ML20206F313
| ML20206F313 | |
| Person / Time | |
|---|---|
| Site: | Cook, 05000000 |
| Issue date: | 02/04/1985 |
| From: | Gido R LOS ALAMOS NATIONAL LABORATORY |
| To: | Tinkler C Office of Nuclear Reactor Regulation |
| Shared Package | |
| ML20204G644 | List: |
| References | |
| CON-FIN-A-7280, RTR-NUREG-1150 Q-9-85-L-73, NUDOCS 8704140213 | |
| Download: ML20206F313 (341) | |
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pebruary 4, 1985 Los Alamos NationalLaboratory Q-9-85-L-73
= s'" K553 Los Alamos NewMexico87545 " " (505) 667-8097, FTS 843-8097 or (505) 667-3133, FTS 843-3133 dIY0 i
Mr. Charles G. Tinkler
- Program Technical Monitor and Project Manager l Containment Systems Branch Mail Stop P-1000 US Nuclear Regulatory Commission Washington, DC 20555 Dear Mr. Tinklart
SUBJECT:
Kevised FIN A7280 Letter Report Enclosed please find five (5) copies of the FIN A7280 program letter report
" Containment Hydrogen Analysis Review of D. C. Cook " revised in accordance with the consents made by R. Palla of the NRC.
Please contact me if there are any questions. Sincerely,
&/aA.Ak Richard G. Gido -
Safety Code Development RCGtwr Enct as cited Distributiont l R. L. Palla, Jr., NRC (w/ enc.) ! J. H. Mahaffy, Q-9 (w/ene.) CRM-4, MS A150 File (RGG)
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) . l e CONTAllNEAT HYDROGEN ANALYSIS RhVIEW OF D.C. COOK (Revised) (FIN A7280 Letter Report) by R. G. Gido l ! October 1984 (Revised January 1985) Energy Division , Los Alamos National Laboratory University of California ]
! Los Alamos NM 87545 i
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- 00hTENTS -
1 ABSTRACT................................................................
- 1. IhTRODUCT10N...................................................... 1
!!. COMPARE-H2........................................................ 2 III. INPUT M0 DEL....................................................... 2 IV. RESULTS........................................................... 3 V. CONCL'USIONS & RECO M ENDAT10NS..................................... 8 ACK N0k'L EDG EMEhT . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9 REFERENCES.............................................................. 9 APPENDIX A - COMPARE-H2 MODEL ......................................... 17 APPENDIX B - INSTRUMEl(T HEAT-SINK M0DELS............................... 46 APPENDIX C - REPRESENTATIVE PARAMETER TRANSIENTS. . . . . . . . . . . . . . . . . . . . . . . 53 APPENDIX D - DETAILED TIME VARIATION OF PRESSURE DIFFERENCES '
j ACROSS DOORS AND FANS..................................... 140 APPENDIX E - ICE-CONDENSER HYDROGEN CONCEhTRATIONS FOR CALCULATIONS NOS. 4 & 5...................'............... 177 i i l l i ( I 2 l l t , I ___________________.._,,=m.. -. _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - . _ _ __
ABSTRACT Presented are containment hydrogen-burn analyses in support of US NRC licensing activities on the D.C. Cook nuclear plant, which uses ar. I ice conder.ser. The ar.nlyses are primarily in the form of calculatiers , of degraded-core accidents using the hydrogen-burn version of the l COMPARE code. The calculations focused on the sensitivity of calcula- I tions to assumptions influencing the pressure differences (AP) across I i' the ice-condenser doors and the recirculation f a r.s . The contain-ment-atmosphere pressures and temperatures, and equipment temperatures were also of interest. Calculatior.s were performed for two versions of the S2D accident sequence,- which is defined as a small-break loss-of-coolant accident (LOCA) with failure of the emergency core cooling system (ECCS). Sensitivity studies were performed to investigate the effect of engireering safety features (ESF), burn parameters, fan failure, and the use of two spray drop sizes instead of just one size. A plant specific input deck and detailed description of the models used are provided. I. INTRODUCTION As a result of the experience derived from the TMI-2 accident, notably the generation and release of a large quantity of hydrogen to the containment. the NRC requires that certain reactor containments be equipped with a system to con-trol excessive amounts of hydrogen that may be generated. Because these re-quirements deal with accidents beyor.d the scope of normal design basis accidents. additional calculational methods had to be developed by industry and . the NRC. In that regard, the COMPARE3 code was modified so that cor.firmatory calculations could be performed to predict the containment atmosphere thermody-namic response during degraded-core accidents. The modified code, designated l l COMPARE-H2, has the capability of modeling hydregen combustion as well as models l for simulating various containment systems, e.g., sprays and far coolers. Earlier analyses using the COMPARE-H2 code were performed in support cf NRC licensing activities associated with the installation of igniter systems ir. the Sequoyah, McGuire and Watts Bar ice-condenser containments.*** j The objective of this report is to extend prior efforts by performing COMPARE-H2 analyses of degraded-core accidents for the D.C. Cook nuclear Plar.t in support of NRC licensing activities. The calculations focused on the ser.si-tivity of calculations to assumptions influencing the AP across the ice-conder.s-er doors and the recirculation fans. Also of interest were the contain- l 1 I
- r, - . _ _ - .* . - - . _ ._ - - .-- .
2-ment-atmosphere pressures and temperatures and instrument tesperatures. Calcu-lations used two different versions of the S D2accident sequer.ce. itt' (XMPARE-E2 The hydrogen-burn version of the COMPARE code used in this analysis evolved from the NRC subcompartment analysis code COMPARE.'*3 COMPARE has been evaluated (a) by comparison with test results,'~' (b) by comparison with other codes.'' and (c) analytically. In addition, COMPARE has been .used to predict the pressure-temperature response _ of ice-condenser' and large dry con-t ainments ' 3 ' " and to establish procedures for performing subcompartment analy-ses. COMPARE-H2, the hydrogen-burn version of COMPARE has been used to perform confirmatory analyses for the Sequoyah, McGuire and Watts Bar plants"* ar.d this report. Most of the features added to COMPARE to make COMPARE-H2 are described The 4 in Ref. 4. The radiation heat-transfer models are described in Ref. 17. code input requirements are described in Ref. 18. III. INPUT MODEL i The purpose of this section is to describe the COMPARE-H2 D.C. Cook 1 ice-condenser containment model used to estimate the AP across the ice-cor. der.ser doors and the recirculation fans, the containment-atmosphere pressures and ~ temperature response, and the equipment temperatures that might result from the j burning of hydrogen. A detailed discussion of the model, including a descrip-tion in the form of an input listing, is presented in App. A. - The models are I based primarily on the description of the models used by the utility for their CLASIX-code hydrogen-burn analyses.:o-23 In addition to the representatfor. of the containment configuration and containment safety systems (e.g., sprays), l i other significant input models include those that simulate the safety-equipmer.t thermal response and the input for the mass and energy release to the contain-t ment.
. Two different hydrogen, steam and energy release transients to simulate versions of a S D accident sequence were considered. The 52D seq'vence, a small
- 2 4
break LOCA with ECCS injection failure, is represented by the releases desig-l
.r, - - _ - - - - . _ . . n6 .. ._ ,_ _ _- S- 3; __ 1- .
- , _ . _ _ , . _ , . - - - _ _ _ , . . . - ~ _ _ _ - . _ _ . _ . - _ _ _ . _ _ _ _ _
nated 5 2D and S 2DA. The blowdown data are tabulated and plotted in App. B of Ref. 6. Equipment -heat -sink models, based on Refs. 24 and 25. are described in detail in App. B of this report.
, Figures 1 and 2 show the COMPARE-H2 model in terms of nodes (volumes),
flow paths (junctions), and door locations. The Fig. 2 model was used or.ly for Calculation No. $ where the ice condenser was represented by six modes. All other calculations used four nodes. Basic assumptions used in the code are that the modes are homoger.cous ar.d in thermodynamic equilibrium ar.d that the flow between nodes can be estimated with relatively simple flow models. Note that the homogeneous mixing assumptior. is ar. impcrtant requiremer.t to justify usir.g the code for analysis of distrib-uted ignit ion systems. However, homogeneity throughout the containment is r.ot necessary for the ignition system to be successf ul. The available evidence indicates that m.ixing processes in an ice condenser containment are adequate to allow simulation with a code such as 004 PARE-H2 if containment systems such as fans and sprays are operating. In particular. Ref. 26 establishes that the level of turbulence generated by the sprays is high enough to promote good mixir.g in the upper com.pa r t m.e r.t . The additier.al turbulence caused by fans, natural convection, and jets contributes to the well-mixed condition. Also, hydrogen mixing tests conducted at the Hanford En- l gineering Development Laboratory demonstrate reasonably good mixing in the lower coepartment of an ice-condenser plant that has forced recirculation because of the air-return fans. IV. RESULTS i The calculations performed are described in Table I and the calculated l l results are sumnarized in Tables 11 and III. Additional calculated results are ' presented in Apps. C through E. Appendix C presents, for Calculations No. 1, 2 and 7 (a) tabulations of the " events preceeding the occurrence of maximum AP and pressure levels and (b) plots of parameter variations with time. These calculations were chosen to represent approximately all of the calculations performed. The parameter plots
4 are for node concentrations, instrument temperatures, and code atmosphere pressures and tertperatures. Appendix D presents detailed plots of the AP across the doors and recir-
~
culating fans because these AP are of prikary interest for the D.C. Cook plart. The.. plots are preliminary because finer [ detail can be obtained. see App. D. However, the App. D information can be used to identify where detailed ir. forma-tion is needed and the calculations carl be rerun with the plot t i ng-inforrea-tion-saving specification modified. Appendix E presents ice-condenser hydrogen ccccentrations for Calculatiens Nos. 4 and 5, which lost the recirculating fans af ter the first upper plenum burn. Calculation No. 4 had four ice-conder.ser nodes and Calculation No. 5 had six nodes. The concentrations are of interest because sufficiently high concen-trations might result in detonations. The following subsections discuss the results for specific areas of ir.ter-l est. A. Blowdown Calculations Nos.1 and 2 used the 5 D2 and S DA 2 blowdowns, respectively. The blowdowns are described in Ref. 6. Tables 11 ar.d 111 show that the S)DA-blowdown calculation had a much higher dead-ended region to fan-accumulator room AP and a much higher upper plenum cable-in-conduit instrument ma xircum The higher AP resulted from the S 2DA calculation having n temperature. dead-ended region burn whereas the SD calculation had none. The higher 2 instrument temperature resulted from the S 2DA calculation having 71 upper plenum burns and 10 lower cortpartment burns whereas the S2 D calculation had 39 upper and zero lower compartment burns. Note that the high maxircum instrurtent tertperature is too high probab!v be- l cause the model used to represent the instrument did not account for the poor contact between the cable and the conduit, see Sec. G below. I l B. Engineered Safety Features - Calculation No. 3 (0.5ESF) represented the minittum engineered safety
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features (ESF) case, that is, one of two spray and one of two recirculation fan trains were used. Calculation No.1 (1.0ESF) had full ESF, that is, both spray and both fan trains were used. l l r - w ~ s --
1 I I i Both calculations had 10 lower compartment burns but the 0.5ESF calcula-tion had 29 upper plenum burns and the 1.0ESF calculation had 39 upper plenum burns. However. the total hydrogen burned was essentially the same f or both calculations. As a result, the amount of hydrogen consumed in each burn and the temperature and pressure levels were higher for the 0.5ESF calculation, see Table Ill. The pressure differences were similar for both calculations, see Table 11. The Sequoyah analysis' had a similar correlation in that the minimum. ESF case had more hydrogen consumed in each burn and higher temperature and pressure levels. C. Fans Off After First Upper Plenum Burn Calculations Nos. 4 and 5 simulated a fan failure af ter the first upper plenum burn with Calculation No. 4 having four ice-condenser nodes and Calcula-tion No. 5 having 6 nodes. The number of nodes was varied to study the sensi-tivity of the calculated results to the noding. The result of primary interest in these calculations is the maximum hydrogen concentration in the ice condenser, see App. E. The calculated results show that (a) a hydrogen concentration of 17 % is reached and (b) the six-node model resulted in slightly higher concentrations than the four-node model. Use of the six ice-condenser node model is recommended for additional analyses that investigate hydrogen concentration in the ice condenser. D. No Upper Plenum Burn Initiation This calculatior. (No. 6) was done to establish possible maximum door and fan AP. Unfortunately, the calculation " bombed" at 5330 s. the time of the first (and only) dead-ended region burn. However, the limited results confirm that high AP will result if upper plenum burn initiation is prevented, see Table II. Note that the high AP resulted probably from use of the basic burn param-eters (Table A-X) except for the prevention of upper plenum burn initiation. As a result, burning was initiated in the upper compartment (a large volume with a large amount of hydrogen) at a high concentration (8 % with a burn fraction of 85 %). This observation is confirmed by the calculation discussed in the next section. i l i
l l l l E.' Burr Initiation Only in Upper Compartment The purpose of this calculation (No. 7) is to use a combinatior. of burr. parameters that might be considered mechanistic with some cor.servatise to produce high pressure levels, which might produce maxir.um AP. This approach was used in the Sequoyah analysis' to produce possible umximum pressures. However. for the D.C. Cook analysis, the major results of interest are the ice-cordenser door ar.d fan AP. The burn-parameter combination used is (a) burn initiation in the upper i.e., burr.i r.g cor.partment at 5 %, (b) upper compartment burn fraction of 40 t, to a minimum of - 1.2 %, (c) burn initiation by propagation in other nodes at 4.1 %, and (d) burn fraction in other nodes of 100 %, i.e., burnir.g to a mir.irum of - 0 S. The mechanistic aspects of the burn-parameter combination are that the best-estimate burn-initiation concentrations ar.d burn fractior.s are low. see Refs. 4 and 26. For example, burn initiation in the upper compartment is estimated to occur at < 4.5 4 with a burn fractior. of < 10 5. Conservative aspects of this burn-parameter combination are that (a) burn initiation is allowed to occur only in the u'pper plenum, which has the largest
- volume and therefore the most hydrogen to burn, (b) burn initiatier. in the upper plenum is at 5 %, which is higher than the expected value of < 4.5 t, (c) an upper compartment burn fraction of 40 % is used, which is higher than the expected value of 10 % (for initiation at < 4.5 %), and (d) a burn fraction of 100 % is used for the hydrogen initiated to burn by propagation, which could be much lower, e.g., only 10 % would burn if initiation was at 4.5 5.
Calculation No. 7 resulted in high peak pressures, temperatures ar.d AP, see Tables II and III. F. Two-Drop-Sizes Calculation No. 8 used two drop sizes of 309 and 810 um whereas Calcula-tion No. 1 used one drop size of 710 um. The results of Table II and III show little difference between the calculated results for the two calculations. fj i
'l
7-G. Cable-In-Conduit Conductior. Model The simple model used for most of the calculations assumed that the cable is in good and . complete contact with the conduit, see App. B. Actusily. the contact is limited and the energy transfer to the cable would not be very ef f ec-tive, see App. C of Ref. 6. Reference 6 also investigated approximately the effect of the poor con-tact. The results indicate that the cable temperatures would be much lower for the more realistic contact, The reduced contact simulation was very approximate and the geometry should be analyzed in two dimensior.s, e.g. , by usir.g the AYER conduction code.2' H. Pressure Differences vs Time Results of primary interest in this ar.alysis are the pressure differerces ( AP) across the ice-conder.ser doors and fans. Table 11 preser.ts the maxinum AP and their time of occurrence. In additior., App. D preser.ts detailed plots of the AP variation with time near the tin.e of occurrence of the maximum. AP. The detailed variation might be useful to evaluate the mechar.ical loadirg of the ice-condenser doors and the fans.
. The App. D plots are not as detailed as could have been obtained because the AP information was not saved for plotting at every time step calculated so that data-storage limitations were not exceeded. However, the App. D plots -
could be used to establish which detailed plots are of interest and the calcula-tion could be rerun over the relatively short time interval of ir.terest with the plotting information saved in greater detail. I. Ir.strument Temperatures The instrument temperatures for this analysis. see Table III. are nuch lower than those obtained in the Watts Bar analysis' because D.C. Cook plant has sprays in the upper compartment, lower compartment and the fan-accumulator room whereas the Watts Bar plant has sprays in the upper compartment only. Based on the analyses of thisf report. the instrue.ent temperatures for the D.C. Cook plant do not exceed the temperature levels of concern (~ 325 f). l l i ! + '/
.* - m e _ __
.g.
V. CONCl.USIONS AND REC 041ENTMTIONS The D.C. Cook analyses for this report resulted in the followirg cerclu-sions and recommendations. Note that emphasis is giver. to the calculated pressure differences ( AP) because these are of primary concerr. for the D.C. Cock plant.
- 1. A higher dead-ended region to fan-accumulator room AP occurs with the S2 DA blowdown than with the S D2 blowdown because the S DA-blewdowr 2 calculation had a dead-ended region burn and the S D-blowdown y calculation did r.ot.
- 2. The use of minimum engineered safety features (ESF) results in similar AP. slightly higher maximum pressures and slightly higher maximum temperatures than when maximum ESF is used. The peak contair.mer.t pressure calculated for these cases is - 27 psia.
- 3. Failure of the recirculation far.s after the first upper pier.um can re-sult in the buildup of hydrogen concentration in the ice condenser to a level of - 17 %.
- 4. Initiation of burns in the upper compartment results ir. high AP across ice-conder.ser doors and fan components because the upper compartner.t is large and, therefore, contains a large amount of hydrogen that could be burned. Upper compartmer.t burnir.g resulted from use of the following burn-parameter combinations:
A. Basic utility burn parameters with the prevention of burn initiation in the upper plenum. B. Mechanistic burn-parameter combination with some conservatise to produce maximum pressures.
- 5. Use of two instead of one drop size had a small effect or. the AP ar.d other aspects of the transient, such as. the number of burns and the peak pressure and temperature.
- 6. Instrument temperatures do not reach maximums of concern because there are sprays in several compartments.
- 7. Consideration should be given to use of a two-dimensional conductior.
code to estimate the cable-in-conduit instrument temperature response when instrument temperatures based on onedimensional models are foer.d to be unacceptsbly high. i I y , --g,-- - - - - - - - - - - .- _. _ m
ACKNOWLEDGEMENTS Greatly appreciated are the guidance and critical evaluations of
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C.' G. Tinkler and R. L. Palla, Jr. Systems Branch, U.S. Nuclear Contfainment Regulatory Commission. . i REFERENCES
- 1. R. G. Gido C. I. Grimes, R. G. Lawton, and J. A. Kudrick, " COMPARE: A Computer Program for the Transient Calculation of a System of Volumes Conr.ected by Flowing Vents," Los Alamos Scientific Laboratory repert LA-NUREG-64SS-MS (September 1976).
J. S. Gilbert, R. G. Lawton, and W. L. Jensen,
- 2. R. G. Gido,
" COMPARE-MOD 1: A Code for Trar.sient Ar.alysis of Volumes with Heat Sinks. Flowing Vents, and Doors " Los Alamos Scientific Laboratory re-port LA-7199-MS (March 1978).
- 3. R. G. Gido G. J. E. Willcutt, Jr., J. L. Lunsford, and J. S. Gilbert.
" COMPARE-MOD 1 code Addendum." Los Alamos Scier.tific Laboratory Report NUREG/CR-1185, LA-7199-MS, Addendum 1 ( August 1979).
- 4. R. G. Gido, and A. Koestel, " Hydrogen Burn Analyses of Ice-Condenser Containmer.ts." Los Alamos Natioral Labora tory report NUREG/CR-3276.
LA-9749-MS (July 1983).
" Safety Evaluation Report Related to the Operation of Sequoyah Nuclear 5.
Plant. Units 1 and 2," US Nuclear Regulatory Comission report 4 NUREG-0011, Supplement No. 6 Docket Nos. 50-328 (December 1982).
- 6. Attachment titled " Containment Hydrogen Analysis Review of Watts Bar,"
to letter Q-7-84-376 to Charles G. Charles Tinkler. U.S. Nuclear Regulatory Comission, from Richard G. Gido. Los Alamos (August 1, 1954).
- 7. R. G. Gido, J. S. Gilbert and W. L. Jensen, " Containment Ice-Condenser Analysis Using the COMPARE code." ANS Topical Meeting on Thermal Reac-tor Safety, pp. 3-719 through 3-732 Sun Valley, ID (July 30-August 4. I 1977). l
- 8. W. S. Gregory J. R. Campbell, R. G. Gido, and A. J. Webb. "Compariser of COMPARE /RELAP3 Subcompartment Calculations with Battelle-Frar.kfurt
. C-Series Test Results " Los Alamos National Laboratory report %
NUREG/CR-2177, LA-8866-MS (June 1981). '; R. G. Gido, W. S. Gregory, P. E. Littleton, and
- 9. J. W. Bolstad, G. J. E. Willcutt, J r. , " Comparison of COMPARE M00-1 Subcomparteer.t ,
- Calculations with Battelle-Frankfurt Containment Tests," Los Alamos ?
e
- - , _ _ _ - - - - - - % - - - . Af --: a
Laboratory report NUREG/CR-1817, LA-8615-MS (Decem-Scientific ber 1982).
- 10. E. S. Idar, J. F. Livre, and R. G. Gido. " Comparison of COMPARE ard ,
BEACON Subcompartroent Analyses of Battelle-Frankfurt Cor t a i r.rre r t Tests," Los Alartos National Laboratory report NUREG/CR-2849 LA-9461-MS (January 1983). E. S. Idar, .R. G. Gido, J. F. Lime, and A. Koestel,
- 11. M. W. Burkett.
"Cortparison of COMPARE and BEACON Reactor-Cavity Analysis." Los Alarnos National Laboratory report NUREG/CR-3305, LA-9776-MS (April 1964).
! 12. R. G. Gido and A. Koestel, " COMPARE Contair.rtent Subcortpartr.er.t Ar.aly-sis Code Evaluation " in Proceedings of the International Meeting on' 3 Thermal Nuclear Reactor Safety, Chicago, IL ( August 29 - Septertber 2. 1982) US Nuclear Regulatory Commission . report NUREG/CP-0027 (February 1983) Vol. 3, pp. 1583-1598.
- 13. D. E. Lamkin, A. Koestel, R. G. Gido, and P. W. Baranowsky, "Contain-etent Main Steam Line Break Analysis for Equipeter.t Qualification." Los Alamos Scier.tific Laboratory report NUREG/CR-1511. LA-6305-MS
- (June 1980).
- 14. R. Gido, D. Larckin, and A. Koestel, " Mechanistic Dry-Pressure-Contain-4 inent LOCA Analysis," Los Alamos National Laboratory report NUREG/CR-2548, LA-9460-MS (July 1982).
l
- 15. R. G. Gido, J. S. Gilbert. and C. G. Tir.kler, "Subcortpartreent Aralysis
' Procedures " Los Alartos Scientific Laboratury report LA-6169-MS (De-certber 1979).
- 16. W. V. Turk, R. G. Gido, and C. Y. Li, "Containcent Reactor Cavity Sub-l cortpartreent Analysis Procedures for a Boiling Water Reactor," Los 4 Alarcos National Laboratory report NUREG/CR-2633, LA-9277-MS (May 1982).
! 17. At tachttent titled " Radiation Heat-Transfer Modification to the Hydrogen-Burn Version of COMPARE" to letter Q-7-83-533 to Charles G. Tinkler, U.S. Nuclear, Regulatory Commission, frort Richard G. Gide. i Los Alamos (Noven:ber 21,$1983).
?
- 18. Attachinent titled " Input Description for the Hydrogen-Burn Version of.
i
- )
COMPARE" to letter 0-7-84-214 to Charles G. Charles Tinkler. , U.S. Nuclear Regulatory Comrtission, from Richard G. Gido, Los Alartes
- (May 2, 1984). , > 19. "The CLASIX Computer Program for the Analysis of Reactor Plant Con-tainment Response to Hydrogen Release and Deflagration," Offshore Power Syst ercs Report No. OPS-36a 31 (non proprietary): OPS-07135 4 (proprietary) (October 1981).
J e . -- - - . .
- 20. At tachment s No. 1 and 2 of Letter AEP:NRr:00540E (Docket Nos. 50-315 and 50-316) to Mr. Harold R. Denton, US Nuclear Regulatory Comiwior from R. S. Hunter Indiana & Michigan Electric Co. (July 2, 1951).
- 21. Attachments No. 1, 2 and 3 of Letter AEP:NRC:00500H fl h ket Nos. 50-315 and 50-316) to Mr. Harold R. Denton, US Nuclear Regulatory Comission from R. S. Hunter, Indiana & Michigan Electric (September 30, 1961).
Co.
- 22. Attachments No. 1 and 2 of Letter AEP:NRC:00$00L (Docket Nos. 50-315 and 50-316) to Mr. Harold R. Denton, US Nuclear Regulatory Comissior.
from R. S. Hunter, Indiana & Michigan Electric Co. (December 17, 1962).
- 23. At tachments No. 1 and 3 of Letter AEP:NRC:00500M (Docket Nos. 50-315 and 50-316) to Mr. Harold R. Der. ton, US Nuclear Regulatory Comission from M. P. Alexich, Indiana & Michigan Electric Co. (March 30,1964 J.
- 24. Attachment 4 of letter to Director of Nuclear Reactor Regulation Atter. tion: Ms. E. AJensam, Chief. Licensing Branch No. 4. Divistor of Licensing, U.S. Nuclear Regulatory Comission. Washingter., DC 20555.
from L. M Mills, Mar.ager, Nuclear Licensing. Ter.r.e s s e e Valley Authority (December 1, 1981).
- 25. Personal comunication. C. G. Tir.kler, Containmer.t Systems Brar.ch, U. S. Nuclear Regulatory Comission (May 29, 1964).
- 26. R. G. Gido, and A. Koestel, "Pararr.eters for Containment Hydrogen-Burn Analysis," Joint ANS/ASME Conference er. Design Construction and Opera- .
tion of Nuclear Power Plants, Portland, OR (August 5-8, 1984).
- 27. G. R. Bloom, L. D. Muhlestein, A. K. Postma, and S. W. Claybrook.
" Hydrogen Mixing and Distribution in Containment Atmospheres."
Westinghouse Hanford Co. report for Electric Power Research Institute EPRI NP-2669 (March 1983).
- 25. Interim Staff Positior. on Environmental Qualificatior. of Safety-Related Electrical Equipment," U.S. Nuclear Regulatory Comis-sion report NUREG-0588, Rev. 1 (July 1961).
- 29. R. G. Lawton, "The AYER Heat Conduction Cortputer Program," Los Alamos Scientific Laboratory report LA-5613-MS (May 1974).
, 30. Personal comunication, R. L. Palla, Containment Systems Brar.ch. U.S.
Nuclear Regulatory Comission (June,1982).
- 31. R. G. Gido, " Liner-Concrete Heat Transfer Study for Nuclear Power Plant Containments," Los Al arr.os Scientific Laboratory report LA-7089-MS (January 1978).
L
TABLE I D.C. COOK CALCULATION SUWARY
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CIlc1984 Time No. Date ID BD Max. s Comrrents 9/23 C1 S D base case w max ESF, utility burn parameters. 1 S)D 6500 2 2 9/23 C2 S)DA 6500 S)DA base case w max ESF, utility burn para-meters. 3 9/29 C3 Calc No. 1 w min (0,5 x max) ESF, S)D 6500 4 9/29 C4 S)D 5000 Calc No. 1 w fans off after 1st UP burn. 5 9/29 C4A SD 2 5500 Same as No. 4 w 6 ice codes. 6 10/8 C5 SD 5330 Same as No. I w UP burn initiation prever.ted to 2 obtain possible max AP's across fans and doors. 7 10/8 C6A S)D 6500 Initiation of burning in UC only with IV/0=5. MV/0-3, propagation w PV/u=4.1, MV/u-o.oul. 8 10/8 C11 S)D 6000 No. I w 2 drop sizes, 4 , NOTES - Abbreviations used include BD = Blowdown. EST a Engineered Safety features. ID = identifier, IV/0 m hydrogen concentration (percent) for burn initiation. ma x a ma x i r.urt , min a scir.imum, MV/O a minim.um hydrogen concentration (percent) for burning, PV/0 m hydrogen concentration (percent) for propagation of burning, AP a pressure difference, UC a Upper Compartment, UP a Upper Plenum, w a with. All calculations used NUREG-0588" guidelines for heat transfer and condensed-mass removal, which included the use of the saturation tertpera t ure ar.d a revaporization fraction of 8%. Heat transfer by thermal radiation used in all calculations. 1
- - - . ._ r-.--- - __ _ __ _ _ _ . -
TABLE II D.C. COOK DOOR AND FAN RO(IMUM PRESSURE DIFFERENCES
/ TIME OT OCCURRENCE
- - - - - Doors - - - - - - - -
LID TDD - - - - - - - Fan - - - - - - - Calc IDD LC to FA No. LP to LC UP to 11 UC to UP UC to FA DE to FA 1 +1.4/4974 +1.9/4975 +0.8/4576 +1.9/4975 +0.5/4645 ' +0.7/4643
-1.3/4575 -1.4/5586 -0.9/5586 -0.7/4972
-0.9/4643 -0.3/5820
~ 2 +1.4/4476 +2.6/4474 +2.3/4474 +2.5/5818 +10.6/5816 +0.8/5818
-3.6/4474 -1.9/5816 -0.6/5616 -0.9/5617
-1.8/5517 -1.1/4595
+0.8/4758 +1.5/6210 +0.5/5039 +0.8/5459 3 +0.9/5986 +1.2/48S1 -0.7/4504
-1.0/5459 -0.3/4660 -1.2/4757 -0.8/4756 -1.4/5697
+0.8/4576 +3.2/4965 +0.9/4949 +0.8/4949 4 +3.2/4966 +1.3/4955 -1.4/4947
-0.6/4824 -0.3/4590 -1.3/4575 -1.6/4947 -1.6/4947 5 +4.5/5119 +2.1/4969 +1.0/5463 +4.7/5119 +9.3/5435 +1.1/5463
-1.0/5463 -0.3/4966 -1.3/4574 -1.8/5463 -1.7/5094 -1.6/5094 6 +4.7/5330 +S.5/5327 +2.5/5323 +12.8/5330 +0.5/4630 +0.7/4970
-0.9/5253 -0.3/5113 -1.1/5323 -1.4/5113 -0.6/5113 -0.6/5234 1
7 +4.8/5019 +13.0/5017 +1.5/5007 +7.1/5020 +7.4/5014 +1.6/5020 i
-4.7/4735 -1.8/4735 -2.2/5006 -1.4/5317 -1.3/4727 -2.1/5020 .
8 +1.2/5840 +1.5/5312 +0.8/4582 +1.5/5312 +0.5/4877 +0.8/5309
-0.9/5591 -0.3/4951 -1.3/4581 -1.3/5190 -0.8/5190 -0.7/5191 NOTES -
Abbreviations used include DE = Dead-Ended region. ESF = Engineered Safety Features. ID = Identifier. FA a Fan-Accumulator room. Il a Top node of ice. which is below the IDD. IDD = Intermediate Deck Doors. LID = Lower Compartment. Ir.let Doors. l TDD a Top Deck Doors. LC = Lower Compartment. UC = Upper UP a Upper Plenum.
TABLE III D. C. COOK CALCULATED RESULTS Comp Instrument Max T. F Comp Max Calc Max T. F - - - - - Heat Sink No - - - - - P. psia Ice Left No UP/LC 20 21 22 23 24_ 25 26 Low /Hirh 4/ Tine, s Notes 1 1362/ 56 212 225 271 226 225 225 25.0/ 43/ - 949 25.8 6500 2 1379/ 64 213 226 380 226 226 226 24.6/ 35/ A 338 25.5 6500 3 1485/ 57 238 254 277 273 232 241 26.8/ 44/ - 549 27.0 65(x) 4 1179/ 42 212 225 76 228 225 225 26.8/ 56/ - 258 27.0 5000 5 1543/ 46 219 238 166 261 225 227 25.2/ 53/ - 1238 34.3 5500 6 1356/ 40 211 226 57 226 226 226 28.6/ NA B 937 41.5 7 1698/ 46 212 225 107 226 225 225 32.0/ 35/ - 1114 38.0 6500 8 1365/ 52 212 225 230 226 225 225 24.7/ 46/ - 945 25.0 6000 NOTES -
- 1. Abbreviatior.s used include Calc a Calculation. Conp a Compartrent.
LC a Lower Compartment. Max e Maximum. Min a Minimum. NA a Not Availa-ble. P = Pressure. T = Temperature, UP = Upper Plenum.
- 2. Description of instrument heat sinks: 20 m Igniter assembly ir. UP:
21 m Barton transmitter casing in LC; 22 m Cable in conduit in LC: 23 m Cable in conduit in UP; 24 a 0.125 inch aluminum in LC: 25 m 0.667 inch aluminum in LC; 26 m 0.250 inch steel in LC. A. S2DA blowdown. B. Calculation " bombed" at 5330 s. l
VOLUME NO. (TYP CAL) NY 4 Nv 3 S UC
;9 NV7 Il e
i 13 NV 6 NV9 It FA . . ,, JUNCTION iy NO (TYPICAL) 7 g l NV to ! (9) 8 *
- 8 y5 IN30ATES DOOR 1 15 (TYR)
I f v ~ I4 2 yy , LC 1l NV2 LF 1 LTD Fig. 1. D.C. Cook four-ice-condenser-node containment model used for all but Calculation No. 5. Shown are the nodes (volumes) are: and flow paths (junctions). The node designations region; FA E Fan-Accumulator room; DE E Dead-Ended 11 E Ice-Condenser compartment first 1/4 from the top, 12 = next 1/4 from the top, etc; LC E Lower Compartment; LP E Lower Plenum; UC E Upper Compartment; UP E Upper Plenum. Doors shown are IDD E Intermediate Deck Doors, LID E Lower Inlet Doors, and TDD E Top Deck doors, with TDD locking open when they reach full open area and the other doors allowed to open and close. m _
WOLUME No. (TYPICAL)
- j
- NV 4
- yy y UC 5 ,
69 8 NV 7 Il s 4 i 13 N V d, yy 9 U It
-(as T JUN
- TION NO.(TYPICAL) !
p 7 (g) g5 IND*JTES DOOR 3 lE (TYP) ' s W it . ig
- 14 n NV l l
LC
/ l N v it Ny t L' '
16 l LIO Fig. 2. model used only D.C. Cook six-ice-condenser-node containment for Calculation No. 5. Shown arenode the nodes (volumes) and designations are: flow paths (junctions). The room; DE E Dead-Ended region; FA E Fan-Accumulator 11 5 Ice-Condenser compartment first 1/6 from the top, 1 12 5 next 1/6 from the top, etc; LC E Lower Compartment; LP I Lower Plenum; UC E Upper Compartment; UP = Upper Plenum. ; Doors shown are IDD E Intermediate Deck Doors, LID E Lower ' Inlet Doors, and TDD E Top Deck doors, with TDD locking open when they reach full open area and the other doors allowed to open and close.
, . -c. _....- . -,, ..- - . , ~ _ --,.~..__..,_,-,_~.._.-,_~_.._%,u.-_. -
, , . g . _ . . - _ _ , _ _ _ _ . _ _
1 APPI'NDlX A COMPARE-H2 MODEL The purpose of this appendix is to describe models used for the calcula-tions performed. A. Sample Ir.put Figure A.1 is the input deck used fc.r Calculatier. No. I ard is representative of the input decks used for all calculations. Understarding of the input details requires analysis of the parar.eter values by comparison with the ir.put description provided in Ref. 18. B. Time Steps Stable calculatier.s resulted for the Watts Bar* calculatior.s if the ratio of the ir.ertial term (flow ter.gth/ flow area = L/A) was greater thor. five (f t). As a result, large time steps could be used if large L/A were used. However, the high ir.ertial term might affect the result it that -aridly occurring pher.er.- er.a would not be represer.ted. An acceptable and efficient cor.bination of tire step ar.d L/A is determined usually by experimer.tation. Checks should be made te provide assurance that the time steps are adequate to accour.t for the physi al pher.cmenL of ir.terest, which was dor.e for the Watts Bar analysis.' r
~
The approach described above was used for the D.C. Cook analysis because of its similarity to the Watts Bar cor.tainmer.t. Specifically. -time steps of 0.020 s with L/A of 0.5 - 10. ft were used until just before burr.ir.g started. Then, actual L/A (0.005 - 1. ft) and time steps of 0.0025 - 0.005 s were ir.troduced interactively ar.d used for the remainder of the calculatior.. i . e. . when burning is occurring. C. Volume Specificatior. The volur.e (node) specification is described in Table A-1, includir.g ini-tial pressures, terperatures and component masses used as initial cor.ditior.s for the calculations. These conditions were based on the initial conditions speci-fied in Ref. 20 except for the 40 F ice-conder.ser temperatures, which are 32"I' in the reference. The use of 32"F resulted in r.ur.erical instabilities because water properties have a large change at this ter.perature. . 1
- Figures 1 and 2 show how the nodes cormr.unicate with each other through junctions. Figure 1 was used for all but Calculation No. 5 and the remairder of the model discussion is for the Fig. I configuration. Figure 2 was used or.ly f0 -Calculation .No. 5 and input pararreters for this configuration are readily determined from the model discussion for the Fig. I cor. figuration.
D. junction Specifications The junctions shown in Fig. I are described in Tables A-II through A V. Most of the junction details were obtained from Refs. 20-23. However. the ice-condenser flow areas were obtained from Ref. 30 ar.d some loss coefficier.ts were estimated. especially those in the direction opposite to that caused by the f a r.. E. Bl owd owr. The Sy D and Sy DA blowdowns used for these analyses are described ir. App. B of Ref. 6. The Sy D blowdown is specified also ir. Ref. 23. F. Heat Sinks The heat sir.ks are described in Tables A-VI through A-VIII. Except for the instrument and some detail of the ice-condenser heat sinks. this informatier. is based or. Refs. 20-23. Reference 31 was used to provide guidance for the cor.- crete depth and number of finite-element increments used to represent the heat sir.ks. The instrument heat-sink models are described ir. App. B. G. Bourdar9 Conditiors . Those used ir.clude:
- 1. Energy transfer to the ice is based on (a) an empirical heat-transfer coefficient. (b) the difference between the atmosphere and ice surface temperatures ar.d (c) no cor.densed mass removal. which is assumed to be er.trair.ed by the flow over the ice.
- 2. Radiation with the sinks lumped together and represented as a blackbody enclosure.
2
- 3. The Uchida condensing heat-transfer correlation as described ir. Refs.
ar.d 18 with additier.al exp,lginatior. giver. below. Energy transfer based on the Uchida condensing heat-transfer correlatier. was determined with the temperiiture difference AT = T, - T, or AT = T, - Ty . re-where T,. T, and T, are the saturation, atmosphere nr.d wall temperatures, spectively. The saturation terr.perature must exceed the wall temperature for the d i
- ]
- s. J
l condensir.g coefficient to be applied. If T, > T, a heat-transfer coefficier.t of 2.0 Btu /(hr f t: F) and AT = T, - T, are used. This procedure follows the guidelines of Ref. 26 in that the conder. sing heat flux qc = he (T, - T,) is replaced by a natural cor.vec t i or, heat flux qn=hn (T, - T ,) when qn *****d5 9c. where h e is the Uchida conder. sing heat-transfer coefficient and h, is the user-specified natural-convectien i heat-t ransfer coefficient, for which a value of 2.0 Btu /(hr f t *F) was used. This is done in recognition of energy transfer continuing because T, exceeds T, even though AT = T, - T,goes to zero as T, approaches T,. The condensed mass removal associated with the ' condensed energy removal depended on whether T, or T, was used and the user-specified value for the revaporizatior. f raction. Rev Frac. Rev Frac re'preser.t s the f rac t i or. of cor.- densed mass that is assumed to vaporize and returr. to the atmosphere. For the calculations of this report, a Rev Frac of 0.08 (84) was used as recommer.ded ir. Ref. 28. H. Radiatior Parameters Radiation was modeled for all of the calculations based on the black-
- body-enclosure approach described in Ref. 17. Specific input details for the model are presented in Table A-IX. These details were obtained from Ref. 23.
I. Sprav Parameters Spray-system parameters based on Ref. 20 are presented in Table A-X. The fall heights were estimated from the drop fall time given in the referer.ce and an assumed t e rm.i na l velocity of 500 ft/ min, which is represer.tative of the values obtained in previous calculations.'** For Calculation No. 8. half of the flow rate for each of the sprays had a drop size of 810 um ard the other half had a drop size of 309 u m. All other calculatior.s used a single drop size of 710 um. J. Burn Parameters The basic burn parameters used for most of the calculations of this repert are presented in Table A-XI whi,ch were used in the ar.alyses of Refs. 20-23. However. parameters were specified to prevent burn initiation in the upper plenum for Calculation No. 6 and parameters were specified to allow burn initiation in the upper compartment only for Cal lation No. 7. j _ . . , . t ,_ _ _ . - - _ .,- - - _
TABLE A-I VOLUME SPECIFICATION Total Component massa Ib Press., Temperature, (PartialPress.b,psy,) No. Volume 103 ft $ psia CF M09 N9 02 Name 139.6 13340.6 4102.3 Lower (0.19) (11.67) (3.14) Compart- 1 249.681c 15.00 110. ment 0.8 1522.5 468.5 Ice-Cond. 40. (0.01) (11.81) (3.18) Lower 2 24.7 15.00 Plenum 1.6 2895.3 891.7 Upper (11.81) (3.18) Plenum 3 47.01 15.00 40. (0.01) 128.2 39115.0 12039.8 Upper (11.77) (3.17) Compart- 4 681.283 15.00 75. (0.06) ment 23.9 3346.5 1032.1 Dead- (0.13) (11.71) (3.16) Ended 5 61.105 15.00 98. Region 30.7 2929.5 900.8 - Fan- (11.67) (3.14) Accumulator 6 54.828 15.00 110. (0.19) Room 0.7 1337.3 411.5 1/4 Ice- 40. (0.01) (11.81) (3.18) condenser 7,8, 21.695 15.00 (Typical)d 9,10 88396. Totals' 1205.387 NOTES - a Initial water masses, relative humidity and hydrogen masses are assumed to be Molecular weights (1b,/(1b,*nole)) used are 18.0 (H2 O), 28.0 (N2 ) and 32.0 (02 )* c Spray, rain and condensed water is assumed to accumulate in Vol No I until the 3 ft3 is reached. ginimum value of 151.x10 Initial conditions assumed to be the same as those for the lower and upper , l plenum. 3 total and ;
- The Sequoyah" and Watts Bar6 ice-condenser analyses used 1191.5 ft l 88228.7 lb , which are different from the D.C. Cook values by 1 % and 0.2 %, re-l spectively.
l
- TABLE A711 D.C. COOK JUNCTION DESCRIPTIONa
# Area d ~ Actual
- D NV2(Node)c .ft2 L/A, ft ~1 Door f NJ NVl(Node)c 2(LP) 990.* 0.1 LID 1 1(LC) 2 2(LP) 8(14) 931.2 0.005 -
3(UP) 1326. 0.005 IDD 3 7(11) 4 7(II) 3(UP) 20.* 0.1 - 5 3("P) 4(UC) 2040.* 0.1 TDD 6 3(UP) 4(UC) 20.* 0.1 - 7 4(UC) 1(LC) 2.2* 1.0 - 8 1(LC) 5(DE) 40.* 0.3 - 9 1(LC) 6(FA) 308.* 0.3 - 10 4(UC) 6(FA) g g. 11 5(DE) 6(FA) g g - 12 1(LC) 6(FA) g g - 13 9(12) 7(11) 982.5 0.01 - 14 10(13) 9(12) 982.5 0.01 - 15 8(14) 10(13) 982.5 0.01 - NOTES a See Fig. 1 for additional information; b NJ = junction nunber; e NV1 = one of the nodes connected by the junction, NV2 = the other node.d Node is the designa-
- tion assigned in Fig.1 (e.g., LC E Lower compartment). Areas with an are from Ref. 20 and other values are f rom f rom Ref. 30;
- All junctions use the in-which use the fan-flow option.
ertial-flow option except for junctions 10 - 12,f LID, IDD, TDD are the lower L/A is the inertial-effect term, see Ref. 2; inlet, intermediate deck, top deck doors, respectively. The TDD will lock into its full open position when reached.
