ML20069Q173
| ML20069Q173 | |
| Person / Time | |
|---|---|
| Site: | Point Beach  |
| Issue date: | 12/03/1982 |
| From: | Boccio J BROOKHAVEN NATIONAL LABORATORY |
| To: | Eberly R Office of Nuclear Reactor Regulation |
| Shared Package | |
| ML20069Q176 | List: |
| References | |
| CON-FIN-A-3703 TAC-11079, TAC-11080, NUDOCS 8212080481 | |
| Download: ML20069Q173 (23) | |
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/>SSOCIATED UNIVERSITIES, INC.
A:
- a. :.2 U Upton Long Island. New York 11973 (516) 282s Departrnent of Nuctect Energy FTS 666' -7690 December 3, 1982 Randall Eberly O.S. Nuclear Regulatory Commission Chemical Engineering Branch - Mail Stop P-302 Washington, D.C. 20555 Re: Fire & Probabilistic Model Reviews - FIN A-3703
Dear Randall:
Enclosed is our review of Wisconsin Electric Power Company's (WEP) fire modeling and analyses used to justify exemption requests of their Point Beacn facility to 10 CFR 50, Appe'ndix R.
If you have any questions regarding this subject matter or our previous four reviews, please feel free to call gither me or Ors. Charles Ruger (FTS 666-2107).
Yours truly, q
John L. Boccio, Group Leader Reliability & Physicai Analysis JLB/sm Enc.
cc: R. Bari w/o enc R. Ferguson, NRC, w/ enc R. Hall w/o enc-W. Kato
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EVALUAT10N OF THE A:.ALYTICAL FIRE i*.0DELING BY THE WISCON51N ELECTRIC POWER CCMPANY (UEP)
IN THEIR JUNE 1982 REPORT
" POINT BEACH NUCLEAR PLANT, UNITS 182, RESPONSE TO 10 CFR 50, APPENDIX R FIRE PROTECTION OF SAFE SHUTDOWN CAPABILITY" Charles J. Ruger Department of Nucl. ear Energy Brookhaven National Laboratory Upton, N.Y. 11973 1.
INTP0 DUCTION This report contains our evaluation of the fire-modeling methodology em-ployed by the Wisconsin Electric Power Company (WEP) in their June 1982 re-port, " Point Beach Nuclear Plant, Units 1&2, Response to 10 CFR 50, Appendix R Fire Protection of Safe Shutdown Capability." As an alternative to the requirements specified in Section III.G of Appendix R to 10CFR50, WEP puroorts
- to provide analyses that justify exemption from these requis ements in particular plant fire areas.
Br efly, the general approach take'n by the licensee in this regard is to i
calculate the energy needed to damage redundant cables in a given plant area employing conservative assumptions in the attendant model, and then to calcu-late the minimum amount of combustibles that would be necessary to provide i
such energy, also employing in the analysis a set of conservative assumptions.
The underlying thesis is to demonstrate that, regardless of what administra-tive controls are assumed, the amount and type of combustibles, as determined via analysis and/or heuristic arguments, that are necessary to damage the re-quisite cables will simply not be found in the plant area under investigation.
A'nore detailed description of the WEP approac.h is contained herein.
In this connection, the overall scope of our evaluation is to assess that (1) the method employed is technically sound: (2) the overall approach will yield realistic or conservative results; rqd (3) the end use of the results is
- valid, i
We start our detailed review of the reference' submittal by "first de-l
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scribing in more depth than above the fire modeling process employed by WEP.
This is followed by some of our general thoughts on the complexity of the fire-phenomena modeling and some key items we consider as forming the foundation of our appraisal.
Sections 4 & 5 give our overall evaluation of the WEP approach based upon a detailed critique, which is provided.
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St".'wy 0F THE WEP FIRE E0h.P.
w ess The general approach taken by '.Ep is to identify the minimum quantity and gecmetry of liquid hydrocarbon spill wnich would exceed the damage criteria for the electrical cables of interest.
This is accomplished in the following manner:
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(1) Identify the electrical cables of interest, their specifi' cations, geometry, and the dimensions of the plant area.
(2) Identify the fixed and transient l'iquid hydrocarbon materials of con-cern.
(3) Calculate the minimum quantity of the fuels of interest and the asso-ciated fire geometry (location, area, and depth) necessary to exceed the damageability criteria for the identified electrical cable through the following mechanisms:
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a) Stratification b) Radiation c) Buoyant diffusion plume impingement For the purposes of analysis, the effects of actual room geometry, floor slope, and equipment layout' are ignorgd.and the presence of a perfectly horizontal floor, free of fire inhibiting equipment, is assumed. Also, the ef-fects of pipes and ventilation systems in diverting the flow of hot gases, abscrbing incident heat flux, or blocking the free passage of radiation to the cables of interest, is ignored.
4 The objective of the analysis is to demonstrate the equivalent protection i
of plant passive fire protection measures alone to that protection afforded by Appendix R.
Thus, wherever possible, the process so described ignores assump-tions regarding " credible" quantities of transi.ent combustibles or the value of, administrative controls and attempts to present fire protection in terms of quantities of different fluids.
The basic fire models used are presented in Appendices A.1 to A.8 of the submittal.
Indluded therein are data on heat release rates and descriptions of the mathematical models employed for calculating the ceiling layer heat flux, buoyant diffusion plume growth, thermal radiative heat flux, a method for determining the size of thermal shields, a heat conduction model, a model for heat transfer inside a cabinet, and a switch radiation model.
Section 4 of the submittal provides a general discussion of the methodology used to support the exemption request.
