i

I An Assessment of Correlations

Between Laboratory and

Full-Scale Experiments for the

FAA Aircraft Fire Safety Program

Part 6: Reduced-Scale Modeling

of Compartments at Atmospheric

Pressure

U S. DEPARTMENT OF COMMERCE

National Bureau of Standards

National Engineering Laboratory

Center for Fire Research

Washington, DC 20234

March 1983

100

, U56

82-2598

1983

Sponsored by:

U S. Department of Transportation

federal Aviation Administration

hnical Center

tic City Airport, NJ 08405

NATIONAL BUREAU OF STANDARDS

JJBRAUT APR r jap

AiuVo^iC v - 7*

NBSIR 82-2598

f 9 * AN ASSESSEMENT OF CORRELATIONS

BETWEEN LABORATORY AND

FULL-SCALE EXPERIMENTS FOR THE

FAA AIRCRAFT FIRE SAFETY PROGRAM,

PART 6: REDUCED-SCALE MODELING

OF COMPARTMENTS AT ATMOSPHERIC

PRESSURE

W. J. Parker

U S. DEPARTMENT OF COMMERCE

National Bureau of Standards

National Engineering Laboratory

Center for Fire Research

Washington, DC 20234

March 1983

Sponsored by:

U.S. Department of Transportation

Federal Aviation Administration

Technical Center

Atlantic City Airport, NJ 08405

U.S. DEPARTMENT OF COMMERCE, Malcolm Baldrige, Secretary

NATIONAL BUREAU OF STANDARDS, Ernest Ambler. Director

TABLE OF CONTENTS

Page

LIST OF TABLES lv

LIST OF FIGURES v NOMENCLATURE vil

Abstract i

1. INTRODUCTION x

2. SCALING RULES 5

3. COMPARISON OF THE RESULTS OF FULL- AND REDUCED-SCALE

ROOM FIRE TESTS WITH COMBUSTIBLE LINING MATERIALS 8

3.1 Room Corner Fire Tests at NBS 3

3.2 Navy Berthing Compartment Fire Tests at NBS 9

3.3 Room Fire Tests at Underwriters Laboratories, Inc 12

3.4 Mobile Home Fire Tests at NBS

3.5 Room Fire Tests of Navy Hull Insulation at NBS 13

3.6 Room Fire Tests at NBS Involving Low Density Cellular

Plastic Foams 34

3.7 Room Fire Tests at NBS Involving Plywood 15

3.8 Room Corner Tests at Upjohn 15

3.9 Room Fire Research Tests at NBS Using Fiberglass Insulation . . 15

4. ANALYSIS i7

5. PHYSICAL MODELING OF AIRCRAFT CABIN FIRES 23

6 . CONCLUSIONS 27

7. ACKNOWLEDGMENT 27

8. REFERENCES 28

iii
LIST OF TABLES

Page

Table 1. Compartment Fire Tests 31

Table 2. Comparison of the Maximum Upper Air Temperature in the

Full and Quarter Scale Navy Berthing Compartment Tests 32

Table 3. Comparison of Room Fire Tests at UL with Quarter Scale Model

and Other Laboratory Tests 33

Table A. Summary of Full- and Quarter-Scale Compartment Fire Tests

of Navy Hull Insulation 34

Table 5. Comparison of Times to Flashover in Full- and Quarter-Scale

Room Fire Tests in the NBS/NRCC Cooperative Program 35

Table 6. Comparison of Times to Flashover and Flameover in Fulland Quarter-Scale Fire Tests of Rooms Fully Lined

with Plywood 36

Table 7. Comparison of Air Temperatures and Heat Fluxes in the

Full- and Quarter-Scale Room Fire Tests with the Room Fully

Lined with Fiberglass 37

iv
LIST OF FIGURES

Page

Figure 1. Comparison of Air Temperature Histories in Full- and

Quarter-Scale Room Corner Tests with Lauan Plywood

Walls and Gypsum Board Ceiling 38

Figure 2. Comparison of Air Inflow Velocity Profiles in Full- and

Quarter-Scale Room Corner Tests with Lauan Plywood

Walls and Gypsum Board Ceiling 39

Figure 3. Upper Air Temperature History in Full-Scale Navy

Berthing Compartment 40

Figure 4. Upper Air Temperature History in Quarter-Scale Navy

Berthing Compartment 40

Figure 5. Comparison of Upper Air Temperature Histories in Fulland Quarter-Scale Mobile Home Bedroom Tests 41

Figure 6 . Comparison of Upper Air Temperature Histories in the

Full- and Quarter-Scale Fire Tests of the Room

Fully Lined with Plywood 42

Figure 7. Comparison of Upper Air Temperature Histories in the

Full- and Third-Scale Fire Tests of the Room Fully

Lined with Asbestos Cement Board 43

Figure 8 . Comparison of Upper Air Temperature Histories in the

Full- and Third-Scale Room Corner Tests of a Foil-Faced Polyurethane 43

Figure 9. Comparison of Upper Air Temperature Histories in the

Full- and Third-Scale Room Corner Tests of a Foil-Faced Polyisocyanurate 44

Figure 10. Comparison of Upper Air Temperature Histories in the

Full- and Third-Scale Room Corner Tests of a Sprayed-On Polyurethane 44

Figure 11. Comparison of Mid-Ceiling Temperature Histories for

Full- and Quarter-Scale Fire Tests in the Room

Lined with Fiberglass at a Scaled Burner Heat

Release Rate of 140 kW 45

Figure 12. Comparison of Upper Air Temperature Histories for

Full- and Quarter-Scale Fire Tests in the Room

Lined with Fiberglass at a Scaled Burner Heat

Release Rate of 140 kW 46

v
LIST OF FIGURES (cont.)

Figure 13. Comparison of Air Temperature Profiles in the Middle of

the Full- and Quarter-Scale Rooms Lined with Fiberglass.

Figure 14. Comparison of Air Temperature Profiles in the Doorway of

the Full- and Quarter-Scale Rooms Lined with Fiberglass.

Figure 15. Comparison of Air Velocity Profiles in the Doorway of

the Full- and Quarter-Scale Rooms Lined with Fiberglass.

Page

47

48

49

Figure 16. Comparison of Scaled Air Velocity Profiles in the Doorway of

the Full- and Quarter-Scale Rooms Lined with Fiberglass. . 50

vi
NOMENCLATURE

a Flame heat transfer coefficient qjVT* (kW/m2, K)

A Area of room (ra2

) A Area of wall (m2

) w

A

c Area of ceiling (m2

)

A^ Area of flame (m2

) A^ Area of floor (m2

)

B Expansion coefficient of gases - 1/T (K *)

0^ Heat capacity of air (kJ/kg»K)

C Heat capacity of room lining material (kJ/kg*K)

D Diameter of pool (m)

E Heat release rate per unit mass of oxygen consumed (13.1 MJ/kg)

f Function Dependence

Fr Froude number - u

2 /gD (dimensionless)

g Acceleration due to gravity (9.8 m/s 2

)

h Height of doorway (m)

h^ Distance from top of doorway down to the neutral pressure plane (m)

H Height of room (m)

H

c Heat of combustion of fuel (kj/kg)

K thermal conductivity of the wall (kW/m*K)

i Flame height above the pool (m)

L Length of room (m)

m’ Mass flow rate per unit width of the air in the boundary layer (kg/m*s)

cl Heat flux per unit area measured at the center of the ceiling (kW/m2

)

q^ Heat release rate per unit area in the flame (kW/m2

)

q'* q" for full scale room

*1 * q" Heat loss rate per unit area of the flame (kW/m2

)

Li

q^ Average heat release rate per unit area of ceiling (kW/m 2

)

vii
NOMENCLATURE (continued)

5" Average heat release rate per unit area of wall (kW/m2

)

Q Total rate of heat release (kW)

Q Rate of heat release from burner (kW)

D

Q' Rate of heat release per unit width of flame (kW/m)

r ratio of the mass of oxygen to the mass of fuel consumed in

complete combustion

S Scale factor

T Temperature (K)

T , Adiabatic flame temperature (K) ad

T Free stream temperature in room and ambient temperature surrounding

the pool fire (K)

T* T - T (K)

T*

ad

T*

1

U

u

00 Uw

v

w

W

y

M.C.

