ML20078L262
| ML20078L262 | |
| Person / Time | |
|---|---|
| Site: | Farley  |
| Issue date: | 06/30/1992 |
| From: | FAUSKE & ASSOCIATES, INC. |
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| Shared Package | |
| ML20078K997 | List: |
| References | |
| FAI-91-44, NUDOCS 9411290129 | |
| Download: ML20078L262 (54) | |
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FAI/91-44 1
FAitI2Y NDCUAR FIANT UNITS 1 AND 2 PHEN 00ENOIDGICAL EVAIEATION SIMMAltY Ott STEAN EIFIDSIONS IE SUPPORT OF THE INDIVIDUAL PLANT EIAMINATION Submitted To:
Southern Nuclear Operating Company Birmingham, Alabama O
Prepared By:
Fauska & Associates, Inc.
16W070 West 83rd Street Burr Ridge, Illinois 60521 (708) 323-8750 June 1992 l
O 9411290129 941109 DR ADOCK 05000348 d n !' a r.
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4 PDR
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L) as m cr Steam explosion phenomena are evaluated for both in-vessel and ex-vessel events as potencia.' mechanisms for containment failure under severe potential causes for radioactive accident conditions and, the refore,
as releases to the environment.
The issue for in vessel steam explosions is whether an explosion of sufficient magnitude to fail the reactor vessel, with consequential failure of the containment, could occur.
This was addressed by evaluating the fundamental physical processes required to create an explosion of such magnitude.
The analysis closely follows the IDCOR assessment of this phenomena and concludes that explosions of this magnitude could not be established within the confines of the Farley reactor vessel.
This is in agreement with the findings of the NRC sponsored Steam Explosion Review Group (SERG) which concluded that the likelihood of an in-vessel steam
/%
explosion leading to containment failure (alpha mode failure) was very unlikely.
Ex vessel steam explosions have been addressed by considering both the potential for rapid steam genera, tion as a result of the explosive interac-tion and the shock waves that could be formed and propagated to the containment boundary.
These analyses clearly indicate that sufficient steam ove rpre s sure to challenge the reactor containment integrity would not be achieved under any realistic conditions.
In addition shock waves that could be produced by explosive interactions, when propagated to the containment boundary, result in overpressure values which are well within the steady-state design basis of the containment boundary.
Consequently, the assessment of steam explosions for the Farley reactor system results in the conclusion that neither in-vessel nor ex-vessel events would lead to conditions which approach the containment capability.
As a
- result, steam explosions are not included as events in the Farley IPE plant response trees.
_. D\\
TABI2 OF CGITENTE
..Q rass ABSTRACT i
11 TABLE OF CONTENTS e
iv LIST OF FIGURES.
t I
LIST OF TABLES v.
........ 1-1 1.0 PURPOSE 21 2.0 PHENOMENA
. 2-1 2.1 Description 2.1.1 Controlling Physical Processes 2-1 2.1.2 Relationship to Containment Failure
.................24 Mechanisms and Modes bi
.\\_/
2.1.3 Relationship to Source Tern 2-5 2.2 Industry Experience With Steam Explosions
. 2-6 2.2.1 Nuclear Incidents
. 2-6 2.2.2 Non-Nuclear Explosion Boiling Studies
. 2-7 2.3 Experiments 2-9 2.3.1 Early Sandia Thermite and Corium Experiments
..... ?-?
2.3.2 Aluminum Water Experiments 2-13 i
2.3.3 Liquefied Natural Gas and Water Experiments 2-15 1
l 2.3.4 FAI Thermite Experiments 2-15 2.3.5 Sandia FITSB Tests 2-17 2.3.6 Summary 2-27 l
C s
~~
b t
111 -
i O
TABI2 OF CONTENTS (Continued)
B ZAsa 2.4 Analysis 2 27 2.4.1 Effect of System Pressure on Steam Explosions 2 27 2.4.2 Steam Explosion Models 2-28 2.4.3 Possible Mechanism for Maximum Steam Generator 2-30 Rate 5
2.4.4 Shock Waves 2-32 3-1 3.0 METHODOLOGY.
3-1 3.1 In Vessel Steam Explosions 3-1 3.2 Ex-Vessel Steam Explosions 3.2.1 Pressure Rise Due to Rapid Steam Generation.
3-2 l
3.2.2 Shock Waves 3-3 4.0 PIANT SPECIFIC APPLICATION 4-1 4.1 Issues.
4-1 4.1.1 In-Vessel Steam Explosions 4-1 4.1.2 Ex-Vessel Steam Explosions 4-1 4.1.3 Uncertainty Considerations 4-2 4.2 Conclusions 4-4 i
5.0 ~
SUMMARY
51 1
6.0 REFERENCES
61 t
i k
7
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- iv -
LIS7 0F FICURES Figure No.
Zgg,g l
2-1 Behavior modeled in VASH-1400 2-2 2-2 Comparison of predicted pressure time behavior from VASH-1400 (400 pa particle size) and available experimental results for steam explosions 2-3 2-3 Measured debris-vater energy transfer rates 1
from EPRI sponsored Mark I liner tests 2-16 2-4 FITS containment chamber.
2 19 2-5 FITS 2B chamber air pressure 2-23 1
2-6 FITS 3B chamber air pressure 2-24 27 FITS 7B chamber air pressure (no camera data) 2-25 2-8 FITS 6B chamber air pressure (saturated water) 2-26 2-9 Debris dispersion configuration 2 31 2-10 Comparison of shock wave pressures for TNT and point source explosions 2-33
)
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4Pihiffo#T
1-1 m
1.0 PURPOSE A steam or vapor explosion refers to a boiling process in whit'h steam or vapor production occurs at a rate larger than the surrounding media can acoustically relieve the resulting pressure increase, leading to the forma-tion of a shock wave.
In previous studies evaluating the public risk asscciated with severe accident sequences, such as the Reactor Safety Study
[NRC, 1975), steam explosions within the primary system have been considered as a potential mechanism for violating both the primary system and the containment, thereby generating a direct release path for fission products.
The in vessel steam explosion con.sidered was theorized to result from the following chain of events:
1.
loss of water from the core resulting in fuel overheating and
- melting, 2..
the catastrophic collapse of the core debris into the water remaining in the lower plenum, O
3.
an instantaneous fine scale intermixing of the core debris and I
- water, 4
rapid heat removal from the core material and expansion of the steam assinst an assumed continuous, overlying liquid slug, 5.
impact of this liquid slug on the reactor vessel head with sufficient energy to rupture the head, and 6.
ejection of this missile with sufficient velocity tc fail the containment vall upon impact.
In NUREC-1116
[NRC, 1985), the NRC sponsored Steam Explosion heview Group (SERG) provided recommendations regarding the likelihood that an in-vessel steam explosion could cause containment failure.
The main conclusion of the group report was:
" Based upon the probability estimates summarized above, the consensus of the SERG is that the occurrence of a steam explosion of sufficient energetics which could lead to alpha acde containment failure has a low probability.
This conclusion is reached despite the expansion of differing opinions on modeling of basic steam explosion sequence phenomenol-ogy."
1-2
(
Ex vessel steam explosions may also potentially occur in the progres-
\\- '
sion of ' a severe accident should debris be discharged from the reactor vessel into a pool of water.
Within a containment, the occurrence of a steam explosion would impose shock waves on submerged surfaces and subcom-partment walls.
These must be evaluated to determine if the resulting loads could challenge the integrity of interior walls and the containment bound-ary.
In Generic Latter 88-20 (NRC, 1988), the NRC identified steam explo-sions as a potential containment failure mechanism that should be assessed as part of an IPE. Both in-vessel and ex-vessel steam explosions have been postulated as a potential mechanism for early containment failure,. possibly-with an elevated release location.
