ML19345D441
| ML19345D441 | |
| Person / Time | |
|---|---|
| Site: | Sequoyah  |
| Issue date: | 11/10/1980 |
| From: | Dalzell W, Gillis J EMVWALKF, KIDDE, INC. (FORMERLY WALTER KIDDE & CO., INC.) |
| To: | |
| Shared Package | |
| ML19345D436 | List: |
| References | |
| PSR-914, NUDOCS 8012150113 | |
| Download: ML19345D441 (175) | |
Text
O DETERMINATION OF IGNITION PERFORMANCE CHARACTERISTICS OF GLOW PLUG HYDROGEN IGNITOR I
FOR WESTINGHOUSE ELECTRIC CORPORATION PITTSBURGH, PENNSYLVANIA REPORT NO. PS R-914 Issued: November 10, 1980 Prepared by: / ^ <r[...// />'/
l Warner G. Dallell Te't o Engineering Supervisor P.otection Systems Division '
- t Approved by: .
_ ///a hosepf . Gillis Wanage -Explosion Protection Systems Protection Systems Division
, A l t& lit;f!O NW FENWAL INCORPORATED : ASHLAND, MASSACHUSETT5 w - ,,w..- = 5c .'- THis DOCUMENT CONTAINS
! _. P00R QUAtlTY PAGES ,
I 8032150(\h \W -
I
Report No.
PS R-914 age 1
SUMMARY
A series of tests have been conducted to ascertain the ignition capablity of a special glow plub ignitor in various mixtures of hydrogen, air and steam. Comparison of the test results, e.g. pressure and temperature transients due to com-bustion of hydrogen, with previously published information has shown good agreement. The performance of the glow plug ignitor in igniting hydrogen mixtures has been consistent with the lit-erature and satisfactory in all respects.
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Report No.
PSR-914 Page 2 RESULTS Test 'H 2
Steam v 6 P, No. (%) Added (Ft/S ec) (PS I) 1 12 No 0 53.00 2 8 No 0 33.00 3 8 No 0 3.00 4 12 Yes 0 66.00 5 8 Yes 0 22.60 6 12 Yes 0 72.00
- 7 8 Yes 0 16.25 l
8 12 Yes 5 67.50 l 9 12 Yes 10 65.00 10 10 Yes 10 53.70 11 10 Yes 5 52.70 l 12 12 Yes 10 58.75 13 12 Yes 0 60.00 14 8 Yes 0 30.00 H
2 Hydrogen Test Concentration (%)
HO2 Steam Added (Yes - No)
V _- Air Velocity at Glow Plug (Ft/S ec)
OP -
Maximum Pressure Increase (PS I)
Detailed Results are Shown in Table No. 1.
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Report No.
PSR-914 Page 3 APPARATUS Tests were conducted in a 1000 gallon spherical test vessel having a pressure rating of 500 psig with the capability o'f being heated to 350 F. The vessel is constructed of carbon steel with a stainless steel liner.
The outside surface of the vessel was insulated with 3 inch thick fiberglass insslation. Thic insulation had an aluminum foil face which oriented away from vessel, i
Mixing of the various gaseous components was acccmplished by means of a small shaded pole electric motor fan. This fan had a 4 inch diameter blade with an air moving capacity of 200 CFM.
Steam was supplied to the test vessel from an electrically heated boiler which was self-regulated to maintain a pressure of 40-50 psig. A r.anually operated ball valve was positioned l ,
between the boiler'.and the test vessel.
The temperature of the test vessel was controlled by a thermocouple controller which had its sensing element in a well inside the vessel and approximately 18 inches from the vessel wall.
. +
The temperature of the test vessel was sensed and recorded from a thermocouple which was approximately 12 inches below the geometric center of the vessel.
m FENWAL INCORPORATED : ASHLAND, MASSACHUSETTS th .J wehe, it.dde & Co.,.ay. sac.
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Report No.
PS R-914 Page 4 APPARATUS (Cont'd)
The temperature of the test vessel wall was sensed and re-corded from a thermocouple which was silver soldered to th'e ves-sel inside wall at a point approximately 12 inches below the equator.
Transient pressures were monitored by means of two strain l guage-type pressure transducers, the output of which are fed to a Consolidated Electrodynamics Corporation tecording oscil-lograph. Timing markers were electronically superimposed on the oscillograph chart, providing a time base to facilitate the i
determination of the rate of pressure rise. One transducer was calibrated to read relatively low pressures resulting from margin-al pressure transients and the other was calibrated to read higher pressures resulting from more complete combustion.
l A mercury manometer was used to measure pressures during 1
the loading of gaseous components by the partial pressure method.
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Samples for gas chromatograph analysis were taken from the test vessel, through a cooling / condensing chamber into a 500 ML glass sample bulb. A vacuum pump and various valves were used so as to be able to draw the sample first into the cooling /
condensing chamber and then into the glass sample bulb.
Air flow across the glow plug (when specified) was provided by a small shaded pole motor electric fan placed on an adjustable horizontal mount.
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s Report No.
PS R-914 Page 5 APPARATUS (Cont'd)
Precise positioning of the fan was done each time air flow was specified by measuring the air flow at the glow plug with an Alnor Series 6000-P Velometer and moving the fan accord'ingly.
This fan had a 4 inch diameter blade with an air moving capacity of 200 CFM.
The temperature of the outside wall of the glow olug box was sensed and recorded from a thermocouple silver soldered centrally on one of the vertical box walls.
The temperature that might be experienced by the glow plug transformer was sensed and recorded from a thermocouple which was silver soldered to a bracket which was similar to the trans-former bracket and mounted inside the glow plug box in a similar location. (Used in tests No.1 and No. 2) .
The gas temperature of the interior of the glow plug box ,
was sensed and recorded from a thermocouple suspended inside l the box. (Used in tests No. 3 through No. 14).
l l c l All thermocouples were 24 gauge iron constantan welded junction with teflon insulation.
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This apparatus is shown diagramatically in Figure No. 1.
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Report No.
PS R-914 Page 6 PROCEDURE Vessel temperature was stabilized at the specified test temperature.
l Barometric pressure, relative humidity and ambient temper-ature were read and recorded.
Air, hydrogen and' steam (when specified) were added accord-ing to the appropriate partial pressure.
The vessel contents were mixed for approximately five min-utes.
The gas sampling apparatus was evacuated and the pre-burn l gas sample was drawn into the cool.19/ condensing changer and 1
i held for 2-3 minutes. The gas sample was then transferred to the glass sample bulb.
The mixing fan was stopped for approximately two minutes.
l The glow plug was energized.
The post-burn gas was sampled in the same manner as pre-
! viously described. -
1 The pre-burn and post-burn gas samples were analized by laboratories having gas chromotography capability. Gases from FENWAL INCORPORATED : ASHLAND. MASSACHUSETTS ,
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Report No.
PS R-914 Page 7
, PROCEDURE (Cont'd) l tests No. 1 through test No. 5 were analized by:
Arnold Green Testing Labs Inc.
! 6 Huron Drive l
Natick, Massachusetts Gases from tests No. 6 through test No. 14 were analized I by:
Dynatech R/D Company 99 Erie Street I
Cambridge, Massachusetts l
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Report No.
