus uehanna Position Remarks (b) Forces in the vi-cinity of the jet front shall be computed on the basis of Formula 2-12 and 2-13 of NEDE-21730. The local velocity, U , . and accel-eration, U , are to be conserva-tively calculated by the methods of NEDE-21471 from the potential function:

-3 8%'. U. . V w Cos e 2

where:

r & 0 m spherical co-ordinates from jet front.

U.

j jet velocity from NEDE-21730.

initial vol-ume of water in the vent.

(c) After the last fluid particle has reached the jet front a spherical vortex Rev. 5, 3/81

Page 16 TABLE 1-4 NRC Acceptance Criteria Iead Plant Position Generic Iong Term NUREG 0487 Su lement No. 1 Zimaer DAR Amendment 13 Pro ram Position Sus uehanna Position Remarks continues propa-gating. The drag on structures in its vicinity can be bounded by using the flow field from the formula for g above with U. as the jet fred ve-locity from NEDE-21730 at time t

= tf.

2. SRV Quencher Jet Loads This load may be ne- SRV quencher )et loads may be ne- Accept NUREG 0487 - Sup- Accept NUREG 0487- Accept NUREG 0487-glected for those glected beyond a 5'ylindrical plement No. l. Supplement No. 1 Supplement No. l.

structures located zone of influence.

outside a zone of X-quencher - Accept influence which is NUREG 0487.

a sphere circums-cribed around the quencher arms. If there are holes in the end caps; the radius of the sphere should be increased by 10 holes diameters.

(Confirmation during long term program required).

B. IDCA/SRV Air Bubble Drag Leads.

Rev. 5, 3/81

Page 17 TABLE 1-4 NRC Acceptance Criteria Lead Plant Position Generic Long Term NUREG 0487 Su lement No. 1 Zisrner DAR Amendment 13 Pro ram Position Sus uehanna Position Remarks 1~ IOCA Air Bubble Loads No change from NUREG 0487 Documented in plant unique Documented in Documented in Subsec-DAR's. plant iinique DAR's. tion 4.2.1.7 of SSES based on 'alculate DAR.

the analytical model of the bubble charg-ing process and drag calculations of NEDE-21471 until the bub-bles coalesce. After bubble contact, the pool swell analytical model, together with the drag computation procedure NEDE-21471 shall be used. Use of this methodology shall be subject to the following cons- grams~

traints and modifi" cations:

a. A conservative (a) No change (a) Position documented (a) Accept NUREG- (a) Following the Document-estimate of bub- on page 5.4-8 of 0487. Long Term Pro- ed in Sub-ble asycaetry of Zisraer DAR. section shall be added 4.2.3.2 of by increasing SSES DAR.

accelerations and velocities computed in step 12 of Section 2.2 of NEDE-21730 by 10'. If the alternate steps 5A, 12A and 13A are used the ac-celeration drag shall be directly REV. 6, 4/82

f.

Page 18 TABLE 1-4 NRC Acceptance Criteria lead Plant Position Generic Long Term NUREG 0487 Su lement No. 1 Zimmer DAR Amendment 13 Pro ram Position Sus uehanna Position Remarks increased by 10$

while the standard drag shall be in-creased by 20$ .

b. Modified coeffi- (b) Accept lead plant position docu- (b) Position documented on (b) Following Lead (b) Following Lead (b) Addressed cients C'rom mented in Attachment l.k of the page 5.4-8 of Zimmer Plant Position Plant Program. in Subsec-accelerating flows Zimmer FSAR with the following DAR. and evaluating tion 4.2.

as presented in modifications: NUREG 0487- 3.3 of SSES DAR.

Kenlegan 8 Carpen- (1) Use C =C -1 in the FA formula. Supplement No. 1 ter and Sarpkaya (2) For non-cylindrical structures modifications.

references shall use lift coefficients for ap-be used with trans- propriate shape of CL = 1.6.

verse forces in- (3) The standard drag coefficient cluded, or an upper for pool swell and SRV oscil-bound of a factor lating bubbles should be based of three times the on data for structures with standard drag coef- with sharp edges.

ficients shall be used for structures with no sharp cor-ners or with stream-wise dimensions at least twice the width.,

c. The equivalent uni- (c) Accepts lead plant position. (c) Position documented on (c) Following Lead (c) Following Long (c) Addressed form flow velocity page 5.4-8 of Zimmer Plant Position. Term Program. in Subsec-and acceleration DAR. tion for any structure 4.2.3.4 of SSES DAR.

or structural seg-ment shall be taken as the maximum values "seen" by that structure not the value at the geometric center.

Rev. 5, 3/81

Page 19 TABLE 1-4 NRC Acceptance Criteria Lead Plant Position Generic Iong Term NUREG 0487 Su lement No. 1 (Zimmer DAR Amendment 13 Pro ram Position Sus uehanna Position Remarks

d. For structures that (d) Accepts Lead Plant position. (d) Position documented (d) Following Lead (d) Following Long (d) Addressed are closer together on page 5.4-8 of Plant Program. Term Program. in Subsec-than three charac- Zimmer DAR. section teristic dimensions 4.2.3.5 of of the larger one, SSES DAR.

either a detailed analysis of the interference ef-fects must be per-formed or a conser-vative multiplica-tion of accelera-tion and drag for-ces by a factor of four must be per-formed.

e. If significant (e) No change from NUREG 0487. (e) Position documented (e) Following lead (e) Following Long (e) Addressed blockage from on page 5.4-9 of Plant Program. Term Program. in Subsec-downcomer brac- Zimmer DAR. tion ing exists rela- 4.2.3.6 of tive to the net SSES DAR.

pool area, the standard drag co-efficients shall be modified by con-ventional methods (Pankhurst &

Holder reference).

f. Formula 2-23 of (f) No change since NUREG 0487. (f) Accept NUREG 0487. (f) Accept NUREG (f) Accept NUREG (f) Documented NEDE-21730 shall 0487. 0487. in DAR, be modified by Subsection replacing M by 4.2.3.7.

Tables 2-1 and 2-2.

Rev. 5, 3/81

Page 20 TABLE 1-4 NRC Acceptance Criteria Lead Plant Position Generic Long Term NUREG 0487 Su lement No. 1 Ziasaer DAR Amendment 13 Pro ram Position Sus uehanna Position Remarks

2. a. SRV ramshead air (a) No change since NUREG 0487. (a), Documented on Page (a) N/A (a) N/A bubble loads. 5.4-9 of Zinger DAR.

b. SRV quencher air (b) No change since NUREG 0487. (b) Documented on Page (b) T-quencher sub- (b) Following Long bubble loads. 5.4-9 of Zimmer DAR. merged structure Term Program T-quencher- methodology is loads may be comp- presented in uted on the basis Susquehanna DAR, of the above rams- Section 4.1.3.

head bubble pres-sure and assuming the bubble to be located at the .

center of the quen-cher device having a bubble radius equal to the quen-cher radius.

X-quencher - loads X-quencher methodo-may be computed on logy being developed basis of the 'he by Burns & Roe.

above ramshead meth-odology using bub-ble pressure cal-culated by the methods of NEDE-21061-P, Rev. 2 for the X-quen-cher.

C. Steam Condensation Drag Loads.

Review will be conducted No change since NUREG 0487. Documented on Page 5.4-9 Plant unique meth- Plant unique methodo-on a plant unique basis. of Zissner DAR. od being develop- logy documented in DAR ed. Subsection 4.2.2.5.

PAF:cvc 34P-B REV. 6, 4/gg

CHAPTER 2

SUMMARY

TABLE OF CONTENTS

2. 1 LOAD DEFINITION

SUMMARY

2. 1.1 SRV Load Definition Summary

2. 1.2 LOCA Load Definition Summary 2.2 DESXGN ASSESSMENT

SUMMARY

2.2.1 Containment'tructure and Reactor Assessment Summary 'uilding 2.2. 1. 1 Containment Structure Assessment Summary 2.2.1.2 Reactor Building Assessment Summary 2.2. 2 Containment Submerged Structures Assessment Summary 2.2. 3 Piping Systems Assessment Summary

2. 2.4 Equipment Assessment Summary 2a 2 ~ 5 Electrical Raceway System Assessment Summary 2.2.6 HVAC Duct System Assessment Summary Rev. 2, 5/80 2-1

This Design Assessment Report contains the SSES adequacy evaluation for dynamic loads due to LOCA and SRV discharge.

Rev. 2, 5j80

2,1 LOAD'EFINITION

SUMMARY

2-1.1 SRV Load Definition Summary Hydrodynamic loads resulting from SRV actuation fall into tvo distinct categories: loads on the SRV system itself (the discharqe line and the discharge quencher device), and the air clearing loads on the suppression pool walls and submerged structures.

Loads on the SRY system during SRV actuation include loads on the SRV pipinq due to effects of steady backpressure, transient vater slug clearing and SRV line temperature. Determination of loading on the quencher body, arms, and support is based on transients resultinq from valve opening (water clearing and air clearing),

valve closinq and operation of an adjacent quencher.

Air clearing loads are examined for four loading cases:

symmetric (all valve) SRV actuation, asymmetric adjacent SRV l actuation, single SRV actuation, and Automatic Depressurization System (ADS-six valves) actuation. Dynamic forcing functions for loadinq of the containment walls, pedestal, basemat, and submerged structures are developed using techniques developed in Section 4.1. Loads on the SRU system due to SRV actuation are discussed 'in Subsection 4.1.2, and loads on suppression pool structures due to SRV actuation are discussed in Subsection 4.1.3. A full scale, unit cell test program vas employed to verify SSES unique SRV loadinq as described in Chapter 8.

2. 1. 2 LOCA Load Definition Summary

,The spectrum of LOCA-induced loads on the SSES containment structure is characterized by LOCA loads associated vith poolswell, condensation oscillation and chugging loads, as well as long term LOCA loads.

The LOCA loads associated vith poolsvell result from short duration transients and include downcomer clearing loads, water jet loads, poolswell impact and drag loads, pool fallback drag loads, poolsvell air bubble loads, and loads due to dryvell and wetwell temperature and pres'sure transients. Techniques used to evaluate these loads are described in Subsection 4.2. 1.

Condensation oscillations result from mixed flow (air/steam) and pure steam flow effects in the suppression pool. Chugging loads result from lov mass flux pure steam condensation. The load definitions for these phenomena are contained in Subsection 4 2.2.

Long term LOCA loads result from those wetwell and drywell pressure and temperature transients vhich are associated with design basis accidents (DBA), intermediate accidents (IBA), and small break accidents (SBA) . Their load definitions are contained in Subsection 4.2.5.

Rev. 2, 5j80 2-3

Structures directly aff ected by'OCA loads include the drywell walls and floor, wetwell walls, RPV pedestal, basemat, liner

'late, columns', downcomers, downcomer bracing system, quenchers, and wetwell piping. Their loading conditions are described in Subsection 4.2.6.

Rev. 2, 5/80 2-4

2 2 - DESIGN-A~SS SSHENT. SUN HAH-Y-Design assessment of the SSES structures and components is achieved by analyzing the response of the structures and

~

components to the load combinations explained in Chapter 5. In Chapter 7 predicted stresses and responses (from the loads defined in Chapter 4 and combined as described in Chapter 5) are compared with the applicable code allowable values identified in Chapter 6 and the SSES design vill be assessed as adequate by virtue that the design capabilities exceed the stresses. or responses resulting from SRV discharge and/or LOCA loads.

2.2.1 Containment Structure and Reactor Building Assessment

. Summar h ~

2 2. 1.-1 - Containment-=Structure- Assessment Summary.

The primary containment walls, base slab, diaphragm slab, reactor pedestal and reactor shield are analyzed for the effects of SBV and LOCA in accordance with Table 5-1. The ANSYS finite element program is used for the dynamic analysis of structures.

Response spectra curves are developed at various locations vithin the containment structure to assess the adequacy of components.

Stress resultants due to dynamic loads are comb'ined with other loads in accordance with Table 5-1 to evaluate rebar and concrete stresses. Design safety margins are defined by comparing the actual concrete and rebar stresses at critical 'sections with the code allovable values. The assessment methodology of the containment structure is presented in Subsectio'n 7.1.1.1.

The results of the structural assessment of the containment structure are summarized in Appendix A. The results shov.that the reinforcing bar design stresses and the concrete design stresses are belov the allowable stresses.

2-2.1 Reactor Building- Assessment-Summary

~

The reactor building is assessed for the effects of SRV and LOCA loads in accordance with Table 5-1.

Containment basemat acceleration time histories are used to investigate the reactor building response to the SBV and LOCA loads. Response spectra curves at various reactor building elevations are qsed to assess the adequacy of components in the reactor building. The assessment methodology of the reactor building is presented in Subsections 7.1.1.2..  !

The results of the structural assessment of the reactor building are summarized in Appendix E. The results show that the reinforcing bars and concrete design stresses as veil as the structural steel design stresses are below the allowable stresses.

Rev. 2, 5/8O 2-5

2 2.2 Contalnment Sugmegged Structures Assessment Summary Design assessment of the suppression chamber columns includes non-hydrodynamic as well as hydrodynamic loads. Subsection 7.1.2.2 describes the methodoloqy used to evaluate the columns.

The results are presented in Figure A-59 and indicate a minimum desiqn margin of 11. 4%.

The downcomers are dynamically analyzed per Subsection 7.1.4 for the load combinations given in Table 5-3. A summary of the stresses under various load combinations are given in Figure A-66 and indicates that the minimum design margin is 14% when the loads are combined by ABS and 50% when the loads are combined by SRSS Results from the analysis of the suppression pool liner plate indicate that no structural modifications are required (see Subsection 7.1.3 and 7.2.1. 5).

The oriqinal downcomer and SRV bracing system has been redesigned so that the downcomers and SRV discharge lines are now supported by separate bracing systems. The SRV discharge lines are supported by bracinq connected to the columns, while the downcomers are braced together by a truss system, but no connections exist at the containment or pedestal wall.

Subsections 7. 1.2. 1 and 7. 1.2. 2 document the evaluation of the downcomer and SRV discharqe line bracinq systems, respectively.

Figure A-67 presents the SRV support system's maximum stresses and desiqn margins, while Pigures A-60 and A-61 shaw the design marqins for the downcomer bracing system members and connections, respectivel'y. All stresses are acceptable.

2,2 3 BOP apd NSSS Piping System assessment Summary All Seismic Category I BOP and NSSS piping are analyzed for the LOCA and SRV hydrodynamic loads and non-hydrodynamic loads per Subsections 7.1.5 and 7.1.6. 1. 1, respectively. F gives the'tresses and design margins. for selected BOP Appendix piping systems.

The stress reports for the above evaluation are available for NRC review.

All Seismic Category I BOP and NSSS equipment are evaluated for the hydrodynamic and non-hydrodynamic loads per the SSES Seismic Qualification Review Team (SQRT) Program. For each equipment Purchase Order, 4-paqe SQRT summary forms are prepared documenting the qualification results.

These SQRT summary forms are available for NRC review.

REV. 6, 4/82

2 5 2 5 ~5 IBlectrical Raceway Svstem-Assessment Summary.

Seismic Category I electrical raceway systems in the containment, reactor systems and control building are assessed by the methods contained in Subsection 7.1.8. Loads are combined as shown in Table 5-6. As a result of static and dynamic analysis, determined that high stresses resulted in certain members of a it vas fev support types. These structural members were strengthened or replaced by stronger members to reduce the stresses below the allov abl es.

2,2~6 HVAC Duct System Assessment Summary Seismic Cateqory I HVAC duct system in the containment, reactor buildinq and control building are assessed by the methods contained in Subsection 7. 1.9. Loads are combined as shown in Table 5-2. As a result of structural analysis, it vas found that a few structural members had hiqh stresses but most of the members had adequate margin of safety. The overstressed members vere strenqthened or replaced by stronger members to ensure an adequate margin of safety.

REV. 6, 4/82 2-7

1 1

iiI

CHAPTER 3 SRV DISCHARGE AND LOCA TRANSIENT DESCRIPTION'ABLE OF CONTENTS II 3 1 DESCRIPTION OP SAFETY RELIEF VALVE DISCHARGE 3.1 1 Causes of SRV Discharge 3 1 2 Description of the SRV Discharge Phenomena and SRV Loading Cases 3 2 DESCRIPTION OF .LOSS-OF-COOLANT ACCIDENT 3-2. 1 Small Break Accident {SBA) 3-2. 2 Intermediate Break Accident (IBA) 3 2.3 Design Basis Accident (DBA)

3 0 . SRV DISCHARGE AND LOCA TRANSIENT DESCRIPTION The purpose of this section is to provide a description of the SRV discharge and LOCA events.

A quantitative description of specific SRV and LOCA related loads for SSES is presented in Sections 4.1 and 4.2 respectively.

3-2

3 1 DESCRIPTION OF SAFETY- RELIEF VAJ VE DISCHARGE Susquehanna Unit 1 (and 2) is equipped with a safety relief system which condenses reactor steam in a suppression chamber pool. BV this arrangement, reactor steam is conducted to the wetwell via fast acting safety relief valves and quencher equipped discharge lines. This section discusses the causes of SRV discharge describes the SRV discharge process, and identifies the resultant SRV discharge actuation cases.

3,1.1 . causes of ssv Discharge.

During certain reactor operating transients, the SRVs may be actuated (by pressure, by electrical signal, or by for rapid relief of pressure in the reactor pressure operator'ction) vessel. The following reactor operating transients have been identified as those which may result in SRV actuation:

a. 'urbine generator trip (with bypass or without)

b. Nain steam line isolation valve (MSI V) closure.

c. Loss of condenser vacuum

d. Peedwater controller failure

e. Pressure regulator failure closed

f. Generator load rejection (with and without bypass) q., Loss of ac or auxiliary power

h. Loss of feedwater flow i Trip of two recirculation pumps Recirculation flow control failure decreasing flow

k. Inadvertent safety relief valve opening

1. Control Rod withdrawal error

m. Anticipated Transient Without Scram (ATWS}

A detailed description of these transients is provided in Section 15. 2 of the FSAR.

3. 1.2 Description of the SRV Discharge Phenomena and SRV Loading Cases l

Before an individual safety relief valve opens, the water level in the discharge line is approximately equal to the water level in the pool. As a valve opens, steam flows into the discharge Rev. 2, 5/80

line air space between the valve and the water column and mixes with the air (see detailed evaluation in Chapter 3 of Reference 1, pages 6-12 through 6-14) . Since the downstream portion of the discharge line contains a water slug and does not allow an immediate steam discharge into the pool, the pressure inside the line increases. The increased pressure expels the water slug from the SRV discharge line and quencher. The magnitude of the water clearing pressure is primarily influenced by the steam flow rate through the valve, the degree. to which entering steam is condensed along the discharge line walls, the volume of the discharge line airspace, and the volume of the water slug to be accelerated.

The clearing of water is followed by an expulsion of the enclosed air-steam volume. The exhausted gas forms an oscillating system with the surroundinq water, where the gas acts as the spring and the water acts as the mass. This oscillating system is the source of short term air clearing loads.

As the air-steam mixture oscillates in the pool it also rises because of buoyancy and eventually breaks through, the pool water surface at which time air clearing loads cease. Shen all the air leaves the safety relief system, steam flows into the suppression pool through the quencher holes and. condenses. The SSPS quencher design assures stable condensation even with elevated pool water temperature.

The SRV actuation cases resulting from the transients listed in Subsection 3. 1. are classified, as being one of the following 1

case

a. -

Symmetric (all valve, or AOT) discharge

b. Asymmetric discharge, including single valve discharge

c. Automa tic Depressurization System (ADS) discharge

~ Also considered in the containment design is the effect of subsequent SRV actuations (second-pop), discussed in Subsection 4,1.3. 6.

The symmetric discharge case [otherwise termed the all-valve, or abnormal operating transient (AoT) case] is classified as the type of SRU discharge that would follow rapid isolation of the vessel from the turbine such as turhine trip, closure of all NSIVs, loss of condenser vacuum, etc. As pressure builds up following isolation of the vessel, the SRVs act uate sequentially according to the pressure set points of the valves. This may or may not result i.n actuation of all the SRVs, but for conservatism in loading considerations all valves are assumed to actuate.

Refer to Subsection 4.1.3.1 for discussion of the loads resulting from this all-valve case.

Asymmetric discharge is defined as the firing of the SRVs for the three adjacent quencher devices which "esults in the greatest Rev. 2, 5/80 3-4

asymmetric pressure loading on the containment. This situation is hypothesized when, following a reactor scram and isolation of the vessel, decay heat raises vessel pressure so that lou set point valves actuate. If., during this tame of discharge of decay heat energy, manual actuation of the too other adjacent SRVs that comprise the asymmetric case is assumed, this actuation would result in the maximum asymmetric pressure load on the containment. Subsection 4. 1.3. 2 gives a discussion of the loads resultinq from the asymmetric discharge case.

The single valve discharge case is classified as the firing of the SRV Mhich gives the single largest hydrodynamic load.

Transients that could potentially initiate such a case are an inadvertent SRV discharge or Design Basis Accident (DBA). Refer to Subsection 3.2.3 for a discussion of the I atter possibility.

Subsection 4. 1. 3. 2. provides a discussion -:f the loads resulting 1

from the single valve ~ase.

The ADS discharge is defined as the simuLtaneous actuate'on of the six SRVs associated with the ADS. See Px.gure 1-4 for the location of the quencher devices assciciated sich the ADS valves.

The is assumed to actuate during an lntermedzate Break Accident (ZBA) or Small Break Accident (SBA) . lf an ADS ADS discharge is hypothesized coincident to an lBA or SBA (described in Subsections 3.2.2 and 3. 2. 1, respectively), the ef fects of an increased suppression pool temperature (resulting fromm steam condensation during the LOCA transient) and an< reased suppression chamber pressure (resulting from clearing of tne dryuell air into the pool durinq the transient) are cons idei.ea in uhe calculation of pressure loadinqs for the ADS discharge <<ase. bee Subse.:t ion 4.1.3.3 for f urther discussion oi the loads r e uk'. i ng i om the 1 A DS case.

Rev. 2, S/8p

3 2 DESCRIPTION OF LOSS-OF-COOLANT ACCIDENT This event involves the postulation of a spectrum of pi ping breaks inside the containment:aryinq in size type, and location of the break. For the analysis of hydrodynamic loadings on the containment, the postulated LOCA event is identified as a Sma11 Break Accident (SBA) ~ an Intermediate Break Accident (IBA) ~ or aDesiqn Basis Accident (DBA) .

3 2. 1 Small Break Accident SB~A This subsection discusses the containment transient associated vith small primary system blovdowns. The primary system ruptures ia this category are those ruptures that vill not result in reactor depressurization from either loss of reactor coolant or automatic operation of the ECCS equipment, ie, those ruptures vith a break size less than 0. 1 sq ft.

The follovinq sequence of events is assumed.to occur With the reactor and containment operating at the maximum normal conditions, a small break occurs that allows blovdovn of reactor steam or water to the dryvell. The resulting pressure increase in the dryvell leads to a high drywell pressure siqnal that scrams the reactor and activates the containment isolation system. The dryvell pressure continues to increase at a rate dependent upon the size of the steam leak. The pressure increase lowers the water level in the downcomers. At this time, air and steam enter the suppression pool at a rate dependent upon the size of the leak. Once all the dryvell air is carried ove'r to the suppression chamber, pressurization of the suppression chamber ceases and the system reaches an equilibrium condition.

The dryvell contains only superheated steam, and continued blovdovn of reactor steam condenses in the suppression pool. The principal loadinq condition in this case is the gradually increasing pressure in the dryvell and suppression pool chamber

.and the loads related to the condensation of steam at the end of the vents.

Bm 2.2 Zntetmediate B eak Accident ~IBA This subsection discusses the containment transient associated with intermediate primary system blovdovns. This classification-covers breaks for which the blovdown vill result in limited reactor depressurization and operation of the ECCS, ie, the break size is equal to or slightly greater than 0. 1 sq ft.

Follovinq the break, the dryvell pressure increases at approximately 1.0. psi/sec. This drywell pressure transient is sufficiently slow so that the dynamic effect of the water in the vents is negligible and the vents vill clear when the dryvell-to-suppression chamber differential pressure is equal to the hydrostatic pressure corresponding to the vent submerqence. The Rev. 2 5/80 3-6

resulting pressure increase in, the dryvell vill lead to a 'high drywell pressure signal that will scram the reactor and activate the containment isolation system. Approximately 5 seconds after the 0. 1 sq ft break occurs, air, steam, and water will start to flow -from the dryvell to the suppression pool; the steam vill be condensed, and the air will rise to the suppression chamber free space. The continual purging of dryvell air to the suppression chamber will result in a qradual pressurization of both the wetvell and dryvell. The ECCS will be initiated by the break and vill provide emergency cooling of the core The operation of these systems is such, that the reactor vill be depressurized in

..approximately 600 seconds This vill terminate the blowdovn phase of the transient. The principal loading condition in this case vill be the gradually increasing pressure in the drywell and suppression chamber and the loads related to the condensation of steam at the end of the vents.

3..2.3 D si n Basis Deci sent D~BD An occurrence of events which cou1d result in a DBA (instantaneous rupture of a main steam or recirculation line) is a remote possibility. Since such an accident provides an upper limit estimate to the resultant effects for this category of pipe breaks, it is evaluated without the causes being identified. For Susguehanna, an assumed instantaneous double-ended rupture of a recirculation line causes the maximum drywell pressure and therefore .the governing LOCA hydrodynamic loads.

The sequence of events immediately fo11oving the rupture of a recirculation line has been determined. A drywell high pressure siqnal is almost instantaneously sensed, initiating a scram and containment isolation and siqnaling the HPCI, CS and LPCI to, start. The flow in both sides of the break vill accelerate to the maximum alloved by the critical flow considerations. In the side adjacent to the suction nozzle, the flow vill correspond to critical flow in the pipe cross-section. In the side adjacent to the injection nozzle, the flov will correspond to critical flov at the, 10 jet pump nozzles associated vith the broken loop. In addition the cleanup line cross-tie will add to the critical flov area..- This high rate of flow out of the ruptured recirculation line results .in a dryvell pressure rise of approximately 44 psiq in 14.5 seconds (refer to FSAR Table 6.2-5 and FSAR Fiqure 6.2-2)

This rapid increase in drywell pressure accelerates the water initially in the containment vent system out through the vents.

Immediately follovinq vent water clearing, an air/steam bubbles start to form at the dovncomer exits. Initially, the bubble pressure is essentially equal to the current dryvell pressure.

As the flow of air/steam from the drywell becomes established in the vent system, the initial vent exit bubble expands, thus accelerating upward the suppression pool water above the vent exits. The steam fraction of the flow is condensed, but continued injection of drywell air and expansion of the air 3-7

bubble results in a.rapid rise in the suppression pool surface knovn as pool swell.

Followinq the pool svell and fallback, there is a period of high-steam flov rate through the containment vent system. For large primary system ruptures, reactor blovdovn and, therefore, vent

'team co'ndensation last for approximately 60 seconds.

Shortly af ter a DBA, the ZCCS pumps (HPCX, CS, and LPCI) automatically start pumping condensate storage tank water or suppression pool vater into the reactor. pressure vessel. Mithin 40 seconds all the ECCS pumps are at rated flov. This floods the

,reactor core until water starts to cascade into the drywell break. The time at which this occurs would depend upon break from'he size and location. Because the dryvell would be full oX steam at the time of vessel floodinq, the sudden introduction of cold vater causes steam condensation and drywell depressurization.

When the -dryvell pressure falls below the suppression chamber pressure, the drywell vacuum relief system is actuated and air from the suppression chamber enters the dryvell. Eventually, sufficient air returns-to the drywell to equalize the pressures.

Similarly, small differential pressures between the dryvell and the suppression chamber can be produced if the containment spray system is actuated condensing steam in the dryvell Follovi.nq the vessel flooding and drywell/suppression chamber pressure equalization phase of the accident, suppression .pool water vill be continuously recirculated through the core by the ECCS pumps. The energy associated with the core decay heat will result in a slov= heatup of the suppression pool. The suppression pool temperature is controlled by the RHR heat exchangers The capacity of these heat exchangers is such that the maximum suppression pool temperature .increase is reached after several hours. The suppression pool can experience a peak temperature of

-approximately 200~F under vorst case conditions. The post LOCA containment heatup and pressurization transient is terminated vhen the RHR heat exchanqers reduce the pool temperature and containment pressure to nominal values.

The primary loads on the containment generated by a DBA are the pressure build-ups in the drywell and suppression chamber, and the loads resultinq from the various modes of steam condensation at the vent ends. The hiqh rate of system depressurization resultinq from a DBA militates against the .firing of an SRV; however, for conservatism a single SRV discharge is considered coincident with the DBA for containment structural loading purposes.

CHAPTER 4 LOAD DEFINITION TABLE OF CONTENTS 4.1 SRV Loads {See Proprietary Section),

4 2.1 LOCA Loads Associated With Poolswell 4.2 1 1 Metvell/Drywell Pressures During Poolsvell 4.2 1 2 Submerged Boundary Loads During Vent Clearing 4.2.1. 3 Dovncomer Mater Jet Load 4 2 1.4 Poolsvell Air Bubble Load 4.2.1.5 Poolsvell Asymmetric Air Bubble Load 4.2.1 6 Poolsvell Impact Load 4.2.1. 7 LOCA Air Bubble Submerged Structure Load 4.2.1.8 Poolswell Drag Load 4.2 1.9 Poolsvell Fallback Load 4 2.2 Condensation Oscillations and Chugging Loads 4 2 2.1 Containment Boundary Loads During Condensation Oscillations 4.2.2.2 Pool Boundary Loads Due to Chugging

4. 2.2. 3 Single Vent Lateral Load 4 2.2. 4 Multivent Lateral Load 4.2.2. 5 Submerged Structure Loads Due To Condensation Oscillations and Chugging 4.2.3 'esponse to NRC Criteria for Loads on Submerged Structures 4.2.3.1 Introduction

4. 2.3.2 NRC Criteria III.D. 2. a.l Bubble Asymmetry

4. 2.3.3 NRC Criteria III.D.2. a.2 Standard Drag In Accelerating Flow

4. 2.3.-4 NRC Criteria III.D.2.a;3 Segmentation of Structures 4;-2.3. 5 NRC Criteria III.D.2. a.4 Interference Effects 4;-2.3. 6 NRC Criteria III. D. 2. a. 5 Blockage In Downcomer Bracing 4.2 3. 7 NRC Criteria III.D.2. a.6 Formula 2-23 of Reference 13 4.2 4 Secondary Loads

4. 2.4. 1 Downcomer Friction Drag Loads

4. 2.4. 2 Sonic Waves 4.2.4.3 Compressive Wave 4 2 4.4 Fallback Loads on Submerged Boundaries 4.2.4. 5 Vent Clearing Loads on the Downcomers 4 2.4.6 Post, Poolswell Waves 4 2 4.7 Seismic Slosh 4 2.4 8 Thrust Loads 4.2.5 Long Term LOCA Load Definition 4.2 51 Design Basis Accident {DBA) Transient 4.2.5.2 Intermediate Break Accident (IBA) Transients Rev. 2, 5/80 4-1

4. 2. 5. 3 Small Break Accident (SBA) Transients 4.2.6 LOCA Loading Histories for SSFS Containment Components 4 2 6 1 LOCA Loads on the Containment Wall and Pedestal 4 2 6 2 LOCA Loads on the Basemat and Liner Plate 4- 2.6- 3 LOCA Loads on the Dryvell and Drywell Floor 4.2.6 4 LOCA I.oads on the Columns 4 2.6. 5 LOCA Loads on the Dovncomers 4.2 6 6 LOCA Loads on the Dovncomer Bracing 4.2 6.7 LOCA Loads on Wetvell Piping 4.3 Annulus Pressurization 4.4 Fiqures 4.5 Tables Rev. 2, 5/80 4-2

CHAPTER 4 gIGUQQS Humber Tg~t e These figures are proprietary and are found in the through proprietary supplement to this DAR.

4-37 4-38 SSES Short Term Suppression Pool Height 4-39 SSES Short Term Qetwell Pressure 4-40 SSES Pool Surface Velocity vs Elevation 4-40a Poolswell Acceleration Time History 4-41 Pool Boundary Load During Vent Clearing 4-42 This Figure has been Deleted 4-43 SSES Poolswell Air Bubble Pressure 4-44 Poolswell Air Bubble Pressure on Suppression Pool Mails Used, for SSES Analysis i 4-44a Condensation Pressure Porcinq Function (Met 6 Dry Mells)

(This fiqure has been deleted)

Symmetric and Asymmetric Spatial Loading Specification (This fiqure has been deleted) 4-46 SSES Drywell Pressure Response to DBA LOCA 4-47 SSES Metwell Pressure Response to DBA LOCA 4-48 SSES Suppression Pool Temperature Response to DBA LOCA 4 49 SSES - Drywell Temperature Response to DBA LOCA 4-50 SSES Suppression Pool Temperature Response to IBA 4- 51 SSES Plant Unique Containment Response to the IBA 4-52 Typical Mark II Containment Response to the SBA 4-53 SSES Components Affected by LOCA Loads 4-54 SSES Components Affected by LOCA Loads REV. 6, 4/82 4-3

gIG~U@S {Cont. )

Number Title 4-55 LOCA Loading History for the SSES Containment Wall and Pedestal 4-56 .LOCA Loading History for the SSES Basemat and Liner Plate 4-57 LOCA Loading History for the SSES Drywell and Drywell Floor 4-58 LOCA Loading History for the SSES Columns 4-59 LOCA Loading History for the SSES Downcomers 4-60 LOCA Loadinq History for the SSES Downcomer Bracing System 4-61 LOCA Loading History for SSES Wetwell Piping

6) 4-62,a-f Chugging Pool Boundary Loads {These figures have been deleted) 4-62,q6h Dynamic Downcomer Lateral Loads Due to Chugging 4-62,i-m Typical Wave Notion Due to Seismic Slosh 4-63 These Figures are Proprietary thru 4-66 REV. 6, 4/82

CHAPTER 4 TgBLgS ~

gumQer ~ g~te 4-1 These tables are proprietary and are found thru in the proprietary supplement to this DAR 4-15 4-16 LOCA Loads Associated with Poolsvell 4-17 SSES Dryvell Pressure 4-18 SSES Plant Unique Poolsvell Code Input Data 4-19 Input Data for SSES LOCA Transients 4-20 Component LOCA Load Chart for SSES 4-21 Wetvell Piping LOCA Loading Situations 4-22 Seismic Slosh Wave Height REV. 6, 4/82

4 0 LOAD DEPXNXTION See the Proprietary Supplement for this section Rev. 2, '5/80 -'4=6

4 '- ~ >> LOCA LOAD DEPINITIOH

~

Subsections 4.2.1, 4.2.2 and 4 2.3 discuss the numerical definition of loads resulting from a LOCA in the SSES containment. The LOCA loaves are divided into five groups. l2 (1) Short term LOCA loads associated with poolswell (Subsection, 4.2.1) .

(2) Condensation oscillations and chugging loads (Subsection 4.2.2) .

(3) Submerged Structures Loads (Subsection 4. 2. 3)

(4) Secondary Loads (Subsection 4.2.4) .

(5) Long term LOCA loads (Subsection 4. 2. 5) .

The application of these loads to the various components and structures in the SSES containment is discussed in Subsection 4.2.6. )2

4. 2. 1 LOCA LOADS ASSOCX ATBD HITH POOLSMBLL A description of the LOCA/Poolswell transient is given in Section 3.2.3 of this Desiqn Assessment Report. The LOCA loads associated with poolswell are listed in Table 4-16..A discussion of these loads and their SSES unique values follows.

4.2.1 1 Qetwelg/D~gwell P~essu~es during Pool@well The drywell pressure transient used for the poolswell portion of the LOCA transient (< 2.0 sec) is qiven in Table XV-D-3 of Reference 7., A portion of this table is reproduced herein -as Table 4-17. This drywell pressure transient includes effects of pipe inventory and reactor subcooling and is the'lowdown the highest possible drywell pressure case for poolswell. This drywell pressure transient is calculated using the method 2 documented in Reference 56.

The short term poolswell wetwell pressure transient resulting from this drywell pressure transient is calculated by applying the poolswell model contained in Reference 8. The equations and assumptions in the poolswell model were coded into a Bechtel computer proqram and verified against the Class 1, 2 and 3 test cases contained in Reference 9. This verification is documented in Appendix D to this report. Inputs used for the calculation of the SSES plant unique poolswell transient are shown in Table 4-

18. The short term wetwell pressure transient calculated with the poolswell code is shown in Piqure 4-39. The short term wetwell pressure peak is 56. 1 psia (41.'4 psig).

Reference 46, Subsection XII.B.3.d. 2 formulates a methodology for, determininq the maximum diaphraqm uplift P to be used for design assessment. This, hP is based on following relation:

hPUP = 8.2 44!F (PSI) 0<F< O.l3 hPUP ~ 2.5 (PSE) F>0.13 RGK. 2, 5/80 AB AP VS V~D(AV) 4-7

where: AB break area:

AP net pool area; AV total vent area VS initial wetwell air space volume; and VD drywell volume For SSES (see Tables 4-18 and 4-19):

AB 3 53 ft>

AP 5065 ~ 03 ftz AV 257.52 ft2 VS t49,000 ft~

UD 239,600 ft>

Inserting into the above equation yields:

F=0168>013 This qives a maximum upli'ft hP of 2.5 PSXD. However, as required by NUREG 0808, a more conservative uplift BP of 5.5 PSID will be used for desiqn.

4. 2.1. 2 Submerged Bougdagg. Loads Du~rin Uent Clearing The submerged get formed by the expulsion of the water leg in the downcomers creates a vent clearing load on the basemat and on the submerged wetwell walls. This loading is defined by Reference 57 as a 24 PSI overpressure statically applied with hydrostatic pressure to surfaces below vent exit with a linear attentuation to zero at pool surface (see Figure 4-41) . This load is applied during the vent clearing.

The NRC, in Supplement No. 1 to NURE6-0487, accepts the above 24 PSI overpressure for the vent clearing load for those plants where (mhL)/f (AP /A ) UD~ ] 55 with m = mass flow in vents -lb/sec VDg= drywell Volume ft~

h = enthalpy of air in vents btu/lb L = submergence

/AU= pool area ft to vent area ratio Ap For SSES, the various parameters are:

m = 17,900 lb/sec UDg= 239, 850 f t~

h = 194 btu/lb L = 12 ft Ap /Av= 5065/257 Substituting into the above gives:

f (17,900) (194) (12) {257) ]/f (5065) (239 850) ) = 8 8 REV. 6, P/82 4-8

Thus, for SSES, the 24 PSI overpressure specified for the air clearing load is acceptable.

4. 2. 1. 3 LOCA - Jet Loads

~

During the vent clearinq stage induced velocity and acceleration fields are created in the suppression pool producing drag forces on submerqed strctures. The oriqinal methodology employed to predict the drag forces is contained in Reference 12 {often called the Hoody get model) and is an analytical representation of an unsteady water jet lischarqinq into a suppression pool.

The get is made up of constant velocity fluid particles traveling at the speed at which they exited the discharge pipe. The jet front is described as the locus of points which a particle overtakes the one exiting immediately before it. No velocities or accelerations are defined in the fluid external to the jet-Reference 46, subsection III.D.1.a proposed that velocity and acceleration be predicted throughout the pool using the potential function of a sphere at the jet f rout. A molification of the load calculated at get impinqement was also required. The Acceptance Criteria was a simple method to determine a bounding get load for all structures below the downcomer exits.

The Hoody ]et model was clearly derived for gets with constant or linearly increasing acceleration. However, the vent clearing transients predicted for Hark II plants typically have an acceleration, increase greater than linear. Strict applicaton of Reference 12 leads to unrealistic mathematicl results. Two

'interpretations of the results are possible depending upon the time base employed. Examining the get in>>real time>> (t in Reference 12) a jet can be seen with two independent fronts traveling at different speeds at different locations which coincide only at the point of jet dissipation. On the other hand, if we use the "exit time>> ( v ) as a basis the jet reverses and moves backward in both space an'd "real time>> before dissipation. Clearly neither of these observations is of much use in calculating loads on structures.

To overcome the difficulties of using this model, an alternative methodology has been formulated. The jet front will be described by the motion of the particle having travelled the farthest at any instant in time. This will be identical to the Hoody jet motion for jets with linearly increasing acceler'ation but will yield a single continuous velocity and acceleration time history even if the acceleration increases more rapidly.,

A sphere is then placed at the jet front generating a potential flow described by the following function:

4

-3

- ,8'0 U.V coso jw r where r and 8. are the spherical coordinates from the sphere center to some position in the suppression pool with 8 measured REV. 6, 4/82 4-9

from the get direction, U. is the velocity of the sphere determined by the veloci+ of the particle having traveled the farthest at the instant in time the drag forces are being computed and V~ is the initial volume of water in the vent.

The local velocity U'., and acceleration, Uare then calculated f rom the above relation by the methods of Ref erence 14. the local velocity and acceleration are known the drag forces Once are computed from Reference 13 as follows:

F U v A

g CAUDx ~n'p 8 .

2g C whe;e F> is the acceleration drag, U n is the local,accelera tinn field normal to the structure, > is the acceleration drag volume for flow normal to the structure, p is the fluid density, ~ is the standa"d draq, <D is the drag coefficient for flow norma% --;.o the structure, A is the projected structure area normal to U n and U is the local velocity field normal to the stzuctu: e~

When the ]et is predicted to dissipate the sphere is traveling at the final jet velocity at the point of maximum jet penetration.

This condition is used as the final load calculation point. The final get velocity is that of the get front gust before the last particle leaving the vent reaches the jet front. The velocity of the last particle is disregarded.

4. 2. 1.4 Boundary Loads Durina Poolswell During the poolswell transient, the high pressure air bubble which forms in the vicintiy of the vent exit creates an increase in pressure on all suppression pool boundaries below the vent exit as well as those walls which it is in direct contact.

Boundaries which are above the bubble location and up to the point of maximum pool elevation also experience increased pressure loads corresponding to the increased pressure in the wetwell airspace as well as the hydrostatic contribution of the water slag.

Ref erence 46, Subsection III.B. 3. b methodology for specif ication of these loads uses the Poolswell Analytical Nodel to determine the maximum values of bubble pressure and wetwell airspace

,pressure. The analysis takes the maximum pool elevation as 1.5 times the initial submergence. Using this data, a static loading is applied to the containment structure as follows:

1. for the basemat uniform pressure equal to the maximum bubble pressure superimposed on the hydrostatic load correspondinq to a submergence from vent exit to the basemat; Rev. 2, 5/80 4-10

II

2. for the containment walls below vent exit maximum bubble pressure plus hydrostatic head corresponding to vertical distance from vent exit:

3. for the containment ~alls between vent exit and maximum pool elevation-linear variation between maximum hubble pressure and maximum wetwell airspace pressure;

4. for the containment walls above maximum pool elevation maximum wetwell airspace pressure.

The pressure distribution used for the SSES analysis is shown in Figure 4-44.

4~2.1.5 Poolswegl As~mmetrig Ai~pubb1e Load The methodology used in the proceeding subsection assumes that the air flow rate in each downcomer is equal leading to a symmetric loadinq of the containment boundary. Reference 46 has expressed concern that circumferential variations in the downcomer air flow rate can occur due to dyrwell air/steam mixture variation that would result in variations in the bubble pressure load on the wetwell wall.

