NBSIR 75-710

A Small-Scale Enclosure for

Characterizing the Fire Buildup

Potential of a Room

W. J. Parker and B. T. Lee

Center for Fire Research

Institute for Applied Technology

National Bureau of Standards

Washington, D. C. 20234

June 1975

Final

Prepared for Ship Damage Prevention and Control

Naval Sea Systems Command

Department of the Navy
Parker, W. J., Lee, B. T., A small-scale

enclosure for characterizing the fire

buildup potential of a room, Proc^ Meeting

to Honor Clay Preston Butlef' on the

Occasion of his 70th Birthday, Is^xperimental

Methods in Fire Research, Menlo.Park, CA

May 9-10, 1974, pp. 43-72 (Stanf<;vJ; Research

Institute, Menlo Park, CA 1974.

2

492

fillip 76'7/o
NBSIR 75-710

A SMALL-SCALE ENCLOSURE FOR

CHARACTERIZING THE FIRE BUILDUP

POTENTIAL OF A ROOM

W. J. Parker and B. T. Lee

Center for Fire Research

Institute for Applied Technology

National Bureau of Standards

Washington, D. C. 20234

June 1975

Final

Prepared for Ship Damage Prevention and Control

Naval Sea Systems Command

Department of the Navy

U.S. DEPARTMENT OF COMMERCE, Rogers C.B. Morton. Secretary

NATIONAL BUREAU OF STANDARDS. Richard W. Roberts. Director

CONTENTS

Page

1. INTRODUCTION 1

2. PREDICTION MODEL 2

3. EXPERIMENTAL DETAILS 7

4. CONCLUSIONS 13

5. ACKNOWLEDGMENT 14

6. REFERENCES 14

iii

A SMALL-SCALE ENCLOSURE FOR CHARACTERIZING

THE FIRE BUILDUP POTENTIAL OF A ROOM

W. J. Parker and B. T. Lee

A 0.76 by 0.76 m (30 by 30 inch) enclosure with

a 0.61 m (24 inch) high ceiling was used to model

some fires ina3x3x2.4m (10 x 10 x 8 ft)

burnout room. Temperatures, oxygen concentrations,

air velocity, and conductive and radiative heat

fluxes were measured. The highest average air

temperature in the upper part of the room was taken

as a measure of the fire buildup potential of the

room. Upper air temperatures attained in the model

were similar in most cases to those in the fullscale compartment. From energy balance considerations this air temperature was related to the oxygen

depletion in the room and was shown to correlate

well with the oxygen content of the combustion gas

and air exhausting from the model and full-scale

room fires.

Key words: Fire growth; fire tests; flashover; room

fires; scale models; thermal radiation.

1. INTRODUCTION

A major concern at the present time in the fire research

is the questionable ability of laboratory test methods to

measure the fire hazard of building materials under actual

use conditions. The difficulty is probably not so much in

the standard test methods used, as in their interpretation

and application and in the lack of a suitable analytical

model for the fire buildup dynamics in rooms or compartments

of fire origin. Ideally, a test method should quantitatively

determine a relevant fire property of the material such as

ignitability , heat release rate, or flame spread rate under

meaningful simulated exposure conditions. A procedure for

predicting the fire buildup in a room, would then use the

information obtained from appropriate laboratory tests

together with a fire growth model derived from basic combustion principles. Such a procedure could help formulate more

meaningful standards.

A complete analytical description of the fire growth in

a room is too complex for exact solution. It is then necessary to make many simplifying assumptions and approximations

which require that the analysis be performed in conjunction

with experiments on room fires.

1
Full-scale room fire tests are expensive and time consuming. A prediction method could be developed more economically by predicting room fire behavior in a laboratory size

enclosure provided only that the fires in the small size

enclosure simulate those in the full size room; i.e., the

major phenomena are the same and have roughly the same

degree of importance. This correspondence could be verified

if approximate scaling relationships could be established

between the room and the small enclosure.

