ML20078L295
| ML20078L295 | |
| Person / Time | |
|---|---|
| Site: | Farley  |
| Issue date: | 06/30/1992 |
| From: | FAUSKE & ASSOCIATES, INC. |
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| References | |
| FAI-91-56, NUDOCS 9411290147 | |
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Text
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FAI/91-56 FARLEY NUC12AR PLurr UNITS 1 AND 2 PHENOMENOIDGICAL EVAIDATION
SUMMARY
ON MOLTEN CORE-CONCRETE INTERACTIOtt IN SUPPORT OF THE INDIVIDUAL PLANT FIAMINATION Submitted To:
)
Southern Nuclear Operatin6 Company Birmingham, Alabanna Prepared By:
Fauske & Associates, Inc.
16WO70 West 83rd Street Burr Ridge, Illinois 60521 (708) 323-8750 June 1992
(
9411290147 941109
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PDR ADOCK 05000348 p
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A.BSTRACT d
Phenome ological issues on molten core concrete interaction (MCCI) have support of the Farley Individual Plant Examination (IPE) been examined in to synthesize MCCI knowledge from Program. The approaches taken were (1) experimental data, computer codes and analytical tools, and, (2) to develcp a failure criterion to determine if such a postulated phenomena could chal-lenge containment integrity at the Farley Nuclear Plant.
For Plant Farley, basemat penetration due to molten corium concrete postulated containment failure mechanism. If molten core interaction is a debris breaches the reactor vessel and contacts concrete surfaces, molten corium concrete interaction can occur.
During postulated severe accidents Farley, water pools may be present in the containment at locations where at core debris might collect.
The steaming due to boil-off of this water plus contributions due to containment heating by the debris would over-pressurize l
)
the containment and induce containment failure prior to basemat penetration by MCCI given no containment heat removal system is available, If contain-
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ment heat removal is available, the existing water pools can be recirculated (contain=ent sprays) or refluxed (fan coolers) to cool the debris and prevent containment failure or delay it sufficiently that recovery or acci-dent management ac'; ions could also be successful in establishing a safe sequences such as station blackout or transients stable state.
For some with no injection, no containment sprays and no containment heat removal, If the cavity could become dry shortly after reactor failure.
the reactor debris is not dispersed from the cavity at vessel failure in such cases, and no water sources are recovered, the containment basemat could undergo MCCI.
For the Farley Nuclear Plant, MCCI would most likely occur (if at all) in the cavity region of the containment. MCCI will erode concrete in both the sideward and downward directions, but a failure criterion is based on the downward direction because the downward erosion speed is much faster than the sidevard erosion speed.
The containment is considered failed when the depth of concrete erosion is equal to the combined depth of the cavity n
floor and the containment basemat.
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1 To -eutimate the extent of concrete erosion with~ time, a technique suitable for hand-calculations is presented' along with experimental and analytical results relevant to MCCI This technique expresses conse:vation of mass and energy for MCCI and proceeds in a piecewise fashion to account for. changes in core debris and chemical phenomena over time. The technique is used to find the following: the time interval during which the zirconium.
in the debris pool is being oxidized, the depth of erosion during the zir-conium oxidation phase, and the time to containment failut e by MCCI.
Inputs to the calculation technique stem from t.he analytical and experimental, results presented herein.
Appropriate physical properties required for MCCI calculations (densities, specific heats, etc.) are tabulated herein. Othe'r-parameters, most notably the sideward erosion rate to downward erosion rate
- ratio, are assigned values based upon the-BETA and SWISS test series
-described here.
The Farley IPE plant specific parameter file is used to provide values for the geometry of the cavity and the corium debris itself.
l Calculations were performed with this technique for the " dry" case where the debris cannot be covered permanently by an overlying pool. This scenario is best illustrated by the station blackout (SBO) case. In the SB0 sequence. MCCI begins about two hours after the reactor vessel failure.
These' assumptions lead to a containment failure time due to MCCI of 62 hours7.175926e-4 days <br />0.0172 hours <br />1.025132e-4 weeks <br />2.3591e-5 months <br /> after the initiating event for the dry case. Sensitivity runs indicate that.
this value is a highly conservative lower bound.
For the case with the debris covered, calculations were not performed -
because of the large body of experimental data which has demonstrated the ability of water to rapidly quench molten debris. Once the debris has been j
to maintain it coolable, quenched, water can ingress into the debris bed assuming the overlying pool is replenished. Also, MAAP runs show minimal concrete ablation in those accident sequences where the debris remains covered.
In principle, calculations can be performed for the " wet" case in f
the same manner as for the dry case to, say, consider sensitivity studies of l
ingression.
These calculations are not performed here since they would simply demonstrate large times to failure, i.e.,
hundreds of hours, for basemat failure by MCCI.
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Based on these calculations, MCCI can be excluded as' a containment failure mechanism, relative to other containment failure mechanisms. For the station blackout " dry" case, containment failure would occur due. to over-pressurization long before failure due to MCCI, as demonstrated by-f I
fundamental analyses as well as MAAP results. For the case where the debris is covered, MCCI would be halted by water ingression into the debris bed.
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-ab TABI2 QF CONTENTS i
)
ZA&E
)
i ABSTRACT.
TABLE OF CONTENTS iv LIST OF FIGURIS vi LIST OF TABLES vi 1.0 PURPOSE.
1-1 2.0 PH ENOMENA.
2-1 2.1 Description 21 2.1.1 Controlling Processes 2-2 2.1.2 Relationship to Containment Failure Mechanisms and Modes 2-4 7-_s 2.1.3 Relationship to Source Term.
.24 2-5 2.2 Experimental Results.
2.2.1 SVISS
Sustained Heated Metallic Melt /
Concrete Interaction With Overlying Pool.
.26 2.2.2 BETA:
Large Scale Metal / Oxide Interaction Experiments.
26 2.2.3 SURC:
Sustained-Urania Concrete Interaction Experiments.
. 2-14 2.2.4 Summary of Experimental Results.
. 2 18 2 18 2.3 Analytical Approaches 2.3.1 CORCON.
2-18 2.3.2 WECHSL 2-22
. 2 23 2.3.3 DECOMP 2.3.4 VANESA.
2-24 2-24 2.3.5 MET 0XA.
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,7 g
bl TABIE OF COffrENTS (Continued)
EA&R 3-1 3.0 METHODOLOCY 3.1 Step 1:
Define Failure Criteria.
32 3.2 Step 2:
Determine Overall Time Interval to Exceed Most Limiting Failure Criterion.
3-3 4.0 PLANT SPECIFIC APPLICATION 4-1 4.1 Issues 4-1 4.1.1 MCCI Failure Criterion (Step 1) 4-1 4.1.2 Time Interval to Exceed the Failure Criterion (Step 2).
4-4 4.1.3 Uncertainty Considerations 4-9 4.2 Conclusions 4 11
-s V
5.0 SUKHARY 5-1
6.0 REFERENCES
6-1 APPENDIX A:
Derivation of Concrete Erosion Calculations.
A-1 APPENDIX B:
Code Input and Calculations.
B-1 APPENDIX C:
MCCI Calculation Source Code
.. C-1 (s,)k i
1 1
- Vi O.
r LIST OF FIGURES Pirure No.
f_ art 21 The SWISS experimental apparatus 2-7 2-2 BETA experimental apparatus.
2-8 2-3 Results of the BETA V1.8 experiment.
2 12 2-13 2-4 Results from the BETA V2 experiments series.
2-5 SURC-4 experimental apparatus.
2-16 26 SURC-3 experimental apparatus.
2-17 4-1 Reactor vessel supports for farley 4-2 4-2 Farley reactor cavity compartment 4-3 O
I t
V' LIST OF TABLES Table No.
Paze 2-1 BETA Experimental Matrix 2 10 2-2 SURC Test Matrix 2-15 2-3 Summary of Selected Core-Concrete Interaction Experiment Initial Conditions.
2-19 2-4 Observed Eroded Concrete Depth and Mass.
2-20 31 Molten Core-Concrete Interaction Calculation Data.
3-4 s
1-1
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1.0 PUltPOSE The potential failure of a containment building due to molten core debris attack on its concrete floor has been a subject of concern for the Nuclear Regulatory Commission (NRC)
Staff since the Reactor Safety Study
[NRC, 1975) identified it as the mechanism for two possible containment failure modes.
One concern is that molten core debris ejected from a failed reactor vessel would come into contact with the containment floor and even-tually erode a large enough volume of concrete that either the reactor cavity walls would lose their load-carrying capability or the basemat would be penetrated and core debris would exit the containment. Also, debris attack would generate steam and noncondensable gases that would contribute to the potential for containment failure by over-pressurization.
In (NRC, 1988), the NRC recommends t: et the extent of molten core-concrete interactions (MCCI) be assessed to determine its potential to cause containment failure. The objective of this paper is to develop
)
a strategy t
for accounting for containment failure due to MCCI in the source term por-tion of the Farley Nuclear Plant IPE.
In this paper, only containment failure due to concrete ablation during MCCI is considered.
Over-pressurization caused by noncondensable gases Senerated during MCCI is not considered here.
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(J 2.0 rumsosoma l
2.1 - Desertation Thermal attack of structural concrete in a containment building can occur during a core melt sequence if the molten core debris breaches the reactor vessel and contacts the concrete surfaces in the containment build-ing.
In a PWR plant, the concrete surface that experiences the most severe thermal attack typically is the cavity floor.
The interfacial heat transfer between the core debris and concrete drives the thermal decomposition and erosion of the concrete.
The thermal attack on the concrete can be broken up into three different phases:
a short term, localized attack as debris leaves the reactor pressure vessel; an aggressive attack by high-temperature debris immediately after the core material is released from the primary system; and a long-term attack in which the debris temperature would remain i
essentially constant and the rate of attack is determined by the internal heat generation.
Immediately after vessel failure, debris is discharged from the vessel into the reactor cavity.
This material, which may be molten, induces an aggressive localized jet attack upon the concrete surf ace, given the absence l
of water in the cavity. A pool in the cavity could greatly mitigate this.
l
- attack, depending upon its depth. The thermal attack is localized to the area below the failure.
Estimates of this attack based on analyses in
[IDCOR, 1985]
show the depth to be 3.94 to 7.84 inches (10 to 20 centimeters),
depending upon the primary system conditions at vessel failure.
Af ter the jet attack, the reactor cavity floor region may be covered by i
I high-tempe rature debris which aggressively attacks the concrete substrate.
Free water, bound water, and gases generated by concrete decomposition are then released.
The gases stir the melted material and promote convective heat transfer between this material and the concrete.
The combination of the sensible heat added to the concrete, the endothermic chemical reactions involved in releasing water vapor and decomposing the concrete, and the
22
((l' latent heat of fusion for melting the substrate extracts a considerable amount of energy from the high temperature melted material.
In fact, the aggressive attack generally requires more energy than is generated by the decay power.
Such additional internal heat generation can result from the oxidation of metallic constituents within the melted material by the steam and possibly by carbon dioxide released from the concrete substrate.
Typically, the high temperature, aggressive attack is driven by internal heat generation and secondly, to a lesser extent, by the initial stored energy.
During the long-term attack, the debris remains at an essentially constant temperature, and the rate of attack is determined by the difference between the internal heat generation and the heat losses to the containment environment.
Without water, these heat losses are principally due to con-vection and radiation, and are somewhat influenced by the natural convection of high temperature gases throughout the containment.
The resulting con-crete attack rate is much reduced from that typical of the high-temperature
(~N attack phase and occurs over a much longer interval.
The noncondensable
-/
gases generated during this period contribute to long term overpressuriza-tion of the containment.
2.1.1 Contro111nr Physical Processes The major physical phenomena affecting the extent of concrete erosion by core debris are closely interrelated and therefore difficult to separate.
For the purpose of this discussion, these phenomena are identified as fol-lows:
rate and amount of core debris expulsion from the reactor vessel; melt-water dynamic interaction and melt spreading; configuration of the debris mass on the concrete; depth of the melt bed; and the quenching effect of water.
The following are brief discussions of how the controlling phenomena influence this issue.
The core debris flow rate from a failure ir. the bottom of the reactor vessel determines the initial velocity with which the melt spreads.
As the initial spreading rate decreases, there is an increase in the time for heat rx transfer from the melt to the concrete floor
- and, if present, overlying i
i
23 b,,-)
C water pool.
Thus, the initial spreading rate and and the energy losses from the melt influence the excent of initial melt spreading.
The initial extent of melt spreading is also very dependent on the effects of the water / debris dynamic interaction on the heat transfer from l
1 the molten debris.