PS R-914 Page 9 Legend for Table No. 1
%H 2 -
Hydrogen Test Concentration (%)
Tv -
Vessel Test Temperature ( F)
V -
Air Velocity At Glow Pub (Ft/S ec)
Baro -
Barometric Pressure (mmHg)
R/H -
Relative Humidity (%)
T amb - Ambient Temperature (OF)
P air - Partial Pressure (mmHg) Of Air Loaded PH 2 - Partial Pressure (mmHg) of Hydrogen Loaded PHO2 - Partial Pressure (mmHg) of Steam Loaded Ty - Glow Plug Box External Wall Maximum Temperature ( F)
T2 - Vessel Internal Wall Maximum Temperaturd I F)
T 3
- Glow Plug Box Internal Maximum Temperature ( F)
T - Vessel Air Maximum Temperature ( F)'
4 Tign - Time From Energizing Glow Plug to Ignition (S ec)
Tp - Time From Ignition to Maximum Pressure (S ec) l
/k P - Maximum Pressure Increase (psi) .
Hp -
Pre-burn Hydrogen Concentration (%)
Np - Pre-burn Nitrogen Concentration (%)
Op -
Pre-b'Irn Oxygen Concentration (%)
Ha - Post-burn Hydrogen Concentration (%)
l Na - Post-burn Nitrogen Concentration (%)
Oa - Post-burn Oxygen Concentration (%)
Su -
Burning Velocity (Cm/Sec)
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APPENDIX 0 (OMITTED) s l
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,i APPENDIX P 2
SENSITIVITY ANALYSIS 1
A sensitivity analysis has been performed by Westinghouse /0PS to evaluate the effects of ignition criteria and containment safety 4
system performance on centainment response. Calculations were i
performed using the Westinghouse /0PS CLASIX code. The sensitivity 1
analysis is discussed in Appendix U and CLASIX is discussed in
, Appendix T.
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P-1
APPENDIX Q ANALYSIS OF EFFECTS OF TMI-2 EVENTS The accident at the Three Mile Island nuclear facility resulted in core damage and the production of sJ snailcant quantities of hydrogen as described in Volume I, section 4.3. Estimates of the hydrogen burned in the containment have ranged from 452 pounds to as much as 1,160 pounds. A reasonable estimation of the burned hydrogen appears i
to be scmewhere between 500 and 600 pounds. This would represent between a 30- and 40-percent metal wat,er retetion at Sequoyah from figure A-4 of Volume I and would result in a concentration of 9-125 by volume (neglecting the presence of steam for conservatism). The containment pressure response to the hydrogen burn at TMI was a pressure spike of approximately 28 lbs/in g, as described in Volume I.
Figure Q-1 repeats the pressure response for convenience. However, it has been noted (reference Q-1) by investigators that the te=perature rise at TMI was only on the order of 50 F as measured at several containment locations. Figure Q-2 shows the recorded temperatures. Cons *ve calculations of the temperature and pressure response of the Sequoyah containment design to a similar hydrogen deflagration have been compared by others in recent weeks 1
to the actual data obtained at TMI. TVA does not believe at this point that such a comparison is valid since the analytical models are not yet sophisticated enough to perform a truly best estimate burn calculation (see Section VIII.E and Appendix F). It is expected that application of the same conservative calculation mechodology to the TMI system would yield results far in excess of those observed e
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at the Three Mile Island facility, particularly in the temperature comparison. Such a comparison should provide insight into the importance of each physical phenomena neglected in present analytical models and will provide some indication of the response that could be expected from the ice cendenser pressure suppression containment to a hydrogen deflagration. Therefore, TVA is working toward this type of comparison.
Plant differences also exist between TMI and Sequoyah. In particular, the nuclear steam supply system at IMI we' =anufactured by the Babcock
& Wilcox Company and consists of a tworloop design with once-through steam generators. Sequoyah employs a four-loop U-tube steam generator
- of Westinghouse design.
Differences also exist in the containment design beyond the obvious pressure suppression versus dry containment concept design. A study is in pecgress to examine the TMI events in detail for determining the approximate response of systems that could reasonably be expected at Sequoyah if similar events occurred. Low power tests have also been run at Sequoyah to demonstrate the ability of the system to achieve and sustain natural circulation heat removal at power levels typical of decay heat following a small break LOCA. Positive results were obtained in these in-plant experiments. Plant operations personnel have also been trained to recognize events such as occurred at TMI and to take appropriate action to maintain adequate core ecoling at all times.
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'o perfors the above analytical tasks, TVA is also pursuing the procurement of outside training on the use of appropriate computer codes for analysis of accidents beyond the design basis. The MARCH computer code written by Battelle Columbus Laboratories appears to be a good candidate for this type of training anj accident simulation.
In anticipation of obtaining MARCH or a similar tool, work has been initiated on the generation of appropriate plant parameters and accident sequences for study. Results from these studies will be used to improve the hydrogen mitigation system as appropriate.
References Q-1 Analysis of Three Mile Island - Unit 2 Accident, NSAC-1 Appendix HYD, July 1979
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APPENDIX R OPERATING PROCEDURES To reflect the installation of the Interim Distributed Ignition System (IDIS) at Sequoyah unit 1, the emergency operating procedurer bsd to be revised. The existing emergency operating instructions (E0I's) were checked for procedures that would; (1) Conflict with the operation of the IDIS.
(2) Prevent the IDIS from accomplisning its function, and
( ~J) Lead to a more severe accident scenario than those previously evaluated because the procedures did not include instructions for energizing and deer 4ergizing the IDIS.
After reviewing all of the Sequoyah emergency operating instructions,
'!VA did not find any reccedures,that conflicted with the expected operation of the IDIS or led to a more severe accident.
Emergency Operating Instruction 1A (E0I-1A), " Loss of Reactor Coolant," however, would prevent the IDIS from accomplishing its function. E0I-1A instructs the operator to " place the (postaccident)
H2 purge system in service . . . if the aontainment atmosphere reaches 35 by volume . . ." The postaccident purse system is designed as a backup to the redundant hydrogen recombiner system. The system consists of a single penetration in the primary containment wall equipped with two normally closed, remote manually operated isolation valves, one on either side of the containment wall, and one pneudatically operated annulus purge exhaust valve located within the annulus. With these valves open, a flow path is established 3
D 1
between the primary containment and annulus which will permit purging of the containment for hydrogen control subsequent to a LOCA.
The operation of this system, however, is based on an assumption of a maximum 5-percent metal-water reaction which will not produce the much larger volumetric quantities of hydrogen that a degraded core accident could. Therefore, allowing t.'is system to purge could lead to unnecessary release of radicactivity to outside the containment instead of making full use of the IDIS to burn off the hydrogen and keep as much radioactivity inside the containment as possible. The E0I-1A instructions, therefore, had to be modified to eliminate oneration of the H 2 purge system while the interim ignition system is installed.
Procedure Modification .
In addition to' revising E0I-1A to eliminate operation of the H E"#8' 2
system. TVA also had to modify EDI-0, "Immediate Actions and Diagnostics," to ensure initiation of the IDIS before it would be needed. Table 1, attached, lists the procedures that required modification in order to reflect che addition of the interim distributed ignition system.
Special Note :
l None of the modifications to procedures listed in Table 1 are to be incorporated into the procedures or implemented prior to NRC approval of the system and these proposed E0I changes.
l R-2
j APFENDIX R
! OPERATING PROCEDURES TABLE 1 PROPOSED EMERGENCY OPERATING INSTRUCTION CHANGFS E0I No. Change Text E0I-C Add to Section II.B 11. Energize power supply to U-1 controlled hydrogen ignition system i by closing breakers 10 , 11 ,
and 12 in Standby Lighting Cabinet LS-4 Tnear CCs surge tank).