TABLE A-111 i D.C. COOK JUNCTION LOSS COEFFICIENTSa
- - - Loss Coefficients d___ Ref. 20 EXLK12 ENLK21 EXLK21 NJb NVl(Node)c NV2(Node)c ENLK12 1.05 1.00 2.05 2(LP) 1.05 1.00 1 1(LC) 0.0525 1.00 f 8(14) 1.62 0.0525 2 2(LP) 1.00 1.62 0.0525 f 7(II) 3(UP) 0.0525 3
1.00 0.50 1.00 f 4 7(11) 3(UP) 0.5 0.45 1.00 1.45 0.45 1.00 5 3(UP) 4(UC) 1.00 0.50 1.00 f 3(UP) 4(UC) 0.50 6 0.50 1.00 1.5 1(LC) 0.50 1.00 7 4(UC) 3.20 1.00 4.2 5(DE) 3.20 1.00 8 1(LC) 3.20 1.00 4.2 6(FA) 3.20 1.00 9 1(LC)
- - - - - - Fan - - - - - -
10 4(UC) 6(FA) .
- - - - - - Fan - - - - - -
11 5(DE) 6(FA)
- - - - - Fan - - - - - -
12 1(LC) 6(FA) 0.0525 0.0525 0.0525 f 9(12) 7(11) 0.0525 13 0.0525 0.0525 0.0525 f 10(I3) 9(12) 0.0525 14 0.0525 0.0525 f 10(13) 0.0525 0.0525 15 8(14) NOTES a See Fig. 1 for additional information; b NJ = junction Nodenumber; is the designa-e NV1 = on the nodes connected by the junction, NV2 = the other nodeg ENLK12 and EXLK12 are tion assigned in Fig.1 (e.g., LC E Lower compartment). l the entrance and exit coefficients for flow from NV1
- All to NV2, junctions use respectively.
the iner-l ENLK21 and EXLK21 are tial-flow option except for junctions for flow from NV2 to NV1.10 - 1[, which use! Estimated. L/ A is the inertial-ef fect term, see Ref. 2; I g- s - - -
l TABLE A-IV i D.C. COOK FAN FL71 TO THE FAN-ACCUMULATOR ROOM (FA) FROM THE UPPER COMPARTMENT (UC), DEAD-ENDED REGION (DE) AND LOWER COMPARTMENT (LC) Path Per Train Press. Diff., Fan Flow, inches water efm x id' l
<-140. 16.72
-140. 16.72
-120. 15.47
-80. 12.64
- 40. 8.939
- 20. 6.321 0.0 5.300.
1.0 5.05 2.0 4.75
- 3. 4.45 4.0 4.15 4.5 3.97 ,
5.0 3.80 6.0 3.42 6.5 3.10 6.8 e 2.50 6.9 1.60 6.9 0.0
>6.9 0.0 NOTES
- 1. There are two trains. 2. The fan flow for each path is obtained by applying the pressure difference for the path to the table and multiplying the fan flow from the table by the factors 0.9569 (UC-FA), 0.0350 (LC-FA) and 0.0024 (DE-FA).
- 3. In Ref. 23 the f ans were started at 760 s, which was 570 s af ter the contain-ment sprays were activated at 190 s (because the containment pressure reached 3 psig). For this analysis, Cales. I and 2 used a fan-on time of 760 s. However, the pressure reached 3 psig at 129.1 s and 94.6 s for Cales.1 and 2, respec-tively. As a result, subsequent 2S D calculations have a fan-on time of 699.1 s (129.1 + 570.). Subsequent S2 DA calculations should use a fan-on time of 644.6 s (94.6 + 570.) .
~
- TABLE A-V D.C. COOK ICE-CONDENSER DOOR PARAMETERS
~ .................. D00R----------------------
Inte rmediate Top Lower Deck Inlet Deck Parameter 89 89 Maximum Open Angle, 55 degree Minimum Differential 5.5 1.15 Pressure for Maximum 0.0069 Opening, psi 1326. 2040. Maximum Flow Area, ft2 990.
- 20. 20.
Bypass Flow Area, ft 2 0. 3 Minimum Differential - 0.005 Pressuretojnitiate - Door Opening , psi
*Not used.
I i l i 1 I l l l l i i I
TABLE A-VI D.C. COOK PASSIVE HEAT SINKS (HS) COMPARE HS No./a Ref. 23 No./ Node Element Temp., oF/ No. of Coordin- Thick- Thickness Area, ft 2 Material Elements ate, in. ness.in. in ._ 1/1,3&4/UC Paint 2 0.012 0.12 0.006 Carbon Steel 3 0.372 . 0.36 0.12 75/31370 Cont. Coeff. 1 0.372 0. O. Concrete 8 2.772 2.40 0.30 concrete 5 7.572 4.80 0.96 Concrete 2 19.572 12.00 6.00 2 0.012 0.12 0.006 2/2/UC Paint 75/310 Carbon Steel 4 0.372 0.36 0.09 Concrete 15 1.80 1.80 0.12 3/5/UC 75/350 cont. Coeff. I 1.80 0.0 0.0 Carbon Steel 3 2.16 0.36 0.12 Cont. Coeff. 1 2.16 0.0 0.0 concrete 8 9.72 7.56 0.95 l concrete 15 2.40 2.40 0.16 l 4/6&7/UC Concrete 10 7.20 4.80 0.48 75/29841 Concrete 5 15.60 8.40 1.68 WTv_----_-.___._-________a---_-_-__.______--__
TABLE A-VI
". . D.C. COOK PASSIVE HEAT SINKS (HS) (Cont.)
COMPARE HS No./a Ref. 23 No./ Node Element Coordin- Thick- Thickness Temp., oF/ No. of ness.in. in. Elements ate, in. Area, ft 2 Material 0.012 0.012 0.006 Paint 2 5/8/LC 0.372 0.360 0.07 Stainless 5 110/540 0.012 0.012 0.006 Paint 2 6/9&l0/LC 0.444 0.432 0.07 Stainless 6 110/3819
- 0. O.
Cont. Coeff. 1 0.444 2.844 2.40 0.30 Concrete 8 7.644 4.80 0.96 concrete 5 25.644 18.00 9.00 Concrete 2 0.012 0.012 0.006 Paint 2 7/11&l2/LC 2.412 2 40 0.16 Concrete 15 110/38581 7.212 4.80 0.49 concrete 10 i 25.212 18.00 3.60 Concrete 5 4.00 4.00 0.40 Insulation #1 10 8/13/LP 12.00 8.00 0.80 Insulation #1 10 80/19100 12.00 0.0 0.0 cont. Coef. I 12.75 d.75 0.13 Steel 6 I l 1 ry:---_-_--_--_--__._---____.--____-_-__-__-_-_-__ -----__.._:_-__---_____2___--____. 2___--___.
4 TABLE A-VI D.C. COOK PASSIVE HEAT SINKS (HS) (Cont.) COMPARE HS No./a Ref. 23 No./ Node
---- Element Temp., oF/ No. of Coordin- Thick- Thickness Area, ft2 Material Elements ste. in. ness,in. in.
- 9/14/LP Insulation #2 10 4.00 4.00 0.40 j 80/13055 Insulation #2 10 12.00 8.00 0.80 Cont. Coeff. I 12.00 0. O.
4 Concrete 8 16.00 4.00 0.50 1 Concrete 2 24.00 8.00 4.00 10/15/LP Paint 2 0.01 0.01 0.005 40/3336 Concrete 15 2.41 2.40 0.16 i Concrete 5 4.01 1.60 0.32 j 4 11/16/UP Paint 2 0.012 0.012 0.006 1 _____ 40/9453 Carbon Steel 3 0.264 0.252 0.08 Carbon Steel 8 1.264 1.00 0.13 Carbon Steel 20 8.534 7.27 0.36 . 12/17/DE Paint 2 0.012 0.012 0.006 s i 98/6590 Carbon Steel 5 0.612 0.60 0.12 Cont. Coeff 1 0.612 0.0 0.0 concrete 8 3.012 2'.40 0.36 I Concrete 5 7.812 4.80 0.96 Concrete 2 25.812 18.00 9.00 i 4 I
TABLE A-VI
'. . D.C. COOK PASSIVE HEAT SINKS (HS) (Cont.)
e i - COMPARE HS No./a ' Ref.~23 No./ 8 Node Element Coordin- Thick- Thickness Temp., oF/ No. of ness,in._ in. ate, in._ Area, ft 2 Material Elements 0.006 2 0.012 0.012 13/18/DE Paint 2.412 2.40 0.16 Concrete 15 98/16789 7.212 4.80 0.48 Concrete 10 17.172 9.96 1.99 Concrete 5 0.012 0.012 0.006 Paint 2 14/19/FA 0.612 0.60 0.12 Carbon Steel 5 110/5640 0.612 0.0 0.0 Cont. Coeff. 1 3.012 2.40 0.30 Concrete 8 7.812 4.80 0.96 . concrete 5 25 812 18.00 9.00 Concrete 2 0.012 0.012 0.006 Paint 2 15/20/FA 2.412 2.40 0.16 Concrete 15 110/10134 7.212 4.80 0.48 Concrete 10 j 18.492 11.28 2.26 Concrete 5 10 6.0 b 6.0 D 0.60 16-19/None/Il-14 Ice 35/73250
- 20-26/None/ Instrument Heat Sinks, see Table B-I. l 1
i i
-- - )
l l I
~
TABLE A-VI D.C. COOK PASSIVE HEAT SINKS (HS) (Cont.) l NOTES
- The information in this column includes the COMPARE heat sink (HS) num-ber / The Ref. 23 HS numbers encompassed / The Fig. I node that contains this HS, where: DE E Dead-Ended region; FA E Fan-Accumulator rooms; II, 12, 13, 14 E Ice first 1/4 from top, second 1/4 from top, etc.; LC E Lower Compartment.;
LP I Lower Plenum; UC E Upper Compartment; UP E Upper Plenum. Also given are the initial temperature / Area. b Radius. 4
.1'
_30_ TABLE A-VII D. C. COOK MATERIAL DEPENDENT PASSIVE HEAT SINK PARAMETERS
- M Material Value Parameter Concrete, Carbon Steel, Paint, Ice 0.9 Emissivity Stainless Steel 0.4 Paint on Steel (UC) 0.21 Thermal Conductivity Paint on Concrete 0.087 k,{Bru/(hr.ft. F)]
Concrete 0.84 Stainless Steel 9.87 Carbon Steel 27.3 Steel in LP 26.0 Concrete in LP 0.8 , Insulation #1 in LP 0.15 Insulation #2 in LP 0.2 Ice 2.25 Volumetric Heat Capacity Paint on Steel in UC 29.8 Paint on Steel in LC, DE, FA, UP 14.7 oc, [ Btu /(ft3 . F)] Paint on Concrete 29.8 Concrete 30.2 Carbon & Stainless Steel 59.2 Steel in LP 26.0 Concrete in LP 28.8 Insulation #1 in LP 2.75 Insulation #2 in LP 3.66 Ice.(SIlb,/ft3) 27.15 l 10.0 Contact Coefficient All [ Btu /(hr.ft2 .oy )) f! - 1 i I I ' */ ! ? l u _.
TABLE A-VII (Cont.) D. C. t00K MATERIAL DEPENDENT PASSIVE HEAT SINK PARAMETERS
- Material Value Parameter b Paint on Steel in UC 0.016 Minimum Increment Size 6y, (in.) Paint on Steel in LC, DE, FA, UP 0.023 Paint on Concrete 0.10 Concrete 0.32 Stainless 0.078 Carbon Steel O.'129 Insulation in LP 0.055 Ice 0.055 aInstrument heat-sink parameter values are given in App. C.
~
b Used 6y(in.) = 0.060 [a(f t2/hr)x6 t(s)]1/2 based on Ref. ??, where a = k/(oc) and a 6 t of 10 s was used. a j' r NN
- i
- ?
t r g
l l l l l TABLE A-Vill D. C. COOK Irl:-BED PARAMETERS Parameter Value Initial Ice Mass, M3 2.37 x 10' lb, Initial Ice Heat-Transfer Area A; 2.93 x 10 5 ft: Heat of fusion,'AE/AM 245 Btu /lb, . Initial Net Free Gas volume 86780 ft' l Notes: During the t rar.s i ent , the node ice heat-sir.k area reductior. (AA) ard ice u c-partmer.t code volume ircrease (AV) during a time step are based or. the foreclae A '. AA(ft ! ) = AE x = 499 x 10-* AE(Btu) AE M; x g or, AA(n ) = 439 x 10- " x AE(J) , ar.d [ AV(ft') = 0.23 AA(ft ) . i or, AV(m') = 0.070 AA(m ) , where AF is the energy added to the ice over the tine step ar.d the ice is assumed to lose ler.gth with the radius of 5.52 in, retained.
A 9 s i ,s
.g 4 ,
)
l l cr -
TABLE M-IX D.C.COOKRADIAT(ONPARAMETERSa Compart-ment with Radiation Beam
/No.b Heat Sinks Includede Length, ft UC/4 1 - 4 & Spray 59.0 LC/1 5 - 7 & Spray
- 25.0 LP/2 8 - 10 8.5 UP/3 11 8.5 DE/ 5 12 and 13 8.5 FA/6 14, 15 & Spray f 8.5 11/7 19 1.0 a Radiation effects are based on all heat sinks being lumped into an equivalent bl'ackbody enclosure as described in Ref. 17. -
b Sei Fig. 1: DE E' Dead-Ended Region; FA E Fan-Accumulator rooms; 11 E Top 1/4 Ice; LC E Lower Compartment; LP E Lower Plenum, UP E Upper Plenum; UC E Upper Compartment; UP E Upper Plenum. c See Tables A-VI.and A-VII. d Spray area of 23252 f t 2/s based on 4000 gpm of 701 um drops.
- Spray area of 10464 f t 2/s based,on 1800 spa of 701 um drops.
p f Spray area of 30.69 ft 2/s based on 528 gym of 701 um drops. 4 e # 4 e d [ s' W r
TABLl' A-X D. C. COOK SPRAY SYSTI:M PAReil:TT.RS
- - - Cor. partner.t u ..
Pa rame t e r - tr L(' FA Drop Diameter, um 701. 701. 701. Flow Rate (for two trains), gpr 4000. 1800. 525. Temperature, F 125. 125. 125. Fall Height.b ft 75. 40.5 11.8 Spray Area.C ft /s 23252, 1086.t . 3069. Notes: a See Fig. 1: FA a Fan-Accumulator rooms; LC e Lower Compartmer.t: UC a Upper Compa r t mer.t : b Fall height determined from Ref. 20 drop fall times and a drop t e rn.i r; 1 velocity of 500 ft/ min, which was obtained typically in previous calcula-tions.*-* C Spherical area of drops. l w
.i 4
t TABLE A-XI a D. C. COOK BASE-CASE BURN AND PROPAGATION TIMES i - - - Burning - - - - - - - - - - - Propagation to NV - - - - - -
......From ...... ......From IK ..... i b' p' p' p' p' i
NVI Name2 b'ft . s IJI Name2 ft. s IKI Namd ft. s 4 UC 58.2 9.7 3 UP 6.6 1.1 1 LC 58.2 9.7 3 UP 10.8 1.8 4 UC 6.6 1.1 7 11 8.4 1.4 7 11 8.4 1.4 3 UP 19.8 3.3 9 12 8.4 1.4 9 12 8.4 1.4 7 11 8.4 1.4 10 13 8.4 1.4 10 13 8.4 1.4 9 12 8.4 1.4 8 I4 8.4 1.4 , , 8 14 8.4 1.4 2 UP 19.8 3.3 10 12 8.4 1.4 Ui 2 LP 26.2 4.36 1 LC 22.8 3.8 8 14 8.4 1.4 1 LC 22.8 3.8 4 UC 58.2 9.7 6 FA 22.8 3.8 5 DE 14.5 2.42 1 LC 22.8 3.8 6 UC 22.8 3.8 6 FA 15.5 2.58 1 LC 22.8 3.8 4 UC 22.8 3.8 NOTES l a Times from Ref. 21 Case C CLASIX I9 analysis, except for ice-condenser values.
- I Volume (node) number, see Fig. 1.
2 UC E Upper Compartment; UP 9 Upper Plenum; II E Ice top 1/4, 12 = Ice next 1/4, ete; LP i Lower Plenum; , LC E Lower Compartment; DE : Dead-Ended region; FA - Fan-Accumulator room. 3 L'U,L b p based on 6 ft/s flame speed and burn and propagation times, i.e., th' Ept . i
- LIS11NG Of les*uf Ca805 99 20 al 30 3S 40 45 60 SS 40 SS 70 TS 80 Lint v1 LIsot No COL S to i 9 C/s/520/al/TS/ Set
- O 87 0 i 2 3 90 99 9 24 3 0 t 3
4 0000- t5000. .SCO 80 03999 3 24000000 9 4 4 0900. .020 S000 09000 S S OSCO. .020 20000 20000 6 S 9000 .020 25000 2S000 7 7 16000- .020 2S000 2$000 *1vt LC S 8 249689. IS. 110, O. 139 6 0. 13340 4 4802.3 0.
- tv2 LD 9 0.0 0. 1522 5 448 9 0 9 24700. 19. 40. 0-999.7 0 -tv3 u8 to 47010. 15. 40. O. f.S 0. 2005 3 15 to 128.2 0. 39995 0 12039 8 0. twe UC 11 401203. 99. TS. O. 12 91909. 98. O. 23.9 0. 3244.9 9032.9 0. *tvl of 13 12 IS.
30.7 2929.9 900.8 0. two Fa 13 Sette. SS. 190. O. 0.
-tv7 It to
- 40. O. 0.7 C. 1337.3 d e t. S 0.
to 21495. 15 0 +1ve le 15 95 31695. 15. 40. O. 0.7 0. 1337.3 ott S
-tyt 12 16 16 21695. 19. 40. O. 0.7 0. 9327.3 aft.S 0.
tyt0 13 17 17 21695. 19. 40. O. 0.7 0. 1337.3 ott S 0. te tifLC 10 3 1 2 9 181LP 19 19 990. 1.0S t.00 1.09 1.00 0. 9 20 SIFL* 20 3 2 8 0 2 File 29 28 939.2 0.62 .0525 .052S 1.00 0. .9 31 Fit 22 22 3 7 3 9 setup 23
- 23 1324. .0929 9.00 9.42 .0925 0. .9 41J3 24 24 3 7 3 0 1. 4feep 25
.90 1.00 .90 1.00 C.
29 20. Situp 24 26 3 3 4 5 SFTuc 27
.45 9.00 45 1.00 0. 1.
27 2040 41JS 28 28 3 3 4 0 1 688=8 29 29 20 .90 1.00 .50 1.00 0. TIFbC 30 30 3 4 9 0 7FtLC at
.90 .50 1.00 C. 10.
31 2.2 9.00 elFLC 32 32 3 1 S O 1. eftDE 33 33 40. 3.20 1.00 3.20 9.00 C. tifLC 34 34 3 1 6 0 97tte 35 35 300. 3.20 1.00 3.20 1.00 C. 1 4 4 0 t tolFN 36 34 S 10F F u 37 37 9.9930 760. 100000. tilfN 30 38 9 1 6 0 t 9 tFF0 39 39 .07te 760. 100000. 9288N 40 40 9 5 8 0 t 42FfL 48 41 .0048 760. 100000. #21F2 42 42 3 9 7 0 .7 13F?e 42 43 982.S .0525 .0525 .0525 .0925 O. 44 tests de 3 to e o 45 45 982.S .052S .0929 .0925 .0525 O. .7 taff2 allfe de de 3 0 to 0 .7 19Ft3 47 47 902.9 .0929 .092S .0525 .0929 0. FT 40 . 48 19 13 49 49 187200. at0000. 167200. 140. 154700. *120. 63210. 20. 46 50 90 824400. 80. 89390. =40 LINT eso COL *S to 15 20 25 30 39 40 45 90 SS 60 SS 70 TS 80 LINENS i i l l 1 Fig. A.I. COMPARE-H2 D.C. Cook input deck card images for the Calcula- f tion No. I calculation, which had the S D2 blowdown. 1 l l i I
j i i l LISTING 08 DPUT CaeOS l L 38st uo COL S to tt 3090600. 29 301.35 40 47900. 49 90 SS 2. to et TO TS 30
?9 LisstNOSt 52 Ss $3000. O. d.S 90 93
- 3. 4t600 4. 39700. 43 99 93 93 44600 4 31000. 4.9 S3 38000. 9. 34200. S.99 te-98 S4 94 2S000. 8.0 16000. S.9 0 99 SS 99 0. 10000. L10 6 54 De LIO M 97 4 S8 97 SS. .0069 300 4 Se 500 W 19 4 TDC G 60 99 49 9.5 60 700 M 61 9.99 3 S2000 of 69 49. 3 4 7 0 42 43 t IS O t it 190.5 000?00. Wto 1*2 83 0. 197.3 99e700. 3972. W90 3 e Se
$2200. 3100. 93.93 494?O. 45 e4 28't. de es 20420. wSo S 6 3a.02 42620 4420. 28.40 we0 T 3 64 SS 3404 3700. 19.42 2tS20.
De e?S2. 44.42 96500. 9999. we0 9 90 41 14.07 99030 4940. 5.253 We0tta92 40 GT 4092. 7204. 4.040 a693. SS 1042. 4.110 $300. St. Me; t 2 49 O. es. 3480. O. MBO 3 4 70 SS 0 4096. .28 9982. 70 3904 .04t3 St. HBO S 4 ft
.740 199. 4192. 1.07 ??t.
MWO T 0 72 79 4420. 4230 .223 995. 72 5700. .4 30 St2. Det, Me? 9 90 73 73 4640. 960 S4S. 4960. .et? M5089 74 74 0070. .0367 599. 100 9*2 75 O. 3000. O. O. 76 TS 0. O. O. 8000 Eso 3 8 76 ette. O. 9003 4828. f t? S 4 ** 4?00. 5376. O 7000 *0
?? 4152 0 It*
- 18 7000 O. 1935 4 3 a e Det.? 19 3 3 7 3 4 79 9 2 3 1 2 i e 4 9 s.te 30 6 5 1 S S t 80 6 4 5 e 19 17 et I
et 2 3 3 2 9 9 Of 32 82 C00m 1.3.4 - UC Pt 93 03 9 4 e t 09 Se 84 3t370- .0 20000. .9 0. f. es t SS SS 9 0 .012 11 POSuc 04 to .29 9 29.8 et 2 et 07 3 0 .372 TS. CS?L et 88 21.3 9 59.2 et.3 39 et 9 0 .372 75. CCQtt to 90 to. , et.e 99 98 0 0 3.??2 75. CONC 92 92 .04 9. 30.3 St S 93 93 $ 0 7.572 15. CONC 98 94 .94 9. 30.2 et.s 95 - 95 2 0 19.512 15. CONC 96 94 .Se 9 30.2 02 91 91 C00m 2 - UC P2 94 90 9 3 4 t 02 99 20000 .9 0 9. 99 380. .0 e2 t 900 900 1 0 .012 19. 30 39 80 49 SO SS 60 SS 70 75 to LlutNO LINEN 0 COL.S to 99 20 25 Fig. A.1. (Cont.)
LISilesG Of tesPUT Cae05 to il 20 25 30 39 40 45 90 SS 60 49 70 TSPolvC to Lisstee0 t0 tiesteso tow S 29 s 82 2 t'. a tot .3s 1. 79 903 402 4 0 .372 Clit 104 103 27.3 9. 89.2 03
- UC P2 IOS
- t04 C00m S 4 9 03 504 106 1 S 20000. .9 0. t.
83 9 807 106 390. .0 t to 19. CO**C 100 907 SS O SC9 100 .84 1, 30 2 83 2 1.00 79 CC0tf 990 tot 1 0 999 190 10. 23 3
- 79. CSfL 112 19I 3 0 2.94 183 ft2 27.3 9. 99.3 83 4 tid 0 2.14 79. CC0tF 113 1 83 S ttS 114 to. TS. tte itS S 0 9.72 C0eet 187
.84 1. 30.3 04 tto P4 198 197 C00m 4.7
- tJC 999 1 3 4 9
- 9. 04 tit 20000 .9 C. 84 1 120 199 39041. .0 74 121 120 il 0 2.40 C0 esc 122 139 .84 1. 30.2 es 2 1.20 15 C0*sc 123 922 90 0 424 123 .04 t. 30.2 94 3 19.60 ft. C0asC 125 824 S O 134 f25 .94 t. 30.2 OS 127
- LC PS 124 C00m 8 9 OS 120 137 1 2 4 4.
30000. .9 0. eS*t 429 (28 S40. .0
.012 sto. POSWC 930 829 9 0 est 130 .21 9. 14.7 et 2
.312 110. SSTL 132 936 8 0 833 932 9 87 9. St.2 04 P4 f34 133 CODE 9.40
- LCt t 06 43S 834 8 6 f. 934
.0 20000. .9 0. es-t tal 3819. P05WC 137
#34 9 0 .012 etc.
830 137 .29 1. 44.1 e6 2 444 tto. 55ft 139 138 0 0 140 139 9.57 t. St.2 e4 3 tot 940 1 0 444 110. CC0tF 94 4 set tot 10. 110. 943 942 8 0 2.044 Comec e44 143 .84 s. 30.2 e4 9
?.444 110. COnec 945 944 S O e44 tal .Se 1, 30.3 et 4 647 tes 2 0 29 444 180. ConsC 30.2 07 tot 147 .04 1. 149 .
Pt 940 C00W 19.12
- LC 9 47 190 149 1 4 8 1.
20000. .9 0 190 Selet. .0 uwie C06 3 30 is 20 as 30 3s 0 es 30 se eO el to to eO uwi O I 1 Fig. A.1. (Cont.) 1 u-
LIS?ttes tp tesPu? CaeOS 40 45 to SS to el 70 75 to Ltestuo LlostND C04*5 to tS to 29 30 39 81 9 158 ett t 0 .012 990. POCONC 152 192 .007 9 29 8 87 2 153 993 el 0 3.413 too. CONC 954 the .34 1, 30.2 87 3 tSS 199 to O ?.292 110. CONC 194 194 .44 4. 30.2 e?.a 137 It? S 0 39.212 #10. CONC 158 498 .94 9. 30.3 09 199 199 C00u 13 . LP PS 160 140 4 4 2 i
- 9. De 161
.0 30000. .9 0.
tot 19100. se.t is2 962 to 0 4.00 D0. INST 163 tes 99 t. 2.79 eg.2 tee tea to 0 13.00 00. INST 145 165 .45 t. 2.?S eg.3 ist 164 0 12.00 80, 947 0 CCott to? 10. et e tot 448 4 0 12.79 00 CStL 909 tot 27.3 9. 99.2 09 970 170 C004 to
- LP P9 t?t 110 t S 2 1
- 1. 09 972 172 130SS .0 20000. .9 0.
et.t os s 173 to O e.00 to INS 2 974 974 .30 9. 3.44 89 2 175 t?S to 0 12.00 90 IN12 976 914 .20 9. 3.46 39 3 t?' t?' s O 12.00 90 CCott **8 478 10. et 4 579 979 0 0 96.00 80. CONC 100 180 .84 1 30,2 tot 89 5 109 2 0 34.00 80 CONC ' 102 902 .Se 9. 30.2 Oto 133 903 C00m 15 a LP P90 tes 104 1 3 2 e Oto tel 195 3336 .0 20000. .9 0 t. sto t 106 ISS 2 0 .01 80 80 CONC tt? 107 .04? 9 29.8 802 908 108 IS O 2.49 40 CONC tot 199 .84 9. 30.2 se0 3 $90 tDO S O d.0 40. CONC 196 198 .08 9 30.2 Ott sti 192 C00m 16 - WP Pet 993 193 9 4 3 9 089 198 994 9453. .0 20000- .9 0. f. ett t 195
- 195 9 0 .012 40. P0lNUC 196 196 .31 9. 14.7 est.2 19?
197 3 0 .26s a0 CSit 190 199 37.3 9 99 2 est.3 999 199 0 0 9.244 80 Cltt 200 200 27.3 9 99 2 30 3S 80 45 SO SS to SS 70 75 80 LINfMO LINtNO C0t*S to 95 20 29 Fig. A.I. (Cont.) O n_ -
l i LISilNG OF input CaeOS LlostNO COL S to tS 20 29 30 SS 40 49 SO SS 40 SS TO TS 80 LINENO
. tot 20 0 0.534 80. i est 4 20s 202 27*3 9. 59 2 i CSTL 202 203 C00m 97
- DE 0 92 203 See t 4 S 9 0 #t2 204 205 8590 .0 20000. .9 0.' t. Ot2 205 206 1 0 .012 98. 512 1 206 207 .29 9. 14.7 P054uc 201 208 5 0 .St2 94.
al2 2 200 209 27.3 9. 59.2 j Cltt 209 290 9 0 .492 94. 312 3 210 2tt to. CCott 211 att S 0 3.012 94 812 4 282 293 .04 9. 30 2 CONC 293 214 5 0 7.812 98. 892 5 214 att .84 1 30.2 CONC 215 216 3 0 25.812 te. 892 6 2'E 217 .04 9. 30.2 CON: 2t? 298 C00m 10
- Of 013 210 219 t 4 g 4 pt3 299 220 18789. .0 20000. .9 0. 1. 013 220 228 1 0 .012 94. et3 1 22t 222 .081 1. 29 8 PCCONC 222 223 15 0 2 442 te. 883 2 223 224 .se t. 30 2 CON: 24 22S 10 0 7.212 90. 893 3 225 226 .94 1. 30.2 CONC 224 227 5 0 17.972 98. #t3 4 227 228 84 9 30 2 CONC 228 229 C00m 19 F4 Ota 229 230 1 6 e t #14 233 239 S440. .0 20000. .9 0. t. Old 22t 232 1 0 .012 190. sta t 232 233 .29 1 14.7 POSNUC 233 23a S 0 .492 150. 884 2 23a 235 27.3 1. 59.2 CSTL 23S 23e t O .412 190. R*4 3 236 237 10. CC0ff 23' 230 0 0 3.012 too. ste 4 23e 239 .44 9. 30.2 CONC 239 240 5 0 7.882 110. eta S 280 249 .44 9. 30.2 CONC 2at 382 2 0 29.882 110. R88 6 242 283 .84 9 30.2 CONC 243 See C00m 20
- Fa Ott ase tal 9 4 9 9 pts 24S Set 10134 .0 20000. .9 0. 1. 01S 246 247 1 0 .012 tio. AtS t 241 240 .047 1. 29.8 POCOsec 248 -
289 19 0 2.492 110 atS 2 249 250 .84 t. 30.2 CONC 250 Lltet NO COL *S to tS 20 at 30 35 40 4S SO SS 40 el 70 19 80 LINEN 0 Fig. A.1. (Cont.) m..-
LISilhG Of INou? CeWOS 29 30 39 60 49 50 SS 90 SS 70 75 to Liut 40 Ltut to COL S to SS 20 StS 3 259 2St to 0 7.392 180. C0we 212 2S2 .84 9. 30.3 est 4 2S3 O 98.492 990, 253 S CONC 214 254 .Se 9. 30.2 0 #6 2SS 25S la
- WOL 8
- T2 O P e6 2S4 254 2 e 2 O 1
.9 C. f. O to 2S?
297 073250. 08.0000 40 00 e to t 2St 290 to 0 00.0000 39.0 S 14 2 299 259 2.29 Sa. .502 0 17 240 ISO 83
- WOL to
- T3 9 17 249 369 2 1 to 3 0 0 0 97 242 242 073290. 08.0000 40.00 .9 0. f.
e t? t 243 243 10 0 00.0000 39.0 S 91 2 264 264 2 25 S4 .502 , 0 IS 245 24S 32
- WCL 9
- to P to 244 364 2 1 9 4 0 0 O is 247 247 073250. 06 0000 40.00 .9 0, t.
e to t 240 240 to 0 00.0000 3S.0 5 te 2 tot 269 2.2S S4 .502 O 99 270 270 It
- WOL T
- TS 0 P 99 27s 279 2 9 7 5 0 0 99 2'2 272 073250. OS 0000 40.00 9 0. 1.
e 99 9 273 213 to 0 00.0000 35.0 S4. .902 5 99 2 2?a 278 2.25 0 20 2?S STS 3GNttle ASStuntv Uo/STL STL CU P to 274 2?S 9 4 3 1 0 0 2'? t.te .9 0. 1. 0 20 2?? 1 O. e to t 2?e 2?S 2 0 13 040. O. 5 20-t 279 279 27. 1. 99 280 040. O. B 20 2 290 to O t.09 S 20 2 att 208 2?. 9 99. # 20 3 202 282 S 0 3.26 040. O. 282 St. S 20 3 203 220. 1 A 2C 4 204 208 S O 4.09 040. O.
$ 20 4 295 205 2? 1. 99 296 296 gat?ON Teawgeg77tB CAStkG LC/>w? STL 0ft 0 P 28 287 2s? 8 2 1 1 0 288
- 9. O. 1.ft .9 0. 9 0 29 200 e 29 9 289 289 9 0 0650 900. O. 290 28.39 $ 29 f 290 .0767 1. e 2t 2 - 29 '
29' S 0 .2565 900. 292 S 22 2 292 27. t. 99. 0 22 293 . 293 Catt! IN COM0Ulf-LC/STL-etWO State CU P 22 294 294 2 e t t 0 0 0 22 29S 29S 9 .304 9.04 .9 0 9 O. # 22 1 294 2 96 4 0 979 100. 297 27. 9. St. 5 22-9 29? 900. O. e 22 2 299 290 S 0 .128 5 22 2 299 299 9610 9 36. e 22 3 300 300 S 0 .00 0 100. O. L t ht MO COL S to 15 to 25 30 SS 40 AS SO SS DO 45 70 75 00 LINf v2 l 1 l l Fig. A.I. (Cont.) l l l l i l l l t _ l
o L151last OF INPUT CaeOS tS 20 29 30 3S 40 45 SO SS to SS TO TS to L Isot No LIsot eso COL S to $ 22 3 309 30s 4146 t. 24.38 e 22 4 302 302 2 0 .0 100. O. S 22 4 303 303 230 9. St. 0 23 30a 304 Catti les CDesDuit UP/Stk.estee. State.CU p 23 305 30S 2 4 3 4 0 0 306
.304 1.85 .9 0. 1. 0 23 306 9. e 23 9 307 307 4 0 . tit 040 O. 3^B St. 5239 30s 27. 1.
e 23 2 309 309 S O 924 040. O. 5232 310 390 .ttte 1. 34. 3et 040 O. e 23 3 Set S 0 .001 5233 382 312 .9964 t. 20.30 e 23 4 313 393 3 0 .0 040. O. 314
- 1. St. 5234 314 230 0 2a 359 3tt .92 SIN AL IkSYe.*LC P 24 3i6 Sto t 9 1 0 0 3t?
t
.12S 9.16 .99 0. 9. 0 24 317 4. e 24 3'0 SIS 2 0 0, too. O. 399
- 4. 36.6 $ 24 319 105.1 02S 320 320 .467tN AL INSte.*LC 9 0 P 29 32' 325 1 9 9 0 02S 322 322 1. .467 1.14 .99 0. 1.
e 25 323 323 3 0 C. t00. O. 5 25 324 324 806.9 9 36.6 325 0 26 32% .250!N STL INSte.*LC P 26 324 326 1 1 1 f 0 0 327
.9 0. 1. 0 26 32' t. .250 f.16 e 26 325 320 3 0 C. 800. O. 329 84.4 5 28 329 30. f.
- 1. Tt UCNIO 330 330 2 0 8 0 O. O. 2. 9.
339 2 IS .022 1. O. 10000. 72 MSte 339 9 f3 MSt? 322 332 1 19 to .022 1. O. 50000. O. 90000. 18 MSIS 333 333 9 to 13 .022 1. it Mste 334 13 3 .022 1. O. 10000, 314 1 0 vt .025 339 339 7 7 9 9 0 0 27 2130 t 1000 900 t v2 336 334 0000. 9000. 8. 337 337 t s. 10 40. t 330 338 1 6. 90. 40. 2 339 339 1 6. 10. 40. 3 340 340 t S. to. 40. 4 349 349 9 8. 10. 40. S 342 342 t 8. to. 40. 8 343 o 343 4 4. 10. 40. 7 344 344 3 s. O, 1900. t 34S 349 3 8. O. 1900. 2 346 346 2 4. O. 1900. 3 347 347 2 S. O. 0600. 4 34e 348 2 S. O. 0600. S 349 . 349 3 6. O. 1900. 6 350 390 2 e. O. iS00. 7 LINteso COL-S to il to 25 30 34 40 45 SO SS to GS 10 75 80 LINteso Fig. A.1. (Cont.)
. f-4
. .,1
- i y n -
l 1
- LIS? test Of IM*UT Cae05 LItst MO COL *S 90 19 30 29 30 39 40 49 SO SS 60 SS TO 79 80 LINtNO 199 391 3 6. O. 30000 t 352 352 3 S. *2000. 90000 3 3S3 393 3 8. *2000. 10000 93 398 39 4 3 6. *0000. 12000. 3 399 3SS 3 6 *2000. 90000 5 3SS 394 3 4. *S000. 4000. 0 3S?
3S? 3 S. *2000. 90000. to 3SS 390 3 S. *2000, 90000. 99 S. *2000 90000, 12 399 359 3 340 360 4 S. etS. 'S. t 3 349 36 9 4 6. etS. IS. 3 3 30 2 342 4 S. ett. 19. 3 7 343 SS 3 4 8. *tt. 99. 3 4 6, 15. 4 See Sta a *tS. t 349 36 5 e S. ett. 99. t S 366 366 4 S. *19. 99. t 6
- 19. 4 4 347 367 4 8. *IS. 360 340 4 6. *tS. 15. S S 349 fan COOL 36 9 3.26 129. ele. 30000. .05 0 atSP 370 370 3 R2pc 379 319 4 79. 4000. 709. ,
372 9 40 1000. 70s. 82tc 3'2
- 12. 520. 709 R2ra 373 373 6 S 8t MR 378 3?d 3?S 379 900. 9. 2. 3. 4 5. 4. Suu8 376 S Sete S Gett 9.00f4 5 Sett S.00ft S DetS S 0014 Se map 3*4 3?? 9.22ES 29 N28 3??
00 3?S .00 .08 .012 9. .05 .0 1 .95 9. 22 9 LC 370 379 90 .00 .002 9. .09 .0 9 .SS t. 22 2 Le 379
.09 .0 .SS t. 22 3 US 300 300 .00 .08 .012 1. 1 22 8 UC 38' 309 .00 .08 .012 9. .05 .0 t. .99 t. 302 302 .09 De .092 9. .05 .0 9. .55 9. 22508
.09 .0 22484 303 303 .00 .00 .012 9 9. .9% t.
388 304 9.0 .08 .092 9. .09 .0 9. .99 9. 22 7 ft 309 90 .00 .012 1. .09 .0 9 .55 t. 22 0 le 309 306 10 ,00 .012 9. .09 0 9. .SS 9. 22 9 12 306 307 9.0 .00 .012 9. .0S .0 1 .SS t. 22 9033 307 4 4 3 80 S.70 3.00 23 I LC See 300 9
- 309 9 0 9. 4 36 3 to t.40 22 2 Le 309 390 4 ? t. t.00 f.10 9.40 23 3 up 393 398 3 9 t. S.70 1.10 S TO 23 8 UC 39' 392 1 6 1 2.42 3 00 3.00 23 $ Of 392 393 4 4 1. 2 SS 3 to 3.00 23 4 Fa 393 398 3 5 9. 9.40 3 30 9 80 23 ? II 398 .
39S 2 to t. t 40 t a0 3 30 23 0 le 394 396 ? to 1 1.40 t so t.a0 23 9 If 394 397 9 0 9 1.80 9.40 9.40 23 1013 397 390 +725000 asc a 390 399 4 9 2 3 4 *t 0 0 0 0 0 St. 23252 e t UC 399 400 t S 8 ? *t 0 0 0 0 0 0 29. 90448 82 LC 400 LINENO COL *S to 19 20 25 30 39 40 49 90 SS 60 SS TO 79 30 LININO l 4 i l Fig.,A,.I. (Cont.) s i
/
- 'e'l I
l l I
. . = , - - . - . - - __. . ..-. - - . . , ~ .
e so I tilflesG 08 lesPut Caec$ Llasteso COL *5 to il 30 36 30 at 40 45 to 99 GO St 90 19 80 - Lliet te 40t 2 8 9 to 0 0 0 0 0 0 0 0.9 33 69 40s 402 3 99 0 0 0 0 0 0 0- 0 0 09 Se u* 40: 803 9 12 13 0 0 0 0 0 0 0 0 0.9 39 f t 403 I 404 4, it *0 0 0 0 0 0 0 0 0.9 3000. Se e, f la. 404 40 t.o i O O O O O O O O e 9.0 409 6t0 co. t to it 30 at 30 39 40 et to SS to et te 99 .0 ti e.e l i i J i i I 4 I Fig. A.1. (Cont.) 4 r i s i ! I O f , ..-r- .- = ., - . . < - - - , , ~ - - , . . , . - - - . . , . . . . . , ,, - .,, - 35- -,e . 4 .-- . , . , _,._...~r. ,,- , . ~ ,,
NOTES - I
- 1. The time steps indicated above (0.020 s) were reduced before burning a,etually occurred after about 4500 s. )
- 2. The flow inertia term (L/A) were reduced to there approximately actual values before burning acetually occurred after about 4500 s.
- 3. Some of the parameters were modified interactively for sensitivity-study calculations. Examples of parameter changes made for the D.C. Cook analysis include the burn parameters, fan-flow table multipliers and spray flow rates.
- 4. Plot limits specified above were modified for some of the plots presented.
Fig. A.1. (Cont.) e
- m. - -
APPENDIX B INSTRUMENT HEAT-SINK MODELS _ The instruments modeled for the D. C. Cook calculations are described in Table B-1 and Figs. B.1 through B.3. All models are one dimensional because this is the only capability available in COMPARE. Multidimensional analyses could be performed with a code like AYEM9 with boundary conditions based on COMPARE. The igniter assembly, Barton transmitter and cable-in-conduit instrument descriptions were obtained f rom Ref. 24. Reference 25 provided de-scriptions for the 0.125-in. aluminum, 0.667-in. aluminum and 0.250-in. steel instruments. The igniter assembly instrument one-dimensional cartesian representation is along the vector X shown in Fig. B.1. The model selected should result in higher internal temperatures for a tr,nsient with increasing ambient temperatures. The App. B temperature plots for this instrument show the innermost temperature variation with tir.e. The Barton-transmitter-casing ins.rument one-dimensional cartesian repre-sentation is along the vector X shown in Fig. B.2. The model selected should result in higher internal temperatures for a transient with increasing ambient temperatures. The App. C temperature plots for this instrument show the - innermost temperature variation with time. The cable-in-conduit instrument one-dimensional cylindrical representation is along the vector R shown in Fig. B.3. The model selected should result in higher internal temperatures for a transient with increasing ambient temperatures. The App. B temperature plots for this instrument show the inne rmost temperature variation with time. In the model the conduit steel is assumed to be wrapped around the cable and does not account for the actual con-tact between tte conduit and cable being small. In Ref. 6, a 10% contact area was approximated by increasing the cable-in-conduit properties by the factor 10. This approximation is based on the temperature response being dependent on h6/k, the Biot number, and a = k/pe, the thermal dif fusivity. Therefore, a reduction of h by a factor of 10 is equivalent to increasing k by a factor of 10. Howev-er, oc must be increased by 10 to preserve a. A better approach would be to l l l N-~
r modify COMPARE so that the heat-transfer coefficient could be multiplied by a factor to acccunt-for the reduced contact.
.The 0.125-in. aluminum, 0.667-in. aluminum and 0.250-in. steel instruments are simple one-dimensional cartesian representations with the outside surface temperatures plotted.
r l l l l l l i i l
*r i l l I l
l \ l l l
TA3LE B-I INSTRUMENT HEAT-SINK MODELS Coordin- Thick-No. of Mater- ate, ness, Description Layer Elements ial in in HS #20 - Igniter Assembly, 1 2 Steel 0.13 0.13 in UP (NV #3), Cartesian, See Fig. B.1. 2 10 Steel 1.01 0.88 3 5 Copper 3.26 2.25 4 5 Steel 4.01 0.75 MS #21 - Barton Transmitter 1 5 Paint 0.0065 0.0065 Casing, LC (NV #1), Cartesian, 2 5 Steel 0.2565 0.250 See Fig. B.2. HS #22 - Cable In Conduit, 1 4 Steel 0.304 0.133 in LC (NV #1), Cylindrical, See Fig. B.3. 2 5 Rockhide 0.171 0.045 3 5 Silicon 0.126 0.045 4 2 Copper 0.081 0.081 HS #23 - Cable In Conduit, 1 4 Steel 0.304 0.133 in UP (NV #3), Cylindrical, 2 5 Rockhide 0.171 0.045 See Fig. B.3. 3 5 Silicon 0.126 0.045 4 2 Copper 0.081 0.081 HS #24 - 0.125 Aluminum, 1 2 Aluminum 0.125 0.125 in LC (NV #1), Cartesian HS #25 - 0.667 Aluminum, 1 3 Aluminum 0.667 0.667 in LC (NV #1), Cartesian HS #26 - 0.250 Steel, 1 3 Steel 0.250 0.250 in LC (NV #1), Cartesian
TABLE B-1 (Cont.) MATERIAL PROPERTIES Thermal Conductivity. Density x Specific Heat. Mat'erial HS
- k. Btu /(hr*f t F) ac. Btu /(ft 3 F)
Steel 20,21,22,23 27. 59. Copper 20,22,23 230. 51. Paint 21 0.0767 28.3 Rockhide 22.23 0.1518 36. Silicon 22,23 0.1156 , 28.31 Alumir.um 24.25 105.1 36.6 Steel 26 30. 54.6 NOTES HS = Heat Sink LC a Lower Compartment . NV = Volume Number UP = Upper Plenum Instrument initial temperatures were set uniformaly to the temperature of the compartment in which they are located. 1 I D- " ,.
