ML20078L329

Plant Phenomenological Evaluation Summary on Probability & Consequences of Deflagration & Detonation of Hydrogen in Support of Ipe

ML20078L329
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Site:	Farley
Issue date:	06/30/1992
From:	FAUSKE & ASSOCIATES, INC.
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FAI-91-58, NUDOCS 9411290163
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{{#Wiki_filter:. b I i OG i i FAI/9)-58 i FARLEY NUCIIAR PLANT UNITS 1 AND 2 PHENOMENOLDCICAL EVALEATION

SUMMARY

ON THE PROBABILITY AND CONSEQUENCES OF DEFIACRATION AND DETONATION OF HYDROCEN IN SUPPORT OF T11E INDIVIDUAL PLANT EXAMINATION + b Submitted To: i Southern Nuclear Operating Courpany Birmingham, Alabama Prepared By: Fauske & Associates, Inc. 16WO70 West 83rd Street Burr Ridge, Illinois 60521 (708) 323-8750 June 1992 O 9411290163 941109 PDR ADOCK 05000348 P. PDR fik't!!f4.4 b J: i

-i. p.v) ( as m cr This phenomenological evaluation summary assesses the susceptibility of the Farley containment to failure due to hydrogen deflagrations and detona-tions that may occur during postulated severe accidents. The failure of the containment could provide a pathway for the release of fission products. In particular, the possibility of an early containment failure due to hydrogen deflagration or detonation is of key interest when assessing potential source terms. The assessment concludes that the postulated containment loadings due to the deflagration of combustible gases in the Farley containment will not cause failure of the Farley containment structure. This conclusion is based on a bounding assessment of the containment pressurization potential assum-ing complete oxidation of Zirconium from the core and of metallic constituents (iron, chromium. and nickel) of the lower core plate. This 7m (v) assessment c onse rva tively ignores the existence of inerting due to steam produced in the course of damaging the reactor core. The limiting calcula-tion also ignores the possibility of incomplete combustion and the effectiveness of containment heat sinks including active and passive heat removal capability. The likelihood of detonations by either direct energy deposition or deflagration to detonation transition (DDT) have been as-sessed. It is concluded that =uch detonations will not challenge the integrity of the Farley containment. Thu: failure of the containment due to deflagrations and detonations in the Tarley containment will not be included as a separate node for either early or long term containment failure in the Farley plant response trees. Recommendations are provided regarding sensitivity studies that will be conducted as part of the Farley IPE to address phenomenological uncer-tainties and the specifics of sequence-dependent timing of interacting phenomena. These sensitivity studies will be used to demonstrate the impact r~N U

~. il - .'f7 of phenomenological uncertainties and confirm this papers conclusion regard-ing the capability. of the Farley containment -to withstand potential deflagration and detonations. Additionally, .it is recommended that the plant specific. analyses (MAAP calculations) be reviewed to specifically i summarize local effects of potential hydrogen deflagration ~ in the various containment regions. The potential locations of hydrogen deflagrations and detonations should be identified so that they may be used to assess any possible impact on the survivability of equipment, instrumentation, and containment penetrations during a severe accident, O I e

'lii - ',)*-- + TABLE OF CONTENTS lA&R 4F 1 i ABSTRACT... iii TABLE OF CONTENTS... LIST OF FICURES v .i LIST OF TABLES vi ... 1-1 l 1.0 PURPOSE. 2.0 PHENOMENA. 2-1 2-1 2.1 Description. 2-2' 2.1.1 Physical Processes t i 2.1.2 Relationship to Containment Failure Mechanisms and Modes . 2-10 2.1.3 Relationship to Source Term. . 2..... 2.2 Experiments. . 2 11 -2.2.1 Hydrogen Deflagrations . 2-11 . 2-14 I 2.2.2 Hydrogen Detonations 2 14 2.2.2.1 Intrinsic Detonability... 2.2.2.2 Influence of Steam and Temperature. . 2 19 . 2 19 i 2.2.2.3 Initiation of Detonations. 2.3 Analyses. . 2-22 2.3.1 Hydrogen Deflagration. . 2-22 . 2 24 2.3.2 Hydrogen Detonation... l l ) _. )

.~. . -. - ~.. iv

• TABIE Of CCFrElfts (Continued)

ZA&R~ f 3-1 '3.0 METHODOLDGY. j 3-1 3.1 Flammability. 3-2

3. 2' Combustion Completeness 3.3 Combustion. Pressure Rise.

3-2 1 3.4 Evaluation of DDT Potential. 3-2 -41 4.0 PLANT SPECIFIC. APPLICATION 1 4-1 4.1 Bounding Assessment of Hydrogen Deflagration. 4.2 Assessment of Hydrogen Detonation Potential. 4-10 l 4-15 j 4.3 Uncertainties 4 18 4.4 Conclusions. 5.0

SUMMARY

5-1. ) 6-1. l

6.0 REFERENCES

5 i APPENDIX A: Calculation of H Air-Steam Composition A-1 2 I B1 APPENDIX B: Post-Combustion Pressure Rise Calculation APPENDIX C: Summary of DDT Potential Evaluation from l C 1-i NUREC/CR-4803. i [ r f } 5 e -e

.y ,5 ((_} LIST OF FICURES f_[gure No. Eagg 2-1 Minimum ignition energy for hydrogen deflagrations (reproduced from Camp, 1983) 24 2-2 Comparison of ignition source energies (reproduced from Fauske & Associates. 1990a) 2-5 2-3 The flammability floor domain for upward flame propagacion for H -Air H O (vapor) mixtures. The 2 2 flammability limit curve is superimposed on the isobaric contours of calculated adiabatic explosion pressures (reproduced from Tieszen, 1987). 2-6 2-4 Theoretical adiabatic, constant-volume combustion temperatures of hydrogen-air-mixtures (reproduced from Sherman, 1981). 2-7 2-5 AICC pressures for various containment initial conditions (reproduced from Shermam. 1984) 29 2-6 Degrees of combustion in hydrogen-air-steam mixtures (from Liu, 1981). 2 12 f3 (,) 2-7 Combustion completeness for Nevada Test Site premixed combustion tests (reproduced from Ratzel, 1985). 2 13 28 The effect of steam addition at intermediate hydrogen concentration (reproduced from Kumar, 1984). 2-15 29 Comparative pressure profiles for three 8% (nominal) hydrogen combustion tests having different pre-combustion steam concentrations. Numbers in parenthesis are the steam concentration (reproduced from Ratzel, 1985) 2-16 2 10 Measured values (McGill, Sandia) of the detonation cell width (1) as a function of hydrogen concentra-tion (reproduced from Tieszen, 1987) 2-18 2-11 Detonation cell width as a function of the equivalence ratio for various steam concen-trations (reproduced from Tieszen, 1957) 2-20 2-12 Detonation cell width as a function of temperature (reproduced from Tieazen, 1987). 2-21 2-13 FLAME apparatus DDT results (reproduced from Sherman, Tieszen, Benedick, 1989). 2-23 j 1

\\ . vt - \\ LIST OF FICURES (Continued) Firure No. g I 4-1 Vertical section of Farley containment 4 14 1 e i i O O

- vtl - e,. [ ) LIST OF TABLIS V Table No. Page 4-1 Inputs for Assessment of Hydrogen Deflagration in Farley Containment 4-2 4-2 Typical Zion like Plant Severe Accident Results 4-4 43 Typical Results for Farley Station Blackout 4-6 4-4 DDT Scaling Assessment for Farley Compartments with Transverse Venting 4 11 4-5 DDT Scaling Assessment for Farley Compartments with no Transverse '.'enting 4-12 m 6 m

11 l i (* 1.0 PURPOSE b -1 i The potential failure of a nuclear power plant containment building due f to an energetic hydrogen burn has been the subject of technical exchange between the Nuclear Regulatory Commission (NRC) Staff, NRC contractors, and-l the nuclear industry. Discussion has been motivated by-the concern that hydrogen evolved during a core damage event could cccumulate in the contain-ment building and be ignited. Such an event occurred in containment during j the TMI-2 accident. -If the combustion was energetic enough to fail the l containment, the timing (i.e., before or shortly af ter vessel failure) and l uncertainties in location of the containment shell failure have two poten-tially _important ramifications regarding the radiological source ters. First, natural fission product deposition mechanisms in the containment I would not have sufficient time to significantly affect (reduce) the masses l of fission products that could be released through the failure location.

Second, fission products carried by gas flows from the contaiusent could result in radiological releases to the environment without-the benefit of.

the fission product removal capability of the containment spray droplets, j the water pool in the containment submerging the core debris, or steam driven Stephan flow in the containment. The objective of this paper is to develop a strategy. to account for ] postulated hydrogen combustion-induced failure of the containment when developing the source term portion of the IPE Level 2 Analysis for the l I Farley Nuclear Plant. The occurrence of hydrogen combustion and/or detona-tion will be assessed to determine the potential for containment failure. i I i 4 O

2-1 g f \\, 2.0 PHENOMEIA i i f 2.1 Description i Hydrogen combustion, or burning, is the result of a chemical reaction { l between gaseous hydrogen and gaseous oxygen. The products of such a reac-tion are steam and energy; the energy is liberated as light and heat. j Necessary conditions for hydrogen combustion are the presence of the right i amounts of hydrogen and oxygen, and the presence of an ignition source or a trigger. l Given a volume filled with only hydrogen and oxygen gases, minimum .} concentrationr, of hydrogen and oxygen are required for hydrogen combustion to occur. A mixture containing too little hydrogen to burn is called " lean", while a mixture containing too little oxygen to burn is called " rich". The presence of another gas that does not participate in the com-bustion reaction (such as steam or nitrogen) also acts to inhibit the O j occurrence of combustion. As the concentration of the inert gas increases, j the threshold concentration of lean hydrogen combustion increases, while the threshold concentration of rich hydrogen combustion decreases. Extensive research has been performed for numerous combinations of hydrogen,

oxygen, and inert gases to map out the hydrogen and oxygen concentrations that are combustible.

