ML20129D760

Rev 2 to Volumetric Flowrate as Function of Time from Drywell to Torus (& Return)

ML20129D760
Person / Time
Site:	Browns Ferry
Issue date:	06/13/1996
From:	Leaver F, Metcalf J POLESTAR APPLIED TECHNOLOGY, INC.
To:
Shared Package
ML19353D888	List: ML20129D562 ML20129D639 ML20129D737 ML20129D760 ML20129D786 ML20129D810 ML20129D819 ML20129D844 ML20129D874 ML20129D883 ML20129D891 ML20129D898
References
PSAT-04011H.01, PSAT-04011H.01-R02, PSAT-4011H.1, PSAT-4011H.1-R2, NUDOCS 9610250129
Download: ML20129D760 (36)

Text

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4 PSAT 04000U.04 -

Attachment 3 PSAT Calculation 0401IH.01

" Volumetric Flowrate as a Function ofTime from Drywell to Torus (and Return)"

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Rev: 01@3 4 CALCULATION TITLE PAGE CALCULATION NUMBER: PSAT 04011H.01 4

l CALCULATION TITLE:

" Volumetric Flowrate as a Functrua of Time from Drywell to Tonis (and Return)"

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4 REASON FOR REVISION: Nonconformance Rot j O -InitialIssue N/A 1

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! 1 - To add Appendix B and Appendix C and to correct cover sheet N/A for Appendix A 2 - To update Reference 3 and expand Reference 9 N/A 4

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l PSAT 04011H.01 Page: 2 of19 Rev: Oh 3 4 Table of Contents j Section P. age 3

Purpose 2

Methodology 2

Assumptions 3

. References 6

Calculation 7

i Results 18 Conclusions 19 i

Appendices: A "Use of a Uniform Sweep-Out Rate During the Release Phase" - 3 Pages B " Impacts of Transient Heat Conduction" - Non-Safety Related - 3 Pages C " Comparison to Severe Accident Analyses" - Non-Safety Related - 11 Pages '

Purpose The purpose of this calculation is to specify the volumetric exchange rates between the Browns Ferry drywell and the torus during two periods of the problem: during the fission product release (gap release phase from 30 seconds to 1830 seconds and early in-vessel release phase from 1830 seconds to 7230 seconds - see Table 3.6 ofReference 1) and after the fission product release phase (7230 seconds until 30 days which is the end of the dose calculation interval from j

Reference 2). During (and immediately after) the fission product release phase the flow is only I from the drywell to the torus and may be referred to as the " sweep-out" rate.

Methodology i

i In order to specify the volumetric sweep-out rate, it is necessary to know the quantity of water remaining in the vessel after the DBA blowdown, the thermodynamic state in the drywell, and the rate at which steam is produced from the core debris in-vessel up to and including the point in i time where the core-debris quench is complete (assuming that to be shortly after 7230 seconds, 1

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PSAT 04011H.01 Page: 3 of19 Rev:@l 2 3 4 the end of the in-vessel release phase). Beyond 7230 seconds + the reflood time, the containment l

is assumed to be well-mixed, but a mixing rate must be specified to reflect that assumption. '

A manual calculation is shown belo'w which: .

• Quantices th; minimum water mass remaining in the vessel after DBA blowdown, o Determinos a minimum steaming rate for that remaining water, and e Calculates the volumetric flowrate rate (drywell to torus) that corresponds to that steaming rate and to the final quench of the core debris.

Assumptions Assumption 1: Reactor vessel reflood occurs at 7230 seconds, terminating the release and quenching the core debris.

Justification: This assumption reflects the position that Reference 3 takes with respect to the release phases of Reference 1. Reference 3 references an NRC position taken on the advanced light water reactors in Reference 4, which is:

"In a forthcoming paper, the NRC staff will indicate that for evaluation of design basis accidents (DBA) for evolutionary and passive light-water reactor designs, only the releases associated with the gap and early in-vessel release phases will be used. The inclusion of the ex-vessel and late in-vessel releases'are considered to be unduly conservative for DBA  !

purposes. Such releases would only result from core damage accidents with vessel failure and core-concrete interactions."  ;

This NRC position, as extended to operating reactors by Reference 3, means that vessel failure is not to be included in the DBA. This position also implies, then, ,

that debris coolability must be re-established at about the time of the end of the in- l vessel release phase; otherwise, reactor vessel failure would likely follow. "

Assumption 2: Containment is well-mixed following the core debris quench at 7230 seconds + the time to reflood.