Either of these characteristics could have substantial effects on the consequence evaluation for hypothetical accident sequences. The objective of this report is to evaluate the poten-tial for in vessel and ex-vessel steam explosions to threaten containment integrity in the Farley IPE.
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2.0 PHENOMEM6 -
t 2.1 Description Explosive interactions between higher and lower temperature liquids have been encountered for decades in metal foundries as well as the pulp and paper industries.
Experience has shown that these accidents can result in significant damage to typical industrial components (furnaces, casting pits, recovery boilers, etc.)
as well as to light industrial buildings.
Human casualties have also occurred as a result of'these events, but the major r
hazard to operating personnel from these events has generally been burns resulting from hot molten material dispersed by the explosive interaction.
such non-nuclear experiences, destructive steam explosions In addition to have been observed in the BORAX {Deitrich, 1965) and SPERT [ Miller, 1964]
test reactors as well as in the SL-1 accident (SL 1].
In all of these three test reactor configurations, the destructive explosion followed a rapid
(-
30 ms) reactivity insertion that was sufficient to melt both uranium-aluminum alloy fuel and aluminum cladding.
Figures 2-1 and 2-2 provide some visualization of the physical processes considered for an in-vessel steam explosion in (NRC, 1975).
2.1.1 Contro111ne Physical Processes For large scale steam explosions to occur inside or outside a reactor
- vessel, large fractions of hot molten material must be very finely frag-mented and intermixed with the water on the time scale of the explosion.
Such processes were envisioned in [NRC, 1975), but, in addition, rapid heat transfer was calculated in the supercritical and superheated steam regions, and the resulting energy transfer was delivered to a postulated overlying liquid slug that covered the interaction zone. Without the slus transmis-sion mechanism, the pressure-time curves shown in the Reactor Safety Study
[NRC, 1975) would have been insufficient to rupture the reactor pressure vessel, which was the mechanism envisioned as causing containment failure (a mode failure).
i
2-2 O
Core Debris
$ f 0 *o g f
o
=
o o OoOo O
Water
- oOo o o 0
O O
0o o (A) initial separated (B) catastrophic Failure Configuration
& Instantaneous Mixing f
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'O$o' !O Co j0,Sjugo 2
J*
o
- Nd%#
l Interaction g
Zone
'q 'qfec}',X (C) Sustained Energy (D) Slug impact Transfer & Slug Acceleration Figure 2-1 Behavior modeled in WASE-1400.
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2-3 O
i 90 g
Pressure Reauired for Vessel FaBure Pressure Above Exploding Mixture (WASH-1400) 70 N
g 60 2
-)
~
y 50 Pressure in Exploding C
Mixture (WASH-1400) m
/
1 m
40 w
l CC' O
30 Thermogna_m3 Crttjal Pressure for Water 20 l
8 Measured Pressure-T1me Data 10; -
for Steam Explosions I
I o
O 10 20 30 40 50 80 70 TIME, msec Tigure 2 2 Comparison of predicted pregaure-time behavior from UASH 1400 (1.575 x 10 in particle size) and available experimental results for steam explosions.
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l V
The key physical factors which determine the magnitude of steam explo.
sions within BWRs and PWRs include the energy required to rupture a reactor pressure vessel, the amount of core material needed to provide such an i
energy release, the fragmentation of the hot material in the water, the mixing energy requirements when the material is finely fragmented and rapidly inter 1sixed during the explosion, the size of an external trigger to initiate the explosion, the propagation characteristics for the coarsely fragmented system, the likelihood of having a water slug over the reaction zone to transmit the energy in a coherent fashion, and the ability of this slug to be transmitted through upper core structures within the reactor pressure vessel.
Each of these factors must be sufficient to create an event of enough magnitude to rupture a reactor pressure vessel; the failure single factor to achieve the proper conditions will preclude an event of a of such magnitude.
2.1.2 Relationshin to Containment Failure Mechantsus and Modes O
Both i-vessel and ex-vessel steam explosions have been postulated to O
be early con.ainment failure mechanisms that would occur immediately follow-ing slumping of the core material into the reactor vessel lower head.
The largest potential for the occurrence of an in-vessel steam explosion would exist during a core melt sequence with a low primary system pressure.
The largest threat to containment integrity as the result of an ex vessel steam a relatively coherent explosion would exist for a core melt sequence with pour of molten material at vessel failure into a water pool.
Three actual containment failure mechanisms are considered to be encom-passed by containment failures induced by steam explosions.
For in vessel steam explosions, a missile (e.g., the reactor vessel upper head) would have to be created with sufficient energy to pierce the containment.
For ex-vessel steam explosions two possibilities are considered:
the blast could weaken the cavity walls sufficiently that the vessel moves and tears out one or more containment penetrations, or the generated steam could overpres-surize the containment.
For the first mechanism the failure area has been assumed as a large break in the containment wall (i.e.,
on the order of several square feet).
If containment penetrations are torn, it is expected
i 2-5 I
i that the failure area would be small, more like a " leak before break" condi.
tion.
A failure resulting from containment overpressure would be dictated by the rate of pressure rise and the total mass of steam generated.
Typically ~this would be a " leak before break" response, but, as will be discussed, the anticipated rate of pressure increase from a steam explosion is less than that associated with a desi n basis large break 14CA.
5 2.1.3 Relationshin to Source Term Containment failure resulting from a steam explosion would influence the expected fission product source term for a sequence by providing a large gas flow path out of the containment shortly after vessel failure. The effect on the source term.
- however, would strongly depend on the availability of water inj ection or sprays into the containment during a sequence.
For sequences in which containment sprays would be available before the explosive interaction, removal of airborne aerosols would be very e f fec tive.
So much so that release to the environment would likely not be much greater than the release of the noble gases. However, such an early U
containment failure would generally increase the source term for accident sequences without containment injection or sprays, since the airborne fis-case with the sion product concentration would be much greater than for a sprays operating.
- Also, the relatively large expected failure size would cause a rapid blowdown of the initially available airborne fission products to the auxiliary building or environment, thereby reducing the. fission product retention effectiveness of the containment.
On the other hand, fission products entering the containment atmosphere after the blowdown would experience little driving force from the containment to the auxiliary building or the environment. Thus, fission products evolved by long-tern revaporization within the reactor vessel would be subject to the naturally-occurring deposition mechanisms in the containment.
For sequences in which vessel inj ection was restored after vessel failure or containment sprays initiated, the expected fission product source term would be somewhat reduced. Delivering water into the containment via either means would cool the containment atmosphere and contribute to a p
reduction in the magnitude of long-term fission product revaporization. For V
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.k /
those sequences which could experience a lower head failure, use of either s
containment injection method would assure the debris is covered by water.
For an upper head failure both vessel injection and containment sprays would again act to accomplish this objective.
2.2 Ina'=try s5merience With steam Exnlosions i
2.2.1 thaclear Incidents 6
The explosion model used in the Reactor Safety Study (NRC, 1975]
resulted prinetpally from concerns generated by the low pressure BORAX and SPERT destructive experiments and the SL 1 accident. Reactor conditions leading to the SL-1 accident (SL-1] and the destructive transients in BORAX
[Deitrich, 1965]
and SPERT [ Miller, 1964) were produced in a fundamentally j
different system than that representative of a postulated severe accident in a commercial LVR.
It is not only important to realize these differences, but it is essential to understand the resulting implications for the
}
phenomenon as well. These basic differences are delineated below.
- 1. ' All three destructive events were produced by power excursions in which the core was driven to molten conditions in 30 msee or less.