These maps are referred to as flammability limits. flammable hydrogen and oxygen mixture, a trigger is necessary Given a to initiate burning. Typically a spark is sufficient to ignite a flammable mixture. The required trigger energy decreases as the hydrogen gas tempera-i ture increases, until a threshold temperature is reached. Above the I threshold, the hydrogen is energetic enough to self-trigger (or auto-ignite) combustion of a flammable mixture. i Given a volume containing hyd'iogen and oxygen concentrations within the i flammability limits, and a trigger that initiated combustion, the magnitude of energy release depends on the mass of hydrogen consumed by the chemical reaction. The combustion of one Ib-mole of hydrogen releases 1.04 x 106 O

.2-2 BTUs of energy. Ignition of gas mixtures with hydrogen concentrations near the flammability limits have too little hydrogen or oxygen for the flame front to. propagate throughout the entire volume and consume all of the available hydrogen; such burns are called " partial" burns. As the hydrogen concentration moves away from the flammability limits, more complete burning of the reactants occurs. Above threshold hydrogen and oxygen concentra-tions, ignition will cause complete consumption of the reactants or " global" .j burns. Global burns yield the maximum energy release that can be obtained from hydrogen combustion. 2.1.1 Physical Processes Two types of hydrogen combustion reactions are pertinent to an IPE: deflagration and detonation. Deflagration is a combustion process in which the combustion front moves at subsonic velocity with respect to the unburned gas, while detonation is defined as sonic or supersonic propagation of the combustion front. This distinction is important because the pressure in the deflagration cannot exceed the adiabatic constant volume process value O (adiabatic, isochoric, complete combustion, or AICC). In a detonation, transient overpressures can exceed this value by a factor of two or more, and pressure can vary significantly across the detonation front. Pressure deflagration because the flame moves is uniformly distributed during a slowly with respect to pressure waves. The transient overpressure as-sociated with a detonation lasts only briefly, so ctructures may be able to withstand detonations when the impulsive load is not e/cessive. Factors which determine the type of combustion reaction are concentra-tions of fuel (hydrogen and carbon monoxide), oxidant (oxygen in air), and inertant (nitrogen, steam or carbon dioxide, in air), initial temperature and pressure, containment geometry, turbulence level, and combustible mix-ture ignition scurces. Composition and the initial thermodynamic state impose limits to both flammability and detonability, while geometry and turbulence can determine the. potential for detonation. Turbulence also e nhance s the burn completeness. In order for combustion to occur, the gas mixture must be flammable and there must be an ignition source. Relatively feeble sources such as static discharge and sparks can cause ignition.

2-3 'OV Figure 2-1 presents the minimum ignition energy for hydrogen deflagrations. For the case of 134 by volume of hydrogen in air approximately 0.07 mil-P lijoules are sufficient to ignite a hydrogen burn. Figure 2-2 compares various potential ignition sources. energies including a match. A match burning for one second can release ten joules of energy which is over three orders of magnitude more energy than that required to initiate a hydrogen burn. High temperrtures lead to slow volumetric oxidation, and very high hydrogen temperatures (about 1340 F (1000 K)) can cause autoignition. Autoignition is most likely to occur for sequences that release hydrogen rich and very high temperature gases from the reactor vessel into non. inerted compartments. It may also be possible that gas heating induced in the reactor cavity by dried-out core debris could result in very high tem-perature hydrogen mixtures being delivered to the lower compartment.

Again, if the recipient compartment is not inerted, autoignition may occur.

The well mixed average containment temperature is not hot enough during severe accident sequences to cause autoignition. O Classical limits for flammability and AICC maximum equilibrium final pressure are presented in Figure 2-3 [Hertzberg, 1981). The region of concern generally lies below the line of stoichiometric mixtures, in which hydrogen is the limitir-reactant. The minimum amount of hydrogen necessary i for combustion is ,_ightly over 44 in dry air. The minimum oxygen con-centration necessary for combustion is about 5% in dry air, corresponding to about 754 hydrogen. Addition of steam to any mixture of hydrogen and air would reduce the hydrogen volumetric concentration and increase the required threshold concentration for combustion. Figure 2-4 presents the AICC overpressure ratio resulting from combus-j tion in air, and indicates classical deflagration and detonation limits [Sherman, 1981). Figure 2-4 shows the lower limit corresponding to upward flame propagation (as indicated in Figure 2-3), and a higher limit cor-responding to downward flame propagation (i.e., against the buoyancy forces acting on the flame). Also, the hydrogen concentration required for O

~ 24 O 6 i i i i i g i i i 4 2 ~ es PRESSURE :1 ATM w J3 1 j .8 3 .6 .4 awzw .2 Z 9 O4 g .1 g .08 3 .06 33 E .04 3 .02 0 I I I I I I 1-0 10 20 30 40 50 60 70 HYDROGEN IN AIR, PERCENT BY VOLUME Figure 2-1 Minimum ignition energy for hydrogen deflagrations (reproduced from Camp,1983). O

2-5 O Recu - FOR DETONAT10N db COFMDTHE AT 13 4 H2 M' 1 12 KV ARC M l 480 V ARC M ERS j ARC t o E IGNITERS e... MATCH kO Y ^ 8 10 10 10* 10' 10 10 10' 10 10 10 10' 101b 1b JOULES /SECOND Figure 2 2 Comparison of ignition source energies (reproduced from Tauske & Associates, 1990s) d O

l' 26 9 _r> ~\\ 10 0 KEY' i - --- I imits of flommcbility Ici149'C), 9g'!'. Isotors of constcnt edic st'c ! esplation pressures,4tm) (Initial temperature at steam 80l soturotton) I cim ,, 3,,,en;,,,,,,,,,,,,,,, N'\\ 2 i i#r E* ** i' J - 70 3 mistures g g \\ L k 6C } 5 { 50 6 \\ H2 steem mistures z h 8 40 -7 \\ o N. h \\ %g \\, \\

• 30 7

F1cmmchte Nonflommoble p 7 j, 6 20 - 5 Ste chiometric /, - g line / 10 ~- Limit (u:wordh _ --[ -4,,,,~~_'- .m.d' 2 i /, N -_i i, i i s O 10 20 30 40 50 60 70 80 90 100 A00E0 WATER VAPOR, volume percent Figure 2 3 The flaminability floor domain for upward flameThe propagation for H, Air H 0 (vapor) sixtures. 3 flammability limit curve is superimposed on the isobaric concours of calculated adiabatic explosion (reproduced 'from Hertzberg, 1"981). pressures O

O O O g g 1 INITIAL CONDITIONS Q-I STOICHIOMETRIC T, = 2 9 8 K = 2 5'.C 8 RATIO N 8 ~ P, = I A T M = 0.1013 M P s g 100% RELATIVE LU HUMIDITY g-e D 7 I oe i W G~ ~ 0-6 DETONATION LIMITS = J< HE5 ~ l W E 3 4 DOWNWARD PROPAGATION LIMITS ~ -= W W E I Q-3 _J z i= UPWARD PROPAGATION LIMITS I a 1 m I a I a i n I I n i a I g 0 8 16 24 32 40 48 56 64 72 INITIAL HYDROGEN CONCENTRATION, VOLUME PERCENT Figure 2-4 Theoretical adiabatic, constant-volume combustion pressures of hydrogen-air-mixtures (reproduced from Sherman, 1981).

2 '- 8 detonation is shown as higher than the threshold for downward flame propaga-

tion, However, detonation limits shown in this figure are too simplistic for reactor applications because detonation limits have been shown to be dependent on scale and temperature in a systematic fashion.

The effect of steam on post-combustion pressure is quantified for various initial saturation conditions as a function of hydrogen concentra-tion in Figure 2 5 [Sherman, 1984]. In this figure, the initial pressure is calculated by adding the partia-pressure of hydrogen and steam to a humid air mixture originally at.14.7 psi (1.0 ata) total pressure and 80.6 F (300 K). A method for calculation of approximate pre-combustion conditions is shown in Appendix A. The post-combustion conditions for non saturated j conditions, or f~r mixtures containing carbon dioxide and/or carbon monoxide, can be easily found with iterative solution of the energy equation as shown in Appendix B. Generally, combustion is incomplete for hydrogen concentrations in dry air less than 84 to 9%, the downward flammability limit. Also, addition of O steam reduces the combustion completeness for lean mixtures. Therefore, the pressure calculations indicated in Figures 2-3 to 2-5 are upper bounds. Details are discussed in Section 2.2. t Detonations are more difficult to achieve than deflagrations. An initiation mechanism must exist, and the gas mixture must be intrinsically a detonable (i.e., able to sustain a detonation once initiated). Initiation mechanisms are energy deposition and deflagration to detonation transition (DDT) during flame acceleration. Intrinsic detonability is a function of both thermodynamic variables (pressure, temperature, and gas composition) and geometry. e DDT is the most likely mechanism for initiation of detonations. in a containment because known energy sources are not large enough for direct 1 initiation to occur. Also, a hydrogen concentration of at least 15% in dry air is probably necessary for DDT in a containment. The required concentra-tion for DDT increases with steam addition. Initiation of detonations is discussed in more detail in Section 2.2. i

29 0 7 4 i i i T' = 4 0 0 K 's ~ 2220 w== s T, = 37 5 K 38 7' = 3 5 0 K 2020 " 20 T, = 3 23 K g" 7' = 3 0 0 K 1820 E 3g o m 1620 2 16 m 1420 14 N' 1220 5 12 a.2 O h to 1020 2 8 820 0 8 INITTAL CCHOITIONS Uo STEAM PARTIAL PRESSURE 820 6 2 CORRESPONOS TO SATURAT10N i o AT TEMPSRATURE 7, l 420 E 4 Pm(300 K) = CJ7 ATy e i Pnge(300 K1 = 0.03 ATM g I t i f f I f f f f f ' - 220 12 18 20 24 28 32 38 40 44 48 52 4 8 HYOROCEM MOLE FRACTION (PERCEMT) l l Figure 2-5 AICC pressures for various containment initial conditions (reproduced from Sherman, 19A4). O

~. 2 10 t' t \\, A mixture of 13t hydrogen in dry air at standard temperature and pres. sure (STP) is intrinsically detonable. The required concentration decreases quickly with increasing temperature, but increases even more rapidly with steam addition. Mixtures with graeter than 304 steam may be immune to detonations. Intrinsic detonability is also discussed in greater detail in Section 2.2. 2.1.2 Itelationshin to Containment Failure Mechant sus and Modes Hydrogen combustion is a potential threat to containment integrity because of the increase in pressure and temperature during the event. During severe accidents, there are several potential sources of hydrogen. These sources include in-vessel oxidation of zirconium and stainless steel during core over heating and relocation. Ex vessel sources include zir-

conium, chromium, and iron oxidation during core-concrete interactions; zirconium oxidation during high pressure debris dispersal (DCH) and long term relocation of degraded core materials following vessel-failure. The

/'" quasi static pressure load on the containment is limited to the AICC pres-

sure, and is generally lower due to both incomplete combustion and heat transfer to structures during the event.

