ML20116L301

Provides Responses to Technical Issues Identified in Re Review of Draft SER on GE Advanced BWR Design

ML20116L301
Person / Time
Site:	05200001
Issue date:	08/10/1992
From:	Quirk J GENERAL ELECTRIC CO.
To:	Ward D Advisory Committee on Reactor Safeguards
References
NUDOCS 9211180239
Download: ML20116L301 (170)

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GE Nuclear Energy fcfi1.7 . ;9 August 10,1992 Mr. David A. Ward, Chairman Advisory Committee on Reactor Safeguards U.S. Nuclear Regulatory Commission Washington, D.C. 20555

Subject:

GE's Responses to the Issues included in the April 13,1992 ACRS Letter Refer ence: David A. Ward Letter to Mr. James M. Taylor,

" Review of the Draft Safety Evaluation Reports on the GE Advanced Bcaling Water Reactor Design," April 13,1992

,y The purpose of this letter is to provide you with the L seral Electric responses to i the technical issues identified in the referenced letter. We have provided responses to issues 1,2,3,5,7,9,10,11,12 and 13. The staff will provide responses to issues 4,6 and 8.

I hope the attached responses are helpful in clarifying the matters raised by you and the other memoers of the Committee.

Sinc ely, j

. F. Quirk, Program Manager ALWR Certification Attachment JFQ/j cc: ACRS NRC E 1 J. C. Carroll D. M. Crutchfield C. W. Dillmann I Catton R. C. Pierson A. J. James W. Kerr C. Poslusny M. A. Ross ,

T. S. Kress C. D. Sawyer

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", , C. Michelson H. E. Townsend

- P. G. Shewmon 1 ,

'J. E. Wilkins }'

C. J. Wylie 17~129 ,b 9211180239 920010 .I PDR ADOCK 05200001 1 I A PDR (~

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(o) ISSUE 1 -CONTROL Bull. DING FLOODING The proposed ABWR plant design locates the Reactor Building Cooling Water (RBCW) ,

System at the lowest elevation in the control building, with the essential 250 V dc battery rooms and the main control room at a higher elevation, but still below ground. l Our concern with this arrangement is the potential for control building flooding due to an unisolated break in the Reactor Senice Water (RSW) System which provides cooling water from the Ultimate Heat Snik (UHS) to the RBCW System. The proposed UHS is a ground-level spray pond which we assume to be at building grade and likely to contain sufUcient water to flood the control building.

The staff should obtain sufEcient information on the interface and conceptual design ,

of the RSW System and UHS to support an adequat evaluation of the flooding potential. The staffs evaluation should include consideration ofisolation valve arrangements, the feasibility of and time available for response, and the assumption of a single active component failure during the recoonse. The design information and flooding analysis should be included in the SSAR.

ISSUE 1 RESPONSE The control building is a seven story building. It houses in_ separate areas, the control room proper, control and instrument cabinets with power supplies, closed cooling water pumps and heat exchangers, mechanical equipment (HVAC and chillers) necessary for building occupation and environmental control for computer and control equipment, and the steam tunnel.

IIIGII ENERGY LINE BREAKS The only high energy lines in the control building are the mainsteam lines (28 inch).

and feedwater lines (22 inch) which pass through the steam openings into the control-building from the steam tunnel. The tunnel is sealed in the reactor building and open in the turbine building. It consists of reinforced concrete with 1.6 meter thick walls.

lE t Any break in a mamsteam or a feedwater line will pressurize the steam tunnel with L steam and a flow of water.

1-1 L -

E The rate of blowdown from a mainsteam line break will cause a pressurization of the-steam tunnel (10.9 psid) and cause most of the steam to vent out of the tunnelinto the turbine building. Any water produced by the break will flow down the steam tunnel into the reactor building or the turbine building. .

The rate of blowdown from a feedwater line break will cause severe flooding inside the steam tunnel plus a pressurization of the steam tunnel (6 psid) Gravity or the pressurization transient will cause all of the flow to leave the control building and flow to the lower elevations of the reactor building or turbine building portion of the steam 3

tunnel. The reactor building portion of the steam tunnel can hold nearly 1000m of water. The turbine building portion of the steam tunnelis open to the inside of the turbine building.

Water or steam cannot entrr the control building. SSAR Section 6.2.3 provides a description of the subcompartment pressurization analysis performed for the steam tunnel.

MODERATE ENERGY LINE BREAKS Moderate energy water senices in the control building comprise 28-inch senice water lines,18-inch cooling water lines, Sinch cooling water lines to the chiller condenser, Ginch fire protection lines, and Sinch chilled water lines. Smaller lines supply drinking water, sanitary water and makeup for the chilled water system.

Floor drains and curbs are provided to route water to the basen. nt floor away from control or computer equipment. In those areas where water infusion cannot be tolerated, the access sills are raised.

Senice equipment rooms may build up limited amounts of water from leaks from one .

of the following systems:

Water System Inside Control Building SSAR Section Reactor Service Water System 9.2.15 Reactor Building Cooling Water System - 9.2.11 HVAC Emergency Cooling Water System 9.2.13 HVAC Normal Cooling Water System 9.2.12 Fire Hose Stands 9.5.1 12

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() Watertight doors are used when necessary to prevent the spread of moderate energy pipe failures from one division to another. Raised sills will be used to prevent intrusion of water into control areas due to fire suppression activities or cooling water leaks. Control room responses to these various levels of flooding may extend from system isolation and correction to reduction of plant load or shutdown, but control room capability is not compromised by any of the postulated Gooding events.

Maximum flooding may occur from the leakage of a 28-inch service water line. A leakage crack of 16in2 has been assumed in one 28-inch line. The senice water system is a low pressure syster ' rated at less than 100 psig). Conservatively assuming 100 psig pressure at the break . tion, this leakage crack will leak water at a rate of 3150 gpm.

Early detection by alarm to control room personnel will limit the extent of flooding.

An additional alarm at a higher level will warn control room personnel and automatically isolate the leaking division of senice water and shutoff the pumps.

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( A senice witer leak is limited to line volume plus operator response time of 30 seconds times the leakage rate. The assumed operator response time is 30 minutes to close the senice water isola' ion valves and turn off the pump in the affected senice water division after the first alarm. Also, the water in the broken pipe is conserva-tively taken to continue to Dow freely out of the broken pipe independent of the elevation of the ultimate heat sink. Two motor driven safety-related isolation valves are provided on each line to allow a single valve failure with a pipe crack. Normally dosed and administratively monitored water tight doors are provided to confine the water to a division. Assuming an operator response time of 30 minutes and allow up to 4km of senice water pipe between the control building and the ultimate heat sink, the design base leak results in 5m of water in the effected divisional basement room.

The failure of an 8 inch cooling water line in the mechanical rooms governs the floor drain system design. Total release from the HVAC emergency chilled water system or reactor building cooling water system will be limited to the system inventory and surge tank volume. Elevation differences and separation of the mechanical functions from the remainder of the control building prevent propagation of the water to the

() control area. All leakage will flow into the basement. This event will cause a lesser flood than a senice water leak.

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, .3 Flooding events that may result from the failure of the fire fighting systems within

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the control building do not inhibit plant safety. There are no sprinkler systems in the control building, llose and standpipes are located in the corridors. ,

B.eferences . SSAR Subsection Subcompartment Pressurization Analyses 6.2.3 -

Flooding Analyses 3A.1.2 General Arrangement Drawings 1.2 Ultimate lleat Sink Conceptual Design 9.2.5 RCW, HECW, and IINCW System Descriptions 9.2.11,9.2.13, and 9.2.12

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ISSUE 2 ADEOUACY OF PHYSICAL SEPARATION Pipe breaks, internal plant flooding, and external events such as fire are of major concern if their effects cannot be confined in order to protect required safe shutdown equipment. We believe that the key to confinement is the provision of appropriate-separation barriers. However, a classical barrier such as the 3-hour-ra'.ed fire barrier wall and its penetrations (e.g., doors and dampers) may not, ofitself, be suflicient to ensure separation under (a) the combined efTects of pressure, heat, and smoke from a fire, and the Gooding which results from fire mitigation, (b) the clTects of pipe whip, jet impingement, or compartment pressurization due to pipe breaks, or (c) the influx of water and hydrostatic 1sessure buildup due to internal floods.

We believe that the SSAR should describe and the staff should evaluate the adequacy of proposed separation barriers for the full range of events and conditions for which separation must be ensured. We continue o recom-mend that systems required for safe shutdown not share a common Heating, Ventilating and Air Conditioning -

(HVAC) System during normal plant opera-tion. The secondanj containment HVAC System for the AllWR is such a shared system.

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[SSUE 2 RESPONSE SSAR REFEkENCE PART (a)

The combined effects of pressure, heat, smoke, and flooding from fire mitigation are dealt with in the following manner:

General i

1. Th. plant is laid out-to minimize the number of penetrations 9.5.1.0.1 ben een divisional areas. 9.5.1.0.2 L

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2. Fire barrier floors, milings, and walls in general are concrete 9.5.1.0.3(1) -

and are at least six uches thick.

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([ 3. - Fire barrier walls which are of special construction are required- - 9.5.1.0.3(2) :

to be ol'an approved construction bearing a UL (or equal) label I

for a three hour fire rating. .

Columns and support beams are required to be_ of reinforced 9.5.1.0.3(7)T

• concrete construction or enclosed or coated to provide a three 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br /> rating if of steel construction.

Heat and Smoke

1. Fire doors, electrical penetrations, piping penetrations, HVAC 9.5.1.0.3(3,4 & 5) duct penetrations through fire barriers must be tested to ASTM El19. A hose st' ream test is required.

2. Fire dampers are required for HVAC ducts penetrating fire: - 9.5.1.0.3(6) barriers between safety 4 elated areas of different divisions. a ln j.] 3. The number of HVAC ducts penetrating fire barriers between 9.5.1.0.3(6) -

safety-rel: ted divisions has been limited to six.

4. The penetration seals are bncked up by die HVAC systems 9.5.1.0.3(8) when operating in their smoke removal mode in that they 9.5.LO.6 maintain the pressure differentials such that leakage is into the -

zone with the fire. Backflow of hot gases and smoke into the i^

fire area without the fire is also prevented by the maintained pressure differential.-

h. essure

1. Except for set.ondary containment, pressure buildup limited by 9.5.1.0.6 directly venting all fire zones through exhaust plenums -

'directly to the atmosphere with. no' intervening automatic closing fire dampers, b

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(f ' 2. Secondary containment provided with power venting by ventilation system in smoke removal mode to limit pressure 9.51.0.6.

Table 9.5-4 buildup. Capacity is 57,500 CMH (rubic meters per hour) or -

approximately 33.000 cubic feet per minute.

3. Steady state burning rates are less than t equal to the ventilation controlled rate. Low loss exhaust plenums will allow .

exhaust Hows from individual rooms at rates of 3 to 4 times the ,

ventilation supply rate so that room pressurization from heat is limited to smallincreases. (Exhausting a room at a ventilation exhaust ;o supp1v Dow ratio equal to e' ratio of the absolute temperatures results in zero pressure rise in the room.)

4. The type of penetration seals used (foamed silicon rubber and Chico A compound) possess large thermal lag characteristics.

This makes them invulnerable to the combined effects of small .;

leaks and momentary pressure surges due post flash-over O

g ventilated burning which may produce short pressure risc conditions. The momentary pcessure increases will be limited -

by the exhaust system, also.

Floodine from nre mitication

1. Door sills, Door drains, and augmented drain paths channel Src . Appendix 19R suppression water to safe accumulation areas. The adequacy of  !

this system is connrmed by the flooding PRA which assumes complete rupture e liquid containing lines such as_ fire

protection stand:. vhich supply the manual hoses and .

j sprinklers.

2. Floor penetrations are curbed to a height above the maximum level due to flooding in any area.

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3. Other than water tight doors, there are no penetrations through nre barriers which are subjected to hydraulic heads due to Gooding, if any are identified the penetrations are required to be tested for the hydraulic head they experitnce.

PART (b)

The effects of pipe whip, jet impingement, or compartment pressurization due to pipe breaks in secondary containment are discussed in SSAR Section G.2.3. There are no other areas in the reactor building or control building which contain high energy lines.

PART (c)

The influx of water and buildup of hydrostatic pressure due to internal floods are covered in the flood analysis SSAR Section 3.4 and in the probabilistic analysis for floods in Appendix 19R.

A V With respect to the concern that the secondary containment has a single ventilation supply and exhaust duct and that the fire barriers within secondary containment may fail due to loss of a portion or all of the ve , ilation system within secondary containment, the following apply:

1. GE agrees with thc ancept of separate divisional HVAC systems for eacii safety disision and have provided them for the reacto huilding outside of secondary containmen. and the coa.rol building equipment areas.

Cooling oa a divisional bacs has been provided within secondary containment.

.cntilation and smoke venting is provided by a system with *.hree load groups of.

s ?ive devices and single exhaust and supply ducts. The smoke removal system maintains the pressures in the divisional fire areas not experiencing a fire positive with respect to the area experiencing a fire.

2. The secondary containment isolation valves could clase as a result of fire induced failures of the exhaust or refueling area radiation monitors. GE does not consider

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v however, for the following reasons:

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a) The lhe uea isolation valves, attached Figure 91, close and isolate each divisional fire area.

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b) The AllWR design meets three of the six criteria given in the Fire Induced Vulnerability Evaluation Methodology (FlV10 for screening a conunon fire  ;

boundar) from consideration of possible failure. Meeting any one of the criteria screens the boundary from any further consideration for failure.

The three criteria and the method in which they are met in the AllWR design are:

1. "lloundaries that consist of a 2-hour or Shour rated fire barrier on the basis of barrier effectiveness." The AllWR barriers are rated 3-hour.

ii. "lloundaries that consist of a 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br /> rated fire barrier with a combustible loading in the exposing compartment <80,000 litu per sq. ft. on the basis of barrier effectiveness and combustible loading." The AllWR design is i lbr Shour rated fire barriers and limited to a combustible loading ofless than (G1,000 litu per sq. ft.

iii. "lloundaries where automatic fiie supprestion is installed over combustibles in the exposing compartment on the basis that this will prevent fire spread to the adjacent compartment." A sprinkler system is:

installed in the truck entry area where exposed combustibles are most likely to occur in secondary containment.  ;

3. The la>out of the divisional fire areas within secondary containment minimires the probability of failure of fire barrier penetration seals between disisions if the smoke removal system fails. Advantage is taken of the tendency of hot gases from a fire to accumulate at the ceilings of the highest rooms to which they are vcnted. The hot layer that Ibrms there creates the highest ambient temperature and the highest pressure difTerential to surrounding areas. If barriers between fire -

areas are going to fail it is therefore most likely to be in the hot ceiling layer of the fire area experiencing the fire. Due to the nature of the AllWR design within secondary containment, hot layer accumulation for the divisional areas do s not 7

occur at the location of a fire barrier boundary between two divisions, except for pocketing between floor support beams. (Figure 95 llustrates this condition.)

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For division 3 within the A11WR secondary containment, hot gases would be l]

- vented up to the operating deck by the open flood drain holes in each floor. The gases would thercr ore accumulate under the operating floor roof at elevation 49700. This is well above elevation 31700, which is the highest point for Division 2 within secondary containment. It is even farther above elevation 23500, which is  !

the high point for Disision 1. Therefore, the probability of a fire in Division 3 -

resulting in the failure of penetration seals in fire barriers to Divisions 1 or 2 is minimized. i For a fire in Disision 2, the prol: ability of barrier failure to Division 1 is minimized because the hot layer in Division 2 accumulates at elevation 31700, which is above the boundary with Disision 1 at elevation 23500. - The probability of failure to  ;

Division 3 is minimized by the fact that the zteam tunnel HVAC room is between Disisions 2 and 3 between elevations 27200 and 31700, where the hot layer would accumulate in Division 2.

For Disision 1, the hot gases would accumulate below the floor at elevation 23500 which is above the highest common barrier with Disision 2 at 18100; lletween .

elevations 23500 and 18100, Division 1 is separated from Division 3 by the steam turnel. These two layout features minimize the probability of failure of Division 1 ,

fire barriers to either Disisions 2 or 3.

4. Systems within secondary containment are designed to operate with loss of ventilation and some amount ofleakage of smoke through barriers. This is accomplished in two ways.

a) Cooling of rooms containing equipment required for safe shutdown is accomplished with safety related room coolers.

b) Motors, scaled valve operators, and scaled instrument sensors and transmitters are located in secondary containment. Smoke will not have a }

significant short term effect on this type of equipment? There are no multiplex units required for safe shutdown located inside secondary containment.

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ISSUE 3 PM)Tl:CTION OF ENVIRONMENTAlIN SENSITIVE EOUIPMENT The AllWR makes extensive use of emironmentally sensitive equipment (including solid-state electronic components) for essential protection, control, and data transmis-sion functions. Such components are known to be susceptible to adverse emironmental changes, particularly temperature extremes. We are concerned that a number of these wnponents may be located in plant areas where postulated events such as pipe breaks, fire, internal flooding, or loss of room cooling may create an adverse emironment.

The response of such components to the emironmental change may be unpredictable

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and lead to unacceptahic system int-ractions or responses. Such emiromnents need to be identified in the SSAR to ensure apptr .niate emironmental qualification of the equipment.

$ SUE 3 RESPONSE The design philosophy for essential digital equijnnent of the AllWR empha-sites the .

a development of qualified and tested equipment that is proven to be protected against

, thermal effects and other failure causes for the intended llVAC emironment with an added margin for safety. The attached Tables 3-1, SN,3-3 and 3-4 list the identity and location of this emironn.cntally sensitive equiprnent (other than the main operator control console and process computer). _

The SSAR, in Section 3.11, defines the environmental conditions with respect to limiting design (onditions for safety related electrical equipment, and documents the qualification methods and procedures employed to demon-strate the capability of this equipment to perform safety-related functions when exposed to the emironmental conditions in their respective locations. Appendix 31 of the SSAR gives the equipment qualification emironmental design criteria and describes the environmental condition parameters for the various plant zones.

ABWR equipment that performs essential protection functions is installed only in areas with mild environments. Ilowever, to ensure reliable operation under unfavorabic conditions, esser:ial AllWR equipment activated by microprocessor-based or other solid-state electronic components utiliics a defense-in41epth concept that results in severallevels of protection against the effects of adverse environments.

3-1

i SEPAldTION AND DISTRilitTTION OF COMPONENTS At the system level, four separated. independent and redundant divisions of protection logic and sensors are established. In the Reactor lluilding, the data acquisition electronics and actuator control electronics for each di,ision are located in separate,

" clean", areas with controlled environments. In the main control room area, the divisions of protection logic are also separated and operate in a con' rolled environment.

All data communications between the main control room area and Reactor lluilding is via liber optic cables l'inal trip outputs to the scram pilot valve solenoids and Main Steam Isolation Yalve (MSIV) solenoids are hardwired). The fiber optic cables will have a fire-retarding protective covering per IEEE 383.

Separation of essential divisional instrumentation conforms to IEEE 384 and Regu!atory Guide 1.75. Three-hour fire barriers are provided between equipment rooms in the Reactor Iluilding. These fire barriers are not prosided in the main control room; there is, however, considerabic separation provided by the location of the divisional back panels around the main control area and non-divisional panels (see Figure SI), in

() addition, the effect of total equipment loss is minimited by the availability of the Remote Shutdown System.

All critical actuation functions for safety related systems depend on a 2-out-of-4 voting scheme in each of the four divisions. The trip status of the four redundant sensors for each measured variable is transmitted among all divisions of logic over wctrically-independent, liber optic cables. For reactor trip (scram) and MSIV closure, the trip outputs of each disision oflogic are also voted as 2out-of-4 so that only a confirmed trip results in action by the final actuators.

In both cases of 2-out-of-4 voting, a failed or degraded chaynel can be removed from i

senice by manually bypassing the logic to a 2-out-of-3 condnion. Operation in the presence of various degraded environments is as follows:

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1. If adverse condidons, such as fire, flood or loss of HVAC occifr in one protection division in the Reactor lluilding, the division will be isolated from the trip logic by means of the bypass function.

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2. Loss of any one divisional equipment cabinet in the main control room would not disable all safety functions. The fail safe (de-energireao-operate) reactor trip and MSIV closure functions would still be available after loss of power or signal transmission. Even if the failure caused the trip output devices to stick closed, the system would effectively remain in a 2out of 3 condition. Loss of ECCS initiation outputs in any one division is mitigated by the availability of other  ;

emergency cooling systems (independent ECCS in Div.1. 2, and 3),

3. In a control room accident that results in the failure of multiple protection divisions, the Remote Shutdown System and manual scram capability separate from the electronic instrumentation proside emergency backup capability, ERROR DETECTION Microprocessor based equipment can fail in complex and subtle ways, but only errors in the critical data path must be trapped in order to prevent false outputs to the equipment actuators.

O Continuous, on-line, self-diagnostics within each microprocessor based controller ,

proside error detection oflost or corrupted data or broken cables throughout the safety-system channels. If a fire or flood in a Reactor Building area damages the data transmission capability of a division ofinstrumentation, the errors are detected and the

failure is ar,nunciate,d ir the control room.

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Error detection capat,lity includes data checks (reasonableness, bounds checking),

RAM and ROM ci,ecki, program flow checks and program timing via watchdog l

timers. Sptem hard vare is abo monitored for shoued, open, and oscillating inputs l

and outputs, and high or low t h wer supply voltages.

In addition to direct detection of data errors, fire detection desices (smoke detectors and

y. prc<luct-of-combustion detectors) located in the divisional Emergency Electrical Eqtnpment Rooms in the Reactor Building will alert the operator in the main control-room to fire in the area.

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ROHUST COMPONENT DESIGN All solid-state components for the protection systems v.ill be qualified for Class lE senice and a 190 year qualified radiation life in the location where they are installed. Low power semiconductor technology (CMOS, low power Schottky, etc.) will be employed to the maximum extent possible. All equip-ment will be designed to operate with loss of liVAC; cooling fans are not specified as part of the cabinet design. Normal operation is presently specified from 5-50 *C (75 *C test) and 0-90% Ril, noncondensing. Ilowever, all Class lE equipment will be designed with Mll SPEC semiconductors that meet Mll STD 883C.

These components are hermetically-scaled and will operate to 110 *C (125 *C test).

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Table 3-1. Division I Essential IkC Faulpment Eauipnwnt l>ocation Environment Remote 51ultiplexing Reactor fluilding Div. I Cooled by essential Div, i Units with: Emergency Electrical HVAC.

- Analog signal Equipment Room (outside Non radioactive area.

conditioning secondary containment). No piping through area.

- A/D converter Protected by Shour fire

- Contact closure inputs barriers.

and outputs

- hiux controlinterface Fiber optic multiplexing lictween Reactor liuilding Non radioactive areas.

cables and main control room in No cables run in primary Div. I cable trays or containment.

conduit.

Control Room hiain control room Div.1 Cooled by two redundant

.\fultiplexing Units with; back panel area. safety-related liVAC-

- Seriall/O systems.

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Parallel I/O Siux control interface Safety System Logic and hiain control room Div. I Cooled by two redundant Control cabinets with: back panel area, safety-related HVAC Digital Trip Aladules systems.

Trip 1.ogic Unit Separated from other

- Safety System Logic - divisi ms by mam control Units area and non-divisional Output logic Units panels.

- Load Drivers Ilypass Control Unit Surveillance Test Controller Fiber optic interdivisional hiain control room - optical Cooled by two redundant data links transmission medium safety related 'HVAC provides electrical and systems, physical isolation.

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Table 3-2. Division II EuentialIkC Ihulpment Jiquipment 1.ncadon Environment Remote hiultiplexing Reactor fluilding Div.11 Cooled by essential Div,11 Units with: Emergency Electrical llVAC.

Analog signal Equipment Room (outside Non-radioactive area.

conditioning secondary containment). No piping through area.

A/D converter Protected by Shour fire Contact closure inputs barricts.

and outputs

$1ux control interface Fiber optic multiplexing lictween Reactor lluilding Non radioactive areas.

cables and main control room in No cables run in primary Div.11 (able trays or containment, conduit.

Control Room hiain control room Div.11 Cooled by two redundant hiultiplexing Units with; back panel area. safety related llVAC Senal l/O systems.

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hiux control interfitte Safety System 1.ogic and Stain control room Div.11 Cooled by two redundant Control cabinet with: back panel area. safety.related ilVAC Digital Trip hlodules systems.

Trip 1.ogic Unit Separated from other Safety System 1.ogic - divisions by main control Units area and non-divisional Output logic Units panels.

1.oad Drivers Ilypass Control Unit Surveillance Test Controller Fiber optic interdivisional hiain control room - optical Cooled by. two redundant data links transmission medium safety related 11VAC provides electrical and systems.

physical isolation.

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V Inhle 3 3. Diyhhm Ill Essential I&C Eaulpment f.quipment location Environment itemote Multiplexing Reactor liuilding Div.111 Cooled by essential Div.111 Units with: Emergency Electrical HVAC.

Analog signal Equijnnent Roorn (outside Non radioactive area, conditioning secondary containtnent). No piping through area.

- A/D converter Protected by 3 bour fire

- Contact closure inputs barriers.

and outputs Mux control interface Fiber optic multiplexing lictween iteactor lluilding Non radioactive areas.

cables and main control room in No cables run in primary Div.111 cable trays or containment.

conduit.

Control Room Main control room Div,111 Cooled by two redundant Multiplexing Units with: back panel area. safety related liVAC

- SenalI/O systems.

- ParallelI/O p -- Mux control interfitcc V

Safety System Imgic and Main control room Div.111 Cooled by two redundant Control cabinets with: back panel area, safety-related ilVAC Digital Trip Modules systems.

- Trip logic Unit Separated from other Safety System 1.ogic - divisions by main control Units area and non-divisional Output logic Units panels.

1 oad Drivers '

liypass Control Unit  :

Surveillan(c Test Controller Fiber optic interdivisional Main control room - optical Cooled by two redundant data links transmission medium safety-related liVAC provides electrical and systems, physical isolation, V

m) 3-7 m

1. Y%

• V Inble 3-1. Division IV EssentialIkC Eculpment Enniomem location Environment Remote Multiplexing Reactor fluilding Div. IV Cooled by essential Div.11 Units with: area (outside secondary liVAC.

- Analog signal containment, adjacent to, Non radioactive area.

conditioning but physically isolated No piping through area.

A/D converter from, Div. I Emergency Protected by Shour fire Contact closure inputs Electrical Equijnnent barriers.

and outputs Room).

. Slux contrci interface Fibu optic multiplexing lictween Reactor lluilding Non radioactive areas.

cables and main control room in No cables run in primary Div. IV cable trays or containment.

conduit.

Control Room Main control room Div. IV Cooled by two redundant >

Multiplexing Units with; back panel area. safety rel' a ted liVAC

- Serial I/O systems.

Parallel 1/O Mux control interface k Safety System 1.ogic and Main control room Div. IV Cooled by two redundant Control cabinets with; back panel area. safety related 1IVAC Digital Trip Module systems.

Trip 1.ogic Unit Separated from other Output logic Units divisions by inain control 1.oad Drivers area and non divisional Ilypass Control Unit panels.

Surveillance Test Controller Fiber op:ic interdivisional Main control noom - optical Cooled by two redundant data links transmission medium safety related llVAC provides electrical and systems.

physical isolation.

I p.

1 1

hd

i fm ISSUE 4 REVIEW OF CHil.1 ED-WATER SWTEM

+ 3 -

' Q} .

The ABWR uses large chilled *ater systems to provide essential environmental cooling, which in turn includes cooling of the solid state electronic components.

Because there was no SRP for chilledwater systems, the staff used other guidance such as SRP Section 9.2.2 (Reactor Auxiliary Cooling Water Systems) when the safety evaluation was perfortned, llowever, this guidance is not appropriate for the evaluation of refrigeration systems.

The NRC stair needs to evaluate the performance of chilled water systems under varying accident heat loads and during loss <>f-offsite power events, and to consider their ability to restart and function after a prolonged station blackout. The DSER sections which should evaluate the performance oflarge chiller packages do not i address these issues, We believe they should.

ISSUE 4 RESPONSE Response to be provided by staH'.

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,Q ISSUE 5 USE OF LEAK ilEFORE-BRIMK 51ET110DOLOGY

%J lt is our understanding that GE will not propose the use ofleak before-break methodology for the AllWR standard plant. Thus, the DSER should be reviewed to ensure that consideration is given to pipe break effects for all systems and locations.