Tad

- T»

(K)

Ceiling temperature (K)

T* for full-scale test (K)

Velocity in boundary layer or in plume above pool (m/s)

Free stream velocity (m/s)

Wind velocity (m/s)

Volumetric air flow rate (m3 /s)

Width of door (m)

Width of room (m)

Distance along surface (m)

Flame length in room (m)

Mass fraction of oxygen in the free stream

Distance from surface (m)

Thickness of lining material

viii
NOMENCLATURE (continued)

AT Temperature rise of flame just above the pool (K)

6 Boundary layer thickness (m)

<f> Q

2 /gD 5 BAT

X Heat loss rate to room surfaces divided by upper air temperature

rise (kW/m2 *K)

v Kinematic viscosity (m2 /s)

p Density of the hot gas (kg/m3

)

Density of air at ambient temperature or in the free stream (kg/m3

)

p Density of room lining material (kg/m3

) s

a Stefan Boltzmann constant (5.67 x 10 11 kW/m2 !

4

)

ix

AN ASSESSMENT OF CORRELATIONS BETWEEN LABORATORY AND

FULL-SCALE EXPERIMENTS FOR THE FAA AIRCRAFT FIRE SAFETY PROGRAM,

PART 6: REDUCED-SCALE MODELING OF COMPARTMENTS AT ATMOSPHERIC PRESSURE

W. J. Parker

Center for Fire Research

National Bureau of Standards

Washington, DC 20234

Abstract

The temperatures, heat fluxes, air velocities, and times

to flashover are compared between a number of previously

reported full- and reduced-scale room fire tests, The model

tests are usually similar but somewhat less severe than their

full-scale counterparts. A simplified analysis is presented

to account for the lower temperatures observed in the models.

Some recommendations are made with regard to physical modeling

of the aircraft postcrash fire.

Key Words: Aircraft fires; fire tests; flashover; reducedscale model; room fire tests

1. INTRODUCTION

The large quantities of aviation fuel carried aboard commercial

aircraft have the potential of causing large pool fires after an otherwise

survivable crash. Thermal radiation and the ingestion of flames through

the open doorways can quickly lead to the buildup of untenable conditions

in the aircraft cabin. The time available for evacuation depends on wind conditions, the size and strength of the fire, the geometry of the

cabin and its openings in relation to the fire, and the materials

comprising the seats and lining materials of the cabin, as well as miscellaneous items brought aboard by the passengers. Ideally, these

1
parameters should be identified and formulated into a mathematical model

which could calculate an escape time. Although a considerable amount of

effort has been put into the development of mathematical models for room

fires [1]*, and to a lesser extent for aircraft cabin fires [2,3,4],

such an analytical model is still a few years away. At the present

time, we must rely on full-scale tests to provide the needed information.

These are costly to run and are particularly difficult in the case of

ambient wind conditions which are difficult to achieve under controlled

conditions in a test bay large enough to house the fuselage of a widebody aircraft. Tests outside are at the mercy of the short-time variability of the winds. It is natural to inquire whether reduced-scale

models can be used to reduce the test to manageable size where wind

velocities can be controlled and many tests can be economically run to

explore the above variables.

There are two reasons to run reduced-scale tests. The first is to

predict what will happen in full-scale, and the second is to provide an experimental basis for developing and checking out the analytical

models.

The most exact form of physical modeling is pressure modeling [5].

If all of the lengths are scaled as the negative 2/3 power of the

pressure, both the Reynolds number and the Froude number are preserved.

In that case the Navier Stokes equations take on the same dimensionless

form for all scales. If radiation can be neglected or can be assumed to

be proportional to the burning rate and if the fuel can be assumed to be

a simple evaporating solid such as Polymethylmethacrylate, the heat and

mass transfer will also scale properly. However the presence of radiation

and char forming solids in real room fire situations has caused some problems with pressure modeling. A pressure chamber required to model a compartment opening into a large air supply plenum would be expensive.

^Numbers in brackets refer to the literature references listed at the

end of this report.

2
Furthermore the requirement that the thickness of the lining material be

reduced by the negative 2/3 power of the pressure would make it costly

to test compartments with composite lining materials such as laminates

and honeycombs which would have to be specially fabricated for the test.

Pressure modeling will not be discussed further in this report , but the

reader is referred to the work of Alpert et al. [6] for more information.

Atmospheric modeling though less exact provides a less expensive

alternative. A large fire cannot be scaled down exactly at atmospheric

pressure because it is impossible to maintain both Reynolds and Froude

numbers constant. The atmospheric models attempt to reproduce the same temperature in the model as in the full scale compartment by requiring

that the total rate of heat production* the total rate of heat transport

by convection through the doorway, and the total heat loss to the

compartment lining materials all be reduced in the same proportion.

Waterman at the Illinois Institute of Technology Research Institute

(IITRI) [7] examined the use of atmospheric modeling of the fire buildup

in rooms with upholstered chairs and couches and with noncombustible

walls and ceiling. These rooms had an opening in one wall. In order to

have the burning items or their gas simulators occupy the same relative area and have the same burning rate per unit area in the model as in the

prototype, it was necessary to require that the total rate of heat

2 release be proportional to the square of the scale factor (S ) . Since

3/2 the ventilation is proportional to wh , where w is the width of the

doorway and h is the doorway height, Kawagoe [8], it is necessary that

3/2 2 wh 'v S so that the air flow is reduced in the same proportion as the

heat release rate. Based on the testing of various alternatives

3/2 Waterman chose to make w ^ S so that h ^ S . Because of the relatively

3/2 small spacing above the door it was necessary to make B. % h ^ S where H is the height of the room. The following scaling rules were

then applied in the study of flashover conditions at IITRI.

3
L 'v W ^ S

3/2 H * S J/

w 'v S

• . -3/

2

fuel supply 'v S

where L, W, and H are the length, width, and height of the room while w

and h are the width and height of the opening.

While scale factors as small as one-eighth worked satisfactorily,

it was recommended that a lower limit of 0.6 m (2 ft) be used for the

height of the model room. The comparison of the temperatures, heat

fluxes, and gas concentrations between the full-scale and model tests

was impressive using only propane gas burners as the fuel source. This

type of scaling would be expected to work satisfactorily for combustible

furnishings but not for combustible walls, since the wall area is

proportional to the scale to the 5/3 power rather than the square of the

scale factor as specified for the fuel input.

The reduced-scale modeling at the Factory Mutual Research Corporation

(FMRC) by Heskestad at atmospheric pressure has used geometrical modeling

of both the room and doorway dimensions [9]. Since the air flow rate is

3/2 5/2 proportional to wh or S where S is the scale factor, the total

* 5/2 heat release rate, Q, is made proportional to S .By making Q

'v 5 / 2

S the flame height is properly scaled. The geometrical modeling

2 insures that the Froude number, u /gL, remains constant since u 'v Jh. ^ /s . However, since the area of the burning items must be proportional to

5/2

S , this scaling procedure is not applicable to burning wall and

2 ceiling surfaces which are proportional S .

The comparisons between the reduced- and full-scale room fire tests

for combustible linings described in this report are based on a set of

scaling principles, originally rejected at IITRI for furniture fires,

but used at National Bureau Standards (NBS) because they have been found

to be the best suited when combustible walls are involved. These scaling

4
rules will be described in the next section. The strengths and weaknesses

of this scaling procedure are brought out in the report. The emphasis

will be on NBS studies of fire development in compartments.

2. SCALING RULES

The quarter-scale model room used in these experiments was based on

a set of scaling rules which attempt to produce the same average gas

temperature in the upper part of the room as the full-scale test [10],

These rules are derived from a quasi-equilibrium energy balance where

the rate of heat production, Q^, by a burner located in the room plus

D

the rate of heat production of the walls and ceiling is equal to the

heat losses through the lining materials of the room and that radiated

through the doorway plus the heat carried out of the doorway by the hot

gases. This balance can be written

+ q"A + q"A = X (T-T ) +

B htf w c c 00 PC V (T-T ) , P

00

( 2 - 1 )

where q^ and q” are the average heat release rates of the wall and

ceiling and A^ and A^ are the respective areas of involvement; T and T^ are the absolute temperatures of the air in the upper part of the room

and the ambient air, respectively; X is the total rate of heat loss by

conduction through the walls and ceiling and by radiation to the lower

part of the room or through the doorway divided by the temperature rise,

T-T^; and p, C^, and V are the density, heat capacity, and volumetric

flow rate respectively of the exhaust air. The temperature rise is,

therefore, given by

T-T00 Qb

+ fow +

5c

A

c pC V + X p

The scaling rules

the numerator and

A^, so that

for the model are obtained by dividing the terms in

denominator of the right-hand side by the floor area,

( 2- 2 )

5
( 2 - 3 )

By preserving the ratios in the brackets as the scale is reduced, the

temperature rise, and, hence, the severity of the fire, remains the same. (Qg/A^) is maintained constant by keeping the gas flow to the burner

proportional to the floor area are kept constant by

using geometric scaling, and X/A^ also is kept constant by geometric

scaling if the fires are similar and the heat transfer coefficients are the same in full- and reduced-scale.