In order to predict the magnitude of fire severity

there must be some quantitative measure of the level of fire

buildup in a room. A suitable measure appears to be the

temperature of the hot air layer below the ceiling. The hot

air layer, as the term is used in this paper, includes the

flaming and nonflaming gaseous pyrolysis and combustion

products. When this upper air temperature reaches 500 "'C there is rapid pyrolysis and ignition of all combustibles in

the upper part of the room. When it exceeds 700 °C there is

sufficient radiation transmitted from the heated upper

surfaces into the lower part of the room to cause ignition

of all combustible materials. This condition is sometimes

referred to as flashover. The above critical temperatures

are based on observations during the course of these experiments .The fire buildup potential of a room can, therefore, be

considered as the highest temperature that the hot air layer

might achieve during the course of a fire.

This paper discusses a simplified prediction method,

tests in an experimental small-scale enclosure and a comparison of tests conducted in both a full-scale room and the

small-scale enclosure.

2. PREDICTION MODEL

The upper air temperature in a room of given size and

enclosing surfaces of prescribed physical and thermal properties can be estimated by means of an overall energy balance for the enclosed space. In this treatment we are only

concerned with the growth of the fire up to the time of

flashover. The room is assumed to be divided into two

temperature regions with the higher air temperature, Tg, in

the upper part of the room and the ambient air temperature,

Tq, in the lower part of the room. Furthermore, there is a

continuous inflow of cool air into the lower portion of the

room and gaseous combustion products exhausting from the

upper portion through a single open doorway. There is no

other opening (door or window) or forced ventilation. Under

2
quasi-steady state conditions the heat balance equation for

the room can be written as

Rate of

heat

generated

Rate of

heat

Rate of Rate of

heat heat

convected + through

the

absorbed + radiated

by the through

interior the

(1)

opening surfaces opening

The rate of heat generated by the combustible contents

("fuel") is an effective value which is more readily determined for gaseous and liquid fuels than for solid combustibles

for which the effects of volatilization, melting and charring

are somewhat uncertain. If the mass flow rate and heat

capacity of the exhausting combustion products are assumed

to be the same as those of the incoming ambient air, the

convection loss through the opening may be expressed as

DCV (T - Tq) , where D is the density of the ambient air; C

is the heat capacity of air at the temperature Tg,* and V is

the volumetric flow rate of the incoming air which may be

supplied by the ventilation system or be induced by the fire.

The error in this assumption decreases as the amount of available (excess) air increases, and is much less important for

gaseous and liquid fuels. The amount of heat absorbed by the

interior linings is a function of the temperature difference,

(Tg - Tg) , where Tg is the temperature of the hot surfaces in

the upper portion of the enclosure. However, it is assumed in

this model that the heat flux to the boundary surfaces in the

upper part of the room is approximately proportional to the

temperature rise of the upper air, (Tg - Tq) , with a constant

of proportionality, L. Radiation from this upper air layer to

the room surfaces is neglected because of the assumed low

emittance of the air, and so the heat transfer from the hot

air to the enclosure is assumed to be confined to the upper

walls and ceiling. The boundary of this heated region is

defined by the depth of the hot air zone. Then the rate of

heat absorbed by the interior linings becomes LA-^{Tg - Tq)

where A is total area of the hot upper surfaces. The rate of

heat generated within the enclosure is QB, the product of the

effective heat of combustion of the burning material and the

mass burning rate, respectively. If the radiation loss through

the doorway is neglected, eq (1) becomes

The heat loss from the hot air to the upper surfaces of

the room is equivalent to the sum of the conduction heat

loss through the upper walls and ceiling, and the radiation

loss from the latter surfaces to the lower part of the

enclosure

u

QB = DCV(Tgo- T ) + LA (T

u g

(2)

3
When steady-state conditions are approached the conduction loss through the interior coverings can be represented

by the heat conduction equation.