For each fire area identified as not being in compliance with Section III.G.2 of Appendix R, a fire hazards analysis is contained in. Section 5 of the sub;nittal.
The discussions provided ir these
.two sections, along with each of the Appendices, comprises the scope of our review.
The following section dese ibes the BHL review philosophy.
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- _ iC %t RE*.'IEL FHILOSOPHY For our appraisal, some generai thougnts are deened warranted on the com-plexity of fire phenomena and the state of fire science with regard to en-closure fire development.
Computer models of enclosure fire development appear capablA of predicting quantities of practical importance to fire safety, provided the model is sup-plied with the fire-initiating item's empirical rate of fire growth a'nd the effect of external radiation on this rate. ' As a science, however, we cannot predict the initiating' 11iem's growth rate due to relatively poor understanding of basic combustion mechanisms.
Questions and doubts have even been raised regarding the ability to predict the bur-ing rate of a non-spreading, hazardous scale fire in terms of basic measurable fuel properties.
- However, while awaiting development of meaningful standard flammability tests and/or-more sound scientific predictions, realistic " standardized" fire test procedures should continue to be formulated for empirically measuring the rate of growth of isolated init.ating items, the attendant fire plume,*its develepment within an enclosure, and the convective and radiative heat loads to " target' combustibles.
Thus, in lieu of large-scale computer codes.to assess the firs hazard in an onclosure, we define " state-of-the-art" for the purposes of this evaluation as one which incorporates a unit-problem approach to seven general components of the fire considered relevant in understandiq, at least on heuristic principles and pracmatic efforts, the phenomena of fire.
The following list may b4 obvious, but*, 'ih the framework of this unit-probleni approach, how one considers the complex heat flux and material flux interactions within the fire-modeling methodology forms the general basis for our appraisal.
The seven components and the various,important interactions are:
The burning object receives rediative and convective heat.from the com-e busting plume and radiative heat from the hot ceiling layer and pos- -
sibly the ceiling.
The' combusting plume (or flame) receives volatile species from the e
j burning object.
It receives air (which may be preheated and vitiated ir. Oxygen) from the cold layer.
When the upper point of the flame ex-tends into the hot layer, overall burning may be modif'ad. Room geome-try, non-combusting obstacles, and burning object location influence pitme development.
The hot layer will be influenced by natural and forced ventilation, by e
the heat and gas combustion products produced by the flame, and by heat losses to the enclosure walls, ceiling, and other objects.
Al so,
transient combustion within the hot cejling layer has been observed
- d may be considered an interaction with the flame.
Transient com-bustion in the hot layer could be due. to excess pyrolyzate from the burning object (both solid firebrands and gaseous inconplete products of combustion).
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The tarcets are heated by radiation (and also convection for an upward e
spreading fire), coning from the combusting plume, the hot layer, and possibly the ceiling (if the hot layer is transparent to ra'diation).
Ignition of a target increases the cverall thernal energy content within the enclosure.
lhe enclosure geometry (ceiling and walls) is heated by convection and i
e rcdiation,from all burning objects, and the hot ceiling layer.
5 e The vents influence the mass flow rate of oxidizer and the radiative and convective components of thermal energy loss.
Positive feedback is a critical part of the fire growth phenomenon and its accountability within the licensee's submittal has,also been a factor in our evaluation.
(Granted, each form of interaction has a characteristic time or physical dimension associated with it, which would provide a measure of its relativeimportance.) A matrix of the more important items, which we fee! are crucial for subsequent discussion in the licensee submittal, is provided in-Appendix A.
4.
StudARY EVAltJATION OF THE WEP APPROACM As a concept, the overall methodology repre.sents, in part, a technically sound and conservative technique for assessing the potential hazard presented by exposure fires to electrical cables.
The modeling tools used in assessing the relative value of existing separ-ation, afforded by the plant configuration in passively protecting plant safe-shutdown systems from the effects of exposure fires, consists in employire the following unit-models:
e ' pool-fire plume model pool-fire induced stratification model e
e pool-fire radiation caodel j
e fire-induced electrical cable damage criterion l
e thermal shield analysis e finite element heat conduction model thermal analysts of cabinet / panel internals
'o e
panel switch radiation model e
The u *t-pr lem approach employed, together with the correlations and electrical cable damage criterion, can be classified as most current and
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methodologically consistent with what-is being suggested in the open litera-ture.as a viable approach for ass _essing the fir.e haza.rd_pote_ntial associated 4
with cable tray fires.
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.:, in..cs; re -;ts, oa finc ::..; method e..p;oyed to be technically w..od and the overall apprcach, if applied properly (as described subsequent-ly) cc';1d yield realistic and conservative results for asessing the thermal environment in the fire area.
However we question the validity of the concept as applied in demonstrating the equivalence of the protection provided with the requirements of Appendix R, Section Ill.G.2.
This is based upon the following general observations:
(1) The use of an electrical damage criterion, in conjunction with the stratification model described, is not valid because the model pro-vides a correlation that is based only on the consideration of the effect of a single exposure fire on the ensuing thermal energy content within the enclosure.
Accordingly, the model/ damage criterion is not uniformly valid when. cables, either in the fire plume or in the drat-ified layer, are in the process of burning, thereby adding thermal energy to the enclosure.
To be consistent with the experiments conducted to establish the stra-tification model, the model/ damage criteri.n could only be considered valid when piloted-ignition, in lieu of electrical failure damage cri -
teria, is employed.
Establishing a time for piloted ignition would be such that the additional heat released by the onset of cable ignition would be small compared to the exposure fire, thereby making the stra.
tification model valid within the tiine frame.
l On the other hand, when the damaged cables are concletely enclosed in conduit, the electrical damage criterion may be sufficient.