Justification: Once the core debris is quenched in-vessel, the production of steam and non- i condensible hydrogen will cease. Steam condensation in the drywell (in particular, if drywell sprays are actuated) will cause a return of non-condensibles and radioactivity from the torus airspace to the drywell. Since the details of the

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I Rev:@l 2 3 4 primary containment thermal-hydraulic conditions during the remainder of the thiny day dose calculation interval are not known precisely, it is reasonable to effectively consider a "one control volume" containment; i.e., a containment that is 1 well-mixed. This is consistent with current practice. '

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Assumption 3

Following the DBA (recirculation suction large break LOCA) the water mass remaining in the vessel is that corresponding to coolant at operating  :

conditions in the volume below the bottom of active fuel, depressurized at ,

constant enthalpy to atmospheric pressure with steam being released from the vessel.

L Justi6 cation: This assumption yields a conservatively small value for the water mass remaining in the bottom of the vessel after blowdown. All water above the bottom of the core is assumed to be removed at its operating state with no change in phase and ,

no liquid remaining Then, the remaining coolant is assumed to flash all the way i down to atmospheric pressure. In reality, coolant would flash throughout the vessel as the vessel depressurizes, leaving more liquid in the bottom of the vessel ,

then the above assumption would permit. Moreover, the coolant would only flasi  ;

down to a pressure corresponding to that of the containment which would be i greater than atmospheric pressure.

While it is true that the volume described above includes some of the annulus between the vessel and the lower shroud that would be drained by the recirculation break, it does not include the jet pumps up to their inlets and the corresponding volume within the core. Therefore, it is a conservative estimate of the volume that could remain water-filled with a recirculation suction line broken.

Assumption 4: In order to calculate the steaming rate from the core debris, it is assumed that the fraction of the core participating in the boil-off of the water mass remaining in the bottom of the vessel increases uniformly from zero at 1830 seconds (end of the gap release phase) to 50% at 7230 seconds (end of the in-vessel release phase).

Justification: This assumption is based in part on Assumption 1. At the end of the in-vessel release all of the core debris will be quenched, both that which has relocated to the lower pan of the vessel and that remaining in the original core region. For conservatism, the debris remaining in the core region is neglected in the calculation of the steaming rate during core degradation; only the assumed 50% of the core debris which relocates to the lower pan of the vessel and its interaction with the residual water (Assumption 3) is included in the quantification of the steam production during the in-vessel release phase.

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The exchange rate between the drywell and the torus is assumed to be constant during the release phase (up to 7230 seconds).

1 Justification:

j This assumption is slightly non-conseivative because it overestimates the removal j rate from the drywell early in the release phase. However, it does simplify the analysis; and for relatively low removal rates (of the order of one per hour) the

underestimate of the late removal compensates nearly completely for the overestimate of the early removal. A further demonstration of the adequacy of this

assumption is presented in Appendix A.

i Assumption 6:

The final core debris quench requires the time it takes minimum ECCS (one j core spray pump) to refill the core region, and it involves only the energy j stored in the one-half of the core debris assumed not to relocate to the lower part of the vessel. l Justification: i i Leaving one-half the core uncovered for a period of 7230 seconds (less the i

blowdown / core uncovery time) results in core debris left in the core region with l

significant stored energy. The restoration ofminimum ECCS will remove this j

stored energy at a rate determined by the coolant injection rate (drawn from the suppression pool) and the rising water level (reflood rate). To determine the

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reflood rate, the ECCS injection rate must be reduced by the rate of steam i

s production. The rate of steam production in this analysis corresponds to a low estimate of stored energy in only one-half of the core debris.

i Reference 5 indicates that the sweep-out rate corresponding to the final core debris quench would be expected to be of the order of 10 drywell volumes per hour.

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References j

Reference 1: Soffer, L., et al., " Accident Source Terms for Light-Water Nuclear Power Plants",

} NUREG-1465, Februrv 1995 4

Reference 2: DiNunno, J. J., et al., " Calculation ofDistance Factors for Power and Test Reactor

Sites", TID-14844, March 1962 i Reference 3: leaver, D. E. and Metcalf, J. E., " Generic Framework for Application ofRevised l Source Term to Operating Plants", EPRI Interim Report TR-105909, EPRI l Research Project 4080-2, November 1995 Reference 4: SECY-94-300, " Proposed Issuance ofFinal NUREG-1465, ' Accident Source Terms for Light-Water Nuclear Power Plants' ", December 15,1994 Reference 5: Leaver, D. E., et al., " Licensing Design Basis Source Term Update for the

Evolutionary Advanced Light Water Reactor", DOE /ID-10298, September,1990 s

l Reference 6: PSAT 04000U.03, " Design Data Base for Application of the Revised DBA Source j Term to the TVA Browns Ferry Nuclear Power Plant", Revision 0 j Reference 7: Babcock and Wilcox, Steam Its Generation and Use. New York,1963 i

Reference 8

McAdams, Heat Transmincion. McGraw-Hill, New York,1942 '

I Reference 9: NRC Generic Letter 88-20, " Individual Plant Examinations for Severe Accident  !