Such strong reactivity transients are not possible in commercial power reactors.
2.
The specific core designs of these reactors could be brought to supercritical conditions by the withdrawal of a single control rod.
In these transients, a control rod was rapidly withdrawn which caused a nuclear excursion with sufficient energy deposition to melt the fuel clad plates.
l l
3.
Each of these three reactors was fueled with thin uranium-aluminum alloy fuel plates clad in aluminum. Thus, the fuel l
and water were uniformly cremixed on a fine scale in the as-fabricated geometry.
No additional melt fragmentation was required to accomplish the explosive energy release.
4.
Since the reactors were essentially at room temperature prior to the excursion, the vessels were filled with cold water except for a small freeboard volume at the top, i.e.
a long, coherent overlying liquid slug was in place prior to the reactivity insertion.
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t 2-7 i
LG l
internal geometry'was very simple and open, which i
5 The vessel t
provided little attenuation or dispersion of any slug move-ment.
I Vith these initial conditions, the configuration established was essen -
cially an inertial layer of water above an expanding layer of water, as assumed in the Reactor Safety Study. The essential feature-of the strong reactivity transient is that it brought the fuel and clad to melting before this configuration could substantially change.
Given these particular characteristics, a slug impact following a steam explosion with the core e
would indeed be the expected chain of events. However, this is fundamen-tally different than an initially separated state of high temperature molten core material and saturated water existing at an elevated pressure 'with substantial internal structure to prevent catastrophic collapse, intimate mixing, and slug formation.
2.2.2 Non-Nuclear hnlosion Boilina Studies In a number of industrial operations the possibility exists of contact-ing two liquids - -one hot and relatively nonvolatile and the other cold and l
E volatile.
Should such an event occur, boiling would occur in a sufficiently short time scale that the surrounding medium cannot relieve the expansion l
acoustically and a shock wave, i.e., an explosion, forms. Accidents of this nature have been given various names, e.g., explosive boiling, rapid phase transitions (RPTs), vapor explosions, thermal explosions, fuel-coolant interactions (FCI), etc.
They have been observed in a number of industrial operations, e.g., when water contacts molten aluminum (or other metals),
molten salts or paper mill smelt, or when cryogenic liquids such as 1RC (liquefied natural gas) are spilled into water.
In the first two examples
}
noted above, water is the volatile liquid which explosively boils whersas in the last example the cryogenic liquid plays the role of the volatile, boil-ing liquid and water is then the " hot" fluid.
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2-8 O
1 2.2.2.1 Smelt-Water Ernlosions Studies of molten salt-water explosions were carried out because in-dustrial accidents involving these reactants have taken place.
Emphasis has been placed on events occurring in the paper industry where molten smelt is
'produced in the recovery boilers.
This smelt is a mixture of, primarily, sodium chloride, sodium carbonate, and sodium sulfide. The smelt tempera-ture is much higher than the critical point of water (- 1520 *F (1100
'K) compared to 750 'F (647 'K)).
Severe explosions have taken place when water inadvertently contacted molten smelt.
Laboratory investigations
[Krause, et al., 1973), [Shick, 1980) into the mechanism of smelt-water explosive boiling events have been primarily useful in delineating the effect of smelt composition on the sensitivity of the salt in producing explosive boiling.
For example, pure molten sodium carbonate has never led to explosive boiling. Addition of either sodium chloride or sodium sulfide, or both, leads to smelts which are more prone to explosive boiling.
Investigators experimented with many. additives both to the smelt and to the water in an attempt to obtain less sensitivity.
Most had little or no effect.
2.2.2.2 Melt-Water Interactions The metals processing industries, particularly those producing aluminum, have also been plagued by explosive boiling incidents.
Alcoa has carried out several test programs
[Lemmon, 1980), [Hess, et al., 1980]
directed primarily at effecting means to prevent such accidents in casting plants.
In most tests, molten aluminum was dropped into water and the subsequent events recorded. Many variables were studied such as water temperature, drop height, nozzle diameter, etc.
The principal result of these investigations was to show that water containers, suitably coated with an organic-based paint, would not lead to explosions when molten aluminum was spilled into the container. Use of such paints in aluminum plants has indeed reduced the frequency of explosions, but many still occur.
In a j
large number of accidents, the quantity of water was quite small, e.g., when
[
" vet" aluminum ingots were loaded into melting furnaces containing molten
29 t/
In contrast to this fact,.few, if any, serious events have oc-curred when small quantities of aluminum were contacted with a large mass of water.
Since laboratory tests were often carried out in the latter fashion, most of these have not resulted in explosive interactions.
2.2.2.3 Other In industries dealing with " reactive" metals, such as t.itanium, zir-conium, etc.,
only a few serious explosive boiling events heve been documented.
In most of these, a significant quantity of molten metal has contacted water and, simultaneously, there has been some external shock such as an electrode falling into the metal-water mixture.
In the few knowt.
incidents, damage has been severe, but quite localized.
Due to the reactive nature of the metal, however, subsequent hydrogen fires have often com-pounded the problem and led to extensive damage.
2.3 Experiments A wide range of laboratory scale and large scale experiments relating to vapor explosions have been performed over the past thirty-five years.
The laboratory scale experiments constitute an extensive literature base and are reviewed in detail in [FAI, 1982] and [IDCOR, 1983].
2.3.1 Early.Sandia 'rbermite and Corium Fxneriments large scale steam explosion experiments have been carried out at the Sandia laboratories in three different test series, two using an iron-thermite mixture
[Buxton, et al.,
1979],
[Mitchell, et al., 1981] to simulate the degraded core material and the other using both an iron-thermite and a corium thermite
[Buxton, et al., 1980] which had a higher melting temperature and is a more realistic simulant of the anticipated debris character.
In both of these experimental series, artificial triggers (explosive detonators) were used to initiate the interaction in some tests.
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OV-In tha first set of experiments (Buxton, et al.,
1979), the iron-thermi.e melt was discharged directly into a 2.95 ft (0.9 meter) diameter vessel filled with water.
For all those experiments carried out with an artificial trigger the water was at the ambient temperature, assumed to be 70
- F (295 K).
(In these experiments the ambient pressure was always slightly less than 14.7 psi (0.1 MPa).)
The melt temperature resulting from the thermite reaction is approximately 4400*F (2700 K) and results in reae-tion products of metallic iron (Fe) and aluminum trioxide (A1 0 ).
The 2 s melting temperature for the aluminum trioxide is approximately 3680*F (2300 K) and that for metallic iron is about 2780*F (1800 K).
Consequently, solidification of either of these constituents requires a substantial decrease in temperature and the resulting fragmentation process could con-tinue as the melt cools, i.e.
lower temperatures reduce the film boiling steam generation rate and allow finer particulation.
To externally trigger an explosive interaction, a 1.41 x 10~
lbs (0.64 g) charge of high ex-plosives was used.
Some of the experiments had a considerable delay before the explosion was initiated, i.e. over 3 sec.
Many tests in this experimen-tal series observed the presence of spontaneous trigger events as well.
A second test series was performed at Sandia (Buxton, et al., 1980) with a different test vessel (3.94 ft (1.2 m) internal diameter) and molten f
material generated from both iron-aluminum oxide thermite and a corium A+R thermite. This latter reaction had products of uranium dioxide, zirconium dioxide, nickel oxide, stainless steel, and molybdenum. The minimum li-j quidus temperature for this mixture is reported to be 4526*F (2770 K), which is considerably greater than the 3680*F (2300 K) temperature for aluminum j
oxide.
Boiling steel would limit the maximum temperature for the corium reaction to 5066*F (3070 K).