If the interactions between the melt and water cause the debris to disperse, or cause violent oscillations at the melt-water inter-
- face, these interactions might lead to heat transfer rates far in excess of heat transfer rates associated with critical heat flux boiling.
If not, the melt-water interface would undergo film boiling rather than nucleate boil-ing.
After the molten debris has initially spread out, two types of debris configurations are possible:
a discontinuous porous debris bed composed of discrete particles; or a continuous slab of partially molten pool.
The debris bed configuration has been shown to occur when sufficient water is available to quench the debris as it leaves the reactor vessel.
Otherwise, a continuous slab configuration typically occurs because there is less water e
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available than would be necessary to ensure debris fragmentation.
It is possible that a debris bed can evolve into a continuous debris slab.
If the porosity in the bed limits the bed dryout heat flux below decay power heat removal requirements, the debris bed would heat up and eventually melt into a continuous debris configuration.
The debris configuration strongly af-fects the quenching capability of overlying water.
The depth of the debris also affects how effectively energy is removed from the debris.
If it is postulated that the initial debris layer is so thick that the rate of heat removal from the debris is exceeded by the internal heat generation rate, then the debris would reheat and ultimately remelt.
Further spreading of molten or partially molten debris would then occur, aided by gas agitation from concrete erosion if the concrete were sufficiently heated.
The debris would spread until the heat removal rate was sufficient to freeze it, or was constrained by the cavity.
In (NRC, 1988),
the NRC has stated that a debris layer less than 9.84 in (25 cm) in depth may be considered to be coolable.
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24 g(3)
The last phenomena to discuss is the quenching effect of water.
For a 1
debris bed, the physical mechanism of cooling is water ingression into the bed with outflow of steam from the bed.
The coolability limit for debris beds is a hydrodynamic limitation within the bed itself that strongly depends on the porosity, and is fairly iidependent of the debris bed depth.
For a thin debris slab or shallow pool, conduction is an effective heat transfer mechanism and the slab can cool quickly with little or no cracking.
However, for a thicker slab or deeper pool, coolability requires cracking of the slab or overlying crust and ingression of water into the debris.
Such cracking would be expected to occur as a result of the volume reduction associated with debris cooling and phase change, as well as sparging of of fgases produced by any thermal attack of concrete.
2.1.2 Relationshio to Containment Failure Mechanisms and Modes Extensive erosion of concrete by high-temperature core debris is a potential late containment failure mechanism that would be expected to occur n
many hours after reactor vessel failure and debris release into the contain-ment.
Two actual failure mechanisms are considered possible as a result of concrete erosion:
penetration of the containment basemat; and sufficient deterioration of the load-carrying capability of the cavity walls that the reactor vessel moves and causes gross mechanical failures of penetrations for piping connected to the reactor vessel.
Both of these containment failure mechanisms would be expected to result in large containment failure areas (on the order of several square feet).
2.1.3 Relationshin to Source Term long-term molten core-concrete attack leading to containment failure influences the exTected fission product release for a sequence by providing a large gas and/or corium flow path out of the containment long after vessel failure.
The effect on the source term, however, depends on the presence of water in the cavity during a sequence prior to containment failure.
For a dry sequence, since the airborne fission product concentration in a closed containment tends to decay with time due to naturally-occurring fission product retention mechanisms, a late containment failure due to MCCI
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generally results in modest source terms, relative to early failures.
For-gross mechanical failures of penetrations for piping,
- however, the' rela-tively large. expected-failure size assures a rapid blowdown of the initially available airborne fission products to the environment.
Fission products entering the containment atmosphere after the blowdown would experience
]
little driving force from the containment.
Thus, fission products evolved by long-term revaporization within the reactor vessel'would be subject to the naturally-occurring deposition mechanisms in the containment.
Fission i
products released ex-vessel during MCCI would also be subject to the same deposition mechanisms.
1 Basemat penetration by MCCI is a late containment failure mode that leads to a large failure area.
A rapid blowdown is not assured because the soil underneath the basemat provides an obstructed release path.
This also R
results in very different offsite consequences, relative to other contain-mont failure modes, because the release is to underlying soil rather than the atmosphere.
Soil provides a very effective fission product scrubbing mechanism, but a basemat failure can potentially contaminate the local water O-table.
For a wet sequence, the expected fission product source term would be i
reduced relative to a dry sequence due to the capture of fission product aerosols by the overlying pool and/or containment spray droplets.
2.2 Ernerimental Results A wide variety of molten debris-concrete interaction experiments have been performed at Sandia National Laboratory (SNL) and Kernforschungszentrum Karlsruhe (KfK).
In general, these experiments were undertaken to validate models and understanding of MCCI phenomenon. The SNL experiments considered either initially metallic or oxidic melts. Metalli experiments at SNL were usually performed with a stainless steel melt charge, while oxidic experi-ments were performed with a uranium dioxide and zirconium dioxide melt charge. Two SNL tests have employed both oxides and zirconium metal.
The KfK BETA series of experiments featured stainless steel and concrete oxide melt charges.
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2.2.1 SWISS
Sustained Heated Metallic Malt / Concrete Interaction with Overivinn Pool s
A schematic representation of the SWISS experimental apparatus is shown i
in Figure 2-1.
The SWISS experiments performed at SNL consisted of two test runs designed to obse rve the effect of an overlying pool on MCCI.
The l
l SWISS-1 test involved 99.2 lb (45 kg) of stainless steel initially at 3005.6 l
F (1925 K) inductively heated over a period of about 40 minutes with one 6
power-interruption (Blose, 1987).
An average of about 0.205 x 10 Bru/h (60 kW) net power was input to the melt charge by the induction source. About t
I 32 minutes after the melt was teemed into the crucible,- water was poured over the melt.
Erosion was essentially one-dimensional, downward into a 8.5 j
in (21.6 cm) diameter limestone common sand concrete plug, because a non-ablating MgO sidewall was used.
A value of 95099 Btu /h ft2 (300,000 W/m )
j 2
was suggested as an average sideward heat flux lost to the Mg0 crucible walls.
The concrete was a limestone-common sand variety essentially identi-
- cal to a default type used in the CORCON code [Muir, 1981]
and typical of
/
many U.S. plants.
Test SWISS-2 was similar to test SWISS-1 except that water was con-l i
tinuously poured onto the debris immediately after the melt teem, and two 2
2 power interruptions occurred. A'value of 63400 Btu /h-ft
'(200,000 W/m )
j heat flux to the Mg0 sidewall was suggested on the sideward heat losses.
i I
2.2.2 BETA:
Larre Scale Metal / Oxide Interaction Exneriments The BETA test series consisted of a series of large scale metallic melt l
experiments conducted at KfK research center ( Alsmeyer, 1986).
These ex-periments were originally conceived and designed to provide reliable experimental data for verification of severe accident programs, such as the 5
WECHSL code [Reimann, 1981].
i The BETA tests typically consisted of large, high temperature metallic
- melts, initially at 3632*F (2000*C) or higher, produced by thermite reac-tion.
These melts were usually poured into an instrumented siliceous concrete crucible of 15 in (0.38-m) inside diameter (see Figure 2-2)..
The 7y
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2-1 O
Ouartz Port rTop Hat Gate Valve j
=
1 EL. 1 1. 5 8 f t.
k 5j h
h Instrumented Flow Tube
- /
=b 8all Valve Aerosols Gas Sampling Expansion Chamber Flow Rate
% Pneumatic Cylinder Melt Penetrator Mel Generator (Mgo)
,H EL.
- 5. 8 8 f t.
10 Zirconia insulation 5j L-induction Co6l Stainless Steet/
8.485 in.
(99.21 lb.)
d, Water inlet Water Exit Cructble (MgO)
- EL. 0. 7 2 2 f t.
h
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0.0 ft.
Concrete Delivered Melt Figure 2-1 The S'JISS experimental apparatus.
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- Molten Pool-Material Sampling &
Thermocouples Movie & TV Cameras Periphery Systems-Infrared Pyrometer Thermite Mass Guages allo Pyrometer Thermrte Jet Pyrometer p
- Exhaust Gas T
rhermocouoles AY%
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Pressure Trasnsducers Mass Spectroscope A
~
A g Aerosol Detector input Power CpnCf 8t8 CTUCible --
Heat Losses
{Thermocouples Water Condensate Balance Mosture Detectors
, Ught Sensors arn.sessie a.A
@ Concrete Crucible Thermite Reaction Tank
@ induction Coll Container for Measurement Probes
@Offgas System rigure 2 2 SETA experisantal apparatus.
2-9 n
interactions were sustained by inductive heating.
The BETA experimental 6
facility has the capability to supply up to 6.48 x 10 Btu /h (1.9 MV) of inductive heating power, which allowed observations of the various physical and chemical processes occurring during molten melt concrete interactions for an extended period of time under quasi-steady state conditions.
Typical j
melt charges in the BETA tests were composed of 661.4 lba (300 kg) of steel and 330.7 lba (150 kg) of oxides.
The steel components typically included Fe, Cr, and Ni.
The oxides useu were aluminum oxide (Al O ),
silicon oxide
)
23 f
(SiO ), and in some cases calcium oxide (Cao).
Table 2-1 summarizes the 2
BETA experiments conducted over the span of two years, from 1984 to 1986.
The composition of the silicate concrete utilized in the experiments is somewhat similar to the basaltic concrete variety used in the United States.
1 BETA V3 series was an exception as it employed limestone common sand con-crete.
No fission product simulants were employed in any BETA tests, l
Major results of BETA tests are (Alsmeyer, 1987)):
1.
The temperature of the melt pool dropped rapidly from an initial temperature close to 3320.6 F (2100 K) to a temperature slightly above the mixture melting temperature, even when the inductive heating power was very high.
2.
The experiments are characterized by the dominance of downward erosion in high power tests, and to a lesser extent in low power tests.
3.
Dispersion or entrainment of the metal layer into an overlying oxide layer was observed in the high power tests.
It appears that the process of dispersion is driven by the high gas fluxes evolving from the concrete.
l 4
When Ca0 was added to the melt pool, no entrainment of metal into the
>xide layer was observed.
Apparently, Ca0 serves to lover the viscosity of the oxide layer, and thus, the metal-oxide layers remained stratified.
It is noted that all low l
power experiments resulted in separated metal and or.ide layers.
l
2-10 v
I_able 2-1 BETA EXPERIMElfrAL MATRII Text Melt Power /(Btu /hr)
Power /kW Remarks V0.1 Iron 0
0 Test V0.2 Iron 1.36 400 of V0.3 Iron + 0xide 5.80 1700 Facility V1.1 Iron Pulsed Pulsed Pour failed V1.2 Iron + 0xide Pulsed Pulsed Lorentz-force excluded V1.3 Steel + 0xide 3.41 1000 V1.4 Steel 0
0 Transient V1.5 Steel 1.54 450 V1.6 Steel + Oxide 3.41 1000 V1.7 Steel + Oxide 5.80 1700 V1.8 Steel + 0xide 6.48 1900 No dispersion (Ca0)
V1.9 Steel + Oxide 1.36-0.68 400-200 cao added V2.1 Steel + Oxide 0.41-0.51 120-150 V2.2 Steel + O.ide 0.17-0.31 50-90 Ca0 added V2.3 Steel + 0xide 0.82 240 Cao added V3.1 Steel + 0xide 5.80-8.53 1700-2500 US Lime /Comm Sand, Heating from 0-66 s only V3.2 Steel + 0xide 1.36->3.41 400 >1000 US Limestone, 30 min heating V3.3 Steel + 0xide 2.05-0.68 600-200 US Lime /Comm Sand, 60 min heating V4.1 Steel + Oxide 3.41-1.02 1000-300 19.685 f t dia crucible pL)
2-11
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6 i
In the BETA V1.8 experiment, 6.48 x 10 Btu /h (1900 W), the highest net inductive heating power of the series was used, corresponding to an internal power density more than a factor of ten higher than that of decay power.
A melt charge composed of 771.6 lbm (350 kg) of steel (82%
Fe, 104 Cr, 8% Ni) and 286.6 lba (130 kg) of oxides (70% Al 0, 30% Cao), produced 23 by a thermite reaction, was poured into the interaction cavity at an initial temperature of 3266.6 F (2070 K).
The observed downward erosion rate was approximately 0.04 in/s (1 mm/s), resulting in a total eroded depth of about 15.75 in (40 cm) in 7 minutes.
Sideward erosion was stated as only a few centimeters (see Figure 2-3).
Experimental results showed that the chromium metal had oxidi::ed con -
pletely in about one minute, and hand calculations indicate that ga.:
evolving only from the concrete floor was insufficient for complete oxidt.-
tion of the chromium.