Vulnerabilities - 10CFR50.54(f)", November 23,1988 I

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Calculation i

Minimum mass of water remaining in va==al post-DBA blowdown

Reference 6 provides the following:

e Volume in-vessel, below BAF - 4100 ft' (Item 3.26) 4 e Reference pressure for determination of coolant mass - 1015 psia (Item 8.9) i e

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, Liquid specific volume at reference pressure - 0.02166 ft'Abm (Item 8.10)  !

The mass assumed to remain in the vessel prior to the flash is:

Mass = 4100 ft' / 0.02166 ft'Abm = 1.89E5 lbm j i

i The enthalpy for saturated water at 1015 psia =

4 hr @ 1000 psia + 0.15 (hr @ 1100 psia - hr @ 1000 psia) =

4 l- 542.4 BTUAbm + 0.15(15 BTUAbm) = 544.7 BTUAbm based on Exhibit I data from i

Reference 7

! Using this enthalpy, the fraction flashed to steam at constant enthalpy, x, can be determined from the following expression (evaluated using data from Exhibit 1):

x(h, @ 14.7 psia) + (1-x)(hr @ f4 7 Psia) = 544.7 BTUAbm 4  !

1150.4x + 180.1 - 180.1x = 544.7 l l

4 970.3x = 364.6 i l

x = 0.3 8  !

The minimum mass of water remaining in the vessel after the assumed flash is:

(1-x)(1.89ES lbm) = (0.62)(1.89ES) = 1.17E5 lbm which is based on Assumption 3.

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PSAT 04011H.01 Page: 8 of 19 Rev:@l 2 3 4 Exhibit 1 l

i TABLE 2. SATURATION: PRESSURES l e

Abe Press. Specific Volume Enthalpy Entropy Internal Energy Abs pr.ss.

tb Temp Sol. Set. Set. Set. Set. Set. Set. Set. tb Sq in. F Ugwid Veper Ligwid Eve , Yeper Llqvid Evep Yeper Liquid Evep Yeper Sq in.

p t vs se hr hvo he se sto so ut use p we 1.0 101.74 0.01614 333.6 69.70 1036.3 1106.0 0.1326 1 A456 1.9782 69.70 974.6 ' 1044.3 2.0 126.08 0.01623 173.73 1.0 >

93.99 1022.2 1116.2 0.1749 1.7451 1.9200 93.98 957.9 1051.9 2.0 3.0 141.48 0.01630 118.71 109.37 1013.2 1122.6 4.0 0.2006 1.6855 1.8863 109.36 947.3 1056.7 8.0 152.97 0.01636 90.63 120.86 1006.4 1127.3 0.2198 1.6427 1.8625 120.85 5.0 939.3 1060.2 4.0 162.24 0.01640 73.52 130.13 1001.0 1131.1 02347 1.6094 1.8441 130.12 933.0 1063.1 5.0 4.0 170.06 0.01645 61.98 - 137.06 996.2 1134.2 7.0 0.2472 1.5820 1.8292 137.94 927.5 1065.4 6.0 176.85 0.01649 53.64 144.76 992.1 1136.9 0.2581 15586 1.8167 144.74 922.7 1067.4 7.0 SA 182A6 0.01653 47.34 150.79 988E 1139.3 9.0 0.2674 1.5383 1.8057 150.77 918.4 1069.2 8.0 188.28 0.01656 42.40 156.22 985.2 1141.4 0.2759 1.5203 1.7962 156.19 914.6 1070.8 9.0 10 193.21 0.01659 38.42 161.17 982.1 1143.3 0.2835 1.5041 1.7876 161.14 911.1 1072.2 10 14.496' 212.00 0.01672 26.80 180.07 970.3 1150.4 0.3120 1.4446 1.7566 180.02 8973 1077.5 14.696 15 213.03 0.01672 26.29 181.11 969.7 1150.S 20 0.3135 1.4415 1.7549 181.06 896.7 1077.8 15 227.96 0.01683 20.089 196.16 960.1 1156.3 0.3356 1.3962 1.7319 196.10 885A 1081.9 20 30 250.33 0.01701 13.746 218.82 945.3 1164.1 40 0.3680 1.3313 1.6993 218.73 869.1 1087.8 30 50 287.25 0.01715 10.498 236.03 933.7 1169.7 0.3919 1.2844 1.6763 235.90 856.1 1092.0 40 281.01 0.01727 8.515 250.00 924.0 1174.1 0.4110 1.2474 1.6585 249.93 50 645.4. 1095.3 00 292.71 0.01738 7.175 262.09 9153 1177.6 70 0.4270 1.2168 1.6438 261.90 836.0 1097.9 60 302.92 0.01748 6.206 272.61 907.9 1180.6 0.4409 1.1906 1.6315 272.38 827A 1100.2 s