This may introduce additional structural protection and environmental qualification requirements in the SSAR.

3SSUE 5 RESI.11SSE The GE/NRC Staff position on leak-before break (LBB) methodology and its application is documented in Summary of Meeting fleid on lumnber 9-10,1991, dated January 3,1992. This position is restated below:

L GE is not seeking approval of Lilll for any specific piping for design certification.

2. The detailed design of ABWR for design certification is performed on the basis

\ of nm assuming LilB for any piping.

3. GE has documented in the SSAR, however, that a combined operating license (COL) applicant, who references the AllWR certified design, may still apply to the NRC for approval of LBB for selected piping by submitting an LBil report if approved, the COL applicant can then eliminate certain hardware and structures and equipment qualification requirements as noted below.

4. GE has documented in the SSAR a detailed technical process for an COL applicant's use in preparing an LBB report and applying for NRC approval of LBB.

The process includes criteria for materials selection, emironment (e.g.,

intergranular stress corrosion cracking, water hammer, thermal fatigue, etc.),

leak detection, and fracture mechanics. .

5. GE has included an appendix in the SSAR to provide detailed guidelines to a COL applicant in implementing the LBB process. These guidelines identify the necessary invetigations, analyses and evaluations, and propose methods for..

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(

analyses that could be acceptable to the NRC staff. Also, included in the appendix-5-1

.y. .- - y & _y. .r. p

an Illi evaluations of two example piping systems to serve as guidelines for Q(%

preparation of an Lilli report.

6. The stafT has reviewed and audited the Lilli process and methodology in the SSAR and the appendix. The staff will include its findings in the FSER.

The Lilli approach presented in the SSAR is not used to replace or relax existing regulations or criteria pertaining to the design bn es of (a) emergency core cooling system, (b) containment system, or (c) emironmental cualification.110 wever, consistent with General Design Criterion 4 (GDC-4), the design bases for dynamic quali0 cation of mechanical and electrical equipment may exclude the dynamic load or vibration effects resulting from postulation of breaks in the 1.lllkjualified piping.

The piping that will he approved by the NRC stalT ror its 1.1111 behavior based on the COL applicant's Lilll report is called the Lillk]ualified piping. Further, consistent with GDC-4, the design bases for structures and equl[nnent may exclude the dynamic effects (such as missile generation, pipe whipping, pipe break reaction forces, Jet impingement forces, decompression waves within the ruptured pipes, and compartment, subcompartment and cavity pressurizations) resulting from postulation of breaks in the Lilll-qualified piping.

The SSAR identities all requirements of structural protection and environmental qualification as a result of postulation of breaks in any piping regardless of whether or not it will be Lilll-qualified.

O 5-2'

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L/

ISSUE 6- USE OF INTEGRAL.1 OW-PRESSURE TURlilNE ROTORS .

i in our July 18,1991 report to you, we reconunend that the staff resiew the issues - '

involved with the use ofintegral low pressure (1.P) turbine rotors. It is our under-standing that this new design for LP rotors will be used for the AllWR, (Rotors of this -

type are being used in rotor replacement progranis at currently operating plants.) The -

practice of turbine manufacturers has been to bore the centerline of this type of rotor to remove impurity inclu.,;ons. We were concerned that the use of unbored rotors was I being contemplated. The Electric Power Research Institute (EPRI) has recently added a requirement in its Advanced 1.ight Water Reactor Utility Requirements Document (URD) that 1.P totors be center bored. .

b ISSUE 6 RESPONSE  !

Response to be provided by staff.

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O V' ISSl'E 7 CAVITY Fl OOR ARFA IlENEATH REACTOR VESSEL l l

i i

The cavity area beneath the reactor vessel is slied to meet the EPRI URD specification ]

of 0.02 m2 /MWt. The AllWR design includes flooding of the cavity. Little consideration has been given to how this should be accomplished. There is little evidence that the planned cavity area willlead to quenching followirig flooding or that the AllWR flooding plans will not lead to ex vessel steam explosions. Further attention needs to be given in the SSAR as to when and how fast the cavity should be flooded in order to avoid exacerbating a core melt accident ifit should occur.

ISSt'E 7 RESPONSE .

A variety of issues were conside ed when assessing the lower drywell response to severe accidents. The primary challenges in the lower drywell are the potentials for core concrete attack and for ex vessel steam explosions. An additional containment t

challenge is the potential for a substantial temperature rise in the containment which could damage the penetration seals and cause excessive leakage from the containment. These severe accident challenges must be balanced with the practical V considerations of normal plant operations. Additionally, there was a desire to make the passive response to a severe accident as benign as possible since several

- equipment failures and/or operator errors are necessary precursors to any ,

hypothesired severe accident in the AllWR.

This idea was implicit in the EPRI URD specification of 0.02 m2 /MWt. Allowing ample space for the debris to spread increases the probability that the debris is coolable given the uncertainties in phenomena. A considerable effort was undertaken to assess the debris coolability for the AllWR design. A summary of that study is given later in this response. Also important to the issue of debris coolability '

is ensuring that molten core debris does . at enter the sumps in the lower drywell. To this end, a conceptual design for a corium shield which prevents debris ingression into the sumps was developed. The corium shield is also discussed later.

The logic described above was also the basis for the selection of severe accident features. The containment overpressure protection system (COPS) was added to the design to preclude structural failure of the containment. In the COPS design a

/3 . rupture disk opens a hardened pipe to the containment stack when the wetwell

\ 'J pressvie exceeds 90 psig. This ensures that any fission product release is scrubbed.

CElb92-16 71 -

Among other benefhs, this considerably reduces the uncertainty in the offsite dose which could result from a severe accident whh continued core concrete interaction.

If the vessel fails and debris is relocated into the lower dr>well, covering the debris with water is highly desirable. Not only does flooding the debris allow for the possibility of quenching it, but it also prevents the temperature in the or>well from rising to a level at which penetratioa seals could be damaged, leading to excessive leakage. Unfortunately, there is also a risk of steam explosions damaging the l containment when water comes into contact with molten debris. A scoping study, }

described below, was developed to determine the magnitude of this threat. }

Operational considerations preclude the option of running the reactor with the lower dr)well flooded. The lower drywell must be flooded after a severe accident has-been successfuliy diagnosed. To provide for flooding, an AC independent water addition system was introduced to the design by allowing for the cross-tle of the fire protection system to the RHR system. Connections are also provided for the use of a fire truck to pump water into the containment via this same system. The firewater addition system allows the operator to add water to either the vessel or the upper dr>well sprays if necessary. This system, while using active pumps, is diverse iam the 73

-. ECCS network, which must be presumed to have failed if a severe accident is hypothesized. After vessel failure, the addition of water through the LPFL lines will flood the lower drywell, and provide for debris cooling. In the SSAR, it was determined that the failure of this system is dominated by operator error, Given the long time available for the operator to initiate the system, the failure rate is estimated to be 0.01 per demand. Thus, any debris which accumulates in the lower dnwell will dmost always be flooded via firewater addition.

In the very unlikely event that the firewater system is not initiated in a severe accident, a device which is independent both of power and of the operator is desired.

The ABWR has a passive lower drywc4 flooder which opens when the tempergure in the lower drywell begins to increase after vessel failure. In a severe accident with vessel failure and no active injection, the core debris which falls into the lower dr)well will eventually dry out. At this time, the debris will begin to transfer heat to the gas and structures in the lower drywell via conduction and radiation. A passive lower drywell flooder has been designed to open a pathway from. the suppression pool to the drywell when the temperature exceeds 5_00 F. The design considerations J

CEB-9246 7-2

n .

,$' for the passive Dooder are discussed below after the phenomenological N tonsidera; ions have been discussed.

DEBRIS COOL-W' rIY IN LOWER DRYWELL i

, appendix 19E of the SSAR discusses core concrete interaction. in part cular, in Settion 19E.2.1.16, it is stated that the core debris will be quenched preventing sutaanthi concrew ablation due to operation of the firewater system or the passive flooder. Even if the Dooder was assumed to fail, water from the cappression pool would flood the lower dr)well after 8 inches of rad),o ablation had occurred. The late t.ddition of water would quickly terminate the core concrete interaction.

Since the original ABWR PRA was submitted there has been continued research in the areas of debris coolability and core concrete interaction. Recent experiments performed at Argonne as part of the MACE program have indicated that debris cooling may be limbed due to crust formation. Analysis was performed to investigate the uncertainties associated with debris coolability in the lower drywell of the ABWR.

The details of this study are provided in Attachment 7A to this response. The experimental data was examined for applicability to and implications for the ABWR.

Next, the issue of debris coolability was decomposed into the controlling parameters and a decomposition event tree (DET) was developed. Deterministic evaluations were then made to quantify the impact of the phenomenology represented in the tree.

Finally, sensitivities to key assumptions were inves.tigated. The conclusions from the uncertainty analysis are summarized below:

1. For core melt sequences that release core material into the containment,90% result in no significant CCI. Virtually no sequences have dry CCI.

2. Even for the low frequency cases with significant CCI, radial erosion does not threaten the structural integrity of the pedestal.

3. The fission product release mode is dominated by operation of the COPS for the case with CCI The release magnitude, which occurs at about 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> for the dominant sequences, cannot be distinguished from a case with no CCI.

O G

CElb92-16 7-3

p 4. Experimental results indicate that sufUcient upward heat transfer d to an overlying watci pool would exist in the AllWR lower drywell to cool the debris.

PROTECTION OF LOWER DRYWELL SUMPS The AllWR has two drain sumps in the periphery of the lower dowell floor which could collect core debris during a severe accident if ingression of the debris is not preverded. If debris enters the sump, a debris bed would form which could be thicker tl.rn th. bed <m the lower dowell floor. Debris coolability becomes more uncertain as the thickness cf a debris bed increases. Thsefore, a protective layer of refractory bricks, a corium shield, has been proposed to prevent the ingression of core debris in the lower drywell sumps.

The equipment drain sump gathers any water leaking from valves and piping. Since the water entering this sump is piped to it, the corium shield for the equipment drain sump is solid except for the inlet and outlet piping which would go through its roof. The walls of this shield need only be thick enough to prevent elevated debris temperatures from degrading the shield structural support.

O The floor drain sump is designed to collect water which falls onto the lower dowell floor. Its corium shield would be similar to that for the equipment drain sump except that it must have channels at the floor level to allow water to flow into the sump. The design restrictions for this shield are more challenging than those for the equipment drain sump. Tne debris must initially be froren in the channels before it ingresses into the sump, and in the long term, it must remain frozen. The basis for sizing the floor drain sump is given in Attachment 7B to this response.

For the short term considerations, a shield material must be selected such that the initial contact temperature of the debris with the shield is less than the melting point of the shield material. This prevents ablation of the shield. The channel height and the thickness of the corium shield must then be determined in order that the debris frecies in the channel before reaching the sump.

For the long term, the shield height above the dowell floor is chosen so that it is not covered with debris. Additionally, the upper shield height is constrained so that it is capable of removing all of the decay heat generated in the channel after a specified S

Q time. The depth of the lower shield wallis then selected such that the debris / shield CEB-9246 7-4

~ interface remains below the debris meltir. temperature and the concretehhield

wall remains below the concrete mehing te mperature. For this calculation all of the decay heat generated in the channelis assumed to be absorbed by the uoper and ,

lower shield walls until the specified time if the outer surfaces of the shield are considered to be insulated. >

A sample calculation was performed using a refractory brick for the shield wall. The  ;

design concept was shown to be feasible. In the sample calculations, upper and lower shield wall heights of 0.4 m were found to be acceptable. This diould not have any-significant impact on the lower drywell arrangement. ,

POTENTIAL FOR CONTAINMENT FAILURE DUE TO EX VESSEL IUEL-COOLANT INTERACTIONS Challenges of the containment during a severe accident may result from fuel coolant interactions (FCI). FCI, often referred to as steam explosion, may occur either at the 1 time of vessel failure when corium and water fall from the lower plenum of the vessel, ,

or when the lower drywell flooder opens after vessel failure has occurred. Fhe critical time censtants for a steam explosion were considered in section 19E.2.3.1 of the SSAR. This analysis concluded that the critical rates for heat transfer and energy

-dimersal indicate a large scale c: cam explosiont c bich could damage the containment will not occur. Nonetheless, a scop.g stu@, provided as Attachment 7C to this response, was performed ~to determine the magnitude of steam explosion which could potentially damage the ABWR containment. ,

7' ,

Before any calculations were performed, several experiments which have provided insights to steam explosions were examined. It was noted that the only. experiments-which have led to catastropr.ic damage are not applicable to the ABWR. In particular,-

the FCI shich_ recently occurred in the BETA facility iri Germany appears to have ;

L~ been caaed by intimate' mixing of the debris and water as a result of the very,

[ . confined geometry of the experimental con 6guration. In contrast, the ABWR lower - '

- drywell is very spacious.-The features of the ABWR were also compared to current..

3 L~ Loperating plants to indicate Re rblative resistance of the ABWR to steam explosionsi h . Four potential failure modes were ' considered in the scoping study. The transmission = ,

of a shock wave through water to the structure may damage the pedestal. Similarly, al ,

ify shock wave through.the airspace can cause an impulse load; However, since the gas is!

U ,

CEB 92-461 7-5 ,

1 V - - e., --,-i + , w eg n , w se

l 7 compressible, the shock wave transmitted through the gas will be much smaller than

( that which can be transmitted through the water. Therefore this mechanism was not considered explicitly. Third, loading may he caused by slugs of water propelled into containment structures as a result of explosive steam generation. Finally, the rapic steam generation may lead to overpressurization of the dr>well.

The limiting FCI challenge was found to be the impulse loading of the pedestal through the water. A simple clastic. plastic analysis of the pedestal structure was performed. It was determined that the ABWR pedestal can withstand a TCl invohing 9.5% of the core inventory. Experimental studies on FCI extrapolated to the ABWR geometry indicate that less than 3% of the debris could possibly participate in a steam explosion in a stratiHed condition where water is added to debris. Therefore, it is concluded that FCI is not expected to damage the ABWR containment.

LOWER DRYWELL FLOODER CONSIDFRATIONS As noted in the introductory discussion, almost all of the severe accident scenarios which result in core debris in the lower dowell are expected to be Hooded using the firewater addition system. The lower dr>well Hooder would be called upon only if the Os operator failed to initiate the firewater system (or,less probably,if an additional diverse system failure occurs). This directs the consideration of a means of Hooding the lower drywell which is completely independent of the operator or initiation logic. The passive Hooder included in the ABWR design meets this goal. Details of the sizing considerations for the lower dr>well Hooder are contained Attachment 7D to this response.

The lower dnwell Hooder opens as a direct consequence of the containment challenge for which water is desirable. If the debris in the lower dowell dries out and begins to heat up, the temperature of the gas in the lower dr)well will also incicase. If the upper drywell temperature were to rise above 500 F (533 K), the containment penetrations could be degraded and begin to leak. Therefore, the lower dowell Hooder is designed to open within about 10 minutes of the lower dowell temperature reaching 533 K. Since there will be a temperature difference between the upper and lower drywells, the passive dooder prevents high ;emperature failure of the containment.

I U

CEB-92-16 76

(3 The passive Hooder valves are not expected to be used in the lifetime of a plant.

b Therefore, there is a significant mcentive to limit the cost of the system. Cour inch valves may be commercially available. Therefore, this sire was selected as tn. basis for the design. The number of valves was chosen to allow the rapid flooding of the lower dr)well to its equilibrium level. The presence of the ten wetwell / drywell connecting vents in the ABWR design suggested the use of 10 passive flooder lines. Analysis performed for the AllWR configuration indicates that the lower dr)well will be flooded in 21 minutes if decay heat is considered. However, if the debris temperature has risen, as would be expected, the debris must also be quenched by the incoming water from the passive flooder. The maximum (transient) heat flux for water over debris which has been observed in debris coolability experirnents is 2 MW/m2. This heat flux would quench the entire debris bed within minutes. For a conservative treatment, 2 MW/m2 is assumed to be the constant heat aux added to the pool of water forming in the lower drywell. If nine of the flooder lines are presumed to open, the lower drywell will be flooded to its equilibrium height in less than 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br />. This isjudged to be sufDciently rapid since the generation of steam during quenching will cool the containment.

The potential implications if only one valve opens were also considered. Ifit assumed that only one valve opens, then the debris will not be quenched with the flow rate throug' one valve. Since the flooder lines are located on the periphery of the lower drywell, a localized region of the debris would be quenched. But, before water could flow across the entire lower dr)well region all of the water would be boiled off. Thus, at some distance from the open valve, the core debris would continue to heat up. This would lead to the opening of additional flooder lines until sufDclent water was being added to the containment to quench the entire surface of the debris, an a water pool begins to form. Once a significant pool of water begins to develop the opening of additional valves will not enhance the coolability, if a crust has formed on the debris which limits .he heat transfer, adding additional water will not increase the heat transfer from the debris. Therefore, there is no concern that operation of one lower dowell fusible line will induce long term core concrete attack which would not otherwise have occurred.

SUMMARY

The containment challenges related to the lower drywell design have been

'v considered. The design of the corium shield surrounding the lower dowell sumps CElk92-16 77

+

i p prevents the ingression of debris into the sumps. This allows any core debris which V

reacnes the lower dr>well to be spread over the entire floor area of the lower dr>well.

Studies were performed which indicate there is only a small probability that the debris would not be coolable given the uncertainties in the phenomenology. Even if >

the debris were not cooled,it has been demonstrated that the implications of continued core concrete interaction on the containment performance are not significant. The operation of the containment overpressure protection systern (COPS) would not occur for about 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />, and the release of fission products from the containment is indistinguishable from the sequences with no core concrete ,

interaction.

The potential for fuel coolant interactions (FCI) was also considered. A scoping calculation indicates that the AllWR can withstand an FCI with 10% of the total core- i mass participating. Experimental esidence indicates that the maximum mass of core debris which could participate in an FCI in the AllWR design is SE Thus, FCI is not -

a significant threat to the ABWR containment design.

Flooding of the lower dr)well in a severe accident is accomplished by either the ac independent firewater addition system or the passive lower drywell flooder. The firewater addition system is diverse from the ECCS injection network The failure of ,

the system is dominated by the failure of the operator to initiate. Since there is a long interval for the operator to start the system the probability of successful operation is high.

Since the failure of the operator to initiate dominates the failure of the firewater -

addition system, a passive lower dr)well flooder was implemented which does not rely on the operator. The passive flooder will open within a few minutes of the lower dr>well semperature exceeding 500 F (533 K). Thus, overtemperature failure of the penetration seals in the upper dr)well is precluded. The flooder is sized to allow the .

rapid quenching of the debris, and flooding of the lower drywell to its equilibrium level within a few hours at most. The opening of one fusible plug before the others open will not lead to degradation of the mtem performance.

Containment challenges which affect the lower dr)well have been addressed. There are ample means of providing water to the lower drywell. The containment design has been shown to withstand a fuel coolant which involves a debris mass more than three times the credible mass which could participate in an FCI. The arrangement of CEIL 92-16 78

the lower drywell lirnits the potential for core concrete interact - given the current I state of phenomenological uncertainty. Thus, it has been demonstrated that the design of the A11WR lower drywell is very robtist.

n kb l

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Cell-92-16 7-9 1

NITACllMENT 7A l

This material is extracted from a June 30 transmittal to the staff 1 m with small differences in the format.ing. This information will

( ) be incorporated into the ABWR SSAR at a future date.

v X.3.2 Debris Coolability and Core Concrete Interaction Appendix 19E of the ABWR PRA discusses core concrete interaction.

In particular, in Section 19E.2.1.3.6,it is stated that the core debris will be quenched preventing substantial concrete ablation due to operation of the passive flooder. Even if the flooder was assumed to fail, water from the suppression pool would flood the lower drywell after 8 inches of radial ablauon had occurred. This conclusion was based on available experimental information and the work performed in IDCOR Subtask 15.2 (Reference 1).

Sir.cc the original ABWR PRA was submitted there has been continued research in the areas of debris coolability and core concrete interaction.

Recent experiments performed at Argonne as part of the MACE program have indicated that, due to crust formation, debris cooling may be limited.

Th a section will investigate the uncertainties associated with debris coolability in the lower drywell of the ABWR. The investigation will begin with a look at applicable experimental data. Next, the issue of debris coolability will be decomposed into the controlling parameters and followed by the develo ruent of a decomposition event tree (DET). After (qj creation of the DET, c eterministic evaluations will be made to quantify the end points of the tree. Finally, sensitivities to key assumptions will be investigated.

X.3.2.1 Applicability of Experiments to ABWR Several experiments have been carried out to investigate the influence of an overlying water pool on debris coolability. The critical parameter that appears to dominate the behasior in several of the experiments is the formation of a stable crust. This crust is found to prevent substantial water ingression and, therefore, debris cooling. The major criticism of these experiments is that, due to their small scale, a stable crust is preferentially formed. This limitation makes it quite diflicult to extrapolate the results to a large reactor cavity. The MACE tests at Argonne have attempted to address this weakness by investigating larger cavity designs.

The following provides a brief summary of several debris coolability experiments.

A i t

V 1. Final Report on Core Debris Coolability, IDCOR Task 15.2 CEB-92-46 7A-1

Theofanous and Saito - 1980 (Reference 1)

\

m

/ Experiments were performed with liquid nitrogen and water and liquid nitrogen and Freon 11. Crust formation was observed at low gas velocities but found to become unstable at high sparging rates. It was observed that as the gas velocity increased to a magnitude typical of core-concrete interaction, the heat transfer rate increased by a factor of ten. The heat transfer rates were found to approach those associated with critical heat flux.

Greene 1988 (Reference 2)

Tests were run with liquid metals with water and Freon Ril. Gases were injected in the melt, it was observed that the water / melt interactions were generally unstable and that the upward heat transfer increased with gas velocity. The typical upward heat transfer rates were found to be 6 times greater than the classical f>crenson correlation.

FRAG (Reference 3)

This series of tests performed at Sandia National Laboratories used 3 mm diameter steel spheres heated and placed in a 20 cm diameter concrete crucible. Tests were performed both with and without water addition. Iloth limestone and basaltic concrete types were investigated. The limestone tests showed that a stable crust made of

'1 concrete and steel formed that kept the water from penetrating the J rest of the debris bed. The basaltic concrete allowed for some water penetration. The conclusion from these tests was that core-concrete attack continued even in the presence of water and that a substantial 4, amount of steel oxidation took place.

SWISS (Reference 4)

These tests, also performed at SNL, involved the interaction of molten steel on limestone concrete. The steel was heated at approximately five times the expected reactor decay heat levels. There appeared to be no

l. An Integrated Suucture and Scaling Methodology for S' vere Accident Technical Issue Resolution, to be published as NUREG/CR, Draft 199t.

2. Greene, G.A., C. Finfrock and S.B. Hurson, "Phenomenological Studh s on Molten CoreConcrete Interactions," Nuclear Engineering and Design, 108,167177, IC1

3. Tarbell, M.W., D.R. Bradley, R.E. Blose, J.W. Ross, and D.W. Gilbert, "Susnined Concrete Attack by low-Temperature Fragmented Cere Debris," NUREG/CF $024, SANDS 2 247G R3,R4, July 1987.

4. Blose,R.E. J.E. Gronager, AJ. Suo Antilla, and J.E. Brockman, " Sustained IIcated Metsnic/ Melt Concrete Interactions with Overlaving Water l'ools," NUREC/CR-4727,

/ SAND 851546 R3, R4, R7, July 1987.

CEB-92-46 7A 2

, oa e NS ,

4 violent melt-water interactions and the melt did not quench. There- '

l7)-

V was a stable crust that was faund to attacn to the MgO sidewall. Typical upward heat flux.was 800 kW/m2. There was also information from the expenment that the overlying water pool provided substantial aerosol

, scrubbing (DFs of 10 30).

Mark i Shell Failure Experiments (Reference 1)

Several experiments were carried out Fauske and Associates to investigate the influence of water on debris coolability and specifically to observe drywell shell heatup. Iron-alumina thermite was discharged onto a concrete slab pre-floode 1 with water. The initial heat transfer was found to be quite high (20 times CHF) and leveled off at about 800 kW/m2 later. .

MACE (Reference 2)

A series oflarge-scale experiments are being pe-formed at Argonne National Labo:atory investigating the coolabil.a of molten corium by_

water during iu interaction with concrete. The M ACE program has attempted to 1) employ prototypic corium melt matenals,2) et. ploy prototypic concrete types,3) obtain realistic melt temperatures,4) obtain realistic MCCI initial conditions,5) include prototypic chemical and internal heating, and by the increased size,5) ensure applicability to reactor cavities.

O In the scoping test, a high initial heat removal was observed. The crust that was formed was found to be supported by the electrodes. There were periodic melt eruptions through the crust that lead to substantial melt quenching. However, the melt did not completely cool and continued to erode concrete. One of the major difficulties with the test was that there were larger that prototypic heating rates.

The next test, M1, was performed on November 25,1991. The major difficulty with this test was that not all of the material melted initially

~

and the sintered region on the top kept the water from penetrating the melt. Low melt-water heat transfer rates were observed. Concrete attack continued with the debris not cooled.

'1 R. Henry, " Experiments Relating to Drywell Shell Core Debris Interaction," BWR -

Mark 1 Containment Workshop, Baltimore; MD, February 24-26,1988. See also Malinovic, B., R. Henry, and B. Schgal, " Experiments Relating to BWR Mark I Drywell Shell Core Debris Intera:tions," ASME/AIChE National Heat Transfer Conference, Philadelphia, August 1989.

2 ACE Program Phase D: Melt Attack :,nd Codability Experiments (MACE) Program,

[/

y presentation by B.R. Sehgal at CSARP meeting. May 1992.

CEB-92-46 - 7A-3

The most recent test, MiB, corrected the problems encountered with

.('N M1. The melt temperature was observed to. decrease steadily to near the-V concrete liquidus temperature after the water was introduced.

Concrete ablation was found to continue but at a reduced rate (a few mrn/hr). The post-test examination showed that there were large holes in the top surface.

The experiments described above are insufficient to enable a full understanding of debris coolability is re lower drywell of the ABWR. Some insights can, however, be extracted - following shows the observed upward heat flux for three of the tesa.

SWISS -

800 kW/m2 Mark 1 Shell Test - 800 kW/m2 MACE Scoping -

600 kW/m2 One of the major reasons why these tests are not prototypic is that, due to their small scale, they promote a stable crust formation. The larger scale MACE tests shoiild generate some useful insigats.

X.3.2.2 Description of Event Tree Analysis X.3.2.2.1 Debris Coolability A decomposition event tree (DET), shown in Figure 7Al, was

! developed to assess the likelihood of debris coolability. This section l describes the branch points and the quantification of this DET.

X.3.2.2.1 Fraction of Debris in Lowe- Dr>well Early (COR_DW E) _

This event assesses the initial debris mass which relocates to the lower drywell soon after vessel failure. The amount of _ debris which enters the lower drywell early is dependent on the amount of debris molten in the lower RPV head at the time of RPV failure and on the amount of entrainment of the debris from the lower drywell. However, for simplicity, L debris entrainment to the upper drywell was conservatively neglected in

i. this analysis. For consistency with the DCH analysir, two regimes are -

l considered for the fraction of the core inventory which is molten in the RPV at the time of RPV failure (see section X 1.1.4 (of the DCH analysis, submitted to the staff on April 20]). These regimes are:

Low 0 - 20% (nominal 10%) 0.9 High 20 - 40% (nominal 40%) 0.1 CEB-92 46 7A-4

p X.3.2.2.2 . Amount of Initial Debris Superheat (SUPERHEAT)

This event is used represent the initial debris temperature when the debris first contacts the lower drywell floor. It is also used as a surrogate to represent the additional metal / water reaction heat production associated with a high metal to oxide ratio in the debris. Superheated debris or debris with a high metal content is expected to be more difficult to quench initially and to experience faster initial concrete crosion. In the deterministic CCI analysis discussed in section X.3.2.1 the low superheat cases are represented by (molten) debris at the U Zr-O eutectic melting temperature (approxiniately 2500 K). High superheat was taken to be temperatures in the range 300-500 K above the melting temperature. This was represented in the deterministic analysis by increasing the amount of steel added to the melt prior to vessel breach.

