ML20138C220
| ML20138C220 | |
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| Site: | 05200003 |
| Issue date: | 03/12/1996 |
| From: | Berthoud G FRANCE |
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| NUDOCS 9704300070 | |
| Download: ML20138C220 (7) | |
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ANL-RE -~S. s0RRELL a',0 0 3 1F06/96 FRI 16:s4 FAI 1 630 2s2 4790 Comments about the document a Lowtr Head Integrity undtr In vessel Steam Explosion Loads e by T.G. Theofanous et al G. Berthoud - CEA/Grenoble This document presents an analysis of the potentiality oflower head failure of the AP600 resu from a Steam Explosion. The conclusion that the risk is negligible (a physically unreasonable quite convincing and is based on:
- 1. the fact that water will be saturated and at I bar due to complete depressurisation to th containment pressure and that these conditions will lead to large and rapid voiding w favorable for large S. E.
- 2. the fact that we have permanent blockages at the bottom of the core that will impeed a relocation through the core support plate
- 3. the fact that relocation will occur sideways through the reflector and core barrel and Steam Explosion will occur in a 3D geometry without any large constraint allowing sustained pressure
- 4. the fact that - even if reflooding is taken into account when the melt will be ejected sidewa will have enough time to heat the added water up to saturation and so to prevent goo The validity of the conclusion is then linked to the validity of the above four arguments.
As for the first argument,there is no doubt that water will be saturated as far as refloodl taken into account, the fact that the pressure will be atmospheric cannot be discusl justified in another report (IVR Report - table 7.3) however, I think that this h !
voiding will be less important at pressures a little bit hig droplet experiments of Nelson in Sandia. I l
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r The fact what we have up to now no evidence of explosion in experiments using reactor like l
i materials (Krotos) (and that this is due to the non occurrence of good mixing) is stressed by the l
authors. But once again these Krotos experiments are performed at i bar pressure while in Faro experiments at pressures of 50 and 20 bars, with saturated water some mixing was obtained. In near future a Faro experiment us ng initially saturated water at 5 bar will be performed and we will I
then have an indication of the quality of mixing at small pressure. )
I will now go through the different chapters trying to analyze the justifications which are presen to support the crucial arguments mentioned previously.
Chapter 2: Problem definition and over all approach
. It is mentioned that it is only recently that pressures in the kbar range were observed experimentally in constrained one dimensional geometry. However, I think that a press .l the order of the kbar amplitude and millisecond duration was meuured in the Sandia FITS-RC2 experiment which was well vented (initially open at the top and later vented at the bottom vessel left the ground). But this was obtained using iron-alumina thermite and subcooled water.
. Anotner important argument is that u because of extensive voiding, we need only be conce about the first relocation event and only for early trigger in it n. This seems to bejusti6ed by the ,
premixing and explosion calculations presented later but I wonder why, after a first e water is sloshing back a second event cannot occur at about the same location where the structur will has been already dynamically loaded and eventually already deformed by the first event.
Chapter 3: Structural failure criteria In this chapter, it is stated that a the time duration of the loads of interest here is less tha structural frequency n, so it is expected that a peak strain would be basically independ of the pressure pulse shape n. However, nothing is said about the estimated value o frequency of the R.P.V. which seems to me to be of the order of magnitude of some msec so far from the load duration.
However, all the analysis is made with the analytical solution of Dufley and Mitcheli which a short pressure pulse 9 and allows to evaluate the plastic equivalent strain with in strain rate effect by formula (3.2). But the comparison of the analytical results to ABAQUS calculations shows that the analytical relation gives conservative results for the plastic stra evaluation (fig. 3.2).
