ML19209C856

Design Assessment Rept, Revision 6, Revised Pages

ML19209C856
Person / Time
Site:	LaSalle
Issue date:	10/31/1979
From:	COMMONWEALTH EDISON CO.
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ML19209C853	List: ML19209C852 ML19209C856
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NUDOCS 7910180380
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LSCS-MARK II DAR Rev. 6 10/79 TABLE OF CONrENTS (Con t ' d )

PAGE 4.2 CONTAINMENT LINER ASSESSMENT 4.2-1 4.2.1 General Description 4.2-1 4.2.2 Loads 4.2-1 4.2.3 Design Load Combinations 4.2-2 4.2.4 Acceptance Criteria 4.2-2 4.2.5 Analysis 4.2-3 4.2.6 Conclusions 4.2-4 4.3 MAIN DONNCOMER VENT ASSESSMENT 4.3-1 4.3.1 Loads Imposed on Downcomers 4.3-1 4.3.2 Design Load Combinations 4.3-3 4.3.2.1 SRV Actuation Load Combinations 4.3-4 4.3.2.2 LOCA Associated Load Cometnations 4.3-4 4.3.3 Acceptance Criteria 4.3-4 4.3.3.1 Primary Stress Intensity Limits (Equation 9 of NB-3652) 4.3-5 4.3.3.2 Primary-Plus-Secondary Stress Intensity Limits (Equation 10 of NB-3652) 4.3-5 4.3.3.3 Cumulative Damage Usage Factor (NB-365 3. 5) 4.3-5 4.3.4 Downcomer Stress Analysis Methods 4.3-6 4.3.4.1 Static Analysis 4.3-6 4.3.4.2 Response Spectrum Analysis 4.3-7 4.3.4.3 Transient Force Analysis by Mode Superposition 4.3-7 4.3.5 Maximum Downcomer Piping Stresses 4.3-7 4.3.6 References 4.3-9 4.4 PLATFORM AND ACCESS HATCH ASSESSMENT 4.4-1 4.5 SUPPRESSION CHAMBER PIPING AND PIPING SUPPORT ASSESSMENT 4.5-1 4.5.1 Loads on Wetwell Piping 4.5-1 4.5.1.1 Piping Lelow the Pool Surface 4.5-1 4.5.1.2 Piping in the Pool Swell Zone 4.5-2 4.5.1.3 Remaining Wetwell Piping 4.5-2 4.5.2 Load Combination 4.5-3 4.5.3 Acceptance Criteria 4.5-3 4 5.4 Analysis of Section III of ASME B&PV Code Equipment and Components in the Netwell 4.5-4 4.5.4.1 Section III of ASME B&PV Code Pump 6 Section Strainers 4.5-4 4.5.4.2 Valves 4.5-4 iii 7910180 3 80 i178 ,62 d

LSCS-MARK II DAR Rev. 6 10/79 TABLE OF CONTENTS (Cont'd)

PAGE 4.5.4.3 Piping Penetrations 4.5-4 l6 4.6 DRYWELL PIPING AND EQUIPMENT ASSESSMENT 4.6-1 4.6.1 Loads on Drywell Piping 4.6-2 4.6.2 Load Combination 4.6-2 4.6.3 Acceptance Criteria 4 6-2 4.6.4 Qualification of Drywell Equipment and Components 4.6-2 4.7 SUPPRESSION POOL WATER TEMPERATURE MONITORING SYSTEM 4.7-1 4.7.1 Purpose 4.7-1 4.7.2 High-Temperature Steam Quenching Vibration 4.7-1 4.7.2.1 Plant Transients 4.7-2 4.7.2.2 Abnormal Events 4.7-2 4.7.2.3 Primary System Isolation 4.7-3 4.7.2.4 Stuck-Open Relief Valve 4.7-3 4.7.2.5 Automatic Depressurization System (ADS) 4.7-3 4.7.3 Transients of Concern 4.7-3 4.7.4 Design Basis 4.7-5 4.7.5 General System Description 4.7-6 4.8 DOWNCOMER VENT BRACING 4.8-1 4.8.1 General Description 4,g_1 4.8.2 Loads on Bracing 4.8-2 4.8.3 Design Load Combinations 4.8-2 4.8.4 Acceptance Criteria 4.8-2 4.8.5 Downcomer Bracing Analysis 4.8-2 4.8.6 References 4.8-3 A.0 CONTAINMENT RESPONSE TO DYNAMIC LOADS A.0-1 A.1 STRUCTURAL RESPONSE TO SAFETY / RELIEF VALVE DISCHARGE LOADS A.1-1 A.l.1 Analytical Model A.1-1 A.l.2 Method of Analysis A.1-2 A.l.3 Response to All Valve Discharge A.1-3 A.l.4 Response to ADS Valve Discharge A.1-4 A.l.5 Response to Two Adjacent Valve Discharge A.1-5 A.l.6 Response of Single Valve Discharge A.1-5 1178 ?63 iv

LSCS-MARK II DAR Rev. 6 10/79 TABLE OF CONTENTS (Cont'd)

PAGE A.2 STRUCTURAL RESPONSE TO LOCA LOADS A.2-1 A.2.1 Analytical Model A.2-1 A.2.2 Method of Analysis A.2.1 A.2.3 Response to Jet Impingement Loads A.2-2 A.2.4 Response to Cyclic Condensation Loads A.2-3 A.3 DRYWELL FLOOR ANAI.YSIS DUE TO DOWNCOMER WHIP A.3.1 Analytical Model A.3-1 A.3.2 Response of Drywell Floor Due to Downcomer Whip A.3-1 A.4 COMPUTER PROGRAMS A.4-1 A.4.1 DYNAX A.4-1 A.4.2 FAST A.4-2 A.4.3 KALSHEL A.4-2 A.4.4 TEMCO A.4-3 A.4.5 PIPSYS A.4-5 A.4.6 RSG A.4-6 B.0 RESPONSE TO NRC QUESTIONS B.0-1 B.1 QUESTION OF JUNE 23, 1976 B.1-1 B.2 QUESTIONS OF JANUARY 19, 1977 B.2-1 B.3 QUESTIONS OF JUNE 30, 1978 B.3-1 C.0 LA SALLE DESIGN BASIS VS. NRC LEAD PLANT ACCEPTANCE CRITERIA C.0-1 C.1 COMPARISON

SUMMARY

C.1-1 C.2 ACCEPTANCE CRITERION II.A.2 C.2-1 C.2.1 Time Phasing of Bubble Dynamic for Multiple Valve Actuations C.2-1 C.2.2 References C.2-3 C3 ACCEPTANCE CRITERION III.A.1 C.3-1 C.3.1 LOCA Water det Loads C.3-1 C.3.2 References C.3-3 C.4 SUBMERGED STRUCTURE METHODOLOGY C.4-1 6 1178 264 V

LSCS-MARK II DAR Rev. 6 10/79 TABLE OF CONTENTS (Cont'd)

PAGE D.0 FURTHER ANALYSES D.1-1 D.1 FLUID STRUCTURE INTERACTION (FSI) D.1-1 D.l.1 Original FSI Considerations D.1-1 D.l.2 Generic FSI Study D.1-1 D.l.3 La Salle FSI Analysis D.1-2 D l.4 References D.1-3 1178 265 vi

I3CS-MARK II DAR Rev. 6 10/79 LIST OF TABLES (Cont'd)

NUMBER TITLE PAGE 4.1-4 Margin Table for Reactor Support for All Valves Discharge 4.1-10 4.1-5 Margin Table for Base Mat for 2 Valves Discharge 4.1-11 4.1-6 Margin Table for Containment for 2 Valves Discharge 4.1-12 4.1-7 Margin Table for Reactor Support for 2 Valves Discharge 4.1-13 4.1-8 Margin Table for Base Mat for ADS Valves Discharge 4.1-14 4.1-9 Margin Table for Containment for ADS Valves Discharge 4.1-15 4.1-10 Margin Table for Reactor Support for ADS Valves Discharge 4.1-16 4.1-11 Margin Table for Base Mat for LOCA Plus Single SRV 4.1-17 4.1-12 Margin Table for Containment for LOCA Plus Single SRV 4.1-18 4.1-13 Margin Table for Reactor Support for LOCA Plus Single SRV 4.1-19 4.1-14 Margin Table for Drywell Floor for SRV and LOCA Loads 4.1-20 4.2-1 Summary of Containment Wall Liner Plate Stresses /Strans for All SRV Cases 4.2-5 4.2-2 Summary of Containment Wall Liner Anchorage Load / Displacement for All SRV Cases 4.2-6 4.3-1 Load Case Design Data Suppression Pool Downcomers 4.3-10 4.3-2 Downcomer Piping Maximum Combined Stresses 4.3-11 4.5-1 Suppression Chamber Piping 4.5-5 6 A.1-1 Dynamic Soil Properties A.1-6 D.1-1 FSI Amplication Factor D.1-4 D.1-2 Margin Table for Base Mat for All Valves Discharge D.1-5 D.1-3 Margin Table for Base Mat for 2 Valves Diccharge D.1-6 D.1-4 Margin Table for Base Mat for ADS Valves Discharge D.1-7 D.1-5 Margin Table for LOCA Plus Single SRV D.1-8 D.1-6 Margin Table for Containment for All Valves Discharge D.1-9 D.1-7 Margin Table for Containment for 2 Valves Discharge D.1-10 D.1-8 Margin Table for Containment for ADS Valves Discharge D.1-ll D.1-9 Margin Table for Containment for LOCA Plus Single SRV D.1-12 viii 1178 266

LSCS-?! ARK II O.' R Rev. 6 10/79 design will be reviewed for these new loads. During SRV actuation, and after a LOCA, these lines will be subjected to dynamic loads due to building excitation.

4.5.2 Load Combination All load combinations given in Table 6-1 of the DFFR will be considered. The individual loads resulting from the various postulated events will be combined in a fashion similar to that discussed in Section 4.3 of this report.

4.5.3. Acceptance Criteria All piping in the wetwell will be analyzed 11. accordance with the design rules of ASME Section III. Depending on the piping function and the loading combination being con-sidered either upset, emergency, or faulted condition allowables will be used.

In addition, essential systems will meet the following cri-teria for assuring functional capability:

For all piping and classes where D/t < 50, D = Outside Diameter t = Wall Thickness Class 1 Piping For tees and branch connections, 53rvice Level D 2.0 S for austenitic steel For all other pi-ing, Fervice Level C 1.5 S for austenitic steel Y

Class 2,3 Piping For tees and branch connections, Service Level C For elbows, Service Level B, or i178 267 4.5-3

LSCS-MARK II DAR Rev. 6 10/79 Service Level C, when .8B is substituted 2

for .75i and the lower of 1.8 S r 1.5 S h y is used for Service Level C allowables.

For c ar red and straight pipe and all other piping, Service Level B For D/t > 50, 6 Stress indices B 2, B 2b, and B shall be divided by 2r (1. 3-0. 006 D/t) (1.033-0.000 33T) for ferritic mater-ial, and (1. 3-0. 006 D/t) for other materials where: D = outside diameter, t = wall thickness, and T = metal temperature.