A g, m mm mmw.w_u um s w. s
. f g
./ .
.Q e ts :
1 (%Sig & b , b\t* q
) @ 5 +4e.1 d i X N N
' Cager Csh!W
" ' " " -~='
E cswe 4 q - '
\ wA 3
i 3 A,.
) 8
'A i 2 i 7,M_ $h_..
_ _j TR 0 o,y) ,} 2.[ y@ l y ,,,'s sW' . l l Fig. B.l. Igniter assembly instrument cross section. Model used in COMPARE is one-dimensional cartesian along vector X.
l 4 Pasn 0.cokC" Ok.,k gfly E f L
- e '
3,35d l ' I V ,. I l
\ ?
Cs k n Mr s%l v l b Kd b N C.1SG h l
'~'
0, ' R
, 4 O %GO
.sl os 1.7 s "
Fig. B.2. Barton transmitter instrument cross section. Model tsed in COMPARE is one-dimensional cartesian along vector X. a
r- -- e o.*DS
. 1r
~
l ] swt l" . I
. 1.05 G ,
& \%k g b WYsp k$y u q, ,
0 0. t tez* a O " t es O ."M 3 Fig. B.3. Cable-in-conduit instrument. cross section. Model used in COMPARE is one-dimensional cylindrical along vector R. I l l 1 l l i k _ '
APPENDIX C REPRESENTATIVE PARAMETER TRANSIENTS This appendix characterizes the calculated results by preser.t i r.g the i results for Calculations No. 1, 2, and 7. Table C-1 tabulates the ever.ts , preceeding the occurrence of maximum AP and pressure levels. Figures C.1 through C.3 are plots of the variation with time of the node (a) component volu-metric (mole) fractions, (b) inst rument temperatures, (c) pressures, a r.d (d) temperatures. Figure 1 can be used to obtain the corresporder.ce betweer. node nur.ber and compartner.t e.g., volume 1 is the lower compartment. Addition-al calculated results are provided in Apps. D and E. The position (opening) of the doors when the maximum AP occurs across them is of interest because the position in conjunction with the AP may affect the mechanical loadir.g of the doors. Although the position of the docrs is not an output parameter, the following can be said about the door positions based on (a) the equation used to determine door position from the AP across the decr. which is proprietary to Westinghouse, (b) the equation parameters specified in Table A-V, and (c) the comments given in Tabte C-1: 1
- 1. The top-deck doors (TDD) are always locked open af ter the first uppe plenum burn, see Table C-l.
- 2. The intermediate deck doors (IDD) and the lower irlet doors (LID),
which do not lock open, will close whenever there is a AP to close them based on the equation used, which has no accour. ting for door inertia. Therefore, these doors will be closed when the maximum AP to close the i doors occurs. Of course, this applies to the TDD if they are not locked oper..
- 3. If the maximum AP in the opening direction is less than the Table A-V l value to result in the door being full open, the door position can be determined from the AP value, the door equation and the door equation parameters given in Table A-V. , i l
i ! i t l l 1 1
.-- - , -~ - - - - ~ - - - - - ~ - - - - - ~~~--~ ~~~~ ' ' ~ ~ ~ ' ~ ~ ~ ' ~
TABLE C 1
- EVENT SEQUENCE FOR CALCULATIONS 1, 2 & 7 CALCULATION 1
,- Time, s - - - - - - - Event - - - - - -
g cown Synbol Value urit% Cowserts Problem 4575.21 0.0 UP IB 6.46 als/s 1st UP & calc burn. 75.24 0.03 UP-UC LO - - TDD locked open. 75.29 0.08 UP-UC @ 1.3 psi Max UP UC AP. 76.30 1.09 UC-UP @ 0.8 psi Man UC-UP AP. 4642.32 0. 0 LC IB 8.74 mis /s 1st LC burn. 42.86 0.54 LC-LP @ 0. 9 psi R x LC-LP AP.
@ 0.7 psi Max LC-FA AP.
42.90, 0.5S LC-FA 43.03 0.71 UP IB 5.57 mis /s 3rd UP burn. 44.72 2.40 UP TB 1.69 s IBT = 1.80 s. 45.04 2.72 LC PP 25.1 psia Peak LC press. 45.05 2.73 LC TB 2.73 s IBT = 3.80 s. 45.07 2.75 DE PP 25.0 psia Peak DE press. 45.09 2.77 FA PP 25.0 psia Peak FA press. 45.43 3.11 DE-FA @ 0.5 psi Max DE-FA AP.* 4765.31 0.0 LC IB 8.65 mis /s 2nd LC burn. , 65.70 0.39 UP IB 6.44 mis /s 5th UP burn. 67.15 1.84 UP TB 1.45 s IBT = 1.80 s. 68.12 2.81 LC TB 2.81 s IBT = 3.80 s. 68.14 2.83 LP PP 25.8 psia Peak LP & cale press. 68.15 2.84 14 PP 25.4 psia Peak 14 press. 68.15 2.84 13 PP 25.2 psia Peak 13 press. 68.16 2.85 12 PP 25.1 psia Peak 12 press. 68.16 2.85 11 PP 25.1 psia Peak 11 press.
/
'l
. ~ _ _ - - _ _ _ - _ _ - _ - _ _ _ _ _ . - _ .
TABLE C-1 (Cont. ) EVENT SEQUENCE FOR CALCULAT10hh 1. 2 & 7 CALCUL ATION 1 (Cont. )
. . Tine, s - s - - - - - Event - - - - - -
Problem rono Svehol Value urits comnents et 4968.74 0.0 UP IB 5.94 als/s 9th UP burn. 69.85 1.11 UP TB 1.11 s IBT = 1.60 s. 71.31 2.57 LC IB 8.75 als/s 4th LC burn. 72.17 3.43 FA-LC MD 0. 7 psi Max FA-LC AP. 73.68 4.94 UP IB 5.88 nis/s 10th UP burn. 73.82 5.08 LP-UP MD 1.1 psi Max LP-UP AP. 74.01 5.27 LC TB 2.70 s IBT = 3.60 s. 74.04 5.30 LP LC MD 1.4 psi - Max L1D AP.* 74.87 6.13 UP TB 1.19 s IBT = 1.8) s. 74.90 6.16 UC PP 25.3 psia Peak UC press. 74.92 6.18 UC-FA MD 1. 9 psi Max UC-F4 AP. i 74.95 6.21 UC-LC MD 1.9 psi Max UC-LC AP. 75.06 6.32 UP-11 MD 1. 9 psi Max IDD AP.
. 75.07 6.33 UP PP 25.4 psia Peak UP press.
75.07 6.33 UP-LP MD 1.9 psi Max UP-LP AP. 5585.20 0. 0 LC IB 8.29 nis/s 9th LC burn. 85.30 0.10 LC-DE MD 0. 6 psi Max LC-DE AP. 85.49 0.29 LC-UC MD 1.0 psi Max LC-UC AP. 85.56 0.36 FA-UC MD 1. 4 psi Max FA-UC AP. 85.56 0.36 FA-DE MD 0. 9 psi Max FA-DE AP. 86.45 1.25 UP IB 5.77 nis/s 22nd UP burn. 87.84 2.64 UP TB 1.39 s IBT = 1.60 s. 87.88 2.68 LC TB 2.68 s IBT = 3.80 s.
$$20.01 0. 0 LC IB 8.18 n.ls/s 10th LC burn.
~
20.41 0.40 11-UP MD 0. 3 psi Max 11-UP AP. 20.42 0.41 UP IB 5.18 nis/s 26th UP burn. 22.68 2.67 LC TB 2.67 s IBT = 3.80 s.
<l 22.77 2.76 UP TB 2.35 s IBT = 1.60 s.
j
. V
_t
TABLI: C-1 (Cont. ) EVENT SEQUENCE FOR CALCULATIONS 1. 2 & 7 CALCULATION 2
..- Time, s - - - - - - - Event - - - - - -
Prhblem g* Corre Syrrbol Value units comrre r.t
- 4473.81 0.0 UP IB 6.77 mis /s 1st UP & calc burr..
73.83 0.02 UP-UC LO - - TDD locked open. 73.96 0.15 UP-11 2 2.6 psi Max IDD AP. 74.01 0.20 UP-UC 2 3. 6 psi Max UP-UC AP. 74.02 0.21 UP-LP m 3. 6 psi Max UP-LC AP. 74.47 0.66 UC-UP @ 2. 3 psi Max UC-UP AP. 74.52 0.71 LP-UP @ 2.1 psi Max LP-UP AP. 74.89 1.05 UP TB 1.05 s IBT = 1.60 5. 75.96 2.15 LP-LC MD 1.4 psi Max Llli AP. 4593.67 0.0 UP IB 6.05 mis /s 3rd UF burr.. 94.61 1.14 UP TB 1.14 s IBT - 1.50 s. 94.89 1.22 11-UP MD 1.1 psi Max 11-UP AP. s i i. I b i l
TABLE C-I (Cont.) EVENT SEQUENCE FOR CALCULATIONS 1. 2 & 7 CAtrULATION 2 JCont. ) )
.- - Time, s - - - - - - - Event Problem gt Cono Symbol Value ,, units Commerts 5814.28 0. 0 DE IB 5.65' mis /s 1st and only DE burn.
I 15.70 1.42 LC-UC MD 1.8 psi Max LC-UC AP. 15.92 1.64 FA-UC MD 1. 9 psi Max FA-UC AP. 16.02 1.74 UP IB 4.97 mis /s 47th UP burn. 16.21 1.93 DE TB 1.93 s IBT = 2.42 s. 15.21 1.93 DE PP 35.1 psia Peak DE & cale press. 16.21 1.93 DE-FA MD 10.6 psi %.* DT-F4 AP. 16.48 2.20 13 PP 25.4 psia Peak 13 press. 16.54 2.26 14 PP 25.1 psia Peak 14 press. 16.84 2.56 LC PP 24.6 psia Peak LC press.
- 16.92 2.64 LC-LP MD 1.8 psi Max LC-UP AP.
17.02 2.74 FA PP 24.7 psia Peak FA press. 17.15 2.87 FA-LC MD 0. 9 psi Max FA-LC AP. 17.51 3.23 12 PP 25.3 psia Peak 12 press. 17.51 3.23 LC-DE MD 0. 6 psi Max LC-DE AP. . 17.61 3.33 Il PP 25.5 psia Peak 11 press. 17.62 3.34 LC-FA MD 0. 8 psi Max LC-FA AP.* 17.67 3.39 UC-FA MD 2.5 psi Max UC-FA AP. 17.77 3.49 UP TB 1.75 s IBT = 1.80 s. 17.87 3.59 UC PP 25.4 psia Peak UC press. 17.90 3.62 UC-LC MD 2. 0 psi Max UC-LC AP. 18.17 3.89 FA-DE MD 0. 6 psi Max FA-DE AP. 18.30 4.02 UP PP 25.5 psia Peak UP press.
, 18.83 4.55 LP PP 24.9 psia Peak LP press.
1 I i i l ll
TABLE C-I (Cont.) EVENT SEQUENCE FOR CALCULATIONS 1, 2 & 7 CALCULATION 7
. Time. s - - - - - - - Event - - - - - - ,
Problem at Como Synbol Value units Comments . 4722.61 0. 0 UC IB 4.65 mis /s 1st UC & calc burn.
-23.72 1.11 UP IB 10.18 mis /s Propagatior. from UC.
23.75 1.14 UP-UC LO - - TDD locked oper.. 24.63 2.02 UP TB 0.91 s IBT = 1.80 s. 26.42 3.81 FA IB 3.34 mis /s Propagatior. from UC. 26.59 3.98 FA-DE MD 1. 3 psi Max FA-DE AP. , 28.35 5.74 FA TB 1.93 s IBT = 2.56 s. 30.22 7.61 LC IB 11.12 mis /s Propagatior. from FA. 30.49 7. S S UP IB 2.74 mis /s Propagttier. frem UC. 32.85 10.24 UC TB 10.24 s IBT = 9.70 s. 32.97 10.36 LC TB 2.75 s IBT = 3.80 s. 33.79 11.18 11 IB 7.95 mis /s Propagation from UP. 34.49 11.88 11 TB 0.70 s IBT = 1.40 s. 4 34.52 11.91 11 IB 2.16 mis /s Propagatior. from UP. 34.53 11.92 Il-UP MD 1. 8 psi Max Il-UP AP. 34.56 11.95 LC-LP MD 4.7 psi Max LC-LP AP. 35.53 12.92 11 TB 1.01 s IBT = 1.40 s. . 35.92 13.31 12 IB 6.34 mis /s Propagation from 11. i f
. i__l.____
._ 7
( l
- i l
5 l l { l TABLE C-I (Cont.) ! EVENT SEQUENCE FOR CALCULATIONS 1. 2 & 7 j CALCULATION 7 (Cont.)
- - Time. s - - - - - - - Event - - - - - - !
Problem at Como Symbol Va lue urits remnents 5005.29 0. 0 UC IB 4.63 mis /s 2nd UC burn. _ 06.39 1.10 UP IB 10.79 als/s PropagationfromNC. 06.47 1.18 UP-UC MD 2. 2 psi Max UP-UC AP. 06.62 1.33 UC-UP MD 1. 5 psi Ma x UC-UP AP. 07.28 1.99 UP TB 0.89 s IBT = 1.80 s. 09.09 3.80 FA IB 3.35 mis /s Propagatfor. from Ur. 11.03 5.74 FA TB 1.94 s IBT = 2.58 s. 12.90 7.61 LC IB 10.26 mis /s Propagatier. from FA. 12.90 7.61 DE IB 5.00 mis /s Propagation from FA. 13.05 7.79 UP IB 2.68 mis /s Propagation from UC. 14.37 9.08 DE-FA MD 7. 4 psi Max DE-FA AP. 14.89 9.60 DE TB 1.99 s i IBT = 2.42 s. 14.89 9.60 DE PP 38.0 psia Peak DE & cale press.* 15.45 10.16 LC PP 32.1 psia Peak LC press. 15.56 10.27 LC TB 2.36 s IBT = 3.80 s. 15.57 10.28 FA PP 32.1 psia Peak FA press. . 16.38 11.09 11 IB 8.80 mis /s Propagation from UP. 16.64 11.35 UC TB 11.35 s IBT = 9.70 s. 16.69 11.40 13 PP 33.7 psia Peak 13 press. 16.72 11.43 14 PP 33.8 psia Peak 14 press.- 16.73 11.44 LP PP 33.9 psia Peak LP press. 17.05 11.79 11 TB 0.70 s IBT = 1.40 s. 17.11 11.82 UP-11 MD 13.0 psi Max IDD AP.* 17.11 11.82 11 IB 2.22 mis /s Propagation from UP. 17.15 11.86 12 PP 34.8 psia Peak 12 press. l 17.16 11.87 UP-LP MD 9.1 psi Max UP-LP AP. 17.16 11.87 Il PP 35.0 psia Peak Il press. 17.17 11.88 UP PP 35.1 psia Peak UP press. 17.24 11.95 UC PP 33.8 psia Peak UC press. 18.42 13.13 UP TB 5.34 s IBT = 1.80 s. 1 l
4 7
., 60-TABLE C-I (Cont.)
i EVENT SEQUENCE FOR CALCUL4TIONS 1. 2 & 7 C4LCULATION 7 (Cont.)
. . Ti me , s % - - - - - - Event - - - - - -
, Problem M Como Symbol Value units comererts
?.5d18.52 13.23 12 IB 6.26 mis /s Propagatfor. from II.
, i 19.02 13.73 LP-LC E 4.8 psi Max LID AP.* 19.40 14.11 12 TB 0.88 s IBT = 1.40 s. 19.57 TB 2.46 s 14.2,6 Il IBT = 1.40 s. 19.68 14.39 UC-LC. m 6. 7 psi E x UC-LC AP. 19.68 14.39 FA-LC E 2.1 psi Max FA-LC AP. 19.86 14.57 UC-FA 2 7.1 psi 4 x UC-FA AP. 19.89 14.65 LC-FA 2 1. 6 psi % x LC-FA AP. 25.75 20.46 LP-UP MD 1.2 psi Max LP-UP AP. 1 s w 5308.90 0. 0 UC IB 4.50 mis /s 3rd UC burr..
; 10.00 1.10 UP IB 9.25 mis /s Propagatior. from UC.
10.94 2.04 UP TB 0.94 s IBT = 1.80 s. 12.70 3.80 FA IB 3.30 mis /s Propagatier. frort UC. 14.63 5.73 FA TB 1.93 s IBT = 2.58 s. 16.50 7.60 LC IB 9.40 mis /s Propagatior. front UC. 16.50 7.60 DE IB 4.03 rel s/s Propagatior. frort UC. - 16.86 7.96 FA-UC E 1.4 psi I Max PA-UC AP.* 16.91 8.01 UP IB 2.78 mis /s Propagatior. frort UC. 16.96 8.06 LC-UC @ 1. 0 psi Max LC-UC AP. 18.54 9.64 DE TB 2.04 s IBT = 2.42 s. 19.23 10.33 LC TB 2.73 s IBT = 3.80 s. 19.43 10.53 UC TB 10.53 s IBT = 9.70 s. 6136.79 0. 0 LC-DE MD 0. 8 psi Max.LC-DE AP. 1 i s
?w L _s
' f
3 . I i l TABLE C-I (Cont.) EVENT SEQUENCE FOR CALCULATIONS 1, 2 & 7
~
NOTES Problem - Time from beginning of calculation of significant occurrences, such as, maximus AP. ! M - Time of event from an initiating event that precedes significant occurrences, such as, maximus AP. -' Comp - Compartment (s) involved in event where: DE E Dead-Ended region; FA E Fan-Accumulator rooms; Il I top 1/4 of Ice; 12 I second 1/4 of Ice etc.; LC E Lower Compartment; UC E Upper compartment; UP I Upper Plenua. Symbol _ - Spnbols used to describe event where: IB I Initiate Burn; LO E TD0 Locked Open; MD E Maximum pressure Dif'ference; i PP E Peak pressure; TB E Terminate Burn. " Value - Values of parameters associated with event. Units - Units for value associsted with event where: als/s E initial burn rate moles /s; s I seconds between burn initiation and burn termination. Coments - Comments regarding event where: IDD E APE Pressure Difference; Intermediate Deck Doors; IST I User specified born time (s); LID E Lower Inlet Doors; taportant TDD consideration. design E Top Deck Doors;
- designates a pressure or A P that may be an 1
l 6 ..
l l 1 l l l C/8/520/RC/TS/8Rr o-NITROGCN 0-STEfM l 9 , ,,, , , , , g.. i a b.
,#d d
g. 4 d = a =
; - = =
; a l : : :
s.e mute ame.e zus.s gaus.s mate asso.s vann.e sens , Tittl51 l Fig. C.1.a Calculation No. I lower compartment (volume 1) nitrogen and steam mole fractions. l l l l
9 C/9/520/RE/TS/8Rr o-5 TIMES O UGEN 0-5 TIMCS HTDROGEN
. 9 4
ir
- . g-s- f
- b - .
U, . E g-5 3- -
~
o a a a a e a a a d, e a n , e a a s.s min.s mis.s mis.e mio.o saim.o exis.o min.s ses.e Titt 51 Fig. C.1.b Calculation No.1 lower compartment (volume 1) oxygen (x 5) and hydrogen (x 5) mole fractions. i l l I o- -
l l 1 1 l i i l i C/B/S20/RC/TS/8Rr a-NITROGCN 0-STER 1 9 g.. l
- g. .
F g-g . . . . . . . . . . . . . . . s.e sais.s amin.o mis.o do smio.o suis.e sein.e man o TirCISI Fig. C.1.c Calculation No. I lower plenum (volume 2) nitrogen and steam mole fractions.
C/B/52D/RC/TS/88F o-5 TIMCS OXTGEN o-5 TIES HYDROGEN 9,, 2 4- - ( f . n g.. ur . g . 2 . I - - -
- , : same aus.e nas.s seus.o a.e asus.s anos.o ans.o e s.o TIE 151 Fig. C.I.d Calculation No. I lower plenum (volume 2) oxygen (x 5) and hydrogen (x 5) mole 3 fractions.
l 1 1 1 1 i C/B/520/RC/TS/8RF o-NITROGEN 0-STERrt cip , 2-- . 2-- . I - , s.e inito mim.e asie.o mio.o sein.o emin.o min.e umn.e TitCISI i Fig. C.1.e Calculation No. 1 upper plenum (volume 3) nitrogen and steam mole fractions. f ., I
,)
# /
1 __ _ _ - _ _ _ _ - 5
C/8/S20/RC/TS/8Rf' ' o-5 TITS 0xtGEN o-5 TIMES HYDROGCN
- T . . . . .
3 .
. h t
j.. 1 Iir
\
rt I lf . pV
\\
lj liy
- d--
k6 . I ' n Qll d . lr -
= ! , .
d- - l . l! i 1 i . d. o.s nais.o ase.s seio.o ein.o wis.o esis. mis.e same TittlSi l Fig. C.1.f Calculation No.1 upper plenum (voljune 3) oxygen (x 5) and hydrogen (= 5) mole fractions. d , 7 s') J e-
se C/8/520/RC/TS/8Rr a-NITROGD4 o-STERr1 9 . . . . . . . . . . . . . . . a1Q . a.. g.- 4* . f s.. ! I
- .. :- w .
l i 4 . 5. . . . . . . , , . . . sais.e mime aus.e 1 s.s sein.e min.o mis.e ein.o sein.o TifEIS) Fig. C.I.g compartment (volume 4) nitrogen and steam mole Calculation No. I upper fractions. l
I e i C/B/S2D/RE/TS/8Rr g o-5 TI!1CS OXYGEN 0-5 Tl!CS HYDROGEN 9
\ . . . . . . . .
- g. .
g.. 2 .
- 4. .
sen.s mio.s mis.e ais.e win.o seks mis e s.e uda e TifEt$1 Fig. C.1.h Calculation No. 1 upper compartment (volume 4) oxygen (= 5) and hydrogen (= 5) mole fractions. t l l
-- - e. _yn. ,
l l l i l l C/8/S20/RC/TS/8T o-NITROGEN 0-51EAPI ,
't. . . , , , . . . . . . . . .
gi r e. 4 s. g g. . h g . N .
- g. .
j . 2 haitti , , s.e mis.e amin.o mis.s mio.o sein.o enia.o nia.e ame.e Tite 151 Fig. C.I.i Calculation No. I top 1/4 ice (volume 7) nitrogen and steam mole fractions. i i i I
C/8/52D/ME/TS/8RP o-5 TIMES OtYGEN o-5 TIMES HYDROGCN
- . , , , , , , h.
g.. g..
- j. .
8 . g . 2 . e - - d: ' nos.s seas.e sess.s ens.e a.e tsun.s aan.e zus.o aan.o TirEISI 1 l Fig. C.I.j Calculation No.1 top 1/4 ice (volume 7) oxygen (x 5) and hydrogen (x 5) mole
' fractions.
1 I I 1 1 ! l l l l l
i C/8/S20/RC/TS/8Rr o-IGN!TCR RSS.-lPI201 o-8RT. TRS. CS.-LC(21 *-CBL. IN CO@.-LCt22
+-CBL. IN CONO.-UPI23 l
g-- . sg.. . g . .
):f .
gk'
~
~
y m ~ f
~
(< . a -_ e . d a a a a a a a a a a a a a a a s.s isin.s amis.e mis.o eso.o mio.o esis.e mis.e sumo TirttSI Fig. C.1.k Calculation No.1 innermost temperatures for (e) igniter assembly in the upper plenum (UP, volume 3), (o) Barton transmitter in the lower compartment (LC, volume 1), (4) cable in conduit in the LC, and (+) cable in conduit in the UP. l l
l i 1 1 l l i C/8/520/MC/TS/ W o .6671N RL-LCl251 a .250lN STL-LCl261
. o .1251N RL-LCl241 . . . . . . .
g . . . . . . g.. sg. . y;.
. m e ~. .
E_ (< a
. a a a e a a e a a a
, m is.s a a d a a esis.e esme sein.s mis.s mis.e eis.e min.e s.s TitEISI Fig. C.1.1 Calculation No. 1 lower compartmenC instrument temperatures - (e) 0.125 in.
aluminum, (o) 0.667 in. aluminum, (a) 0.250 in. steel.
/:'
i ej
}
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, t '. - - . - *- _ ~ , .- _ __ _
4 i C/8/523/RC/T5/8M' o -YOL.tPC 1 W , , . . . . . . g.. .
- g. .
g.. .
. u o G
k-- % - do n I a a e a a a a a a e e a a a s.e mis.e aus.o mis. ein e sein.o asia.o nin.e sus.e , Tlft:51 l Fig. C.I.m Calculation No. I lower compartment (volume 1) pressure. O u
I l I C/B/SEMIC/TS/sW' o-VOLUPC 2 ( , , , , . . . . . . . . . . g.. g.. . g.. . u
.. Nh .
do a a a a I a a a a a a a a a a a s.e mie.s amis.e min.e ein e sein.o asio.e sein.e sums Tittt$1 Fig. C.1.n Calculation No. 1 lower plenum (volume 2) pressure.
-. y -- .--- - - - - - - - - -- -- + - - e n
1 c C/8/$2Dntr/TS/IIRr e-V0ust 3 i , , . . . . . . g.. . l . . o g.. . i" n
. . p
] h .
g< > . I a s e a e a a e a e a a e a e s.e sein.o amin.o suis.s ais.o suino suis.e suis.e aute fifCISI Fig. C.I.o Calculation No. I upper plenum (voltpe 3) pressure. 0 I l l
)
l l l
se C/8/$ 2/RC/ W o-Y0Lurt 4 g , , , , , , , , . , , . . . 9 g . g.. . 9 - R" n 1
.. h .
. i 1
JA >
=
. 1 1
h a a e a a a a a a a a a e a a s.o mio.e amie.o mim.e ein.o suis.o ase.o nia.e sun.e TIPEISI
]
Fig. C.I.p Calculation Na. 1 upper compartment (volume 4) pressure. 3 l l l l l 1 [ I
k-
. ,/ Il 1
~
J C/B/SzintE/TS/lRF o-Y0uPC 5 g . . , , , . . , , . . . , , M.. g .
- g. .
o g.. . k . do n a a e e e e a a a a e a a a s.e mis.s mis.o min.s eine sein.o sak.o mis.o aus.o TifEISI Fig. C.1.q Calculation No. I dead-ended region (volume 5) pressure.
, ,, , - . , ,p.w-. e n -+-n-- ,,,...,.w.e . - . - - - o e ~-e - --- ,. -- 4e ,-.-,e,n--,-r,,-,
C/5/$20ntE/TS/W o-vtLtFC 6 g . . . . . . . . . . . . . . . g.. . g.. .
- g. . .
o
- h. -
go . W . . . . . . . . . . . . . . . s.e amis.e amin.o suis.e ein.o sein.o esis.o mis.e aus o TitCESI Fig. C.1.r Calculation No. 1 fan-accumulator rooms (volume 6) pressure.
c CMntC/TstlW e -v0LLrt 7 ( . . . . . . . . . . ..... g.. . 9 - R"
- g. . .
o gn- -
> % Ohh :
p,o - I e a e a a e a e a a a a e a a s.s wis.s amis.e mio.e es's.o mio.o amin.o min.e sun.o TIEISI Fig. C.1.s Calculation No. I top 1/4 ice (volume 7) pressure. e 9 1 .$ # .' "- a _ , - - -_w--
1 l t 1 l 1 C/B N
, e-YOLUPE I l . . . . . . . . . . . . . . .
g.. . e g.. . g.. o . lg... . e g.. . I e g a e a e a a e a a a a a a a a e.e sein.e amis.e smio.e < mis. sein.e suis.: sein e suus.e TIFEt51 Fig. C.1.t Calculation No. I lower compartment (volume 1) temperature.
,,- +- - ,
l 1 i i I C/8/52Vitt/TS/8Rr
, e-v0L.trC 2 l . . . . . . . . . . . . . . .
g.. .
- g. .
ll g.. . g.. . _OMhW Q
.o
- g. . . . . . . . . . . . . . . .
s.e esis.e mio.o mie.e ein.o sein.o omin.o mia.s aus.o . TIftt51 l l Fig. C.I.u Calculation No. I lower plenum (volume 2) temperature. l l l l l l 1 1
e. C/8/52)Mtt/TS/tF
, a-vourt 3 l . . . . . . . . . . . . . . .
g.. , l
. ll /
9 ;I i I l- j ll
!! I!
[i" . i
- ll - .
l l a;
. i g..
i . l' l bl lC q asis.e esia.e seine esis.e sein.s aus.e s.e asia.s amis.o TIPEt$1 l l Fig. C.1.v
- Calculation No. I upper plenum (volume 3) temperature. l I
l
~ - - - - -- - .-_ _ _ _ . - _ _ ,_
oo
- i C/8/52 Vite /TS/gRr e-Voupt 4 g . . . . . . . , , , , . . .
t-I
\ g- -
1 l I I
'. ? i R"
e g.. l
. i.0 .
wi il l u ! MI Yf
~
I 1
. a a a f
g i a a a a a a a a a a a l s.e sais.s mis.s ais.e gale.e seino seio.s nin.e male TitCISI ( i Fig. C.1.w , Calculation No. I upper compartment (volume 4) temperature. 4 L j - i i t l w-
C/8/521/E/75/EF e-vourt 5 l , , , , , , . . . . . . . - -
. ~
g.
. ~
g.. g.. n '
~~
hh;(
~
e 1 { 3 , , . . . . . . . . . ain.e asie.e eine suisse sun.e suuke aume e.e sein.e T!!EISI fig. c.1.x Calculation No. I dead-ended region (volume 5) temperature. i
. */
- /
C/B/$2)N
, e-vourt 6 l . . . . . . . . . . . . . . .
e
- g. . .
f e
- g. .
- s 1 e
- g. .
Ie g. . e
- g. .
tilith ti 1 _ r t ; e a a a a a 4 i d e a a a a a e e a a , a.e asia.e amis.e smie.e ein.e sein.e emio.o sais.e aus.e ' l TifEtS1 Fig. C.I.y Calculation No. 1 fan-accumulator rooms (volume 6) temperature.
-.-- . -- 4y ')"--
,y= , gg .
t
\.
~'
! i
(
) l 1
' i l
i l C/8/52nt/TS/er e-voust 7 (.. e
- g. . .
M Ig..
- g. . .
e
- g. . .
f
. MMul -- ^
s.e suis.e asis.e suis.e guis.e saiike esis. se ama.e TIFEtS1 c. Fig. C.1.z Calculation No. I top 1/4 ice (volume 7) temperature. i r .-, :,.. , ,
(.
=
's i
w
\
1 i
.~..
j;%
- h. '
'4 I
- 1 C/8/52DR/RC/TS/8RF
, e n-NITROGEN 0-STERN
- o e
g.. .
\
t, . un% : , t' 2 - Y}- g w_= -
- g. .
i . . g . . . ............ s.e mim.e asio.o mim.e ein.o sim.o esis.o nin.e aus.s TitCISI ';
' 5 Fig. C.2.a Calculation No. 2 lower compartment (volume 1) nitrogen and steam mole ,
fractions. ,
' )
1 o
. 't ;;.7., . - _ .i.., .
,e
.w.-
i
/
1 t I I C/B/52DA/8tC/TS/8Rr o-S Tit 1CS OtTGEN o-5 TIT S Nf0ROGEN s a
", J .
, g. ,
= .
6
- EaJ g.
d> n - t a- ~ l e e d: e e e e i a e , a a l l s.s senO.0 J550.0 300.0 gue.0 95 2.0 3500.0 34D0.0 OED.0 TITISI Fig. C.2.b Calculation No. 2 lower compartment (volume 1) oxygen (x 5) and hydrogen (= 5) mole fractions. 9
. . . . . . _ . . i _ i. . _ . . ... ., i -
ii
/
/-
. t :
l'
/
w/ C/B/520A/RE/TS/BRr c-NITROGEN 0-STEM a-i r a-m . a- - _ - = y M . s-d i
. a a a n s e i n, d a a e n a n, _
us.o ano.o ano.o sum.o m m.s son.o e.o ion.o ano.o T!fEISI Fig. C.2.c I Calculation No. 2 lower plenum (volume 2) nitrogen and steam mole fractions. )
--, , - - - - - . - , , , , , ~ ~ - , . , . . . , - , ,. ,+, - - - - , - . - - - .
i l l l I C/B/S2DA/RC/TS/8Rr , o-5 Tirts cxYGCN o-5 T[MCS HYDROGEN 9 i - 6 y d- , E .
- g. .
M. g
. e a a a a a a a a a d- 1 e a ime.o mie.o zio.o mio.o mio.o suio.e m m.o asso.o e.e TifEtSI Fig. C.2.d Calculation No. 2 lower plenum (volume 2) oxygen (x 5) and hydrogen (x 5) mole fractions.
O M"- .a
Os C/B/520A/RE/TS/8Rr o NITROGEN 0-STERM o , , , , , gu g.. E . u Y d-G'. N.. g . smedar...- g - . . . s.e unio.o mic.o mic.o d.o mic.o seio.o amo.o som.o Ilt1CESI Fig. C.2.e Calculation No. 2 upper plenum (volume 3) nitrogen and steam mole fractions. O
l l 1
~
l l l 1 l 1 C/B/S P /RC/T5/ err o-5 TIMES HYDROGCN 3 0-5 TIMES OrTGEN N , . . . , , , . u P
. - a -
4: a su n.e ansLo e.e sono.o anuc.o zu.o ano.c sum.o esmo.o TIMCISI Fig. C.2.f Calculation No. 2 upper plenum (volume 3) oxygen (= 5) and hydrogen (= 5) mole i 1 fractions. l l 1 1
I 4 C/9/52@/RC/TS/W e-N!TRO*iN o -STEM e . . .
~
ci1( J -
. A_ _
?. ..
C . u .
'g. ,
d -
'.t..
=
+ f.
enn.o ass.s austo use.s enc.e smo.o s.s esse.s ase.o TifE151 Fig. C.2 3 and steam mole nitrogen Calculation No. 2 upper compartment (volume 4) fractions. l - i
)
C/8/S P /RC/T5/8Rr 0 5 TIMCS HYDROGEN Ao-5TirtS0xTGEN g e 6-( w - Ys-d.
- g. .
I; a a ses.o sus.e m e.e se,e e.e seus.e asso.o mas.e eso.e 11tEl51 Fig. C.2.h Calculation No. 2 upper compartment (volume 4) oxygen (= 5) and hydrogen (= 5) ' sole fractions.
l l C/S/527/RE/TS/890 0 NITROGEN O-$7EM o , , . , , , f11 g. w . u - Y a-g . g. e i _ s: . , , ano.o sec.o anc.o seno.e aus.e s.o sent.o anc.o see.s TIMEISI Fig. C.2.1 ! Calculation No. 2 top 1/4 ice (volume 7) nitrogen and steam mole fractions. O 4 b-
I i l i i l l l l C/B/520A/RC/TS/89r e-5 T!E S 0xyGtw o-S TINCS HYDROGEN N.
~
o . . .
*O.. -
O"
- E .
u h;. . d* .
- l 9.. .
- g. . . . . .
emo o l mio.o som.o sein.o feixi.e s.o isic.o mis.o nic.o TitCISI Fig. C.2.j Calculation No. 2 top 1/4 ice (voludh 7) oxygen (x 5) and hydrogen (= 5) mole fractions. d: 1
. .)
# /
l , - ,_. ~ . .- _.
l C/B/52DA/RC/TS/89~ o lGN!TER ASS.-UPI201 o-BRT. TRS. CS.-LCl21
- CBL. IN C00.-LCl22
* * - C,BL . IN C.OC.-U*t23 , . .
g . . , , , g. L g.
. g.
E - (d
. f .
- s.s iam.s muc.o anc.s anc.o susc.o aus.o res.o enum.o it'EISI Fig. C.2.k Calculation No. 2 innermost temperatures for (e ) igniter assembly in the upper plenum (UP, volume 3), (o) barton transmitter in the lower compartment (LC, volume 1), (a) cable in conduit in the LC, and (+) cable in conduit in the UP.
l l l
) . s f
i C/B/523R/RC/TS/ W o .6671N RL-LCl251 a .250tN STL-LCl261 o .125tN AL-LCt241 o . . . g . g. t g. g . ' w y'
^
E .
. h'.
F ,' 9 . 34 e . . . . . . . . a . . . . . . unio.o anko mis.s aan.s amio.o nic.e amo.o s.s senc.e TIMCIS) Fig. C.2.1 temperatures - lower compartment instrument Calculation No. 2 aluminum, (f.,) 0.250 in. steel. (*) 0.125 in.saluminum, (o) 0.667 in.
i l l l l i i l
-100-C/8/S D /RC/TS/ W e-v0LUPC 1 gi . . . .
. g.
M - R-g.
~
- g. .
e . do
. a e a a k a i e i e a a a a a e nia.e aus.s min.s asio.s sim o aim.o min e esis.o e.s 11051 Fig. C.2.m Calculation No. 2 lower compartment (volume 1) pressure.
1
- l I :
1 i I j i l I i j.- . , . - - . -
,, t , - . , , . , , - --
-y.' , ' ~ _ - . ,e', -, -- - -.
-101-C/8/523R/MC/TS/8mr e-vtn.t# C 2
( , . . . . . .. ....
- g. .
- g. . .
- g. .
1. kN A ! . so . W . suo o ano.s unn.o enn.e m e.e same
- s. ines.e amo.o Tittl51 Tig. C.2.n Calculation No. 2 lower plenum (volume 2) pressure.
t 9 _. _ . . . _ ._y _.,_ _ , , - . _ , - _ , _ . . , # w _ ,. , . . _ . - , ,
4 t
- t l} 102 J
w/ C/8/52PLME/T5/gRr e.v0L.UPC 3
, e .
s, , . . . , , . . . . . . M. g
- g. , ,
K-
!. j 7==% :
- v. .
1 . . j f( > a a a a a l N a a a a a a a i a a a.e sais.s asio.e sie e eim.s use.e emin.e mite oms,e 11tCt51 i t Tig. C.2.o Calculation No. 2 upper plenum (volume 3) pressure. O l 4 I I f I
, , - , . - -c.. ..,-.,n--- - . . . - - . - , , - , - . . - . . . . - , , , . - . . - _ . , ,, -,, - , - - - - - - - . , . - . . - . -- . . , -
-2., a -- . - - -a a - -, , - a 5
-103-P O
4 4 4 1
- C/8/32FURE/TS/W a-VOLUME 4
. a e '
y , ,
, e e > >
z-l . 4 1 m . Gj. E k e
- i f4>
i . J I e a a a a a
. , . i i I i , , , ,
T T F I summe es.: ante Y W Y see.s sum.s sum.s J s.e issc s . am.s TI C 51 I i Fig. C.2 9 c j Calculation No. 2 upper compartment (volume 4) pressure. s t I 4 1 i i 1 )
-104-C/8/520RnE/TS/8F e-v0Lurt 5 9 . . . .
3 . . . M.. g
- g. .
. g-1.
v-. go . g . , , , , , , , , . , ,,,, e.e sais.s amin.s mim.o eine nic.o sais.e nus.s ans.e TitCISI Fig. C.2.q Calculation No. 2 dead-ended region (volume 5) pressure. i b u
4 l l
-105- l I
1 i l 1 1
- l 1
1 C/8/520R/RC/TS/8Rr e-v0LUMC 6 g . . . . . . . 9 . g .
~
g..
~
9 . g .
~
- g. .
f1 >
~
a a a e e a . N e a a a e a e a a s.: isso.e ase.o suo.c eso.s use.s men.e ans.s ses.s T1011 Fig. C.2.r Calculation No. 2 fan-accumulator rooms (volume 6) pressure. h- -
-106-C/B/52DMtC/Ts/gRr e -vourt 7 i , , , , , , . . .
g.. . g.. . 9 g . . h_. go . a e n e a e a e a e a a a a a s.e mim.o ain.o mie.o mio.e mim.o amis.e mio.e suas i TitCt51 1 Fig. C.2.s Calculation No. 2 top 1/4 ice (volume 7) pressure. G
. - ~ , . . , - - - , . , . . , - , , - - . , ----,--,-r, , , - , _ . , - - - - - - - - - .,----w- --r------, - - - ,
-107-C/8/527t/RE/TS/tPtr
, e-v0LUPE 1 l . . . .
g. M - g.. e . lg..9
- g. .
- g. .
p / -- . it g . . . . . . . . suic.e mio.o unio.e esis.o seine semi.e e.s isim.s amic.o TifEt51 Fig. C.2.t Calculation No. 2 lower compartment (volume 1) temperature.
)
-108-l 1
.e l C/8/S20R/RC/TS/ F o-VaupC 2 l , , , . , , . . , , . . .
e -
. 2-g.
g. os
- g. ' [ :
s [
=o - * - . . -
e > > m m.e sum e s.e mm.e asso.o suo.o ano.o sa:.o sum.o TitCISI Fig. C.2.u Calculation No. 2 lower plenum (volume 2) temperature. i f
l l
-109-I 1
C/8/S D /RE/TS/gRr
. o-YOLUNC 3 9
E f$' . E J . g. r l' ' .
.1, E".s o ' son: o man.o mie o eso.e mim,, ,,;,,, ,,g , , , , ,
71rC151 Jig. C.2.v Calculation No. 2 upper plenum (volume 3) temperature. i
- l e .._ _ . - . . . _ - . . , _ . . - - . . . - . . - . _ _ . . - . - _ - . - . . . _ . . . - _~ . _ . . -
1
-110-i c/s/ sam /Renstant
, s-v0Lu t 5 . .
g g.. l "
=
( .. {l" , e
- g. ,
g .. l g & A 8'E8 I y, ' d,, ' mie e a!s. a n.e s e.e suis.e Titt.151 Fig. C.2.w Calculation No. 2 upper compartment (volume 4) temperature. 1
-111-I i
I C/B/52M/MC/T5/8RIF
. o V0t.UPC 4 , . . .
g , , , .
=.
y i . g. a {I' . g.. Q% ' 8;( - A A E O I g a g a e a e mie.e aio.e mino esis.e me.e same e.e esis.s amis.o Titt 51 l 1 Fig. C.2.x Calculation No. 2 dead-ended region (volume 5) temperature. I 6 9 , - - , , , . , - - - -, .- -
-s - -, -- ,. s. -. - ---- - - - - - --- -
-112-C/5/$2)RMIC/T5/gRr
, e-V0uPC 4 l , , , . . . . . . , , . .
e
! N_-
I g. u g. .
)
- l f e .
g. g. 7
. a a a a a a a a gj a a a a a a a
- s. sais.e mis. suis.e sic.e usin.e min.e suis. suute Tl'Et51 Tig. C.2.y Calculation No. 2 fan-accumulator rqoss (volume 6) temperature.
e I 4 I i I i w.
. -113-9 C/5/52M/ME/TS/8mr
, e WC1.LFC 7 l . . . . , , , , , , , .
9 . l-j- f- - E - Wo . E
.Fd.
g- - I _ d - ines.e aus.o mas.e eso.o sum.e ses.e m m.o ens.e e.s TitEtSi Fig. C.2.s v Calculation No. 2 top 1/4 ice (volume 7) temperature. i
l l .
.l>
-114- ,
/
. /
.. /
J C/B/520/REITS/89r/69 e-NITROGEN o-5TE8Pt e , . . . . 1 - e. g , i
- e y a- -
E . k 4
- J g g.
E,!
.e ,
*g d-o .
e - = ' l e "
- : sum.o suuke suuke ame.e as:.c ann.o sun.e e.e mus.o TifEt31 Tig. C.3.s and steam mole lower compartment (volume 1) nitrogen Calculation No. 7 fractions. .
O n
- . . - _ _ - _ _ _ - - - - . ., _ .t - - - . . -
$ l f>
-115- 1
\
. '. l t ,
) !