(2) The above observation notwithstanding, the electrical failure tests that form the basis for the damage criterion employed were obtained from test observations on the short circuiting of a 70V signal.
Volt-ages in plant cables < ould be much higher than this and could concef v-ably cacse earlier damage than indie.ated'by the experimental tests.
'(3) An intrinsic limitation of the stratification model in attempting to l
show equivalency in protection provided is the independancy of the l
correlation to lateral separation distance.
In effect, the model would show that the local thermal environment to redundant horizontal l
cable trays, situated within the stratified layer at the same height l
above the floor, would be identical, regardless of the horizontal sepa-l ration between each tray, all other pertinent data being equal.
t (4) Neither the models employed, nor the methoctiogy used, c.onsider the in-creased heat flux that exposure fires can generate when located near walls and corners.
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Only liqu:c,... n 111s are ccasidered.
Tne possibility of excess pyrolytate resulting f rom insulation degradation or from initiating fires resulting from the burning of solid combustibles, which could enter into and subsequently burn within the stratified layer, has not been investigated.
.s (6) Errors in the data listed, needed in establishing the ha'zards as-sociated with high fire-point liquid hydrocarbons, provides signific-ant doubts when used with the analyses described, as to conclusions drawn that such liquid spills do'not present a significant fire hazard when spilled on concrete.
(7) Fires initiated at locations other than on the flour t. ave not been addressed.,
(8) The non-linear optimization methodology used to determine the minimum amount of liquid fuel required to cause electrical damage to both re-dundant and safe-shutdown systems is not presentad.in sufficient de-tail to allow for audit calculations or appraisal.
(9) The Rayleigh numbers of the postulated fires are far beyond the range for which the plume impingement model is valid.
(10) An error has been' found on the.. thermal shield analysis, which, if corrected, would alter the limits placed on the wake velocity and temperature defects incorporated ir establishing the size of shield required for protecting cables imnersed within the fire plume.
(11) It is not clear wnich radiation heat transfer model is used in the' a 11ysis or from where the configuration factor is obtained.
5.
DETAILED EVALUATION OF THE WEP APPROACH The basic fire models are presented in Appendices A.1 to A.8 of the sub- ' - - - -
mittal. These appendices include data on heat release r.tes and models for-ceiling layer heat flux, buoyant diffusion plumes, thermal radiati]n, a method for determining the size of thermal shields, heat conduction, internal cabinet heat transfer, and cabinet switch radiation.
Section 4 of the submittal pro-vides a general discussion of the methodology used for the exposure fire The fire hazards analysis of analyses which support the exemption requests.
each fire area identified as not being in compliance with Section III.G.2 of Appendix R is contained in Section 5 of the submittal.
These sections are now discussed further with regard to modeling, assumption uncertainties, and ap-plication of the methodology.
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Appendix A.1 of the submittal describes a basis for selecting liquid hydro-carbon heat release rates, based on the current state of knowledge in fire sc ances.
Values of the heat of combustion, vaporization rate, and heat re-lease rate, are given for acetone, lubricating oil and heptane.
The assump-tion that ventilation is always sufficient to provide ideal fuel-oxygen ratios leads to the use of a conservative upper bound for the heat release rate.
Also, conservative ' asymptotic values (large scale fires) for steady-sta(e mass loss rate per unit area are used, i.e., the fire is assumed to reach steady-state conditions immediately.
The use of laboratory-scale generated, actual heat of combustion data by Tewarson is also conservative since the most ef-ficient combustion achievable in the laboratory is employed in the analysis.
5.2 Review cf Appendix A.2 Appendix A.2 of the submittal is based on the correlation of flewman and Hill l for the convective and radiative heat flux in the stratified ceiling hot gas layer developed by a pool fire within an enclosure.
The heat flux is related to the room's dimensions, the target height above the floor, the fuel's flammability parameters, and the room ventilation rate.
This correlation should be adequate for evaluating the heat flux dra to.
pool exposure fires.
However, it should be pointed out that one conclusicn reached froa, the data in Reference 1 and carried over into the correlation, namely that horizontal heat flux variations are minima {, is not in acreement with some other authors 2-5 In these references, data and theory,3 2
show that, for radial distances from the fire plume axis greater than ?dO". of the ceiling height, the heat flux decreases with radial distance to the -1/3 power.
However, in re-examination of Figure 7 of Reference 4, the heat flux appears to have a radici dependency to the -1.25 power.
This is shown in Figure 1 provided herein.
To Turther check this difference, we utilized the heat transfer coefficient parameter, he, presented by Veldman, et al (Re-ference 15) in their Figure 14.
This shows a radial dependercy for this para-meter to the -0. 6 power which, when applied to the -2/3 power correlation ~
presented by Alpert in Reference 2 for the maximum plume temperature differ-ence, AT, yields in concert a radial power law dependency of approximately
(-1.27), which is in close agreemeat with the -1.25 power indicated in Fig.1.
These works conside.- a quiescent enclosure while fleunan and' Hill include l
forced ventilation in most of their tests.
However, since tiewman and Hill's heat flux data for no ventilai. ion fall in the center of You and Faeth's data 4 for radial distances closer than 20". of the ceiling height (no radial' dependence), the neglect of the decrease in heat flux with radial distance by tiewman and Hill should yield a conservative result.
This also tends to show no benefit to horizontal cable separation for radial distances closer than 20".
of the ceiling height.