K

u

L (T - T ) = (T - T, ) (3) e g o X s b

where is the thermal conductivity of the solid lining

materials in the upper part of the room, is the temperature of the unexposed (outside) surface of the lining material, X is the thickness of the material, and L is the heat

. c loss coefficient by conduction.

If the surface temperature approaches the critical

value needed for flashover, the radiation loss from the

upper surfaces to the lower part of the room becomes dominant. It is assumed that the temperature in the lower part

of the room remains at T

expressed by o

Then, the radiation loss can be

L (T - r g = f eo (4)

where is the heat loss coefficient by radiation, f is the

view factor between the hot surfaces in the upper walls and

ceiling and the exposed surfaces in the lower part of the

room, e is the emittance of the room surfaces, and a is the

Stefan-Boltzmann constant.

From eqs (3) and (4) one can obtain the following

expression for L,

K

u

T - T, s b

T - T

L g o.

f £0

T - T

g o

4 4 - T

5 o

(5)

This expression for L is used in conjunction with eq (2)

or its further development to estimate the maximum temperature that could be reached in a particular room.

Equation (2) can be rearranged to give the temperature

rise

^^g ^o^ " 1 + LA /(DCV) ^ u

4
Assuming no side reactions in the combustion process,

eq (6) can be expanded to yield

(B/Bg) (QBg/DC^V) (CyC)

*^o^ ~ 1 + (cyC) (L/DC^) (AyA)/(V/A)

where Bg is the stochiometric burning rate, Cq is the heat

capacity of the air at ambient temperature, and A is the

floor area. Equation (7) contains some important groupings

as illustrated in eq (8):

Ye* (Cq/C)

(^g " ^o^ " 1 + (CyC) (L/DC^) (AyA)/(V/A)

where B/Bg is identified as the fraction of the oxygen

depleted from the incoming air and is denoted by Y, and the

quantity (QBg/DC^V) is a characteristic temperature rise,

0*. Y has a maximum value of unity and can be measured

during the course of the room fire tests.

The ratio of the heated area to the floor area, A^/A,

will vary with the room configuration, and the location and

severity of the fire. For a 3x3x2. 4m high room with

the heated air layer extending down to one-half of the

ceiling height, it would be 2.6. For larger rooms it will

be less than this. It is a variable that needs to be

examined as part of the study.

V/A is a ventilation parameter equal to the volumetric

inflow divided by the floor area. For forced air ventilation this would be specified by the designer. For open

doorways it would be proportional to the WH-^/^ factor,

where W and H are the width and height of the doorway,

respectively.

For ease of calculation, 0* is based on methane.

Substituting the calculated values of 0* and C/C^ for air

into eq (8) , the temperature rise for the freely burning

fire in a room will be given approximately by

(rn _ rn ^ = 2470Y , ^^g -^o^ 1 + 705L (AyA)/(V/A) '

5
The numerical values used for calculating the characteristic

temperature rise, 0*, are

3 Q=46.8xl0 J/g for methane,

D = 1.32 X 10"^ g/cm^.

C. = 1.01 g°C

and

B /V = 0.751 X

10""^ g/cm^

for 9.52 volumes of air needed to burn one volume of fuel.

C/Cq is 1.07 at the 700 °C temperature assumed for flashover.

Equation 9 separates the expression for temperature up into

convenient factors which can be measured individually during

the room fire tests.

Y is measured as the oxygen depletion fraction and can

be expressed as

'. ^ " B~ " Dc\ 0*" " DC e* (V/A) ^^i^A^/^) (10)

S O O 1

The summation which is over all of the combustibles involved

in the fire, focuses down on the fire test requirements.

The heat release rate per unit area, q^^, must be obtained

with a heat release rate calorimeter, and the area involved

in the fire, Aj^, requires information from some kind of

flame spread test. Both q^^ and A-^ depend on the radiation

field and this should be reflected in the test methods.

Whether there is an i-th term depends on the ignitability of

the material which is the basis for yet another test.

This expression displays some scaling possibilities.