30 70 312.03 0.01757 5.472 282.02 901.1 1183.1 0.4531 1.1676 1.6207 281.76 820.3 1102.1 30 90 320.27 0.01766 100 4.896 290.56 894.7 1185.3 0.4641 1.1471 1.6112 290.27 813.4 1103.7 90 327.81 0.01774 4.432 198.40 888.8 .1187.2 0.4740 1.1286 1.6026 298.08 807.1 1105.2 100 l 120 341.25 0.01789 3.728 312.44 877.9 1190.4 140 0.4916 1.0002 1.5878 312.05 795.6 1107.6 120 353.02 0.01802 3.220 324.82 868.2 1193.0 0.5069 1.0682 1.5751 324.35 160 363.53 0.01815 785.2 1109.6 140 >

100 2.834 335.93 850.2 1195.1 0.5204 1.0436 1.5640 335.39 775.8 1111.2 140 373.06 0.01827 1.532 346.03 850.8 1196.9 03325 1.0217 1.5542 345.42 767.1 1112.5 180 200 381.79 0.01839 1.288 355.36 843.0 1198.4 0.5435 1.0018 1E453 354.68 759.0 1113.7 200 250 400.95 0.01865 1.8438 376.00 825.1 1201.1 0.5675 0.9588 1.5263 375.14 740.7 1115.8 250 300 417.33 0.01890 1 5433 393.84 800.0 1902.8 0.5879 0.9225 1.5104 392.79 350 724.3 1117.1 300 431.72 0.01913 1.3280 400.69 794.2 1203.9 0.0056 0.8910 1.4966 408.45 709.6 1118.0 350 400 444.59 0.0193 1.1613 424.0 780.5 1204.5 450 0.6214 OA630 1.4844 422.6 695.9 1118.5 400 456.28 0.0195- 1.0320 437.2 767.4 1904.6 0.6356 0.8378 1.4734 435.5 683.2 1118.7 450 500 467.01 0.0197 0.9278 449.4 755.0 1204.4 0.6487 0.8147 1.4634 447.6 671.0 1118.6 500 550 476.93 0.0199 0.8422 400.8 743.1 1203.9 000 0.6608 0.7934 1.4542 458.8 659.4 1118.2 550 486.21 0.0201 0.7698 471.6 731.6 1203.2 0.6720 0.7734 1.4454 469.4 648.3 1117.7 600 700 503.10 0.0205 0.6554 491.5 709.7 1201.2 800 0.6925 0.7371 1.4296 488.8 627.5 1116.3 700 518.23 0.0209 0.5687 509.7 688.9 1198.6 0.7108 0.7045 1.4153 506.6 607.8 1114.4 80C 900 531.98 0.0212 - 0.5006 526.6 668.8 1195.4 0.7275 0.6744 1.4020 523.1 589.0 1112.1 900 1000 544.61 0.0216 0.4456 542.4 649.4 1191.8 0.7430 0.6467 1.3897 538.4 1100 571.0 1109.4 1000 556.31 0.0220 0.4001 557.4 630.4 1187.8 0.7575 0.6205 1.3780 552.9 553.5 1106.4 1100 1200 567.22 0.0223 0.3619 571.7 611.7 1183 4 1800 0.7711 0.5956 1.3667- 566.7 536.3 ' 1103.0 1200 577.46 0.0227 0.3293 585.4 593.2 1178.6 0.7840 0.5719 1.3559 580.0 519.4 1099.4 1300 1400 587.10 0.0231, 0.3012 598.7 574.7 1173.4 0.7963 0.5491 1.3454 592.7 502.7 1095.4 1400 1600 596.23 0.0235 0.2765 611.6 556.3 1167.9 2000 0.8082 0.5269 1.3351 605.1 486.1 1091.2 1500 635A2 0.0257 0.1878 671.7 463.4 1135.1 0A619 0.4230 1.2849 662.2 403.4 1085.6 3000 2500 668.13 0.0287 0.1307 730.8 360A 1091.1 -

3000 0.9126 0.3197 1.2322 717.3 313.3 1030.6 2500 695.36 0.0346 0.0858 802.5 217J 1020.3 0.9731 0.1885 1.1615 783.4 189.3 972.7 3000 8306.2 705.40 0.0503 0.0503 902.7 0 902.7 1.0580 0 1.0580 872.9 0 872.g 3206.2 l

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- Minimum steaming rate for the remaininn water e