C/B r520/'tC/TS/F/68I o-5 TIMCS OITGCN o-5 TIMCS HYOR03CN 9 , , , , , , . , , , , , g. Y6
.. f I
- g. . ,
h Q.. e g: . . . . . . . , , .' i . . o.o ism.o muc.o not.o aum.o smo.o ess.o Mus.o ammo TIMCt51 Fig. C.3.b Calculation No. 7 lower compartment (volume 1) oxygen (= 5) and hydrogen (= 5) sole fractions. 3 6 g l
. _ _.e._.__.: _ .,
., l l
-116- I i
C/B/523/8tt/TS/8er/6n c-hliROGCh o $7CAN g. l 3 , g g.
- g. ,
( a!
. 8 i ' a a e n
,a e a a a n I ' '
aus.o ans.e same e.e use.e men.o me.o ano.e unc.s 11 0 51 ,s I Fig. C.3.c . i Calculation No. 7 lower plenum (volume 2) nitrogen and steam mole fractions. x e
*[ } .
t [\ s
-a.
l
1
-117- l l
i l l l l l i C/E '370/RC/TS/8F/6* o-5 TIMCS OKIGCN 0-5 TIMCS HIDROGCN
- g. .
4. 2- . l Q 2 . l
\
I. . . . . . . mic.o aic.o us:.o sum.o s.o amo 1 3' - s.e asia.o aNo.o
' TIMCISI l Tig. C.3.d Calculation No. 7 lower plenum (volume 2) oxygen (= 5) and hydrogen (= 5) sole )
fractions. I T'
-118- i i
d 4 C/B/52D/F/TS/94r/6R c NITROGth o-SitH i l .: , , , , . . av Y a-4 . i ul}. m - i s.. same saim.o rain.o anno amie.o mzz.o amic.o s.s saic.c TIMCISI Fig. C.3.e Calculation No. 7 upper plenum (volume 3) nitrogen and steam mole fractions. l . I I 1 b
1 l l
-119-C/B/S20/#E/TS/89"/59 e-5 TIMCS HYDROGEN 3 o-S T!MES 0xTGEN T, , , , , ,
r - o- g
~
gi 6 . E . . u / Y a- i m .
. l i
a l
/
I l . . . . . .
- g. , . . .
susto anc.o anc.o exc.c sec.o ene.o nac.s l s.o isse.n , Tint 51 l 1 Fig. C.3.f Calculation No. 7 upper plenum (volume 3) oxygen (x 5) and hydrogen (= 5) mole e fractions. 9' 1 -
l I
)
I
-120-e.
C/B/520/F./TS/8F/6R c-Nitr#,CN 0-STCfM 61( . 6-6 . w= ' d g..
- /. . . .
asio.o mim. eim.o sein.o emin.o nin.o aan.e s.s iam.o 117C151 Fig. C.3.g and stes:n mole Calculation No. 7 upper compartment (volume 4) nitrogen fractions. D
l l l ). )
-121-l i
l l C/B/520/RC/TS/8er/68 , o 5 TIT S NTO% I.,CN 5 0-5 TIMCS 0xTGCN
*\
N ( .
- f .
. r
[ r'
$ a- .
h;.. k ' 2
/
e ..... d: - s.e ism.c anc.c snac.o em.o scos.o sanc.o nec.s suo.e TINCIS1 Fig. C.3.h Calculation No. 7 upper compartmenk (' volume 4) oxygen (x 5) and hydrogen (x 5) sole fractions. e d . I c)
-/
e p- m _. ,
l 1
-122- l C/B/520/RC/TS/8Rr/69 o-NITR3 GEN o-STCAN
= 5 3 5 8 5 a e a 4 em u u
^
av %,3 - g.. u
$ 6-d 2 .
I a eo man e s.s saim.o acc.o mic.s am. sum.o ein.o T1051 Fig. C.3.1 Calculation No. 7 top 1/4 ice (volume 7) nitrogen and steam mole f ractions. l l f 1 7 -- ,
-123-I i
l C/B/520'E/T5/9F/69 o-E TIMrs OrYEN c-5 TIMCS 6tTDR03CN M_ . . . . T. . . . . ' e 9- . WI v . 2 ly.. l i . f u *
$ 2-- f. -
s j . ( d . l ./ i
> / y (
- i a--
v I l e ..... I g . . . . . . . . . . . mio.o mic.o aim.o suic.o anim.o nia.o anno s.o inic.o TIMEISI Fig. C.3.j Calculation No. 7 top 1/4 ice (volume 7) oxygen (= 5) and hydrogen (x 5) mole fractions. l l I l
4
-124-l l
C/B/S2C/*C/TS/Bor/6A o-lGNITER 955.f!201 o-BRT. TRS. C5.-LCl218-CBL IN COND.-LCl22 9 -t2.. IN CO T.-U'12,3 . . . . . . g , , . 9 . g. M - s g. g . h! = _ gR-E
$4
-s E - - a
=
l susto mue.o zoo.o exa.s smo.o esso.s fees.o s.s isso.c TitEISI , Fig. C.3.k Calculation No. 7 innermost temperatures for (*) igniter assembly in the upper plenum (UP, volume 3), (o) barton '. transmitter in the lower compartment (LC, volume 1), (S) cable in conduit in the LC, and (+) cable in conduit in the UP. l
-125-C/B/520/RC/TS/8F/6R o .125tN ML-LOI241 o .6671N RL-LCt251 a .250lN STL-LCl261 O
p- . . . . . . . . . . 9 . g.. M s g.. h w
", h. ~ ~; - . .
a p . M . g<
. e a a e a
,a a e a a e a s.s isse.s ame.o mm.s em.s ses.s see.s m m.s men.o
' TIMCISI Fig. C.3.1 Calculation No. 7 lower compartment instrument temperatures - 2 (*) 0.125 in.saluminum, (o) 0.667 in. aluminum, (t.e) 0.250 in. steel.
I g. t fl -126-
)
-. /
C/8/S2D/RE/TS/ EFT /6A o-ytt.tyt I 9 , , . . y , . . 9 g 9 . R-G-
- 4 { si- = . -
} ,
so i , , , seie.o ano amm.o mic.o suio.o ain.o suic.o s.o sois.o TirEI51 Fig. C.3.m Calculation No. 7 lower compartment (volume 1) pressure. ; e 9 s - , - , - , , , -e s. .n - - a
-127-J C/B/52D/itt/TS/w/6A o-VOLUPC 2 e . . .
i . . . . . . , , , - a M. g . 9
- g . .
i . .
- g. , .
YW .- _ _ ) ,
?
. do . d o N n i e a e a e e a a e a a a n s.s inc.o mis.o sein.o ain.o saio.o suis.o stim.s suso.o - TITtS1 .4
- Fig. C.3.n
- Calculation No. 7 lower plenum (volume 2) pressure.
I J e i 1
. - . . . . ~ n . . r _,
l i
)
-128-C/8/52D/RC/T5/8Rr/6R o -v0LtF C 3 i . .
= -
. g-g..
. =.
g ll W* 4> i . . . . . . . . s.e sein.o amm.o mic.o aim.o mic.o emis.o niz.s amm.o TIEt51 Fig. C.3.o Calculation No. 7 upper plenum (volume 3) pressure. 9 -- e p , e
-129-C/9/520/RC/TS/ w /6A n-va.tK 4 e . . .
g . . M.. g e - g.. M. g . C
- g. =4" _
j n 55 4) e . i . . . ; . . . . . . met. eente use.s ses.e vues.e aus.s s.e iam.o Juan.e 71 4 151 Fig. C.3.p Calculation No. 7 upper compartment (volume 4) pressure. u
I l i l
-130-C/8/$2D/RC/75/gRr/6A o-v0LtPC 5 9 . . . . . .
3 . , . . 9 . g o . g.. g. e. e . go
. e a a a n , a n a a a s n a
,n see.o geno.s sec.o sus.o sus.s sente s.0 Hec.o anc.c TitEt51 Fig. C.3.q
- Calculation No. 7 dead-ended region (volume 5) pressure. d 9 i _ /L __ _ -. _ ,, -_ - , . . . . _ _ . _ - _ , . .,---__.m. p _ ..--
i
-131-C/8/520/RC/T5/WE/6R o-v0LLMC 6
( , . . , , . . , , , , .
- f. =
o g.. e g.. . B d' - - . _
/ i e
g< >
)
e i W . , . . . . . . . . . . . . . i
- s. iein.o amis.o sein.o mic.o su's.o suis.o nine aus.o I TitCISI Fig. C.3.r Calculation No. 7 fan-accumulator rooms (volume 6) pressure.
I l l l
9 l
-132-C/8/52D/RC/15/ W /5A e-YOLUT 7 e
( , , . . . . . . . 9 - K-o - R"
= .
M-l 1
~
h R- __
} ,
s* 4 > i . . . . . . . . smo.o exe.o s o.o sus.o ano.e men.o s.s tar e m uc.o itTt51 Fig. C.3.s Calculation No. 7 top 1/4 ice (volume 7) pressure. 9
P i 1
-133-C/8/$20/RC/TS/8Rr/69
, a-voufC I l
- f. .
M g.. .
=
g.. . g..
~
A%y%
~
\
d i i e.o scut.o anc.o anc.c anc.o sane.o anun.o nahe eso.e T!rE151 Fig. C.3.t calculation No. 7 lower compartment'Tvolume 1) temperature. J
/
'/
/
, , , , ,,- ., e,. , --
I l
-134-i l
C/8/52D/RC/Ts/3Rr/6R
, e-V0uPC 2 l . .
9 . g. 9 . g. o . g. e . Akkd g.
.o
[ e e a e a a y a a a a e a e e i ano.e snee.o son.o pas.o enum.o e.e senc.s aan.o mec.o T1 0 51 l l Fig. C.3.u Calculation No. 7 lower plenum (volume 2) temperature.
1 l 1 i
-135-
)
i I C/B/S20/RC/TS/8Rr/69
, e-v0LLFC 3 l , , , , , , . . . . . .
=..
g g..
- g. .
g.. . g.. g hLL
.'".s s sein.o asio.o mic.o asio.o win.o esim.o min.e sun.s TITISI I
Fig. C.3.v l Calculation No. 7 upper plenum (volume 3) temperature. l 1
l
-136-i I
C/8/530/RE/TS/8Rr/6A o-VOL.tPC 4 e . li . . . . . . n . g-l- fN" . e . (-- v M
. e a e a a e i e a a a a f n a a sein.o sein.s aan o mio.o mic.e esin.o mim.o s.e ses.s TI C SI Fig. C.3.v Calculation No. 7 upper compartment.(volume 4) temperature.
f I 4-_ g,p- e 7=. - y 9
1
-137-C/B/52D/RC/T5/gRr/69
. o-ta. Lyt 5 g . . . . . . . . . .
g.. . g.. . ll g . . E9 g g..
. f -
a_d e ' - - - ; - a e.e ans.e anc.o ane.e ano.e uns.o anno.e m e.e ame T K ISI l Fig. c.3.x Calculation No. 7 dead-ended region (volume 5) temperature. I l l 1 l l l \ .\
M
/
j/ -338-
/
1 C/8/52D/RC/75/lRr/64
, e-VOLL7E 6 l . . . . . . . . . . . . . .
g..
- g. .
g f. . c . o g.. g..
,r - WLLh .
sum.o suun.e nun.e unus.e e.e seus.e amt.o anc.s anLo T!!CISI Fig. C.3.y Calculation No. 7 fan-accumulator rooms (volume 6) temperature. 1 I e e
. . . --- ,. _----. - _ . - - . . _ . . . - - . . - - . -~ _ . - - - . . -
i
-139- !
B . i I C/8/520/RC/T5/ w /6R
, e-YOL M 7 l , , , , , , , , , , ,
M g.. . i
- g. . .
l $$' l g g.. .
!, e g.
g- . . . . . . . . . . . . . . . e.s sais.o men.o sein.e ain.o mis.s esia.e min.s ame.e TitC151 Fig. C.3.: r Calculation No. 7 top 1/4 ice (volume 7) temperature.
f
-140- ,' APPENDIX D DETAILED TIME VARI ATION OF PRESSURE DIFFERENCES ACROSS DOORS AND FANS This appendix preser.ts detailed variation of the pressure differences ( Al')
i across doors ar.d far.s. Such detail might be used to evaluate the capability of the doors and fans to withstand the mechanical loading resulting from the pressure differences. In particular, detailed transient informatior. may be needed to determine the loading variation with time. 4 The plots are not as detailed as could have been obtained because the AP's for each time step were not saved so that storage limitations were not exceeded. However, the plots presented can be used to estimate the time interval (s) where j more detailed plots are of interest. Ther., the calculation could be restarted with the AP saved to obtain a more detailed plot, e.g. , every other time step. The figure captions include the correct calculated maximum AP and time of occurrer.ce from Table 11. This ir. formation car. be used to evaluate the adequacy of the plotted result and the possible need to rerun the calculation to obtair. f greater detail. Also, the plots show with dashed lines the estimated pressure plot did not show
!. profile to the correct maximum AP when the actual 1
satisfactorily the correct AP. The AP signs may be different in the plots than in Table. 11. This is simply a matter of how the difference is taken. For example. (a) in Table 11 the lower inlet doors (LID) AP loading of interest is between the lower pier.ue (LP) and the lower compartment (LC) and (b) the plots show the LID AP between the LC and the LP, which is opposite to the direction taker. in the table. Plots are not presented for Calculations Nos. 5 and 6. Calculation No. 5 j i has unexplained file difficulties, and Calculation No. 6 did not end normally to produce the plot files required. i 4 e 4 i I r P , . , - > , . , - - - - - - , - - - - - - - ~ - . . - - - - - - - , - - , - . - , , - - - ,. -
- - . - - - - - - - . . - - - , . - - - ~ -
-141-C/l/SE/NC/75/8RP o-YOL3C 1 70 2 a ' = - - - .
. ~
.5 -
D N _ W , 8 ' 0 *. . j . E. k .
- g. .
a * *
- a a a a a a a e enke e s.e
. l w i.e wa., ,3,, A, ,pg, ,;,,, ,,,,
110 51
\
l Fig. D.1. a Calculation No. 1 AP across the lower inlet doors located between the lower cor-partment (volume 1) and the lower plenum (volume 2). A maximum AP of -1.4 psi against the doors (i.e., in their closed position) occurred at 4974.03 s. 1 1 1
I
-142-en C/t/520/RE/TS/8RF o-v0umt 3 TO 7
*. +
f%
. '\ .
W-i
.... \
. s
. . s
* .. ~
ess.: ein.e 44.s da.e da.e dos is.e o n.s err.e , Tittt51 ( Fig. D.1.b Calcula tion No. 1 AP across the ir.termediate deck doors located betweer the upper plenure (volume 3) and the top 1/4 ice (volurne 7). A ma x imurr AP of 1. 9 ps : against the doors (i.e. , in their closed positior.) occurred at 4975.05 s. v_.
1 i
-143-I l
i
. \
l I C/1/520/RC/TS/ W ,' o-vou.rt 3 TO 4 d . . . . . . . . . . . . . . .
*e .. .
+
V.
~
g
#I .
b, f )
&= N f O g
Nf' - t I 4 4- . o 1 s f : : : : : : : ess.: eur.s aus.e as ea.e es.s es.e e s.s eu.e TIEt51 i Fig. D.I.c s Calculation No. 1 AP across the top deck doors located between the upper pler.urr (volume 3) and the upper compartner.t (volume 4). A maximum AP of -0. 5 psi oc-3 curred at 4576.30 s. which was af ter these doors were locked open. ar.d 1.3 psi at 4575.29 s.
, ,. . . . , , , , , ,---,a
*~ -- , , , , . . -
N % 't % k , l
-144-1*
+
4 '
)
l l
,A '
C/1/520/RE/TS/str o-VOLtrE 4 70 6 i s! . . . . .
. ~
4-d.. .
- X ,
- 4
- - 5 4..
, V f es.e wLe o m.e w r.e
- eaa.a e s.a e s.e wa.a ws.e TifEISI d Fig. D.1.d Calculation No. 1 AP across the fan compor:>,r.4 - i c< ed between the upper cor-partment (volume 4) and the fan-accumulator e . ion.. rilume 6). A maximun AP of 1.9 psi occurred at 4974.92 s. i -
,t 1
a r s ? A 1
,- - .. _a.
i
-145- l l
l 1 1 l
\
. C/1/520/RC/TS/0Rr e-vourt 510 5 at . . . . . . . .
~
e. a .
~
i . 4
/s
_e .
\
- m- , y ,_
~
' ~ .
- a. . .
- . 4 4
es.e eu.s eu.s oo.s one es.s es.e ee.e es.s
- TirEI51 Fig. D.1.e ,
Calculation No. 1 AP across the far. coreponents located between the dead-erded region (volume 5) and the fan-accurculator roorts ( volurte 6 ).~ A maximurt AP of ; 0.5 pl occurred at 4645.43 s. f- l
\
i
- 1 f!.
I 3 '/ J
, s ,
e' 1 F 4 -A ' I l I I. . --
-146- )
~
l
- \
C/1/520/RE/75/W
. e-roust ! TD 6 d . . . . . . . .
g..
._ e .
o- - g .
~
- g. .
~
- l :
W l ees. eu.s eu.s oo.e e u.e es.e es. ee.s om.e Tiftt$1 Fig. D.1.f Celculation No.1 AP across the far components located between the lower cer-partment (volume 1) and the fan-accurculator rooms (volume 6). A reaxieurt AP cf 0.7 psi occurred at 4642.90 s. I i k __
)
l
-147-I I
f $. C/2/52tM/RC/T5/W . o-v0LUNC I TO 2
\ . . . . . . . . . . . . . .
7 e . 4
?
w ".. . o.- u
\ l
}* . t j
'I fa \ .
a-. d . 9 Y : : . : l : w s.c w2.s Mrs.s we.c ees.e w s.e m .e gen.e gen.e T!!CtSI Fig. D.2.a Calculation No. 2 AP across the lower inlet doors located between the lower corr-partmer.t (volume 1) and the lower plenum (volume 2). A maximum AP of -1.4 psi against the doors (i.e. , in their closed positier.) occurred at 4475.95 s. i l 1 l l i I i l W
1 i
-145-(
1
. l
_ 1 I C/2/520RntE/T5/8Rr s-v3.tmc 3 70 7 . e . . . . . . . . . . g- . . . .
+ .
#1 e il d' 11
- I 8I
*. I u
N I " b' .I g e_ = u- , t
~
f q- . e .. j edr.e 445.4 +EPe.e edte ede.e eds.e esis.e
' eers.e eda.e TilCISI Fig. D.2.b Calculation No. 2 AP across the ' intermediate deck doorsAP A a:aximurr. located of 2.6 psibetween upper plenum (volume 3) and the top 1/4 ice (volume 7). 4473.96 s.
against the doors (i.e. , in their closed position) occurred at s l I t U
1 l l l
-149-C/2/53DR/NC/75/ W +
s-WOLUT 3 TO 4 g 9 - p . II
. 31 .
- a. .
II '
\\
u ". . I{ W 8I g .
- t. - I 5.- g u '
h l
' ' g
. I .
- g. . ,
+ .
e a a i n e a a a W* a a e a e a e ode,e we,e wi.e eda.e eds.e ode.e ,Js.o esis.e edr.e 11Et51 Fig. D.2.c Calculatior. No. 2 AP across the top deck doors located betweer. the upper plerur-(volume 3) and the upper compartment (volume 4). A maximum AP of -2.3 psi oc-curred at 4474.47 s. which was af ter these doors were locked open. ar.d 3.6 psi j l
' at 4474.01 s.
e k -- - ,-
4 -150-f jl
/
/
C/2/$2DVRC/75/8Rr m -V0u2E 4 70 6 I e.. g E ci- - h. e - - , , . , , , , f-site * . . sis.s mir.o mia.s mis. mia.s mai.e mit.s mis.s TIEts) Fig. D.2.d Calculation No. 2 AP across the far. coreponents located between the upper cor-partment (volume 4) and the fan-accurtulator roosts (volume 6). A nmxieurt AP of 2.5 psi occurred at 5817.67 s and -1.9 psi at 5815.92 s. l l I l l 1
-151-C/2/520R/RC/75/ F ,
e-v0LuPC 5 TD E 9 . o . g.. g ";.. . g . . o. g ,;. . . 3 . . r i 9 , , , , , , , , ,
,4', , , , ,
mas.s mit.e mis.s mis.o mir.e sia.e mis.e mis.e nu.e fifEISI Fig. D.2.e , Calculation No. 2 AP across the fan comper.ents located between the dead-er.ded i region (volume 5) and the fan-accumulator rooms (volume 6). A maximum AP of l 10.6 psi occurred at 5816.21 s and -0.6 psi at 5816.17 s. J l l l l l l l 1
-152-C/2/$2F/RC/TS/8Rr e-YOLtFC 1 TO 6 d . . . . . . . .
g.. I f * ..
-A
- y. .
4 mir.o sia,o mis.s sh.e an.e ma.s mie.o mis.s mis.s 11tEISI Fig. D.2.f Calculation No. 2 AP across the fan coeponer.ts located betweer. the Icwer cer-pa r t e.er.t (volurr.e 1) and the fan-accumulator roons (volume 6). A sa x imurt AP of 0.8 psi occurred at 5817.62 s and -0.9 psi at 5817.15 s.
~
f, ,s - . . , _ ,? p . _ , - - . --. --
I
-153-e, C/3/$20/RC/T5/8Rr/.5C$r e-vourt i TO 2 d . . . . . .
a.. E . b* . co 4-.
. e a e a a e a a a M* 8 a a a e a seks seks soie sais.e suur.e sake aus.e susi.s seks f!MCISI Fig. D.3.a Calculation No. 3 AP across the lower inlet doors locatedA between maxircumthe AP lower of -0.9cer-psi partment (volurce 1) and the lower plenum (volume 2).
against the doors (i.e., in their closed positior.) occurred at $966.45 s.
-154-C/3/520/RC/TS/gRr/.5C$r i e-VOLUPC 3 70 7 . . . .
at . . . . ei- .
- T g- $g f ii De "
E; . fy- .
- y. .
me e ' l me as ama.e as f e s.: esi.s ei.e . ..s TirEISI Fig. D.3.b Calculation No. 3 AP across the internediate A maximum AP ofdeck 4661.31 s. 1.2 psi doors upper plenum (volume 3) and the top 1/4 ice (volume 7 m_--_---
.r. , - - - r---
I l
-155-C/3/520/MC/TS/8Rr/.St5r n-VOLLM: 3 TO 4 9
R 5 5 g e 5 5 5 5 5 5 5 5 3 3 5 4 . e +
.:- - Il 1 .
. g e la E-
. t / .
- g. ...
x . y.. . .
= ,
f : l J en.e ele.e en.e es.e 99.8 FELS 9 58.8 9 55.0 9 88.8 ' TirC151 l Fig. D.3.c Calculation No. 3 AP across the top deck doors located between the upper plerte (volume 3) and the upper comparteer.t (volume 4). A maxieue AP of -0. 8 psi oc-curred at 4757.50 s. which was af ter these deers were locked oper.. and 1.2 p-
, at 4756.74 s.
i 1 e h-
l
-156- !
l
~
C/3/520/RC/TS/8Rr/.505r e-v0Lt.FC 4 TO 6 9 . . . . . . . . . . . . . . . 4
- o. -
~
f .. 9 1 nam.s sas.: saur.a aan.e sauto mia.s sain.s mia.e sta.e T! ret 51 Fig. D.3.d Calculation No. 3 AP across the fan components located between the upper com-partment (volume 4) ar.d the fan-accueulator rooms (volume 6). A maximum AP of 1.5 psi occurred at 6209.69 s. l 1 i t i l I I
-157-C/3/52BME/T5/8t/.5CY o-voust 5 70 6 e .
a . . . . . . a..
'g s _
e - iis e : ki' " 4
- l :
T l sBF.0 W2.9 WILO 945.8 9u.8 SELS sex.e sus.e sus.o TitCL53 Fig. D.3.e Calculation No. 3 AP across the fan cortponer.ts located between the dead-er.ded region (volume 5) and the fan-accumulator rooms (volume 6). A maxircum AP cf 0.5 psi occurred at 5039.47 s. : l' e r! - 1
.j l
l 1 l n -
l
-158- )
i C/3/520/MC/T5/enr/.5C5r e-vouK 1 TO E e . . . . . . . w . . . . . . . . e - n-4
/t .
/
/
. / ', .
- f;- -
4 .
. h-e * ' * * '
- l f w.s sen.s tems tas.s m .s ten.s te os u s.e tes.
TilEI51 Fig. D.3.f Calculatior. No. 3 AP across the far. comporer.ts located betweer. the lower coe-partner.t (volume 1) and the f an-accueulater rooms (volurre 6). A scaxieue AP cf 0.6 psi occurred at 5168.69 s. _ . _ . _ . . , , . - - - -. . - , _ ._y ._ , . . _ , _ - - . , , .
-159-D i
F C/4/520/RC/TS/0Rr j e-v0amt i TO 2
- a. .
h" . c e. . e . a.
* , . . . . . , , , , . Wdi.e eur.e ans.s Mi.e aks an.o ear.e ain.e ers.e Tiff.t$1 Fig. D.4.a Calculation No. 4 AP across the lower inlet doors located between the lower com-pa r t rrer.t (volume 1) ar.d the lower pier.um (volune 2). A r.axireur. AP of -3. 2 psi against the doors (i.e.. in their closed position) occurred at 4965.65 s.
l l w _
. , - . _ _ , . , _ , - . . _ .- . . . ,- . _ - + , y- - - - - -, --
-160-C/4/520/RC/TS/8RF e -vou.MC 3 TO 7 .
- w. . . . .
e a- .
/
ah 's . s g.9 - ~ b - e. g g.. . 1 - b . h . em.e aiko sis.s air.s me s.s s.s ab.e aks d.e Tiff.151 Fig. D.4.b located betweet the C'a l cul a t i er. No. 4 AP across the intermediate deck doorsA maxleum AP of 1.3 psi 4954.92 s. ~ upper pier.urt (volurce 3) and the top 1/4 ice (volurre 7), 6 I
, . . _ - - - _ - . _ ~ . .
.. - - - . . . - _ ~ _ , _
-161-C/4/2/RE/TS/Wif" o-v0Lurt 3 TO 4 e
g . . . . . . . . . . . . . . . g..
+
e et -
.: -- gq W ~
ii ' g t E
' e s.e e s.e es.s st.e ds.s da.e dt.e ein e me TitEt$1 Fig. D.4.c Calculation No. 4 AP across the top deck doors located between the upper pier.ue (volume 3) and the upper cortparte.er.t (volurte 4). A maximum AP of -0.6 psi oc-curred at 4576.30 s which was af ter these doors were locked open, and 1.3 psi 3 at 4575.29 s.
7 , .- ,,, - - - - - , a g - -- .- ,
-162-
/
- f. ..
- l
. 1
.Il l
, 1 J
t/4/520/RC/TS/8RF - o-v0LUME 470 6 . . . 1 . g . w ";. ~ W
- t. .
isc o .. . j game eh.e ars.e y , , . s.s esis.o es.o < sir.s asr.s ais.s Tirtt51 Fig. D.4.d between the upper con-Calculation No. 4 AP scross the fan (volurte 6). Alocated components maximure AP of partment (volume4964. 4) and 87s. the fan-accurnulator roorts 3.2 psi occurred at
.-.m,,. .-. wy , _ - , - - , , - -
,., y , _ -
-163-C/4/520/fC/T5/gRr e-v0LLPC 5 70 6 e . . . . .
w , . . . . . . . .
- a. .
g .:- 4 l\
~
f - I \ e. 5 :
\ i *
\ l
.. \ ,l .
\
' gi . 4
~
.. =
. a a a e a a a a a d* a n a a e a es.s ois.e s.e eis. eis.e ais.e d.e esa.e essLe TIFCISI I
Fig. D.4.e Calculatior. No. 4 AP across the fan cortponents located between the dead-er.ded regior. (volume 5) a r.d the f an-accurnula tor roores (volume 6). A naximurt AP cf 0.9 psi occurred at 4945.90 s ar.d -1.6 psi at 4947.26 s. l l l l l i l l 1
, - . --, ,, . - . . , , . , , - , s
l
-164- 1 1
I C/4/$30/RC/T5/ W a.vou.FC 1 TO 6 d . . . . . . . . . .
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' #8L8 eiLe O.e ek.e eis.e ein.e guit.e sia.e aske Tirtt$1 Fig. D.4.f Calculation No. 4 AP across the fan coreponents located between the lower com-partment (volume 1) and the fan-accumulator rooms (volume 6). A seximum AP of 0.5 psi occurred at 4945.87 s and -1.4 psi at 4947.23 s.
i l l l b_
-165-C/7/520/RC/T5/0er e-v0LUT I TO 2 d . . . . . . . . . . . . . . .
- g. .
5
- f. .
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e d* e a a a e i a e a a a a a a a i wit.e mis.e mis.s wit.e mis.e wise sais.e mit.e use.e 11E151 4 Fig. D.S.a ' Calculation No. 7 AP across the lower inlet doors located between the lower cer - partreent (volume 1) and the lower pier.um (volume 2). A maximurn AP of -4.6 psi against the doors (i.e., in their closed position) occurred at 5019.02 s.
,. . - _ .._ _ ,_ _ _ _ _ _ _ . _ _ - - 4 ,_ _.
166-E/7/52 Vile /TS/W e -v0LUT 3 70 7 d . . . . . . . . . . . t- gg li . 3- l ', - 1 I I i f ..
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a a a 1 a a a a a A a a a a a mate site mite eit.e sia.e site min.e amt.e mute fift 51 Fig. D.S.b Calculation No. 7 AP across the intermediate deck doors located between the upper plenue. (volume 3) and the top 1/4 ice ( volue.e 7 ). A maximun AP of 13.0 psi against the doors (i.e., in their closed position) occurred at 5017.11 s.
1
-167-1 C/7/520/RC/TS/ W e-v0LUPC 3 TO 4 g , , , , , . , , , , , ,
g. W , n E" '{ - n .
- g. ,
Y : tea.s sus.e tape.c ten.e tems.e sur.e aske aus.e me.e TIMCt51 i Fig. D.5.c Calculation No. 7 AP across the top deck doors located betweer the upper plerum (volume 3) ar.d the upper compartment (volume 4). A maximum AP of -1.5 pst ec-curred at 5006.62 s. which was af ter these doors were locked oper., ard 2.2 pei
' at $006.47 s. ;
i l l l l l
,f . . - - - . . ..-4 .- ,,, . , . , _ _ , _ _ _ , _ _ m, . _ . , - , , _ _ _ - . , . _ , _ _ .
1
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- mis.o mis.s us. snai.e umLe uit.s uits mis.s wit.s TitCL51 Fig. D.5.d Calculatier. No. 7 AP across the far. cor.por.er.ts located betweer. the upper cor-partner.t (volurne 4) ar.d the far.-accur.ulator rooms (volume 6). A manieue AP cf 7.1 psi occurred at 5019.86 s.
t
-169-C/7/520/RC/T5/ err e v0LUPC 5 70 6 9
e , , , , , , , , . . . , , , . 9 - 3- , j.. . g . . We
- g. . .
Y . : : : tees.e use.e mis.e uta.e mis.e sie.e mes.e mes.e nr.e Tinct $1 Fig. D.5.e Calculatior. No. 7 AP across the fan coreponents located between the dead-er.ded regior. (volume 5) a r.d the for. acpazulator rooms (volurte 6). A maximue AP cf 7.4 psi occurred at 5014.37 s.
) .
, 1
l l 4 l 2
-170 C/7/520/RC/TS/ W n-v0amt I TO 6 9
e . . . . . . . . . . .... 9
- p. .
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- g. .
d mit.o mis.o mis.o mit.o mio.o mis.o mis.s min.o maa.e 4, Tiric151
\
Tig. D.5.f Calculation No. 7 AP across the fan components located betweer. the lower cen-partment (volume 1) and the fan-accumulater rooms (volume 6). A maxieun AP of 1.6 psi occurred at 5019.90 s and -2.1 psi at 5019.66 s. Q. .
^\ l -
-111-
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' i C/8/S20/RC/TS/8F/20 o-vourt 1 TO 2 w- . . . . . . , , . . . .
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j. mee m e.e sum. tem.s sair.o sus.: ses.e seni.e sena.e TitCISI Fig. D.6.a Chiculation No. 8 AP across the lower inlet doors located betweer. the lower con:partment (volude 1) and the lower plenum (volume 2 ). A maximum AP of
-1.2 psi against!,the doors (i.e.. i r. their closed posi t i er. ) occurred at 5640.83 s. -
h 5 i
\
-172- >
t i
,I
.'N C/8/5EMtE/T5/F/20 4
o-WOLUPC 3 TD 7 g . . . . . , . .
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ain.s sue.e TI C 51 Fig. E.1.
Hydrogen concentratier.s for Calculation No. $ with six ice nodes where - (*) first ( of si x ) r. ode f rort the top (volume 7). (o) third (of six) r.ede f rort the top (volurte 10). ( A) fif th (of six) node frort the top (volume 11). t l
,o
1 l
, -160- ' C/5E/IEC/TS/gprM/61 o-20r6 ritM TOP, W 9 o-40F6 r1tM TOP, W 8 a-60r6 T1tti itP, W 12 2 . . . . . .
g . . . . . . . . . d . 4 b
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; =
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i Fig. E.2. Hydroger. concer.trations for Calculation No. 5 with six ice r. odes where - (*) second (of six) node from the top (volurr.e 9). (o) fourth (of six) node frer the top (volume 8), (A) sixth (of six) mode from the top (volume 12). - l
W h INDIANA & MICHIGAN ELECTRIC COMPANY t '
' P.O. BOX 16631 COLUMBUS, CHIO 43216 t =
October 10, 1983
- AEP:NRC:0500K Donald C. Cook Nuclear Plant Unit Nos. 1 and 2 j Docket Nos. 50-315 and 50-316 j License Nos. DPR-58 and DPR-74 (
RESPONSES TO REQUESTS FOR INFORMATION ON HYDROGEN COMBUSMt. T/C 7NTROL Mr. Harold R. Denton, Director Office of Nuclear Reactor Regulation U. S. Nuclear Regulatory Consission Washington, D. C. 20555 , l
Dear Mr. Denton:
)
l This letter and its Attachments provide additional information on i hydrogen combustion and control during degraded core accidents for the i Donald C. Cook Nuclear Plant Unit Nos. 1 and 2. More specifically, the l O information contained herein is heing ,rovided as a ,areia1 res,onse to three (3) Requests For Information transmitted to Mr. J. E. Dolan i (Indiana & Michigan Electric Company) by Mr. S. A. Varga (NRC) . These Requests For Information are dated July 30, 1982, September 16, 1982, , I and August 10, 1983. Attachment 1 to this letter presents a brief overview of the present status of our efforts to address your staff's concerns with l regard to hydrogen control for ice condenser centainments. In l particular, ninety-three (93) issues of concern have been identified . from the three (3) Requesta For Information referenced above. Twenty-one (21) of these issues have previously been addressed via our sukaittals dated October 15, 1982 (AEP:NRC 0500J), and December 17, 1982 (AEP:NRC:0500L). An additionat thirty-five (35) issues are addressed in Attachments 2 through 5 to this letter. Attachment 1 Presents our plans to respond to the remaining thirty-seven (37) issues. Attachment 2 to this letter provides responses to twelve (12) of the issues identified in Mr. S. A. Varga's July 30, 1982, Request For Information. Likewise, twenty-one (21) of the concerns raised in the September 16, 1982, Request For Information are addressed in Attachment 3 to this letter. Attachment 4 contains a copy of the report entitled " Fog Inerting 1 Analysis For PWR Ice Condenser Plants." This report is provided in { response to Question 10 of the September 16, 1982, Request For ! I Information. .
- A340100040 6aav2p PDR ADOCK 05000315
- 4 P ppg \
2-99 ,
~- - - - _. _ _ _ _ _
. .,, Mr. Harold R. Denton AEP:NRC 0500K 1
/
/. -
t I, Attachment 5 cohtains a copy of a paper presented at the Second International Workshpp on the Ispact of Hydrogen on Water Reactor Safety (Albuquerque, New Mexico, October 3-7, 1982). This paper, entitled " Fog Inerting Criteria For Hydrogen / Air Mixtures," is provided in response to Question 11 of the September 16, 1982, Request For Information. . - This docianent has been prepared following Corporate procedures which incorporate a reasonable set'jof controls to ensure its accuracy and completeness prior to signature by the undersigned. Very truly yours, I e P. A ich Vice President 1 MPA/ cam Attachments
. cc John E. Dolan - Columbus W. G. Smith, Jr. - Bridgman R. C. Callen G. Charnoff .
. E. R. Swanson - NRC Resident Inspector, Bridgman
\a i
P l
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=
- - - - - . - . . , , - e .
.. .e ' se-s a - . m e- s . s.m a + .- . . - - - - - - . - - - - - - - - - - - - - - - -
e e e i - FOG INERTING ANALYSIS FOR PWR ICE CONDENSER PLANTS BY S. S. TSAI CORE AND CONTAINENT ANALYSIS NUCLEAR SAFETY DEPARTMENT . WESTINGHOUSE ELECTRIC CORP. O i I s , f ! NOVEMBER 1981 - f ....m '
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- PDR ADOCK 05000315 P PDR 0430Q:1 h( , ,-; !
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I
r: ABSTRACT The recent hdrogen burn test conducted at the Lawrence Livermore ,, National Laboratory has raised the NRC and the ice condenser plant xid owners concern about' fog inerting probability and consequ'ences tii ice condenser plants. The present investigation is aimed at resolving this : fog inerting issue. In this report, major fog formation 'and removal ,, mechanisms that exist in the post-accident ice condenser containment are
' CC identified and quantified. Methodologies have been developed for pre-dicting fog formation and removal rates and for predicting fog concen-trations in various compartments in an ice condenser containment. ;
This methodology development has resulted in two computer programs, FOG and F0GMASS. The FOG computer program employs the Hijikata-Mori boun-dary layer fog formation theory, and calculates the fog formation rates due to boundary layer and bulk stream condensation. The computer pro-gram FOGMASS solves the mass conservation equations for fog droplets and calculates the fog concentrations in various compartments. Both compu-ter programs have been used to predict fog concentrationsde the { Sequoyah, McGuire, and D. C. Cook containments, using the CLASIX output
- data for a S D 2
accident sequence. In order to utilize the calculational results frce the study, a fog
~
inerting criterion has been estsbitsbed. This criterion uses the hdro- , gen concentration, volume mean drop size, and fog concentration to define the fog inerting regime. For a given hydrogen concentration, the minimum fog inerting concentration was found to vary with the square of the volume mean drop size. This criterion has been verified by the Factory Mutual recent fog inerting test data. The application of the fog inerting criterion to the three ice condenser plants shows that fog inerting would not exist in the upper and lower compartments. Fog inerting in the ice condenser upper plenum at hdro- { gen concentratons at which glow plug igniters are designed to operate is very unlikely. . ( I 0430Q:1
- e + m D _- - . -. - , - -
TABLE OF CONTENTS i .- Title Page Section ABSTRACT i TABLE OF CONTENTS 11 LIST OF TABLES iv , LIST OF FIGURES - V 1.0 . BACKGROUND 1-1 2.0 IN TRODUCTION 2-1 3.0 FOG GENERATING MECHANISMS 3-1 , IN AN ICE CONDENSER CONTAINMENT - 3.1 Fog Generated by Break Flow 3-1 3.1.1 Amount of Fog Generated 3-3 by Break Flow - 3.1.2 Drop Sizes Generated by Break Flow 3-5 3.2 Nucleation of Fog Droplets in Containment 3-6 Atmosphere ({} 3.2.1 Nucleation Theories 3-7 3.2.1.1 Clas~sical Theory of 3-7 Homogeneous, Nucleation 3.2.1.2 Heterogeneous Nucleation 3-9 Theory
^
3.2.2 Fog Formation Conditions 3-10 3.2.3 Conditions for Fog Formation Near 3-12 a Cold Surface 3.2.4 Rate of Fog Formation 3-15 3.2.5 Fog Drop Sizes 3-19 3.3 Fire Mist Droplets From Containment Sprays 3-19 f 4.0 FOG REMOVAL ECHANISMS IN AN ICE CONDENSER 4-1 CONTAINENT 4.1 Settling Due to Gravity 4-1 4.2 Agglomeration 4-2 4.3 Vapori zation 4-2 ( 4.4 Removal by Spray Drops 4-3 4.5 Other Removal Mechanisms 4-3 II 0430Q: 1 t ., -- , - _ . , , -- -- --, r,. , - , - - - ,-+~ , ,. s - < , . - .
~~ ~
. _ . . . . . . _ . _...~ E ~~~~' ~ ~~- ~' E _E -
- TABLE OF CONTENTS (Continued)
U - Section Title Page 5.0 FOG INERTING CRITERI A 5-1
. 5.1 Previous Work. 5-1 5.2 Present Theory 5-2 5.3 Verification of Theories by Experiments 5-6 6.0 ASSESSENT OF FOG INERTING PROBABILITY IN ICE 6-1 4 CONDENSER CONTAINENTS
- 6.1 Determination of Volume Fraction of Fog 6-1 Droplets in Ice Condenser Containment Subcompartments 6.1.1 Calculation of a bmak 6-5 l
j 6.1.2 Calculation of e cond 6-6 6.1.3 Calculation of u set 6-6 6.1.4 Calculation of m sp 6-7
, 6.2 Fog Inerting Probability in the Sequoyah 6-7 Plant 4 O 6.3 Fog Inerting Probability in the McGuire 6-23
. Plant 6.4 Fog Inerting Probability in the D. C. Cook 6-37
! Plant 6.5 Ef fect of Fog on Global Costustion
~
6-50 7.0
SUMMARY
AND CONCLUSIONS 7-1 ACKNOM.EDGMENTS 7-3
- REFERENCES R-1 APPENDIX A A-1 3
APPENDIX B B-1 l ( 0430Q:1 111
- - . = .
- -~= -- == .
j - [ LIST OF TABLES j] Tabl e No. Title Page l l l 6.1 F0G Input Data for Sequoyah Lower Compartment 6-18' l
- w. /
6.2 FOG Input Data for Sequoyah Ice Condenser 6-19 6.3 Geometric Data for Sequoyah Containment 6-20 MARCH Prediction of keactor Coolant Mass and 6-21 6.4 Energy Release Rate for the S D 2 Sequence-3 6.5 Intercompartmental Flow Rates (f t /sec) 6-22 Predicted by CLASIX for Sequoyah 6.6 FOG Input Data for McGuire Lower Compartment 6-33 5-34 O s7 roG raPut oata for acGuire tc. coad a er 6.8 Geometric Data for McGuire Containment' 6-35 6.9 Intercompartmental Floit Rates (f t3 /sec) 6-36 Fredicted by CLASIX for McGuire 6.10 FOG Input Data for D. C. Cook Lower Compartment 6-46 6.11 FOG Input Data for D. C. Cook Ice Condenser 6-47 , 6.12 Geometric Data for D. C. Cook Containment 6-48 6.13 Intercompartmental Flow Rates (ft3/sec) 6-49 Predicted by CLASIX for D. C. Cook , i iv 0430Q:1
f.IST OF FIGURES Figure No. Title Page 3.1 T-S Otagrim for Reactor Coolant Of scharged 3-4 From Break 3.2 Vapor Pressure and Temperature Profile Near 3-14 a Cold Surface
. 3.3 Formption of Fog Near a Cold Surface 3-16 3.4 Drop Size Of stribution Predicted by Nefburger 3-20 and Chien
. 3.5 Particle Size Of stribution for 50 PSI Pressure 3-21 Drop Across Nozzle No.1713 4.1 Terminal Velocity as a Function of Orop Radius 4-5
[) in Steam-Air Atmospheres 4.2 Agglomeration Rates in Air Between Equal-Sized 4-5 Orops 5.1 - Minimum Ignition Energies and Quenching Distance 5-3 for Hydrogen-Oxygen Inert Gas Mixtures at Atmo-spheric Pressure 5.2 The Ef fect of Droplet Spacing on Flame Quenching 5-4 5.3 Schem'atic Representatioh of Temperature Profile 5-7 Through the Flame Front f 5.4 The Parameter /eg u as a Function of 5-7 (Y u - I Ilf 'l for Different Values of Xeg t y 04300: 1 . I 4-
._ .. ._. .-. - ..e- _ _.-- - - -
4 LIST OF FIGURES (Continued) Figure No. Titl e Page l 5.5 (K) crit 't at the Flamability Limit as a 5-8 ) Function of (Yu - Y f)/eg 5.6 Comparison Between Theories and Factory Mutual 5-10
, Fog Inerting Expericients on 4.76 Percent H2 5.7 Comparison Between the Present Theory and 5-11 Factory Mutual Fog Inerting Experiments on 7.2 Percent H 2
5.8 Comparison Between the Present Theory and 5-12 Factory Mutual Fog Inerting Experiments on 7.9 Percent H 2 6.1 Sequoyah CLASIX Containment Model 6-8 6.2 Fog Formation in TVA Sequoyah Lower Compartment 6-10 6.3 Fog Formation in TVA Sequoyah Ice Condenser 6-11 6.4 Fog Concentration in Sequoyah Containment 6-14 6.5 McGuire CLASIX Containment Model , 6-24 i 6.6 Fog Fonnation in Duke McGuire Lower Compartment 6-25 6.7 Fog Formation in Duke McGuire Ice Condenser 6-26 6.8 Fog Concentration in McGuire Containment 6-29 6.9 D. C. Cook CLASIX Containment Model 6-38 t vi 0430Q:1
- ~
t_ .