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4 On the ;*.n:r ha a, Referer.ces 2 2ra 5 show that if the exposure ih e is near a wall or in a corner, the ceiling temperatures increase as if the fire heat release rate is increased by a factor of 2 and 4 respectively.
There-fore, care must be taken in applying the !!eaman and Hill correlation for ex-posure fires in the vicinity of walls or corners so that non-cons,ervative re-suits are not obtained.
The submittal does not use the Newman and Hill correlation exactly as pro-sented in Reference 1.
Instead, a modified fem as given on page A.'2-4 is j
used. Apparently, this was done to extend the correlation at ventilation rates greater th.an those for which measurements were taken in Refe.rence 1.
This fact, coupled with the unrealistic cooling behavior of the original i
Newman and Hill correlation at higher ventilation rates as shown in Figure A.2-2, leads to the need for the modified correlation, which continues the data trend to higher values of ventilation.
This modified correlation is more conservative than the original.
Since the labeling of Figure A.2-2 is somewhat confusing, it is replotted as Figure 2 (attached) with the modified correlation on page A.2-4 included.
The correlation is not valid if secondary fires occur, or if excess pyrolyzates burn in the stratified layer.
5.3 Review of Accendix A.3 Appendix A.3 of the submittal dcscribes a turbulent, buoyant diffusion plume model which is essentially the ciudsical Morton-Taylor model.
The experiments of Stavrianidis6 are considered along with his correlations for critical height, (height to which plume correlations are valid), and virtual source height.
The heat flux correlations of You and Faeth4 for the stagnation region (r/H < 0.2) and the ceiling jet gre also presented.
The correlations are for Rayleigh numbers of 109 to 1024, whereas the fires 17 discussed in Section 5 of the submittal have Rayleigh numbers of about 10 There should be some defense of this extension.
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These represent state-of-the-art correlations for hydrocarbon pool-fire pl'umes.
However, there are several errors, most likely typographical, which in the buoyancy should be corrected.
First, the exponent of the factor Fa expressions on pages A.3-2 and A.3-3 should be 2/3 rather than 1/3.
A review of You and Faeth's work' yields the following comments concerning the heat flux correlation on pages A.3-8 and A.3-9 of the submittal. The Greek symbol v appearing in tN Rayleigh number is defined as the kinematic viscosity, not the radial velocity.
The heat flux correlation appearing on the bottom of page A.3-9 is valid in the ceiling jet, outside the stagnation regior (r/H >
C.2) for free-flame height to ceiling-height ratios up to 2.5, as evidenced by the data in Figure 7 of Reference 4.
The radial dependence in the correlation should be to the -1.25 poser a3 explained in the review of Appendix A.2.
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The radiant heat transfer frc= a high-temperature, turbulent, buoyant 'if-d fusion plume is discussed in Appendix A.4 of the submittal.
A classical ap-pro'ach based on the Stefan-Boltzmann law is used.
A uniform gaseous tempera-6 ture of 1255* K is assumed based on the work of Stavrianidis.
It is not clear which correlation for flame height is used, although Stavrianidis has a correlation for hydrgcarbon which is consistent with data.
However, passing mention of Steward's work is all that is found in this Appendix.
Effective values for gaseous and soot emissivities ar~e used, with a value of 0.1 being
.An expression for the gaseous emissivity (a, which is dependent taken for soct.
on the gaseous temperature, the partial pressure of CO2 combustion product), and the mean beam 'ength is presented.
These classical expressions and assumptions are acceptable as the present state of knowledge in radiant heat transfer.
However, there is some confusion about the definition of mean beam length
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on pages A.4-5 and A.4-7, where it is defined as a fraction of the electrical cable diameter.
The mean beam length cannot be a functicn of the target receivingtheradiation,butmustbeageomegricpropertyoftheflame producing the radiation.
Hottel and Sarofim have shown that the average mean beam length for a target at the flame boundary (very conservative) is well approximated by Lm = 3.5V /Af f
where Vf is the flame volume, and Af the flame bounding area.
Less conser-vatively for targets far remosed from the flame, a somewhat better approxima-tion 9 for l is 0.9 times the ratio of the effective flame volume to the e
flame area projected on a vertical plane.
It is not clear if this expression was used in the determination of the needed gaseous emissivity in the calcula-tion of radiant heat transfer, or whether a value of 0.2 was used as mentioned
' in the main body of the submittal. Also, calculations for a cylindrical flame, usirg the above mean beam length, give approximately the same heat flu'x
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results as the expression on page A.4-7, with D equal to the fire diameter.
Therefore,' the use of cable diameter in the submittal may only be a documenta-tion error.
A typographical error does exist on page A.4-6, where both the factors 0.131 ana 0.94D should be raised to the 0.412 power.
Also in need of clarification is the nature of the configuration factor used to obtain the fraction of the heat flux delivered to a target point by the assumed radiant right cylinder. The equation on page A.4-7 contains this factor but no mention is made as to what values are used or from where they,
are obtained.
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In Appendix A.5 of the sub.uittal, an analysis is presented which is used to provide a basis for determining the required size of baffles used to protect a vertical stack of trays from convective heating due to, direct impingement of an exposure fire plume.
A d3ta correlation 10 based on the turbulent wake behind a blunt body is used to obtain an expression for the required baffle width in terms of the downstream extent of the zone,to be protected.
The condition that the velocity be reduced to 20 percent of the-free stream value was used as a protected zone boundary definition.
- However, it is then impli,ed that the temperature reduction (defect) in the wake is linearly proportional to the velocity defect.