4.5.4 Analysis of Section III of ASME B&PV Code Equipment and Components in the Wetwell 4.5.4.1 Section III of ASME B&PV Code Pump Suction Strainers Since the strainers were only purchased recently and have not been fabricated yet, the loads discussed in Subsection 4.5.1 will be factored into their design.

4.5.4.2 Valves There are no valves in the wetwell area.

4.5.4.3 Piping Penetrations The penetration anchor design will be reviewed when the re-vised pipe loads are completed. However, since the penetra-tions were originally designed for pipe rupture loac s it is anticipated that their designs will be satisfactory.

I178 ?68 4.5-4 6

TABLE 4.5-1 SUPPRESSION CilAMBER PIPING TERMINATES NUMBER PENET RATI ON ABOVE BELOW LINE/ DESCRIPTION OF LINES SIZl: E LEVATI ON WATER WATER REMARKS Main steam SRV discharge 18 12 733 ft 7 in. X Drywell floor penetra-tion RilR pump suction 3 24 681 ft 2 in. X L

RilR suppression pool spray 1 4 727 ft 6 in. & X Two supplies to one 730 ft 6 in. Q spray ring j u

R11R relief from heat >

w e exchanger 2 8 704 ft 6 in. X 6$

RiiR test return 2 16 704 ft ') in. &

701 ft 0 in. a X Pump C combines with >

Pump I. line outside conta nment RIIR re.ief valve discharges 6 2 703 ft 6 in., X 704 ft 6 in. &

701 ft 0 in, LPCS pump suction m 1 24 681 ft 2 in. X LPCs pump test return 1 14 703 ft >

in. m LPCS discharge side relief valve discharge 4 701 ft 0 in.

1 X $

N LPCS suction side relief U valve discharge 1 2 701 ft 0 in w

CD e x-sO

TABLE 4.5-1 (Cont'd)

TERMINATES NUMBER P EN ET RATI ON ABOVE BELOW LlN E/DESCRIPTI ON OF LINES SIZE E LE'.'AT I ON WATER WATER REMA RKS

!!PCS pu.To suction 1 24 681 ft 2 in. X llPCS test return 1 14 703 ft 6 in. X llPC.c suction side relief valve discharge 1 2 704 ft 6 in. X RCIC pamp suc ti >n 1 8 688 ft 6 in. X e m

o RCIC turbine e> 1aust 1 10 704 ?t 6 in. X m 6 3 u RCIC suction sit'e relief y w valve discharge 1 6 704 ft 6 in. X X 8

m RCIC vacuum pump discharge 1 l i; 704 ft 6 in. X c

RCIC minimum flow 1 2 701 ft 0 in. X $

Drywell equipment drain 1 4 718 ft 6 in. X Drywell sump penetra-tion and containment wall penetration Drywell equipment drain 1 4 718 ft 6 in. X Drywell sump penetra-cooling tic' and containment g wall panetration <

Drywell f loor d rain 1 4 718 ft 6 in. X Drywell sump penetra-

• tion and containment wall penetration o D.

w W

CD a ss N'

CD

TABLE 4.5-1 (Cont'd)

TE RMIN ATES NUMBER P EN ET RATI ON ABOVE BELOW LI N E/DE SC RI PTION OF LINES SIZE E LEVATI ON WATER WATER REMARKS Water temperature instrumentation 18 1/2 733 ft 75 in. X Drywell floor penetra-tion Air temperature 2 1 X Drywell floor penetra-instrumentation tion C o

m Main steam nigh flow 16 5/4 733 ft 7 ';. in. & X Drywell floor penetra- 6 i n s t ru me n'_a t i on 725 ft 6 in. tion and containment h wall penetratica @

w b Reactor recirc. flow 19 3/4 733 ft 75 in. & [

instrumentation 725 ft 6 in. X Drywell floor penetra- g tion and containment >

wall penetration

• RilR shutdown cooling 4 3/4 733 ft 7 in. & X Drywell floor penetra-instrumentation 725 ft 6 in. tion and containment wall penetration Nuclear Boiler 3 3/4 733 ft 75 in. & X Drywell floor penetra- x 725 f t 6 in. tion and containment C Instrumentation wall penetration .

m Drywell pneumatic 4 3/4 733 ft 75 in. & X Drywell floor penetra-instrumentation 725 ft 6 in. tion and containment -

wall penetration {w W

M CD e \>

N

TABLE 4.5-1 (Cont'd)

TEPNINATES NUMBER PENETRATION ABOVE BELOW LINE/ DESCRIPTION OF LINES SIZE ELEVATION WATER WATER REMARKS 6

IIPCS instrumentation 1 3/4 733 ft 7\ in. & X Drywell floor penetra-725 ft 6 in. tion and containment wall penetration PWCU in s t ru me nt a t ion 2 3/4 733 ft 75 in. & X Drywell floor penetra-725 ft 6 in. tion and containment wall penetration e tn O

L/1

= 0 0

lc O

H W

w M

CD e %>

N N

LSCS-MARK II DAR Rev. 6 10/79 4.6.1 Loads on Drywell Piping In addition to the loading cases of weight, thermal expan-sion, seismic, valve actuation, etc. which have already been considered in the piping and equipment design, two new dynamic loads will be considered. These new loads result from the building acceleration response spectra resulting from the main steam SRV discharge and LOCA as discussed in Sections 3.2 and 3.3 respectively.

4.6.2 Load Combination All load combinations considered will be consistent with those given in the DFFR. The individual loads will be combined in a fashion similar to that discussed in Section 4.3 of this report.

4.6.3 Acceptance criteria All piping in the drywell will be analyzed in accordance with the design rules of Section III of the ASME B&PV Code.

Depending on the piping function and the loading combination being considered either upset, emergency, or faulted con-dition allowables will be used.

In addition, essential systems will meet the following cri-teria for assuring funtional capability:

For all piping and classes where D/t < 50, D = Outside Diameter t = Wall Thickness Class 1 Piping For tees and branch connections, 6 Service Level D 2.0 S for austenitic steel For all other piping, Service Level C 1178 273 4.6-2

LSCS-MARK II DAR Rev. 6 10/79 1.5 S for austenitic steel Class 2, 3 Piping For tees and branch connections, Service Level C For elbows, Service Level B, or Service Level C, when .3B is substituted 2

for .75i and the lower of 1.8 S h r 1.5 S For curved and straight pipe and all other piping, 6 Service Level B For D/t > 50, Stress indices B 2, B 2b, and B shall be divided by 2r (1. 3-0. 0 0 6 D/t) (1.033-0.000 33T) for ferritic mater-ial and (1.3-0.006 D/t) for other materials where: D = outside diameter, t = wall thickness, and T = metal temperature.

4.6.4 Qualification of Drywell Equipment and Components (Section III of the ASME B&PV Code)

The effect of the new load combinations on Section III of ASME B&PV Code equipment and components will be dealt with as follows:

a. The piping reactions on the equipment will be maintained within the equipment vendor allowables by providing additional pipe restraint if needed.

b. The seismic qualification reports for all equipment and components mounted to the floor will be reviewed and checked using the new load combinations.

c. For Seismic Category I valves, the dynamic 4.6-3 t 7b /k

LSCS-MARK II DAR Rev. 6 10/79 coefficients will be computed from thc new piping analysis which considers all LOCA and SRV related dynamic loads. The valves will be qualified to meet these dynamic coefficients or a new piping support arrangement will be developed to reduce the dynamic coefficients to acceptable levels. In addition, active valves will undergo a review to see that the stress allowables are also met.

1178 ?75 4.6-4

LSCS-MARK II DAR Rev. 6 10/79 C.0 LA SALLE DESIGN BASIS VS. NRC LEAD PLANT ACCEPTANCE CRITERIA This appendix provides an assessment of the current design basis for the La Salle County Station against the NRC " Mark II Generic Acceptance Criteria for Lead Plants" of September 18, 1978. This comparison and the information provided, reflects the Mark II Lead Plant positions discussed with the NRC staff in meetings on October 19, 1978, December 12, 1978, February 13, 6 1979, and July 26, 1979. The positions assume that the Newmark/

Kennedy Criteria for use of the SRSS method of load combination will be accepted. In areas where the La Salle position differs from the NRC Acceptance Criteria, support will be provided by Mark II Owners Group Tasks and by La Salle unique efforts as appropriate.

I178 276 C.0-1

MARK II OWNERS GROUP LOAD OR PHENOMENON LOAD SPECIFICATION NRC, pr. VIEW STATI's LA SALLE POSITION ON ACCEPTANCE CRITERI A

1. LOCA-Rel at ed Hydrodynamic Loads A. Submerged Boundary Loads 33 psi over pressure added to loc a l Acceptable Acce p t ab le . The Mark 11 program has provided a realistic During Vent Clearing hydrostat ic below ve nt exit (walls assessment of wall loads based on 4T results (General Electric and basemat ) - linear attenuation let ter MFN -030-79, Mr . Sobon to Mr. Stolz, March 20, 1979). 6 to pool sur f ace.

B. Pool Swell Loads

1. Pool Swell Analytical Mode l a) Air Bubble Pressure Calculated by the Pool Swell Anal- Acceptable ,,,_,

y ti ca l Mode l ( PS AM) used in cal-culation of submerged boundary loads.

C b) Paol Swell Elevat ion 1.5 x submer gence. KRC Criteria 1.A.1 Acce p t able Q eg c) Pool Swell Veloc it y Velocit y history vs. pool eleva- NRC Criteria 1.A.2 Acce pt abl e

^

L. tion predicted by the PSAM used to

/, com pu t e impact loading on small The impact of a 10% increase in pool swell velocity will [

structures and drag on gratings be assessed. Although the assumptions used in the Pool o between initial pool sur f ace and Swell Analytical Model are already erry conservative g maximum pool elevat ion and steady- and eliminate the need for any additional f actors, the state drag between vent exit and resulting calculated load increase should not require maximum pool elevation. Anal- design changes since there are --'- a minimun of component s ytical velocity variation used up in the pool swell region of t'.e wetwell.

to maxitaun ve locit y. Maximun velocity applies therea f ter up to maximum pool swell.

d) Puol Swell Acceleration predicted by the PSAM. Acce pt abl e Acce l erat i on Pool accelerat ion is utilized in the calculat ion of accelerat ion drag loads on submer ged component s 3 during pool swell . $

---* e) Wetwell Air Wetwell air compression is cal- Acce pt abl e _ , , ,

%%j Compre s s ion culated by the PSAM. De f i nes the ,.

pressure loading on the wet we ll R g)g) boundary above the pml surface 7 during paol swell .

4%J%

N N

MARK 11 OWNERS GROUP LOAD OR PRENOMENON LOAD SPECIFICATION NRC REVIEJ STATUS LA SALLE POSITION 04 ACCEPTANCE CRITERI A f) Drywell Pressure Pl an t un ique . Utilized to PSAM Acceptable if 3ased Acceptable.

History to calculate pool swell loads. on NEDM-10320. Other-wise plaat uni pue reviews re qu i r. .d .

2. Loads on Submerged Maximum bubble pressure predicted Acceptable _

Boundaries by the PSAM added uniformly to local hydrostatic below vent exit (we lls and basemat) linear attenua-t ion to pool sur f ace. Applied to walls up to maximum pool swell elevation.