This loading condition's calculated by statically applying the maximum air bubble pressure obtained from the PSALM to 1/2 of the submerged boundary and statically applying 120% of the maximum bubble pressure to the other 1/2 of the" submerged boundary. The pressure load on the basemat and wetwell walls below the vent exit is the sum of the air pressure and the hydrostatic pressure.,

For the portion of the wall above the vent exit, the pressure increase due to the air bubble is linearly attenuated, from the bubble pressure at the vent exit to zero at the pool surface.

This increase is then added to the local hydrostatic pressure to obtain the total pressure. The time period of application of the load is from the termination of vent clearing until the maximum swell height is reached.

4. 2. 1. 6 Pop lswell Impact Load-Any structure located between the initial suppression pool surface (Bl. 672 ) and the peak poolswell height (El. 690'-2",

~

see Figure 4-38) is subject to the pool swell impact load. As documented in the response to NRC Question 020.68 the poolswell

~

maximum elevation is determined hy the poolswell Analytical Hodel with a polytropic exponent of 1.2 for wetwell air compression to a maximum swell heiqht which is the greater of 1.5 vent submergence or the elevation corresponding to the drywell floor uplift* hP determined from the equation documented in Subsection

4. 2.1.1 (2.5 PSlD) . For SSZS, usinq the design drywell floor uplift BP=2.5 PSID leads to the greatest poolswell height and yields 1.51 times the initial vent submergence., Since all grating is removable only "small" structures as defined in Referenc'e 10a, Subsection 4.2. 5. 1 are subject.to poolswell impact loads.

REV. 6, P/82 4-1 1

Poolswell impact loads of <<small<< structures are determined as specified in Reference 46, Subsection III.B.3.c.1. An SSES plant-unique velocity vs. elevation curve has been generated with the poolswell model {see Figure 4-40) . The velocity curve is conservatively increased by a 1.1 multiplier and used to calculate the impulse per unit area, pulse duration and maximum impact pressure at the component~ s elevation. The peak pressure is then used to define a versed sine shaped hydrodynamic loading function P

-max (1-cos2'9t/T)

(

2 where: P = pressure acting on the projected area of the structure; Pm~ = the temporal maximum of pressure acting on the projected area of the structure; t = time;

~ = duration of impact The loadinq function corresponds to impact on rigid structures.

In actuality, the structures being analyzed may be more flexible, resultinq in the pressure pulses, during impact, being modified by the motion of the structure. To account for this, the hydrodynamic mass of impact is added to the mass of the impacted structure vhen performinq the structural dynamic analysis.

During the dryvell air purge phase of a LOCA, an expanding bubble is created at the dovncomer exits. These rapidly expanding bubbles eventually coalesce into a <<blanket<<of air vhich leads to the pool swell phenomena. The bubble charging process creates fluid motion in the suppression pool which causes drag loads on the submerged structures.

The submerged structure draq loads due to air clearing, prior to pool swell, are'alculated in the same manner as the drag loads due to CO and chuqqinq presented in Subsection 4.2.2.5. However, the chugqing and CO sources are replaced with a source representinq the bubble qrovth prior to pool svell. This source is derived from the oriqinal 4T data. All sources are assumed in-phase {87 sources) .

4. 2. 1. 8 ~ Pools we+1 Dr~a Load.

Subsequent to bubble contact all bubbles are assumed to coalesce into a blanket of air.and the poolswell drag loads are due the rapidly accelerating upward slug of vater and acts in the vertical direction only {except for lift forces which act in the traverse direction to flow). The one dimensional pool svell model is used to predict the vertical flov field. Once the flow field is known the drag forces are calculated by the methods of Reference 13 modified by the methodology presented in Subsection REv. 6, 4/82 4-1 2

4.2.3. This load applies to any structure located between the elevation of the vent exit and the peak poolswell height. The duration of the drag load begins vhen the vent clears except for structures vhich are oriqinally not submerged. Por structures which are not submerged,, the drag load duration is based on the slug transient time (Reference 10a, page 4-78, step 3).

4. 2. 1.9- Poolgwegg Fa+lbgck Load-After the termination of poolswell the slug of water falls under the influence of gravity causing drag forces on structures lcoated between the peak poolsvell height and the vent exit. The motion of the vater is described by 2 the following equations:

H(t) = H max gt /2 v (t) = gt O'PB(t) = q where q is the acceleration constant, H(t) is the height above initial water level at time t, ~x is the maximum svell height, and t is time starting with t = 0 at maximum swell height The drag load is then calculated from the methods of Reference 13 modified by Subsection 4.2.3 of the DAR. The loading stops when 8 (t) has fallen below the structure or when H (t) has returned to normal vater level vhichever is calculated to occur first.

4.2.2 Condensation gscilJatjggs agd~Ch gqing Loads Condensation oscillation and chugging loads follow the poolswell loads in time. There are basically three loads in this secondary time period, i.e. ~ from about 4 to-60 seconds after the break.

"Condensation oscillation< is broken dovn into tvo phenomena, a mixed flow regime and a steam flov regime. The mixed flow regime is a relatively high mass flux phenomenon which occurs during the final period of air purging from the drywell to the vetwell when the mixed flov throuqh the downcomer vents contains some air as well as steam. The steam flow portion of the condensation oscillation phenomena occurs after all the air has been carried over to the wetwell and a relatively high intermediate mass flux of pure steam flow is-established.

"Chuqqing" is a pulsating condensation phenomenon vhich can occur either'follovinq the intermediate mass flux phase of a LOCA, or during the class of smaller postulated pipe breaks that result in steam flow through the vent system into the suppression pool. A necessary condition for chugging to occur is that only pure steam flovs from the L'OCA vents. Chuqginq imparts a loading condition to the suppression pool boundary and all submerged structures.

In Revision 2 of the DAR we stated that the DPFR CO andchugging

~

steam condensation boundary load definition (see Appendix A to Reference 21 and Reference 16) would be compared with the LOCA steam condensation load definition derived from the GKN II-5 test .

data to evaluate the conservatism of the DPPR load. Subsections 9.6. 1.1 and 9.6. 1.2 document this comparison.

Rev. 2, 5/80 4-'l3

As a result of this comparison and the possible schedule delays associated with licensing SSES based on the DPPR load, PPSL decided on April 1, 1982 to terminate the re-evaluation of SSES based on the DPFR load and re-assess SSES with the GKN II-H load definition.. Subsection 9.5.3 documents the GKN II-N load definition. For chugging, both a symmetric and asymmetric load case are consi.dered, while for CO, only a symmetric load- case is considered.

For plant evaluation, PPSL does not define a separate CO and chugging load definition, as with the Nark II Owners.

the acceleration response spectra (ARS) generated for the LOCA Instead, steam condensation phenomena for combination with the other dynamic loads (i.e., SRV (ADS) ~ seismic, etc.) is the so-called LOCA load, which represents an envelope of the ARS curves generated for both the GKM-IIH CO and chugging load definition, and symmetric and asymmetric load cases (see Subsection 9.6.1.1).

Subsection 7.0 provi'des the results of the re-evaluation .of the SSES plant to the LOCA steam condensation load derived from the GKM-IIH test data.

4.2.2. 1 Containmegt Boundary L~ods Due To Condensation Oscillations This subsection has been deleted.

4.2.2.2 . Pool- Bounda~Loads Due to C~hu ging This subsection has been deleted.

4. 2.2. 3 Downcomer Lateral goads The chugging load imparted to the downcomer is taken from Reference 47. This reference specifies two sinusoidal dynamic loads used when evaluating downcomer lateral bracing systems.

The durations and amplitudes specified are 3ms, 30 kip and 6 ms, 10 kip {as shown in Pigures 4-62G S H).

However, in response to the NRC's concerns with the Mark II single vent lateral load, SSES is re-evaluating the downcomers with an extrapolated single vent lateral load of 65 Kips and 3 msec time duration for faulted conditions. Subsection 9.6.3 verifies the conservatism of this load based on a statistical analysis of the GKM II-M bracing force data at 10-5 exceedance probability.

Nultivent lateral loads due to chugging are presently being evaluated by the methodology documented in letter report <<Method of Applying Nark II Single Vent Dynamic Lateral Load to Mark II Plants with Multiple Vents," transmitted to the NRC on April 9, 1980 under Task A.13.

4-14

2.2.5 -

Submerged Structure Loads Due to Condensation Osc llation and Chu 99 in 9 Condensation Oscillation and chugging induce flows fields in the suppression pool causing draq loads on the submerged structures (i.e., SRV lines, downcomers, etc.) . The methodology for calculating these draq loads to be combined with the other design basis loads is presented below.

The force on a submerqed structure is the sum of an acceleration force F~ and an unsteady drag force FD

'+

FD Hnder certain conditions the pressure gradient is of sufficient magnitude so that the submerged structure force is essentially the acceleration drag force. In order for this to be true, the Strouqhal Number must be sufficiently large.

I For the SSES submerged structures and the flow fields induced by chuqginq and CO, the Stroughal Number is sufficiently high that negligible error will be incurred by ignoring the unsteady drag force.

The submerqed structure drag force can be approximated by the integral of the pressure field P@ over the structure surface:

F p@QS'K where: P@

= determined by the equations for potential flow K = hydrodynamic mass factor For a linear isentropic fluid where the velocity is everywhere small compared to the sonic speed c, the equations for potential flow reduce to the acoustic wave equation (Reference 65). Thus, the pressure field alsp satisfies the acoustic wave equation.

Thus, for calculatinq the SSES submerqed structure drag load due to CO and chugging, the above expression is used, sith the pressure P@ , as a function of time and position, calculated by the IQEGS/MARS acoustic model of the SSES suppression pool. The pressure P< is calculated in an analagous manner as the svmmetric wall loads (see Subsection 9.5.3.4. 1) for each source, except that the pressures are~calculated at the submerged structure surface locations instead of the containment boundary.

For each structure beinq analyzed (i.e., column) a pressure time history (PTH) is calculated for every 60o increment circumferential around the structure at each elevation correspondinq to a nodal point of the structural model. Thus, for each node point elevation, six pressure time histories are calculated. )his is repeated for each source. These sets of PTHs, calculated for each source, are then integrated across the structure~s surface to give resultant force time histories for structural analysis.

REV+ 6, 4/82

The force time histories are then multiplied by a hydrodynamic mass factor, K, of 2 to account for the modification of the flow field due to structure's presence.

4.2.3 Response to NRC Criteria for Loads On Submerged Structure

4. 2. 3. 1 Introduction In October 1978 the NRC published NUREG-0487, Nark II Containment Lead Plant Program Load Evaluation and Acceptance Criteria. It addresses the load methodologies proposed by the Mark II Lead Plant Program for determining LOCA and SRV hydrodynamic loads.

NURFG-0487 was highlv critical of the lead plant position for determining submerged structure loads and stipulated very conservative alternative loading criteria. The following subsections will present the NRC submerged structures acceptance criteria and the corresponding Nark II response.

4 2.3.2 NRC Critegia III,D.2~a~1- Bubble A~smmetrg A conservative estimate of asymmetry should be added'y increasing acceleration and velocities computed in Step 12 of Section 2.2 of Reference 13 by 10%. If the alternative steps 5A, 12A, and 13A are used the acceleration drag shall be directly .

increased by 10% while the standard drag shall be increased by 20%

Response: These criteria are acceptable.

4-. 2.3 3 NRC. Criteria III.D.2 a.2: . Standard Dr~a In-Accelerating Plow-

~

The drag coefficients C for the standard drag contribution in steps 13, or 13A, 15 of section 2.2 and step 3 of section 2.3 of Reference 13 may not be taken directly from the steady state coefficients of Table 2-3. Hodified coefficients C from accelerating flow as presented in References 49 and 50 shall be used with transverse forces included, or an upper bound of a factor of three times the standard drag coefficients shall be used for structures with no sharp corners or with streamwise dimensions at least twice the width.

response:

4 The three references show that in oscillating flows the standard drag coefficient for cylinders can exceed the steady flow value.

Values of C in excess of 2.0 were observed while steady state values (for cylinders) never exceed 1. 2. The NBC's position is

. interpreted to mean that neglecting the unsteady effect on standard drag coefficients will be nonconservative in some cases.

method is presented in Reference 51, Appendix A to account for unsteady effects on standard and acceleration drag during various phases of the LOCA and SRV transients. Also included are methods to estimate transverse forces due to vortex shedding.

REV. 6, 4/82 4-16

Subsequent to reviewing the methodology contained in Appendix A of Reference 51, the NRC in Supplement No. 1 of NUREG.-0487, required several modifications to the methodology for determining the unsteady drag coefficients.

A review of the SSBS pool swell "and fallback drag load calculations indicates that SSES has incorporated 'these modifications into their calculations. Drag coefficients- are not required for calculating the submerged structure drag loads due to air bubble charging prior to pool swell, and the drag loads due to chuqqing and CO, since these loads are calculated using the pressure time histories at the structure locations (see Subsection 4. 2. 1.7 and 4. 2. 2. 5).

4.2.3.4 NRC Criteria 'XXX.D.2.a 3: S~ementat jon of St~uctugeg

~ ~

The eguivalent uniform flow velocity and acceleration for any structure or structural segment shall be taken as the maximum values "seen" by that structure, not the value at the geometric center.

Response

For structures submerged in a non-uniform flow field, the velocity and acceleration will be a function of position along the structure. The NRC's criterion is interpreted to mean that the velocity and acceleration should be taken at the,end of the segment closest to the disturbing source instead of the geometric center. For. certain restrictions on segment length, 'the error in the calculation of draq using the velocity and acceleration at the geometric center is very small. This is demonstrated for acceleration drag in Reference 51, Appendix B and -for standard drag Reference 51, Appendix C. Appendix' also contains a discussion that shows that neglecting end effects in drag calculations is conservative.

I 14 The computation of drag forces on submerged structures independent of each other {as presented in Reference 13) is adequate for structures sufficiently far from each, other so that interference effects are negligible. Interference effects can be expected to be insignificant when two structures are separated by more than three characteristic dimensions of the larger one. Por structures closer together than this separation, either detailed-analysis of interference effects shall be performed or a conservative multiplication of both the acceleration and standard drag forces by four shall be performed.

REV. 6, 4/82 4-17

Response

Interference effects can 'have a significant effect on drag forces. A modification to the calculational procedure is proposed to account for interference. Reference 51, Appendix D describes the proposed me'thod for standard drag vith the exception that the free stream velocity used vill be that at the structures geometric center in all cases. Reference 51, Appendix presents the proposed method for acceleration drag.

4 2,3 6 NRC Cgjtegga III D.)~a~5: ~ Blockage In Dovncomer Bracinq A specific example of interference which must be accounted for is the blockage presented to the motion of the water slug during pool swell due to the presence of downcomer bracing systems. Xf significant blockaqe relative to the net pool area exists, the standard drag coefficients shall be modified for this effect by conventional methods (Reference 52) .

Response

Blockage effects on the pool swell drag loads produced on the dovncomer bracing system vere accounted for by using the methods in Reference 87.

4,2,3,7 pRC CrjteRiR III~.2.3.6 .~oreulR 2-23 of Reference 13 Pormula 2-23 of 'Reference 13 shall be modified by replacing MH with ~PB Q where yA is obtained from Table 2-1 and 2-2. This is then consistent'vith the analysis of Reference 14.

Response

This criteria is acceptable.

Seconda~ Load The previous subsections have identified and specified loading methodologies that result in significant containment dynamic loads. In addition, several pool dynamic loads can occur which are considered secondary when compared to the previous loads or because the containment and related equipment response is small when sub)ected to them. The following subsections identify the secondary loads and the load criteria to be applied to the SSZS containment.

4. 2.4. 1 ~ Dovncomer Friction Dyad Loads Priction Drag loads are experienced internally by the dovncomers during vent clearing and subsequent air/or steam flow. In addition, the dovncomers experience an external drag load during poolsvell. Usinq standard drag force, calculation procedures these loads are determined to be 0.6 and .3 KIPS per downcomer, respectively and are not considered in the structural evaluation of the containment.

4-1 8

4. 2.4. 2 Sonic. Qaves-Immediately following the postulated instantaneous rupture of a large primary system pipe, a sonic wave front is created at the break location and propaqates through the drywell to the vent system. This load has been determined to be negligible and none is specified.

4.2.4.3 Compressive ~H ye The compression of the air in the drywell and vent system causes a compressive wave to be generated in the downcomer water legs.

This compressive wave then propaqates through the pool and causes a differential pressure loadinq on the submerged. structures and on the wetwell wall. This load has been evaluated and is considered neqliqible.

4,2.4.4 Pallback Loads on Subm~e @ed Boundarj,es During .fallback "water hammer" type loads could exist water slug remained intact during this phase.

if the However available test data indicates that this does not occur and the fallback process consists of a relatively gradual settling of the pool water to its initial level as the air bubble <<percolates<<upward.

This is based on visual observations durinq the EPRE tests (Reference 32) as well as indirect evidence provided by a careful examination of pool bottom pressure forces from the 4T, EPRX, foreiqn licensee and Narviken tests. Thus these loads are small.

and will not be considered.

The expulsion of the water leq in the downcomers at vent clearing creates a'ransient water get in the suppression pool. This get formation may occur asymmetrically leading to lateral reaction loads on the downcomer. However, this load is bounded by the load specification durinq chugging an'd will not be considered for containment analysis.

4.2.4.6 Post.Poolgwe+1 Qgyes Reference 46 indicates the potential for containment loading due to post p'oolswell waves impinqinq on the wetwell wall and internal components. Per the response'o Question 8020.8 documented in Appendix A to Reference 10a, this load is considered negligible when compared to the other design basis loads.

4.2.4.7 Seismic Slosh Seismic slosh loads are defined as those hydrodynamic loads exerted on the suppression pool walls by water in the suppression pool durinq a seismic event. Although these loads are expected to be small in comparison with other hydrodynamic loads such as those associated with air/steam SRV discharqe and LOCA poolswell 4-19

and steam condensation loads, they have been calculated for the SSES containment evaluation,'s reguested by the NRC in NUREG-0487 The methodology used to calculate seismic slosh loads for the SSES containment is the SOLA-3D computer code, developed at Los Alamos Scientific Laboratory for multi-dimensional fluid flow analyses, including seismic slosh (Reference 71 and 72). The code has been used for seismic slosh analysis previously, where a toroidal MK I BQR suppression pool was approximated by an annular geometry, and excited by a simulated sinusoidal seismic event.

Results of this analysis are reported in Reference 73. It was demonstrated that SOLA-3D could be used to describe suppression pool water motion for a seismic excitation applied to the containment structure.

The seismic slosh analysis for SSES suppression pool has been patterned after the annular suppression pool analysis described.

in Reference 73, with appropriate SSES suppression pool and containment parameters used. The results of calculations are pressure,-time histories, caused by vater wave motion, to be applied to suppression pool boundaries in manner and location similar to the method used for SRV and LOCA hydrodynamic loads.

Generally, water motion above the quiescent suppression pool surface. causes "vave loads> and water motion below causes "inertial loads.> The inertia loads vill alvays appear to be larqer than the wave loads because the normal hydrostatic load would be included below the vater surface. at 24 ft. submerqence in cold vater, the hydrostatic(Porheadexample, vould be sliqhtly more than 10 ps'iving a 10 psi bias to the inertia loads at pool bottom.)

Some numerical results of the calculations are shown in Table 4-22 for the selected locations in the suppression pool. As can be observed'hese pressures are small relative to those calculated for the other hydrodynamic loads. Pigures 4-62 i, g, k, and m show typical wave motion at the four containment locations in Table 4-22.

4 2.4 8 'hrust Loads Thrust loads are associated vith the rapid venting of air and/or steam throuqh the downcomers. To determine this load a momentum balance for the control volume consisting of the drywell, diaphragm floor and vents is taken. Results of the analysis indicates that the load reduces the downward pressure differential on the diaphragm.

4.2.5 -Long Term LOCA Logd Definition The loss-of-coolant accident causes pressure and temperature transients in the drywell and wetvell due to mass and energy released from the line break. The dryvell and wetwell pressure and temperature time histories are required to establish the

'REV. 6 4/82 4-20

structural loading conditions in the containment because they are the basis for other containment hydrodynamic phenomena. The response must be determined for a range of parameters such as leak size, reactor pressure and containment initial conditions.

The results of this analysis are cont'ainment initial conditions.

The results of this analysis are documented in Reference 7.

The DBA LOCA for SSES is conservatively estimated to be a 3.53 ft~ break of the'ecirculation line (Reference 7). The SSES plant unique inputs for this analysis are shown in Table 4-19.

Dryvell and vetwell pressure responses are shown in Figures 4-46 and 4-47 (extracted from Reference 7) . These transient descriptions do not, however, contain the effects of reactor subcoolinq. Suppression pool temperature response is shown in Fiqure 4-48 (Reference 7). This transient description also does not contain the effect of reactor subcooling. Dryvell temperature response is shown in Piqure 4-49 and similarly does not contain the effects of pipe inventory or reactor subcooling.

4.2.5.2 -Intermediate Break-Accident K lIBAl Transients The vorst-case intermediate break for the Nark II plants is a main steam line break on the order of 0.05 to 0.1 ft~.

Suppression pool temperature response is shown in Piqure 4-50.

Dryvell temperature and wetvell and dryvell pressures for the SSES IBA are shown in Piqure 4-51.

A.2.3.3 Bsal3, BBe~g Accident ~BBA~Tn nsients.

At this time plant-unique SBA data for SSES is not available.

The vetwell and drywell pressure and temperature transients for a typical Hark II containment are used to estimate SSES containment response to these accidents. These curves are shown in Figure 4-17 (extracted from Reference 10).  !

The various components directly affected by LOCA loads are shown schematically in Piqures 4-53 and 4-54. These components may in turn load other components as they respond to the LOCA loads.

For example, lateral loads on the dovncomer vents produce minor reaction loads in the drywell floor from vhich the downcomers are supported. The reaction load in the drywell floor is an indirect load resulting from the LOCA and is defined by the appropriate structural model of the downcomer/drywell floor system. Only the direct loadinq situations are described explicitly here. Table 4-20 is a LOCA load chart for SSES. This chart shows vhich LOCA loads directly affect the various structures in the SSES containment desiqn. Details of the loading time histories are discussed in the follovinq subsections.

Rev. 2, 5/80 4-21

4 2,6,1 -LOCA goads og the Containment Sall and Pedestal Figure 4-55 shows the LOCA loading history for the SSES containment wall and the BPV pedestal. The wetwell pressure loads apply to the unwetted elevations in the wetwell; and addition of the appropriate hydrostatic pressure is made for loads on the wetted elevations. Condensation oscillation and chuqging loads are applied to the wetted elevations in the wetwell only. The poolswell air bubble load applies to the wetwell boundaries as shown in Fiqure 4.44.

4.2.6.2 LOCA. Loads on the Basemat, and Liner Plate Figure 4-56 shows the LOCA loading history for the SSES basemat and liner plate. Wetwell pressures are applied to the wetted and unwetted portions of the liner plate as discussed in Subsection 4.2.6. 1. The downcomer water get impacts the basemat liner plate as does the poolswell air bubble load. Chuqging and condensation oscillation loads are applied to the wetted portion of the liner plate.

Piqure 4-57 shows the LOCA loading history for the SSES drywell and drywell floor. The drywell floor undergoes a vertically applied, continuously varying differential pressure, the upward component of which is especially prominent during poolswell when the wetwell air space is highly compressed.

4~2~6,4 LOCQ broads og the Columns Pigure 4-58 shows the LOCA 3oading history for the SSES columns.

Poolswell drag and fallback loads are very minor since the column surface is oriented parallel to the pool swell and fallback velocities. The poolswell air bubble, condensation oscillations and chuqqing will provide loads on the submerged (wetted) portion of the columns.

4.2.6.5 LOCA goads on the Downcomers Piqure 4-59 shows the LOCA loading history for the SSES downcomers. The downcomer clearing load is a lateral load applied at the downcomer exit (in the same manner as the chugging lateral load) plus a vertical thrust load. Poolswell drag and fallback loads are very minor since the downcomer surfaces are oriented parallel to the pool swell and fallback velocities. The poolswell air bubble load is applied to the submerged portion of the downcomer as are the chugging and condensation oscillation loads.

4.2.6 6 *LOCA-~~-o ds o 9 the Dow 3~X co e Braci 39 Pigure 4-60 shows the LOCA loading history for the SSES downcomer bracing system. This system is not, subject to impact loads since it is submerged at elevation 668 ' As a submerged structure it Rev. 2, 5/80 4-22

is subject t'o poolswell drag, fallback and air bubble loads.

Condensation oscillations and chuqging at the vent exit will also load the bracing system both through downcomer reaction (indirect load) and directly through the hydrodynamic loading in the suppression pool.

4 2.6 7 LocA'goods og- Wetwell~ipj~n Figure 4-61 shows the LOCA loading history for piping .in. the SSES wetwell.. Since the wetwell piping occurs at a variety of elevations in the SSES wetwell, sections may be completely submerged, partially submerged, or initially uncovered. Piping may occur parallel to poolswell and fallback velocities as with the main steam safety relief piping. For these reasons there are a number of potential loading situations which arise as shown in Table 4-21. 1n additio, the poolswell air bubble load applies to the submerged portion of the wetwell piping as do the condensation oscillation and chugging loads.

Rev. 2, 5/80 4-23

~43 $$ NULUS P Bg SS URIZ ATION The RPV shield annulus has the recirculation pumps suction lines passing through .it (for location in containment see Figure 1- 1).

The mass and energy release rates from a postualted recirculation line break constitute the most severe transient in the reactor shield annulus. Therefore, this pipe break is selected for analyzing loading of the shield wall and the reactor pressure vessel support skirt for pipe breaks inside the annulus. The reactor shield annulus differential pressure analysis and analytical techniques are presented in Appendices 6A and 6B of the SSES Pinal Safety Analysis Report (PSAB).

Rev. 2, 5/80 4-24

Figures 4~1 through 4-37 aqd Figure 4>>,62 ape proprietary aqd are found iq, the propre.etaay supplement to this DAR.

V

~ ~

18 18.17 ft Above initial elevation of suppression pool Cm before LOCA

(= 17. 56 ft. above ool-surface evel at moment UJ of vent clearing)

I U

X UJ X

DK I

I D

2.'C UJ I UJ UJ UJ D ~','

, can,

< ~dd/C C

0. 00 0. 20 0.40 0.80 0. 80 1. 00 TIME AFTER VENT CLEARING (SEC) sd lid' Rev. 2, 5/80 k ~ I USQUEHANNA STEAM EI.ECTRIC STATION 1

UNITS t AND 2 DESIGN ASSESSMENT REPORT SSES SHORT TERM SUPPRESSION POOL SURFACE HEIGHT FIGURE 4 38

60 56.1 podia 50 40 P P

0:

30 20 10 T30766-1 104-77 0.893 sec, 0

0.00 0.25 0.50 0.75 1.00

'IME AFTER VENT CLEARING (SEC)

Rev. 2, 5/80 SUSQUEHANNA STEAM ELECTRIC STATION

~ UNITS 1 AND 2 DESIGN ASSESSMENT REPORT SSES SHORT TERM WETWELL PRESSURE F I FIGURE 449

30 29.35 fps 25 20 G

7.

I O

15 IL K

0 0

10 730766.1 10.3 77 672 677 682 687 692 ELEVATION (FT)

SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT SSES POOL SURFACE VELOCITY vs ELEVATION FIGURE 440

wWt '

COMPUTER RUN T30766.'l I10/3/77l EV CJ CO I

IL R

O I

K IJJ LII

-100

-150 0.0 1.00 TIME AFTER VENT CLEARING (SEC.)

Rev. 2, 5/80 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT POOLSWELL ACCELERATION TIME HISTORY FIGURE 4-40a

PEDES-HWL TAL EL. 672'-0 Q3 CONT.

WALL '

EL. 660'-0

Q2 EL.:648'-0'-

Ql 24 + 14.7 + 10.4, = 49.1 psia Q2 '24 + 14.7, + 5.2 = 43.9 psia Q3 0 + 14.7 + 0'= 14.7 psia Rev. 2, 5/8p SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REFORT

.SSES VENT CLEARING

'PRESSURE DISTRXBUTION FIGURE 4

I THIS FIGURE HAS BEEN'ELETED Rev. 2, 5/80 SUSQUEHANNA STEAM ELECTRIC STAYIOiM UNITS 1 AND 2 DESIGN ASSESSMENT REPORT pi SSES JET IMPINGEMENT AREA (NATER CLEARING)

FIGURE 442

55.0 T30766.1 10.3.77 50.0 co 45.0 D

40.0 CO tC 35.0 30.0 0.00 0.25 0.50 0.75 1.00 TIME AFTER VENT CLEARING (SEC)

SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT SSES POOLSWELL AIR BUBBLE PRESSURE FIGURE 4-43

4 0 i'\

II

r ,,>>r>>

CONTAINMENT COLUMN PEDESTAL'I WALL PI EL. 690 -II HWL .0" EL. 672 P2 EL 660-0 PS Pg EL6lB-0 BASEMAT Pj 56.67PSIA P2-" 4I.96 PSIA Rev. 2, 5/80 PS-" 52>>36PSIA SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT AIR BUBBLE PRESSURE ON SUPPRESSION POOL WALLS FIGURE 4%i

This figure has been deleted REV. 6, 4/82 SUSQUEHANNA STEAM ELECTRIC STATION UNITS1 AND2 DESIGN ASSESSMENT REPORT CONDENSATION PRESSURE FORCING FUNCTIONS FIGURE 4-<<A

This figure has been deleted REV. 6, P/82 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT SYMMETRIC AND ASSYMMETRIC SPATIAL LOADING SPECIFICATION FIGURE 445

120 80 CO 0

CO UJ CC Lll K

~ 40 m

pe Um cn 0

Q z zz Cll pO zx co K Ill Q CO 9 m~m IVI en~2 P I o~ m~m ZZp

+m +0m 0

~ ~O 0 -20 40 80 80 m

oS oc Q TIME (SECONOS) 0 R

120 80 40 0

0m m CO Z

9 Z CO~M m~m y) CO CO ~

Z I

+0m py O 0

XI -I 0 20 40 60 80 m

TIME (SECONDS) 0 O I

I 0

Z

I' n

1'

200 150 U

UJ O

I-0 O

100 0

C g~ Ch CO 0

z m lTI a

mM z zz pgC 2 c+

gzrn 0 37mm~

Cl C

~ m+~

ChCn I +

omg nm~ m2'm Q 0) 0 ZZr

-IOm 0Z 50

~~A m 20 40 60 80 0 0 mO o TIME tSECONDS)

~ I CO 0

0 Z

REFER TO FIGURE 6.2-3 OF THE FSAR (DRYWELL CURVE)

SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT DRYWELL TEINPERATURE RESPONSE FOR DBA LOCA F IGURE 449

150 140 130 120 0

O 0

D CO 110 n

C cD 100 m ~

m CD x 4p CD Cl Q Z CD Z Z c c CD pc+

ch Z ch 90 CO IIl A1 I fll 0 200 600 1000 1200 1400

+ ~

M Cll m TIME (SEC) m>2 m f

~0m z 0z Q

CD

~

Ill M O ITI Q O A Q co Q

CCI 0

Z

SEE FSAR FIGURE 6.2-14 Ia} CONTAINMENTPRESSURE RESPONSE FOR INTERMEDIATE BREAK AREA SEE FSAR FIGURE 6.2-15 (b} DRYWELL TEMPERATURE RESPONSE FOR INTERMEDIATE BREAK AREA Rev. 2, 5/80 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT SSES PLANT UNIQUE CONTAINMENT RESPONSE TO THE IBA FIGURE 4-51

1 30 DRYWELL 20 cc

, D M

Ch ill WETWELL cc c

10 1OO 101 102 103 1O4 1O5 TIME (sec)

(a) CONTAINMENTPRESSURE RESPONSE FOLLOWING SMALL BREAK DRYWELL 200 D

I- WETWELL 100 I-1O0 101 102 10 1O4 1O5 TIME (sec)

(b) CONTAINMENTTEMPERATURE RESPONSE FOLLOWING SMALL BREAK (LIQUID BREAK)

SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT TYPICAL MARK II CONTAINMENT RESPONSE TO THE SBA FIGURE 4-52

CONTAINMENT WALL 0

0 , 0 00 00 0o 0

~ 08'0 Q OOWNOOMER M 0

Q gE o

000 0.00 0

0 0R QN 0 0 00 0 0 F00~

0 DOWNCOMERS F0~0 COLUMNS 0 000 s 0 0 00 000 WETWELL 00 PIPING NOTE:

DOWNCOMER BRACING IS ONLY PARTIALLYSHOWN IN THE INTEREST OF CLARITY.

LETTERS INDICATE SRV QUENCHERS R'"'V . 6, 4/82 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT SSES COMPONENTS AFFECTED BY LOCA LOADS FIGURE 4.S3

I 'I I

B.O. SLAB

'I.' EL. 700'- 3" B.O. HYDROGEN VACUUM BREAKER R E COMBINE R E L. 692'-1" I I EL. 691'.0" T.O. PLATFORM EL. 691'-0" MAXIMUMPOOL SWELL Pgge7tt MAXIMUMPOOL SWELL HEIGHT= 1.51 X MAX VENT SUBMERGENCE

=

18.17' HIGH WATER LEVEL EL. 672'-0"

~

I BRACING NORM WATER LEVEL E L. 668'-0" E L. 671'-0" MAXIMUMVENT SUBMERGENCE

~12'-0" B.O. VENT PIPE EL. 660'-0" DIAPHRAGM SLAB WETWELL SUPPORT COLUMN PIPING 12'"

3I Qtl T.O. SLAB

~

EL. 648 I 0II RI'.V . 6, 4/82 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT SSES COMPONENTS AFFECTED BY LOCA LOADS FIGURE 4-54

WETWELL/DRYWELL P&T DURING POOLSWELL %

WETWELL/DRYWELL P&T WETWELL/DRYWELLP&T DURING DBA LOCA ""

DURING LOCA ~~

POOLSWELL AIR BUBBLE" MIXED FLOW STEAM FLOW I GGING C P ~i+~ I C P ~~~a

~ ~ I C ll xm Q 0Ill q 0 ~

Ql Q~Q 0 VENTS PEAK FALLBACK B LOWDOWN m z zZ BREAK CLEAR POOLSWELL HT. COMPLETE COMPLETE ygO pc+ Z Zm+cn 0 Ch Ch m+

Ch m

0.6863 1.5793 2.6371 15.0 25.0 60 U ~ CO Ch TIME (SEC)

'Q ZI+ m+m Z m

MWcn ~O fm QZO fll " DBA ONLY rFg 0 O "% IBA OR SBA z~ CO i~" EITHER DBA, IBA OR

"'" DBA AND IBA ONLY SBA 0

z

WETW EL L/DRYWE LL P&T DURING POOLSWELL" WETWELL/DRYWELL P&T WETWELL/DRYWELLP&T DURING LOCA ~~~

DURING LOCA """"

DOWNCOMER WATER JET LOAD ~

POOLSWELL AIR BUBBLE 4 I

MIXED FLOW STEAM FLOW IICHUGG(NG +++

C.O. "" ""

)

I C.O.

0 Q 0 m Q x rn a zz

. I O VENTS PEAK FALLBACK BLOWDOWN Xp BREAK p ill zc+z cn CLEAR POOLSWELL HT. COMPLETE COMPLETE 2 p CO CO p ~+m 0.6863 1.5793 2.6371 15.0 25.0 60.0

@co 0 m co CO TIME (SEC)

~ Cl P c-+a P. > m mzg zom

~mx H~O m DBA ONLY HQ A DBA AND IBA ONLY z~

U

'ITHER DBA, IBA OR SBA

~~" ~ IBAOR SBA 0

Z

WETWELL/DRYWELL POT DURING POOLSWELL 0 WETWELL/DRYWELL PST WETWELL/DRYWELLPRT DURING LOCA ""~

DURING LOCA ""

PEAK BREAK POOLSWELL HT.

c0m 1.5793

~l O m TIME (SEC) zoo o

rn n 2 z

g +@ 2

~c+

zgO cnz~

M mtnm Pn~g

)~a <+m mz<

2Um

~XIX I K CO 4 Og+

Om0Zl m ~~

DBA ONLY IBA OR SBA

~fI o CO i~i EITHER DBA, IBA OR SBA I

0 z

POOLSWELL DRAG LOADS POOLSWELL AIR BUBBLE LOADS %

FALLBACK LOAD "

MIXED FLOW STEAM FLOW I

CHUGGING ~""

lCHUGGING C 0 I C 0 I 0

I 'ill C m

~aoo

~

gnmo m m

rzr PqOmr rn n 22 2C > BREAK VENTS CLEAR PEAK POOLSWELL HT.

FALLBACK COMPLETE B LOWDOWN COMPLETE fTl g W2 ch 0.6863 1.5793 2.6371 15.0 25.0 60.0 gQchW mO co+m m vl m ZCO m TIME (SEC) y) no+

g()Q CO E>m m2r 20 m r cn~~ +~ 0 CHC goO m ONLY ii DBA cn <~'> 0 O CO DBA AND IBA ONLY I

"~~ DBA, IBA AND SBA O

2

DOWNCOMER CLEARING LOAD ~

POOLSWELL DRAG LOAD "

POOLSWELL AIR BUBBLE LOAD 4 FALLBACK LOAD %

MIXED FLOW I STEAM FLOW I <<~i I CHUGGING CO ~~ II CO ~"

A gm 0 Op. O m XJ ~ m co

~O 0 Z VENTS PEAK FALLBACK B LOWDOWN Ill l- 2 Z BREAK CLEAR POOLSWELL HT. COMPLETE COMPLETE eo 0 yC +

cnZ co lll Q CO 0.0 1.5793 25.0 co m~ I 0.6863 2.6371 15.0 60.0

~ ~ Ol m O> CO m

~Q 2+

Qx

+ +Um c

~ O O~ CO

<o K g7 m

0 O

~ DBA ONLY "i DBA AND IBA ONLY m~

Zl  %%

DBA, IBA AND SBA

• 0 Z

POOLSWELL DRAG LOAD ~

POOLSWELL AIR BUBBLE LOAD ~

FALLBACK LOAD ~

MIXED FLOW STEAM FLOW ICHUGGING C.O."" I C.O."

I 0

C 0

m m

X PJ gl H g VENTS PEAK FALLBACK -

BLOWDOWN 37 X ~ BREAK CLEAR POOLSWELL HT. COMPLETE COMPLETE P ill C 2.C o 0

-4>

2'O 0.6863 1.5793 2.6371 15.0 25.0 60.0

+ ITl Q m~m CO Q cn- ~ CO ch O + TIME (SEC)

~~A Fi m~m Ch g ~ +0m ~ O.

4 DBA ONLY mKcn gO+OO m g7 0

~ DBA AND IBA ONLY O O DBA, IBA AND SBA mC I Ol 0

z

POOLSW EL L IMPACT LOAD ~

(applied sequentially in cases where both loads occur)

POOLSWELL DRAG LOAD ~

POOLSWELL AIR BUBBLE LOAD ~

FALLBACK LOAD ~

MIXED FLOW I STEAM FLOW I CHUGGING """

C.O ~~ I C.O ~~

0 0 C

+,C

~'0 0 m

m X

~A m>

CO z VENTS PEAK FALLBACK B LOWDOWN zpC-+ BREAK . CLEAR POOLSWELL HT. COMPLETE COMPLETE Qa ftt COZ CO

~ m+m CO 0.0 0.6863 1.5793 2.6371 15.0 25.0 60.0 U COCO ~

CO

~>m zz!

+0m ~ 0 Cg r 0, m 0 DBA ONLY

~i DBA AND IBA ONLY 2m O O zc P.