EDI-1A Add to Section II.G 2. Ensure centrolled hydrogen ignition system is in service per E0I-0, section II.B.11.
E0I-1 A Delete from Section Place H2 purge system in service as II.00.5 follows: ,
- a. If the containment atmosphere reaches 3% by volume, place the H purge system in service per S$1-83.1.
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APPENDIX S TVA DEGRADED CORE INVESTIGATION PROGRAM In November 1979, in the course of the TVA Nuclear Program Review of the Sequoyah Nuclear Plant and the Report of the President's Commission en the Accident at Three Mile Y= land, the TVA Task Force on Nuclear Safety recommended that "TVA . . . (4) Study ways to contain larger amounts of hydrogen and to backfit feasible features into the Sequoyah design."
To fulfill that commitment, the Nuclear Engineering Branch of the Division of Engineering Design was charged with coordinating a study and making a report on the effects of postulated degraded core accidents on the Sequoyah Nucloar Plant and or design features to mitigate those effects. The study and draft report were completed on April 15, 1980, and included input from Westinghouse, Burns and Roe, Sargent & Lundy, Stone & Webster, and TVA's Office of Power.
The study began with a summary of the original design bases of t>e Sequoyah Nuclear Plant, including accident sequences, analytical tools used to calculate phenomena associated with the accidents, and features and equipment required to mitigate the accidents. Then, the Three Mile Island accident significance was assessed by investigating the accident phenomena and their ramifications on the
( current design bases for Sequoyah. An analysis was performed to compare the current risk posed by Sequoyah with Surry, a WASH-1400 reference plant. Accident sequences beyond the design bases were l
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identified and investigated. Since the calculational tools to analyze these accidents were of uncertain accuracy and limited availability, hand calculatiens were performed to scope the magt:ltudes of accident parameters. Information from the Battelle Columbus computer code MJLECH was then obtained to estimate the =agnitudes and rates of the parameters. The present margins in the containment safety systems for degraded core accidents were evaluated. Later, a number of cencepts for study were proposed to mitigate accident pheno =ena by preventing or minimizing hydrogen combustion (e.g., centrolled ignition, Halen inerting) or by increasing the containment capacity for overpressure events (e.g., filtered venting). Another analysis was performed to evaluate the risk posed by the Sequoyah Nuclear Plant, assuming the mitigating features were installed. Conceptual designs by TVA and others of the mitigations were also evaluated on the bases of physical effectiveness, technical feasibility, additional risks, reliability, cost, and schedule. The report of the study recommended that further investigation of both degraded core accidents and their mitigations be undertaken before commitments were made to specific mitigation devices. However, controlled burning of hydrogen and postacaident inerting with Halen were suggested as the most promising mitigations for further study. Nit rogen inerting .3as rejected after a thc rough study, mainly because of the er.treme hazards it would pose to personnel during the almost daily containment entries required for the inspection and maintenance of ice condenser and other vital equipment. Filtered venting was rejected because of the large radiation dose to the public that would result from venting the containment during a severe accident.
l CA L m _
To act on the recommendations of this report and to continue to fulfill TVA's Nuclear Program Review commitment, a Degrad" Core Task Force was established in June 1980. Organized in the Nuclear Engineering Branch, it is composed of seven nuclear, mechanical, chemical, and electrical engineers whose two-year mission is to further investigate degraded core accident sequences, their analyses, and their mitigations. The primary efforts in mitigation are to support the installation of igniters at Sequoyah to allow controlled burning of hydrogen and to evaluate the use of Halon to suppress hydrogen combustion. Other mitigation concepts such as catalytic reco=biners are also being investigated. ,The primary efforts in accident analysis are to develop MARCH-type code packages capable of realistic calculations of accident parameters. Since overconservative or arbitrary comouter models may be counterproductive to determining the need for mitigation or to the selection of optimum mitigations, code improvements will be made by the incorporation of realistic assumptions or physical models bench-marked to present and planned experiments. Special efforts are also being made to analyze the containment structural integrity for uniform and localized pressure effects and to analyze and qualify necessary equipment inside containment for accident environmental effects.
A compre'hensive effort related to all of the above work is the technical support of utility input during future NRC rulemaking on degraded core licensing issues.
HYD.2 c2
1 i
1 APPENDIX T CLASIX PROGRAM DESCRIPTION CFFS. ORE POWER SYSTEMS e
4 l
l l
I.
t l
1.0 Introduction As a result of the incident at Three Mile Island, Offshore Power Systems began the development of a computer program to investigate the response of an ice-condenser containment to a degraded core condition and the subsequent ignition of the hydregen released to the containment. Because of the compartmentalized nature of the ice condenser containment, an analytical model representing a number of volumes would be required. None of the programs available at the time represented anything other than steam and air. To adequately track the migration of hydrogen and eventually convert hydrogen and oxygen to steam and heat, each constituent must be individually considered. It was concluded that none of the existing programs ,
could be readily modified and a new program was required. The new program was designated CLASIX to reflect the analysis of Class IX ar,1 dents.
2.0 Analytical Model The analytical model selected for the CLASIX program is snown schematically in Figure T-1. The containment can be represented by up to six volumes interconnected with appropriate flow paths.
l A seventh volume is included for representation of a blowdown volume for venting the containment. Volume 1 in the figure characterizes the lower compartment in the .'olume of the containment below the divider deck and outside the ice condenser doors but exclusive of any further subcompartmentalization such l
l as Volumes 5, 6, and 7. Volume 5 characterizes the equipment and l instrumentation volume off the lower compartment. Volume 6 can l
represent the pressurizer shed and Volume 7 the steam generator l
l I T-1 l
dog house. Volume 2 is the single ice condenser volume in that an ice heat sink may be specified. The upper plenum of the ice condenser between the intermediate doors and the top deck blanket os t
is modeled as part of the upper compartment. Volume 3 represents (QlEn the upper compartment and generally has the largest net free volume.
The flow paths from Volume 1 to 2 and Volume 2 to 3 have representation of the lower inlet doors and intermediate docrs respe tively. The flow path from Volume 3 to 4 permits specification of a rupture disc, submergence of the discharge under water and specification of a length of water which must be expelled before gas discharge may begin.
Also shown in Figure T-1 is a recirculation fan. The ran say be turned on and off at optio'nal times during the transient. This 7
original fan model took suction'from the upper compartment and discharged into Volumes 1, 5, 6, and 7 in optional proportions.
Such a flow pattern is not entirely correct for Sequoyah since the hydrogen skimmer system actually would draw flow from Volumes l
1 5, 6, and 7 and discharge into Volume 1. The code has since been r
modified to permit specification of fan flow from any compartment to any other compartment with up to nine fan flow paths.
However, since the skimmer system was designed to prevent the long "+ra buildup of hydrogen during design basis accidents and constitutes only 0.55 percent of the tot-1 air return ran flow in l
Sequoyah, it is felt that this flow' pattern change is negligible l
O l
i and would not significantly affect the results or conclusions of Q 12.'o the original degraded core transient analysis.
The most significant capabilities of the CLASIX program are listed in Figure T-2. The vent, ice condenser, recirculation fan and the ice condenser doors were discussed above. Each of the non-condensible gases and the steam inventories in each compartment are calculated individually. The steam properties are from the computerized A3ME steam tables and the steam may be saturated or superheated. Sprays are* optional in all volumes except the ice condenser. The sprays are individually controlled by tables specifying time, flow rate, temperature and heat transfer c'.efficients. The spray drop size and fall time may be varied.