The ratio of the involved area to the floor area, A^/A,

should be preserved. This requires geometrical scaling

since the i-th combustible may be a wall. However, (V/A)

must also be preserved so one must make adjustments of the

dimensions of the openings to secure the proper airflow.

The height above the doorway is critical to the phenomena

taking place in the room so the height has been scaled

geometrically, and Wh3/2 ^as been controlled by changing the

width of the doorway.

6
The scaling rules used in this work are as follows:

1. All dimensions are proportional to the scale factor

except for:

a. the width of the doorway which is proportional

to the square root of the scale factor, and

b. the thickness of the materials which remains

the same

2. Fuel content and fuel surface area are proportional

to the floor area.

3. Air supply rate is proportional to the floor area.

4. Time is the same in the model as in full-scale.

3. EXPERIMENTAL DETAILS

A quarter-scale model was built of a 3x3x2. 4 metre

(10 X 10 X 8 foot) burnout room with a 2 metre high and a 0.9 metre wide (80 x 35 inch) doorway. The doorway width

was reduced by one-half. The model was made of steel and

lined with low-density asbestos cement board (ACB)

Thermocouples were embedded in the ACB to measure the

heat conduction into the walls and ceiling. The floor was

independently supported on a load cell to measure the burning rate of the combustible contents. The other variables

measured were the vertical temperature profiles in the

center of the model and in the doorway, the oxygen and CO

concentration of the outgoing air, the horizontal component

of the inflow air velocity profile, and the radiation levels

incident on the floor. Along with the radiometer on the

floor, crumpled-up newsprint served as an ignition indicator.

A series of corner tests were run inside of the fullscale burnout room [1].-^ Two 1.2 x 2.4 metre (4x8 foot)

wall panels of the specimen material formed one corner in

the rear of the room. The ceiling above the corner was

lined with another 1.2 x 2.4 metre panel of the same or a

different material. Various fire exposure sources were used

including wood cribs and waste containers. The variables in

the test series were the type of wa!* I material, ceiling

material, and fire source.

Numbers in brackets refer to the literature references

listed at the end of this paper.

7
Two of the above tests were repeated in the model,

which used a simulated wood crib in the form of a gas burner.

The burner was a steel box with a mineral wool cover to even

out the flow. The gas flow was adjusted to produce onesixteenth of the heat release rate of the 6.4 kg (14 lb)

wood crib based on a value of 15.1 x 10^ j/g (6,500 Btu/lb)

for wood. The area of the top surface of the burner was

one-sixteenth of that of the wood crib so that the fuel flow

rates per unit area were the same.

With this simplified type of scaling we are mainly concerned with trying to reproduce upper air temperatures as

closely as possible. Figure 1 shows a comparison of the

time histories of the air temperature at 1 inch from the

. ceiling and at the mid-height of the room for full-scale and

model enclosures with lauan walls and gypsum board ceiling.

The maximum inflow air velocity which occurs near the

bottom of the doorway should be proportional to the square

root of the height of the doorway, therefore, the full-scale

velocities have to be divided by two for comparison with the

model. Figure 2 shows this inflow velocity profile at the

doorway for enclosures having both gypsum board walls and

ceiling. The velocities were measured with a hot wire anemometer in both cases.

The similarities have been encouraging on a very limited

number of tests. We made some comparisons with more materials in the early part of the project using the IITRI

scaling [2] in which the height of the compartment and the

height of the opening were proportional to the 2/3 power of

the scale factor. The agreement was not as good in that

case [3]

The model enclosure was next applied to a realistic

fire risk situation. The prototype compartment fires were

those conducted under the Navy's habitability program of

making berthing spaces more comfortable for the crew without

an accompanying significant increase in the fire risk potential.