I CO I

~~~ DBA, IBAANDSBA C) z0

i This figure has been deleted REV. 6, 4/82 SUSQUEHANNA STEAM ELECTRIC STATION UNITS1 AND2 DESIGN ASSESSMENT REPORT CHUGGING POOL BOUNDARY LOADS FIGURE 4-62 A 6 B

This figure has been deleted RHV 6, 4/82 SUSQUEHANNA STEAM ELECTRIC STATl08 UNITS 1 AND 2 DESIGN ASSESSMENT REPORT CHUGGING POOL BOUNDARY LOADS fIGURE 4-62 C & D

Thi.s figure has been deleted REV. 6, 4/82 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT CHUGGING POOL BOUNDARY LOADS PIGURE 4-62 E 6 F

10 O

hC 5

CD CD 6 MSEG.

0 5 10 15 TIME (SEC.)

FIGURE 4-62g DYNAMIC DOWNCOMER LATERAL LOADS DUE TO CHUGGING 30 20 CD CD 10 3 MSEG.

0 5 10 15 TIME (MSEC.)

FIGURE 4-62h DYNAMIC DOWNCOMER LATERAL LOADS DUE TO CHUGGING Rev. 5, 3/81 SUSQUEHANNA STEAN ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REF(MT DYNAMIC DOWNCOMER LATERAL-LOAD DUE TO CHUGGING FIGURE 4-62 G & H

WRVE HEIGHT (t=2i Z=2i IZPL)

CS I%

III 4J LP CXg I

.03 R.OR A+02 le02 l.OI I oOI I .Ol I o0! I .00 I .00 0 .00 X-TIHK ISECI REV. 6, 4/82 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND.2 OESIGN ASSESSMENT REPORT TYPICAL NAVE l'lOTION DUE TO SEISMIC SLOSH

IJA VE HE1GHT t l =2i 7=3'M1 ~ I ZPL)

O cog

~0

~ l

~a g gled 4J

~I A

5.03 2.02 0.02 ~ .02 0.01 10.01 1 .01 1 ~ 01 1 .00 1 .00 2 .00 X-TtIIE INC)

REV. 6, 4/82 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT TYPICAL l'IAVE 1'10TION DUE TO SEISMIC SLOSH FIGURE 4-62J

FAYE HEIGHT t t'= t H1 ~ J=2. I J'P 0@00 4e0l Oo01 0 ~ OI I 01 I ~ Ol I ~ OI I 00

~ I 00 1 ~ 00 X- T1HE (SEC I REV. 6, 4/82 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESlGN ASSESSMENT REPORT TYICAL WAVE IibTION DUE TO SEISNIC SLOSH flGURE g-62K

LJ Fl V E HE IG HT ( t= lM 1 ~ Z = O'H 1 ~ t ZP L )

CI I

O~

Xg W

u.8 l/l I

n nv

~

bo03 lo02 4 ~ 03 0 ~ 02 So01 I Dl I oDI I oDI I o00 'I o00 2 ~ 00 X-TLHK ISECI SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 OESIGN ASSESSMENT REPORT TYP ICAL NAVE I'lOTION DUE TO SEISMIC SLOSH FIGURE 4 fjQ

Tables 4-1 through 4-15 are proprietary and are found in the proprietary supplement to this DAR.

~ ~ r--

TABLF. 4-16 LOCA I OADS ASSOCIAT ED PITH POOLS PELL Load

1. Retvell/Drywell Pressures during Poolsvell

2. Poolswell Impact Loads

3. Poolsvell Drag Loads

4. Dovncomer Clearing Loads

5. Dovncomer Ra ter Jet Load

6. Poolsvell Air Bubble Load

7. Poolsvell Fallback Load Rev. 2, 5/80

TABLE 4- 17 SSES DHYWELL PHESSURE T~ime seconder'.

00000 15. 46 0 00195 15.

18'5.21

0. 00208

0. 00586 14.79 0 0645 '

18. 17 ~

127 21. 16 0 252 26. 61

0. 502 36 52 0 627 38 26"

0. 658 37. 71

1. 057 42-09 1 867 48.43

1. 900 48. 54 2 119 48. 73

TABLE 4-18 SSES PLANT UNIQUE .POOLSMELL CODE .INPUT DATA fthm Down comer Area (each) 2 96 Suppression Pool Free Surface Area 5065 ft Maximum Downcomer Submergence 12.00 ft Downcomer Overall Loss Coefficient 2.5 N umber of Dow nc ome rs 87 Initial Metwell Pressure 15.45 psia Wetwell Free Air Volume 149,000 ft~

Vent Clearing Time 0.6863 sec Pool Velocity at Vent Clearing 3. 0 'f t/sec Initial Drywell Temperature 135oZ Initial Drywell Relative Humidity 0. 20 Rev. 2, 5/80

TABID 4-19 INPUT DATA FOR SSES IOCA TRANSIENTS Drywell free air volume 239~ 600 ft ~

{including vents)

Wetwell free air volume 149i000 ft>

Maximum downcomer submergence 12. 0 Downcomer flow area (total) 256.7 ft Downcomer loss coefficien+

Initial drywell pressure 15.45 psia Initial wetwell pressure 15.45 psia Initial drywell humidity 405 -55to Initial pool temperature 90~F Estimated DBA break size 3.53 f t~

Number of vents Initial mass of steam in vessel 24,500 ibm Initial mass of saturated water in 674,000 ibm vessel Minimum suppression pool mass 7.6x106 ibm Initial vessel pressure 1,055 psia Vessel 8 internals mass 2,940,300 ibm Vessel 6 internals overall heat 484.9 Btu/sec F transfer coefficient Vessel and internals specific heat 0 12 3 Bt u/ibm OF Initial control rod drive flow 10.83 ibm/sec Initial steam flow to main turbine 3 931. 5 1 bm/s ec RCIC 6 HPCI (HPCS) flow initiation 489.5 in level, distance from vessel "0" Rev. 2, 5/80

Table 4-19 /Continued}

RCIC 6 HPCI (HPCS) flow shutof f 581.5 in level {normal water level), distance from vessel <<0<<

Rated RCIC flow rate to vessel 83.4 ibm/sec Rated HPCI -(HPCS) flow rate to vessel 695 ibm/sec RCIC shutoff pressure 165 psia HPCI (HPCS) shutoff pressure 165 psia Condensate storage tank enthalpy 48 Btu/ibm CRD enthalpy 48 Btu/ibm Initial power level 3.23x10~ Btu/sec Peedwater enthalpy 78 Btu/ibm Cleanup system flow 36.94 ibm/sec Cleanup system return enthalpy 413.2 Btu/ibm Initial vessel fluid enthalpy 573 ~ 1 Bt u/ibm RHR heat exchanger <<K<< in pool 306 Btu/sec oF cooling mode RHR heat exchanger steam flow in 25 lbs/sec condensing mode I

RHR heat exchanger flow in pool 1390 lbs/sec cooling mode RHR heat exchanger outlet enthalpy 108 Btu/ibm in condensing mode Service water temperature 90 oF Rev. 2, 5/80

TABLE 4-20 COMPONENT LOCA LOAD CHART FOR SSES LOAD STRUCTURE DIRECTLY AFFECTED 1 2 3 4 5 6 7 8 9 10 1] 12 13 Containment Wall X X X X X Pedestal (incl. interior) X X X X X X Basemat X X X X X X X X Liner Plate X X ,X X X X X X Drywell Floor X X X Drywell X X X Columns X X X X, X Downcomers X X X X X X X Downcomer Bracinq X X X,X X Wetwell Piping X X X X X X LOAD LEGEND 1 Wetw ell/Drywel1 pressure d urinq poolswel1 2 Poolswell impact load 3 Poolswell drag load 4 Downcomer clearing load 5 Downcomer water jet load 6 Poolswell air bubble load 7 Fallback load 8 Hiqh mass flux condensation load 9 Medium mass tlux condensation load 10 Chuqqing load 11 Wetwell/Drywell PGT during DBA 12 Wetwell/Drywell PGT durinq IBA 13 Wetwell/Drywell PGT during SBA

TABLE 4-21 HZT HELL PIPING LOCA LOA DING SITUTATEONS

~Pi ~in Configuration LOCA Load to be Applied 4

1 Completely Submerged (a) ver tical skin drag load only (C<)

(b) horizontal drag load (CD) 2 Partially Submerged (a) vertical skin drag load only (C<)

3 Initially Uncovered (a) vertical skin drag load only (Cf )

(b) horizontal impact load, then drag load (CD)

Table 4-22 Sloshing Wave Height of Max. HF2 (2 2) HF3g (2,17) HBK2g (7t2) HBK3g (7g17)

Height I =,2g J,=2 I = 2, J = 17 I=7, J=2 I= 7t J = 17 sec ~

25.40 14.0 (1.40) 9.90 25.80 (1.80) 17.50 25.60 (1.60) 12.90 25+95 (1.95) i" Fig..4-62i Fig. 4-62) Fig. 4-62k Fig. 4-62m Note: ~ = Shows Location

() = Inside bracket is the net wave height fran the initial position 24 ft. frcxn the bottan of tank.

I = Mesh numbers on the radius from inside to outside.

J = Circumferential division numbers.

REV. 6, 4/82

CHAPTER 5 LOA D COMBINATIONS POR S TR UCTUR ES P XP XNG A ND EQUXPHENT TABLE- Og COPTQQTS 5 1 CONCRETE CONTAINMENT AND REACTOR BUILDING LOAD COMBINATIONS 5 2 STRUCTURAL STEEL LOAD COMBINATIONS 5 3 LINER PLATE LOAD COHBINATIONS 5 4 DOMNCOHER LOAD COMBINATIONS 5 5 PIPING'UENCHER, AND QUENCHER SUPPORT LOAD COMBINATIONS 5.5.1 Load Considerations for Piping Inside the Dryvell 5.5.2 Load Considerations for Piping Inside the Metvell 5.5.3 Quencher and Quencher Support Load Considerations 5.5.4 Load Considerations for Piping in the Reactor Building 5 6 NSSS LOAD COMBINATXONS 5 7 BALANCE OP PLANT (BOP) EQUIPMENT LOAD COMBINATIONS 5 8 ELECTRICAL RACEWAY SYSTEM LOAD COMBINATIONS HVAC 5 9 DUCT SYSTEM LOAD COHBINATIONS 5 10 FIGURES 5 11 TABLES RZV. 6, 4/82 5-1

CHAPTER 5 gIGUQQS Numher- T~ile Piping Stress Diagrams and. Tables 5-2 Piping Stress Diagrams and Tables 5-3 Piping Stress Diagrams and Tables 5-4 Piping Stress Diagrams and Tables Rev. 2, 5/80 5-2

CHAPTER 5 TA~BES Number Title 5-1 Load Combinations for Containment and Reactor Buildinq Concrete Structures Considering Hydrodynamic Loads 5-2 Load Combinations and Allowable Stresses for Structural Steel Components 5-3 Load Combinations and Allowable Stresses for Downcomers 5-4 Load Combinations and Allowable Stresses For Balance of Plant (BOP) Equipment Load Combinations and Allowable Stresses for NSSS Equipment and Piping 5-6 Load Combinations and Allowable Stresses for the Electrical Raceway System 5-7 Load Combinations and Allowable Stresses for HVAC Ducts and Supports Rev. 7, 6/82 5-3

5 0. - LOAD COHBXNAPXONS- POR SXRUCNURES PXPXNG AND~EUIPHENX To verify the adequacy of mechanical and structural design, i,t is necessary first to define the load combinations to which structures, piping, and equipment may be subjected. In'ddition to the loads due to pressure, yeight, thermal expansion, seismic, and fluid transients, hydrodynamic loads resulting from LOCA and SRV discharge are considered in the design of structures, piping, and equipment in the drywell and suppression pool. This chapter specifies how the J.OCA and SRV discharge hydrodynamic loads will be combined with the other loading conditions. Por the load combinations discussed in this chapter, seismic and hydrodynamic responses are combined by the methods specified in Reference 10

! Subsection 5.2 2 and Reference 10 Section 6 3.

Rev. 2, 5/80

5 1 CONCRETE CONTAINMENT AND REACTOR BUILDING LOAD COMBINATIONS The loads on the containment, internal and reactor building concrete structures ar'e combined to assess the structural inteqritv in accordance with the design load combinations given in Table 5-1; The factored load approach is used in the design and analysis of the structural components. The load factors adopted are based upon the degree of certainty and probability of occurrence for the individual loads as discussed in Ref 10, Subsection 5.2.2. The time sequences of occurrence of the various time dependent loads (as presented in Figures 4-55 throuqh 4-61, for example) are taken into account to determine the most critical loading conditions.

5-5

5 2 STRUCTURAL STEEL LOAD CONBTNATIONS The load combinations for structural steel in the containment and the reactor building are given in Table 5-2. These combinations apply to "the 'suppression chamber steel columns, the downcomer bracing, and the reactor building structural steel 5-6

5 3 LINER PLATE LOAD COMBINATIONS The liner plate and anchorage system are designed for the load combinations listed in Table 5-1 except that all load factors are taken as<unity.

Rev. 2, 5/80 5-7

5' DOMNCOHER LOAD COMBINATIONS r

Load combinations for the dovncomers are given in Table 5-3.

These load combinations are based on the load combina'tions given~

in Table 6-1 of Reference 10.

Rev. 2, 5/80

5 5- - PIPINGg~UENCHER~ AND. gUENCHER- SUPPORT LOAD COHBZNATXONS LOCA loads considered on piping systems include poolsvell impact loads, poolsvell drag loads, downcomer vater jet loads, poolswell air bubble loads, fallback drag loads, condensation oscillation loads, chugging loads, and inertial loading due to acceleration of the containment structure produced by LOCA loads. Loads due to SRV discharge on piping systems include vater clearing loads, air clearing loads,-fluid transient loads on SRV discharge piping, reaction forces at the guencher, and inertial loading due to the acceleration of the containment structure produced hy SRV discharge loads.

The load combinations and the acceptance criteria for piping systems are given in Table 6-1 of Reference 10.

5.5 1 Load Considerations for Piping Inside the Dryvell Piping systems inside the dryvell are subjected to inertial loading due to the acceleratiern of the containment produced by LOCA and SRV discharge loads in the vetvell. The SRV discharge pipinq in the dryvell is also subjected to fluid transient forces due to SRV discharge.

5.5.2-- Load Considerations for Piping Inside the Wetvell All pipinq in the vetvell is subject to the inertial loading due to LOCA and SRV discharge.

Drag and impact loads due to LOCA and SRU discharge on individual pipes in the vetwell depend on the physical location of the pipinq. Other SRV discharqe and LOCA loads applicable to piping in the vetwe11 are discussed in the paragraphs that follow.

Piping systems located below the suppression chamber water level are shovn on Figures 5-1 and 5-2. These lines are located outside of the jet impingement cone of the downcomer. Tn addition to the inertial loads, these piping systems are subject to air bubble loads, condensation oscillation loads, and chugging loads due to LOCA and SRU operation. The'RV piping, guencher, and quencher support are also subject to fluid transient forces due to SRV discharge.

Piping systems vithin the poolswell volume are shown on Figures 5-2, 5-3 and 5-0. All horizontal runs of these pipes are above the suppression chamber water level. ,The folloving loads, in addition to inertial loads, act on these systems:

'a 0 The horizontal runs of pipe belov elevation 690 '2'~,

experience poolsvell impact , poolsvell drag and fallback drag loads.

Rev. '2, 5/80

b. The vertical portions of pi'pe in the vater belov elevation 690 ~ -2" experience poolsvell drag and fallback drag loads.

5.5.3- guencher and guencher Support =Load Considerations The quencher and quencher supports are subjected to the folloving hyd.rodynamic loads in addition to the pressure, veight, thermal, and seismic loads:

a. Unbalanced loads on the quencher due to SRV vater clearing and, air clearing transients, irregular condensation, and steady state blovdovn

b. Drag loads due to SRV discharge and LOCA

c. SRV piping end loads

d. Inertial loading due to the acceleration of the containment produced by SRV discharge and LOCA.

5.5.Q Load Considerations for Piping in the the ReactoR Baildincn The effects of the inertial loading due to acceleration of the containment produced by SRV discharge and LOCA loads vill be evaluated for this piping.

Rev. 2, 5/80 5-1 0

5 6 -VASSS LOAD COHBINATIOHS The load combinations used for the evaluation of the VASSS piping and equipment are contained in Table 5-5.

REV. 67 4/82 5-1 1

Load combinations for seismic category I equipment located vithin the Containment, reactor and contro1 buildings are assessed for the 1oad combinations shovn in Table 5-4.

REV. 6, 4/82 5-12

I

5. 8 ELZCTQICAQ gACgHAY SYSTQN LOAD COtfBZNATIONS The load combinations for evaluating the Electrical Raceway System are given in Table 5-6.

RZV. 6, 4/82 5-13

5- 9 HVAC DUCT SYSTEN LOAD COMBINATIONS

~0 The load combination for the HVAC duct system are given in Table 5 7&

Rev. 7, 6/82 5-14

ELEVATION LINE NO. GTY SYSTEM PENET NO. DIM. X REST. EL 24"-H 8 B-110 RHR X-203A,B,C & D 660'0" 658'-1" 656'2" 23lt CORE 16"-HBB-104 X-206A 5 B 659'6" 658'1" 658 -8" 18" SPRAY 6"-HBB-102 RCIC X-214 654'-1 0.1/8" 654'-1" = 652'-1" 13 7/8" 854'-I" 16"-HBB-109 HPCI X-209 655'.6" 654'-1" 652'8" 2'-1-1/2" 654'1"

,I

,I ~

i; ~

r

~;

~ ~

+~0 qo qG pif'p:

>o

'b 51.656 FIGURE A 3i BII 7go H j 0.

G 5t+It 33o GO+

8.. pi 2'-6"

~ I 656" FIGURE C 51.

50o 0'jg 0 ~

'9o-30'N SUSQUIEHANNA STEAM ELECThlC STATION UNITS 1 ANO 2 FIGURE B DESIGN ASSESSMENT REPORT tWNQ STIR&

DIAS%~ AO TAOLES FIOUhf 5-1

FIG. TYPE OF ELEVATION LINE NO. QTY SYSTEM PENET NO. OIM. X REST. EL NO. PENET A 668'4" 12"-H 8 8-101 RCIC X-215 SLEEVE 674'-3" 659'4" 2t 9lt 659'9" 24"-HBB-108 657'4" Bll 688'4" HPCI X-210 SLEEVE 674'-1" Ql 658'-1" 3'.3 1/8" 10"-HBB-120 RHR X-246A Bt 8 SLEEVE 674'-0" 666'-6" 667'4" 3'4 5/8" 673'3" 6"-H 8 D-186 RHR X-226A & B SLEEVE 673'-3" 665'4" 1'6" 668'4" LATER 2"-HBB-114 RCIC X-216 EMBEDDED . 667'-3" LATER LATER LATER 4"-EBB-1 02 HPCI X-211 EMBEDDED 677'4" 665'4" 1

~ Qll 668'4" CORE 665'4" IN Qll 668'4" 4"-HBD-183 SPRAY X-208A 8t 8 EMBEDDED 673'.3" CORE 676'W" 10"-HBD-183 X-207A 8I 8 EMBEDDED 685'-1" 665'4" 1

~ gtt 677'4I" SPRAY 668'W" 685'.1" 18"-HBD-185 RHR X-204A 5 B SLEEVE 685'-1" 666'4" 2t 3I ~

670'4" 673'.10" LATER 3"-H 8 8-108 HPCI X-244 EMBEDDED 670'4" LATER LATER LATER LATER 2"-HBB-101 RCIC X-245 EMBEDDED 673'-0" LATER LATER LATER LATER 2"-H 8 B-101 RCI C X-217 EMBEDDED 673'-0" LATER LATER LATER io 0 EL A o o'P

~

ELA J

450 H 1'-3 5/8" ONE HORZ TWO DIR REST.

P HORZ REST.

ANCHOR ANCHOR 0/4/

ELB EL 8 DIM. X DIM. X

,O. //

D ~

FIGURE A FIGURE B FIGURE C 0 . id

~

b,o' ~ b.

EL A ELA DIR TWO DI R ITWO HOR Z REST.

REST l I DIR NOVIEHAMNQSTEAM ELECTRIC STATION lTWO HORZ REST. LRIITS 1 AND 2 EL 8 ELB DIM. X DIM. Z DESIGN ASSESSMENT REPORT'RIE P.II STkKNa o ~

WNMAKGAIR) TAOl.Ka FIGURE D FIGURE E PIOunIE Sh.

e 0

A 12"~1O1 &1'W' 12"~101 10 KJ.S.DP.N.O.KA.H. $

51'W'I EEVE PENETRATION E L 704'4I" I

I

'I I I I I EL 894'4" TWO DIR HORZ tk ANCHOR EL 694'W"

! TORSIONAL REST. MAX POOL EL HEIGHT EL. 699' HIGHNATER LEVEL EL 672'A" EL 668'4" 1WO DIR 4'a' HORZ REST. o

~

.'AEL A

EL 649'4" TWO DIR HORZ8c TORSIONAL REST.

FIGURE A FIGU REB BOTTOM SUPPRESSION POOL EL 648'Z" Rev. 2, 5/80 SUSQUEHANNA STEAM ELECTRIC STATION UNITS1 AND2 DESIGN ASSESSMENT REPORT PIPING STRESS DIAGRAMS AND TABLES FIGURE 5-3

FIG. QTY LIIOE ICO. SYSTHI KL A KL N IEL C RAD Y DIN. X EST.

NO. KL Pent% -5 10"- NS'4" ON'W'2'4 34" 1i%0" &F4" RAD Y ELC 24 VERTICAL5 4 AXIALREST.

ELB VERTICAL REST.

POOLSWELL EL 69D'-2'.

EL A I

HIGH WATER EL 672'-0" DIM. X EL 648'4" Rev. 2, 5/80 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT PIPING STRESS DIAGRAMSAND TABLES, FIGURE 54

TABLE 5-1 LOAD OMBINATIONS POR CONTAINMENT AND REACTOR BUILDING CONCRETE STRUCTURES AND

+CONTAINMENE' LINER PLATE /CONSIDERING HYDRODYN AMIC) LOADS Load Single Equation Condition D L P R V SRV~<i AOZ ADS ASYM Valve LOCA<>>

o o %s B A A Normal 1

w/o Teap. 141710 1.5 XCi >

2 Normal

.w/Temp. 1 0 1.3 1.0 1.0 1.0 1% 3 3 Normal Se v. Env. 1.0 1.0 1 0 1 0 1 0 1 25 1 25 4 Abnoraal 1.0 1.0 1 25 1.0 1.0 - 1 25 4a Abnormal 1.0 1.0 1. 25 1 0 1.0 - 1 0 5 Abnormal Sev. Env. 1010 1 0 1.0 1 1 X X 5a Abnormal Sev. Env. 1.0 1.0 111010-10 Normal 6

Ext. Env. 1.010101010 1 0 1.0 7 Abnormal Ext..Env. 1.0 1 0 1 0 1.0 1.0 1.0 1.0 1 0 X I 7a Abnormal Ext. Env. 1.0 1.0 1 0 1.0 1.0 1 0 1.0 1 0

+ Por liner plate the coefficients are unity.

Rev. 2, 5/80

Load Dcscri tion D ~ Dead Loads Eo Operating-Basis Earthquake I. ~ Live Loads Ess ~ Safe Shutdown Earthquake Pn ~ SBA or IBA (LOCA) Pressure Load Po ~ Operating Prcssure Loads RA ~ Pipe Brcak Tempcraturcs Reaction Loads To = Operating Temperature Loads PA = DBA (LOCA) I'rossurc Load Ro ~ Operating Pipe Reactions TA ~ Pipe Break Temperature Load SRV ~ Safety Relief Valve Loads RV ~ Reaction and jct forces associated with the pipe break Notes:

l) X indicates applicability for the designated load combination.

2) For the columns designated AOT, ADS, ASYH, and Single Valve, only one of the four possible columns may be included in thc load combination for any one equation. For example, in Equation l either AOT or ASYII may be considered with the other loads but not both AOT and ASM simultaneously.

3) LOCA includes chugging, condensation oscillation, and large air bubble loads.

4 ReVs 2, 5/80

Table 5-2 LOAD COHBENATIOHS AND ALLOWABLE STRESSES FOR COHPONENTS (Suppression Chamber Columns, STEEL'TRUCTURAL Downcomer Bracing ~ and Reactor Building Structural Steel.

Stress Equation Condition Load Combination Limit Normal D+L+SRV "F w/o Temp.

Normal D+L+T +S R V 0

F S

w/Temp.

N or mal/ 0+L+T0 +E+SRV 1.5 F Severe Normal/ D+L+T +E '+SRV 0

1.5 F Extreme Abnormal D+L+P+ (To+T a) + R ~

(Note 1)

+SRV+LOCA Abnormal D+L+P+ (T +T )+R+E (Note 1)

Se vere +SRV+LOCk Abnormal/ D+L+P+ (T +T a ) +H+E' (Note 1)

Extreme +SRV+LOCA Note 1: Xn no case ..hall the allowable stress exceed 0.90F>

in bending, 0.85F in axial tension or compression.

and 0.50P in shea%. Where the design is governed by requi ements of stability (local or lateral buckling), the actual stress shall not exceed 1.5F S Rev. 7, 6/82

Table 5-2 (Cont'd)

No tations Notations:

Allowable stress according to the AISC, "Specification for the Design, Fabrication, and Erection of Structural Steel for Buildings,<~ dated 1969, Part l.

DeaQ Load Live Load Thermal effects during normal operating conditions including temperature qraQients and equipment and pipe reactions.

Added thermal effects (over and above operating thermal effects) which occur during a design accident.

Desiqn basis accident pressure load Local force or pressure on structure due to postulated pipe rupture including the effects of steam/water jet impingement, pipe whip, and pipe reaction.

Load due to Operatinq Basis Earthquake.

Load due to Safe Shutdown Earthquake.

SRU Safety relief valve loads.

LOCA Loads due to Loss of Coolant Accident conditions (chugqinq, condensa tion oscillation, or large air bubble loads) .

Minimum specified yielQ strength Rev. 2, 5/8p

Table 5-3 LOAD COMBINATIONS AND ALLOWABLE STRESSES FOR DOWNCOMERS

~Eaaaioa Condition Load Combination Primar Stress Limit Upset D+P +SRY o ALL Emergency D+Po+SRVALL+E 2.25 Sm Emergency D+PSBA+SRVADS+E+LOCA(SBA) 2.25 S Faulted D+P +SRVALL+E' Faulted D+ PIBA+SRYADS+E+ LOCA (IBA) 'm A (orP Faulted D+P )

+SRPADS+E'+LE(SBA or IBA) 'm Faulted D+PA+E'+LOCA(DBA) 'm Notations:

S m

Maximum allowable stress according to Table 1-10.1, Ref. 29.

Dead weight of the downcomer p

o Pressure differential between drywell and suppression chamber during normal operating condition.

SBA Pressure differential between drywell and suppression chamber during SBA.

IBA Pressure differential between drywell and suppression chamber during IBA.

Pressure differential between drywell and suppression chamber during DBA.

S RVAz z

Dynamic lateral pressure and inertia load due to the discharge of all 16 safety relief valves simultaneously.

SRV Dynamic lateral pressure and inertia load due to the discharge of all 6 ADS safety relief valves simultaneously.

Load due to Operating Basis Earthquake Et Load due to Safe Shutdown Earthquake LOCA Loads due to chugging, condensation, oscillation, or air bubble loads. The governing applicable loading case should be considered. The loads should include:

1. Lateral load at the tip of the downcomer

2. Horizontal and vertical inertial loads

3. Submerged structures loads Rev. 2, 5/80

TA BLE 5-4 LOAD COHBXNATXONS ~AD ALLOMABLE STRESSES

~Euation Cond ition- Load Combination Stress Limit 2 Normal D+L+ SR V Fs v/o Temp 6 pr.

Normal D+L+T+P+SRV Fs w/Temp 6 pr.

Abnormal/Severe D+L+T+P+E+SRV+LOCA l 5F Abnormal/Extreme D+L+ T+P+E'+SR V+LOCA l 5F where F Allowable stress for normal conditions s

D Dead Load Live Load Pressure loads during operating conditions including pressure gradients and equpment and pipe reactions.

Thermal effects during normal 'operating conditions including temperature gradients and eguipment and pipe reactions.

Loads due to operating basis earthguake E l Loads due to Safe Shutdown earthguake SRV Loads due to Hain Steam Safety relief valve operation LOCA Loads due to Loss-of-Coolant Accident occurrence.

REV. 6, 4/82

TABLE 5-5 LOAD COMBINATION AND ACCEPTAHCE CRITERIA FOR ASIDE CODE CLASS 1g 2 AND 3 HSSS PIPING AND EQUIPMENT Design Evaluation Load Combinafion- Bys j,s Basis Leve~l N + SRV Upset Upset (B)

N + OBE Upset Upset (B)

N + OBE + SRV Emergency Upset (B)

N + SSE + SRV Faulted Paulted* (D)

N + SBA + SRV Emergency Emergency* (C)

H + IBA + SRV Faulted Faulted'Service Paul ted~ (D)

N + SBA + SRV Emergency Emergency+ (C)

N + SBA + OBE + SRV Faulted Paulted* (D)

N + IBA + OBE + SRV Faulted Faulted~ (D)

N + SBA/IBA + SSE + SRV Faulted Faulted+ (D)

N + LOCA++ + SSE Faulted (D)

LOAD DEFINITION LEGEND Normal (N) Normal and/or abnormal loads depending on acceptance criteria.

OBE Operational basis earthquake loads.

SSE Safe Shutdown earthquake loads.

SRV Loads associated with Safety Relief Valve actuation.

REV. 6, 4/82

LOAD'ONBIMATIOH TABLE (Cont.)

LOC Al The loss of coolant accident associated with the postulated pipe rupture of large pipes (e.g., main steam, feedwater, recirculation piping) .

LOCA2 pool seell ~d ag/Sallback loads on piping and componentslocated between the main vent discharge outlet and the suppression pool water upper surface.

LOCA3 Pool swell ~im act loads on piping and components located above the suppression pool water upper surface.

LOCA4 Oscillating pressure induced loads on submerged pipinq and components during condensation oscillations.

LOC A5 Buildinq motion induced loads from chugging.,

LOCA Vertical and horizontal loads on main'vent piping.

LOCK Annulus pressurization loads.

SBA The abnormal transients associated with a Small Break Accident.

IBA The abnormal transients associated with an Intermediate Break A cci den t.

All ASIDE Code Class 1, 2, and 3. piping systems which are required to function for safe shutdown under the postulated events shall meet the requirements of HRC's 'dInterim Technical Position Punctional Capability of Passive Components" - by MEB'.

The most limitinq case of load combination among LOCAl throuqh LOCA>

REV. 6, 4/82

L 0'

TABLE 5-6 LOAD COMBINATIONS AND ALLOMABLE STRgSSJS-'FOR THE-El'gCTRICATA RACEMAY SYSTEE Load Combination- //lovable Stresses 1 D+L+SRV F 2 D+I +E Note 2

3. D+E ~ +SRV+LOCA Note 2 NOTES:

1.

2.

'or notations, see Table For detailed discussion, 5-2.

see Subsection 3.7b.3.1.6.1 of the SSES FSAR.

REV. 6, 4/82

TABLE 5-7 LOAD COMBINATIONS AND ALLOWABLE STRESSES FOR. HVAC DUCTS 'AND SUPPORTS~

Ducts Load Combination Allowable Stresses 1 D+L+SRV >s

2. 9+PM +SRV FS 3 D+PT 5FS 4 D+PM + E 25FS ¹

5. D+PM +E+SRV Note 1

6. D+PM +E'+SRV Note 1 7 D+PM +PA +E~+SRV+LOCA Note 1

8. When protection against tornado depressurization is required.

D+PP+W D +SRV+LOCA Note 1

9. For ducts inside drywell of containment, the following additional load combination is also appliable:

D+HA +Pp +P A +E +SRV+LOCA Note 1 Duct Supports 1 D+L+SRV FS 2 D+E 25Fs 3 D+E+SRV 1 'ote 4 D+E'+SRV+LOCA Note

¹ This value shall be F~ for transverse and longitudinal bracxnq and their connec tions.

Note -1: In no case shall the allowable stress exceed 0.90+ in bendinq, 0.85F> in axial tension or compression, and 0.50+ in shear. Wh'ere the design is qoverned by requirements of stability (local or lateral buckling),

the actual stress shall not exceed 1.5FS Rev. 7, 6/82

TABLE 5-7 (Cont.)

Notations Dead Load Live Load Po Duct Normal Operating Pressure Load

.P.

T Duct Test Pressure Load Desiqn Basis Accident Pressure Load Duct Maximum Operating Pressure Load, excludinq PA 6 PT, e.g., Pan Cutoff Pressure Load E "Operatinq Basis Earthquake" (OBE) load E t "Safe Shutdown Earthguake" (SSE) load MD Tornado Depressurization Load HA Forces due to thermal expansion of HVAC ducts under accident conditions SRV Safety Relief Valve Loads (Hydrydynamic Loads)

LOCA Loss of Coolant Accident Loads (Hydrodynamic Loads)

Allowable Stress for Steel, governed by AISI or AXSC Codes, as Applicable Yield Strength for Steel (ASTM'specification minimum)

Rev. 7, 6/82

CHAPTER 6 DESIGN CAPABILITY ASSESSMENT CRITERIA TABLE OF CONTENTS 6 1 CONCRETE CONTAINMENT AND REACTOR BUILDING CAPABILITY ASSESSMENT CRITERIA 6.1.1 Containment Structure Capability

'ssessment Criteria 6.1.2 Reactor Building Capability Assessment Criteria 6.2 STRUCTURAL STEEL CAPABILITY ASSESSMENT CRIT ERXA 6 3 LINER PLATE CAPABILITY ASSESSMENT CRITERIA 6 0 DOWNCOMER CAPABILITY ASSESSMENT CRITERIA 6 5 PZPINGiQUENCHER AND QUENCHER SUPPORT CAPABILITY ASSESSMENT CRITERI'A

6. 6 NSSS CAPABILITY ASSESSMENT CRITERIA 6.7 EQUIPMENT CAPABILITY ASSESSMENT CRIT ERIA 6.8 ELECTRICAL RACEWAY SYSTEM CAPABILITY ASSESSMENT CRITERIA 6 9 HVAC DUCT SYSTEM CAPABILITY ASSESSMENT CRITERIA

/

Rev. 2, 5/80

6 0 DESIGN CAPABILITY ASSESSMENT CRITERIA The criteria by which the design capability is determined are discussed in this chapter Design of the SSES is assessed as adequate when the design capability of the structures, piping, and equipment is greater than the loads (including LOCA and SRV discharge) to which the structures, piping, and equipment are subjected. Loadinq combinations are discussed in Chapter 5. The marqins by which design capabilities exceed these l.oadings are discussed in Chapter 7, Design Assessment.

Rev. 2, 5/B0

6 1 CONCRETE CONTAINNENT AND REACTOR BUILDING CAPABILITY ASSESSNENT CRITERIA The acceptance criteria detailed in the SSES .FSAR Section 3.8.1.5 have been usM to assess the structural integrity of the containment and internal structures. No changes are made in these acceptance criteria when the effects of the dynamic SRV discharge and LOCA loads are included.

The acceptance criteria for Seismic Category I structures presented in the SSES FSAR Subsection 3.8.4.5 have been used to assess t'e structural integrity of the reactor building and its components.'o change is made in these acceptance criteria when tge effects of the dynamic SRV discharge and LOCA loads are included.

The allowable stresses for structural steel in the containment and the reactor building are given in Table 5-2. These criteria apply to the suppression chamber steel'columns, the downcomer bracing, and the reactor building structural steel.

6-4

6 3- LINER PLATE CAPABILITY ASSHSSHFNT CRITERIA The strains in the liner plate and anchorage system fields and anchors) from self-limiting loads such as dead load, creep, shrinkage, and thermal effects are limited to the allosable values specified in Table CC-3720-1 of Reference 30, and the displacements of the liner anchorage are limited to the displacement values of Table CC-3730-1 of Reference 30.

Primary membrane stresses in the liner plate and anchorage syste m

( tf elds and anchors) from mechanical loads such as SRV discharge and chugging are checked according to Subsection NE-3221.1 of Reference 29. Primary plus secondary membrane plus bending stresses are checked according to Subsection NE-3222.2 of the same code. Fatigue strength evaluation is based on Subsection NE-3222.4 Allocable design stress intensity values, design fatigue curves, and material properties used conform to Subsection NA, Appendix I of Reference 29.

The capacity of the liner plate anchorage is limited by concrete pull-out to the service load allovables of concrete as specified in Ref erence 31 I Rev. -2, 5f80

6,0 DONNCO~NR CAPABILITY ASSESSNFNT CRITERIA The allowable stresses for the downcomers are given in Table 5-3.

These allowable stresses are in accordance with Reference 29; Subsection NE. As permitted by Subsection NE-1120 for NC components, the downcomers are analyzed in accordance with Subsection NB-3650 of Reference 29; however, the lower allowable stresses, Sm, f rom Table j:-10. 1 for MC components are used when performing the analysis.

Rev. 2, 5/80

6 5 PIPZNGi QUENCHER, AND QUENCHER SUPPORT CAPABILITY

~: KSSESSH~ET .CRXTEBXA Piping in the containment and reactor building. is analyzed in accordance with Reference 29 Subsections NB3600, NC3600, and, ND3600 for, the loading -described in Subsection 5.5.

The quencher is designed in accordance with Reference 29, I Subsection NC3200,for loading discussed in Subsection 5.5.3., The quencher support is designed in accordance with Subsection NP3000 of Reference 29.,

6-7 Rev. 2, 5/80

6 6 - NSSS CAP ABXLETY ASSESSMENT CRZTERIA The capability assessment criteria used for the analysis of NSSS pipinq systems, reactor'ressure vessel (RPV), HPU supports, RPV internal components and, floor structure mounted equipment are shown in Table 5-5, Load Combinations and Acceptance Criteria.

Table 5-5 is in agreement with a conservative general interpretation of the HRC technical position, "Stress Limits for ASME Class 1, 2 and 3 Components and Component Supports of Safety-Related Systems and Class CS Core Support Structures Under Specific Service'Loading Combinations."

Peak response due to related dynamic loads postulated to occur in the same time frame but from different events are combined by the square-root-of-the-sum-of-the-squares method (SRSS) . A detailed discussion of this load combination technigue is presented in Reference 80.

REV. 6, 4/82 6-8

6-7 BALANCE OF PLANT~BOP) EQUIPHENT CAPABILITY 'ASSESSNENT CRITERIA

6. 7. 1. 1- Seismic Category I BOP eguipment located within the containment, reactor and control building are assessed for load combinations shown in Table S-4; In these load combinations, seismic and hydrodynamic loads are generally combined using the absolute sum method.

6712 However, for the <<marginal<< cases the responses of the

<<dynamic<< events (Seismic, SRV, LOCA) are combined by the square root of the sum of the squares (SRSS) method before adding these values to the other loads by the absolute sum (ABS) method. The maximum loading effects of both the horizontal and vertical directions are considered as arisinq from simultaneous excitation in all three principal directions for all combinations involving dynamic loads as detailed in Subsection 7.1.7.4 1.3.

6 7.2 Testing 6.7.2.1 When equipment is qualified by testing, the test motions have simulated the combinations and damping. The eguipment have remained operational and functional, before, during and after such tests.

{a) OBE alone 1/2% damping (b) SSE alone 1% damping (c) SRV alone 2% damping (d) LOCA alone 2% dampinq (e) OB E4SR V+LOCA 2'5 d,amping (f) SSE+SRV+LOCA 2% damping 6-7 2 2 Cases (a) and (b) are covered in the FSAR. Cases (c) f and (d) are covered in the test evaluation or (e) and (f) . Test requirements are depicted by tests response spectrum (TRS) for a qiven damping value. Equipment is deemed to he qualified if the equipment did not fail or malfunction during the test and the TRS envelope the required response Spectrum (RRS). The RRS for cases (e) and (f) are obtained by combining the response spectrum of the individual components of each event by adding the larger of the horizontal responses to the vertical responses on an absolute sum basis. However, for marginal cases the sguare root of sum of the squares (SRSS) method is allowed for the individual dynamic events and components.

REy. 6, 4j'82 6-9

6 8 ELECTRXCAQ gACBMAY S YST Bll CA PABXLITY ASS ESSMgÃT CR XTERXA The allowable stresses for the Blectricl Race@ay Systea are contained in Table 5-6.

REV. 6$ 4/82 6-10

6 9 ~ DUCT SYSTEM/ C~A$ BILITX- ASSESSllgNT C RITRRIA

~

The allowable stresses for the miscellaneous steel for the HVAC duct system are given in Table 5-7.

Rev. 7, 6/82 6-1 1

CHAPTER 7 DESIGN ASSESSMENT TABLE- Og CONTENTS ASSESSMENT MET HODOLOGY 7 1-1 Containment and Reactor Building A ssessment Methodoloqy

7. 1.1 1 Containment Structure 7.1 1.1 1 Hydrodynamic Loads 7 1-1 1 1.1 Structural Models 7 1 1.1 1 2 Damping 7 1 1 1 1 3 Fluid-Str uctur e Interact ions 7 1 1 1 1.4 Supplementary Computer Program

7. 1. 1. 1. 1. 5 Load. Application 7.1 1.1.1.5 1 SRV Discharqe loads 7.1 1.1.1 5 2 LOCA Related Loads 7 1 1 1 1.6 Analysis .

7.1-1.1.1 6 1 Response Spectrum Analysis 7~ 1.1 1.1 6 2 Stress Analysis 7~ 1 1.1 2 Seismic Loads 7.1 1.1 3 Static and Thermal Loads 7.1 1.1 4 Load Combinations 7 ~ 1 1.1 5 Design Assessment 7 1 1 1.6 Equipment Hatch 7.1.1 1 6 1 Structural Model 7 1.1 1.6 1 Loads and.Load Combinations 7 1 1 1 6 3 Design Assessment 7.1 1.2 Reactor and Control Building 7 1.1.2 1 Hydrodynamic Loads 7.1.1 2 1 1 Structural Model

7. 1. 1. 2. 1. 2 Load Application 7~1 1 2 1 2 1 SRV Discharge loads 7 1 1.2 1 2 2 LOCA Related Loads 7.1 1 2 13 Analysis 7 1.1 2 1 3.1 Response Spectrum Analysis 7 1.1 2 1 3 2 Stress Analysis

7. 1.1 2 2.2 Seismic Loads

,.7-1-1. 2-3 Static and Thermal Loads 7-1 1 2 Load Combinations 7.1.1. 2.5 Design Assessment 7 1-2 Structure Steel Assessment Meth odology 7~ 1 2 1 Dovncomer Bracing 7.1 2 1 1 Bracinq System Description

7. 1.2.1. 2 Structural Models 7 1.2. 1.3 Loads 7-1.2.1 3 1 SRV Discharge Loads 7 1 2 1 3 2 LOCA Related Loads 7 1 2 'l 3 3 Seismic Loads 7 1 2 1 3 4 Static 6 Thermal Loads 7 1 2 1 4 Load Combinations

7. 1.2. 1. 5 Design Assessment 71.2 2 SRV Support and Column REV. 6, 4/82 7" 1

7 1 2 2.1 Description of SRV Su pport Assem bies and Suppression Chamber Columns 7 1 2 2 2 Structural Models 71.223 Loads71-223 SRV Discharqe Loads 7 1 2 2 3.2 LOCA Related Loads 7 1 2~2 3 3 Seismic Load

7. 1.2-2.3 4 Static Load 7.1.2 2 3.5 Load Combinations 7.1.2 2.3.6 Design Assessment 7 1 2 3 Openings ia Containment Liner 7 1 2 3 1 Equipment Hatch-Personnel Air Lo ck 7.1232 CRD Removal Hatch, etc.

7.1-2 3 3 Refueling Head 6 Support Skirt 7 1 3 Liner Plate Assessment Methodolo gy 7.1.4 Downcomer Assessment Methodology 7 1.4 1 Downcomer System Description 7.1 4 2 Structural Model 7 1 4 3 Loads and Load Combinations 7 1.4 4 Design Assessment 7 1 4 5 Fatigue Evaluation of Downcomers in Wetwell Airspace 7.1 4 5.1 Loads and Load Combinations Used for Assessment 7.1 4 5 2 Acceptance Criteria 71 453 Method of Analysis 71454~

Results and Design Margins 7 1 5 BOP Pipinq and SRV System Assess ment Methodology 7 1 5 1 Fatigue Evaluation of SRV Discha rge Lines in Wetwell Air Volume 7 1.5 1 1 Loads and Load Combinations Used for Assessment 7.1 5.1 2 Acceptance Criteria

7. 1. 5. 1. 3 Methods of Analysis 7.1 5 1 4 Results and Design Margins 7 1 6 NSSS Assessment Methodology 7 1.6.1 NSSS Qualification Methods 7 1.6.1 1 NSSS Pipinq 7 1.6 1 2 Valves 7 1-6. 1. 3 Reactor Pressure Vessel, Support s and Internal Components

7. 1.6 1 4 Floor Structure Mounted Eguipmen 7 1 6 1 4 1 Qualification Methods

7. 1.6. 1 4.1 1 Dynamic Analysis 7 1 6 1 4 1 1 Methods and Procedures 7 1.6. 1 4 1 2 Testinq 7.1 6 1 4 1 3 Combined Analysis and Testing 7 1 6 142 Computer Programs 7 1.7 Balance of Plant (BOP) Eguipment Assessment Methodology 7 1 7 1 Hydrodynamic Loads 71.711 SRV Discharge Loads71-712 LOCA Related Loads 71.72 Seismic Loads 7173 Other Loads 7 1 7 Qualification Methods 7.1 7 1 7 4.1 7.4 1 1 Dynamic Analysis Methods and Procedures e-REV. 6, 4/82 7-2

7. 1.7. 0 1. 2 Appropriate Dampinq Values 7.1.7.Q 1 3 Three Components of Dynamic Motions

7. 1.7. 4 2 Testing

7. 1.7. 4 3 Combined Analysis and Testing 7 1 8 Electrical Raceway System Assessment Methodology 7.1.8.1 General

'. 1.8.

7.

7.

1.8.

1.8.

2 2.

2.

1 2

Loads Static Loads Seismic Loads 6 7 1.8 2 3 Hydrodynamic Loads 7 1. 8.3 Analytical Methods

7. 1.9 HVAC Duct System Assessment Nethodology 7.2 DESIGN CAPA BXLITY MARGINS

7. 2.1 Stress Marqins 7 2.1. 1 Containment Structure

7. 2. 1. 2 Reactor and Control Building 7.2.1.3 Suppression Chamber Columns 7.2.1. 4 Downcomer Bracing

7. 2.1. 5 Liner Plates

7. 2. 1. 6 Downcomers 7 2.1.7 Electrical Raceway System

7. 2.1. 8 HVAC Duct System 7.2.1.9 BOP Equipment 6

7.2. 10 1 NSSS Equipment 7.2.1. 11 NSSS and BOP Piping

[s 7.2.2 Acceleration Response Spectra 7.2.2. 1 Containment Structu e 2 7.2.2. 2 Reactor and Control Building 7.2 3 Containment Liner Openings 7.2.3. 1 Equipment Hatch.- Personnel Airlo"k

7. 2.3-2 CRD Removal Hatch, etc.

7. 2.3. 3 Refueling Head and Support Skirt P

7.3 FIG URFS Rev. 8, 2/83

CHAPTER 7 FIGURES Number 'Fit le 7" 1 3-D Contaiment Finite Element Model,(ANSYS MODEL) 7" 2 Equivalent Modal Damping Ratio vs. Modal Frequency For Structural Stiffness Proportional - Damping 7-3 Finite Element Soil - Structure Interaction Model 7-4 Containment Response Analysis Containment Stress Analysis Finite Element Containment Equipment Hatch 'Hodel 7-7 Reactor Buildinq Response Analysis 7-8 Reactor Building Stress Analysis 7-9 Downcomer Bracinq System - Plan View 7-10 Downcomer Bracinq System Connection Details 7- 11 Downcomer Bracinq System Computer Model 7- 12 SRV Support System Plan View 7-13 SRV Support System Details 7-14 Finite Element ModeL of Column 7- 15 Finite Element Model of Column 7-16 General Arrangement - Personnel Lock 7- 17 Equipment Door Details 7-18 CRD Hatch Details 7-19 Refuelinq Head Details 7-20 Liner Plate Hydrodynamic Pressure Due to Chugging 7-21 Liner Plate Pressure - Normal Conditions 7-22 Liner Plate Hydrodynamic Pressure Due to Chugging and SRV 7-23 Liner Plate Pressure - Abnormal Condition 4

7-24 Downcomer with Vacuum Breaker and Detail of Cap Rev. '6, 4/82 7-4

PIGURQS (Cont.)

Number T tie 7-25 Dovncomer Without Vacuum Breaker 7-26 Location Where Dovncomer k

Fatigue Analysis was Performed 4

REV. 6, 4/82 7-5

CHAPTER 7, TABLES Number Title 7-1 Maximum Spectral Accelerations of Containment Due to SRV and LOCA Loads at l% Damping 7-2 Naximum Spectral Accelerations of Reactor and Control Buildings Due to SRV and LOCA at 1% Damping 7-3 Usage Factor Summary of Dovncomers 7-4 Usage Factor Summary of SRV Discharge Lines 7-5 Dovncomer and Bracing System Nodal Prequencies REV.', 4/82 7 6

7 0 DESIGN ASSESSMENT Loads on SSES structures, piping, and eguipment are defined in Chapter 4. The methods by which these loads are combined are discussed in Chapter 5.'he criteria for establishing design capability are stated in Chapter 6.

This chapter describes the assessment of the adequacy of the SSES design by comparing design capabilities with the loadings to which structures, pipinq, and components are subjected and demonstrating the extent of the design margin. The f irst section of this chapter discusses the methodology by which design capability and loads are compared. The second section summarizes the results of these comparisons.

.Rev. 2, g/80 7-7

7 1 ASSESSMENT METHODOLOGY 7.1.1 Containment and gegctog Bgildi~n Assessment gethodologg-

7. 1.1. 1. 1 Hydrodynamic Loads The dynamic analysis for the structural response of the containment and internal structures due to the SRV discharge loads and LOCA loads is performed using the finite element 61 ~ I method. The ANSYS (see Ref erence 75 and 76) finite element computer program was chosen for the transient dynamic analysis.

Piqure 7-1 shows the ANSYS finite element model. Beam elements and spar elements are used for the stabilizer truss. Lumped mass elements are used for the RPV internals and suppression pool fluid. Sprinq-damper elements are used to model the rock foundation. The ANSYS model includes a total of 761 elements and 200 dynamic deqrees of freedom.

The soil structure interaction is taken into consideration by modelling the soil usinq a series of discrete springs and dampers in three directions as shown in Pigure 7-1. The properties of the discrete springs and dampers are calculated based on the formulae for lumped parameter foundations found in Reference 33.

The validity of this soil model is proven by comparing the results with those of an independent model which represents the soil by finite elements.

7.1.1.1.1.2 ~Dan ing-Structural Damping The equations of motion for a discretized structur'e must include a term to account for viscous damping that is linearly proportional to the, velocity. The equations of motion for a damped system are.

IN] [Pj + Ic] fr] + IK] [r/= f>(<)J where t'C] is the viscous damping matrix.

A viscous damping matrix of the form ICj = a [M] + B gj ~ 'as used (Refer'enc'e '53) .

Where ~ and B are proportionality constants which relate damping to the velocity of the nodes and the strain rates respectively. This damping matrix leads to the following relation between a and B'nd the damping ratio of the ith mode Ci:

c, = /~. + B /2 REV. 6, 4/82 7-8

where wi is the natural frequency of the ith mode.

usual- case of only structural damping, a = 0 and therefore

'or the 2C i /w,i Since only a sinqle value of g is permitted in the ANSYS input, the most dominant natural frequency of the structure is selected for the computation of g {See Reference 54) .

A value of p equal to 0.00063 is used in the ANSYS model which corresponds to structural modal. damping of approximately 4 percent of critical at 20 Hz which is the most dominant natural frequency of the-structure.

Fiqure 7-2 shows modal damping ratio versus modal frequency for structural stiffness-proportional-damping.

b. Soil Sprinqs and Radiation Damping The elastic half-space theory as described by Reference 33 (BC-TOP-4A Rev. 3) were used to compute the values of the Spring Constants and dampers in the horizontal and vertical directions (KH, K~, CH 6 C> ). The following parameters to represent the rock foundation: are'sed G = Shear Modulus of foundation medium 1 154 x 103 KSX u =Poisson's ratio of foundation medium 0 3 V

S

= Shear wave velocity

= 6180 ft/sec From which we get the following:

KH = 3.37 X 106 K/in C

H

= 1.57 X 104 K-sec/in K = 3.96' 106 K/in v

Cv = 2.72 X 104 K-sec/in The above lumped foundation springs and dampers were then distributed'to every node on the basemat according to the tributary area.

Rev. 2, 5/80 7-9

7.1.1 1.1.3 Fluid-Structure Interaction For the application of SRV loads described in Section 4. 1, a finite element model of the containment vas developed in vhich the suppression pool vater vas included. The water mass constitutes only one seventh of the total mass of the reinforced concrete structure. The model used considers fluid-structure couplinq by lumping the vater mass in the suppression pool at each nodal point of the wetted surface. The weighted area approach is considered to determine the fluid mass at each node of the suppression pool.