* 3/2 Since V « wh , where w is the width of the doorway and h is its

height, V/A^ can be maintained constant by making h proportional to the

scale factor and making w proportional to the square root of the scale

• 3/2 factor. Actually V ^ wh^ , where h^ is the distance from the top of

the doorway down to the neutral pressure plane. However, if scaling is

successful, h^ is the same fraction of h in the two scales so that the

above relationship holds. The wall above the doorway traps the hot

combustion products from the fire and is critical to the phenomena

taking place in the room, so that this height was chosen to be scaled

geometrically. For quarter-scale modeling, the doorway width then is

half of its full-scale value while the other dimensions are only one quarter, except for the thickness of the wall and ceiling materials.

The wall and ceiling materials must be of the same thickness as in the

prototype, to insure that the heat losses per unit area remain the same for the same interior air temperature. (Small differences in heat

losses are expected since the convective heat transfer coefficient will

be less in the model due to lower air velocities.) This is of great

practical value since materials are tested in the thicknesses available

in the marketplace, and composites do not pose an additional fabrication

problem. This is summarized by the following set of scaling rules.

6
All Dimensions ^ Scale Except

1/2

o Doorway Width (Scale)

o Wall and Ceiling Thickness Same as Full Scale

Fuel Supply Rate 'v Floor Area

Air Supply Rate ^ Floor Area

Time Same as Full-Scale

However, the following problems are encountered with the scaling:

1. since the lateral flame spread rate does not change with scale,

the area covered by flame is relatively too large in the model;

2. the flame heights are observed to be too high in the model;

3. the convective heat transfer coefficient is too low in the model

since air velocities are proportional to the square root of the scale;

4. radiation from the upper part of the room is scale-dependent when

the hot air layer is semi-transparent and a vertical temperature

gradient exists within the layer;

5. flame radiation is quite scale dependent;

6. the increased size of the doorway opening required to scale the

volumetric air flow rate permits slightly greater heat losses and

slightly less heat release from combustible walls; and,

7. during the final approach to flashover, the volume expansion of the

upper layer may account for a significant fraction of the outflow.

This rate of expansion is proportional to the cube of the scale factor,

whereas the buoyancy driven flow is proportional to the square of the

scale factor according to the above rules.

7
Nevertheless, in many tests at NBS with different interior finish

materials from low density foam plastics to high density cellulosics,

both the maximum temperatures reached and the times to flashover* and

flame out of the doorway for the full-scale and the model rooms have

been similar, but usually somewhat longer times and lower temperatures

have been observed in the model. Duplication of the full-scale tests

have been reasonably good which encourages the study of the fire development phenomena on a more economical scale. Presently, some empirical

adjustments to the scaling rules are being examined in an effort to

achieve a higher level of agreement.

3. COMPARISON OF THE RESULTS OF

FULL- AND REDUCED-SCALE ROOM FIRE TESTS WITH COMBUSTIBLE

LINING MATERIALS

3.1 Room Corner Fire Tests at NBS

The first test at NBS of the modeling procedure described in

section 2 was of a room comer test in which the wall in the rear corner

of the full-scale room was formed by two 1.2 x 2.4 m (4 x 8 ft) panels

of lauan plywood. These panels were 6.4 mm (11/64 in) thick. The

remainder of the 2.9 x 3.1 x 2.4 m (9.5 x 10.5 x 7.9 ft) room was lined

with 16 mm (5/8 in) thick Type X gypsum wallboard. There was a 0.90 m

(35 in) wide by 2.0 m (80 in) high open doorway and a 6.4 kg (14 lb)

wood crib in one rear corner as an ignition source. The full-scale test

was part of a series [11] in which the wall and ceiling materials and

the ignition source size were varied in order to compare the performance

of materials in a room fire with their performance in various laboratory

fire tests. This series of tests is reviewed in reference [12] as part

of the correlation of ASTM E 84 with full-scale room fire tests. The

quarter-scale model used a natural gas diffusion burner whose linear

*Flashover is defined here as the time at wt^ich the incident heat flux

on the center of the floor reaches 20 kW/m which is sufficient to

ignite light combustibles.

8
dimensions were one-quarter of those of the wood crib and whose gas flow

was adjusted to release heat at one-sixteenth of the average heat release

rate of the wood crib. The effective heat of combustion of the wood

during the burning phase was taken to be 15 MJ/kg (6500 Btu/lb).

Figures 1 and 2 give a comparison of the air temperature and doorway

velocities between the two scales. These figures give a good indication

of the rough agreement that can be expected with this type of scaling.

3. 2 Navy Berthing Compartment Fire Tests at NBS

Reduced-scale modeling was next applied to a Navy berthing compartment

[13]. 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 thirteen full-scale tests shown in Table 1 was conducted in the 3.1 x 3.1 x 2.1 m (10 x 10 x 7 ft) burnout room with a 0.7 x 1.9 m (27 x 75 in) doorway. The door was either closed, open, or partly open during each test. The contents of the compartment 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 a rapid

growth of the fire and thus produce the worst conditions.

Ten of these tests were 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. However, the 76 mm (3-in) thick neoprene mattress could barely be

squeezed along with the other bedding into the space between the tiers

9
of the bunk. This would have constricted the airflow by an unacceptable

amount. A compromise was made by using a 25 mm (1-in) thick mattress in

the model, feeling that it would be satisfactory at least in the early

part of the test.

Table 1 shows the range of conditions covered by the 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 as variations in the subsequent tests.

Figure 3 shows the temperature history of the upper air as determined by a thermocouple 25 mm (1-in) down from the center of the ceiling

in the first three tests with 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 850°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 might be expected of the model.

Table 2 provides a comparison of the maximum gas temperatures 25 mm

(1 in) below the center of the ceiling during the full- and reducedscale tests. They are listed in order of increasing temperature in the

full-scale test. They can be divided into three groups. When the

temperature was less than 200°C in the full-scale test, it was less than

200°C in the model test. When the full-scale test temperature was between 200°C and 300°C the temperature in the model was also in the same range. However, when the temperatures were above 300°C there were

large differences between the two scales. In two cases the temperature

in the model was higher, and in the other two cases the temperature in

the full-scale compartment was higher. These variations may be largely

due to the complexity of the furnishing array where minor differences

can determine whether or not the fire will progress to the next stage.

Some statistical variation of the results would also be expected if

these tests were repeated on the same scale. It is significant that in

all four tests in the last group the temperatures in the model were

above 450°C indicating that critical conditions were being approached.

This suggests that a repeat of test 8 could lead to flashover.

In these tests, the reduced spacing between the tiers of the bunk

in the model due to the thickness of the bedding materials brought the

flames closer to the tier above thus increasing radiation feedback.

This was counteracted to some extent by the restriction in airflow and

by the earlier burnout of the mattress which had a smaller thickness

than that required by the scaling rules. Thus, there are some limitations

in the ability of the model to predict the full scale fire performance

of complex furnishing arrays.

11
3 .

3

Room Fire Tests at Underwriters Laboratories, Inc.

In 1974 an extensive series of tests with over 30 different types

of materials were run in the tunnel, wall, corner, and room configurations

at the Underwriters Laboratories, Inc. (UL) [14] in order to examine the

relationship between the ASTM E 84 flame spread classification (FSC) and

the performance of the interior finish material in a variety of building

enclosure geometries. Only five materials were run in the 2.4 x 3.6 x 2.4 m (8 x 12 x 8 ft) room with a 9.1 kg (20 lb) wood crib. Although

the FSC correlated with the corner tests it did not do as well with the room tests as seen in Table 3. Materials S and A caused room flameover

much sooner than the plywood even though they had a very low FSC.

Neither the ASTM E 84 tunnel test, the corner tests, the ASTM E 162

flame spread test [15], or the NBS [16] and OSU [17] heat release

calorimeters correlated with the room fire tests. The correlation

between the full scale room fire tests and some previously unpublished

quarter-scale room fire tests also run at UL using a gas burner were satisfactory for the four materials tested on the reduced scale. One

reason that the times were shorter in the model was that the burner was turned on to its maximum gas flow at the beginning of the test while the

wood crib required a considerable fire buildup time.