Two cases were considered in the DET analysis. The first case represents sequences with a small amount (10% of core inventory) of molten debris in the lower plenum at vessel rupture and the second case represents large amounts of debris (40%).

For the case of a small debris mass in the lower RPV, it is likely that either

1. Vessel failure occurred fairly quickly after core slump into the q lower plenum (MAAP type failure model), or that V 2. The debris in the lower plenum was initially quenched by residual in the lower plenum and that RPV failure occurred later after the water was boiled away and the debris started to reheat (BWRSAR type failure model).

For these situations it is judged likely that the debris temperature will be at, or near, its melting point. Hence, the following probabilities were assigned for this case:

Case 1 Small Debris Mass in Lower Drywell Early Low (Superheat) 0.9 High (Superheat) 0.1 For the case of a large amount of molten debris it could be expected that this resulted from a delayed failure of the RPV allowing more debris to flow into the lower plenum (MAAP model) or for melting and heating of quenched debris already relocated to the lower plenum (BWRSAR model).

For both situations the extended time to vessel failure could result in higher molten debris temperatures a. RPV failure. It is unclear what the actual debris temperature would be for this case. Hence, probabilities of 0.5 are assigned to each branch to represent this large uncertainty.

CEB-9246 7A Case 2 Large Debris Mass in Lower Drywell Early (m[

v-Low (Su}. rheat) 0.5 High (Superheat) 0.5 X.3.2.2.3 Debris Quenched Early (QUENCH._E)

The probability that long term debris cooling will be established is greatly increased if the initial debris pour is quenched soon after being expelled from the vessel. Initial quenching of the debris implies either that the debris has been fragmented to sizes which allow cooling, or if the debris is a continuous " pool" that it is sufSciently shallow to allow cooling by conduction through the layer of solid debris.

The ABWR design makes it extremely unlikely that water will be in the lower dowell prior to RPV failure. Most of the core damage probability is the initiated from a transient. These would not result in water in the lower drywell at the time of vessel failure. Only a LOCA n the reactor drain line would resuh in water entering the drywell. All other LOCAs blow down into the upper drywell (which drains directly to the suppression pool). Hence, water which enters the lower dowell coincident with the expelled debris must come from residual RPV inventory or from in-vessel injection systems which are operating at (or are initiated at) RPV failure. For a MAAP-type melt progression the water contained in the lower plenum is nearly full of (k.)' water at the time of vessel failure. Thus,70,000 kg of water is available to quench the debris.

In addition, water may enter the lower drywell at the time of vessel failure via the passive flooder. If water from the vessel does not enter the cavity, the debris will rapidly heat the lower dnwell, and the flooder will open quickly. For a BWRSAR type melt progression model, there will not be water in the lower plenum at the time of vessel failure. In this case the lower dowell will heat up quickly and the passive flooder will open. A calculation was performed with a modified version of MAAP-ABWR which simulates the BWRSAR melt progression model, described in subsection X.3.2.6 Case LATE indicates that the flooder will open about 30 minutes for this case. Thus, it is very likely that water will be available to quench the initial debr;s expelled from the sessel.

The major parametersjudged to impact the probability ofinitial debris quenching are 1) the mass of debris in the lower drywell following RPV failure,2) the availability of water in the lower dowell, and 3) the initial temperature of the debris. The mass of debris retained in the lower drywellis determined in a preceding event. The initial debris temperature is also determined in a prior event. The source of water depends on the the presumed core melt progression model as described above. In a MAAP type o melt progression the initial availability of water is assured. For a BWRSAR

(') model the water comes e.ither from injection systems which begin to inject at vessel failure or from the operation of the flooder which is considered in CEB-92-46 7A-6

l 1

the next node. Since no credit will be taken for early quenching if a V(S significant amount of debris enters the cavity before lower drywell flooding occurs, the o-der of this question and the late cavity flooding question (CAVWAT_L) question is not important for the BWRSAR case.

Four cases were defined in the DET. These case are:

Case 1 Small Debris hiass and Low Superheat For this case approximately 24000 kg of molten debris are released from the RPV at vessel failure. Since the debris has a low superheat and the debris depth is ven shallow ( < 5 cm) it is highly likely that the debris would be initially quenched.

Quench 0.99 l No Quench 0.01 Case 2 Small Debris hiass and High Superheat As for case 1 approximately 24000 kg of molten debris are released from the RPV at vessel failure. In this case the debris has a high superheat and although the debris depth is very shallow ( < 5 cm) it is somewhat less likely that the debris would be initially quenched by residual RPV coolant p inventory for this case than for case 1.

v Q,uench 0.95 No Quench 0.05 l

Case 3 Large Debris hiass and Low Superheat For this case approximately 94000 kg of molten debris are released from the RPV at vessel failure. The debris depth in this case would be relatively shallow ( < 15 cm ). Since the debris pool is relatively shallow and the debris superheat is low it isjudged that it is likely to be initially i, quenched by residual RPV coolant inventory. 1 Quench 0.75 No Quench 0.25 Case 4 Large Debris hiass and Low Superheat

]

As for case 3 approximately 94000 kg of molten debris are released from the RPV at vessel failure and the debris depth would be relatively shallow ( < 15 cm ). However, despite that the debris pool is relatively shallow, the debris superheat is high and it is judged to be indeterminate )

p whether or not the debris will be quenched by residual RPV coolant-V inventog.

CEB-9246 7A-7 l

p Quench 0.50 U 0.50 No Quench X.3.2.2A Water Enters Cavity Late (CAVWAT_L)

This parameter is used to represent the longer term addition of water to the lower drywell. The lower drywell water addition systems which are considered are the dicsel driven firewater system, any vessel injection which is available late in the accident and the passive Gooder. The passive Gooder system will begin to inject water when the fusible valves located at the ends of the pipes near the dowell floor melt. The fusible ulves on the passive Gooder system are assumed to open when the lower dowell gas tem aerature reaches 500 F (533 K). Assuming a BWRSAR melt progression mode, the fusible valves on the passive Gooder system would open in approximately 30 minutes. For a MAAP type melt progression model the water in the lower drvwell is first boiled off. The debris then begins to heat up. If the debris is quenched during the early boil off phast the debris must reheat resulting in approximately 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> to flooder actuation. If the debris was not quenched early, the flooder opens about 30 minutes after the debris bed dries out. Within these time periods operation of the fire water injection system is also considered. This event is sorting type event quantified (either 0 or 1) based on prior branch decisions in the CET.

Time Remaining Core Debris Falls into Cavity (COREDROP) p)

(

X.3.2.2.5 This event assesses the timing of the entry of the remaining debris into die lower dr>well relative to the timing of the addition of water (i.e.

from the passive flooder or firewater system). If the majority of the debris is held up in the vessel until after war addition begins then debris cooling is substantially more likely than if the bulk of the RPV debris enters the  ;

lower drywell prior to water addition. MAAP calculations indicate that the residual RPV debris will melt and fall into the lower drywell very slowly after vessel failure. This behavior is also typical of BWRSAR type calculations (Reference 1)

Two cases are considered in the quantificat5n of the event. The timing of residual RPV debris entry into the lower drywell is considered to be sensitive to the extent of the accident progression in-vessel at the time of vessel failure. For the case of a small amount of molten debris in the lower RPV plenum at RPV failure (Event 1 in this DET) it is inferred that RPV failure has occurred relatively "early" in the in-vessel accident progression process. Conversely, for a large amount of molten debris in the lower RPV plenum at RPV failure it is more likely that the in-vessel accident

1. S R. Greene, S.A. Hodge, C.R. livman, M.L. Tobias, 'The Response of BWR Mark !!

Containment to Station Blackout Severe Accident Sequences", NUREr/CR-5565, Q("T ORNL/TM 11548, May 1991.

CEB-92-46 7A-8

n progression is further advanced at the time of RPV failure. Consequently, it would be expected that for the case of small initial debris pours the timing

(~) betweer. vessel failure and later debris pours would oc delayed relative to the case of large initial debris pours. Based on insights from ABWR specinc MAAP analyses r.nd from a review of BWRSAR calculations for other BWR sequences the following branch probabilities were estimated.

Case 1 Small Debris Mass in Lower Drywell Early After Late Injection 0.9 Before Late Injection 0.1 Case 2 Large Debris Mass in Lower Dr>well Early After Late Injection 0.5 Before Late Injection 0.5 X.3.2.2.6 Heat Transfer Rate to Overlying Water (HT_ UPWARD)

This event assesses the longer term steady state heat transfer rate -

which characterizes upwards heat transfer from the debris. Three regimes are considered; 1) heat transfer limited by hydrodynamics in an overlying water pool (CHF limit),2) heat transfer limited by film boiling to an

~O V

overlying water pool and 3) heat transfer limited by conduction through a debris crust on the upper debris surface. Nominal values of the heat -

- transfer rate used in the deterministic CCI model'to characterize these three heat transfer regimes are 900,300 and 100 kW/m 2, respectively.

The conduction limit represents conditions where a crust forms on the surface of the debris and water cannot penetrate into the debris bed.

The use of a 100 kW/m 2heat aux is believed to be very conservative. If the debris is not quenched and core concrete interaction occurs, the upper crust will thin to a condition where the upward and downward heat fluxes -

are nearly equal. This will lead to a heat flux much higher than 100 kW/m2 Therefore, this value will lead to very aggressive core concrete interaction.

The hydrodynamic limit represents cases where water can penetrate into the debris bed allowing a much greater effecti.ve debris / coolant heat transfer area. Under these conditions 'he heat transfer rate is limited.by the ability of the water to penetrate the debris bed. The use of 300 kW/m2 is much lower than the typical heat fluxes observed in the experiments performed to date.

The film boiling regime is selected to represent an intermediate heat transfer rate where,- for example, the crust is unstable allowing water to penetrate the debris bed in a limit _e d fashion. The early phase of the q experiments indicate a heat Dux well in excess of 900 kW/m2 before the formauon of a crust.

V CEB-92-16 - 7A-9

/"N - Four cases were identified for quantification. These cases are-U described below.

Case 1 Large Debris Mass in Lower Dr>well Early, Debris Initially Quenched and Residual Core Debris Enters Lower Dr>well After Flooding -

This case is considered the most favorable set of conditions for establishment of a particulated debris bed which would be conducive to water ingression and coolability. The initial phase of the interaction is characterized by large amount of debris which is initially quenched in the lower drywell. Prior to the entry of tlm residual RPV debris the lower drywell-is flooded resulting in the residual debris pouring into a pool of water which is likely to lead to fragmentation, quenching and the establishment of a particle bed. Consequently, a high probability is assigned under these conditions to an upwards heat flux characteristic of a particle bed with water ingression.

900 kW/m2 0.95 306 kW/m2 0.045 100 kW/m2 0.005 Case 2 Small Debris Mass in Lower Drywell Early, Debris Initially Quenched and Residual Core Debris Enters Lower Drywell After Flooding This case is considered to represent nearly as favorable a set of conditions for establishment of a particulated debris bed as was Case 1. In contrast to Case I however, the initial phase of the interaction is characterized by only a small amount of debris which is quenched in the lower drywell. Hence, a larger amount of debris enters the lowe'r drywell after RPV failure than for Case 1, Prior to the entry of the residual RPV debris the lower drywell is flooded resulting in the residual debris pouring into a pool of water which is likely to lead to fragmentation, quenching and the establishment of a particle bed. Consequently, as for Case I a -

relatively high probability is assigned under these conditions to an upwards heat flux characteristic of a particle bed with' water ingression.

900 kW/m2 - 0.9 300 kW/m2 0.09 100 kW/m2 0.01 Case 3 No Initial Debris Quench and Residual Core Debris Enters Lower Drywell After Flooding This case is considered less favorable for establ:shment of a

,-A particulated debris bed which would be conducive to water ingression and .

,. V coolability. The initial phase of the interaction is characterized by failure a

L CEB-92 46 7A 10

to quench the debris soon after RPV failure. However, prior to the entry of

'_] the residual RPV debris the lower dywell is flooded resulting in the residual debris pouring into a pool of water which is likely to lead to V-fragmentation of this debris. However, since the initial debris pour was not-quenched, long term establishment of a coolable particulated debris bed is somewhat uncertain. Consequently, a lower probability has been assigned for the most favorable debris bed configuration compared with Cases 7 and 2.

900 kW/m2 0.5 300 kW/m2 0.4 100 kW/m2 0.1 Case 4 Residual Core Debris Enters Lower Dr)well Prior to Flooding This is considered the least favorable set of conditions for establishment of a particulated debris bed which would be conducive to water ingression and coolability. For this case the bulk of the residual core debris enters the lower drywell prior to lower drywell flooding. This could lead to formation of a molten pool undergoing concrete attack. Later water addition, instead of particulating the debris may lead to crust formation limiting the ability of water to penetrate into the debris.

O 900 kW/m2 0.1 b 300 kW/m2 0.6 100 kW/m2 0.3 X.3.2.2.7 Core Debris Concrete Attack (CCI)

This event characterizes the nature of the debris concrete attack.

Three branches are considered. The No CCI branch represents cases where the little or no debris concrete attack would be expected. Wet CCI represents cases where CCI occurs in the presence of an overlying water .

pool and Dry CCI is for cases where the lower drywell was not flooded.

Case 1 Lower Drywell Not Flooded The. Dry CCI case occurs for all sequences where both active injection l- and the passive flooder fail to supply water to the lower dgwell after vessel failure. Under these conditions Dry CCI is assured.

No CCI 0.0

-Wet CCI 0.0 Dry CCI 1.0

,q O

CEB-92-46' 7A m

+

. .I 4.

[L' ' Case 2 Lower Dr)well Flooded, Upward Heat Transfer 900 kW/m 2

- For casec where the lower drywell is flooded MAAP analysis and supplemental hand calculations indicate that if the upward heat transfer is :

above about 300400 kW/m2 then the debris bed will be coolable.

No CCI 1.0 Wet CCI 0.0 .

Dry CCI 0.0 Case 3 Lower Drywell Flooded, Upward Heat Transfer 300 kW/m2 For cases where the lower dr)well is flooded MAAP analysis and sup alemental hand calculations indicate that if the upward heat transfer 'is . -

in t ie range of about-300 kW/m2 then the debris bed should be coolable. j Since this case represents a range of upward heat transfer regimes-(200-400 "

kW/m )2 and the lower part of this range may not in all cases be coolable "

the following probabilities were assigned. .

No CCI 0.75 -

+

Wet CCI 0.25 l' - Dry CCI 0.0 -

Case 4 Lower Drywell Flooded, Upward Heat Transfer 100 kW/m2- D For cases where the lower drywell is flooded MAAP analysis and . .

supplemental hand calculations indicate that if the upward heat transfer is. '

below about 200 kW/m2then the debris bed will not 'ae coolable.

g No CCI 0.0 L ..

Wet CCI 1.0 Dry CCI 0.0 L

X.3.2.2.2 Pedestal Resistance to CCI

., - This section. describes the decomposition' event tree'(DET) analysis .

H 'used to assess the probability of pedes'al failure as a result of radial core - _

g concrete (CCI) attack inithe lower drywell after reactor vessel failure. The{

! DET is shown in Figure 7A 2.~ Pedestal wall failure is considered to be i b sensitive to 1) the nature of the CCI (i.e. whether wet or dry),2);whether the; debris ~ spreads from lower drywell into suppression pool following radial penetration lthrough the pedestal wall to _the wetwell'drywell. connecting; l3-g . vents, and 3) the extent of radial erosion compared.to downward erosion.

? -.The lower drywell will be flooded in most cases as a result of either active <

CEB-92-46 - 7A-12 3 '

yd _

i p injection systems such as the firewater addition system or via passive Q . injection ~through the lower drywell flooder.

X.S.2.2.2.1 Core Concrete Attack (CCI)

This event characterizes the nature of core concrete attack, Three branches are considered.

No CCI Wet CCI Dry CCI .

1 The No CCI branch represents cases where there is little or no concrete-attack. Wet CCI represents cases where CCI occurs in the presence of an overlying' water pool. Dry CCI is for cases where the lower day,' ell was not L flooded. The rate of CCI is higher for cases with dry CCI.

This event is a sorting type event which assigns a probability of 0 or 1 depending on the branch taken in the previous CET event. i X.3.2.2.2.2 Suppression Pool water floods lower drywell after downcomer 3 penetration (SP IhCRESS) l p  !

' g~ This event assesses if suppression pool water will flood into the lower drywell after the erosion front reaches the wetwell dr>well connecting vents. The vents are imbedded in the pedestal. If 25 cm of the pedestal concrete is eroded, the ablation front will reach the inner surf ace of the connecting vents. It is considered quite likely that this will result in water ingression and flooding of the lower drywell.

This event is only significant for Dry CCI sequences where the lower J drywell is not initially flooded by either active injection or the passive flooder. The probabilities are assigned based on judgement.

SP Ingression 0.95 No SP Ingression 0.05 '

X.3.2.2.2.3 Debris Flows from lower drywell to suppression pool after a downcomer penetration (WW DEB) _ {

This event assesses whether a significant amount of the molten debris will flow from the lower drywell into the-suppression pool following j penetration of the wetwell drywell connecting vents. After 25 cm of radial 1 erosion the ablation front will reach the inner surface of the downcomers.. 1 The floor of the lower drywellis above the bottom of the connecting vents, j

'n which, in turn, are above the floot of the wetwell. Thus, once the f!

..Q downcomers are breached, a flowpath exists from the lower drywell into the CEB-92 7A .

a

~ suppression pool. Flow of a significant portion of the molten debris into the suppression pool will increase the debris surface area in contact with

-(v) water and decrease the debris depth in the lower drywell. Although there is a great deal of uncertainty in this behavior, it is considered fairly likely that the debris will flow into the suppression pool.

WW Debris 0.7 No WW Debris 0.3 X.3.2.2.2.4 Ratio of radial to axial erosion (RAD _ EROS)

Given that CCI is occurring, this event assesses the ratio of the radial concrete crosion to the downward erosion. Three branches are considered 1/5,1/3 and 1/1. CCI experiments have generally demonstrated significantly more downward concrete penetration than radial penetration.

It is hypothesired that radial erosion is limited because the concrete decomposition gasses establish a gas film between the debris pool and the -

concrete walls.' This gas film acts to insulate the concrete sidewalls, and to-convect debris heat upwards limiting the heat transfer to, and ablation of the concrete sidewalls. Conversely, the gas film at the bottom surface of the pool would be unstable due to the heavier overlying debris pool. The density difference would cause the lower gas film to_ collapse, allowing contact of the debris with the concrete. This difTerence in gas film behavior would limit the sideward heat transfer compared to the downward O heat transfer.

O In the BETA series of debris concrete experiments conducted at the KfK research center in Germany, downward erosion rates exceeded sideward erosion rates by a factor from 3 to greater than 5. For example, in the high power CCI experiment BETA V1.8, the downward erosion was measured to be approximately 40 cm and the sideward erosion was only about 2 cm (1/20 sideward to downward erosion ratio). For the low power experiment V6.1, the downward crosion was 35 cm and the sideward erosion was 10 cm (1/3.5 ratio).

Based on the CCI experiments, and the generally accepted model described above, it seems appropriate sa assume that downward erosion is strongly favored 'over sideward crosioi . Consequently, larger probabilities are assigned to the 1/5 and 1/3 branches than for the 1/1 branch.

However, since some residual uncertainty remains as to the appropriate assumption for the extent of radial erosion for large reactor scale situations, a probability of 0.1 is assigned to the 1/1 erosion branch.

~

Radial to axial erosion ratio 1/5 0.45 Radial to axial erosion ratio 1/3 0.45 i i

g Radial to axial erosion ration 1/1 0.1 q l- O l i

j

~

CEB-92-46 . 7A-14 ,

i il X.3.2.2.2.5 Pedestal Failure (PED)

(

This branch assesses the probability of pedestal failure as a result of

, excessive radial concrete crosion of the lower drywell pedestal wall.

Structural analysis of the pedestal indicates that the loads can be supported without yielding if only the outer shell and 15 cm of the steel webbing remains intact. Thus, for a total wall thickness of 1.7 m, the lower limit for the amount of radial erosion which can be sustained without pedestal structural failure is 1.55 m. However, since the total depth of the pedestal is 1.7 m, erosion to the full 1.7 m depth will obviously result in pedestal failure. Additional discussion of the pedestal strength under radial concrete crosion is presented in subsection X.3.2.4.

Analyses were performed to estimate the extent of concrete erosion in the lower drywell under a variety of conditions. The results of these analyses are summarized in subsection X.3.2.5. Four cases were considered in the DET for quantification cf pedestal wall failure. These cases are described below.

Case 1: Debris flows into suppression pool after downcomer penetration This case represents sequences where a substantial amount of the core debris relocates into the suppression pool after downcomer penetration.

O This is represented by deterministic calculations FMX100, FMXCSP and D NFlood. The calculations indicate that the increase in the pool surface area results in either a coolable debris configuration, or greatly reduced radial erosion rates. Consequently, the likelihood of sufficient radial penetration to fail the pedestal in this case is considered to be remote.

No Pedestal Failure 1.0 Pedestal Failure 0.0 Case 2: Wet CCI with no Debris flow into the suppression pool after downcomer penetration .

For sequences where CCI was predicted to occur in the presence of an-overlying water pool with no debris relocation to the suppression pool, the maximum amount of downward concrete erosion at 50 hours5.787037e-4 days <br />0.0139 hours <br />8.267196e-5 weeks <br />1.9025e-5 months <br /> was 1.55 m (Case FMX1P). Using this value for the amount of axial erosion, the radial erosion depth is estimated for the three cases. Comparing this value to the pedestal capability of 1.55 m, the following estimates are made for the probability of pedestal failure:

RAD _ EROS 1/5 1/3 1/l' No Pedestal Failure 1.0 1.0 0.5 3

(d Pedestal Failure 0.0 0.0 0.5 CEB-92-16 7A-15

b N Case 3: Dry CCI with no debris flow into su,pression pool and no late suppressiori pool water ingression into the .ower drywell This case represents case DRY in the deterministic analysis, in this cace the debris is assumed to remain dry for the entire duration of the accident. No flow of either water or debris through the wetwell dr>well connecting vents is presumed to occur when the ablation front reaches the '

the vents. For this case the axial ablation depth at 50 hours5.787037e-4 days <br />0.0139 hours <br />8.267196e-5 weeks <br />1.9025e-5 months <br /> was calculated to be 2.5 m. Using this value to estimate the radial erosion depth for the three radial to axial crosion ratios, the split fractions are are assigned based on the pedestal capability:

RREROS 1/5 1/3 1/1 No Pedestal Failure 1.0 0.99 0.0 Pedestal Failure 0.0 0.01 1.0 Case 4: Dry CCI with no debris How into the suppression pool and late suppression pool water ingression into the lower drywell The case in which the debris is initially dry, but becomes flooded with ,

water after the ablation front reaches the wetwell drywell connecting vents is considered to be slightly better than Case 3. In this case the debris is

. V(,) assumed to remain in the lower drywell throughout the period of CCI.

Therefore, the split fractions assigned are:

RAD _ EROS 1/5 1/3 1/1 No Pedestal Failure 1.0 0.99 0.5 Pedestal Failure 0.0 0.01 0.5 X.3.2.3 Deterministic Model for Core Concrete Interaction As described above, several key parameters influence the potential for concrete erosion in the presence of an overlying water pool. An analytical-tool was selected to investigate the impact that these parameters have on CCI, containment pressurization, opening of the over-pressure protection system, and possible fission product release. MAAP-ABWR was selected since, with a few minor code modifications, it was capable of investigating the key parameters identified in the DET. MAAP-ABWR allowed the impact of parameter variations to be carried out through containment pressurization and fission product release.

A few simple code modifications were made to allow the user to control the debris coolability and to siinplify the specification of the severe accident scenario. These changes are summarized below.

O) R CEB-92-46 7A-16

1

1. Subroutine PLSThi was modified to alto,v the user to specify the.

[A>Y ' upward heat flux. Niodel parameter, FCHF, was redefined to be the upward heat flux in watts /m2, All other debris-to-water heat transfer mechanisms were disabled in PLSThl.

2. The following actions were added to the hfAIN routine:

- If lower dr>well gas temperature exceeds 533 K - o}ien passive flooder If radial erosion exceeds 25 cm - allow debris to spread to wetwell and allow water to flood the lower drywell

- If radial erosion exceeds 50 cm - fail drywell with an area of ADW1.EK (user input)

If upper drywell wall surface temperature exceeds 533 K - begin to leak out of the upper drywell as specified in subsection 19F.3.2.2.

The major assumptions included in' the htAAP analysis were:

1. CCI experiments have generally demonstrated significantly more downward concrete penetration than radial penetration it is-hypothesized that radial crosion is limited because the concrete

/D decomposition gasses establish a gas film between the debris V_ pool and the concrete walls. This gas film acts to insulate the concrete sidewalls, and to convect debris heat upwards. limiting the heat transfer to, and ablation of the concrete sidewalls.

Conversely, the gas film at the bottom surface of the pool would be unstable due to the heavier overlying debris pool. The density difference would cause the lower gas film to collapse, allowing contact of the debris with the concrete. This difference in gas film behavior would limit the sideward heat transfer compared to_-

the downward heat transfer.

In the llETA series of debris concrete experiments conducted at-the KfK research center in Germany, downward crosion rates exceeded sideward erosion rates by a factor from 3 to greater than

5. For example, in the high power CCI experiment IlETA VI.8,' the downward erosion was measured to be approximately 40 cm and the sideward erosion was only about 2 cm (1/20 sideward to downward erosion ratio). For the low power experiment V6.1, the downward erosion was 35 cm and the sideward crosion was 10 cm -

(1/3.5 ratio).

Ilased on the CCI experiments, and the generally accepted model described above, it seems appropriate to assume that the ratio of radial to axial attack is 1/5. However, this parameter is included C') as a parameter in the DET for pedestal erosion since the ratio is still uncertain.

CEB-92 46 7A 17 1

li

\~

(7 Since MAAP assumed tint radial and axial penetration were

(/ identical, the axial ablation numbers were multiplied by 1/5 to obtain an estimate on the radial attack depth,

2. The heat transfer from the debris to the water was assumed to be equal to the user specified value throughout the transient. -

Other than the changes described above, the standard MAAP-ABWR code was used to quantify the CCI decomposition event tree.

X.3.2.4 Pedestal Strength The configuration of the ABWR pedestal is shown in Figure 7A-3. The width of the pedestal is 1.7 m. The design consists of two concentric steel cylinders connected by steel web stiffeners. Ten wetwell-drywell connecting.

vents run through the annular region between the cylinders. The remainder of the space is filled with concrete. If significant core concrete attack occurs, the strength of the pedestal could be compromised as the pedestal is eroded. The strength of the pedestal after it has undergone erosion is examined to determine the maximum erosion depth allowable to ensure that the pedestal does not collapse.

The pedestal is designed based on the maximum stress obtained in the steel plates. The strength of the concrete is neglected. The allowable

O stress in the steel plates is 0.6 times the yield strength, neglecting C temperature. The calculated stress without seismic loads in the ABWR pedestal is 0.4 times the yield strength.

For design analysis the largest single load is the accident temperature.

If core concrete interaction were to take place as a result of a severe accident, the inner plate of the pedestal would melt..Without a continuous inner plate the moment induced by the differential temperature .

disappears. It is expected that any temperature induced moments acting along the stiffeners will be. strain limited.- Therefore, they will not reduce the capability of the outer plate, in order to estimate the allowable ablation depth, the seismic and thermal loads are removed and the remaining loads are calculated. No attempt was made to take credit for the relocation of fuel from the vessel onto the floor of the drywell. The strength of the remaining concrete is neglected.- The loads are compared to the yield strength of the remaining pedestal steel. Therefore, this calculation corresponds roughly to a senice level C type of calculation.

The results of the calculation show that the outer shell of the pedestal-plus 15 cm of the web sdffeners are required to maintain the pedestal loads below yield This limit i:. used as a consenative estimate of the pedestal ultimate capability after erosion. The total pedestal width is 1.7 m.

O

.a Therefore, pedestal integrity is ensured for ablation depths up to 1.55 m.

CEB-9246 7A-18

~. -

p X.3.2.5 Application of CCI Model to ABWR V

The deterministic code used for investigating core-concrete interaction in the ABWR was described in Section X.3.2.3, This section will describe the evaluations that were made to support the quantification of the CCI decomposition event tree.