In this second test series, external triggering was also induced by
)
explosive detonators, but two different sizes were used. One was the same as that employed in the first iron thermite test, i.e. 1.41 x 10~3 lba (0.64 g) of PETN, and the other was a detonator plus a lead-covered explosive cord (0.76 m) in length and containing 1.32 x 10~2 lba (6 g) of PETN.
2.49 ft This second method represented a much more energetic trigger than that used in the thermite tests.
In fact, the pulse duration for the corium A+R event
2-11 A(,)
in Run 59, which used this larger trigger, vss not much different than that represented by the trigger alone. Also the measured work (- 33 Btu (30 kJ))
was less than the work released by the high explosive (- 39 Btu (35 kJ)).
Explosions were observed with the iron-thermite as initiated by both spon-taneous and artificial triggers.
However, with the corium A+R melt only one mild explosion is reported and this was triggered by the 1.32 x 10-2 lba (6 g) PETN external trigger.
The time delay before the trigger is fired (- 1.3 sec) is longer than the time required to cool the coarsely fragmented par-ticles to the liquidus temperature.
The fact that the trigger was needed to mix a considerable fraction of the melt down to an explosive size scale is indicative of the difficulty encountered in making such materials undergo a the rmal explosion.
A major part of this difficulty is due to the rapid cooling and freezing of the corium particles as described above.
The results of 17 tests were reported [Mitchell, et al.,1981] for both ambient and high pressure initial conditions and also with and without an external trigger.
The two experimental series were designated as Melt Delive ry and FITS (Fully Instrumented Test Series).
The Melt Delivery C/
experiment consisted of 12 tests all run at atmospheric initial pressure s
without an external trigger.
These resulted in eight self-triggered explo-sions -- two in the water coolant before t'ae melt impacted the rese rvoir bottom and six when the melt contacted the reservoir base. The initial FITS experimental matrix consisted of five runs -- three at atmospheric and two at an elevated pressure.
Two explosions resulted -- one at atmospheric pressure and in the free stream before the melt hit the reservoir bottom, and one at elevated pressure (- 1 MPa/150 psi) which was initiated by an external trigger when the melt was lying on the reservoir base.
The purpose of the Melt Delivery test series was to develop an effi-cient means of delivering the melt into the coolant and that of the FITS experiment was to determine the mechanical work output from such explosions.
The melt used in both these experiments was iron-thermite (Fe-Al 0s) which 2
had an initial temperature of - 4400*F (2700 K).
Melt masses of 1,32 - 11.8 lba (0.6-5.38 kg) were employed which resulted in coolant-melt mass ratios of 366-37 respectively, and the conversion ratio of mechanical work to melt f
initial thermal energy was reported to be about 1-3t.
These tests resulted k
2-12
^
f b).
in eight self-triggered explosions, but explosions were not observed for water-melt mass ratios of 83-113.
The explosivity and initiation site were apparently sensitive to melt mass and shape in the 15 non-externally triggered tests at ambient pressure.
There were no self triggered explosions when the initial melt mass was less than approximately 4.0 lba (1.8 kg), and there were nine such explosions for the initial melt mass greater than - 6.6 lba (3 kg). This was interpreted as a threshold for an appropriate melt-water mixture to produce a thermal explosion.
Five different phases were observed in these experiments:
(1) melt entry, (2) pre-mixing, (3) triggering, (4) propagation, and (5) expansion of the interaction products. The melt was observed to start coarse fragmenta-tion and pre-mixing virtually upon entry into the water.
The triggering and propagation phases of the thermal explosions were observed to start at the leading edge of the melt, both when the melt was still in the free stream and also upon impact with the coolant reservoir base.
Propagation of the Oy event was observed to start at the base of the melt and propagate through the mixture at - 656 - 1969 ft/s (200-600 m/s).
This phase was considered to be complete at the start of the expansion of the melt-water mixture.
to assess the The FITS-A experiment was conducted in a closed vessel influence of an external trigger (1.4 x 10~
lbu (635) og PETN, 3.6 Btu (3.8 KJ)) on high pressure cut-off of thermal explosions.
In this test series, the Fe-Al O melt (4.3 - 11.8 lba, 5072*F (1.94-5.38 kg, 3073K)) was poured 2 3 from - 3.9 ft (1.2 m) into tap water (198 - 498 lba, 50 - 77 *F (90-226 kg, 10-25'C))
producing a 1 - 1.5 ft (0.3-0.45 m) long melt mass at entry and three at coolant-melt mass ratios of 41 80.
Five tests were performed ambient pressure (12 psi (0.083 MPa)) and two at an elevated pressure (148 and 158 psi (1.02 and 1.09 MPa)).
Two explosions resulted -- one at ambient pressure which was self-triggered in the free stream and one at 1.09 MPa which was externally triggered.
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2-13
'l d(3 Tests FITS-4A and 5A are of particular interest because they were designed to investigate the effect of high pressure cut off on steam explo-sions.
In the FITS-4A test, 9.5 lba (4.29 kg) of melt was delivered to 498 lba (226 kg) of water at 77'T (25'C) in an ambient pressure of 148 psi (1.02 MPa) without an external trigger and did not produce an explosion. This benign result was explained on the basis of a dispersed and cooled melt at entry into the coolant.
The FITS-SA tert was performed at a system pressure of 158 psi (1.09 MPa) and was essentially a rerun of FITS-4A but with an external trigger.
The FITS SA run did produce a thermal interaction after being initiated with the detonator (3.6 Btu (3 s KJ)), but the mixture had not self-triggered after 0.44 see at which time all the melt was on the bottom of the coolant reservoir.
2.3.2 Aluminum-Water Erneriments Large scale tests have also been carried out for an aluminum water system where either external triggers [Ieng, 1957), [Hess, et al., 1980),
[Lemmon, 1980), (Higgins, 1955), [Higgins, 1956) or a shock tube configura-tion [ Wright, et al.,1966] have been employed.
In (leng,1957), large scale molten aluminum-water experiments were performed to investigate the manner in which steam explosions could be triggered. The reference test, which repeatedly produced explosions, in-volved the discharge of 22.8 kg (50 lba) of commercially pure molten aluminum into a clean, mild steel container partially filled with water at 78.1
'F (12.8-25.6'C).
In contrast to chemical temperatures of 59.5 explosions, no flash or fire could be detected either during or after the explosions.
The following parameters were varied:
1.
discharge rate and mass, 2.
drop height, 3.
water depth, and 4.
aluminum and water temperatures.
Also, different water additives, solid surfaces, and surface coatings were employed in the experiment. It was concluded that three requirements must be met to produce an aluminum. vater explosion:
2-14 1.
Molten metal in considerable quantities must penetrate to the
~ '
bottom surface of the water container.
2.
A triggering action must occur on the container bottom surface when it is covered by the molten metal.
3.
The water depth and temperature must lie within certain ranges..
In (Hess, et al.,
1980), tests were performed to study the level of external stimulus (a hammer impact) required to initiated explosive interac-tions in aluminum-water systems.
For th6se experiments, - 48.5 lba (22 kg) of molten aluminum was poured into a square container 0.98 ft (0.3 m) on a side.
The molten aluminum temperatures varied between 1346 'F (730*C) and 1436 'F (780*C) with the water temperature variation being from 37.4
'F (3*C) to 89.6
- F (32*C).
In the experiments, 4 sec elapsed between the entry of melt into the water and the hammer impact, thereby allowing much of the material to be accumulated on the bottom of the container instead of as individual particles in the water.