For the application of the LOCA steam condensation loads, based on the containment vali pressure time histories calculated by the acoustic methodology (see Subsection 9.5.3.4. and 9. 5.3.4.2),

1 the water mass was excluded. The exclusion of the vater-mass is due to the fact that fluid structure interaction was already considered during the pressure time history calculations (Reference 65) .

7 1.1.1. 1 4 Supplementary Computers Programs Supplementary computer programs vere used for preprocessing and postprocessinq of data qenerated for or by the ANSYS computer program.

A preprocessing program called CHUG vas developed to convert the pressure time history forcing functions into concentrated force time history forcinq functions acting at the associated nodes of the ANSYS model. The program writes the nodal forces onto a

.file for processinq by ANSYS.

A postprocessor program vas developed to calculate the acceleration time history. This program is called DISQ.

reads the structural response displacement time histories It qenerated from ANSYS displacements, scans the maximum displacements and generates the acceleration time histories using the Past Pourier Transformation method.

Bechtel inhouse computer program MSPEC vas used to compute the acceleration response spectrum obtained from DISQ. The program also performs plottinq and broadening of the spectrum.

A computer program ENVLP was developed to generate envelopes of a number of spectrum obtained from MSPEC.

Computer program FORCE vas developed to scan the maximum absolute stresses qenerated by ANSYS stress pass. A further explanation of FORCE is found in Subsection 7.1.1.1. 1.6. 2.

Verification of CHUG, DXSQ, ENVLP and PORCE are available for revie v.

REV. 6, 4/82 7-10

7. 1. 1. 1. 1. 5 Load Agplica tion 7.1 1.1.1.5.1- SRV Dischagge Loads The SRV loads have been defined in Section 4.1 based on KQU -SRV Traces 476, 82 and 35.

To obtain the maximum response of the containment due to bubble oscillation, a vide range of frequency content of the forcing function is considered.

The ranqe of frequencies specified by KQU is betveen 55% and 1104 of the frequencies of the three original traces as present in Subsection 4.1.3.5.

Based on the natural frequencies and the mode shapes of the primary containment as shown in Appendix B-1 five different

~

frequencies in the range specified are selected. in order to obtai.n the maximum structural response. The five frequency values are considered for each of the three original KRU pressure-time history traces vhich result in fifteen pressure-time histories to be considered.

As described in'ubsection 4- 1.3, four pressure distributions depending upon the number of valves actuated are considered; i.e., <<:All valve, ADS, asymmetric, and single valve". Hovever, the aziquth distribution on the periphery indicates that the all valve case qoverns the ADS case for, the symmetric loading and the aEvmmetdic case governs the single valve case fot the asymmetric loathing.'herefore, the design assessment is based on only tvo cases,'.e., "symmetric and asymmetric>>.

7.1 1.1,4.5.2-. T.OCA Related Loads =-

The LOCA loads are based on LOCA steam condensation tests performed by 5'raftvek Union AG (KQU) at their GK5-Il-li Se4tion 9.0 describes the test facility, test matrix, test'acility.

test results and the GKh-XX-M KOCA load definition developed tv re-evaluate SSES for chugging and condensation oscillation.

7. 1. 1. 1. 1. 6 Analyses 7.1.1.1.1.6.1 Time" History Analysis

~

The structural finite element model of containment as outlined in Subsection 7.1.1.1.1.1 is solved by <<Reduced Linear Transient Dynamic Analysis" of the AHSYS computer program. The description of the analysis and the data input are contained in References 75 and 76, respectively.

Por each set of pressure time histories, based on the analytical procedure in,Piqure 7-4, acceleration response spectra vere qenerated at 52 dynamic degrees of freedom in the containment.

Nodal point response spectra generated from several load 7-11

conditions/traces vere enveloped into one set of floor response spectra curves vhich represent SRV and LOCA.

The response spectra vere generated in two pairs of damping values, the low and the high dampings. The lov dampinq values are 0.5, 1 2 and 5 percent of critical, and the high damping

~

values are 7, 10, 15 and 20 percent of critical. The peak frequencies of the spectra are broadened by 15% and 20% for low and hiqh dampinq values, respectively.

Appendix B contains the above response spectra for lov damping values at 9 locations.

7.1.1.1.1.6.2 Stypsis Anallysis The AHSYS computer proqram (stress pass) is used to compute the force and moment resultants due to SRV and LOCA related loads. A postprocessor proqram called>>FORCE>> is developed and used to scan for the maximum absolute values of f orces and moments in the azimuth dire" tion.

A mul.tiplier factor for the force and moment resultants due to SRV loads has been established to cover f or a11 the range of f requencies as specified in Subsection 7. 1. 1. 1. 1. 5. 1. The followinq procedure is used to establish the multiplier.

A statistical analysis of all the forces and moments obtained from the three traces vith varying frequencies in the range specified is performed. Trace number 82 is taken as the base to establish a multiplier factor to cover the'ther 2 traces and the variation of frequencies since highest stresses at most cross-sections.

it is observed to develop the A multiplication factor of 1.7 is established to be applied to the resultant forces and moments from Trace 082 SBV discharge loading.

The forces and moments due to Chugging and Condensation Oscillation (CO) loads are considered. From the response spectra plots of Chuqqinq and CO loads, and 303 were the controlling cases.

it was found that KQU Sources 306 Therefore, these two load cases have been analyzed for stresses in containment. The displacement-time histories obtained from the GKN-II-5 load defini ion (see Subsection 9.5.3) are inputted to AHSYS computer model. A post processor program called SCALE was used to scan for the maximum values of forces and moments in the azimuth direction for each load case. Por the containment sections shovn in Figure A-2, the envelope of force resultants for all the load cases was inputted to the CECAP computer analysis (Re'fer to Plov Chart, Piq. 7-5, for further information).

7.1.1.1.2 Seismic Loads

.Seismic loads constitute a significant loading in the strucutral assessment. The same seismic loads as those used in the initial buildinq design are used. In that design, a dynamic analysis vas made using discrete mathematical idealization of the entire Rev. 8, 2/83 7-1 2

structure usinq lumped masses. The resulting axial forces, ~

moments, and shear at various levels due to the Operating 'Basis Earthquake and the Safe Shutdown Earthquake are used (see section 3.7 of FSAR). The effects of the seismic overturninq moment and vertical accelerations are converted into forces at the elements.

As required by NOREG 0487, the effect of sloshing on the containment due to horizantal and vertical SSE is invetigated by performing a time-history analysis. As described in Subsection 4.2.4.7, pressure time histories due to seismic slash vere ~

qenerated for'. input to the ANSYS model shovn in Figure 7-1.

The response spectra qenerated from the seismic slosh load ~are presented in Figures B-51 to B-58. By inspection ~ the peaks are small.

7 1.1.1.3 Static and Thermal goads The loads under consideration are the static loads (dead load and accident pressure) and temperature loads (operating and accident temperature) vhich are all axisymmetrical.

a To analyze the above static loads, an inhouse computer program PANEL is used. Moments, axial and shear forces are computed by FINEL in an uncracked axisymmetric containment model.

finite-'lement b The operatinq and accident temperature gradients are computed using NE 620 computer program (Bechtel procedure is discussed in Subsection 3.8.4. of the program).'his 1

PSAR ~ 'L ce The results from a, b and the dynamic/seismic analysis are combined and applied to a containment element. The element contains data relative to rebar location, direction and quantity and concrete properties. Mithin that wall element an equilibrium of forces and strains compatibility is established by allowinq the concrete to crack in tension.

Xn this way the stresses in the rebar and concrete are determined. The program used for this analysis is called CECAP. For. further explanation, see Figure'-5.

7. 1. 1. 1. 4 Load Combinations All load combinations from 1 through 7a as presented on Table'-1 have been analyzed. This was done under step c of Subsection 7.1.1.1.3 above. If all the 'SRV actuation cases and chugging-asymmetric-loadinq along with other loads are to symmetric and .be considered, 41 loading combinations would have to be assessed.

I Some of these load combinations have been eliminated by inspection since they are not governing. The f ive basic load combinations which have been assessed and presented in this report are 1, 4, 4a, 5a and 7a.

REV. 6, 4/82 7- 13

The reversible nature of the structural responses due to the pool dynamic loads and seismic loads is taken into account, by considering the, peak positive and negative magnitudes of the response forces and maximizing the total positive and negative forces and moments govern'ing the design.

Seismic and pool dynamic load effects are combined by summing the peak responses of each load by the absolute sum (ABS) method.

This is conservative and the square root sum of squares {SRSS) method is more appropriate since the peak effects of all loads may not occur simultaneously. ~

However, the conservative ABS method is used in the design assessment of the containment and internal concrete structures in order to expedite licensing.

7,1 1,1. 5 Design Assessment Material stresses at the critical sections in the primary containment and internal concrete structure are analyzed using the CECAP computer program. Critical sections for bending moment, axial force and, shear in three directions are located throuqhout the containment structure. The liner plate is not considered as a structural element. The CECAP program considers concrete cracking in the analysis of reinforced concrete sections. CECAP uses an iterative technique to obtain stresses considering the redistribution of forces due to cracking and in the process it reduces the thermal stresses due to the relieving effect of concrete crackinq. The program is also capable of describinq the spiral and transverse reinforcement stresses directly. The input data for the program consists of the uncracked forces, moments and shears calculated by FINEL, ANSYS, and seismic analysis. The loads are then combined in accordance with Table 5-1 with appropriate load factors.

7. 1.1.1.6 Pauioment Hatch There are two equipment hatch openinqs in the containment dryvell vali at approximately El. 723 ft. The openings are 180~ apart and have a diameter of approximately 12 ft. Concrete and rebar stresses around the local hatch area were assessed.,

7. 1.1. 1. 6. 1 Structural Model Pigure 7-6 shows the STARDYNE finite element model that vas developed for analysis of the drywell wall around the hatch opening. The model consists of a section of the drywell wall, diaphragm slab, and vetwell wall vith all boundaries at least two hole diameters avay from the edqe of the opening. All loads can be considered as symmetric about the opening centerline, thus only one half of the openinq was modeled. The model uses quadrilateral plate elements vith both membrane and bending stiffnesses. Uncracked sections vith concrete 'material properties vere used. Loads were applied statically and boundary conditions were chosen to be consistent, with the type of loading applied (Ref. BC Topical Report 45).

REV. 6$ 4/82 7-14

7. 1,1. 1. 6. 2 Loads and Load Combinations Load combinations are as per Table 5-,1. Hydrodynamic loads applied to'he model'boundaries were taken from the force and moment results of the ANSYS containment model described in Section 7.1.1.1.1. Seismic loads were taken from force and moment results of the containment model as given in Section

7. 1.1.1.2. Temperature was considered for the worst case wall qradient. 'I 7.1 g 1 3 1.6 3 Design Assessment Four criti"al sections around the hatch opening were used for assessment. Moment and force resultants from the STARDYNE model were innut to computer proqram CECAP. (CE987) to determine stresses in the concret'e" and rebar..

7.1.1 2 Reactor and Control Buildings

7. 1. 1. 2. 1 Hydrodynamic Loads

7. 1. 1. g. 1. 1 Structural. odel M

The construction of the SSES reactor buildinq is such that no d irect couplinq with the containment occurs. A 2 in. separation joint is kept between the containment structure and the reactor buildinq at all levels where the two structures abut, except at the base slab where a cold joint exists. This arrangement minimizes the transfer of any direct dynamic response to the reactor buildinq from the containment, where the SRV discharge and LoCA related hydrodynamic loads originate.

The horizontal motions of the containment are considered to be fully transferred to the reactormotions building through the cold joint at base slab; but the vertical are attenuated to account for the transfer through the rock under the tvo'structures.average The attenuation has been accounted for by using the,veighted acceleration time histories at different points avay from the containment and to the end of the reactor building boundary. The veiqhted averaqe acceleration is defined as:

n i~1 i N

in vhich ai is the individual acceleration. C >i is the free f ield area on which the acceleration acts and i is the veiqhted average coefficient.

This averaqe time history is applied as an input motion to the reactor buildinq dynamic model. The finite element soil-structure interaction model used fo' the attenuation ~study is shown in Pi'qure 7-3.

Rev. 6, 4/82 7-15

The mathematical models of the reactor and control buldings

'onsist of lumped masses connected by the linear elastic members.

Using the elastic properties of the structural members, the V 4

representative stiffness values for the models are determined.

The models used for desiqn and hydrodynamic load assessment proqrams prior to January 1, 1983, forNorth-South, East-Pest, and Vertical directions are shown in Piqures C-1 C-2, and C-3

~

respectively in Appendix C'. {These models are the same as

~

those used for the seismic analysis prior to January. 1, 1983.}

Subsequently, revised reactor and control building dynamic models for the North-South, East-Pest and Vertical directions have been utilized in desiqns, qualif ications and assessment programs. In t he months preceedinq January 1, 1983, the models were revised as a result of discrepancies in some of the original modeling assumpt ions and representations. Using the rev ised models, a nev set of response spectra was generated. Safety related structures. systems and components that were designed/qualified to response spectra from the previous models were assessed to the

, revised response spectra. Appendix L provides a discussion of the modelinq changes, revised response spectra and a,description of the assessment program.

7~1. 1.2. 1. 2 Load Application 7.1.1.2.1.2.1 - SRV Discharae Loads The axisymmetric,and asymmetric SRV,discharge loadings u ed in the reactor building assessment are described in the chapter 0. 1 of this report. Durinq the axisymmetric loading, only the gross vertical motion of the base slab is transferred to the reactor building. Therefore, the broadened response spectra curves for axisymmetric loadinq qiven in Appendi:ces AC'nd 'L're vertical direction only However, durinq the asymmetric loading, for

'gross vertical motion as well as the gross horizontal motion of the base slab are considered in developinq the vertical and horizontal response spectra curves for the reactor building. The vertical motions are attenuated and the horizontal motions are directly transmitted to the Reactor/Control Building foundation, refer to 7. 1. 1.2. 1. 1 The broadened response spectra curves for asymmetric loading given in the Appendices 'C'nd both vertical and horizontal directions.

'L're for Three different pressure-time history traces (Figures 0-28 through 0-30 of Chapter 0) are used for generating response spectra curves at the base of. reactor building over a wide range of frequencies, i.e., 55% to 110$ of the oriqinal.

7. 1. 1 2. 1.2.2 LOCA Relat.ed Loads Loadinqs associated with Loss of Coolant Accident (LOCA) are briefly hescribed'n 7.1. 1. 1.1.5. 2. The gross vertical and horizontal motions of the Containment,~base slab due to symmetric and asymmetric load conditions are transferred to the Reactor/Control Building. The vertical motions are attenuated 7-16 0

and the horizontal motions are directly transmitted,to the Reactor/Control Buildinq foundation, refer to 7.1.1.2.1.1.

7. 1. t. 2. 1. 3 Anggyses

7. 1. 1. 2. 1. 3~1 Time Hi y to gy, An glygis

~

To develop floor response spectra, a time history analysis of Reactor/Control Buildinq was performed using three separate lumped mass models which simulate the E-W, N-S, and vertical responses. The models are shown on Figures C-1/L-1, C-2/L-2, and C-3/L-3. The analytical procedure is presented in the flow chart in Figure 7-7.

The structural or modal dampinq used in the transient. analysis of the Reactor/Control Buildinq for hydrodynamic loads due to SRV and LOCA is 4 percent of critical dampinq. Based on Regulatory Guide 1.61, this is the damping value recommended for reinforced concrete structures for OBE condition. As this value is used for both Hpset condition (load combinations including OBE) and Faulted condition (load combinations including SSE) it is considered to be conservative.

Like in the containment nodal point response spectra generated

~

from several load conditions/traces were enveloped into one set of floor response spectra curves which represented SRV and LOCA.

For analyses utilizing the models pr sented in Appendix C, the dampinq values included in generatinq the floor response spectra and broadening of the peak frequencies of the spectra are the same as in the containment structure.

Appendix C contains the floor response spectra based on original models for low damping values for SRV and LOCA. Appendix L contains the floor response spectra based on the revised models for low dampinq values for SRV and LOCA.

7.1.1.2.1 3.g Stress Analysis The largest responses at the reactor buildinq base due to all the hydrodynamic loadings are used to obtain forces and moments in the members of the reactor building. The damping values are 2%

and 5% for load combinations involving OBE and SSF/LOCA respectively. For the first part of the, analysis, the Bechtel Program CF, 917 is used to do the modal analysis for the vertical, the East-West and the North-South directions. The results of these analyses are used for input to the Bechtel Program CE 918.

Another input to proqram CE 918, is the envelope of the acceleration response spectra of the gross motion time-histories due to KWU Sources 303, 305, 306, 309 and 314, symmetric and asymmetric load cases. These are obtained from steps 12 and 15 of Fiqure 7-4. The analysis determines member axial forces, shear forces, and bending moments. The analytical procedure is presented in the flow chart in Fiqure 7-8 The following load cases are considered.

0 Rev. 8, 2/83 7-17

1. Condensation-Oscillation vertical .for 2$ and 5% dampings.

2a. SRV vertical symmetric and asymmetric for 2% and 5%

dam pin qs.

2b. SRV North-South asymmetric for 2% and 5% dampings.

2c. SRV East-West asymmetric for 2% and 5% dampings. Case 2c involved four separate conditions depending on the positions of the Reactor Buildinq crane.

3a. LOCA vertical symmetric and asymmetric for 2% and 5%

dam pin qs.

3b.- LOCA North-South symmetric and asymmetric for 2% and 5%

dampinqs.

3c. LOCA Fast-West symmetric and asymmetric for 2X and 5%

dampinqs.

The combined forces and moments in the members of the models presented in Appendix 'C'ue to LOCA, SHV, and seismic loads for both 2'%nd 5% damping values in each of the vertical, Fast-lest, and North-South directions were determined (see Figures E-23 thru E-32). The stress analysis for the revised models is discussed in Appendix L.

The reactor buildinq superstructure steel was analyzed separately using a 3-D finite element lumped mass model. The model is shown in Figure E-21. The bridqe crane and crane girders were also modeled. The dynamic analysis was done usinq the time-history method for seismic loads and response spectrum method for hydrodynamic loads with Bechtel computer program BSAP. Member forces and moments were qenerated for several different crane and trolley positions. In qeneral, the members experienced their hiqhest stresses when the bridge cranes were positioned such that the maximum possible tributary load is distributed to the columns. The critical case is when bridge crane bumper strikes on one side of the superstructure during SSE or OBE. The results are described in Subsection 7. 2. 1.2.

The refuelinq pools and qirders were analyzed separately using a 3-D finite element model. The structure contains the surqe tanks vault, fuel shipping cask storaqe pools spent fuel storage pool, reactor well, and the steam dryer and separator storage pool.

For refuelling conditions, all compartments are considered full of w'ater with the exception of the surge tanks vault, which is empty. For operatinq condition, only the spent fuel storage pool and'the fuel shippinq cask storage pool are full,of water while the remaininq compartments are empty. Water mass was lumped at the compartment floors for the dynamic analysis.

The dynami" analysis was done using the response spectrum method with the computer program STARDYNE. Static and thermal analyses were also performed on STARDYNE program.

Rev, 8, 2/B3 7-18

'I The analysis was performed for critical load combinations which were established by inspection. The results are described in sub sect ion 7. 2. 1. 2.

The box sanction columns supportinq the refueling pool girders were included in the finite element model of the refueling pool analyzed above. The displacements and reactions obtained from

'he above model were used to assess the'structural strength and stability of the columns.

7. 1. 1.2. 2 Seismic Loads The seismic analysis methodology is discussed in the subsection 3.7b.2.1 of the FSAR.

7.1.1~2. 3 St atic ggd Thygma.l goads The static loads are discussed in the subsection 3.8.4.4 of the PSAR.

7. 1 1.2. 4 Load Combinations All individual loads are combined with the appropriate load factors as shown in Table 5-1.

Steel stru"tures are checked for the load combination listed in Table 5-2.

I

7. 1. 1.2. 5 Design Assessment Critical sections. for bendinq moment, axial force and shear in all three directions are located throughout the reactor building.

Design capability at the critical sections is determined and then the desiqn capability is compared with the actual forces and moments acting on the sections under all the load combinations.

This comparison yields design margins. The desiqn margins are discussed in Section 7. 2. 1. 2 Rev. 6, 4/82 7-1 9

7 1. 2 . Structural, Steel Assessment methodology

7. 1. 2. 1 Downcomer Bracing

7. 1. 2.1. 1- . Bracing System Description There. are 87 downcomers which extend vertically f rom the r 'I diaphragm slab to El. 660 -0" in the wetwell, which is

~

approximately 12 feet below normal water level. The five vacuum breaker downcomers have been capped (see Figure 7-25), however, with reqards to the bracing system, these five downcomers still at provide vertical and lateral support, since they were capped the downcomer exits. Downcomers are 24" O. D. pipes with 3/8 inch wall thickness, and are embedded in the diaphragm slab.

Downcomers are separated into four independent quadrants. At El.

668'-0" all downcomers within a quadrant are tied together laterally with a bracinq system consisting not of 6 inch O.D. XX-connected to either strong, pipes. The bracinq members are the wetwell wall or pedestal, thus eliminatinq stresses due to thermal expansion and wetwell wall displacement during hydrodynamic loads. The downcomers support the bracing vert ica l ly. The bracinq connections consist of 1 /2 '~ ri.ng plates and vertical stiffeners. The SRVD lines are not connected to the bracinq. Figures 7-9 and 7-10 Sheets 1-3 show a plan view of the bracinq system and the bracinq connection details, respectively.

7. 1 2.1.2 Strugtugal models A 3-D STARDYNE finite element model of both the bracing and downcomers was developed for analysis of both the downcomers and hracinq. The worst case quadrant of the four was chosen for modelinq (3 ADS lines in the vicinity of the quadrant). The chosen quadrant extends from containme'nt radial of 345~ to radial of 66.7>. This quadrant consists of 23 downcomers modeled as pipes and havinq fixed boundary conditions at the diaphragm slab.

Bracinq members are modeled as pipe elements between downcomers usinq the actual brace member lengths. Beam connector elements extend from the node at the center line of each downcomer to the end of the brace member. Connector elements have equivalent section properties chosen so as to match stiffnesses determined analytically from the finite element model of the bracing connections described later. A lumped water mass consisting of two times the downcomer or bracinq pipe volume (one time for the virtual mass effect and one time for; the contained fluid) is used for nodes below the water level to account for the effect due to fluid-structure interaction. The model consists of 323 nodes, 251 pipe elements, 88 beam elements, and 276 dynamic degrees of freedom for reduced eigenvalue solution (STARDYNE HQR) . Total weiqht considered in the model is 214.5 kips. Figure 7-11 (Sheets 1 F. 2) shows the model.

A separate RSAP finite element model was developed for assessment of the bracinq connection and downcomer in the vicinity of the connection. Figure 7-11, Sheet 3 shows the model. A section of the downcomer at the brace "levelt"is modelled with plate elements.

7-20 Rev. 6, 4/82

Boundari'es'of'he dovncomer vere taken sufficiently far avay from the ronnection to eliminate their influence. The connector plates, top partial plate's, main ring plates, vertical stiffeners,'nd top rinq plates veremember'orces modeled vi'th plate elements.

(see Figure 7-13, Sheet 3j. Brace from the STARDYNE dovncomer and bracinq analysis vere used as input loads for the assessment of the connection shovn in Figure 7-10, Sheet

3. The BSAP finite element model vas also used to determine the stiffnesses 'of. the connector elements used in STARDYNE.

7~1. 2.1.3'oads The basis for a'll hydrodynamic loads considered, is given in Sections 4 and 9.

7 1.2. 1.3.

~ & 1 SRV Discharge Loads SRV actuation results in fluid pressure loads actinq on the containment, dovncomers, and bracinq. All loads are based on KMU Traces 76, 82, and 35. With respect to the dovncomers and bracinq, two different types of loads can be defined. One type consists of inertia l'oadinq. This is movement of the containment structure due to SHV'luid pressures acting directly on the containment. 'he response spectrum method is used for analysis of this loadinq by applying the diaphragm slab spectra (El. 702 ~-

3w, see Appendix B) due to SRV to the STARDYNE model.

l The second type 'of loads are described as submerged structure loads. The'se loads are due to the direct fluid pressures acting on the downcomers and bracinq. As described in Subsection 4.1.3.7.3, potential flov theory and the method-of-images vere used to calculate the load time histories for each downcomer in the model. These were applied to the STARDYNE model and a linear transient dynamic analysis vas performed.

7. 1.2.1. 3. 2 LOCA Related Loads Durinq a LOCA several types of loads act on the dovncomers and bracinq. Tvo of these are inertia and submerged structure loads.

These have the same definition as for the SRV case and the analysis is performed in the same manner. This consists of the response spectra method 'for inertia load analysis and linear transient dynamic analysis for submerqed structure loads.

Subsection 4.2.2.5 describe the methodology for determining the downcomer drag loads due to CO and chugging.

The containment response spectra generated for CO and chugging were determined by the methodology documented in Subsection 9.5.3.

In* addition to the above loads, a dynamic lateral load due to chuqqinq at the downcomer tip also occurs. For analyzing multiple dovncomers in a quadrant, the generic multi-vent lateral load definition documented in Subsection 4. 2. 2. 4 is used.

Rev. 6, 4/82 7-2 1

Xn addit'ion,-'as "requ'ired by the NBC, a single vent impulse'ith a.

65 kip am'plitude'nd 3 msec duration is applied one time per LOCA event to any sing'ledowncomer. This is a low probability event and is'only used to'how that. the dovncomer would not fail for one such loadinq:. '*:

For 'both types "of-'i'p loads, several linear transient dynamic .

analyses were 'performed. Loads vere applied in directions, so as to maximize force's and moments in the downcomers and braces.

Air clearing in the dovncomers during a LOCA also produces poolswell drag and fallback loads on the bracing. This load occurs before:Chugging and CO and need not be considered in combination vith those LOCA loads. Bechtel Nuclear Staff defined the pressure time history loads on the braces and they vere analysed locally for these 'loads (see Subsection 4. 2. 1.7) . An overall equivalent static load on the bracing system vas applied to the STARDYNE -model.

K

7. 1. 2~1 3. 3 Seismic Loads 4

The diaphragm slab'response spectra developed for OBE and SSE as descrihed in S'ubsection 3.8.1.4 1 of the PSAB were used as input to the STARDYNE model "to obtain resultant forces in the dovncomers" and'racing.

Tn addition to the inertia loading, seismic sloshinq in the

.,uppression pool imparts loads on the, dovncomers and bracing (see Subsection'.2.4'.7) . The sloshing freguency is very lov and static loads 'ba'sed on the sloshing fluid pressures vere applied to the STAR'DYNE model.

7. 1. 2. 1. 3.4 St'a'tgc and Thermal Loads The dead load of the dovncomers and bracing is considered. The LOCA condition results in the worst temperature loading 4-52, Section 4). A maximum temperature of. 180oP is used (Bef.'igure with 65~ being taken as the stress free condition.

7.1 2.1.4 Load Combinations Load combinations and allowable stresses are in accordance with Subsection 5.2. The stochastic loads, i. e., seismic inertia, and the inertia and submerged pressure loads of SRV and chugging are comhined by'SRSS method. The chugginq lateral load is defined as a sinqle impulse and is added by absolute sum method. The seismic: sloshinq loads are added by absolute sum method due to their lov frequency wave. All the static loads are combined by absolute. sum method. 'Poolswell is not combined with other LOCA loads since it preceeds them (see Subsection 4.2.1) .

7. 1. 2. 1. 5 Desian* Assessment-The results from 'the three dimensional STARDYNE model of the bracing an'd dovncom'ers are" combined to determine the total stress 7-22 Rev. 6, 4/82

due to both axial forces and moments. A comparison betveen the calculated combined stresses and allovables is made and the stress'argins are given in Appendix A.

7. 1. 2. 2 SR V Suooort and Column

7. 1. 2. 2. 1 Description of SR V Support Assemblies and Suppression Chamber Columns In the suppression pool, there are three types of support confiqurations to laterally brace the SRV, discharge lines; tvo are at El. 666~ and'the third is at El. 667'. Each type of support assembly consists of tvo horizontal bracinq members and at least one knee brace member. The support assemblies are connected from the SRV discharge lines to th'e adjacent column (or columns) vith 4-inch diameter double extra stronq pipes.

The support assemblies restrain the SRV discharge lines in a horizontal direction but not in vertical direction. The general plan of. these support assemblies is shown in Figure 7-12 and member conne"tion and the details are shown in Pigure 7-13.

The suppression chamber columns are 42 inch diameter pipes with 1-1/4 inch wall thickness. The columns are attached at the diaphragm slab't El. 700'nd at the basemat at E1. 648'.

7. 1. 2. 2. 2 St ructurgl Nodels a 0 The columns were independently analyzed for static and dynamic loads. The analytical methods used for non-hydrodynamic loads such as dead, live, pressure, temperature, seismic and pipe rupture loads are described in the FSAR, Section 3. 8. 3. 4. 5.

h. For the hydrodynamic SRV loads, the ANSYS computer program vas used. Por the 'hydrodynamic LOCA related loads HASTRAN computer program was used. A typical column model is shovn in Figure 7-14. The total length of the column is divided into beam'lements vhich are gained at, node points. An effective vater mass due to submergence vas also considered.

Dynamic horizontal forces vere applied to the column at the node points belov the water. Time-varying forces and moments in the column vere calculated for each element.

Co Another finite element model vas developed in vhich the SRV lines, the SRV support assembly and the column were included. SRV and LOCA related submerqed structure, loads as veil as the inertia effects from the dynamic loads vere considered From this analysis, the SBV discharge pipe's reactions at the support locations were obtained.

b The assessment of the columns is based on the combination of loads obtained from a, b, and c above. The assessment of the SRV support assembly is based on loads obtained in paragraph c above.

Each of the support types is analyzed separately.

Rev. 6, 4/82 7 23'

Tn order to Be'term'ine the local stresses in the vicinity, of the support assembly on the column wall, the column was elements. modeled withthe NASTRAN computer program using plate finite The model is shown in Fiqure 7- 15.

7. 1. 2. 2. 3 Loads The support assemblies of the SRV discharge lines are submerged structures. They are sub)ected to direct pressure loads from air bubble etc., the reactions from the SRV lines due to SRV discharqe loads, and the inertia loads due to the building response from dynamic loads. Thermal loads are due to increase in pool t'e m pe ra ture during LOCA.

7. 1. 2.2. 3. 1 SpV Discharge Loads The horizontal SRV discharqe pressure-time histories are considered as actinq on the columns, the SRV discharge pipe and the support assemblies. The vertical SRV discharge pressures are considered as acting on the support assemblies alone.

The reactions from the SRV lines obtained from Subsection 7.1.2.2. 2.c are applied to the end of. the SRV support members for computation of lonqitudinal member forces. The direct hydrodynamic pressures due to SRV actuations are applied statically perpendicular to the SRY support members, with a dvnamic magnification fa tors. The SHY hydrodynamic pressures are determined as, defined in Subsection 4.1.3.7. This is done for the romputation of moments and shear forces in the members.

The inertia forces from building responses due to SRV discharge load are also included by usinq the response spectra results shown in Appendix B.

,'lember forces and moments obtained from direct appliration of SRV d ischarqe pressures, reaction forces of SRV pipe line, and th' inertia buildinq responses are combined by absolute sum.

The SRV submerqed structure load definition is based on Subsection 4.1.3.7.

7. 1.2.2.3.2 LOCA Relapsed Loads During a LOCA, several phenomena cause hydrodynamic loads on the SRV support assemblies. The manner in which the LOCA related loads are applied to the SRV support assemblies is exactly the same as described for the SRV loads in Subsection 7. 1.2.2.3. 1 The LOCA related loads used for the bracing are used for the SRV support assemblies, except the lateral tip load due to chuqginq is eliminated.

Among the LO"A related loads, poolswell load and fallback load occur before Chuaqinq and CO and need not be considered in combination with those LOCA loads. The pressure time history loads, due to pool swell. for the SBV assembly supports, were 7-24 Rev. 6, 4/82

determined by linearly reducing the pressure time history, due to poolswell, for the downcomer bracinq, by the ratio of the d ia meters.

7. 1~2. 2. 3 3 Seismic Load The seismic loads on the coupled structure of SRV lines, support assemblies, and columns were obtained by dynamic analysis using the response spectra developed for OBE and SSE as described in Subsection 3.8.1.4.1 of the FSAR.

7. 1. 2. 2. 3. 4 Stat ic Load The dead load, thermal load and bouyancy of the support assemblies were considered.

7.1.2.2.3.5 Load Combinatjons The load combinations and allowable stresses are in accordance w ith Subsection 5. 2.'lthough the loads on the bracinq system under consideration act in a random horizontal directions, each individu'al load is applied to the system in the worst possible direction to find the maximum resultant forces.

7 1 2.2. 3.6- Desian Assessment The combined stresses due to axial forces and bending moments were determined for all bracinq members. Comparison between the resultinq calculated stresses and the allowable stresses has been made. Resultinq stress marqins for the bracing members and their connections are tabulated in Appendix A.

7,1,2,3,1 Equipment Hgtch-Personnel Air Lock The portion of the equipment hatch-personnel air lock not backed by concrete w'as reevaluated for additional loads due to hydrodynamic effects {SRV and LOCA) . This reevaluation was performed by Chicago Bridge and Iron Company (CBI) under subcontract from Bechtel. The general arrangement of the personnel lock is shown in Figure 7-16.

The personnel air lock doors are designed to withstand a pressure of 55 psig in the containment vessel. The door mechanism is designed to seal the door aqainst an internal pressure of 5 psig.

For reevaluation, CBI used their computer program E781 for stati" analysis of shells. The proqram is based on Reference 77.

Equivalent static loads were considered for seismic and hydrodynamic cases usinq peak spectral accelerations. CBI used the hydrodynamic spectra as <given in Appendix C. Design Load combinations qiven in Table 5- 2 were used with modifications for forces on the structure due to thermal expansion of pipes under'ev.

7-25 6, 4/82

accident conditions. Stress limits specific% in the ASME code were used.

CBIis model vas divided into 2 parts:

The first model comprised the 1>> thick cylinder and the 3" thick flanqe extendinq to the parting joint. An axissymmetrical configuration, vas used since the shape of the containment vessel at its intersection with the equipment hatch is conical. No restraints at the junction with the containment vessel vere considered.

The second model included the 3>> thick flange beyond the parting joint, the conical head and a portion of. the personnel lock extending from the interior bulk head to an'appropriate distance beyond..

Kt the flange interface, the seismic, SRV, LOCh jet and pressure

~

loads have a tendency of prying open the door. A meridional force is, therefore, required to permit relatively small radial Reflections and rotations at the interface This force was applied as a restoring force at the parting joint in the form of a meridional force and a transverse shear. Relative displacements were evaluated to assure leaktightness.

The ma jor dead load contribution is in the airlock. Theref ore, dead loads and loads from seismic accelerations vere applied to the second model, as discontinuous loads at the center of gravity o f the air lock.

Loads due to SRV, Seismic and LOCA cases were combined by SRSS.

7 1.2 3.2 CRD Removal Hatch~ Suppression Chamber Access Hatch And Eauioment Hatch These hatches vere subcontracted to CBI for design and analysis for additional SRV and LOCA loads. Designs vere performed manually in accordance with Bechtel specifications and appropriate design codes. Details of the CRD removal hatch and equipment hatch are qiven in Figures 7-17 and 7-18.

7.1.2,3. 3 Refuellina Head and Support Skirt Reevaluation of the refuelling head and support skirt vas performed by CBI under subcontract from Bechtel. Piqure 7-19 shows the refuelling head.

CBI s program

~

E 781 vas used for the static analysis. For dynamic analys is, equivalent pressures from the peak response spectra at El. 778.8 ft. vere used. The static and dynamic stresses vere then combined as per Table 5-2 of this report.

Leak tiqhtness of the flanged joint vas investigated for the various loads and suitable pre-stress was recommended to prevent separation of the flange joint components.

7-26

7,1,3 Liner plate Assessment Methodology FSAR Su'bse" tion 3.8.1 provides a description of the liner plate and anchoraqe system for the containment.

The analysis of the liner plate and anchorages for nonhydrodynamic loads is in accordance with Reference 18.

For the analysis of the liner plate and anchorage for hydrodynamic suction loads, the contributing load on the liner is that due to the net "neqative" pressure.

The loads considered for this assessment are KWU Chugging, KWU SHV, hydrostatic pressure and wetvell air pressure.

Figure 7-20 presents the maximum negative pressure due to KWU chuqqinq vhich vere scanned from the symmetric and asymmetric load conditions of Sources 303, 305, 306 and 309. As can be noted from Piqure 7-20 'race 306 gives the maximum neqative pressure on a.ll locations.

The maximum neqative pressure due to the actuation of all SRV s ~

is -7.8 psi.

The hydrostatic pressure of 24'ater gives 10.4 psi pressure on the base slab liner plate.

The wetvell air pressure is 25 psi due to a small break LOCA.

For normal condition the combination of hydrostatic pressure and the actuation of all the SRV s is considered. The distribution ot this pressure is shown in Figure 7-21.

For abnormal condition, the combination of KWU chuqqinq, SRV, hydrostatic pressure and wetwell air pressure is considered. The phasinq of SRV and. chuqqing events is obtained by aligning the maximum suction peaks. These events are combined by direct addition of pressures as demonstrated in Fiqure 7-22. The total net peak pressures for the abnormal condition are tabulated in Fiqure 7-23. Point 1 in this figure does not lie on pressure boundary and thus, is not critical.

The assessment of liner plate is found in Subsection 7. 2.1.5.

7. 1. 4 Dovncomer Assessment Methodology

7. 1. 4. 1 Downcomez System Description Tn the wetwell, there are 87 downcomers, 82 of vhich function as dry veil vents durinq a LOCA. The other 5 provide vetvell to drywell pressure relief through the tvo vacuum breakers in series mounted on each of them. These five downcomers are capped at the bottom end to protect the vacuum breakers from the cycling due to

~ chuqqinq. Appendix K provides the assessment of cappinq five of 7-27 Rev. 6, 4/82

the eighty-seven downcomers as a fix for VB cycling during c huqging.

Downcomer layout, location of vacuum breakers and the cap arranqement are shown on Piqures 7-9, 7-24 and 7-25, respectively.

7. 1. 4. $ Structural Nodel The downcomers are modeled with the bracing system as described in Subsection 7.1.2.1.2.

The downcomers with the vacuum breakers are included in the STARDYHE model.

An additiona'1 3-D model was developed in which not only the bracinq system and downc'omers as described in subsection 7.1.2.1. 1 were included, but also the vacuum breaker, the vacuum breaker support and a column. This was done in the same quadrant as described in Subsection 7.1.2. 1.1.

7. 1 4. 3 Loads and Load Combinations Loads affectinq the downcomers are the same as those described in Subsection 7 1.2.1.3. Load combinations are given in Table 5-3.

The SRSS sum is used for the dynamic loads, except for the chuqqinq lateral and seismic sloshing loads which are added by absolute sums as. described in Subsection 7. 1. 2.1. 4.