X.3.2,5.1 Sequence Selection The MAAP-ABWR code, as modified for this application, allowed for a great amount of flexibility in analyzing the impact of key parameter variations on core-concrete attack. The following lists the key parameter variations that were investigate'd:

Upward heat transfer to overlying water pool Mass of debris discharged from vessel Mode of fission product release from containment

- Flooding of lower drywell resulting from radial penetration of vertical connecting vents ,

Debris spreading related to radial penetration of vertical connecting vents The base case sequence selected to investigate core concrete interaction was the low pressure loss of injection scenario. This event was -

initiated by a transient with the assumption that all injection was unavailable. The RPY was depressurized manually when the core level dropped below 2/3 core height. Without coolant injection, the core melts and slumps into the lower vessel head. Local penetration failure occurs and the debris is discharged into the lower drywell. Table 7A 1 provides a chronology of the events up until the vesselis failed.

Table 7A-2 defines each o." the sequences analyzed and provides a summary of the results. The first column gives the case designation along with reference to specific notes.~ Columns two through four provide the relevant sequence definition information. For purposes of demonstration, all cases were executed for the dominant sequence, a low pressure loss of injection sequence with a containment pressure at the time.of vessel breach of approximately 1 atm. The upward heat flux was varied between 100,300, and 900 kW/m .2 A value of 100 kW/m2 was selected to approxima_te the heat transfer associated with a stable crusi formation where the upward loss is controlled by conduction of heat through the crust. A value of 300 kW/m2 was selected to represent limited water ingression into the debris bed with the upward heat transfer being controlled by film boiling. The largest nlue used represents critical heat flux limits for debris cooling.

- Further discussion of these values is included in subsection X.3.2.2.6 s

l CEB-92-46 7A-19

1 1

As run in its standard manner MAAP-ABWR calculates that 60% of the L j j total core inventory was released from the vessel. The remaining 40% was calculated to be held up in the core with the decay heat being radiated to the vessel wall and convected into the upper dr>well. The 40% remaining behind is typically the outer peripheral bundles which have low decay heat.

To support the DET quantification additional cases were run assuming that 100% of the core was discharged from the reactor vessel. This has two major .

innuences on the containment behavior. Without the peripheral bundles in the core, the drywell heatup is reduced. Second, the added core mass on the lower drywell floor will influence the calculation of core-concrete attack, debris coolability and containment pressurization.

X.3.2.5.2 Summary of Results Table 7A-2 summarizes the results of the deterministic analyses for the ABWR. The following general conclusions are indicated by these results:

1. For all sequences with successful operation of the flooder, radial concrete erosion was less than the structural limit described in subsection X.3.2.4. Radial attack does not pose a significant challenge to containment.

2. For sequences with operation of the containment overpressure protection system, due to suppression pool scrubbing, the fission product release is dominated by noble gas.

3. Release times for cases with the passive flooder are on the order of 20 hours2.314815e-4 days <br />0.00556 hours <br />3.306878e-5 weeks <br />7.61e-6 months <br /> after the initiation of core damage (defined as onset of melting).

L 4. The extended time period between vessel breach and rupture disk

[' actuation-(cr containment failure) provides for a substantial reduction in '

the amount of fission product released from containment.

5. Using experimentally-based values for the upward heat transfer (Section X.3.2.1) would result in debris cooling in the ABWR and early-termination of the core concrete attack. Therefore, the lower bound for upward heat transfer is conservatively assumed to be 100 kW/m2. This is done in order to obtain substantial concrete crosion and demonstrate the l

robustness of the containment design if the debris is not quenched.

j 6. For the dominant scenarios with successful operation of fire water-to provide water to'the debris, the time from onset of melting to fission l

product release is 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> from the beginning of the accident for all upward heat transfer rates.

L A set of plots for case FMX100 case are included in Figures 7A-4a through 7A-4e. This case demonstrates long term core concrete interaction, but is otherwise typical of the conditions analyzed.

i-

!= CEIk92-46 7A-20 ,

L

X.3.2.5.3 Initial Concrete Attack due to impinging Corium Jet At vessel failure, core material is discharged from the RPV'onto the Door of the lower drywell At low RPV pressures, the discharge rate of the debris is controlled by grasity and the vessel breach area in the lower head.

From analyses performed for FCI calculations, subsection X.2.7.6.2.2.2, it is assumed that ten gnetrations failed. This results in a maximitrn corium discharge rate of 6000 kg/s. The total failure area is 0.145 m 2. Assuming a density for corium of 8000 kg/m 3, a discharge of 6000 kg/s cormsponds to a corium velocity of 5 m/s. The following calculation estimates the initial concrete attack depth resulting from this impinging corium jet.

The model from the MAAP subroutineJET (Reference 1) was used to compute the concrete attack from an impinging jet of corium. The stagnation point heat transfer coefficient between the corium jet and the concrete is approximated by the expression.

Nu - - 1.14 6 k (7A-1) or FPc u

.-- m c

[ h = 1.14k'*IDjet Pcm where:

kcm = Corium thermal Conductivity pcm = Comum viscosity ue= Velocity of the corium stream impinging on the Door Djet = Diameter of thejet h = Heat transfer coeflicient pcm = Corium density Nu = Nusselt number Re = Reynolds number The corium velocity at the cavity floor is given by, u c= uo y gt raji (74 3)_

O

'Q l. MAAP 3.0 B Computer Code Manual, EPRI NP-7071 CCML, Volume 2. November 1990.

CEB-92-16 7A-21

c l

where

(]

LJ.

uo = Velocity of the cerium expelled from the reactor vessel g = Acceleration of gravity and trai si defined by 1 2 uo trei + -gtfg3 = zy (7A-4) where zyis the elevation of the reactor vessel above the lower dgwell floor.

A crust of frozen corium forms on the concrete and the ablation process is the same as at the reactor vessel penetration. Thus, the concrete ablation velocity is given by h(Tcm -Tcnp) ucn =

Pcn c pcn(Tc np- To)+ Tcn (7A-5) where Tcm = Bulk corium temperature pcn = Concrete density c pcn = Concrete specific heat ,

Ycn = Concrete latent heat Tenp = Concrete melting temperature _

To = Initial concrete temperature Substituting the corium velocity and the ABWR specific geometrical parameters into the above equations. results in an ablation rate of approximately 1 cm/sec. With the debris being discharged over 5 second,,

the resulting ablation depth is 5 cm. This would only occur in the central portion of the lower drywell, and would in no way threaten the integrity of the structures.

3.2.6 Sensitivity to Various Parameters Also included in Table 7A-2 are other analyses that address possible

• sensitivities to modelling assumptions. These results are described below.

Case DRY

-(N This case was run _ssuming that the passive flooder did not open and

'V that, even after radial penetration of the vertical vent pipes, water was not 1

l CEB-92-46 7A-22

hd N

Q introduced into the lower drywell. The drywell began to leak at about 20 hours2.314815e-4 days <br />0.00556 hours <br />3.306878e-5 weeks <br />7.61e-6 months <br /> and resulted in a slow, low magnitude, release of fission products.

(/

Case DWFAIL This case is identical to case FMX300 except that the drywell was assumed to fail at the COPS set point. Due to the long time between vessel breach and containment failure, the fission products settle out very effectively and the result is a low magnitude release.

Case FMX1P This case was identical to case FMX100 except that the debris is assumed to not spread into the wetwell after penetrating the vertical connecting vents. The results indicate no sensitivity to this assumption.

The radial attack at 50 hours5.787037e-4 days <br />0.0139 hours <br />8.267196e-5 weeks <br />1.9025e-5 months <br /> is 31 cm for a ratio of radial to axial attack of one to five.

Case NFLOOD This case was identical to case ABWR100 except that the firewater addition system and passive flooder were not operational. Therefore, the debris was initially dnj. After 25 cm of radial erosion, the debris was assumed to spread into the wetwell and water from the suppression was Q introduced into the lower drywell. The results indicate more concrete V erosion with the COPS actuating at 17.4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br /> compared to 19.1 hours1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br />.

Case FIRE

^

This sequence was identical to FMX100 except that the firewater system was used to add water to the debris. Due te the addition of cold water, the pressurization of containment due to neam was reduced and the COPS was not predicted to open until 24.6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br /> as compared to 17.6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br /> for the case with passive flooder operation.

The overall conclusions from the sensitivity analyses are that the ABWR containment design is quite insensitive to the uncertainties associated with core concrete interaction. The concrete crosion rates are -

consistent with other published results (Reference 1) and do not pose a serious threat to containment integrity, Operation of the COPS provides for a scrubbed release of the fission products and greatly limits the risk to the public.

Case LATE This sequence was identical to case DRY except for a delayed vessel failure. The RPV was assumed to fail after all of the water in the lower f~

v-

1. NL: REG /CR 5565, op. cit.

CEB-92-46 7A-23

plenum had boiled away and the debris heated up to the eutectic melting (qj point (2501 K). Vesel failure occured at 5.3 hours3.472222e-5 days <br />8.333333e-4 hours <br />4.960317e-6 weeks <br />1.1415e-6 months <br /> into the sequence as compared to 1.5 hours5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br /> for the base case. Since there was no water discharged with the core debris at vessel failure, the gas temperature quickly increased to above the flooder actuation temperature. The flooder was assumed not to work for this case. The purpose of te run was to obtain an estimate of the time period between vessel failure and flooder actuation.

The MAAP analysis conservatively assumes that the gas mus must reach 533 K before the flooder can open. In this case it took about c. hour before the grc reached 533 K. Factoring in the difference between me wall surface and the gas temperature, the flooder would be expected to open within 30 minutes after discharge of the core debris. All other aspects of this run were similar to the DRY case.

X.3.2.'i Impact on OfTsite Dose The effect of the maximum core concrete interaction source term on a relaease with operation of the rupture disk is shown in Figure 7A-5. The cases with rupture disk are the only risk significant release categories which would be impacted by core concrete interaction (The other sequences are cases with early containment failure due to DCH.) As the Figure 7A-5 clearly shows, CCI does not have significant impact on the offsite dose.

X.3.2.8 Conclusions This section investigated the impact of core <oncrete interaction on the ABWR containment response. First, detailed DETs were developed to address all of the key parameters that influence CCI. Then, several determinstic analysis were carried out to support quantification of the trees. The following summarizes the important conclusions of the CCI investigation:

1. For the dominant core melt sequences that release core material into the containment,90% result in no significant CCI. Virtually no sequences have dry CCI.

2. Even for the low frequency cases with significant CCI, radial erosion remains below the structural limit.

3. The fission product release mode is dominated by operation of the containment overpressure protection system. The release, which occurs at about 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />, is not distinguishable from a case with no CCI.

4. Experimental results indicate that sufficient upward heat transfer to an overlying water pool would exist in the ABWR lower drywell to cool the debris.

CEB-92-46 7A 24

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SUMMARY

OF CCI DETERMINISTIC ANALYSIS FOR ABWR g

tr!

Y

$ Debris Mass Radial H2 h Containmen Upward at Vess. Fail Attack Generated Time of FP Fission Product Release Fraction Press. at Heat Trans. (Frac. of Tot. at 50 hrs. at 50 hrs. Release Mode of from Containment Case # essel failui (kw/ma2) Inventory) (me'.c 3) (Kg) (hours) Release PG Csl S-ABWR100 1 Atm. 100 0.6 0.22 1813 19.1 COPS 1.0 2E-06 3E-09 ABWR300 1 Atm. 300 0.6 9E-07 122 23.3 COPS 1.0 2 E-10 ' 2E-12 ABWR900 1 Atm. 900 0.6 7E-06 122 23.2 COPS 1.0 3E-11 2E-12 FMX100 1 Atm. 100 1.0 0.25 2130 17.6 COPS 1.0 1E-06 1E-08 FMX300 s Atm. 300 1.0 7E-03 154 19.3 COPS 1.0 1 E-08 3E-15 FMX900 1 Atm. 900 1.0 7E-04 111 19.1 COPS 1.0 1 E-08 2E-14 FMXCSP (1) 1 Atm. 100 1.0 0.25 2126 15.7 COPS 1.0 4E-07 3E-10 SENSITIVITY RUNS CR( 1 Atm. N/A 0.6 0.50 4990 19.8 DWT 0.34 4E-03 1E-05 DWFAIL 1 Atm. 300 1.0 7E-03 154 19.3 DWF 1.0 8E-04 2E-10 elRE 1 Atm. 100 1.0 0.25 2131 24.6 COPS 1.0 SE-06 4E-10 FMX1P (2) 1 Atm. 100 1.0 0.31 2762 17.6 COPS 1.0 1 E-06 1 E-08 NFlood (3) 1 Atm. 103 0.6 0.25 2127 17.4 COPS 1.0 SE-07 SE-10 LATE (4) 1 Atm. NA 1.0 0.31 2697 20.0 DWT 0.23 6E-01 9E-09 Notes:

COPS - Containment Overpressure Protection System DWT - Drywell Leakage occurs through penetrations DWF -- Drywell Failure (0.0973 m^2)

(1) - FMX100 Run with five times steel mass (2) - Penetration into connecting vents does not cause debris spread

$-(3)- - Flooder not operational, Radial attack results in penetration to WW and debris spread

@ (4) - Vessel failure. assumed to occur after lower plenum water boiled away and debris reheats

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e

. t

. ' e 2 r cd m n

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. . ' o u f l F

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. ' i a t e 1

. ' a H l

. ' u d 1

c r e . '

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. ' e r

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I

_ _:___~__ __ _-

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ogn n one n ~ oor om* o, n U age-F ._

e2xeg Leaaa

, u

,i l  ;;-  ;!  :  :-;, t i; 1  : 2

l O

Figure 7A-5 Whole Body Dose at 1/2 Mile as a Probability of Exceedence 1 _ __ , ._ _ _ . - - _ _ . _ _ _

$ 'ix y ..

\\  ;

- \\

$ case 2

= x O! o.1 d \m 9

O

's \

k \

a \

g O

C

' 0.01-0 25 5 10 15 20 EARLY WHOLE BODY DOSE IN REM O

ATTACIIMENT 78

,o$ .

His material was submitted to the Staff on August 7,1992.

l v

). This information will be incorporated into the AllWR SSAR at a future date.

D. Prevention of Molten Debris Ingression into Lower Drywell Sump D.1 1ssue During a hypothetical severe accident in the AllWR, molten core debris may be present on the lower drywell (LD) floor. The EPRI ALWR Requirements Document specifies a floor area of at least 0.02 m2/MWt h to promote debris coolability. This has been interpreted in the ABWR design as a requirement for an unrestricted LD floor area of 79 m2.

The ABWR has two drain sumps in the periphery of LD floor which could collect core debris during a severe accident if ingression is not prevented. If ingression occurs, a debris bed will form in the sump which has the potential to be thicker than the bed on the LD floor. Debris coolability becomen more uncertain as the thickness of a debris bed increases.

The two drain sumps have different design objectives. One, the floor drain sump, is designed to collect any water which falls on the 1.D floor. The other, the equipment -.

drain sump. collecu, was:r leaking frorn valves and piping.

D.2 Proposed Design D' A protective layer of refractory bricks - a corium shield - could be built around the-sum as to prevent corium ingression. The shield for equipment drain sump would be solic except for the inlet and outlet piping which would go through its roof. The shield for the floor drain sump would be similar ext.ept that it must have channels at floor level to allow water which falls onto the LD floor to flow into the sump. The height of the channels would be chosen so (hat any molten debris which reaches the inlet would frecie before it exited and spilled into the sump. The width and number of the channels would be chosen so that the required water flow rate during normal reactor operation is achievable. A sketch of a concept for floor drain sump shield is shown in Figure 711-1.

The walls of the equipment drain sump shield (solid shield) only have to be thick'

- enough to prevent the elevated debris temperature from degradmg the shield l internal structural support. The walls of the floor drain sump shield (channeled l shichl) must be significantly thicker so that molten debris flowing through the l channels has enough residence time to ensure debris solidification. ,

l L Both shields would extend above the LD floor to an elevation greater than the expected maximum height of core debris. Thus, no significant amounts of debris will collect on the shield roofs. The solid shield will be placed directly on top of the LD floor. ' The channeled shield will have refractory bricks embedded into the LD -

floor beneath the shield to prevent core-concrete interaction involving the molten

(

U]

debris in the channels.

DIIM 9207 7tbl

n

( ) The analyses presented in sections D.4 and D.5 proside a basis for sizing the U

proposed design of the floor drain sump corium shield.

D.3 Success Criteria for Proposed Design For the proposed design to be considered successful, it must satisfy the following requirements.

a) Melting Point of Shield Material Above Initial Contact Temperr.ture The shield wall material shall be chosen so that its melting temperature is greater than the interface temperature between the debris and the shield wall.

b) Channel Length The length of the channels in the shield must be long enough to ensure that a plug forms in the channel before debris spills into the sump. The freezing process is expected to take on the order of seconds or less to complete.

c) Shield lleight, Iluw, Above lower Drywell Floor The shield height above the lower dnwell floor shall be chosen to ensure long term ,

debris solidification. The freezing process will be complete during the time frame -

when the shield walls are behaving as semiinfinite solids. In addition, the shield Abl must be tall enough to prevent debris from accumulating on the roof of the shield, d) Shield Depth, li iw, Below lower Dnwell Floor The shield depth of the below the lower dr>well floor shall be chosen to ensure long term debris solidification.

c) Water Flow Rate The total flow area of the shield channels shall be great enough to allow water flow a rates sated in the Technical Saecifications without causing excessive water pool formation in the lower dr>wel..

Section D.6 contains example success calculations for these requirements, except c),

for a chosen channel height of I cm.

D.4 . Analysis of Shield Freezing Ability Heat transfer and phase change analyses are presented in this section to determine.

the feasibility of a channeled shield to prevent molten debris ingression into the -

floor drain sump. Two time frames were considered. First, a freeze front analysis was performed for early times.(seconds or less) to determine the time rer uired to form a plug. The long term ability of a plug to remain solid was determined using a steady state analyses.

.( .

DBM-9207 7B f') D 4.1 Assumptions

\ f The major assumptions invoked in the analyses and their bases follow.

1. Molten debris enters the channel with negligible super heat.

Molten debris interacts with structural material (steel, concrete, etc.) and the lower drywell ernironment as it passes from the vessel, contacts the LD floor and spreads to the shield. This interaction depletes the molten debris of any super heat and can result in cutectic fonnations. The melting temperature of core debris which has undergone little interaction is approximately 2500 K.

Significar.t interaction with the concrete floor reduces the debris melting temperature to approximately 1700 K.

2. During the freezing process, the temperature profile of the solidified debris rapidly obtains its steady state value.

This assumption introduces little inaccuracy because: a) the heat conduction coefficient in the solidified debris is significantly larger than that of the shield material, and b) the depth of the solidified debris is considerably less than the height of the shield.

3. Heat transfer within the channel and shield is one-dimensional.

pV The height of each channel is much less than its length. The heat transfer in the shit ld material is low enough that any heat transferred from debris contacting the shield wall outside of the channel does not affect the temperature along the channel untillong after a plug has formed. Any heat transfer to the shield material between adjacent channels enhances the debris freezing process.

4. The shield wall acts as a semi. infinite slab with an initial temperature of 330 K during the initial freezing process.

The properties of shield cause it to be a poor conductor of heat. The penetration depth during the short duration of the freezing process is on the order of a ten millimeters. The small increases in LD temperature prior to the presence of core debris does not significantly alter the shield temperature from its value during normal plant operation.

5. Core debris is not expected to enter the LD until at least two hours after accident initiation. This places decay heat level at approximately one. percent of rated power.

Core debris will not enter the lower drywell before about two hours for any credible severe accident, see ABWR SSAR section 19E.2.2.

t 73 6. The decay heat genmation in the debris is negligible compared to the rate of j )- latent heat generauon during the freezing process.

J DBM-9207 7B , ,

O This assurnption was verified during the analysis.

V 7. The thermal conductisity and thermal diffusivity of debris in solid and liquid phases are the same.

D,4.2 Irdtlal Freerhig of Molten Debris in Channel .

If the floor drain sump shield fulfills its design objective, a debris plug will form in the channel before molten corium has a chance to traverse the channel and reach

• the sump. Mo! ten debris enters the channel at a significantly elevated temperature (2500 K to 1700 K) compared to the shield wall (~ 330 K). The walls absorb heat from the debris because of the large temperature difference. Since the debris contains -

negligible su aer heat, any heat loss by the debris results in freezing. Freeze fronts start at the channel walls and move toward the center of the channel. The leading edg, of the frecie front will stay at the melting temperature of the debris. The heering process is symmetric about the cen'erline of the channel because the same amount of heat is transferred through each wall while they are behaving as semi-

!niinite slabs. The channel walls behave as semi-infinite slabs during the freezing process because the heat conduction rate through the wall material-is low compared to the release rate oflatent heat. A sketch of the freezing process is shown in Figure 71k2.

a) Freezing Time f~') The temperature profile in the crust, assuming it quickly reaches its steady state V shape,is (Reference 1) 2' T - Tr,m x T 1

,rc (x) = 9 c 1- x + - + , + Tr m 2 1, 2 2kr( Lcs 2

(7[pl) where Tc(x) is the temperure within the crust x is the crust coordinate measured from the crust centerline q is the heat density of the crust 1x is the half - thickness of the crust kr is the thermal conductivity of debris T. is tlx interface temperature between the wall and debris Tr,m is the melting temperature of debris .

l The energy balance at the freeze front is L

L Frank P. Incropera and David P. DeWitt, Fundamentah of Heat and Afass Transfer, 2nd l Ed., John Wiley and Sons,1985, pp. 854 l

l-l DBM-9207 7B-4

l l

1 a

?

V qi := -k r dx n- ts g where qj'h is the laten heat Ut'x.

The latent heat Hux is ,

q"h

• dt Pcmhih (ilks) where xc is the crust thic kness ,

t is time pcm is the density of debris h ih is the debris latent heat of fusion.

Combining these two equations, evaluating the temperature gradient and rearranging yields

'#= Tr,m - T,)- ci .

dt pcmh th .Xc *

. (7 p

/ '\

U This is a non-linear, non-homogeneous, firstorder differential enuation Before effort is expended to solve it, the relative magnitudes of the terntr cantaining the crust thickness will be determined to see if either one dominates.

The initial interface temperature between the wall of the channel and the debris can be approximated by assuming both the debris and the shield wall behave as semi-in0 nite solids. The resulting temperature will be somewhat less than the actual interface temperature because the freezing process will force the crust to stay closer to its initial temperature than it would if it were an semiinfinite solid body only experiencing conduction. The contact temperature between the debris and the channel wall assuming semiinfinite bodies is (Reference 1)

Tr,m Q(kpc). m + T;y(kpc),

.T.= -

](kpc)cm +M.(kpc),. (7;g) where c- is specific heat cm represents debris material properties w represents wall material properties .

G

1. Glen E. Mycts, Analytical Methods in Conduction Heat Transfer, Genium Publishing Corp. , Schenectady, NY,1987, p. 202 DBM-9207 7B-5 :

i Using the debris properties found in the AllWR SSAR Table 19E.217 (Important Parameters for Steam Explosion Analysis) and representative wall properties found in Table 71bl, the interface temperature is estimated to be 1390 K.

'llie debris energy generation density can be found by assuming a decay heat level and a total amount of corium. The density is 4, QthPtm m em (7343) where Qm is the decay heat level mm is the total mass of corium, 235 Mg .

t Evaluating this two hours after accident initiation (decay heat level equals approximately one percent of rated power) yields q = 1.5 x '@ MW .

The two terms inside the brackets in equation 71b4 can now be evaluated. For a channel height of I cm (x .c max = 0.5 cm) and a debris melting ternperature of 1700 K. ,

these values are 0 2 Tr,m -T,) = 1.86 x 10 W / m 5 2 q = 3.8 x 10 W / m .

Therefore, the term containing the temperature difference across the crust is much larger than the one containing the heat generation rate. The temperature profile in the channel system ignoring energy generation in the debris is shown in Figure 7B-2. Equation 71H c:m be simplified to

• C= f f

dt pcmh th cx (T ,m -T ) (7gy7)

Sching this equation with the initial condition that x (t=0) c = 0, reveals 2k g(Tr,m - T.)t x=3 c .

) Pcmhth (71k8)

This equation can be rearranged to determine the time required to freeze debris in a channel of height Ho. The freezing time is O

DBM 9207 7B-6

[V t t "" '

llo Pemh th _ _

8k g(Tr,m - T.) ,

3 (73g)_  ;

b) Interface Temperature, T.

The interface temperature between the debris and the channel wall can be '

determined by equating the heat flux from the crust to that which the crust can absorb. The heat flux from the crus. is 9"ruit " -kf b O 5"./2 (7tyjo) which enluates to

+

4"ruit =

2 x (Tr,m -T,)

c (71y11)  ;

As shown presiously, the temperature difference term dominates the energy generation term in this ec uation for small channel heights. Therefore, the crust  :

heat flux can be simplificcl to O 9"ru,i = h(T,,, .. T,) .

xc

- (71bl2)

Inserting the expression for xc in equation 71k8 and rearranging yields k gpcm h th (Tr,m -T,)

9,',"' '

) 2t (71F13)

The heat flux absorbed by the channel wall can be approximated by that which a semi-infinite solid body can absorb. This flux is (Reierence 1) kw(T,-T i )

. q" =

'/Mw t (71kl4) '

where crw is the thermal diffusivity of the wall material -.

Equating 71bl3 and 7B 14 produces an equation governing the interface temperature.

It is 7 -

s t

1. - Incropera and DeWitt, op. cit. p. 203. -

. DBM 9207 - 7Ik7 u .-

e

< W2 T, - Ti ,

nk,pQgo w JTf,m - T:

y (

2k 2w ,

(71ki5)

Solving this equaticn for T, using the quadratic formula >icids 2

-(co-2T;)i (c - o2T;)2 -4(T _i c pr m)

,T, =

2 (7Ibl6) where co represents the square of the right hand side of equation 71kl5.

Negative solutions of this equation are physically impossible. For a Tr,m of 1700 K and a Ti of 330 K, the inteiface temperature is 1560 K. Similarly, the interface temperature is 2180 K for Tf ,m = 2500 K and Ti = 330 K. The other solutions to equation 71 15 were negative which is physically impossible.

Since this temperature is higher than the value for two semiinfinite solid bodies coming into contact, the dominance of the temperature difference term in equations 71b4 and 7B ll should be reverified. The heat generation and temperature difference ~

terms for a interface temperature of 1560 K and channel half height of 0.5 centimeters are k 5 a

-qTr,m -T,) = 8.4 x 10 W / m' xc

(; *1 = 3.F x 10 5 W / m2 .

2 Even though the dominance is not as great as before, the temperature difference term is still signincantly greater than the heat generation term and the assumptions made previously are still valid.

D.4.2 Required Channel Length to Inswe Freezing The propagation rate of the freeze front was determined in the previous section.

This allowed determination of the time to com aletely freeze the debris in a channel of specined height. A simple tpproximation of the channel length required to provide this residence time is tiv.. product of the initial molten debris velocity and the freezing time. This approximation would predict shield dimensions considerably larger than actually required. A more realistic channellength can be obtained by considerir.g the reduction in channel flow area as debris freezes. In the remainder of this section, the following parameters will be determined: a) debris veloQ, it channel entrance ; b) channel area decrease resulting from debris freezing; c) average channel debns velocity; and finally d) the required channel length to insure plug formation at the channel entrance before corium ingression into the sump.

DBM.9207 71k8

l i

1 a) Debris Velocity at Channel Entrance The possibility exists that mohen debris will not even enter the channel after it has  !

come into contact with the shield wall. Debris which is spreading across the lower dr)well floor will have at least a thin crust formed on its leading edge if the flow energy of the advancmg debns front is not great enough to break this crust and overcome surface tension on the length scale of the channel height, debris will not enter the channel. Unfortunately, the physics of crust formation is not currently understood well enough to support this argument without a great deal of uncertainty.

The entrance velocity will be governed by the height of conium outside of the ,

channel. Assuming that the debris spreads uniformly across the lower dowell floor, the height of debris can be obtained 'ay integrating the volumetric expulsie's rate of corium from the vessel divided by the floor area of the lower drywell. A conservative, overprediction of debris depth can be obtained by multiplying the maximum expulsion rate by time and dividing by area. The upper bound of the ex was shown in section X 2.7.6.2.2 (submitted to the NRC onJune to be 30,192) pulsio 6000 kg/sec.