3. Impac t Loads NRC criteria 1.A.6 Acce p t a n t e. Although the criteria is unnecesarily con-servative inves tigations i ndicate t hat, d ue to the size a) Small biructures 1.5 x Pressure-Veloc ity cor rela- and f requency of structures in the La Salle pool swell g tion f or pipes and I beams. zone , t he design loads used are es ns er va *.ive with respect n to the NRC Acceptance Criteria. It should be noted that Y Constant duration pulse.

analyt ical work per f ormed by Sargant & Lundy utilizing f n the PSTF (Pressure Suppression Test Fac il it y) data for 5 circumf erential target s ind ica t e s that the DFFR s pe-L cification is conservat ive for the size and freqaency of structures in the La Salle Pool Swell Zone. Tests E per formed by EPRI ( EPRI No. NP-793, May 1978) to deter-mine fl at pool impact on rigid and flexible cylinders are also in good agreement with DFFR. The NRC Acceptance l6 Criteria utilized an assumption (1-beam impact duration I is inversely proportional to velac it y) which is incon-sistent with theory and experiment al evidence. Never-the les s , the NRC Criteria have Seen used to assess structures in the pool swell zone and these struc tures can withstand the conservat i ve criteria.

b) Large Structures None - Plant unique load where Plant unique review applicable. where applicable NRC Criteria T . A.3 Acc e p t ab le . La Salle has na grating in pool swell area. Y c) Grating No impact load spec ified. P s vs. open area correlat ion an3r*8 o-velocity us elevation history from the PSAM. o D

suumgib CO is> u CO

MARK II OWNERS CROUP LOAD OR PHENOMENON LOAD SPECIFICATION NRC REVIEW STATUS LA SALLE POSITION ON ACCEPTANCE CRITERI A

4. Wetwell Air Compression a) Wall Loads Direct application of the PSAM Acceptable calculated pressure due to wet-well compression.

b) Diaphragm Upward 2.5 psid NRC Criteria I.A.4 Acce pt able Loads

5. Asymmetric Load None NRC Criteria I.A.5 La Salle was a seseqd for an asymmetric load of the vent clearing pressure (22 psig) applied over a 180* sector of the wetwell wall. This load was applied from the base mat to the wetwell floor. The pool hydrostatic load (12 psig at the base mat with linear decrease to zero at water sur f ace) was superimposed on the asymmet ric load. The La Salle design can accommodate this conservat ive load. General Electric has provided an analysis showing that this asymmetric load will actually be less than g 10% of the maximum vent clearing pressure (CE letter n n

~

MFN-076-79, Farch 16, 1979). Based on a subsequent Y analysis by Brookhaven National Laboratories, the NRC  %

-$ revised this criterion to 20% of the smuimun vent clearing E Pressure (NRC/MK 11 Owners Meeting, July 24-25, 1979, -

~

Bethesda, Maryland ). La Salle has been assessed for an asymmetric load five time; greater than the present $

criterion.

C. Steam Condensation and Chugging Loads

1. Downcomer Lateral Loads a) Single Vent Loads 8.8 KIP static NRC Criterna I.B.! Accept able b) Multiple Vent loaJs Prescribes variation of load NRC Criteria I.B.2 Accept able per downcomer vs. number of downcomers. m

2. Submerged Boundary ,

Loads a) High Steam Flux Sinusoidal pressure fluctuation Acceptable , _ _ _ ,

R Loads added to locel hydrostatic. 3 Aplitude uniform below vent

'%%j exit-linear attenuation to pool (3gy surface. 4.4 psi peak-to peak amplitude. 2, 6, 7 Hz frequencits.

8 %s}

N W

MARK II OWNERS GROUP LnAD OR PHENOMENGN LOAD SPECIFICATION NRC REVIEW STATOS LA SALLE POSITION ON ACCEPTANCE CRITF RI A II. SRV-Related Hydrodynamic Loads A. Pool Temperat ure Limite None spec ified NRC Criteria 1;.1 and Acce pt able f or KWU and GE f our arm 11.3 q.cncher Quencher Air Clearing Mark 11 plants utilizing the NRC Criteria II.2 La Salle can accommodate these l oa d s . The first four SRV 6

Loads quencher use an interim load spe- dis ch ar ge cases listed in the NRC Acceptance Criteria cification consisting of the rams have been assessed. In addition, a simultaneous valve head calculational procedure. actuation case is considered. The cases considered Merk 11 plants utilizing the four and the phising involved were discussed with the NRC are quencher use quenche- load in the December 12, 19 78 mee t i ng . This material is methodology described i. DFFR. documented in Section C.2.

Analytical models have been used to predict forcing f unction f requencies f or t he load cases considered. Because of t he g wide range of discharge conditions considered the frequency ,

range used exceeds the 4-11 Hz. range s pec i fi ed . A i 9

g-range used exceeds the 4-11 Hz. ranga s pec i f ied . A presenta-tion on the impact of modifications to the SRV f requency

(

M e range was given in the February 13, 1979 meet.ng. Results l r.

of an assessment of the SRV T quencher frequency range were lg presented at the July 26, 197*, meeting. Document at ion o f these results is being provided.

{

An addit iona. demonst rat ion of the conservatism of the lead plant Co. ( approach SNRC-374, Marchhas been documented 30, 1979 by Long Mr. Novarro (Island LILOD)to Lighting Mr. S. A. Varga NRC transmitting a report entitled, "Justificat ion of Mark II Lead Plant SRV Load De finit ion.'

In plant tests will be run to demonst rate the adequacy and con-servat ism of t he design loads.

B. Que ncLe r Tie-Dawn Loads E

1. Quencher Arm Loads y (a) Four Arm Quencher Vertical and lateral arm loads Acc e p t able ,,,,

developed on the basis of bounding -.

assumptions for air / water dis-  !!

charge from the quencher and ein- 2 servat i ve combinat ions of maximum /

minimum bubble pressure acting on

• • sj the quencher.

O ss)

D CD

MARF 11 CVNERS GROUP LA SALLE POSITION ON ACCEPTANCE CRITERI A NRC REVIEW STATUS LOA 9 SPFCIFICATION LOAD OR PHENOMENON Review Continuing Acceptable. These loads will be calculated using the (b) 6WU T Quencher KWU "T" quencher not included in methodology and assumptions described in DFFR for four arm Mark 11 O.C. Program. T quencher que nche r s , as recommended in the Acceptance Criteria. 6 arm loads not specified at this The KWU T quencher methodology was used to verify the cen-time. servatism of this approach.

2. Quencher Tie-Down Loads Acceptable ,,,,,

(a) Four-Arm Quencher includes vertical and lateral arm load transmitted to the base-mat via the tie downs. See II.C.I .a above plus vertical transient wave and thrust loads.

Thrust load calculated using a Ver-standard momentum balance.

tical and lateral moments for h air or water clearing are cal- y culated based on, conservative clearing assumptions.

g Review Continu;ng Acceptable. These loads will be calculated using the ,,

KWU "T" quencher noi. included methodology and assumptions described in DFFR for four M (b) KWU "T" Quencher nethodology and assumptions described in DFFR for four

[ c in Mar k II 0.G. progr am. T The KWU T-quencher methodology was used to verify the l6 I

g quencher tie-down loads not conservatism of this approach. i specified at this time.

III. LOC A / SRV Submer ged S t ruc t ur e Loads A. LOCA/SRV Jet Loads See Sec t i on C . 3 f or LOCA Jet Loads; SRV Ramshead - RA. l6 Methodology based on a quasi-one- NRC Criteria III.A.I

1. LOCA/ Rams head SRV Jet Loads dimensional model.

Open Item. The spherical tone of influence de fined in No loads spec ified for lead plants. NRC Criteria III.A.2 the Acceptance Criteria is not approppriate f or the

2. SRV-Quencher Jet Loads Model under development in long- two ari quencher. A zone of influence for each arm will ,

term prograa. be defined as a cylinder with an axis coincidental  ;

with the quencher arm. The length of the cylinder will -

he equal to the length of the quencher arm plus 10 end o' cap hole diameters. The radius of thenocylinder is expected structures are within to be quite sma ll . However, because o 5 feet of the quencher arm. 5 feet willofbe assumed. Sn.ce no the quencher, the NRC 23 Structures are located wit 5in 5 feet

"""* Criterion III.A.2 is now satisfied.

N CD is>

~

MARK 11 OWNERS GROUP LA SALLE POSITION ON ACCEPTANCE CRITERIA LDAD SPECIFICATION NRC REVIEW STATUS LOAD OR PHENOMENON

8. LOCA/SRV Air Bubble Drag Loads NRC Criteria III.B.I. The NRC Acceptance Criteria required modification to the

1. LOCA Air Bubble Loads The methodology follows the LOCA present methodology in several areas. The lead plante air carryover phase from bubble have addressed these concerns. Generic documentat ion charging, bubble contract, pool will be provided in a Mark II Owners Group submittat.

rise and pool fallback. The For La Salle County Station, these items have been addressed drag calcula* ions include standard as follows:

and accelera- on drag component s.

a. Bubble asymmetry - The NRC accpetance criterion recommends a 10% increase in velocity and acceleration to accommodate potential asymmetries in the LOCA and SRV air bubbles. La Salle feels that this added f actor is unnecessary given the available data and the conservatism already included in the calculations.

In support of this position, it is best to discuss C LOCA and SRV air bubbles separately.

O The LOCA charging air bubble transient is driven by P the pressurization of the drywell. The limiting g 7* case is a double-ended guillotine break of a recircula-tion line. In spite of the I md probability of this e type of break, additional conservatisms are inc luded o in the calculation of the mass-energy release rate. $

Af ter the bubble begins to form, observations from tests indicate no significant asymmet ries when the bubble expands (EPRI report NP-441, April 1977). Drag 6 coefficients are also conservatively determined as outlined in Appendix l-K. Loads measured in tests show that LOCA air bubble loads are greatly over-predicted by analytical models.

The SRV air bubble loads have been demonstrated to be very conservative by the Caorso in plant tests (CE letter KFN-090-79 to Mr. J. F. Stolz, March 28, 1979) and the KTG single all tests in Germany (as reporte d by T Pennsylvania Power & Light). In addition, multiple "

valve phasing has been assessed for very conservative o-unrealistic conditions. Finally, as noted for the LOCA air bubble, drag coef ficient s have been determined con s e rv at i ve ly . ((

---e In view of the conservatisms identified here and

~. the conservatism of the current methods when compared to test data, no additional multipliers are being applied

'J to velocities and accelerations to simulate postulat ed

()C) bubble as ymmet r ies .

s 's?

N

MARK 11 OWNEPS GROUP LA SALLE POSITION ON ACCEPTANCE CRITERI A LOAD SPECIFICATION NRC REVIE'J STATUS LO AD OR PHENOMENON

b. Standard Drag in Accelerating Flows - Addressed in Sr: tion C.4

c. Nodalization of Structures - Addressed in Section 0.4.

d. Inter ference Ef f ects - AJdressed in Section C.4.

e. 'erference in Downcomer Bracing - Does not apply La l a Salle.