The impact level determined in the experiments of (Hess, et al.,1980] was 0.176 Btu (186 J) and the authors s
L suggested that perhaps only half of this was actually transmitted to the mixture due to inherent losses within the impact on the wall and the trans-mission of the energy to the coolant.
Other aluminum-water experiments have been carried out by Lanunon
[Lemmon,1980) and Higgins (Higgins, 1955),
[Higgins, 1956] where molten material has been poured or injected into weer and an explosive interaction was initiated by a strong external trigger. For those experiments reported in
[Lemmon, 1980),
triggers up to 1.1 x 10 2 lba (5 g) of primacord were used.
A No. 6 blasting cap was employed by Higgins in his experiments.
Scoping calculations for the specific experimental configurations used in these references result in an assessment that the external stimulus was orders of magnitude greater than that required to rapidly six the materials on an explosive time scale.
Another type of aluminum water experiment of note is the shock tube experiment described in [ Wright, et al., 1966).
In these
2-15 and a cover gas. To carry out these tests, the recovery gas was evacuated a
1 and the diaphragm was ruptured allowing the atmospheric pressure to ac-celerate a water slug resulting in a strong, direct impact of the cold water q
column on a molten aluminum surface Large interaction pressures for these events were measured in the water column.
1 2.3.3 Liauefied Natural Cas -i Water Erneri-nts Large scale tests [Koopman, et al.,
1981]
have been performed with Liquefied Natural gas (LNG) and water using material volumes approaching those of interest for the reactor accident case.
In these tests, water is the hot fluid and LNG (mostly methane) is the cold liquid which undergoes the explosive vaporization.
This fluid pair is similar to the corium water system in that the interface contact temperature is far greater than the thermodynamic critical temperature of the LNG, making explosions difficult to initiate.
Long delay times were provided in an attempt to accumulate substantial quantities of LNG below the water surface.
The magnitudes of
{}#
the explosions obtained represented the interaction of only a small fraction Ns of the LNG injected.
2.3.4 FM Thermite Erneriments l
i Two sets of experiments have been performed at FAI in which 44 lba (20 kg) of molten iron-thermite was injected into water. The first [Malinovic, et al.,1989) was performed to study the role of water in protecting the Mark I containment liner under severe accident conditions while the second
[FAI,1990] addressed the influence of water during a high pressure melt ejection.
Both of these represent conditions which could cause ex-vessel steam explosions and both facilities were instrumented sufficiently to evaluate the steam generation rates resulting from these interactions.
Interpretation of the rate, in terms of a heat flux based upon the projected floor area where the interaction occurs, provides a means of 1
applying the results to a reactor / containment system.
Figure 23 il-lus trate s the measured heat flux to the overlying water pool in the Mark I O
O O
O 2
iii;iiiiiiiii;iiiiiiiiiiiiiiiiiiiiiiiiiiiii jiii y 10 E TEST 8 h
A TEST 9 O TEST 10
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CONDUCTION SOLUTION 1
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i g
A i
_i iniiii,ii iii n Iiii i t i n in lii ii5 i n iIiiich niii>>>>
10 o so 100 150 200 250 350 350 400 450 500 TIME (sec)
Figure 2-3 Measured debris-water energy transfer rates from EPRI sponsored Mark I liner tests.
2 17 A
experiments when the test apparatus was instrumented to detect the energy transfer to the test box walls. All tests show a very high energy transfer 6
rate within the first few seconds, the value being between 6.3 x 10 and 9.5 6
and 30 MV/m ), which subsequently decreases to about 2
x 10 Btu /h ft2 (20 6
(0.9 MW/m ) after the debris is frozen.
In this set of 2
0.28 x 10 Btu /h-ft2 experiments, 11 tests were performed 10 of which had water available in the
. simulated containment prior to the discharge of the molten iron thermite.
6 6
In all 10 experiments, rapid energy transfer rates (6.3 x 10
- 9.5 x 10 (20-30 MV/m )) were observed when the debris was discharged into Btu /h-ft2 2
the water.
FAI direct containment heating experiments [FAI, 1990) also had suffi-cient instrumentation to estimate the steam generation rates when debris was discharged from the simulated RCS into the reactor cavity and subsequently up onto the containment floor.
Table 2 1 summarizes the information for these experiments in terms of the energy transfer rate in the cavity for the three experiments in which water was available (DCH-1, DCH-2, and DCH-4) and f
also for the energy transfer rates from the debris to the water as the debris was discharged onto the containment floor. Values are also given for estimated additional energy transfer due to the transfer into the steel structural heat sinks in the simulated containment lower compartment. These additional energy transfer rates should be summed with those determined from the containment compartment pressurization rates.
As illustrated by this table, the energy transfer races are large and comparable to those observed in the Mark I tests.
These rates are an order of magnitude greater than those typical of the critical heat flux (CHF) for a horizontal upward facing surface.
2.3.5 Sandia FITSB Tests Later Sandia FITS tests provided sufficient pressure transient informa-I tion to evaluate the average steam generation rate resulting from explosive interactions.
Steam generation rates can then be divided by the cross-sectional area of the FITS vessel to determine the effective heat fluxes.
Figure 2-4 taken from [Mitchell, et al., 1986) shows a cross-section of the O
.m
4
--*ui 2-18 I
Table 2-1 EFFECTIVE HEAT FLUX MEASUREMENTS FOR DEBRIS-WATER INTERACTIONS Initial Intermediate ung Term Test Pressurization Period Quenchina 2
M Btu /h-fc2 MW/m 2 2
2 M Btu /h-ft2 MW/m M Btu /h-ft MW/m DCH-1 4.75/13.3*
15/42*
3.49/6.02*
11/19*
2.7/5.5*
8.5/17.5*
DCH-2 2.22/10.78*
7/34*
4.12/6.66*
13/21*
2.31/5.17* 7.3/16.3*
DCH-3 N/A N/A N/A N/A 1.27/4.12*
4/13*
DCH-4 1.27/9.83*
4/31*
3.49/6.02*
11/19*
N/A/2.85*
N/A/9*
- Contribution from the heat sinks added to the vaporization calculation.
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O Figure 2-4 FITS contaitument chamber.
2-20 i
p(l FITS facility.
In this test series, about 18.6 kg (41 lbm) of molten ther-mite was poured into water test containers located in the FITS chamber and the resultant pressure history in the chamber gas space was recorded.
Table 2-2. which was also taken from (Mitchell, et al.,1986), summarizes the test conditions and obse rva t ions made vith respect to explosive interactions.
Figures 2-5 through 2-8 illustrate the pressurization of the gas space, the first three with initially subcooled water and the last with saturated water.
While only some of the experiments had explosive interactions, the j
principal focus is on the net steam generation rate created by the explosive interaction.
The large steel vessel is considered to be pressurized with steam with the realization that this also increases the potential for con-densation on the vessel walls.
The tests shown in Figures 2-5 through 2-8 are those with the largest vessel pressurization.
A comparison of these figures also shows that the time to the peak pressure is approximately 1 sec for these tests, even though the path to this pressure may differ somewhat.
(Test FITS 7B experienced about 90% of the pressure increase in the first second with the remainder occurring over the next 3 secs.)
1 The average steam generation rate can be estimated by using the perfect gas equation.
dE _ RI AH dt V dt l
where each variable has the standard meaning.
As an average representation, I
1 assume that the gas space pressure increases 0.35 MPa (51 psi) in 0.5 sec.
8 (5.6 m ) (Marshall, 1986]
- and, if 8
The volume of the FITS vessel is 198 ft an average gas temperature of 260 'F (400*K) is assumed, the steam gener-ation rate is 2.65 lbm-moles /sec (1.2 kg-moles /sec),
which is a mass addition rate of 47.5 lbm/sec (21.6 kg/sec).