The series of nine full-scale tests shown on table 1

was conducted in the 3 x 3 x 2.4 metre (10 x 10 x 8 foot)

burnout room in which the contents consisted of a three-man

bunk with bedding and a three-man locker stuffed with cotton

waste. The bedding included a neoprene mattress, cotton

sheets, a wool blanket and a chicken feather pillow. The

lining materials and ventilation conditions were varied from

test to test. Ignition was by 800 ml of ethyl alcohol in

the middle of the lower bunk. The bedding materials were in

considerable disarray in order to promote as rapid a growth

as possible of the fire, so to produce the worst conditions.

8
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These tests were all duplicated in the small-scale

model using a scaled-down bunk and locker. The bedding was

reduced in area by a factor of 16. Scaling requires that

the thicknesses stay the same because of the importance of

the heat transfer within the material. For example, if

incident radiation levels are the same in the model and the

prototype as they should be, the thickness must be the same

to produce the same surface temperature.

Keeping the materials the same thickness is a great

convenience except for the 3-inch thick neoprene mattress

which could barely be squeezed along with the other bedding

into the space between the tiers of the bunk. This would

constrict the airflow by an unacceptable amount. A compromise was made by using a 1-inch thick mattress in the model,

feeling that it would be satisfactory at least in the early

part of the test.

Table 1 also shows the range of conditions covered by

the nine tests. The first four have progressively increased

ventilation. In all but the fourth test the bunk had closed

ends and back. In that test the back and ends were removed

allowing easy flow of air across the bunk. The standard set

of materials, which included a high melting temperature

polyamide carpet, vinyl coated aluminum bulkhead panels, and

a low density acoustical tile overhead, was used on these

first four tests. Then the lining materials were varied one

at a time. A partial doorway opening and curtains over the

bunk openings were also included in the subsequent tests.

Figure 3 shows the temperature history of the upper air

as determined by a thermocouple 1 inch down from the center

of the ceiling in the model. In the first two tests the

door was closed. The forced ventilation was on in the

second test. Neither led to temperatures above 200 °C. On

the third test with the door open the air temperature reached

nearly 500 °C. This temperature would probably have been

maintained for a while or even increased if the full mattress

thickness could have been used. Except for a localized bunk

fire this was not very spectacular.

Figure 4 shows the same three tests in the full-scale

room. The first two tests again exhibited temperatures less

than 200 °C. The third test barely reached 500 ''C in the

initial stage but it maintained the high temperature long

enough to ignite the cotton waste in the locker in the upper

part of the compartment, causing a very severe fire reaching

a final maximum temperature of nearly ^00 °C.

10
Qualitatively one might say that the model did not

scale well for test 3 (a big fire in the burnout room — not

much in the model). However, both temperatures approached

500 °C which is a danger point. One became fully involved;

the other did not. It could have been deduced from the

model test result that the potential for flashover was

there. This is all that the model is required to do.

Although the scaling relationships including the fixed

thickness requirement for the materials suggest that the

model would operate in real time, the same rate of flame

spread as in the prototype would mean an earlier peak fire

involvement in the model. For rapid fire spread upward,

this time difference for maximum fire involvement may not be

large (see f ig. 1) . However, for situations where much of

the fire spread is along the horizontal plane, the peak fire

development could occur much sooner in the small enclosure

than in the prototype room. This is evident from figures 3

and 4. The peak temperatures prior to locker involvement,

if any, occurred at 6 minutes in the model and 13 minutes in

the full-scale for test 3.

Table 2 provides a comparison of the temperatures 1

inch below the center of the ceiling and the oxygen depletion

fractions at the top of the doorway between the small- and

full-scale tests of all nine situations. The entries exclude

the first peak due to loose sheet burning and later peaks

due to ignition of the locker materials if they occur.

Except for Tests 1, 8 and 4, the agreement is about as good

as could be expected. In Test 1 there may have been sufficient air leakage into the closed compartment to raise the

full-scale temperature. In test 8 the agreement between

full-scale and model was quite good until the model suddenly

increased its involvement very late in time. Test 4 was the

only one which permitted cross flow between the tiers of the

bunk. The reduced spacing in the model due to the thickness

of the bedding materials brought the flames in closer contact with the tier above. This would cause higher temperatures on the underside of the above tier and increase radiation back into the bunk, resulting in a more intense fire,

which might account for the 8 0 °C higher temperature in the

case of the model.