The transient load during,a period of milliseconds and may be unable to detonation lasts only for a cause containment failure, though equipment damage may still be possible. Because flames may accelerate during a deflagration, equipment damage may occur due to a local impulse loading. The temperature load only lasts.from tens of seconds to a few minutes, with the peak being brief and decay being sharp. From the phenomenological description in Section 2,1.1, it is clear that hydrogen combustion can.only occur when both sufficient hydrogen and i oxygen are present, and when insuf ficient steam is present to prevent flam-

mability, and when an ignition source exists.

Therefore, hydrogen combustion is not a possible containment failure mechanism for inerted containments. This includes inerting cue to steam which is often a direct and natural consequence of the postulated accident sequence and the progres-sion to core overheating and damage. Also, the containment load is highly i dependent upon initial conditions which influence combustion completeness l 4

2 11 1 i and pressure rise. The re fo re, the existence of a combustion event does.not ascomatically imply. containment failure. The maximum pressure must be compared against a containment failure criterion (see Section 4). 2.1.3 Relationshir to Source Term. Combustion influences the source term ~ primarily through containment l failure. That is, if hydrogen combustion can result in containment failure, then the time of combustion becomes significant. In general, combustion can-i potentially increase the source term because containment failure can occur i earlier through this mechanism than through slow pressurization due to loss of decay heat-removal. Early containment failure ~ generally increases the source term because the airborne fission product concentration tends to decay with time in an intact containment. In summary, the delay time be-tween fission product release from fuel (either in-vessel or ex vessel) and release to the environment (due to containment failure) may be decreased due to hydrogen combustion. O Because combustion influences containment gas and structure tempera-tures, a secondary influence on fission product revaporization occurs. This effect -is usually minor in the containment, and is quantifiable with knowledge of the structure's temperature history. 3 2.2 Experiments l 2.2.1 Hydromen Definerations t 7 In the preceeding section, a discussion of-ideal' adiabatic pressures l resulting from hydrogen deflagrations was presented. In practice, at low hydrogen concentrations, this pressure limit is not achieved because combus-l tion is incomplete. As illustrated in Figures 2-6 and 2-7, incomplete for hydrogen concentrations below the downward flammability j burning occurs limit, but above that limit combustion is fairly complete. The effect of s te am addition is also shown in'these figures. There is close agreement t between these sets of deflagration data despite a significant disparity in. i i

.u <s(j, 100-O V 'O / 90 / / O / 80 Q) / h 70 00 O Of O 3 f e O % STEAM; LITERATURE p 60 [O i O o % STEAM g O V 5 - 7 % STEAM a / O 7-10% STEAM 3a 0 oO O ro-15 % STEAM e o O i40 / x I s i 8 ao VlI;,l1 \\ 5 I u. 20 / I i 10 O V,O i i i o3 ( 5 s 7 8 9 to li 12 HYDROGEN CONCENTRATION (%-VOWME) Figure 2-6 Degrees of combustion in hydrogen air steam mixtures (from Liu, 1981).

2 13 v) 5 g 3 Standard Tests '@ Steam Laden Tests $ Tests with Spray Systems 00 Operative ~~.:.k. :, ^ N v

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l f f y 0 i 4 6 8 C 2 M HYDROCEN COtGNTRATON ( 7. ) Tigure 2 7 Combustion completeness for Nevada Test Site premixed combustion tests (reproduced from Ratzel, 1985). ) O

2 14 tN (_,) geometric scale of the vessels used in each experiment (i.e., 0.0763 ft8 (.002 m ) and 72324 ft8 (2048 m ), respectively) [Liu, 1981; Ratzel, 1985]. 8 8 In both cases, it is evident that addition of steam has an effect on the completeness of combustion, shifting the required hydrogen concentration to a higher value as more steam is added. Steam affects the combustion completeness, flame velocity, heat

capacity, and emissivity of the combustible gas mixture which, in turn, reduces the resultant system pressure rise.

Figures 2-8 and 2-9 illustrate this reduction in combustion pressure as a result of increasing the relative 8 s concentration of steam for two different size systems 222 ft (6.3 m) and 72324 ft8 8 (2048 m ) spheres [Ratzel, 1985; Kumar, 1984]). Both sets of data were taken with initial hydrogen concentrations of 8%. The pressure rise ratio is reduced by about 50% in the large apparatus, and by even a greater factor in the smaller apparatus. In each case, combustion was only about 38% complete for the highest steam addition test. 2.2.2 flydroren Detonations 2.2.2.1 Intrinsic Detonability The lowest value of hydrogen concentration for intrinsic detonability is now understood to be dependent upon geometric scale. Increasing scale allows the possibility of detonations at lower hydrogen concentrations. For

example, the detonability limits shown in Figure 2-4 were based on observa-tions in a small apparatus. Recently, detonability has been obse rved for mixtures of 13% hydrogen in dry air at 212*F (100*C) and 9.5% hydrogen in dry air at 212*F (100*C) [Kumar, 1984] in a much larger apparatus (16.93 in (43 cm) diameter tube).

In both cases, these detonations were initiated by large explosive charges. The National Research Council reached the conclu-sion that mixtures of 9 to lit hydrogen might be detonable based upon these experiments with hot, dry mixtures driven by explosive charges. Of course, explosive charges do not exist in a containment. (" t\\

O O O ~ Bottom Ignition Hydrogen Content 81 200 - Inttial Temperature 100*C ~ 3' initial Pressure 98 kPa S n =

• 150

[ \\ Dry H /Ai' ~ 20 2 o i sm I / L. / .\\ O-r. 100 ! s' i !l N. b 5 j 151 Steam j j Il \\ [N w 301 Steam g %.~%... l ' ~.... O i i i i i a i i o O 4 8 12 16 20 Time After Ignition, s Figure 2-8 The effect of a, team addition at intermediate hydrogen concentration (reproduced from Kumar. 1984).

2 16 l 3 O JS l mSPO4 (5 7.) o 3 v W /\\ ~ g 23 j\\ NTSPO5 (317.) Uiy I \\ l .!'..N 1 i .\\ 1 i :.*i .. '..N l O ig ..... N' ~N ~ l;f msPM (39 7.) ....., ~..... s ;. n' u ?: ~~~~~.... I j t i I I 1 20 0 20 40 40 80 W TEE AFTER STANT OF COWUSTION ( S ) Figure 2 9 Comparative pressure profiles for three 84 (nominal) hydrogen combustion tests having different precombustion steam concentrations. Numbers in parenthesis are the steam concentrations (reproduced from Ratzel. 1985). O 1

H2 17 i ( The ability for a detonation to be sustained or to propagate has 'been empirically found to be closely related to an intrinsic property of the mixture known as the detonation cell width, A [Tieszen, 1987; National Research Council, 1987; Berman, 1986). The value of A is lower for mixtures which are more easily detonable (hereafter termed more sensitive-mixtures). Detonations have a three-dimensional cellular structure formed by multiple in:eractions of transverse waves and the main shock front. The structure is ob servable from the " fish-scale" pattern left on a smoked foil by shock wave in:ersections. The lowest stable detonation wave mode, called the singlehead spin mode, can be related to a tube diameter D through the rela-l t.on A - r D [ Ties en, 1987]. For less sensitive mixtures.- where A is g larger and A/w > D, detonation is still possible given a sufficicntly strong initiator. For an open, unconfined cloud, the detonation criterion is more strict than for tube geometry. The minimum cloud diameter d, is related to i the detonation cell width A - d,/6.5 [ National Research Council, 1987). Measured cell widths for mixtures of hydrogen in dry air at 77'F (25'C) are shown in Figure 2-10. A minimum value of A occurs near stoichiometry t O (29.7% hydrogen). For leaner mixtures A increases rapidly, indicating a decrease in sensitivity. Detonation cell widths are unifo rmly lower at higher temperatures, indicating greater mixture sensitivity, that is, in-creased intrinsic detonability. These are plotted in terms of equivalence

ratio, denoted by 4 or e,

which is the molar ratio of hydrogen to air divided by the same quotient for dry air at stoichiometry. The equivalence ratio 4 can be written as SXg (2 1) 4*1.x .x 2.387. The equivalence ratio for 13% hydrogen in dry air is. where S 0.357. This value is unchanged by addition of steam to the dry mixture because the overall H and 0 m le fractions decrease in the same propor-2 2 tion. l 1 l

2-18 100 1 I l l l Twbe tMas= tar 5=oned resti pr,sswee osctitations (en) 8 ~ m V y l 'y 9 20 ~ 3 b a 9 2E. 5 a O 5 i 0 6 = 8 2-l i I I i 10 to M so 50 60 2 Figure 2-10 Measured values (McGill, Sandia) of the detonation cell vidth (1) as a function of hydrogen concentra-tion (reproduced from Tieszen, 1987).