The velocity in the channel without area reduction due to debris freezing can be conservatively overpredicted by ignoring frictional effects. This velocity is y e(t) = ]2gAz(t) (7gy)7) where ve is the ve!ocity at the entrance of the channel g is the gravitational acceleration constant Az is the height of debris in the lower dr>well .

Expanding debris height yields e

"M v (t) = I Pcm A id (71pl8) where ih m is the maximum ejection rate of corium from a failed vessel 2

A Id is the floor area of the lower dnwell (70 m minimum) .

b) Channel Area Decrease Resulting From Debris Freezing Since the entrance velocity is assumed to remains constant, the mass flow rate of corium in the channel decreases in time due to the area reduction resulting from debris freezing. A conceptual picture of this area reduction process in shown in Figure 7Ik3. Conservation of mass requires that the mass tiow rate of corium entering the channel per unit length is constant throughout the channel. The mass flow rate at the entrance of the channel and at the location downstream where the debris front hasjust arrived is DB.49207 7Ik9 l _ . . , _._ _ _ _ . . _ .

i mi(t)= pcm V e(t)H i (t)= pcmVo(t)Ho (7ipig) where sh i is the time vaging mass flow rate per unit width at the entrance of the thannel Hi is the time vaning entrance flow height of the channel vo is the time vaning velocity at the downstream location in the channel where mcstten debris hasjust arrived Ho is the unobstructed height of the channel .

This equation requires that vo(t)= ""(O H

Hi (t) o (71k20)

The enuance flow height is Hi (t)= Ho-2x cit). (7tygi)

Inserting the relationship for xc found in equation 71b8 into this expression yields 8k((Tf,m - T.)t o

11;(t)= H -)

pcmh th (7Ik22)

The product of this equation and the width of the shield channel describes the reduction of channel inlet flow area with time.

c) Average Channel Debris Velocity The velocity of the leading edge of molten debris in the channel can be obtained by combining equations 71k20 and 71k22, it is

< s 8k((Tr,m -T,)t vo(t)= v e(t) 1- 1 , ,

Hoj pcmhth L s (7Ik23)

The average ve!.; cit) of debris bet veen the entrance of the channel and the leading edge of molten corium is t

O DIN-920 < /Ibl0

l 1

tvo(tidt V(t)= " .

tdt  ?

O (71k24)

Evaluating this integral yields I Y(t)= an d ""t ,

Ho (7425). l where .

T 2gth yn y ,4 _ ,

5jpcm^ld '

u, = 5 . 2k r(Tr.m - T.) .

3. pcmhth This is the average velocity of the molten debris into the sh8cid channel..

d) Required Channel Length to insure Frecting The channellength required to ensure a plug forms at the channel entrance before debris spills into the sumps is

-Iq,ce,e=Y(tgee,e)t freeze

=atc V2 k'

, anb" 1 2 D "' ',

Ho (7&26)'

D.5 Long Term Ability of D6ris to' Remain Solid -

initial debris solidification was considered in section D.4. The rec uirements for-

keeping the debri s .6 the channel frozen for an extended period on time (at least 24

hours) will 've determined in this section. The height of the upper (above the lower .

dr7well floor) ihield wall and depth of the lower (below the lower drywell floor) '

shield. wall.will be specified.

D.S.1 - Upper Shield Wall (Above Lower Drywell Floor) -;

L-.

The roof of the upper shield wall should be free, or at least nearly so, of debris to _

L /' . provide long term cooling to the debris frozen in the channel. No significant amount of debris will' splatter on the roof during ejection from th- vessel because the DBM 3297 i 7&ll

, - - ~ , , . . - , , n , ..,..m, - . .-n., . . _ , . ,,, .+.., :n -, ,.d.. +b

i f N.

(') sump is near the periphen of the lower dowell. To prevent any debris from flowing on top of the shield roof, the shield should be taller than the maximum possible debris pool depth in the lower dr>well This requirement is given by y mc m . tot P(m A id, min -

(71k27) where mcm ,to is the total amount of corium ,235 Mg 2

A ldmin is the minimum floor area of the lower dr>well, 79 m .

Evaluating this expression yields Huw 2 0.33 m. (7B-28)  ;

in the long term (at least minutes after debris solidification), the lower dr>well will be filled with either saturated steam or water. Heat transfer from the shield to the environment is less effective when steam is present. Therefore, only steam will be considered in the remained of this analysis. A shield wall sized to perform its function when steam is present will also perform its function when water fills the lower drywell. O p

t/

The maximum steam temperature in the lower dowell is that of saturated steam at the ultimate containment pressure (180 psig). The steady state heat flux through the upper shield wall is k*

qEw = Huwi (T -To)

- (7B-29) where q", is the steady state heat flux through the upper shield wall Hw u is the height of the upper wall Ti is the temperature of the upper wall in contact with debris To is the temperature of the upper wallin contact with the lower dowell emironment.

Natural convection governs the temperature of the wall in. contact with the lower drywell environment. The heat flux from the top of the wall can be written as.

. 9uw = E(To -T id ) (7330)-

where h is the natural convection heat transfer coeflicient Tid is the temperature of the lower dr>well emironment.

The natural convection heat transfer coefficient depends on' the Rayleigh number.. ,

The Rayleigh number is DBM-9207 7B 12.-

~

rm \

gP(T,,- Tid )L.

()_ ,

vu '

(71k31) where Rat. is the Rayleigh number is the therrnal expansion coellicient of steam

= 1/T assuming id ideal gas behasior v is the kinematic viscosity of steam a is the thermal diffusisity of steam L, is the characteristic length of the shield top .

The characteristic length of a horizontal heated plate is one-half its width (Reference 1). The floor drain sump is approximately one meter wide; therefore, the characteristic length of the shield roofis 0.5 meters. Evaluating the Rayleigh number for saturated steam at ultimate containment pressure (180 psig,190 C) >ields 8

Rat, = 5.5x10 K-1(T o - Tid) *

(7Ik32)

For 107 s Rat, s 1011, the Nusselt number for an upward facing, heated plate ~

undergoing natural convection is (Reference 2)

I Nui, = 0.15Rag,A . (7;p33)

TNe average natural convection heat transfer coefficient is

_. k h = -Nu 13 (7tp34) where k is thermal conductivity of steam in the lower drywell.

Combining equations 7B-30 and 71k32 through 71b34 yields q"w = 8.79(To -T id ) W /mW (71k35) which can be rearranged to

1. Incropera and DeWitt, op. cit., p 43S5.

2. Ibid.

DBM-9207 - 7Ik13

33/4

[u') To= Ty +

e N.,""

g

( 8.79 W / m2 , (71b36)

Inserting this into equation 71b30 3icids

~ '

r ,, 33/4 T-T i id- p K q",= Fluw (8.79 W /m j

, , 7; This equation can be solved iteratively to determine the heat flux which can be transferred through the upper wall of a given wall height. The wall height requirement for transferring a given heat flux is I e ,, 33/4 I liuw 5 ," T i - Tid-- .j K '

quw (8.79 W /m ' .

7&38)

The decay heat level in the AllWR 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> after accident initiation is approximately 0.6E The volumetric heat generation rate of debris at this time cara be determined ;

using equation 71M. It is 0.9 MW/m3. The c i-bris/ wall interface teinperature which will guarantee that the debris remains frozen is 1700 K. The temperature of saturated p> - steam at ultimate containment pressure is 190 C. The height of the the upper shield t" wall which will transfer all of the heat generated in the channel foc these conditions is k*

liuw 5 '

8.52 W /m2K (71L39)

If the upper shield wall ' satisfies this inequality it will be capable of transferring all of .

the heat generated by debris in the channel and, as a result, guarantee long term debris solidification even if the lower drywell has not been flooded. To be acceptable, the height of the shield wall must satisfy the inequalities in equations 71k39 and 71k28.

D.5.2 Lower Sideld_ Wall (Below Lower Drywell Floor)

One side of the lower shield wall is in contact with debris and the other is in direct contact with the basemat. The basemat is constructed of concrete. A conservative estimate of the lower shield wall depth can be made by assuming that concrete acts -

like a perfect insulator. Thus, no heat is allowed to pass from the shield wall to the L basemat. The boundary condition between the debris and the wall is conservatively assumed to be constant heat flux. The initial burst of energy into the shield wall l- caused by debris freezing has ample time to distribute itself throughout the wall.

l ,e With these boundary conditions, the temperature distribution in the lower wall can i ( be determined analytically.

l DBM 9207 7IkI4

l (N, The analytical solution will provide a means for determining the time required for U cach of the interfaces to reach their a!!cwable temperature limits for a given heat flux. The wali/basemct interface temperature should not exceeded the melting point of concrete (1450 K). Continued debris solidificatiun is guaranteed if the v.all/basemat interface temperature does not exceed 1700 K. The wall will be sired so that the limits are not exceeded during the first 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> after initial debris solidification. The upper shield wall will be sized so that it can transfer the full decay heat load after 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> has elapsed, as discussed in the previous section.

The temperature distribution in a slab subjected to constant 1. eat flux at one surface (x = li w) i and insulated at the other (x = 0) is (Reference 1)

~

Ty (x,t)-1,, w = - 4E" +9E*"I*< *- ed'" '/3 S cos" Pwc3,,wHg kw 611 g n n=1 n HN (71M0) where Ty is the temperature distribution in the lower shield wall T,w i is the adjusted initial temperature of the shield wall 97, is the heat flux through the lower shield wall ,

~

cp,w is the specific heat of the shield wall Hy is the depth of the lower shield wall below the lower drywell floor.

The maximum temperature at each interface is achieved as t * . The maximum temperatures at the wall / debris interface, Tw/d. and the wall /basemat interface, Tiwf t,, are T

w /d = Ti ,i, + PwC t

+44 pw HN 3k w.

g and qfy t 3"w H9 T i w j t, = T ,9 + Pw Cgw Hiw 6k" 2)

The heat' flux through the lower wallis bounded by one half of the heat flux generated in the channel when the sizes of the upper and lower wall are comparable.

The actual heat flux will be less because the upper wall is free to convect to the lower drywell environment and will accept more heat flux than the lower wall. The maximum heat generation in the channel corresponds to a decay heat level of one-l x) b 1. H.S. Carslaw andJ.C,Jeager, Conduction of # cat in Solids, 2nd Ed., Oxford University Press,1959, pp. I12 3.

DBM 9207 71bl5

percent. Since decay heat decreases with time, using the maximum value bounds the temperature response of the lower shield wall. Using equation 746, one half of the heat flux generated in the channel is

, Q 1%dhPcm II 91/2than

• ym cm " (9iw)limitirig *

(71M3)

The initial tem aerature of the shield wall should be adjusted to account for the energy it absor as during the debris freezing process. If both shield walls have the same thickness the adjusted temperature is T;*y = Ti + _ p rh th li .,

2pwcp,wll iw

{7;g4)

Equations 7441 through 7444 can be used to determine if a chosen lower shield wall depth will satisfy the requirement of keeping the debris in the channel frozen for at least 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />. After 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> has elapsed, the upper shield wall will be able to remove the entire amount of heat generated in the channel, see section D.5.1.

The process for determining an acceptable wall depth proceads as follows. First a  :

wall depth is chosen which is comparable to the u er shield wall height. Then, the q adjusted initial temperature and heat loads are cal lated using equations 78-44 and j 7B43 respectively. The interface temperatures at 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> are determined by equations 71M1 and 71M2. If Tw/d < 1700 K and Tw/b < 1450 K, the chosen depth is acceptable. If not, a new depth is chosen and the process repeated until an acceptable depth is determined. An example of this procedu.c is given in part d) of the next section.

D.6 Example Calculation The sizing requirements for the floor drain corium shield were set forth in sections D.4 and D.5 based on a chosen channel height. An example sizing exercise is presented in this section. The selected channel height, lio, is orie centimeter.

l Representative shield wall material properties are shown in Table 7&l.

f-a) Melting Point of Shield Material Above Initial Contact Temperature l

l- The initial contact temperature between the debris and the channel wall given in-L equation 7416 is

-(co-2T i )i}(c -2T)2 o i -4(T;2 - c oTr,m T, =

2 (7B-16) l where V

DBM 9207 - 7416

a en kf6m llhUw ,

2k w >

The parameters required to evaluate this equation are Ti the initial temperature of the shield wall, 330 K

'i r,m debris freezing temperature, ranges from 1700 K to 2500 K  :

kr debris thermal conductisity , 30 W / m2 K pcm the density of corium . 9000 kg / m 3 5

hu, debris latent heat of fusion, 2.7 x 10 J/kg aw thermal diffusisity of the shield wall material, a representative 4 2 value of 1.48 x 10 m / sec will be used kw thermal conductisity of shield wall material, a representative value of 4 W / mK will be used . ,

Evaluating equation 7Ibl6 yields interface temperatures of 1560 K and 2180 K for -

~

debris melting temperatures of 1700 K and 2500 K, respectively. The melting temperature of the representative shield material is over 2200_K: therefore, it passes G this test.

V Channel Length b)

The equations needed to determine the channellength requited to ensure that a plug is formed at the entrance of the channel before debris spills into the sump are 7B-9 and 7B 26. These equations combine to give _

3/2

• ab oo 2 I

frecic

• of hecie heetc

%%cre 2

Ho Pcm hlh 8kf(Tr,m -T )

,,4 I2grh s.e, 5]pcm Ald 2kf(Tf,m -T.)-

O V

b=5 o

3 Pcmhih DBM 9207 71kl7 _

i O 'Be maximum length results when Tr.m = 1700 li The contact teruperature was sho previously to be 1560 K for this frecring temperature. The other parameters required to evaluate these equations are rii m the ejection rate of debris from a failed vessel, 6000 kg / sec (conservative maximum )

Ald. min is the minimum floor area of the lower dowell, specified as 0.02 m2 / MWth in the EPRI ALWR Requirements Document ;

it is equal to 79 m2.

Using these parameters, the plug formation time is 7.2 seconds and the required channel length is 1.06 meters. This length was determined using a highly conservative corium discharge rate. The analysis assumed a constant discharge rate equa' to maximum discharge rate pred ted using a highly conservative model. The actual discharge rate will be lower. If ti e length requirement is highly restrictive, the .

discharge rate could be refined with additional effort.

c) Shield Height lluw, Above lower Dawell Floor  :

The height requirements for the upper shield wall are given in equations 71k28 and 7B-39. These equations are H uw 2 0.33 m (71b28) and k*

Huw s 2 8.52 W /m K (71p39)

For a wall conductivity of 4 W/mK, these inequalities require 0.33 m 5 H uw s 0.47 m . (71b46)

A height of 0.4 metbrs is chosen, d) Shield Depth, Hiw, Below Lower Dnwell Floor The lower shield wall should be sized according to equations 7B-41 through 71b44.

An initial height of 0.4 meters is chosen to begin the determination of acceptability.

The adjusted initial temperattire of the ~ + ver shield wall accounting for energy absorption during debris freezing is O

I 1

DBM 9207 7B-18

l

( Ti ,a = Ti + E51""

Pwc p.,H9

= 341 K. (7n44)

The limiting heat flux through the lower wall is Q.1%.dh Pcm H (9Ew)1imiung 2m em 2

= 7520 W / m . (71b47)

The wall / debris, Tw/d, and the wall /basemat, Tw/b, interface temperatures are given by o", t + qfw Hy T

i w/d = T ,i, + Pw c pw H9 3k w p

and  :

qfw t - q[w Hy-O T w/ b = T ,9 + i Pw c pw H iw 6k w 2)

Evaluating these expressions yields: Tw/d = 1190 K and Tw/b = 820 K. Since these temperatures meet the requirements for long term debris solidification (Tw/d < 1700 K and Tw/b < 1450 K), the chosen wall depth is acceptable, e) Summary of Shield Requirements A proposed floor drain sunn corium shield with a specified channel height of one cenumeter and wall materia'. properties shown in Table 71k1 will prevent corium ingression into the sump if it meets the following requirements.

Minimum. melting point of shied material: 2180 K.

Channel Length: 1.06 m.

Height above lower drywell Door: 0.4 m.

Depth below lower drywell Door: 0.4 m.

D.7 Detailed Design Issues During detailed design of the ABWR, the exact shield material and shield O dimensions will be chosen. The requirements for the shield are stated section D.3.

Example calculations of the requirements are shown in section D.6. Interference 1

DBM-9207 7B-19

i

~v Table 7B 1 Material Properties of a Representative Refractory Brick and Concrete Property Representative llrick* Concrete Melting Temperature (K) > 2200 1450 Density (kg/m3) 2700 2300 T' 1al Conductivity (W/mK) 4 1.3 Specific IIcat (J/kgK) 1000 2 Thermal Diffusivity (m2/s) 1.48 x 104 7.5 x 10 7 (3'

O

• Mark's Standard Handbookfor MechanicalEnginem, 8th Ed., Theodorc llaumeister, Editor in Chief, McGraw liill llook Company,1978, pp. 4171 to 4177.

(3 '

\

, L) l D11M-9207 71k20

,~

kJ -

Lower Dr)well Corium Shield

. T Y I 8

F:

?

,.e Channel ,

1. _. ., - 3

.- C e

. O n.

y,w Mahdows Sump Basaltic Concrete Fill r

t Containment Boundary a) side view Channels l

. . . = - +

- .m l

..= .=-

v

? -

M v - 6.e.

I . F u . .. .

Basaltic Concrete Fill ~

l b) front view l

-q Figure 7B 1

(/ Conceptual Design of Lower Dr)well Floor Drain Sump Shield i

DBM 9207 71M1

lrl l'ilt>"t{lrp,Il!I\'\

x,mp . ,.y ~D yr.Q.,

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=

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NITACE.fENT 7C This materialis extracted from aJune 30 transmittal to the staff f

• with small differences in the formatting. This information will be incorporated into the AllWR SSAR at a future date.

X.2.7 fuel Coolant Interactions Fuel toolant interactions were addressed in the early assessment for .

the AllWR response to a severe accident. Subsection 19E.2.3.1 examined the hydrodynamic limitations for steam explosions and concluded that there was no potential fm a large scale steam explosion. The pressurization of the containment ham non explosive steam generation was calculated in the analyses for the accident scenarios. The following sections examine the '

available experimental data base for its relevance to the AllWR _

configuration, and provide a simple, scoping calculation to estimate the ability of the AllWR containment to withstand a large, energetic fuel coolant interact'an.

X.2.7.1 Introduction Challenges of the containment during a severe accident may result

. from fuel coolant interactions. Iloth the impulse and static loads are ',

considered here. Fuct coolant interactions (FCI) may occur either at the

<ime of vessel failure when corium and water fall from the lower plenum of the sessel, or when the lower drywell flooder opens after vessel failure has occurred.

I The critical time constants for a steam explosion are considered in 19E.2.3.1. This analysis cone'.ades that the critical rates for heat transfer and energy dispersal indicate a large scale steam explosions which could damage the containment will not occur. Nonetheless, this study was performed to examine the potential impact of a large steam explosion on

the AllWR.

I Several experiments which have provided insights to steam explosions are examined, and features of the AllWR are compared to previous plants to

' indicate the relative resistance of the AllWR to steam explosions. A scoping i calculation is also performed to estimate the size of steam explosion the AllWR could withstand.-

Four potential failure modes are considered. The transt.:ission of a

~

shock wave through water to the structure may damage the pedestal.-

Simi!arly, a shock wave through the airspace can can cause an impulse load. However, since the gas is compressible, the shock wave transmitted through the gas will be much smaller than that which can be transmitted R through the water. Therefore this mechanisrn is not considered here, y Third, loading is caused by slugs of water propelled into containment structures as a result of explosive steam generation. Finally, the rapid steam

!g i

generation may lead to overpressurization of the drywell.

CElk92-46 7C-1

-4 J

. , . . < , ,, . , . ,,m, .

X.2.7.2 Applicability of Experiments A large number of experiments have been perforr 1. to beu.er understand FCI. Most of the~ experiments nave been , rformed at bench scale with simulant materials. Freon Water and Liquid Nitrogen / Water systems are often used. While these experiments are necessary to understand the underlying p& sics of FCI, they are not directly applicable to the reactor condition. However, there are also several er.periments performed with metal and oxides which provide Sht to the potential for energetic FCI in a sev-re accident.

Other experiments, performed er different reasons, also yield some insights to FCI. Some experiments performed for debris coolability and

{ core concrete interaction studies added water to the debris. With one -

notable exception, these experiments did not result in an energetic 't e 1, Finally, one experiment was performed to examine the impact of a water solid reactor cavity on direct containment heating. In the following settion each of these ex acriments is examined for the insights into FCI and applicability to the ABWR.

X.2.7.2.1 Fuel Coolant Interaction Tests investigations into energetic fuel coolant interactions and steam explosions date back to 1950. Early experiments, including those by Lang (References 1 and 2) and Higgins (Reference 3),identiSed the g 9 requirements for considerable mixing of the molten debris and water.

Higgins and Lemmon (Reference 4) noted that the debris must be superheated and that the violence of the explosion increased with the melt L:mperature. Unfortunately, the triggers used in many of these experiments were very large. Thus, information sbout the propagation and energetics of L

_ the e opeiments is not applicable to reactor conditions, _

A wide vaiety of experiments have been performed to investigate steam explosions. This section discusses results from selected experiments.

s Most of the :xperiments are prototypic of the reactor condPion wherein t dehns falls into a pr~ existing pool or water. Fhe implications of these h

1. G. I mg, " Explosions of Aluminum and Water, " Aluminum Company of America, Ne. insington, Pennsylva, d. COA Report 2 50 33, August 1950.

F

2. G. ag, "Explosions of Molten Aluminum and Water," Metal Progress, Volume 71,

=

p. i , May W57.

U

3. II.M. Higgins,"The Reaction of Molten l'ranium and Zirconium Alloys with Water,"

h Aerojet Report 20: v2, Azusa, California, April 1955.

A.W. Lemmon, " Explosions of Molten Aluminum and Water," 1.icht Metals.1980, (C.

~

4.

Minn, Ed.h P. 817, Proceedings of Technical Sessions Sponsored by TMS Light Metals Committee at the 190th AIME Annual Meeting.

CE492-46 7C-2 m

yq experiments on the potential for large, energetic FCI in the ABWR are also dit:ussed.

One of the important parameters in determining the potential challenge to the contamment from a steam explosion is the duration of the pressure pulse. Buxton and Benedick (Reference 1) performed a large series of experiments using iron alumina thermite. The pressure traces for-these experiments indicate an explosive pressure pulse of about 5 msec.

The final, intermediate scale test performed at Sandia (Re rerence 2) usr A a corium thermite mass to simulate the materials which might be tn -:al of a severe accident. As in the Buxton and Benedick experiment, the duration r the pressure pulse in these experimeats was about 5 msec.

Three shakedown test were performed using iron-alumina thermite with water in the crucible, in all of the tests spontaneous, self-triggered -

explosions occurred. In contrast, all four of the corium tests were externally triggered which resulted in one run with a " weak explosion" and one with a

" mild explosion". Two hypotheses were p oposed to explain these results:

The non-condensible gasses generated by oxidation stabilized the film boiling blanket, making it less susceptible to triggering.

The UO 2and ZrO superheat 2 was only about 300 K. It is possible that the debris froze before the trigger was initiated before the trigger initiated. This would prevent fine fragmentation of the p@t. debris.

t-Both these hypotheses have important implications for applicati on to the severe accidents, Presuming a BWRSAR-type melt pro Jession, the early pour of debris from the vessel would be metallic, in this case stabilization of the gas film around the debris could prevent a large mass of molten material from participating in a steam explosion. On the other hand, the superheat associated with a large oxidic melt is_ typically less than a few hundred degrees. Therefore, it is likely that the surface of the debris droplets would freeze. This would slow the heat transfer to the coolant and -

a steam explosion would not occur.

X.2.7.2.2 Experiments with a stratified system -

In some of the recent expt.riments performed to examine core concrete interaction, water has been added to the debris. As discussed in subsection X.2.7.6.2.2 the probability of a large amount 'of water in the 1

1. LD. Buxton and W.B. Benedick. " Steam Explosion EfTiciency Studies," Sandia National Laboratory, SAND /79-1399, NUREG/CR-0947, November 1979.

LD. Luxton, W.B. Benedick and M.L Corradini, " Steam Explosion EfTiciency l

p- 2.

Studies: Part !! Corium Experiments," NUREG/CR 1746, SAND /80-1324 Sandia l -(

National Laboratory, October 1980.

I l CEB-9246 7C-3

- ~.

O V

lower drywell at the time of vessel failure is very small. After core debris is introduced to the lower drywell it is flooded either by active systems, or the l

passive lower drywell flooding system. Therefore, this is the most probable configuration for a large FCI event in the ABWR.

Far fewer experiments have been performed in a strat.fied geometry than in the configuration of debris poured into water. Wmk by Bang and Corradini (Reference 1) used triggered Freon / Water and Liquid Nitrogen / Water systems. In these studies the interaction zone for the vapor explosion is less tnan 1 cm thick. Assuming this depth is representative of reactor material, this would lead to the conclusion that less 3% of the ABWR core inventory could participate in an FCI evemt.

In some of the recent experiments perfonned to examine core concrete interaction, water has been added to the debris. The MACE and WETCOR tests added water to a pre-existing pool of debris. These tests _

involved fairly large masses of molten simulant to which water was added.

Thus, the initial condition is a stratified pool in which water lies over the core debris. The materials and masses of the experiments are summarized in Table 7C-1. No energetic fuel coolant interactions were observed to-occur in the stratified confip ation. The experiments typically indicated an early heat transfer phase in which the heat fluxes w re on the order of 1.5 to 2 MW/m2. Late:, presumably after the formation of a crust above the molten debris pool, the heat fluxes decreased. These heat fluxes are OV considered in subsection X.2.7.6.2 below in bounding the non-explosive steam generation rates.

X.2.7.2.3 BETA V6.1 Recently, a steam explosion occurred in the BETA facility. Experiment V6.1 was intended to represent the Bibilus reattors. These reactors have an annular pool of water around the pedestal cavity. BETA V6.1 was designed to determine the impact of these water pools on corium concrete interaction. The configuration of V6.1 is shown in Figure 7C-1. The system consisted of a concrete crur ble with an annular water pool which was vented back to the inner crucible via a small path. Molten iron alumina -

thermite was introduced into the cavity which was then allowed to ablate.

The debris eroded the concrete in the approximate shape shown in -

Figure 7C-1. The superheat of the melt was very high since there was no water on the debris. Eventually, the sideward erosion caused the debris to reach the annular water pool at one local point. Instants later an explosion l occurred. The bottom of the crucible was sheared off. There was severe I damage to the facility. All of the instrumentation was destroyed and the l melt injector was thrown several meters up, damaging the ceiling.

l

1. K.!!. Bang and M.L Corradini, " Vapor Explosions in a Stratified Geometry," Nuclear l Science and Technology, Volume 108, Number 1, May 1991, l

l --

CEB-92-16 7C. l -

qq v

The energy required to do the damage has not yet been determined, However, the structure surrounding the test facility was fairly weak, unprotected sheet metal. Although the doors were blown o aen they were not damaged. Therefore, it is believed that the pressure spiie may not have been very large.

The symmetry of the damage to the Scility indicates that the explosion was very syrnmetric. There was very little irregularity in the i shearing of the bottom of the crucible. Thus, it is difficult to believe that l the explosion began on one side of the crucible and propagated sideward. 1 An alternate hypothesis has been proposed (Reference 1). When the debris

~

I penetrated to the annular pool the steam generation ratc increased. Since the annular compartment vents back to the center of the crucible via a small line, the pressure increased and water was forced back into the debris. The debris was still highly superheated at this time. The confinement of the system allowed for intermixing of the debris and water ,

and prevented the pressure from being relieved. Thus, the damage caused I to the system was not a result of a shock wave, but rather due to simple i pressurization of a confined region. ]

The steam explosion observed in the BETA facility is not applicable to the ABWR system. Although suppression pool and vent system of the ABWR-is located in an annulus around the lower drywell, there is adequate vent area to relieve the pressure in the wetwell drywell connecting vents. In fact, the BETA configuration is also much more restrictive than the Bibilus reactor it was intended to represent. This restrictive condition resulted in ingression of water into the melt. Since the ABWR configuration has much more vent area water ingression will not occur.