Another mitigating factor is investigated: the effect ofload localization which shows that impulse, the equivalent plastic strain is smaller when the loading is smaller. Use o then made by assuming that a fraction B of the impulse is used for bending energy so thi
- effective impulse n is applied for the evaluation of the equivalent plastic strain. We is a material and geometric a constant a but I have not found any indication ofits evaluatio the results shown on fig. 3.8 and fig. 3.9 are obtained. As fig. 3.9 is used to evaluate the loads calculated in Chapter 6, I think that it should be a little more explained. ,
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1*:/06/96 FRI 16:5s FA.t 1 630 352 47s0 ANI. RE m S. SoRRELL 200s It also.seems to me that the localized loadings are applied on the axis of the hemisphere (see table 1 3.2); does the fact that these localized loadings will occur on the side of the hemisphere with
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singularity where the sphere is linked to the cylinder will modify the conclusions we can draw from fig. 3.9.
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N.B. There are some errors in table 3.2 as for the value of Ao which does not correspond to 4 I 1
as written in the caption Chapter 4: Quantification of melt relocation characteristics This is a very important task as most of the boundary conditions for premixing and explosion calculations are obtained from such an evaluation.
- The downward relocation path (arguments 2) is not envisaged: a we expect this path to be blocked by molten cladding and the blockage be robust n. This a expert judgment a is supported by the large heat sink associated with the large amount of < cold a materials in the lower part of the core. As it is said that due to be big stainless steel ref!cctor, the first relocation will be delayed compared to what occurred in TIM and that at this time, we will have a large oxide pool, it is important to know if this molten pool will reach the region of the lower fission gas polonium where the heat sink is not very large and where we can have a breakdown of the supporting r
material. However, in the paper, the blockage is said to occur in the region of the 7 cm a lower Zr
. plug and lowermost spacer grid m. Some calculations are presented to show that the pluggi of this region by melt with negligible superheat is of the order of seconds. For this calculation I have some trouble with formula 4.2 where, as for me, A is not the same as in the Carslaw and Jaegger text book but I did not try to perform the calculation. We can also make another rem if we have some breaking down of the fission gaz plenum region, when the molten pool arrives we may have superheated molten material from this pool that with flow in the lower blockage re for some times before plugging. It would be interesting to know what amount of molten materials i
can be transfered in the lower plenum through the holes in the core support plate before plugging of the passages in the blockage region. As for this plugging time- which is enacial to support argument 2 - it would be interesting to see more realistic calculations including the influe the interface thermal resistance between the crust and the solid wall that will slow freezing process and then increase the plugging time.
!
- As for the blockage coolability:
- the stable blockage thickness should be sent.ible to the radiation factor fr which is set to 0.7 without any explanations
- the cooling of this blockage is ensured for about 100 mn which is the time required to vapo the water which fills all the volume between bottom of active core and bottom of core support L
plate. It is later estimated that meltthrough of the reflector by the molten oxidic pool wil between 76 and 91 mm according to the amount of oxidation (80 to 95 mm in the calculation a without a preheating). If we add the time require to melt through the core barrel, we get timing of the release of the same order of magnitude than the insurrance of blockage co As all these calculations are order of magnitude ones, I think that argument 2 (no downward l relocation path) may be questioned.
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A.NL-PI .b. S.' SoRRELL 4 000 1**/06/96 FRI 16:36 FA.I L 630 2s2 4780
- Molten pool formation I
- In the initial heat up calculation, are the retlector and core barrel in contact everywhere as it is shown fig. 4.8 and 4.97 In that case the cooling effect will be overestimated and the melt superheat underestimated.
2
- During the transient heat up calculation, what happens to '.he molten cladding and how the calculation with the effective thermal conductivity is enventually modified?
e Molten pool calculation Such a calculation is performed for the oxidic and metallic pool, and there is a crucial l
1 hypothesis which is the presence of a stable oxidic crust at the upper surface of the oxidic 2 In the document, it is mentionned that it is assumed that the clad drains but is it fully true?
j Cannot we have some metals included in the moving down oxidic pool? What happen to the
' pan of fission products which are released at fuel melting? Will they modify the molten behaviour for the stability of the upper crust and the evaluation of the differents fluxes?
l 4
- Melt through and melt release calculations It is said that rapidly the metallic pool will melt through the reflector but it is assumed that the 4
metal u will be gradually draining a into the space between the flats of reflector and the core barrel. Cannot we have some kind of metallic jet impacting on the core barrel with some rapid L
' meltthrough leading to a steam explosion between metals and water?