- Thus, gases released from decomposition of the sidewall must have entered the debris pool and contributed to chromium oxidation.
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observed to be clearly separated from the metal The oxide layer was layer and well mixed with concrete slag, thus entrainment of metal into the 6
oxide layer was not observed.
In a similar high powered (~ 5.8 x 10 Beu/h) test V1.7 in which no Ca0 was added, complete entrainment of metal into the oxide resulted.
The BETA V2 series are low power experiments designed to investigate crust formation and its influence on the melt-concrete interaction. The l
i V2.1 test consisted of the melt charge composed of 661.4 lb (300 kg) steel (90% Fe, 10% Ni) and 330.7 lb (130 kg) oxides (70% Al 0 30% M0 ) initially 23 2
at 3632*F (2000*C).
The melt pool was sustained at an induced power of approximately 4.09 x 10 to 5.12 x 10 Btu /h (120 to 150 W) for more than 100 minutes.
This power level is representative of long-term decay power.
A total of about 13.78 in (0.35 m) downward and 3.94 in (0.10 m) sideward erosion was observed, with an average 0.0024 in/s (0.06 mm/s) downward erosion rate (see Figure 2-4).
In the experiment, sidevard erosion stopped at about 1000 seconds after a depth of 3.94 in (10 cm) was attained.
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2 12 O
Axial 400 E
l E
i
/ v~1mm/s-0 200 e0 T
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W 100
,I O
-c.ei,
s, 0
100 200 300 400 500 TIME, sec Tigure 2 3 Results of the BETA VI,8 experizanc.
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500 E
400 E
Axi 300
,e g
O
/9 g
200 o
3 Radial 100:
o I
O o/
O 1000 2000 3000 4000 5000 TIME, see Figure 2 4 Results from the BETA V2 e9eriments series.
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2 14
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o 2.2.3 SURC: Sustained-Urania concrete Interaction urneriments Sandia National Laboratories is conducting an ongoing. series 'of Sustained Urania Concrete (SURC) tests.
Table 2-2 identifies the SURC experiment matrix, and at the time of this writing, experiments SURC 3, SURC-3A, and SURC-4 are complete.
i The interaction crucible for the SURC 4 experiment
[CSNI, 1989]
con-sisted of an Mgo annulus of 15.75 in (40-cm) inside diameter with a basaltic concrete plug at its base.
A schematic of the SURC 4 experimental apparatus is illustrated in Figure 2-5.
The test charge for SURC-4 consisted of 440,9 4
lb (200 kg) SS-304 (approximately 73% Fe, 19% Cr, 8% Ni) inductively-heated in the interaction crucible to approximately 2690.6 F (1750 K) prior to the onset of erosion.
Fission product simulants were employed to simulate r
fission product release, consisting of 1.1 lb (0.5 kg) Te, 2.58 lb (1.17 kg) of La 0, 2.71 lb (1.23 kg) Ce0, and 2.43 lb (1.1 kg) Bao.
i 23 2
5 A net 2.05 x 10 Btu /h (60 kW) inductive heating power was applied to O
the test charge for the duration of the test and a value of 31700 Btu /h-ft j
2 (100,000 V/m ) was suggested as representative of the heat loss through the Mgo wall.
Because of the relatively inert nature of the Mg0 annulus sidewall to the molten debris, only one-dimensional downward erosion was possible. After quasi-steady concrete erosion was initiated, 44.1 Ib (20 kg) of zirconium were instantaneously added to the molten debris to observe the effect of oxidation.
Observed erosion was between 9.65 in to 10.83 in 5
(24.5 to 27.5 cm) of concrete.
{
The melt charge for SURC-3 [Bradley, 1987), composed of 110.2 lb (50 kg) of SS 304 (734 Fe, 196 Cr, 8% Ni), was inductively heated to ap-l proximately 3050.6 F (1950 K) prior to being teemed into the interaction crucible.
The interaction crucible consisted of an Mgo annulus of 7.87 in (20-cm) inside diameter and a limestone-common sand concrete plug base.
A schematic of the SURC-3 experiment apparatus is shown in Figure 2-6.
A net 1.024 x 10 Btu /h (30 kW) inductive heating power was applied to the melt 2
charge for the duration of the experiment. A value c,f 31700 Btu /h-ft j
2 (100,000 W/m ) was suggested as representative of heat loss through the Mg0 O
i 2 15 y
Iable 2-2 SURC TEST MATRIX Test Melt Charge Mass (1b)
Mass (kg)
Concrete Water i
UO -ZR0 -ZR 551 250 LIMESTONE NO 2
2 2
UO -ZR0 -ZR 551 250 BASALTIC NO 2
2 l COMPLETED l 3
STEEL-ZR 110.2 50 LIMESTONE NO 3/s STEEL ZR 110.2 50 LIMESTONE NO
(~
4 STEEL-ZR 441 200 BASALTIC NO v
5 UO -ZR0 -ZR 551 250 LIMESTONE YES 2
2 6
UO -ZR0 -ZR 551 250 BASALTIC YES 2
2 7
STEEL-B C 441 200 LIMESTONE NO 4
8 STEEL-B C 441 200 BASALTIC NO 4
O
2 16 i
O Crucible Cover (Mgo)g k
l Water Coofed
\\
Aluminum Containment i/
Vessed i
l (2 Sections) l Interaction Crucible l
(MgO)
% 304 Stalnless
'N/A VD/
Steet Tube Induction Coll k
h x% t s
l Argon Purge i
0-Ring h
h i
N!
13 l
1:~Pf'$d.h gPower Connections MgO Bricks %
[
ig O
k Power Feed 1
/ 4
'~
"O ##
Basalt Concrete-a.
Slug Instrumentation Feed Scale 1/10 ng,yg pon
. m eoc iis a.a rigure 2 5 SURC 4 experimental apparatus.
O
2-17 o
\\
1 2 Color Pyrometer j
Quartz Port i
d Gate Valve vFlow Tube Top Hat J
Expansion Chamber l
Pneumatic Cylinder Gas Sample Tube l
en enetrator O
Melt Generator (Mgo)
Zirconta insulation Stainless Steel (110.23 fb)
Thermocouple s Type X & C)
Zirconium --
A Delivery Tube Aerosol Sample Tube Zirconium Metal Crucible (Mgo)
Solenoid Delivered Melt Induction Coil l
Type K Thermocouplesf
\\(Alumina Tube ype C Thermocouples)
Conerteta nuoone os Fig *re 2-6 St.*RC-3 experimental apparatus.
^~
'2 18 sall.
Because of the relatively inert nature of the Mgo annulus sidewall to the molten debris, only one-dimensional erosion was possible. After quasi-steady concrete erosion was initiated,-3.09 lb (1.4 kg)- of zirconium were instantaneously added to the molten debris.
Observed erosion was about 13.8 in (35 cm) of concrete.
2.2.4 S'
-rv of Wunerimental Results In summary, we can compile the results from several experiments with different scales, different debris material masses and composition, and different power levels to develop a basis on which to provide a method for assessing MCCI and potential failure times (see Tables 2-3 and 2-4).
These data have been used to benchmark the core-concrete interaction models used here and in the MAAP code (DECO'P) [EPRI, 1990; Fauske & Associates, Inc.,
A 1990).
These data also provide a basis for estimating the fraction of the debris' downward heat flux to be 0.5 and for estimating the ratio of.
sideward to downward erosion to be 0.2.
(
2.3 Analytical Anoroaches Results of numerous predictions for the experiments described above have been reported in the literature. However, rather than discuss the results of all of these calculations, the principal computer codes used to obtain the results will be discussed.
The principal computer codes used for.
thermal-hydraulic calculations related to core-concrete interactions are
- CORCON, WECHSL, and DECOMP; the principal computer codes used for aerosol release calculations related to core-concrete interactions are VANESA-and MET 0XA.
2.3.1 00EOGK Sandia National Laboratory's CORCON [Muir, 1981] is incorporated in the NRC's MELCOR code [ Summers, 1990) and the Source Term Code Package (STCP)
[Cieseke, 1986). CORCON models the two-dimensional attack of concrete by core debris during severe accidents in light water reactors.
It assumes an immsdiate separation of the molten core into immiscible layers; namely, a metellic and an oxidic layer.
As the MCCI proceeds, the decomposed concrete t
?
?
O O-O Table 2-3 SilMMARY OF SELECTED CORE-CONCRFTE INTERACTION EXPERIMENT INITIAL CONDITIONS b
Diameter Debris Metal Oxide Temperature Input Test (m/in) lleight' M as Mass (K/F)-
Power (m/in)
(Kg/lh)
(Kg/lb)
(kW/Bru-h) -
L SWISS-1 0.216/8.5 0.16/6.3 45/99.2
-/-
1925/3005.6 55/187660 SWISS-2 0.216/8.5 0.16/6.3 44/97
-/-
1925/3005.6 60/204720 SURC-1 0.216/8.5 O. I 8/7.09 50/I10.2
-/-
1970/3086.6 30/102360 SURC-4 0.40/15.75 0.20/7.87 200/440.9
-/-
1746/2683.4 60/204720 BETA VI.8 0.38/15.0 0.78/30.7 350/771.6 130/286.6 1970/3086.6 1900/6482800 BETA V2.1 0.38/15.0 0.72/28.3 300/661.4 130/286 6 2173/3452 120/409440 I
'Appmaimate initial collapsal height.
6Either at melt time or at start of crositm.
1
..,.w e
2 m
m
n, 2-20 p,
Table 2-4 OBSE3tVED 30DED CONCRETE DEPnl AND MASS Experiment Eroded Concrete Depth Eroded Concrete Mass in (cm) lbm (kg)
SWISS 1 6.7 (17) 31.5 (t%.3)
SWISS 2 6.7 (17) 31.5 (14.3)
BETA VI.8 side 0.78 (2) 229.5 (104.3)b down 16 (40) 942.9 (428.6)c BETA V2.1 side 3.9 (10) 200.9 (92.3) down 14 (35) 2223 (1010.4) gs N-SURC-4 10 (26) 165 (75.1)
SURC-3 14 (35) 55.7 (25.3) 3
" Mass derived from concrete density (p - 2260 kg/m ) multiplied by estimated volume of concrete eroded, bLow estimate based on 1-D eroded depth only.
"High estimated based on 2-D eroded depth and with assuming cylindrical shape of molten corium.
C\\
V
2-20 s
i y
l Table 2-4
}
OBSERVED ERODED CONCRETE DKPTH AND MASS i
Experiment Eroded Concrete Depth Eroded Concrete Mass in (cm) lba (kg) i SWISS-1 6.7 (17) 31.5 (14.3)
SWISS-2 6.7 (17) 31.5 (14.3)
BETA V1.8 side 0.78 (2) 229.5 (104.3)b down 16 (40) 942.9 (428.6)*
BETA V2.1 side 3.9 (10) 200.9 (92.3)
{
down 14 (35) 2223 (1010.4)
-s g SURC-4 10 (26) 165 (75.1) 1 SURC-3 14 (35) 55.7 (25.3) l 3
" Mass derived from concrete density (p - 2260 kg/m ) multiplied by estimated volume of concrete eroded.
1 Low estimate based on 1-D eroded depth only.
- High estimated based on 2-D eroded depth and with assuming cylindrical shape of molten corium.
r l
+
i
2 21 i
Ds t
forms a second oxide layer.
The orientation of layers depends on their a configura-relative density.
During the early stage of concrete attack, tion of heavy oxide-metal-light oxide is possible.
As the MCCI proceeds, l
{
the heavy oxide layer is diluted by the molten concrete (slag) and even-point where the density of the heavy oxide layer is less tually reaches a than that of the metallic. At this time, the heavy oxidic and metallic layers are assumed to flip instantaneously, and the heavy oxidic layer combines with the light oxidic layer.
CORCON also assumes a gas film exists at the molten core / concrete interface. Heat transfer from the molten core l
pool to concrete is governed by the convective and radiative processes across the film.
In CORCON, the concrete ablation is calculated based on i
steady-state one-dimensional energy balance at the molten core / concrete interface.
The decomposition heat of concrete is calculated based on the user specified concrete decomposition temperature. Transient heat conduc-l tion into the concrete is not modeled.
CORCON also assumes that, upon solidification, crusts will form at one or more interfaces with the interior of the layer remaining liquid.
If part or all of a layer becomes frozen, energy can only be transferred by heat conduction through crusts, which is ordinarily far less effective than j
convection.
Because of internal heatin5 of debris and the fact that cooling cannot continue unless heat losses exceed sources, freezing is largely self-limiting.