NRC Criteria III.B.2 a. Neglecting Standard Drag - Standard drag is calculated

2. SRV-Rams Head Air The methodology is based on an and included for all submerged structure load calculations.

Bubble Loads analyitcal model of the bubble charging porcess including bubble b. LOCA Bubble Criteria - See Section C.4.

rise and oscillat ion. Ac ce l er a- g tion drag alone is considered. ,

NRC Criteria 111.B.'. TL- bubble location and radius will be defined appropriately

3. SRV-Quencher Air No quencher drag model provided for fot T-quenchers. Bubbles are located near the arms. The $

lead plant s. Lead plants propose bubble size is predicted from the line air volume. E n Bubble Loads interim use of rams head model (See s Ill .B.2 above ). Model will be [

1 developed in long-term program. t Lead plant load s pe- Described in La Salle Clost e Report C. Steam Condensation Drag No generic load methodology provided. Generic model under eification ano NRC Load s review will be con-de velopment in long-term porgram.

dacted on a plant unique basis with confirmation in long-term program using generic model.

E.

O

'2 N

CO as) s Q

U

LSCS-MARK 11 DAR Rev. 6 10/79 APPENDIX C.4 SUBMERGED STRUCTURE METHODOLOGY TABLE OF CONTENTS PAGE C.4-1 C.4 SUBMERGED STRUCTURE METHODOLOGY C.4-1 C.4.1 Introduction C.4.2 Drag and Lift Coefficients for Unsteady C.4-2 Flow C.4-2 C.4.2.1 General Considerations LOCA Charging Air Bubble C.4-7 C.4.2.2 Pool Swell C.4-9 C.4.2.3 C.4-ll C.4.2.4 Fallback SRV Air Bubbles C.4-13 C.4.2.5 C.4.3 Interference Effects on Acceleration Drag C.4-14 6 Method of Analysis C.4-14 C.4.3.1 C.4.3.2 Two Stationary Cylinders (Real Cylinders) C.4-15 C.4.3.3 Stationary Cylinders Near a Plane Boundary C.4-17 (Rea' and Imaginary Cylinders)

Total Acceleration Drag Force C.4-18 C.4.3.4 Practical Application C.4-19 C.4.3.5 Model/ Data Comparisons C.4-21 C.4.3.6 Interference Effects on Standard Drag C.4-22 C.4.4 C.4.4.1 Interference Between Two Cylinders of C.4-22 Equal Diameter C.4.4.2 Interference Between More Than Two C.4-23 Cylinders of Equal Diameter C.4.4.3 Drag on Small Cylinder Upstream of Large C.4-27 Cylinder C.4.4.4 Standard Drag on Smaller Cylinder Down-stream of Large Cylinder C.4-30 Standard Drag on the Large Cylinder C.4-31 C.4.4.5 Structures of Non-Circular Cross-Section C.4-31 C.4.4.6 C.4.4.7 Interference Between Non-Parallel C.4-31 Cylinders 1178 284 C.4-i

LSCS-MARK II DAR Rev. 6 10/79 TABLE OF CONTENTS, Cont.

PAGE C.4.5 The Uniform Velocity and Acceleration 6 Evaluated at the Geometric Center of a C.4-32 Structure La Salle-Specific Parameters C.4-36 C.4.6 References C.4-37 C.4.7 1178 285 C.4-ii

LSCS-MARK II DAR Rev. 6 10/79 APPENDIX C.4 SUBMERGED STRUCTURE METHODOLOGY LIST OF FIGURES NUMBER TITLE C.4-1 LOCA Air Clearing Velocity C.4-2 LOCA Air Clearing Acceleration C.4-3 Top View of 36' Sector of a Typical Mark II Suppression Pool C.4-4 Schematic View of Section A-A of Typical Mark II Suppression Pool C.4-5 Model/ Data Comparisons C.4-6 Cylinder Locations C.4-7 Flow Around Unequal Cylinders 6

C.4-8 Top and Side View of Structure and Source Location C.4-9 Location of Horizontal Submerged Structure in the Suppression Pool C.4-10 Top and Side View of (Second) Structure and Source Location C.4-ll Location of Vertical Submerged Structure in the Suppression Pool C.4-12 Horizontal Component of Velocity Along a Structure Segmented into 4, 8, and 32 Sections C.4-13 Vertical Component of Acceleration Along a Structure Segmented into 4, 8, and 32 Sections C.4-14 Radial Component of Velocity Along a Structure Segmented into 4, 6, 8, and 32 Sections C.4-15 Radial Component of Acceleration Along a Structure Segmented into 4, 6, 8, and 32 Sections 1178 286 C.4-iii

LSCS-MARK II DAR Rev. 6 10/79 C.4 SUBMERGED STRUCTURE METHODOLOGY C.4.1 Introduction Mark II suppression pools are expected to experience fluid motion as a result of safety / relief valve (SRV) actuation and postulated loss-of-coolant accident (LOCA). The veloc-ity and acceleration fields will cause drag loads on sub-merged structures in the pool. These loads have been cal-culated in the past by assuming the existence of a uniform flow field and applying steady-state standard and accelera-tion drag coefficients. The NRC Lead Plant Acceptance Criteria (NUREG-0487, October 1978) pointed out that this method might not be conservative under certain flow condi-tions and for certain structure geometrics. Tnis appendix explains the methods used by the lead plants to ensure 6 that the design loads are conservative.

Subsection C.4.2 presents a method of evaluating the correc-tion to both standard and acceleration drag coefficients in unsteady flow and a method to evaluate the transverse (lif t) force in this flow. Subsections C.4.3 and C.4.4 describe the effect of neighboring structures on che sub-merged structure drag loads. This method provides mc1ified drag coefficients for the range of geometrics existing in 'fark II plants.

Eubsection C.4.5 presents the results of sensitivity st.ud-ies which verify the adequacy of the nodalization used in predicting loads. These studies show that increasing the number of points at which loaas are calculated will not make significant changes in the result.

1i78 287 C.4-1

LSCS-MARK II DAR Rev. 6 10/79 C.4.2 Drag and Lift Coefficients for Unsteady Flow C.4.2.1 General Considerations Drag and lift loads on submerged structures in the suppres-sion pool due to the LOCA charging air bubble, pool swell, fallback, and the SRV air bubble are considered. In calcu-lating these loads on submerged structures, acceleration and standard drag and lift coefficients are used whenever they are applicable to a specific situation. The effects of unsteady flow on the above mentioned coefficients are treated in this subsection, and interference effects, if present, are addressed in Subsections C.4.3 and C.4.4. The steady-state drag coefficients are corrected appropriately to include the effects of unsteady flow.

Because the majority of the available data for unsteady flow have been developed for a cylinder, the discussion provided 6 herein is in terms of a cylinder. The structures in the La Salle suppression pool are all cylindrical. If different-ly shaped structures were present, a cylinder with an equiv-alent diameter would be used for calculations. This ap-proach is conservative for drag load calculations, with the possible exception of the prediction of initiation of vortex shedding. Lift forces due to vortex shedding on sharp-edged structurer would be calculated conservatively based on avail-able data.

Possible shapes of submerged structures in a Mark II suppres-sion pool include circular cylinders, box beams and I-beams.

Equivalent diameters can easily be determined for these structures.

First, to determine the unsteady effects on a submerged structure, a circular cylinder with an equivalent diameter is considered. Then, from the appropriate literature, the drag and lift coefficients due to unsteady flow for cylinders C.4-2

LSCS-MARK II DAR Rev. 6 10/79 are determined, and a drag coefficient multiplier is calcu-lated as the ratio of the unsteady drag coefficients to the steady-state drag coefficients. Finally, these multipliers are applied to the steady-state drag coefficients of the particular submerged structure of interest. However, in computing the loads, the actual dimensions of the structure are utilized in order to properly determine the loads. The lift coefficients are determined utilizing actual lift data for cylinders for the applicable transient conditions.

To determine the equivalent diameter, the length of the par-ticular structure is cor. served and an equivalent diameter is then determined by circumscribing a circle about any structure (Reference 1, pg. 4-14). For a cylinder, the equivalent diameter is equal to the diameter of the cylinder.

If a box beam is consider +1, its equivalent diameter is: 6 2

D # + A EQ $

+ l a

v h-b-For an.I-beam, the equivalent diameter is:

A D = a +b EQ T l i j a v

HbH The nondimensional numbers on which drag coefficients depend are based on the equivalent diameter. They are:

UD m gg

a. Reynolds number - Re = y C.4-3 1178 289 -

LSCS-MARK II DAR Rev. 6 10/79 UT m

b. Period parameter - K = D EQ D

EQ

c. Strouhal number - S = g m

where:

U = maximum velocity at the location of the loaded m

structure during the transient, D = equivalent diameter, EQ T = period of flow oscillation, f = vortex shedding frequency, and v = kinematic viscosity.

6 In order to properly determine the effect of unsteady flow, a drag coefficient multiplier is defined as the ratio of the unsteady to steady-state drag coefficients:

f =

g D

M1 m C M

where:

C = steady-state standard drag coefficient, D

C = unsteady standard drag coefficient, D1 C teady-state acceleration drag coefficient, M

C gy = unsteady acceleration drag coefficient,

= standard drag coefficient multiplier, and f

d f = cceleration drag coefficient multiplier.

m 1i78 ~290 C.4-4

LSCS-MARK II DAR Rev. 6 10/79 These factors are based on drag coefficients for circular cylinders. Once they are determined for cylinders utilizing the appropriate data, the factors are applied to the steady-state drag coefficients for the particular structure of interest. These steady-state drag coefficients are described in Reference 1. The factors are applied in the following manner:

C

• E C D3 d D2 C

• M3 m M2 where:

C D2

= steady-state standard drag coefficient for the particular submerged structure, C = corrected unsteady standard drag coefficient, D3 6 C

M2 steady-state acceleration drag coefficient for the particular submerged structure, and C = corrected unstea3y acceleration drag coeffi-M3 cient.

The lift coefficients are determined utilizing the actual lift data for cylinders for the applicable transient condi-tions.

However, if submerged structures of unique shapes are en-countered in the suppression pool, they need to be considered on a plant-unique basis.

Finally, when all the coefficients are determined, the stan-dard and acceleration drag forces are calculated based on the corrected drag coefficients and the actual structure dimensions. The sum of these forces is the in-line force.

Il78 ?9l C.4-5

LSCS-MARK II DAR Rev. 6 10/79 The transverse force consists only of lift. The following -

equations present the contributions of the standard and ac-celeration drag forces to the in-line force:

p ,CD3 Ap U(t) lU(t)l 2g c p ,CM3 S PU(t)

A C

IN-LINE S+ A where-F = standard drag force, S

A = acceleration drag force, 6

IN-LINE = total in-line force, C = corrected unsteady standard drag coefficient, D3 C = corrected unsteady acceleration drag coeffi-M3 cient, A = projected area of submerged structure, Vg = volume of submerged structura, p = fluid density, U(t) = velocity in the in-line direction, and b(t) = acceleration in the in-line direction.

The determination of the lift force is considered separately in each of the following sections.