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- From the above analysis it is concluded that when the oxidic melts through the reflector, ther no metal on it and that failure of the core barrel occurs soon after. First, it would be interesting to evaluate the time required for core barrel meltthrough (if there is an open space between i- two of them).
But there is another problem if the space between the flats of the reflector and the core ba already filled with the metals from the metallic layer, how the oxidic pool can rapidly go
~t he core barrel. This situation may be a promoter for downward a:ation if this added metal may increase the time for meltthrough.
(
- As for the location and size of the failure, most of the information is obtained from expert judgment and should be fbrtherjustified:
- the failure a is eapected a to be local azimuthally and very near the top of the oxidic poo would agree with that statement as even, if the calculation is 2D cylindrical, once a flat w l
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, LW O6 96 FRI 16:57 FAI t 030 252 4730 the rapid relocation will impeed failure on other Gats. But I would not be able to give any probability for 2 quasi simultaneous failure, or 3..
. for the size of the breach, it is said that 0.4 m u would appear geometrically a good upper bound on the first breach width n and that a u 10 cm axial gap is believed to be conservative n.
- there is no mention of the rapid increase of the size of the breach during the melt release as it has been observed in experiments. However, as only short duration premixing-triggering scenario are taken to be ofinterest, this enlargment would not be important. But, if we take into account steam explosion occuring when water is sloshing back after a first event, this has to be taken into account.
Chapter 5: Quantification of Premixtures Given the melt release conditions (flow rate, location, temperature and composition), the premixtures are calculated with an improved version of PM-ALPHA which is now 3D and inclu melt fragmentation law (which was lacking up to now) as it is recognized that it is interesting a to know the distributions of the melt length scale n. However, this fragmentation law is not described and this should be donc and justified as fragmentation is responsible for voiding (a the rate of voiding increasing rapidly with the rate of breakup n). I would also like to know why the a breaking law is operative only for u long as the coolant has a void fraction ofless than 50% n. If the fragment was always operative, voiding would be larger so there must be a good reason for doing so but I do not see why. .
The melt entrance conditions into water are also specified and not calculated:
- entrance velocity whose evaluation is correct
- distribution of the melt u over an effective radial width of 10 cm n with a melt volum evaluated to get the correct mass flow rate. This distribution is crucial in determining the amou vapour which is produced as the larger the entrance area, the longer you are in the film boil regime in which the steam production is at maximum. This behaviour was observed in MC3D recalculations of FARO tests, where a doubling of the pressure increase (linked with vapour production) was obtained with a doubling of the diameter of the melt flow. Recently CHYNE ,
recalculations showed the same trend.
- initial droplet diameter which is set to 20 mm (a large enough value to represent a minimally up melt stream). This parameter is also important for vapour production. It would be intere see sensitivity calculations with diameters varying from 10 to I cm.
I am not so sure that the melt will be transformed in a droplet population before entering the water.
We may have a large melt stream on the wall with subsequent fragmentation into the wate a different law than droplet fragmentation. Would it lead to a u benign evolution n as it is mentionned. This is again an expert judgment.