7. 1. 4 4 Design Assessment Reference 30 is used for checkinq the downcomer stresses due to the load combinations qiven in Table 5-3.

7.1 4 5 Fatigue Evaluation of Downcomers In Qetwell Air Volume In an effort to evaluate the steam bypass potenti'al arising from a failure of the downcomers in the wetwell air space, a complete fatigue analysis of the same has been performed'. Sp. cifically, the analysis was performed where the downcomers penetrate the diaphram slab as shown in Piqure 7-26. This analysis considered all the cyclic loadinq acting on the downcomers and is in accordance with the applicable portions of ASNE Code. This evaluation is considered supplemental and does not displace the original desiqn basis for these lines as set forth in the appropriate FSAR/DAR sections.

7. 1.4.5.-1 'Loads and Load Combinations used for Assessment The downcomers are subject to numerous dynamic and hydrodynamic loads .from normal, upset, and LOCA-related plant operating conditions. For purposes of fatique evaluation, the following loads are include: {1) All significant thermal and pressure transients. (2) All cyclic effects due to the hydrodynamic loads includinq SRV actuations, CO and chugging. (3) Seismic 7-2 8

effects. A description of each of these loads is provided in the appropriate DAR sections. The determination of load combinations as veil as number and duraction of each event is obtained from the applicable sections of DFFR, and PSAR.

7. 1.4 5.2 Acceptagre Cgitqria The design rules, as set forth in the ASME Boiler and Pressure Vessel Code,Section III, Subsection NB vere utilized for the fatique assessment. When required, allowables for fatigue stress evaluation vere based on Mill certification reports for d owncomers.

7. 1.4. 5. 3 Methods of Analvsis The SRV discharqe lines and downcomers in the wetvell air volume, were .analyzed for the appropriate load combinations and their associated number of cycles. The combined stresses and correspondinq equivalent stress cycles were computed to obtain the fatique usage factors in accordance with the equations of Subsection NB-3600 of the ASIDE Code.

7. 1.4.5. 4 Results and Design Margins The cumulative usaqe factors for the various loading conditions for the dovncomer (see Figure 7-26) are summarized in Table 7-3.

7.1.5 BOP Pining and I SRV Systems Assessment Methodology The BOP pipinq and SRV systems were analyzed for the loads discussed in Section 5.5 using Bechtel computer programs NE101 and WE632. These programs are described in PSAR Section 3.9.

Static and dynamic analysis of the piping and SRV systems are performed as described in the paragraphs belov.

Static analysis techniques are used to determine the stresses due to steady state loads and/'or dynamic loads having equivalent static loads. The drag and impact loads are applied as equivalent st,atic loads.

Response spectra at the pipinq anchors are obtained from the dynamic analysis of the containment sub jected to LOCA and SRV loadinq. Pipinq systems are then analyzed for these response spectra following the method described in Reference 19.

Time history dynamic analysis of the SRV discharge piping subjected to fluid transient forces in the pipe due to relief valve opening is performed usinq Bechtel computer code ME632.

7,1 5,1 Patigue Evaluation of SRV Discharge Lines in Metwell Air Volume Xn an effort to evaluate the steam bypass potential arising from a failure of the SRV discharge line in the vetvell air space, a complete fatigue analysis of the same has been performed.

Rev. 6, 4/82 7-29

Specifically, structural analyses of all the SRV discharge lines from. the diaphragm slab penetration to the quencher vas performed. Fatique evaluation of fluedhead penetration, elbovs and 3-way restrainst attachment to pipe vas done. This analysis considered all the cyclic loadinq acting on the SRV discharge lines and is in accordance with the applicable portions of ASME Code. This evaluation is considered supplemental and does not displace the original design basis for these lines as set forth in the appropriate FSAR/DAR sections.

7 1.5.1 1 Loads hand Load Combinations Used fog Assessment The SRV discharqe lines are subject to numerous dynamic and hydrodynamic loads from .normal. upset, and LOCK-related plant operatinq conditions. For purposes of fatigue evaluation, the following loads are included: (1) All significant thermal and pressure transients. (2} All cyclic efforts due to the hydrodynamic loads includinq SRV actuations, CO and chuqging and (3) Seismic effects. A description of each of these loads is provided in the appropriate DAR sections. The determination of

-load combinations as veil as number and duration of each event is obtained from the applicable sections of. DFFR and FSAR.

7. 1. 5. 1. 2 Acceptance Critegia The desiqn rules, as set forth in the ASIDE Boiler and Pressure Vessel Code, Section II1:, Subsection NB were utilized for the fatigue assessment, When required, allowables for fatique stress evaluation were based on Mill certification reports for SRV discharge lines.

7. 1.5. 1. 3 Methods of Analysis The SRV discharqe lines, in the wetwell air volume, were analyzed for the appropriate load combinations and their associated number of cycles. The combined stresses and corresponding equivalent "tress cycles were computed to obtain the fatigue usage factors in accordance with the equations of Subsection NB-3600 of the ASME Code.

7. 1.5.1.4 Results and Qesjgn margins The cumulative usage factors for fluedhead, 3-way restraint attachment to pipe and elbow are summarized in Table 7-4.

7. 1.6 NSSS Assessment Methodology "Safety related" General Electric Company supplied NSSS piping and equipment located within the containment and the reactor and control buildinqs are subjected to hydrodynamic loads due to SRV and LOCA discharqe effects principally oriqinatinq in the suppression pool of the containment structure. Section 4.1 and 4.2 describe the methodologies used to define these SRV and LOCA loads, respectively. The NSSS piping and equipment are assessed to verify their ade'quacy to withstand these hydrodynamic loads in 7-30 Rev. 6, 4/82

combination with seismic and all other applicable loads in accordance with the load combinations given in Table 5-5.

The structural system responses for the SRV and LOCA suppression pool hydrodynamic phenomena are generated by Bechtel Power Corporation usinq defined forcinq functions. These structural system responses are transmitted to General Electric in the form of (1) broadened response spectra and (2) acceleration time-histories at the pedestal to diaphram floor intersection and the stabilizer elevation.

The response spectra for pipinq attachment points on the reactor nressure vessel, shield wall and pedestal complex (above the pool area) are generated by General Electric, based upon the acceleration time-histories supplied by Bechtel Power Corporation, using a detailed lumped mass beam model for the eactor pressure vessel internals, including a, representation of the structure. For the assessment of the NSSS primary pipinq (main steam and recirculation) a combination of General Electric and Bechtel developed response spectra are used as input responses for all attachment points of each piping system. For

~ he assessment of the NSSS floor mounted equipment, except the reactor pressure vessel, the broadened response spectra supplied directly by Bechtel are used.

The acceleration time-histories and the detailed reactor Pressure vessel and structure lumped mass beam model are used to generate the forces and moments ac+inq on the reactor pressure vessel supports and internal components. These forces and moments are used for the GE assessment of reactor'ressure vessel supports a nd internals.

The structural system response for the LOCA induced annulus pressurization transient asymmet'ric pressure build up in the annular region between the biological shield wall and the reactor pressure vessel is based on pressure time-histories supplied by Bechtel. These pressure time-histories are combined with get eaction, jet impinqement and pipe whip restraint loads for the assessment. A time-history analysis is performed resulting in accelerations, forces and moment time-histories as well as response spectra at the piping attachment points on the reactor pressure vessel, shield wall, pedestal, pressure vessel. supports and external components (see FSAR Appendices 6A and 6B) .

7. 1.6.1 NSSS Qualification Nethods

7. 1.6.1. 1 NSSS Piping The NSSS piping stress analyses 'are conducted to consider the secondary dynamic responses from: (1) the oriqinal design-basis loads includinq seismic vibratory motions, (2) the structural system feedback loads from the suppression pool hydrodynamic events, and (3) the structural system loads f rom the LOCA induced annulus pressurization from postulated feedwater, recirculation and main steam pipe'breaks.

7-31 Rev. 6, 4/82

Lumped mass models are developed by General Electric for the NSSS primary pipinq systems, main steam and recirculation lines.

These lumped mass models include the snubbers, hangers and pipe mounted valves, and represent the major balance of the plant branch piping connected to the main steam and recirculation systems. Amplified response spectrum for all attachment points with'in the piping system are applied: i.e., distinct acceleration excitations are specified at each piping support and anchor point. The detailed models are analyzed independently to determine the piping system resulting loads (shears and moments) for:

1) ea"h desiqn-basis load which includes pressure, temperature, weiqht, seismic events, etc.,

2) the bounding suppression pool hydrodynamic event; and

3) the annulus pressurization dynamic effects on the unbroken pipinq system.

Additionally, the end reaction forces and/or accelerations for the pipe mounted/connected equipment (valves and nozzles) are simultaniously calculated.

The pipinq stresses from the resultinq loads (shears and moments) for each load event are determined and combined in accordance with the load combinations delineated in Table 5-5. These stresses are calculated at geometrical discontinuities and compared to ASNE code allowable determined stresses (ASIDE Boiler and Pressure Vessel Code, Section III-HB-3650) for the appropriate loading condition in order to assure design adequacy.

Computer codes used to perform the VASSS piping stress analysis are described in FSAR Section 3.9. 1.2.

7. 1.6. 1. 2 Valves The reaction f orces and/or accelerations acting on the pipe mounted equipment when combined in accordance with the required load combinations are compared to the valve allowables to assure design adequacy. The reactor core pressure boundary valves are qualified for operability during seismic and hydrodynamic loading events by both analysis and test. This qualification is unique for each valve.

7.1.6.1.3 Reactor pressure Vessel~ Supports and Tnternal Components The boundinq load combinations for seismic, hydrodynamic and annulus pressurization forces are established within each Ii acceptance criteria range (upset,,emergency a nd faulted) . At the initial analysis step, the loads are conservatively combined using the maximum vertical forces with the maximum horizontal shears and moments from all combinations within each acceptance criteria range. These conservative maximum loads are then compared to generic bounding forces originally used to establish 7-3 2 Rev. 6, 4/82

the component desiqn. When the combined calculated. forces are less than the desiqn forces, then the component is deemed adequate. When the calculated forces are greater than the design forces, then the increased stresses are compared to the material allowables. When the calculated stresses are below, the material allowables, then the design is deemed ad'equate. If the increased stresses are above the material allowables, then the specific load combination is identified and another stress analysis is conducted usinq refined methods, component adequacy.

if required, to demonstrate the I

In certain cases, component test results are combined with analyses to assess component adequacy Fatigue evaluations of the Reactor Pressure Vessel, supports and internal components are also conducted for SRV cyclic duty loads. 'he equipment is analyzed for fatique usaqe due to SBV load cycles based upon the loading durinq the SRV events. SRV fatigue usage factors are calculated and combined with all other upset condition usage factors to obtain a cumulative fatique usage factor.

Computer proqrams used to conduct RPV component analyses are described in FSAR Section 3.9. 1.2.

7.1.6.1.4 Floog Structure amounted Equipment

7. 1.6.1. 4.1 Oualif ication Methods

~he adequacy of, the design of the equipment is assessed by one of the followinq:

a. Dynamic analysis

b. Testing

c. Combination of testinq and anal'ysis The choice is based on the practicality of the method depending upon function. type, size, shape, and complexity of the equipment and the reliability of the qualification method..

In general, the requirements outlined in IEEE-344-75, Reference 55, are followed for the qualification of equipment.

7. 1. 6. 1. 4. 1. 1 Dyna mic Anal ysi s 7.1.6.1.4.1.1.1 methods and Procedures The dynamic analysis of various equipment is classified into three groups accordinq to the relative rigidity of the equipment based on the magnitude of the fundamental natural frequency described below.

(a) Structurally simple equipment - comprises that equipment which can be adequately represented by a one degree of freedom system Rev. 6, 4/82 7-33

(b) Structurally riqid equipment Comprises that equipment whose fundamental frequency is:

(i) greater than 33 Hz for the consideration of seismic loads, and, (ii) qreater than the hiqh frequency asymptate (zpA) of the required. response spectra (RRS) for the consideration of hydrodynamic loads (c) Structurally Complex equipment Comprises that equipment whirh cannot be classified as structurally simple or structurally rigid.

The appropriate response spectra for specific equipment are obtained from the response spectra for the floor at which the eauipment is located in a buildinq for OBE, SSE and hydrodynamic l oads. This includes the vertical as well as both the N-S and E-horizontal di.'rections. For equipment which is structurally simple, the dynamic loadinq (either seismic or hydrodynamic) consists of a static load corresponsing to the equipment weight times the acceleration selected from the appropriate response spectrum. The acceleration selected corresponds to the equipment's natural frequency, if the equipment's natural frequenry is known. If the equipment's natural frequency is not known, the acceleration selected corresponds to the maximum value of. the response spectra.

For eguipment which is structurally rigid, the seismic load rnnsists of a static load corresponding to the equipment weight times the acceleration at 33 Hz, selected from the appropriate

'response spectrum and the hydrodynamic loadinq consist of a static load corresponding to the equipment weight times the accelerations at the ZPA selected from the appropriate response spectrum.

For the analysis of structurally complex equipment,. the equipment is idealized by a mathematical model which adequately predicts t he dyna mi" properties of the equipment and a dynamic analysis is performed usinq any standard analysis procedure. An acceptable alternative method of analysis is by static coefficient analysis for verifying structural inteqrity of frame type structures that can he represented by a simple model. No determination of natural. frequencies is made and the response of the equipment is assumed to be the peak of the response spectrum. This response is then multiplied .by a.static coefficient excitationof 1.5 to take into account the effects of both multifrequency and multimode response.

7. 1. 6. 1. 4. 1. 2 . Testing In lieu of performinq dynamic analysis, dynamic adequacy is estahlished by providing dynamic test data. Such data must conform to one of the followinq:

Rev,. 6, 4/8> 7-34

Performance. data of equipment which has been subjected to equal or greater dynamic loads (considering appropriate frequency range) than those to be experienced under the specified dynamic loadinq conditions.

2 ~ Test. data from comparable eguipmen't'reviously tested under similar conditions, which has been subjected to equal or greater dynamic, loads than those specified.

3. Actual testinq of equipment in operating conditions simulating, as closely as possible, the actual installation, the required loadinqs and load combinations.

A continuous sinusoidal test, sine beat test, or decaying sinusoidal test is used when the applicable floor acceleration spectrum is,a narrow band response spectrum. Otherwise, random motion test (or equivalent) with broad frequency content is used.

The equipment to be tested is mounted in a manner that simulates the actual service mounting. Sufficient monitorinq devices are used to evaluate the performance of the equipment. Pith the appropriate test method selected, the equipment is considered to be qualififed when the test response spectra (TRS) envelopes the "equired response spectra (RRS) and the equipment did not malfunction or fail. A new test does not need to be conducted if equipment requires only a very minor modification such as additional bracinqs or change in switch model, etc., and proper iustification is qiven to show that the modifications do not jeopardize the strength and function of the eguipment.

7 1.6.1.4.1.3 Combined Analysis and Testing There are several instances where the qualification of equipment hy analysis alone or testing alone is not practical or adequate because of its size or its complexity, or large number of similar confiqurations. In these instances a combination of analysis and testing is the most practical. The following are qeneral approaches:

(a) An analysis is conducted on the overall assembly, to determine its stress level and the transmissibility of motion from the base of the eguipment to the critical components. The critical components are removed from the assembly and subjected to a simulation of the environment on a test table.

(b) Experimental methods are used to aid in the formulation of the mathematical model for any piece of equipment. Mode shares and "frequencies are determined experimentally and incorporated into a mathematical model of the equipment.

7.1.6.1.4.2 Computer Programs Computer programs used to conduct equipment analyses are described in FSAH Section 3. 9. 1. 2.

7-35 Rev. 6, 4/82

7,1,7 Balance of plant ~4 lBOPl Equi@ment Assessment Methodoloqv Seismic Cateqory I BOP equipment located within the containment and the reactor and control buildings are subjected to hydrodynamic loads due to SRV LOCA discharge affects principally ori'ginatinq in the suppression pool of the containment structure.

The equipment and equipment support are assessed to verify their adequacy to withstand these hydrodynamic loads in combination with seismic and all other applicable loads in accordance with the load combinations given in Section 5.7.

7.1 7.1 Hydrodynamic loads

7. 1. 7. 1. 1 ~ SR V Disrharqe Loads Loadinqs associated with the axisymmetric and asymmetric SRV discharges are described in Chapter 3 and 4 of this report.

Acceleration'esponse spectra at the various elevations where the equipment are located have been generated for all appropriate pressure history traces (Figures 4-28 thru 4-30 of Chapter 4) for damping values of 1/2%, 1%, 2%, and 5%. These have been enveloped into a single curve for each of the above damping values. Such enveloped curves are "qenerated for each of the N-S, E-W and vertical directions. These curves form the basis for the SRV loads for equipment assessment.

7. 1. 7. 1. 2 LOC A Re la ted Lo ad s Loadings associated with loss-of-coolant accident (LOCA) are described in Section 4.2. Acceleration response spectra at various elevations where the equipment are located have been generated for the LOCA lo ads for damping values of 1/2%, 1%,,2X and 5%. These have been enveloped into a single curve for each of the above dampinq valu es. Such enveloped curves are generated f or each of the N-S, E-W an d vert ical dir ect ion s.

These curves form the basis for the LOCA loads for equipment assessment.

7 1.7.2 Seismic Loads The details of: seismic input and seismic loads are discussed in Section 3.7 of PSAR. The effects of both operating hasis earthquake (OBE) and safe shutdown earthquake (SSE) are considered. These loads are provided in the form of Acceleration response spe" tra at each floor for dampinq values of 1/2%, 1g, 2%

and 5% for each of N-S, E-W and vertical directions.

7. 1 7. 3 Other Loads In addition to hydrodynamic and seismic loads, other loads such as dead loads, live loads, operatinq loads, pressure loads, thermal loads, "nozzle loads and equipment piping interaction loads, as applicable, are also considered.

Rev. 6, 4/82 7-36

7.1 7.4 oualification Methods The adequacy of the design of the equipment is assessed by one of the folowinq:

a. Dynamic analysis

b. Testing under simulated conditions

c. Combination of testing and analysis.

The choice is based on the practicality of'he method depending upon function, type, size, shape, and complexity of the equipment and the reliability of the qualification method.

In qeneral the requirements outlined in IFEE-344-75, Reference 55, are followed for the qualification of equipment.

7. 1. 7. 4. 1 Dynamic Analysis 7.1.7.4.1.1 Methods and Procedures The dynamic analysis of various equipment is classified into three groups according to the relative rigidity of the equipment based on the magnitude of the fundamental natural frequency described below.

(a) Structurally simple equipment.- comprises of that equipment which can be adequately represented by one degree of freedom system.

(b) Structurally riqid equipment Comprises of that equipment whose fundamental frequency is:

(i) greater than 33 Hz for the consideration of seismic loads. and, (ii) qreater than 80 Hz for the consideration of.

hydrodynamic loads.

(c) Structurally Complex equipment Comprises of that equipment which cannot be classified as structurally simple or structurally rigid.

When the equipment is structurally simple or rigid in one direction but complex in the other, each direction may be classified separately to determine the dynamic loads.

The, appropriate response spectra for specific equipment are obtained from the response spectra for the floor at which the equipment is located in a building for OBE, SSE and hydrodynamic loads. This includes the vertical as well as both the N-S and E-horizontal directions.

Rev. 2, 5/80 7-37

For equipment which is structurally simple, the dynamic loading (either seismic or hydrodynamic) consists of a static load correspondinq.to the equipment weight times the acceleration selected from the appropriate response spectrum. The acceleration selected corresponds to the equipment' natural frequency, if the equipments s natural frequency is known. 'Xf the equipment~s natural frequency is not known, the acceleration selected corresponds to the maximum value of the response spectra.

For equipment which is structurally riqid the seismic load consists of a static load corresponding to the equipment weight times the acceleration at 33 Hz, selected from the appropriate response spectrum and the hydrodynamic loading consist of a static load correspondinq to the equipment weight times the acceleration at 80 Hz., selected from the appropriate response spectrum.

For the analysis of structurally complex equipment, the equipment is idealized by a mathematical model which adequately predicts t he dynamic properties of the equipment and a dynamic analysis is performed usinq any standard analysis procedure. An acceptable alternative method of analysis is by static coefficient analysis for verifying structural inteqrity of frame type structures such as members physically similar to beams and columns that can be represented by a simple model. No determination of natural frequencies is made and the response of the equipment is assumed to be the peak of, the response spectrum at damping values as per Section 7.1.7.4. 1.2. This response is then multiplied by a static coefficient of 1.5 to take into account the effects of both multifrequency excitation and multimode response.

7. 1. 7. 4. 1. 2 A ppropgi ate Damping Values The followinq dampinq values are used for the design assessment:

1) Load Combinations involvinq OBE but not hydrodynamic loads 1/2%

2) Load Combinatiosn involvinq SSE but not hydrodynamic loads

3) Load Combinations involvinq hydrodynamic

. loads, or seismic and hydrodynamic loads 2%

If the actual then dampinq value of the equipment is different (from test results) these actual values are used.

7. 1.7.4. 1.3 Thee@ Components of Dynamic Notions The responses such as internal forces'tresses and deformations at any point from the three principal orthogonal directions of

~,he dynamic loads are combined as 'follows:

Rev. 2, 5/80 7-3 8

The response value used is the maximum value obtained by adding the response due to vertical dynamic load with the larger value of the. responses due to one of the horizontal corresponding dynamic load by the absolute sum method.

7.1.7 4.2 Testing Tn lieu of performinq dynamic analysis, dynamic adequacy is established by providinq dynamic test data. Such data must conform to one of the followinq:

1. performance data of equipment which has been subjected to equal or greater dynamic loads (considering appropriate frequency range) than those to be experienced under the specified dynamic loadinq conditions.

2.. Test data from comparable equipment previously tested under similar conditions, which has been subjected to equal or grater dynamic loads than those specified.

Actual testinq of equipment to the reguired load combinations while simulatinq the actual field installation.

A continuous sinusoidal test, sine beat test, or decaying sinusoidal test is used when the applicable floor acceleration spectrum is a narrow band response spectrum. Otherwise,, random motion test (or equivalent) with broad frequency content is used.

The equipment. to be tested is mounted in a manner that simulates the actual service mountinq. Sufficient monitorinq devices are used to evaluate the performance of the equipment. Hith the appropriate test method selected, the equipment is considered to be qualified when the test response spectra (TRS) envelopes the required response spectra (RRS) and the equipment did not malfunction or fail.- A new test does not need to be conducted if equipment requires only a very minor modifications such as' additional bracinqs or change in switch model etc. and proper justification is given 'to show that the modifications do not jeopardize the strength and function of the equipment.

7.1.7.4.3= Comhined Analysis and Testing There are several instances where the qualification of equipment hy analysis alone or testing alone is not practical or adequate because of its size, or its complexity, or large number of similar configurations. Xn these instances a combination of analysis and testinq is the most practical. The following are general approaches:

4 (a) An analysis is conducted on the overall assembly to determine its stress level and the transmissibility of motion from the base of the equipment to the critical components. The critical components are removed from the assembly and subjected to a simulation of the environment on a test table.

7-39 Rev. 6, 4/82

1V Experimental methods are used to aid in the formulation of the mathematical model for any piece of equipment.

and frequencies are determined experimentally and Node'hapes incorporated into a mathematical model of the eguipment.

I 7.1.8 Electrical Raceway System Assessmen t Methodology

7. 1. 8. 1 General The FSAR Subsection 3.7b.3. 1.6 provides a detailed description of the electrical raceway system design methodology. The analysis and desiqn of supports or Plectrical Raceway Systems for non-hydrodynamic loads are in accordance with Reference 3.7b-7 of the PSAR. SRU discharqe and LOCA loads are considered similar to seismic loads by usinq appropriate floor response spectra for the hydrodynamic loads. A damping value of 7'%f critical is used For all raceway systems for abnormal/extreme load condition and a damping value of 3% of critical is used for normal load condition involvinq SRV discharge loadinq only.

7. 1. 8. 2 Loads

7. 1.8.2. 1 Static Loads The static loads are the dead loads "and live loads- Por cable trays, the weiqht of the cable is considered to be 45 lhs/ft and a concentrated live load of 200 lb. applicable at any point or cable tray span, is used.

7. 1. 8. 2. 2 Seismic Loads The details of the seismic motion input are discussed in Section 3.7 of the FSAR. The effects of the operating basis earthquake (OBE) and the Safe Shutdown earthquake {SSE) are considered.

7. 1.8 2.3 Hydryrlynamic Loads The details of the axisymmetric and asymmetric SRV discharge loads, as well as LOCA loads includinq condensation-oscillation and chugqing are discussed Section 4 0 The enveloped acceleration response spectra at each floor for N-S, E-W, and vertical directions have been generated and widened.

These curves form the basis for the hydrodynamic load assessment of the electrical raceway system. Examples of the response spectrum curves for the containment and Reactor and Control builrlinqs are presented in Appendices 8, C and L.

7. 1. 8. 3 Analytical Nethods Cable tray systems are modeled as three dimensional dynamic system consistinq of several consecutive supports complete with cable trays and lonqitudinal and transverse bracing. The cable tray properties are determined. from the load deflection tests.

7-4 0 Rev. 8, 2/83

Nember points are modeled as spring elements having rotational stiffness with known spring values as determined from the test xesults.

Composite spectra are developed by enveloping the broadened floor response spectra for critical floors for seismic, SRV and LOCA loadinq conditions. The desiqn spectrum is obtained by adding these response spectra curves by the squax'e root sum of the squares method. The composite response spectra curves are obtained for vertical and two horizontal directions.

Acceleration values utilized in the desiqn are determined from the composite response spectra with the consideration of a t 20%

frequency variation at the fundamental frequency of the cable t ray system.

Nodal and response spectrum analyses are performed utilizinq "Bechtel Structural Analysis Program" (BSAP) which is a general purpose finite-element computer program. The seismic and hydrodynamic responses are combined by the square xoot sum of the squares method The total response due to the dynamic loads is calculated by determining absolute sum of vertical response and only the larger response of the two horizontal responses.

Dead and live load stresses are determined from a static analysis of a plane frame 'model usinq BSAP computer proqram and these results are combined with those from the response spectrum analysis. Por normal load condition, SRV discharge stxesses are proportioned from the xesponse soectrum analysis of SSE plus SRV discharge plus LOCA loads according to their spectral acceleration ratios at the fundamental frequencies. Several different support types which are widely used have been analyzed hv these method s.

An alternative method for analyzinq other support types which occur less frequently, uses lonq hand calculations by a response spectrum analysis technique. The support may be idealized 'as a single deqxee of freedom system. 1n qeneral, the maximum peak spectral accelerations were used in the analysis. In some cases where the stresses are critical, a more ref ined value for the acceleration response was used corresponding to the computed system fundamental frequency and considering a frequency variation as explained earlier in this section. The vertical and horizontal seismic responses axe combined according to Subsection 3.7b.2.6 of the PSAR. The member stresses are kept within the elastic limit.

7. 1~9 HVAC Duct System Assesspent methodology

'Phe SRV discharqe and LOCA are considered similar to seismic loads by usinq appropriate floor response spectra qenerated for the CO, chugging, and SRV loads described in Section 4.0.

dampinq value of 5'.4 of critical is used for load combinations involving SSE, SRV discharge and LOCA loads. Mhile a dampinq value of 3% of critical is used for load combinations involving 7-4 1 Rev. 8, 2/83

OBE and/or SRV discharqe loads. For a discussion of the seismic and hydrodynamic loads input for HVAC duct system assessment, refer to Subsections 7.1.8.2.2 and 7.1.8.2.3, respectively. The HVAC duct system had been analyzed by the alternative method described in the Subsection 7. 1.8.3 by determininq the f undamental frequencies of the system in three directions. The inertia forces are determined from the composite spectra described in Subsection 7.1.8. 3 to establish member forces and moments due to hydrodynamic as well as seismic loads.

Rev. 8, 2/83 7-4 2

7 2 ~ DESIGN CQQABILTY QARGIMS-7 2. 1 Stress Margins Stresses at the critical sections for all of the structures described in Section 7.1, pipinq and equipment are evaluated for all the loadinq combinations presented in Section 5.0. The stress margin is defined as (1 - stress ratio) x 100 stress ratio = ~ Cn ~ fn Fn where, Actual Stress Allowable Stress C = Amplification Coefficient 7.2. 1. 1 Contaj,nment Structure The results from the structural assessment of the containment structure are summarized in Appendix A. Piqure A-2 shows the desiqn sections in the basemat, containment walls, reactor pedestal, and the diaphraqm slab which were considered in the structural assessment. The tables in Appendix A give the calculated desiqn stresses and margins for load combination Equations 1, 4, 4a, 5, 5a, and 7 (as listed in Table 5- 1) .

The followinq observations are made from a review of the structural stresses. The calculated stress level is very low for load combination equation No. 1 (an upset condition) i.e.,

reinforrinq bar stresses are less than 20 ksi. In general, among all the applicable load combinations'he most critical load rombination is No. 7a. The maximum reinforcinq bar design stress is predicted as 47. 24 ksi, which occurs in a wetwell section on the outside face helical bars when usinq the absolute sum (ABS) method. This given a minimum stress margin of 12.5% (see Pigure A-29) .

However, the calculated maximum reinforcinq bar design stresses are relatively low in the reactor pressure vessel pedestal, diaphraqm slab, and the base slab, as they are less than 18 ksi, 34 ksi, and 45 ksi respectively. The maximum principal concrete compressive stress occurs at the base slab and is calculated as 4280 psi. Thus, all the reinforcinq bar design stresses are.

below the allowble stresses. It should be noted that the allowable stresses on which the margins are based, are related to t he minimum specified strength. The actual quality control test results for the reinforcinq bars and concrete show the material strenqths to be higher than the minimum specified and therefore, the margins are actually greater than calculated.

Rev. 6, 4/82 7-4 3

In qeneral, the concrete stresses were found to be low except at section 27 in the containment basemat (see Figure A-2), where the concrete stress in compression exceeded the maximum allowable stress in five load combinations out of six that were considered in this report. However, under each load combination the concrete is in triaxial compression at Se'ction 27. Under the worst load case, the>>hydrostatic>> component of the stress is 2830 psi and the  !'deviatoric>> component is only 1392 psi.

Because of this larqe hydrostatic component, the concrete compressive strain is much smaller than the value of 0. 003 in/in permitted by the codes. The concrete, therefore, has a very 3arqe strain margin before failure will commence. It must also he emphasized that not only the actual strength of the placed concrete is hiqher than the minimum specified, as indicated in the paraaraph above. but.that the concrete continues to qain strength after placement. The increase in strength at the end of five years could be as much as 20'%ver the 90 days strength.

Therefore, the locally high compressive stresses in the concrete at Section 27 are deemed acceptable.

7. 2.1.2 Reactor and Contgol Buildinq The results of the structural assessment of the Reactor and Control Building are summarized in Appendices E and L. The analytical results presented herein and in Appendix>>E>> are based on analyses performed using the structural models shown in Appendix>>C>>. The assessment results based on analyses performed using the revised structural models (as discussed in Subsection

~

7.1.1.2.1.1) are presented in Appendix>>L" Pigures concrete F,-1 through E-22 show the desiqn sections in the basemat and the structure composed of floor slabs, shear walls, blockwalls, refuelinq pool qirders, as well as floor structural steel and superstructure steel, which were considered in the structural assessment. The sections selected for assessment were considered to be most critical based on previous seismic calculations. The tabl'es in Appendix E qive the calculated design stresses and marqins for the critical load combinations equations 1 and 7a of Table 5-1 and equations 1 and 7 Table 5-2. The other load combinations do not govern.

In the case of floor slabs, the calculated stress levels, in qeneral, are very low for slabs above El. 683.0 ft.5-1 The aoverninq load combination is equation 1 of Table (normal condition) and the reinforcing steel stresses are significantly less than 20 ksi. For slabs below El. 683.0 ft. also, the qoverninq load combination is equation 1 of Table 5-1. The maximum reinforcinq steel stress was 49.79 ksi, which occurs in the reactor building slab at El. 645.0 ft. (see Figure E-33)

The selected floor sections for the review and assessment are qiven in Figures E-1 through E-6.

In the case of shear walls, the maximum rebar stress was 43.25 ksi, and the minimum stress marqin is 20'5 (see, Figure E-34) . The assessed elements are given in Figures E-1, E-3, E-4, P.-7, and E-A.

7-44 Rev. 8, 2/83

In the blockvalls the calculated maximum reinforcing bar design stress is 30.6 ksi for load combination equation 7a {see in Figure F.-35) . The minimum stress margin for compressive stress the concrete is 22% The blockwall elements revieved for assessment are shovn in Piqures E-9 throuqh E-16.

In the case of Reactor Buildinq structural steel (see Pigure The E-

36) ~ load combination Eq. 7 of Table 5-2 generally governs.

maximum bendinq stress was found to be 31.9 ksi which is less than the allowable value. This stress occurs in a beam at El.

719.1 ft. In the other cases the stress marqins are 29% or more.

The structural steel elements selected for assessment are given in Piqures F,-17 throuqh E-20.

A three-dimentsional lumped mass model vas qenerated for determininq the dynamic response of the Reactor Building Crane Support Structure. This model is shown in Figure F.-21. Equation 7, Table 5-2 serves as the qoverninq loading combination.

Selected members as qiven in the model vere assessed for structural inteqrity and stability. The design margins for structure and crane girder are 0% (see Figure E-37). This condition is reached hy lettinq the rails deform in such a vay that the crane bumper strikes aqainst one of the rail girders.

The assessment of the Refuelinq Pool Girder shows that the maximum rebar stress was 51.7 ksi and the desiqn margin is 4g (see Figure E-38). The elements selected for assessment are shovn in FiqureE-22.

hs shown in Figure E-38a, the box section columns supporting the refuelinq pool were found to have adequate strength for resisting dead, live, and dynamic loads including seismic {OBE, SSE), SRV, and LOCA loads imposed by the refueling qirders. Equation 6 was found to be the qoverninq euqation for columns. The strength of the box se"tion columns is summarized under elements 41 and 42.

The minimum design margin is 38%.

7.2.1.3 SRV Support Assemhlies and Suppression Chamber Columns The stresses at critical sections of, the SRV support assemblies and the suppression chamber columns were calculated separately for the load combinations in Table 5.2. The. maximum stresses are governed by load combination 7a for both the SRV support assemblies and columns. The results of the SRV support assembly analysis are shovn in Figure A-67. The lowest stress margin of SRV support system which includes all bracinq members and connections is 21.7%. On the other hand, the maximum stresses in column (42 inch diameter pine), at the top and bottom bolt anchorages are shown in Fiqure A-59. The lowest stress margin in the column structure is 11. 4%.

7. 2. 1. 4 Downcomer Bracing Stresses in the bracinq members and connections vere checked usinq the load combinations and allowable stresses as given in 7-4 5

Table 5-2. Dynamic loads were combined on the basis of the SRSS method. Combined axial and bendinq stresses were investigated for the most hiqhly loaded members. 'Bquations 1, 3, 4 and 7 qovern for the brace members with the desiqn margins= as indicated in Figure A-60. Por the connections, equations 2 and 7 are critical and the resulting desiqn margins are shown in Figure A-

61. All bracinq members and connections are adequate.

7. 2.1.5 Liner Plate For the normal load condition, the liner plates do not experience any net neqative pressure as can be observed from Figure 7-21.

For the abnormal load condition, the maximum net negative pressure on the pressure boundary portion of the liner plates occurs on the containment'all, at point 8 of Piqure 7-23, and is

-6.39 psi. Since this is an impulse load of .004 seconds duration and the liner plate is supported every 2 feet, the stress in the liner plate is 12.5 ksi, well below the allowable.

There is a marqin of 51% for pullout of the embedded T steel sections that support the liner plate.

The liner plates on the base slab are supported by embedded M4x13 structural steel members every 10 feet. The maximum negative net pressure on the base slab occurs at the corner. The magnitude is

-5.12 psi. However, due to liner plate connection on the corner between base slab and containment wall, the neqative net pressure does not cause a, bending problem in. the liner plate and no pullout problem on Q4x13 sections. The liner plate located away from the corner described above, do not experience negative pressure.

7. 2; 1. 6 Downcomers A list of downcomer and bracinq system modal frequencies and participation factors is given in Table 7-5 The fundamental system mode is at a frequency of 1.8 Hz, which is a cantiliever type of mode for all downcomers moving together. Downcomer stresses were checked according to ASIDE Code Section NB3652 using load combinations in Table 5-3. Stresses and design margins are given in Figure A-66.

7.2.1.7 Flectrical raceway System It of is apparent from the analysis that high stresses are a result responses due to horizontal inertia loads. During the normal load condition, stresses under SRV discharge are generally low.

However, for the abnormal/extreme load condition, certain members required strengthening to relieve high stresses. After i mplementinq these modifications, the resultant stresses do not exceed .the allowable stresses in any member of the electrical raceway system supports. The modifications to electrical raceway systems are a result of the assessments performed using the structural models shown in Appendix "C". The assessment results based on analyses performed using the revised structural models 7-46 Rev. 8, 2/83

{as discussed in Subsection 7.1.1.2.1.1) are presented in Appendix "L".