A closer review of Reference 10 indicates that experimental data and theoretical results based on Taylor's assumption of turbulence, rather than Prandtl's theory of free turbulence, results in the wake-temperature defect being equal to the square root of the velocity defect. Therefore, a shield which 1,imits the velocity to 20% of the free stream velocity, will only reduce the temperature to 45% of its free stream value.
This is less conservative than implied in Appendix A.S.
- 5. 6 Review of Appendix A. 6 Appendix A.6 discusses the solution of the two dimensional heat conduc-tion equation on page A. 6-5 for transient heat conduction in a solid by means The ac'uPacy of this method depends upon a of the finite element method.
c judicious choice by the analyst of element shape, nadal positions, interpola-ting function, and also a final judgement as to the acceptability of the temperature profiles.
There is no discussion of boundary conditions (e.g.,
how are the radiative and convective boundary conditions on both the heated surface and back face applied), the effect of neglecting the third spatial dimension, or the acceptance of the analyst's judgement.
However, the issue is not how to solve the equation, but rather, how WEP should demonstrate that the complex heat conduction processes taking place during a fire can be adequately modeled by the equation.
It is stated on page A. 6-8 that a switch arid an undervcitage relay might be representative of a broader class of the many different types of components that may be found mounted--in electrical switchgear and cabinets in a nuclear power plant.
No evidence is given in support of this judgement.
5.7 Review of Appendix A.7 This appendix contains a brief discussion of a one-dimensional heat trans-fer model for computing temperatures of objects inside a cabinet or panel.
Again we feel that WEP should discuss the limitations of the model.
For example, the, back wall in Figure A.7-1 appears to be exposed to a constant ambient temperature during the fire.
This may not be valid in general.
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Part of a t. cat radiction term is '
missing f rcm Eqn.1 on page A.7-2; the numbers at the bottom of page A.7-3 are for the product of density and heat capacity (pc in the model and not c as in 3
the text); and the units in parenthesis should be BTU /in -R for both steel and air.
j S.8 Review of Appendix A.8 This appendix contains a model for thin-wall temperat'ure response.
Under the thin wall condition, there are no body temperature gradients and heat l
received diffuses, instantaneously through the material.
This simplifies the mathematics of hest conduction and also affords treatment of more complex systems.
As a practical measure, a plate is considered thermally thin if the l
temperature difference across its thickness at a given instant is less than some prescribed value.
However, the thin-wall approximation may possibly not be valid for bakelite since its thermal conductivity is much less than that for steel.
The model calculates the response of a thin plate exposed to a i
radiant heat input while reradiating to a constant sink environment.
The equation is solved by a commonly used fourth-order Runge-Kutta method.
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5.9' Review of Chapter 4 Chapter 4 of the submittal outlines in.very general terms the methodology -
used in the fire hazards analysis of Chapter 5.
Due to this generality, only I
two comments are made here, viz, 1) the ventilation assumption and 2) the ignitability of high fire point hydrocarbon spills.
i The assumption is made that there is always sufficient ventilation to sup -
port an optimum stoichiometric fuel / air ratio and to maintain the compartment desmoked.
This'results in conservative estimates of the heat release rates.
Also conservatism is imparted in the analysis as a result of the neglect of attenuation of radiant energy due to smoke.
However, nowhere is consideration given for the possibility of secondary fires stemming from the ignition of the
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products of incomplete combustion, elsewhere in the enclosure.
i liquidhydrocarbonsbasedontheworkofModaki{ibilityofhig.firepoint The analysis in Section 4.3.2 on the combus is significant for evalua-ting the magnitude and duration of the external heat source necesary for ignition of postulated spills in the plant.
Note that the expression in the submittal (T on the right hand side represents time; on the left hand side T represents temperature) is only the leading term of Modak's expressions.
For thick spills this term is the classical solution for a non-transparent medium, with the additional terms necessary for semi-transparent oils.
For thin l
spills, the leading tem represents the condition where the spill depth ap-proaches zero.
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.;2 Sere are some serious err:rs i-TSbles 4-1.v.! /-2 cf the submittal.
In Table 4-1, tne values of thermal :mu::ivity and voi./#.ric neat capacity listed for concrete are actually th:2 values for copper given in reference 11.
Additionally, the units of thermal conductivity have been interpreted incor-rectly from reference 11.
Table 4-1 should read:
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Concrete (273* K)
- 1. 8 x 10-3 2.10 x 103 Liquid Hydrocarbon (300*-600*K) 1.25 x 10-4 1.90-x 103 5
This is an error of 109 in Aj of the hydrocarbon.
Whether this erroneous value was actually used in calculations is not clear.
The use of the correct parameters in the leading term of Medak's rela-tionships for a 10: minute exposure duration results in external heat fluxes considerably lower than presented in Table 4-2.
We calculate based upon the correct data tha following which should be compared with Table 4-2 on page 4-18 of the submittal.
Thin Spill Thick Spill Lubricating Oil-Flash point (489*K)
(Pennzoil 30-40)
- 20. 56 kW/m2 5.15kW/m2
-Ignition Temperature (650*K)
'37.98 9.52
~
Heptane-Ignition Temperature (487*K) 20.41 5.11 Comparing the values in these two tables leads one to believe that the con-clusion in the submittal, namely "that high fire point liquid hydrocarbons are, in actuality, not significant fire hazards when spilled on concrete" sh'ould be reconsidered in light of these corrected heat flux values.
5.10 Review of Chapter 5 The fire hazards analysis of individual fire areas is discussed in Section 5.
This section also addresses specific assumptions which are very important to the analysis, su d as the cable damage criterion, and the non-ignitability of lubricating oil.
The safety injection, containment spray, and emergency feedwater~ pumps lubricating, oil is not considered as a source of combustibles in the analysis.