As the melt enters the vessel, or non-explosive) would expel the dynamic interactions (either explosive melt and water from the lucite test vessel.
To provide an equivalent basis for comparison with the FAI/EPRI Mark I tests, the steaming rate should be represented as a heat flux using the cross sectional area of the FITS vessel
"'T
(- 19.4 ft2 (1.8 m )).
Using this area, the average heat flux from the melt 2
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O
2-27 6
(_,)
2 to the water is about 8.6 x 10 Btu /h-ft2 (27 MW/m ),
i.e. a value in close agreement with that observed in the Mark I experiments.
2.3.6 Summary In summary, the results from significant scale experiments with greatly different geometries can be compiled to develop a basis on which to provide interpretation for the containment response due to rapid steam generation by dynamic interactions.
Specifically, dynamic interactions should be con-6 6
sidered with steam generation rates from 3.2 x 10 to 9.5 x 10 Btu /h fts (10 to 30 MW/m ).
The projected area of the compartment floor should be 2
used as the pertinent value for determining the total energy production rate.
This can then be used to determine if the uncertainties in this range provide for any substantial change in the overall accident progression or in the accident management decisions that would be exercised in such events.
2.4 Analysis
(/
2.4.1 Effect of System Pressure on Stsaa Explosions Severtl experimental investigations have focused on the effect of elevated prissures.
These include studies using Freon-22 as the working (exploding) f uid [ Henry-Fauske, 1979] as well as water [Hohman, et al.,
1979]
and
[Hohman, et al., 1982]. These test series employed fluid pairs which had been demonstrated to explode in a reproducible manner, such that a single parameter (pressure variation) would be meaningful.
Increasin5 the system pressure was observed to prevent explosions in all three studies.
Those performed without exte rnal triggers found that a reduced pressure (ratio of the pressure and the thermodynamic critical pressure for the working fluid) of 0.05 was sufficient to prevent explosions.
This cor-responds to a pressure of 150 psi (1 KPa) for water.
It was also obse rved that the explosions became somewhat less efficient as the ambient pressure increased.
In fact, in the water experiments, no explosions were obse rved for system pressures of - 75 psi (0.5 MPa).
n)
[
/
2-28
[
Experiments performed with external triggers found that explosive
\\
interactions could be induced at somewhat higher pressures, but a reduced pressure of 0.10 was found to suppress explosions even with ve ry strong external triggers.
For water this is a pressure of - 300 psi (2 MPa).
These pressures are well below those for the small LOCA and transient accident scenarios.
Therefore, this single experimental observation is sufficient to address the issue of in-vessel s te am explosions.
Explosive triggers. do not exist in a reactor system.
Hence, the set of experiments most relevant for the IPE analyses are those without explosive external triggers.
These will be used in the application to the reactor system, which would reduce the limiting pressure even further.
For those primary system conditions where explosions could be in-itiated, the assessment of the threat to the RCS integrity needs to evaluate the magnitude of the explosive interactions and the capability of the inter-action to transfer an impact loading to the RPV walls and upper head.
This was treated in the IDCOR Program [IDCOR, 1983] in terms of (1) the maximum molten mass and water which could be intimately mixed, (2) the efficiency of the explosion and (3) the capability of transferring an impact load to the RPV upper head.
Through these evaluations it was concluded that a suffi-cient melt-water mixture could not be established to approach the energy yield necessary for challenging the vessel integrity.
In addition, no efficient energy transfer mechanism could be found which could transmit the necessary impact load to cause failure of the RPV upper head.
2.4.2 Steam Frnlosion Models Numerous models have been proposed to explain the primary steps in the occurrence of a steam explosion.
However, no consensus on modeling of steam t
explosions has emerged to date.
The wide variation in views exhibited by the members of the SERC underscores this fact.
Since the NRC states that no truly predictive mechanistic model exists- [NRC, 1988),
this section is limited to a short overview of the two types of models that exist.
These basically address the mechanisms for fragmentatinn of the lower volatility O
2-29
/aic) material and the mechanisms for providing the intimate liquid-liquid con-tact; the former is required to obtain the characteristic larger heat transfer area, while the latter is required for the characteristic rapid heat transfer rate.
Since fragmentation of the hot material is considered to be a necessary (but not sufficient) condition for a large scale explosive interaction, a
rather extensive experimental and theoretical effort has been devoted to the understanding of this process. The fragmentation models may be grouped into the following four general categories:
1.
hydrodynamics models: which treat effects between the molten material and coolant independent of thermal conditions; 2.
violent boiling models:
fragmentation induced in the molten material via the disruptive forces and associated with bubble growth and collapse including spontaneous nucleation; 3.
thermal stresses theories:
molten material breakup as a consequence of surface quenching and solidification; and
[
4 entrapment / gas release theories:
rapid phase change of an entrapped species resulting in sudden expansion and fragmenta-tion of the molten material.
From metal water experiments it appears that an essential precursor to a vigorons interaction is the establishment of a stable vapor film between the fuel and coolant.
The interaction is triggered by the destabilization of the vapor film, allowing extensive liquid-liquid contact.
In general, the models developed to explain the mechanism allowing liquid-liquid contact can be caughly classified into two broad types:
(1) boiling models, which depend upon the rapid production of vapor after liquid-liquid contact is established, and (2) hydrodynamic models, which depend on the breakup of high temperature material due to the large relative velocity between fuel and coolant after collapse of the vapor layers due to arrival of a pressure wave.
Some recent models consider that both types may be present at the same time.
There are also purely parametric models which are concerned with the consequences, but not the mixing process physics, once a set of initial conditions is assumed.
,O v
30 0
2.4.3 Possible Machanism for Maximim Stamm Censration Rate The information presented in Section 2.4.2 was taken from a wide variety of experimental information and provides a substantial data base'for describing the maximum melt-water steam generation rate in containments.
One can provide a theoretical basis for heat fluxes in the range of 10.4 x 6
Beu/h fts (30 MW/m ) for a system with co-dispersed debris and water as 2
10 depicted in Figure 2-9.
A steam velocity sufficient to levitate and separate the water droplets from the high temperature dense debris is given by 4
3.7 go 'p ( - p)
U E
p 8
where g is the acceleration of gravity, a is the steam-water surface tension and p and p represent the saturated water and steam densities respec-g tively.
If thi.s - is considered to be the maximum steam production rate which could exist without separation of the water droplets from the co-disperse configuration, then the heat flux associated with the vapor production rate is given by 4jgo(p g/A ~ 3.7 h p
g-p gg where h is the latent heat of vaporization.
Substituting the appropriate g
values for steam and water at 1 ata into this expression results in a value 6
of 10.4 x 10 Btu /h ft8 (30 MW/m ); a value in agreement with those obsersed 2
in the various experiments. Hence, the major ramification of an explosive interaction could be the co-dispersion of melt and water which then con -
tinues to transfer energy and vaporize water into the containment atmosphere at a rate limited by the ability of the water droplets to remain as part of the co-dispersed medium.
O
2 31 i
9 s k o
o.
o o
Debris Particulates o
o Water Droplets j.
o o O
o o
o o
o o
o o
a (a) Co-Dispersed Configuration t
O e.o e
e e o
o 0
O e
o i
o o
o o
O '
i
- o oa o
,.o.o.a.
o e
e
\\
(b) Configuration if the Droplets are Fluidized
=ue= e4 Figure 2-9 Debris dispersion configuration.
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2 32 e
(g.
1
)
2.4.4 Shock Waves
.Modeling of the shock waves induced by steam explosions is only_ neces-sary if it is conceived that these would challenge the containment integrity.