A simple quasi-steady-state crust formation model was included in j
the CORCON code to calculate the thickness of the crust and the rate of heat transfer from the partially solidified core debris to concrete.
In order to i
model the crust formation process, some assumptions were made in CORCON to i
calculate the liquidus and solidus temperatures of the metallic and oxidic mixtures.
For the metallic mixture, a simple fit to the iron chromium-nickel ternary phase diagram was used.
The presence of metals other than Fe, Cr and Ni was ignored.
CORCON treats the oxidic mixture as a pseudo-j binary system in which the fuel oxides (UO2 + Zr0 ) f rm ne " component" and 2
concrete and steel oxides form the other: the two components are assumed to 1
form an ideal solution in both the liquid and solid phases, i
In CORCON, the oxidation reactions between the metallic constituents and the concrete decomposition gases are assumed to proceed to equilibrium I
i
2-22 V
concentrations defined by minimization of the Gibbs free energy for 38 chemical species composed of 11 elements.
Carbon is among the 38 species considered in the chemical reactions calculation of CORCON and therefore, the coking reactions are modeled by CORCON.
2.3.2 VECllSL
- WECHSL, KfK's stand-alone code
[Riemann, 1981),
models the two-dimensional concrete attack of the core debris during severe accidents in light water reactors.
It assumes the existence of an underlying metallic layer covered by an oxidic layer of fuel, metal oxides, and/or the concrete decomposition products.
Heat transfer from melt to concrete is modeled in analogy to boiling phenocena.
A film model, discrete bubble model, or transition boiling model for heat transfer are used according to the exist-ing gas flow and inclination of the interface.
When freezing processes become important due to cool-down of the melt,
/
crust formation is modeled starting from the melt / concrete interface and possibly resulting in a fully frozen layer.
Crusts are assumed to be perne-able to the gases.
Transient heat conduction into the concrete is not considered in the WECHSL code. The solidus and liquidus temperatures of the metallic mixture in the VECHSL code are calculated by a simple fit to a chromium nickel iron ternary phase diagram.
The same set of equations is used in the CORCON and WECHSL codes to predict solidus and liquidus tempera-tures of the metallic mixture.
In VECHSL, the liquidus and solidus temperatures of the oxidic mixture are determined either by a " binary" phase diagram or by a user input table.
In VECHSL, the gases released from the decomposing concrete may oxidize the constituents of the metal layer. The metal oxidation reactions take place in the order Zr, Cr, and then Fe.
This means that no Cr-oxidation is considered as lon8 as Zr is present in the melt, etc, The rate of oxidation i
for Cr and Zr is limited by the gas supplied from the concrete.
A tempera-ture dependent equilibrium constant is used to determine the oxidation rate of Fe.
bv
2 23 2.3.3 DECONF DECOMP is a subroutine from the MAAP code [EPRI, 1990).
It considers the' debris to be either a solid cylinder or a molten pool surrounded by a
- crust, depending upon its energy.
Crust growth / shrinkage based on energy balance describes the solidification process occurring within the molten
- debris, and temperatures are determined from phase. diagrams based on the composition of the debris. Transient conduction calculations are carried out in the concrete floor, sidewall in contact with debris, and upper wall (surroundings), and concrete ablation is allowed in all heat slabs.
For the case that the debris is partially molten, the heat transfer coefficient at the molten pool / crust interface is a user defined constant.
The liquidus and solidus temperatures of the metallic phase of the mixture are determined as follows:
(i)
The CORCON fit to the iron-chromium-nickel ternary phase diagram is used to determine the liquidus and solidus of the O
steel mixture containing Fe, Cr and Ni.
(ii) The Fe-Zr binary phase diagram with Fe replaced by stainless steel is used to obtain the liquid and solidus temperatures of Zr-steel mixture.
l In DECOMP the treatment of liquidus and solidus temperatures of the oxidic phase is similar to the CORCON treatment of oxides.
The chemical reaction model of DECOMP considers chemical equilibrium by a Gibbs free energy minimization technique.
Chemical equilibrium is calcu-lated in METOXA, another subroutine from the MAAP code, by solving the algebraic relations for mass action and element balances..An ideal solution model is used with three phases:
liquid metal, liquid oxides and 5ases.
Non-ideality is allowed for some compounds through user-input activity coefficients.
A Newton-Raphson technique is employed with a reduced set of algebraic relations, known as basis equations.
The equations are solved to yield the molar abundances of particular species important to a stable,
2 24
<-'s s
i efficient solution.
Reactions and element balances for the basis set are
'~
given in [EPRI, 1990].
2.3.4 VANESA The VANESA code calculates the release of fission products and struc-tural material during MCCI. VANESA models the vaporization of melt species into gases which are produced from concrete decomposition.
The ther-mochemistry and kinetics of this process are modeled mechanistically.
As the gases exit the melt, aerosol formation from bubbles breaking the melt surface and from the condensation / nucleation of vapors is modeled empiri-cally.
The corium is modeled as a layered two-phase system; an oxidic layer above a dense metallic layer which is in contact with the concrete basemat.
The reaction of CO and HO with the maj or metallic constituents are 2
2 evaluated to determine the equilibrium oxygen potential.
This oxygen poten-
[ )\\
tial is assumed to hold for the oxide phase and is used to calculate the s
equilibrium vapor pressures of species in the M 0-H ternary phase diagram where M is the element of interest.
The chemical species considered in the VANESA code are given in [ Lee, 1985).
A kinetic analysis, which considers condensed phase transport, transport across the gas / melt interface and gas phase transport, is then performed to estimate the amount of material trans-ferred from the melt to gas bubbles.
2.3.5 MET 0XA The fission product chemistry of MET 0XA, a MAAP subroutine, is described above in Section 2.3.3.
In METOKA, aerosol generation occurs only through chemical reactions.
Mechanical aerosol generation is not modeled.
all capable of producing MCCI simulations that agree These codes are i
with experiments given they are used by people experienced in the respective code.
In particular, the heat balance on the debris that apportions the distribution upward, sideward, and downward is a key modeling consideration
/"~N l
2 25
('T-)
and must be properly described via the code model inputs and boundary condi-tions.
This was demonstrated in an international standard problem that employed the SURC-4 data (CSNI, 1989). The predicted erosion depths for the various codes were shown to be in good agreement with the experimental results.
i Ov O
i l
i
3-1 7
3.0 MET 110D01DGY The aggressive attack on concrete by molten core debris may lead to a late containment failure if actions are not taken to cover the debris with water and thereby arrest the attack.
Even with the appropriate response, deep core debris beds may not be coolable.
In (NRC, 1988), the NRC states that experimental evidence exists that suggests core debris beds with depths greater than 9.84 in (25 cm) may not be coolable.
A simple, stand-alone calculation procedure is included in the method to estimate containment failure time due to MCCI.
Although MAAP, or a comparable integral code, should be used to address all aspects of MCCI for a sequence (including non-condensable gas generation, ex-vessel fission product release, etc.),
the method provided here can be used to estimate containment failure time due to concrete erosion for basic sequences.
The procedure outlined here has three advantages:
(1) it obviates lengthy MAAP
(]
- runs, (2) it provides a convenient means of investigating ablation phenomena, and (3) it is scrutable to independent reviewers.
In the event the debris is not coolable, the calculation method to be used to determine the possibility of containment failures caused by molten core-concrete interaction is based on the following simplifying assumptions:
Erosion of the upper portion of the cavity wall by radiative heat transfer is negligible compared to the direct attack of the core debris.
Erosion proceeds equally at all locations covered by debris.
Zirconium is the only important source of chemical energy in the core debris.
The ratio of the rates of advance of sidevard and downward concrete attack is constant.
O
32 The evaluation method consists of the following steps:
Step 1:
Define failure criteria for basemat and reactor cavity walls.
l t
Step 2:
Determine the time required for the concrete attack by j
core debris released from reactor vessel to exceed the most limiting failure criterion identified in Step 1.
'i Step 3:
Based on extent of concrete attack, select the ap-propriate treatment of containment failure caused by molten core concrete interaction.
3.1 Sten 1:
Define Failure Criteria Core debris-concrete interaction is hypothesized to be able to cause containment failure either by penetrating the cavity floor,
- liner, and basemat or, by weakening the reactor vessel supports sufficiently that the O
reactor vessel and attached piping move and tear out associated penetrations through the containment.
For the latter failure type, the failure criterion must account for the weakening of the cavity wall as this wall's thickness decreases due to erosion, as well as the associated weakening of the wall l
due to the dehydration of the concrete that-ocr:urs ahead of the erosion i
front.
The-criterion for failure of the cavity wall may be stated function-ally as follows:
i i
, it {1 - f ] x (3-1) x y
p t
where:
x,,, - total concrete erosion in the sideward direction, ft, x - width of the cavity wall, ft, p
fy - minimum fraction of cavity wall required for reactor vessel support.
l
~
3-3 Note that for a large, dry PVR the width of the cavity wall is large. Values for f must be in the range 0 < f s 1; a value of 0.5 for f would increase-y y
g stresses in the containment wall by about a factor of two and would repre-i sent. removal of design margin corresponding to a safety factor less than two in. the original design (since the dehydration front slightly decreases the f
load bearing fraction of the wall cross sectional area).
For basemat penetration, the failure criterion is simply:
e,d * *f * *b I3'2) f where:
x
-t tal e nerete erosion in the downward direction, ft.
e,d g - thickness of the cavity floor, ft, x
x - thickness of the basemat, ft.
b 3.2 Sten 2-Determine overall Time Interval to Exceed Most Limitina O
Failure Criterion The calculation method presented below is based on [Plys, 1987].
Typical values for data required by the calculation are provided in Table 3-1.
The calculation proceeds in a piecewise continuous fashion to account for changes in the governing phenomena as time passes.
The first phase of the erosion process is considered to be an interval during which unoxidized zirconium in the core debris is assumed to react with steam and carbon dioxide liberated by the concrete erosion.
During the second phase of the erosion process, the chemical reaction energy is considered negligible and the concrete decomposition enthalpy reduced to reflect the assumed lack of chemical reactions.
The calculation method proceeds by determining whether containment failure would be predicted to occur before the zirconium in the core debris bed is depleted.
If containment failure occurs first, then the time at which containment failure would occur is obtained by a straightfor-ward calculation.
If zirconium depletion would occur first, then the i
calculation procedure becomes more complicated.
First, the time interval to O
i
3-4' Table 3-1 MOLTEN CORE-CONCRETE INTERACTION CAlfUIATION DATA P
Molar Veights: A
- 91.22 lb/lb mole zr A
- 44 lb/lb mole CO2
^H O - 18 lb/lb mole 2
Heats of Reaction: Q
- 1.2898 x 10' Btu /lba (3 x 108 J/Kg)
N0 (pg,, 19g7) 2 y
Q
- 1.1608 x 10 Btu /lbs (2.7 x 10 J/Kg) 1 C02 (no coking assumed) [Plys, 1987]
Limestone /
Concrete Procerties Common Sand Basaltic Limestone en (1bm/ft3) 143.59 143.59 143.59 Q
p 4
A' (Btu /lba) 1539 1242.5 1780
[ Lee EI}d Kazimi, 1985) i f
.047
.049
.041 H0
[Leeandkazini, 1985) i f
.22
.015
.357 C0
[LeeandK$zimi, 1985)
'i Miscellaneous [Plys, 1987]
A
+ "p slag
~
en cn c
- 0.23885 Btu /lba F p,,7 AT - 932*F during zirconium reaction
- 392*F during iron reaction O
1 3-5 l
i 1
A reach zirconium depletion is determined.
Then, an iterative calculation is required to determine the predicted depth of concrete erosion corresponding to zirconium depletion.
Finally, the additional concrete mass that must be eroded to cause containment failure and its corresponding time interval must be determined.
To determine whether containment failure would be predicted to occur before zirconium depletion occurs, the masses of concrete that would be eroded at these times are calculated and compared. To estimate the mass of concrete that has been eroded when the zirconium is completely oxidized, the number of moles of zirconium available for reaction (N ) must be deter-mined:
zr/A (3-3)
N
-m where m
- mass of unoxidized zirconium in the core debris contained in the cavity, ibm, A
- m lecular weight of zirconium, lb/lb-mole.
zr The governing chemical reactions are taker to be:
Zr + 2H O - Zr02 + 2H2 + 2QHO I
)
2 2
2 2+C' (3-Sa)
Zr + CO
- Zr0
+ 2QC0 2
C+CO - 2C0 (3 5b) 2 where:
- heat of reaction per mole of gas produced, Btu /lb-mole QH 0'9CO 2
2 (J/Kg-mole).