.. -_. __ . -...a-----------------...--
.s =
. l
' 1
- LIST OF FIGURES (Continued)
Figure No. Title Page 6.10
- Fo'g Fon' nation in AEP Cook Lower Compartment 6-29 l 6.11 Fog Fonnation in AEP Cook Ice Condenser 6-40
, Fog Concentration in D.C. Cook Containment 6-43 6.12 i
O b l
. t 4
g O i i vif [ i 0430Q:1
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1.0 BACKGROUND
The incident at Three Mile Island has demonstrated that a significant amount of hydrogen.could be generated during cort degradation. Thi,s - experience raised NRC concern about the safety of nuclear power plants, in terus of their capability to control hydrogen during severe acci-dents. Since ice condenser plants have a relatively small volume and low containment design pressure, the problem is magnified. The refort, the NRC has requested the ice condenser plant owners to study hydrogen control methods for use in their plants. In this regard, the Tennessee Valley Authority (TVA), Duke Power and American Electric Power (AEP) have proposed the use of glow plug igniters at various locations inside j their ice condenser containments to ignite hydrogen at low concentration. 9 Recently, the NRC requested Lawrence Livermore National Laboratory (LLNL) to carry out experiments on these igniters to determine their effecti veness. In these experiments, two tests with high steam concentration seemed to indicate that substantial fog formation could
- O occur when saturated steam is discharged into a unheated vessel and under some conditions fog could effectively preclude hydrogen from
; combustionIII. ,
The LLNL tests raised some doubts about the effectiveness of glow plug igniters under fog formation conditions. In a recent review of hydrugen related issues for ice condenser plants,(2) the NRC has raised several questions concerning the prvbability and consequences of fog formation and steam supersaturation in ice condenser plants. In response to the NRC questions, TVA, AEP, and Duke established experi- . i mental and theoretical analysis programs to study the fog inerting prvb-lem. The experimental program was contracted to Factory Mutual. The experiments were designed to test glow plug igniter's performance under different fogging conditions. At the same time, the plant owners requested Westinghouse to perform fog inerting analyses for the Sequoyah, McGuire, and D. C. Cook plants. This report presents the results of the Westinghouse studies. , l l 0430Q:1 1 -1
~ __ .
2.0 INTRODUCTION
From the post-test analysis of the LLNL hydrogen burn tests, it appears that substantial fog formation occurred inside the test vessel. Ge n-erally, fog droplets are only few microns in diameter. These sizes of ' droplets have potential to prevent' a 'lamable f gas mixture from combus- ! tion or quench a propagating flame. This is because these sizes of droplets vaporize very fast (on the order of miliseconds), absorbing an enormous amount of the heat released from combustion if a substantial quantity of these ~ droplets is present in the atmosphere. In comparison, large water droplets in the range of few hundred microns or larger (e.g. spray droplets) have no inerting effect on conbustion(20) and hence have insignificant effect on glow plug igniter's performance. There-fore, the,present analysis will be concentrated on the generation and ! removal of fog (mist), and its impact on the glow plug igniter system.
. There are a number of fog generation and removal mechanisms present in a post-accident ice condenser containment atsosphere. The fog generation mechanisms include fog generated by the break flow (if it is two-phase),
fog formation near the ice and structural heat sink surfaces (since the surface temperatures could be well below the dew point), and fog genera- ' tion due to homogeneous and heterogeneous nucleation in condensing bulk - streams. 1
. The fog removal mechanisms include gravitational settling,'agglomera-tion, vaporization and removal by spray droplets. In order to estimate the post-accident fog concentrations in ice condenser containments, these competing mechanisms must be studied, and evaluated. To solve this problem, it requires a numerical integration of the mass conserva-tion equations for the mist droplets in the various containment subcom-partments. By making some simplifying assumptions the transient fog concentration in the various subcompartments have been estimated.
3 i ( l 0430Q:1 2-1 __--___-.n-__ _ _ _ _ _.-. ,
. i .
c l 1 The analysis presented here/ considers all the fog removal and generation
- t. .
sechanisms previously descFibed. In addition, it considers the fog l entrainment in the intercpapartmental flows (including fan flows) in the j fog mass conservation equations. In order to perform this analysis it l was necessary to use CLASIX results for a S 2D event as boundary condi- . tions to the problem. ' l
.s
- In addition to calculation of fog concentrations in various containment compartments, it was necessary to establish a fog inerting criterion. A fog inerting criterion has been proposed by Berman et al., which pre-dicts the minimum fog concentration required to inert a given hydrogen concentration and given volume mean fog drop size. This criterion seems to overpredict the minimum fog inerting concentration, when compared with experimental data. A more realistic fog inerting theory is presented in the pres 2nt study.
l
. The fog inerting methodology, analysis, and results are presented in the following sections of this report. Sections 3 and 4 present the method-ology for calcuiating the fog fonnatica and renoval rates. Section 5 -
gives the fog inerting criteria, and Section 6 presents the results. [) 1 ( 0430Q:1 2-2 e -
. a ' , 3.0 FOG GENERATING MECHANISMS IN AN ICE CONDENSER CONTAINMENT The inerting capability of fog droplets depends on their sizes and con-centration in the containment atmosphere, as well as the hydrogen con-centration. This section is intended to ' identify various fog generation ,
mechanisms present in an ice condenser containment and to determine the I
~
drop sizes and the rates of fog generation from these mechanisms. Three . fog generation mechanisms are discussed in this section and the dominant fog generation mechanisms are identified. 3.1 FOG GENERATED BY BREAK FLOW l The post-LOCA containment atmosphere is most likely to be a drop-laden atmosphere. The large-scale sis.ulated LOCA experiments conducted to date have directly or indirectly confirmed the presence of two-phase atmospheres. For example, Marvikken(3) and Battelle - Frankfurt experiments were instrumented to measure fluid densities and water C. levels in various parts of the containment. Therefore, fog generation by the break flow cannot be neglected. The following discussion of this
~
phenomenon pertains to small LOCAs. , In the early stage of a small LOCA transient, a substantial . portion of the primary coolant discharged from the break will remain as liquid. Because of the superheat and high velocity, this liquid will be framen-ted by aerodynamic forces and homogeneous nucleation mechanism into' small droplets. These droplets are expected to be entrained by the intercompartmental and fan flows and spread to other parts of the ice condenser containment. During their travel throughout the containment, the fog droplets will be removed by pravitational settling, sprays, and vaporization. The fog generation period lasts until the water level in the reactor vessel falls to the break elevation and the break flow is no longer two-phase. For the particulfr S D2 sequence analyzed by, l ( l 0430Q:1 3-1 -
-- - ~
3 - . .
.. ... . . . _ . . _ _ _ _ . . __ _ _ _ _ u ..._
. i CLASIX,(5) this period lasts for about 36 minutes and about 4.2 x
' 0
, 10 lbs of water has been discharged into the lower compartment during this period of time.
After the water level 'in the mactor vessel falls below the bma'k eleva-tion, the bmak flow rate is substantially mduced. The flow is essen- - tially steam and no fog dmplets will be generated. As a msult, the lower compartment becomes superheated af terward. Fog droplets may vaporize during their travel through this compartment and substantial mmoval of mist droplets are expected. Large suspended drops generated by the bmak flow will be mmoved very quickly by gravitational settling and impingement. For the drops larger than 20 u, the removal rate is high and complete mmoval only takes a few seconds. For the smallest drops (less than 1 u) the terminal veloc-
, ity is so small that they virtually remain suspended in the atmosphem indefinitely. The only effective removal mechanisms for these sizes of ,
drops are vaporization, and collision with larger spray drops. The , weight fraction of these sizes of drops is estimated to be 1 per- ! cent (3) generated by the break flow. The population of these small drops can ini:mase if nucleation of embryos occurs in a saturated atmo-sphere. i i 1 I i i i l l 0430Q:1 j 3,2 Q -.
~_
~ ^ ' ~~
b l
, 3.1.1 AMOUNT OF FOG GENERATED BY BREAK FLOW As discussed previously, the existence of a two-phase drop-laden iegime has been observed e'xperimentally. In a small LOCA, flashing of primary coolant at the break and subsequent vaporization of blowdown liquid represent a series of changes of thermodynamic states. Since the reac-tor coolant pressure is high, the break flow will be choked. The accel-eration of primary coolant to the break location is essentially an isen-tropic proces's.,in which the pressure decreases to the point at which substantial homo'geneous nucleation occurs. When the flow leaves the break, the liquid is framented by both the aerodynamic forces and the nucleation mechanism into small fog droplets. These fog droplets con-tinue to vaporize, because of the superheat in the droplets, untti a thermodynamic equilibrium state is reached. Because of the high super-heat and large aerodynamic forces, it is expected that the fog droplets ~
generated are very small. This vaporization process is essentially isenthalpic. The existence of a two phase drop-laden regime can also be explained by
- use of a T-S diagram for steam as shown in Figure 3.1 (Figure 1 of Reference 6). It is shown in this figure that the blowdown liquid goes through a series of thermodynamic states, starting from the subcooled ifquid state 8,. The liquid expands isentropically from the subcooled state B, to the state 81 at the break, where a two-phase critical flow is developed. At the same time, temperature changes from T, to T.3 After leaving the break, the droplets continue to vaporize because of excessive superheat until finally an equilibrium state B2 is reached at which the droplets are in thermal equilibrium with their surroundings. This vaporization process is essentially isenthalpic. At this time, the droplet temperature drops to T2 and the atmospheric temperature also rises to T2 . For a small LOCA, the equilibrium tem-perature varies with time. According to the CLASIX analysis of the Sequoyah plant, the lower compartment gas temperature rises quickly from 100*F to approximately 200*F and then stay at this temperature for an extended period of time. Using these temperatures as final equilibrium
[ l l 0430Q:1 3-3 l u -- - - _ - - . - _ . - - - _- - - -. ._ _= - -
I Kg/cm 2 100 10 "1 0.1 0.01 , l l I I i 600 450 - 400 SM g y 400 -
- 350 3MD
- Il -
g ' o 300 - s
** ~
1, a l 200 1# '
/ \ l 150 - 1 T2 1/ , \\ \\\ '\ \ ~
I I
\ s ^2
[ a2 \ \A \ A0
~
> \- [\ .N \ l
. O I !\ !
0.0 0.2 0.4 0.6 0.8 X=1 t e FIGURE 3.1 T-S DIAGRAM FOR REACTOR COOLANT DISCHARGED FROM BREAK G .l
temperatures for water droplets, the weight fraction of water droplets ( in the break flow is approximately 50 percent, which is consistent with l
' I the MARCH calculations (7) of the break flow rate and its energy
! release rate.
The discussion given above is valid only when the initial state of the break flow is subcooled or saturated liquid. After the water level inside the reactor vessel falls below the break elevation, the break ; i flow will be steam. The moisture content of the steam will be very low. ( l
- even though isentropic expansion may lead to homogeneous nucleation and I subsequent condensation in the vapor stream. Depending on the super-saturation that can be achieved in this isentropic expansion, a conden- ;
sation shock is possible when critical supersaturation is reached. l However, it is believed that the fog droplets generated by homogeneous I nucleation in this supersonic jet is negligible as compared to other fog l generating mechanisms. Hence, it will be neglected in this present l analysis. Therefore, the fog generation by the break flow is considered i possible only when the water level in the reactor vessel is above the i break elevation. O According to the MARCHI7) calculation at 2172 seconds into the acci- ; [ dent, the water level inside the reactor vessel falls below the break ; } elevation for the S D 2 case analyzed in Reference 7. By this time , j i approximately 421,000 lbs of water has been discharged from the break ! and 56 percent of this discharged fluid, i.e., 236,000 lbs, will be sus-
- pended in the atmosphere as condensate. However, most of these droplets will later be removed by gravitational settling, sprays, and vaporiza-f tion.
3.1.2 DROP SIZES GENERATED BY BREAK FLOW f- The flashing jet experiment conducted by Brown and York (8) has indi-
, cated that the drop sizes produced by flashing liquid are small. They
{ derived a correlation for the linear mean drop size based on .the test data. The correlation shows that the mean drop size is inversely pro- l portional to the Weber number and it decreases linearly with increasing l 4 l ( j 04300:1 3-5 i 1 1
,4, _ _ . _ _ - _ , - , _ . . _ _ . _
- ~~ :_..
~ ~~
._.E ' _ _
< l -
superheat. However, this ' corr /lation is applicable for ifquid superheat I less than,75*F and it can not extrapolated to the large superheat of the reactor coolant. However, some conclusion concerning the drop sizes produced by blowhown of the reactor coolant can be drawn for this condi-tion. The break flow ha's much larger superhe'at and Weber number than those used in this experiment; therefore, the drop sizes produced by the break flow should be much smaller than 50ir observed in this experi-ment. Gido and Koeste1 I9I have developed a method for estimating the drop size leaving the fragmentation / evaporation zone of a blowdown jet. This model is based on the assumption that drops with an internal temperature s difference of less than SK will escape fragmentation. This model has been verified by the low superheat data of Brown and York. Application of this method to the LOCA condition shows that the maximum attainable drop ' size is 79 (this means that any drop size larger than 7p will not escape framentation by homogeneous nucleation). The corresponding mean drop size is about 4g, based on the observation of the. largest drop size . and mean drop size in the experiment reported in Reference 8. However, this volume mean drop size is not used in the pdesent analysis. Instead, the present analysis uses 10 u mean drop size,. considering the drop agglomeration effect. 3.2 NUCLEATION OF FOG DROPLETS IN CONTAINMENT ATMOSPHERE Nucleation of water embryos from the homogeneous vapor phase plays an important role in mist generation in ice condenser plants. Nucleation , is a process by which tiny water embryos or condensation nuclei are , formed from a pure vapor phase at a rapid rate. In incipient homogene- ; ous nucleation, the local gas temperature drops below the dew point ; corresponding to the local steam partial pressure and some degree of l local supersaturation is needed. The degree of supersaturation needed to start nucleation depends on the nisaber of condensation nuclei present 1 in the containment. These condensation nuclei could be very small water droplets or dust particles. If sufficient number of condensation nuclei ( 0430Q:1
. 3-6 ,
1 h.
. < f.
exist, supersaturation could be small. It is likely that the ice con-I , denser contalment contains a subtantial number of dust particles such-that little supersaturation is needed for nucleation. This section is devoted to the discussion of fog fomation by homogene- l ous or heterogeneous n'uct'ation. e The classical nucleation theories are i used to explain the nucle'ation phenomenon. i s
~
3.2.1 NUCLEATION THEORIES The process of nucleation of an embryo water drop is important in under- ,' standing the mechanism of fog formation in ice condenser plants. Two types of nucleation process, namely, homogeneous and heterogeneous nucleations, and their theories will be discussed i Section 3.2.1. l 3.2.1.1 CLASSICAL THEORY OF HOMOGENEOUS NUCLEATION
~
When an embryo droplet, usually assumed spherical, is formed from con- l densation of water vapor molecules, its free energy changes. The change of free energy can be expressed as AG = 4 r2 e - (4/3) wr3 nL KT in (P/Po) (3.1)
\
where e is the surface free energy per unit area,.or surface tension, r is the drop radius, P is the vapor pressure, Po is the saturation j pressure at the droplet' temperature, nt is number of molecules per unit volume, K is the Boltzman, constant, and T is the drop temperature. The supersaturation S, is defined as P/P,, Equation (3.1) represents a free energy barrier to the growth of the drops at a given suprsaturation. At < maximum AG, the critical radius r* can be obtained from Equation 3.1 asl t . r* = n gKT in (P/P,) i
. h l
0430Q:1 3-7
. n . . . _ . . . .- . - - . _ . .- . .. .. - . . - . . -
. *,e *
- l l
= i
)
The drops of the c'itical, r size can be considered as condensation nuclei ]4 , since at this size the drops will grow with no change in free energy. This critical size reprt;ents an aquilibrium size at which. a supersatu-
. rated vapor at vapor pressure P is in equilibrium with this critical drop at a lower saturation pressure P,. However, this equilibrium mode is unstable. For example, if a drop of the critical size origi-nally in equilibrium wit'h the surrounding vapor suffers a sudden small g increase in size due to condensation, then (if the drop temperature does not change), Equation 3.2 shows that the equilibrium pressure, P, on its' y
surface will decrease. Therefore, the actual vapor pressure will then be greater than the equilibrium value and further condensation wiil occur. This is why the drop of this critical size is called condensa- l tion nucleus. [ The nucleation 'ract of critical-sized embryos can be obtained from the s kinetics of a nonequilibrium distribution of embryos. The classical , nucleation theory (10) shows that ther e is a very sudden increase in the nucleation rate when past a certain critical value of supersatura-h tion. An extensive validation of the nucleation theory was conducted by 1
. Yolmer and Flood IllI in an experiment in which a number of vapors were expanded to visible condensation in a cylinder. The observed critical supersaturations agreed suprisingly well with theory in nearly all cases, includting water vapor.
Critical condensation nuclei sizes typically range from 10 to 100 ! atoms. These sizes are considerably smaller than the mean free path of the vapor molecules and therefore the rates of mass and heat transfer at
'the #op surface cannot be predicted by bulk transport theories. In this case, "the kinetic theory of gas should be used to predict the rates of mass and heat transfer at the drop surface.
Starting from the Linetic theory of gas and the energy conservation equation, the rate of growth of a condensation nucleus was obtained by Hill et al.(10) It was found that the growth rate is on the order of 10-3 ft/sec. Therefore, it hakes '
'o nly about 1 milisecond for the ,
condensation nucleus to grow to a fog droplet size of 1 u. 0430Q:1 3-8 n+. _
. . .. - - - _.. . ... :-. 1 a
3.2.1.2 HETEROGENEOUS NUCLEATION THEORY
't .
Another mechanism of forming embryos is heterogeneous nucleation on foreign particles that could suspend in the containment atmosphere. These particles may serve as nucleation sites for vapor and thus enhance the nucleation rate. The source of foreign particles in the containment following core degradation could come from fission product aerosols and dust particles. The size distribution of these particles are important l because the supersaturation required to form embryos depends on particle sizes. A typical size distribution of atmospheric aerosols is that of Junge(12) , taken from surveys made near Frankfurt A.M., Germany. The surveys found that the size range of dust particles is from 0.01 to 1
- u. In the range from 0.01 to 0.5 u, there are between 100 and 10,000
- particles per cubic centimeter. A majority of particles have sizes smaller than 1 micron. At the smallest size of 0.01 u, the critical supersaturation is about 1.02 and at the largest size the supersatura-tion is only 1.001.
The other source of aerosol particulates is fission. products. During normal operation, the primary coolant contains very little fission pro-ducts. However, a large release of fission products, such as the gap release, could occur at about the same time the hydrogen releases. The amount of fission products released to the containment depends on acci-dent scenarios. The distribution and transport of fission products in the containment can be predicted by the CORRAL code (13). The size l distribution of fission products .in the containment can be extrapolated from the CSE experimentsII4) . These experiments indicated that soon I af ter fission product release, the mean particle diameter was 15 u. A few hours later, the mean diameter decreased to about 5 u because of l settling of large particles onto the floor. These sizes are substan- l tially larger than those of dust particles and therefore, critical ! j supersaturation is even smaller than values. quoted above for ttie dust l particles. l [ l l l 0430Q:1 3-9
The atmospheric aarosols consist of particulates of various sizes, vari-ous chemical components, and various electrostatic charges. The aerusol particulates could be solubie or insoluble in water. All these proper-ties could affect the required supersaturation for nucleation. i l In the case of insoluble particulates, the contact angle, d, betw4en the ' embryo and the particle surface is important. If the particle is com- l plately wettable, d = 0, it forms a base on which a small amount of water can. fonn a drop of large radius of curvaturt and thus satisfy the Hemholtz equation (Eq. 3.2) at a much lower supersaturation than would be the case if same number of molecules fonn a drip with a particle cort. Fletcher (15) developed a relationship between the supirsiatura-tion and drup radius for several values of contact angle, assuming that the particle is spherical. Competely wettable, a particle of 1 micron or so, when covered with a film of water, is theoretically at the crit-feal radius, and it needs only 1.001 critical supersaturation. The post-accident containment atmosphere is likely to contain a substan-tial amount of aerosol particles. These particles will act as condensa-tion nuclei and therefore, little supersaturation is required to prt-Q cipitate condensation. 3.2.2 FOG FORMATION CONDITIONS l Fog formation in a mixture of vapor and noncondensible gases has been of interest to meteorologists, and turbine and condenser designers. Fog is formed by homogeneous or heterogeneous nucleation as i. result of tem-perature drop below the dew point (sometimes with conconnitant pressure drop). During the temperature drop, a local gas element will go through a series of thermodynamic states. Eventually, a state is reached at which incipient fog formation occurs. Some degree of vapor supersatura-tion is needed to precipitate fog formation. The vapor supersaturation , at which rapid nucleation of vapor first appears is called critical supersaturation. The critical supersaturation, in general, is a P !( 0430Q:1 3-10
-- - a s swas l
1
- e. ,
i functA of temperature, vapor properties, mixing time (if a mixing
- process is in,cived), and concentration and sizes of foreign particles.
! The critical supc;aturation data for water has been given in Reference 15. l Fog formation in an ice condenser containment as a result of homogeneous or heterogeneous nucleation could occur: (i) inside the thermal boun- l dary layer near a cold surface, (ii) in adiabatic or nearly adiabatic expansion of vapor jet, and (iii) in mixing of a hot vapor stream with another cooler gas. Surface cooling may create a region of local supersaturation within the thermal boundary layer, even though the bulk stream is still super-heated. If the local supersaturation reaches the critical supersatura-tion, incipient fog formation will commence. This condensation mecha-
- nism may exist in any compartments within the containment especially in the ice condenser where ice temperature is well below the dew point.
When a high speed vapor - noncondensible gas mixture jet goes through an adiabatic or nearly adiabatic expansion, the gas mixture temperature and l pressure will drop rapidly such that condensation may occur somewhere in ' the expansion process. This is the case when a hydrogen-steam mixture i ! jet exits from a break at a supersonic speed. The jet experiences a rapid expansion and if critical supersaturation is reached, condensation shock may occur somewhere within the expanding jet. This condensation mechanism can only occur in a compartment in which the hydrogen-steam mixture jet exists. Condensation in a fast expanding vapor - noncondensible gas jet is a , I localized phenomenon. Usually very little moisture is generated in the expansion process even if a condensation shock does exist. Therefore, the present study does not attempt to treat the condensation shock as a source of fog formation. i ( 04300:1 3.j j l r - -
. -. . . .: . .. .- l
..e, .
~
[ l'he third mechanism, conde/nsation due to mixing, may exist in a compart-t .. ment where a hot hydrogends, team mixture mixes with a relatively cold containment atmosphere. During the mixing process, local critical
. supersaturation within the mixing gas could.be reached and condensation. .
would ensue. This mechanism could exist in'the lower compartment in which relatively cold gas from the upper / compartment is returned by the
~
deck fans and mixed with the hot humid air. Thus, the afxing of cold and hot vapor streams will be treated in the , present study.\ However, only bulk condensation is considered. That is, it is not intended to compute the temperature profile to predict the 4 local condensation rate. Instead, the bulk gas is assumed at one uniform temperature, and bulk condensation will occur when mixing results in saturation conditions. This is consistent with the CLASIX code assumption of uniform gas temperature. Because of time restriction, it is almost impossible to treat all the . 4 condensation mechanisms. However, major condensation mechanisms will be O identified and treated in the present study. Before entering into the discussion of the methodology to calculate the fog formation rates from various fog formation mechanisms, a discussion of fog formation conditions is necessary. Since the bulk condensation approach for the mixing process has been adopted, the fog formation conditions for the mixing process are simply that critical supersatura-tion is reached in the bulk stream. For practical purposes, the crit-ical supersaturation is assumed to be one since it is likely that plenty of coridensation nuclei exist in the atmosphere before mixing condensa-tion takes place. 3.2.3 CONDITIONS FOR FOG FORMATION HEAR A COLD SURFACE Fog starts to form at a fast rate near a cold surface when local vapor ! supersaturation reaches the critical supersaturation. Near the cold surface, a thermal boundary layer is formed, within which local vapor pressure and saturation pressure vary. Typical vapor pressure and l l 0430Q:1 3-12 O
, l l
l temperature profiles, when the incipient homoger.ous nucleation first j (. - appears, are shown in Figure 3.2. It is seen in this figure that when j the local vapor pressure reaches the critical vapor pressure there is a sudden appearance of fog in the boundary layer due to the fast nuclea- ;
. tion rate. Rosner and EpsteinIIII' have derived fog fonnation condi-
- tions near a cold surface, assuming that the local vapor pressure curve is tangent to the critical vapor pressure curve at the fog incipient point. A mort general fog fonnation criterion was given by Hijikata and Mort (17) 9,
- i. p > (h) (3.3) where AW=W -
W, . aT = T - T, i and the weight fraction of condensing vapor, W, can be related to the i partial pressure of the condensing vapor Py as C (Py /p) (My/Mg ) N" (*} 1-(P/p) (1 - M /M y g) where P = total pressure
- My = vapor molecular weight M
g
= noncondensible gas molecular weight Equation (3.3) may be rewritten as n>2 (3.5) wMm ,
n=2h/(h) ( 04300:1 3-13
. ... . .. . - - . . . ....:.- .._.-~..a 21123 3 Pv, crit (T )
A
.- I .
I Pv Pv,eq(T ) Pv, = 1 1 11 I ll I 11 I ll l Pcrit (Tw)( ll l 1I I Pv, w ll l Y 0 =
, SUPERSATURATED ,
REGION
- k "
THERMAL SUPERHEATED BOUNDARY LAYER g g OR SATURATED ll ! ! VAPOR SUPERSATURATED REGION 9 I U I T"
< # sc I FOG 3 5 I D.P.lt I '
i e IT I a 8 : To I I g I I I I I
** I I
__t Tw l ~ 6 6 e FIGURE 3.2 YAPOR PRESSURE AND TEMPERATURE PROFILES NEAR A COLD SURFACE ( 3-14 m -
.. ..'~ . . . . _ . . _ - . ___.-_.....'...
'}~
The parameter n is used in the following section to calculate the fog ^l -
- formation rate. It will be demonstrated that when n < 2, no fog forma-tion is possible.
3.2.4 RATE OF FOG FORMATION NEAR A COLD SURFACE As has been discussed in Section 3.2.3, fog will form near cold surfaces (e.g., in the ice condenser early in the transient.) As ~ discussed in Section 3.2.1, once water embryos are formed it takes only a few mili-l seconds for them to grow to the micron size. Af ter these micron size < fog droplets are formed, it needs very little supersaturation for fur- ) ther growth. Therefore, in the present analysis, it is- assumed that vapor and droplets are in thermal equilibrium and local vapor pressure . f I is equal to the local saturation pressure. This section is concerned
! with the transport of these micron-size fog droplets within the thermal boundary layer.
i I The boundary layer fog formation rate can be determined using the Hijikata-Mori theory IIII of fog formation in the thermal boundary r layer. It was assumed that a thin liquid film, having a thickness of a g on a cold surface, coexists with a gas-droplet flow in a two-
~
phase boundary layer of thickness a outside the liquid film as shown in i Figure 3.3. l ^ It was further assumed that the saturathon condition exists within the two-phase boundary layer and the boundary layer approximation is applf-cable. Numerical solutions were obtained for the mass fraction of fog droplets, Yo , at the gas-liquid film interface. The fog droplet flow rate at a distance X along the plate may be expressed in terms of Y as < . (a m g=Lp), Yudy (3.6) l l1 - i 0430Q:1 3-15 l . e - 1
o . ) l I s
~
P
> F, .
usin Fiou . Tw ut . , i. WOG ~-
'T '
[ } II #' -
- l0 -
l J '
/
/ ,7
- O O -
Two Phase y ,[ y [ , , Boundary layer , EV* Io I If
/ -
I nterf ace '
/O/////////////////////////////////////// \ '
Liquid Film Cooling Surface - 1 1 FIGURE 3.3 FORMATION OF F0G NEAR A COLD SURFACE ' i i 4 i-
where Y = mass fraction of fog droplets in the boundary layer t
#g = fog droplet density
~
^
oy
= vapor density '
8 = noncondensible gas density 9 I o
,tly=odov 8 + #g) y = coordinate perpendicular to the plate s = fog boundary layer thickness ,
s L = width of boundary layer
" av a + og ,
using the boundary layer approximations Y=YO (1 - y/a) (3.7) u=u - Z (j (au ) - y (8Z-)3} (3.8) u a (x) = a x 1/2 (3,9) 8u = a (x) (1 - () (3.10) where a = known constant i ( = known constant . U, = free stream velocity , l l 1 i I , l
!l' l
04300:1 3-17 l 1
' ~
~, _ _- __ , _ _ _ _ _ _ _ _ __ _ __ _
Substituting Eqs. (3.7) thrg gh (3.10) into Eq. (3.6), we have the rate j
. of fog formation [
g il l 025 ' U. 0'.25 , ( 3.1'1)
"r
- SL 8 Y o 3 s
Derivation of expressions for a, Y o, and C is given in Appendix A. Even though boundary layer fog formation may occur in any containment subcompartment, the fog formation rate is likely to be small except in the ice condenser. For fog formation in the ice condenser, L is the total length of the periphery and x is the height of the ice bed. During fog formation in the boundary layer, heat transfer to the cold
~
surface will decrease the bulk fluid temperature. If the bulk fluid temperature drops below the dew point corresponding to the free stream vapor pressure, then bulk stream condensation could occur. In this case, it is assumed that the boundary layer
- thickness, a, will grow so ,
thick that LaU becomes the gas volumetric flow rate Q through the con-V densing compartment. This is a very conservative assumption in terms of
- the fog formation rate. Under this assumption Equation (3.11) becomes 0.25 0.025 [ -(3.12) a =pQY ,
cond o j1-g g3, g)3 where scond is the sum of boundary and bulk stream fog formation rates.
~
I l t 0430Q:1 3-18
-- ~- - - - -
3.2.5 FOG DROP SIZES As mentioned earlier, when homogeneous nucleation commences, a large number of condensation nuclei are formed and they grow to the micron
, size within a few. milliseconds. In het,erogeneous nucleation, fog drop- ,
lets grow on aeresal particles, which are usually less than 1 u. 'In any case, the final drop sizes are determined by the atmospheric conditions f with which the drops are in th'ermal equilibrium. l Neiburger and Chien(18) studied the growth of cloud drops by condensa-tion and calculated droplet size distribution based on a cloud cooling rate of 6*c/hr. The initial size distribution of condensation nuclet (sodium chloride) were chosen to correspond to available observations as [ shown in Figure 3.4 (designated as O second). The calculated drop size 4 distributions at 3000 and 6000 seconds art shown 'in Figure 3.4. It is l seen that the sizes of fog droplets range from 0.01 u to 20 u. The j - volume mean drop size is 8 u at 3000 second. The volume mean drop size for homogeneous nucleation is expected to be smaller than this value. ] . Fogs of volume mean drop sizes ranging from 9 to 14 p(30) have been - observed to exist in a natural enviroment, e.g. valley. In the prtsent { study, a volume mean fog drop size of 10 u is chosen for fog deposition i
~
and inerting calculations. l l 3.3 FINE MIST DROPLETS FROM CONTAIM4ENT SPRAYS ! The contairment sprays produce fairly large drop sizes. A. typical con-tairment spray nozzle, e.g., Spraco 1713 nozzle, produces the size dis-l tribution as shown in Figure 3.5, using a pressure difference of 50 psi l across the nozzle (19). It is seen that kater droplets produced from ! contairunent range from 100 m to 2000 n. These large drops have little effect on hydrogen combustion and flannability ifmits, as already demon-l strated in the Fenwal tests (20) and more recent tests at Factory Mutual (21) . To affect the combustion characteristics of a hydrogen i mixture, the drop' sizes have to be' smaller than about 20 u, namely in
~
the fog drop size ranges. Since containment sprays essential 1y do not l produce drops in this size range, containment sprays will not be con-l sidered as a means to produce fog droplets. Rather, it will be con-( sidered as a means to remove the fog droplets. l i l , i ! 0430Q:1 3-19 l 1.. -
- - . _. = . - - - - - - --.
^ ^
.. ~... " ^
* --~ ~~ . . . . . -
j r I l I e s B . .t , t i ; i n .. i s isiti i a a 460
= i . si.u
= =
=
= = g m .
= 0 A
.,,. E A i- =
=
3
=
6
- i ios S
=
=
iv .
= =
g 8 M
* ,, M 4
M e J ese + g to8 y g g 5 . 2007 3 2 _
; l 1 .
3
'** a , s
=
=
~
- M E o _
io 8 $ _ :
=
io-! n . E 5 : I =
- I ~
i = t I. [
= ,
s
- is 3 g E .
l = . = M M fh 1 g h h h f'I l
- io 8 to ' ios io' io8 r-e FIGURE 3.4 CLOUD OROP SIZE DISTRIBUTION PREDICTED BY NEIBURGER AND CHIEN 8 .
1 I f 3 20 i r, .- - - - - - . . . - , , . , , - - - - - - , . - , + . , - , - - . , . . .---..--c
-----1.. ., - . - _ . , -, .-
^
. .^ . . . , . . . - - . - . . . . . _ . . .
e ' s, , a m N 8 t a 4
-t n5
}
O O
\ 80 s -
M N N o
\ C .
9 W= o N 2 N 4 $ N Om N
' O us C e z X a
-C E M 4J V h
ENN1 "E 8 L%%NN g-E $ m ANN %NNNN e E O h%N%%%\%9 gE - A%NN% %NNNN%9 E EN%NNNNNNNNNNNNNN%NNNNN , i
- L%NNNNNNNNNNNNN%NNNNNNNN%NNNNNNN%NN1 i mNNNNNNNNNN'%%NNNNNNNNNNNNN%NNNNNw , i -.
ENNNNNNN%NNM - O 6" h n W
~
w N h w t t n .J j (J l l m b O C C C O O O O E ] W N no e w N < ch. 4
- .C S31311.HYd 30 H38HAN m
= M LaJ K
D (A b I 3-21
,ey y
e , .,~ , --- ,---,
. __ r--
4.0 FOG REMOVAL ECHANISMS IN AN ICE CONDENSER CONTAINMENT In Section 3, the mechanisms of generating fog droplets were discussed. After these droplets are generated, they can be removed from the con-
. taf ruient. atmosphere by gravitational settling, vaporization, containment.
sprays, and impingement on structures. They can also coalesce with other drops during collision and form bigger drops. These bigger drops could easily settle out of the atmosphere under gravity. These fog droplet removal mechanisms will be discussed in this section. 4.1 SETTLING DUE T0*GRAYITY Drop removal rates due to gravitational settling depend strongly on drop radius. The removal rate increases if nearly with drop terminal veloc-ity, drop concentration, and settling area. The relationship may be expressed as i . I4 1) m set =Yt"A O where n is the mass of mist droplets per unit volume, and A is the set-j t1ing area. The terminal velocity, Vt , is a strong function of drop radius and the relationship is shown in Figure 4.1. It is seen that the terminal l velocity is approximately a linear function of drop radies in both lami-l I nar the turbulent regimes. For a 1000 u drop, its terminal velocity is : above 1 m/s, while for a 10 u drop, which is the typical fog drop size, 1 f ts terminal velocity is only about 1 cm/s. Therefore, there is very I little removal by gravity for fog droplets. i l 4 4-1 0430Q:1 I een l
. .. .
- j l
4.2 AGGLO ERATION , ( i
- Af ter the fog droplets are produced, the droplets will undergo changes in the number density and size distribution with time, when drops col-
** ' lide with each other' a'nd coalesce'. The agglomeration Nte (No. of par-ticle per unit volume per unit time) has been found to be proportional l to the square of the drop population density and the coagulation mecha-nisms dependent rate constant K(22) ,.
For drops larger than 1 u, the dominant mechanism is the difference in velocities between drops in adjacent streamlines. This is usually termed the velocity gradient coagulation. For drops smaller than 1 g, the velocity gradient effect becomes small, and drops are brought ] together by Brownian motion. This leads to greatly different agglomera-J tion rates for different initial drop sizes. A typical agglomeration rate as a function of drop size in a moderately turbulent atmosphere is l shown in Figure 4.2. In Figure 4.2, the sharp rise of the agglomeration ! rate with drop diameter larger than 1 u implies that the larger drops agglomerate quickly to the maximum stable size supported by the atmo- !h spheric turbulence. The agglomeration rates for drops less than 1 u are l Very small. ' Since most of the fog droplets are in. micron size ranges, the agglomeration rate is not large. It is assumed in the present analysis that the initial 4 u blowdown mean drop size will grow to 10 g !, (See Section 3.2.5) . Agglomeration as a separate mechanism for fog growth has been conservatively neglected. l i l , 4.3 YAPORIZATION Fog droplets suspended in the containment atmosphere is considered to be i in thermodynamic equilibrium with the surrounding gas. When the sur-rounding atmosphere becomes superheated or when the droplets are entrained into a superheated subcompartment, it can undergo vaporization i ' or condensation. 4 f ( i . 4-2 0430Q: 1
EE
~ ^ -- ~
_ _. 1 . . _ _ _ . ._ _._ . . _ _ _ . .
. e, .
A l l ? l In, the present analysis, it/is assumed that water vapor and af st drop-II ' lets are in thermal equilithium at all times. Therefore, the amount of vaporization or condensation will be determined by the thermal equilib-rium state reached by the vapor and drops. In other words, it is not intended to model ' heat transfer between the. drops and the surrounding ; gas, and thus determine the vaporization rate. This is a good assump- ' tion for the small fog drop sizes. 4.4 REMOVAL B,Y SPRAY DROPS As mentioned above, the containment spray droplets range from 100 u - ,
- 2000 u, which are substantially larger than the fog droplets. If fog droplets enter the spray zone, they will probably be removed by the spray droplets by colliding with them, since the spray drop mass is much larger than the fog drop mass. A simple analytical model is used in the d
- present study which assumes that all the fog droplets residing in the spray zone will be swept by the sprays to the floor with the spray drop removal efficiency E. The spray removal rate may be expressed as .
C (4.2)
"sp = E Qsp Mc I"sp V I where E = spray drop removal efficiency ! Q sp = volumetric flow rate of sprays n .= volume fraction of spray droplets in-the spray zone sp M = mass of fog in compartment volume V c
4.5 OTHER REMOVAL E CHANISMS Another similar mechanism for fog removal is the formation of droplets
- in the ice condenser. These droplets which would be generated in the
! ice bed when the ice melts, would fall through the ice bed, and remove
- fog droplets from the flow through the ice condenser. This large quan-tity of water would be effective in removing fog droplets. However, due j i
to difficulty in modeling this removal mechanism, it is conservatively l l neglected in the present analysis. ! I l ) 0430Q:1 4-3 1 I
-- ---.-- -. - - --.___-. - - - _ . . --. ~ . - - -_ --.-
. al
-- ~-
l In addition to the removal mechanisms mentioned above, fog can also be t - removed by impacting structural surfaces. Due to the inertia of fog droplets, substantial fog removal by impacting structural surfaces could occur, when the drop-laden mixture flow passes through long, narrow, curved paths, such as ice basket flow path's, and fan ducts. Moreover, the centrifugal force exerting on the fog droplets, when they pass , through the fans, could cause the fog droplets to firpact the blade sur-faces or other parts of the fans. These removal mechanisms are believed to be significant; however, they are conservatively neglected in the present analysis. It is, therefore, believed that the present analysis
.f s very conservative.
C 1 i , I e
/
I l
/
( . 4-4 .
, 1kr
- 0430Q
- 1 , '
. .s. . '
i
.e, .
21273 1 I -
- TERMINAL DROP FALLING VELOCITIES IN 10 - STEAM AIR ATMOSPHERES
',[r+ !
*p+ 4 .
1.0 -
,/
a -
=
3 - C - 0.10 =- TUR8ULENT - 3 MINAR REGIME
- REclME
~
HATCHED REGION INDICATES: 50 <Re <55
. 0.01 I ' I IIIII I I I I'III '
O.001 0.01 0.1 1.0 DROP RADIUS (CM) FIGURE 4.1 TERMINAL VELOCITY AS A FUNCTION OF DROP RADIUS Q IN STEAM-AIR ATMOSPHERE E - GRADIENT g I AGGLOMERATION 3 =_ h10 _
, , 3,5 CM'3 e
g i g....i l g - k 102 =- BROWNIAN g O AGGLOMERATION [ n = 105CM*3 g - 2 1 NET g 10 F RATE
@ E /
3,0 i i ii viiil ei f isirl i i rivist 0.01 0.1 1.0 10 DROP DIAMETER (uM) , FIGURE 4.2 AGGLOMERATION RATES IN AIR BETWEEN EQUAL-SIZED DROPS 4;5 re , . _ _ _ . _ _ _ _ . _ . _ , _ _ . _ , . _ _ _ _ _ _ . _1 _ _ _ _ __ ___ __ .
- .;_ _ . . . . . . _ _ . _ _ ___~ . .L __ . :__ _ . ._.
.. s .
5.0 FOG INERTING CRITERI A o . Recent %drogen burn experiments conducted at Lawrence Livemore Labora-tory indicated that substantial fog fomation could occur when saturated
. steam is discharged.into an unheated vessel. It appeared that this fog .
I prevented a glow plug igniter from successfully igniting the hydrogen mixture in the vessel. The ability of fog in inhibiting and quenching of hydrogen combustion can be explained as follows. The fog droplets suspended in the %drogen-air-steam mixture act as a heat sink that could absorb a large amount of coeustion heat, greatly reducing the pressure and et'mperature rises resulting from %drogen combustion. If droplets are sufficiently small such that they could vaporize inside the thin ( lam) flame front, the flame may be quenched or inhibited. For a flame speed of 2 m/s, the drop residence time is of the order of 0.5 x 10-3 s econds. (24) In .such a short period of time, the droplets of initial radius less than about 4 u will vaporize entirely in the flame front. The quenching of a propagating flame is also governed by the distance between droplets. As the droplets become closely packed, the total { droplet surface area available for energy loss increases. A critical spacing between droplets exists such that a large ' fraction of the heat released is absorbed, thus preventing flame propagation. This critical spacing is known as the " quenching distance", which is usually deter-mined by propagating flames in tubes. 5.1 PREVIOUS WORK The effectiveness of fog droplets in inhibiting or quenching a flame depends on its quenching distance, was determined by Berman et al.I24I as dq = [4y/S] crit (5.1) where Y is the gas voline and S is the heat transfer surface area. For a hydrogen-air mixture, the data on the quenching distance is shown in ( 0430Q:1 5-1 U- ,,_,n ,,,v ,,?,-, ...,-,n,,e,---nn,,,,. -n,n-
. :: . =.._.. .....;._
e *. . l 4 i i' l 1 ( - Figure 5.1. In the suspended fog droplets, this volume-to-surface ratio (i.e., V/S) is equal to I d (1 - n) i n where d is the mean droplet diameter and n is the volume fraction of water. When four times this ratio approaches the quenching distance, a critical drop'let diameter can be obtained as d * (5.2) c 1 Using this criterion for- quenching a flame, for a given volume fraction
- of water and gas composition, qd can be determined. The critical l
droplet diameter then can be determined from the above equation. The drop sizes less than the critical drop size is capable of quenching a flame. A plot of Eq. (5.2) for two hydrogen concentrations is shown in Figure 5.2. 5.2 PRESENT THEORY The previous theories do not model the heat transfer and combustion 4
- processes occurring between the burned gas and the suspended droplets.
I A new theory has been developed, which models the heat loss and combus-l tion. 1 l i ,( ' ! 0430Q:1 5-2 =- . ,
.. _ _ - . _ - ~ - _ _ . - _ _ _ -. .
.- e. .
( - 4
* ,n = , . ,
l l
~~
l i i n . . . *~
. ./i y ./
- ! ,i nI i i l f /l / M I l\ % l i I7 # //l =
l _
, r % -1 ,'i'
, .x,x . . .ev, . . .