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Volumetric tiowrate (drvwell to torus) corresnonding to minimum steaming rates The two volumetric flows ofinterest can be determined assuming the drywell is steam-filled at 41.7 psia and near-saturation (based on Reference 6, Item 8.1). From Exhibit 1:

I v, = v, @ 40 psia - (1.7 psi /10 psi)(v, @ 40 psia - v, @ 50 psia) s v, = 10.5 -(1.7/10)(10.5 - 8.5) y, = 10.2 ft'/lbm Volumetric flow corresponding to 4.4 lbm/sec = 4.4(10.2) = 45 cfs (to be used from 1830 see to 7230 sec)

Volumetric flow corresponding to 31.9 lbm/sec = 31.9(10.2) = 325 cfs (to be used from 7230 see to 7890 sec)

For a drywell volume of 159000 ft' (Reference 6, Item 3.1) the quench flowrate of 325 cfs -

- corresponds to a drywell sweep-out rate of 7.4 per hour, comparing favorably with the 10 per hour rate given in Reference 5. This rate is sufficiently high to permit it to be used to characterize the *well-mixed" behavior of the containment beyond the core debris quench. '

A question that could be raised regarding the volumetric sweep-out rate is the effect of condensation in the drywell on the correspondence between the minimum sweep-out rate and the minimum steaming rates; i.e., could condensation decrease the sweep-out rate for a given steaming rate. The answer is two-fold. First, Appendix B discusses the fact that condensation would not be expected during core degradation because of heat-sink saturation

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' Rev:@l 2 3 4 during and immediately after blowdown. This explanation in Appendix B, howeve i

Related because it is not necessary to the defense of the position that neglecting condensa conservative. It is true that drywell condensation could decrease the sweep-out rate; but condensation also brings about diffusiophoretic removal of aerosol. Since the Reference I term is dominated by aerosol, this is an important effect. Ifone considers the expression for diffusiophoretic aerosol removal in Reference 3 (recognizing the drywell is steam-filled),

reduces to:

Removal rate = Steam Condensation Rate / Steam Density And this expression is the same as one would obtain for the volumetric sweep-out rate of the drywell if the steam generated in the drywell were flowing into the toms instead of the drywell. Therefore, the two phenomena are essentially equivalent; and as a matter radionuclide removal efficiency would be expected to be greater for diffusiophoretic de than for flow to the torus because of pool bypass and the difficulty of scrubbing small aeros Therefore, steam condensation in the drywell, to the small extent it may occur, can be Results The volumetric flows to be used for the exchange between the drywell and the torus are as follows:

l From t=0 to t=1830 seconds: Flow from drywell to torus = 0 (no source term for first 30 seconds, no steare.ing during gap release)

Flow from torus to drywell = 0 (no return flow dt-ing release Phase)

From t=1830 to t=7230 secs: Flow from drywell to toms = 45 cfs = 1.6ES cfh

' Flow from torus to drywell = 0 (no return flow during release 4 phase)

From t=7230 to t=7890 secs: Flow from drywell to torus = 325 cfs = 1.2E6 cfh Flow from torus to drywell = 0 (no return during core debris quench)

From t=7890 seconds to end: Flow from drywell to torus = 1.2E6 cfh (mixing flow- no

, scrubbing)

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Rev: Oh 3 4 Flow from torus to drywell = 1.2E6 cfh (mixing flow) I A comparison of these results to similar results for severe accident analyses of various sources is provided in Appendix C. It is useful to review these comparisons because these comparisons confirm the behavior discussed in this calculation. However, Appendix C is not Safety-Related because the results presented above do not depend on any of the Appendix C observations.

Conclusions The flow from the drywell to the torus during the core degradation is about one drywell volume per hour. This is comparable to other natural removal rates. This value, by itself, will decrease the average radioiodine concentration in the drywell during the core degradation by about a factor of 1.6 if referenced to the Reference I source term (without removal) or by about a factor of 3.0 if referenced to the Reference 2 source term. See Appendix A.

The flow from the drywell to the torus during the final core debris quench is about seven and a half drywell volumes per hour, but it only lasts for 11 minutes. The final core debris quench will decrease the radiciodine in the drywell by about a factor of four (i.e.,1/e##"*D.

i These effects combine with suppression pool scrubbing (of the flow from the drywell to the torus) s and with aerosol sedimentation to yield significant decontamination of the containment atmosphere.

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J Rev:@2 3 4 APPENDIX A APPENDIX TITLE:

"Use of a Uniform Sweep-Out Rate During the Release Phase" SAFETY-RELATED APPENDIX: Yes CALCULATION NUMBER: PSAT 0401IH.01 CALCULATION TITLE:

" Volumetric Flowrate as a Function of Time from Drywell to Torus (and Return)"

Purpose The purpose of this appendix is tojustify a uniform sweep-out rate from the drywell to the torus during the release phase from essentially t=0 to t=120 minutes.