Figure 2 10 taken from (Class, 1974) illustrates the decay of substantial shock waves in air as the shock wave e xpands.
A slope cor-responding to a pressure amplitude decay proportional to 1/r8 is also included for reference and provides a reasonable assessment of the decay characteristic for strong waves.
If anything, the higher amplitude portion of the curve decays faster than this simplified representation.
If an
\\
interaction zone size is postulated along with a maximum pressure for the interaction, this type of decay can be applied to the Sandia FITS experi-ments to compare the measured shock wave pressures in these tests with this decay characteristic. Table 2-2 summarizes the experimental conditions for the FITSB series, including the size of the test chamber in which the ther-mite and water were mixed. As an interaction zone, half of the square dimension is used as the radius for the initial calculation. Also, for the
("'/
g peak pressure achieved in the interaction zone one half the critical pres-
\\--
sure (- 1450 psi or 10 KPa) is used since this corresponds to a condition in which the critical size bubble embryos equal the size for thermally dominated bubble growth (Henry, et al.,1979).
For pressures greater than this value, the vapor cannot be produced.
Other experiments have shown this value to be an upper bound of the pressure that can be achieved when the system is not tightly constrained.
The expansion from the interaction zone out to the diameter of the FITS vessel, 2.5 ft (0.76 m) radius, is performed following the approximation shown in Figure 2-10.
Since only three different size vessels were used in the eight experiments, only three different shock wave pressures incident on the FITS vessel wall are calculated by this approximate method.
These are illustrated in Table 2 3 for the different experiments.
As illustrated, this technique substantially overestimates the measured pressure at the FITS vessel boundary.
This is not surprising since the curve shown in Figure 2-10 is compared to a chemical explosion which is typically more energetic and has a stronger shock wave than those generated by steam explosions.
O I
- J
2 33 i
O 1000_
3
~
N Point Source
~
N (1/r)2 1
100 g-
=_
~
6
\\
~
s CO 10 1
1 5
5 4
o TNT 1 g-g g
E 5
b
\\
t
- j l ' I t til ih,L 0.1 1 1 188811-1 I
0.1 1
10 r,ft.
1 l
l Figure 2-10 Comparison of shock wave pressures for TNT and point source explosions.
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1 2-34 i
I~)
l
%)
l l
Table 2-3 CHAMBER AIR PRESSURE DATA FROM FITSB (Times From Melt Entry)
Steam Exclosion Phase i
Calculated Explosion Pressure Peaks Pressure Expc.
(s)
(MPa/psig)
. (MPa/psig)
Peak 1stl l
2nd 1st l
2nd i
4 1B 0.144 0.282 0.095/13.78 0.197/28.6 1.6/218 I
48
-0.029_
0.146 0.020/2.9 0.500/72.5
.1.6/218 88 0.017 0.144 0.01/1.45 0.373/54.1 1.6/218 25 0.087 n.o.2 0.220/31.9 n.o.
1.6/218 3
35 0.081 n.o.
0.440/63.8 n.o.
0.7/87 65 n.o.
n.o.
n.o.
n.o.
0.9/116 i
78 10.20 n.o.
0.01/1.45 n.o.
0.7/87 95 0.102 n.o.
0.210/30.45 n.o.
1.6/218 l
STime taken from start to pressure rise.
Zero time taken from average of two active melt position sensors 2.5 ca above water surface.
2Not observed.
l L
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6
3-1 O
f 3.0 METHOD 01DGT j
The disposition of containment failure due to steam explocions relative to the plant response (PRTs) involves separate approaches for in-vessel and ex-vessel steam explosions, respectively.
The IDCOR work, which is consis-i tent with the recommendation of the SERG in [NRC, 1985), forms the basis for the treatment of in vessel steam explosions.
Results of analyses performed in accordance with significant scale experiments and expansion characteris-tics of shock waves form the basis for the treatment of ex vessel steam explosions.
This section describes the methodology for addressing steam explosions.
3.1 In-Vessel Steam Exnlosions Fundamental experiments (see Section 2.4.1) show that the initiation of steam explosions is very sensitive to pressure levels and is prevented at system pressures beyond 10% of the primary system normal operating pressure.
As a result, there is no threat to the RPV integrity by this phenomenon in
\\
accident sequences that do not meet this criterion.
For sequences whic'.2 result in a depressurized
(<
1 MPa/150 psia) primary system at the time molten core debris would be expected to flow into the lower plenum, the approach utilizes the IDCOR analysis (IDCOR, 1983].
With the mechanistic evaluations for the melt-water masses which could interact and the assessment of the energy transmission capability, there is no set of credible conditions which could approach conditions sufficient to challenge the vessel head integrity. This is consistent with the conclu-sions of the NRC sponsored Steam Explosion Review Group.
3.2 Ex-Vessel Steau Exnlosions As discussed in Section 2, ex-vessel steam explosions could occur and may be a major mechanism for quenching of core debris should it be dis-charged from the reactor vessel. There are two aspects to be addressed:
(1) the overpressure in the containment due to rapid steam generation and (2) the shock waves which could be created by the interactions.
\\
3-2 Ch V
3.2.1 Pressure Else Due to Manid Steam Ceneration Based on the possible mechanism for maximum steam generation rate postulated in Section 2.4.3, the steam generation rate due to the explosive interaction of debris and water can be written as 1
(3 1) fg
?
where in SI units, E,
- steam generation rate (kg/s) q",
- heat flux due to explosive interaction and based on pool area 2
- 30 MW/m A
- water pool cross sectional area where explosive interaction p9g 2
occurs (m )
h
- latent heat of vaporization of water g
- 2.25 MJ/kg The pressure increase (AP) due to this rapid steam generation, using the ideal gas law, is given by RT E at (3-2)
AP N 0 v s 2
where i
R
- universal gas constant - 8314 J/kg-mol *K s
V
- total containment free volume (a )
b0 - molecular weight of water - 18 kg/kg mol 2
at
- explosive interaction time (s) r
3-3
. sm, N,)
3.2.2 Shock Waves l
l The 1/r2 decay of shock waves from the interaction zone, through air to the containment boundary was described in Section 2.4.4.
The two basic parameters are (1) the 10 MPa maximum 2ttainable pressure (PIZ) in the interaction zone, and (2) the dimensions of the interaction zone.
Assuming the interaction zone to be a sphere, we can find the radius of the sphere that contains steam equivalent to the amount generated at the rate a, during the interaction time period at (typically on the order of milliseconds) as S
8 E
wR:A. b0 P (3,3) 3 I
77 2
or 1/3
,a At RT
~
IZ O
O IZ 2
The 1/r2 decay law, then, gives the impact pressure at the containment wall j
as
'R P
P X,
yg (3-5).
~
ew where P,- impact pressure at containment wall
(
X,, - distance from center of interaction zone to containment wall.
t 1
t
4-1
.,m 4.0 FIANT SPECIFIC APPLICATIQE 4.1 Issues 4.1.1 In-Vessel Steam Explosions According to conclusions made in Section 3.1, there are no conditions which could lead to vessel rupture due to an in-vessel steam explosion.
Consequently, in-vessel steam explosions leading to containment failure are not included as a node on the Farley plant response trees.
4.1.2 Ex-Vessel Steam Explosions l
4.1.2.1 Pressure Rise Due to Ranid Steam Generation In order to evaluate the pressure rise from Equation (3-2), the inter-action pool area (Ap997) and the interaction time (at) must be predetermined subj ect to some uncertainty.
However, even when conservative values for.
A,,y and at are used, this still results in an insignificant pressure rise p
that will not threaten the containment integrity. As mentioned in Section 2.3.5, most interaction times observed in experiments were about 1 second.