3.4 Mobile Home Fire Tests at NBS

A series of fire tests was conducted for the Department of Housing

and Urban Development in the bedroom of a single-wide mobile home and in

a quarter-scale model of that bedroom [18]. The objective of the test

was to evaluate the relationship between fire buildup in the reducedscale and full-scale enclosures and thus determine the feasibility of

using a reduced-scale model test to assess the potential contribution of

particular combinations of interior finish materials to fire growth in a mobile home.

The model tests simulated the phenomena of fire growth and flashover.

However, the time to flashover occurred later in the model than in the

12
full-scale bedroom. Flashover was taken as the time at which the level

2 of radiation at the center of the floor reached 20 kW/m which is just

sufficient to ignite light combustibles. Several modifications to the

model were examined, but while showing decided improvements, none adequately corrected the time delay to flashover. These modifications

included (1) the use of propane instead of methane as the burner fuel to

increase the radiation, (2) the use of a triangular burner designed to

increase the flame contact with the walls, and (3) the use of a lower

and wider door opening to increase the thickness of the gas layer above

the door while maintaining the same. In spite of some long delays

in flashover time, the rise in upper air temperature during the early

part of the test was indicative of flashover conditions in the fullscale test.

The temperature history plots in Figure 5 illustrate the relatively

rapid rise in the upper air temperature in the model up to around 500°C

and the rather long delay time prior to flashover.

3 . 5 Room Fire Tests of Navy Hull Insulation at NBS

In the development of a screening test for the fire performance of

hull insulation for the Navy [19], Lee and Breese found that strict

conformance with the scaling rules did not result in flashover in the

model for two PVC acrylonitrile butadiene closed cell foams even though

flashover was observed for the corresponding full-scale room fire tests.

In order to provide a more severe fire exposure in the model, the area

of the wall above the door was increased to 1.4 and 1.8 times its scaled

value. This resulted in lowering the doorway height to 0.93 and 0.86

times its scaled value. The width was adjusted in each case to maintain

3/2 the scaled value of wh . The results of a number of these tests are summarized in Table 4. The lowering of the doorway height increases the

area of the wall exposed to the high temperature gas layer and also

affects the heat transfer to the surface. The amount by which it was

13
lowered was determined empirically. No theoretical basis has yet been

established. Even with these adjustments, the model tests still resulted

in times to flashover as much as twice as long as their full-scale

counterparts. The temperatures were somewhat higher in the full-scale room when the fire buildup was large and somewhat higher in the model

when the fire buildup was low.

3 . 6 Room Fire Test s at NB S Involving Low Density C ellular Pl ast i c Foams

A series of room fire tests was conducted at NBS with fiberglass,

a 65 percent mineral and 35 percent cellulosic fiber insulating board,

and five rigid cellular plastics covaring a large range of FSC in a 2.9

x 3.2 m (9.5 x 10 ft) room 2.4 m (8 ft) high with a 0.74 m (29-in) wide

and 1.9 m (76- in) high doorway [20]. These materials fully lined the

test room during a cooperative research program with the National

Research Council of Canada (NRCC) on the ASTM E 84 tunnel test. The

ignition source was a natural gas diffusion burner located in a rear corner of the room and having a net heat release rate of 79 kW, which is

equivalent to the burner in the ASTM E 84 tunnel. The time to flashover

in the room is compared with that in the quarter scale model in Table 5.

Data on the ASTM E 84 flame spread classification (FSC) and the peak

heat release rate measured in the NBS heat release rate calorimeter [17]

2 at an external radiant exposure of 20 kW/m are included. Plywood (see

section 3.7) and rigid PVC nitrile rubber foam (see section 3.5) are added to the table in order to extend the range of materials. Although

they were not part of this test series, the test room and the gas flow

rates to the burner are about the same for these materials. It is noted

that the times to flashover in the model correlate better than the FSC or the heat release rates with the full-scale behavior. The ranking of

the materials in the quarter-scale ro^m fire tests is identical to that

for the full-scale tests with flashover times on the order of 50 percent

greater than for the full-scale counterparts. The fire buildup for low

density plastic foam material D in the full scale room rapidly reached a peak and died back within 60 s. It finally went to flashover at 368 s

14
when the corner above the burner reignited. The model followed the

initial buildup and decay but did not reignite.

3. 7 Room Fire Tests at NBS Involving Plywood

Two unreported demonstration tests were then run a week apart in

November 1976 of a 3.1 x 3.1 x 2.4 m (10 x 10 x 8 ft) room fully lined

with 4 mm (5/32 in) thick plywood using a natural gas burner in the rear corner with a net heat release rate of 79 kW (4500 Btu/min) . Each

full-scale test was followed by a quarter scale model test on the same

day. As seen in Table 6, the times at which the flames first extended

beyond the doorway were remarkably close between the model and prototype

run on the same day, even though there were differences between the

full-scale tests run on different days. A comparison of the temperature

histories in the upper part of the room for the two scales is shown in

Figure 6 for test 2.

3.8 Reduced Scale Room Corner Fire Tests at Upjohn

The Upjohn Company [21] has also utilized this scaling procedure to

model the 2.4 m x 3.7 m (8 x 12 ft) height ICBO [22] room fire test.

They used a scale factor of one-third rather than one-fourth. A premixed

flame of propane and air was programmed to provide a similar heat

release rate and flame height history as the 13.6 kg (30 lb) crib used

in the full-scale tests at the Underwriters Laboratories. However, the

gas flow rate to the burner was reported to be approximately 12 percent

higher than the scaled value. This was apparently done to make the

3/2 model test slightly more severe. No reduction in wh was made to

account for the air introduced by the premixed burner. Three cellular

plastic foams: a 25 mm (1 in) foil-faced polyisocyanurate, a 25 mm (1

in) foil-faced polyurethane, and a 51 mm (2 in) sprayed-on polyurethane,

were compared between full- and one-third-scale tests. The gas temperature

rise at 13 mm (0.5 in) and 25 mm (1 in) below the center of the ceiling

in the model and full-scale tests, respectively, agreed within + 12

15
percent which they considered acceptable for a screening test. In the

test with an asbestos cement board blank, the temperatures were as much

as 30 percent higher in the model. The temperature histories are shown

in Figures 7-10.

3.9 Room Fire Research Tests at NBS Using Fiberglass Insulation

In order to investigate the observed differences between the two

scales and the effect of the lintel height in the model, Lee conducted a set of full- and quarter-scale room fire tests with the walls and ceiling

lined with fiberglass [23]. The room was 3.1 x 3.1 x 2.3 m (10 x 10 x 7.5 ft) high with a 0.73 x 1.93 m (29 x 76 in) doorway. Three different

doorway heights were used in the model. The burner gas flow rate was varied to represent different degrees of fire buildup. Full- and quarterscale comparisons were made with the burner in the center of the room,

against the rear wall, and in the corner. The test results summarized

here are limited to the corner location for the burner and to two heat

release rates by the burner. In full-scale test P2, the heat release

rate was 140 kW, and the flames were restricted to the region around the corner where the burner was located. In the reduced-scale counterparts,

the flame extended somewhat beyond that but did not reach the thermocouple tree at the center of the room. In full-scale tests P12, the

heat release rate was 460 kW, and the flames covered the ceiling but did

not extend beyond the doorway. In the corresponding model tests, they

extended beyond it. The thermocouples at the center of the ceiling and

102 mm (4 in) below it were outside of the flame zone in the P2 series

of tests and were in the flames in the P12 series. The temperature versus time of the center of the ceiling and of the air 102 mm (4 in)

below it for full-scale test P2 and at 25 mm for its reduced-scale

counterparts are shown in Figures 11 and 12. MI refers to the scaled

doorway height, Mil to 93 percent scaled, and Mill to 86 percent scaled.