Thus, by Equation (3 4) two moles of gas are produced for each mole of zirconium reacted:
36 N - 2N (3 6) g zr The number of gas moles produced per unit mass of concrete eroded (n ) is given by:
i f
f (3 7)
E
^H 0 +A n
CO 2
2 where:
- e neret4 gas mass fractions for steam and carbon dioxide, fH,0'fC0 2 olecular weights of steam and carbon dioxide, lb/lb mole.
^H 0*ACO 2
2 The total concrete mass that would be eroded durin6 P ase 1 (men,zr} '*7-h then be estimated as:
(3-8)
~ "g/"g men,zr The mass of concrete that would be eroded at the time either of the failure criteria are satisfied (see Equations (3-1) and (3-2)), may be estimated using Equation (A-6b) from Appendix A:
d *e,s/#s (1ba)
(3-9a) en,s " #cn m
s cn.d ~ #cn 's d *e d (Ibm)
(3 9b) where:
3 p
- e nerete density, lbm/ft j
en A
rate of total to downward volumetric erosion rate, f,-1-+r,[d i
average area subjected to downward erosion over the' erosion time d
interval of interest, ft t
r, - ratio of sidevard rate of erosion to downward rate of erosion.
O 1
37 failure i
are less than men,zr, then containment If either m "cn.d en,s i
would be predicted to occur before the zirconium in the core debris mass was fully oxidized.
The time interval of interest may be estimated using Equation (A 8) and the following expression for the integrated core decay power (El-Vakil, 1978]:
- 2 0.74 o 74 t
- t (3-10)
Q de - 0.128 Q, DK C 1 where:
r Q
~ *
- d***7 P "*#'
DK Q, - core full power. Btu /h (or %'),
start of time interval of interest after core shutdown, second, t
y end f time interval f interest after re shutdown, second.
t 2
For this case, the time at which the uncoolable core debris bed was formed corresponds to the start of the time interval, while the time that the containment fails is the end of the time interval.
Substituting Equation
?
(3-10) into Equation (A-8) and rearranging yields the desired end time:
,1.35 i
- 7 ' 8 "cn (1 ~
S !( s en))!S f ! cn I3'11) t2~
q r o q where failure "en ition# "cn,d, whichever corresponds to the containment m
~
en cond fraction of core debris bed power assumed to enter concrete, f
q total erosion enthalpy, Btu /lba, including sensible heat to A
concrete melting, sensible heat to corium temperature, slag heat of fusion, and chemical reaction energy addition, total heat of reaction to oxidize zirconium, Btu /lba, see Q
Equation (A 4).
AG i
j
3-8 i
If containment failure will be predicted to occur after zirconium depletion, then the length of time required to fully oxidize the zirconium in the core debris must be determined, before the containment failure time can be estimated. This requires that the depth of erosion corresponding to j
m be determined, This is done using the following equations in an en,zr iterative fashion:
i
- d,zr ~"en,zr!(#cn d)
(3 12a) s (3-12b)
- d,zr
~#
s,zr s
where:
- depth of eroded concrete when zirconium is depleted, ft, x
d,zr x,
- sidevard erosion distance when zirconium is depleted, ft.
The factor f, is determined by:
O f,-1+r,(A,/Ad]
(3*13) where:
A, - area subjected to sideward erosion, ft
~ ""* #E' "" "* f ^ !^d ver time interval f interest.
A,/Ad s
The averaged values used in Equations (3-12a) and (3-13) are defined as follows for this time interval of interest:
f
'A A
l #+A (3-14) s.zr A /A s
2 A, do d, z r, (3 15)
(Ado + Ad,zr]
i d
If a rectangular geometry is used to approximate a cavity, for example, the initial values for sidevard and downward erosion areas, A,, and Ado'
- '"E*"'
I tively, are:
.-y
39 (3-16a)
A,,-(2L,+2W]h, where i
L,,W, - original cross sectional dimensions of the cavity, ft, h, - original depth of the debris on the cavity floor, ft, (3-16b)
A
~bo "o do The values of the sideward and downward erosion areas to be used in the second iteration cycle and all subsequent cycles may be expressed as:
s, z r] {h, + xd, zr)
(3'178)
~ ( o + 2W
+ 8x A
s.zr and d,zr ~ho+2xs,zr)ho+2xs.zr)
(3-17b)
A O
l Similar expressions would be determined for other uncoolable. debris bed configurations.
The steps in the iteration procedure involve estimating and. updating 5d-change. T start the as the estimated values for xd,zr ""
s,zr and A,/Ad
^do
" ^ !^d.
Next, calculate a iteration, use A f#
^
do d'
so s
value for f, using Equation (3-13), and then obtain first estimates of xd,zr and x fr m Equations (3-12a) and (3-12b), respectively The revised j
s,zr values for erosion distances then should be used to start the next iteration cycle.
Update values for A,/A ' 's'
""d using (in order) Equations (3-d d
17a), (3-17b), (3-14), (3-15), and (3 13). Then, calculate new values for using Equations (3-12a) and (3-12b).
Satisfactory conver-x
""d d,zr s,zr gence is obtained when the value of x changes less than 0.1% from the d,zt previous iteration.
Once convergence is reached for the erosion distances, the end time of the interval required to deplete unreacted zirconium in the core debris bed f
3-10 im 1
U-should be estimated using Equation (3 11), with m
identified as the en m 1
appropriate value for the factor a To then determine the containment failure time, the additional mass of concrete that must be eroded after zirconium is fully oxidized to reach the failurecondition(am must be found.
This value may be contaircent determined as follows:
(3*18) 0"cn ~ 0* #cn s d where.
ax - distance erosion front must advance to cause containment failure, m.
l The value for ax in Equation (3-18) should be selected as the minimum of ax,, and axe,d; (3-19a) ax
~*
e,d e,d ~ *d,zr (3-19b)
~ (*
AX e,s s,zr s
e,s where ax
- distance erosion front must advance in downward di ection to e,d reach containment failure criterion for downward direction, ft, Ax
- distance erosion front must advance in downward direction to e,s reach containment failure criterion for sideward direction, ft.
t be used in Equations (3-13) and (3-18) should The values of \\,/Ad ""d d
be as follows:
A,4 - f (Ad,zr*^dcf]
(
0)
C+
(3*21) j A,/Ad
~
,d,zr d,cf, l
1 O
where:
3-11
(%
nal area of une lable core debris bed at time of d'cf - cr ss-secti 2
A containment failure, ft
{
l sideward erosion area o uncoolable core debris bed at time of A"I = containment failure, ft As an example, the values for A and A f r c ntaiment sump geometry s cf d,cf would be given by:
s,cf 2L, + 2W +8(x
+##*)(o**d,zr+ax]
(3-22a)
A s,zr s
(*s,zr+#ax]
(3-22b)
(*s, zr + # 0*)
A
+
~
s d,cf o
s o
Once am has been determined, the predicted time of containment failure may en be found from Equation (3-11),
using the time at which zirconium was as m
,Q set equal to zero, and the appropriate value depleted as ty, am r
of A,,from Table 3-1.
p The result of the preceding calculation is an estimate of the least d
time required for molten core-concrete interaction to exceed a containment failure criterion.
If the criterion stated by Equation (3-1) is the most limiting condition, then containment failure is assumed to be at a penetra-tion as a result of vessel movement; fission products then would be released from the containment building to the environment.
If the criterion stated by Equation (3-2) is the most limiting condition, then contairment failure is assumed to be caused by basemat penetration; fission products and/or core debris would be released from the containment building to the soil under-neath the containment basemat.
O
41 4.0 FIANT-SPECIFIC APPLICATION 4.1 Issues 4.1.1 NCCI Failure Criterion (Sten 1)
Examination of Farley Nuclear Plant shows that containment failure at containment penetrations, caused by erosion of the cavity walls and _the embedded structural steel columns (Figure 4-1) that support the reactor vessel, will not occur as early as the meltthrough of the base mat.
Core debris of about 1 ft depth (see Equation (4-1) for this estimation) would accumulate on the cavity floor at the 80' elevation assuming that it is not swept out of the cavity.
At this depth, the debris will begin to erode sideways into the containment foundation and away from the center of the containment, as well as downward, but at a much faster rate, into the con-tainment foundation.
As the corium erodes the concrete in a sideward
~'\\
direction, it can travel a fair distance without compromising the ability of
[Y the cavity walls to support the vessel and its associated piping.
This is-because of the massive a ount of concrete in the cavity walls relative to a small sidevard erosion d.
_nce.
Failure of the vessel support at the 120' elevation, which is 40 ft higher than the debris pool, would require sidevard erosion of a large fraction of the concrete cavity vall thickness (the cavity is 5 ft thick along the cavity keyway but 9.8 ft thick below the reactor vessel).
The thickness of the cavity floor is 1.1 ft and the basemat thickness is 7.75 ft.
The downward erosion distance required for the containment failure is, therefore, 8.85 ft (2.7 m).
This distance is comparable to the sideward erosion distance required for containment failure.
It is not possible for the corium to travel such a distance sidewards without first penetrating through the entire basemat because the sideward rate of erosion would be much less than the downward rate of ero-sion, as demonstrated by the experimental evidence discussed above, basemat penetration is, therefore, the MCCI failure criterion for the Farley cavity configuration.
42 O
,,U, r.
sbbi WTUT GUTLET O
s
- g..
4 1 1
-tw+
j l l !UPPerf
~
latt!
n rr
(
j istfT d
I U' u IL OU TW ABCMM B4LT1 ynt opgnings tatAA LMS h
Lp!
04 TUT SUPPMT tatt o
b U
u-
- ~
~
$8PP987 1488 d
$UPfttT ime""
O O
$8PMIT PA4 oQQo REACTOR VESSEL SUPPORT Figure 4 1 Reactor vessel supports at Plant Farley (taken from Farley System Descriptions OPS-40301B Figure 7).
's
i 4-3 1
I a
/
r
/
/
O g
4 i;s e-: ;
' V (.,f.; y g. 3. a, I
L**.
,hJ
\\,
i:
4, s
I
/l' i
e 5
./\\
p
[(
)~l i
j
@-- y g
L ".* -
w
' j_ I 7
e A
1 w
.h 4
i t
1 l
/
.%q 7
N-
~.
n a. y j
a M.i ma
_u.
_wLw_
t f..
m m..
semos cury-m 4r at s r--s e m 3._ y,w f
a m mas L_.W.
h
- i
..t h w e i,". gg"
=
l Em C.J,'*' " '
s_
I,..
h. Y j,j aus
]# f
=
e E **= ~
3g 5-~
- I4. _ _. *' W
_._Ji a_
II **
- .a N
Y_.g. _ _
t' :
pVydl -
= m m.e L=v-w L
i
_m u.,
.7
, [.y
- =
'E WW a.*,"N,
j a
am cm A
c.,...
._h w-
.._ z, _. N u %~
y stenw m 't ID "'*. " ***
Y Figure 4-2 Farley reactor cavity compartment (adapted from Bechtel drawing D-176105 Rev. 6) 1 i
i
44 O
4.1.2 Tf== Interval to Exceed the Failure Criterion (St R_21 f
i The time to exceed the failure criterion is determined by the procedure outlined above, subject to a key assumption the entire core, with its full initial inventory of zirconium and fission products, is expelled instan-taneously at the start of the time interval.
Vessel failure and core-concrete attack can be delayed for many hours after reactor shutdown if the operator can remove decay heat via the secondary side heat sinks or inj ec -
tion with either charging pumps or low head safety injection pumps. By neglecting potential mitigating operator actions, the decay heat of the debris and the heat transfer to the concrete are overstated. The expulsion of molten debris from the failed vessel would take some finite. time period that depends upon vessel pressure at the time of failure. However, for a PVR, this time period is small when compared to the duration for core-concrete attack (many hours, or even days), so it can be thought of as instantaneous.
Although some fraction of the core could remain cooled in-vessel while the bulk of the core is expelled, it is convenient'to make the conservative assumption that the entire core is expelled with its full initial inventory of fission products and zirconium. In many accident sequences, a substantial fraction of the zirconium can be oxidized in-vessel, as much as, say, 504, as opposed to being exidized during corium-concrete attack. This potentially decreases the duration of the zirconium oxidation phase and slows corium.
concrete attack overall, since the chemical energy of the corium is
]
decreased.
This possibility is not considered here. Moreover, a substantial fraction of the the core's initial fission product inventory will not reside in the corium attacking the basemat.