Hydrogen, nitrogen and heat addition tables are available in Volumes 1, 6, and 7. There are break flow tables available in the same volumes anc the break flow may be subcooled, in the saturated region or superheated. Burning control of the hydrogen is available in all compartments.
The burn control parameters are shown in Figure T-3 When the volume percent of hydrogen exceeds the value for ignition, r
i burning is ast.umed to occur provided that the volume percent of oxygen exceeds the value for ignition.
m
- t
Once the hydrogen is ignited, the flame will proceed at some velocity thus establishinE a burn time for the hydrogen in tb-volume and a propagation delay time or elapsed time for the flams to reach an adjoining volume. Since the ignition source may require a relatively high hydrogen concentration, the flame may propagate to adjoining compartments at much lower concentrations.
Depending upon the hydrogen concentration at ignition, only a 4
fraction of the hydrogen may burn. Finally, because of depletion of oxygen below a minimum concentration, the flame may be extinguished.
3.0 Program Description A flow diagram of the CLASIX program is shown in Figure T-4. The
. program has a restart capability. During any analysis with CLASIX, the instantaneous conditions can be written, periodically, to a restart file. By reading this file in a subsequent run, the transient c&n be picked up at the time of the restart file and the transient continued with altered input. The first decision in the program is whether this is a restart run or the start of a new transient. Whether or not it is a restart run, a complete input edit is written. Once this edit is complete, the finite difference integration cycle is initiated.
The output options include a short output which prints the total pressures and the temperatures of all compartments on a single line. A long form output is a detailed printout of containment conditions and consists of a full page of information. A restart file may also be written and all three forms may be intermingled.
mh
After each printout a test for completion of the transient is made, Based on the flow path parameters and the differential pressures, a volumetric flow rate is calculated. Then, based on the source volume conditions, the individual constituent = ass flow rates and enargy flew rates are determined.
The spray flow rate, temperature and film coefficient of heat transfer are linearly interpolated frca input tables. There is no difference in the treatment of the' containment spray water prior to, during, or after a hydrogen burn. If the inlet spray temperature is above the saturation temperature corresponding to Q l 2.o.
the total pressure of the compartment, a sufficient quantity of Ql2d the spray will vaporize to bring the spray into thermodynamic equilibriu with the comparment. This is assumed to occur instantaneously. The droplet size, mass rate of flashicg and the energy addition rate to the gas phase are thus determined.
Knowing the droplet size, temperature, heat transfer coefficient and the residence or fall ties of the spray, heat transfer rates I
to the gas can then be calculated. If the initial spray
! temperature is below the saturation temperature, only heat transfer rates need to be calculated. During those periods of t
time when the gas temperature is above the saturation I
temperature, heat transfer will result in vaporization at the l surface of the droplet so that there is a corresponding mass and energy addition rate to the gas phase of the volume.
t P_e
[
Heat transfer takes place according to Q=H.A. T where Q is the rate of heat transfer, H is the film coefficient of heat transfer, A is the surface area of the droplet and T is the temperature difference between the compartment atmosphere and la the droplet. The heat transfer may be positive or negative.
()12f
, During evaporation, the change in mass of the droplet is calculated and used in evaluation of the surface area for heat transfer consideration. The fall time or total evaporation of the droplet is utilized to terminate the heat transfer to a given mass of spray. Thus, the amount of spray water in the air at the time of initiation of the burn is considered.
The heat addition rate to the volume is linearly int (rpolated from the input tables. Nitrogen and hydrogen addition rates and their corresponding temperatures are also interpolated from the input tables. The hydrogen and~ nitrogen addition tables in CLASIX require three entries for each point. These entries are
- tme, flow rate and temperature. (See Table U-2.) Using the sl cific heat at constant pressure, the energy addition rates are calculated. The brea.k flow rate and energy flow rate is linearly interpolated from the input tables. If the break flow enthalpy is above the saturation enthalpy corresponding to '.ba total pressure of the break volume, the rates are added dirtctly to the gas phase. If the break flow enthalpy is below the saturatiot enthalpy, the break flow is assumed to flash instantly to water and vapor at the saturation enthalpies giving corresponding flow and energy rates.
m.C
, The flow entering the ice condenser flows over the ice, transferring heat to the ice, condenses steam and melts some of the ice. The current version of CLASIX retains the melted ice in tha single-volume ice condenser node along with the remaining ice in the bed. Thus, no significant free volume increase in the ice C)llb condenser node is concidered. The CLASIX model for heat transfer is the same as that in LOTIC, but the correlations are based on vapor content in air. Therefore an air equivalent of the nitrogen, hydrogen, oxygen mixture of non-condensibles must be approrimated. The ice bed heat trans'rer correlation is based on the ratio of the mass of condensibles to the mass of non-condensibles with a range from 10,000 Btu /hr-ft -F for pure steam to 72 Btu /hr-ft -F for pure non-condensibles. This latter (kI2-j value is considered conservative and the former value is comparable to published values in standard references for similar conditions.
i Based on the constituents in the volume and the input burn control parameters discussed above, it is determined whether or not ignition occurs. If ignition occurs, clocks are initialized to calculate propagation delay time to adjacent compartments; burn rates and amounts of hydrogen to be burned are also i
calculated. If the propagation time from some adjacent compartment has expired, conditions are checked to see *f the criteria for propagation are satisfied. If so, tte same calculations as for ignitios. ..re completed.
-n
_ _ ._ ,_- _ . _ . . __ m _ _
\
A maximum time step from the input is determined for comparison with an internally calculated stable time step.
With all the rates of change calculated, equations for the conservation of total energy in the compartment and the conservatien of the masses of the individual constituents are written in finite difference integration form. Knowing tha total I
internal energy of the compartment and specific volume of the vapor, a temperature and partial pressure of the vapor can be assumed and steam properties determined from the steam tables.
Based on these values the total energ'y can be calculated and compared to the actual values. Iteration leads to a ccavergence within acceptable limits.
With the converged results, parameters are updated for the next time step. The finite difference integration is continued until the transient has been completed.
4.0 Summary The long form printouts provide sufficient information so that hand calculations of summations were used to verify conservation of mass and energy. Comparison of printed output with the ASME steam tables confirms that the steam properties are properly evaluated by the iterative scheme discussed above.
l t
T.8 t
li VENT VOLUME
/
3 r 3 UPPER COMPARTMENT k 2 ^"
1 d L 2
ICE CONDENSER g
1 1
LOWER COMPARTMENT e--
,= __
l
/ t h~ t 7 l 6 5 STEAM i
PRESSURIZER DEAD I l
GENERATOR VOLUME
! ENDED I
- VOLUME l 1
+ VOLUME
+ i 1
L_-_____tL______1 ~
I CLASIX MODEL FIGURE ~1 i
(SUPERSELE3)
T-9
4 CLASIX CAPABILITIES
- 1. VENT FROM UPPER COMPARTMENT -
- 2. ICE CONDENSER
- 3. RECIRCULATION FAN c
4.
DOORS-LOWERINLETAADINTERMEDIATE
- 5. INDIVIDUAL REPRESENTATION OF 0 ' H , N AND H O 2 2 2 2 6.