The best agreement is obtained for strict adherence to the scaling

rules. The vertical temperature profiles for the center of the room and

the center of the doorway are shown in Figures 13 and 14 for the same

tests. Again, the agreement was best for strict scaling. Figure 15

16
compares the doorway velocities for the same tests. In Figure 16, the

velocity in the model (M2I) is multiplied by 2, which is the reciprocal

of the square root of the scale factor, and compared with the velocity

in the full-scale test. The maximum scaled inflow velocity was about 20

percent higher than the full-scale value, but the outflow velocities

were quite close. Table 7 shows the heat fluxes incident at the center

of the ceiling and the temperature of the hot gas layer 25 mm (1 in)

below it. In neither full-scale test P2 nor its quarter-scale counterparts does the flame from the burner reach the thermocouple. However,

the relatively larger flames in the models may be responsible for their

larger heat fluxes at the center of the ceiling. In the case of fullscale test P12, the flame covers the entire ceiling as it does for its

quarter-scale counterparts M12I, M12II, and M12III. In this case, the

heat flux to the surface is considerably higher than for the models. As

seen in column 6, the heat flux to the surface divided by the black body

radiation based on the gas temperature below it is relatively constant

(0.6 + 0.06) where there is flame impingement. This would suggest that

the increased heat transfer in the full-scale case is related to the

increase in temperature rather than a dramatic increase in the effective

emmissivity. These ratios are higher for the nonflame impingement case presumably due to a higher ratio of convection to radiation. No attempt

was made to separate the convective and radiative components of the heat

flux in these measurements.

4. ANALYSIS

The data from the fiberglass lined rooms in section 3.9 suggest a possible explanation for the reduction in severity of the quarter-scale

test. Because of the relatively longer flame lengths and larger flame

areas in the model, the heat release rate per unit area of the flame is

smaller. Therefore, the temperature and the rate of heat loss per unit area of the flame would be lower and there would be a lower rate of heat

transfer to the surface. This reduction in heat transfer would result

in a lower mass loss rate or heat release rate per unit area of the

17
specimen mater: al. The lower heat release rate per unit area of the

specimen nay be compensated for to some extent in the early growth of

the fire by the larger scaled burning area.

However once the flames cover the ceiling in the full scale room,

they will be extending beyond the doorway of the model for the same scaled heat input and this extension will not be contributing to the

temperature rise in the room.

A simplified analysis will be presented here to qualitatively

account for the dependence of the flame height and the flame temoerature on scale. The full-scale test P12 will first be idealized by assuming

that the -- k' is released from a line burner at the 3.0 m long base cf

the rear wall in order to simulate one dimensional flame spread.

The flame length is calculated by the following model suggested by

Ouintiere [21] . The velocity distribution u(y) in a turbulent boundary

layer on a flat plate with forced flow is given by [25],

where u is the free stream velocity and y is the distance from the

surface. The boundary layer thickness is given by

where is the distance along the surface and v is the kinematic viscosity.

The mass of air flowing in a unit width of the boundary layer is given

by

(4-1)

(4-2)

p u 5 . ao ao (4-3)

18
The heat released per unit width of the flame up to the distance x assuming complete consumption of the oxygen is given by

Q'

Q'

• i n • t m = E m.c. rt m

a 0- a i

p »

u<

= 0.32 P uCO CO 4/5 1/5 4/5

E y.

(0.366)

(4-4)

where H is the heat of combustion of the fuel, r is the ratio of the

c mass of oxygen to the mass of fuel consumed in the reaction, yn

is the

U

2 mass fraction of oxygen in the free stream, and E is the heat released

per unit mass of oxygen consumed (13.1 MJ/kg [26]). Thus, the flame

length is given by

m oo x"5/4 "1/4 - 1 tv x“5/4 ,

*

, .5/4

x

f = (0.32 pJ v u

b

(E yQ

) (Q ) (4-5)

Taking the following values for the properties,

E = 13.1 x 10^ kJ/kg p =1.2 kg/m^ v = 20 x 10 ^ m^/s y

* 5/4 x£ = 0.00218 (Q')

3/ /u

I 00

2

(4-6)

The flame length for the idealized model has the following scale dependence,

(Q') 5M

x£ ^ 1 Oi f u

/ 4

.3/4 (4-7)

This expression is derived by assuming that the total heat release rate

is proportional to the square of the scale, the width proportional to

the scale, and the velocity proportional to the square root of the

% scale. Hence^Jje flame from the burner in the model will be too high by

a factor of — = S or 1.41 for 1/4 scale.

Next assume that the temperature, T, of the flame is given by

T , - T ad

T , - T

ad «

(4-8)

0.233 ,

19
where the drop in temperature from its adiabatic value, T , , divided by

ad

the increase in the adiabatic temperature over the free stream temperature,

T , is equal to the ratio of the rate of heat losses, q", to the rate of

heat production, q”, in the flame. It Is assumed that the heat capacity

is independent of temperature. Then

or letting T* = T - T and T* = T - T 00 ad ad 00

(4-9)

(4-10)

The adiabatic flame temperature for methane is 2236 K, and if the

ambient temperature is taken to be 293 K, then T*^ = 1943 K. The area

of the back wall and ceiling of the fiberglass lined room was 15.9 m^

2

(171 ft ) and the total heat release rate of the gas burner during test

P12 was 460 kW yielding an average heat release rate per unit area of 29

2 kW/m for the flame. The heat flux to the ceiling tabulated in Table 7 is really to a water-cooled heat flux gage. The heat flux to the

ceiling surface would be much less due to the smaller temperature

difference between the hot gas and the surface. In estimating the heat

losses from the flame, the actual net heat flux into the fiberglass

ceiling at 78 seconds can probably be neglected in comparison to the

heat radiated into the lower part of the room. Likewise, the rate of

heat absorption of the walls will be low so that a large portion of the

2 radiation will reach the floor level. If the 13 kW/m measured by a heat flux gage located at the center of the floor is taken as a lower

limit for the losses, then 1092°C is the upper limit for the temperature

according to equation (4-9). The temperature was 820°C measured with a 0.51 mm (20 mil) bare thermocouple 102 mm below the center of the

ceiling. The temperature would be expected to vary some with distance

along the ceiling.

20
Next we determine how this temperature might be expected to change

with scale. The heat release rate from the burner Q is proportional to

D

the area, A, of the room. To the extent that there is a similar involvement on the two scales, the total rate of heat release, Q, will also be

proportional to the area. (This assumption will be violated to some

extent by differences in both the heat release rate per unit area of the

surface and the area of the surface that is covered with flame at a

given time. The assumption will be better early in the test before the

involvement of the lining materials becomes large.) However, in the

experiments discussed here, the total rate of heat release is from the

burner and it is proportional to A on both scales so that

Q E Q % A ^ S

2

(4-11)

D

as assumed in equation (4-7)

The heat release rate per unit area of the flame is given by

«f " ^

' slM (4’12)

Then the heat release rate per unit area of the flame on any scale is

given by

• ii z-,1 / 4 . = S (4-13)

21
where q” is the heat release rate per unit area of the flame in the

full-scale test. Next make the approximation that the rate of heat loss

per unit area from the flame is proportional to the temperature rise.

That is

q" = aT*

JLj (4-13)

Equation (4-10) can now be written

T* = T* (1 - aT* / q '' S

1 ^ 4

) ad f ^ (4-14)

For the full-scale test the temperature rise T* is given by

T* = T*

d

(1 - aT*/

q^ ) (4-15)

Since a/q'' is unknown but assumed to be constant between scales, it can

*1

be eliminated by combining equations (4-14) and (4-15) so that

. 1/4 T*

T* ~ 1/4 » 1 1 + (S ' -1)T*/T*

1 ad

(4-16)

1/4 where T*

d = 1943, S = 0.707 for the quarter-scale* test, and T* was observed to be 800 K. Then T*/T* = 0.80 and T* would be expected to be

643 K. The observed temperature increase for the model with strict

scaling was reported to be 610 K.

Obviously, the oversimplification of the theory, the idealization

of the experiment, and the uncorrected radiation errors in the temperature

measurement do not permit confidence in the fortuitous agreement reported

here. However, this reasoning does indicate and the experiments verify

that one should expect a reduced effective flame temperature and a lower

heat transfer rate to the pyrolyzing surfaces in the model and, thus, a less serious fire. Until the heat release rate in the compartment due

to the combustible linings becomes comparable to the heat release rate

22
by the burner, the average upper gas temperature developed in the model

with strict scaling should be similar but the flame temperature and the

heat transfer to the flame covered surface should be greater in the

full-scale case.

5. PHYSICAL MODELING OF AIRCRAFT CABIN FIRES

In this discussion, two types of Froude scaling will be considered,

both of which scale the aircraft fuselage geometrically. The type

presently used at FAA [27] also scales the doorways geometrically and

will be referred to as "geometrical scaling". The type used in the NBS

compartment modeling will be referred to as "proportional scaling" in

the sense that the induced air flow is proportional to the area of the

compartment. This is accomplished by making the doorway width proportional to the square root of the scale and the height proportional to

the scale.