Fission products are distributed j
throughout the primary system and containment compartments in a manner that depends upon the severe accident progression. Volatile fission products initially present in the corium can be vaporized or entrained to form 1
aerosols which are transported throughout the containment. The net effect of j
these mechanisms is to reduce the mass of fission products and decay heat in the corium as it attacks the cavity floor.
O i
4-5
,. m
\\
\\
/
A small computer code was developed to expedite the repetitive calcula.
v tions of the procedure and allow for further sensitivity calculations.
The data required were taken from Table 3-1 and the Farley MAAP parameter file (Fauske 6 Associates, Inc.,
1991).
(Code input data and results are presented in Appendix B, and the MCCI source code listing is included as Appendix C.)
For a large, dry PVR, the initial condensed debris depth in the cavity is given by:
+
+
+ 0.10 (4-1)
,u ZR SS,
SS h, -
C
- where, 1
M is the mass of UO in the debris bed (181,200 lbm or 82,200 kg),
2 M
is the mass of zirconium in the debris bed (38,230 lbm or 17,340 g
kg),
M is the mass of the lower core plate (5240 lba or 2380 kg),
CSP V
M is the mass of the lower head (58,000 lba or 26,300 kg),
g p
is the density of UO2 (10,100 kg/m )*
u 3
p is the density of zirconium (6570 kg/m ),
ZR p
is the density of stainless steel (8238 kg/m ), and, g3 2
2 A
is the cavity floor area (472.7 ft or 43.9 m ),
c Only 10% of the lower head mass is used because the size of the vessel failure will be small in relation to the radius of the lower head.
In
- addition, this calculation uses the assumption that none of the debris has been swept out of the cavity and dispersed into the lower cor'artment. This overstates MCCI in the cavity and it is clearly conservative for the case j
where the cevity is initially dry.
The cavity floor is modeled as a rectangle to avoid the necessity of dealing with the complexities of the actual geometry.
This is achieved by maintaining the perimeter and the area of the modeled rectangle the same as j
the perimeter and the area of the cavity floor, respectively.
Given the cavity floor area (A ) and the cavity floor perimeter (P ), the dimensions c
l 4-6 i
s
)
of the equivalent rectangle can be found from the fellowing two require-ments:
LW
-A (4-2) l oo e
and P
L
+W E
(4 3) o o
2 i
Solving the above two equations gives 0.5 r
16 A P +
P e
e c
(4,4)
L o
4 and 2
16 A p
p e
e c
(4 5)
W 4
For Plant Farley, applying A - 472.7 ft2 (43,9,2) and P - 106.8 ft c
c (32.5 m) yields L - 42.2 ft (12.9 m) and W - 11.2 ft (3.40 m).
i o
o l
Othar parameters required for the procedure, such as the ratio of sideward rate of erosion to downward rate of erosion (r,) and the fraction of core debris power to concrete (f ), are derived from the experimental q
evidence cited above. The choice of values for these parameters is discussed below.
-i i
A ratio of the sideward rate of erosion to the downward rate of erosion must be specified based on the BETA experiments and engineering judgment.
Clearly, the lower bound for r, of zero will result in the maximum downward erosion rate.
- Also, the downward erosien area will remain constant during the attack for an r, of zero, resulting in very simple expressions for ancn' i
If r,
is assumed to be zero, an extremely conservative containment failure
(}
time can be quickly estimated by hand.
4-7 e
small values of r, greatly increase containment failure
- However, even time relative to the r, - O assumption because f, is made larger than 1. The assumption that r,
- 0 is quite convenient, but it results in unrealisti-cally conservative answers.
Experimental evidence suggests that the ratio
~
of erosion rates is non zero, but much less than one.
From Table 2-4, the BETA experiments indicate an r, of 0.05 for the V1.8 test and 0.29 for the l
V2.1 test.
The V2.1 test was scaled to reflect long-term decay heat, in contrast to the V1.8 test which reflects a heat generation rate an order of magnitude larger than d cay heat. As a result, 0.2 will be used as a value for r,.
The fraction of core debris bed power assumed to enter the concrete (f ) must also be specified. A value of 1 is overly conservative because it q
neglects radiation and/or convection heat transfer from the the top surface of the debris, while a value of 0, on the other extreme, denotes no concrete attack.
It is assumed f is zero, i.e., MCCI is not a relevant containment failure mechanism, if a deep overlying pool of water can be maintained on the debris.
- [his assumption is based on experimental results demonstrating the ability of water to rapidly quench molten debris and ingress into debris beds to maintain them coolable indefinitely. If the pool cannot be main-tained, MCCI will begin after the pool has dried out.
t Note that this assumption reduces the complexities of innumerable severe accident sequences to just one consideration - is the debris covered by a deep overlying pool for an indefinite duration? If not, f - 0.5, based q
on the methodology of [Fauske & Associates, 1985).
The following section considers which accident sequences will not have an overlying pool. The value of f would actually vary from sequence to sequence as a function of accident progression and operator actions. Similarly, for a given severe accident sequence, f would be a function of time due to operator actions, j
9 accident progression, fission product transport, etc. For now, we focus on
.j the simple " dry" case where f has a constant value of 0.5 after debris bed q
dry-out.
]
to calculate Finally, an initial time for MCCI (ty) must be specified the time to exceed the failure criterion.
Even in those cases where a water
48
(
pool cannot be maintained (due to equipment failure, operator error, etc.),
MCCI will not-begin until the debris bed has dried out.
The Farley contain-ment design prevents any water from spilling into the cavity from the lower compartment, and dry out will occur after the remaining water in the lower head, expelled at reactor vessel failure, has boiled away.
For Plant
- Farley, the cavity dry out time has been calculated for a station blackout sequence to occur at about 6.0 hrs after scram, based on typical MAAP 1 - 6.0 hours0 days <br />0 hours <br />0 weeks <br />0 months <br />.
analysis for a TM1.5 with seal LOCA.
Therefore, t Substituting these values into the code developed specifically to i
evaluate the expressions presented in Section 3.0, and using a nominal full 9
power for Farley of 9.05 x 10 Beu/h (2650.W),
results in a containment j
failure time of about 62 hours7.175926e-4 days <br />0.0172 hours <br />1.025132e-4 weeks <br />2.3591e-5 months <br /> after scram.
It is important to stress that this value represents a conservative lower bound for the time to fail the l
Farley containment as a result of molten core-concrete interaction, P
For the dry cavity case, a station blackout is considered because ECCS and containment heat removal systems will be inoperative. In an SBO, MCCI will begin s'.tortly after reactor vessel failure, but the containment will also pressurize as the containment atmosphere is directly heated by the debris or As steam is generated by any debris which has been entrained to the lower compartment.
Containment pressurization by steaming in the lower
.i 117 compartment, as calculated by MAAP, indicates containment failure (P psia) about 30 ' hours af ter the initiating event - long before the time to reach the MCCI containment failure criterion at 62 hours7.175926e-4 days <br />0.0172 hours <br />1.025132e-4 weeks <br />2.3591e-5 months <br />.
If all the core debris is retained in the cavity, typical MAAP predictions indicate contain.
ment failure due to non-condensible gas generation and slow containment l
heating at about 60 hours6.944444e-4 days <br />0.0167 hours <br />9.920635e-5 weeks <br />2.283e-5 months <br /> after the initiating event-.
However, even if containment over-pressurization is hypothetically ruled out, the station
[
blackout would have to last several days for failure by MCCI.
Otha nise, l
recovery actions could be taken to halt the ablation by covering the debtis with subcooled water.
The wet cavity case is any sequence where ECCS injection and contain-ment heat removal are available to keep the debris bed covered. Iow pressure recirculation with RHR heat exchangers operating could be used to cover the l
49 p(
debris via the vessel hole, and remove decay heat from the containment. If low pressure recirculation were available, however, core damage would not have occurred in the first place, so this means of arresting MCCI is con-sidered only as a possible recovery action.
In addition to pressure suppression functions, containment fan coolers will condense steam and thereby maintain a supply of water to cover the debris orovided that water can find its way from the lower comoartment to the cavity.
Containment failure due to MCCI is precluded for the wet case because of the ability of water to quench molten debris and insress into debris beds to render them coolable.
If a pessimistic view of ingression is taken for the wet case (: hat is, ingression cannot occur) calculations could be per.
formed, just as teey were performed for the dry case, by selecting an appropriate power fraction into the concrete (f ). However, these calcula-tions would show failure times of hundreds of hours, and we would therefore revert back to the same conclusions as the dry case.
p 4.1.3 Uncertainty Considerations The results presented above are subject to uncertainties.
In par-ticular, the fraction of core decay heat power (f )
is subject to large q
uncertainty as a result of two considerations:
(1) the heat transfer mechanisms from the debris by radiation and convection to overlying water or
- gas, and, (2) the transport of volatile fission products from the debris by vaporization and/or entrainment to other parts of the containment. The decay heat power fraction to the concrete was taken from [Fauske & Associates, 1985), but no effort was made to account for fission product release from the debris. Fission product release is now accounted for in a simple manner to scope out its potential impact and demonstrate the conservatisms of the aforementioned results.
This phenomenon essentially decreases the fraction of decay heat power to the concrete, so fission product release in the cavity can be addressed by varying the parameter f.
Fission product release and transport is dependent upon severe accident progression, timing, operator actions, etc.
A few KAAP runs can be examined to find a range for the fraction of decay heat energy removed from the I
T
4 10
)
OV debris by transport of volatile fission products.
This will improve the accuracy of f. The MAAP User's Manual (Fauske & Associates, 1990) contains q
input decks and tabular output for three runs made with the " Zion like" parameter file: a station blackout with no operator actions, a small LOCA without the recirculation mode, and a large LDCA without the recirculation mode. Tabular output for each of these runs gives the fraction of core decay heat in the cavity at various times. For the blackout. case, 39% of the decay heat is in the cavity at 40 hours4.62963e-4 days <br />0.0111 hours <br />6.613757e-5 weeks <br />1.522e-5 months <br /> into the accident, 32% is in the lower compartment, and 294 is in the form of released fission products. For the large LOCA, 72% of the decay heat is in the cavity at 30 hours3.472222e-4 days <br />0.00833 hours <br />4.960317e-5 weeks <br />1.1415e-5 months <br /> into the accident and 28% is in the form of released fission products. And finally, for the small LOCA, 71% of the decay heat is in the cavity at 44 hours5.092593e-4 days <br />0.0122 hours <br />7.275132e-5 weeks <br />1.6742e-5 months <br /> and 29%
is in the form of released fission products, Based upon these examples, it seems reasonable to assume that in the long term, 70%, rather than 100n, of the decay heat is in the cavity.
Calculations have been made assuming that the cavity floor could be O-opposed to the actual geometry which closely modeled as a rectangle, as resembles a rectangle placed next to a circle. The actual geometry is not far removed from a rectangle,
- and, perhaps, the initial cavity floor geometry may be " forgotten" as the melt attacks the concrete. As the debris attacks the concrete and spreads sideways at a uniform race, the debris geometry resembles the initial geometry less and less.
This leads to the supposition that as long as the initial floor area and the perimeter are accurate, the actual geometry of the cavity floor may not be significant.
Additional uncertainty considered here is the impact of the assumed constant erosion rate ratio (r,).
To simplify the calculation, r,
was assumed constant.
Furthermore, the average value of r, which was taken from experimental data ranges from 0.05 (in V1.8 test) to 0.29 (in V2.1 test) depending on the heat supply rate in the experiments. This differs by a factor of 6.
It is, therefore, important to take into accour -
the sen-sitivity of the containment failure time to r,.
In general, the smaller value of r, means more heat rate will be available to erode downward, the-reby reducing the failure time in the downward direction.
When r,- 0.1 is assumed in the calculation, the containment failure time for Farley i
4 11 i
l decreases to 52 hr provided that all other parameters remain the same.
The Farley containment failure time, predicted by the stand alone model,
's i
expected to be more conse rvative (i.e.,
earlier) than the failure time predicted by MAAP.
4.2 Conclusions The analysis presented here shows that MCCI can be excluded from con-sideration as a significant late containment failure mechanism.
This is not meant.to downplay the significance of MCCI in the Farley - Level 2 PRA, or suggest that containment failure due to concrete melt-through cannot occur under any circumstances.
Rather, it is clear that relative: to other con-tainment failure mechanisms, MCCI will occur so late in time that: (1) the more rapidly developing-o t,he r,
containment will have failed due to mechanisms, or (2) mitigating actions will almost surely have taken place to arrest MCCI before reaching the containment failure criterion of basemat melt-through.
- Also, relative to other failure mechanisms, the source term for a basemat melt through would be small because of the failure time (very
'~'
late) and location (below ground).