The plant design features would limit the corium dispersal area during the high pressure vessel blowdown to no more than one fourth of the lower com-partment floor.
If this interaction is assumed to occur over 25% of the entire lower compartment floor due to corium entrainment following the high pressure vessel blowdown, the steam generation rate would be 1300 kg/s and the pressure rise within the containment would be only 0.004 MPa (0.6 psi).
The above calculation uses A,,7 - 100 m 2 2
(1057 ft ), V - 58500 as (2.066 x p
s 10 fe ) [FAI, 1992), T - 390 K and at 1 second.
To accommodate the uncertainty in at due to a longer vessel blowdown time, at may be increased I
as part of a sensitivity study.
Even if At - 10 seconds is assumed, the predicted pressure rise would not be greater than 0.04 MPa (6 psi).
- Thus, ex-vessel steam explosions would not challenge containment integrity by overpressure.
O
i 42 l
I
\\
C 4.1.2.2 Shock Wave Imoact
)
The possibility of ex-vessel steam explosion inside the cavity of Farley is ruled out because of the 16 ft tall " curb" formed by the instru-ment tunnel vall below the seal table.
This prevents water from flowing from the lower compartment floor into the cavity, thus keeping the cavity dry during any sequence until the reactor fails.
The possibility of a steam explosion due to the interaction between debris particles entrained during the high pressure vessel blowdown (from the dry cavity) and the water pool on the lower compartment floor is un-likely. The design of the Farley cavity and the instrumentation tunnel will minimize the entrainment of debris from the cavity to the lower compartment.
Debris from the cavity must make two right angle turns and collide with the secondary shield wall in order to reach the water pool on the lower compart-ment floor.
Most debris will not make its ve)- to the lower compartment floor.
Heat loss from the debris to walls during the entrainment process pg will also reduce the debris temperature and, therefore, reduce the O
likelihood of a steam explosion.
If a steam explosion is assumed to occur, it will most likely be in the region of the lower compartment between the instrumentation tunnel curb and the secondary shield wall.
The shock wave that would result from the explo-sion will strike the secondary shield wall first and therefore will not directly impact on the containment wall.
The impact of a steam explosion in the lower compartment can be as-sessed as follows.
Using the same steam generation rate as when evaluating the pressure rise due to rapid steam generation (i.e., In, - 1300 kg/s), with 0.1
- second, P
dt 10 Ma (l W pW aM T 580 K ( m 7) 77 (corresponding saturation temperature), Equation (3-4) yields R OM m yg (3.1 ft).
With X,, - 7.8 m (25.5 f t), approximately the distance from the midpoint between the curb below the seal table and the secondary shield wall to the containment wall, Equation (3-5) gives P
- 1.33 x 105 Pa (19.3 ew psi).
Therefore, the pressure load would be much less than the containment
,m ultimate capacity.
Note that this calculation does not even consider the
,v)
4-3 (m) breakup of a shock wave due to the secondary shield wall and still shows V
only a small impact on the containment wall.
4.1.3 Uncertainty Considerations As discussed in this document, in [NRC, 1985) and in [IDCOR, 1983), a specific chain of events must occur before an in vessel explosive interac-tion coul d challenge the RPV integrity.
The failure of any link in this chain would prevent interactions which could challenge RPV integrity.
Evaluationa cf each link in the chain concludes that the only one which-could be achieved is an explosive event, but then only when the RPV pressure is very low, i.e.
typically less than 300 psia (2 MPa).
Even if an explo-sion would occur, the mass of material, the efficiency, the slug formation and the slug transmission could not be realized to any significant degree.
Given (1) the extent of individual analyses performed on each link in the chain of events, (2) the extent to which each is not satisfied in a realis-tic analysis and (3) the imrartance that each be satisfied before the RPV integrity could be challenged, it is concluded that there is no realistic combination of uncertainties which could make such an event credible.
Ex-vessel explosive interactions could occur for sequences which would progress to vessel failure.
Here again, the extent of the interaction (and damage potential) is determined by the mass of molten material involved, the efficiency of the interaction and the decay of the shock waves as they propagate from the source.
In this evaluation we have assumed.
an efficient interaction since the shock waves are calculated to decay as for TNT, and
. an unimpeded expansion of the shock waves to the containment bound-ary.
These global representations are sufficient to determine if ex vessel explo-sions could challenge containment integrity.
- However, best estimate analyses would consider:
O
o 44 O
a much smaller interaction zone, e
'V s
much weaker shock waves for steam explosions than for chemical explo.
sions, and
+ the breakup of shock waves by the structures.in the lower containment compartment, in particular the concrete shield walls separating the lower. compartment from the annular compartment.
With the conservative assessments used for these individual evaluations, it is concluded there is no realistic combination of uncertainties which could result in a credible threat to containment integrity.
4.2 Conclusions There is no credible set of circumstances in which ex-vessel steam explosions could challenge the containment integrity.
The re fore,
ex-vessel steam explosions are not included as a node in the Farley IPE plant response trees.
I O
f
5-1
~
(,)
5.0
SUMMARY
The influence of in vessel and ex-vessel steam explosion events on the potential for containment failure has been addressed for the accident condi-tions of interewt.
The evaluations for in-vessel events closely parallel those performed as part of the IDCOR program and result in a conclusion that the slumping of molten debris into the RPV lower plenum could not result in sufficient energy release due to a steam explosion to threaten the vessel integrity.
- Hence, such interactions could not lead directly to containment failure (alpha mode) and a consequential release of fission products to the environ-ment.
This is in concert with the conclusion of the NRC sponsored Steam Explosion Review Group.
Evaluations of both steam generation rate and shock waves induced by ex-vessel explosive interactions show that these would not be of sufficient strength to threaten the containment integrity.
While such explosions could
(_/
occur, the principal result of these events would be to rapidly cool debris and slightly pressurize the containment.
Neither of these would be suffi-cient to challenge the containment integrity for any realistic accident conditions.
N As a result of the evaluations performed for both in-vessel and ex-vessel events and the conclusion that these would not lead to challenges of the containment integrity, steam explosions are not included in the Farley IPE plant response trees.
O,U
6-1
6.0 REFERENCES
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and Benedick, W.
B.,
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i
- Hohmann, H.,
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l O
i
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(_
- IDCOR, 1983,
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Lemmon, A. W., 1980,
" ?.x p l o s j o n s of Molten Aluminum and Water",
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Philadelphia, PA, AIChE Symposium Series, Vol. 85, No. 269, pp. 217-222.
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B, W.,
- 1986,
" Hydrogen: Air: Steam Flammability Limits and Combustion Characteristics in the FITS Vessel",
NUREG/CR-3468, SAND 84-0383.
(m)
Miller, R. W., Sola, A and McCardell, R. K., 1964, " Report of the SPERT 1
\\/
Destructive Test Program on an Aluminum, Plate-Type, Water Moderated Reactor", IDO-16883.
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Corradini, M. L. and Tarbell, W. W.,1981, " Intermediate Scale Steam Explosion Phenomena:
Experiments and Analysis",
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(1986),
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FITSB Series", NUREC/CR-3983, SAND 83-1057.
NRC, 1975, " Reactor Safety Study," WASH-1400, NUREG/75-0114.
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" Individual Plant Examination for Severe Accident Vulnerabilities 10CFR50.54(f)",
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" Concentration-Gradient Trigger Mechanism for. Smelt-Water Explosions", Paper presented at the American Paper Institute Annual Recovery Boiler Committee Mtg., Chicago, IL, 30-31.
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n
~
f
'63
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L. C. and Zivi, S.
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l l
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i 4
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