Reference

NBS

PUBLICATIONS AlllDb 033700

NBSIR 84-2912

The Thermal Response of Aircraft

Cabin Ceiling Materials During

A Post-Crash, External Fuel-Spill,

Fire Scenario

Leonard Y. Cooper

U.S. DEPARTMENT OF COMMERCE

National Bureau of Standards

National Engineering Laboratory

Center for Fire Reserch

Gaithersburg, MD 20899

October 1984

100

.U56

84-2912

1984

Sponsored by:

U.S. Department of Transportation

Federal Aviation Administration

_9chnical Center

tiantic City Airport, NJ 08405
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VAnomZ Rrrpr.,y

NBSIR 84-2912

THE THERMAL RESPONSE OF AIRCRAFT

CABIN CEILING MATERIALS DURING

A POST-CRASH, EXTERNAL FUEL-SPILL,

FIRE SCENARIO

Leonard Y. Cooper

U S. DEPARTMENT OF COMMERCE

National Bureau of Standards

National Engineering Laboratory

Center for Fire Reserch

Gaithersburg, MD 20899

October 1984

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
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TABLE OF CONTENTS

Page

List of Figures v

Abstract 1

1. INTRODUCTION 2

1.1 Purpose of the Investigations 2 1.2 Description of the Tests 2

2. AN ANALYSIS OF THE THERMAL RESPONSE OF THE CABIN CEILING

MATERIAL 6

2.1 Components of the Surface Heat Flux 6

2.2 Radiation from the Doorway 8

2.3 Captured External Fire Product Gases - An Equivalent

Buoyant Source 9

2.4 Radiation and Convection for the Seating Fire 9

2.5 Convective Heat Transfer From a Combined, Equivalent

Source of Buoyancy 11

2.6 Radiation Between the Lower Celling Surface and the FarField Cabin Surfaces 13

2.7 Heat Transfer from the Upper Ceiling Surface 13

2.8 The Boundary Value Problem for the Ceiling, and the

Method of Its Solution 14

3. CALCULATION OF THE RESPONSE OF THE CEILING IN THE POST-CRASH TEST

SIMULATION 15

3.1 Calculations for the Test 111 Ceiling Response 15

3.3 Comparisons Between^Somputed and Measured Temperatures .... 16

4. AN ANALYSIS OF THE THERMAL RESPONSE OF THE NEAR-CEILING

THERMOCOUPLES 17

4.1 A Model Equation for the Temperature of the

Thermocouples 17

4.2 Solutions for T^ in the Test 111 Scenario 20

4.3 Comparison Between Computed and Measured Temperatures - An Optimum Choice for the Value of 8 20

5. PREDICTING THE POST-CRASH TIME-TO-IGNITION OF CEILING

CONSTRUCTIONS IN A FULLY SEATED CABIN 21

5.1 Using the Algorithm to Estimate a Measure of Post-Crash

Cabin Life Safety 21

-iii-
TABLE OF CONTENTS (continued)

Page

5.2 Estimates of Post-Crash Fire Growth in Arrays of Cabin

Seats 22

5.3 Post-Crash Response of a Honeycomb Celling Material - Estimates of Tlme-to-Ignltlon 23

6. SUMMARY AND CONCLUSIONS 24

7. ACKNOWLEDGMENTS 25

8. REFERENCES 26

9. NOMENCLATURE 27

-Iv-
LIST OF FIGURES

Figure

Figure

Figure

Figure

Figure

Figure

Figure

Figure

Figure

Figure

Figure

Figure

Figure

Figure

Page

1. A schematic of the test setup 29

2. Measured doorway heat flux vs time (lower

fluxraeter upper fluxmeter ); Test 111

( background) 30

3. Measured doorway heat flux vs time (lower fluxmeter);

Tests 104-111 31

4. Measured temperature vs time at position 1; Tests 104-111 32

5. Measured temperature vs time at position 2;

Tests 104-111 33

6. Measured temperature vs time at position 3;

Tests 104-111 34

7. Measured temperature vs time at positions 1 ( ),

2 ( . . ), and 3 ( ); Test 111 35

8.

A simplified version of the post-crash fire scenario ... 36

9.