In order to maintain the same temperature in the model as in the

prototype, the heat produced, the heat absorbed by or conducted through

the room linings, and the heat passing out of the doorway must all be

reduced in the same proportion. For geometrical scaling, the air flow

The 3/2 5/2 through the doorway is proportional to wh and, thus, to S

heat release rate or burning area and the heat losses through the lining

5/2 materials must also be proportional to S and, hence, the heat loss

1/2

per unit area through the lining must be proportional to S . This

2 would be accomplished if Kp C 'v- S and Z ^ K/p C where Z is the

s p s p

S

r s thickness, K is the thermal conductivity, p is the density and C is

^ P

s the heat capacity of the room lining material. For proportional scaling,

2 the air flow and the burning area are proportional to S , so that Kp C

^ ^s and Z should both be independent of scale. If combustible lining

materials are involved, it is essential to have proportional scaling.

23
The fact that the radiation into the compartment from an external

fire is proportional to the doorway area (wh) while the air flow is

3/2 proportional to wh “ will be referred to as the "doorway dilemma" since

their ratio needs to be independent of scale for proper modeling in

either of these cases.

Unless the "doorway dilemma" is counteracted, it will result in

radiation levels—and, thus, temperatures—which are too high in either

model. Note that both of these models obey Froude scaling only so long

as the temperature profiles in the compartment are similar on both

scales. In that case, the velocity is proportional to ^h and h is

proportional to the scale.

External to the fuselage there is also a problem of maintaining the

Froude number. For large pool fires, the heat release rate per unit area becomes independent of the diameter. Thus,

2 2 Q

'x, D ^ S (5-1)

since the diameter of the pool, as well as the fuselage, is taken to be

proportional to the scale (S) in these experiments. According to de Ris

[28], the heat release rate (Q) of the pool fire can be written as

pC T*

P

(5-2)

where p is the density of the fire plume gas, is its heat capacity,

T* is its average temperature rise over ambient, D is the diameter of

the fire, and u is the average upward velocity of its plume gas.

Thomas [29] gives the following expression for the flame height, l,

D

(5-3)

24
where B is the expansion coefficient of gases, 1/T; AT is the temperature

rise of the flame just above the pool; and g is the acceleration due to

gravity. For geometric similarity of the flame:

gD

5 BAT

= constant = 4> (5-4)

Combining (5-2) and (5-4), the Froude number is given by

2 Fr = gD

16({>BAT

,Vc

P

V

(5-5)

which can be considered to be independent of scale. Furthermore, from

(5-3) the flame height (£) is proportional to scale:

Z ^ D ^ S (5-6)

However, this requires that

Q * D

5 ' 2 * S5/2 (5-7)

as seen from (5-4)

.

Because of the conflict between (5-1) and (5-7) , it is impossible

to maintain Froude scaling, as pointed out by Eklund [27], as long as the same liquid fuel is used for the model and the prototype. However,

if a gaseous fuel were used in the model, it could be controlled so that

5/2

Q ^ S and similarity of the flame and the Froude number could be

maintained. If the emissivity of the gas flame in the model were less

than that of the flame from the aviation fuel in the full scale fire,

the radiation into the small scale compartment would be reduced. This

would help to counteract the "doorway dilemma" discussed above. Hence,

the fuel gas should be selected on the basis of its radiating properties.

25
The ambient wind must scale as

"W

'v D

1 ^ 2

'v S 1 ^ 2 (5-8)

to satisfy Froude scaling.

In light of the above discussion, the following set of scaling

rules are recommended.

(1) The fuselage and fuel source dimensions should be proportional

to the scale except for the thickness of the fuselage, which should

be Independent of scale.

(2) The doorway heights should be proportional to scale.

(3) The doorway widths should be proportional to the square root of

the scale.

(4) The rate of heat release (Q) of the simulated pool fire should

be proportional to the scale to the 5/2 power.

(5) The emissivity of the flame should be proportional to the square

root of the scale.

(6) The wind velocity should be proportional to the square root of

the scale.

The above set of scaling rules apply to the whole system. If the

impact of the lining materials on the fire growth is to be singled out

for evaluation, then the problem can be uncoupled by replacing the

burning seats with a gas burner adjusted to produce the same scaled rate

of heat output as in a full scale cabin fire. That rate could be

determined from mass loss rate measurements in the full scale fire along

with effective heat of combustion measurements of the seats in a large

scale calorimeter such as the NBS furniture calorimeter [30], Proportional scaling should be used.

26
6. CONCLUSIONS

Until mathematical modeling becomes sufficiently advanced to make

reliable predictions based on fire property data, the reduced-scale

physical model probably represents the best indication of full-scale

behavior of any laboratory test method. However, the model is generally

less severe in terms of maximum temperatures and times to flashover than

the full scale test. Thus, important conclusions found from the model

tests need to be verified by a limited amount of full-scale testing.

The reduced-scale model can also be helpful in the development, verification,

and refinement of the mathematical models which should be able to predict on small- as well as full-scale.

Proportional scaling (i.e., the area of the doorway times the

square root of its height is proportional to the area of the enclosure)

is required when the interior surfaces are involved in combustion. It

also is necessary in order to take heat conduction losses through the

lining materials properly into account unless the lining materials can be changed with scale as required by geometrical scaling. The burning

of interior items, such as seats, can be done with geometrical scaling,

5/2 provided the burning areas are adjusted to be proportional to S .

The proportionally scaled quarter-scale room fire test has always

been somewhat less severe than its full-scale counterpart, and a plausible

explanation is presented in this report. Provided the fire is not

ventilation limited, the geometrically scaled test would be expected to

be somewhat more severe than the proportionally scaled test because of

the reduction in air flow. Neither type of scaling can be counted on to

exactly duplicate a full-scale test.

7 . ACKNOWLEDGMENT

This work was supported by the Federal Aviation Administration as part of a more comprehensive review of the correlation between fullscale tests and laboratory fire test results. Mr. Richard Hill was the

contract monitor.

27
8. REFERENCES

[1] Levine, R. S., "Mathematical Modeling of Fires," NBSIR 80-2107, Nat. Bur.

Stds., Washington, DC (September 1980).

[2] Howell, W. , Coordinator, "Workshop on Mathematical Fire Modeling,"

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[3] Reeves, J.B. and MacArthur, C., "Dayton Aircraft Cabin Fire Model,

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[4] Quintiere, J. and Tanaka, T., "An Assessment of Correlations Between

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[6] Alpert, R. L. , Matthews, M. K. , and Modak, A. T., "Modeling of Ceiling

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[7] Waterman, T. E. , "Room Flashover—Scaling of Fire Conditions," Fire

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[9] Heskestad, G. , "Modeling of Enclosure Fires," Fourteenth Symposium

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[10] Parker, W. J. and Lee, B. T. , "A Small-Scale Enclosure for Characterizing

the Fire Buildup Potential of a Room," NBSIR 75-710, Nat. Bur. Stds.,

Washington, DC (June 1975)

.

[11] Fang, J. B., "Fire Buildup in a Room and the Role of Interior Finish

Materials," TN 879, Nat. Bur. Stds., Washington, DC (June 1975).

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.

[13] Lee, B. T. and Parker, W. J., "Naval Shipboard Fire Risk Criteria

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76-1052, Nat. Bur. Stds., Washington, DC (September 1976).

28
[14] Castino, T.G., et al. , "Flammability Studies of Cellular Plastics

and Other Building Materials Used for Interior Finishes", Subject

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Part 18, American Society for Testing and Materials, Philadelphia, PA.

[16] Parker, W. J. and Long, M. E. , "Development of a Heat Release Rate

Calorimeter at NBS," Ignition, Heat Release, and Noncombustibility

of Materials, ASTM STP 502, American Society for Testing and Materials,

p. 135-151 (1972).

[17] Smith, E.E., "Measuring Rate of Heat, Smoke, and Toxic Gas Release,"

Fire Tech. (1972) 8, p. 237.

[18] Klein, D. K. , "Reduced-Scale Modeling of Mobile Home Fires: A Progress

Report," NBSIR 81-2333, Nat. Bur. Stds. , Washington, DC (December 1981).

[19] Lee, B. T. and Breese, J. N. , "Submarine Compartment Fire Study—Fire

Performance Evaluation of Hull Insulation," NBSIR 78-1584, Nat. Bur. Stds.,

Washington, DC (May 1979).

[20] Parker, W. J. , "Comparison of Selected Room Fire Tests with ASTM E 84 and

Other Laboratory Test Methods", A cooperative program with the National

Research Council of Canada. NBSIR, Nat. Bur. Stds. (to bpublished)

.