SATURATED AND SUPER-HEATED STEAM
- 7. SPRAYS
- 8. H2, N AND HEAT ADDITIONS 2
~
- 9. BREAK FLOW
- 10. BURN CONTROL FIGURE 2 T-10
BURN CONTROL
- 1. v/o H2 IGNITION
- 2. v/o H 2 PROPAGATION
- 3. o/o H2 CONSUMED
- 4. v/o 02 IGNITION i
- 5. v/o 02 SUPPORT COMBUSTION
- 6. PROPAGATION DELAY TIME
- 7. BURN TIME 1
FIGURE 3 T-u
START YES NO RESTART o
READ RESTART FILE READ INPUT A
READ INPUT CHANGES I __
I WRITE INPUT EDIT 1
11 WRITE OUTPUT g
STOP7 1I CALCULATE MASS AND j i ENERGY RATES BETWEEN COMPARTMENTS i
E CALCULATE SPRAY FLOW RATE AND HEAT TRANSFER RATES d
HEAT ADDITION RATES,
)
l HYDROGEN ADDITION RATES (
14ITROCEN ADDITION RATES FROM TABi.ES I
BREAK FLOW AND ENERGY FROM TABLES I
CALCULATE FTASHING ands k .
{ HEAT TRANSFER RATES
( TO ICE. MELTING j RATES r
i l 2 l
l l
FIGURE 4 T-12 i
2 1r DETERMINE IGNITION, BURN RATE, PROPAGATION I
DETERMINE MAXIMUM TIME STEP FROM INPUT
' SIMULTANEOUSLY CALCULATE STABLE TIME STEP & ITERATE TO CONVERGED STEAM PROPERTIES AT END OF TIME STEP BASED ON RATES o
UPDATE MASSES, INTERNAL ENERGY & PRESSURES OF EACH COMPARTMENT
+
I I CHECK FOR DEPLETION OF RYDROCEN OR OXYGEN IN COMPARTMENTS
~WHERE BURNI!iG EXISTS i
l l 'I l
1 4
l l
l 1
FIGURED (Con't.)
i
-T-13
i l
APPENDIX U
SUMMARY
OF ANALYSES OF ICE CCNDENSER CONTAINMENT RESPONSE TO HYDROGEN BURN TRANSIENTS OFFSHORE POWER SYSTEMS 1
I
1.0 Introduction A series of analyses have been performed to study ice condenser containment response to hydrogen burn transients for an accident sequence similar to the TMI-2 accident. The particular sequence stucc ed is that designated as S2D in WASH-1400. This is a small break loss of coolant accident (LOCA) accompanied by the failure of emergency core cooling injection.
The S D transient may be divided into three phases. The first phase is the period from accident initiation to the beginning of hydrogen generation. This period is 'similar to the small break LOCA transient. Existing calculational techniques such as the Westinghouse Long Term Ice Condenser Containment Code (LOTIC) are currently used for analysis of the containment response during this period.
Tro r sond phase is the period from hydrogen initiation through the end of core melt. This phase, which progresses through various stages of degraded core conditions, is representative of the TMI-2 accident. Analyses of the containment response during this phase of the transient are performed usiE3 the CLASIX computer program which was developed by Offshore Power Systems.
These CLASIX analyses are the subject of this report and are discussed in the following pages.
The third phase of the transient is the period following vessel melt.
This period is beyond the scope of these analyses.
i U-l t _ _ _ _
J 2.0 CLASIX Base Case Analysis The parameters for the base case analysis were selected by the consideration of (1) available experimental data on hydrogen burn characteristics, (2) the potential effect of the containment engineered safeguards systema parameters, and (3) the containment geometry. The mass and energy releases from the break (steam, hydrogen, and fission prcoucts) were based on calculations by Battelle Columbus Laboratories using the MARCH code. These parameters are summarized in Tables 1 through 3 The conditions inside the containment prior to the onset of hydrogen generation, including subcompartment volumes, temperatures, air and steam partial pressures, and ice mass, were determined from LOTIC analyses and the MARCH generated blowdown. These paramsters are summarized in Table 4.
For the base case, it was assumed that ignition would occur at a hydrogen concentration of 10 percent by volume (V/o). Consistent with ignition at 10 /o, the flame was assumed to propagate at 6 feet per second. Propagation was assumed to occur to any connected volume with a hydrogen concentration of at least 10
/o. Ignition in and propagation to any compartment with less than 5 */o oxygen were suppressed. Consistent with ignition at 10 */o, it was assumed that complete combustion occurs in any compartment in which a burn is initiated either by ignition or by propagation. These parameters are summarized in Table 5.
Burn times were established by assuming that, when the above Q l 3 l=
conditions for either ignition or propagation in a given U-2
ccmpartmer.t were met, burning proceeded through a characteristic length selected to represent the flame path in that compartment.
Dividing the characteristic length by the flame speed results in a burn time which is used as input to the CLASIX program.
1 Internal to the program, once the conditions for ignition are 1
satisfied, the number of pounds of hydrogen to be burned is divided by the burn time to arrive at the rate of burn which is used as a constant until the hydrogen is censumed or the oxygen is exhausted.
For the base case, ignition was assumed to occur at a single point within each compartment: the center of the lower compartment, the circumference of the dead-ended region, or at the top of the upper cpapartment. These characteristic lengths are best-estimate flame paths for burns that originate wi, thin a compartment. Considering the number and location of igniters, ignition by two sources in close proximity would have a neglible effect on the burn rate. And although CLASIX is based on the assumption of uniform mixing in each volume, there will be some non-uniformity in the distribution of the constituents as well as some nonuniformity in charactaristics of the ignition sources, making the probability small that two igniters, a significant distance apart, would cause ignition simultaneously. Even if ignition occurred simultaneously at two points at the extreme distance within a compartment, the effect would be the same as
, doubling the flame speed. This was studied in the sensitivity analysis (see cases JV914 and JV915). Increasing the flame speed l by a factor of two resulted in an increase of 1.5 psi in the peak TT S
pressure within the containment. The base case characteristic lengths are conservative for burns in a compartment that were propagated from an adjacent once since the flame front must traverse the entire length of the compartment. t g g6
".*he spray system was modeled to provide a constant flow of 6000 gallons of 125 F water per minute to the upper compartment. The spray drops were assumed to have a diameter of 680 microns, a fall time of 10 seconds, and a heat transfer coefficient to the upper compartment atmosphere of 20 BTU /hr ft F. The spray system is automatically initiat' i 30 seconds after the containment reaches 3 psig. In t.:is transient, spray initiation occurred prior to the beginning of the CLASIX analysis. The spray parameters are summarized in Table 6.
l l
The fan system was modeled to provide a constant flow of 80,000 cubic feet per minute from the upper compartment to the lower compartment. About 0 55 percent of the fan flow was directed to the dead-ended region of the lower compartment to represent the hydrogen skimmer system. (See Section T-2.0.) The fan system is initiated 10 minutes after the containment reaches 3 psig. In this transient, ran -initiation occurred prior to the beginning of l the CLASIX analysis. The fan parameters are summarized in Table 6.
The fan system flow rate is based on whether the fans are assumed O on or off and is used as a constant regardless of transient Gl2 c conditions. An air return fan head / flow correlatior, will be U-4 i
l
incorporated into the CLASIX program and is scheduled for completion in the first quarter of 1981. In the interim, the effects of transient pressures during both lower and upper compartment burning have been qualitatively evaluated.
During a bur- in the lower compartment, the increased lower compartment pressure would tend to decrease the air return fan flow.
However, just after completion of the burn, the relatively cold air from the upper compartment entering the lower compartment would tend to cool the lower compartment and decrease che lower compartment pressure and consequently tend to increase fan flow. These effects only occur during and shortly after a burn in the lower compartment and tend to be offsetting.