1 1 1 Al 1 1/A I I I I I W, I I I . 1 :
, , , , > r- 12
- is
, tu 1 i kh k N 9 4 0 0 4 l J $ [I
\ \'l I i i l 1 '
/ s7 /
f _ ! i IW II I 2#r j t + == ! FIG. 5.1 MINIMUM IGNITION ENERGIES AND QUENCHING DISTANCE FOR HYDROGEN-0XYGEN INERT GAS MIXTURES AT ATMOSPHERIC PRESSURE I i i > f e , 4 i l . 5-3 l 1 ._--._--_ _ _ ___ .-...__ -.....__ _ __._.. _ .--.._.-__.-_. _ _ _ . . _ . _ _ , - - _ _ . _ . _ . _ . _ . - . _ . . _.---__.#
~ . , .. . . _ .. . ._ _ _ _._ _ . . - - _ _ . . _. _
i . s. .
- j I
i . f - -- I o I i -
*i*"* CO mbuST.. ion 4
, . P^* *, A, .
) 8 . , . r2 . l- - -- . 1
- , n .
I I 1 l . ,
. l H - mg -- .
' l- -
, .. g,,;.y.,
-- Droo. le t ..,ogs..ws th. _; w.=
1 . 7= =._a
,-lamo.vej meze_.cu.dioWf--63s; l l
! Y ! w =gropaga, . .m- , , ..
-N- , , .
i i l m_. w . ; ; f f l
. / .
- Il i., . . ..
' ' L a._ _. '
i
; m_c,le -fractiori.-o y ,,,,,,,1o s f-fl 2 :
!.i "o. . _,, em / z
. is . v.s ?. .
i d'L. 1 am - 4
. o._ _
o
- i. .
(A , 1 i q- ._ i 1 . , . f. 7
. 4.
r
- . i.e . h+ '
' ' ' ' i i .. Ed ' ,
2 >< , , , .
- 1 < .
4 ,
. f . f s,s
/
. s '
I' s 1 p -
.u . s
,
- l
,
- 1 .--
4,
. t
., / --
, /
1 .
,' .. X.-=--3 0 g;;,.' . .
r . I V_. _ f .
' # 4G " JeD9 NI J U[ / ' '
. ~
. - s * *
- f.;.
\ .. _ j -.- _. f . . .. l l
/ , e
' * , _, /_ . y ' . . ! ..
..,' l ,
't f t 0, 4 2. . .
..S 42 n n.
n-M ' e n .Wh I , . 3 i
' WATER CONTENT.-.4- -(by Lumej
; i. .
. . # . 1
; ... i i .
j , . . ... . l . i k(,! FIGURE 5.2 THE EFFECT OF DROPLET SPACING ON FLAME QUENCHING 4 4 i 5-4
~ ~ - . - - - , - _ _ _ _ _ _ _MWde wwww ym
Consider 'a hydrogen / air / steam / mist droplets mixture in which a flame is
. propagating. The flame may be divided into three zones: heating zone,
. reaction zone, and post-reaction zone as shown in Figure 5.3. The l unburned gas at temperature T, move in the reacton zone with the
. Iaminar burning velocity 5,. If the unburned gas density is o,;
then the constant mass flow rate a is equal touu p I . W un k ned gas is heated to ignition temperature Tg and burned in the reaction zone to reach the flame temperaturef T . The fog droplets will act as a heat sink that reduces the flame temperature. The problem has been formulated and sol.ved by von Karman(25) . In his fomulation, three energy equations, which' incorporate the heat loss tems, were written for the three zones described above. The solution to these equations yields the following relationship i u
- 1 2 2 Ke g = {1 - exp ( 7 u ) (Y, - Yf)
] x [1 + / g ,{4 gf,)2 J x f1f gj g where = og dp (79 - TuI/9 1 i
/ UP/ Iw i
Keg
= (S/dw)o p g l
t a the ratio of heat loss rate per unit volume to the I heat release rate by chemical reaction per unit i volume l I l q = heat of combustion
)
C = mean specific heht . P f i l 4( ! 0430Q:1 5-5 l 1
I = heat conductivity I w = reaction rate (mass of fuel consumed per unit time ! per unit volume) l l
= hydrogen mass fraction in the heating zone j Y,
Y = hydrogen mass fraction in the reaction zone f a = ou Su A plot of Eq. (5.3) is shown in Figure 5.4. It is seen that for a' given Key, there is a minimum value of (Y, - fY )/e g . Below this sint-
; num value, there is no solution for the q u. Therefore, this value is considered as the flammability ifait. At the flammability 1
limit, the value of Key can be determined from Figure 5.4 or from Eq. (5.3) as 1 I I/'t) (5.4) i (K) crit 'l " f III u ~ f O
'A plot of (K) crit '1 as a function of (Yu " IfI/*i is shown in Figure 5.5. Equation (5.4) may be expressed as ,
2 f Iu -I fi n , 9 P '" 3" IYu ~III f 't ) (5.5) 2 f 12 1 (T9 - T,) 2 Detailed derivation procedure for Eq. (5.5), is given in Appendix 8. Using the data on Su from Reference (26) we can calculate the right l
- hand side of Eq. (5.5) for a given composition and initial gas tempera-j ture.
! 5.3 YERIFICATION OF THE0 RIES BY EXPERIIENTS i ! Experiments have been conducted at Factory Mutual to study the effects !( of water fog density, droplet diameter, and temperature on the lower j 04300:1 ,i 14 ---
,-e w. . - . . ~ . . . + - - -e# .,,, .wew..,--,.c,mm-e,, ww-,,.-w-.r-----.-,,_-+a 1 F - . g. e.w,se,- , + 2 -.
s *. e ( - r. remeen.r monote
- v. Jb .sr,.r,va-m mm m mm ,.o ,..
ms- m e on = ,, FIGURE 5.3 SCHEMATIC REPRESENTATION OF TEMPERATURE PROFILE THROUGH THE FLAME FRONT' O % g g .n.e o f ine&@
;f .um.o.m C ,
A.Ogo _ lp C (
.n,.oso -
0 I k o-Ng N o n. c n M e u a (r.-r,Ve, l l
~
FIGURE 5.4 THE PARAMETER Kt u AS A FUNCTION OF (Yu-Y)/0pFOR t DIFFERENT VALUES OF K0t ( 5-7 u ,
~
l, O - - - j -
.1 j- 0.3
- i 1
r 4 1 l o.2 - - i . n' . t l m I E U>. .i i .I _ - 1 . o.i -
+
,i .
. I-o.o I 1 I I I l l 0 1 2 3 4 5 6 7 s ,
(Yu - Y )/0 g g . n FIGURE 5.5 (K) crit ~0. t AT Tile FLAPfMBILITY LIMIT AS A FUNCTION OFu(Y - Yf )b. t 1 - i,-
. i I
. ~ .
g s-1' 2
e flamability limit of hydrogen-air-steam mixtures (21) The results 1
- indicated that most of the fog nozzles tested at 20*C only changed the Itait from 4.03 volume percent to 4.76 percent, corresponding to fog concentration in the range ,of 0.028-0.085 voltane percent, and average drop size ranging from 45-90 aferons. For the 50*C case, the lower 1
' ' flammability Itmit increases to 7.2 percent, corresponding to 0.01-0.04 volume percent of fog and 20-50 micron average drop sizes. The results ;
demonstrated that the fog inerting effect is more pronounced at small drop sizes. , Figures 5.6 through 5.8 show the comparison between the test data and j the theoretical predictions. For this comparison, the present theory used the free stream temperature to calculate the thermodynamic proper-ties used in Equation (5.5). This yielded somewhat higher fog concen- !
; trations than those calculated by use of the mean of the flame and free !
I stream temperatures. In Figures 5.6 and 5.7, the data suggests a linear ' relationship between the volume corcentration and volume mean drop size f' l ! on the log-log plot. It also suggests that the minimum fog inerting i ! corcentration varies approximately with the square of the volume mean drop size. In this regard, the present theory is consistent with the i, data while the Berman et al. theory is not. - { The present theory is in good agreement with the Factory Mutual data at 4.76 percent H2 ; .however, it overpredicts the minimum fog ine'r ting l concentration at 7.2 percent H2 . The cause of this discrepancy is still unP/twn. The discrepancy may be caused by the uncertainty' of the q, data.' The following discussion supports this claim. The fog droplets : are very small and they vaporize very fast in a flame. Therefore, the I fog droplets' behave as steam except for their larger heat absorption I I capability. When the fog droplets vaporize, they absterb the heat of -
- vaporization which is much larger than the steam sensible heat,. . Typ-
- feally, the heat of vaporization of water is abdut 10)0.StuAs and the !
s average specific heat of steam in the temperature range of irxerest is ts about 0.48 8tu/lb. It is well known that a Wdrogen flame carmot propa-j gate in steam higher thait about 54 percent in a steam-air mixture. At 7.9 H2 , the adiabatic flame temperature is about 1240*F and therefore i ! , . 5-9 I 0430Q:1 , s _-_ -' " T * ~r:W ".~ ~ .: ~~
. .. . c
< [.,
/
- i. .
- l y
. . . ..t.. .
. ,.. . . . . ..g ; . . .....
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- OSpreco 2020-1704 U -
vSpraco 1806-1605 a
. a -
^
u . .
. E 10-2 ;_ ,
m - -
, C s
~
o - NON-FLAKdBLE ZONE - m T , . 4 -
' C'
~ ~
PRESENT THEORY 5 - - g .
- O E v
iO-5 . w n -
=-
5 - .s s FLM9%BLE ZONE - .*,( - u - , s -
- g. ..
BER.*MN ET AL. - THEORY ft - 10'"
' I ' ' ' ' ' ' 'l i ' ' ' ' ' ' '
IQ 10 0 200 t YOLUME MEAN DIAMETER, MICRONS 4 FIGURE 5.6 COMPARISON BETWEEN THEORIES AND FACTORY MJTUAL FOG INERTING EXPERIMENTS ON 4.76 PERCENT H 2 t 5-10 b
$ 4 Y
4
. ~=,-,
l l l 10'3 i i , , , , , , 9 - O Sproco 2163-7604 - 8 - 7 Sproco 2020-1704 7 . OSonicore 035H - 6 - 2 E5 - -
"t Non-Flammable
&4 -
Zone m f
.C .
g3 -. Flammable -
~
'"'""'"' ~* "*
O s>= . 5 - . y 82 -
% Hg in Air At~50*C-8 L&.
- O ig s i i , i i i , ,
io ao so 40 50 so 70 e.o so, r VOLLPI MEAN DIAMETER, MICRONS.' FIGURE 5.7 COMPARISONBETWE(NTHEPRESENTTHEORYANDFACTORY MUTUAL FOG INERTING EXPERIMENTS ON 7.2 PER6ENT H2 ( -.- 5-11 '
~
S $ e
- h
-e. ,--m . _ _ _
f 10'
- /
i i . ,
.l. . , ,.
} f 5 X 10' EE NON- FLAM. M. ,_A.B.. L E. . ZON E ,J n , l l x --
E /- PRESENT THEORY m + W u. s . oN ' g .
. i . .
. . i. . ., .... ..... ... ... ... ...
- - 3_ .:!: :::: l%.;' ::: ;;' ; ,' FLAMABLE . ZONE, .ll l ,': 'll, l
+-
u 10 I i = /
, o M
9
~ -
' N t yy s
j s . . .J . a.,t.
. .m . ., . . ,1 ..., , , , .
-- 1
& 4_
- 5 X 10 5 a FACTORY MUTUAL DATA q
03j v 7 ON, 7.9% H9 I N Ar I,R . , , , - , , . . , . , 1 . . 8 u. t .
. . , im
?
4 e' . . , ,.o u.. . , .,, ,, , . .., , [' l 1 , I 'tf I' I ,8. 4 i ei ,l l . g .i. .
..,.f
-4 ' ' "" "" "" ' ' ' ** '" "- - '"' '"- '
- 10 50 100 500 . 1000 <
VOLUME MEAN DIAMETER, MICRONS 4 i FIGURE 5.8 COMPARISON BETWEEN THE PRESENT THEORY AND FACTORY MUTUAL FOG INERTING EXPERIMENTS ON 7.9 PERCENT H 4 2 O a 5-12 6 e
- ., ,_ , ri
f the increase of the steam sensible heat is about 540 Stu/lb. Conse-quently, for the same amount of fog droplets and steam, the fog droplets heat absorption capability is about 1.9 times higher. This means that the fog concentration which is equivalent to 22.1 percent steam in steam I and air is capable of inerting 7.9 percent H 2
. This fog inerting concentration wa's calculated to be 1.61 x 10-4 To inert 7.2 percent H2 , a minimum fog concentration which corresponds to about 21.3 per-cent steam in steam and air is required. This gives a minimum fog inerting concentration of 1.56 x 10-4 for 7.2 percent H2 These
. estimates show that the present predictions are reasonable and conserva-4 tive. The present theory is conservative because it neglects convective and radiative heat transfer and thus underpredicts the heat loss. The estimates are consistent with Factory Mutual data on 7.9 percent H2 but not on 7.2 percent H2-It should be noted that in the tests three fog concentration measuring l . techniques were used. These three techniques gave substantially dif-l ferent results. The discrepancy is at least one order of magnitude l- dif ference. The fog concentration data presented in Figures 5.6 through 5.8 were obtained from one of the techniques. In view of the uncer- -
tainty of the data, care must be exercised in using them for inerting analysis purposes. They should be used in conjunction with the present fog inerting criterion in the assessment of fog inerting potential in the ice condenser plants. Some uncertainty also exists in the present fog inerting theory. The uncertainty associated with the underpredic-tion of the heat loss and temperature dependence of the thermophysical properties is estimated to be +63 pert:ent. It should also be pointed out that the Factory Mutual data and the pre-sent theory can only predict the minimum fog inerting concentration. To t insure hydrogen burn in all directions in the ice condenser upper , plenum, further work in this area may be required. 4 l !(
~
l 0430Q:1 5-13 l I u __ _ _ _ =. ,
l
. I l
6.0 ASSESSENT OF FOG INERTING PROBABILITY
- IN ICE CONDENSER CONTAINENTS i i
i As discussed in the previous sections, there exists several mechanisms I~ .- of generating and removing ' fog droplets from the ice condenser contain- - ment. In addition, fog droplets are also transported from one subcom-partment to another by entrainment in the gas stream. The fog entrain-ment rate is difficult to assess without knowing the velocity field and drop size distribution. For simplifying purposes, it is presently assumed that, 'the mass fraction of af st droplets in the intercompart-mental and fan flows is the same as that within the subcompartment from which the flows are originated. This is a good assumption since the fog droplets are small. The amount of fog droplets in a subcompartment depends on all these mechanisms. The total amount of fog droplets is important in determining the volume fraction of suspended condensate in a subcompartment. This volume frac-tion, in turn, is used in the fog inerting criteria to determine whether a particular hydrogen mixture composition formed in a subcompartment at l any time is flamable or not. In other words, by knowing the hydrogen concentration and the mean' fog drop size, we can determine whether the calculated volume fraction of fog droplets is high enough to prevent the mixture from combustion. i 6.1 DETERMINATION OF YOLUME FRACTION OF MIST DROPLETS IN ICE CONDENSER CONTAINMENTS l Consider a subcompartment in the ice condenser comtainment as shown in Figure 6.1. There exist several mechanisms by which mist drops can be generated or removed. Fog droplets can be generated by homogeneous or heterogeneous nucleation in the thermal boundary layer and/or in the bulk stream and they can increase in size by condensation or decrease in size by vaporization. The rate of generation of af st droplets by con-densation and their continued growth (or shrinkage due to vaporfzation) is represented by icond. The other mechanism of generating utst drop-1ets considered in this analysis is the primary coolant discharge from the break and the rate of generating fog droplets from this mechanism is t 0430Q:1 6-1 l l _- . . _ . _. __. , .n .
~
4 represented by ibreak. Two feg droplet removal mechanisms are consid-l i, ered in thf s analysis: one is gravitational settling and the other is removal by containment spray. The fog droplet removal rate by gravita-tional settling is represented by i set and that by spray is represen-ted by 6,p. In addition to the generating and removal mechanisms discussed above, the mist droplet concentration in a subcompartment is
.also affected by the intercompartmental and fan flows. In the intercom-partmental and fan flows, the mass fraction of fog droplets entrained is n and the gas mixture flow rate is 6. Therefore the rates of fog drop-lets mass into and ou't of a subcompartment are Z n9, A9 , and Znout "out, respectivey1' . It should be noted that Znin "in and Z n out bout include the fog mass entrainment rates in all the ,
intercompartmental and fan flows into and out of a subcompartment. The mass conservation equation for the fog droplets in a subcompartment ] , cay be expressed as di dt "in in ~ "out out + " break + "cond ~ " set ~ "sp (6.1) O . . where { is a summation over all the flow paths. In Eq. (6.1), if cccM is negative, then it becomes the rate of vaporization. Eq. (6.1) can be integrated to give the total mass of condensate at time t
~
! t . . .
~
i Mg (t) =of ([nin "in "out "out + " break ; dt
+ "cond ~ " set ~ "sp 4
~
- i
- i 1 + "* i E"out~1 out"*
~
f Ib"in "in break I
+Econd - " set ~ "s ) at, (6.2) 1 1
( 0430Q:1 6-2 4
- . - -- e-- ..- - . . , , ,- . > , , - + - . - .
=
l - The present analysis will emp1gy the CLASIX calculations of containment j i transientduringasmallLOCA.[.j In the CLASIX analysis, the entire ice condenser containment is usualky divided into five or six subcompart-ments for analysis purposes. Temperatures, total pressure, steam l partial pres'sures, and intercompartmental flow rates are calculated j during transients. This information is used in Eq. (6.2) to determine fog droplet mass. When applying Eq. (6.1) to each individual subcompartment, we have the following fog ma'ss conservation equations in finite difference form: Upper Compartment NC It + 'tl " NCItI+(E"in"inIt)
-["out"outIt) + kC,condItI I'*3I .
O . .
-m VC, set (t) - mVC,sp(t)] gg Lower Compartment
\C It + 't) " NCItI+(E"in"in I
-E"out"out * "LC, break (t) (6.4) l
~
C cond(t) - kC, set (t) - kC,sht) ) at ( l 0430Q:1 6-3
,a_j
r = . l Ice Condenser Upper Plenum ,
)
( - l Mgp(t + at) = MVP(t) + ([nin "in(t) - [ nout "out(t) 1
*kP,condIt)~kP, set (t)} at Ice Condenser Lower Plenum .
MLP (t + at) = Mtp(t) + ([nin "in(t)
. (6.5)
~ b"out out(t) + kP,cond(t) - kP, set (t)) at Dead Ended Region MDE (t + at) = MDE(t) + ([nin "in(t) (6.6) c ~b"out out(t) + DE,cond(t) - mDE, set (t) at i s 4
~
Fan / Accumulator Rooms
- YA (t + at) = YA(t) + ([qin "in(t) (6.7)
~ ["out "out(t) + mFA.cond(t) l I
~
A, set (t)-hA,sp(t))at 8 In the present analysis, the fog concentrations in the intercompart-mortal and fan flows are assumed to ]he the same as those in the compart-ment from which the flows are orginated. - 1 These rooms were analyzed only for the D. C. Cook plant (See Figure 6.9). 0430Q:1 6-4
In the equations given above, the intercompartmental and fan flow rates . are provided by CLASIX calculational results. The fn and iout i procedures of calculating fog droplets generating and removal rates are based on the discussions in the previous sections and the detatis are ,
, ~
given in'the following sections. 6.1.1 CALCULATIONOFkREAK To date little experimental data is available to estimate the amount of fog droplets generated'by the break flow. For a large LOCA, Almenas and Marche 11o I3I estimated that 13 percent of the total' blowdown drop population (by weight) has drop radius range from 1 u to 20 u and only 1 percent less than 1 u. This estimate is somewhat larger than the 4 u mean drop size sited in Section 3.1.2, which is believed to be conserva-ti ve. Since we are only interested in fog drops smaller than 20 u, and only these drops can remain suspended in air untti the time when the hydrogen is released, we assume that the estimate of Almenas and Marche 11o is Q applicable in small LOCAs and 14 percent of the suspended liquid are fog, droplets which have a potential inerting effect.- The fraction of reactor coolant discharged from the break remains as suspended liquid has been determined in Section 3. Knowing the break flow rates from a computer code such as MARCH, we can calculate the amount of liquid suspended in the atmosphere. Then from the drop size distribution we can calculate the amount of fog droplets suspended in the atmosphere. Defining the blowdown rate as k, the liquid fraction of the break flow as (b, the fraction of fog droplets smaller than 20 g as fb ' we have a E (0*0I break " Ib b i( ) 0430Q:1 6-5 ) . ~ .
, . = - . . _-
^ ^ ^ ~
. . . . . . . T.: ;. n..- ' ~^. . . _ _ ~ ^
e In the present analysis fb = 0.14 is used. f becomes b zers when the ( , water 1.evel in the reactor vessel falls below the break elevation. 6.1.2 CALCULATIONOFACOND As discussed previously, acond is the rate of formation of mist dmp-lets by nucleation, condensation, or vaporization. Nucleation of fog
- droplets can take place in the thermal boundary layer and in the bulk fluid. We conservatively assume that little supersaturation is needed for nucleation in the bulk stream and fog will form when the bulk stream steam partial 'pressurt reaches the saturation steam pressure correspond- 4 ing to the gas stream temperature. Therefore, the bulk stream fog formation rates can be determined from the equilibrium thermodynamic
~
states of the gas mixture. The boundary layer fog formation rate can be determined using the > Hijikata-Mori theory (17I of fog formation in the thermal boundary layer as discussed in Section 3.2.4. The fog formation rate in the thernal boundary layer and the bulk stream is given by Eq. (3.12). Boundary layer and bulk stream fog formation rates will be calculated !. for the ice condenser and lower compartment. A computer program called F0G has been developed to calculate acond* This computer program requires input of the volumetric gas flow rate, { gas and wall temperatures, total pressurt, and steam partial pressure. This information can be obtained from the CLASIX output. ; 6.1.3 CALCULATION OF M SET i ! The rate of settling of the fog droplets depends on their terminal velocity, concentration and compartment cross sectional area. The l droplet terminal velocity is a function of drop size. In the present . stu(y, Equation (4.1) will be used to calculate the fog gravitational ! settling rate. - - t i 6 0430Q:1
l
* \
' I 6.1.4 CALCULATION OF l13p l l
Tiie mass of a fog droplet is much smaller than that of a spray droplet. . Therefore, when a spray droplet collides with a fog droplet, the fog droplet will coalesce with the spray drop and fall to the sump. In the - present study, the fog removal rate by sprays is given by Equation (4.2). It is expected that the ' spray drop collection efficiency is very high, and therefore a 100 percent drop collection efficiency is assumed in the analysis. A sensitivity study is needed to be carried out to study the effect of E on the volume fraction of fog droplets. A computer program called FOGMASS has been developed to solve Eqs. (6.3) through (6.7). This program uses a finite difference numerical scheme j to carry out integration. This program takes input from FOG and CLASIX 1 output data. Specific output data from CLASIX are time histories of gas j temperature, wall temperature, total pressure, steam partial pressure, j and intercompartmental and fan flow rates. O s.2 r0G IncaTIaG ra0sAsILITv In Tac Scou0ria eLaur
- l. The computer codes, F0G and FOGMASS, were used to perform fog inerting j.
j analysis for the Sequoyah plant. F0G was used to calculate the rates of l fog formation due to boundary layer and bulk stream condensation in the - l Sequoyah ice condenser and lower plenum. Then these fog formation rates j were used in FOGMASS to compute the fog concentrations in each of the Sequoyah containment subcompartments. i To ccepute the fog formation rates in the ice condenser upper plenum and lower compartment, some output data from the Sequoyah CLASIX analy-sis (27) are needed. These data include time histories of gas tempera-ture, wall temperature, total pressure, and steam partial pressure in sach containment subcompartment, as well as the intercompartmental and ! fan flow rates. In order to utilize the CLASIX output data, the ice condenser containment is subcompartmentalized in the FOGMASS program in exactly the same manner as in Reference 27. The subcompartmentaliza-4
.. tion model used in the Sequoyah CLASIX analysis is shown in Figure 6.1.
In this study only the S D2 accident scenario has been analyzed. 04300:1 6-7
- . - _ . , . - . , . . r, . --- , , , , , . . , - - , . - .@-
. . . . . - -~~~ ~
( - FIGURE 6.1 SEQUOYAH CLASIX CONTAINMENT MODEL ICE ***----== CONDENSER _
' UPPER UPPER ~
j ] PLENUM COMPARTMENT !
' - f A I A :
i
- I I
ICE 8ED l
+ _____,
s s IM _ 4 4 CONDENSER t DEAD
- LOWER 4 LOWER ENDED P W UM
~
< .O l COMPARTMENT REGION 6
"" ~~~
AIR RETURN PAN / HYDROGEN SKIMMER SYSTEM
- i FLOW PATH CONTAINS DOORS .
T ' FLOW ALLOWED IN 8OTH DIRECTIONS ! m FLOW ALLOWED IN ONE DIRECTION
........... SPRAY HEADER -
~
( . 6-8
--ge - , -r - - - - - - , -, _ ~ ,ee- , - - , - , . -- - - - - - .
=-
. I 4
The F0G input data for Sequoyah S 2 0 Case 1 are given in Tables 6.1 and
,6.2, and the calculational results are shown in Figures 6.2 and 6.3. In Figum 6.2, the fog fomation rate in the lower compartment is shown.
Fw the first few hundred seconds the wall temperatum is lower than the dew point corresponding to the/ steam par'tial prissure and therefore fog starts to fom" After about 600 seconds, the fog formation rate becomes negligibly small since the wall temperature is only a few degrees below the dew point. Thert is no fog formation in the lower compartment after about 1800 seconds. The fog formation rate in the ice condenser is shown in Figu're ,6.3. It is seen that the fog formation rate in the ice condenser is muc'h larger than that in the lower compartment. It increases with the ice condenser steam flow rate and reaches a peak of 14 lb/sec at about 1800 seconds. The fog formation rate in the ice condenser' then begins to decriase and is low at the time of significant hydrogen release. The nine fog fomation rates in the lower compartment and in the ice condenser are input to FOGMSS in a tabular form and there is a built-in interpolation scheme in FOGESS to obtain values for the intermediate time steps. FOGMSS computes the rate of fog generation by the break flow, the fog I settling rate due to gravity, and the fog removal rate due to sprays, as ' well as the rates of fog entrainment by intert:ompartmental and fan flows. The input data needed to calculate each of these rates are dis-cussed as follows. The rate of reactor coolant release to the containment and the coolant - enthalpy were obtained from the MRCH outputII for a small LOCA. The quality of the break flow was calculated using the enthalpy and the lower compartment gas temperature. According to the MRCH predic-tion I7I the disci.arge of liquid by the break flow into the lower com-partment lasts for only 2172 seconds. Afterward, the water level in the reactor vessel drops below the break elevation and the fluid discharged j k l t t 0430Q:1 6-9 l
'd.
0*0005
~
z i_ .
-r
+
x '
< j a .
1
- g. . . .
o - 2 , 0'000* ., x W . r o a . z - o o 0*000E o w D
- c:!
r e
~
O z o 0000e$ - u-i
. - ~
z w x v, . o - : 6 w a , 1 o 1 0'0001 ~ i l
/
/
/ ,,
4 0'0 O O O O o Q l o o o a o a o O O o o o
- e. <> < m m - o.
/;r I e o e
o o 1
- o o .
(33S/81) ,31V8 N 011V W 8 03 003 6-10 l 1 l y s.. - , - . - _ _ q
- I ~
0*0005 - 1 i
= :
w 6 - C 3 x
- 0*0004 5
= w a 5 a
w v E u 5 Ti e 0*000E 3 CT . g"
. w .
~
y O 5 -
~
z j
~ ' - .E x
/ 0*0002 2 e
C> -.
. -
- w %
< w z
- 8 m .
C w - Led M, z e o e N N 0*000L "
~
w
" !!!r
~N E N
- j 0*0 e e e = 8 8 E 3 5 5 E. *. S E E, .
g & ~ c
- A =
'" i i ~ -
(33gf gm livu. N011YHu03 003 ; 6,-11 t
i .
- from the break is essentially steam. Therefore, in the present study,
! , , it is assumed that no fog is generated by the break flow after 2172
- seconds.
i
, For fog removal by gravitational settling, a volume mean drop size of .
10 g was assumed. The terminal velocity of a 10 m drop is about 1 cm/sec. Because of this low terminal velocity, gravitational settling is not an effective fog removal mechanism. The assumption of 10 g volume mean drop size is therefore conservative, considering the fact that for a few thousand seconds the drop agglomeration mechanism would be able to increa'se volume mean drop size substan' 'y. It should also be noted that a smaller volume mean drop size mean at the minimum fog inerting concentration would be reduced and thus makes the present analysis conservative. Furthermore, no consideration was given to the deposition of fog on the walls and vertical surfaces of the structure, or for fog removal in the fan flows when it passes through ducts and fans. All the assumptions mentioned above make the present analysis very conservative. The containment geometric data needed in computing ! the settling rate are given in Table 6.3. O For fog removal by sprays, a spray flow rate of 9500 gpa was used for Sequoyah. According to the Sequoyah CLASIX analysis (27), the spray's are initiated at 142 seconds. A volume fraction of sprays (volume of , j sprays divided by volume of the spray zone) of 3.3 x 10-4 was.used, which was obtained using a spray drop fall height of 107 ft, a spray l zone volume of 485,500 ft 3, and a volume mean drop size of 700 u. .As 4 previously discussed a spray removal of a 100 percent was used. In Figure 6.1, the directions of the intercompartmental fTows are shown. The intercompartmental flow rates for the six. flow paths and
; nine time steps were obtained from the OPS CLASIX analysis and are given i in Table 6.5. The present analysis considers the intercompartmental I flows' as the mechanisms of transporting fog from one compartment to I ano ther. It was assumed in the present analysis that the fog concen-1 trations in the intercompartment flows are the same as those in the compartments from which the flows are originated.
4 - 0430Q:1 6-12 1 & m
.g s
1 I
- It is seen in Figure 6.1 that two trains of the air return fan and hydrogen skimer system take suction from the dead ended region and from the upper compartment and discharge into the lower compartment. The fans are initiated at 712 seconds. The fan head-flow curve reported in -
Reference 27 wss used to compute the fan flow rates. Fan flow rates of 3 3 1645 ft /sec and 10 ft /sec were used for the air return fan and the hydrogen skimer system, respectively. These flow rates were calculated using average ap's betweert the upper compartment and the lower compart-ment, and between the dead ended region and the lower compartment. It was also assumed.that the fog concentrations in the fan flows are the same as those in the coigartments from which the flows are originated. The results of the F0GMASS calculation are shown in Figure 6.4. It is seen that for the first few hundred seconds the fog concentrations in the lower compartment, ice condenser lower and upper plenues are about the same and increasing. At about 700 seconds, the lower compartment fog concentration reaches its peak of 2.2 x 10~4 . Afterward, the l intercogartmental flows transport more fog droplets out of the lower g compartment than are generated by the break flow and condensation and, therefore, the lower compartment fog concentration decreases. However, the upper plenum fog concentration keeps rising until about 900 seconds, l due to an increasing fog formation in the ice condenser and more fog , entrained in the intercompartmental flow into the upper plenum. The upper plenum fog concentration reaches its peak of 5.4 x 10-4 at about 900 seconds. The lower plenum fog concentration is almost the same as the lower compartment fog concentration because of little difference in the intercogartmental flow rates into and out of the ice condenser lower plenum. Therefore, these two volumes behave as a single volume in terms of fog concentration. At 2172 seconds, the break flow in the lower compartment stops genera-ting fog and, therefore, the fog concentrations drop sharply there-after. The effect is more pronounced for the lower compartment and lower plenum fog concentrations. The highest fog concentration exists in the ice condenser upper plenum while the lowest exists in the upper compartment. The effect of sprays on the upper compartment fog conceri-( tration is clearly seen in Figure 6.4. At 142 seconds, the sprays are - 0430Q:1 6-13 . l __ _ _ - , _ _ , ., _ __ _ _ _ .I
, o ,
0'0009 c., ,. ,.s l e y o', P ,is [ c. c, c. , , , ,,, I-- O e " t er o s - b ". d xm znm e., ,
/
a
.L p
I 0'0005
*IC <
i h P i':1
- C /
zI y1 e om* uz" , t d I.1 n W
=a o - 1 0 000g
< c. W < r
- >- ep CCO 2 w j!, en i
o .=, ,s C = 1 w o e' <- es e u. .- e v, .- ,s gu O-I f
-o , 0*000E i
f z es 7 :e. ; o lL =u w
, )
f ,* O g& et
-z=
x r I, , , -- y - Q , <
.=,
lll'l* Q 11 - $ w u ir u
/
0*0002 wu j z c: , ..., - a m . owI - umo n, ... d . w
<.2 ez a w , ,' 1 I -
e n _.: w - c. , H 1, . 1,h \
-l s
0*0001
' h *- --
M .... , s -
.ll.
( g n n ,
=,--
W
'n.,
}
n p :,c,
=.- ..,
/ -
O'O m n n v w u x a.o w w u:r ~ ~ e .am em m . CC C CCC CC CCC CC a : : CCC QC CC"3D
- : : : : : :: : CC: C CCC, Cr -u mm ww w
""CC CC r =.v")
C'X"" CC C CmN C CCC - wwCC w - ww CC CCCCC - ww CCC CC C CCQCC QCC cc CCC . . . . . QC CCC . . . . . QC CCC . .. .. oc CCC . . . cc . . Q . CCe Ce Cem ee e u- mm % mm % mn.: % mm - ( X I H E **.U / 0 2 H E **.LO N O 11Y 81N 3 J N 0 J 9 03 6-14
m - - turned on and the upper compartment fog concentration drops sharply until about 600 seconds. At about 600 seconds, the upper compartment fog concentration starts to increase again because the intercompartmen-tal flow into the compartment increases sharply at that time. A peak . . concentration-of 7 x 10-6 in the upper compartment is reached at about 1200 seconds. Hydrogen starts to release into the containment at about 3804 seconds, i according to the MARCH calculation (27) . It reaches 4 volume percent at about 4300,'.4400, and 4670 seconds in the lower compartment, upper l plenum, and upper compartment, respectively. At 4300 seconds, the calculated icwer compartment fog concentration is 9.7 x 10~I, which is about an order of magnitude smaller than the ! minimum fog concentrations required for inerting 4 percent H . At 2 4670 seconds,' the upper compartment fog concentration is 1.35 x 10-6 , i which is about a factor of five smaller than the minimum fog concen- . tration required for inerting 4 percent H2 *. At the times of reaching Q 8.5 percent H2 , the fog concentrations in the lower and upper compart- l ments are even lower than the figures given above. Therefore, it is l concluded that the fog concentrations in the lower and upper compart- ! ments are too low to have any inerting effect. The use of the present , theory on fog inerting also leads to the time conclusion. However, at .4400 seconds, the calculated fog concentration in the upper plenum is 6.1 x 10-5 which is higher than the Factory Mutual fog inerting data extrapolated to 10 u drops and the present theoretical prediction. The data shows that in order to inert 4.76 percent Ng the fog concentration must be 8.4 x 10-6 or higher for 10 u volume mean drop size. At 4600 seconds, the upper plenum hydrogen concentration reaches about 7 percent and the fog concentration is 5.5 x 10-5, Again, an extrapolation of the Factory Mutual data to 10 g shows that fog concentration of 2.1 x 10-6 or higher is required to inert 7.2 percent H 2 . In comparison, the present' theory on fog inerting pre- l dicts 1.02 x 10-4 for 7.2 percent H
- 2 k
The fog inerting criterion used is described in Section 5.2. 0430Q:1 6-15 w= , , - , ,
.v ,
~
1 h Therefore, it appears that it is passible to inert 7 psrcent H2 but l 1 unlikely. However, at 8 percent H2 in the upper plenum, which occurs at about 4650 seconds, the fog concentration is 5.5 x 10-5, which is too low to inert 8 percent H2. An extrapolation of the Factory Mutual 8 percent H2 data to 10 m volume mean drop size and the present pre- , ,, diction give 1.9 x 10-4 and 1.2 x 10-4 for the minimia required fog inerting concentration, respectively. Therefore both the theory and the extrapolation of test data show that fog 'inerting will not occur in the upper plenum. The glow plug igniters which have been installed in the Sequoyah con-tainment were designed to burn hydrogen lower than 8.0 pen:ent. As discussed previously, no fog inerting ef facts will be expected in the Sequoyah lower and upper compartments. Therefore, the glow plug igni-ters are expected to function as designed in these, two compartments. It may be possible that fog present in the ice condenser upper plenum may prevent the glow plug igniters from igniting hydrogen below 7 percent. However, it seems very unlikely that the same igniters would fail to ignite 8.0 percent H2 as designed, considering the fact that consider-g able conservatism has been exercised in the present analysis. Sensitivity studies of the spray removal efficiency and the fraction of blowdown droplets smaller than 20 u for the Sequoyah plant have been performed. A case of 10 percent spray removal efficiency was run using FOGMASS. The calculational results showed that the fog concentrations in the lower compartment, lower plenum, and upper compartment at 4600 l seconds were increased approximately by a factor of 10. However, these concentrations are still too low to inert 8 percent hydrogen. In com-parison, the fog concentration in the upper plenum is increased by only 20 percent because the concentration at this time is primarily deter-mined by the fog formation rate in the ice condenser. This increase is too small to change the conclusion given previously on the inerting probability in the upper plenum. Another case in which all the blowdown droplets were assumed to be smaller'than 20 u was run using FOGMASS. The calculational results showed that at 4600 seconds the fog concentra-tions in the lower compartment and lower plenum were increased by 15 k i ! 6-16 0430Q:1 [ , i - m . - _ _ - -
f percent while the increases in the upper plenum and upper compartment wirre negligibly small. The insensitivity of the fog concentrations to the parameter of the fraction of blowdown droplets smaller than 20 a is due to the effectiveness of the spray removal. At 4600 seconds, almost . .- ' all the blowdown droplets are removed by the sprays. The sensitivity studies showed that the fog concentration in the upper plenum at the time of significant hydrogen release is not sensitive to the spray removal efficiency and the fraction of blowdown droplets smaller than 20 a. O
~
l i I ( 6-17 ' 0430Q:1 . l
e , TABLE 6.1 FOG INPUT DATA FOR SEQUOYAH LOWER COMPARTENT I - Steam Lower Compartment Gas Wall Total Partial I
. . Gas Flow Rate' Temp. Temp-. Pressure Pressure
~
3 Time (sec) (ft /sec) (*F) (*F) (psfa) (psia) 60 1404.5 150 118 16.7 5 610 646.7 215 202 21.6 15.3 1210 3157.2 188 176 20.4 8.9 1810 3115.5 188 176 20.5 8.8 2410 2913.7 180 173 20.1 7.5 3010 2871.7 179 169 19.9 7.2 3510 2739.3 178 169 19.9 6.9 4010 2755.9 175 164 19.4 5.5 4510 2848.8 197 173 19.8 4.8 0 . I I I l l l ( 04300:1 6-18 l I
-e - - -- - - , -n -
, , TABLE 6.2 FOG INPUT DATA FOR SEQUOYAH ICE CONDENSER Steam
~
Ice Condenser Gas Ice- Total . Partial . Gas Flow Rate Temp. Temp. Pressure Pressure 3 Time (sec) (ft /sec) ('*F ) (*F) (psfa) (psia) 60 1082 120 32 16,6 2.5
- 610 -
96.4 132 32 21.8 2.3 1210 ' 2654 186 32 20.4 8.1 1810 2799 188 32 20.5 8.8 2410 2679 182 32 20.0 7.6 3010 2629 179 32 19.9 7.2 3510 2502 178 32 19.9 7.0 4010 2594 171 32 19.4 5.7 4510 2628 187 32 19.8 4.7 0 . 1 i 4 l i 6-19 . i 0430Q:1 e- -.c, -, - - . - a - * .,
1 l s TABLE 6.3 GE0ETRIC DATA FOR SEQUOYAH CONTAINENT , Yolume (ft ) Floor Area ( f t2) 1 Lower Compartment 289,000 . 5,410 . l
- i Ice Condenser Lower Plenum 24,200 3,100 Ice Condenser -
Upper Plenum - 47,000 3,200 Upper Compartment 651,000 10,390 Oead Ended Region 94,000 3,350 , O i e r ( 0430Q:1 6-20 l'
. 2
O' , o TABLE 6.4 MARCH PREDICTION OF REACTOR COOLANT MASS AND l ENERGY RELEASE RATE FOR THE S D 2 SEQUENCE Time H O Mass Release Rate H O Energy Release Rate *
* ~2 2 (seconds) (1bm/sec) (Btu /sec) ,
l 5 0.0 197.2 1.167 x 10
, 2172 190.5 1.097 x 10 5 ,
5.230 x 10 4 2478 44.85 3180 53.53 6.547 x 10 4 4 3804 34.82 4.262 x 10 4 4428 21.40 2.842 x 10 4 4752 48.42 5.558 x 10 4 5700 19.42 2.182 x 10 4 6012 14.07 1.583 x 10 3 6960 5.253 5.989 x 10 3 7062 4.718 5.388 x 10 7206 4.060 4.693 x 10 3 ,j 0 l
~
l l 1 I i 6-21 0430Q:1 p -
3 TABLE 6.5 INTERCOWARTENTAL FLOW RATES (ft /sec) 4 PREDICTED BY CLASIX FOR SEQUOYAH Time Flow From Flow From Flow From Flow From. Flow Fms , l 1 (sec) LC to LP LP to UP UP to UC UC to LC DE to LC 6.001El 1.175E3 1.082E3 7.029E2 -9.905El -1.304E2 6.100E2 3.580E2 9.641El -3.931El -2.113E1 -2.676E2 1.210E3 2.864E3 2.654E3 1.272E3 -1.838E2 -1.094E2 1.810E3 2.828E3' 2.799E3 1.323E3 -1.79X 2 -1.088E2
. 2.410E3 2.695E3 2.679E3 1.375E3 -1.502E2 -6.855El 3.010E3 2.654E3 2.629E3 1.407E3 -1.634E2 -6.326El 3.510E3 2.528E3 2.502E3 1.352E3 -1.643E2 -4.699El 4.010E3 2.613E3 2.594E3 1.537E3 -1.095E2 33.348E1 ,
4.510E3 2.694E3 2.628E3 1.627E3 -1.106E2 -4.426El I' 0 i g-.
/
i l i j f l 6-22 0430Q:1
, . . . , , , - . . - . . - . . , , . , . .-,_,s... s. ,e m -
,. - - . . s
0 U e !
. I i ~
6.3 F0G INERTING PROBABILITY IN THE McGUIRE PLANT
, ~ -
The computer codes, FOG and FOGMASS, were used to perform fog inerting j analysis for the McGuire plant. FOG was used to calculate the rates of , fog formation due to boundary layer and bulk st'reae condensation in the - l McGuire ice condenser and lower plenum. Thehthesefogformationrates were used in F0GMASS to compute the fog conc'entrations in each of the l McGuire containment subcompartments. . ! I To compute the fog formation rates in the ice condenser upper plenum and lower compartment,'some output data from the McGuire CLASIX analy-sis (28) are needed. These data include time histories of gas tempera-ture, wall temperature, total pressure, and steam partial pressure in each containment subcompartment, as well as the intercompartmental and , fan flow rates. In order to utilize the CLASIX output data, the ice condenser containment is subcompartmentalize,d in the F0GMASS program in exactly the same manner as in Reference 28. The subcompartmentaliza- . tion model used in the McGuire CLASIX analysis is shown in Figure 6.5. In this study only the S D2 accident scenario has been analyzed by
) CLASIX for McGuire.
The F0G input data for McGuire S2 D Case 1 are given in Tables 6.6 and 6.7, and the calculational results are shown in Figures 6.6 and 6.7. In~ , Figure 6.6, the fog formation rate in the lower c:mpartment is shown. For the first few hundred seconds the wall temperature is lower than the dew point corresponding to the steam partial pressure and therefore fog starts to form. The fog formation rate is low because the wall tempera-ture is only a few degrees below the dew point. Fog formation in the lower compartment becomes zero after about 600 seconds. The fog forma-tion rate in the ice condenser is shown in Figure 6.7. It is seen that the fog formation rate in the ice condenser is much larger than that in the lower compartment. The fog formation rate increases with the ice condenser steam flow rate and reaches the first peak at about 1510 sec-onds. Then the rate decreases because of the decrease in the steam flow rate. The fog formation and the steam flow rates start to increase again at about 2510 seconds. The fog formation rate reaches the second 4 0430Q:1 6-23 p _ . -
- __ ~ - ~ ' ' - - -
*. , I ,'
l FIGURE 6.5 MCGUIRE CLASIX CONTAINMENT MODEL (-
)
+
McGUIRE CLASIX MODEL
,i'
, ( .
I
;~ _ + - - --
, i l
' CONDENSER m UPPER i
UPPER COMPARTMENT ., j PLENUM i ! \4 g ' A A !, s e i l I i ICE BED 8 I l
]fs ARF !