Figure 5 shows a plot of temperature rise versus oxygen

depletion for the nine tests. Only oxygen depletion data

corresponding to peak upper air temperatures are shown.

However, a data display similar to figure 5 also occurs when

the oxygen depletion versus temperature history is plotted

for an individual test. Where there were two successive

peaks in the temperature history, as in test 9, the oxygen

depletion values for both temperature maxima were indicated.

11
Table 2. Comparison of the Maximum Temperatures

and Oxygen Depletions Between the Full-Scale and the Model

Test Size Temperature

( C)

Oxygen

Depletion

(percent)

1 Full 190 3 .

Small 105 3.5

2 Full 170 7 .

Small 165 8.0

6 Full 200 8 .

Small 213 10

5 Full 232 14

Small 203 13

7 Full 244 15

Small 250 12

8 Full 270 13

Small 560 22

9 Full 402 24

Small 397 27

3 Full 506 40

Small 463 23

4 Full 578 39

Small 660 50

12
The open circles denote model data and the solid circles

denote full-scale data. This plot demonstrates the effect

of the thermal losses. The solid line is the theoretical

curve assuming no heat losses while the dashed line is the

theoretical curve taking radiative heat losses into account

with certain assumptions. It assumes that all losses are by

radiation into the lower part of the compartment. It uses a

typical ventilation parameter of 5 cm/s measured for the

open door situation. The product fA /A is assumed to be

equal to unity as it would be if the^surface had an emissivity of unity. For the assumed emissivity of 0.8 it will be

only slightly less. With these assumptions the theoretical

relationship between Y and (Tgo- T ) can be written as

Y =

(T - T ) g o

1 + DC (T - T ) g o

4 4 s o

(11)

which reduces to

(T - T ) _ y = g o C_

0* C

(12)

when the radiation losses are neglected,

values used in the calculation are:

The numerical

and

0*

C

o

C

o

D

= 2,640K,

= 1.01-^

= 1.01 + 10^(T - T )-^,

g o gC

T =

0.8,

300K,

1.32 X 10"-^g/cm-^,

V/A = 5 cm/s

4. CONCLUSIONS

The general agreement between tne full-scale and model

data has demonstrated the usefulness of the model as a tool

for characterizing room fire development. It can facilitate

the development of a prediction method for fire buildup.

13
which in turn would help establish more meaningful fire

standards. Equation (9) provides a useful breakdown of the

prediction method for upper air temperatures into separately

calculable and measurable factors. The correlation between

the temperature rise of the upper air and the oxygen depletion is reasonably good. The temperature versus oxygen

depletion plot also provides a means of comparing the factors

(Tg - T^) , Y , L, and V/A for consistency.

The main thrust of future studies with the model should

deal with principles governing the area of involvement. A.,

which is required in the expression for Y.

6 . ACKNOWLEDGMENT

This work was sponsored in part by the Naval Sea Systems

Command, Department of the Navy, under the direction of

Mr. E. Cherowbrier, technical agent for the Naval Ship

Engineering Center.

7p REFERENCES

[1] Fang, J. B. and Gross, D., Contribution of Interior

Finish to the Growth of Fire in a Room, Fire Safety

Research, Proceedings of a Symposium on Fire held at the

National Bureau of Standards on August 1973, Nat. Bur.

Stand. (U.S.), SP 41] (Nov. 1974).

[2] Waterman, T. E. , Determination of Fire Conditions

Supporting Room Flashover, Final Report IITRI Project

M6131, DASA Contract No. DA-49-146-XZ-475 (Sept. 1966).