Considering that the duration of the burn in the base case is of the order of seven seconds compared to more than 100 seconds between curns, the total effect is believed to be small. Since a decrease in total fan flow by a factor of two in the sensitivity studies (see case JV902) resulted in an increase in the peak pressure oi* only one psi, the anticipated effect of intercompartmental pressure differentials during lower compartment burning on air return fan flow, and consequently on peak pressure, is believed to be negligible.
Although including a fan head / flow curve may have a negligible effect on lower compartment burns, a more realistic fan flow could have a significant beneficial effect on pressure response to an upper compartment burn. The pressure increase cannot be relieved through the ice condenser and consequently has a much higher peak value and slower rate of decay than a burn in the TV C
lower compartment. Inclusion of a fan head / flow correlation will increase the flow to the lower compartment and decrease the peak pressure and increase the rate of decay in the upper compartment.
Consequently, the present =ethod of analysis for Q 12.c upper compartment burning is conservative.
The ice condenser lower inlet and intermediate deck doors were modeled based on the door representation in the Westinghouse Transient Mass Distribution (TMD) analyses for ice ccndenser plants.
Both types of doors were modeled to act as check valves, preventing reverse flow. I sign data for the ice condenser doors are available which indicate the differential pressure required to fully open the doors. Using this as input to the CLASIX program, it is assumed that the angle of the door opening is proportional to the pressure force on the projected area of the door. In equation form, O 0 AP Cos 0 AP7 Cos 0 7 Q 12 d where 9 is the angle of door opening, AP is the differential press'ure and the I refers to reference conditions. The lower inlet doors were restricted to a maximum opening of 55 degrees for this small break transient. Since relatively low pressures are required to fully open the _ lower inlet and intermediate deck doors in Sequoyah, even relatively large errors in the l approximation will have negligible effect on the pressure response of the containment. The door parameters are summarized g in Table 6.
U-6
Flow parameters for the flow paths between the various subcompartments are based on typical ice condenser containment geometry and are consistent with similar parameters used by Westinghouse in TMD calculations. Flow path parameters are summarized in Table 7.
The results of the base case CLASIX analysis, identified as JV900 in Tables 8 and 9, indicate the hydrogen will be ignited in a series of nine burns in the lower compartment over a period of about 2300 seconds beginning about 5,000 seconds after accident initiation. One of the burns propagates into the ice condenser.
Each burn in the lower compartment consumes about 100 pounds of
, hydrogen and the burn in the ice condenser consumes about 37 pounds of hydrogen, giving a total burn of about 900 pounds of hydrogen. For the first burn, calculated peak pressures were 26.5 psia in the lower compartment and 28.5 psia in the ice condenser and upper compartment, with a pre-burn pressure of 22.5 ps';. Subsequent burns resulted in successively lower pressure peaks. Peak temperatures of 2200 F, 1220 F, and 150 F were calculated in the lower compartment, ice condenser, and upper compartment, respectively. Only small differential pressures occur across containment structures during the transient.
l As a result of the action of engineered safety features, such as the ice condenser, air return fans and upper compartment spray, the pressure and temperature peaks were rapidly attenuated i
between burns with pressure returning to the pre-burn value approximately two minutes after the burn. At the end of the l
l ,, -
b transient, 7080 seconds after accident initiation, 650 pounds of hydrogen remained distributed in the containment at a concentration insufficient for ignition and 300,000 pounds of ice remained in th. ice condenser.
3.0 CLASIX Sensitivity Studies To determine the effects of ignition criteria and safeguards performance on containment response to hydrogen transients, a number of sensitivity studies were performed on the parameters that have been judged to have the greatest potential impact.
In the first sensitivity case, identified as JV901 in Tables 8, 10, and 11, the hydrogen concentrations for ignition and propagation were reduced from 10 */o to 8# /o. Based on experimental data, the. burn fraction was also reduced from 1.0 to 0.5. All other parameters were the same as those used in the base case. In this case there was a series of seventeen burns in the lower' compartment, eight burns in the ice condenser, and one burn in the upper compartment. Although the number, magnitude, and distribution of the burns varysconsiderably from the base case, the total amount of hydrogen burned and the peak containment pressure during the transient do not vary appreciably from the base case. Peak temperatures in the lower compartment and ice condenser are considerably lower than in the base case due to the smaller magnitude of the individual burns in these compartments.
In the second sensitivity study, identified as JV913 in Table 8, the hydrogen concentration for pecpagation was reduced from na
l0 /o to 8 */o while the concentration for ignition was kept fixed at 10 */o. For consistency with experimental data. the 8 */o burns were restricted to a 50 percent burn fraction while the 10 */o burns had complete combustion. All other parameters were identical to those used in the base case. In this case there was a series of three burns initiating in the lower compartment. All three burns propagated to the upper compartment and tto of the burns propagated to the ice condenser. Again, although the number, magnitude, and distribui.icn of the burns varies considerably from the previous cases, the peak pressures and temperatures are similar to those calculated above.
In the third sensitivity case, identified as JVTC4 in Table 8,
~
the hydrogen concentration for propagation was again reduced from 10 */o to 8 /o while the concentration for istlition was kept fixed at 10 */o. For conservatism, all burns were assumed to have complete combustion. In addition, the fan flow rate was i
reduced from 80,000 cubic feet per minute to 40,000 cubic feet per minute. The results for this case are similar to the JV913 I
case, with an increase peak pressure and temperature in the upper l compartment due to the greater magnitude of the upper compartment burn.
In the fourth sensitivity case, identified as JV914 in Table 8, the flame speed was increased from 6 feet per second to 12 feet 7
per second. All other parameters were identical to those used in the base case. The results of this case were very similar to l
it O I_
4 those for the base case.
In the fif th sensitivity case, identified as JV915 in Table 8, the flame speed was increased frem 6 feet per second to 12 feet per second, the hydrogen concentrations for ignition and propagation were reduced from 10 V/o to 8 #/o, and the burn fraction was reduced from 1.0 to 0.5. All other parameters were identical to those used in the base case. The results of this case are similar to the first sensitivity caso.
The flame speed of ' feet per second assumed for the base case is Ql3C consistent with reported data for laminar flame propagation. The-12 feet per second used for the two sensitivity cases agrees with upper-bound estimates for laminar burning.Section IV.C refers to Rockwell/ Atomics International experiments reported in reference IV-2 concerning the dramatic effects of pressure reduction when water sprays are used during hydrogen burning.
(See Figure U-la.) These experiments also measured effective flame speeds of approximately 5 feet per second during the ,
nonspray tests and approximately 50 feet per second during the spray tests. The increase was attributed to turbulence created by the sprays. Conditions created by the containment sprays at Sequoyah would tend to increase the effective flame speed from laminar into the turbulent range. However, the reported experimental result of approximately 50 feet per second probably overestimates the effect at Sequoyah since the Rockwell experimental droplet populatior. density was about an order of 1
J-10 L
Uh hh, eg 7.a l b]
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ta
~ i' I I I l l l l i 8 7,f x.
20 -
T
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9% H2 f7 i
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8 15 - V% WITh 78 gom H2O 13.6 I
% V.
,3.2 g N 13.2 "'
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! 0 2 4 6 8 10 12 14 16 18 ==
TIME (sec) l l l l l l l l l 20 -
l l
1 11% H2 3
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\ WITH 74 gom H2O
@15 -
f 13.8 N % __. _ -. - - - _ _ _ _
13.0 10 ' b ! ! !