Approach

• The approach is to set up a spread-sheet wherein:

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A release of 5% radioiodine is introduced over 30 minutes with no removal, and e

An additional 25% is added over 90 minutes using (1) no removal, (2) removal at a constant rate (" lambda") of one per hour, and (3) a linearly increasing removal rate beginning at zero and increasing to two per hour at the end of the 90 minutes.

The percent airborne is plotted and the integral under each of the curves is also calculated. The area under the curve (in %-minutes)is indicative of the release that would occur from the drywell for a constant leak rate and no decay. An assumption of no decay is acceptable since I-131 is the dominant radioiodine nuclide and it has a half-life of 8.1 days compared to the two-hour duration of this calculation.

Results The results are shown on Figure A-1. The accuracy of the spread-sheet can be checked by observing the slope of the calculation for any percent airborne. For example, for the increasing

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j PSAT 0401IH.01 Page:A2 ofA3 i Rev:@l 2 3 4 lambda case the maximum airbome percent (about 13.1%) is reached at about 84 minutes. At 84 minutes the variable removal rate would be:

O + 2 x (84 min -30 min)/ 90 min = 1.2 / hour I

The removal in terms of Whour would be: '

1.2 x 13.1 = 15.7 Whour = 0.261 %-min j

This is almost exactly the addition rate (0.278 %-n in) which explains the zero slope.

i As another example, the constant removal rate case ends with an increasing slope of about 0.3 W 6 min or 0.05 Wndn with an airborne percent of about 13.7%. The removal rate at this percent would be:

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1/ hour x 13.7% x 1/60 hours / minute = 0.228 Wmin l l The net increase would be:

i 0.278 Wmin (added)- 0.228 Wmin (removed) = 0.05 Wmin

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The results in terms of areas under the curves is shown on the figure. Note that the area under '

the constant removal curve is only 5% less than the area under the increasing removal curve. This i

shows that using a constant removal rate to approximate the increasing removal rate is acceptable, at least for the case oflimited removal (i.e., one per hour). A larger removal rate would increase this difference and make the constant removal rate approximation increasingly non-conservative.

l It is also of interest to note that either of the removal cases are about a factor of 1.6 better than the no-removal case.

1:

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4 4 . PSAT 04011H.01 Page: B1 ofB3

Rev: 0@ 3 4 l APPENDIX B i

APPENDIX TITLE-

. " Impacts of Transient Heat Conduction" i

SAFETY-RELATED APPENDIX: No l CALCULATION NUMBER PSAT 0401IH.01 1

i CALCULATION TITLE:

i s

" Volumetric Flowrate as a Function of Time from Drywell to Torus (and Return)"

Purpose The purpose of this appendix is to show (1) that the drywell shell is likely to saturate thermally well before significant 6ssion product release begins and (2) that tl e reactor vessel will still retain a signi6 cant amount ofsensible heat at the time the fission product release begins. The Srst finding supports the view that little condensation will be occuning in the drywell during core degradation and the second supports the view that neglecting sensible heat transfer from the '

l vessel shell is a significant conservatism when considering steam generation during core degradation and the associated purge flow from the drywell to the torus.

Approach  !

The approach involves estimating the equilibration time for transient heat transfer to the drywell shell and from the vessel shell and comparing those times to the start of the bulk of the fission product release to the containment (i.e., the start of the in-vessel release phase at t=30 minutes).

Exhibit 1, taken from " Principles of Heat Transfer" by Kreith, constitutes the basis for these estimates.

The drywell shell assumed data is as follows:

L = 0.125 A (assumed thickness of shell = 1.5 inches) 0 = 0.5 hours5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br /> (30 minutes - time before start of bulk of fission product release) a = 0.5 A2/ hour (thermal diffusivity for carbon steel) h = 100 BTU /A2-F-hr (typical steam condensation heat transfer coefficient when noncondensibles are present) k, = 26 BTU /A-F-hr (thermal conductivity for carbon steel)

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PSAT 04011H.01 Page

B2 ofB3 Rev: 0@2 3 4 The vessel shell data is assumed to be the same except L = 0.75 ft (9" thickness). The surface j heat transfer coef5cient of h = 100 BTU /hr-ft2-F is also representative of heat transfer from a I

surface to liquid water.

4 Results For the drywell shell, Bi = 0.5 and Fo = 16 at 30 minutes. From Exhibit 1, Q/Q,is essentially 4

unity indicating that all heat transfer that can occur (for a given tenperature difference) will have l . occurred by this time; i.e., the shell is thermally saturated. The shell would be 95% saturated by

] the time Fo = 8; i.e., by about 15 minutes.