In light of the lower revised values of required heat flux in Table 4-2, (a thick spill of oil wigh a flash point of 450 F would only require an external flux of about 5.3kW/m for 10 minutes to ignite), this assumption should be reconsidered.-
I s
o 13 N ch:;ic cc:.uar tu ;.:tential of the co.:astibility of the produ:t;. ;.1 pyrolysis of the cables.
For instance, the PE/PVC cable nas carbon monoxide and gaseous hydrccarbon yields 17". and 4". of the mass loss rate, respectively.
These products can collect in the ceiling -layer and result in a secondary fi re.
However, the stratification model is not valid for such secondary fires.
On the other hand, if the cables are completely enclosed.*in conduit, these combustion oroducts need not be considered.
The next consideration is the important one of selection of a cable dama'ge criterion.
The analysis focuses on the minimum conditions necessary to cause a 1 qs of cable function' through piloted electrical failure as defined by 9
Leen.
The chofce of the electrical failure appears to be somewhat less conservative for two reasons.
First, as stated by Tewarsonl4, cable damage first appears as insulation / jacket degradation, then piloted. ignition and then electrical failure.
Since Appendix R states that cables.should be free from fire damage, it would be more conservative to use the insulation / jacket degradatior, failure mode as a cable damageability criterion.
Also, page 4-3 of the submittal orders the stages of fire damage as offgassing, electrical failure, then ignition.
There should be some explanation of this ordering.
Secondly, the electrical failure tests of Lee were based on short cir-cuiting a 70V signal.
However, voltages in plant cables are usually much higher than this and could conceivably cause earlier damage than the tests indicated.
We note that Lee 13 tested two types of PE/PYC cables,. designated by h'im as Samples 5 and 6.
The electrical failure indices used by WEP are those associated with the latter sample due to the Qct that it exhibits a larger slope, which is a measure of the critical energy, than the former sample.
However, referring to Fig. 5-15 of Reference 13, it should be no.ted tht.t for external heat fluxes of 70 kW/m2 or less, the trend of the data indicates that Sample 5 exhibits earlier electrical failure than that shown for Sample 6 for the same incident heat flux.
Accordingly, the use of Sample 6 as thc referenced cable would yield non-conservative estimates within the aforenoted heat flux range.
i The point we are mak'ing here is that one should be careful' in the choice of referenced cabie utilized in the analyses due to the fact that the data can exhibit crossover for the same insulation / jacket material.
l However, it appears likely that WEP is making an unstated assumption, which would result in the cable with the largest slope in Figure 3-15 of Reference 13 being the most easily damaged.
This assumption is that cables are damaged at all haat fluxes, not just at heat fluxes above the critical heat flux as indicated in Reference 13.
This 'would result in all curves of l
Figure 3-15 of Reference.13 being shifted so as to pass through the origin.
The cable with the largest slope would then be damaged at an earlier time foe the same incident flux.
This neglect of the critical flux is conservative,
,. hut it is inconsistent with the data obtained.in Reference 13.
(.-
l
s "nother f6ctor in applyinn tne.eG udology is the assumption of instantly achievir.; a steacy-state, over-venti'.a.cc comaustion condition.
Assuming steady-state conditions are reached W.ediately conservatively maximizes the heat release from the exposure fire.
' We new discuss the application of the unit models given in the. Appendices
. to the specific fire areas.
The submittal st.ates that a "back calculation" approach is used which calculates the smallest quantity of fuel to'cause both redundant divisions to just exceed the damage criteria.
It is stated that
" classical optimization techniques for non-linear functions" are used.
However, this methodology. is not explained sufficiently to be reproducible.
The methodology description does not state which equations and minimization techniques are used.
Each. result should at least state the heat flux that each mechanism (plume impingement, stratification, radiation) delivers to the cable.
Stffice it to say that for Fire Zones wherein cables of concern are routed in conduit, the electrical failure criteria may be appropriate if indeed there are no other intervening combust.ibles within the area in question.
The stratified ceiling layer heat flux model has been discussed in Section 5.2, which reviews Appendix A.2.
It appears that the submittai uses a method which considers the transient heat flux model and ignores the critical hett flux aspect of the damage criteriom This conclusion is based purely on the results of various calculations since no description of the details or method are given in the submittal. The assumption pf cable damage occuring below the critical heat flux is extremely conservative when the critical heat flux is a substantial percentage of the maximum, steady-state heat flux.
The stagnation' plume impingement model on page A.3-8 is discussed next.
Calculations for configurations representative of those in the submittal yield rather, low values of heat flux.
The reasons for these low values are now given.
17, which The Rayleigh number for the plume impingement modg)is about 10 is.far beyond the range of correlation (109 < R ' < 10 given in Ap-l a
l pendix A.3.
The question remains as to why the plume impingement model yields--
such a % value for stagnation point heat flux.
Responding to fhis query -
required on in-depth examination of the relevant reference, viz., Reference 4..
We have already' alluded to some concerns regarding the use of the correlations presented in this reference in our critique of Appendix A.3 To re-emphasize, we feel tha{/5.e heat flux parameter behaves like (r/H) -1.25
'n rather than like (r/d)-
What concerns us here is that the experimental data used to obtain the correlations needed in the plume-impingement model are based upon tests performed with a sub-scaled apparatus.
One is therefore resorting to correlations based upon experiments performed at a maximum height on the order
'of 1 meter and applying these laws in areas where the ceiling heights could be an order of magnitude larger.
Also, the experiments considered flame heat.re-lease rates up to 35g0 W, while the present fires have heat release rates of l
about a factor of 10 larger.