2 19 t /"g () For the 16.93-in diameter (43-cm diameter) heated detonation tube (HDT) apparatus described in [Tieszen,. 1987] the critical. cell width is x x 16.93 - 53.16 in., which corresponds to roughly 134 hydrogen from extrapolation of the curve s in "Figure 2 10. Thus, the observation of a detonation at 9.5% hydrogen [ Stamps, 1987) demonstrates that high explosives can induce detona-tions in less sensitive mixtures. In that experiment 0.2337 lbm (106 grams) of high explosive, or about 474 Btu (0.5 MJ), was the trigger size. 2.2.2.2 Influence of Steam and Temperature The detonation cell width increases dramatically with the addition of steam as shown in Figure 2-11 [Tieszen, 1987] for mixtures at 212*F (100*C) Thus, steam as a diluent makes detonation more difficult to achieve. As temperature increases, detonation becomes easier as seen by the decrease in cell width in Figure 2-12 [Tieszen, 1987). Comparing the two figures, one can clearly see that in this temperature range the steam inerting effect is far more pronounced than the heating effect on detonation cell size. i O 2.2.2.3 Initiation of Detonations Detonations are much more likely to be initiated by deflagration to detonation transition (DDT) due to the size of ignition sources present in containment. The energy required for a detonation can be compared with energies of various ignition sources as shown in Figure 2-2 [Fauske & Associates, 1989). The largest possible ignition source in a containment, a 12 kv arc, results in a maximum are energy of 37900 Btu (40 MJ) over four cycles. As shown in Figure 2-2, a 12 kv are produces a peak rate of energy deposition that is about two orders of magnitude lower than the value for initiation of a planar detonation of 134 hydrogen in a confined tube- (0.176 lbm (80 grams) of high explosive per reference [ Shepherd, 1985)). All the l ignition sources indicated in Figure 2-2 are sufficient to cause a deflagra-tion. O

2 20 0 1000 , 1. 0 o h h m E E a w 100 3 g ,g 3 g O a ~ ~ O 8= 0 6e 10 O O o O o O 0% H2O Sandle HOT ~ O 10% H2O 1 a 20% H2O \\ o 30% H2O Shepherd Model 1 0.1 1 10 Equivalence Ratio Figure 2 11 Detonation cell width as a function of the equivalence ratio for various steam concen-O trations (reproduced from flessen, 1987).

4*46 ] 450 4 4 i o Po = 1 atm HDT DATA o /o = 42 mol/m' 400-e Po = 1 atm SHEPHER0_ ~ e /o = 42 mol/m' MODEL [ 350 - E \\ X 300-i .c ~ U 250 - 3 =e a O 200-C i o ee 150 - c e o e O O 100 - 0 50-0 l i 260 280 300 320 340 360 380 i Initial Temperature (K) Figure 2 12 Detonation cell width as a function of temperature (reproduced from Tieszen, 1987).

2-22 The lowest hydrogen concentration for which DDT has been obse rved is + 15% [ She rman, et al., 1989). The apparatus used was the FLAME facility at Sandia, which is a half scale model of an ice condenser upper plenum, 8 ft (2 04 m) high, 6 ft (1.83 m) wide, and 100 ft (30.5 m) in length. To promote turbulence, obstacles were placed along the interior of this long straight rectangular channel. The 154 low limit corresponds to a case with no top venting and periodic obstacles every 6 ft-(1.83 m). In a case with no obstacles, 254 hydrogen was required, as shown by Figure 2-13, and for a case with obstacles but 50% top venting, 20% hydrogen was required. It is very difficult to relate the detonation cell width to a necessary or sufficient criterion for DDT because other characteristic lengths of the geometric configuration are influential. A case of attempted scaling of DDT by A is reported by [Berman, 1986] who compares two sets of experiments [ Stamps, 1987], one in a 3.28 ft by 3.28 ft (1 m by 1 m) channel and another in a 9.84 ft by 9.84 ft (3 m by 3 m) channel. The smaller experiments were performed at-stoichiometry, and the larger experiments were performed by lower concentrations in the belief that scaling with A could occur At a 01. concentration of 21% hydrogen (4 - 0.63), for which A(d .63)/A (4 - 1) - 3, DDT occurred. This supports the hypothesis that in similar geometries DDT may occur for mixtures with similarly scaled cell widths. In both cases the apparatus included obstacles and was unvented. 2.3 Analyses 2.3.1 Hydroren Deflarration i Two types of analyses are-required for quantification of hydrogen l deflagrations: flammability and pressure rise. The current state of-the-art analytical tool for severe accident analysis is the combustion model developed by the DOE Advanced Reactor Severe Accident Program (ARSAP) [Fauske & Associates, 1989b] which has been incorporated into HAAP 3.08. i

2 23 i so: 00@ oRD 8 'i O DEFLAGRATION - NO OBSTACUES $ DETONATION - NO OBSTACt.ES o O DEFLAGRATION - OBSTACLES 3-f: E DETONATION - OBSTACLES e 20 - z l10-00O O %~ m m 0 5 10 b 20 T5 35 HYDROGEN MOLE FRACTION, % Figure 2-13 FIAME apparatus DDT results (reproduced from Sher =an, Tieszen, Benedick, 1989). O

2 24 ) In the absence of a more 'detaileo model, the flammability limit diagram '(Figure 2 3) can be used to estimate flammability. ' Mixture mole fractions can be estimated as shown in Appendix A. The combustion completeness can be estimated from Figure 2-7. .The combustion pressure rise can be estimated .throu6h lthe methods discussed in Appendix 8. Section 3.3 outlines an as. sessment methodology based on these appendices. 2.3.2 Hydromen Detonation No publicly available computer programs exist to predict the detonation cell width, the essential parameter for estimation of mixture detonability, Programs have been developed by NRC for internal use- [ Shepherd, 1985) and ~ agree fairly well with existing data, as shown in Figures 2-10 and 2-11. An empirical technique for judgement of detonability [Sherman and Borman,1987)- is described in Section 3.4 i O

31 'r s lV 3.0 METil0DOLOGY The methodology for determining the potential for hydrogen combustion consists first of identification of sequences in which combustion might occur and second estimation of the combustion pressure rise. The potential for detonations is then considered. Finally, the load to containment must be compared with the containment failure criteria. The MAAP code can be used to study hydrogen combustion and the resulting containment load as an integral part of assessing severe accident sequences. Section 3.3 is ap-plied in Section 4.1 in a simplified and conservative manner to assess the potential for containment overpressurization due to hydrogen burn for the Farley containment. 3.1 Fl ammabili ty Identification of sequences in which combustion might occur requires g calculation of the gas composition (mole fraction of all components), tem-(~'I 's-perature, and pressure histories for all sequences. These states are then compared with a flammability limit diagram such as Figure 2-3. Fortunately, many sequences can be ruled out if one or more of the following conditions are met: less than 4% H2 *1**78' greater than 55% H O when H Present (steam inerting), 2 2 less than 5% 0 when H Present. 2 2 If an ignition source exists, then combustion is assured when the flam-mability condition is met. Otherwise, combustion may never occur. However, random ignition sources could cause ignition at any time. Therefore, when mixtures are flammable and no explicit ignition source exists, a sequence should be evaluated both with and withcut combustion. V i

i 32 ~3.2 Combustion Completeness A model can be used to quantify combustion completeness. or the data of Figure 2 7 will suffice for an estimate. An overestimate of combustion completeness will lead to an overprediction of final pressure, and therefore it may result in a conservative estimate of containment overpressure failure or conservatively early containment failure times. 3.3 Combustion Pressure Rise r Combustion pressure rise is determined through solution of the energy equation as shown in Appendix B. The steps to follow are: 1. Determine the initial gas composition (number of moles of each gas). 2. Determine the combustion completeness. / 3. Determine the final gas composition. 4 Determine the heat of combustion. r 5. Solve for the final temperature. 6. Solve for the final pressure. 3.4 Evaluation of DDT Potential The potential for DDT in current commercial and a'dvanced light water l reactor containments has been evaluated using a procedure for engineering judgement by Sherman [Sherman and Borman, 1987). The procedure assumes that the potential for DDT can be evaluated based on the mixture intrinsic flmm-mability (detonation cell width)- and type of geometry. Five classes of mixture sensitivity are defined ranging from class 1, most detonable and near stoichiometry, to class 5, least detonable, hydrogen mole fraction less than 13.5 in dry air. Five classes of geometry were defined ranging from

I 3-3 f r h class 1, most conducive to DDT featuring large geometries with~ obstacles and partial containment, to class 5, unfavorable to DDT featuring large scale and complete unconfinement or small scale spherical geometry with central ignition and no obstacles. i The mole fraction and equivalence ratios in Table C-1 are shown for dry f hydrogen air mixtures. For mixtures that include steam, experimental. data [ and code calculations are used to calculate the corresponding cell widths. Figure 2-11 shows dependence of cell width on hydrogen and steam concentra-tions.