As for jet fragmentation calculations with THIRMAL, I cannot trust them if the fragmentation governed by Kelvin Helmholtz type calculations. Moreover,in FARO experiments with 10 cm jet, it took more than 2 m of water to break the jets in a 50 bar atmosphere for which void smaller, l
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Chapter 6: Quantiikation of explosion loads Nothing is said : bout the parameters used in ESPROSE-m 3D but as the trigger uses a 100 bar
! steam release, we may think that the hydrodynamics fragmentation law will be correct. Due to the small amount of melt involved in explosion calculations, there is no probfem with the energetics of the explosion and we are only interested in dynamical loadings of the RPV. This is done by the estimation of the impulse and of the local area ofloadings from ESPROSE-m results. I have some d is estimated page 6.3 from the area evolution as shown on fig.
problems to understand how
- DS 6.5.a.
In the text, it is said that peak impulses are around 0.1 and 0.2 MPa.s with effective area around 0.1 l
m2 (which gives g- 0.15) and from fig. 6.5 e where I find a 0.2 Mpa.s impulse I dol 2
understand how I get Ao - 0.1 m from the area evolution which is shown.
i Apart from this problem, ifI accept the figures mentionned in the text, I agree when coming b l
! figure 3.9 that there is no risk for I ~ 0.2DSand 10 ~ 0.15 e as '3D - 0. This is confirmed by th ABAQUS calculations of the two most energetic explosion calculations obtained by PM-ALPHA l
j plus ESPROSE-m.
i.
As a conclusion, I can say that if we accept the scenario which is retained by the authors, I think th:
j
- whatever my remarks about premixing quantifications - the AP 600 RPV cannot be challenge ,
L steam explosion. However, I would like to have more established confirmations of this s
- mechanistic calculations when possible or parametric calculations when it is not. J
]
The main thing to be confirmed is the impossibility to have a downward relocation i.e )
l
- the possibility to have a break down of the lower fission gas plenum rather than a contin
) . draining. This will give a sudden access of the core support plate holes to the melt f
- 1 l
- the influence of the already relocated metallic pool on the oxidic release. It may take a longer ti l l to break through and the blockage integrity may then be challenged. The influence of an resistance between the oxidic solid crust and the wall - specially at the top of the pool - will als
)
participate to an increase of the time of break through and of the evaluation of melt su Other branches for the scenario should also be evaluated:
- the possiblity of the metallic melt to rapidly go through the core barrel leading to metal-w l steam explosions the possibility to have steam explosion at later times (in the oxidic case) when wate back aner a first small scale (i.e low energetic) event.
As for reflooding scenario, the fact that water will be closed to saturation should also be eval r *
- '0 ' ---- - . _ . _
M L-RE *** 5. SORRELL 4.002 A'2 /13/ 96 THC 15:54 F42 1 630 252 4730 .
TECHNISCHE UNIVERSITkT MONCHEN LEHRSTUHL A FOR THERMODYNAMIK Prof. DrAng. Dr..Ing. E.h. E Mayinger 80333 Munden AreswoGe 21 ihwmod f em.'s A . rechoshe Un,vme Menden 80290 MGnchen TEL lC69) 2l05 3435 2l05 3436 Director Ta n 522s54 % d of Reactor Engineering Division %c89) 2105 u5t Dr. L. W. Deitrich ""
Argonne National Laboratory 9700 South Casa Avenue Neue Telefonnummer
] j" IIII" i' 8043' (089)289-16215 (089) 289-16218 Fax Unw 0*m
__ Unw Zeichen komm the Zei6.n 03.12.1996 M/ba Review on the report 00E/1D-10541
" Lower Head Integrity under in-Vessel Steam Explosion Loads" l l
Dear Dr. Deitrich,
hich is titled Please find e'acloud rny review comments on the report DOE /ID-10l
" Lower Head integrity undst in-Vessel Steam Explosion Loads". l l
Sincerely yours,
~ -
k w N~I, REACTOR ENGiMEERING DN RECEIVED F, _/
-DIRECTO. T 3 07F CI-ggg ,
Prof. Dr.-Ing. Dr. Ing.E.h. F. Mayinger i ACT:Ct4: _ _ . , ,
p.- ca.iarTE- L (4NJA,,.M ~13M G _.
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