1178 ?92 C.4-6

LSCS-MARK II DAR Rev. 6 10/79 C.4.2.2 LOCA Charging Air Bubble The LOCA charging air bubble is considered to be a non-os-cillatory accelerating flow. It is readily observed from the velocity and acceleration time histories of the tran-sient (Figures C.4-1 and C.4-2) that the trancient exhib-ited an increasing velocity and high positive values of acceleration. In addition, the fluid flow never reverses and the acceleration is nearly constant. This is true for all locations in the suppression pool. Comparing typical velocity time histories (Figure C.4-1) to acceleration time histories (Figure C.4-2), one can observe that the accelera-tion is the major contributor to the drag _ cad, since the velocity is Emall. The velocity and acceleration time his-tories were generated utilizing the LOCA charging air bubble model described in Referente 1.

The geometric configuration that was utilized to determine the LOCA charging air bubble t*ansient on a submerged struc-ture is shown in F gures C.4-3 and C.4-4. A 36 segment 6 of a typical Mark II suppression pool was utilized which

ontained 10 downcomers. The LOCA cha_ging air bubbles were used at these downcomer locations to determine the velocity and acceleration time histories on a vertical submerged structure in the pool.

For this transient, Reference 2 is utilized in determining the standard and acceleration drag coefficients, which are conservatively taken as 1.2 and 2.0, respectively.

Due to the low velocities and small duration of this tran-sient, lift due to vortex shedding is not present. In Refer-ence 2, ohe author indicates that for unsteady flow, no lift force due to vortex shedding is present for small-period parameters. In addition, the author states that for a fluid starting from rest, a certain finite time is required for separation to occur if vortex shedding is to be present.

1178 293 C.4-7

LSCS-MARK II DAR Rev. ( 10/79 With the in formation in Reference 2, it was determined that the time required for separation to occur was longer than the duration of the LOCA chargin'g air bubble transient (see the following discussion) . Therefore, lift due to unsteady flow effects is not considered for this transient.

Lift Due to Vortex Shedding According to Reference 2, the separation acessary for vor-tex shedding to be present does not occur until a fluid has movte aistance:

S = 0.293D where:

S = traveled distance, and D = cylinder diar.eter. 6 The authors also state in Reference 2 that:

S=fVt where:

V = velccity, and t = time for separation to occur.

Assuming a representative case for LOCA charging air bubble velocity in a Mark II suppression pool, the maximum pipe diameter required for separation to occur can be determined.

A linear velocity increase from 0.0 to 3.0 ft/sec was as-sumed to closely resemble the actual transient velocity with a transient duration of 60 msec. I  ! grating the velocity, a traveled distance S of 0.09 foot was determined. Utiliz-ing the above mentioned equations, the traveled distance C.4-8 l}[h 294

LSCS-MARK II DAR Rev. 6 10/79 In translates to a maximum pipe diameter of 3-1/2 inches.

inches that other wcmds, for a pipe with a diameter of 3-1/2 experienced a velocity transient increasing linearly from This 60 msec.

0.0 to 3.0 ft/sec, separation would occur at has is the time when the LOCA charging air bubble transient ended.

Following this procedure for other piping wl"hin the suppres-sian pool experiencing the LOCA charging air bubble trans-ient, it can be concluded that lift due to vortex d5edding is not present and does not need to be considered for Mark II pool geometries.

C.4.2.3 Pool Swell Pool swell is regarded as being an oscillatory flow, with the pool swell duration considered to be the half period 6 of the flow field. This flow is exhibited by experimental For Reynolds numbers data, namely, the EPRI and 4T tests.

in the subcritical region, the drag coefficients are deter-mined from Reference 3.

These drag coefficients are depen-However, if the Reynolds dent only on the period parameter.

number is in the supercritical region (> 4 x10 ) , the steady-state standard drag coefficient reduces from 1.2 to 0.71 To determine the unsteady standard and ac-(Reference 4).

celeration drag coefficients for flow at the supercritical Reynolds numbers, Reference 5 is utilized which correlates and the pipe rough-the Reynolds number, period parameter, ness to both the standard and acceleration drag coefficients.

The correlations for smooth pipes are used since these cor-relations best represent the structures within the Mark II suppression pool.

In addition, lift due to vortex shedding is considered.

Reference 6 provides the necessary information to determine lift loads.

As before, the lift coefficient is based on the period parameter and the Reynolds number where both are C.4-9 ll[g )g

LSCS-MARK II DAR Rev. 6 10/79 evaluated at the maximum velocity observed in the transient.

Moreover, the vortex shedding frequency must also be defined.

Reference 5 provides a correlation of fr, the relative fre-quency, to the period parameter and the Reynolds number evaluated at e maximum velocity.

, f f =2 r f y

where:

f = relative frequency, r

f y

= vortex shedding frequency, and f = oscillating fluid frequency.

However, if the relative frequency falls out of the range shown in Figare 21 of Reference 5, Reference 6 indicates 6 that a Strouhal number of 0.3 should be utilized at high Reynolds numbers.

When determining the lift force due to vortex shedding, the maximum amplitude is based on the maximum velocity the struc-ture sees (Reference 6):

2 p _C ApU sin 2 nf y t 2g c

where:

F g = lift force (transverse to flow direction) ,

Cg = lift coefficient, A = projected area of structure, p = fluid density, C.4-10 b

L SC S-M.'.RK II DAR Rev. 6 10/79 U = maximum in-line velocity the structure observes, m

f = vortex shedding frequency, and y

t = time.

The vortex shedding frequency is specified f rom the correla-tion mentioned previously. With the maximum amplitude and frequency, the lift force is defined. The lift force varies with time and is sinusoidal in nature for viscous lift.

In this case, the acceleration drag force considers the effect of gravity and is determined in the following manner:

,Vg [C M3 I DI -9l F

A 9C where: 6 FA = accelcration drag force, V g = volume of submerged structure, C = corrected unsteady acceleration drag coetfi-M3 cient, b(t) = acceleration in the in-line direction, g = gravitational acceleration, and p = fluid density.

C.4.2.4 Fallback Fallback is considered to be a constantly accelerating flow.

It is arsumed that fallback behaves as a falling water slug.

For this case, the standard and acceleration drag coeffi-As with pool swell, cients are determined from Reference 2.

C.4-ll ll[b hh[

LSCS-MARK II DAR Rev. 6 10/79 the lift coefficients due to vortex shedding are also deter-mined. The lift coefficient is given by Reference 7 as Cg = 1.0 for Ka'rman vortices. In order to determine the vortex shedding frequency, Reference 7 indicates that a Strouhal number of 0.22 should be utilized. This is also substantiated by Reference 8. The force is then determined as:

C g ApU sin 2 xf y t L 2g c

where:

Pg = lift force, Cg = lift coefficient, 6

A = project area, p = fluid density, U = maximum in-line vel city the structure observes, m

f y

= vortex shedding frequency, and t= time.

In this case, the acceleration drag force is determined in the following manner:

C p = H 9 S P A

9C where:

F g = acceleration drag force, C g = hydrodynamic mass coefficient (C g =C g = 1),

1178 298 C.4-12

LSCS-MARK II DAR Rev. 6 10/79 V g = volume of submerged structure, 9 = gravitational acceleration, and a = fluid density.

C.4.2.5 SRV Air Bubbles SRV air bubbles are conside'..ed to be of oscillatory nature.

Reference 3 is utilized to determine the unsteady standard and acceleration drag coefficients. These drag coefficients 6 are based on the period parameter evaluated at the maximum velocity. However, if the unsteady drag coefficients are less than the steady-state drag coefficients, then the steady-state drag coefficients are utilized for load deter-mination. If any lift is present, Reference 3 is used, which also bases the lift coefficients on the period parameter.

This reference mentions that no lift is present for period parameters less than 5. The drag loads are determined as described in Reference 1.

Ii78 299 C.4-13

LSCS-MARK II DAR Rev. 6 10/79 C.4.3 Interference Effects on Acceleration Drag When submerged structures are closely located in a flow field, they can interfere with one another, affecting the accelera-tion drag. The proximity effect can be accounted for by either 1) utilizing actual data that is presented in Refer-ences 12, 14, and 15, 2) performing a detailed analysis, or 3) applying a conservative factor of 4 on the accelera-tion drag.

A detailed method is presented for determining interference effects of nearby cylinders and/or a boundary on the acceler-ation drag for circular cylinders and is based on References 9 through 14.

According to the method, the interference effect between any two stationary cylinders can be completely determined by six force coefficients which are functions solely of the 6 radius ratio and the relative spacing. For the case of more than two cylinders, the total proximity effect on a given cylinder may be approximately obtained simply b-j superim-posing each interference effect between cylinder pairs.

C.4.3.1 Method of Analysis The following assumptions are considered in the method:

a. Two-dimensional potential flow without separa-tions and wakes is considered.

b. The velocity and acceleration in the flow field are the same as those seen locally by the sub-merged structure (cylinder).

c. The containment and pedestal walls are consid-ered as plane boundaries.

1178 300 C.4-14

LSCS-MARK II DAR Rev. 6 10/79

d. Coordinate system

+x : radially outward from reactor pressure vessel centerline.

+y : vertically upward.

+z : by right-hand rule parallel to the plane boundary.

Origin: at the center of cylinder in question.

C.4.3.2 Two Stationary Cylinders (Real Cylinders)

If the p-th cylinder is isolated in a free stream, the hydro-dynamic (acceleration drag) force per unit length of the cylinder is:

2 F =2pw A U pg p =n 6

where:

F pg = acceleration drag force per unit length, p = fluid density, A = radius of p-th cylinder, and p

U an = acceleration normal (in-line) to p-th cyl-inder.

When the n-th cylinder is in the vicinity of the p-th cyl-inder, as shown in Figure 1, the change in the force of the p-th cylinder,AF s:

pn, AF pn " # " ^p an (

l ~l)**E(I"1)-C 2exp[i(2S p -a y)]

+ pap lU-n (C3 + 4)exP(iB pn)-C5 expli(3Bpn -2a 2)3

-C xp[i (2a ~

6 2 ""I l  !

1178 501 I C.4-15

LSCS-MARK II DAR Rev. 6 10/79 where:

U=n "

a = Arg (U ), angle of acceleration with respect to z-axis, a

2

=

Arg (Uan ' # 9 z-axis, and S = angle between the line through centers of cyl-pn inders and the z-axis.

=

n Cy = 1+2 L b l,2j+1

]=1

~

C n_ {bl,2j+1 2

=2(_A A p j=1 3

f 1,2k-1 2,2j 6 k=1 j=ll pnA _ 9 1,2k-1 _A n 9 2,2j l p j 4= = b l,2k-1 2,2j-1 C = 4r ^n 4 A p / k =1 j=1[L pn - 9 1 ,2k 2,2j-1 f9 Y 2 = =

C =

f A n)

}

{ [bl,2k-1 b 2,2j-1 A 3 5

4 r(Ap/ k=1 j =l[r'pn _ 9 1,2k-1 _ _n A A p q2,2j-1 g

/A n b 2,2j

=

l,2k C 4nt A t A 3 6

gp j=1 k=1 pn _ 9 1,2k A - A_ n_ q 2 , 2 ],

i

@ ( )

1178 302 C.4-16

LSCS-MARK II DAR Rev. 6 10/79 where:

b y,y =1 2,1 " I b l,m " rm12 2,m-1 9 1,m 2

b 2,m =b 1,m-1 2 ,m 9 fE *1 in which:

91 ,1 = 0, 92 ,1 = 0, A 6 g . p for j)2, y,3 -_ g

_g -

pn n 9 2,j-1 A

q2 r 3' -

n f r j>2, L _A pn p 1,3-1 A = radius of the n-th cylinder, and n

L pn = distance between the centers of cylinders.