Hence, the mixture class in Table C-1 corresponds to cell widths; I

the equivalence ratios / mole fractions shown are illustrations for the dry case. Class 1 mixtures are extremely detonable. They are very likely to undergo DDT in most geometries of interest. Class 2 mixtures are slightly l less likely to detonate. Class 3 mixtures have been obse rved to undergo transition in geometries which favor flame acceleration. Detonations have + been propagated through Class 4 mixtures, but to date, DDT has not been obse rved for a hydrogen concentration less than 15%. Class 5 mixtures are unlikely to undergo DDT, although a detonation has been propagated in a l 13.5% H in air mixture at STP. 2 t The procedure for estimating DDT potential is thus as follows: 3 1. De te rmine the mixture class from Table C-1 using Figures 2 10 and 2-11. l 2. Determine the geometric class as guided by Table C-2. 3. Find the result class in Table C-3, 4 Assign a probability to this result from Table C-4 I t i 1 c

4-1 (q) 4.0 PIMr SPECIFIC APPLICATION Section 3 of this paper describes a method for assessing the effect of hydrogen deflagration or detonation within the containment during the course of a severe accident. This section will describe the application of that methodology to the Farley IPE analysis. A bounding assessment of contain-ment pressurization due to hydrogen burns will be made to assess their potential as a failure mode for the Farley containment. 4.1 P,oundlne Assessment of Hydronen Deflacration The methodology presented in Section 3 for combustion pressure rise is applied in this section. In order to demonstrate the robustness of the Farley large dry containment a very conservative set of assumptions are made in the following assessment. The initial conditions and inputs used for this assessment are presented in Table 4-1. The values for the net contain-c ment volume, zirconium mass, and lower core plate mass were taken from the Farley parameter file. Since steam acts as an inertant, it is conservative a completely dry atmosphere in the containment immediately prior to assume to an accident. Sten 1: Containment Hydroren Composition Following the method described in Appendix A the moles of air (n ) g initially in the containment is calculated, using Eq. (A-1) and Table 4-1, to be 4872 lb moles - (2214 kg moles). This number represents the amount of air present under normal operating conditions, i.e., prior to a severe accident. The amount of nitrogen does not change during the severe acci-dent, but the mass of steam does due to the initiating event and condensation on active and passive heat sinks. The mass of hydrogen may increase over time as a result of the severe accident and zirconium oxida-tion with steam. O \\ l V

4'. 2 ' O V Table 4-1 INPUTS FOR ASSESSMDC OF HYDROCEN DEFIACRATION IN FARLEY CONTAINMDR t CONTAINMENT g Volume (V) 2.067 x 108 fes (5.85 x 10* m ) 8 Initial Pressure (P,) 14.7 psia (1.013 x 105 Pa) Initial Temperature (T,) 120*F (322 K) i Relative Humidity 04 Initial Air Mass (n ) 4872 lb, mol (2214 kg-mol) C.QEE Hass of Zirconium 38,230 lb, (17,377 kg) Mass of Iron, Chromium, and Nickel in Lower Core Plate 4368 lb, (1985 kg) PHYSICAL PROPERTIES Constant volume Heat Capacity j Gn (at 1160*F or 900 K)- Nitrogen 0.195. BTU /lb, R (0.818 KJ/kg-K) Oxygen 0.183 BTU /lb, R (0.767 KJ/kg K) Steam 0.43 BTU /lb, R (1.8 KJ/kg-K) Hydrogen 2.53 BTU /lb,-R-(10.59 KJ/kg K) ) i Heat of Reaction [Villiams, et al.,1987) REACTIONS (BTU /lb -mol or MJ/ke-mol reactant) 2+ 2-HO 1 04 x 105 or 242 H 2 2.57 x 105 or 98 Zr + 2H O

• Zr02 + 2H2 2

Cr + f H O - f Cr 023+fH 0.86 x 105 or i)0 2 2 Fe + H O - Feo + H 0.12 x 108 or 28 2 2 Ni + H O

• Nio + H 0.01 x 105 or 2.5 2

2 O q

'43 I (* k The containment hydrogen composition will vary with accident sequence. In order to perform a bounding assessment,. typical conditions for three severe accident sequences predicted by the MAAP code for a Zion like PWR plant have been selected and are summarized in Tables 4-2. The tempera-

tures, pressures and H mass from' the core are taken to represent the 2

containment conditions right after reactor vessel failure. Additional mass of H was pr duced due to molten core-concrete interac-2 tion (McCI) over a long period of time starting from the debris dryout time after the reactor vessel failure. However, as'shown in Table 4-2., a sig-was pr duced by MCCI only during the station nificantly large. amount of H2 blackout sequence. The availability of water from the engineered safety ~ system prevented the debris from attacking the concrete during the smal1 and. 1arge IDCA sequences. and thus, minimized the H Production. 2 Among the three typical severe accident sequences shown for a Zion-like

plant, the station blackout would produce the overall largest mass of hydrogen and, the re fo re, is the most likely sequence to form a flammable gas mixture.

Similar arguments can be applied to Farley as well; the station black-out sequence would produce the largest amount of hydrogen gas. The significant difference from the sample calculation of Table 4-2.is the shorter dryout time after debris is expelled out of the Farley vessel. This is because of a 16 ft (4.9 m) tall " curb" which. prevents water on the lower compartment floor from flowing into the Farley cavity during the station blackout. In the sample calculations of Table 4-2, this curb does not l exist. The debris dryout time in the cavity of Farley depends on the amount of water remaining inside the RPV at the reactor vessel failure time. A i fraction of this water may remain in the cavity after the vessel blowdown. If the debris dryout time is shorter, MCCI will start producing hydrogen at an earlier time. During IDCAs at Farley, water can be delivered to the I cavity after reactor failure through the low pressure injection system, thereby preventing the occurrence of MCCI-induced hydrogen generation. O

44 I i <~\\- '\\sl I Table 4-2 TYPICAL ZION-LIKE SEVERE ACCIDENT'RESULTS( } Containment at RPV Mass H Mass H % Zr 2 2 Failure Time From Core From MCCI(2) Oxidation Accident Sequence Temperature Pressure Small LOCA 242 *F 34.8 psi 617 lbm 0.2 lbm 32 390 K 2.4 x 10s Pa 280 kg 0.1 kg Large LOCA 241 *F 40.6 psi-485 lbm 4.8 lbm 25 O 395 K 2.8 x 106 Pa-220 kg 2.2 kg ,%) Station Blackout 278 *F 34.8 psi 540 lba 1287 lba 93 410 K 2.4 x 105 Pa 245 kg 580 k6 ( )These-results are derived from the PVR sample problems provided in the MAAP User's Manual-(Modular Accident Analysis Program (MAAP), MAAP 3 0B User's Manual, March 16, 1990. ( ) Total hydrogen produced during molten core concrete interaction (MCCI) at time of containment failure. t O

45 l 1 1 Jrs - k,,s) To be conservative in the assessment, it is assumed that large uncer-tainties exist. in the overall mass of H2 generated and the amount of steam at the time of interest for the three. sequences. .Howeser, this can be overcome by assuming in core hydrogen production due to 100% oxidation of all zirconium and metallic constituents of the lower core plate. This highly ' conservative amount of H is assumed to enter the containment com-2 partments'at the time of reactor vessel failure. In light of this l assumption, the differences among ths three sequences become insignificant. Considering one of these sequences will suffice. Here, the station blackout sequence will be used for sample calculations. Both hydrogen and steam increase with time as MCCI in the cavity and steaming of water on the lower compartment floor proceed. As noted above, the total amount of H at the time of reactor failure will be assumed at the 2 high limit.of 1840 lb (836 kg) equivalent to 100% oxidation of Zr in the core and metallic constituents of the lower core plate. However, the amount of inerting steam will be conservatively limited to the level at the reactor failure time. The containment temperature and pressure at RPV failure are O. taken to be 280 *F (410 K) and 40 psia (0.28 MPa), based on typical Farley MAAP results and consistent with or more severe than the Zion-like condi-tions shown in Table 4-2. Appendix A can now.be used with these data to calculate the mole fractions of steam and hydrogen in the containment prior to hydrogen deflagration. The results are shown in Table 4 3. The containment. cond'- tions at the time of vessel failure show containment inerting by steam. This can be seen by referring to Figure 2 3 and observing that these condi. tions fall in the non-flammable region; hydrogen in air - X - 8.8% and H,w added water vapor - X - 44%. g 4 b 'l s

4-6 (('}) Table 4-3 Typical Results for Farley Station Blackout Before Accident Air initially in containment (n ) 4870 lb-mol (2210 kmol) A N initially in c ntainment 3850 lb-mol (1750 kmol) 2 0 initially in c ntainment 1020 lb-mol (465 kuol) 2 After RPV Failure H fr m 100% oxidation of Zr 838 lb-mol (381 kmol) 2 H fr m 100% oxidation of Fe, Cr, Ni 82 lb-mol (37 kmol) 2 of lower core plate Containment temperature before combustion (T ) 280 *F (410 K) p Containment pressure before combustion (P ) 40 psi (2.8 x 105 Pa) P 22.3 psi (1.53 x 105 Pa) nc P, (-P -Pnc) 17.8 psi (1.23 x 105 Pa) p H O in containment air 4630 lb-mol (2100 kmol) 2 H O from combustion 920 lb-mol (418 kmol) 2 X 0.44 3 X 0.088 HW X 0.159 H,d E **1 N c mbustion 563 lb.mol (256 kmol) 02 "fE'# 2 Total gas and steam after combustion 9960 lb mol (4530 kmol) l Containment pressure after combustion 112 psi (7.73 x 105 Pa) Containment temperature after combustion 1700 *F (1200 K) / \\ V j

47 ( ) Steo 2: Combustion Comoleteness v As observed in Step 1, steam inerting is anticipated in a majority of severe accident sequences. The operation of contairment heat removal sys-tems is included in the typical LOCA sequences. In the case of a station

blackout, no active heat removal systems are available.