0 - 2 4 6 8 to 12 14 16 18 TIME (sec) 4008-4518 Figure U-la Shock Tube Pressures for Flame Test at 1.0 atm, With and Without Water Spray
- AI-73-29
- 3 n
magnitude greater than in the Sequoyah upper ccmpartment and the long, narrow shock tube geometry used in the experiment has a much greater L/D ratio than a representative space inside the Sequoyah containment where burning might occur.
At any rate, the two flame speed sensitivity cases described above showed that doubling the flame speed (thus halving the associated burn times) resulted in very moderate pressure increases. It has been estimated that increasing the flame speed in CLASIX to 50 feet per second would not result in large pressure increases since the corresponding burn times would have been decreased by less of an increment than in the two sensitivity cases. Since the CLASIX code currently lacks structurr.1 heat sinks, burning in the lower compartment (that occurs during the best-estimate base case) would be relatively unaffected by a decrease in burn time. The major impact of a reduced bu"n time would be fel't for a situation involving burning in the upper compartment since the time for heat transfer to the containment spray water would be shortened. This effect any be estimated by examining the sensitivity case where no spray operation was assumed (JV910). Even for this limiting case, the transient pressure only peaked at approximately 35 psig, substantially less than the containment yield capacity of 46 psig. In addition, the CLASIX tabular treatment of containment spray performance neglects any beneficial physical effects of deflagration pressure suppression (see Section IV.C) other than simple heat transfer. Some of these additional etfects are aesponsible for the marked reduction in pressure response as UJ2
i indicated in the referenced experiments. (See Figure U-la.) Ql3 c In the sixth sensitivity casa, the fan flow rate was reduced from 80,000 cubic feet per minute to 40,000 cubic feet per minute with all other parameters identical to the base case. The results of this transient are almost identicaA to those fcr the base case.
In the seventh sensitivity case, the fan flow rate was reduced to zero in the CLASIX part of the analysis (i.e., from the beginning of H2generation). All other parameters were identical to the base case. In this case, steam and hydrogen from the break push air out of the lower compartment reducing the oxygen supply below V
the minimum 5 /o required for burn initiatien. As the transient-cantinues, hydrogen accumulates in the upper compartment and eventually ignites there.
This burn propagates to the ice condenser. The combined upper compartment and ice condenser burns cause a redistribution of the containment atmosphere, adding oxygen to the lower cempartment and hydrogen to the upper compartment. With the ar.dition of oxygen to the lower compartment there is an independent lower compartment ignition which also propagates to the ice condenser and forces more hydrogen into the upper compartment in which hydrogen is still burning. The net result is a burn of approximately 1200 pounds of hydrogen in total, of which 860 pounds burn in the upper compartment. The peak calculated pressures are 92 psia in the upper compartment, 86 psia in the ice condenser, and 46 psia in the lower ccmpartment.
?? 11
In the eighth sensitivity case, the ice condenser drain temperature was increased from 32 F to 132 F with all other parameters identical to those used in the base case. The results of this case are almost identical to the base case results except for the ice remaining at the end of the transient. This indicates that drain temperature is important to the ice '
emdenser efficiency but not to its effectiveness.
In the nintn sensitivity case, the initial ice mass was reduced by 1.17 x 10 pounds. Thus the ice mass input to the CLASIX part of the analysis was reduced from 1.67 x 10 pounds to 5 x 10 5 pounds with all other parameters identical to those used iri the base case. This is a non-mechanistic study to determine the total effect of ice on a hydrogen transient. In this transient, ice melt out ccourred during the second of a series of seven burns in the lower compartment'. The peak containment pressure was about 10 psi higher than in the base case.
The remaining sensitivity cases were performed varying spray parameters. Since spray operation is much more important for upper compartment burns than for lower compartment or ic'e condenser burns, these variations were performed for the upper compartment burn in the first sensitivity case, i.e. , the case with ignition and propagatior at 8 */o hydrogen and 50 percent burn fraction, identified as JV901.
i.
In the first spray sensitivity casa, the spray heat transfer U-ll+
coefficient was reduced from 20 Btu /hr ft2 F to 2 Btu /hr ft F.
In the second spray sensitivity case, the spray temperature was increased from 125 F to 180 F. In the third spray sensitivity case, the spray flow rate was increased from 6000 gpm to 4700 gpm. In the fourth spray sensitivity case, the spray flow rate was increaseJ from 6000 gpm to 9400 gpm. In the fifth spray sensitivity case, the spray flow was reduced to r.ero in the CLASIX part of the analysis. In the sixth spray sensitivity case, the spray drop diameter was reduced from 680 microns to 400 microns. In the last spray sensitivity case, the spray drop diameter was increased from 680 microns to 1000 microns. The results of the spray system sensitivities are summarized in Table l
- 11. These results indicate that while spray temperature and some minimum flow rate are important, the remaining parameters are relatively unimportant for containment response to hydrogen burn transients. In the spray temperature sensitivity case, the increased spray temperature resulted in an increased ambient I pressure. The increased pressure required additional hydrogen to achieve the hydrogen concentration required for ignition so that i
l a greater quantity of hydrogen was consumed in each lower compartment burn, and an upper compartment burn did not occur.
An adt tional special analysis was recently performed to determine the amount of hydrogen which could be tolerated to burn h in the upper compartment using conservative assumptions including a pressure limit of 46 psig for the containment shell. This i-amount of hydrogen is a function of containment conditions peier M
i . ,
to the burn, containment safeguards operation and concurrent events as well as the flame speed and containnerit y assure.
Assuming (1) a pre-burn atmosphere typical of the S2D transient in which the upper compartment is at 8.6 psig and 122 F, (2) a
, conservative constant air return fan flow of only 80,000 cfm l
j from upper compartment to lower compartment, (3) 9,400 gpm upper compartment spray, (4) no concurrent burns elsewhere in the containment, and (5) a flame speed of 12 feet per second over a 90-foot length; a burn of at least 450 pounds of hydrogen (8.6 percent by volume) could be tolerated in the Sequoyah containment upper compartment without exceeding 46 psig. This calculation conservatively does not include the effects of the containment walls acting as heat sinks, additional fan flow due to tne
, pressure differential between the upper and lower compar ?snts, and additional spray flow from the RHR spray system.
Ql5 4.0 Summary and Conclusions The calculations described above represent the first attempt to perform a realistic assessment of hydrogen transients in an ice i condenser containment. The geometry and engineered containment l
safeguards parameters used in the study were based on the i Sequoyah containment design. Therefore the results are directly applicable to the Sequoyah plant.
l l
The results of this study indicate that over a wide range of cases, the Sequoyah containment pressure response to hydrogen transients would not cause containment failure.