Additionally, there was no water on top of the debris before penetration into the annulus, Thus, the molten debris in V6.1 was highly superheated. This is contrasted to the situation in the ABWR. The ability to -

use active _ systems, such as the firewater addition system, and the presence of the passive lower drywell flooder virtually ensure that there will be water above the debris in the ABWR. The area of the ABWR lower dowell is also very large which enhances the coolability. The uncertainty analysis of subsection X.3.2 indicate there is a low probability that significant core concrete attack will occur. Therefore, the initial contact mode obsened in V6.1 is unlS-ly.

Ev n it CCI occurs and the pedestal is eroded to the wetwell dryw 1 connecting vents. The presence of water above the debris will cause a en -

to form. The temperature on the lower surface of the crust will be at t 4 melt point of the debris. Withia any molten region,' the' debris temperatura -

will be nearly equal to the melt temperature due to convection in the debris

l. M.L Corradirii, personal comrnunication, June 24,1992.

CElv92-46 7C-5

pool. Thus, the addition of any water to the molten pool will cause the debris to freeze and a steam explosion will not occur.

The conditions which led to the explosion at the BETA facility is not prototypic of the ABWR. Due to operation of the Gooder there is a small likelihood that the debris will ablate the side wall and enter the wetwell drywell connecting vents. This is demonstrated in subsection X.3.2. Even if the debris does penetrate the pedestal to the connecting vents, the vent area in the ABWR is sufficient to relieve the steam generation caused by the imtial contact of water and debris. Thus, water would not be forced into the melt as occurred at BETA. Fintily, the superheat of the melt at the BETA facility was very high, whereas the superheat of any deb:is which contacted water in the ABWR would be low. Thus, debris would be easily solidified, reducing the heat transfer to the water and preventing rapid steam generation. Thus, the explosion in V6.1 does not indicate that containment damage will occur in the ABWR as a result of FCI.

X.2.7.2.4 High pressure melt ejection experiments Sandia performed a series of experiments tc. examine the influence of water pools on the behavior of high pressure melts in a Zion like cavity (Reference 1). Two configurations were examined. In the SPIT-15 test debris was injected into a closed acrylic box. This allowed for visualization of the phenomena. In the SPIT-17 and HIPS experiments a Zion like cavity was constructed. The basic configuration of the SPIT-17 and HIPS experiments O is shown in Figure 7C-2. The SPIT-17 cavity was made of aluminum while V the HIPS experiments used reinforced concrete cavities.

In all of the experiments water was present in the cavity at the time of melt ejection. The inertia of the water prevented venting of the cavity. Thus, the steam generation in the cavity forced the region to pressurire and the structures were destroyed before gas now from the end of the structure could relieve the pressure in the cavity, it is interesting to compare these erperiments to BETA V6.1. In both instances it appears that large pressure spikes were created when the debris and water were tightly confined. This early confinement keeps the water and debris in close contact, and seems to lead to the fragmentation of the hot molten material which is a necess,ry precondition for steam explosions.

The results of this experiment are not applicable to the ABWR configuration. The lower dryw 'l is not Gooded with water and there is ample venting of the region. The extreme damage observed in these Q

\g

l. W.W. Tarbell, et al., " Pressurized Melt Ejection into Water Pools", NUREG/CR-3916.

SAND 84-1531, Sandia National Laboratories, March 1991.

CEB-92-16 7C-6 I

7]

U experiments appear to be consistent with that in BETA V6.1, both in the mode and magnitude of the damage to the facilities.

X.2.7.3 Explosive Steam Generation Thit section presents a bounding analysis of the maximum steam .

generation rate which can occur for a given mass of corium interacting with water.

X.2.7.3.1 Phenomenology Corium interactions with water can result in rapid steam generation.

The rate of steam generation can be limited by the amount of corium or water present. Maximum generation for a given amount of corium occurs when enough water is present to completely quench the corium. Corium mass, surface area, temperature and heat transfer coefficient dictate the maximum rate when ample water is available. ,

Two configurations are possible for quenching in the ABWR. First, ,

corium can exit the vessel when the lower dqwell contains significant amounts of water. Corium exit from the vessel can be either by a slow pour (small vessel breach) or by a sudden drop (catastrophic failure oflower y vessel head). Second, corium can enter a dry lower drywell and form a pool.

Subsequently, the lower drywell is flooded with water and the debris is quenched. This situation, commonly refereed to.as a stratified geometry (U3 steam explosion, is the expected configuration for any large FCI in the 4 ABWR.

Molten core debris is expected to be discharged from the vessel close i to its liquidus temperature,2600K. Therefore, the maximum temperature in  :

either the pour or stratified geometries will be 2600K. The actual temperature will be lower due to heat loss by the debris prior to interaction:

.with water. In the pour case, corium will transfer heat to the air surrounding the vessel as it falls. Any residual water in the lower dowell, as well as concrete beneath and air above the debris pool will abso'rb heat in the stratified geometry.  ;

For rapid steam generation to occur in either situation, the ejected corium must break up into small particles. The analysis presented in.

19E.2.3.1.4 demonstrated that corium breakup in the ABWR will be driven -

j by Taylor instabilities. The smallest particles formed will be approximately '

2.5 mm based on the Taylor critical wavelength. Debris breakup in the stratified geometry will also be governed by Taylor instabilities. -4 Crust formation.will hinder debris breakup. Since corium is expected to exit the versel near its liquidus temperature, any heat loss should-contribute to crust formation. Furthermore, the outer debris surface will

,, freeze rapidly after encountering water. Freezing will hinder further. .

'-} droplet division because.more energy will be required to fracture the outer crust than it does to overcome the liquid surface tension. This, in part,

4 CEB-9246 7C-7 .i 1

i

.(;, Y explains why self-triggering can be obsened with some highly superheated-Al metals, but is much less likely with molten core debris.

X.2.7.3.2 ~ Ilounding Analysis Moody, et al., (Reference 1) determined the maximum steam generation rate during FCI based on a simplified thermal-hydraulic r.cthodology. The steam formation rate from a single corium' droplet assuming heat transfer to saturated water is HA ,(To - T ) .,fi, g.d h'8 (7Gl) where m e.a is the steam formation rate H is the heat transfer coeflicient A 4 is the surface area of a corium droplet To is the droplet surface temperature

(

(

T. is the saturation temperature of water at the ambient pressure h, is the latent heat of vaporization for water t is the time from beginning ofinteraction b 'is the thermal response time.

Heat transfer from the droplet to the surrounding is dominated by convection and radiation. The heat transfer coefficient is o

H = H , + H 'Ta=- H T.)

' + ((T"' -- (7gg)

T') c where He is the convective heat transfer coefficient-H, is the is the radiative heat transfer coefficient

'(O d '

l. F.J. Stoody, R. hturalidhnan, S.S. Dua, " Assessment of ExNessel Steam Pressure Spikes in BWR hiARK 11 Containments" 17th Water Reactor Safety Information hieeting, St.' REG /CP4104, Ort,1989.'

CElk92-16 - 7C-8

(~j o is the Stefan-Boltzmann constant c is the emissivity of the droplet.

Due to the high temperature of corium, convective heat transfer from the surface of the particle will be in film boiling regime. The maximum convective heat transfer coefficient that can be expected is that of-enhanced film boiling, which is 390 W/m K. 2 The emissivity suggested for use in MAAP (Reverence 1) for corium is 0.35 This value will be used for this analysis.

If a mass of corium, Mc, interacts with water and breaks up into droplets of average radius, r, the number of droplets, N, will be given by f '

N4 nr' = M' t 3 s pc (703) .

where P. is the density of corium.

The total steam generation rate of N corium droplets is m , = N rh , ,, = is , ,,,, e" 7g)

A where the maximum generation rate is tg Ib ,,,, = 3M ,H(Te - T )

g Pch ,, r-(7G5)-

This is the maximum steam generation rate that can occur for a given amount of corium broken up into small droplets in a large body of saturated water.

X.2.7A Impulse Loads Rapid steam generation can produce a shock wave which imparts impulse loads to containment structures. Energetic FCis, however unlikely, _

may occur in the lower drywell of the ABWR. Water in the lower dgwell,-

which must be present for rapid steam generation, can transmit shock waves from the site of FCI to the walls of the pedestal. Shock waves which pass into the gas space above the water will be rapidly damped due to gas compressibility and will not represent any threat to containment integrity.

If the impulse load is large enough, the pedestal will fail causing the vessel to tip. Tipping of the vessel would most likely lead to tearing of the containment penetrations. The scoping analysis presented in this section

!O

1. MAAP User's Manual, op. cit.

CEB-92-16 7C-9 L

i

. . - . . .- ~ . - - ~ . . .- . -- ,

(

\

estimates the amount of corium which can participate in a FCI without exceeding the impulse load capability of the pedestal.

X.2.7,4.1 Maximum Impulse Pressure Moody (Reference 1) determined the maximum pressure increase at-the site of FCI based on the steam generation rate given in equation (7C 5).

His analysis applied the Rayleigh bubble equation to'a single steam bubble with an equivalent volume of the many bubbles formed during interaction with N corium droplets of radius, r. Because the volume varies as r3 , this results in overestimation of the rate of bubble expansion The bubble expansion rate dictates the pressure rise. Therefore, this analysis bounds the pressure generated by the maximum steam generation during FCI.

The maximum pressure increase of a single submerged steam bubble above the ambient pressure during its formation at the generation rate given in equation (7C-5) is 2 *I/3 g - 3 , mas AP,,,, = 0.178 pi R,,

(7CAi) where Pi is the density of saturated water at the ambient pressure R: is the Universal gas constant for steam Ro is the starting radius for steam bubble growth.

The starting radius for bubble growth can be estimated by a spherical volume equal to the corium volume plus the total volume of water it

vaporizes which in equation form is 4

-nR* 3

= M f- + M ,c ,(T,; - T )

3 p, hgpi where P, is the density of corium c, is the specific heat of corium.

The maximum pressure predicted by equation (7C-6) is shown in Figure 7C-3 for participating corium masses from 0 to 30,000 Kg. The required corium properties were taken from Table 19E.2-17. The steam and l

O 1. Moody, et al., op. cit.

l CEB-92-46 7C-10 g

n water properties are saturated conditions at two atmospheres. Two  !

E atmospheres is a likely containment pressure at vessel failure for the ABWR.

The peak pressure during impulse loading of the ABWR pedestal resulting from fuel coolant interactions should be bounded by the pressure shown in Figure 7C-3. The pressure predicted by equation (7C-6) is conservative because of the assumptions which went into its creation.

Furthermore, this is the pressure at the site of FCI. The pressure experienced by the pedestal wall will be reduced because the shock wave has to pass through some amount of water before it impinges on the wall.

The pressure will decay as e2 as it moves away from the source (Reference 1).

X.2.7.4.2 Impulse Duration The main difference between energetic fuel coolant interactions (steam explosions) and non-energetic interactions is the time in which the energy stored in the corium is transferred to the coolant. Short transfer times,.on the order of milliseconds indicate explosive reactions. Longer times are indicative of non-energetic interactions. Several fuel coolant interaction experiments invohing cotium simulates were reviewed in subsection X.7.2.1. Pulse widths were observed to be of the order 5 ms or less for FCI.

X.2.7.4.3 Pedestal Capability L

Detailed calculations of the capability of the ABWR pedestal to withstand impulse loading have not been performed. However, a simple.

elastic plastic calculation can provide a capability which can be used for scoping analysin This estimate can be compared to the maximum pressure expected during a FCI for a given amount of participating corium and the impulse duration. The pedestal in Grand Gulf (MARK lil containment) was analyzed in NUREG-1150 (Reference 2) with regards to its ability to withstand pressure rpikes generated by steam explosiors. Since the ABWR pedestal is expected to be at least as strong as that of a MARK 111, the impulse capability of the Grand Gulf pedestal can also be used for comparison.

l l

l

l. M.L Corradini, et al., "ExNessel Steam Explosions in the .NIARK 11 Containmen;",

NUREG-1079 Appendix C. December 1985.

O V

2. " Severe Accident Risks: An Assessment for Eve U.S. Nuclear Power Plants", NUREG--

i150 Appendix C.

CElk92-46 7C-11 l

X.2.7.4.3.1 - Elastic Plastic Calculation A failure limit estimate bases in a simple clastic plastic calculation has been performed by Corradini (Reference 1), The assumptions made in this analysis are:

1. The pedestal wall is thin compared to its diameter;

2. The pressure loading is uniform both spatially and temporally;

3. Failure is based on a strain criteria of (failure stress / yield stress) equal to 10;

4. The pedestal wall is considered to be free standing.

The resistance to deformation, R., of the pedestalis y .

R "' =

R, (7C-8)

Uv is the yield stress of the pedestal wall where A. is the thickness of the pedestal wall

.O R. is the radius of curvature of the wall.

The natural period of the pedestal, T, can be calculated from 2

p,R.

T = 2x)E. (7gg) where P. is the wall density E. is the Young's Modulus of the pedestal.

Since the pedestal is a composite structure, the determination of each of these parameters can be quite complicated. A conservative estimate of the resistace to deformation and the natural period can be obtained by using the following parameters:

Oy = 175 MPa (value for the A441 steel plates which define the boundarit., of the pedestal)

O 1. Corradini, et al., op. cit.

1 l

CEB-92-16 7C-12

fs I') A. = 6 cm (total thickness of the two A441 steel plates V which define the boundaries of the pedestal, ignores steel webs an'd concrete fill)

R. = 6.15 m (average radius of the pedestal)

p. = 2,400 Kg/m3 (density of concrete fill between steel plates)

E. = 200 GPa (typical value of steel).

Using these parameters yields: Rm = 1.7 MPa and T = 4.2 ms.

The maximum response of clastic plastic one-degree systems (undamped) due to rectangular load pulses is shown in Figure 7C-4, The ratio of pulse duration, td, to natural period is the horizontal axis. The- ,

strain criteria, , forms the vertical axis. The relationship between these two axis parameters is given by a series of curves defined by the ratio of resistance to deformation, Rm, to the average pressure of an impulse, Fi.

The amplitude of the square pulse can be conservatively estimated by the maximum pressure rise expected during a FCI, APmax, which is calculated in subsection X.2,7.4.1.

As discussed previously, the impulse duration of a FCI is expected to be pb approximately 5 ms. The ratio of td/T for this duration is 1.2. Using this ratio and a strain criteria of 10 yields a Rm/Fi of approximately 1.0. This, implies that the pedestal can withstand a APmax of 1.7 MPa.

The maximum ratio ot Rm/F iin Figure 7C-4 is 2.0. Using this ratio, the maximum pressure rise the pedestal can withstand is estimate to be 0.85 -

MPa. The uncertainty in pulse duratio~n (assumed to be-5 ms) is irrelevant for the maximum ratio of Rm/Fi ecauseb it is obtained for pulse durations much greater than the natural period of the pedestal.'

This simple clastic-plastic calculation predicts that the pedestal can withstand a maximum pressure during a fuel coolt.nt interaction of 0.85 MPa. The amount of corium which must participate in a FCI to achieve this pressure can be obtained from' the analysis presented in subsection X.2.7.4.1 and summarized in Figure 7C 3. The amount is 22,400 Kg. The ABWR contains 235,000 Kg of corium. Therefore, the ABWR pedestal can withstand a FCI invoking 9.5% of the corium inventory.

7~q_

U CEB-92-46 7C-13

~

d

X.2.7.4.3.2 Comparison to NUREG-1150 Grand Gulf Pedestal Tlw ability of the Grand Gulf pedestal to withstand steam explosions was considered in NUREG-1150 (Reference 1). The smallest impulse load expected to fail the pedestal was reported to be 3.5 psi-sec (0.024 hiPa-sec).

This limit can be used for comparison to the ABWR because the ABWR pedestal is expected to be sturdier than that of a STARK 111. For a pulse duration of 5 milliseconds, this impulse corresponds to a square wave pressure of 4.8 51Pa. This value is significantly higher than the pressure predicted by the clastic plastic scoping analysis. Alternatively, the pressure predicted by the clastic plastic analysis (0.85 SIPa) can be applied for 28 milliseconds before a impulse load of 3.5 psi-sec is exceeded. Both of these comparisons imply that the clastic-plastic analysis bounds the impulse load required to fail the pedestal.

X.2.7.4.4 Capability of the ABWR to Withstand Pressure Impulse The ABWR pedestal has been shown in this scoping analysis to be capable of withstanding a peak pressure of 0.85 h!Pa during a steam explosion. The amount of corium required to produce this pressure impulse during a fuel coolant interaction was shown to' be 22,400 kg. This represents 9.5% of the ABWR corium inventory. This is more than three times the maximum amount of debris which coulc; participate in an FCI event based on the observations discussed in subsection X.2.7.2.2.

.. Therefore, the ABWR pedestal is very resistant to the impulse loading which could occur in a severe accident. This failure mechanism need not L

be considered further in the containment event trees or the uncertainty analysis.

X.2.7.5 Water hiissiles Submerged steam formation resulting from fuel coolant interactions can be rapid enough to propel an overlying liquid mass. Impact loads can be imparted to containment structures if the liquid mass (water missile) is ejected from the water pool with a great enough velocity. Although a prediction of water missile impaction does not imply damage, additional analysis would be needed to assess the structural response. The maximum height to which a water missile can rise will be determined in this section for a given amount of participating corium. The rise height will be compared to the distance between the expected water surface of a pre-Hooded lower drywell and the bottom of the reactor vessel to determine if-damage to the containment could occur. No other structures are I. considered because damage to them will not lead to containment failure.

L

\

! . kJ[')

1. St' REG.ll50, Appendix C, op. cit.

I-l CEB-9246 7C-14 l

[] v X.2.7.5.1 Maximum Rise Height Moody (Reference 1) used the steam generation rate determined in Section X.2.7.3.2 to predict the upward propulsion velocity and elevation characteristic of a water missile. The maximum velocity that a water missile can obtain is the maximum radial expansion rate of the steam bubble formed during FCI. This expansion rate is 5 R , T.,6 , ,,,, * * ,

R~ = 3 -

5,2 4xp,Ri -

(7C-10) L where Ri is the equilibrium steam bubble radius. It is equal to

-1 /3 4 M ,c,(T,, - T=)

R =

3n h '8p8 (7C-11) ~]

where p, is the vapor density.

Balancing the kinetic and potential energies of a water missile yields

.G Ay m. = Ri l, V, 2g (7C-12) 1 y

l 1 It -

l where Aym, is the maximum nse height a missile will rise above the water i surface and g is the acceleration of gravity.

Maximum missile rise heights are presented in Figure 7C-5 for i

participating corium masses of 0 to 30,000 kg.

L. )

l X.2.7.5.2 - Available Rise Height j l

l' The water level in the lower drywell will not he greater than suppression pool water level during a severe accident. The normal water level of the suppression pool is 6.05 meters below the bottom of the reactor i vessel. Consequently, a water missile can rise approximately six meters before encountering any structure who's damage could lead to

-] 4 containment failure.

X.2.7.5.3 Capability of ABWR to Withstand Water Missiles j

s The amount of corium which can participate in a FCI in the-ABWR .

and not generate a pressure impulse which is expected to fail the

/ \

N ,A '

l. Moody e al., op. cit. l l

CEB-92-46 7C 15

l L

r- containment is 22.4 Mg. This amount _of corium will produce a water-missile which will rise 1.75 meters, see Figure 7C-5. This rise height is significantly lower than the available rise height of 6 meters. Therefore, the pedestal will fail from impulse loading before the required amount of corium participates to elevate a water missile even to the bottom of the -

reactor vessel. For this reason, water missiles are not expected to play a role in determining if the ABWR containment fails due to fuel coolant interactions.

X.2.7.6 Containment overpressurization The final element of this study focuses on the pressurization of the containment which may occur during periods of rapid steam generation which may occur when corium is being quenched. In the highly unlikely event of an ABWR core melt which leads to vessel failure, the corium will fall into the lower drywell. There are ten connecting vents which join the lower dnwell, the upper drywell and the wetwell as shown in Figure 7C-6.

The pressure suppression containment prevents large increases in containment pressure by sparging the steam through the connecting vents to the suppression pool which condenses the steam. However, if the pressure rise is extremely rapid,' the vents may not be able to' clear before the containment is damaged.- At even higher steam generation rates, the area from the lower drywell to the upper dgwell could be too small and a pressure difference between the drywell regions could occur, failing the

_q lower drywell. This analysis determines the steam generation rates for different limits on FCI. The maximum rate is then compared to the Q containment pressure capability to assess the potential for containment damage as a result of overpressure during an FCI event.

X.2.7.6.1 Methodology This calculation compares the pressurization due to rapid quenching of corium to the pressure capability of the containment. Two non-explosive steam generation limits are considered. If there is sufficiently large water mass, then the quenching of corium will provide the steam generation limit. If the mass of water limits the steam spike then the steam generation will be less than, or equal to, the water flow mto the lower drywell. The impulse pressure limited mass, calculated in subsection X.2.7.3.1, is also considered.

If the : is no water in the lower drywell at the time of vessel failure then the n.aximum rate of steam generation at some later point in time is the rate at which water is introduced into the lower drywell. If there is still water in the lower plenum at the time of vessel failure, as predicted by MAAP, then this source of water could react with the corium in the lower drywell. Water addition could also occur via the passive flooder, the use of the firewater addition system or by means of ECCS recovery. Each of these possibilities will be examined to determine the maximum rate at which water could be added to the lower drywell.

J CEB-92-46 7C-16

[ For most of the core melt sequences in the ABWR PRA there will not be water in the lower dr>well at the time of sessel failure. Nonetheless, an ev.duation will he performed assuming that coriurn falls into a pre-existing pool of water and is quenched instantaneously. This will provide a limit on the peak containment pressure which could result from quenching of debris as it falls into the lower dr)well. For the AllWR, the vast majority of sequences with vessel failure occur at low pressure. Therefore, gravity is the driving force for the flow of corium from the lower head of the vessel to the lower dr>well. Both hiELCOR and hiAAP predict that the vessel fails at the penetrations for low pressure melts. After the initial hole is formed, the note ablates due to the flow of hot corium. In order to determine the sensitivity of the ABWR containment to rapid steam generation 40% of the total l'02 mass is assumed to be molten at the time of vessel failure. This value is consistent with the upper limit for molten debris used in the uncertainty analyses for direct containment heating.

Two potential limits for pressurization due to steam generation are considered. First, the pressurization of the lower dowell is determined using the limit of the vent area from the lower drywell to the upper drywell, This determines any limits for considering that the upper and lower drywell regions have good communication, and will respond similarly to the pressurization. Second, the response of the pressure suppression system is evaluated. Dowell >ressurization rates are used to determine the vent clearing response wh ch is in turn used to determine the peak i-containment pressure as a function of the pressurization.

X.2.7.6.2 51aximum steam generation rates The first step in determining the peak pressures that_may result from fuel coolant interactions is to determine the maximum steam generation rates. The steam generation can be limited either by the available water or the available corium. Both of these possibilities will be considered separately.

X.2.7.6.2.1 Water added to debris There are four potential sources of water addition to the lower drywell.

First, in a hiAAP type core melt progression, there may be water in the lower plenum at the time of vessel failure. After the corium fall _s into the lower dr>well, the water will follow through the ablated hole in the lower plenum. Second, the lower drywell passive flooder opens when its fusible L material melts. Water from the wetwell is then driven by graity _into the-lower drywell. Third, the firewater system may be used to add water to either the vessel or the upper drywell. In either case, water will eventually flow into L the lower drywell at the firewater injection rate. Finally, if the ECCS is recovered, these systems could be used to inject water into the vessel which again will flow into the lower dr)well.

,O CEB-92-16 7C 17

n)

(

v X.2. 7.6.2.1.1 Water inventory from lower plenum If there is water in the lower plenum at the time of vessel failure, then it will fall into the lower dr>well after the corium. Under these conditions the flow will be driven by gravity through the ablated vessel failure. The expected failure mode for a BWR is penetration failure (Reference 1). A parametric study was performed in subsection X.3.2 to determine the final, ablated area resulting from different numbers of CRD penetrations. The final area varied from 0.06 m2 for 10 penetrations failed to 0.08 m2 for one penetration. In order to bound the flow of water into the lower plenum a value of 0.1 m2is used which results in a maximum mass Dow rate of 1020 kg/s.

X .2. 7.6.2.1.2 Passive Flooder Flow The passive Gooder is composed of ten pipes connecting the lower drywell to the suppression pool with fusible material at the lower drywell end which opens when it reaches a specified value. This is shown schematically in Figure 7C-7.

The flow from the wetwell into the lower drywell is driven by the difference in the water height, h, between the connecting vents and the Gooder. The flow rate is given by o

) th = pA42gh y q,55 where:

th = Water ma.ss flow into the lower drywell (kg / s) p = Density of water (kg / m')

A = Total area of passive flooders (m ')

g = Acceleration of gravity (9.81 m / s')

h = Driving head of water (m)

The maximum flow through the passive flooder would occur when the pressure difference between the wetwell and the drywell was suflicient to open the vacuum breakers and the suppression pool is cold. Assuming a suppression pool temperature of 30 C, p = 996 kg/m3 The total area of the passive flooders is A = 0.081 m2 Assuming that the poolis at the high water level, the height of water above the passive flooder is h= 4.753, which yields a maximum flow rate of m= 780 kg/s.

f3 f '

V 1. J. L Rempe, Light Water Reactor Lower Head Failure Analysis (Draft), NUREG/CR.

5642, December 1990.

CEB-9246 7C-18

X.2.7.6.2.1.3 ECCS and Firewater Flow

.(~]-

V Th- ECCS and firewater system are both capable of adding water to the vessel which would flow into the lower dowell. The firewater system does not rely on AC power, so it is availabic even during a station blackout event.

The ECCS is dependans on AC power, and thus will not be available during station blackout, but could inject water during recovery late in a severe accident. The ECCS system has a now rate far greater than the firewater system. Therefore, no determination of the firewater flow is necessary. The

• maximum ECCS flow will be bounded by the runout Dow of the ECCS pumps. The actual flow will be somewhat smaller due to the flow losses at higher velocities when all of the pumps are operating simultaneously.

There are two HPCF systems each with runout flow of 3800 g (230 kg/s) and three LPFL systems with now of 4200 gpm (265 s). Thekg/pm RCIC system is not considered since the vessel will be depressurized. The total water addition rate to the lower drywell is 1250 kg/sec.

X.2.7.6.2.2 Corium pour from vessel into pre-existing pool of water X.2.7.6.2.2.1 Probability of preflooded lower drywell (The following information was provided to the staffin the hiay 16-17, 1990 meedng summarized by Dino C. Scaletti onJune 8,1990 and in the letter from P.W, hiarriott to Charles L Niiher " Response to NRC/GE i ( Alay 16-17,1990 hiceting Discussion Topics" dated August 9,1990.]

l 1

l The configuration of the ABWR containment, shown in Figure 7C-6, limits the potential for water to be in the lower dnwell at the time of vessel failure, ne vessel skirt is solid and there are no active injection systems in the lower drywell. Therefore the only possible sources of water to the containment are the wetwell dowell connecting vents, the passive flooder and the vessel itself.

The wetwell drywell connecting vents connect the upper and lower drywell regions to the suppression pool. The connecting vent is a vertical channel which has a horizon:al brauch which lead to the lower drywell.

Therefore,in order for Dow from the upper drywell to enter the lower drywell it would have to fall almost 9 m down the. connecting vents, then turn to enter the lower dnwell. This is not viewed to be a credible scenario.

For the water level in the wetwell to rise sufficiently to overflow into the connecting vents, approximately 2.2E6 kg -(4.8E6 lbm) would have to be added to the containment. If the EPGs are followed, this would occur only.

if injection was aroviding injection from an external source in the event that flow from tae suppression pool was not available. This implies that the L only available injection sources are the firewater system. Assuming the larger firewater system flow, the time required to reach the spillover level is-O..

d approximately 11 hours1.273148e-4 days <br />0.00306 hours <br />1.818783e-5 weeks <br />4.1855e-6 months <br /> which should allow ample opportunity for recovery of systems which do not add water to the containment. Furthermore, CEIF92-16 7C-19

,, O procedures tell the operator to continue core cooling with outside V injection if there is no core damage. Therefore, it isjudged that the probability of a sequence with overflow from the wetwell to the lower drywell leading to core damage is vanishingly small.