Furthermore, a limited amount of water may be present in containment to inert it since no injection systems are operable during a station blackout. However, the water inventory in the accumulators alone is sufficient to inert the con-tairment. Given an inerted containment no combustion would occur. Nevertheless, the bounding asses,sment provided in this section will conser-vatively assume complete combustion. Steo 3: Final Cas Cocoosition The gas composition changes following the hydrogen combustion. The entire hydrogen mass is assumed to be burned so the final gas composition is free of hydrogen. The containment oxygen inventory is reduced by 460 lb-(j moles during the burning of the 920 lb-moles of hydrogen. At the same time, the steam mass in containment is increased by 920 lb-moles. The net result is a small reduction in the number of moles of gas in the containment fol-Iowing the complete combustion of the hydrogen and assuming no condensation of the steam. For the station blackout, the ratio of total moles of gas af ter combustion to total mole s of gas before combustion is 0.957. Steo 4: Heat of Combustion The heat of combustion is simply calculated by multiplying the heat of reaction of one Ib mole of hydrogen combining with oxygen to produce steam by the number of lb-moles of hydrogen. In Table 4-1, the heat of reaction is given as 1.04 x 105 BTU per Ib-mole of hydrogen. Oxidation of 100% of the =irconium inventory and the metallic constituents of the lower core plate produced 920 lb moles of hydrogen. Thus. 9.57 x 107 BTUs of energy are released by the complete combustion of this hydrogen. /'~'s L) 4 l

. - _7 48 Steo 5: Post-Burn Temperature l I The pressure in containment is bounded by calculating the adiabatic isochoric complete combustion (AICC) of the hydrogen. This adiabatic cal-culation ignores the presence of any passive or active heat sinks in the Farley. containment. Thus, all the combustion energy is used to heat the containment atmosphere and produce the largest possible pressure increase. 'l This calculation assumes that all the hydrogen produced during-the severe-accident accumulates in containment and burns all at one time. This ignores the possibility of hydrogen burning as it is released either by auto-ignition or in localized regions should the flammability limits be l satisfied. If the hydrogen is burned as it is released, the passive and active heat removal available in the Farley containment would avoid sig-nificant pressurization. The energy balance for a constant volume burn (see Appendix B) can be used to calculate the post-burn temperature. The result (Table 4-3)'for the station blackout is a post-burn temperature of 1700*F (1200 K). Steo 6: Post-Burn Pressure 1 The post-burn containment pressure for the AICC calculation can be estimated per Equation B-2 in Appendix B. This results in an estimated post burn containment pressure of 112 psia for the station blackout se-l quence. This bounding assessment has been based on. very conse rvative i assumptions. The assumptions include a large source of hydrogen (oxidation of 100% of the core zirconium and the metallic constituents of the lower core plate), no credit for containment heat sinks (passive or active), and no credit for steam inerting. Nevertheless, the resulting post-burn pres-sure is still within the Farley containment's ultimate capacity of 117 psia, as calculated in the Farley Phenomenological Evaluation Summary on Containment Overpressurization. r It is concluded that hydrogen deflagration in the Farley containment will not fail the containment boundary. Thus, the plant response trees do not need a separate node for hydrogen deflagrations. c t F --+n- - - .n, -enn --..,a

49 -{) 4.2 Assessment of Hydroren Detonation Potential i The initiation of a hydrogen detonation by direct deposition of energy j in the gas mixture is discussed in Section 2.2.2.3 (Initiation of Detonations) of this paper. A large energy source (explosive charge) is required to initiate a detonation even in the extreme conditions postulated in Table 4-3 (15.9% hydrogen on a dry basis). Figure 22 shows that the most energetic ignition sources found in reactor containments are at least several orders of magnitude too small to trigger a hydrogen detonation.

Thus, it is concluded that hydrogen detonation by direct energy deposition is not possible in the Farley containment.

The detonation of hydrogen by a transition from a deflagration (DDT) due to acceleration of the flame front is discussed in Section 2.2.2.3. The potential for a DDT induced detonation in the conta1 ament will now be as-sessed. This potential will be assessed as a function of two variables; i.e., reactivity of the mixture and geometric configuration. The assessment o of reactivity of the mixture in given dimensions will utilize scaling of \\j detonation cell width as discussed in Section 2.2.2. The geometric con-figuration assessment will utilize the method presented in Section 3.4 (Evaluation of DDT Potential). The assessment will assume that H2 gas is unif rmly mixed with s tea 2n and air in and between all compartments. For WCA sequences, the post lDCA air mixing system would provide good mixing between containment regions and would minimize hydro 5en pocketing. During a station blackout (i.e., a total loss of AC power), air mixing by natural circulation can be expected.

However, some hydrogen pocketing may occur in the steam generator enclosures, particularly in the area around the reactor coolant pumps.

This is due to the compartmentalized design of these enclosures, as well as the location of the pressurizer relief tank in the lower compartment. The occurrence of hydrogen pocketing can be considered as a deviation from the presumably well mixed conditions. However, it should be noted that during a station blackout with a seal LOCA, not only the local concentration of hydrogen but also the concentration of steam (an inertant) in the steam generator enclosures could be higher than the containment average. O

4-10 l g s 1 's_s' The detonation cell width is estimated as follows. Equation (2-1) is used to calculate an equivalence ratio (d) to estimate the detonation cell I size based on the postulated extreme conditions in the Farley containment. These unrealistic, but extreme, conditions are a dry containment (X - 0) 3 and 100% oxidation of the zirconium and the lower core plate (leads to XH,d - 0.159). These conditions result in 4 - 0.45 and per Figure 2-11 a detona-tion cell width of approximately 0.10 m. This detonation cell width could be used to define a scale factor to extrapolate between the FLAME facility (small scale) and the reactor scale. The results from the FLAME facility (see Figure 2-13) show the requisite hydrogen concentrations (15% and 204) to have DDT for both transversely vented and unvented geometries. The equivalence ratio that applies to these tests results in detonation cell widths of 0.66 ft (0.2 m) (15% H ) and 0.082 ft (0.025 m) (20% H ). Scale 2 2 factors can be calculated by forming the desired ratios of detonation cell widths. For example, the scale factor for 15.9tm H fr the transversely 2 unvented case would be 0.33 ft/0.66 ft (0.10 m/0.2 m) - 0.5 and for the transversely vented case would be 0.33 f t/0.082 f t (0.10 m/0.025 m) 4 ("'} Such scale factors are then applied to the dimensions of the FLAME facility 2 (6 ft x 8 ft duct) to estimate the minimum channel size needed at reactor scale in order to produce a DDT. The results for DDT assessment are shown in Table 4-4 for compartments with transverse venting, and in Table 4-5 for compartments without transverse venting. The lower section of the lower compartment (El. 105' 6" 129') and the annular compartment can be considered as channels with transverse venting because flame propagation in either one of these compart-ments will experience side venting due to the opening in the walls separating the two compartments (as shown in Figure 4 1). Furthe rmore, flame propagation in the annular compartment will also be " vented" due to the increased width of the annular compartment at the 129' El. Figure 4-1 provides insight into the lower and annular compartments configurations. Both the lower section of the lower compartment (El. 105' 6" - 129') and the annular compartment are assigned geometric class 4, which refers to geometries unfavorable to flame acceleration (according to Table

l 4 11 (y-N Table 4-4 t DDT SCAllNG ASSESSMENT FDR FARIEY COMPARTMENTS j WITH TRANSVERSE VENTING Bases: 1) Regions with transverse venting

2) Assume 15.9% hydrogen 3)

Safety factor of 2

4) Minimum steam inerting, 10% steam II)

Scale Factor: Ratio of Detonation Cell Widths x x Inerting Factor g, g Factor A 15.9% 1 x 10 0.10 m0.025 m

• 1 x

x 10 2 3 2 20% - 20 (applied to FLAME facility dimensions - 6 ft x 8 ft) Minimum channel size required at reactor scale in order to accelerate flame to DDT: 20 (6 ft x 8 ft) - 120 ft x 160 ft Farley Key Dimensions: Width Height Geometric (ft) (fti Class Lower Compartment @ EL. 105' 6" - 129' 27 19 4 Annular Compartment @ EL. 105' 6" - 129' 13 24 4 Annular Compartment @ EL. 129' - 155' 47 26 4 CONCLUSION: NO POTENTIAL FOR DDT III See Figure 2-11. O p .m-

4 12 i Table 4-5 i DDT SCALING ASSESSMENT FDR FARIEY COMPARTMENTS WIT 11 NO TRANSVERSE VENTING j i Bases: 1) Regions with no transverse venting 2) Assume 15.9% hydrogen 3) Safety factor of 2 4) Minimum steam inerting, 10% steam. l Scale Factor: Ratio of Detonation Cell Widths x x Inerting Factor ( ) 3,f,ty Factor xfx10 15.9% 0m x x 10 - 15% - 2.5 (applied to FLAME facility dimensions - 6 f t x 8 f t) Minimum channel size required at reactor scale in order to accelerate flame to DDT: 2.5 (6 f t x 8 ft) - 15 f t x 20 f t Farley Key Dimensions: Width Height Geometric (ft) (ft) Class Steam Generator Compartment @ EL. 152' - 166.5' 3-4 ft 15 ft 3 Steam Generator Compartment 152' 22 ft 10 ft 2 @ EL. 129' CONCLUSION: DDT is possible but unlikely ( }See Figure 2 11. O

4 13 3/ m h C2 in Appendix C).

Also, as shown in Table 4 4, the dimensions of the lower compartment and the annular compartment are much smaller than the j

calculated minimum channel size required at a reac or scale in order to accelerate a flame to DDT. Therefore, there is no potencial for DDT to occur in these regions. DDT assessment results for regions with no transverse venting are summarized in Table 4-5. Minimum channel sizes required for DDT at a reac-tor scale, based on 15.9% with 104 steam inerting, are relatively small compared to the containment dimensions. The steam generator compartments from El. 129' to El. 152' form vertical setM ght channels. Although both ends are open, there is no transverse venting. This configuration, accord-ing to DDT scaling (Table 4-5) and the assigned geometric class 2 for flame acceleration (from Table C-2), tends to support DDT if flames from the lower level of the lower compartment would propagate upward into these enclosures.