7. 2 1 8 HVAC Due/ System Similar to<the analysis of the electrical raceway system, the analysis of the HVAC duct system demonstrated that. most of the support members have actual stresses lower than the allowable stresses. However, certain structural members required strengtheninq to relieve hiqh stresses under the abnormal/extreme load conditions The strengtheninq of HVAC duct supports are a result of the assessments performed usinq the structural models shown in Appendix "C". The assessment results based on the analyses performed usinq the revised structural models {as discussed in Subsection 7. 1.1. 2. 1.1) are presenteR in Appendix II LII

7. 2. 1.9 BQP Equipment All Seismic Category I BOP equipment are re-evaluated for the hvdrodvnamic and non-hydrodynamic loads {see Subsection 7. 1.7) via the SSES Seismic Qualification Review Team {SQRT) program.

For each BOP equipment, 4-paqe SQRT summary forms have been prepared documentinq the re-evaluation of that equipment. In some cases, modifications were required to reduce the stresses below the allowables. The modifications to BOP equipment are a result of the assessments performed using the structural models shown in Appendix "C~~. The assessment results based on analyses performed usinq the revised structural models {as discussed in Subsection 7.1.1.2.1.1) are presented in Appendix "L".

In response to SER Open Item I 11, the BOP SQRT summary forms requested by the NRC were formally submitted on February 25, 1982

{

Reference:

PLA-1024). The remaining BOP SQRT summary forms are available for review.

7. 2. 1.10 NSSS -

gguipment All Seismic Category I NSSS equipment are re-evaluated for the load combinations qiven in Table 5-5 via the SSES SQRT program.

~or each NSSS equipment, SORT summary forms are prepared document.ing the re-evaluation of that particular equipment. The assessment results based on analyses performed using the revised structural models {as discussed in Subsection 7. 1. 1.2. 1. 1) are presented in Appendix "L".

The NSSS SQRT summary forms requested by the NRC will be formally suhmitted to the NRC under the SSES SQRT program. All NSSS SQRT summary, forms are available for review.

7. 2. 1. 11 NSSS and BOP . Piping As documented in Subsection 7. 1.5 and 7.1.6.1.1, all Seismic Cateqory I BOP and NSSS piping have been analyzed for hydrodynamic and. non-hydrodynamic loads per the load combinations 7-47 Rev. 8, 2/83

qiven in Subsections 5.5 and 5.6, respectively. As a result of this evaluation, many modifications vere required to maintain the stresses below the allowable values. Appendix P provides a summary of the stresses and desiqn margins for selected BOP piping systems based on analys'is results for the structural models shown in Appendix>>C". The above required modifications are a result of analyses performed usinq the Appendix "C" structural models. The assessment results based on analyses performed usinq the revised structural models (as discussed in Subsection 7.1.1.2.1.1) are presented in Appendix>>L>>.

The results of the above evaluation are documented in stress reports, which are available for NRC review.

7 2. 2 Acceleration Response Spectra

7. 2 2. 1 Containment Structure The method of analysis and load description for the acceleration response spectra generation are outlined in Subsection 7.1.1.1.1.6.1. Appendix B contains example acceleration response spectra for SRV, condensation oscilation and chuqqinq, and seismic sloshinq load cases. From a revidw of the SRV and LOCA acceleration response spectra curves the maximum spectral accelerations are tabulated in Table 7-1 for 1% of critical.

damping.

7.2.2.2 peactor. and Control Building The methods of analysis and load application for the computation of the acceleration response spectrum in the reactor and control building are described in Subsections 7.1.1.2.1.1 and

7. 1. 1.2.1.2. Appendix>>C>> contains the acceleration response spectra fo" lov dampinq values for SRV and LOCA load cases based on analyses performed using the structural models shown in Figures C-1, C-2 and C-3. Appendix>>L>> contains example response spectra qenerated using the revised structural models, as discussed in Subsection 7.1.1.2.1 1. From a reviev of the SRV and LOCA acceleration response spectra curves based on the models presented in Appendix "C>>, the maximum spectral accelerations are tabulated in Table 7-2 for 4% of critical damping.

7 2.3 Containment Liney Openings 7.2 3. 1 Pguipment Hatch-personnel Air Lock Stresses in the equipment hatch-personnel ai" lock vere all vithin allowable limits. Hovever, as a result of the nev loads, bolt pre-load had to be increased from 65 to 72 kips to maintain acceptable levels of displacement at the flanged )oint. The resultant equivalent radial load applied at the bearinq on the hinge support results in a minimum safety factor of 3 at ultimate for the roller and race.

7,2.3 2 CRD Removal Hatch~ Suppgession Chamber Access Hatch 7-4 8

and 'gguj,pmyyt Patch CBI's analysis indicated, no stresses in excess of the specified allowable limits for the additional loadings considered.

7. 2. 3..3 Re fue linc Head and Sunnort A ~

Skirt The refuelinq head and flanqe were found to have no stresses exceedinq allowable limits. The only effect of the new loads applied was to increase bolt pre- stress f rom 161 to 200 kips to maintain leaktiqhtness at the flanqed joint. Pigure A-33.1 gives the stress marqins in the refuelinq head and the flange.

Rev. 6, 4/82 7-49

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-28.40 KWU 306 31.39 KWU 306 REV. 6, 4/82 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT LINER PLATE HYDRODYNAMIC PRESSURE DUE TO CHUGGING FIGURE 7-20

PEDESTAL CONTAINMENT WALL HYDROSTATIC

+10.4 pai 24'10.4 pai.

BASE MAT SRVgI 18' 7.8 paI 6'7.8 paI TOTAL 18EV.

6, 4/82 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT LINER PLATE PRESSURES NORMAL CONDITION FIGURE 7-21

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DESIGN ASSESSMENT REPORT LINER PLATE HYDRODYNAMIC PRESSURE DUE TO CHUGGING AND SRV FIGURE 7-22

0 eataiaeant Wall Pedestal b+14e88'21@73 RE~30 ~ 00'ls ~ 30 0+44 ~

00'OZNT ZN 'FZCURE 3 4 6 7 CHUCCZNC 62e16 -26 42 24+74 26.85 26 ~ 69 32,72 28,40 31 ~ 39 SRV Trace 76 5 ~ 76 - 7 ~ 80 7 ~ 80 7 ~ 80 7 ~ 80 7 ~ 80 7.09 3.05 Hydrostatic Se76 10.40 10 '0 10.40 10.40 10 '0 6.82 3.05 Wetwell pressure 25 F 00 25 F 00 25+00 25.00 25 F 00 25 00 25.00 25+00 due to SBA or IBAD NET PRESSURE -37.16 lo18 2 86 Oi75 0.91 ' ~ 12 3.69 6.39 aWetvell pressure due to DKL is 34'si.

REV. 6, 4/82 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT LINER PLATE PRESSURE ABNORMALCONDITION FIGURE 7-23

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Table 7-1 MAXIMUM SPECTRAL ACCELERATIONS OF CONTAINMENT DUE TO SRV AND LOCA LOADS AT 18 DAMPING TYPE OF LOAD NODE 'LEVATION MAXIMUM STRUCTURAL LOAD CASE DIRECTION NUMBER SPECTRAL FREQUENCY ACCELERATION ( ) Hz Axisymmetric Vertical 841 7780-9-3/4" 1.088 15 SRV Horizontal 135 672 I Qll 1. 58 38 Asymme tric Vertical 252 702'-3" 0.83 40 Horizontal 131 672'-0" 0.875 38 Axisymme tric Vertical 235 7020-3" 1.80 54 CHUGGING Horizontal 131 672'-0" 8.5 30 0 Asymmetric Vertical 235 7020-3" 1.56 54 Horizontal 131 672<0" 7.1 30 (CO) Axisymmetric Vertical 850 731'-3-1/4" 1.0 ll Horizontal 131 6721-0" 1.97 30 REV. 6, 4/82

Table 7-2 MAXIUM SPECTRAL ACCELERATIONS* OF REACTOR AND CONTROL BUILDINGS FOR 4X OF CRITICAL DAMPING STRUCTURA TYPE OF LOAD NODE SPECTRAL FREQUENCY LOAD CASE DIRECTION NUMBER ELEVATION ACCELERATION ( ) Hz Axisymmetric Vertical 25 697'-0" 1.7 15 Horizontal NA NA SRV Asymmetric Vertical 25 697 I Ofl 0.35 15 Horizontal 37 683 I Oll 0.35 25 (E-W)

Axisymmetric Vertical 25 697'-0" 3.5 15 Horizontal 37 683'-0" X.O 25 (E-M)

L CHUGGING Asymmetric Vertical 25 697'-0" 2.7 15 Horizontal 36 I 670I-O" 2.1 75 (E-M)

(CO) Axisymmetric Vertical 23 870'-0" 1.85 Horizontal 37 683'-0" 1.0 25 (E-W)

• These accelerations are based on a review of the acceleration response spectra presented -in Appendix C.

Rev. 8, 2/83

0 0'able USAGE FACKR SU%MR OF 7-3 DCNNQCHERS EMEZGBKY/FAULTEDGCNDITICH SBA IBA or SM KBA

+ CBE + SRVl ipressure ~ Pressure iPressure

+ SRVl + SRV2 'Thermal 'Thermal ~ Thermal

+ SRV2 + amo Transient Transient Transient

~Steam Flow 'Steam Flow iSteam Flow

+ CHUG + CHUG + orna

+ SRV

+ SRV SSE

+ SSE At diaphragm location 0.0083 Oe608 0.774 0.774 0.791 ~ 782 mates: 1) SRV is a combination oi direct loads and building response loads.

2) Ggig is the maximum chugging load (direct load and building response).

3) The calculation is based on ASME,Section III, 1979 Smmer Addendum.

4) The combination of + OHJG, + SRV and SSE or CBE is by SRSS.

')5) Thermal and pressure loads are canbined with 4) by absolute sun.

SRVl

7) SRV2 is submerged structure load.

is building response load.

REV. 6, 4/82

TABLE 7-4 MAXIMUM CUMULATIVE USAGE FACTORS FOR SRV DISCHARGE LINE CALCULATED CODE COMPONENT CUMULATIVE USAGE ALLOWABLE CUMULATIVE FACTORS USAGE FACTORS Flued Head 0.46 1.0 3-Way Restraint 0.51 1.0 Elbow (Line P) 0.56 1.0 REV. 6, 4/S2

Table 7- 5 DOWNCOMERS AND BRACING SYSTEM MODAL FREQUENCIES FREQ. HEIGHT PARTICIPATION FACTORS MODE (HZ) HORIZ-X HORIZ-Y VERTICAL 1 1.84 0 '20 l. 274 2 1.84 -1.278 0 ~ 321 3 2.53 0.001 -0 ~ 013 4 6.58 0.001 5 8 '4 0.001 -0.002 6 9.95 -0.001 0.001 7 13.27 0 004 -0 002 -0 '02 8 14-05 -0.001 0.004 -0.002 9 14 '5 0.001 '-0 F 001 0.004 10 15.12 0.003 0.002 -0.001 11 15.17 -0.007 0.006 12 15 27 0.002 0.001 13 15 38 0.003 -0.008 14 15 ~ 44 -0.001 0.003 -0 007 F

15 15 '6 -0.003 -0.001 0 '02 45 15.7S 0'02 -0 '12 46 15.76 -0.004 0.001 0 '04 47 17.44 0.010 0 '21 48 17.44 -0.504 0.006 49 17.50 0. 023 -0.116 50 17.78 0 015 0.126 93 45.05 -0.072 0.460 94 45.14 -0 '16 -0 '59 95 45.33 -0.005 -0 '27 96 45 '2 0.007 0.256 REV. 6, 4/82

P li 0

P

CHAPTER 8 SSES~UENCHER VERIPICATZON TEST Chapter 8 is proprietary and is found in the proprietary

'supplement to this DAR.

CHAPTER 9 SSES LOCA STHAil CONDENSATZON VERIFICATION TEST GKi'l-IXi1 TABLE OF CONTENTS 9 0 GKM IIN TESTS 9~ 1 XNTRO DUCTION 9 1.1 Purpose of Test 912 Test Concept 9.1 2.1 Unit Cell Approach, 9 1 2 1 1 Single Cell Theory 912 21 Simulation ofChamber SSES Parameters (Hetvell) 9 1 2.2-2 Suppression 9.1.2.2 3 Vent Pi.pe 91224 Pool Internals 9 2 TEST FACILITY AND INSTRUNENTATION 9 2-1 Physical Configuration 9 2 1.1 Steam Accumulator and Discharge Line (llain Steam Line Break) 9 2 l.2 Steam Buffering of the Steam Accumulator (Recirculation Line Br ea k) 9 2 9 2 2 l 3 Test Tank Instrumentation 9.2. 2 1 General Description 9 2 2.2 Instrumentation Identification 9 2 2 3 Operating Instrumentation 9 2 2 4 Test Instrumentation 9 2 2 S Visual Recording 9 2 2 6 Inspection and Calibration of the measuring Instrumentation 9227 Analysis o f Measure ement Errors 9 3 TEST PARANETERS AND i'lATRIX 9 4 TEST RESULTS 9 S 'ATA ANALYSIS AND LOAD SPECIFICATION 9 6 VERIFICATION OF THE DESIGN SPECIFICATION 9 7 FIGURES 9 8 TABLES Rev. 3, 7/80

CHAPTER 9 FIGURES 2

Number Title 9-1 Test Stand Schematic Diagram Test Tank Coordinate System and Vest Instrumentation Test Instrumentation 9-5 Test Instrumentation 9-6 Bracing Configuration 9-7 Bracing Design 9-8 Quencher Dummy 9-9 I-Beam Design 9-10 Data Recording: Schematic Block Diagram 9-11 Calibration of the Sensors and Registration Instruments 9-12 Time Interval for Calibrations, Checks and Adjustments 9-13 Calibration System 9-14 Physical Calibration of tho Pressure Vransducers P6.1 ..P6.8 by Lowering of the Mater Level in the Pool 9-1Qa Calculated SSES Vent Steam Mass Flux vs.

Time RCL Break 9-lQb Calculated SSFS Vent Mass Zlux vs. Time Full MSL Break Figures 9-15 thru 9-151 are contained in the Proprietary Supplement.

Rev. 3, 7/80 9-2

CHAPTER 9 TABI ES Nueher Title 9-1 Comparison of Fixed Parameters 9-2 Operating Instrumentation 9-3 Test Instrumentation GKN II-8 Test Matrix 9-5 Test Parameter Tables 9-6 thru 9-11 are contained in the Proprietary Supplement.

Rev. 3, 7/80 9-3

9. 0 GKM IIM TESTS 9 1 INTRO DUCTION The NRC in NUREG 0487, "Mark II Containment Lead. Plant Program II Load Evaluation and Acceptance Criteria~~, accepted the Mark Owners load definition,for condensation oscillation but with regards to the specified frequency range cautioned that: "Some modification may be required to correct for the difference in vent configuration between the 4T (Temporary Tall Test Tank)

II facility and the prototypical Mk Containment." The Hark II Owners then proceeded to run several series of small scale tests to investigate the effect of vent length on the condensation oscillation load. Results f rom these tests proved inconclusive.

It was then decided that the most expedient way to resolve the questions associated with vent length effects was to run a series of full-scale tests in a facility with a prototypical vent configuration.

The Mark IX Owners Group selected the GE 4T facility to run this new series of full-scale tests. In addition, it was decided by PPSJ to conduct a series of transient steam blowdown 'tests in a modified GKM II test tank in Mannheim, Germany. This chapter presents a description of thi" test program, the results from these tests and a comparison of the results with the design specification used on the SSES containment.

9.1.1 Purpose of Test The load specification for the LOCA steam condensation events for the SSES is based on the results of the tests performed in the first quarter of 1976 at the 4T test tank in the GF. Pressure Suppression Test Facility. These load definitions are provided in Section 4.2 of the SSZS DAR. In order to resolve NBC concerns regarding the differences in vent length between the 4T tank and the prototypical MK II containment and to verify the LOCA steam condensation load specification used on SSES, it was decided by PPSL to conduct this series of tests.

9 1.2 Test Concept The concepts used to design and perform the tests were:

1) Use of a conservatively defined single cell

2) The close prototypical simulation of the downcomer system parameters Rev. 3, 7/80

9.1 2.1 Unit Ce11 A~roach 9.1 2.1.1 Si~nle Cell Theory For a gas bubble oscillating in a free vater space, the vater mass coupled to the bubble is alternately accelerated and decelerated. During this process the overpressure and underpressure amplitudes decrease vith increasing distance from the bubble. Mhen a solid wall is placed near the oscillating bubble, the water acceleration is restricted in the direction of the vali and the decrease in pressure amplitude in the direction of the wall is less. This effect can be expressed mathematically by replacing the bubble by a potential source and accounting for the wall by the method of images. The effects of the real source and the image source are added for each point of the flow field.

For the case in which a bubble is enclosed in a narrow vater space, closely surrounded by solid valls and a solid bottom vith a free vater surface at the top, the water space belov the. bubble is for all practical purposes unmoved. Only the vater'o'lume above the bubble is free to oscillate Conseguently, the pressure gradient in the lower water space is nearly zero, vhile the pressure amplitude above the bubble decreases with increasing proximity to the vater surface, until surface.

it is zero at the vater Analytically, the case in which a planar field oX uniform strength sources are all acting in phase is the same as the case in which solid valls exist between each of the individual sources. The single cell test configuration used at GK5-IIH simulates this extremely conservative case of parallel sources acting in phase with the same strength.

9.1.2 2 Simulation of SSES Parameters The following section provides a description of those parameters that vere simulated in the GK."1-III test facility. A single cell corresponding to the SSES is simulated at actual scale in the GKN-IIM test stand. The single cell consists of a vent pipe with proportionate dryvell and suppression chamber. A comparison of the plant and test parameters is given in Table 9-1.

9.1.2.2.1 Drywell The volume of the drywell part of the test tank corresponds to the proportionate volume of the drywell in the plant. The dryvell valls are preheated to temperatures of about 143 OC (corresponding to 4 bar saturated steam) in order to avoid significant steam condensation. As a result, the mass flow values in the test are higher than in the plant, vhere greater condensation on the dry well internals and walls is possible.

Since the drywell of the test stand consists of a volume without any major internals, the air is .flushed over just as fast, and probably even somevhat faster than in the plant.

Rev. 3, 7/80

9 1 2 2. 2 Suppression Ch amber QMetvel1)

Like the dryvell volume, the free air volume of the suppression chamber also corresponds to the proportionate value in the plant.

As a result, the pressure build-ups in the test tank and in the SSES containment are equal.

The ratio of surrounding water surface to the cross-sectional area of the vent pipe varies in the plant as a function of the pipe s position. Theoretical and experimental investigations shov that the condensation loads decrease with increasing area ratio. Therefore, the single cell with the smallest area ratio at the containment wall vas simulated in the test stand. Its area of 3 77 m~ (40.7 ft.~) is clearly less than the mean value in the SSES (5.64 m~); 60.7 ft.~. This adds considerable conservatism to loads measured in the test stand.

Due to the decreased volume of vater relative to the mean value in the plant, there is a greater heating of the vater in the chamber during the tests than would be expected in tsuppression he pla nt.

The volume flexibility of the suppression chamber walls is less than or equal to the plant value of 0.6 x 10-~ m~/bar (37.2 in~/bar) relative to the single cell.

9.1.2 2 3 Vent Pipe The vent pipe has practically the same dimensions and the same distance from the bottom as in the plant. Previous test series and also theoretical considerations have shovn that the condensation loads vary somewhat with the submergence depth of the pipe.

For small submergence depths, the loads first increase rapidly with increasing depth and then approach a limiting value asymptotically. Therefore, the tests are performed at the highest value of. submergence depth, 3.66 m (12 ft), occurring in the plant.

/

The vent pipe braces have a stiffness greater than or equal to the maximum value of 770 x 106 N/m (4386 kips/in) occurring in the plant and are located at the same position as in the plant.

9.1 2 2 4 Pool Internals To be able to determine the load on a perforated-pipe quencher of the depressurization system located near the vent pipe in the suppression chamber during the condensation processes, a quencher arm having the actual dimensions is installed in the test tan'k.

The quencher arm with central member is velded to the inner cylinder in the pool at a distance of 1.1 m from the bottom.

Rev. 3, 7/80 9-6

To determine the vertical loads produced by the condensation on steel structures in the water region, an I-beam (ASCl 8 10 x 45) is arranged horizontally between the vent pipe outlet and the pool surface (6.3 m from the bottom).

Rev. 2, 5/80 9-7

9 2 TEST FACILITY AND INSTRUMENTATXON 9.2 1 Physical Configuration The test configuration as constructed is typically illustrated diagrammatically in Figure 9-1. The entire test system consists of:

'he steam accumulator (GKM Designation: Condensate Accumulator S 6),

The arrangement for steam buffering of the steam accumulator (GKN Designation: Feedwater Tank 3202),

and The actual Test Tank {GKN Designation: Condensate Accumulator S 3) .

The test set-up simulates the pressure suppression system of the reactor plant in a so-called single cell (one vent pipe with proportionate drywell and suppression chamber) at actual scale.

From a tank {S 6) which is filled partially with saturated steam (simulating the reactor pressure vessel), steam flows via a discharge line and flow orifice into the actual test tank (S 3) which is subdivided into a drywell and a suppression chamber 9 2 1 1 Steam Accumulator and Dischargee'ine gl'lain Steam Line Break)

The Condensate Accumulator S 6 in Shop I of GKN, with a of about 120 m~, is used to simulate the reactor, pressurecapacity vessel ia the test stand; see Figures 9-1 to 9-5. Bef ore test start, this accumulator is filled with water and steam in a saturated condition The pressure is 20 bar or less, depending on the requirement of the relevant tests.

Between the accumulator S 6 and the actual test tank there is mounted an HD 400 pipe as a discharge line; cf. Figures 9.1 9.5. Located in this line is an isolating slide valve, quick-opening valve and a standard orifice for flow-rate limitation in accordance with the simulated break size. By using orifices of different diameter and by specifying the pressure water filling of the condensate accumulator, the blowdownandtransient is set. Besides flow-rate limitation, the standard orifice is also, .

used for flow-rate measurements.

Before test start, the discharge line is sealed at the entrance into the test tank (S 3) by a rupture disk combination ND 400 with support pressure (nitrogen) . The rupture disks expose the flow cross-section in a .few milliseconds at test start.

Rev. 3, 7/80 9-8

9 2 1 2 Steam Buffering of the Steam Accumulator ~Recirculation Line Break/

The blowdown from an assumed pipe break inside the containment

{BCL break) results in a relatively high level, short term constant mass flow rate. Hovever, using the test stand as set-up in Subsection 9.2.1.1 leads to a steadily decreasing mass flow rate.

In order to simulate this situation under the given conditions of

,the test stand, the assignments of the individual tanks was changed so that tank B 202 was used as the actual accumulator.

The tank S 6 was then used as a buffer tank which is continually connected to the GKM superheated-steam network. At the beginning of the test, this tank is connected directly to the discharge line to the test tank. Within 10 to 20 seconds after test start, this connection is broken by means of a quick-closing valve in accordance vith the prescribed mass flow rate variation and the test proceeds as described previously until pressure equalization is achieved in tank B 202 and test tank S 3.

9 2 1.3 Test Tank The condensate accumulator S 3 is used to simulate the SSES containment and is constructed as shown in Figures 9-1 to 9-5.

The upper portion of the tank is the drywell and the lower portion is a partially waterfilled suppression chamber. The following volume subdivisions result:

Drywell S 3 {with pipe portion of the suppression chamber at high water level in the inner tank) 75. 6 m~ [3 Suppression chamber air space with completely filled annular gap and high water level in the inner tank 47 m~ (3 Water filling of the inner tank in the S 3 at high vater le ve1 26 m~ I3 This subdivision conservatively simulates the SSES "single cell."

The bottom of. the drywell serves as the diaphragm floor where the vent pipe is attached.. The vent pipe is identical in length, diameter and wall thickness to the plant version.

In the lover part of the test tank, the simulated suppression pool, a thick-walled inner cylinder made of steel, vas installed.

The installation of this inner cylinder satisfies tvo requirements resulting from the specified similarity to the plant. First, the water volume is reduced to simulate the smallest plant single cell and second, the wall thickness of 100 mm results in a stiffness which corresponds to that of the concrete valls in the plant. The vent pipe bracing stiffness and location is very closely prototypical of the actual SSES as built arrangement.

The partition vali between drywell and suppression chamber is provided vith swing-check valves for protection of the test Rev. 3, 7/80

stand. The steam inflow at the upper end of the vent pipe is simulated in a representative manner by the installation of the correct vent riser and jet deflector plate.

The drywell region of the test tank is provided with an electrical heating system on the outside wall The initial temperature of the wall and thus the condensation of steam inside the drywell can thereby be contzolled.

Besides comprehensive instrumentation, viewing ports are mounted on the test tank in the air region and water region of the suppression pool, making it possible to observe the processes with a television camera and high-speed cameras. To permit good f ilm quality, demineralized water pool.

is used to fill the suppression 9 2 2 'Xnstrumentation Instrumentation is provided for controlling the test sequence, determining the prescribed measurement quantities, and recording them.

9.2 2 l General Description The instrumentation u'sed in the GKM-IIN test facility consists of operating instrumentation and test instrumentation. The purpose of the operating instrumentation is to control the test sequence and monitor the test stand. The test instrumentation ensures the recording of all data of significance for evaluation of the phenomena which occur during steam condensation.

Details on the operating instrumentation are given in Subsection 9.2.2.3 A detailed description of the test instrumentation can be found in Subsection 9. 2. 2. 4.

9.2.2.2 Instrumentat ion Identif ication The measurement transducers are identified hy a system of letters and numbers. Each identification starts with a letter or letters describing the type of transducer:

p for Pressure Transducer T for Temperature Sensor (Thermocouple)

L for Mater Level Measurement DG .for Displacement Gage SG for Strain Gage I for Electrical Impulse Signal LP for Level Probe LC for Load Cell AF for Air Fraction OR for Oxygen Rate Pollowing these letters is a number which characterizes the mounting location or measurement location in the test stand. Por Rev. 3, 7/80 9-10

that purpose, the test stand is divided into different System Groups as follows (see Fig. 9-1):

System Group 1 steam lines to the accumulator S6 and to the feedwater tank B 202 and in the feedwater tank B 203 System Group 2 feedwater tank B 202 System Group 3 steam accumulator S6 System Group 4 steam supply to the test stand System Group 5 instrumentation of the proportionate drywell with the vent pipe System Group '6 suppression chamber The System Groups 1-4 contain the operating instrumentation, while groups 5 and 6 designate the test instrumentation.

After this identification number there is a decimal point which separates this number from the running numbers of the transducers.

9 2.2 3 Operating Instrumentation The purpose of the operating instrumentation (see Table 9-2, Figures 9-1, 9-3, and 9-4) is to monitor the steam accumulator, feedwater tank and steam lines. The signals- from the measurement transducers are read by a process control computer and recorded.

This computer is a part of the operating instrumentation. All data are stored on magnetic tape and can be printed out or plotted after each test. Before test start, the process control computer compares the recorded measurement signals with prescribed setpoint valves and prints them out. If .the measurement value differs from the setpoint value by a prescribed percentage, that measured value is identified in the printout.

The operating instrumentation concentrates on the measurement of pressures, temperatures and water levels in the steam accumulator, stea,m lines and feedwater tanks.

9 2. 2. 4 Test Instr umenta tion The test instrumentation (see Table 9-3 and Figures 9-3 to 9-8) records all the data needed to evaluate the phenomena occurring during steam condensation and the resulting loads in the pool and also the data needed to determine the steam flow rate in the discharge line. The dynamic pressure loads and accelerations are measured at several points in the pool. The forces occurring at the vent pipe bracing and on submerged structures in the.

suppression pool are recorded by strain gauges. The pressure build-up in the vent pipe is measured at several points. In addition, level probes are installed on the vent pipe so as to be Rev. 3, 7/80 9-11

able to record the dy namic behavior of the water surf ace. The strains on the pipe are measured at two places on the vent pipe.

100 mm below the bracing (see Figure 9-5) and approximately 100 mm below the gusset plate bracing arrangement simulating the diaphragm slab (see Figure 9-3) . Pressure and temperature measuring points in the air space of the suppression chamber and in the proportionate accumulator provide information about the variation of pressure and temperature during the tests. Two differential-pressure measuring points in the water region of the suppression chamber record the air bubble fraction in the pool.

At the upper end of the vent pipe there was a measuring point for the continuous sampling of the steam to determine the air content. The measurement system for continuous sampling is provided by SRI International.

The data is recorded on magnetic tape in analog form by means of carrier-frequency amplifiers and dc amplifiers. This ensures that high-freguency measurement signals are recorded with proper frequency and amplitude. The data is reduced later by a computer. Simultaneously with the recording on magnetic tape, most of the measurement points are also recorded on Visicorders.

That type of recording makes it possible to get a quick 'look at important measurement variables shortly after each test. At the same time, a few selected transducer channels of the 'test instrumentation are recorded additionally at the process control computer. This procedure makes it possible to perform a guick and simple summary evaluation of that data after each test.

Each measurement chain consists of a transducer, connection cable, amplifier (carrier-frequency or dc amplifier), balancing unit and recording unit (see Figure 9-10).

The utilized pressure tra nsducers have a measuring diaphr agm and a foil strain gage system which is directly connected to the diaphragm. All pressure transducers in the water region of the suppression chamber have an exposed measuring diaphragm with direct contact to the surrounding water. Earlier studies by KMU have shown that this type of transducer is hest suited for recording higher-frequency pressure oscillations with correct frequency and amplitude.

The measuring diaphragm for pressure transducers P4 1, P5. 1 P5 5 and P6.9 required protection from the hot steam. This was accomplished by means of a short water-filled pipe which connects the tr'ansducers to the measurement site. The remaining pressure transducers did not require protertion.

9 2.2 5 Visual Record~inc The processes in the water region of the suppression chamber are recorded optically on film by a high-speed camera and on video tape by a television camera.

Rev. 3, 7/80 9-12

The cameras are mounted outside the tank and observe the processes by means of bul1~s eyes. Several underwater searchlights axe installed in order to ensure satisfactory lighting of the end of the vent pipe.

A uniform electrical reference signal ensures time correlation between all the data acquisition systems. P Rev. 2, 5/80

9 2 2.6 Inspection and Calibration of the Neasuri~nc Instrumentation The calibration and the electrical and physical checking of all sensors before, during and after the tests were performed in accordance with the Test and Calibration Specifications.

Figure 9-11 shows diagrammatically the physical calibration of the transducers, the setting and calibration of the amplifiers and recorders, and the quality inspection of the transducers.

The time intervals stipulated for these inspections and calibrations per the Inspection and Calibration Procedures are given in Figure 9-12. Figure 9-13 shows the chain of the calibration system from the National Standards of the Physikalisch Technische Bundesanstalt (PTB) to the measuring instruments.

An additional physical inspection of the pressure transducers in the water region was performed by incrementally lowering the water level and comparing the measured pressure to the known hydrostatic pres ure at the transducer location Rith a few exceptions, the 88 sensors used in the tests were fully operational for the duration of the tests. On December 10, 1979, the pressure transduce" P 5.4 failed. Xt was replaced by a new transducer for the subsequent tests. After initial difficulties with the continuous 0 measuring device, a modification of the sampling arrangement resulted in satisfactory performance. At a few level probes, the insulators were damaged by parts of the rupture-disk diaphragms heing carried along by the steam flow. Those level probes were re placed. The strain gauges of measuring point SG 5.1 had to be replaced on November 14, 1979 due to too low insulation resistance.

The final inspection of the sensors after the completion of the test pxoject showed a fully operable instrumentation system.

9.2 2.7 Analysis of Neasurement Errors Based on the information fxom the manufacturers of the measurement instruments, KRU's own investigations, and taking into considexation the experience gathered in similar test projects, the maximum measurement errors for the individual transducers are as follows:

Pressure transducers P 6.1 ... P 6.8 Linearity error and hysteresis error of the transducer 0.5% of 10 bar = 1.25% of 3 bar 1. 25$

Sensitivity error relative to 40 K temperature difference 0. 75%

Error of the measuring amplifier 0. 5$

Rev. 3 7/80 9-14

Error of the balancing unit and the recorder 0 5%

Maximum total error + 3%%u of the measured value Pressure transducer P 4.1 Linearity error of the transducer 0 3g of 50 bar = 0.75$ of 20 bar 0. 75'5 Reproduction error of the transducer 0.1% of 50 bar 0. 05 bar Sensitivity error relative to 10 K temperature difference 0 1%

Error of the measuring amplifier 0 5%

Error of the balancing unit and the recorder 0. 5'7a maximum total error ~+ 0 ~ 05 bar + 1.85% of the meamured value~

Pressure Transducers P 5.1~P 5.5~ P 6.9 Linearity error of the transducer 0.3% of 20 bar = 1.5]i of 4 bar 1 5X Reproduction error of the transducer 0.1% of 20 bar 0.02 bar Sensitivity error relative to 40 K temperature difference 0. 4X Error of the measuring amplifier 0. 5%

Error of the balancing unit and the recorder 0. 5%

Maximum total error +0.02 bar + 2.9% of the measured value Pressure transduers P 5.~2 P5 3~ P 5.4 Linearity error of the transducer 1% of 10 bar = 2.5% of 4 bar 2. 5Fd Sensitivity error relative to 40 K temperature difference 2 pO Error of the measuring amplifier 0 5%

Error of the balancing unit and the recorder 0 5%

Rev. 3, 7/80 9-15

Maximum total error +5 5% of the measured value Differential-pressure transducers P 4. 2~ P 5. 6~ AP 6 1~ AP 6 2 Linearity error of the transducer 0. 5%

Sensitivity error relative to 10 K temperature difference 0 2%

Error of the measuring amplifier 0. 5%

Error of the balancing unit and the recorder 0. 5%

maximum total error +25 of the measured value Displacement transducers DG 6. 1 .. 6. 5 Error of the transducer Error of the measuring amplifier 0 5%

Error of the balancing unit and the recorder 0. 5%

Maximum total error +2% of the measured value Acceleration transducers AG 6.1~ AG 6.2 Linearity error of the transducer 0. 75%

Sensitivity error relative to 10 K temperature difference 0-2%

Error of the measuring amplifier 0. 5X Error of the balancing unit and the recorder 0 total error 5'aximum

+ 2X of the measured value Strai~n gules SG~ LC Tolerance of the k-factor Influence of temperature on the k-factor Error of the measuring amplifier 0. 5%

Rev. 3, 7/80 9-16

Error of the balancing unit and the recorder 0. 5X Maximum total error + 5% of the measured value Temperature measuri~n ~pints Error of the transducer 1 K Error of the measuring amplifier 0- 5%

Error of the balancing unit and the recorder 0. 5%

Maximum total error + 1 K + 1% of the measured value Repeated recalibrations yielded far better results than indicated by the list of errors.

An overall inspection of the pressure transducers in the water region by incremental lowering of the water level (see Subsection 9 2.2.6) yielded maximum deviations of approximately +0.005 bar and -0.003 bar from the nominal value.

The deviations are illustrated as a frequency distribution in Figure 9-14. They are characterized by a Gaussian distribution.

In order to record the high frequency process with correct frequency and amplitude, the measurement chains were designed for the dynamic. range anticipated during the tests. The dynamic range was limited by the carrie frequency measuring amplifier to approximately 1.4 kHZ, which was substantially less than the 10 kHZ eigenfrequency of the pressure transducers. The magnetic tape recorders did not impose any limitation with a frequency cut-off of 2- 5 kHz The frequency cut-off of the Visicorders was determined by the utilized galvanometers. They were at 1 kHz for all the high-frequency measuring points. The frequency characteristics of the individual galvanometers was inspected before the tests.

9 3 TEST PARAMETERS AND MATRIX The test matrix provided for twenty-two tests with eleven different parameter combinations (see Table 9-4) . Earlier test series indicate that the strength of the condensation events is very highly stochastic and can differ for tests with identical boundary conditions. In order to largely rule out any erroneous correlation of measurement values with the parameters, each test is repeated once.

Four different line breaks were investigated:

Rev. 3, 7/80 9-17

2 the complete break-off of a recirculation loop (RCL break)

3) the complete break-off of a main-steam line (full MSL break) two other steam-line breaks corresponding to 1/3 and 1/6 of

3) the full MSL break area.

For the RCL break, the break flow consists of both liquid and steam flow. A portion of the liquid flashes into steam and together with the steam from the break gives the total steam flow into the suppression pool. FSAR Table 6 2-9 presents the break steam flow and break liquid flow, together with their associated enthalpies at various times during the RCL break. FSAR Figure 6.2-2 shows the drywell pressure response for the RCL break.

This data was used to calculate the fraction of liquid break flow that flashes into steam (assuming thermodynamic equilibrium), and the corresponding total vent steam flow. Figure 9-14a shows the SSES calculated vent steam mass flux vs. time f or the RCL break.

The RCL tests were run to match this curve as closely as possible (see Subsection 9.4.l.l.l).

For the full MSL break, the break flow is also comprised of both liquid and steam flow. Again, a portion of the liquid flashes into steam and combines with the steam from the break to give the total vent steam flow. FSAR Table 6.2-10 gives the break steam flow and break liquid flow, as well as their associated enthalpies at various times during the full MSL break. FSAR Figure 6.2-11 shows the drywell pressure response for the full MSL break. This data was used to calculate the fraction of liquid break flow that flashes into steam (assuming thermodynamic equilibrium), and the corresponding total vent steam flow.

Figure 9-14b plots the SSES calculated vent steam mass flux vs.

time for the full MSL break. The full MSL tests were run to match this curve as accurately as possible (see Subsecton 9411)

For these larger break transients the range of low mass flow densities is passed through very rapidly. In the event of smaller breaks the blowdown times are distinctly longer. The 1/3 and 1/6 MSL breaks were chosen to investigate longer blowdown tzansients. Their break sizes were selected so that, if required, it is possible to, compare the results with data known from earlier tests series.

The test matrix provides for tests at initial water temperatures of 24OC, 32~C and 55~C (75~P, 90~F and 1300P). The value of 320C corresponds to the mean temperature which is maintained by the cooling system of the suppression pool during normal plant operation The emphasis on the tests at 32OC is explained the fact that no clear dependence of the condensation loads on by the water temperature was observed in previous test series. The temperatures 24oC and 550C were taken from the limits of the operation field of the pressure relief system of the plants.

Rev. 3, 7/80 9-18

The amount of air flushed over from the drywell influences the backpzessure in the suppression chamber and also the composition of the air-steam mixture flowing through the vent pipe.

Host of the tests are performed with the same (proportionate) amount of air as in the plant. The steam is introduced in such a manner that it can mix in a mostly homogeneous manner with the air. By introducing, cool air to the drywell just before the beginning of the test, the air temperature is brought to a temperature corresponding to that in the plant. To investigate the ef feet of a possible incomplete steam-air mixing, individual tests are performed with reduced air content in the drywell. In those tests, cool air is not introduced into the drywell. The air temperature is then raised by means of the drywell wall heating system mentioned previously in Section 9.2. Thus, the mass of air is decreased by about 15%.

A detailed listing of the test parameters and operating conditions measured before and after each test is contained in Table 9-5. The following parameters are compiled in this table:

Test duration Bottom clearance and submergence Water temperature in the test tank Temperature of wall and air in the drywell Mater volume in the accumulator S 6

'Pressure in the accumulators S 6 and B 202 Pressure in the drywell and in the air space of the suppression chamber Air content in the drywell Diameter of the flow limiter.

The initial and final values were obtained from the computer listings (see Subsection 9.2.2.4). The air temperature in the drywell was not read from the listings >>before test," but rather they were obtained from a listing just after the shutdown of the ventilator connected to the drywell some time before the beginning of the test.

Por the water temperature and the air temperature in the drywell, the mean value was formed from the corresponding measuring points.

At the end of the test, the water temperature after the mixing of the pool was indicated.

Rev. 3, 7/80 9-19

9-4 IEST RESULTS See the Proprietary Supplement for this section.

Rev. 2, 5/80 9-20

9. 0 GKM-IIM STEAM BLOWDOWN TESTS 9 5 DATA ANALYSIS AND LOAD SPECIPICATION See the Proprietary Supplement for this Section.

Rev. 2, 5/80 9-21

9 0 GKM-ZIM STEAM BLONDOWN TESTS 9 6 VERIFICATION OF THE DESIGN SPECIFICATION See the Proprietary Supplement for this Section.

Rev. 2, 5/80 9-22

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Calibration Steps for all Recorders Rev. 3, 7/80 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT INTERVALS FOR CALIBRATIONS, CHECKS AND ADJUSTMENTS FIGURF 9-12

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P6.8 BY LOWERING OF THE WATER LEVEL FIGURE

350 300 k g/Z2 S 250 200 150 100 50 10 20 30 S 40 TIME Rev. 3, 7/80 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT CALCULATED SSES VENT STEAM MASS FLUX VS ~ TIME RCL B REAK FIGURE 9-1 4 a

250 200 kg/m s 150 100 50 10 20 30 40 50 s 60 time Rev. 3, 7/80 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT CALCULATED SSES VENT STEAM MASS FLUX VS'XME-FULL MSL BREAK FIGURE 9-14b

TABLE 9.1 COMPARISON OF FIXED PARAMETERS GKM II M SSES Test Vessel Single Cell (As Built Values)

'3 Drywell Free Volume, m 77.9 75.6 (Including Vent Pipe at High Water Level) 3 Wetwell Free Air Volume, m 48.4 47 (High Water Level)

Drywell/Wetwell Air Volume Ratio 1.61 1.61 (High Water Level)

Free Pool Area, M2 Small Cell at Containment Wall 3.7 3.77 Mean Value 5.8 Vent Pipe Dimensions Length, m 13.86 13.76 Outer Diameter, mm 610 610 Wall Thickness, mm 9.5 10.0 Vent Pipe Submergence, m 3.66 3.66 (High Water Level)

Vent Pipe Clearance, m 3.35 3.63*

(Exit to Pool Bottom) to 3.54 Distance Between Bracing 2.44 2.44 and Vent Opening, m Volume Flexibility of 0.6 0.6 Wet Containment Walls, dm /bar

• At the Deepest Point Rev. 5, 3/81

Rev. 3, 7/80 Table 9-2, (1 of 2)

OPERATING INSTRUMENTATION

)}C Transducer Data Recording Place Transducer Meosuring - Point Marking Measuring - Localion Type Measuring Stock40. Measuring OPS Control Test Ran e Am ifier Channel Station Focilit Pressuro in the superheated stean line to the PG vith I' rene'te 25 bar 20 nA