The passive flooder is designed to open when the temperature in the lower drywell airspace reaches 533 K (500 F). This temperature is slightly less than the temperature of the steam in the vessel under normal operating conditions. However, any potential break flow would cool by i flashing as it reaches the lower dr)well. Therefore, the passive flooder will not open until after vessel failure.

Thus, a LOCA in the bottom head of the vessel as the only source of water which could be present in the lower dr)well at the time of vessel failure. All of the menetrations in the lower head are small, and any loss of coolant accident tarough them is classified as a small break LOCA. A .

conservative estimate of the core damage frequency for events initiated by LOCAs in the bottom head is the frequency of all small break LOCAs which lead to core damage for the ABWR. Examining Table 19D.4-1

[ updated version sent to Glenn Kelly on May 21,1992), the fraction of all core damage events initiated by a small LOCA is about 0.2% Because this value is very small it isjudged that steam explosions which could result from a preflooded cavity will not have a significant impact on risk.

X.2.7.6.2.2.2 Steam generation rate for pr2 flooded lower drywell For the ABWR it is very unlikely that there is water in the lower drywell at the time of vessel failure. Thus, steam generation is usually limited by the availability of water. However, * :re may be sequences for which there is ample water, and the limitation on the steam generation rate is the energy of the quenching corium. Thus, it is prudent to determine the maximum steam generation from this limit if there were a large water supply available. A large mass of water is assumed to be present in the lower drywell for this portion of the analysis.

A wide number of analyses have been performed to determine the mode of vessel failure. While there are still some uncertainties in the -

details of the analysis, the work perfo-ed to date provides overwhelming indication that a BWR vessel fails at the penetrations-(References 1 and 2).

Once there is some flow through a penetration, the molten material will-begin to ablate the hole. Since there is little change in the driving force for the flow of molten material, the maximum flow rate will occur when the hole size is maximized as the mass is exhausted.

1. J. L Rempe, BWR Lower Head Failure Assessment for CSNI Comparison Exercise, EGG-EAST-9609, April 1991.

2. S, A. Ilodge and L J. Ott, Failure Modes nr & BWR Reactor Vessel Bottom llead, ORNL/M-1019, May 1989.

CEB-92-46 7C-20

in some 51ELCOR type analyses, the corium quenches in the lower

.(O) plenum of the vessel. It subsequently heats up and causes vessel failure.

Therefore, there is little corium molten at the time of vessel failure. The flow rate of corium from the vesselis limited by the rate at which the corium melts in the vessel. Using a MAAP type analysis the corium does not-quench in the lower plenum. Thus, there is a large molten mass at the time-of vessel failure. Since this will result in larger flow rates than the MELCOR type model, the 51AAP results will be used to determine the corium Dow rate for this analysis.

MAAP (as well as MEI.COR) uses the Pilch model for the ablation of-the penetration (Reference 1). The velocity of the corium through the vessel failure is approximately constant, therefore the ablation rate of the failure is linear. A series cf MAAP runs were performed which examined the flow rate of molten debris and vessel failure area as a function of the number of failed penetrations. The results of these calculations are shown in Figures 7C-8 and 7C-9. The maximum rate of debris tjection from the vessel is about 6000 kg/sec. Assuming this material quenches as it_is-ejected, the steam generation rate is about 2800 kg/sec.

The experimental heat Dux observed when molten core debris simulants are poured into water is on the order of 1.5 to 2.0 MW/m2based on the floor area. Using the upper bound on the experimental observations, the maximum steam generation rate for the ABWR is

. 80 kg/sec. This is far below the value determined above for the

( instantaneous quenching of debris for a bounding debris pour rate.

i X.2.7.6.2.3 Explosive steam generation rates Based on the examination of the impulse loading calculation of X.2.7.4.3.1, the ABWR can withstand the shock wave which corresponds to-22.4E3 kg of debris. The maximum steam generation rate associated with-this amount of debris is 4100 kg/sec (see subsection 2.7.3.2).

X.2.7.6.2.4 Maximum steam generation The maximum steam generation rates for each of the mechanisms described above are summarized as shown in Table 7C-2. Based on these

( results the limiting scenario is the maximum steam explosion fromlthe'-

scoping study. Therefore, even though this event is far large than the :

L expected steam generation rate, the containment pressurization will be estimated using this value.

l. MAAP User's Manual,' op. cit.

i CEB-92 7C-21

X.2.7.6.3 Containment pressurization The containment peak pressures may be calculated based on the Dow rates determined above. The results given below are for the most restrictive pressurization rate. Three limits are considered. The first condition is the flow rate of steam from the lower dqwell to the upper dnwell. Second, the time pencd before the suppression pool vents open must be considered.

Finally, the quasi-steady condition of flow from the da well to the wetwell through the suppression pool is considered.

X.2.7.6.3.1 Dnwell Connecting Vent flow Consideration of the flow through the drywell wetwell connecting vents is important to ensure that there is adequate vent area to allow the upper and lower dowells to communicate freely. If the flow is restriced a significant pressure difference could exi3t between the upper and lower drywell regions.This could potentially result in Icaci drywell region failure, even though this region has a much higher ultimate strergth than the dowell head (see Appendix 19F). Using the maximum steam generation rate, and an effective area of about 11.25 m2 in the drywell wetwed connecting vents, the pressure difference between the upper and lower .

drywell regions is less than 0.15 MPa (21 psid).

X.2.7.6.3.2 Vent clearing A

U If the drywell pressure is higher than the wetwell pressure at the time of the FCI, then steam flow to the wetwell can begin immediately. However, if the vents are not open, the pressure must accelerate the water in the vents to allow steam flow. During this interval the pressure in the drywell will tise quickly. Since the pressure difTerence between the upper and lower dowell regions is small, the entire drywell volume may be considered when calculating the pressure rise during this perio'd.

Assuming that the initial dqwell and wetwell are at equal pressures maximizes the time for vent clearing. The time to vent clearing is -

calculated based on analysis by Moody (Reference 1). This model requires the pressurization rate for the drywell A constant ramp rate is deterrained by assuming a steam generation rate and using the ideal gas relationship for steam. The pressure rise in the dowell due to steam generation is then

~

calculated using the pressurization rate and the time to vent clearing.

Using the maximum steam flow rate, a pressure rise of _0.26 MPa (38 psid) is-calculated, i-L l

O v 1. Frederick Moody, Introduction to Unsteady Thermolluid Mechanics,1990.

CEB-9246 7C-22 l

L

/3 X.2.7.6.3.3 Horizontal vent flow ,

V After the vents have cleared, steam will begin to flow from the dr>well to the suppression pool. The drywell pressure during this time is equal to I the wetwell pressure plus the flow and water heads. Using conservative l assumptions, and the maximum steam flow rate, the dnwell wetwell pressure difference is found to be 0.16 MPa (23 psid).

X.P.7.6.4 Summary of overpressurization limits Based on the calculations presented above, the maximum pressure rise in the lower drywell due to fuel coolant interactions occursjust before the wetwell drywell connecting vents clear. At this time a pressure spite in the lower drywell of 0Al MPa (59 psi) may occur. FCI events of the magnitude considered here occur when there is a large mass of unquenched debris which comes into sudden contact with water. In the ABWR this only occurs early in the course of a severe accident when the wetW1 pressure is well below the COPS setpoint, typically about 30 psia (0.2 MPa). Even if the wetwell pressure were near the COPS setpoint of 90 psig (0.72 MPa) the lower drywell would be below its estimated ultimate capability of 180 psig.

Therefore, FCI leading to failure of the lower drywell is not a credible event.

Concerning the upper drywell region, a conservative calculation based '

on the maximum steam generation rate given in Table 7C-2 indicates that p

t the maximum pressure in the upper drywellis the wetwell pressure plus 38 psi. Again, considering that FCI events of the magnitude considered here occur when there is a large mass of unquenched debris which comes into sudde,. contact with water, the drywell will be well below even the -

service level pressure (97 psig). Therefore, one would not expect drywell failure as a result of FCI.

The only FCI event one could hypothesize to occur late in the accident is the recovery of ECCSjust before containment' failure. However,in the ABWR design the passive flooder ensures that there is water above the -

debris. The addition of ECCS water will not cause increased heat transfer from the molten debris. Therefore, FCI leading to containment failure late in a severe accident has been ruled cut by design.

The rapid steam generation rates which can occur due to bounding fuel coolant interactions do not lead to failure of the containment structure or opening of the rupture disk in the ABWR. Therefore, no further consideration of steam generation rates is required.

A V

CEB-92-46 7C-23 L

Table 7C-1 Core Concrete Interaction Tests with Water Addition to Debris Experiment Simulant Debris hiass (kg) Water Addition hiACE N10 UO2 Zrog - Zr 130 Flooded after attack started NIACE hil UO2 - ZrO2 - Zr 400 "ooded after attack started, upper crust was not fully molten hiACE hilB UO2 - ZrOg - Zr 400 Flooded after attack started, no crust above debris WETCOR Al 2O3 - Ca0 34 Water added a-1 liter /sec O

O CEIk92-46 7C-24

O Table 7C-2 Maximum steam generation for steam spikes Water limited cases Flow from lower plenum at the time of vessel failure 1020 kg/s Passive flooder 780 kg/s Recovered ECCS 1250 kg/s Debris limited case Debris falling into cavity is quenched instantaneously 2800 kg/s Experimentally limit for debris poured into water 80 kg/s Eyplosive steam generation Scoping result for shock wave capability 4100 kg/s O

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A'ITACilMENT 7D j x This material is extracted from aJune 30 transmittal to the staff with small editorial changes. This information will be (b) incorporated into the AllWR SSAR at a future date.

X.4.2 1.ower 3rywell Flooder This sectian prmides the basis for sizing the lower drywell flooder.

The system is described in detail in Section 9.5.12 of the AI!WR SSAR.

X.4.2.1 System 5 ~4ernents The lowei le,w!! u n n. o s prmides water to quench any coie ,

debris which re, M > Er ;)w veel into the 1.D and to establish a water pool above the deu.u C. enching a required to prevent or mitigate core concrete interaction (C(.1). The overlying water pool ensures that any fission products releaser from the debris bed will be scrubbed.

The flooder system is comprised of ten piping lines. Each 1me will originate in one al the ten drywell to wetwell connecting vents. The vents are arrange d symmetrically around the perimeter of the LD. The flow through each flooder line will be initiated by the triggering a fusible disk at the line exit (LD side). Since four inch diameter fusible disks may be commercially available, the flooder line diareter was chosen as four inches. The flow rates through the flooder lines will be calculated in the

\ next section.

X.4.2.1 Flooder Flow Rate The flow rate through the flooder system will be governed by the flow area line, the hydrostatic driving head and head losses in the lines.

The flow area depends on the diameter of the flooder lines and the number of lines that are participating. Assuming that one flooder fails to operate, the flow area is Ar=Ejn 4 d r (7D-1)

= 0.073 m2 where dr= diameter oflines (0.1016 m,4 in), and nr = number oflines (9).

The location of the flooder line entrance below the suppression pool surface determines the hydrostatic head, see Figure 7D 1. Due to steaming of water in the drywell, the drywell pressure is greater than the wetwell lp Q pressure and the water level in the DW to WW connecting vents is assumed to be depressed to the bottom of the first row of horizontal vents. This CEB-92-46 7D-1

i l

leaves a hydrostatic head, Az, of 0.375 meters to the inlet of the flooder lines.  :

Form and frictional head losses decrease the flow through the flooder  !

lines. Form losses are due to entrance and exit efTects as well a the 90o cibow and valve. A loss coeflicient, k, of 3 conservatively accounts for all the head losses in the flooder system. ,

Applying flernoulli's equation to steady, irrotational flow and asmming that the level of the suppression pool does not change (since the surface area of suppression pool is much greater than that of the lower dr>well) yields 2gAz n" 1+ k (7112)

= 0.099 m3/ scc where v n is the total volumetric flow rate through 9 lines and g is the acceleration of gravity. For a liquid density of 980 kg/m 8, this corresponds to a mass flow rate, in of 97 kg/sec.

XA.2.2 Time to Fill Lower Drywell Water flow that enters the lower drywell will provide cooling to the debris bed as well as establish an overlying liquid layer. Neglecting the subcooling of the flooder water, heat transfer from the debris bed to the water will result in vaporitation. The vaporized flooder flow is f

Q up = h rgp i;q (73y3) where v ap is the volume rate at which flooder water is vaporized  ;

Q is the heat transfer from the debris bed to the flooder water hrg is the latent heat of vaporization for water pi;q is the density of water.

The amount of flooder flow which can contribute to filling the lower i dnwell is f nH = 0 uomi - tevap (7D4)

The time to fill the lower drywell is CElb92-16 7D-2

l t nii = V nii / vnti (7D-5) where Vnti is the volume of the lower dowell below the flooder exit. The flooder exit will be 1.15 meters above the lower dr)well floor. The surface area to the lower dr>well floor is 88.25 m2. Thus, Vntl=101.5 m.

Hooder actuation is expected to occur approximately five hours after reactor scram during most severe accident scenarios. The decay heat level at this time is approximately one percent (1%) of the rated power.

Assuming the entire core relocates to the lower drywell, the debris bed will have a decay heat generation rate QDil, of 39.26 htW. If all of this heat is transferred to the flooder water, the rate and time to fill the lower drywell are 0 (;ij ,911 = 0.080 m3/sec tgigi ,91g = 21 minutes.

The maximum heat flux from the surface of debris that has been experimentally observed (X.2.7.2.2) is 2 MW/m2 The lower drywell has a h' surface area of 88.25 m2 . Thus, the maximum cooling rate of the debris bed, Qmax,is 177 MW For this heat transfer rate, the rate and time to fill the lower dr)well are 3

i ntt . min =0.016 m /sec t nt! . max =1.8 hours9.259259e-5 days <br />0.00222 hours <br />1.322751e-5 weeks <br />3.044e-6 months <br />.

In practi:e, this high heat ik x is not expected to be maintained as the debris is quenched. Nonetheless, the time to fill the lower dr>well to the elevation of the flooder exit will be bounded by these two values,21 minutes and 1.8 hours9.259259e-5 days <br />0.00222 hours <br />1.322751e-5 weeks <br />3.044e-6 months <br />. This difference in timing will not have a significant impact on the fission product release from the containment since the c3uenching of the debris will force any fission products generated during this time into the suppression pool.

X.4.2.3 Consequences of One Flooder Line Opening First Core debris that enters the lower dnwell will be distributed fairly uniformly. The lower dr)well floor was designed so that debris spreading would not be hindered. The temperature of the lower dowell air space and structures should be even more uniform because of convective and radiative O se t<< ester trem a 8<is - teri i. ceiae< <esiees iiiteea te 8sers mere 1, cat than warmer ones resulting in temperature equalization.

1 1

CElk92-46 7D3

( llowever, if highly non uniform debris dispersal occurs, it has been

( postulated that one flooder line could open and its operation could delay or even prevent the other lines from acuvating, in the worst physical case, the initiation of one flooder line causes a crust to form without completely c uenching the debris. The crust limits heat transfer from the surface of the oebris bed. Core concrete interaction (CCl) will occur if surface heat transfer is reduced enough.

CCI results in large quantities of gases being formed under the surface of the crust. The gases will increase in pressure due to continued generation until the crust ruptures or they escape from the edges of the bed. In either case, the gases will pass from the debris bed into the LD airspace. The passage will be unobstructed gasses exiting the debris above the water elevation or through an overlying layer of water. Since only one flooder line is activc the water layer, if it exists, will bc :hin and no significant amount of heat will be transferred from the gas to the liquid.

Concrete melts at approximately 1500ll The released gases from core concrete interaction will be at least at this temperature. Higher temperatures will be obtained by the gases as they interact with debris material in their exit. Thus, gases enter the LD air space a very high temperature. The CCI gases will increase the temperature of the 1.D airspace. More flooder lines will become active as the lower drywell temperature increases. For this reason, the activation of a single flooder

( line is transient at worst and is not expected to adversely afTect the operation of the other lines.

X.4.2.4 Valve Opening Time The fusible , lug valve is designed to open when the lower drywell temperature reac aes 533 K. The fusible materialis made up of an alloy mixture of two or more of the following metals: tin, silver, bismuth, antimony, tellurium, rinc and copper. Alloy contents are chosen so that the pisg melts when its temperature reaches 533K. Different mixtures will be used to reduce common mode failure effects.

The melting points of the individual metals are as fobows:

Metal Wting Point (K)

Antimony (Sb) 903 Bismuth (Bi) 544 Copper (Cu) 1356 Silver (Ag) 1233

( Tellurium (Te) 722 C

CEB-92-46 7D-1

Tin (Sn) 505 Zinc (Zn) 692 The basic configuration of the fusible plug valve is shown in Figure 7D 2. The plastic cap has a melting point wLich is much lower than that of the fusible plug. The plug needs only to melt by an amount occupied in the annular groove which is 2.0 mm deep before it is expelled by hydrostatic pressure. The stainless steel and tcilon disks are expe' led with the plug.

The valve opening time is the time required to melt the fusible metal in the annular groove. To estimate the opening time, a c .'culation has been made for a pure bismuth plug. For the purpose of this calculation, inaterLd and thermal properties of bismuth were used as it has the closest melting point to 533K.

The following material and thermal properties of bismuth were used (Reference 1):

latent heat of fusion: 12.0 cal /gm 1.atent heat of vaporization: 204.3 cal /gm Density: 10.03 yn/cc at 300 C Heat capacity: 0.034 cal /gm/C at 271 C Thermal conductivity: 0.041 cal /sec/cm/C at 300 C The following stainless steel 304 material and thermal properties were used:

Melting temperature range: 1400-1420 C l

Thermal conductivity : 0.0389 cal /sec/cm/C @l00 C l

0.0513 cal /sec/cm/C @500 C Specific heat: 0.00916 cal /gm/C @ 0-100 C Density: 8.03 gm/cc Heat transfer from the surrounding stainless steel pipe to the plug is by conduction. Heat transfer from the lower drywell region to the stainless l steel pipe by convection and radiation. Heat transfer from the bottom of the valve was neglected. Using these assumptions and the material

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I Engineerin'g Materials llandbook, First Edition,1958.

CElk92 46 /11 5

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-D- properdes listed above, the valve opening time was calculated to be about 10 1: O minutes. This is a representative time from when the lower drywell gas space reaches 533K until the flooder line becomes active.  ;

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A/ ISSUE 8 ADFOUAW OF THE AllWR PRA is is irnpossible to determine whether the PRA submitted by the applicant uill be adequate for a safety determination absent information on how it is to be used by the staff. In our February 14 1992 report to the Conunission on the l'se of Design Acceptance Criteria during 10 CFR Part 52 Design Certification Criteria during 10 CFR Part 52 Design Certification Reviews, we commented on the need for guidance on the use of PRA in the review of new plant designs. At this point the applicant has sul>

mitted a PRA, a contractor has performed an extensive resiew, and the staff has prepared a DSER, llowever, the use of the PRA in the design certification process is still undelbed.

Presumably, the results of the PRA will be used in the course of the staffs determina-tion that the design is expected to produce a nuclear power plant that has an appropriate response to severe accidents, in the Severe Accident Policy Statement, the Commis-sion indicated that a PRA would be required for each new design, and that the results of this PRA would be part of the information which would guide the staffin its deter-V mination that a design is adequate to deal with severe accidents. The policy statement published in the &deral Recister of August 8,1985 also states that "Accordingly, l

within 18 months of the publication of this Severe Accident Policy Statement, the stafT will issue guidance on the form, purpose and role that PRAs are to play in severe accident analysis ar d decision making for both existing and future plant designs. . ."

The Statement says further, "The PRA guidance will describe the appropriate combination of deterministic and probabilistic considerations as a basis for severe accident decisions."

The staff has yet to produce the promised guidance. We urge that the staff formulate a set of criteria that it plans to use in making severe accident decisions. This should include the way in which the results of a PRA are to be used in the process (notjust whether the PRA has been donc properly),

l ISSUE 8 RESPONSE l

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,, Response to be provided by staff.

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l

JFiUE 9 CONTAINMENT llYl)ROlWNAMIC 1.OAI)S i

Air <learing loads on containment structures are the result of a complex process resulting from the drywell air being forced into the wetwell by the primary system blowdown. The water in the vent system is pushed down and out until the horizontal vents are cleared. The water-clearing inocess produces ajet of water into the suppression pool which causes a load on the outer part of the wetwell wall. This water clearing is followed by an air steam mixture which creates a large bubble as it exits into the pool. The steam condenses but the air expands, forcing the water above it up into the wetwell air space. The wetwell air space is compressed due to the rnomentum of the water in the layer above the hubble.

The wetwell air space will he sult jected to an energetic two phase eruption as a result of -

the air-clearing process. The vacuum breakers which are in the vicinity will he i exposed us this environment unless protected. The SSAR should d: scribe what the ensironment will be and what protective measures, if any, are needed to ensure sursivid of the vacuum breakers. If a vacuum breaker does not close, the suppression pool is hypassed and the wetwell/ drywell pressures will rise at a rate dictated by the capability of some means other than the suppression process (e.g., containment sprays) to remove heat and rondense steam. The SSAR should contain an analysis of such a L situation.

The early work to address problems arising from analyses of the Mark 1,11, and 111 contaiments is not sumcient to address similar procmses that will occur following a 1.OCA in an AllWR containment. The AllWR is dilTerent for two reasons: (a) the -

volume of the wetwell air space in the AllWR is approximately that of a Mark 11, and (b)-

the impact of the air-clearing loads will he allevia'ed somewhat because the expected h10wdown flows'are much smaller than those eq:cted in a Mark I or Mark 11.

Nevertheless, the combination of a much smaller wetwell and the lower mass flow from the break have not received sumcient attention to be written off hy the staff or GE .

without further analysis or experimerc Al investigation. We are not aware of any testing .;

of the AllWR tyne geometry. We helieve there are sufficient differences in both-geometry and 1.OCA characteristics to requite further evaluation of the air-clearing -

phase of the 1.OCA by more extensive analysis and/or experimental investigation.

I L

9-1

.. - ,___ ,. _-_ _ _ _ . . _ _ _ _. _ -m_._

OPESTION 9 Rl'SPONSI.

AllWR containment, which employes pressure suppression system similar to that employed in cailier liWRs is expected to experience pool swell response phenomena (causing slug of water accelerating upwardly) similar to that observed in earlier llWR (ontainments (Mark I,11, and 111). During this pool swell phase, wetwell region will be subjected to significant pressure loading and drag loads on internal structures.

POOL SWEli PilENOMENON Following a postulated loss-of coolant accident (LOCA) in the dr)well, the rapid increase in di)well pressure will accelenate the water initially standing in the dr>well-to+ctwell connecting vents irito the suppression pool. During this vent clearing process, the static pressure at the exit of the vents will correspond to the containment air space pressure, plus the appropriate hydrostatic pressure, and plus the pool momentum pressure. The pool momentum pressure is a result of the pool being accelerated upward by the incoming water from the vent system.

After the waien is ricared from the vents, air-steam mixture from the dowell flows into the suppression paol creating a large bubble as it exits into the pool. As the air and steam flow from the drywell becomes established in the vent system, the initial vent exii bubble will expand to the suppression pool hydrostatic pressure. The r 1 fraction of the flow will be condensed but continued injection of dr)well air ant. . pansion of the air hubble will force the water above it up into the wetwell air space. The ligament of water above the expanding bubble is accelerated upward by the difference between the bubble pressure and the wetwell airspace pressure acting above the pool surface and against the aucleration due to gravity. The rising water ligament thins until the air bubble breaks through, signaling the end of the pool swell proper and the beginning of froth formation.

During the early phase when the pool swells in a bulk mode (i.e., a slug of water is accelerated upward by the expanding air bubble), structurca close to the initial pool surface will be subjected to impact loads as the rising pool surface impact; the lower surface of the structures. In addition to these initialimpact loads, these same structures will be subjected to drag loads as water flows past them. Equipment and piping in the suppression pool will be subjected to drag loads. Structures at locations above the maximum pool swell height are likely to be subjected to froth impact loads, 92 )

AllWR CONTAINhiENT POOL SWELL PilENOhfES'ON in view of the fac t that AllWR design utillies pressure suppression system similar to that used in earlier llWRs. AllWR design also is expected to experience pool swell phenomenon similar to that experienced in earlier llWRs. Table 91 shows key parameters for AllWR, Alark 11, and hlark 111 containments, llowever, the pool swell phenomena in the AllWR design is expected to be milder than that observed in earlier designs (Stark 11 and 51 ark 111) . This is because,in AllWR i) the expected maximum blowdown flows into the pool are much smaller than those expected in hf ark 11 and

.\tark 111 designs, li) and the pool to vent area ratio is much larger than in hlark 11 and Atark 111 desigi,s. The largest primary system pipe break area in the AllWR design is about i ft2 compared to about 3 ft2 pipe break area in h! ark 11 and h! ark 111 designs, and the pool-to-vent area ratio in AllWR is about 38 compared to about 20 in hiark 11 and about 12 in 51 ark 111.

Considering that the ABWR design einployes closed wetwell air space (similar to hiark 11 design) and horizontal vent system (similar to hiark 111 design), AllWR pool swell phenomena would exhibit pool swell features of both hiark 11 and Atark III designs. The presence of a closed wetwell air space is expected to produce relatively a slow pool swell proc ess. This can be attributed to the fact that the increasing pressure in the air space i

during pool swell will exert a decelerating efket on the rising pool, similar to that observed in .51 ark 11 tests. The presence of horizontal vents is expected to produce pool swell response with slanted pool surface in AllWR design, similar to that observed in

.\tark til tests, see Figure 11-1. The cause of slantness of the pool surface Iwith horizontal vents) can be attributed to the fact that the expanding air bubble at the vent exit does not penetrate the entire pool width.

The above discussion will lead to conclude that AllWR design is expected to experience pool swell response substantially milder than that determined and de rmed for hlark 11 and hiark 111 designs. A milder pool swell response will imply lower maximum pool swell height, lower maximum pool swell velocity, and lower maximum wetwell air space pressure in ABWR design.

ABWR POOL SWELL ANALYSIS As stated earlier, AIMR containment pool swell response was determined by using the l same computer code (i.e., PICS.\1) as that used for the hiark 11 pool swell response -

93 >

l

A calculations. A brief description of tais computer code, and technicaljustincation for j using this code (recognizing that 1,0CA blowdown now into the suppression poolin f

'AllWR is via horirontal vents, compared to that via vertical vents in Mark 11) are  ;

provided in the following paragraphs.

i PICSM Code A detailed description of PICSM is provided in GE report, NEDE-21544-P. Key features and capabilitics of this code are summarised below.

Model Description PlCSM is a one-dimensional pool swell model of the dynamic and thermodynamic conditions in the wetwell following a postulated loss of coolant accident (LOCA).- This- .

model approxhnates the pool swell phenomena by a constant thickness water slug, which is accelerated upward by the difference between the air bubble pressure acting below the pool surface and the wetwell g s space pressure acting above the pool surface. ,

The transient bubble pressure is computed using the known drywell pressure history O' (user specified) and a quasi steady compressible vent flow model. 1 Model Assump.tinm

1. The air is assumed to behave as an ideal gas.

2. Following vent clearing, vent flow feeding the. ah bubble can be assumed as i) air only, or 11) a nilxture of air and steam. Assumption of air only flow, referred to as all air carryover, will maximite the mass flow rate of 1 noncondensables and, therefore, maximizes the resultant pool swell response.

3. - The mass flow rate of noncondensables into the bubble is calculated assuming quasi steady state adiabatic flow through a duct with friction,~_

A

4. The air in the dr)well is isentropically compressed and heat transfer to the walls -is conservatively _ neglected;

5. The model allows two different assumptions for modeling the temperature of the bubble: i) a variable temperature equal to the current drywell temperature cy4 3

. . , , 4.,;,._.___.,_m --

m.

throughuat the transient, and li) a constant bubble temperature equal to user specified value. A variable temperature input maximires bubble pressure and, therefore, maximires the resultant pool swell response.

ft followir.g sent clearing, the water above the exit of the vent acceleratts as a slug of constant thickness.

7. Frictionallosses between the water and the confining walls are considered to be negligible.

8. The wetwell air space is polytropically compressed by the oving water slug, lleat transfer to the walls and pool surface is neglected.