However, the steam generator enclosures are constricted above El.152' 4

ft, (Figure 4-1). This results in reduced channel widths of only 3 q which are much smaller than the minimum channel size (15 ft x 20 ft) re- / quired for DDT to occur. Therefore, it is concluded that DDT cannot be supported and sustained through the steam generator compartments. The DDT potential for the Farley containment is now estimated by first determining the mixture class from Table C-1. The dry basis hydrogen mole fraction conservatively estimated in Section 4.1 to be 15.9%, results in mixture class 3 being selected. Next, the geometric class is selected from Table C-2. This is shown in Tables 4 4 and 4-5 for each region under con-sideration. Table C-3 is now used to combine the mixture class and geometric class results to define the potential results. ihe potential result for lower and annular compartments is result class 4 that DDT is possible but unlikely and per Table C-4 the probability of DDT is assigned a value of 0.10. The steam generator enclosures, i.e., the " upper" part of (

4 14 E F *0* lOst LJ ab / l I v'Of71C4L 1 WEL4. % 5 + 9, ,g, y ag -- 4 % a0LAft "Jitaag tL DS .mU as "u:T -(Artit l M M ~ n-N eDatMTat

• JE. t B6 ust.g spg1A

• 1rf EL 08

• O CE ng p

. s., 1 ....l E i TI Jppe% m cal gg gg, ge e m asa a g mer ~--e y ,, a... _.] F ][ 'w on w g]

== 1 m

n.,. o.

l anssaar sr a = d m m.g _ n. . t 2"2.W st y

== ' = = = (L SS'.# (L 05 0* EL 79' ( tu n *. c Figure 4 1 Vertical section of Farley Containment (taken from Farley System

Description:

Containment Structure and Isolation OPS 40302B(SOT)) O

4-15 ,.Y the lower compartment, is a combination of geometric class 2 and 3 regions; the geometric class 3 region will dominate the potential for DDT. The potential results for the steam generator compartments, using geometric class 3 as an overall class, is result class 3 that DDT may occur with 50% probability.

However, the small dimensions in this region preclude the propagation of DDT.

As mentioned earlier, hydrogen (and steam) pocketing may be possible in the steam generator enclosures. For such cases, the hydrogen molar fraction used in the DDT scaling of Table 4-5 could be increased. This would result in an increased likelihood of DDT propagation through the steam generator constricted area.

However, even if DDT did occur, it would not continue into the upper compartment due to the deceleration as the burn expands into the huge upper compartment volume.

The upper compartment can be con-sidered as an unconfined geometry at large scale (geometric class 5) which is very unfavorable to flame acceleration. The potential reesult for the upper compartment is class 5, which is highly unlikely (to impossible) to q promote DDT. Thus, it is concluded that containment failure due to hydrogen detonation is unlikely to occur in the Farley containment. 4.3 Uncertainties The major uncertainties in assessing the potential for deflagration or detonation of combustible gases in the Farley containment relate to the sources of hydrogen and carbon monoxide considered and the flammability of the containment atmosphere. The flammability of the containment atmosphere depends upon the mole fractions of the combustible gases (hydrogen and carbon monoxide), of the oxygen (air), and of the inertants (nitrogen, steam and carbon dioxide), The evolution of hydrogen, and possibly carbon monoxide and carbon dioxide given corium concrete attack, and the evolution of steam are all time dependent functions. The initiating event and se-quence of key events including potential operator actions or recovery activities can influence the evolution of both combustible gases and iner-tant steam in the containment atmosphere. Thus it is prudent to utilize an q integrated code (MAAF) to study the interactions between the several rate b

4-16 q ) dependent processes. The MAAP code will be used to conduct sensitivity studies to address phenomenological uncertainties as part of the Farley IPE containment response and source term quantifications. The MAAP runs should be reviewed and insights derived regarding the likelihood of deflagrations within the several containment regions. Likewise given the prediction of burns during severe accident sequences, the location and duration of the burns should be noted for use in other assessments regarding the performance of instrumentation within the containment and the performance of containment penetrations. The key uncertainties associated with the sources of combustible gases during postulated severe accidents in the Farley containment are dis ussed below. The in-vessel generation of hydrogen due to the oxidatiot. of zirconium during core overheating, degradation, and relocation represent major sources of uncertainties. Considerable research has been conducted both in the p United States and abroad regarding core melt progression and relocation processes. Additional research is in progress and is being planned. A means of addressing the unresolved issues associated with the complex process of core damage and relocation can be addressed using selected parameters available within the MAAP code. The formation and degree of blockage of the fuel channels during core relocation is a major uncertainty. This can be addressed by considering core melt progression with and without blockage being considered. Likewise the initiation of the relocation of core material impacts the degree of oxidation which may be assessed. This can be addressed by varying the eutectic temperature or core melt relocation temperature considered for such assessments. Lastly, the in-vessel hydrogen generation can be impacted by the duration of vessel failure given that the time available for steam production and oxidation of zirconium while the debris is captured within the primary system affects the hydrogen produc-tion. The ex-vessel sources of combustible gases also include some uncer-tainty. In the case of molten core concrete interactions the amount of Ab

N 4-17 Q) _ concrete attack _ depends upon the success in cooling the debris on the con- / tainment floor. The coolability of core debris is an uncertainty which still remains in the technical community. Different rates of debris coolabil hy can and should be assessed as part of the sensitivity studies performed for the Farley IPE. If debris is rapidly cooled such that core concrete interactions are limited, then the ex vessel sources of hydrogen and carbon monoxide will essentially be precluded. It then becomes impor-tant to assure a supply of water for the long term maintenance of the debris coolability. Ex vessel contributions to hydrogen can also be postulated for-high pressure melt ejection and subsequent direct containment heating se-quences. The major uncertainty associated with high pressure melt ejection is the likelihood of the high pressure existing in the primary system at vessel failure. Emergency operating procedures direct the operators to depressurize the primary system under certain conditions while for other sequences the initiating event itself may depressurize the primary system. Given that some degree of pressurization is present in the primary system when the reactor pressure vessel is postulated to fail, the next uncertainty l that needs to be addressed is the degree of debris dispersal and the amount of material available to be dispersed in a high pressure melt ejection f l event. These two parameters are suitable for consideration in the sen-sitivity studies. i Thus it is recommended that in order to insure that these phenomenological uncertainties are completely addressed and documented that they be incorporated as part of the Farley IPE sensitivity studies.

However, the assessment performed in this paper based on bounding analyses is sufficient to conclude that containment failure due to deflagrations and I

detonations is not expected for the Farley design. The sensitivity studies should help confirm this conclusion and provide additional details and insights regarding the containment response to severe accidents for the Farley containment design. 4.4 Conclusions The assessment provided in this position paper for the susceptibility of the Farley containment to postulated deflagrations and detonations of O i

4 18 combustible gases during severe accidents concludes that the Farley contain-ment would not be failed by this postulated containment failure mechanism. j Potential sources of combustible gases were identified and-discussed' and a bounding pressurization calculation was provided. Based on adiabatic .j isochoric complete combustion of the selected bounding hydrogen inventory, the resulting containment pressure is less than the ultimate capacity of the ') Farley containment. It is therefore concluded that even if one assumed'that barns occurred within the Farley containment they would not result in suffi-cient pressurization to fail the containment. Clearly this is a grossly cons e rvat ive assessment as the actual environment during postulated severe C accident events in containment will produce significant amounts of steam which will inert the containment against hydrogen burns or detonations. It was concluded based on the evidence presented in this summary paper that detonation induced by direct deposition of energy in a detonable mixture of containment gases was not a possible containment failure mechanism. Detonation due to a transition from deflagration to detonation was also considered. Two methods were used to assess the likelihood of the transi-tion to detonation in the Farley containment given the existence of a O deflagration. Both of the methods resulted in a very low likelihood of a DDT. Due to the small size of ignition sources required to initiate a deflagration it is far more likely that combustible gases would be consumed within a containment by deflagration rather than a detonation. k' i ? ( f v

5-1 5.0

SUMMARY

Based upon the assessments provided in this paper and the conclusion that the likelihood of deflagrations or detonations in the Farley contain-ment is very low and improbable, it is concluded that the Farley plant response trees need not incorporate a node for containment failure due to hydrogen deflagration or detonation. It is recommended that to address the range of uncertainties and the uniqueness of specific-event ' timing during 3 postulated severe accidents that sensitivities be performed regarding l hydrogen generation and deflagration. The uncertainties and suggested considerations for selecting cases to be included in the sensitivity study 1 are also provided in this paper. l

r 61 .Q-4

6.0 REFERENCES

's-) Be rman, M., 1986, "A Critical Review of Recent Large-Scale Experiments on Hydrogen Air Detonation", Nuclear Science and Engineerine. 93, 1986, pp. 321-347. Camp A., el al., 1983, Licht Vater Reactor Hydrogen Manual, NUREC/CR-2726, 3 SAND 82-1137, Sandia National Laboratories, Albuquerque, NM. Fauske & Associates,

1990a,

" Technical Support for the Hydrogen Contro11 Requirement for the EPRI Advanced Light Water Reactor Requirements Document". DOE /ID 10290, U.S. Department of Energy. Fauske & Associates, 1990b, " Modifications for the Development of the MAAP-DOE Code - Volume III: A Mechanistic Model for Cembustion in Integrated Accident Analysis". Fauske & Associates,

1992, "Farley Nuclear Plant Units 1

and 2 Containment Data Collection Notebook for Individual Plant Examination." Hertzberg, Martin, 1981, " Flammability Limits and Pressure Development in Hydrogen-Air. Mixtures", Proc. Workshoo on the Impact of Hydroren on-Water Reactor Safety. Volume III, NUREC/CR-2017, SAND 81-0661, Sandia .i National Laboratories. Kumar, R. K., et al., 1984, Intermediate Combustion Studies of Hydrozen-Air-Steam Mixtures, EPRI NP-2955, Electric Power Research Institute, Palo Alto, California. Liu, D. D., et al.,

1981, "Some Results of WRNE Experiments on Hydrogen Combustion", Proc. Vorkshoo on the Imoact of Hydrogen on Vater Reactor Safety.

Volume

III, NUREC/CR 2017, SAND 81-0661, Sandia National Laboratories.