[3] Parker, W. J. and Lee, B. T. , Fire Buildup in Reduced

Size Enclosures, Fire Safety Research, Proceedings of a

Symposium on Fire held at the National Bureau of

Standards on August 1973, Nat. Bur. Stand. (U.S.), SP 411

(Nov. 1974)

14
Table 3. Nomenclature

A floor area

burning area of i-th combustible

area of heated surface in upper part of room

B mass burning rate

B mass burning rate required to consume all of the incoming air

C heat capacity of air at the temperature

heat capacity of air at ambient temperature

D density of ambient air

f view factor between the hot surfaces in the upper part

of the room and surfaces in the lower part of the room

H height of the doorway

K thermal conductivity of materials lining upper part of

^ the room

L effective heat transfer coefficient equal to ratio of

heat losses per unit area and the temperature rise,

(Tgo- T )

L heat loss coefficient by conduction

cL„ heat loss coefficient by radiation

Q net heat of combustion of burning materials

q^ heat release rate per unit area of ith combustible

t time

temperature of unexposed surface of lining material

Tg temperature of the hot air in the upper part of room

ambient temperature of air

15
Table 3. Nomenclature (cont'd)

T temperature of hot surfaces in upper portion of room

V volumetric flow rate of incoming air

W width of doorway

X thickness of walls and ceiling

Y fraction of oxygen depleted from the incoming air

e emissivity of surface

0* characteristic temperature rise of air in upper part of

room

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a Stef an-Boltzmann constant

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NBS-114A (REV. 7-73)

U.S. DEPT. OF COMM.

BIBLIOGRAPHIC DATA

SHEET

1. PUBLICATION OR REPORT NO.

NBSIR 75-710

2. Gov't Accession

No.

3. Recipient's Accession No.

4. TITLE AND SUBTITLE

A Small-scale Enclosure for Characterizing the

Fire Buildup Potential of a Room

5. Publication Date June 1975

6. Performing Organization Code

7. AUTHOR(S)

W. J. Parker and B. T. Lee

8. Performing Organ. Report No.

9. PERFORMING ORGANIZATION NAME AND ADDRESS

NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

10. Project/Task/Work Unit No. 4926387

11. Contract/Grant No.

12. Sponsoring Organization Name and Complete Address (Street, City, State, ZIP)

Ship Damage Prevention and Control

Naval Sea Systems Command

Department of the Navy

13. Type of Report & Period Covered

Final

14. Sponsoring Agency Code

15. SUPPLEMENTARY NOTES Order No. N00024-69-F-S177 : Contract No. 6101C-NN531

16. ABSTRACT (A 200-word or less factual sunmtary of most significant information. If document includes a significant

bibliography or literature survey, mention it here.)

A 0.76 by 0.76 m (30 by 30 inch) enclosure with a 0.61 m (24 inch)

high ceiling was used to model some fires ina3x3x2.4m (lOxlOx

8 ft) burnout room. Temperatures, oxygen concentrations, air velocity,

and conductive and radiative heat fluxes were measured. The highest

average air temperature in the upper part of the room was taken as a measure of the fire buildup potential of the room. Upper air temperatures attained in the model were similar in most cases to those in the

full-scale compartment. From energy balance considerations this air

temperature was related to the oxygen depletion in the room and was

shown to correlate well with the oxygen content of the combustion gas

and air exhausting from the model and full-scale room fires.

17. KEY WORDS (six to twelve entries; alphabetical order; capitalize only the first letter of the first key word unless a proper name; separated by semicolons) Fire growth ? fire tests; flashover; room fires;

scale models; thermal radiation.

18. AVAILABILITY Unlimited

I

' For Official Distribution. Do Not Release to NTIS

I I Order From Sup. of Doc, U.S. Government Printing Office Washington, D.C. 20402, SD Cat. No. CH

I ^1 Order From National Technical Information Service (NTIS)

Springfield, Virginia 22151

19. SECURITY CLASS

(THIS REPORT)

UNCL ASSIFIED

20. SECURITY CLASS

(THIS PAGE)

UNCLASSIFIED

21. NO. OF PAGES

23

22. Price

$3.25

USCOMM-DC 29042-P74