1 4
i 51 5.0 SIRetARY This discussion has focused on two simple accident scenarios, i.e., dry j
and wet cavities, to simplify consideration of the innumerable accident that mi ht actually involve MCCI. If the debris cannot be covered
.{
E sequences by. water'for an indefinite period, the sequence is considered-to be dry, j
Other accident sequences involving a more complex set.of~ operator'interven-j tions that-do not fall into either of the two simple classes defined here require individual analysis, such as MAAP runs, to determine exact contain-l ment failure mode and timing, Such analysis, conducted as part of an IPE sensitivity study, would not contradict the basic conclusion that MCCI can i
be neglected relative to other containment failure mechanisms for the Farley l
IPE.
4 i
i t
b I
i 4
i i
i 1
5 0
't t
i I
I 4
f i
h
r 1
I 6-1
,D
6.0 REFERENCES
- Alsmeyer, H.,
- 1986, Beta Exoeriments for Verifyinn the WESCHL-Codes:
Exnerimental Results for Melt-Concrete Interaction, English i
Translation.
Alsmeyer, H, et al., 1987,
" Beta Experimental Results on Melt / Concrete Inte rac tior ;:
Silicate Concrete Behavior",
Proceedines of the Committee on the Safety of Nuclear Installations (CSNI)
Soecialists' Meetina en Core Debris-Concrete Interactions, EPRI NP-5054-SR.
Blose, R.
E.,
et al., 1987, SWISS:
Sustained Heated Metallic Melt / Concrete Interactions with Overivine Water Pools, NUREG/CR 4727, SAND 85 1546.
Bradley, D.
R., and Copus, E.
R., 1987, "Significant Results from SURC 3 and SURC-3A Experiments",
Present?d at 15th Vater Reactor Safety Meetine.
National Bureau of Standards. Gaithersburg.
CSNI, 1989, "SURC-4 Experiment on Core-Concrete Interactions," Committee on the Safety of Nuclear Installations CSNI Report No. 155, Volume 2.
E1.Vakil, M.
M., 1978, Nuclear Heat Transoort, The American Nuclear Society, LaGrange Park, Illinois.
)
EPRI, 1990, "MAAP 3.0B Complete Code Manual", EPRI NP-7771-CCML.
Fauske & Associates, Inc., 1985, " Approximate Source Term Methodology for Pressurized Vater Reactors", FAI/85-58.
Fauske & Associates, Inc.,
- 1990, "Results of 3.08 Design Review Group Comments", FAI/90-22.
Fauske & Associates. Inc., 1991, " Containment Data Collection Notebook".
- Gieseke, J.
A.,
et al.,
1986, " Source tern Code Package, A User's Guide i
(Mod. 1)", Battelle Columbus Division, NUREG/CR 4587, BMI-2138.
- IDCOR, 1985, Grand Gulf Nuclear Station - Interrated Containment Analysis",
IDCOR Technical Report 23.1GG, Mississippi Power and Light ca.
4
- Lee, M.,
and Kazimi, M. S.,1985, Modeline of Corium/ Concrete Interaction, MITNE-267.
Muir, J.
C.,
et al.,
- 1981, "CORCON/ Modi:
An Improved Model for Molten Core / Concrete Interaction", NUREG/CR-2142, SAND 80-2415.
NRC, 1975 WASH-1400, Reactor Safety Study, NUREC/75-0114.
NRC, 1988, letter to All Licensees Holding Operating Licenses and Construction Permits for Nuclear Power Facilities, " Individual Plant 10 CFR 50.54(f)",
Examination for Severe Accident Vulnerabilities O
Generic Letter No. 88-20.
6-2 l
v' Plys, M.
C.,
1987, " Hand Calculations for Core-Concrete Attack and contain--
ment Failure Timing With Applications to lasalle and Grand Gulf",
-Presentation to NUREG-1150 Molten Core-Containment Expert Review Group, Albuquerque, NM.
Reimann, M., and Murfin, W. B., 1981, "The WECHSL Code: A Computer Program for the. Interaction of Core Melt with Concrete". KfK-2890.
Summers, R.
M.,
et al., 1991, "MELCOR 1.8.0:
A Computer Code for Nuclear Reactor Severe Accident Source Ters and Risk Assessment Analyses",
Sandia National laboratories. Albuquerque, NM.
i s
i I
I O
i A1 6
/N V
AFFEND H A' Derivation of Concrete Erosion Calculation i
P The calculation presented below is based on the method presented by
[Plys, 1987). This calculation is based on the following assumptions:
. The mass erosion rate is linearly proportional to the decay heat f
rate plus heat of chemical reactions.
. Only reaction heats from the downward attack are accounted.
. Sideward erosion speed is assumed a constant fraction of downward erosion speed (i.e., r, - U,/Ud
.j
- e nstant).
The energy balance for the eroded concrete according to the first assumption may be written as:
O cn(U^d+U^)~0
(^'1) p A
d ss en where 3
p
- concrete density, Iba/ft A
-t tal erosion enthalpy, Btu /lba, including sensible heat to cn concrete melting, sensible heat to corium temperature, slag heat of fusion, and decomposition reactions.
U - erosion speed, ft/s, A = area, ft h
6 Q - input power, w, including decay power and chemical reactions, d - downward, i
s - sideward, i
The mass balance for the eroded concrete yields:
1 O
t
A2 I
i kJ' t,
ss)d'
(^~
(U^d+
A p
m d
o E
r
(^~
(Q/Acn) o where m
-t tal er ded concrete mass, Iba.
en The input power in Equation (A-1) may be written as:
HO CO 2
2 O
(^~
)
A OH0+ 44
- CO2, 9~f Odk * #cn dd 18 q
2 where f - power fraction into concrete, p
q O
Q
- decay power, Btu /h, dk
- gas mass fraction in concrete, fH 0'IC0 2
2 QH0'kO - heat of reaction per mole gas to liberate H and CO, 2
2 2
Btu /lb mole.
Typical values for the latter quantities are:
Qg 0 = 1.2898 x 10 Btu /lb-mole H O 2
2 5
Q
- 1.1608 x 10 Btu /lb mole CO CO 2
2 for overall reactions to H and CO.
2 If we define:
('T V
i r
-A-3 i
i O
,a H0+a e
O 9
(**'}
Q
~
r 18 44 C0 2
2, and t
UA A
f, - 1 +
-1+r (A 5) sA Then, Equation (A-2a) may be rewritten as:
i 5
i t
(A-6a) d^d men ~ #cn s,
f o
(A 6b) d *d
- #cn s where:
- depth of eroded concrete, ft, x
d 2
i - average surface area of downward concrete erosion, ft d
Equation (A 2b) may be rewritten as:
i 9
d *d (A*7) cn"cn ~I O
A q,
dk q #cn r
o I
Manipulating Equations (A-6b) and (A-7) then yields:
j I
t
( q! cn) 9 de dk
(^
}
"en ~
l-fQ (f,1cn)
- d ~"en!(#cn's d)
(A 9) j (A-10) x, - r, xd O
B-1 O
APPENDIX B Code Input and C.siculations O
O
82 j
1
)
I FARLEY DRY CAVITY CASE (SI UNITS) i 0.5
/ F1 MINIMUM FRACTION OF CAVITY WALL P.IQUIRED FOR VESSEL SUPPORT 1.52-
/ XP WIDTH (THICKNESS) 0F CAVITY WALL (M) 0.0
/ XSP (USED ONLY IN BWR; SET TO 0.0 FOR PWR CAVITY) 0.34
~/ XF THICKNESS OF CAVITY FLOOR (M) 2.36
/ XB BASENAT (M) 17340. / MZR MASS OF UNOXIDIZED ZIRCONIUM (1004-17340 kg) 0.0339 / FH2O CONCRETE CAS MASS MtACTION FOR STEAM
]
0.0067 / FCO2 CO2 12.9
/ LO DEBRIS LENGTH (M) 3.40
/ WO DEBRIS WIDTH-(M) 0.25
/ H0 DEBRIS HEIGHT (M) 0.2
/ RS RATIO OF EROSION RATES l
21600. / TINIT INITIAL TIME (S) (6 HRS) 0.5
/M DECAY HEAT FRACTION INTO CONCRETE j
2.652E9 / QO FULL POWER (W)
[
3.41E6 / LCN DECOMP. ENTHALPY DURING ZR OXIDATION (J/KG) (PLYS, 1987)
AFTER 3.11E6 / LCNP
=
O
-)
1 0
B.3
.y FARLEY DRY CAVITY CASE (SI UNITS)
MCCI POSITION PAPER CALCUIATIONS DECAY HEAT FRACTION IS.
0.50000 EROSION RATIO IS 0.20000 SIDEVARD FAILURE CRITERION IS 0.76000 METERS DOWNWARD FAILURE CRITERION IS 2.7000 METERS OVERALL FAILURE CRITERION IS 2.7000 METERS 190.09 NO. OF INITIAL MOLES OF ZIRCONIUM 380.18 MOLES OF CAS PRODUCED 0.20356E 02 NO. OF CAS MOLES PER MASS OF CONCRETE MASS OF CONCRETE ERODED DURING ZR OXIDATION -
0.18676E+06 8.1500 INITIAL SIDEVARD EROSION AREA 43.860 INITIAL DOWNWARD EROSION AREA FINAL SIDEVARD EROSION AREA - AT FAILURE -
414 77 FINAL DOWNWARD EROSION AREA - AT FAILURE -
161.04 102.45 AVERAGE DOWNWARD EROSION AREA 1.3807 RATIO OF EROSION AREAS AT FAILURE 1.2761 TOTAL TO SIDEVARD EROSION RATE RATIO CONTAINMENT FAILS AFIER CONCRETE FULLY OXIDIZED MASS OF CONCRETE ERODED AT FAILURE CRITERIA
.81190E+06 KC MASS OF CONCRETE ERODED BY ZIRCONIUM OXIDATION
.18676E+06 KC i
0.96799E-01 e CONVERCENCE ERROR IN XDZR FACTOR FOR EROSION AREA BY ZIRC. OXID. -
1.1303 53.790 DOWNWARD EROSION AREA FOR ZIRC. OXID.