[21] Chlem, W. P. , Nadeau, H. G., and Waszeciak, P. H., "Development of a Small

Compartment Corner Test Method for the Preliminary Evaluation of Fire Performance of Cellular Plastics," J. of Cellular Plastics, Vol. 16,

No. 4, p. 212-222 (July /August 1980).

[22] ICBO Acceptance Criteria for Foam Plastics, International Conference of

Building Officials, 5360 South Workman Mill Road, Whittier, CA 90601

(December 1976).

[23] Lee, B. T., "Quarter-Scale Modeling of Room Fire Tests of Interior Finish,"

NBSIR 81-2453, Nat. Bur. Stds., Washington, DC (March 1982).

[24] Quintiere, J. G. , NBS private communication.

[25] Schlichting, H. , "Boundary Layer Theory," 4th Edition, McGraw-Hill,

p. 536-537 (1960).

[26] Huggett, C. , "Estimation of Rate of Heat Release by Means of Oxygen

Consumption Rate Measurements," Fire and Materials, Vol. 4, No. 2,

p. 61-65.

[27] Eklund , T. I., "Feasibility of Froude Modeling a Pool Fire External to

an Aircraft Fuselage," FAA/NAFEC Report No. NA-79-21-LR, Federal Aviation

Administration, National Aviation Facilities Experimental Center, Atlantic

City, NJ 08405 (July 1979).

29
[28] de Ris, J. , "Modeling Techniques for Prediction of Fires," presented at

Conference on Polymeric Materials for Unusual Service Conditions, NASA

Ames Research Center (November 1972).

[29] Thomas, P. H., Webster, C. T. , and Raferty, M. M. , Combustion and

Flame, Vol. 5, No. 4, p. 359-367 (December 1961).

[30] Babrauskas, V., Lawson, J.R., Walton, W.D., and Twilley, W.H.

,

"Upholstered Furniture Burning Rates Measured with a Furniture

Calorimeter", NBSIR, Nat. Bur. Stands, (to be published).

30
Table

1.

Compartment

Fire

Tests

X CO (0 <0 <0 (0 (0 3 3 3 X

XX

X 4-4 X *3 3 3 •3 *3 3 "3 3 o o 3 CJ

CO

X

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•3 -3

X •3 3 3 3 3 3 3 3 X 3 3 3 3 3 3 3 3 3 3 3 3 CO CO CO 3 3 3 CO 3 CD 3 3 3 3 3 O O O o 3 3 O O O O O O O

PQ

X

X rHX CP £

rH rH i—1

rH rH rH rH

CJ

U CJ

U

O o CJ CJ o CJ CJ <j

P’5 rH

co- 4H 3

•a 4H 4H 4H 4-1 op 4-1 13 3 3 3 3 3 3 3 44 4-1 44 44 3 o <!5X 0 0 0 0 5 O O o O o

CO 30

CO 4-4 1 CO

r-l 0) X H 3

Pm

X)3

CO 33

CD

g£ 33ao g

cp o g& Q o O 3 3 3 3 3 3 3 OJ 04 04 04

O

rH rH CP CP CP CP CP

O

CP \ \ X o CJ O O O o o O X X X X 33 3 3 X 00

CO 3 3 X Co 3 X 3 X 4-1 3 3 X X X 3 3 • 3 X 3 O a CJ 3 3 00 3 3 X 3 PH 3 3 3 3 X 3 3 3 4-4 rH Q Q Q O 3 3 X X <4H X X S 3 C4H Pi o 3 3 3 3 3 3 X 3 Pi Pi X 3 X 3 X U X O o O 3 3 3 o CJ CP 3 X 4-4 Pi X 3 3 CP > pH 3 O 3 4-4 3 3 4) 3 3 3 3 PQ O 1 Q CP o CO > 3 3 3 3 3 3 o X 3 3 PH PH CP CP Pi 3 3 3 3 3 X 3 PC O 3 3 3 3 O 3 3 X Pi 3 O 3 3 3 3 > X X PQ 3 <i 3 X <J 3 3 3 3 O 4-4 3 CP 3 X 3 3 4-1 3 3 3 3 4-4 3 3 Pi s 3 X 4J 4-) 4J 4-4 4-1 4-4 fo 3 CD CP Pi 00 e^s CP 3 U a 3 X 3 3 3 3 3 3 4-4 3 O 3 4-4 4-1 4-1 3 P4 3 3 o > 3 3 4-4 CO CO CO CO CO CO •H 3 CJ 3 3 3 3 X 3 CJ X 04 3 6 X X CO 3 3 3 CP CP CP 3 CJ 3 1 3 4-4 3 3 3 3 3 3 3 3 X 3 XS Pi Pi Pi 4-4 CO 3 3 X 3 3 O 3

Pi Pi Pi Pi Pi Pi 3 Pi 3 X 3 3 3 Pi 00 CJ 3 CP o 3 X 3 3 3 3 3 3 Q 3 X i—i

CJ CJ CJ 3 3 X 3 O o 4-4X • 3 3 3 3 3 3 3 > a 3 CJ X X 3 3 3 3 6 3 3 3 3 3 3 X 6 3 PQ X X X 3 >5 X X CP 4-4 3 3 o 3 3 3 3 3 3 PC X o o o 3 3 Pi CJ 3 S'? Pi Pi 3 X 4-4 4-) 4-4 4-1 4-4 4-1 •u X 3 0} 3 o o o 3 CP O 3 3 o 3 3 3 3 4-1 CO CO CO CO CO CO X O a O 13 13 :s 3 O <fl Q PQ 00 O CJ 3 3 O

04 < PQ CO CO

<3 PQ <f s3- lO cO *a*

00 05

K*XO *X3 4-4 3 3 mi in O 3 X 04 OO CO <r uo vO r- 00 05 a H a a a a a a a a a a

X33

Xiao4J4JoX

<U X!

4-1 so

CO3oX

4-1 X3

00 X33N

Pi

4-) 4-1

4-1 3 3 a) B >

xl

00 4-4 3 oo X 3

CD •r4 i—I 3

CJ 003 S O O X

Pi 3

M—I 3

?o

O4-) •H3

00 x

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31

Low

Density

Acoustical

Panels

on

Overhead

Fibrous

glass

Insulation

on

Two

Bulkheads

Vinyl

Coated

Panels

on

Two

Bulkheads

High

Temp

Polyamide

Carpet

Bonded

to

Steel

Deck
Table 2

Comparison of the Maximum Upper Air Temperature in the Full

and Quarter Scale Navy Berthing Compartment Tests

Test Scale Temperature (°C)

2 Full 172

Quarter 168

1 Full 191

Quarter 107

6 Full 200

Quarter 209

3B Full 230

Quarter 202

5 Full 232

Quarter 265

7 Full 2AA

Quarter 250

8 Full 354

Quarter 706

4B Full 570

Quarter 875

9 Full 800

Quarter 531

3A Full 850

Quarter 463

32
Comparison

of

Room

Fire

Tests

at

UL

with

Quarter

Scale

Model

and

Other

Laboratory

Tests

d §

•H

O

cd

u

a) 4) > rH CO o « w

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cd I rH

Fn •"***

CM

co § pq o

25 v£>

cm

u ^0

cd rH rH co CO

O

CO CO • CM

PQ o 00

V

V rH

SC CO

CM

•u 0 O

cd <3 rH <3 Ov 3 S3 CM rH 33 rH

S3 ,5 rH

coO

A

O CO

CMVOW

U

CD>o

CD

pH

n

CDC

m £ om m

CM

CO

VOUO vO

rH

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uorco

co <3

S5 co £

T-V rv

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m pH PH

O CO 8 8 8 8 00 CJu V-/

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CM Ph Ph

CD CO CT> P. • • O CM rH v^ v-/

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CO CM CM CM CM X3

pH r—

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CD T3 CO <3 e> ed SC II II II II O

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33

Fire

Retardant

Treated

Plywood

(12.7

mm)

(0.5

in)
CO3H3U

•Hab

0)6

4-) u3 C

1-5 O 4J CJ 3H3 3

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m

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34

at

25

mm

below

center

of

overhead

at

25

mm

below

top

of

doorway
Comparison

of

Times

to

Flashover

in

Fulland

Quarter-Scale

Room

Fire

Tests

with

Laboratory

Fire

Tests

in

the

NBS/NRCC

Cooperative

Program

x£

0

0) —

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cov

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3

0) *

35

the

gas

flow

to

the

burner

was

increased

by

50%,

flashover

occurred

in

53
Table 6

Comparison of Times to Flashover and Flameover

in Full and Quarter Scale Fire Tests of Rooms

Fully Lined with Plywood

TEST FLAME OUT DOORWAY (sec) FLASHOVER (sec)

Full-Scale #1 190 *

Model #1 188 *

Full-Scale #2 158 156

Model #2 158 185

*Fire was extinguished before flashover occurred.