9. The air velocity in the dowell is sulliciently small so that static and stognation conditiens are equivalent.

tinplicability of PICS 51 to AllWR As noted under model description, PICShi is based on simpic first principle analytical models and provides a conservative simulation of pool swell phenomena in the wetwell region. It should be recogniicd that severity of pool swell response will be strongly dependent upon and determined by the drywell pressure history (which is an external input to the code) during the pool swell phase. Given that AllWR design employes a closed wetwell air space similar to that in 51 ark 11 design and AllWR input drywell pressure history can be determined similar to that for hiark 11, PICShi appears to be a technically adequate model for sinmlating and calculating AllWR pool swell response, as well.

Ilowever, as it was noted earlier, PICShi approximates the pool swell phenomena by a constant thickness water slug (which is representative of hiark 11 pool swell feature) whereas, pool swell phenomena in AllWR design would be expected to experience a variable thickness water slug (because of horizontal vents, representative of hiark pool swell feature). This single difference of constant vs variable thickness water slug, additional studies were performed evaluating PICShi capability for predicting horizontal vent test data from 51 ark 111 tests, in these studies, variable thickness water slug was simulated by modeling a fraction of total pool surface area as constant thickness water slug as input to PICShi. Results from these studies showed that modeling 80% of the 9-5

p total pool surface area as constant thickness water slug produced results which compared well conservatively with the measured test data. PICSM vs Mark 111 test data comparison results are shown in Figures 9-2 and 9-3. Large Test series involved a relatively small and closed wetwell air space (representative of ABWR design), and the Scaled Test (1/3 scale) series involved a relatively very large wetwell air space volume (representative of Mark 111 design).

In view of the above discussion, and recognizing that input data / assumptions in PICSM will be defined in a conservative manner, PICSM model wasjudged to be technically adequate for AllWR application. Considering inherent conservatism in the PICSM modeling assumptions, PICSM model should proc a conservative AllWR pool swell response.

AllWR Pool Swell Response Calculations As noted in the above paragraphs, AllWR analysis was performed with PlCSM using all input data / assumptions conservatively in a manner consistent with the staffs n guidelines denned in NUREG-0808. Initiating LOCA event was determined to be a b postulated instantaneous feedwater line break which resub -d in most severe pressurization condition in the drywell, it should be noted snat drywell pressurization condition will determines severity of the pool swell response. Pool swell response severity depends directly upon severity of the diywcll pressurization. Specifically:

1. A polytropic index of 12 for wetwell air space compression was used,in order to obtain conservative values for pool swell height and swell velocity. A-polytropic index of 1 A for wetwell air space compression 'was used, in order to obtain conservative value for wetwell airspace pressure.

2. Analysis' assumed a variable bubble temperature equal to the current drywell temp (rauue throughout the transient. This assumption maximizes bubble pressure and, therefore, maximizes the resultant pot , w.cll,

3. In consideration of slantness of the pool surface (i.e., variable thickness water slug),80% of the pool surface area was modeled as a constant thickness water q slug. As ermined based on PICSM vs Mark 111 test data comparison results-L/ discussed in above, this assumption' of using 80% of pool surface area is expected to produce conservative results.

Mi

Key results from 'llWR PICSM analyses are summarized below. I

1. Maximum wetwell airspace pressure during the pool swell process did not  !

exceed the drywell pressure, implying no negative pressure differential I loading on the diaph agm floor, see Figure 9-1.

2. Maximum pool swell velocity was determined to be about 16 feet per second, ,

see Figure 9-5. I

3. Maximum pool swell height was determined to be about 21 feet, see Figure 94 l

Engl Swell Hydrodynamic 1.ogli i Structores located at lower elevations, i.e., from 0 to 21 ft above the initial pool surface will be subjected to and designed for impact (by an intact water ligament) and drag loads. These loads will be computed in accordance with the methodology defined in NUREG-0978. Structures located above 21 ft from the initial pool surface will be subjected to froth impact loads. Severity of froth impact loads will depend upon the structure location in the wetwell air space relative to the initial pool surface.

As seen from Figure 9-7, equipment tunnel, catwalk and the WY,'/DW vacuum breaker are likely to be subjected to significant loads due to pool swell. Vacuum breakers which are located well above the expected maximum pool swell height are not expected to be subjected to direct pool swell impact loading, but they are likcly to be subjected to significant froth impact loads. liowever, these vacuum breakers will be adequately protected from possihic froth impact loads, thraugh appropriate design features which will provide appropriate adequate shielding for the vacuum breakers. - These features would include prmiding solid catwalk area (below the breakers) of sufficient measure assuring complete shielding of vacuum bieners from possible froth impact loads.

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Table 9-1 CONTAINMENT CONFIGURATION PARAMETERS: ABWR AND EARLIER BWRs ABWR MARK 11 MARKlli TYP TYP DRW/ ELL VOLUME cu ft 259000 229000 275000 WETWEl.L GAS VOLUME cu ft 210000 165000 1140000 SUPPRESSION POOL VOLUME cu ft 126000 132000 120000 SUPP POOL SURFACE AREA sq ft 4833 5000 5900 30 98 120 NUMBER OF (HORIZ) VENTS TOTAL VENT FLOW AREA sq ft 125 295 495 VENT DIAMETER ft 2.3 2 2.3 VENT CTR LINE SUBMERGENCE tt 12.33 TOP VENT ft 11.48 N/A 7 MIDDLE VENT ft 15.98 N/A 11.5 BOTTOM VENT ft 20.48 N/A 16 POOL TO VENT AREA RATIO 38.5 20 12 LARGEST PIPE BREAK AREA sq ft 1 3.11 3.54 BREAK TO VENT AREA RATIO 0.008 0.01 0.007 BREAK AREA TO DW VOL RATIO 3.86E-06 1.70E-05 1.29E-05 WW GAS TO DW VOLUME RATIO 0.81 0.6 4.1

• CONT, DESIGN PRESSURE DW psig 45 45 30 WW 45 45 15 CONT. DESIGN TEMP. DW deg F 340 340 330 WW 219 275 185 o-8

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i Figure 9-7 ABWR WW/DW VACU U M BREAKER POSITION LOCATION

ISSUE 10- ADEQUACY OF SSAR TREAThfENT OF THE REATOR WATER CLI.ANUP SYSTEM We performed a review of the Reactor Water Cleanup (RWCU) system using our own staff. This system was chosen because it is a non-safety system located outside of primary containment, but inside the building which houses engineered safety features. It uses pipes up to 8-in. nominal diameter whose rupture would result in a LOCA and a source of serious environmental disruption in the building. This system is not seismically qualified or built to quality assurance standards.

Our review identified a number of deficiencies in the SSAR, some of which are listed below:

(a) There is little useful information presented in the SSAR that describes how the Japanese codes and standards used for the RWCU System design can be converten to domestic design standards. The Quality Group classifications for -

certain portions of the RWCU System are inconsistent with the Japanese code--

related classifications s'nown on the Piping and Instrumentation Diagrams. The Safety Class / Quality Group transition between the piping inside primary contain-ment and that outside primary containment is not in accordance with ANSI /ANS safety class standards for BWR fluid systems.

(b) The questionable ability of system' isolation valves to close under large-break.

LOCA conditions has been the subject of extensive NRC testing and a Generic Letter (GL 89-10), However, the SSAR specifies no special performance

! requirements for these valves.

(c) The safety-grade leak detecdon and isolation system which actuates the system isolation valves was not described in detail suflicient to support an assessment of

. its adequacy. ,

(d)' The ABWR PRA did not evaluate as initiating events RWCU System line breaks (or other LOCAs) outside the primary containment. The exclusion of these l

b.caks was based erroneously on an analysis of the effects of suppression pool

,LJ 7-

gc

], bypass events on overall risk. However, toe analysis failed to take into account that the bypass path (e.g., RWCU_ System pipe break) could be the initiator for the core-damage event.

(e) The PRA analysts took credit for the RWCU System as a heat removal system in all sequences wk reactor pressure is assumed to remain high. The analysts assumed that the capacity of the non-regenerative heat exchanger (NRHZ) is adequate to remove the decay heat. The capacity appears to be adequate; however, our calculations indicate that the outlet temperatures on the RWCU System side and cooling water side of the NRHX would exceed the desiga limits for the piping. Furthermore, a temperature sensor between the NRHX and the RWCU System pumps in the present desigr. would automatically isolate . NRHX on high temperature, making it availabic.

The items mentioned above are among a number ofissues that were identified. It is-important for the staff to ensure that the shortcomings of the RWCU System and PRA related portions of the SSAR are not indicative of problems in the remainder of that report.

_%(3) .

ISSI'E 10 RESPONSE PART (a)

Note 11 of SSAR Figure 1.7-1 (P&lD Symbols) provides a correlation between the

_ Japanese code-related classifications shown en the SSAR P&lDs and domestic standards. Unfortunatch, there are two ambiguities in this correlation: one safety designation (SC-3 or NNS) and one seismic category classification (I or NSC). The Note correlations have been modified to eliminate these ambiguities and the corresponding symbols on the P& IDS will be changed as part of our SSAR verification .

aethity. These ambiguities may have resulted in the observation that the Quality Group classifications for certain portions of the RWCU syetem are inconsistent with theJapanese code-related classifications shown on the P&lDs. It should be noted, however, that outboard of the second isohtion valve the pipe is Safety Class 3, Quality

[i

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Group C and Non-seismic Category 1.

10-2

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'1 7

[* It is not clear that the Safety / Quality Group transition between the piping inside primary containment and that outside primary containment is not in accordance with ANSI /ANS-52.1. The transition is between SC-1 and SC-3. This transition of the RCPB penetrating containment (Case 3) is not explicitly included in the examples of the standard. GE believes that the RWCU System does in fact meet the criteria of Section 3.3.2.1 of ANSI /ANS-52.1. . It should also be noted that the staff has not endorsed ANSI /ANS 52.1.

4 PART (b)

The following special performance requirement for all isolation valves will be added to the end of SSAR Subsection 6.2.4,2 (system design portion of containment isolation system subsection):

ti isolation valve closure will be assured by using the latest state of the art -

technology in valve design. Valve actuators will be sized based on demonstrated -

valve design and established disk friction factors. Adequate thrust capability will be developed with sufTicient margin in the actuator and the valve is appropriate to

( ,

N demonstrate acceptability of the valve design for its application, t

PART (c)

The leak detection and isolation system which actuates the RWCU system isolation.

valves is treated in the following sections of the SSAR. .

SSf.R Section Tooic l 5.2.5.1.2 2nd para. Isolation of RWCU system on high MSL temperature 555.1.2 5th para. Intersystem radiation leakage monitoring from~ RWCU' system heat exchangers 52.5.1.2 (f) High differential mass flow rate in RWCU system pipingi .

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Af s SSAR Scotion - Topic 5.2.5.1.2 (g) High radiation in RWCU system heat exchanger discharge lines (intersystem leakage) 5.2.5.2.2 (3) RWCU system differential for RWCU system isolation 5.2.5.2.2 (4) MSL, temperature for RWCU system isolation 5.2.5.2.2 (5) Area temperature for RWCU system isolation 5.2.5.2.3 2nd para. Time delay for RWCU system Table 5.24 RWCU system leak detection and isolation functions for

RWCU system monitored variables x/

O Table 5.2-7 Monitored trip alarms for RWCU system leakage sources Figure 5.2-8 Sheet 7 1.cak detection and isolation instrument engineering-diagram for RWCU system 7.3.1.1.2 ($f (g) RWCU system differential flow monitoring- *-

7.3.1.1.2 (3)(h) RWCU intersystem leakage monitoring-RWCU system temperature monitors in equipment areas--

7.3.1.1.2 (3)(p) 7.3.1.1.2 (6)(b) RWCU system bypasses and interlocks -

7.3.1.1.2 (7)(b) - RWCU system I&C redundancy. and. diversity- _

Figure 7.3-5 Sheet 58 RWCU system mass differential flow trip logic

,r).

Y 95 10 4

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SSAR Section Topic Figure 7.S5 Sheet 59 RWCU system regenerative heat exchanger temperature trip logic Figure 7.3-5 Sheet 60 RWCU system nonregenerative heat exchanger temperature trip logic Figure 7.S5 Sheet 61 RWCU valve room temperature trip logic

, Figure 7 S5 Sheet 62 RWCU system inboard valve isolation logic i L

Figure 7.S5 Sheet GS RWCU system outboard valve isolation-logic-Figure 7.S5 Sheet CA RWCU system injection and purge line valves isolation

,G logic l \/

Figure 7.S5 Sheet 65 RWCU system head spray valve isolation logic GE believes the information delineated above adequately describes the leak detection

- and isolation system.

PART (d)

- The approach taken in the bypass study (SSAR Section 19E.2.3.3) is to consider _the

~

j. presence of a bypass path as an event independent of the events which.cause the core-l~ damage. This approach is acceptable because for large breaks, the system which has failed is not,in general, relied upon to prevent _ core damage. Therefore, no:

consequence of these failures affect the systems preventing core damcge. Similarly, Lnone of the systems associated with the smaller bypass lines are associated with preventing core damage. Therefore, they too are not associated with the cause of the core melt.

l,-

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10-5

( If a break in the RWCU suction line were the postulated 1.OCA, the containment isolation valves would be expected to close, terminating the event. Regarding the failure of the RWCU suction, in combination with failure of the isolation valves to close, the concern is that there may be flooding that could have a high consequence if it leads to an eventual loss of suppression pool and CST inventory or flooding of other ]

ECCS rooms. In the event of this extremely unlikely scenario, existing EPGs would j require depressurization of the RPV which would both slow the break flow and allow .j for possible manual closure of the isolation valves. The system arrangement of the ABWR routes the RWCU lines above the core to avoid a potential siphon of the core inventoty. In the event of an unisolated RWCU line break, lowering the RPV level to below the shutdown cooling suction and depressurizing the RPV would be sufficient to terminate the break flow without causing core damage. This action should be possible prior to any impact on other ECCS equipment. SSAR Section 19E.2.3.3 was I modified to reflect these actions and the required operator actions are being included i in SSAR Section 19D.7 so that procedures or Accident hianagement Guidance addresses the needed action.

J PART (c)

The use of the RWCU as a decay heat removal system has been proposed as an ultimate backup to the 3 RHR systems. This is mainly accomplished by bypassing the RWCU regenerative heat exchanger (RHX) which causes the nonregenerative heat exchanger (NRHX) to remove additional heat. To address the concerns about this proposal the following acuons have been taken:

1. RWCU System Alodification The high temperature isolation feature of the RWCU system has been modified to only isolate and bypass the filter demineralizers rather than the entire RWCU-system. This will be accomplished by closing valves F201(A), F202(A), F201(B) and F202(ll) and opening valve F011. This change still protects the filter demineralizers from high temperatures but leaves the RWCU system available for decay heat removal if needed. The RWCU P&lD reflecting this change will he included in the SSAR.  !

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2. Emergency Procedure Guidelines Modifications The procedure to use the RWCU as an emergency decay heat removal system has been modified to include steps to increase the RBCCW flow through the NRHX. After the RBCCW flow has been increased, the following steps will cause the RWCU to remove a significant amount of decay heat when the reactor system is at high pressure:

a) Bypass the filter demineralizers in the RWCU system, b) Bypass the shell side of the RHX in the RWCU system.

3. Impact of Temperature Changes When the RWCU RHX is bypassed there is a concern about the large increase in the temperature difference across the NRHX. This temperature difference increases to about 440'F. However, the RHX and the NRHX are essentially the same design and the normal temperature difTerence across the RHX is 415'F.

O Thus the increase in the temperature difTerence across the NRHX is acceptable Q during emergency use.

1 There was also a concern expressed about the temperature of the R%CU and the RBCCW piping exceeding design limits. In the heat removal mode the exit temperature from the NRHX in the RWCU line would be about 155'F which is only 5'F above design. Also, the exit temperature from the NRHX in the RBCCW line would be about 216*F which is only about 30*F above design. Both these temperature increases are acceptable during emergency use.

The RWCU return flow to the feedwater will decrease to about 155'F. However, this case is covered by the RHX out of service event which is already analyzed.

The SSAR will be modified to require the COL applicant to demonstrate that the above changes in operating temperatures are acceptable during emergency use.

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ISSUE 11 - PLANT DESIGN LINE AND AGING MANAGEMENT-We recommend that the SSAR clearly define the scope of the 60-year design life for -

the ABWR and describe a program plan for achieving it. This program should include those aging management measures which are necessary to maintain the plant within its design basis throughout its design life. This program should specify ,

the original design and application criteria and, where required, the projected refurbishment or replacement requirements with appropriate rationale, To the extent applicable, the lessons learned from the NRC's Nuclear Plant Aging Research Program, as well as other aging research projects, should be incorporated into this program.

ISSUE I1 RESPONSE A new subsection 1.2.1.3, Plant Design Life Criteria, will be added to the SSAR in the next amendment. ' This subsection, based in part on EPRI URD Volume II, Chapter 1,

. , Paragraph 11.3,is provided below

i

s l .2.L3 Plant Design Life Criteria 7?

The AllWR design accommodates component refurbishment and replacement for early in life failures and obsolescence. Breakthrough type technical developments

- are not needed to achieve a 60-year life.

The COL applicant shall study the design to evaluate the longevity of structures, systems and components and develop a design life understanding, management and classification system in conjunction with the procurement and maintenance-programs. ~ The classification will categorize items (i.e., structures, systems,-

subsystems and components) according to design life capability for developing the strategy to be employed to support the design life requirement.

L- For those components which are expected to be replaced for obsolescence or early failure, the COL applicant shall provide a plan for replacement, refurbishment, and

'p repair activities, as appropriate, to assure the design life of the overall plant. Such V activities shall be scheduled to demonstrate that the plant availability requirements are -

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satisfied, The plan shall include a comprehensive program fa obtaining behavior data and record keeping for evaluating life capability oflong life components based upon their operating history and measurement of their life limiting characteristics, A program is required for suncillance specimens, instrumentation, material condition monitoring, environmental monitoring, etc.

The reactor vessel replacement is not anticipated during plant lifetime and need not be considered within the availability requirements. The reactor vessel is a very conservative design, is subject to a rigid surveillance program during plant life, and is designed for replaceability of internal parts, therefore, replaceability of the vessel itself is not considered necessary. Ilowever, the AllWR design has considered reactor vessel replacement in the development of plant layout and structural design such that, if required, reactor vessel replacement would be possible and would be accomplished williout inadvertent obstacles.

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ISSUE 12 - STATION GROUNDING AND SURGE PROTECTION Chapter 8 of the ABWR SSAR defines the scope of and specifies the requirements for the electrical power systems. The scope is limited to the onsite electrical power _

systems and to the interface requirements with the offsite electrical power systems.

Notably absent are lightning protection, station grounding systems, and surge protection measures which are necessary to protect plant personnel and equipment '

during normal and abnormal conditions. These measures are required to eliminate or reduce electrical shock hazards to personnel, and to protect systems and equipment against damage or misoperation as the result of lightning strikes, switching operations, electrical arcs, short circuits, status electricity, etc. These protective ,

measures and their interface requirements should be included in the SSAR, The AIMR makes extensive use of sensitive solid-state electronic components for essential protection control, and data transmission functions. These components should be protected from extraneous electrical impulses that will damage them or

. cause improper performance To the extent practical, these components should be .

isolated from potential adverse signals that may be transmitted over control or data .

links from remote locations, meteorological stations, switchyards, etc.

- We note that the EPRI URD (Volume II, Chapter 11, item 9, ' Electrical Protective Systems") addresses requirements for these systems. We recommend that these grounding, surge protection, and isolation features be included in the SSAR; ISSUE 12 RESPONSE l:

i J Appendix 8A (copy attached) has been added to the SSAR. It covers station grotmding L .and surge protection, cathodic protection, and electric heat tracing.

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' - SECTION 8A CONTENTS Section Title Page 8A. - MISCELLANEOUS ELECTRICAL SYSTEMS 8A.! STATION GROUNDING AND SURGE ,

PROTECTION 8A.1 1 8A.I.1 - Description 8A.11 8A.I.2 . . Analysis 8A.1-1 8A.13 COL License Information 8A I.1 SA.I.4 References 8A.I.1 8A.2 CATHODIC PROTECTION 8A.2 1 SA.2.1 Description 8A.2 1.

8A.2.2 Analysis 8A.2-1 8A.23 COL License Information 8A.2 1

-8A.2.4 References . 8A.21 8A.3 ELECTRIC HEAT TRACING 8A31 8A3.1 Description . A31

(- SA3.2 - Analysis - 8AJ-1 8A.33 COL Information RAJ-l .

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. 8A.1.1 Description - 8A.1 1 8A.1.2 Analysis 8A.1 1  !

8A.1.3 ' COL License Information 8A.11 SA.1.4 References 8A.1 1 '

ILLUSTRATIONS -

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V 8A.1 Siation Grounding and Surge The lightning protection system covers all major Protection plant structures and is designed to prevent direct - ,

lightning strikes to the buildings, electric power -

8 L1.1 Description equinment and instruments. It consists of air tert .als, bare downcomers and buried grounding The electrical grounding systern is comprised of: electrodes which are separate from the normal -

grounding system. Lightning arresters are prosided (1) an instrument grounding network. for each phase of all tic lines connecting the plant electrical systems to the switchyard and offsite line.

(2) an equipment grounding network for grounding Plant instrumentation _ located outdoors or-electrical equipment (e.g. switchgear, motors, connected to cabling running outdoors is provided distribution panels, cable;. etc.) and selected with surge suppression devices to protect the mechanical components (e.g. fuel tanks, equipment from lightning induced surges.

chemical tanks, etc.),

8A.I.2 Analysis (3) a plant grounding grid, and There are no SRP or regulatory requirements for (4) a lightning protection network for protection of the grounding and lightning protection system. It is structures. transformers and equipment located designed and required to be installed to the outside buildings, applicable sections of the following codes and standards.

The plant instrumentation is pounded through a separate insulated radial grounding system (1) IEEE Std 80, Guide for Safety in AC Substation-comprised of buses and insulMed cables. There is a Grounding h d single point connection to the station grounding grid, (2)(EEE Std 81, Guide for Measuring Earth The equipment grounding network is such that Resistivity, Ground Impedance, and Earth all taajor equipment, structures and tanks are Surface Potentials of a Ground System -

grounded with two diagonally opposite ground

- connections. The ground bus of all switchgear ; (3)IEEE Std 665, Guide for Generation Station assemblies, motor control centers and control Grounding cabinets are connected to the station ground grid .

through at least two parallel paths. One bare copper (4) NFPA 78, National Fire Protection Association's cable is installed with each underground electrical Lightning Protection Code . ,

duct run, and all metallic hardware in each manhole is connected to the cable. This code is utilized as recommended practices only It does not apply to electrical generating A plant grounding grid consisting of bare copper plants.

cables is provided to limit step and touch potentials

- to safe values under all fault conditions. The buried 8A.1.3 . COL License Information

grid is' located at the switchyard and connected to _ _

systems within the buildings by a 500 MCM bare It is the responsibility of the COL applicant to copper loop which encircles all buildings (See ITgure =

perform ground resistanee measurements to-8A.11) determine that the required value of 0.05 ohms or p less has been met and to make additiocs to the .

. The target value of ground resistance is 0.05 system if necessary to meet the target resistance.

L  ; ohms or less for the reactor, turbine, control, service and radwaste buildings. If the target grounding 8A.1.4 References ll resistan:e is not achieved by the ground. grid, l- auxiliary ground grids. < hallow buried ground rods or (1)IEEE Std 80, Guide for Safety in_AC Substation l --

deep buried ground rods will be used in combination Grounding Amendment 21 SA.I. t j

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TAllLE OF CONTENTS Section Title Page 8A.2 Cathodic Protection 8A.2 1 8A.21 Descripton 8A.2 1 8A.2.2 Analysis 8A.2 1 8A.2.3 COL License Information 8A.2 1 8A.2.4 References 8A.2 1 a

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ABWR m uim40 Standard Plant nrv n 8A.2 Cathodic Protection t$) Prepackapd rine type reference electrodes shall r] be permteently installed near poorly accessible s") 80.10escription protected surfaces to provide a means cf monitoring prc,tection level by measuring A cathodic protection system is prosided. Its potentials.

design is plant unique as it must be tailored to the l

site conditions. The COL license applicant must (6) Test stations above grade shall be installed provide a design meeting the requirements listed in throughout the station adjacent to the areas Subsection 8A.23. beiag protected for termination of test leads from protected structures and permanent reference 8.U.2 Analysis electrodes.

There are no SRP or regulatory requirements 8A.2.4 References nor any national standards for cathodic protection systems. The system is designed to the requirements (1) Utility Requirements Document, Advanced listed in Subsection 8A.23. Light Water Reactor, Volume !!, ALWR Evolutionary Plant, Electric Power institute 8A.23 COL License information The COL applicant is required to meet the following minimum requirements for the design of the cathodic protection systems. These requirements are the same as those called for in Chapter 11, Section 9.4 of the Utility Requirements Document issued by the Electric Power Research Institute.

(1) The need for cathodic protection on the entire site, portions of the site, or not at all shall be (r~^)

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determined by anclyses. The analyses shall be based on soil resistivity readings, water chemistry data, and historical data from the f.ite gathered from before commencement of site preparation to the completion of construction and startup.

(2) Wher;large protective currents are required, a shallow interconnected impressed current

, system consisting of packaged high silison alloy anodes and transformer rectifiers, shall normally be used. The rectifiers shall be approximately 50 percent oversized in l anticipation of system growth and possible i higher current consumption.

(3) The protected structures of the impressed current cathodic protection system shall be connected to the station grounding grid.

- (4) Localized sacrificial anode cathodic protection systems shall be used where required to supplement the impressed current cathodic protection system and protect surfaces which are not connected to the station grounding grid or

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! is' are located in outlying areas.

Amendment 2t 8A 2 t i

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\v) TA13LE OF CONTENTS Section Title Page 8A3 Electric Ileat Tracing 8A.31 SA 31 Description 8A31 8A3.2 Analysis 8A3-1 SA33 COL License Information 8A31 l

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[ 80.1 Description The electric heat tracing system prosides freeze protection where required for outdoor service components and fluid warming of process fluids if required, either in or out doors. If the operation of the heat tracing is required for proper operation of a safety related system, the heat tracing for the safety related system is required to be safety related, also. Power for heat tracing is supplied from buses backed by the onsite standby generators. Non safety related heat tracing has access to the combustion turtine generator through the same load group as the components protected. Safety related heat tracing is assigned to the appropriate disision for a source of safety related power.

80.2 Analysis There are no SRP or regulatory requirements for cathodic protection systems. They are required to be des:gned and installed to the applicable sections of the following codes and standards.

(1) IEEE Std 622, Recommended Practice for the Design and lastallation of Electric Heat Tracing Systems in Nuclear Power Generating Stations (2) IEEE Std 622A, Recommended Practice for the Design and lastallation of Electric Pipe Heating Control and Alarm Systems in Nuclear Power Generating Stations 80.3 COL License Information No COL applicant information is required.

8A.3A References The following codes and standards have been referenced in this section of the SSAR.

(1) IEEE Std 622. Recommended Practice for the Design aad Installation of Electric Heat Tracing Systems in Nuclear Power Generating Stations (2) IEEE Std 622A, Recommended Practice for the Design and installation of Electric Pipe Heating Control and Alarm Systems in Nuclear Power .

Generating Stations O

Amendment 21 8AM L _ - . __ -- .- -- - . . - - . _ - - - - - _ - - _ - _ _ _ - - _ - - - - - - - - _ - - - _ - - - - - _ - - - - - - - - - - _ . - _ - _ - - - - - - . - - - - - - - - - - - - - - - - - - - -

JSSUE 13 - CORROSION CONTROL. FOR STRUCTURES The SSAR should include an interfv quacmeu for a corrosion control program to identify the potential for the corrosion of structures and componenta and to letermine the corrective measures to be taken. The program should commence prior to the completion of the detailed design of building substructures and underground installations. The program should consider the potential for corrosion from gahanic direct currents which may Gow as the result of copper ground mats on site, including the electrical switching stations' ground mats. The potential for corrosion of containment building substructures and liners should be considered. The mitigation -

measures may include coatings, wrappings, cathodic pro'ection, electrical bonding, elimination of gahanic currents, or other mitigation means.

ISSUE 3 RESPONSE Section 8A.2 of Appendix 8A (see response to issue 12) has been added to the SSAR oroviding the requirement that the COL applicant provide a cathodic protection system tailored to the site conditions.

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