C.4.3.3 Stationary Cylinders Near a Plane Boundary (Real and Imaginary Cylinders)

A single stationary cylinder in a uniform stream, U=n, is hydrodynamically equivalent to a cylinder moving at a speed

-U=n in a still fluid, except that the former cylinder ex-periences an extra force, p Tr A U =n, from the pressure field that has been created to provide the fluid acceleration, U Thi3 is also true for any number of cylinders if the mn.

cylinders move together.

1178 503 C.4-17

LSCS-MARK II DAR Rev. 6 10/79 When a cylinder is moving in an arbitrary direction (on a line or on a curve) with respect to the plane boundary, it can be considered as two equal cylinders moving symmetri-cally with respect to the plane boundary, since the plane acts as a perfect reflector (mirror) of the hydrodynamic pressure. A similar argument can be applied to multiple cylinders.

Based on the above discussion, the plane boundary can be removed and replaced by an imaginary cylinder of the same size as the original (real) cylinder and located at twice the distance between the real cylinder and the plane boundary away from the original cylinder. Similarly, multiple imag-inary cylinders can be obtained by reflecting those mul-tiple real cylinders near the plane boundary.

Now if the n-th imaginary cylinder is in the proximity of the p-th cylinder, the change in the hydrodynamic force of the p-th cylinder , A Fpn, can be derived from Reference 9 and is:

AF pn ^p I *n I l -II **E I i" l) - 2**P[II"l+2S pn Il

+o A plU n! I 3 C ) exp(is 4 pn)-C 5exp(i3 B pn) -C 6 **E I- N pn If where:

C y through C 6 are the same as described earlier.

C.4.3.4 Total Acceleration Drag Force Summing up each effect from all the surrounding N real and imaginary cylinders, the force on the p-th cylinder is ap-proximately given as:

n 2 . n F

p

=F pg

+ }]

n=1 AF.,n *

"^ p *n

+

n=

AF pn n/p n/p C.4-18 1178 504

LSCS-MARK II DAR Rev. 6 10/79 or:

F = C puA p m l ban ! + v P ^p l "nl where:

C, = acceleration drag coefficient, and C y = convective force coefficient.

C.4.3.5 Practical Application When the increase in force on the p-th cylinder arising from interference effects is calculated, only those real and imaginary cylinders which have a significant contribution should be considered. Significant contributions to the sum-mation equations presented in Subsection C.4.3.4 arise only from thase cylinder pairs within a gap distance of 3D, where 6 D is the larger diameter of the pair being considered.

If the flow is omnidirectional during a specific transient, a magnification factor KM *^Y * ^ "* * * *^* ""*

lCm I that is determined in the range of 0 5ay1180 . This is performcd to account for the interfercnce effect on the acceleration drag. Similarly, to account for the lift force, the maximum lC y lcan be determined by varying a 2 in the range of 0 Sa 21180*. Then the maximum lift coefficient is com-bined with the standard drag coefficient, C D, by the square root of the sum of the squares to include the lift force due to the proximity effect.

The acceleration and drag forces are determined as follows:

2Kgon A 2 jh lp = p nl 9c I 8 505 C.4-19

LSCS-MARK II DAR Rev. 6 10/79

1. p l U, l U-n lFspl ) C + k l2 where:

F = acceleration drag per unit length on p-th cyl-Ap inder, K

g =lCml/2 magnification factor andlCml is the maxi-mum acceleration drag coefficient, p = fluid density, u

p = radius of p-th cylinder, b -n = acceleration in the normal (in-line) direction to the cylinder, U = vel city in the normal (in-line) direction 6

-n to the cylinder, C = standard drag coefficient, D

! L I = maxint.'m lift coefficient, and P = standard drag per unit length on p-th cylinder.

sp The directions of the acceleration and standard drags are the same as those without the interference effect.

However, if the flow field is well defined and the direction of flow known, then the actual acceleration drag and lif t coefficients are utilized. In this case, the lift force is applied in the transverse direction. The equations used are then:

2C pA m p}n 2

"ag = sc 1178 306 -

C.4-20

LSCS-MARK II DAR Rev. 6 10/79 CgpA pjU=nl U =n F =

Lp 2g c

where:

C, = acceleration drag coefficient determined through the analysis, Cg = lift coefficient determined through the analy-6 sis, and F

gp

= lift force in the transverse direction.

C.4.3.6 Model/ Data Comparisons The method has been tested numerically as well as experimen-tally and the results (Figure C.4-5) indicate excellent agreement with both known numerical values (References 9, 10, 12) and experimental data (References 11, 13).

Il78 507 C.4-21

LSCS-MARK II DAR Rev. 6 10/79 C.4.4 Interference Effects on Standard Drag When submerged structures are located closely together in a flow field, they can interfere wita one another causing an effect on the standard drag, actual data presented in the references can be utilized, a detailed analysis can ce used or a factor of four can be applied to the standard drag force.

This section deals with the detailed analysis to determine the interference etfects on standard drag.

Three technical papers (Ref erences 16 through 18) have de-scribed this pnenomenon and presentea experimental data on interference effects. Most of the data presented in refer-ences are applicable to interference between two cylinders.

Reference 17 has presented some data for three cylinders whose axes are coplanar. Cylinder spacing, Reynolas number, and the angle between flow direction and the plane contain- 6 ing evlinder axes were varied in the above investigations.

In many instances the data obtained from the above refer-ences can be applied direcly to Mark II suppression pool conditions.

A procedure has eeen developed to utilize the above data for interference between more than two cylinders. The results have been compared with measured data of three cylinders and are round to be conservative.

C.4.4.1 Interference Between Two Cylinders of Equal Diameter As indicated by the data given in References 16, 17, and 18, tne interference between two parallel cylinders alters the flow direction drag and also induces a lift force nor-mal to flow direction. For two cylinders of equal diam-eter, the interference effect on drag forces is small, and C.4-22

LSCS-MARK II DAR Rev. 6 10/79 in most cases negative (i.e., the drag is reduced due to I interference). The lift force, however, is not always insignificant.

The following bounding values for interference between two cylinders of equal diameter can be utilized without any further detailed analysis. A bounding value of C D

for Reynolds num'a er greater than 8,000 and S/d ratio greater than 0.2 is 1.4, and the bounding value for C g is 1.0.

C.4.4.2 Interference Between More Tnan Two Cylinders of Equal Diameter To evaluate the drag coefficient of a cylinder whicn is interfered by more than one cylinder, the maximum of C Do I C Di; and D Di,j should be used' 6

n Di - Do

]=1 I Di,; ' 'Do}

jfi where:

C = standard drag coefficient for the i-th cyl-Di inder, C = standard drag coefficient for a single cyl-Do inder without any interference, and C

Di,j = drag coefficient of i-th cylinder when it is interfered by Cylinder j alone.

Figure C.4-6 illustrates an arrangement of cylinders. The stanaard drag coefficient of Cylinder 1 would be:

D1

• Do + I Dl,2 -CDo) + I Dl,3 -CDo) + (C Dl,4 -CDo) 1l78 509 C.4-23

LSCS-MARK II DAR Rev. 6 10/79 From this, the maximum value of CDl; CDo; CDl,2; CDl,3; and C Dl,4 w uld be used for the standard drag coefficient.

To evaluate the lift coefficient, the maximum of C gt and C gt,$ should be used:

n C gg = E C

, gi,3

]=1 j/i where:

C gi = lift coefficient of i-th cylinder, and C gi,3 = lift coefficient of i-th cylinder when it is 6 interfered by Cylinder j alone.

From Figure C.4-6, the lift coefficient would be determined in the following manner:

Cgi = CL1,2 + CL1,3 + Cgy,4 The maximum value of Cgy; CL1,2; CL1,3; nd C gt,4 would be used for the lift coefficient.

The above described method yielding interference on standard drag between more than two cylinders of equal diameter (bounding procedure) is illustrated in the following ex-amples:

1178 510 0

C.4-24

LSCS-MARK II DAR Rev. 6 10/79 Example 1 Consider the three cylinder arrangement shown in the follow-ing figure:

d d

U ,

S 2 0

\

l 4

Leth=1,O=60*, and Re = 2.78

• 10 .

For this arrangement.

6 C Dl,2 = Drag coefficient of Cylinder 1 when inter-fered by Cylinder 2 only.

= 1.01; (h=1)

C Dl,3 = 1.02; (h=3)

C Do

= 1.16 [ Reference 4, pg. 341}

( Dl,2 -CDo) + (C Dl,3 -CD1)

Di~ Do C = 1.01 + 1.02 - 1.16 Di

= 0.37

.'. Maximum of CDi; CD1,2; CDl,3; CDo is 1.16

.'.Use C = 1.16 D

From measurements (Reference 17) :

C = 0.97, which is less than the calculated drag D

coefficient.

C.4-25

LSCS-IIARK II DAR REv. 6 10/79 Example 2 Consider the three cylinder side-by-side arrangement.

4 (B e = 2. 7 8

• 10 , f = 1) i 3 d a

S v

2 d a

u U r I d 6 C

Dl,2 = 1.03 (h-1)

C Dl,3 = 1.05 (f=3)

C = 1.16 Do

..CDl-1.16 = 1.03 - 1.16 + 1.05 - 1.16 or C = 0.92 D1 Maximum of CDo; CDl; CDl,2; C Dl,3 is 1.16 Use C D = 1.16 The measured value is 0.98, which is less than the calculated value.

1178 312 C.4-26

LSCS-MARK II DAR Rev. 6 10/79 C.4.4.3 Drag on Small Cylinder Upstream of Large Cylinder Figure C.4-7 presents the flow around cylinders of unequal diameters.

The coordinates of point A (center of smaller cylinder) are

(-a , b) . Let R and R) be the radii of larger and smaller cylinders. The velocity potential of flow around the larger cylinder in the absence of smaller cylinder is 9 = Ux (1 + I x 2+y 2 where:

9 = velocity potential, 6

U = free stream velocity, and R = radius of large cylinder.

If u and v are the x and y components of velocity at point "A" (smallec cylinder absent) then,

~#

u=U 1+R 2 (a + b ) 2) _

2ab R U and v=

(a2+b)2 2 where:

u = the velocity parallel to the free stream velocity at the centerline of the smaller cylinder, and v = the velocity perpendicular to the free stream velocity at the centerline of the smaller cyl-inder.

C.4-27

LSCS-MARK II DAR Rev. 6 10/79 1

In order to use the above velocity correction, it is more convenient to increase the standard drag coefficient and use the corrected standard drag coefficient with the free stream velocity. The standard drag force on the smaller cylinder with interference present is:

C D

1. A (u2,y2)

Fy= 2g c

where:

Fy = standard drag force with interference, C

D

= standard drag coefficient, o = fluid density, A = projected area of smaller cylinder.