For the vessel shell, Bi = 2.9 and Fo = 0.4 at 30 minutes. From Exhibit 1, Q/Q is about 0.5 indicating that about 50% of the sensible heat initially in the vessel shell remains at 30 minutes j with the other 50% having been transferred to the residual water. (Note that during the 30 l seconds or so of blowdown, the Fo would be less than 0.01 and virtually no sensible heat would ,

4 have been transferred). The 50% of the initial sensible heat transferred during the first 30 minutes t

' after blowdown, in terms of actual BTUs, can be estimated by assuming the weight of the portion of the vessel shell in contact with the residual water to be about 60 tons (half of the lower head).

Given this assumption, 50% of the original stored erargy (remembering that the outside is insulated) would be about 2 MBTU. If transferre:i over 30 minutes, the average heat transfer rate would be about 4 MBTU/hr or 1.2 Mw. This is comparable to the initial heat transfer rate '

calculated from the core debris at 30 minutes.

By 120 minutes (end ofthe fission product release to the containment) Fo would be 1.6 and the 50% remaining sensible heat would have been largely transferred to the residual water. If transferred uniformly over the 90 minute interval corresponding to the bulk of the Sssion product release, the transfer rate would be about 1.3 MBTU/ hour or 0.4 Mw. This is about 10 percent of the average heat transfer rate from the core debris assumed in the main calculation.

Based on the above, ignoring the contribution of the sensible heat stored in the lower head after blowdown is a significant conservatism. This heat would produce more than one megawatt of steaming during the first half hour (i.e., during the gap release when no steaming was assumed) and would add about 10 percent to the average steaming rate during the bulk of the fission product release.

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d [ > / b _ o,'b2 ~ "7 0.0 0.o M 2 5 2 5 2 5 2 5 2 5 2 5 E 0.0001 0.001 0.01 0.1 1.0 10.0 m E Biot's Modulus Bi= 5L[ks Fic. 4-9. Dimensionicss heat flow to or from a wall subjected to a sudden change in environmental temperature.

Figure 4-9 is a plot of Q/Q, vs. the Biot modulus for various values of Fo. Here Q represents the total change in internal energy per unit area, il i.e., the amount of heat transferred per unit area in the time interval be- oi 9"s tween 6 = 0 and 6 = 0 in Btu per square foot; Q represents the initial J "

internal energy per unit area relative to the fluid temperature T., i.e.,

} cpl (T - T ). A positive value of Q indicates, therefore, that heat is transferred from the wall to the fluid, while a negative value of Q shows that the direction of heat flow is into the slab.

I l i  !

PSAT 04011H.01 . Page:Cl ofCl1 Rev: 1 34

} APPENDIX C i APPENDIX TITLE:

i

" Comparison to Severe Accident Analyses" SAFETY-RELATED APPENDIX: No i

CALCULATION NUMBER PSAT 04011H.01 CALCULATION TITLE.

" Volumetric Flowrate as a Function of Time from Drywell to Torus (and Return)"

Purpose The purpose of this appendix is to present severe accident analyses done by Battelle Columbus  !

(an NRC contractor) and by TVA, itself, that add support to the estimates of accident progression and thermal-hydraulic behavior for the DBA LOCA that constitute the main part of this calculation.

\

Approach Two Battelle analyses have been done in which the initiating event is a large LOCA. The plant actually analyzed is Peach Bottom, but as can be seen on Exhibit 1 (3 pages) taken from Table 4.1-1 of the Browns Feny Individual Plant Ev=mination (IPE), Peach Bottom and Browns Feny are nearly identical. The Source Term Code Package (STCP) was used for these analyses. l The two Battelle analyses include a recirc suction LOCA with no injection (AE-y, where the y indicates a large, early containment failure) and an interfacing-system LOCA outside containment (so-called V-sequence which involves loss ofinjection, as well, because the line break outside containment knocks out the ECCS). These analyses are documented in BMI.2104 Volume II (July 1984) and BMI-2139 Volume 1 (NUREG/CR-4624, July 1986), respectively. Since in both cases the containment function is ammad to be lost either prior to or very early in the accident progression, it is not useful to look at the containment response. However, a comparison of overall event timing (to the assumptions used in the main part of this calculation for the DBA LOCA) and of primary system parameters is useful.