Y a ICE 4- d DEAD - CONDENSER LOWER' ENDED LOWER 4 - 4 > GION PLENUM COMPARTMENT , O - wf 4 c . . AIR RETURN FAN / HYDROGEN SKIMMER SYSTEM FLOW PATH CONTAINS DOORS , 4' ej, 4 x M P FLOW ALLOWED IN BOTH OIRECTIONS 1 M FLOW ALLOWED IN ONE DIRECTION
\ \ ;
/
........... SPRAY HEADER 't .A
- e. ].
6 o s .i
, 6-24
^
\
! ~
_. . ., _ _ ,,.w. _r _ _ _ _ _m,,
0*000G i i- *P'" . x w . , z
- x. .
c_ . I ' a 0'0004 cc w 2 s c a w x 3 0'000E . . . u .. 1 2
- w I M i
a" .
- O E $
z 0*0002 m C 4 s a H s u
< W -
y g; m
.i o
w - y
- 1
- c- .
n. O'0001
- C 1
1 s 3, .,
- I 3
. s
- r / .
t O O O O o O O O O , O O O O O ) c- O gs a m t\J
. ". o e
o ., o o . o
4 (335/81) 31VB N 0.11V H 8 03 003 l t
6-25
. = = . . .
. o 1
i 0*0005 11 a j w e : z : w :
$ 0*0004 e
v - L.J d / w c: <
/
o o 0*000E v I N -
% w d 5
= =
8 z - Oc -
=
o / . 0*0002
< u I w m e .
o 6 "
\ =
8 N ' 0*0001
- 0*O l
3 e o o e o o O C C C O C D C C O C ' O. e o D C D C ' o s c m o
- N e
D e N Q (33S/81) 31Y B N O 11Y H 80.i 90.1 l
.(
6-26 1 t 1 1 L. - j
,7 m.
. s in lo rc pa nt a d n the ice o nse re i pu to FOGMSS in a tabular form. -
FOGmSS computes the rate of fog generation by the break flow, the fog ; settling rate due to gravity, and the fog removal rate due to sprays, as well as the rates of fog entrainment by intert:ompartmental 'and fan flows. The input data needed to calculate each of these rates art dis-cussed as follows. The rate of reactor coolant release to the containment and the coolant enthalpy were obtained from the mRCH output II for a small LOCA. The quality of the break flow was calculated using the enthalpy and the lower compartment gas temperature. According to the mRCH predic-tion III the discharge of If quid by the break flow into the lower com-partment lasts for only 2172 seconds. Afterward. the water level in the reactor vessel drops below the break elevation and the fluid discharged from the break is essentially steam. Therefort, in the present study, it is assumed that no fog is generated by the break flow after 2172
, seconds.
For fog removal by gravitational settling, a volume mean drop s'ize of 10 u was asstmed. The assumption of 10 u volume mean drop size is con- - servative, considering the fact that for a few thousand seconds the drop agglomeration mechanism would be able to increase volume mean drip size substantially. It should also be noted that a smaller volume mean drop size means that the minimum fog inerting concentration would be reduced and thus makes the present analysis conservative. Furtherwort, no con-sideration was given to the deposition of fog on the walls and vertical surfaces of the structure, or for fog rtmoval in the fan flows when it passes through ducts and fans. All the assumptions mentioned above make the present analysis very conservative. The containment geometric data needed in computing the settling rate art given in Table 6.8. l 6-27 0430Q:1 p _.,-wr- -- ~ -- w
b-For fog removal by sprays, a spray flow rate of 6800 gpa was used for McGuire. According to the McGuire CLASIX analysis (20I, the sprays are initiated at 124 seconds. A volume fraction of sprays (volume of sprays divided by volume of the spray zone) of 3.3 x 10-4 was used. As pre-viously discussed'a spray removal efficiency of a 100 percent effi'ciency
~
was used. In Figure 6.5, the directions of the intercompartmental flows are shown. The intercompartmental flow rates for the six flow paths and eight time steps were obtained from the OPS CLASIX analysis and are given in Table 6.9. The present analysis considers the intercompart-mental flows as the mechanisms of transporting fog from one compartment to another. It was assumed in the present analysis that 'the fog concen-trations in the intercompartment flows are the same as those in the compartments from which the flows are originated. ' Figure 6.5 shows two trains of the air return fan and hydrogen skismer system and the fan flow directions. The fans are initiated at 694 sec-onds. The fan head-flow curve reported in Reference 28 was used to , b compute the fan flow rates. Fan flow rates of 1000 ft /sec 3 and 100 f t /sec were used for the air return fan and the hydrogen skimmer system, respectively. These flow rates were calculated using average ap's between the upper compartment and the lower compartment, and . between the dead ended region and the upper compartment. It was also assumed that the fog concentrations in the fan flows are the same as 1 those in the compartments from which the flows are originated. 1 The results of the FOGMASS calculation are shown in Figure 6.8. It is seen that for the first few hundred seconds the fog concentrations in
, the lower compartment, ice condenser lower and upper plenums are about the same and increasing. At about 600 seconds, the lower compartment fog concentration reaches its peak of 1.6 x 10~4 Afterward, the intercompartmental flows transport more fog droplets out of the lower compartment than are generated by the break flow and condensation and, therefore, the lower compartment fog concentration decreases. However, the upper plenum fog concentration keeps rising until about 800 seconds, k b .
. .l l
0430Q:1 6-28 b
^ ~ ~
O : 0.0100 -- 0.0070 F0G CONCENTR ATION IN MCGUIRE CONTAINMENT 0.0050 ~
!= LOWER COMP. 2=1. C . LOWER PLENUM ~ 3=1. C . UP P E R
- 0.0030 PLENUM 4= UPPER COMP. 5=0E A0 EN0E0 REGION 5 0.0020 x
m 0.0010
- 0.0J07 r .
s 0.0005 yt,__ ,, D 0.0003 # \^ "" " ^ - o 0 0002 # A 4 'l ' l.00E-04 WL 4 '~ ' -
; ~ ~ '
- 7.00E-05 7' 'Fw '
s 5.00E-05 / \\ N b 3.00E-05 ' " ' z 2.00E-05 \k N
/
A 1.00E-05 7.00E-06 &" u
\\.
u A: ^ w
's m 5.00E-06 _,
\i, 'mN z 3.00E-06 / ' N-
"u "
w 2.00E-06 # % %-- ' u 1.00E-06 7.00c-
\ 's.! .' ' +d, 3
T -i f o 5.0050707 - o 3.00E-07
' 2.00E-07 1.00E-07 O O O O O O O O e . . . . .
. O O O O O O O O O O O O O S S -
2 2 8 8 TIllE (SEC) . FIGURE 6.8
l 1 due to an increasing fog formation in the ice condins:r and more fog , 1 entrained in the inten:ompartmental flow into the upper plenum. The l 7 upper plenum fog concentration reaches its peak of 6.4 x 10~4 at about 800 seconds. The lower plenum fog concentration is almost the same as the lower copartment fog concentration because of little difference in , I the intercompartmental flow rates into and out of the ice condenser
- lower plenum. Thertfore, these two volumes behave as a single volume in j tenas of fog concentration.
At 2172 seconds, the break flow in the lower compartment stops genera-ting fog and, therefore, the fog concentrations drop sharply there-after. The effect is more pronounced for the lower compartment and lower plenum fog concentrations. The highest fog concentration exists in the ice condenser upper plenum while the lowest exists in the upper compartment. The effect of sprays on the upper compartment fog concen-tration is clearly s,een in Figure 6.8. At 124 seconds, the sprays are > turned on and the upper compartment fog concentration drops sharply untti about 600 seconds. At about 600 seconds, the upper compartment fog concentration starts to increase again because the intercompart-mental flow into the compartment increases sharply at that time. A peak concentration of 7.5 x 10~ f n the upper compartment is reached at about 1500 seconds. Hydmgen starts to release into the containment at about 3804 seconds, - according to the MARCH calculation (28) . It reaches 4 volume percent at about 4300, 4400, and 4850 seconds in the lower compartment, upper plenum, and upper compartment, respectively. At 4300 seconds, the calculated lower compartment fog concentration is [ 8.4 x 10-7, which is about an order of magnitude smaller than the minimum fog concentrations required for inerting 4 percent 2H . At 4850 seconds, the upper compartment fog concentration is 1.47 x 10-6 j , ! which is about a factor of five smaller than the minimum fog concen- . tration required for inerting 4 percent H 2*. At the times of a
- The fog inerting criterion used is described in Section 5.2.
k ) 0430Q:1 6-30 l i
, , - - , _ , , - _ . . _ _ _ . . _ . . , - _ . , - . _ , , , , , . _ . , - _ , , , . -? . , ,
reaching 8.5 percent H2 , the fog concentrations in the lower and upper compartments are even lower than the figures given above. Therefore, it is concluded that the fog concentrations in the lower and upper compart-ments are too low to have any inerting effect. The use of the present theory on fog.inerting also leads to the same conclusion. . , However, at 4400 seconds, the calculated fog concentration in the upper plenum is 9.8 x 10-5 which is higher than the Factory Mutual fog inerting data extrapolated to 10 u drops and the present theoretical prediction. The data shows that order to inert 4.76 percent H2 the 6 fog ccTentration must be 8.4 x 10 or higher for.10 u voline mean drop size. At 4500 seconds, the up, 'r plenum hydrogen concentration reaches about 7 percent and the fog cs 1 centration is 9.3 x 10-I. Again, an extrapolation of the Factory Mutual 8ata to 10 u shows that fog concentration of 2.1 x 10-5 or higher is required to inert 7.2 percent H2 . In comparison, the present theory on fog inerting pre-dicts 1.02 x 10-4 for 7.2 percent 2H . Therefore, it appears that it is possible to inert 7 percent H2 , but unlikely. However, at 8 per-cent H in the upper plenum, which occurs at about 4600 seconds, the 2 0 fos co# centration is 9.1 x tr5, which is too iow to iaert 8 percent H. 2 An extrapolation of the Factory Mutual 8 percent H2 data to 10
- u volume mean drop size and the present prediction give 1.9 x 10-4 and 1.2 x 10-4 for the minimum required fog inerting concentration, ~
respectively. Therefore, both the theory and the extrapolation of the test data indicate that fog inerting will not occur. The glow plug igniters which have been installed in the McGuire contain-ment were designed to burn hydrogen lower than 8.5 percent. As discus-sed previously, no fog inerting effects will be expected in the McGuire lower and uoper compartments. Therefore, the glow plug igniters are , expected to function as designed in these two compartments. It may be possible that fog present in the ice condenser upper plenum may prevent j the glow plug igniters from igniting ttydrogen below 7 percent. However, it seems very unlikely that the same igniters would fail to ignite 8.5 d . *The fog inerting criterion used is described Section 5.2. i k 0430Q: 1 6-31
- \
percent H2 as designed, considering the fact that considerable conser-
-I 'vatism has been exercised in the present analysis.
O i i 4 I ( 6-32 0430Q:1
as TABLE 6.6 FOG INPUT DATA FOR McGUIRE LOWER COMPARTENT Steam Lower Compartment Gas Wall Total Partial . Gas Flow Rate Temp. Temp. P ressure Pressure
~
Time (sec) (ft 3 73,c) g.F) ( *F) (psia) (psia) t 60 1624.6 1 60 149 16.5 7 510 1248.1 225 215 22.2 18.3 1510 2387.8 205 198 21.9 12.6 l r 2010 2393.8 205 198 22 12.4 2510 1940.7 195 193 21.5 10.4 3260 2055.5 200 195 21.6 10.8 3760 - 1801.7 200 194 21 9.3 4510 1919.3 2 50 222 21.2 7.3 l 0 . l l l ! I l 1 i , i a 1 1 i k I 6-33 0430Q:1 . l e
-' d m
, TABLE 6.7 FOG INPUT DATA FOR McGUIRE ICE CONDENSER Stena Ice Condenser Gas Ice Total Partial
. Gas Flow Rate Temp. , . Temp. , . Pressure Pmssuru ,.
Time (see) (ft 3/sec) (*F) (*F) (psia) (psia) 60 820.5 90 32 16.5 1 510 107.1 130 32 22.2 2.3 1510 1926 190 32 21.9 9.3 2010 1637 193 32 22 9 2510 1145 188 32 21.4 8.6 3260 1630 195 32 21.6 10.3 3760 1514 193 32 21.1 8.1 4510 1464 192 32 22.1 7.1 0 1 i l
- ( <!
;I' l 6-34 /
l 0430Q:1 embiS
- . _ , . - . , - , . - . _ , . - - - - _ . . . - . ,,. -g -. -- .---. . - - _ - . . _ . - .
I TABLE 6.8 GE0lETRIC OATA FOR McGUIRE CONTAINfENT 3 Yolume ( f t ) Floor Area (ft2)
. Lower Compartment 237,400 .
5,410 -
, 1 Ice Condenser [^ -
Lower Plenum 24,200 3,100 i Ice Condense'r Upper Plenum 47,000 3,200 Upper Compartment 670,000 10,390 Dead Ended Region 130,900 3,350 0 . I 6-35 0430Q:1
, * *- . - . - - <- -,.,.-.. -- -- 8D- --- e -, - - - - - -
S --e ' - -v -
y
. .y
' l l
3 TABLE. 6.9 INTERCOMPARTENTAL FLOW !!ATES (f t 73,g) ) PREDICTED BY CLASIX FOR McGUIRE Time. Flow From Flow From Flow From Flow From F, low From (sec) LC to LP LP to UP UP to UC UC to LC DE to LC
- l 6.001El 1.351E3 8.20SE2 5.783E2 -1.198E2 -1.538E2
- 5.100E2 8.716E2 1.071E2 -2.269El -2.863E1 -3.479E2 1.510E3 2.008E3 1.926E3 8.635E2 -1.900E2 -1.898E2 2.010E3 2.010E3 1.637E3 6.869E2 -2.266E2 -1.572E2 2.510E3 1.722E3 1.145E3 4.807E2 -1.410E2 -7.767El 3.260E3 1.713E3 1.630E3 6.666E2 -2.087E2 -1.338E2 3.760E3 1.546E3 1.514E3 7.231E2 -1.289E2 -1.268E2 4.510E3 1.634E3 1.464E3 7.640E2 -1.328E2 -1.515E2 0
i i i l 1 l l ? l l i .i I
- ( ,
6-36 04300:1
a . l 4 6.4 FOG IERTING PROBABILITY IN THE D. C. COOK PLANT I- The computer codes, FOG and FOGMASS, were used to perform fog inerting analysis for the D. C. Cook plant. FOG was used to calculate the rates "* of fog foma' tion due to boundary layer and bulk stream condensation in ~ the D. C. Cook ice condenser and lower plenum. Then these fog formation l rates were used in F0GMASS to compute the fog concentrations in each of the D. C. Cook containment subcompartments.
- l
. . To compute the. fog fomation rates in the ice condenser upper plenum and lower compartment, some output data from the Cook CLASIX analysis (29) } are needed. These data include time histories of gas temperature, wall temperature, total pressure, and steam partial pressure in each contain-ment subcompartment, as well as the intercompartmental and fan flow l rates. In order to utilize the CLASIX output data, the ice condenser i containment is subcompartmentalized in the FOGMASS program in exactly the same manner as in Reference 29. The subcompartmentalization model t used in the Cook CLASIX analysis is shown in Figure 6.9. In this study j only the S 20 accident scenario has been analyzed. O l The FOG input data for Cook SZ D Case 1 are given in Tables 6.10 and { 6.11, and the calculational results are shown in Figures 6.10 and 6.11. l In Figure 6.10, the fog fomation rate in the lower compartment is l show n. It is seen that the fog fomation rate is negligibly small. It i should be noted that the calculation of the lower compartment fog ! concentration in the D. C. Cook plant starts at 600 seconds instead of ! 60 seconds used for the other two plants. The fog fomation rate in the l lower compartmetqt starts to increase at about 4200 seconds because of the increase in the steam partial pressure. It reaches 0.017 lb/sec at t f i
- about 4590 seconds. Fog famation in the lower compartment will stop j af ter 4700 seconds because of the hydrogen burn thereafter. The fog fomation rate in the ice condenser is shown in Figure 6.11. It is seen i I
? that the fog fomation rate in the ice condenser is much larger than j that in the lower compartment. It increases with the ice condenser i steam flow rate and reaches a peak of about 15.6 lb/sec at about 1200 seconds. The fog fomation rate in the ice condenser then begins to , decrease and is low at the time of significant hydrogen release. , 0430Q: 1 6-37 l
- - FIGURE 6.9 D.C. COOK CLASIX MODEL i
rrrrra
. ICE .
COND.ENSER 4 + UPPER m,
i COMPARTMENT l LE UM -
I A i A i l ICE BED
- 1r 1r e
ICE CONDENSER LOWER Y. LO COMPARTMENT pgENUM l
's L
JL ;L l 1 0 ' i t' 1
'r - 'f l t
1 . DEAD ! FAN / ENDED ACCUMULATOR REGION l ROOMS : s. I n A FLOW PATH CONTAINS OCORS 4 : FLOW ALLOWED IN BOTH OIRECTIONS 4 FLOW ALLOWED IN ONE DIRECTION g 39rrrrrrrrr rrra SPRAY HEADER 6-38
, - - - - . . - ~ - . . . . , - . . . .
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, _ - - - - - - - % , vw ,- _ _, +y e. , , - - - - - - , - ym , + = - - . --
The eight fog formation rates in the lower compartment and in the ice ( . condenser are input to F0GMASS in a tabu'lar form. FOGMASS computes the rate of fog generation by the break flow, the fog
'/ settling rate due to gravity, and the fog removal rate due to sprays, as
- well as the rates of fog entrainment by ingercompartmental and fan flows. The input data needed to calculate'each of these rates are dis-cussed as follows. j ,
j The rate of reactor coolant release to the containment and the coolant enthalpy were obtained from the MARCH output II for a small LOCA. The quality of the break flow was calculated using the enthalpy and the l lower compartment gas temperature. According to the MARCH predic-tion I7I the discharge of liquid by the break flow into the lower com-partment lasts for only 2172 seconds. Afterward, the water level in the reactor vessel drops below the break elevation and the fluid discharged from the break is essentially steam. Therefore, in the present study, it is assumed that no fog is generated by the break flow after 2172 seconds. O ' For fog removal by gravitational settling, a volume mean drop size.of 10 a was assumed. The assumption of 10 g volume mean drop size is con- - servative, considering the fact that for a few thousand seconds the drop - agglomeration mechanism would be able to increase volume mean drop size substantially. It should also be noted that a smaller volume mean drop size means that the minimum fog inerting concentration would be reduced i and thus makes the present a'nalysis even more conservative. Further- I more, no consideration was given to the deoosition of fog on the walls l
! and vertical surfaces of the structure, or for fog removal in the fan flows when it passes through ducts and fans. All the assumptions men- 1 tiened above make the present analysis very conservative. The contain-e ment geometric data needed in computing the settling rate are given in 4
Table 6.12. I . ) i i 6-41 0430Q:1 I I . 7 ., _ . , _ . _ _ . , , , _ _ , . _ _ , , - _ _ _ , , , . . . . _ , _ - , , - . _ _ . . . . . . E.w m,. "
,_- .,_.. -,,_.. _ , - _ y . . . , . , - - - . , , _ . , _ . . _ _ , . _ _ , _ _ , . _ . _ . , , _
~ .
- l. .
For fog removal by sprays, spray flow rates of 4000,1800, and 528 gpa ( , were used for the upper compartment, lower compartment, and - fan / accumulator rooms, respectively. According to the Cook CLASIX analysis (29) the sprays are initiated at 141 seconds. A volume l , j fraction of sprays (volume of sprays divided by volume of the spray zone) of 3.3 x 10-4 was used. As previously discussed a spray removal efficiency of a 100 percent efficiency was used. l In Figure 6.9, the directions of the intercompartmental flows are ; ! - shown. The tritercompartmental flow rates for the six flow paths and j eight time steps 'were obtained from the OPS CLASIX analysis and are 4 given in Table 6.13. The present analysis considers the intercompart-mental flows as the mechanisms of transporting fog from one compartment l to another. It was assumed in the present analysis that the fog concen ' trations in the intercompartment flows are the same as those in the lI compartments from which the flows are originated. i Figure 6.9 shows two trains of the air return fan and hydrogen skimer ! system and the fan flow directions. The fans are initiated at 711 sec- !O j onds. The fan head-flow curve reported in Reference 29 was used to compute the fan flow rates. Fan flow rate,s of 1388, 61.76, and 4.13 , l ft 3/sec were used for the flows from the upper compartment, lower ! compartment, and dead ended region to the fan / accumulator rooms, . l respectively. These flow rates were calculated using the ap?s between i the the fan / accumulator rooms and three other compartments. It was also assumed that the fog concentrations in the fan flows are the same as - those in the compartments from which the flows are originated. The re: ult: ef the F0 MASS calculat49 are shown in Figure 6.12. It is seen that for the first few hundred seconds the fog concentrations in the lower' compartment, and the ice condenser lower plenum are high. At i about 140 seconds, the lower compartment fog concentration reaches its i peak of 1 x 10-4 Af ter the sprays are initiated at 141 seconds, the fog concentrations in the lower compartment, upper compartment, and l fan / accumulator rooms drop sharply. However, the upper plenum fog [ concentration keeps rising untti about 1200 seconds, due to an increasing I 0430Q:1 6-42 l 4
- ene sew,-- . . - - . . ..,,-.,,e ,.e----e.v.,, w-,,--e ,--a.. .m.,--,------,----.n. - , . , - . - e e,+-.- e - . , . . ,w- . . , --,s--,,, < ~----,------,-w, , ,
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a fog formation in the ice condenser and more fog entrained in the intercompartmental flow into the upper plenum. The upper plenum fog concentration reaches its peak of 2.4 x 10-4 at about 1200 seconds. After about'1200 seconds, the lower plenum fcg concentration is almost . .
~
the same as the lower copartment fog concentration since the , intercompartmental flows quickly make the fog concentrations in these two compartments uniform. Therefore, these two volumes behave as a single volume in terms of fog concentration. At 2172 seconds,' the break flow in the lower compartment stops genera-ting fog and, therefore, the fog concentrations drop sharply there-after. The effect is more pronounced for the lower compartment and lower plenum fog concentrations. The highest fog concentration exists in the ice condenser upper plenum. The effect of sprays on the upper compartment fog concentration is clearly seen in Figure 6.12. At 141 seconds, the sprays are turned on and the upper compartment fog concentration drops sharply untti about 300 seconds. At about 300 seconds, the upper compartment fog concentration starts to increase again because the intercompartmental flow into the compartment increases sharply at that time. A peak concentration of 9.5 x 10-6 in the upper compartment is reached at about 1400 seconds. Hydrogen starts to release into the centainment at about 3804 seconds, according to the MARCH calculation (2N. It reaches 4 volume percent 1 at about 4350, 4370, and 4700 seconds in the lower compartneent, upper plenum, and upper compartment, respectively. . At 4350 seconds, the calculated lower compartment fog concentration is 10~7, which is about two orders of magnitude smaller than the minimum fog concentrations required for inerting 4 percent H2 . At 4700 seconds, the upper compartment fog concentration is 2.4 x 10-6, which is about a factor of two smaller than the minimum fog concentration required for inerting 4 percent H2 *. At the times of reaching 8.5 g
- The fog inerting criterion used is described in Section 5.2.
04300:1 6-44
percent H2 , the fog concentrations in the lower and upper compartner.ts are even lower than the figures given above. Therefore, it is concluded
- that the fog concentrations in the lower and upper compartments are too low to have any inerting ef.fect. iThe use of.the present theory on fog , . .
inerting also" leads to the same conclusion. , However, at 4370 seconds, the calculated fog concentration in the upper plenum is 6.5 x 10-5 which is higher than the Factory Mutual fog l inerting data extrapolated to 10 g drops and the present theoretical prediction. The data shows that,i order _to. i. nert 4.76. percent H2 the , fog concentration must be 8.4_x 10- orhigherfor10pvolumemean drop size. At 4530 seconds, the upper plenum hydrogen concentration reaches about 7 percent and the fog concentration is 5.5 x 10-5, Again, an extrapolatf ortof the_Eac_ tory _t!utual data to 10 g shows that i fog concentration o(2.1 x 10-5 or higher is. required to. inert 7.2 3 .
- per g H .g In comparison the present theory of fog inerting predicts 1 02 x 10_-4_forl.2_ percent H2 Therefore, it appears that it is 4 possible to inert 7 percent H2 , but unlikely. However, at 8 percent H2 in the upper plenum, which occurs at about 4600 seconds, the fog
- concentration is 5.1 x 10-5 wMch is too low to inert _8 percent I H. An extrapolation of the Factory Mutual 8' percent H 2 data to 10 2
u volume. mean drop size and the present prediction give 1.9 xand 10-4 , 1 1.2 x 10"4Io'r the minimum required fog inerting concentration, r espectively. i The glow plug igniters which have been installed in the Cook containment ! were designed to burn hydrogen lower than 8 percent. As discussed pre-viously, no fog inerting effects will be expected in the Cook lower and f upper coger i.ments. Therefore, thG gicw pl::g ig::iterc are c:pected to function as designed in these two compartments. It may be possible that 3 fog present in the ice condenser upper plenum may prevent the glow plug igniters from igniting hydrogen below 7 percent. However, it seems very unlikely that the same igniters would fail to ignite 8 percent H2 as l designed, considering the fact that considerable conservatism has been
- exercised in the present analysis.
I 1 0430Q:1 6-45 ( 1 w w.y&-- +w,-h,.-pe- - q- -,-+,wir+, -w+-- m--- -v-a--.--- -* -
i . TA8LE 6.10 FOG INPUT DATA FOR D. C. COOK LOWER COPFARTENT
* ~
Steam . Lower Compartment Gas Wall Total Partial
. . Gas Flow Rate' Temp. Temp. Pressure Pressure '
Time (sec) (ft 3f,,e) g .7) g.7) g,,,,) g,,,,) 600 799.4 222 215.2 21.8 17.4 1200 2798.2 190 183.5 20.2 9.4 1800 2805.8 190 180.3 20 9.1
'2400 '2513.6 180 177.2 19.6 7.6 3000 2448.5 178 170.4 19.3 7.2 3600 2359.7 175 169.3 19.2 6.4 4200 2272.3 165 16'1.9 18.8 5.3 4590 2482.7 168 161 19,5 5.8 a
4 0 4 4 L l' ( /[ ! I 0430Q:1 6-46
~
~,
.m- - . . - , -- -- v-._. -,..n.,-,,-- - . . - , - - - - , . . . . . , , . . , , , . , .,,-,.s... ,,,,,,._,4 ,.--,-.,,n ..,y a,.,,, .,n , . -.n.,_ . . , , , , . ,.. , , . - - .
b D TABLE 6.11 FOG INPUT DATA FOR D. C. COOK ICE CONDENSER I - Steas Ice Condenser Gas Ice Total Partial
~
Gas Flow Rate ' Temp. Temp. Pressure Pressure - Time (sec) (ft 3/sec) (*F) (*f) (psia) (psfa) . 600 76 147 32 21.8 . 3.4 e 1200 2548 190 32 2C .1 9.3 1800 2572 188 32 19.9 9.0 2400 2359 184 32 19.7 7.9 3000 2256 187 32 19.3 7.1 3600 2199 175 32 19.2 6.6 4200 2126 166 32 18.8 5.3 4590 2312 163 32 19.8 4.3 O l i ( . I i 0430Q:1 6-47 l u_.
TABLE 6.12 GE0ETRIC DATA FOR 0. C. COOK CONTAINENT Volume (ft3 ) Floor Area (ft2) 249,681 5,410 Lower Compartment
!ce Condenser Lower Plenum 24,700 3,100 Ice Condenser 3,200 Upper Plenum 47,010 Upper Compartment 681,283 10,390 Dead Ended Region 61,105 853 1
Fan / Accumulator Rooms 54,828 2,500 0 . ( ( ! l 0430Q:1 6-48 \
= - - , _ - - - , , , _ . - - - _ - - . - , , . . , . . - - - - - - - . - . ,-ye,_ .-.+._,-,,.y
f . 3 TABLE 6.13 INTERCOMPARTENTAL FLOW RATES (ft /sec) PREDICTED BY CLASIX FOR D. C. COOK Time Flow From Flow From Flow From Flow From Flow From Flow Fros ' (sec) LC to LP LP to UP UP to UC UC to LC DE to LC NA to LC 600 6.387E2 7.60dE1 -4.410E1 -3.74SEi -1.232E2 -1.229E2 1200 2.577E3 2.548E3 1.106E3 -1.740E2 -4.720E1 1.509E3 1800 2.600E3 2.572E3 1.155E3 -1.620E2 -4.381El 1.529E3 2400 2.356E3 . 2.359E3 1.145E3 -1.325E2 -2.512E1 1.595E3 3000 2.273E3 2.256E3 1.178E3 -1.463E2 -2.923E1 1.55X3 3600 2.202E3 2.199E3 1.190E3 -1.334E2 -2.333E1 1.603E3 4200 2.136E3 2.126E3 1.258E3 -1.183E2 -1.802E1 1.642E3 4590 2.346E3 2.312E3 1.400'E3 -1.130E2 -2.371El 1.650E3 0 i l l s l
\
0430Q:1 6-49
* .- -.-,.-,.-..,,--.r -
m .
. 1 6.5 EFFECT OF FOG ON GLOBAL COMBUSTION s f , ,
E In order to assess the effect of fog on the deflagration Itait of hydro- ] gen, which is defined a,s the minimum hydrogen concentration at which the ' 9 flame propagates in all directio'ns, a flame temperature crite'rion which considers fog droplets as a heat sink was used. This criterion assumes that the critical flame temperature of 710*C is still appitcable to a S
,[.
hydrogen mixture which contains fog droplets. For a given fog concen-tration, the heat required to heat a unit mass of the mixture to 710*c y can be calculated. Then the hydrogen concentration needed to supply k this amount of heat, assuming 100 percent combustion, can be deter-n mined. Using this method, the calculated fog concentrations of 5.5 x I 10-5 and 5.1 x 10-5 for the Sequoyah plant at 4650 seconds and for the D. C. Cook Plant at 4600 se.conds, respectively, were found to be C I" capable. of raising the deflagration limit to 10.6 vol. percent H2* comparison, the calculated fog concentration of 9.1 x 10-5for the g McGuire plant at 4600 seconds was found to be capable of raising the deflagration limit to 12 vol. percent H2 . This study shows that in { order to achieve global combustion in the upper plenum, hydrogen concen-O tration higher than 8.5 percent may be required. The effect of increas- ( ing hydrogen concentration required to obtain global combustion in the y. upper plenum should be investigated, h b i i ( , 04300:1 6-50 e em M
----,-,v.
~
7.0 SUMARY AND' CONCLUSIONS ( . The present study has developed a systematic methodology to study the potential fog inerting problem for the PWR ice condenser plants. In the r presint investigation, ma3or fog formation'and' removal mecha'nisest a'e identified and quantified. Theoretical models are developed to predict ' . the fog formation rate due to boundary layer and bulk stream condensa-tion, the fog removal rates due to gravitational settling and contain-i
- ment sprays. The mass conservation equations for the fog droplets in each of the containment subcompertments are solved simultaneously in
~
order to obtain time histories of fog concentration. These equations incorporate fog formation due to condensation, fog generation due to ! break flow, fog removal due to gravitational settling and sprays, trans-I port of f'og by the intercompartmental flows and fan flows. Computer programs FOG and FOGMASS have been developed to ecmpute fog formation rates and fog concentrations in each of the containment subcompart- - ments. These two computer programs have been used to analyze a S ZO accident sequence for the Sequoyah, McGuire, and D.C. Cook plants. The g analyses espicyed output data from the Sequoyah CLASIX analyses. Speci-fically, time histories of gas temperature, wall temperature, total f pressure, and steam partial pressure in each containment subcompartment, as well as the intercompartmental and fan flow rates were used in the present analysis. l A fog inerting criterion has been developed to predict the minimum fog concentration required to inert a given hydrogen concentration and volume mean fog drop size. The present fog inerting criterion has been shown to be in agreement with the Factory Mutual data. The criterion i shows that the minimum fog inerting concentration varies with the square < of the volume mean fog drop size. ! I The present study shows that the fog concentrations in the upper and j j lower compartments of the three plants analyzed are too low to have any 1 inerting effect on hydrogen mixtures. Therefore, the proposed glow plug l igniters are expected to function as designed in these two compart-i ments. It may be possible thr.t fog present in the ice condenser upper ! I , l t 7-1 I
- 0430Q:1 1
l
l l t I plenum may prevent the glow plug igniters from igniting hydr 6 gen below 7 percent. However, it seems very' unlikely t. hat the same igniters would fati to ignite 8.5 percent H2 as designed. It should be recogritzed that the existing theories and data can only . predict the minimum fog concentration for inerting.' Further work may be required to verify the fog inerting theory associated with flame propagation in all directions. O O r ( /f
.;I 7-2 043)Q:1
~ g ,
- e. ,___ _-
9 ) . AC NOWLEDGMENTS 4 I , The author wishes to express his sincere gratitude to Mr. N.J. Liparulo, Ors. V. Srinivas, B. Lewis, and 8. Karlovitz for assistance, sugges-tions, and helpful discusifons, 'particular.ly in the area of the fog , I
- inerting, criteria and the flame temperaturp criteria for fog, to ,
i Messrs. D. F. Paddleford, R. Bryan, F. G. Hudson, and K. Shiu for valdabit consents, to Mr. K. C. Perry, Mr.' S. J. Reiser, and s Ms. R. M. Mariner for providing data on the three ice condenser plants, and to Mr. T. 'J..Miele for providing programming assistance. l [ He also muld like to thank TVA, Duke Power, and AEP for providing the financial support. i I ( j l
/
O -
** l
}
f I 0430Q:1 7-3 w _. -
e ) REFERENCES
- I l
l 1. B. Lowry, "Preifainary Results: A Study of Hydrogen Igniters," l ENN80-45, Lawrence Livermore National Laboratory, November 17,198L'.
- 2. " Additional Questions on Hydrogen Control System for Ice Condenser Plants," NRC memo from L. Rubenstein to R. Tedesco, dated June 26,
,- 1981.
- 3. "The Marvikken Full Scale Containment Experiments," MXB-301 AB Atomenergi, March,1977.
~
- 4. T. F. Kanzleiter, "LOCA Experiments With a PWR Multi-Compartment Model Containment," Trans.1977 LWR Safety ' Conf., Idaho Falls, Idaho, 1977.
- 5. G. M. Fuls, "The CLASIX Computer Program for the Hydrogen Release and Degradation", OPS-07A35, Offshore Power Systems,1981.
0 6. K. K. Almenas, "The Physical State of Post-Loss-of-Coolant Accident ' Contaf fnent Atmospheres," Vol. 44, Nuclear Technology, pp. 411-427, August, 1979.
- 7. "Suomary of Analysis of Ice Condenser Containment Response to Hydro-gen Transients," Offshore Power Systems report No. RP-28A52, Septem-ber, 1980.
- 8. R. Brown and J. L. York, " Sprays Formed by Flashing Liquid Jets,"
Yol. 8, No. 2, A!Ch.E Journal, p.149, May,1962.
- 9. R. G. Gido, and A. Koestel, "LOCA-Generated Drop Size Prediction - A -
Thermal Framentation Model," Trans. Am. Nucl. Soc., 30, p. 371,1978.
- 10. P. G. Hill, H. Witting, and E. P. Demetri, " Condensation of Metal l Yapors During Rapid Expansion," Journal of Heat Transfer, p. 303, i November, 1963. !
I i R-1 0430Q:1 4 h-- - e.
= - . . - . . . _ .
J
- 11. M. Yolmer and H. Flood, Z. Physik Chemie, A170, p. 273,1934.
I -
- 12. C. E. Junge, Advan. Geophys., H. Landsberg and J. Van Mieghem, ed.,
4.1, Academic Press, New York,1958. ,
- 13. R. J. Burian, and P. Cybulskis, " CORRAL II User Manual," Battelle Columbus Laboratories, January,1977. .
- 14. R. K. Hilliard and L. F. Coleman, " Natural Transport Effects on ,
i Fission Product Behavior in the Containment Systems Experiment," BNWL-7457. Battelle-Northwest, Richland, Washington,1970.
- 15. N. H. Fletcher, J. Chem. Phys., 29, p. 572; 31, p.1136,1958.
- 16. D. E. Rosner and M. Epstein, " Fog Formation Conditions Near Cold .
Surfaces," Yol. 28, No.1 J. of Colloid and Interface Sci., Septem-ber, 1968. .
- 17. K. Hijikata, and Y. Mori, " Forced Convective Heat Transfer of a Gas Q With Condensing Vapor Around a Flat Plate," Vol. 2, No.1, Heat 3
Transfer - Jap. Res., pp. 81-101, January,1973. ,
- 18. M. Nefburger and C. W. Chien, " Computation of the Growth of Cloud Drops by Condensation Using an Electronic Digital Computer,"
Geophys. Monograph No. 5, pp.191-209,1960.
- 19. R. M. Kemper, "!odine Removal by Spray in the Salem Station Contain-ment," WCAP-7952, Westinghouse Electric Corp. , August,1972. .
I i 20. N. J. Liparulo, J. E. 01hoef t and D. F. Paddleford, " Glow Plug Ignitor Tests in H2 Mixtures," WCAP-5909, Westinghouse Electric Corp., March 6, 1981.
- 21. R. G. Zalosh and S. N. Bajpai, " Water Fog Inerting of Hydrogen - Air Mixtures," EPRI Project Preliminary Rp.1932-1, September,1981.
k R-2 04300:1 l
- 22. J. M. Marche 11o, "Controk of Air Pollution Source " Marcel Dekker, e ,
Inc. , New York,1976.
- l
- 23. Letter from B. Lewis and 8. Karlovitz, to L. E. Hochrtiter, dated may 5, 1980. .
- 24. M. Beman, et al., " Analysis of Hydrogen Mitigation for Degraded Core Accidents in the Sequoyah Nuclear Power Plant," Sandia draft report, Decen6er 1,1980.
- 25. T. von Kannan, Unpubitshed notes,1956.
- 26. S. S. Tsai, " Flame Temperature Criteria . Tests," NS-CCA-81-039, West-inghouse internal memo, dated June 17, 1981.
- 27. Attachment to Offshore Power System letter PST-NE-109, dated May 22, 1981.
- 28. Attachment to Offshore Power System letter PST-NE-106, dated May 14, O 1981.
- 29. Attachment to Offshore Power System letter PST-NE-218, dated August 6, 1981. .
- 30. M. L. Corrin, J. R. Connel, and A. J. Gero, "An Assessment of Wam Fog - Nucleation, Control, and Recommended Research," NASACR-2477, November, 1974.
I 0430Q: 1 R-3 . e- ,, - y - 2-_-,-, , ' , --
( APPENDIX A COMPUTATION OF Y,AND C IN EQUATION (3.12) The Hijikata rt fog formation theory IIII used the boundary layer approximation for the continuity, momentum, and energy equations. The fog concentration and velocity profiles in the boundary layer are assumed in Eqs. (3.7) and (3.8). Substituting Eqs. (3.7) through (3.10) into the conservation equations, we have
- d + d ( + h Y, + v' =0 (A-1)
- (1 + Yf)
+dC+hY +h(Y,+v' =0 (A-2)
- y + y g + M Y, + v' =-f(1+g)y (A-3)
A (n) + B (n) C + C(n)+hE Y, O -lid-v-v(t+c)'o! o o (^-4) where A (n) = 4 (n + ) n+ ) (n + 4) 3 I" # 0 B (n) = 4 (n + Z) (n + 3) (n + 4) 3 I"
- II C (n) = 4 (n + 3) (n + 5) (n + 6)
AW I
*= " 1 - W, (
2 aW l y , 0 (1 - W,)S, l h P9 04304:1 A-1
*- l an- . . . . . _ _ .-- .... , .
2 i , a U, R = 4 v. -
" a u,
= .
W. = weight fraction of vapor at free stream W, = weight fraction of vapor at wall
=
4W W; , W, S, = Schmidt number
~
v = kinetic viscosity v, = component of the free stream velocity perpendicular to the wall ' h = heat of vaporization fg C = specific heat of non-condensible gas pg
- AT =
T , - T, T. = gas temperature at free stream T, = gas temperature at wall . Equations (A-1) through (A-4) are four algebraic equations for four unknowns, Y,, C, R, and v' These equations have been solved by the computer program FOG. In FOG, the values of Y,, C, and R are , computed and used in Eq. (3.12) to compute the fog formation rate. ; i l l'
/
/
i 9 A-2 0430Q:1
- - .._ l - . . . . . .. . - --l
. . . _ _. _ . _ . . . . _ _ _ _ . . . . . = - - - -
APPENDIX B ( ' 1 DERIVATION OF EQUATION (5.$) l l This appendix gives detailed procedures to derive Eq. (5.5), starting, ' from Eq.' (5.4') f (K) crit 't = f ( (Y, - Y )/eg) f - - (5.4) where the ratio of heat loss rate per urif t voline to the heat release rate by chemical reaction per unit volume, (K) crit, is defined as ( 8-1) Kerit = S/C pw and the ratio of sensible heat to heat of combustion, og , is defined as o g=Cp (Tg - Tu )/q (B-2) To arrive at Eq. (5.5), it is necesary to assme that all the heat loss
<= ettributed to coavectioa heat treasfer to fas dropi ets or oair oae
- O drop size. Under this assumption, the rate of heat loss per unit voltane per degree, 5, my be expressed as .
2 S = ned h where n = number of drops per unit volume d = volume mean drop size h = heat transfer coefficient It is further assimed that the relative velocity between the droplets and the mixture flow is so small that heat transfer coefficient, h, can be approximated by the conduction limit. Under this assumption, Eq. (B-3) reduces to S = ~12ni (B-4) T ( 0430Q:1 8-1
- - - , - - j g e---- DJ e--- ------
- 4. a7 3 4
, J .
TENNESSE2 'V ALLEY ni, THOR s's j CH t- A:.0 0 3.4 T CN *. EG? EI *. 01 j 400 Chestnut Street Tower II ( June 2, 1981 .-<h ! b.? pp F* X C.e " R "'/..
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\
Director of Nuclear Reactor Regulation 5s - s.' e Attention: Ms. E. Adensam, Chief / Licensing Branch No. 4 \?V',-,,[
[* t Division of Licensing u*
U.S. Nuclear Regulatory Co= mission Washington, DC 20555
Dear Ms. Adensam:
In the Matter of the Application of ) Docket Nos. 50-327 Tennessee Valley Authority ,
) 50-328 As required by the Sequoyah Nuclear Plant unit 1 operating license (DPR-77) item 2.C(22)D(3)(b) and the draft Sequoyah unit 2 operating lice'nse, we ara enclosing 20 copies of a report to address the hydrogen
', environment items in the Sequoyah Nuclear Plant Safety Evaluation Report
)
i and the hydrogen control censures research program. ,) The enclosed report is a nonproprietary version. We will submit the proprietary pages under separate cover. If you have any questions, please get in touch with D. L. Lambert at FTS 857-2581. Very truly yours, TENNESSEE VALLEY AUTHORITY v} \p,s L. M. Mills, Manager l Nuclear Regulation and Safety Sworn to pd su ibed before me b thi /'4.
- day of., 6+ d. '1981 g/
1 l h .c % b h n. h e.u ; L u . s l Notary /Public , l My 00::1ssien Expires 5L ~ 1 / / l In lecare '20)
, e am a f
. . i f* j E2. [ 0 0.' * ' ' t., "* O d . t!
J. s w
. s Sandia National Laboratories dete: January 11, 1982 Albuquerque. New Mexico 87185 to:' J. C. Cummings, 4441 dW7. & N Y1p4 l
#wn: C. M. Stone, 5521, P. S. Vntrs, 5523 l meka: Structural Analysis Results of the Sequoyah Nuclear Power Plant's Steel Containment Subjected to a Hydrogen Burn Scenario Ref: 1. Key, S. W., Beisinger, Z. E., and Krieg, R. D.,
"HONDO II, A Finite Element Computer Program for the Large Deformation Dynamic Response of Axisymmetric Solids," Sandia National Laboratories, SAND 78-0422, Albuquerque, NM, Oct. 78.