Under these circumstances, the similitude in I
the turbulent length scales, which phenomenologically describe the diffusion processes of mass momentum, and energy, are markedly different.
This precludes the "universiality" of the correlations employed.
l g.-
~n..
z.- ;
/ Jis factor, we sg:st tne use cf the correlatien presented by
....n Alpert and used).
Basically their c;rrelation, for stagnation point heat flux agrees with that of Referen:e 4 :nce the Rayleigh number dependency is removed, and also agrees with the heat-flux parameter given by Nevnan and Hill l in the limit of zero ventilation rate.
It is not clear what radiation model is being used in the exposure fire effects calculation.
Is it the classical model with a total emi-ssivity of 0.3 as mentioned on page 5.8, or is it the model on page A.4-7 in which the gaseous emissivity is represented as a function of fire diameter.
Note the comments on the, definition of this diameter in Section 5.4.
In any case, both models require an, expression for the configuration factor which is not dis-cussed in the submittal..We suggest a configuration factor such as that given in Reference 16 for a cylinder radiating to a plane surface with a normal per-pendicular to the axis of the cylinder.
Another comment relates to larger size radiation fires.
For large fires in an enclosure, consideration must be given to the radiant heat flux from the stratified smoke layer to the target below.
Such a model is discussed.n Reference 17.
Summarizing this aspect of our review, we are still unsure as to the analytical procedures used by the utility in the back-calculation approach.
Much credit is taken for thermal shields located beneath cable trays.
However, no analysis is described for determining the heat flux to the cables.
The analysis in Appendix A.5 for determining shield width is a phenomenologi-cal model based on analysis including small scale turbulence.
H wever, fires differentiate themselves by large scale turbulence, resulting in convoluted ficus not accounted for in this analysis.
Also, fire pcsition is an important consideration.
If a' fire is not lo-cated directly beneatn a tray or there are adjacent banks of trays, the fire plume may cause camge by radiation to the top or sides of the tray or by convective heat tra.,sfer as the shield-disturbed plume courses its way through 4
the banks of trays.
In Fire Zone 8. a calcu'.ation is mentioned which results in a tray surface temperature of 188*F, which is below ignition temperature.
The compatability of using ignition temperature in lieu of the damage criteria, previously de-fined, has not been substantiated in the submittal.
Therefore, since little detail of the shield analysis is given, and due to errors pointed out in Section 5.5 and censiderations mentio%d here, the credit taken for t' a fire protection of thermal shields should be further scrutinized by UEl-2nd elaborated upon by analysis.
Also, in Fire Zone 8, where 4.4 gallons of acetone was found to damage a single cable, credit is taken for the mitigating effects of cimd-box cable
' trays and Kaowool blankets in contradiction to the analysis basic assumptions.
, This would require a heat conduction analysis through the box using a g
J e
t
,.._ 2... y.,
~
-_ -...... - ---... -. -. -. ~..
o con.wtive heat flux coeffici
- uter boundary ccndition.
This
~
analysis could, in principle, ve 9 r if detailed informaticn regarding box geometry and thermal properties wer2 gt.cn.
Neither was this information detailed in the submittal,~nor was there any discussion pertaining to a hett-conduction analysis.
For Fire Areas 6 and 8', WEP attempts td show that the electrical panels and cabinets used house relays and other electromechanical equicment also provide some level of fire propagation retardancy.
They.also take credit for rapid fire detection and response, which.is in contradiction to the premise of Chapter 4 that no credit will be given to fire protection mitigating systems.
Nevertheless, HEP, considers a 30-second exposure fire (cabinet panel is immersed in flames) using the " Hotbox" model of Appendix A.7, and a 30 second exposure fire using the heat conduction model of Appendix A.6.
We question the application of.these models in relation to Appendix R exemptions since they give only a general understanding of thermal lag and panel. fire resis-tance, and there is no clear discussion of boundary conditions.' Also, there could be a loss of functionability due to thermaT stresses induced by the highly transient thermal environment, (the table on page 5-96 shows large temperature gradients) and additional heating through vantilation ducts.
6.
CONCLUSIONS In our appraisal and review process, w'e have considered the following at-tributes: accuracy, completeness, applicability, and traceability.
Of the four,.we founo traceability, especially in the exemption request and in the optimizat.on technique, to be the most wanting.
Next in the decreasing 4
hierarchal order is completeness, mainly manifested by th'e lack of due con-sideration of other types and locations of exposure fire.
For applicability, we mainly question the use of the cable damageability index employed.
Ac-curacy, in a sense, is linked to the overall traceability of the analy@ sis aji._ w as such, cannot be completely judged.
We, however, do give credit to for utilizing state-of-the-art modeling techniques-(as we have defined); we give credit for their use of reasonable physical data and, in some respects, the degree of conservatism employe.d.
To editorialize for the moment, we feel hard-pressed to judge the overall conservatism.
In some fire phenomena fac-tors, the mode)s and assemptions lead to over-conservatism; in others, non-conservatism prevails.
We think the approach taken by WEP, employing a unit-problem methodology, is technically sound in assessing the impact of liquid pool spill fires, albeit incomplete in appraising the overall fire hazard within an area. Also, in aur estimation, the analysis, its limitations, and the lack of tra.:eability of the submittal, precludes one from demonstrating equivalency teLween pro-posed fire protection features and requirements stipulated in Section 11I.G.2 of Appendix R to 10 CFR 50.
~_ _
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7.