60.115 SIDEWARD EROSION AREA FOR ZIRC. OXID.
1.4699 THICKNESS ERODED BY ZIRC. OXID.
30.056 HOURS AFTER SCRAM TIME FOR ZIRCONIUM OXIDATION i
CONTAINMENT FAILS AT 62.222 HOURS AFTER SCRAM O
l i
)
4 C1 i
i O
APPENDIX C
}
MCCI Calenlation Source Code O
l O
C2 l
O C MCCI POSITION PAPER CALCULATIONS IMPLICIT REAL (A-H,K-Z)
CHARACTER *80 TITL C
C XES
- SIDEVARD FAILURE CRITERIA C
' F1
- FRACTION OF WALL REQUIRED FOR VESSEL SUPPORT l
C XP
- SHORTEST-DISTANCE BE7VEEN DEBRIS AND VESSEL SUPPORT WALL i
C XSP
- VESSEL SUPPORT WALL THICKNESS C
XED
- DOWNWARD FAILURE CRITERIA C
XF
- DRYWELL FIDOR THICKNESS C
XB
- THICKNESS OF THE BASEMAT C
NZR
- NUMBER OF MOLES OF ZIRCONIUM C
MZR
- MASS OF UNOXIDIZED ZIRCONIUM IN THE CORE DEBRIS C
AZR
- MOLECUIAR WEIGHT OF ZIRCONIUM C
NG
- NUMBER OF MOLES OF GAS PRODUCED FOR EACH MOLE OF ZIRCONIUM C
NGC
- NUMBER OF CAS MOLES PRODUCED PER UNIT MASS OF CONCRETE ERODED C
FH2O
- CONCRETE CAS GSS FRACTION FOR STEAM C
FCO2
- CONCRETE GAS MASS FRACTION FOR CARBON DIOXIDE C
AH20, ACO2 - MOL. WEIGHTS FOR STEAM AND CARBON DIOXIDE 1
C MCNS
- MASS OF CONCRETE ERODED SIDEVAYS FOR FAILURE CRITERIA C
MCND
- MASS OF CONCRETE ERODED DOVNVARD FOR FAILURE CRITERIA C
MCNZR
- TOTAL CONCRETE ERODED DURING PHASE 1 - ZIRC. OXIDATION C
RHOCN
- DENSITY OF CONCRETE C
FS
- FACTOR FOR SIDEWARD EROSION RELATIVE TO DOWNWARD EROSION C
- AVERAGE AREA SUBJECTED DOWNWARD EROSION C
RS
- RATIO OF SIDEWARD RATE OF EROSION TO DOWNWARD RATE OF EROSION O
C TFINAL - END OF TIME INTERVAL OF INTEREST C
TINIT
- BEGINNING OF TIME INTERVAL OF INTEREST C
FQ
- R ACTION OF CORE DEBRIS BED POWER ASSUMED TO ENTER CONCRETE C
QR
- TOTAL HEAT OF REACTION TO OXIDIZE ZIRCONIUM C
14N
- TOTAL EROSION ENTHALPY C
QO
- INITIAL CORE POWER C
14,WO HO - INITIAL BED LENGTH, WIDTH, HEIGHT C
XDZR
- DEPTH OF ERODED CONCRETE WHEN ZIRCONIUM IS DEPLETED C
XSZR
- SIDEWARD EROSION DISTANCE WHEN ZIRCONIUM IS DEPLETED 1
C ADZR
- DOWNWARD EROSION AREA C
ASZR
- SIDEWARD EROSION AREA C
ADO
- ORIGINAL DOWNWARD BED AREA C
- ORIGINAL SIDEVARD BED AREA 1
C AD
- AVERACE DOWNWARD EROSION AREA C
ASAD
- AVERAGE AS/AD DURING THIS TIME INTERVAL READ (21,1000) TITL 1000 FORMAT (A80)
READ (21,*) F1 READ (21,*) XP READ (21,*) XSP READ (21,*) XF READ (21,*) XB READ (21,*) MZR READ (21,*) FH2O READ (21,*) FCO2 READ (21,*) 14 READ (21,*) WO Oi READ (21,*) H0
C-3 y
READ (21,*) RS READ (21,*) TINIT READ (21,*) M READ (21,*) QO READ (21,*) LCN READ (21,*) LCNP DATA AZR,AH20,ACO2,RHOCN,QH20,QCO2,XMAX,IMAX w /91.22,18.0,44.0,2300.0,
- 3.0E8,2.7E8,10.0,100000/
C kTITE OUT EROSION RATE AND DECAY HEAT FRACTION INTO CONCRETE TO KEEP TRACK C OF VARIOUS RUNS WRITE (22,49) TITL 49 FORMAT (' ' A80)
WRITE (22,50) M.RS 50 FORMAT (' MCCI POSITION PAPER CALCUIATIONS ',/,
DECAY HEAT } TACTION IS ',G13. 5, ' EROSION RATIO IS ',G13. 5,/)
C SIDEWARD FAILURE CRITERIA XES-(1.0-F1)*XP+XSP VRITE(22,51) XES 51 FORMAT ('
SIDEWARD FAILURE CRITERION IS ',G13.5,' METERS')
C BASEMAT PENETRATION FAILURE CRITERIA XED-XF+XB VRITE(22,52) XED 52 FORMAT (* DOWNVARD FAILURE CRITERION IS ',G13.5, ' METERS')
C FAILURE CRITERIA p
XFAIL-MIN (XES/RS,XED)
IF (XFAIL.EQ.XES*RS)XFAIL-XES VRITE(22,53) XFAIL 53 FORMAT (' OVERALL FAILURE CRITERION IS ',G13.5,' METERS',//)
C ZIRCONIUM DEPLETION OR CONTAINMENT FAILURE 7 C FIND NUMBER OF MOLES OF ZIRCONIUM NZR-MZR/AZR' WRITE (22,61) NZR 61 FORMAT (' NO. OF INITIAL MOLES OF ZIRCONIUM
' G13.5)
C MASS OF CONCRETE ERODED DURING ZIRCONIUM DEPLETION PHASE (MCNZR)
NG-2.0*NZR WRITE (22,62) NG 62 FORMAT (* MOLES OF GAS PRODUCED
' G13.5)
C NUMBER OF GAS MOLES PRODUCED PER UNIT MASS OF CONCRETE ERODED NGC-n120/AH20+FCO2/ACO2 WRITE (22,63) NGC 63 FORMAT (' NO. OF GAS MOLES PER MASS OF CONCRETE
' C13.5)
MCNZR-NG/NGC WRITE (22,64) MCNZR
+
64 FORMAT (' MASS OF CONCRETE ERODED DURING ZR OXIDATION
',G13.5)
C FIND FS AND AD i
AS0-(2.0*LO+2.0*WO)*HO WRITE (22,65) ASO 65 FORMAT (' INITIAL SIDEVARD EROSION AREA
' C13.5)
ADO-LO*WO VRITE(22,66) ADO 66 FORMAT (' INITIAL DOWNWARD EROSION AREA
',G13.5)
IF (XED.LT.XES) ASF-(2.0*LO+2.0*WO&8.0*(RS*XFAIL))*(H0+XFAIL) g IF (XES.LT.XED) ASF-(2.0*LO+2.0*WO+8.0*(XFAIL/RS))*(H0+XFAIL)
C-4
.h O
VRITE(22,67) ASF 67 FORMAT (' FINAL SIDEVARD EROSION AREA - AT FAILURE
',C13.5)
IF (XED.LT.XES) ADF-(LO+2.0*(RS*XFAIL))*(WO+2.0*(RS*XFAIL))-
IF (XES.LT.XED) ADF-(LO+2.0*XFAIL)*(WO+2.0*XFAIL)
WRITE (22,68) ADF 68 FORMAT ('
FINAL DOWNVARD EROSION AREA - AT FAILURE
',G13.5)
AD-0.5*(AD0+ADF)
WRITE (22,69) AD 69 FORMAT (' AVERACE DOWNVARD EROSION AREA
',013.5)
ASAD-0.5*(ASO/AD0+ASF/ADF)
WRITE (22,70) ASAD
' G13.5) 70 FORMAT (' RATIO OF EROSION AREAS AT FAILURE FS-1.0+RS*ASAD WRITE (22,71) FS 71 FORMAT (' TOTAL TO SIDEVARD EROSION RATE RATIO
' G13.5,/)
C FIND MASS OF CONCRETE THAT WOULD BE ERODED AT THE TIME EITHER FAILURE C CRIT.:.RIA ARE SATISFIED IF (RS.GT.O.0) MCNS-RHOCN*FS*AD*XES/RS MCND-RHOCN*FS*AD*XED C CONTAINMENT FAILURE BEFORE ZIRCONIUM FULLY OXIDIZED?
MCNMIN-MIN (MCNS.MCND)
IF (MCNMIN.LT.MCNZR) THEN C CONTAINMENT FAIL OCCURS BEFORE THE ZIRCONIUM FULLY OXIDIZED QR-FH20*QH20/AH20+FCO2*QCO2/ACO2 C SEE EQUATIONS 10 AND 11 0F POSITION PAPER c23456789c23456789c23456789c23456789c23456789c23456789c23456789c23456789 O
TFINAL-(TINIT**0.74+7.8*McNMIN*(1-ly*QR/(FS*LCN))
m /(QO*FQ/LCN))**1.35 WRITE (22.13) TFINAL/3600.0 13 FORMAT (' CONTAINMENT FAILS BEFORE CONCRETE FULLY OXIDIZED',/,
'AT TIME
',G10.5,' HOURS',/)
WRITE (22,14) MCNMIN,MCNZR 14 FORMAT (' MASS OF CONCRETE ERODED AT FAILURE CRITERIA
,G10.5,' KC',/,
' MASS OF CONCRETE ERODED BY ZIRCONIUM OXIDATION.
l
,G10.5,' KC',/)
l GO TO 99 ELSE WRITE (22,15) 15 FORMAT (' CONTAINMENT FAILS AFIER CONCRETE FULLY OXIDIZED',/)
WRITE (22,1<.) McNMIN,MCNZR CONTINUE ENDIF C CONTAINMENT WILL FAIL AFTER ZIRCONIUM DEPLETION C FIND DEPTH OF EROSION CORRESPONDING TO MCNZR C INITIALIZATION MCNZR-2.0*MZR/AZR/(FH20/AH20+FCO2/ACO2) 1 C
C DO Ih0P OVER XSZR,XDZR DO 10 I-1,IMAX XDZR-XDZR+(XMAX/IMAX)
XSZR-RS*XDZR ADZR-(Ih+ 2. 0 *XS ZR) * (W0+ 2. 0*XS ZR)
O ASZR-(2.0*LO+2.0*WO+8.0*XSZR)*(H0+XDZR)
\\
C-5 I
AD-0.5*(AD0+ADZR)
ASAD-0.5*(ASO/AD0+ASZR/ADZR)
FS-1.0+RS*ASAD ERROR-ABS (XDZR-MCNZR/RHOCN/FS/AD)/XDZR IF (ERROR.LT.O.001) CO TO 20 10 CONTINUE 20.
CONTINUE i
VRITE(22,75) ERROR *100.0
' C13.5, 75 FORMAT (' CONVERGENCE ERROR IN XDZR s*)
VRITE(22,72) FS 72 FORMAT (' FACTOR FOR EROSION AREA BY ZIRC. OXID.
',G13.5) i VRITE(22,73) ADZR 73 FORMAT (' DOWNVARD EROSION AREA FOR ZIRC. OXID.
' G13.5)
VRITE(22,74) ASZR 74 FORMAT (' SIDEVARD EROSION AREA FOR ZIRC. OXID.
',G13.5)
VRITE (22,54) XDZR'
',G13.5) 54 FORMAT (' THICKNESS ERODED BY ZTD.C. OXID.
C234567890C234567890C234567890C2345678' 234567890C234567890C234567890C234567890 T0XID-(TINIT**0.74+7.8*MCNZR
/Q*QR/(FS*LCN))
/(Q0* N/LCN))**1.35 VRITE'(22,56) T0XID/3600,
'.G13.5, 56 FORMAT (' TIME FOR ZIRCONIUM OXIDATION
' HOURS AFTER SCRAM'./)
C DETERMINE CONTAINMENT FAILURE TIME XFAIL-MIN (XED XEZR,(XES-XSZR)/RS)
O AS C F-( 2. 0* LO+ 2. 0 *V0+ 8. 0* (XS ZR+RS *XFAIL) ) * (H0+XDZR +XFAIL)
ADCF-(LO+2.0*(XSZR+RS*XFAIL))*(V0+2.0*(XSZR+RS*XFAIL))
AD-0.5*(ADZR+ADCF)
ASAD-0.5*(ASZR/ADZR+ASCF/ADCF)
FS-1.0+RS*ASAD MCN-XFAIL*RHOCN*FS*AD QR-0.0 C BE SURE TO USE CORRECT VALUE OF IAMBDA - EROSION ENERGY C234567890C234567890C234567890C234567890C234567890C234567890C234567890C234567890 TFINAL-(T0XID**0.74+7.8*MCN*(1-FQ*QR/(FS*LCNP))
/(Q0*FQ/LCNP))**1.35 VRITE (22,16) TFINAL/3600.
16 FORMAT (' CONTAINMENT FAILS AT ',G10.5,' HOURS AFTER SCRAM',/)
99 CONTINUE STOP END O
FAUSKE at ASSOCIA113, D0C.
CALCULATION NOTE COVER SHIET f'
t SECTION TO BE COMPLETED BY AUTHOR (5):
\\
Pase E~Af[7I
- 68 Re <ision Number
/
Cale-Nota Number _
Title PHEAlcMGNe LaQ tCA L EVA LU AnoAl SAM Malt Y' OM THE Pac en e it. 7 y' sata (couSEQUENCES of DEFLA GR AncM AMD PETc N AT7C As aP HVDK0G EN Prejeen Fne es Y fitactsA R PLAstr
.TPE Shop onger.2~PGC I 9
\\
l Purpoee: EvhLuATE ms-Po TEk ri A L POR Cc AirA-1NMEtO T F'A t L U ILE DME' 70 HY'bR CG E^l 0AM Bsss n C A.
Results Summary. 77+E FARL.Ey (c k TM( N M E/JT
/5 s*Jc7 E4 S cf PT/ 8 LG" 70 PktLullE pug 70 H, pffL AGR A ncd CR DE Tc N A-TI ON.
(Ref. f.i.9 ef Arrnweb).
CW I
AuNs):
Naete (Print or Type)
Signature Dese O
YL k) l MA-NG p t Lo X SECTION TO BE COMPLETED BY VERIMER(S):
CW Verifm(s):
Name (Pnnt or Type)
Signature Dese kao0Baa/U KN/r2.
- 3. P.
Bua e t arc a V'
SECTION TO BE COMPLETED BY MANAGER:
Approval Responsible Manager:
Si Name (Print or Type)
M. E. EER6Et Lh f
$W
/
4 i
- -