36
Table 7

Comparison of Air Temperatures and Heat Fluxes in the Fulland Quarter-Scale Room Fire Tests with the Room Fully

Lined with Fiberglass

Test*

Scaled Heat

Release Rate

of Burner (kW)

Air Temperature

(°C)**

Heat Flux at

Center of Ceiling

4" (kW/m2

) c

Time to

Flashover

(a)***

4c

P2 140 374 5.2 OO 0.52

M2I 140 365 11.0 00 1.20

M2II 140 428 12.0 OO 0.92

M2III 140 413 11.0 OO 0.92

P12 460 820 51.0 78 0.64

Ml 21 460 630 25.0 138 0.68

Ml 21

1

460 719 33.0 102 0.61

M12III 460 721 31.0 90 0.57

*P = full-scale test

MI = model test with doorway height 0.25 x full-scale

Mil = model test with doorway height 0.23 x full-scale

Mill = model test with doorway height 0.22 x full-scale

**Measured 25 mm below center of ceiling at time of flashover if it occurred;

otherwise at 240°C.

***Time at which heat flux to floor reached 20 kW/m2

.

****Ratio of heat flux to the center of ceiling to the black body radiation

at the temperature in column 3.

37
</> UJ

Do ‘3dnivaadW3i

38

Figure

1.

Comparison

of

Air

Temperature

Histories

in

Full

and

Quarter-Scale

Room

Corner

Tests

with

Lauan

Plywood

Walls

and

Gypsum

Board

Ceiling
39

FigurG

2.

Comparison

of

Air

Inflow

Velocity

Profiles

in

Fulland

Quarter-Scale

Room

Corner

Tests

with

Lauan

Plywood

Walls

and

Gypsum

Board

Ceiling
1000

r--I11

-i1

1000 Do ‘aunivuidMai

oo fc.S O J=

*J

U

in P

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03

H

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60

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eg co l * III J L S

eg §

Go ‘ 3dniVH3dW3I

>>pO 60

LJ 3

CO *H

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SB 4JP

0) 03 p pa 33 £ P 3

03 Z

e-a 03

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3) *H

co

03p3

60

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40
0001

(Go) 38fUWfldW3i

ih-

&M

£

M g

t» *

« s

£ &

E E

o

(Go) auniV83dN31

41

Figure

5.

Comparison

of

Upper

Air

Temperature

Histories

in

Full-and

QuarterScale

Mobile

Home

Bedroom

Tests
0 40 80 120 160 200

TIME (s)

Figure 6. Comparison of Upper Air Temperature Histories in the Fulland Quarter-Scale Fire Tests of the Room Fully Lined with

Plywood

42
1000 to 9

Go ‘auruvHdMSi gniigg-cum

S ^ 0*3 CN! ^

Oo ‘aamvaadimai ONnoo-onm

43

Figure

7.

Comparison

of

Upper

Air

Temperature

Figure

8.

Comparison

of

Upper

Air

Temperature

Histories

in

the

Fulland

ThirdHistories

in

the

Fulland

ThirdScale

Fire

Tests

of

the

Room

Fully

Scale

Room

Corner

Tests

of

a

FoilLined

with

Asbestos

Cement

Board

Faced

Polyurethane
uqordn

-I

nn#

I

uyofdnfA

3 S? § 3 k

Do ‘3dniVU3dN31 3NI1I30-0IW

S

0)

1-4 3

4J

«o

1-4 0)a

0 X3

0) c

E-* <0

I XJu

•H-CH

CO 0)

t-4 1 •U c

•H rH CO CO < rH a) J3 H 4J

)-i 0)

a) M C4 a <U a) 3

a jc

£

5k

33 4J i—

1

o O

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0 5

o a> o i CO •H OS T3

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cO 43 i—

i

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0 3

o a

o CO co O

0)P

3

60

•H

Cl4

JCH

COM

0)a

II H cO

M O

C0 +J a w

0 *H O 33 CJ

<J iH3

5-4 0L

CDa 0) a j=

33 -u

U-4 C

O *H

C4-4 o

CO

4-1 CO

0)

4-1 ai co H ^

3

M c

0) c0 0 5k M O

o o O CO •H s 5k

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03 Pa; td

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co u

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O',

aiM

3

60

•H

44
Figure 11. Comparison of Mid-Ceiling Temperature Histories for

Full- and Quarter-Scale Fire Tests in the Room

Lined with Fiberglass at a Scaled Burner Heat

Release Rate of 140 kW

45
1

3

eo

9

CM

CM

(A

3o ‘3dniV»dM3i

46

Figure

12.

Comparison

of

Upper

Air

Temperature

Histories

for

Fulland

Quarter-Scale

Fire

Tests

in

the

Room

Lined

with

Fiberglass

at

a

Scaled

Burner

Heat

Release

Rate

of

140

kW
Figure 13. Comparison of Air Temperature Profiles in the Middle of

the Full- and Quarter-Scale Rooms Lined with Fiberglass

47
Figure 14. Comparison of Air Temperature Profiles in the Doorway of

the Full- and Quarter-Scale Rooms Lined with Fiberglass

48
PERCENT

DOORWAY

HEIGHT

100-1 m

• a

90- Ok A

•

80- £> A

•

70- 4P A

•

60-

!

- 0*6

1 ao - a®

-50

*° - -40

4® - -30

- -20 • M2 1 o m ii ** - -10 a M3 III a P2

i

-1.0 0 1.0 2.0 3.0

VELOCITY, m/s

Figure 15. Comparison of Air Velocity Profiles in the Doorway of

the Full- and Quarter-Scale Rooms Lined with Fiberglass

49
100

90

80

70

60

50

40

30

20

10

0

Fitiur

VELOCITY, m/s

16. Comparison of Scaled Air Velocity Profiles in the Doorway of

the Full- and Quarter-Scale Rooms Lined with Fiberglass

50
MBS-114A (REV. 6-76)

U.S. DEPT. OP COMM.

BIBLIOGRAPhIC DATA

SHEET

\ 1. PUBLICATION OR REPORT NO. NBSIR 82-2598

2,0ov*.t Accotslon No. 'll

4. TITLE AND SUBTITLE

An Assessment of Correlations Between Laboratory and

Full-Scale Experiments for the FAA Aircraft Fire Safety

Program, Part 6: Reduced-Scale Modeling of Compartments

at Atmospheric Pressure

5. Publication Date March 1983

T. Performing Organization Code

7. AUTHOR(S)

W. J. Parker

8. Performing Organ. Report No. DOT/FAA/ CT-82/160

9. PERFORMING ORGANIZATION NAME AND ADDRESS

NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE

WASHINGTON, DC 20234

10. Project/!ask/Wcrk Onit No.

11. Contract/Grant No.

12. SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (Street, city, State, ZIP) U. S. Department of Transportation

Federal Aviation Administration

Technical Center

Atlantic City Airport, NJ 08405

13. Type of Report & Period Covered

October 1982

15. SUPPLEMENTARY NOTES

| 1 Document describes a computer program; SF-185, FIPS Software Summary, is attached.

16. ABSTRACT (A 200-word or leas factual summary of moat significant information. If document includes a significant bibliography or

literature survey, mention it here.)

) The temperatures, heat fluxes, air velocities, and times to flashover were

compared between a number of previously reported full- and reduced-scale room

fire tests. The model tests were usually similar but somewhat less severe

than their full-scale counterparts. A simplified analysis is presented to

account for the lower temperatures observed in the models. Some recommendations are made with regard to physical modeling of the aircraft postcrash

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)

Aircraft fires; fire tests; flashover; reduced-scale model; room fire tests.

18. AVAILABILITY Unlimited

I | For Official Distribution. Do Not Release to NTIS

I | Order From Sup. of Doc., U.S. Government Printing Office, Washington, DC

20402, SD Stock No. SN003-003-

[X~| Order From National Technical Information Service (NTIS), Springfield, VA. 22161

19. SECURITY CLASS

(THIS REPORT)

UNCLASSIFIED

20. SECURITY CLASS

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21. NO. OF

PRINTED PAGES

59

22. Price

USCOMM-DC