An analysis of the structural response of the Sequoyah nuclear power plant primary containment to a hydrogen burn has been conducted. The analysis was performed using the HONDO II [1] finite element code with the pressure-time information supplied by R. K. Byers, 4444. The investigation studied the response of an axisymmetric steel containment to a triangular impulse traveling along the containment dome. The duration of the impulse at a point was assumed to be 6 ms with an exponentially increasing peak pressure from 58 psi (.4 MPa) at the dome / cylinder intersection to 580 psi (4 MPa) at the top of the containment dome. Figure 1, which you provided, shows a rough schematic of the Sequoyah plant with containment thicknesses and other modeling information. Variation in the containment thickness was modeled in the analysis as were the circumferential stiffeners. Figures 2-4 give information about the applied impulse. The pulse was assumed to have a velocity of 2.71E4 in/sec (688 m/sec) with an arrival time at the top of the containment of 40 ms. The analysis extended for a period of 70 ms. The containment was assumed to be constructed of A516 steel with a Young's Modulus of 30.E6 psi (207 GPa) and Poisson's ratio of .3. The containment was expected to yield, oy = 32000 psi (220 MPa), so a linear hardening slope of 30.E4 psi (2.07 GPa) was used. Available material data indicates the material to accommodate strains of 21 percent in an 8 in. specimen. Figure 5 and 6 show the axial and radial displacements of. selected nodal points around the containment dome. Nodal point 600 is located at the dome / cylinder intersection and nodal point 1390 is located at the top of the dome. The other nodal
.-w-w--._ _4 _ _ . , _ , _ _ . _ , . , - - . - - - , . , - , .y. -n .7
--,( 4., 9 g -g "
_._-.g4g--.-,. _ . , y-9 J5 9.,f,- w y g r-#,_,7,,
- o. .-
.m J. C. Cummings, 4441 January 11, 1982 points are located at uniform spacing around the dome circumference. The maximum deformation occurs at the top of the dome (node 1390) with a movement inward of 12 in (.304 m) and a maximum expansion of 9 in (.23 m). Figure 7 shows the effective stress-effective strain curve for the most highly This element is located at the top loaded element in the mesh.A maximum strain of 5.5 percent is of the containmen' dome.
This is far below the 21 percent allowed for the expected. material. The undeformed containment dome mesh is shown in Figure 8. Figures 9-12 show the deformed shape of the containment dome at selected times. Figures 13-16 show the deformations enhanced by a factor of 5. which high. lights-the deformations. In conclusion, we have studied the str'uctural response of the primary steel containment for the Sequoyah nuclear plant subjected to a hydrogen burn scenario. A traveling impulse of varying magnitude and fixed duration was used to simulate the burn environment. . A maximum strain of 5.5 percent was computed i for the containment which is approximately 25 percent of the allowable failure strain. We conclude that failure of the containment due to ductile rupture is not probable.- Some additional study should be given to the response of the containment between the welded stiffeners where a loss of ductility due to the weld might be expected. These results are preliminary and should not be released outside of Sandia without appropriate review. ' CMS:5521:njm Copy to: 4441 M. Berman 4442 T. E. Blejwas 4442 W. A. Sebrell 4444 R. K. Byers 5520 T. B. Lane 5521 R. D. Krieg 5523 R. C. Reuter . 5521 C. M. Stone 5523 P. S. Veers i 1
, I I
i i
, _r-_.-..e.- . , ,- -_ .-w-e se - e q g m r-1 .4-,1 a 4 . r .y--. -, .. -.--- - , - - - - ._ . -
l l
. 9
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?
rl e Bo1 s 6.
== g
? u *tal rt de3 1 d=f **/ ? Y \ .
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g&!
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e
. sL =lr/ e
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- f** 'T
. Eg 4, i' ej
-r .
.y,9 +
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=L sat s fL 3'i- *~ ' '
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- m_ yti .5 M1 .
mn 2 s .
-_~' ,- "
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,g. 4.:. W.
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--'-+---
t
~
, ,y"*
e E3,. 4 :~e' il
.j-,
.p=
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,M k ,
.g Oa HS -
! s # ,*
~ 1. 73.g n 7 2 <- . _,
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r-2 1 tri sk s w s, . s*
- i
@, m:=- 3
*=. : nL yd g
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;} 7a/ 1 ' s l .
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It il 1 Figure 1. Schematic of the Sequoyah Nuclear Power Plant Containment Structure
a q .. 1 i P MAX = = e I b i
= t e
m a b O. l I i
; Time
- At w Figure 2. Pressure History at a Point l
l l l
]
l l l
- I li .
\
0 .
. 5 5 c e
S . 2 1 2 0
=
t
) t s a e
0 h e
. 0 c m 5 n o i D
( g h n t o g l n A e L n o c i r t A u b i r t s 0 i 5 D . 4 e r u s s e r P 3 e r u 0 g , 0 i
, . 4 F 0 0 0 0 0 5 0 5 1 1 2
n m o. v E = M M E o.
! . . . ,J
e. l 500* I 400 n
; 300.
a w x
=c z
o. 200. 100.
=
Radians 0 uf4 r/2 Figure 4. Angular Position Measured From Dome / Cylinder Intersection I I
- -^- , -
1
' ' l ts .
- e. .
4 H88
- 1000
.. n ggg
-4 <
50 34 -
" . s -** .
" ..ss "
. is * ** . .s" TIME Figure 5. Axial Displacement Histories at Selected Nodal Loc'ations s.s .
s. 4..- 4. s.s . 5. 2.s < 1200 1.. . a.. 1000
. .s .
- . , 1390 U f' _
~
7 600 l
~~
. ., 800 1
= t.s .
2. e
- .. is .
'5
" .ns ***- . s " .sss TipE "s" .
Figure 6. Radial Displacement Histories at Selected Nodal Locations
. g ,
l o
.40eOE+5 <
.5589E+5 tO .5000E+ 5.<
m m E .2588E+5 < sn s .2689E+s < u M w
.1589E+5<
w
.15eBE
- 5 <
5808. ~ i E. .
.36 B. .51 .22 .85 .94 .95
.Ses .s s .s2s .ess .e45 .ess Effective Strain, c Figure 7. Stress-Strain Curve for the Most Highly Stressed Element in the Model. This Element is Located at the Top of the Containment Dome on the Axis of Symmetry.
e . q
. 1 l
l l 4 I-4 I- . I-N . it M
<n e a a a a
R-AXIS One Figure 8. Undeformed Mesh of the Containment Dome.- Element is Used to Model Behavior Through the. Shell Thickness. l l 1 i l l
- v;
- I l
l 1 l N - it X m I
- i. - l I
g.- -
-: a ]
l
- r R-AXIS Figure 9. Deformed Mesh Plot of Containment Dome at
- t = .019 Seconds. Magni fi cation .of Displacements is 1. f r l
'f 4
l I
- e. -
O
*y y y -
y- -r- 1 ~ r y --g-- 4 4 - [ -. + y- yy ,-p-
N - M m m E 2" . l, , . . . ; :.
, 5 i i i : : :
. R-AXIS Figure 10. Deformed Mesh Plot of Containment Dome at t= .036 Seconds. Magnification of Displacements is 1.
. 3. - . - . _ _ . . .
s-1 g. N . it X m 5 s.-
- R-AXIS Figure 11. Deformed Mesh Plot of Containment Dome at t= .0516 Seconds. Magnification of '
' Displacements is 1.
l
. ~
v
' l N .
ez M a
./
- - J r.-
- 4 4 4 1 4 a 4 r
R-AXIS Figure 12. Deformed Mesh Plot of Containment Dome at t= .0684 Seconds. Magnification of Displacements is 1.
. . . , - . - - - - , ,- , e , we e- ,
4 4 X m I
- - f..
l
- 1 I 1 a i a i
~
? R-AXIS Figure 13. Deformed Mesh Plot of Containment Dome at t= .019 Seconds. Magnification of I
Displacements is 5. I l l l l l I l l
7 2 N - i: M sn i- . I
. R-AXIS Figure 14. Deformed Mesh Plot of Containment Dome at t= .036 Seconds. Magnification of Displacements is 5.
1 1
- I 1
'- - -- - - - - . . - . . - , _ , , , . l
l 1 N . i: M
<n
- s. -
\
- i
- a. a a = *. :
. . : : . . l I
y R-AXIS Figure 15. Deformed Mesh Plot of Containment Dome at t= .0516 Seconds. Magnification of Displacements is 5. , 1 l i l p I . _. ,_ ,,.v.__ __ ._ _. _ ,_ __
- i l
l 1 8" g. 1 N i: x ei.e an I-e J* r* .- ; a a 4 a a
? R-AXIS Figure 16. Deformed Mesh Plot of Containment Dome at t= .0684 Seconds. Magnification of Displacements is 5.
g-n -, - ;cr.-
.. y,. , - -
. l m).~.2;~ .
-J %3C !
1 I MISSISSIPPI POWER & LIGHT COMPANY Helping Build Mississippi \ P. O. B OX 16 4 0. J A C K S O N. MIS SIS SIP PI 3 8205 s
,:,- ;; s NUCLEAR PRooUCTloN DEPARTMENT ,
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U.S. Nuclear Regulatory Commission b -. W Office of' Nuclear Reactor Regulation '/ ^'E 2 0 l'.'.b Washington, D.C. (~ 20555 O C; Attention:
- s. -
g.. - .... . 2 4 Mr. Harold R. Denton, Director x 4\' (\
Dear Mr. Denton:
SUBJECT:
Grand Gulf Nuclear Station Units 1 and 2 Dopket Nos. 50-416 and 50-417 File 0260/0756 Report on the Effects of Hydrogen Detonation as the Result of a Hydrogen Generation Event AECM-82/32 The enclosed report documents the ability of the Grand Gulf Nuclear Station (GGNS) containment to withstand the effects of local ( detonations. This analysis was conducted due to serious concern on the part of the Nuclear Regulatory Commission (NRC) regarding this issue. The results of this analysis satisfactorily resolve this area and no further work is planned by Mississippi Power & Light Company (MP&L) with regard to the effects of local detonations. It should be noted that MP&L does not believe that there is a potential for transition to detonation during hydrogen burns resulting from operation of the GGNS Hydrogen Ignition System (HIS). This is based not only on our judgement but on the evaluation conducted by - Bernard Lewis and Bela Karlovitz of Combustion and Explosives Research, Inc. (COMBEX), who concluded that there was no potential for transition to detonation. The results of the COMBEX evaluation are being submitted separately. Yours truly, I i l L. F. Dale Manager of Nuclear Services
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RMS/SHH/JDR:Im
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Attachment ![ / l cc: (See Next Page) ' 0 S' A [K O&co4j,, PDR AESR1
%f Member Middle South Utilities System l
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1, ** . AECM-82/32 MISSISSIPPI POWER & LIGHT COMPANY Page 2 j 1
.I cc: Mr. N. L. Stampley (w/a) . l Mr. G. B. Taylor (w/a) ,
Mr. R. B. McGehee (w/a) ! Mr. T. B. Conner (w/a) Mr. Richard C. DeYoung, Director (w/a) . Office of Inspection & Enforcement U.S. Nuclear Regulatory Commission
~~
Washington, D.C. 20555 L ]. , I 4 s i e L t . l I r l .i AE5R2
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I f A Description of the. A='restigations
. Into the Nature and Effects of g Assumed Hydrogen Detonations in the 1
Grand Gulf Containment st
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I i t r 6 I January 1982 y 't I 3' s s . i l i __monsom m ame;;; PDR ADOCK 05000416 A t PDR 2 opp :
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Table of Cantents 1
.. Section page f
. 1.0 Introduction 11 2.0 Significant Conclusions from Literature Review 2-1 i
e 3.0 Estimating the Pressure-Time Curve 3-1 3.1 Reported Pressures for Hydrogen Detonations 3-1 3.2 Spherical Detonation Waves Versus Plane 3-3 Detonation Waves 3.3 The Pressure-Time Curve 3-4 4.0 Containment Structural Analysis 4-1 5.0 Conclusions / Major Conservatisms 5-l' 6.0 References 6-1 I T 4 l 1
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l U ! r . l l ,-l 1.0 Introduction
! )
l The purpose of this study is to calculate the potential effects of' ** i
'j hydrogen detonations on the Grand Gulf containment building. In this .
- report, the probability of such an event occurring is not addressed. In fact, based upon a review of the available literature and discussions with recognized experts, it is concluded that detonations will not occur ,
for two reasons: -1) the hydrogen ignition system will prevent the g buildup of detonatable concentrations of hydrogen; and 2) clie Mark III containment geometry and environment are not amenable to inducing a
-l; transition to detonation. In spite of the conclusions regarding the
' ; possibility of detonation, this report conservatively evaluates the potential consequences of such an event. k
; The practice of attempting to calculate consequences resulting from j phenomena which are not well understood is analagous to the original i
I treatment of in-vessel steam explosions. In WASH-1400, an in-vessel - steam explosion leading to containment failure was felt to be'of major i
' importance among those events producing a significant radioactive release. The investigators based their judgement on the results of f
several small-scale experiments and accidents (BORAX, SL-1, etc.). Based on this limited knowledge of steam explosions, the authors took the most conservative approach and assumed similar results could be obtained in f( [ the large-scale reactor vessel environment. It was acknowledged that a e i great deal of uncertainty was associated with this scenario, 1 i
- Recognizing that so little was known about steam explosion phenomena, the
- NRC initiated research on this phenomena. With the advent of the Zion / Indian Point sevet:e accident studies in 1979, this research was accelerated. In a recently released Sandia Report (Reference 6), it was concluded that the in-vessel steam explosion scenario postulated in WASH-1400 was much less probable, and much less significant, than f originally thought. -
A similar situation now exists with regard to the potential impact from f hydrogen detonations. True, laboratory experiments in shock tubes (and y other apparatus) have produced detonations with significant pressure i pulses. If the results of these detonations obtained in the laboratory were assumed scalable to containment dimensions and geometries, then concern might be warranted, but the available evidence indicates that large-scale hydrogen detonations of a magnitude sufficient to threaten containment integrity will not occur. The probability of, and consequences from, hydrogen burning still tend to dominate the small fractional risk associated with hydrogen gas generation. i- There are no easily defined scenarios for modelling hydrogen gas detonations in a Mark III containment. Therefore, a procedure was 1, developed which extrapolates from the available theoretical and experimental data to obtain a conservative pressure-time history. ( However, the degree of conservatism can only be determined in a broad, qualitative sense. 1-1
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3 2.0 Significant Conclusions from Literature Review l- The first step in this study was to conduct a review of available ' g literature on ' gas detonations and shock waves. This review, which
- y encompassed a wide variety of references (1-5), led to confirmation of y some previously held judgements, as well as some insights important to o the issue at hand. The most significant aspects are briefly discussed '
i below, with further elaboration presented in subsequent sections of this j study. ;
- i] a. Shock tube geometries are not found in the Mark III containments.
> As expected, the available experimental data on hydrogen detonations
*l does not translate very well.to the Grand Gulf containment. Most 4
i; ~ experiments are conducted under very precise conditions, usually in 1 shock tubes. The volumes in a containment are too large and. ' unconfined to accommodate such detonations and the initial-
, atmospheric conditions are generally not as favorable for detonation y as the experiments. It is a major conservatism to assume that the results of these tests are applicable or scalable-to the containment ; ,- environment.
, b. Blast Waves and detonation waves have different properties. Simply 4
'y stated, a blast (shock) vava is a compression wave moving at
,1 supersonic speeds through a non-reacting medium (air). Blast waves are generated by the detonation of an explosive charge, such as TNT.
A detonation wave can be modelled as a shock wave travelling through
! a combustible gas (hydrogen and air) with a chemical reaction occurring behind the wave front. The detonation wave may be
- generated by a spark, a small explosive charge, or a combustion wave
~
degenerating into a detonation wave. The last case is the harde'st to obtain, requiring very precise geometry and conditions. This is further explored in Section 3.2.
- c. Wave reflection is very important. This study is ultimately concerned with the. interaction between the decor.ation wave and the containment wall. The basic scenario examined is one in which the pressure pulse generated by the detonation wave is the primary 4
, threat to the containment boundary. The temperature spikes which accompany a detonation are of very short duration, and are, q therefore, not a threat to the containment integrity. Assuming for j the moment that detonation waves behave the same as blast waves, this objective ~is simplified to determining the impulse imparted to the wall. The impulse is defined as the area beneath the pressure-time history curve.
y As a shock wave propagates through air, the air in the shock front is compressed. If the surface of an object in:the path of the shock wave is parallel to the direction of propagation, this surface will be exposed to the same pressure (the static pressure) as the compressed air. On the other hand,'if the surface is an obstacle' , standing in the way of the shock wave, then the_ impact of the air l mass may result in a pressure.many times greater than the static I pressure. The static pressure is frequently called the incident or
-undisturbed pressure, whereas the pressure resulting from impact of the air mass is the reflected pressure.
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This evaluation is concerned primarily with the reflected pressure [. and-its corresponding impulse, which can also be many times greater
- than the static imp.ulse.-
- d. Spherical waves versus plane waves. Most of the actual test data reported in the literature comes from shock tube experiments. 'In a shock tube, the shock wave experiences one-dimensional propagation, I
a plane whve. However, most actual shock waves are not one-dimensional, but three-dimensional or spherical. In the case of a g.
'. reactor containment, any postulated shock wave would probably have characteristics closer to a spherical wave than a plane wave.
The. distinction between spherical and plane waves is important when calculating the impulse generated during wave reflection. For plane shock waves, the reflected impulse may be, at most, twice the ine'ident impulse. But for spherical shock waves, the reflected insulse may be many times higher than the incident impulse. Section 3.2 presents a thorough discussion of the dd.fferences.
- e. The difference between spherical detonation waves and blast waves is 4
important. . The discussion in the preceding. sections (2.c and 2.d) 2 assumed that detonation vaves behaved much the same as blast wavas. In actuality, this is not true, and the differences between the two can be significant. While Section 3.0 of this report gives a detailed discussion of'this point, the major difference can be [ summarized as follows: the peak ref1seted pressure of a spherical l detonation wave can be as much as four times lower than that in an 4
" equivalent" spherical blast wave. This difference also results in a significant decrease in the calculated reflected impulse.
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I$ - N n h h 3.0 Estimating the Pressure-Time Curve i This part of the study will. calculate a conservative pressure-time h history which will be assumed to result from a local hydrogen detonation. In this context, the importance of distinguishing between detonation . k,- waves and blast waves, and plane or spherical wave propagation will be y demonstrated. D . F The basic characteristics which differentiate detonation waves from blast waves were given in Section 2.b. Section 2.c further established that (( it the pressure and impulse resulting from wave reflection are the primary
'/ concern in this study. A full appreciation of the differences between detonation and blast waves is gained when determining the respective h[ characteristics, including reflection from a structure, in each j calculation.
3.1 Reported Pressures for Hydrogen Detonations
- ] Table 1 provides a compilation of peak pressure in the incident i detonation wave for various mixtures of hydrogen and oxygen or air.
d! The incident pressure in the detonation wave is the pressure after l!, which the chemical reaction is complete. Among the references noted jj in Table 1, there was generally good agreement on the incident pressure. i f ; Only a few investigators reported measurements of the reflected
; E pressure in hydrogen-oxygen / air detonations. None of the .
~ [; experiments evalt:ated measured both the incident and reflected
; pressures in the same test. This was usually due to limitations in
' [- e the experimental apparatus. Table 2 compiles data obtained ou T, reflected pressure. Additional data is provided in Figure 1.
il ** As can be seen from Table 2 and Figure 1, there is' a substantfal difference between measured and calculated reflected pressures, and
,. there is also a range of reported values among the various
,.! investigators. Some insight can be gained from a closer look at
.f these experiments, and examination of the nature of detonation waves l h in greater detail.
- I l l Gordon (Ref. 2) provides an excellent description of a detonation l p- wave. A detonation wave is characterized as a constant speed shock ii wave followed by a reaction zone of finite width. In the very front
[ of the detonation wave, the combustible gas is compressed, without reaction, much the seme as in a blast wave. The chemical reaction
'i
..I , is initiated by the high pressure and temperature existing in'the
':P compressed gas. As the reaction continues to completion, the , 7j temperature of th'e gas rises considerably, while, conversely, the { pressure drops. For instance, Gordon examined the detonation wave - i"-( ' for a 20% hydrogen-air mixture. At the very front of the wave, when there has been no reaction, he reported a pressure of about 24 ata, and a temperature of approximately 1350*K. At the tell end of the 4 I
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, wave, when the reaction is near 100% complete, the pressure has
, * , dropped to approximately 13 sta and the temperature has increased to , .; 2425'K. This latter state point is the Chapman-Jouget (C-J) point. ,
, Behind the detonatioh wave,- the reaction product gases expand g rapidly, resulting in a precip,itous drop in pressure, temperature, and mass velocity.
a i _ The significance of the chemical reaction and the resulting [* 1 expansion of the gaseous reaction products is manifested when
) calculating the reflection pressure. Gordon estimated the
! reflection pressure which would: result for detonation waves where
} the chemical reaction went to various stages of completion. Table 3
/
summarizes his calculations. In a detonation wave with no chemical reaction 8.=0, in which case, it is no longer a detonation wave,
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! but a blast wave), the reflected pressure would be 153 sta. When i
-[ the reaction goes to 100% completion, the reflected pressure is only 31 ata. In the experiment, Gordon actually measured 44 sta,
) indicating the reaction had gone to about 90% completion.
Results similar to Gordon's were reported by Sokolik (Ref. 3) for
, experiments carried out by Kogarko.. In this case, Kogarko measured the reflection pressure and the detonation wave velocity. Then, for
- comparison, he made calculations of the reflection pressure for the following three cases
}*' a. For a detonation wave (DW) with 100% completion of reaction; l b. For a shock wave with an incident pressure equal to the detonation wave incident pressure (SWp);
For a shock wave with a velocity equal to the detonation wave c. velocity (SWv). i Table 4 gives the results of the experiments. Kogarke's results, like Gordon's, show that the pressures generated during wave reflection are much smaller for detonation waves than for
" equivalent" blast waves. His results also point out that, in 3 out of 4 cases, the equal pressure shock wave model (SWp) provides the closest approximation of the reflection pressure resulting- from a detonation wave. . . The basic conclusion to be drawn from the preceding ' discussion is i that detonation waves are profoundly different than blast waves.
Calculations of reflected pressures for detonation waves must consider the nature of the wave, since detonation wave reflected
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I pressure is much less than would be calculated for " equivalent" - - blast waves. - l t ! 3-2 j .
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. i 1 3.2 Spherical Detonation Waves versus Plane Detonation Waves
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E The pressures reported in the preceding section were from I > experiments which generated plane detonation waves. Another type of ' wave propagation, spherical, is also possible (cylindrical is the 4 _f. third type, but it has properties similar to spherical). .When [. calculating the effects from any shock wave, it may be important to i distinguish between plane or spherical propagation. 1 The relative significance between plane and spherical propagation is !
- j. best shown by a TNT blast example. This can be done by assuming a
[ detonation of a one kilogram charge of TNT located 2 feet from a ! Fg wall. Using the TNT blast characteristic curves in Reference 1, the [ impulse imparted to the wall can be determined for both type
+ propagations. In each case, the impulse due to the incident wave is f! the same, about 1.5 atm-asec. For a plane va've, the reflected
-impulse would be, at most, twice the incident impulse, or about-3.0 l ata-asec. However, if the wave were spherical, the impulse from the f reflected wave would be about 17 ata-asec, almost six times higher j than the plane wave case. Closer to the charge, the difference >
; between the spherical and plane reflected impulse increases (to a maximum of about 32 times), while at distances farther away, the
. difference disappears. In this example, the spherical and plane reflected impulse are equal at about 10 feet.
!I
, The type of propagation is also important for detonation waves. The differences, however, between the. spherical and plane detonation I waves are not as great as in the blast wave case. A plane detonation wave, in a shock tube, will be self-sustaining once initiated. In contrast, Sokolik found that a spherical detonation f will experience continuous attenuation as it propagates. Spherical-e detonations are characterized by abrupt drops in density, pressure, I and r. ass flow velocity just behind the reaction wave front . This
; velocity decreases te zero over a distance equal to half the redius
. of the sphsre enveloped by the detonation front. Thus, 1/8 of the '
total volume is filled with reaction product gases at rest. This characteristic tends to decrease the magnitude of the difference in l *. impulses between the two types of detonation waves (as compared to plane and spherical blast waves). Also, as with the blast wave, at
- large distances from the center the characteristics of the spherical
- detonation wave tend to resemble those of a plane detonation wave.
3 The critical question is whether a hydrogen detonation 'should be i assumed to be a spherical or a plane wave. The most conservative , y approach would be to use spherical wave propsgation; however, it may d not be possible to generate a spherical detonation. Sokolik (Ref. . :. 3) summarized the events necessary to form a spherical detonation , via spark ignition: i.
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"1. the propagation away from the spark of the primary flame front l j leaves part of the chemical energy unliberated; the higher the 3 flame velocity, the greater is this energy. This may be related to a retardation of the reaction in the flame when the ~
h compressions and rarefaction waves reflected from the wall pass L through it; j .? 2. a compression wave reflected from the envelope is focused at-i i- the center of the -charge, causing at this point an explosive {? literation of the remaining energy and the production of a
. spherical shock wave; this type of ignition is analogous to the explosion of a powerful detonator;
'.(( 1 3
- 3. finally.. propagation of the shock wave, related to the liberation of the energy remaining behind the primary flame fror.t. represents the final stage of spherical detonation."
~
} [ , Therefore, without a powerful detonator to initiate detonation, spherical detonation will be difficult _ to generate and impossible t without an enclosure. It is highly improbable that these conditions will exist in a Mark III containment.
^
i t Both plane and spherical propagations will be examined. ) i 3.3 The Pressure-Time Curve i
? A significant part of this study is the selection of the
- pressure-time history to be assumed for the postulated hydrogen
{*;j detonation. It must be emphasized here that this selection prccess j~ is extremely judgemental, and, in some cases, arbitrary assumptions are necessary to define the detonations. Theoretical calculations !, done for this study have been supported and modified with
~
; er.perimental evidence oto ensure a conservative evaluation. For if example, shock ' tube data .has been assumed applicable to the containment despite r'anificant environmental and geometrical
! differences. Blast wave correlations have been used to calculate
, [ detonation wave parameters despite the differences already noted. All of these are significant assumptions; however, the final j conclusion is that the pressure-time curves selected are very .t conse rvative. 1 The data presented in Section 3.1 vill form the basis for the I pressure-time curve. Two concentrations of hydrogen will be i { assumed, 20% and 67%, only because this represents the upper and l ; lower bounds of the data available (see Table 2). l As mentioned in Section 3.1, there is reasonably good agreement on. ,[ l the incident pressure in the detonation wave. For a 20% hydrogen-air mixture, the incident pressure (P )g is approximately 13 ate, and
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for 67% hydrogen plus oxygen (assumed to be the same as air), P =19 g i: ata, i5
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[. . j The reflected pressure (P ) is a re difficult t calculate. Some 4 2 i i values are given in Table 2. Also, in Reference 5, a methodology j for~ calculating the impact pressure developed in hydrogen-oxygen ' 4 detonations is given. A conservative application of this technique 1 : - yielded a ratio of P /P equal to 2.90. This agrees closely with 2 g
'e the ratio calculated in Table 3 (at=.90) of 2.94. If ratios are l 4 taken of the measured values reported in Tables-1 and 2, the value of P2/Pg lies between about 2.9 and 3.5. Therefore, it was conservatively assumed that the peak reflected pressure is 3.5 times d, the incident pressure. This yields values of P equal t 45.5 ate 2
and 66.5 atm for the 20% and 67% hydrogen mixtures, respectively. { Having obtained an estimate of the peak pr~ essure, the next important-t parameter to determine is the time duration of the pressure pulse. The shape of the curve is also very important. Unfortunately,'very
. little data was found on these parameters for hydrogen detonations.
T Therefere, to determine these parameters. TNT blast wave characteristic curves were used. These curves are available in l Reference 1 (Figures 25, 26, 27). The degree of conservatism
} introduced by this model is discussed in Section 5.0.
The resulting pressure curves are shown in Figure 2 with significant parameters summarized in Table 5. For each hydrogen mixture
.( : concentration, two curves are shovn; one is the pressure pulse assuming
.: spherical propagation, the other assumes plane propagation. The significance of the type of p.ropagation is discussed in Section 3.2.
The worst case curve in Figure 2 is curve 1. As expected, it is the 67% hydrogen mixture assuming spherical propagation. The maximum impulse is 21 atm-msec. By comparison, the impulse generated , assuming plane propagation is less than half the spherical case. l* only 9 atm-msec. , 1 I I i f f 3-5 5 .
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j 3 4.0 Containment Structural Analysis-1 The load of the local detonation is a short-duration pressure impulse ' j load. It can be expressed as a pure impulse, denoted by: i = impuls'e = 21 sta-asec = .31 psi-sec. (See Table 5) p For an idealized bilinear behavior system, the required resistance of a
- ; structure can be obtained by
- ,
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- R a = ( 2y - I. (EQ. 5.16 of Ref. 7) where R,= required yield resistance in terms of uniform pressure, psi.
l ,' W = angular natural frequency of structure, rad /sec. 1-1 4 = ductility ratio.' . 5 In the ' case of elastic response (q=1), EQ. 5.16 of Ref. 7 is reduced to a static equivalent pressure of:
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j El " i "'# i Using the above formulations, the dynamic responser of the containment wall, equipment hatch and upper personnel lock are summarized in Table 6. From Table 6, it can be observed that the structural components have i adequate capacities J:o resist the postulated hydrogen detonation. Although the equipmenI hatch and the panels of the personnel lock exceed their elastic limit, the ductility ratio is well within the acceptable value for the short duration loading (Refs. 8, 9). In addition, th,e dynamic increase factor (DIF) for the material yield strength and the actual material yield strength have not been utilized in the analyses. Based on the ductility ratio of 3 for the containment wall structure (Ref. 8) and 10 for steel structures (Ref. 9), the limiting pressure impulse for the containment is estimated to be .375 psi-sec., or 26 ' ata-msec. , J i . l 1 4-1 l
l *.* t 4 h. d 5.0 Conclusions / Major Conservatisms
;l i $ This report demonstrates that the Grand Gulf containment is able to
[4 withstand the pos'ulated t local detonations of hydrogen under a wide
- variety of conditions. This is true even under the very conservative j assumptions established in this report. It is emphasized that the pres.;ure profiles established in Section 3.0 and shown in Figure 2 are f hypothetical; there is no real set of circumstances which could generate 4 '; pressures this severe. The major conservatisms incorporated into this
, study are listed below.
- h 1
o Data and observations obtained from shock tube experiments were , assumed applicable to the Mark III ccatainment geometry and
,: . environment.
4
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- Ig o The worst case pressure pulse assumes a 67% hydrogen mixture. This q curve is actually based on experiments for a 2H2 + 0, mixture. If
- r the mixture were 67% hydrogen + air, detonation may flot occur, since g the upper detonation limit for hydrogen in air is approximately 65%.
, l
}y o In calculating the peak reflected pressure, the maximum ratio of reflected pressure / incident pressure was used (3.5). The ratio is ;. actually closer to 3.0. 'p Q o The worst case pressure pulse assumes spherical wave propagation.
g Spherical propagation yields much more conservative (greater) impulses than plane propagation; furthermore, spherical detonations
, ; are much harder to obtain. '
s o.
! c The impulse due to the reflected detonation wave is overestimated by )l the blast wave curves. The blast curves-do not incorporate the fact j that the gases behind a detonation wave experience a severe i 4 deceleration, eventually coming to rest; thereby reducing the impulse.
, o The impulses given in Table 5 consider only the positive phase of j wave. reflection. There is a negative phase, for both incident and reflected waves. In actuality, the resultant impulse should be ll f f used, which would be less than that given. I j,* o When the containment wall response due to a pressure pulse loading is determined, the peak impulse (from Table 5) is applied over the
>I
- entire wall surface. Realistically, this peak impulse would only be
; , seen in the wall area located at right angles to the wave . propagation. For surface are:s not at right angles, the impulse is 4 ; less.
i: o The effect of varying the initial hydrogen mixture pressure and
. ! temperature was not considered. Initial conditions are usually 1 i atm and 291*K (with some variations in temperature used among the 1
investigators). The primary effect of changes in the initial 3 conditions is to modify the detonation limits, which would have no impact on the results of this study. I 5-1
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!' 6.0 References i
- 1. Granstr'om S.A.;" Loading Characteristics of Air Blasts from '
i i Detonation Charges," Acta Polytechnical, No. 196. 1956.
? l ; 2. Gordon, W.E.; " Third Symposium on Combustion and Flame and Explosion ; Phenomena," p. 579, Baltimore, 1949. '
- 3. Sokolik, A.S.; "Self-Ignition, Flame and Detonation in Cases," i
; Translated 1963 NASA TfF-125, OTS63-11179.
i.
',.' 4. Lewis, B. and G. Von Elbe; " Combustion Flames and Explosions of Gases," 2nd Edition, 1961.
i
- 5. Moyle, M.P. and Churchill, S. W.; " Impact Pressures Developed in l Hydrogen - Oxygen Detonations," Symposium on Shock Waves in Process Equipment Annual Meeting; Chicago, Ill., 1957. *
- 6. " Probability of Containment Failure Due to Steam Explosions Following a Postulated Core Meltdown in an LWR"; Sandia National Labs, NUREG/CR-2214; June, 1981.
~
- 7. Biggs, John M. , " Introduction to Structural Dynamics," McGraw-Hill Book Co., 1964.
- 8. ACI 349-80, Code Requirements for. Nuclear Safety Related Concrete -
Structure Appendix C.
- 9. Nuclear Regulatory Commission, Standard Review Plan. Section 3.5.3, Appendix A. July 1981.
i 6-1 6
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;' Incident Pressures in Hydrogen Detonation Experiments
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Mixture Composition Incident Pressure (ata) Reference s
; Measured Calculated i
-{ 2H2+O2 20.4 18 3
': 18.05 4 il 13.3 19.0 2
- 18 .06 3
; (2H 14.13 .4 2 + O2 ) + 502 4
' 15 .0 3 16.0 14.1 2
.! 13.8 3 j (2H 2 + O2 ) + IN2 17.37- 4 (2H 2 + O2) + SN2 l4*39 4 (2H 16.3 4 2 + O2 ) + SAr (2H 16.3 4 2 + O2) + SHe
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9 17.31 3 20% H - Air 2 13 3 ,' S P "
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io Table 2 b i L Reflected Pransure in Hydrogen Detonation Experiments I Mixture Composition Reflected Pressure (Atm) Reference I
! M. uad Calculated 2H2+O2 7.3 --
3 [ 40 4* 46.7 2 44.8 34.4 2 25% H2 .- Air 31.0
; 20% H 2
. . Air 44.2 2 1
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*In the 2H, + O2 mixture, G rdon stated that the measured reflection pressure I of 40.4 aEm is smaller than expected because of limitations in time g resolution of the instrumentation.
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Figure 1 I Pressure accompanying detonation-wave reflection in hydrogen-air mixtures:
- 1) calculated; 2) measured (according to Kogarko and Zel'dovick, Ref. 3). ,
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j- . s Table 3 lil l 5 I. Theoretical Pressures Before and Af ter Reflection ~ For Different Degrees of Reaction (*) in 20 Percent H2 - Air (Ref. 2)
}
S T- [ P(before) P(after) t oc (ata) P(after)/ (ata) P(before) T 0.0 23.7 .152.9
! 0.9 6.46 '- -
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!f .
Velocity of xerlection Pressure (ata) detonation wave 6 Mixture Composition (DW), m/.sec Heasured DW SWp SWV ' i < f 2H 2820 67.3 46.5 2+O2 99.7 237 CH3 8 + 502 2530 195.5 106.5 200.3 400.7 CH4 + 202 2322 168.4 68.2 203.2 464.5 l CH4 + 202 + 2N2 2030 149.0 54.2 103.6 '207 1 0 s 6 s e 9 l t 9 l 1
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\
e MISSISSIPPI POWER & LIGHT COMPANY Helping Build Mississippi WRMIAHddB P. O. B O X 18 4 0 J A C K S O N, MIS SIS SIP PI 3 9 205 June 25, 1982 NUCLEAR PROoUCTION oEPARTMENT U.S. Nuclear 7egulatory Commission d. % '3 7
^"
Office of Nuclear Reactor Regulation ' - _. . . _ . . _ . . Wash,ington D. C. . 20555 Attention: Mr. Harold R. Denton, Director
Dear Mr. Denton:
SUBJECT:
Grand Gulf Nuclear Station Units 1 and 2 Docket Nos. 50-416 and 50-417 Fire 0260/0756
Reference:
J Cfi-82/26F Local Detonations AECM-82/292 Based upon our review of the Sandia evaluation provided in the above reference, no compromise in containment integrity or the two personnel locks located at l evations 124'-8" and 124'-10" resulted. The results of the containment evaluation is presented in the above i reference. The MP&L evaluation of the air locks is attached for your
.. information.
It is our understanding that this completes the efforts in the local detonation area; and, as satisfactory results have been achieved, no further action is required by MP&L. Yours truly, F Dale Manager of Nuc' lear Services RMS/SHH/JDR:de I Attachment cc: (See Next Page) gDOI ia' re-~ ~ ,._ i*DR ADOCK'056 A 4i6 PDR Member Middle South Utilities System
. . - 7 e ~6 AECM-82/292
~^ #82 MISSISSIPPI POWER O LIZHT COMPANY cc: Mr. N. L. Stampley j Mr. R. B. McGehee Mr. T. B. Conner Mr. G. B. Taylor Mr. Richard C. DeYoung. Director Office of Inspection & Enf'orcement U. S. Nuclear Regulatory Commission Washington, D. C. 20555
-" ' ~ " ~ ' "~ ' ~
O'Reilly, Yg3~caal Adm51sYrhIor '
' ' ~
Mr. . . 'P.. Office ot~ Inspection and Enforcement U.S. Nuclear Regulatory Cummission Region II 101 Marietta St., N.W., Suite 3100 Atlanta, Georgia 30303 2 I k e i I 1 l < i n 9
*5 6
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Attachment to AECM-82/292 Structural Response of Lower Containment and Drywell Personnel Airlocks Due to Local Hydrogen Detonation , t The Grand Gulf lower containment and drywell personnel airlocks'(below the HCU floor level) have been evaluated for structural adequacy regarding the ability 4 to withstand ,a postulated local hydrogen-detonation. The Impulsive Pressure-Time.(IPT) curve associated with the local detonation was defined by ' Sandia National Laboratory and is presented in Figure 1 in terms of absolute ' l pressures. Although the IPT curve in Figure 1 is shown abruptly truncated at , in. -reality, according to
. researcl fifteen milliseconds (msc:))tne W O , M. : merman , urv_e does e_xtend beyo.the nd-15 msec andre.sponsibl_e will. .Sand
- eventually attenuate, to atmospheric pressure. As discussed with Charles i Tinkler of the NRC Containment Systems Branch, this information has been
! accounted for in the present. evaluation by extending the IPT curve in Figure 1 j with a constant pressure line beyond 15 msecs for an additional 1.5 msec and
- attenuating the pressure to atmospheric at 19 msecs.
} The response of each personnel lock has been evaluated by modelling various
! components of the lock as an elasto-plastic, Single Degree of Freedom (SD0F) dynamic system and subjecting it to the given IPT curve. The SDOF system is based on equivalent mass and spring properties' calculated in accordance with
- the method outlined in the reference cited below. The static pressure at which section yielding is initiated in the components is conservatively taken
! as the yield resistance of the SDOF system.
The SD0F. system described above was analyzed for the augmented Sandia IPT curve utilizing a computer program which employs step-by-. step numerical. integration technique to solve the equation of dynamic equilibrium. The results of the analysis' indicate that the subject locks are capable of l withstanding the Sandia pressure impulse resulting from a local hydrogen detonation by mobilizing their yield resistance. The resulting ductility ratios are calculated to be less than ten. .t J
Reference:
Introduction to Structural Dynamics, J. M. Biggs, McGraw Hill { Book Company. - l l l 4 Y 5 6 ee
-- e : n -- n--===-.---.---
.,..A.ttachment to AECM-82/292 ph 1 s-
- a .- 1
( ,N EULERI AN i POINT 11 ( 3.086E+00. 4.115E-03. ^ ---
.- 2.400 . , a i . , , i i . 'i . .
to -
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y , o - .
- 4 204 . : __. . . .. .. . . . . . .. .____-
2.000 - 1.800 - - 1.600 - - 1.400, - - j 2.200 -
~
E 1.00 - -
= :
sn 43 . p .8000 ' - -
- z. .
\ .
.6000 -
.4000 - .
.2000 - -
J
-0.000 '. ' ' ' ' ' ' ' ' ' ' ' '
1.00 : 3.0,0 5.00 7.00 9.00 11.0 13.0 l' T i t'.E (SEC1. '", XIC MD GULF. 0.28 H2-ORY AIR
^
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FIGURE 1 - -- .. - - . . . . . , . . . ~ . . . . , , , . . .
a r_ ll MISSISSIPPI POWER & LIGHT COMPANY
}J l
~
~
Helping Build Mississippi ( P. O. B O X 164 0. J A CK S ON. MIS SIS SIP PI 3 9 2 05 HEAR PRODUCTsON DEPARTMENT 1982 - - U. S. Nuclear Regulatory Commission - Office of Nuclear Reactor Regulation J V , 3 8' Washington, D C. 20555-Attention: Mr. Harold R. Denton, Director
Dear Mr. Denton:
Subject:
Grand Gulf Nuclear Station
- 'i Units 1 and 2 Docket Nos. 50-416 and 50-417 File 0272/0756/L-860.0 Hydrogen Contfol - Local
, Detonations and Drywell Head External Pressure Capacity AECM-82/315 In response to informally transmitted concerns of tha Structural Engineering Branch Reviewer, Dr. C. P. Tan, Mississippi Power & Light Company is
[- submitting information (Attachments 1 and 2) regarding local detonations
~ and drywell head external pressure capacity. This letter supplements the information transmitted via AECM-82/118, dated March 31, 1982, and AECM-82/271, dated June 11, 1982, and confirms information provided to Dr. Tcn on June 30, 1982. _ ,
If you have any further questions, please contact this office. Yours truly,
.J /,
. L. F. Dele ..
Manager of Nuclear Services f1SM/SHH/JDR:mm Attachments cc: Mr. N. L. Stampley Mr. G. B. Taylor Mr. R. B. ficGehee tir. T. B. Conner Mr. Richard C. DeYoung, Director Office of Inspection & Enforcement U. S. Nuclear Regulatory Commission Washington, D. C. 20555
- 72,^'7^1 o j M
%e _-m-
l MIO 12:lPPI POWER & LIGHT COMPANY . l
, AECM-82/315 l .
Page 2 Mr. J. P. O'Reilly, Regional Administrator , Office of Inspection & Enforcement - - 1 U. S. Nuclear Regulatory Commission Region II 101 Marietta Street, N. W., Suite 3100 Atlanta, Georgia 30303 9 0 s
- e e e Y
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, l Attachment 1 to AECM-82/315 ( Leakage through, Containment Personnel Airlock The lower containment perspnnel airlock has been analyzed for a local hydrogen detonation using the Sandia pressure-time curve. The results of the analysis indicate that the maximum lateral defonmation of the door panel on~t.he airlock will be small albeit with a ductility ratio of less than 2.52 For mid panel deformation of this magnitude, the performance of the inflatable seals at the periphery of the , door will not be affected. Considering the very short duration of the pressure pulse, it has-been determined that there would be no leakage out of the containment when both doors on the airlock are in the closed position. Leakage through the interior-door, even when the leak path is postulated to be extremely severe, will be insufficient to overcome the pressurized seal on the exterior door.
~
In fhummary, neither the response of the airlock door due to the postulated local hydrogen detonation nor the response of the seals to the same detonation will initiate leakage out of the containment. e ao # - g s
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_.- . ~ . . .. ..- . . - . . Attachment 2 to AECM-82/315 Capacity of the Drywell Hea_d under External Pressure The buckling capacity of the drywell head, which is comprised of an elleptical shell and a circular cylinder, has.been evaluated using a. conservative formulation developed at the David Taylor Model Basin (DTMB). The critical external pressure for the elliptical head was calculated using the method given in the-DTMB Report 1757," The Effect of Initial Imperfections on the Collapse Strength of Deep Spherical Shells." This method employs nominal dimensions and worst case deviations from nominal dimensions to determine the critical pressure, which was calculated to be 135 psi. The critical external pressure for the cylinder was determined using the method given in the DTMB Report 1639, " Structural Analysis and Design Cqhsiderations for Cylindrical Pressure Hulls." This method also employs
- nominal dimensions and worst case deviations from nominal dimensions to determine the critical pressure. When the calculations were performed for a conservative out-of-roundness based on the out-of-roundness, tolerance given on the manufacturing ' drawings, the minimum differential pressure capability of the cylinde'r was calculated to be 89 psi. However, when a best estimate of the out-of-roundness is used in the formulation, the minimum differential pressure was calculated to be 180 psi.
In summary, the buckling capacity of'.the drywell' Head has been evaluated using j a formulation taking into account the effect of initial imperfections. The lower bound estimate of 89 psi for the buckling capacity of the drywell head is based on a conservative assumption regarding such initial imperfections. [ m l
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