Future analyses may be performed to expand the range of l
U-16 l
L
parameters studied, such as the burn criteria in the S D 2
transient, and to extend the analyses to other transients.
em O l
l l
l I
I l
i U-17
TABLE 1 MARCH REACTOR COOLANT MASS AND ENERGY RELEASE RATES S20 SEQUENCE TIME MASS RELEASE RATE ENERGY RELEASE RATE (sec) (1bm/sec) (BTU /sec) 0.0 197.167 116722.67 2172 130.500 109728.00 2478 44.850 52295.10 3180 53.533 65471.27 3804 34.817 42615.60
- 4428 21.400 28419.20 4752 48.417 55582.33 5700. 19.417 21824.33 6012 14.067 - 15825.00 6960 5.253 5988.80 7062 4.718 5388.34 7206 4.060 It 4693.36
+ re U-18
TABLE 2 MARCH HYDROGEN GENERATION RATES AND TEMPERATURES S20 SEQUENCE TIME MASS RELEASE RATE TEMPERATURE (sec) (ibm /sec) (oF) 0.0 0.0000 61.24 3480 0.0000 61.24 3804 0.0413 66.56 4116 0.2600 1582.29 4428 0.7400 795.45 4752 1.0700 771.47 3 5700 0.4300 611.53
_ . . 6330 0.2233
~
555.39
- 6648 0.16b0 E35.22 6960 0.1167 519.43 8070 0.0367 519.43 i
I i
f U-19 L
TABLE 3 MARCH FISS10!! PRODUCT EtlERGY RELEASE RATES S2D SEQUENCE TIME ENERGY RELEASE RATE (sec) (BTU /sec) 0.0 0.0 3810 0.0 4116 1803 4428 4800 4752 6708 5376 7000 7080 7135 9
l U-20
TABLE 4 SUBCOMPARTMENT PARAMETERS
- CLASIX BASE CASE ANALYSES S20' SEQUENCE LOWER ICE UPPER DEAD ENDED COMPARMENT** CONDENSER COMPARTt1ENT+ , REGION 3
VOLUllE (ft ) 3.03 X 10 5
7.85 X 10 4 6.98 X 10 5 7.81 X 10 4 02 PRESSURE (psia) 2.53 3.70 3.62 3.62 N2 PRESSURE (psia) 9.56 13.98 13.67 13.67
? -H2O PRJSSURE (psia) 6.10 0.40 0.90 0'.90
% .* - . s . > i, i. ; .i y TEMPERATURE (DF) 171 75 98 98 ICE MASS (1bm) 1.67 X 10 6 2
ICE HEAT TRANSFER AREA (ft ) 2.02 X 10 5 BASED ON LOTIC RESULTS INCLUDES HELTED OUT PORTION OF ICE CONDENSER l + INCLUDES ICE CONDENSER UPPER PLENUM i
TABLE 5 HYDROGEN BURN PARAMETERS CLASIX BASE CASE ANALYSES H2 V/o FOR IGNITION *10 H2 V/o FOR PROPAGATION 10 H2BURN FRACTION 1 02 V/o FOR IGNITION 5 MINIMUM 02 v/o TO SUPPORT COMBUSTION 0
- PROPAGATION DELAY TIME LC-IC 12 sec IC-UC 5.5 sec UC-LC b -
- EXCEPT IN THE ICE CONDENSER; ASSUMED NO IGNITION SOURCES AVAILABLE
- BASED ON A FLAME SPEED OF 6ft/sec d
U-22
TAB:' i SYSTEM PARAMETERS CLASIX BASE CASE ANALYSES Spray System Ilow Rate 6000 gpm i Temperature 125 F Drop Diameter 680 Fall Time 10 see Heat Transfer Coefficient 20 BTU /hr ft2 og Initiation Time 52D during LOTIC Air Return Fans
- Flow Rate 80000 cfm i t. Fraction of Flow to DE Compartment. 0.0055 Initiation Time S2D duringLOTIC 1
Ice Condenser Lower Inlet Doors MaximumOpeningAngle(dogrees) 55 Differential Pressure for Maximum Opening 0.0206 psi Maximum Flow Area 840 ft2 i
i Ice Condenser Intermediate Deck Doors Maximum Opening Angle (degrees) 85 Differential Pressure for Maximum Opening 0.493 psi Maximum Flow Area 982.47 ft 2 i
U-23
- n.
1 I
- TABLE 7 1
FLOW RATE PARAMETERS
- i CLASIX BASE CASE ANALYSES 4
LC-IC IC-UC UC-LC DE-LC 2
Flow Area (f t ) .** **
2.2 108.6 Flow Loss Coefficient . 1.12 2.26 1.5 3.0 c:
- Based on TMD Models for ice condenser containments
- Function of door opening
TABLE 8
SUMMARY
OF RESULTS S20 BURN SENSITIVITY STUDIES JVTC4 JV900 JV901 JV913 'S20 JV914 JV915 Base Case (10/8%) (10/8%)
(10%) (8%) J100/50% Burn) (100% Surn) (12 fps LC) (12 fps UC)
- Burns LC 9 17 3 3 9 IC 1 8 2 2 1 UC 0 1 3 1 0 Magnitude of LC s100 d5 $115 90 $100 Burns (Ibm) IC 37 16-45 25-49 80 48
.p UC -
200 s230 430 -
Total H2Burned (1bm) s900 s1050 1100 950 900 H2 Remair.ing (lbm) 4 50 $ 500 450 600 650 Peak Temp. (DF) LC s2200 $1200 $1900 2100 2100 IC *1200 s 700 $ 630 1500 1370 UC $ 150 s 260 s 275 480 160 Peak Press,(psia) LC s26.5 s28.5 29 34 27 29 IC $28.5 s28.5 29 44 30 29 UC 28.5 $30.5 33 53 29 36 5 5 5 6 5 Ice Remaining (lbm) 3x10 3.2x10 4.5x10 5x10 3.2x10 Figures 1-8 9-16 17-24 25-32 33-40 41-48 t
TABLE 9
SUMMARY
OF RESULTS S2D FAN AND ICE CON 0ENSER SENSITIVITY STUDIES JV900 JV902 JV903 JV904 JV905 (Base Case) (1 Fan) (no Fan) .(Drain Temp)
(Less Ice)
- Burns LC 9 9 1 7 9 IC 1 1 2 0 1 UC 0 0 1 0 0 Magnitt le of LC $100 s100 Burr.s';bm) IC sl30 s100 37 s 60 -
% 39 UC - -
Tot'l H 2Burned (1bm) s900 s900 s1200 s850 s950 H2 Remaining (lbm) s650 s650 350 s700 s600 Peak Temp. (OF) LC 2200 s2200 2370 s2400 s2000 IC 1200 sl350 2583 s2000 s1270 UC 150 s 160 1088 s 270 s 150 Peak Press. (psia) LC 26.5 s26.5 46.4 s41 IC s26.5 28.5 s26.5 86.4 s41 s28.5 UC 28.5 s29.5 92.4 s41 s26.5 5 5 6 Ice Remaining (1bm) 3x10 3.7x10 6.3x10 0.0 8.3x10 5 Figures 1-8 49-56 57-64 65-72 73-80 t
TABLE 10
SUMMARY
OF CASES CLASIX SPRAY PARAMETER' SENSITIVITY STUDIES Case Number -JV901 JV905 JV907 JV908 JV909 JV910 JV911 JV912 Spray Flow Rate (gpm) 6000 6000 6000 4700 9400 0 6000 6000 Droplet Size /t 680 680 680 680 680 -
400 1000 Heat Transfer to Drop 20 2 20 20 20 -
20 20 (BTU /hr fta F)
Spray Temperature (F) 125 125 180 125 125 -
125 125 9
-4 .
Figures 9-16 81-88 89-96 97-104 105-112 113-120 121-128 129-136
6 TABLE 16
SUMMARY
OF RESULT 5 CLASIX SPRAY PARAMETER SENSITIVITY STUDIES Case Peak Pressure Peak Tempera ture (psia) (OF)
JV901 30.5 260 JV906 34.0 340 JV907 This case did not have an UC burn.
JV908 33.8 270 JV909 30.9 255 JV910 52.3 930 JV911 31.5 255 JV912 32.2 280 i
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