6 The direction of flow is:

-1 8 = tan

({}

Tne standard drag force without interference is:

C D

o AU F=

2g c

where:

F = standard drag force without interference, and U= free stream velocity.

1178 5i4 C.4-28

LSCS-MARK II DAR Rev. 6 10/79 The ratio of the two standard drag forces yields the follow-ing expression for the correction of the standard drag co-efficients:

C D (interference)

C D (without interIerence) =m 2+n2 where:

l 2 2

-a m=

1+R I (a2+b)2 2 f-3_

, and n= 2ab R (a2+b)? 2 '

If the determined ratio of the standard drag coefficients 6 is less than 1, then a ratio 1 is used. However, if the de-termined ratio is greater than 1, then the determined ratio is utilized.

The standard drag force on the smaller cylinder is then de-termined by:

F C

D AU lUl 1= 2g where:

Fy = standard drag force on smaller cylinder, C

D = the corrected standard drag coefficient for in-terference, and U = free stream velocity in the in-line direction.

1178 315 C.4-29

LSCS-MARK II DAR Rev. 6 10/79 The lift force on the smaller cylinder is determined in the same manner as ::as the standard drag force:

C g (interference)

C g (without interference)

=m 2+n2 where m and n are the same as previously described. Once again, if *he determined ratio is less than 1, then the ratio of 1 is utilized. However, if the determined ratio is greater than 1, then the determined ratio is used. The lift force is then determined by:

Cg p AU lUl Fg -

2g c

6 where:

Fg = lift force, and C g = corrected lift coefficient for interference.

The lift force is applied in the transverse direction to the resultant flow.

C.4.4.4 Standard Drag on Smaller Cylinder Downstream of Large Cylinder lf the smaller cylinder is located downstream of the larger cylinder, then the lift and standard drag ccefficients are evaluated. corresponding to S/d (where S is the distance be-tween the cylinder surfaces and d t.he diameter of the cyl-inders) ratios for both cylinders, assuming equal diameters.

The coefficients are first determined by assuming both cyl-inders equal to the diameter of the small cylinder, and then assuming both cylinders equal to the diameter of the large cylinder. When determining the coefficients, the centerline C.4-30 178 516

LSCS-MARK II DAR Rev. 6 10/79 distance between the two submerged structures is maintained at the actual distance.

Afterwards, the larger coefficients are used for determin-ing submerged structure loads. In addition, if the lift and standard drag coefficients are less than the coefficients without interference, then the coefficients without inter-ference are utilized.

C.4.4.5 Standard Drr.g on the Large Cylinder 6

The method describcd in Subsection C.4.4.4 is used to de-termine the standard drag and lift on the large cylinder.

C.4.4.6 Structures of Non-Circular Cross-Section The methodology of determining an equivalent diameter de-scribed in Section 6.4.2 is used to determine the coeffi-cients due to interference effects.

C.4.4.7 Interference Between Non-Parallel Cylinders To estimate the interference effects between non-parallcl structures, the .ift and standard drag coefficients are de-termined by assuming that the structures are parallel. The distance utilized between them would De tne minimum distance between the two structures. In the same manner, the larger coetticients of either with or without interference effects dte chosen for submerged structure load determination.

I178 5i7 C.4-31

LSCS-MARK II DAR Rev. 6 10/75 C.4.5 The Uniform Velocity and Acceleration Evaluated at the Geometric Center of a Structure In evaluating drag forces, the velocities and accelerations at the geometric center are utilized. These velocities and accelerations are used since the standard and accelera-tion drag coefficients are defined at the geometric center of a structure. However, if a flow field is not axially uniform along a particular submerged structure, the struc-ture is segmented into smaller sections. Then the velocity and acceleration closely resemble uniform conditions over the segment. In addition, the flow is always assumed to be locally uniform as evaluated normal to and on the axis of the structure at the midpoint of each segment. The parameter L/D (where L is the segment length and D the outside diameter of the structure) is utilized to segment 6

a submerged structure in order to provide an adequate repre-sentation of the flow field along the structure. This parameter is a structural criteric a and is used in the analysis of submerged structures. It was determined that a segmentation that yields approximately a 1.0 1 L/D j l.5 is adequate to properly describe the flow field along a structure. This segmentation range is valid for all struc-tures and can be utilized for all transients in the suppres-slon pool. Furthermore, as the outside diameter of a sub-merged structure is increased, the segment length should decrease. When the segment length is decreased, a better representation of the flow field is obtained. If the out-side diameter is less than 1.0 foot, the segmentation should be equal to an L/D of 1.5. However, if the outside diameter is equal to or greater than 1.0 foot, a segmentation equal to an L/D of 1.0 should be used. If the determined number of segments along a structure is a fraction, i.e. not an integer number, then the fraction number of segments should be rounded up to the greater integer number of segments.

1i78 $18 C.4-32

LSCS-MARK II DAR Rev. 6 10/79 In order to determine the adequacy of the segmentation, a computer program was utilized for evaluating the velocities and accelerations along a structure. Two submerged struc-tures were used to assess the adequacy of the segmentation.

The first structure was a 1.0-foot-diameter, 6-foot-long horizontal cylindrical structure. A vertical structure 21 feet long and 3.5 feet in diameter was also used. A bubble was utilized to load both submerged structures, and the method of images was used to provide the velocities and accelerations in the suppression pool. The first struc-ture was segmented into four sections, each 1.5 feet long.

This provided for an L/D of 1.5. Then the structure was segmented into eight sections, each 0.75 foot long, which provided for an L/D of 0.75. The segmentation of the struc-ture is shown in Figure C.4-8. The structure was placed horizontally in the suppression pool, with tl'e bubble located 6 3.5 feet vertically below and 4.13 feet horizontally away from the centerline of the structure. In addition, the bubble was placed 1.4 feet away from an end point of the structure. The location of the submerged structure in the suppression pool is shown in Figurc C.4-9. In order to represent the actual velocities and accelerations along the structure, the structure was segmented into 32 sections.

The second structure was also segmented into four and eight sections, each 5.25 feet and 2.625 feet, respectively.

This provided for an L/D of 1.5 and 0.75 when the structure was segmented into four and eight sections. Moreover, to provide an L/D of 1.0, the structure was segmented into six sections, each 3.5 feet in length. In the same manner, to represent the actual velocities and accelerations along the structure, the structure was segmented into 32 sec-tions. The structure was placed vertically in the suppres-sion pool, and the bubble was located 4.0 feet horizontally away from the centerline of the structure and 6.0 feet from an end point of the structure. The segmentation scaemes Il78 5l9 C.4-33

LSCS-MARK II DAR Rev. 6 10/79 are shown in Figure C.4-10, and the location of the struc-ture in the suppression pool in Figure C.4-ll.

Figures C.4-12 and C.4-13 illustrate the velocities and accelerations along the horizontal submerged structure.

Figures C.4-14 and C.4-15 show the velocities and accelera-tions along the vertical structure for the various segmenta-tion schemes described earlier. One can readily observe from these figures that as the segment length was reduced, the velocities and accelerations were more accurately de-scribed in the suppression pool. Figures C.4-12 and C.4-13 reveal that the velocities and accelerations for the small nutizontal structure do not vary greatly when the number of segments is increased from four to eight. However, Figures C.4-14 and C.4-15 indicate that for a large struc-ture, the minimum number of segments that can be used to 6

sufficiently describe the velocities and acceleration along the structure is six, which translates to an L/D of 1.0.

In order to properly assess the adequacy of the segmenta-tion, standard and acceleration drag loads were determined for each segmentation scheme. When the standard drag load along the structure for both the four- and eight-segmented structures was determined, only a 0.12% change resulted from reducing the segment size. When the acceleration drag load was determined along the same structure, a change of approximately 0.07% was attained when the number of segments was increased from four to eight. However, when compared to the results of segmenting the structure into 32 sections, the largest change in the acceleration drag load of 0.09% was exhibited when the segmentation was in-creased from 4 to 32 sections. A change of only 0.02%

resulted in the acceleracion drag load when the segmenta-tion was increased from 8 to 32. The increase in the stand-ard drag load was 0.15% and 0.03% when the segmentation was increased from 4 to 32, and 8 to 32, respectively.

C.4-34

LSC3-MARK II DAR Rev. 6 10/79 When the standard drag load was determined for the verti-cal submerged structure and compared with the actual stand-ard drag load (32 segments), changes of 14.47%, 1.1%, and 0.11% resulted when the segmentatior. ,Ts increased from 4 to 32, 6 to 32, and 8 to 32, respectively. In the same manner, when the acceleration drag load for the same struc-ture was compared with the actual acceleration drag load (32 segments), changes of 3.5%, 0.23%, and 0.03% resulted when the segmentation was increased from 4 to 32, 6 to 32, and 8 to 32, respectively. Comparing the standard drag load of the structure when segmented into 4 sections 6 with the structure segmented into 32 sections yielded re-sults that inaccurately described the flow field in the suppression pool. As previously noted, a change of 14.47%

resulted when the four-segmented structure was compared with the actual representation. The four-segmented struc-ture yielded an L/D of 1.5, which is considered to be in-appropriate for large cylinders. When utilizing six seg-ments, an L/D of 1.0 was obtained, which showed very good agreement of the standard drag load with the actual load that resulted in a change of only 1.1%. The acceleration drag load did not vary considerably when the segmentation schemes were differed. This shows that a segmentation of a structure yielding a 1.0 1 L/D < 1.5 is sufficient to provide an adequate description of both the velocities and accelerations for design purposes.

1178 321 C.4-35

C.4.6 La Salle-Specific Parameters Submerged Structure Spacing Structures Considered Minimum Spacing ( t)

Downcomer to Downcomer 3'-6" to 8'-0" Downcomer t o Column 4'-6" to 5'-3" Downcomer to SRV Line 2'-3" to 4'-6" Submerged Structure Size StrucLure Diameter Downcomer 24" g Column 36" O l

12" 6 :r O SRV Line g b Flow Characteristics Re H

• max H Phenomena Flow Type Period Parameter

• LOCA Charging Constant Acceleration <0.28 $10 6 g 5 6 LOCA Pool Swell Oscillatory Flow ** 50-75 10 -10 LOCA ballback Constant Acceleration *** 10 6 6

SRV Discharge Decaying Oscillation <4 10

• Period Parameter = Umt /d (Oscillatory Flow) g s/d (Constant Acceleration) ,

N

• O ** Pool Swell is considered a portion of a :.. ngle cell of an oscillatory flow.

• • • Varies greatly with structure size and location. Vortex shedding is y

(

y assumed and lift force calculated.

LSCS-MARK II DAR Rev. 6 10/79 C.4.7 References

1. " Analytical Model for Estimating Drag Forces on Rigid Submerged Structures Caused by LOCA and Safety Relief Valve Ramshead Air Discharges," NEDO-21471, September 1977.

2. T. Sarpkaya and C. J. Garrison, " Vortex Formation and Resistance in Unsteady Flow," Journal of Applied Mechanics, pp. 16-24, March 1963.

3. T. Sarpkaya, " Forces on Cylinders and Spheres in a Sinusoidally Oscillating Fluid," Transactions cf the ASME, pp. 32-37, March 1975.