National Research Council Report, 1987,. Technical Aspects of Hydrogen Control and Combustion in Severe ' Licht-Water Reactor Accidents, National Academy Press, Washington, D.C, Ratzel, A. C., 1985, Data Analysis for Nevada Test Side (NTS) Premixed Combustion

Tests, NUREG/CR-4138, SAND 85-0135, Sandia National Laboratories.

i

Shepherd, J.

E.,

1985,

" Chemical Kinetics and Hydrogen-Air-Diluent-Detonations, Tenth International Colloculum on Dynamics of Exelosions and Reactive Systems, Berkely, CA. l .l

'62 r"5s. ( ) Sherman, M. P.. et'a1., 1981, " Deliberate Ignition and Vater Fogs as H Control Measures _for Sequoyah", Proc. Workshoo on the Imoact of Hydronen on Water Reactor Safety. Volume IV, NUREC/CR 2017 SAND 81-0661, Sandia National Laboratories. Sherman, M. P., 1984, Hydrogen Combustion in Nuclear Plant Accidents ~ and Associated Containment Loads", Nuclear Engineering and Design. 82, p. j 13-24. 1

Sherman, M.

P., and Borman, M., 1987, The Possibility of Local Detonation During Degraded Core Accidents in the Bellefonte Nuclear Plant, l NUREC/CR 4803,-SAND 86 1180, Sandia National Laboratories. Sherman, M. P.', Tieszen, S. R., and Benedick, W. B.,

1989, FLAME Facility, NUREG/CR-5275, SAND 85-1264, Sandia Natior.a1 Laboratories.

Stamps, D.,

1987, Sandia National Laboratories, Handout at the Hydrogen Combustion Peer Review Committee Meeting, Washington, D.C., Called by M..Silberberg, Division of Reactor Accident Analysis, U.S. NRC.

Stamps, D.,

'Benedick, W. B., Tieszen, S. R., 1991, Hydroren-Air-Diluent Detonation Study for Nuclear Reactor Safety Analyses, NUREG/CR 5525, SAND 89-2398, Sandia National Laboratories. Tieszen. S. R., et al.,

1987, Detonability of H -Air-Diluent Mixtures, 2

NUREG/CR-4905, SAND 85-1263, Sandia National Laboratories. )' Williams, D. C., et al., 1987, Containment Loads Due to Direct containment Heating and Associated Hydrogen Behavior: Analysis and Calculations with the CONTRAIN Code, NUREC/CR 4896, SAND 87-0633, Sandia National Laboratories, Albuquerque, New Mexico. ? i

.~ A1 I ( APPEKDLE A ) Calculation of H -air-Steam Composition 4 7 1 i This appendix discusses the approximate. method. for evaluating the containment air mixture composition that. is used to determine its flam-f mability. An initially dry containment atmosphere is ' assumed. It is I assumed that the moles of hydrogen in the dry atmosphere is known and that' the pre burn atmosphere (which includes steam) temperature and pressure are l known. Define - initial dry air moles nA P, - initial dry air pressure. T, - initial dry air temperature, n - moles H added g 2 added P - partial pressure of non condensible gases after H2 nc X - a le fra ti nH in the dry mixture H,d 2 T - pre-burn temperature P, - steam partial pressure at T p P - total pressure at T p p n, - moles of H O 2 X ~ *** N2

• I* f#**EI " ~ "H (ng + n, + n )

/ H,w A X, - vet H O mole fraction - n,/(nH + "s * "A} 2 d - relative humidity V - containment free volume The initial amount of air is given by PV IA~1) n ~ A RTo where 1 i 4 R - 8.314 J/mol K \\ 'l i l

A2 Assuming 214'of which is oxygen, we'have moles of oxygen - 0.21 n, and g moles of nitrogen - 0.79 nA The mole fraction of H in a dry mixture is defined as 2 i N b,d~n (^ +n H + Next, the containment temperature and relative humidity, if known, are used to determine the partial pressure of steam P, - ( P, (T ) (A-3) p The partial pressure of the noncondensibles at this new temperature is found ( from the ideal gas law as O T n^ + nH J P (A-4) T P -P J ' b,d o T, A o ne n The sum of these two partial pressures yields the total containment pres-sure: 1 P -P +P (A-5) p s nc t It is likely that P is known instead of 4. In this case instead of p using Eq. (A 3), Eq. (A-5) is solved for P from the known value of P and s p P given by Eq. (A-4). nc For cases of interest, the ideal gas law may be used for the steam mole fraction without significant error, if the steam inerting has high margin (i.e., the resulting mixture is far from the flammability limit). Therefore the steam mole' fraction is O

A3 ,e () X, - P,/P (A.6) p and the hydrogen mole fraction in the wet atmosphere is b,w~b,d(1-X,] (A-7) Civen the mole fractions, Figure 2-3 can be used to determine if the wet mixture is flammable. The post combustion pressure can be found directly from Figure 2-5 if the initial relative humidity is very close to one. Otherwise, the methods of Appendix B can be used. ExJarpl.g (SI Units) Suppose the containment volume is 10,000 m8 It is at 300 K and not inerted initially. Suppose that 100 kg of hydrogen are produced and put into the containment, and steam is generated such that the atmosphere is at (./' 90s humidity at 330 K. Then the initial number of air moles is - 400.9 Kg-moles (A-8) n ~ ~ A 314 0 o and the dry hydrogen mole fraction is "H (100/2) b,d ~ ~ A + "H (401+(100/2)) n 5 At 330 K (about 57'C) the vapor pressure of water is 0.172 x 10 Pa. so that steam partial pressure is 0 P, - (0.9)(0.172 x 10 ) - 1.548 x 10 (A-10) At this temperature the noncondensible partial pressure is 5 (1.0 x 10 ) ~ l'2 (A'II) P nc 0.889 o

A-4 ,-) and the total pre-burn pressure is 5 P - 1.392 x 10 Pa (A-12) P The steam and hydrogen mole f ractions are therefore 5 X, - 1.548 x 104/1.392 x 10 - 0.111 (A-13) Xg,,- (0.111)(1 - 0.111) - 0.099 (A 14) This mixture is flammable according to Figure 2 3. Since the relative humidity is near 1 Figure 2-5 may be used to find the post-combit:; tion pressure. s [Q,)

B-1 ' f~% APPENDIX B Post-Combustion Pressure Rise Calculation This appendix describes the calculation of the adiabatic isochoric complete combustion (AICC) temperature and/or flame temperature given-initial gas masses, temperature, pressure and energy. The AICC temperature is found by equating the end state energy to the sum of the pre-burn state energy and the combustion energy release. A' constant volume process (isochoric) is thus assumed. P .M c T +Q (I'1) M c T g yg g g yg p b ,1-1 ,postburn ,1-1 ,preburn-

where, Q - heat of combustion, b

M - mass of i-th component, g C - specific heat at constant volume of i-th component, yg T - preburn temperature, p Tg - postburn (AICC) temperature, i-H,0,N,H0. 2 2 2 2 The AICC pressure P is estimated by g n RT g g (B-2) Pg-y

where, i

e r

B2 i V - containment free volume, ss t i I R universal gas constant - 8.314 J/mol-K l g - final gas moles. n 1 l I 1 l i

- -~ i C-1 l 3 l 1( i APPENDIX C w S=- ry of DDT Potential Evaluation from NUREC/CR-4803 t I i -l Table C-1 Classification of Hydrogen-Air Mixtures l at 20*C and I ata Pressure l Hydrogen Cell Hydrogen Mixture Mole ~ Equiv. Width Mole Equiv. Class Fraction Ratio Fraction Ratio t in mm 4 y 1 1 1 24-30 .75-1.0 0.787-0.59 20-15 38-30 1.5-1 ( 2 21-24 .63 75 1.575 0.787 40-20 48-38 2.2 1.5 ? E 3 15-21 42.63 12.60-1.575 320-40 63 48 4.1-2.2 { 4 13.5-15 .37.42 47.24 12.60 1200-320 70-63 '5.6-4.1 5 < 13.5 <.37 > 47.24 > 1200 no data > 5.6 i i I l 1 f i O 5 l

.~ I C-2: i 5 // Table C-2 l Geometric Classes for Flame Acceleration l Ceometric Class 1. Large geometries with. obstacles in the path of the expanding unburned gases. Partial. confinement favors gas expansion past the obstacles. A large tube with numerous obstacles and with ignition going i from a closed end to an open end is an example. Class 1 geometries are the most favorable to large flame acceleration. Geometric Class 2. Geometries similar to class 1 but with.some feature which hinders flame acceleration. Examples would be a tube open on both ends or large amounts of transverse venting. i Geometric Class 3. Geometries that yield moderate flame acceleration but I are neutral to DDT. Examples are large tubes without obstacles, small tubes (several inch diameter) with obstacles. Geometric Class 4 Geometries unfavorable to flame acceleration. Examples are large volumes with hardly any obstacles and large amounts of venting transverse to the flame path, or small volumes without obstacles. DDT will not usually occur in a class 4 geometry. i Ceometric Class 5. Geometries so unfavorable to flame acceleration that not even large volumes of stoichiometric hydrogen-air mixtures are likely to f detonate. The only examples are totally unconfined geometry at large scale, or a small spherical geometry without obstacles and central ignition. l i

C-3 ( (._/ Table C-3 Matrix of DDT Potential Results Highly May Highly Results Likelv Likely Occur Unlikelv Unlikelv Mixture Class. 1 2 1 i 1 Geometric Class Very Favorable 1l 1 1 2 3 4 Favorable 2l 1 2 3 4 5 f\\ Neutral 3l 2 3 3 4 5 ( Unfavorable 4l 3 4 4 5 5 Very Unfavorable 5l 4 5 5 5 5 I 1

C-4 7 N.) Table C-4 classification of the Probability of DDT Result Class 1. DDT is highly likely, p _,99 Result Class 2. DDT is likely. p.,90 Result Class 3. DDT may occur. p.,$o Result Class 4 DDT is possible but unlikely, p ,10 ( ) Result Class 5. DDT is highly unlikely to impossible. P . 01 V (~h V

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