Several large LOCA analyses have also been made by TVA using MAAP3B. These include a recirc suction LOCA with no injection, the same event with recovery of ECCS injection prior to

PSAT 04011H.01 Page:C2 of C11 Rev: 0@2 3 4 vessel failure, and a main steamline LOCA (inside containment) with recovery of ECCS prior to vessel failure. For these analyses the overall timing is compared to the assumptions used in the main part of this calculation; and also, a detailed comparison of noble gas transport in containment is made to investigate the overall thermal-hydraulic behavior of the containment and to further support thc transport analyses and assumptions made in the main body of this calculation Results s

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a PSAT 0401IH.01 Page:C5 ofC11 Rev: 1 34 Exhibit 1, Sheet 1

[ Table 4.1 1 (Page 1 of 31. Basic RCS and Containment Comparison l Plant Name Peach Bottom Browns Ferry

• Type of Reactor BWR/4 BWR/4 Type of Containment Mark i Mark I Reactor Core Thermal Power (Mwt) 3,293 3,293 Number of Fuel Assemblies 764 764  :

Number of Control Rods 185 185 Resetor Vessel inside Diameter (inches) 251 251 Inside Height (feet) 72.92 72.92 Design Pressure (psig) 1,250 1,250 Number of Safety Valves 2 0 Lowest Safety Valve Setpoint (psig) 1,230 N/A Safety Valve Capacity (kib/hr) 925 N/A Safety Valves Vent To Drywell N/A Number of Relief Valves 11 13 Lowest Relief Valve Setpoint (psig) 1,105 1,105

Relief Valves Capacity (ktb/hr) 889 851 i Relief Valves Vent To Suppression Poo! Suppression Pool RHR System l

'O Number of Loops I 2 2 Number of Pumps 4 4 t; Flow Rate per Pump (gpm at,psid reactor 10,000 at 20 10,000 at 0

, . vessel to dryws!!) ~ ~

' Number of Heat Exchangers 4 4 ximum Capacity of Heat Exchanger (Btu /hr) 70,000,000 70,000,000 MR Service Water System

. ber of Pumps 3 8 w Rate per Pump (gpm) 4,666 4,500 C

umber of Pumps 1 1 pacity (gpm at psid) 616 at 1.120 616 at 1,120 ,

ber of Pumps 1 1 Rate per Pump (gpm at psid) 5,000 at 1,120 5,000 at 1,120

i l

l PSAT 0401 IH.01 Page:C6 fCl1 Rev: 34 Exhibit 1, Sheet 2 Table 4.11 (Pape 2 of 3). Basic RCS and Containment Comparison Table Plant Name Peach Bottom Browns Ferry Type of Reactor BWR/4 BWR/4 Type of Containment Mark l Mark I

l. m-LPCI (RHR)

Number of Divisions 2 2 .,

Number of Pumps per Division 2 2

Flow Rate per Pump (gpm at psid 10,000 at 20 10,000 at 0 l reactor to dry vessel)

Core Spray .

Number of Divisions 2 2 Number of Pumps per Division 2 2 l Flow Rate per Pump (gpm at psid) 3,125 at 122 3,125 at 105 (l ll Shutoff Head (psid) N/A - 400 Containment , ,, ,

1 Constructor CBI POM i Drywell Material and Construction Steel Steel i' Drywell Free Volume (ft3) 175,800 159,800 Drywell Design Temperature (*F) 281 281

. Torus Material and Construction: Steel Steel 3 123,000 126,200 TorusMaximum Torus Minimum Free Water Volume Volume (ft ) (ft )3 N/A 127,800 .

Torus Design Temperature (*F) 281 281 l Containment Design Pressure (psig) 56 56 l . Drywell to Torus Vent Configuration Diagonal large- Diagonal large-i diameter vertical diameter vertical

• ' ~, piping venting below piping venting below the water level of the water level of the pool, the pool.

l Drywell Spray (RHR)

! Number of Trains 2 2 -

Flow Rate per Pump (gpm at psid 10,000 at 20 10,000 at 0 reactor to dry vesse!)

(Amendment 8, FSAR) . -j

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Rev: Oh 3 4 Exhibit 1, Sheet 3 i

i Table 4.1 1 (Page 3 of 3). Basic RCS and Containment Comparison Table' '

Plant Name Peach Bottom Browns Ferry l

(type

{

of Reactor BWR/4 BWR/4 i fjType of Containment Mark I Mark I jsecondary Containment I

y Reactor Zone Free Volume below 1,122,000 1,360,000

• Refueling Floor (ft3)

Blowout Panel Design Pressure Hatch Cover (psid) N/A O.25 Refueling Floor (psid) 0.25 0.25 s.

Steam Tunnel (psid) 0.30 0.625 Standby Gas Treatment System Design Flow (Unit 2, CFM) N/A 4,660 Refueling Floor Area (three units)

Free Volume (ft3) 1,314,000 2,601,000 Blowout Panel Design Pressure (psid) N/A 0.35 -

Turbine Buildino Volume (ft3) 2,100,000 5,700,000

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