A~i'.NN!.EDGEMEf1T Tne author wishes to express his appreciation to Drs. John Boccio and Arthur Tingle for their suggestions and discussions relating to the fire-modeling methodology employed by WEP in, their fire-haarcs analysis of the Point Beach facility.
4 8.
REFERENCES Newman, J.S. and 'Hil'1, J.P., " Assessment of Exposure Fire Hazards to Cable 1.
Trays", EPRI-tlP-1675, Electric Power Research Institute, Palo Alto, Ca.,
January 1981.
Alpert, R.L., " Calculation of Response Time of Ceiling-Mounted Fire Oe-2.
tectors", Fire Technology, Vol. 8, '972, pp.181-195.
Alpert, R.L., " Turbulent Ceiling det Induced by Large-Scale Fires", Com-3.
bustion Science and Technology, Vol. 11, 1975, pp. 197-213.
You, H.Z. and Faeth, G.M., " Ceiling Heat Transfer During Fire Plume and 4.
Fire Impingement", Fire and Materials, Vol. 3, No. 3,1979, pp.140-147.
Alpert, R.L. and Ward, E.J., " Evaluating Unsprinklered Fire Hazards", FMRC 5.
J. I. No. 0183 6. 20, Factory Mutual Research Corporation, Norwood, Ma.,
August 1932.
Stavrianidis, P., "The Behavior of Plumes Above Pool Fires", a Thesi's 6.
presented to the Faculty of the Department of Mechanical Engineering of Northeautern University, Boston, Ma., August 1980.
i Steward, F.R., " Prediction of the Height of Turbulent Diffusion Buoyant f
7.
Flames", Combustion and Science Technology, Vol. 2,1970, pp. 203-212.
l Hottel, H.C. and Sarofim, A.F., " Radiative Trar,sfer", McGraw Hill Book 8.
l Company, New York, 19 67.
Orloff, L., " Lip Effects in Pool Fires", Paper No. 33, Canadian Section of 9.
the Combustion Institute, Kingston, Ontario, May 1979.
- 10. Schlichting, " Boundary layer Theory", Seventh Edition, McG~ raw-Hill Book Company, New York,1979.
- 11. Modak, A.T., "Ignitability of High-Fire-Point i.iquid Spill ', EPRI NP-1731, Electric Power Research Mstitute, Palo Alto, Ca., March 1981. -
- 12. Tewarson, A., Lee, J.L. and Pion, R.F., ""Categorizs'. ion of Cable Flam-mability Part 1: Laboratory Evaluation.of Cable Flammability Parameters",
EPRI HP-1200, Electric Power Research Institute, Palo Alto, Ca., October i
.1979.
O
t
\\
-18
- 13. Lc;, J.L., "A S :dy of
- .;;ntility of Electrical Cabies in Simulated Fire Environcents", EPRI HP-1761, Electric Power Research Institute, Pain Alto, Ca., March 1981.
- 14. Tewarson, A., "Damageability and Combustibility of Electric Cables", paper presented at FMRC/EPRI Seminar, Factory' Mutual Conference Center, florwood, Ma., December 9-11, 1981.
- 15. Veldman, C.F., et al, "An Experimental Investigat'on.of the Heat Transfer from a Buoyant Gas Plume to a Horizontal Ceiling, Part 1: Unobstructed Ceiling", fBS-GCR-77-97, June 1975.
- 16. Modak, A.T., 5and Orloff, L., " Fires of Insulations on Tank Exterf ors",
FMRC J. I. fio. 0AGR9.BU, Factory Mutual Research Corporation, flor.cood, tiassachusetts, October 1979.
- 17. Orloff, L., Modak, A.T., and Markstein, G.H.,," Radiation from Stroke Layers", Seventeenth Symposium (International) on Combustion, 1978, pp.
1029-1038.
- 18. Emmons, H.W., "The Prediction of Fires in Buildings", Seventeenth Sym-posium (International) on Combustior.,1978, pp.1101-1111.
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- 57
, c; 1.
Initiating Fire 1.1 Type of ccmbustible (liquid and/Or solid) 1.2 Amount of ccrbustible 1.3 Co.nbustible gecmetry/ orientation 1.3.1 pool spill (confined or unconfined) 1.3.2 solid fuel (vertical /herizontal) 2.
Initiating Fire Location 2.1 Relative to
- target (s)"
~
- 2. 2 Relative to room'gcometry 2.2.1 centrally located
?.2.2 wali E.2.3 ' corner 2.2.4 non-burning obstacles i
2.2.5 height 3.
Ccmbustion/ Pyrolysis Properties Initiating compustible/ target..gcmbustible (transient and/or fixed),
3.1 3.1.1 ignition sensitivity 3.1.2 mass 1..sss rate in pyrolysis 3.1.3 mass loss rate in cembustion 3.1.4 heat flux to surface (radiative & convective & lo_.cs) 3.1.5 excess pyrolyzate 3.1.6 fuel stoichiometry 3.1.7 heat release rate 3.1.8 product generation rate 4
i 4.
. Target Damageability Criteria 4.1 Solid combustibic > (cables) 4.1.1 insulation / jacket degradation ignition (, iloted ard auto ignitioq)'
4.1.2 n
4.1.3 electrical integrity failt. e
- 4. 2 Equipmert (safety eclated) 4.2.1 radiation heat flux 4.2.2 convective heat flux 4.2.3 chemical degradation (frem. products of combustion) 1 G-
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- 6. 2 Hot layer
- 6. 2.1 thickness 6.2.2 heat content 6.2.3 convective, heat transfer
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- 6. 3 Target (s)
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Protection Heasures
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- 7. 2 Detection
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- 7. 4 Administrative Controls 6
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