4. G. K. Batchelor, An Introduction to Fluid Dynamics, pg. 341, Cambridge, England, 1967. 6

5. T. Sarpkaya, "In-Line and Transverse Forces on Cylin-ders in Oscillatory Flow at High Reynolds Number,"

Journal of Ship Research, Vol. 21, No. 4, pp. 200-216, December 1977.

6. T. Sarpkaya, " Vortex Shedding and Resistance in Har-monic Flow About Smooth and Rough Circular Cylinders at High Reynolds Numbers," NPS-59SL76021, February 1976.

7. J. P. Den Hartog, Mechanical Vibrations, Fourth edi-tion, pp. 305-306, McGraw-Hill Book Co., Inc., New York, 1956.

8. J. A. Roberson and C. T. Crowe, Engineering Fluid Mechanics, pp. 338-340, Houghton Mifflin Co., Illinois, 1975.

1i78 523 C.4-37

LSCS-MARK Il DAR Rev. 6 10/79

9. T. Yamamoto, " Hydrodynamic Forces on Multiple Cir-cular Cylinders," Journal of the Hydraulics Division, ASCE, pp. 1193-1210,- September 1976.

10. T. Yamamoto and J. H. Nath, " Hydrodynamic Forces on Groups of Cylinders," Paper No. OTC 2499, presented at the Offshore Technology Conference, Houston, Texas, May 3-7, 1976.

11. T. Yamamoto and J. H. Nath, " Forces on Many Cylinders Near a Plane Boundary," Preprint 2633, presented at the ASCE National Water Resources and Ocean Engineer-ing Convention, San Diego, California, April 5-8, 1976.

12. C. Dalton and R. A. Helfinstein, " Potential Flow Past a Group of Circular Cylinders," Journal of Basic En-gineering, ASME, pp. 636-642, December 1971.

13. T. Yamamoto, J. H. Nath, and L. S. Slotta, " Wave Forces 6 on Cylinder Nearby a Plane Boundary," ASCE Journal of Waterways, Harbors and Coastal Engineering Divi-sion, pp. 345-359, November 1974.

14. T. Sarpkaya, " Forces on Cylinders Near a Plane Boundary in a Sinusoldally Oscillating Fluid," Journal of Fluids Engineering, ASME, pp. 499-505, September 1976.

15. T. Sarpkaya, "In-Line and Tr ansver se Forces on Cyltn-ders Near d Wall in Oscillatory Flow at Hign Reynolds Numbers," Paper No. OTC 2898, presentea at the Offshore Technology Conterence, Houston, Texas, May 2-5, 1977.

16. E. I. Hotl, " Experiments on Flow Around a Pair of Parallel Circular Cylinders," proceedings of the 9tn Japan Nattoral Congress for Applied Mechanics, pp.

231-234, 1959.

1178 324 C.4-38

LSCS-MARK II DAR Rev. 6 10/79

17. C. Dalton and J. M. Szabo, " Drag on a Group of Cylin-ders," Transactions of the ASME, Journal of Pressure vessel Technology, pp. 152-157, February 1977.

6

18. M. M. Zdravkovich, " Review of Flow Interference Be-tween Two Circular Cylinders in Various Arrangements,"

Transactions of the ASME, Journal of Fluids Engineering, pp. 618-633, December 1977.

I178 525 C.4-39

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12.8' 15.O'

- DOWNCOMERS

$ - SUBMERGED STRUCTURE LA SALLE COUNTY STATION MARK 11 DESIGN ASSESSMENT REPORT FIGURE C.4-3 1178 334 TOP VIEW 0F 36 SECTOR OF A TYPICAL MARK II SUPPRESSION POOL

Rev. 6 10/79 i i i i DRYWELL FLOOR  ; i t s

a 7 21.8' I

a l DOWNCOMERS 8.8 --

SUBMERGED 2 STRUCTURE WITH LOAD 3 POINTS I THROUGH 4 7.8' 4

I'O o y v l-l 2.8'M RADil FROM , l 5 O'--*

CENTERLINE OF 24.3'  ;

CONTAINMENT ,

ARE: , 30.8  ;

35.8' LA S ALLE COUNTY ST ATION MARK 11 OESIGN ASSESSMENT REPORT

~

1178 335 FIGURE C.4-4 SCHEf1ATIC VIEW 0F SECTION A-A 0F TYPICAL MARK II SUPPRESSION FOOL

Rev. 6 10/79 a

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Rev. 6 10/79 lx Z-+

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l 2 3 4 Z/D LA SALLE COUNTY STATION MARK 11 DESIGN ASSESSMENT REPORT FIGURE C.4-5 MODEL/ DATA COMPARISONS (SHEET 2 of 4)

Rev. 6 10/79 1 1 1 al n

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O I 2 3 4 5 6 7 y/a j LA SALLE COUNTY ST ATION 3 MARK 11 DESIGN ASSESSMENT REPORT 1178 538 FIGURE C.4-5 MODEL/ DATA C0f1 PARIS 0NS (SHEET 3 of 4)

Rev. 6 10/79

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))[h} MARK ll L4 51GN ASSESSMENT REPORT FIGURE C.4-5 MODEL/ DATA COMPARISONS (SHEET 4 of 4)

Rev. 6 10/79 d 2

                                /
                  /

30* - 3Oo 2d U - o ,_ \

                           /
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FLOW " > "l n d I g LIFT, L i 4 d y DR G,D r LA SALLE COUNTY STATION 1178 540 -- ,, o es,e~ .s s esse ~1 e<ecer FIGURE C.4-6 CYLIflDER LOCATIONS

Rev. 6 10/79 v JL A " R I S

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j j ~/' @ jfj LA SALLE COUNTY STATION MARK 11 DESIGN ASSESSM ENT REPORT FIGURE C.4-7 FLOW AROUND UNEQUAL CYLINDERS

Rev. 6 10/79 TOP VIEW 0F STRUCTURE STRUClon2 SEGMENTED INTO 8 SECTIONS LOAD LOAD LOAD LOAD LOAD LOAD LOAD LOAD POINT POINT POINT POINT POINT POINT POINT POINT N0. 1 NO. 2 N0. 3 N0. 4 N0. 5 NO. 6 NO. 7 N0. 8 I I I I I I I I N NI N N NI N N N,POINTNNNN I H LOAD P0 INT LOAD POINT LOAD LOAD POINT N0. 1 N0. 2 N0. 3 NO. 4 STRUCTURE SEGMENTED INTO 4 SECTIONS 1 FOOT DIAMETER CYLINDER O' SOURCE FROM CENTERLINE LOCATIO4 OF STRUCTURE 4.13 FEET

• --- 1. 4 FT --*

SIDE VIEW 0F STRUCTURE

4.13 FT  :

1 FOOT o DIAMETER CYLINDER 3.5 FT LA S ALLE COUNTY ST ATION SOURCE MARK 11 DESIGN ASSESSMENT REPORT FIGURE C.4-8 2 TOP AND SIDE VIEW 0F STRUCTURE AND SOURCE LOCATIONS

Rev. 6 10/79 TOP VIEW OF 36* SECTOR 36'

                                                      /                    ,

24

                                           /        /
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A

                                  //
                                   /e          SUBMERGEO STRUCTURE               -8*   )
                   ~         O ~ BUBBLE l                                                       O*

RADil FROM CENTERLINE OF 12.8' ,15.0' 19.0 i 35.8 i CONTAINMENT ARE : 13.0 SIDE VIEW OF SECTION A-A a SUBMERGED

                                           / STRUCTURE n              O:      BUBBLE 8.75' y   ir   U LA sat LE COUNTY STATION MARK 11 DESIGN ASSESSMENT REPORT 1178      ,54:,

FIGURE C.4-9 LOCATION OF HORIZONTAL SUBMERGED STRUCTURE IN THE SUPPRESSION POOL

TOP VIEW 0F STRUCTURE Rev. 6 10/79

                                    + 4 . 0 F T -.-

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                                   .         j     .       3.5 FOOT DIAMETER SOURCE CYL IN DER o

SIDE VIEW OF STRUCTURE

                                                      \

LOAD POINT LOAD POINT *

• NO. 1 NO. 1 LOAD POINT NO. 2 LOAD POINT A 21 FOOT VERTICAL NO. 2 STRUCTURE SEGMENTED . LOAD POINT INTO 6 AND 8 SECTIONS _ NO. 3

                                 ,                       . LOAD POINT NO. 4 LOAD POINT ,                  +-

0. NO. 4 SOURCE LOCATION LOAD POINT 4.0 FEET FROM -

                             ~
                                   .                            NO. 6 CENTERLINE OF STRUCTURE LOAD e INT NO. 5                     + LOAD POINT NO. 7
                                                          ~

LOAD POINT \ NO. 6

• LOAD POINT NO. 8 LA SALLE COUNTY STATION MARK ll DESIGN ASSESSM ENT REPORT r FIGURE C.4-10 TOP AND SIDE VIEW 0F (SECOND) STRUCTURE AND SOURCE LOCATION

Rev. 6 10/79 TOP VIEW OF 36' SECTOR 36

                                                        /
                                            /                                   >

18 BUBBLE UBMERGEO STRUCTURE A O RADil FROM CENTERLINE OF CDNTAINMENT ARE: 12.8' 21.O' 25.O' 35.8' SIDE VIEW OF SECTION A- A a I s 21.O' BUBBLE SUBMERGED

                                                        # STRUCTURE L                     U I

6.O'

             ,    u

_L LA S ALLE COUNTY STATION

                                            -        MARK 18 DESIGN ASSESSMENT REPORT FIGURE C.4-il LOCATION OF VERTICAL SUBMERGED STRUCTURE IN THE SUPPRESSION POOL

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    =     5Z 3

Rev. 6 10/79 1.0

                                                             \

O.9 i  ! O.8 I X $ I R o7 i D I l U 0.6 5 g -_ _. _ l_ _ d __ i I > 0.5 l a l 30.4 - l i l l I 4 SEGMENTED $ l STRUCTURE g l l L-

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0.2 l g [6STRUCTURE SEGMENTE 32 SEGMENTED I -- 8 SEGMENTED I STRUCTURE O .1 STRUCTURE SOURCE LOCATION

                  - -125   2"-                             "

0.0 i-O 2 4 6 8 10 12 14 16 18 20 22 DISTANCE ALONG STRUCTURE (FT) LA SALLE COUNTY STATION Jj7g }4g MARK 11 DESIGN ASSESSMENT REPONT FIGURE C.4-14 RADIAL COMPONENT OF VELOCITY ALONG A STRUCTURE SEGMENTED INTO 4, 6, 8 AND 32 SECTIONS

Rev. 6 10/79 1.0 -

                                                                    ~

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_ _ _1_ _ } O.7 ' i i ' ! 8 i 8 1 4 SEGMENTED P i I STRUCTURE

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                            ))[h        jf }                     LA SALLE COUNTY STATION MARK 11 DESIGN ASSESSMENT REPORT FIGURE C.4-15 RADIAL COMPONENT OF ACCELERATION ALONG A STRUCTURE SEGMENTED INTO 4, 6, 8 AND 32 SECTIONS}}