~ 1 stean line stean accuaulator sensor P 1 2 ~ aln stean line P 1 ~ ) Pressure ln the feedvater tank feedvater tank B 20) feedvater tank B 202 P ),1 Pressure In the stean accuaulator, 2/3 stean cone Vater level ln the stean accueulator Barton 1.565 bor 2/6 b )~ 1 cell 20 nA I: ) ~ 2 0.1825bar 20 nA 2/7 Vater Ga0e Pressure In the blovdovn line PG with P A~ 1 remote 25 bar 20 mA before throttle noasl ~

sensor Pressure ln the blovdovn line PG vlth P A.A before the rupture discs renot ~ 25 bar 20 nA 2/4 sensor P A,5 Pressure betveen the tvo rupture ckcs 2/5 Tesperature ln the superheated stean line to the DCA T 1~1 stean line stean accuaulator RTO, 550 C 2/8 T to2 nein stean line RTO 400 C 2/8

• A more exact position indication for these sensors is not mecessary for the usability of the measurement signals and for the test execution

Table 9-2, (2 of 2)

Rev. 3, 7 I 80 OPERATING INSTRUMENTATION Transducer Data Recording Place Transducer Measuring - Point Morking Measuring - Locotion TYpe Measuring Stack-No. Measuring OPS Control Test Ranae AA'40lifier Chorftet Station Fociht Tesperature In the feedvater tank feedvater tank 2 203 ATa 300 C DCA 4/13 T21 foedvater tank 0 202 2/10 T 3~1 Teeperature in the stean accunulator, 2/11 stean zone T32 Tesporature in the stean accunulator, 2/12 vater zone Taoperature for the correction of 2/13 T 3 ' the eater level eeasurenent ln the stean accuaula'tor T 6~1 Tcoperature in the hlovdovn line CTC 250 0C 2/1 4 heforo thc throttle nozzlo Tenperature in the dryvell,cl ihe well 2500C 3/11 T 5~6 parton 5/5 L 6.1 Valor lovel ln tho suppression pool coll 1235 bar 0 ~ ~ 20sA L 6.2 Vator level in the annulus pap 1235 bar 2/15 2

• A more exact position indication for these sensors is not necessary for the usability of the measurement signals and for the test execution

Rev. 3, 7/80 Table 9-3., (1 of 5)

TEST INSTRUMENTATION Measuring Location Transducer Datct Data Recording Tronsducer Meosuring -Point Morking Level Angte Type Measuring Stock- Measuring D PS Visi- Mogne l<c H (mm) Ronge Amp5rier Chonnel corder Tope SG sr, 6.3 Strain In the. tank vali ~

2650 135 seal Gooops/s CFA outside,vertical bridge so G.C Strain In the tank vail ~

265o 135 outside,horlsontal SG 6 5 Strain In the tank vail ~

45 outside,vertical 2650 sr. 6.6 Strain in the tank 2650 45 vali'utside,horizontal so67 Vertical bending strain at 1093 5/7 the quencher duesy so68 Ilorisontal bending strain 1093 5/8 at the quencher dussy SQ G,9 Vertical bending strain at the legs of the test vessel SG 6 ~ 10 Vertical bending strain at the 90 legs of the test vessel X

OG 6.2 Displacesent of the inner 10000 270 2 mm X OG 6,6 cylinder at the crossing place 90 X

OG 6.3 Dlsplacesent of the Inner 0 10000 DG 6.5 cylinder at the crossing place 180 X DG 6 ~C Dlsplacesent at the stiffening ring 6107 90 4/10 Tesperature In the suppression 6800 CTC 150 C DC A 3/0 T 6~I pool,vater cone 180 5200 180 3/1 T 6.2

• 7 0 mm from the middle of the weld seam/quencher arm

• • 100 mm below the weld seam at the leg of the vessel

Rev. 3, 7/80 Table 9-3, {2 of'5)

TEST INSTRUMENTATION Measuring Location Transducer Data Data Recording Transducer Meosuring -Point Mor'king Level Angle TVpe ~ 'tock-Nct Meosurlng Meosuing OPS Vi5<- Mognetic H {mm) Ronge Ampbfier corder Tope T 6,) Teeperature In tho suppression 3651 CTC 150 C DC A 3I2 pool,vater xone T64 2653 180 " 3/3 T 6.7 1097 180 3/4 T 6.8 3IS Tooperature In tho suppression T 6.P pool,air xone,top 1COOO 3/6 T 6.10 Tesperature In the suppression 8010 3/7 pool,air xone,belov 270 BG

+P Pressure ln the blovdovn line cooplete 50 bar CPA 5 I1 C

before throttt ~ noxxle brldpe Differential pressure at the nP C ~ 2 throttle nocti ~ )5 bar 5/2 ap C.) Barton bar 20 nA cell C 5/0

+T C.l Teeperature In the blovdovn line before throttle noxxl ~ CTC 250 C DCA 2/14 Plexo-

.Dynanlc pressure ln the stean electric Char po p )~e a<<cueulator,vater sons 20 bar

~ epllfler transducer BC P 5~ I Pressure In the dryvell cooplete 20 bar CPA brldpo P 5~2 Pressure ln the dovncooor pipe,top 15550 225 10 bar P 5~ ) ~ Pressure" ln tho dovncooer pipe,olddle 10580 270 The arrangement of the sensors required for the steam flow

• • 200 measurement is according to DIN 1952 mm out of center

• • • The sensor was instralled according to the drawing R 523 G 22 1986

Table 9-3, (3 elf 5)

Rev. 3, 7/80 TEST INSTRUMENTATION Measuring Location Transducer Data Data Recording Transducer Measuring-Poinl Morking Level Angle Type Measurngi S fock- Na Measuring DPS Visi- Magnelic H {mm) (0) .Ronge Amplifier Channel corder Tope SG P S,cc Pressure In the dovncoaer pipe,bolov 7320 270 coaplete IO bar CFA bridge P S~S Pressuc' In the dovncoaec pipe,exit 3750 270 20 bar AP 5.6 Pressure differential betvaen dry bac.

veil and suppression chaaber 3,5 Oil 5 ~ I Oxlpene rat ~ ln the dovncoaer pipe I 5290 180 6/3 Indication of the evinp check DG 5 I valve be<<veen dryvell and CFA 3/14 suppression chasber DG 5,2 3/IS TS I Toaperature in tho dryvell,top C T C 260 C D C A

~

3/8 T 5~2 Tesperature in tho dryvell,belov 3/0 T 5 ' Teaperaturo In the dryvellcsuap 3/10 Ieaperatura In the T ST 5 dovncoaer pipe, ~ Iddle I0580 3/I'2 T56 Teaperaturo ln tho dovncoaercexlt 3750 3/13 X SG I Pressure at the suppression pool P 6~

vail,vater zone 6156 180 <<oaplete IO bar CFA 4/o bridpe P 6,2 4155 180 4/I

Rev. 3, 7/80 Table 9-3, (4 o f 5)

TEST INSTRUMENTATION Measuring Measuring Location Transducer Data Data'ecording Transducer Measuring-Poinl Marking Level Angle TYpe Measuring Stock-Nct DPS Visi- Mognetic H (mm) {o) Range Ampbfier Channel corder Tape SC Pressure at the suppress!on pool 4/2 p 6.) vali,vater xone 3651 180 conplete 10 bar C F A brlgge P 6.A 2653 180 4/3 2653 4/4 p 6.6 2653 4/5 P 6,7 1097 180 4/6 p 6.8 4/7 P 69 Pressure ln the suppression 16770 20 bar 5 I5 chaaber,air xone AF 6.1 Alr rractlon In the suppression 4155/6156 ),5 bar 4111 chaabei' vs'tel'one 180 AF F 2 26 53/6156 180 4/12 X l Voter level In the dovncoaer Pipe park 0 CA LP 5 ~ )750 90 plug LP S 2 %0)0 90 LP5) ALSO 90 LP 5*6 5950 90 Ll'.5 7950 90

• 200 mm out of center '

Rev. 3, 7/80 Table 9 3, (5 TEST INSTRUMENTATION Measuring Location .Transducer Data Data Recording Tronsdvcer Meosuring-Point Morking Level Angle TYP e Meosurin g Stock-H Q. M costs'ing DPS Vist- Mognetic H (mm) Ronge Arrp5fier Chonnel corder Tope SCi SG (..1 l.onoi tudi na1 strain,Brac ino 1 6107 counlete 6000 Itm/m CPA 4I8 brl duo sr, 6.0 l.onoi tudina1 strain,oraclna 2 6107 50 4/9 IC 6 Ionds oh.the I Scoot 6522 870 5/9 sa 5,1 8endioo strain in the dovncoaer 6007 90/270 sa 5 8 6007 OI180 SG SG 5 ' 15700 90I 270 semi bridge SG 15700 0/180 SG

• G 6.1 Acreleration of tho inner cylinder center comptete ae59 bridge AG 6, 7010 90 Wafer level in the suppression pool Barton 1.235 bar 0...20 mA 5/6 cell

Rev. 3, 7/80 Table 9-4 GKM II-M TEST MATRIX Test Number 123 6789 10 11 '12 14 15 17 18 19 20 33 34 8 210 (RCL) g 190 (IRSL)

Breat< Size (mm) 8 110 (1/3 HSL) 8 80 {1/6 t>SL) 2~'C {75 F)

Pool Temperature S2'C (90 F) 55'C (130 F) 100% gC Dry~iell Air Content 85 % (approx.)

Repeat Test

Table 9-5, (1 0 f 3)

Rev. 3, 7/80 TEST PARAMETER l<ater Tcnp. at th> Tenp. in tbc lfnter Prcssure Drywall Cond.Chanbcr Drywcll Air Di sn>ntnr at Test Sub tenperature Drywcll,Wall Drywell Air Volunc in S6 Pressure Press>>rc Content the Flew Dnration Start E>>d Star t Spacn in S6 Start Start Fnd Start F.nd Start l<estri ctor Start Start

,le oc oc oc l>a> ba> l>ar bnr l>nv 37 3.6 3.7 34 65 3'l-n "19.8 1' 30 1.0 ~ >

8 210 3.6 3.7 33 66 177 58 t)9 ss19 6 'l 0 3 2 1.0 .8 1OO 2'l 0

3. 6) 3.8 24 27 140 6r 'l9. 7 0 3~1 1~0

~ '>

~ ~ y 100 190 4 73 3.6 3.7 25 59 1/> 2 7.6 19.0 1 ' 3 2 1 ' 100 190 3.6 3.6 32 64 61 7.8 19.8 l 0 3.2 1.0 2.8 100 1')0 3.6 3.6 33 65 62 7.6 19.8 l.o 3.3 1.0 'l 00 81 3.6 3.7 33 70 170 146 7.8 17-3 1 ~ 0 3~O 1.0 5 l90 n 8o 3' 3.7 34 68 160 7s3 17.0 1 0 3 0 1.0 C) 190 74 3 ' 3.6 56 87 143 7.6 19.8 0 3 3 1 0 9 100 190 10 71 3 ' 3.7 55 84 55 7 ' 18.6 1.1 3.4 1.o 2 100 190 Dh a i>ista>>ce fro>> thc Dotton the same pressure in the- B "02 Su)> a S<<boerdence a Attnr )lixing ol the Pool

'1 Rev. 3, 7(80 Table 9-5, (2 of 3)

-TEST PARAMETER Water Tempi at th( Temp. in the Water Prcssure Dry<<nil Cond.Chnmbrr Dry<<(. 1 l A) r i)iamnter at Test tcmperaturc Drywcll Wall l)ryvcll Air Volume in SG Pressure Prcssure Content the Plov Tr st Duration DB Sub Start Start Start Start l(cstrictor Start End Start Space in SG End End Start Star t i~ . oC oC oC oC m 3 1)ar )lat'rar )

a)'2 2') 4 3.6 3 ~ Q 3'Q 1ll3 0.7 1 7 ~ 2i 3 ' 1.0 100 110 12 216 3.6 3.0 35 67 57 0.5 17 F 1 F 1 3. 3 '1.0 r) ~ () 100 110 23 l:2Q 3.0 25 6o 63 9-5 1 7~ 3.3 1.0 r)

<< ~ LJ() '1OC) r) ~ (. 00 3.6 3 0 6" Q 5 17- ') 1 0 I ~

100 25 li13 3. 6) 3-0 33 63 1ll 3 6)7 0.7 27 ' 3-2 1.0 <<r t)() 100 00 16 l(03 3.6 3 0 33 65 139 Gi( Q 5 27 r) 3 ~

~ ) ~ 0 2.00 00 17 ll 3Q 3.6 3. ( 3li 60 170 8.7 17 ' 1.0 >.0 2.0 ~ )) 05 00 ll 32 3.6 3~ .'3 69 173 2/< 5 0.7 27-1 1 0 2.0 1 0 r<<0r 420 3.G 3. 0 52 Qil 137 70 27r2 1.0 3.1 '1 0 ~ 7 2.00 Qo 20 ll 03 3.6 3.6 55 85 l37 6)3 '17 ~ 1 1.0 3.0 1.0 Q 100 i)D ~ Ui) tancc from the Bottom Suh ~ Submergence After Hixin!) of tbc Pool

Rev. 3, 7/80 Table 9-5,(3 of 3)

TEST PARAMETER

<<ates Tcnp ~ at th( To!I>i> in thc tCntcr Prcssure Dlywcll Cond.Chauhcv Drywcl1 Alr llia!!>r.LI'r al.

TosL Drywcl1 Mall Drywall Air Volusc iu Prrssul c Prcssure Contcllt, lhc Flow Tost Duvation DD Suh Tcupcl'n Lure 56 Start Eud Stapt Space 1 n S6 Start Start End Star t, End Start llcstrictor Start Stav t hnv l>nl l>av l ~ I>v 3G 3,6 3,6 51t 85 62 9,8 313 111 9 100 3 t 36 3,6 3 GI 5lt 85 182 59 30,4 19,8 3,It 1,1 3,0 100 210 u>s - ulul ~ >I oo "ron tho Dot Lux>

~

• +tile satde pl esstll-e in till>lD 202 Suh " Suhu>'I'us!ucl!

~ a Aftl.r Nixing of tho Pool

CHAPTER 10 RESPONSES TO NRC QUESTIONS TABLE OP CONTENTS ~

10 1 NRC QUESTIONS 10 1 1 IDENTIFICATION OF QUESTIONS UNIQUE TO SSES 10.1 2 IDENTIFICATION OF QUESTIONS PERTAINING TO THE NRCi S REVIEW OF THE DAR 10 1.3 QUESTIONS RECEIVED DURING THE PREPARATION OF THE SAFETY EVALUATION REPORT {SER) 10 2 RESPONSES 10 2 QUESTIONS UNIQUE TO SSES AND RESPONSES THERETO 10.2. 2 QUESTIONS PERTAINING TO THE NRCiS REVIEW OP THE DAR AND RESPONSE THERETO 10 2 3 QUESTIONS INFORMALLY RECEIVED DURING THE PREPARATION OF THE SAFETY EVALUATION REPORT (SER) AND RESPONSE THERETO 10 3 FIGURFS REV. 6, 4/82 10-1

CH A PTER 10 FEGUQQS Number Title 10-1 This figure has been deleted.

10-2 This figure has been de1eted.

10-3 Special relationship of downcomers and pedestal holes 10-4 Transducer locations for the ten vent pipe configuration 10-5 Transducer locations for the six vent pipe configuration 10-6 Transducer locations for the two vent pipe configuration 10-7 Typical pressure time histories from pressure transducers P20, P25 ... 29 and P134 1 0-8 Typical pressure time histories from pressure transducers P20, P25 ... 29 and P134 1 0-9 Freguency distribution of measured normalized wall pressures 1 0-10 Pool wall pressures at three circumferential vent exit locations 1/6 scale 3 vent geometry 10-11 Pool wall pressures at three circumferential vent exit locations 1/10 scale 19 vent geometry 10-12 Plan locations of transducers for wetwell 10- 13 Locations of pressure transducers for wetwell 10-14 Vent exit elevation pool wall pressures for a chug from JAERI test 0002 1 0-15 Comparison of probability density of the normalized pressure amplitudes from GKM II-M tests 3 ... 10 and JAERI 10-16 Comparison of probability density of the normalized pressure amplitudes from GKM IT-M tests 11 6 12 and JAERI 10-17 Comparison of probability density of the normalized pressure amplitudes from GKM II-M tests 13 ... 20 and JAERI 1 0-18 Comparison of pressur e response spectra of test 21. 2 all valve case and the SSES load definition 1 0-19 Comparison of pressure response spectra of test 21. 2 all valve case and one valve case and the SSES load definiti REV. 6, 4/82 10" 2

figures (Cont.)

Table SSES containment response spectra - KWU SRV¹76 Asymmetric d,irection horizontal 10-21 SSES, containment response spectra, KMU SRV¹76 - Asymmetric direction vertical l 10-22 SSES containment response specttra KWU SRV¹76 Asymmetri direction horizontal 10-23 SSES containment response spectra KWU SRV¹76 Asymmetric direction vertical 10-24 SSES containment response spectra KMU SRV¹76 Asymmetric direction horizontal 10-25 SSHS containment response spectra KMU SRV¹76 - Asymmetric direction vertical 1 0-26 SSES containment response spectra - KMU SRV¹76 Asymmetric direction horizontal 10-17 SSFS containment response spectra KMU SRV¹76 - Asymmetric direction vertical SSES containment response spectra - KMU SRV¹76 - Asymmetric direction horizontal 1 0-29 SSES containment response spectra KMU SRV¹76 - Asymmetric direction vertical 10-30 SSES containment response spectra KWU SRV¹76 - Asymmetric direction horizontal 10-31 SSES containment response spectra KWU SRV¹76 Asymmetric direction vertical 10-32 SSES containment response spectra KMU SRV¹76 Asymmetric Lc direction horizontal 1 0-33 SSES containment response spectra KWU SRV¹76 Asymmetric direction vertical 10-34 SSES containment response spectra KHU SRV¹76 Asymmetric direction horizontal 10-35 SSES containment response spectra KWU SRV¹76 Asymmetric direction vertical 10-36 SSES containment response spectra KWU SRV¹76 Asymmetric direction horizontal REV. 6$ 4/82 10-3

FIGUQgS (Con t. }

Number Title 10-37 SSES containment response spectra KMU SRV476 - Asymmetric direction vertical 10-38 SSES'containment response spectra KWU SRVf76 Asymmetric direction horizontal 1 0-39 SSES containment response spectra KMU SRV$ 76 Asymmetric direction vertical 1 0-40 SSES containment response spectra KMU SRV476 Asymmetric direction horizontal 10-41 SSES containment response spectra - KMU SRV$ 76 Asymmetric direction vertical 10-42 LGS containment response spectra KWU SR V476 - Asymmetric direction horizontal 10-43 LGS containment response spectra - KMU SRV476 - Asymmetric direction vertical 1 0-44 LGS containment response spectra KWU SRV476 - Asymmetric direction horizontal 10-45 LGS containment response spectra KM 0 SRV476 Asymmetric direction vertical 10-46 LGS containment response spectra KWU SRV476 Asymmetric direction horizontal 1 0-47, LGS containment response spectra KWU SRV476 - Asymmetric direction vertical 10-48 LGS containment response spectra KMU SRV476 Asymmetric direction horizontal 1 0-49 LGS containment response spectra KMU SB Vf76 Asymmetric direction vertical 10-50 LGS containment response spectra KWU SRV$ 76 Asymmetric direction horizontal l 0-51 LGS containment response spectra KMU SR V476 Asymmetric direction vertical 10-52 LGS containment response spectra KMU SRV476 - Asymmetric direction horizontal 10-53 LGS containment response spectra KMU SRV476 Asymmetric direction vertical REV. 6, 4/8Z 10-4

FIGURES (Cont.)

Number Tittle 10-50 LGS containment response spectra KWU SRV476 Asymmetric direction horizontal 10-55 LGS containment response spectra '- KWU SRV476 Asymmetric direction vertical 10-56 LGS containment response spectra KWU SRV476 Asymmetric direction horizontal 10-57 LGS containment response spectra KWU SRV476 Asymmetric direction vertical 10-58 LGS containment response spectra KWU SRV476 Asymmetric direction horizontal 10-59 LGS containment response spectra KWU SRV476 Asymmetric direction vertical 1 0-60 LGS containment response spectra KWU SRV476 Asymmetric direction horizontal 10-61 LGS containment response spectra KWU SRV476 Asymmetric direction vertical 10-62 LGS containment response spectra KWU SRV476 Asymmetric direction horizontal 10-63 LGS containment response spectra KWO SRVS76 Asymmetric direction vertical 10-64 Reactor Pressure Transient Case 2.a Without Shutdown Cooling 1 0-65 Suppression Pool Temperature Transient Case 2.a Without Shutdown Coolinq i

~,, i '

~

~

ip ii REv. 6, 4/8? 10-5

CHAPTER 10 TQBQQS gggle 1 0-1 Normalized RNS vent static pressure and variance JAERI data 1 0-,2 Comparison of JAERI/GKH II-5 normalized mean vari'ance I, ~L II III L

t l.'

- 'I ~

L II L;

lI,""'

i4: I

~ I~ ~

~

I~

L L

REV. 6. 4/82 10-6 L

I

10 0 RESPONSES TO NRC OUESTIONS This chapt'er will provide responses to those Nuclear R'egulatory Commission (NRC) questions which have been designated by Reference 10(as amended) to be found in the plant-unique Design Assessment Report, to those questions- for which the response in Reference 10 is inapplicable, to those questions generated from previous NRC reviews of the plant unique DAR, and those question's received durinq preparation of the SER. The NRC questions for which responses will be provided are identified in Subsections to these

10. 1. 1, 10. 1.2, and 10. 1. 3, and detailed resposes questions are found in Subsections 10.2 1, 10.2.2 and '10.2.3.

h REV. 6p 4/82 10-7

The below listed guestions address concerns unique to SSES.

These questions are ansuered in detail in subsection 10.2.1 N020. 26 Primary and Secondary LOCA Loads M 020 27 Inventory Effects on Blowdown M 020. 44 Poolswell Waves and Seismic Slosh N020. 55 SRV Loads on Submerged Structures N 020 58 (1) ~ (2) ~ (3) Plant Unique Poolsvell Calculations N020 59 (1), (3) g (4) Dovncomer Lateral Braces N020 60 Wetvell Pressure History N020. 61 Poolsvell Inside Pedestal N130. 1 Pressure Loading Due to SRV Discharge N130. 2 Load Combination History N130. 4 Soil Nodelinq N130 5 Liner and Anchorage Nathematical Model N130 6 Containment Structural Nodel-Asymmetric Loads N 130. 12 SRV Structural Response RZV. e, 4/82 10- 8

10.1.2 - XDENTIPICATION OF QUFSTIQQS PERTQIQING TO THE NRCis )S REVIFQ OP THE "DAR I

The below. listed questions address concerns generated as a result of the NRC',s review of the DAR. These questions are answered in detail in Subsection 10. 2. 2 Ouestion Number Question Tonic 1 . NUREG-0487 Acceptance Criteria 2 Drywell Pressurization 3 Chuqqing Loads on Submerged Structures 4

5 IBA and SBA for Typical Nark II Poolswell Haves and Seismic Slash Containment List of Piping, Equipment, etc., Subject to Pool Dynamic Loads 7 Applicability of the Generic Programs Tests and Analysis to the SSES Design 8 Time History of Plant Specific Loads 9 Mass and Energy Release 10 "Local" and <<Bulk<< Pool Temperature 11 Suppression Pool Temperature Monitoring System REV. 6, 4/82 10-9

10 1. 3 OUESTIONS 9 RECFXVED DURING THE PREPARATION OF THE The below listed questions were informally received during the NRC~s preparation of the SER. These questions are answered'in detail in, Subsection 10.2.3.

Question Numb SSES LOCA Steam Condensation Load Definition

{SER Item ¹27) 2 T-Quencher Frequency Range {SER Item ¹28)

SSES ADS Load Case {SER Item ¹28) 4 Quencher Bottom Support at Karlstein {SER Item ¹28) 5 Bendinq Moment in the Quencher Arm Recorded at Karlstein {SER Item ¹28)

Suppression Pool Temperature Response (SER Item ¹30)

Local to Bulk Temperature Difference for SSES (SER Item ¹30)

Ouencher Steam Mass Flux (SER Item ¹30) 10-10 REV. 6, 4/82

10 2 RESPONSES A

10 2 1 OUESTIONS UNIQUE TO SSES AND RESPONSES THERETO QUESTION $ 020 Q6 The DFFR presents a description of a number of LOCA related hydrodynamic loads without differentiating betveen primary and secondary loads. Provide this differentiation betveen the primary and secondary LOCA-related hydrodynamic loads. We recognize that this differentiation may vary from plant to plant.

We vould designate as a primary load any load that has or vill

'result in a design modification in any Nark II containment since the pool dynamic concerns vere identified in our April 1975 generic letters.

RESPONSE 8020 g6 The table belov shows the LOCA-related hydrodynamic loads on the SSES containment. Those loads vhich have resulted in containment desiqn modifications are designated as <<Primary Loads." These primary loads result from the poolswell transient.

Drywell floor uplift pressures durinq the wetvell compression phase of poolswell lead to the decision to increase the SSES drywell floor desiqn safety marqin for uplift pressures hy relocating dryvell floor shear ties.

Poolsvell impact, drag, and fallback loads resulted in the relocation of equipment in the SSES wetvell to a position above the peak poolsvell height. Furthermore, the dovncomer bracing system vas redesiqned.

All other LOCA-related hydrodynamic loads are designated as

<<Secondary Loads" since no design modification has resulted from their .presence.

LOCA Load <<Primarv Load" "Secondary Load"

1. Wetwell/Dryvell Pressures x<<~

(During Poolswell)

2. Poolswell Impact Load xc z3

3. Poolswell Drag Load x<>>

4. Downcomer Clearing Load 5 Downcomer Jet Load

6. Poolsvell Air Bubble Load

.7 Pools well Fallback Load x<+)

Rev. 2, 5/80 10-11

LOCA Load <<Primarv Load>> "Secondarv Load"

8. Mixed Plow Condensation Os i ti on c 1 la Load

9. Pure Steam Condensation Oscillation Load

10. C hug qinq

11. Metwell/Drywell Pressure and Temperature durinq DBA LOCA X (Long Term)

12. Retwell/Drywell Pressure and Temperature during XBA LOCA (Long Tern)

13. Metwell/Drywell Pressure and Temperature durinq SBA LOCA (Lonq Term)

Footnotes.

(1) Shear ties chanqed in drywell floor.

(2) Equipment moved in wetwell.

(3) Equipment moved in wetwell. Bracing system redesign.

(4) Fquipment moved in wetwell.

OUFSTXOM N020. 27 The calculated drywell pressure transient typically, assumes that the mass flow rate from the recirculation system or steamline is equal to the steady-state critical flow rate based on the critical flow area of the .jet pump nozzle or steamline orifice.

However, for approximately the first second after the break opening, the rate of mass flow from the break will be greater than the steady-state value. Xt has been estimated that for a

. Nark I containment this effect results in a temporary increase in the drywell pressurization rate of about 20 percent above the value based solely on the steady-state critical flow rate. The drywell pressure transient used for the LOCA pool dynamic load evaluation, for each Nark ZI plant, should include this initially higher blowdown rate due to the additional fluid inventory in the recirculation line.

R'F SPOUSE N 02 0 g7 The drywell pressure transients have been recalculated by GE (Reference 7) with the additional blowdown flow rate produced by

.the inventory effects included in the analysis. The LOCA loads

~

presented in Section 4.2 have been calculated using these Rev. 2, 5/80 1 0-12

recalculated dryvell pressure transients. Specifically, the dryvell pressure transient resultinq from the DBA LOCA including the effects of pipe inventory has been used as input to the poolsvell model.

OUESTION. M020 44 Table 5-1 and Figures 5-1 through 5-16 in the DPPR provide a listinq of the loads and the load combinations to be included in the assessment of specific Mark II plants. This table and these figures do not include loads resulting from pool svell vaves followinq the pool swell process or seismic slosh. We require that an evaluation of these loads be provided for the Mark containment desiqn. ll II RFSPONSE M020 44 Subsections 4. 2.4.6 and 4.2.4.7 provide our response.

QUESTION M020 55 The computational method described in DPPR Section 3. 4 for calculatinq SRV .loads on submerged structures is not acceptable.

Tt is our position that the Mark II containment applications should commit to one of the follovinq tvo approaches:

(1) Design the submerqed structures for the full SRV pressure loads acting on one side of the structures; the pressure, attenuation lav described in Section 3.4.1 of NFDO-21061 for the ramshead and Section A10. 3. I of NEDO-11314-08 for the quencher can be applied for calculating the pressure loads.

(2) Follov the resolution of GESSAR-238'I on this issue.

The applicant for GESSAR-238 NI has proposed a method presented in the GE report, >Unsteady Drag on Submerged Structures,<~ which is attached to the letter dated March 24, 1976 from G. L. Gyorey to R. L. Tedesco. This report actively under review.

's RESPONSE M020 55 Loads on submerqed structures due to SRV actuation are discussed in Subsection 4.1.3.7.

QUESTION M020 58 Relatinq to the pool svell calculations, ve require the following information for each Mark II plant:

(1) Provide a description of and justify all deviations from the DFFH pool swell model. Identify the party responsible for conducting the pool swell calculations (i.e., GE or the AGE) . Provide the program input and REV. 6$ 4/82 10-1 3

results of bench mark calculations to qualify the pool svell computer proqram.

(2) Provide the pool svell model input includinq all initial and boundary conditions. Shov that the model input represents conservative values vith respect to obtaining maximum pool svell loads. In the case of calcul'ated input, {i.e., dryvell pressure response, vent clearing time), the calculational methods should be described and justified. In addition, the party responsible for the calculation (i.e., GE or the AGE) should be identified.

(3) Pool svell calculations should be conducted for each Mark II plant.. The followinq pool swell results should be provided in qraphic form for each plant:

{a) Pool surface position versus time (b) Pool surface velocity versus time (c) Pool surface velocity versus position fd) Pressure of the suppression pool air slug and the wetvell air versus time.

RESPONSE N020.58 (1) A specific response to this question can be found in Subsection 4.2. 1. 1. Verification of the SSES poolswell model is provided in Appendix Section D.l.

(2) Input and discussion of the poolswell model input can be found in Table 4-17, 4-18 ~ and Section 4.2. 1.1.

(3) The requested qraphic results of the SSES poolsvell calculation can be found in Pigures 4-38, 4-39, 4-40, and 4-43.

OUESTXON N020 59 In the 4T test report NEDE-13442P-01 Section 3.3 the statement is made that. for the various Mark II plants a vide diversity'xists in the type and location of lateral bracing between downcomers

'nd that. the bracing in the 4T tests vas designed to minimize the interference vith upvard flow. Provide the following information for each Nark II plant:

A description of the downcomer lateral bracing system.

This description should include the bracing dimensions, method of attachment to the downcomers and walls, elevation and location relative to the pool surface. A sketch of the bracinq system should be provided.

(2) The basis for calculating the impact or drag load on the bracinq system or dovncomer flanqes. The magnitude and Rev. 2, 5/80 10-1 4

duration of impact or drag forces on the bracing system or downcomer flanqes should also be provided.

(3) An assessment of the effect of downcomer'lanqes on vent lateral loads.

f RPSPONSE M020,59 Subsection 7.1. 2..1 describes the SSES bracing system and the methodology for assessinq the adequacy of bracing system.

basis for calculating the impact or drag loads on P

(2) -

The the downcomer bracinq system (El. 668~) and, downcomer stiffener rings (El.: 668'nd El.. 682') is given in Section 4.2. The. magnitude and duration of impact or draq forces on the bracinq system and downcomer stiffener rings is also given in'ection 4.2 j (, 1 (3) This item is not applicable to 'the SSES design.

1 OHESTION M020 60-In the 4T test report NEDE-13442P-01 Section 5.4.3.2 the statement is made that an underpressure does occur with respect to the hydrostatic pressure prior to the chug. However, the pressurization of the air space above the pool is such that the overall pressure is still positive at all times during the chug.

Ve require that each Mark II plantprovide sufficient information regarding the boundary under'pressure, the, hydrostatic pressure, the air space and the SBV load pressure to confirm this statement or alternatively provide a boundinq calculation applicable 'to all Nark XI plants.

RPSPOgSE $ 020 60 This information is provided in Subsection 7. 1.3 of the DAR.

OOESTION M020. 61 Significant variations exist in the Mark II plants with regard to the design of, the wetwell structures in the region enclosed by the reactor pedestal. These variations occur in the areas of (1) concrete backfill of the: pedestal, (2) placement of downcomers, (3) wetwell air space volumes, and (4) location of the diaphragm relative to the pool surface. In addition to variation, between plants, for a qiven plant, variations exist in some of these areas within a given plant. As a result, for a given plant, significant differences in the pool swell phenomena, can occur in these two reqions. Me will require that each plant provide a separate evaluation of pool swell phenomena and loads inside of the reactor pedestal.

RESPONSE M020 61 REV. 6, 4/82 10-15

The SSES pedestal and vetwell area is shown on Pigures 1-1 and

10. 3. Due to the absence of dovncomers in the pedestal interior, no pool swell would be expected in this 'region. There are 12 holes in the pedestal, hovever, eight of which vould allov the flow of vater from the suppression pool to the pedestal during a LOCA. Some downcomers are near the pedestal flow holes, leading to the possibility that air could be blovn through the pedestal holes, which vould lead. to a greater pedestal pool swell than would be experienced by incompressible water flov alone. One would expect the pedestal pool swell to be much reduced from the suppression pool swell due to its relative separation from the suppression pool and the lack of direct charging from downcomer vents. Indeed, 1/13.3 scale model tests of the SSES pedestal design conducted at the'Stanford Research Institute under the sponsorship of FPRI show that the pedestal pool svell is less than 20 percent of the pool svell in the suppression pool (Reference 32). There is no piping or equipment inside the SSES pedestal and, since the pedestal'ool swell is very small, the only load involved due to pedestal pool svell would be a small ~ P across the pedestal due to different water levels between the suppression pool and the pedestal interior. This load is considered in the design of the SSES pedestal.

OUESTION M130. 1 Provide in Section 5 a description of the pressure 1oadings on the containment wall, pedestal wall, base mat, and other structural elements in the suppression pool, due to the various combinations of SRV discharqes, including the time function and profile for each combination. If this information is not generic, each affected utility should submit the information as described above.

gggPONSE M130 1 Chapter 4 describes the pressure loadings and time histories due to SRV discharqe and other hydrodynamic loads.

QUESTION M 130. 2 In DPPR Section 5.2 it is stated that the load combination histories are presented in the form of bar charts as shovn on Figures .5-1 throuqh 5-16. It is not indicated hov these load combination histories are used. In particular, it is not clear vhether only loads represented by concurrent bars vill be combined, and it should be noted that depending on the dynamic properties of the structures and the rise time and duration of the loads, a structure may respond to tvo or more given loads at the same time even though these loads occur at different times.

Also, althouqh condensation oscillations are depicted as bars on the bar charts, the procedure for the analysis of structures due to these loads has not been presented. Accordingly, the

.description of the method should include consideration of such conditions. Also, for condensation oscillation loads and for SRV

'.:.',;':oscillatory loads, include lov cycle fatigue analysis.

Rer. 2, 5/80 10-16

RESPONSE N130 2 The loads vill be combined according to Section 5.0. Section 7.0 describes the assessment methodology and results for the re-assessment of SSES for the hydrodynamic and non-hydrodynamic ~

loads.

OUESTION A

N130 4 Through the use of figures, describe in detail the soil modelling as indicated in DFFR Subsection 5.4.3 and describe the solid finite elements vhich you intend to use for the soil.

RESPONSE M130 4 Soil modelling is explained in Subsection 7. 1.1.1.

OUFSTION N 130 5 Describe the mathematical model vhich you will use for the liner and the anchoraqe system in the analysis as described in DFZR Subsection 5.6.3.

RESPONSE 5130. 5 The mathematical model which vill be used for analysis of the suction pressures is liner and the anchoraqe for hydrodynamic described in Subsection 7.1.3.

'OUFSTION N130 6 In DPFR Subsection 5. 1. 1. 1 it vas stated that the SRV discharge could cause axisymmetric or asymmetric loads on the containment.

In Subsection 5.4.1 an axisymmetric finite element computer program is recommended for dynamic analysis of structures due to SRV loads, and no mention is made of the analysis for asymmetric loads. Describe the structural analysis procedure used to consider asymmetric pool dynamic loads on structures and through the use of f iqures, describe in more detail the structural model which you intend to use.

RESPONSE N130. 6 The dynamic analyses and models used are explained in Chapter 7 OUESTION 8130. 12 Reference is made in DFPR Subsection 5.4.3 to studies of structural response to SRV load. Provide citations for this reference and vhere such studies are not readily available, copies are requested.

RESPONSE N 130 12 REV. 6, 4/82 10-17"

Studies mentioned in DPPR Subsection 5.4.3 are the results of analysis completed for a specific plant at the time of writing of the DPFR. Reference to the,studies was intended to indicate the need for considerinu strain dependent soil properties.. Ror the SSES analysis, Reference 33 is used to detereine the soil constants in the analysis.

10-18

10 2 2 - QUESTIONS PERTQINIQG TO Tgg NRC>

~ S BEVIEM OP THE DAB AND RESPONSE TH EBBTO QUESTION 1 The LOCA and SRV related pool dynamic loads that are currently acceptable to us are discussed in NUREG-0487. Table IV-1 of NUREG-0487 summarizes these Mark II pool dynamic loads. By on Table IV-1 the letter, dated February 2, 1979, you indicated, staff that vill LOCA related dynamic loads acceptable to the be adopted for SSES. Revise the DAB to incorporate this information and provide the same information for the SRV related pool dynamic loads. For both the SRV and LOCA loads indicate the alternative criteria that vill be used for each item for which an exemption is proposed and provide references that discuss these alternative criteria.

RESPONSF.-

See response to Question 021.69 contained in Volume 16 of the SSES FSAB and Table 1-4 of the DAR.

QUEST ION 2 Subsection 4.2.1.1 of the, DAR state that the dryvell pressure transient used for the pool swell portion of LOCA is based on the methodology described in NEDO-21061. Subsection III. B. 3. a. 6 of NUREG-0487 requires that a comparison similar to thos e presented in reference model described in 14 be made if NEDM-10320.

the model Me used require is the differen model t from the prior to completion of review of the pool swell calculations.

~Reference (1) Letter "Response t'o NRC Request for Additional Information {Round 3 Questions," to J. F. Stolz (NRC-DPM) from L.

J Sobon (GE), dated June 30, 1978.

BF.SPONSE See response to Question 021.70 contained in the SSES FSAB.

OUESTION 3 Subsection .4. 2. 2. 2 of the DAR states that the chugging loads on submerged structures and imparted on the dovncomers vill be evaluations evaluated later. Provide the present status of these and the schedule, for your submission of the completed evaluation.

R ESPOUSE See response to Question 021.71 in the SSES PSAR.

QUEST ION 4 Statements are made in Subsections 4.2.3.2 and 4.2.3. 3 of the DAR that plant unique data of the Susquehanna SES intermediate break 10-19

accident {XBA) and small break accident {SBA) are estimated from curves for a typical Mark XX containment. Discuss the applicability of these analyses {e. g., pover level, initial conditions, dovncomer configuration, etc.) to Susquehanna SES.

Qg SQONSE See response to Question 021.72 contained in the SSES PSAR.

OUESTION 5 Provide the information previously requested in 020.44 regarding loads resultinq from pool svell waves following the pool swell process or seismic slosh. Discuss the analytical model and assumptions used to perform these analyses.

RESPONSE

See response to Question 021.73 contained in the SSES PSAR.

QUESTION 6 Provide a list and drawing to identify all piping, equipment instrumentation and structures in containment that may be subjected 'to pool dynamic loads. Xn addition, provide drawings to show the location of access qalleys in the vetwell, the vent vacuum breaker confiquration, wetvell grating, vent bracing confiquration, vent configuration in the pedestal region of vetwell and larqe horizontal structures in the pool svell zone.

R'ESPONSF.

See response to Question 021.74 contained in the SSES PSAR.

Q U X STION 7 Discuss the applicability of the generic supporting programs, tests and analyses to Susquehanna SES design {i.e., FSI concerns, dovncomer stiffners, downcomer diameter, etc.) .

RESPONSE

See response to Question 021.75 contained in the SSES PSAR.

QUESTION 8 Provide the time history of plant specific loads and assessment of responses of plant structures, piping, equipment and components to pool dynamic loads. Identify any significant plant modifications resultinq from pool dynamic loads considerations.

RF.SPONSF

• See response to Question 021.76 contained in the SSES FSAR.

REV. 67 4/82 '0-20

OUESTXON 9 Provide figures showing reactor pressure, quencher mass flux and suppression pool temperature versus time for the followinq events:

('I) a stuck-open SBV during power operation assuming reactor scram at 10 minutes after pool temperature reaches 110~P and all RHB systems operable; (2) same as event (1) above except that only one RHR train a vailable; (3) a stuck-open SRV during hot standby condition assuming 120~P pool temperature initially and only one RHR train available;

{4) the Automatic Depressurization System {ADS) activated followinq a small line break assuming an initial pool temperature of 120~F and only one BHR train available; and (5). the primary system is isolated and depressurizing at a rate.

of 100~P per hour with an initial pool temperture of 120~P and only one RHR train available.

Provide parameters such as service water temperature, BHR heat exchanger capability, and initial-pool mass for the analysis.

RESPONSE

,See response to Question 021.77 contained in the SSES PSAR.

QUESTION 10 With regard to the pool temperature limit, provide the following additional information:

(1) Definition of the ~~local" and "bulk" pool temperature and their application to the actual containment and to the scaled test facilities, if any: and (2) The data base that support any assumed difference between the local and the bulk temperatures.

RESPONSE

See response to Question 021.78 contained in the SSES PSAR.

QUESTION 11 Por the suppression pool temperature monitoring system, provide the followinq additional information:

(1) Type, number and location of temperature instrumentation that will be installed in the pool; and, Rev. 2. 5/80 10-21

(2) Discussion and justification of the sampling or averaging technique that will be applied to arrive at a definitive pool temperature.

See response to Question 021.79 contained in the SSES PSAR.

Rev. 2, 5/80 1 0-22

10.2.3 Questions Received During the Preparation of the Safety Evaluation Renort and Resoonse Thereto QU@STION 1 Qith regard to the SSES LOCA steam condensation load definition, provide the following additional information:

(1) Justification for the interchangeability of the GKM II-M temporal chug strength probability distribution with the spacial variation of chug strengths at SSES.

{2) Justification for not considerinq CO 6 SRV{ADS).

(3) Comparison of the CO measured at 4T-CO with the CO abserved at GKM XI-M.

PgSQONSE 1

{1) The SSES LOCA steam condensation load definition assumes that the chuqs occurring simultaneously at different vent pipes of SSES have different intensities and follow the same distribution of chug amplitudes in time as in the GKM II-M sinqle vent facility. This assumption forms the basis for tvo key elements of the LOCA load definition.

The first element assumes that the average of simultaneously occurrinq chuqs at different vents in SSES is eguivalent to the average of consecutive GKM II-M chugs. Thus, as ~

documented in Subsection 9.5. 3. 1. 2, the random amplitude chuqs at SSES vere replaced with the same chug at every vent which represents the averaqe of consecutive GKM II-M chugs or

<<mean value<<chug.

The second element assumes that the chug amplitude or strenqth at the individual SSES vents are random variables which have the same probability distribution as the distribution of chug amplitudes at GKM II-M. The <GKH,II-M probability distribution vas then applied statistical'ly to an analytical model of the SSES suppression pool to calculate the symmetric and asymmetric amplitude factors. These factors vere then applied to the selected mean value chugs to achieve the desired exceedance probability prior to transportation to SSES for containment analysis {see Subsections 9.5.3.4.1 and 9.5.3.4.2) .

These tvo elements infer that the multi-vent facility is h

composed of many <<single cells> vhose chug strengths vary stochastically and independently of each other. The random nature of chuqqinq is explained qualitatively by looking at the actual hubble collapsinq mechanism. The most plausible mechanism for bubble collapse at the individual vents appears to be the convection in the pool. This means that bubble collapses at indivdual vents are triggered by the local turbulent convection at each vent. Thus due to the REV. 6, 4/82 1 0-23

stochastic nature of turbulence, the time at which rapid condensation and hence bubble collapse is triggered varies from vent to vent. This implies that the'ize of the bubble formed before collapse starts, sill also vary from vent to vent. Therefore, the chug strength will vary from vent to vent. Since, the GKH XI-H tests vere designed to be prototypical of SSES (i.e., same initial pool temperature, same steam flov, etc.), this random variation is expected to be similar for both the GKH II-N single vent facility and the SSES plant.

Additional qualitative data verifying the random nature of chugqinq is provided by numerous multi-vent test programs.

Specifically, the KMU multi-vent concrete cell tests in Karlstein. Creare subscale multi-vent tests and JAEBI full scale multi-vent tests provide multi-vent data of the c hugginq phenomena.

The Karlstein facility investigated the chuqginq phenomena for 2~ 6, and 10 vents at subscale. Each vent in the concrete cell was instrumented with a pressure transducer in such. a way that it was indicative of the chug strength for its respective vent. Piqures 10-4, 10-5, and 10-6 illustrate these vent transducers and the remaining transducers for the 10, 6, and 2 vent facilities, respectively'.

10-7 and 10-8 shov typical pressure time histories

'igures for the pressure transducers mounted near the vent pipes for the six vent configuration. These pressure transducers vere all exposed to a steam environment and, clearly indicate that

,the chug strengths differ by up to a factor of 10.

Zn addition, Piqure 10-9 shovs that the distribution of relative frequencies of the measured vali pressures becomes narrover as the number of vent pipes increases from 2 to 6 to

10. Again, the variation in chug strengths results in a lower global "pressure, amplitude vith increasing number of vents.

This variation in chuq strengths vas also observed in the Crea re'ubscale multi-vent test prog ram. This ob servation vas obtained by examininq the pool sall pressures measured at the three different circumferential locations at the vent exit. All tes't geometries had three transducers located 120~

apart circumferentially at the vent exit elevation. In the multi-vent geometries, each of these pressure transducers vas located close to a particular vent. Therefore, the amplitude of the POP measured at each circumferential location 'reflects to a larqe extent the chuq strength at the vent closest to (since pressure amplitude varies inversely vith the distance it betveen the vent and vali pressure measurement location) .

Por example, only if the chuq strengths at all vents vere identical, vould the peak over-pressure (POP) measured at each of these three circumferential locations be identical.

10-24

Figure 10-10 shows the pool wall pressures at the three circumferential vent exit locations in the 1/6 scale 3 vent geometry. The steam mass flux was 8 ibm/sec ft~ and as determined from the vent static pressures over 80% of the chuqs shown had all three vents participating. This figure shows that the POP's at the three locations are different for individual chugs. Therefore, it can be concluded that the chug. strength varies from vent to vent.

Similar data from the 1/10 scale, 19 vent geometry at a steam mass flux of 8 ibm/sec ft~ are shown in Figure 10-11.

Again, from vent static pressure data for vents closest to each circumferential wall pressure measurement .location, it was determined that all three vents participated in the chugs shown. ,The POP~s at the three different circumferential locations are seen as being different for individual chugs.

Note that the variation of chug strength from vent to vent i' expected to be stochastic to a large extent.. Therefore, it is expected, that for some chuqs, the chug strength at the three vents would be similar.

Additional proof that the chug strengths in a multi-vent facility behave stochastically is given by the JAERI multi-vent test data. There are several pool wall pressure transducers that are located near the exits of different, vents in the JAERI facility. Specif ically, transudcers MWPP-202, 302, 602, and 702 are located at the vent exit elevation next to vents 2, 3, 4, and 7, respectively (see,Pigure 10-.12 and 10-13) . The pressure amplitudes measured by these transducers reflect the chug strengths at vents closest to them.

The variation of chuq strengths at individual vents is shown in Piqure 10-14. The pool wall pressures at the vent exit elevation for a chug occur at 62.5 seconds in JAERI test 0002. In this chug event, a high amplitude chuq occurred at vent 7 as indicated by the large pressure spike at MMPP702.

The other vents had relatively smaller chugs. Keep in mind that the variation of chug strengths from vent to vent is stochastic in nature and that not all pool chugs will exhibit the large variation seen in Piqure 10-14., Nonetheless, varying deqrees of variation in chug strengths from vent to vent were found in all the chuqs from Tests. 0002, 2101, and 3102 for which expanded time traces are available.

So far, we have stated that chuqging is stochastic in nature, and as such the chug strenqths are expected to vary, even though the same thermodynamic conditions exist at each vent (i.e., steam air content, mass flux, bulk pool termperature, etc.). As presented above, this phenomena has been observed in numerous multi-vent test facilities. However, we have not quantitatively verified our assumption of the interchangeability of the temporal chuq strength variations at GKH II-M with the spacially varying chuq strengths at SSPS. Again, the Creare subscale multi-vent test data and REV. 6, 4/82 1 0-25

JAERI test data provide information verifying the conservatism of this assumption. Each will be presented b elo w.

As previously stated, one element of our LOCA load definition replaces the random "amplitude chugs at SSES with the same chug at every vent, which is representative of the mean value data at GKM II-M. The Creare test data coupled with the accepted acoustic methodoloqy provides verification of this assumption. Creare has acoustically modeled the 1/10-scale single and multi-vent geometries and they have derived a source which represents the mean value chug in the 1/10-scale single vent geometry.

They then placed this mean value chug source at each vent location of their acoustic model for the 1/10-scale 3, 7, and 19 vent geometries. For each of the three multi-vent geometries, the pressure time histroy at the pool bottom elevation (same as the transducer location at this elevation in the test geometries) was computed for 20 chug events.

Bach chug event involved selectinq start times for .individual vents randomly within a 20 msec time window. The multi-vent multiplier was then computed based .on the mean POP at the.

pool bottom elevation for the 20 computed chugs. The predicted multi-vent multipliers compared quite favorably with the measured values. Subsection A 5. 2.2 of Reference 66 gives a detailed description of the, analysis and results.

Thus, for subscale multi-vent geometries, the first element of our LOCA load definition is verified.

Final guantitative justification for our key assumption is provided by comparinq the available JAERI full-scale multi-vent data with the GKN II-M single vent data.

There are two sets of JAERI data available that can be used to infer chuq strengths at individual vents in a given multi-vent chug event. The first set is the pool wall pressure data from the pool wall transducers located at the vent exit elevation. In the JAERI test geometry, there were four pool wall pressure transducers-MRPF 202 '02, 602, and 702-located such that each of these transducers is very near the exits of four individual vents. Therefore, the pressure data from a given transducer reflects the chug strength at the vent closest to that transducer.

As previously stated, the data from these wall pressure transducers were used to qualitatively show that the chug strengths vary significantly from vent to vent in a JAERI multi-vent chug event. Unfortunately, since a pool transducer "sees" pressures due to chugs at all vents to varyinq extents, the data from such transducers are not suitable for quantitative evaluation of vent to vent chug stre'nqth variations.

REV. 6, 4/82 10-26

The other set of JAERI data that provides a measure of chug ~

strenqths at the individual vents are the vent static pressure measurements. Pive of the seven vents in the JAERI test facility are instrumented with vent exit static pressure transducers.

The vent static pressure is a direct measure of the >>vent component>> of the chuq-induced pool'wall pressure. Purther, due to desynchronization in a multi-vent geometry, the "vent component" is the dominant component of the chug induced pool pressures observed in multi-vent chugqing. Therefore, the spatial (vent to vent) variation of the vent static pressures in the JAERI multi-vent geometry should provide a reliable estimate of the vent to vent chug strength variation in a multi-vent qeometry.

Individual vent" exit static pressures of 1.125 sec periods are available for 38 chuq events from six JAERI tests, eight chuqs from Test 0002, seven chuqs from Test 0003, six chugs from Test 0004, five chugs from Test 1101, five chugs from Test 1201, and seven chuqs from Test 2101. These chugs were selected from periods 'of high amplitude chuqging in each test. Therefore, this data base covers the worst chugging regions observed in these" JAERI tests.

The indivdual vent exit static pressures for a given pool chug event were processed in the following manner. .First, the rms pressure Pi was computed for each vent static pressure trace. Next, the average rms pressure P was computed. for example, if vent static pressures were available for all the five instrumented vents, the average rms vent static pressure for that chug is:

Pg + P2 + P3 + P4 + P5 Since we are interested in the relative variation in chug strengths between individual vents, the individual rms vent static pressures were normalized by the average rms pressure P ~

'he normalized indivdual rms vent static pressure Pi for the 38 chuqs analyzed are qiven in Table 10-1. Also shown are the values of the normalized variances for the individual vent rms pressures for individual chug events. Note that due to instrumentation malfunctions, for all except one JAERI test, vent exit static pressure data are not available for all five instrumented vents.

Due to small number of vents (at most five) for which vent static pressure data are available, it is difficult to draw meaningful statistical inferences for vent to vent chug strenqth variations from any one individual pool chug event.

REV. 6, 4/82 10-27

Therefore ~ it is necessary to hake an assumption that allows the use of the data from all 38 chug events such that meaningful statistical inferences can be drawn. This assumption is that the normalized statistical distribution of chuq strengths from vent to vent is independent of blowdown conditions. That is, the normalized vent to vent chug strength for all 38 chug events are samples selected from the same statistical population. Note that this is precisely the same assumption made in analyzing the temporal statistical properties of 'the GKM XX-M single vent data (see Subsection 9-5.3.2 1)-

The GKN II-M data that provides a direct measure of the vent component of the chug strength are the pool wall pressure data band pass filtered between 0.5-13 Hz. In this frequency range, the pool wall pressures measured are due to the vent pressure oscillations produced by the chug (see Subsection 9.4. 2 1 2)

As described in Subsection 9.5.3.2.1, the pressure amplitudes of individual chugs were normalized by the sliding mean value over a given time interval. In this way, a normalized data base ref lectinq the temporal variations of chug strengths was obtained for all the GKM II-M tests. Note that again implicit in this procedure is the assumption statistics of the variation of the normalized that the chug strengths is independent of system conditions. As previously mentioned, this assumption was al'so used for combining the JAERI data for 38 pool chug events into a single statistical data base.

The histograms of the normalized chug strengths for the various GKM II-N tests axe given in Fiqures 9-181, 9-182 ~ and 9-183 At this point, we now have a normalized vent to vent chug strength variation data base from the JAERX multi-.vent tests and a corresponding normalized chug to chuq strength variation data base from the GKN II-M sinqle vent tests.

Table 10-2 shows the vaxiance for the JAERI and GKN II-M data bases. The variance for the JAERI data base is the average value of the individual variances shown in Table 10-1 for each of the 38 chug events. The variance of the GKM IX-N data was calculated for the 0.5-13 Hz band passed data plotted in Figures 9-181, 9-182, and 9-183. It is seen that the average variance from the JAERI tests is virtually identical to'he variance from the GKM II-N Pull NSL tests+

and is somewhat greater than the variances from the 1/3 and 1/6 NSL GKM IX-N tests. This implies that the vaxiation of vent to vent chug stxenqths in the JAERI multi-vent tests is equal to or qreater than the chug to chug strength variation observed in the GKN II-N single vent tests.

Figures 10-15 through 10-17 show the comparison of the probability density histograms of the JAERI data and the low

• The full MSL break chug strength statistics were used to develop the SSES probabilistic amplitude factors.

REV. 6$ 4/82 10-28

band passed GKM II-N Full MSL, 1/3 MSL and I/6 MSL data, respectively. Again, the JAEBI and GKN XI-M data histograms are quite similar.

From the above comparisons it can be again concluded that the assumption that the vent to vent variation in chug strenqths in a single vent geometry is equivalent to the vent to vent chuq strength variation in a multi-vent geometry, used in developing the SSES chugging load definition from the GKN II-N sinqle vent test'data is quite reasonable.

Additional verif ication of the conservatism of the SSES LOCA load definition is provided by comparing the wall loads at JAERI calculated with the SSES LOCA load definition with the available JAERI wall load data (see Subsection 9.5.3.5.1).

Figures 9-268 and 9-269 show that the SSES LOCA load definition bounds the available JAERI data by a substantial margin. Please note that the wall loads calculated by the "

SSES LOCA load definition do not include the symmetric amplitude factor and thus represent t>mean value<> chugs.

(2) The Nark II Owners have specified two different CO loads for containment analysis. The first CO load (CO 1) corresponds to the CO occurring at the beginning of a postulated LOCA and the second CO load (CO 2)'orresponds to the reduced 'CO load occurrinq later in the blowdown. For containment analysis,

~ ~

the Owners combine the reduced CO 2 load with loads due to SRV (ADS), on the basis that ADS occurs later in a LOCA

~ gustifyi.nq a reduced 'CO load for the combination CO 6 SRV

{ADS)

However, SSES combines the so-called LOCA loads with SRV

.. (ADS) for containment analysis. The LOCA load comprises the envelop of the responses due to both chugging and CO. Thus, the SSES load combination LOCA 8 SRV (ADS) considers both CO and chuqqinq and is more conservative than the Owneris combination of a reduced CO load (CO 2) with SRV (ADS).

(3) The SSES LOCA laod definition selected one CO pressure time history (PTH No. 14) from GKM II-M as representative and boundinq of the CO at GKN II-M (see Figure 9-177a 6 b).

subsequently, this CO PTH was sourced and applied in-phase to the IWEGS/MARS acoustic model for containment analysis-Figure 9-264 represents the enveloping PSD of PTH No. 14.

Figure 2-1 of Reference 70 presents the envelop for PSD values observed for CO in the 4T-CO tests. These two figures indicate that the PSD of PTH No. 14 from GKM II-M compares favorably with the envelopinq PSD of the CO in 4T-CO.

OUESTION 2 The dominant frequency for the Karlstein T-Quencher Test 21.2 appears to be 8.0 Hz instead of'he 6.8 Hz reported, in Table 8-10 of the DAB. Usinq the multipliers from Figure 8-174 and this 8.0 REV. 6, 4/82 1 0-29

Hz frequency, we get a transposed frequency of 10.6 Hz. This value falls outside of the specified frequency range. A Fourier analysis indicates an exceedance of approximately 70% at this, 10.6 Hz frequency., Please provide justification for the existing load specification frequency range.

RESPONSE 2 As can be seen in Figure 8-188, Test 21.2 does not show a clearly predominant frequency. Me have interpreted 6.5 Hz as the predominant frequency because of the maximum peak occurring in the PSD at that frequency; however, a second peak, only slightly lower than the 6.5 Hz peak, can be seen in that PSD at approximately 8.0 Hz.

To investigate further the significant of Test 21.2 to the acceptability of the Susquehanna T-Quencher load specification, KWU performed a pressure response spectra comparison of the load specification and Test 21.2.

The method of <<weighted traces<<presented to the NRC in the June 13, 1980 Lead Plant Neeting and documented in the KMU Report R 141/141/79 is used for this comparison. Figure 10-18 shows that the Susquehanna load specification bounds the measured pressure time history of Karlstein Test 21.2 representing the all valve case.

Assuminq a maximum predominant frequency in Test 21.2 of 8 Hz and transferrinq the measured data of Test 21.2 to the all-valve and sinqle-valve load case we qet the comparison shown in Figure 10-

19. The pressure response spectra of the Susquehanna load specifications is slightly exceeded by the pressure spectra from Test 21.2 in the frequency ranqe between 10 Hz and 11 Hz. This sliqht exceedance is only related to the single-valve load case and is considered insignificant to the total load specification and in relation to the total data base from Karlstein.

Xn addition, the term>>dominant frequency<< is highly subjective and sensitive to the method chosen for determininq the dominant frequency. Oriqinially, KNU determined the dominant frequency range for the three SSES desiqn traces (KKB Traces 435, 76 and

82) to be 6.5 to 8.0 Hz (see SSES DAR, page 8P-101). This frequency range was based on a PSD analysis of the three traces.

However, for these non-stationary SRV traces, the PSD analysis is sensitive to the time segment chosen for analysis. Using a particular time duration may give one dominant freguency while another may 'give a sliqhtly different dominant frequency.

Subsequently, Bechtel has taken the design traces and performed their own analysis to determine the dominant frequency. They calculated a dominant frequency range of 6~45 to 8.6 9 Hz for the three traces. This frequency ranqe was based on the inverse of the peak-to-peak oscillation time period for the first two peaks.

This was done for both negative and positive peak-to-peak periods.

REV. 6, 4/82 1 0-30

Furthermore, Sarqent 6 Lundy have. determined the dominant frequency range of the three traces to be 6.8 to 8.9 Hz. As can be'een, the dominant frequency varies according to who performs the analysis and the methodology selected.

For containment analysis, the KMU methodology requires that time scale multipliers be applied to the three 'design traces. They from 0.9 (time contraction or frequency expansion) to 1.8 'anqe (time expansion or frequency contraction). When these multipliers are applied to the three design traces, specified frequency, ranges of 3.3 to 8.9 Hz, 3.6 to 9.7 Hz and 3.8 to 9.9 Hz are obtained by usinq the above dominant frequency ranges from the oriqinal traces. Thus, the specified frequency range varies dependinq on the interpretation of the "dominant frequency.

However, regardless of the interpreted dominant frequency range",

the same three traces and time expansion and contration factors are used for containment analysis. Thus, ones opinion of what the dominant frequency range is for the three traces is not as important as the time factors chosen for actually applying the traces to the containment boundary.

With this in mind, Figures 10-20 thru 10-41 illustrate the response spectra qenerated by KWU Trace 476 for SSES. The trace was f reguency expanded and con tract ed by 110% a nd" 55%,

respectively, to give a specified frequency ra'nges of 3.3 to 8.9.

Hz, 3. 6 to 9.7 Hz or 3. 8 to 9. 9 Hz, again, depending on the interpretation of the "dominant frequency".

Figures 10-42 thru 10-63 show the response spectra generated by KWU Trace 076 for the Limerick Generating Station (LGS) . The LGS structural model is essentially identical to the SSES'odel.

However, these spectra reflect the use of frequericy expansion and contraction factors of 125% and 55%, respectively. This gives specified frequency ranges of 3.3 to 10 Hz, 3.6 to 10.9 Hz or 3.8 to 11 Hz. Thus, dep'endinq on the dominant frequency, these spectra reflect the use of the NBC's upper bound dominant frequency of 11 Hz, as required by Supplement No. 1 to NUREG-04 87.

A node by 'node comparison of the two 'spectra shows that the expanded spectral, input used for LGS has negligible effect on the total response contributed by all modes. Thus, this supports the conclusion that an extention of the upper frequency multiplier would have no siqnificant impact on the SSES response spectra an aly sis.

OUE STION 3 The Karlstein tests run with depressed water legs to simulate the ADS load case utilized the lonqest discharge line length for SSES. Is this line length prototypical of the SSES ADS line lengths? If not, what is the maqnitude of the difference between the SSES ADS line lengths I

and the test line length? Xf not REV. 6$ 4/82 1 0-31,.

,prototypical is the data from the ADS tests acceptable for transportation to SSES vith regards to frequency content?

RESPONSE 3 Tests 10.3, 11. ' 12. 1, and 13. 1 are considered representative 1

for the ADS actuation load case. These tests vere all performed vith the long discharge line. No tests vith a short discharge line and a depressed initial vater level (representing ADS conditions) vere performed. These long line tests represent a bounding condition, in that the lonqest discharge line with depressed initial vater level contains the largest possible initial air mass and will therefore produce the lowest possible il pressure osc lati on freq uenc y.'o check whether the'requencies expected from short line ADS actuation fall within our specified frequency range we will transpose the test results from Test 11. 1 to short line conditions.

Table 8B on page 8P-105 of the Susquehanna DAR shows the average frequencies measured during the Karlstein tests. A portion of that table is shown below:

Measured Frequencies (Hz)

Long Clean Conditions ~3 5) +-4 Line Real Conditions Short Clean Conditigns Line Real Conditions 6 5

~Tests vith lov amplitude This data indicates a ratio of approximately 1.3 exists between the frequencies measured in long line tests and short line tests.

Subsection 8.5.3.3.4.6 of the Susquehanna DAR provides the comparison of the '-Quencher ADS load specification vith the Karlstein test results..- When the measured'requency for Test 11.1 was adjusted to account for .back pressure and vater surface area effects the. measured 3 Hz frequency vas raised to 5.7 Hz.

. To check the short line ADS load case ve vill adjust this 5.7 Hz by the 1.3 ratio obtained above. This produces a predominant frequency for the ADS short line conditions of V = 5.7 r 1.3 = 7.4 Hz This frequency lies within the specified frequency range.

REV. 6, 4/82 10-32

Was the quencher bottom support used at Karlstein prototypical of the supports at Susquehanna SES'?

ggSPONQQ 4 The bottom support used in Karlstein is protopical but not identical of those used at Susquehanna. The T-Quencher installed in the Karltsein test tank had the same distance between the bottom of the support and the quencher mid-plane as those quenchers installed at Susquehanna. Therefore, the thermo-hydraulic loading on the quencher supports are the same for the Karlstein test tank and Susquehanna. From a structural point of view, the bottom support used at Karlstein is not identical to those used at Susquehanna in that the supports in the plant are stiffer.

QUESTION 5 In three instances, the bendinq moment in the quencher arm recorded at Karlstein exceeds the specified bending moment. Xs the specified bendinq moment in the quencher arm conservative?

Whyo'gspopsg 5

As shown in Figure 8-153 the measured bending moments transposed to the weld of the quen'cher arm exceed the specified moment in 3 out of a total of 99 cases durinq vent, cleaning. The total load specification for the quencher arm is made up of three components:

a) internal pressure b), bendinq moment c) tempe rature gradient The followinq table lists the specified and maximum measured values for each of the load components.

Maximum Conclgtjog Measured Value Steady State.

Pressure 22 bars 13 bars Internal Temperature 219o C 191 6o C Bending Moment 65 kNm 85 kNm REV. 6, 4/82 10-33

As can be seen, the specified values exceed the measured maximum values except for the referenced bending moments noted above.

As a result of this exceedance, a stress analysis, identical to the one performed for the specified values, vas completed using the above maximum measured values. This analysis shows that the total stress due to the specified loads bounds the total stress due to the maximum measured loads. In addition, a fatigue evaluation of the arm veld was performed using the maximum measured data. The results indicate the veld has a usage factor less than unity, and thus is acceptable.

Explain vhy a single failure vill not disable both the BHR shutdown cooling function and one BHR loop in the suppression pool coolinq mode.

RESPONSE 6 A single failure can indeed disable the RHR shutdovn cooling function and one BHR loop in the suppression pool cooling mode under the following assumptions. Both units are operating at full power when a complete long-term loss of offsite pover (LOOP) occurs. This leads to main steam line isolation and reactor scram. Pollowing the T.OOP all four {4) diesel generators should start to supply power to the ESS busses, hovever, it is assumed that the diesel qenerator OG501C does not start (single fai1ure).

OG501C supplies pover to the ESS busses 1A203 and 2A203+, to the RHR pumps 1C and 2C+, and to the RHR service water pump 1A. Loss of OG501C means that the inboard shutdown cooling isolation valves on both units, 1F009 and 2F009~, loose power to their operators, thus disabling the RHR shutdown coolinq mode. Since these valves are located inside the primary containment, conservativey assumed that they will not be manually reopened.

it is Only the <<B<< loop and the corresponding RHRSM loop of 'the RHR system (in both units) vould be readily available for suppression pool coolinq, using e.g., RHR pumps 18 and 2D+. The >>>>A<< loop of one unit could be made available by manually operatinq four (4) valves {close P048A, open P024A, HV-1210A and HV-1215A) and using RHRSH pump 2A+ and either RHR pump 1A or 2A+. Hovever, a simultaneous operation of RHR pumps 1A and 2A+ is prohibited by electrical interlocks. Thus one of the units would have only one RHR loop available in the suppression pool cooling mode vithout the possibility to switch to shutdown cooling.

This case has ndt been considered in the transients submitted as part of Appendix I of the DAR and may be more limiting. However, a similar but more conservative case vas analyzed as part of a sensitivity study and resulted in a maximum pool temperature of 203oF. The assumptions for this case are indentical to case 2.a (Appendix I, DAR) except that shutdown cooling is not initiated.

Por this case, the curves for reactor pressure vs. time and suppression pool temperature vs. time are found in Piqures 10-64 and 10-65, respectively.

• Indicates Unit I2 component.

REV. 6, 4/82 10-34

As ment'ioned above, this case is similar, but more conservative than the case un'der consideration. The major difference is that reactor water make-up would not be from the feedvater/condensate system but from HPCI (at reactor pressures above approximately 300'psia) and core spray (at reactor pressures below approximately 300 psia), which both take suction from the condensate storage tank,and/or the suppression pool. Thus, water much colder than feedwater would be used for make-up.

This contributes to the reactor depressurixation and leads to less steam being dumped into the suppression pool. The peak suppression pool temperature for this case will therefore be lower than that shown in Figures 10-65.

To confirm a temperature of less than 2030P we have initiated an additional analysis case, whose results are contained in Appendix I {Pigures I-14 and .I-15) .

g 0 EST ION 7 How will PPSL use the LaSalle in-plant test data to establish the local -to bulk hT for Susquehanna S ES'P RESPONSE 7 The following table gives a comparison of suppression pool geometries for LaSalle and Susquehanna SZS:

LaSalle Suscruehanna

'8 k

Su ppr ess ion Pool,I.D. 86 ~ ~8 lt 88~

Pedestal O.D. 30 ~

29 ~ 9fl Suppression Pool Volume 142, 160 ft> 126 r 980 ft>

(Normal Rater Level)

No. of Quenc hers 16 Pool Volume/Quencher 7898 ft~ 7936 ft>

Quencher Submerqence (Normal Mater Level) 21.5 ft 19.5 ft Heiqht of Quencher Center-Line Above Base Hat 5ft 3.5 ft Based on the similarity between Susquehanna and LaSalle the local to bulk ~T established from LaSalle inplant tests is also applicable to Susquehanna. Xn addition, PPCL is continuing to fund the development of computer codes (like Bechtel's KPIX} for the prediction of SRV discharge induced suppression pool mixing processes. The calculated temperature distributions will be compared to existinq {Caorso) and future (LaSalle or Zimmer) in-plant test data.

REV. 6>> 4/82 10-35

Pollovinq satisfactory qualification of the computer codes they can then be used to establish local to bulk temperature differences without test.

QUESTION 8 Shat are the reactor pressures that correspond to guencher steam mass fluxes of 42 ibm/ft~s and 94 ibm/ft~s?

RFSPONSZ 8 The reactor pressures are 163 psia and 369 psia resp ctively.

REV. 6% 4/82 10-36

This figure has been deleted.

REV. 6, 4/82 SUSQUEHANNA STEAM ELECTRIC STATION UNITS 1 AND 2 DESIGN ASSESSMENT REPORT DOWNCOMER BRACING SYSTEM FIGURE'0->

This figure has been deleted.

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~1,00 K

0 K

IAJ IAI 0. 15 O

p 0.50 V)

'- ' cvg Cfy 0. 25 2 mr cnrn m

( CO D

C KO M -I Cyg P m a <ZZO mme 0 m

H z Vl Z+ Z g Z I

OO~Z Z ~ 5*-'. 0. 00 r ICZ ITI 0.1 4 6 8 gp 2 4 6 8 lp.o 2 4 .6 8 gpp FREQUENCY-CPS

~cy2 ~ ~Pm Liner)ck Generat)on Station, Accelerat)on Spectra for ~DELI Z ~lvg CTI<

o Ch m rgl Q

Load Case: SRV ASYMMETRIC - TRACE 76 0 NO Z CAZ 'I U

8- o cn Node: ~21 Direction: vERT 'lev: 236'-2" Ang1e:

rgl O

0.005,0.01,0.02,0.05 By: lr .Date: ~5-SDCheck: ldll Date:

0'Damping:

~5b Bo z

.1.26 till. 00 X

O w0.75 O

O I-f

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m OB m.m O- R z+D.Z

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~ ~ Q) Z I

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g7 Ill

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Load Case:

Node: ~BL SR -

FREQUENCY-CPS Direct)on: HORIZ 0.005,0.01,0.02,0.05 By: 2c.

'l L)mer)ck Generat)on Station, Accelerat)on Spectra ASYMMETRIC TRACE I6 for Date: s~+2 Check:

WETHELL

~

ev: 236 '-2" Angle:

Date: S~l/So 90'aeptng:

C'

• \

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<<IC7 FREQUENCY-CPS m<

O Q L)mertck Generat)on Stat)on, Accelerat)on Spectra for WETNEIL m

I. m -

0) O Load Case: sRv AsYMMETRIc TRAcE 76 0R MO Elev: 236'-2" Angle:

CI CO CIA Node: 295 0)rect)on: VERT 0

Z 0.005,0.01,0.02,05,05 By: gc Date: 5-5-Bo check: tI~ Date: ~5/5 Ba 90'ampteg:

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~ ~C/3 Directton: Elev: 264'-6" Angle: 0 ~

0 NO Z

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~ CO fll Z) ~ ~ gag CO O<m 0 VZ Z

Cyl 0

o Load Case:

Node: 331 Direction: VERT 'lev:

SRV ASYMMETRXC TRACE 76 0.005,0.0l 0.02,0.05 By: K.

264'-6" Date: ~5 oCheck:

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6) K CO m o Load Case: SRV ASYMMETRIC TRACE 76 Z~

0 Z I Node: 335 Dlrecttnn: HORXZ Elevl 264'-6" Angle:

0.005,0.01,0.02,0.05 By: 1'c Date. 5'-5-BgCheck: WL Date: m/C(ga 90'amptng:

0 R

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L)mer)ck Generat)on Station, Accelerat)on Spectra for DRYHELL o rn m Load Case: SRV - ASYMMETRIC TRACE 76 CO Q 0 0

Z NOC7)Z th Node: 995 Direction: vERT Elev: 26a'-6" Angle:

Cyg fll 0 I

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m<m m TRACE 76 O CD

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~Qm m >up Lamer)ck Generation Stat)on, Acceleration Spectra for DRYHELL o m

~

m

-O Load Case: RAE76 20 VlO 2.

m 0 a

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ttl n Load Case: - SYMMETRIC - TRACE 76

~%0 cn 312'-7" 0

Z %CD Z I Node: all Direction: VERT ElevvI: Angle:

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~Q O H pde: 415 0)rection: HORIZ . El ev: '312 '-7" Angle: 90

~Z CflCyg 'I Dampteg: 0.005,0.01,0.02,0.05 By: g. Date: 5W-&0 Check: '4(e Date: 5 5/Bo 0

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m g7 L imeri ck Generation Station, Acce 1 erati on Spectra for CO O Load Case:

~CO CAPl g

Co I

Node: ~~ Di rection:

peeping: p.ppg,p.pI,p.pg,p.pg v R

~c Elev: ~312'- " AngIe:

pate: e-5-Cp check: ~l~ate: ~ </A>

20 py:

HOURS CI 2 CD Q 0

100 CD CL 6

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I m X OI C: O 0)R.Pl

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A Vl ~ Pl O TIl1E AFTER SCRAM SECONDS C>3 0 I"-, I Pl C9 C/l m C) yy 0 f ~ . C: m O

I CO CD c C)

I Pl O C) Z

TABLE 10-1 JAERI DATA NORMALIZED RMS VENT CHUG STATIC PRESSURE JAERI STAR VAR.

VENT VENT VENT VENT VENT TEST IME Q2 sec 0002 8.65 .88 . 13'38 0.99 . 015 2.37 ..87 0.75 .114 6.35 '. 17 ~ 03 0.81 .033 2.65 .99 .29 0.72 .083 4.65 .72 .29 0.98 .080 6.75 .85 .06 1.09 .018 8.80 .85 .09 1.06 .016 0.25 .90 .03 1.07 .007 0003 2.27 ,10 .Ol 0.89 .011 4.10 .83 .07 1.10 .021 5.98 .61 .36 1.04 .141 7.85 .16 .13 0.71 .064 9.90 .64 .05 1.31 .144 1.45 .54 .50 0.97 .232 6.85 .12 .Ol 0.83 .014 0004 9.50 .95 .44 0.61 . 173 0.65 .86 .34 0.79 .089 3.00 .47 .77 '.76

.461 5.20 .41 .35 1.23 .264 9.00 .44 .75 0.81 .453 3.05 .68 .29 1.03 .094 1101 0.40 0.81 .86 1.36 0.97 .061 2.02 0.91 .78 1.21 1.10 .036 4.20 1.3 .68 1.01 0.96 .075 6.25 0.77 .49 . 24. 1.50 .207 8.80 0.89 .54 .42 1.14 .140 1201 7.60 0. 86 1 00 .15 1.00 .013 9.40 1; ll 1.35 1.'20 1.08

.72 0.82 .081 0.93 .23 0.75 . 042 3.00 1.31 0.65 .15 0.90 i.084 4.90 1.22 0.60 .27 0,91 .097 101 5. 80 1.14 0.84 0. 84 .90 1.28 .040 9.75 1.13 1.17 0. 89 .99 0.82 .023 2.00 1.07 0.67 0.98 .89 1.40 .071 3.85 0.89 1.07 1.23 .22 0.60 .072 6.10 2.08 0.56 0.29 .20 0.88 .478 8.15 0.87 0.82 1.10 .30 0.90 .039

00. 1 0.96 0.71 0-. 93 .18 1.21 .041

TABLE 10-2 JAERI/GKMIIM COMPARISON DATA NORMALIZED BASE MEAN VARIANCE JAERI 0.108 DATA GKMIIM MSL DATA 0.107 (0.5-13 Hz)

GKMIIM 93 MSL DATA 0.083 (0.5-13 HF)

GKMIIM I/6 MSL DATA 0.064 (0.5-13Hz)

11 0 REFERENCES

1. Dr. M. Becker and Dr. E. Koch, <<KKB-Vent Clearing with the Perforated-Pipe Quencher" (translated by Ad-Ex, Watertown, Massach usetts), KWU/E3-2796, Kr a ft wer k Union, October 1973.

2. Dr. M. Becker and Dr. E. Koch, "Construction and Design of the Relief System with Perforated-Pipe Quencher" (translated by Ad-Hx), F3/E2-2703, Kraf twerk Union, July 1973

3. Dr. M. Becke , <<Results oif the Non-Nuclea" Hot Tests with the Relief System in the Brunsbuttel Nuclear Power Plant" (translated by Ad-Hx), KWU/R113-3267, Kraftwerk Union, December 1974.

4. Dr. H. Weisshaupl, <<Formation and Oscillations of a Spher:ical Gas Bubble Under Water" (translated by Ad-Ex), AEG-Telefunken Report. No. 2241, Kraftwerk Union, December 1972

5. Dr. H. Weisshaupl and Schall, <<Calcrrlation Model to Clarify the Pressure Oscillations in the Suppression Chamber After Vent Clearing<<(tr:anslated by Ad-Hx), AEG-Telefunken Report. No. 2208, Kraftwerk Union, March 1972.

6. Dr. M. Becker, Feist and M. Burro, "Analysis of the Load

.'leasured on the Relief System During the Non-Nuclear liot Test in KKB<<(transla ted hy Ad-Ex), R 113/R 213/R 314/R 521-3346, Kraf twerk Union, April 1975.

7. Letter, J. W. Millard to M. J. Lidl, <<Susquehanrra 1 6 2:

'lass and Enerqy Release for Suppression Pool Tempera ture Arralysis during Safety Relief Valve and LOCA Transients,<<GB-77-65, March 14, 1977.

8. R. J. Ernst and .'I. G. Ward, "Mack ZI Pressure Suppres iof1 Containment Systems: An Analytical Model of the Pool Swell Phenomenon,<<NZDE-21544P, General Electric Co.,

December 1976.

9. Letter., F. C. Rally to 'lark XI- Technical Steerinq Committee Members, <<Pool Swell .'iodol Test Cases " MKIl-301-E, Auqus t 22, 1977.

10 "Dynamic Forcing Functions In forma ion report (DFFi<),<<Rov.

2, NHDO-21051, Gerreral Electric Co. and Sargerrr. a<<d Lundy Engineers, September. 1976.

10a. "Dyrramic For" ing f rrnction Information Report (DFFR),<<Rev.

3, NEDO-2106l, General Electric Co. and Sargent and Lundy Engineers, Ju>>e, 1978.

Rev. 2, 5/80 11- 1

11. T. Y. Fukushima, et al., "Test Results Employed by GE for BthtR Containment and Vertical Vent Loads, <<NFDE-21078-P, Table 3-4, General Electric Co., October 1975.

12 F. J. Moody, Analytical Model for Liquid Jet Properties for Predictinq Forces on Riqid Submerged Structures, NEDL-21472, General Flectric Co., September 1977.

13 .R. J. Ernst, et al., Mark II Pressure Suppression Containment Loads on Submerqed Structures An Application Systems:

Memorandum, NEDE-21730, General Electric Co., September 1977.

14. 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, General Electric Co., (to be published) .

15. Mark II Phase I, 4T Tests Applications Memorandum, Letter and Report to H. R. Butler (NRC) from J. F. Quirk (GE),

June 14, 1976.

16. M. J. Bilanin, et. al., Mark IZ Lead Plant Topical Report:

Pool Boundary and Main Vent Chugging Loads Justification, NEDE-23617P, July 1977.

17. Marmeatlas (Heat Transfer Data), VDZ (Society of German Engineers), Dusseldor f, 1974.

18. T. E. Johnson, et al., "Containment- Building Liner Plate Design Report, <<DC-TOP-1, Bechtel Corporation, San Francisco, December 1972.

19. "Seismic Analysis of Piping Systems " BP-TOP-1 Rev 2 Bechtel Power Corporation, San Francisco, January 1975.

20. Letter, F. C. Rally to Mark ZZ Technical Stea ring Committee Members, August 22, 1977, MK IZ-301-E,

Subject:

Pool Swell Mode Test Cases.

21. Letter, J. R. Martin to Hark II Owners Group and TSC, MK ZZ-250-E,

Subject:

Condensation Oscillation Excerpts to Applications Memorandum, July 1, 1977.

22 D. Hoffman and E. Schmid, <<Brunsbuttel Nuclear Power Plant List of Test Parameters and Host Important Measurement Results of. the Non-Nuclear Hot Tests with the Pressure Relief System<<(translated by Ad-Ex), R 521/40/77,.

Kraftwerk Union, August 1977.

23. D. Gohel, <<Result of the Non-Nuclear Hot Tests with the Relief System in the Philippsburg Nuclear Power plant<<

(translated by Ad-Ex), R 142-38/77, Kraftwerk Union, Ma rc h 1 977.

11-2 Rev. 2', 5/80

24. D. Hoffman and E. Schmid, "Philippsburg I Nuclear Power Plant List of Test Parameters and Most Important Measurement Results of the Non-Nuclear Hot Tests with the Pressure Relief System" (translated by Ad-Ex), R 521/41/77, Kraft werk Union, August 1977.

25. Klans-D. sterner, "Fxperimental Studies of Vent Clearing in the Model Test Stand'i (translated by Ad-Ex), KHU/R 521-3129, Kraftwerk Union, July 1975.

26. D. Gobel, "KKB Nuclear Start-Up Results of the Tests with the Pressure Relief System" (translated by Ad-Ex), R 142-136/76, Kraf t werk Union, September 1976.

27. D. Hoffman and Dr. K. Melchior, "Cond=nsation arid Vent Clearing Tests in GKM with Perforated Pipes" (translated by Ad-Ex), KrfU/F3-2594, Kraft werk Union, Nay 1973.

28. GE Drawing 761E579, Bechtel Ho. 8856-M1-B11-89

29. ASME Boiler and Pressure Vessel Code, Section ZZI, Division 1, 1974

30. ASME Boiler 2, 1 974.

and Pressure V..ssel Code, Section III, Divi"ion

31. ACI 318-71.

32. R. L. Kiang and B.J. Grossi, "Dynamic Modelling of a Nark IZ Pcessure Suppression System," EPRZ-HP-441, Palo Alto, April 1977.

33. "Seismic Analyses of Structure and Fquipment for Nuclear Power Plants," BC-TOP-4A, Bechtel Power Cocporation, November 1974.

34. MARC-CDC User Information Manual Control Data Corporation, 1975

35. Norse, P. N. and H. Feshbach Nethoils of theoretical physics I NcGraw Hill, New York, Tororito, London, 1953

36. E. Kamcke Differentialgleichungen Losuiigsmethoden und Losungen (Differential Equations Solution Methods and Solutions) Volume I Akademische Verlagsgosellschaf t, Leipzig, 1967.

37. Properties nf Mater and Steam in SZ-Units Springer-Verlag, Berlin, 1969

38. Gobel> KKB hot test results Loads on internal:-. in the pool of he suppression chambec during pressure celief processes 13 No v. 1974; KN U- R 113/203 Rev. 2, 5/80 11- 3

39. Prandtl Stromunqslehre (Hydrodynamics) Vieweq 6 Sohn, Bra unschweig, 1965

40. Werner Tests of mixed condensation with model quenchers KWU-E 3-2593, Nay 1973

41. T. Potna Dehnunqsmessstrei fentechnik (Foil Strain Guage Technology) Philipps-Taschenbucher Tll, 1968

42. NcCandlers Methods Guide for Reactor 'Internal Structure Vibrations Analysis GE Nemo SAR -2A .July 1966

43. Dubbels Taschenbuch fur den Naschinenbau (Dubbels Pocketbook for Machine Construction) Springer, Berlin 1963

44. J. N. Bigqs Introduction to Structural Dynamics, NcGraw Hill, 1964

45. Becker, Gobel, et al. Analysis of the loads measured on the relief system during the KKB non-nuclear hot test KWU-Rll-R31-3346, Ap-il 1975

46. "Nark II Containment Lead Plant Load Evaluation and Acceptance Criteria", Rev. 0, NUREG-0487, U. S. Nuclear Regulatory Commission, October 1978

47. "Dynamic Lateral Loads on a Main Vent Downcomer-Nark II Containment," NEDE-24106-P, General Electric Co., Narch 1978

48. Davis, W. N., NK II Main Vent Lateral Loads Summary Report, NEDE-23806-P, 'General Electric Co., October 1978.

49. Kenleqn, G. H. and Carpenter, L. H. "Forces on Cylinders and Plates in an Oscillating Fluid," NBS J. of Research, Vol. 60, pp. 423-44D, 1958.

50. Sarpkaya, T., "Force on Cylinders and Spheres in a Sinusoidally Oscillatinq Fluid," Trans. ASME, J. of Applied Mech., pp., 32-37, 1975.

51. Chandra, V., Donashovetz, I. and Hsieh, J. S., "Response to NUREG-0487 Crit"ria for Computing Loads on Submerged Structures," SAI-.161-79-PA March 1980.

52. Pankhurst, R. C. and Holder, D. W., "b;ind Tunnel Technique;"

Chapter 8, Pitman and Sons, Ltd., London, 1952.

53. Wilson, E. L, "A Computer Proqram for the Dynamic Stress Analysis of Underqround Structures," USAEWES, Control Report No. 1-175, Jan uar y, 1968.

54. Desai and Abel, "Introduction to the Finite Element Method, ~~

Van Nostroid Reinold Co., 1972.

Rev. 2, 5/80 11-4

55>>IEEE Recommended Practices for Seismic Qualification of Class 1E Equipment For Nuclear Powe" Generating Stations," IEEE Std. 344-1975.

56. A. J. James, "The General Electric Pressure Suppression Containment, Analytical Model,<< GE, July 1971.

57. Let ter FN-080-79, L. J. S obon (GE) to J. F. Stol z (N RC),

Subject:

Vent Clearinq Pool Boundary Loads for Nark II M

Plants 3/20/79.

58. P. W. Huber, A. A. Sonin, W. G. Anderson, "Considerations in Small-scale Nodelinq of Poolswell in BWR Containments,>>

NUREG-CR-'1143, July 1979, Contract No. NRC-04-77-011.

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62. Antony-Spies, P., >>Theorv of the Excitation of Eigenmodes of a Water-Filled Tank by a Callapsing Steam Bubble>>

(t ranslated by Ad-Ex), Technical Report,KWU/R14/77, Sept. ember, 1977.

63. NARC-CDC, User Information Manual, Control Data Corporation, 1976

64. Koch, E. and Sobottka, H., >>KKP 1/KKI Estimate of, the Mitinq Values of the Dynamic Loads on the Pressure Supnression System During Air-Free Condensation at the Vent Pipes", Technical Report KKU/R113/3593, December 1975.

65. "Nark II Improved Chuqqinq Nethodology>>, NEDE-24822-Pg General Electric Company, Nay 1980.

66. >>Single and Multivent Chuqqinq Final. Report>>, NEDE-24300-P, General Electric Company, May 1980.

67. Mark IT Owners Group, "Assumptions for use in Analyzing Nark II BWR Suppression Pocl Temperature Transients Involving Sa fety/Relic f Valve Discharge, <<Revision 1, December 1980.

68. Everstine, G. C., <<A Nastran Implementation of the Doubly Asymptotic Approximation for Underwater Schock Response", Nastran Users's Experiences,. NASA TMX 3428/

pp 207-228, October 1976.

Rev. 5, 3/Sl 11-5

6q. MacNeal, R H., Citerley, R., and Chaiqin, M., >>A New Method for Analyzinq Fluid-Structure Interaction using M.S.C/Nastran>>, Trans. 5th Int Conf. on Structural Mechanics in Reactor Technology, Paper 84/9, August-1979.

70. Mach II Generic Condensation Oscillation Load Definition Report, NEDE-24288-P. General Electric Company, November 1980.

71. C. W. Hirt, B. D. Nichols, N. C. Romero, >>SOLA: A Numerical Solution Alqorithm for Transient Fluid Flows, <<LA-5852/

April 1975.

72. B. D. Nichols, C. W. Hirt, R. S. Hotchkiss, <<SOLA-VOF: A Solution Algorithm for Transient Fluid Flow with Multiple Free Boundaries,<<LA-8355, August 1980.

73. C. W. Hi r t, B. D.

Analysis for N icho ls, L. R. S tein, Pressure Suppression

<<M u Sy i

1t dimensional tems,>> LA-UR 1305, April 1979.

74. Zimmer Nuclear Power Station Unit, Attachment I. k, Amendment 99, Submittal of Revision 61 to the FSAR, September 28, 1979.

75. <<ANSYS Enqineerinq Analysis System Theoretical Manual,>>

November 1, 1977 by Swanson Analysis Systems, Inc.

76. >>ANSYS Enqineerinq Analysis Systems Users Manual" August 1, 1978 by Swanson Analysis Systems, Inc.

77. A. Kalmins >>Analysis of Shells of Revolution Subjected to Symmetrical and Non-Symmetrical Loads<<, Journal of.

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78. A bra hamson, G. R., and Has'hemi, A., "SSES In-Plan Tests to t Measure Submerqed Structure Loads and Pool Frequencies,"

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81. Letter from Roger J. Mateson, Office of Nuclear Reactor Requlations, to Dr. H. Chau, Chairman of the Mark II Owners Group, dated 'February 25, 1982.

Rev. 7, 6/82 11-6

82. T.ett.er f rom G. D. Bouchey to A. Schvencer, <<Desynchronization Nethodoloqy in the Chuqqinq Load Specification,"Rated March 15, 1982, Letter No G02-82-324.

83. Letter from G. D. Bouchey to A. Schwencer,<<Comparison of Structural Response to Symmetric and Asymmetric Chugging and Seismic Toads," dated April 5, 1982, Letter No. G02-8 2-36 2.

84. G. K. Ashley II II and N. N. Howard, <<Understanding Poolsvell in a Nark Type BUR,<<ANS Topical Meeting on Thermal Reacto Safety, July 31-Auqust 4, 1977, Sun Valley, Idaho.

85. B. R. Patel, P. X. Dolan and J. A. Block, Creare TN-307 Report (NEDE-24781-1-P), January, 1980.

Rev. 7, 6/82 1-7 1