The viewfactor between the doorway and a ceiling

element 37

10. The idealized post-crash fire scenario 38

11. Computed Test 111 lower ceiling surface temperatures

(+: 6 = 0.; x: 6=3.0m‘^) and corresponding, measured

near-celling temperatures at positions 1, 2, and 3 39

12. Two extreme configurations for placement of the

near-ceiling thermocouples 40

13. Predicted and measured Test 111 thermocouple

temperatures (6 = 2.0 m^)

.

<>: predicted T^, configuration 1 0: predicted T^, configuration 2

, , .

: measured T^ 41

14. Predicted and measured Test 111 thermocouple

temperatures (6 = 3.0 m^).

<>: predicted T^, configuration 1 0: predicted T^, configuration 2

, , .

: measured T^ 4 2

-V-
LIST OF FIGURES (continued)

Figure 15* Predicted and measured Test 111 thermocouple

temperatures (5 “ 4.0 m^).

<>: predicted T^, configuration 1 0: predicted configuration 2

, , .

; measured • •

Figure 16. Estimate for Q for arrays of polyurethane seats

with and wlthoul^Blocklng layers [13] ( extrapolated

from curves of [13])

43

44

Figure 17. Predicted lower surface, post-crash, temperature of the

honeycomb ceiling material in a cabin outfitted with

polyurethane seat cushions; with ( . ) and without

( ) blocking layers 45

-vl-
THE THERMAL RESPONSE OF AIRCRAFT CABIN CEILING MATERIALS DURING

A POST-CRASH, EXTERNAL FUEL-SPILL, FIRE SCENARIO

Leonard Y. Cooper

Abstract

An algorithm is developed to predict the thermal response of aircraft

ceiling materials during a post-crash fire scenario. The scenario involves an

aircraft's emergency exit doorway which opens directly onto the flames of an

external, fuel-spill fire which engulf a large portion of the fuselage. Data

of near-ceiling temperatures acquired during a series of eight, full-scale,

wide-body aircraft cabin, post-crash test simulations provide indirect validation of the algorithm. These tests involved cabins outfitted with only

single, mockup seats. Two other full-scale cabin tests Involving fire spread

through twenty-one seat arrays with different types of seat construction

provide the input data required to exercise the algorithm in evaluations of

fully outfitted cabins.

Relative to the post-crash scenario, a measure of cabin fire safety is

proposed, viz., the post-crash time-to-ceiling- ignition. This measure would

be used as a surrogate for the post-crash time available for passengers to

safely evacuate the cabin. In this sense, the algorithm is exercised in an

example evaluation of the fire safety of a candidate honeycomb celling

material used together in cabin systems Involving polyurethane cushion

seating, with or without a blocking layer.

- 1 -
1 . INTRODUCTION

1.1 Purpose of the Investigations

The purpose of this investigation is to analyze aircraft cabin ceiling

surface temperature data recently acquired during full-scale test simulations

of post-crash fires. The analysis is carried out with a view toward the

development of a procedure for estimating the temperature histories of overhead aircraft cabin materials subsequent to the ignition of exterior, fuelspill fires. With such a capability it would be possible to estimate the time

for such materials to reach ignition temperatures. This would result in a

rational means of ranking the fire safety of candidate overhead aircraft cabin

materials

.

All tests described here were carried out by staff of the United States

(U.S.) Department of Transportation Federal Aviation Administration (FAA) at

the National Aviation Facilities Experimental Center, Atlantic City, New

Jersey.

1.2 Description of the Tests

The experiments involved a wide-body aircraft cabin fire test series

similar to those post-crash fire test simulations reported previously in

reference [1]*. The test setup simulated a wide-body aircraft with two open

doorways where the fuselage at one of the doorways is engulfed by fire of a

*Numbers in brackets refer to the list of references at the end of this

paper.

-2 -
large, burning, external fuel spill. The fuel-spill fire is simulated by a

fire in a 2.44 m x 3.05 m pan of jet fuel (JP-4) . The threat to the cabin

which would be generated by this external test fire configuration has been

shown in [1] to be representative of the threat which would be generated by

real, large, external fuel-spill fires. No-wind conditions were simulated

during the tests. The actual test article was a surplus U.S. Air Force C133A

cargo aircraft with wings and tail surfaces removed.

The ceiling of the test cabin was made up of panels of 0.0127 m thick,

rigid Kaowool® ceramic fiber board. In the analysis to follow, the thermal

properties of this material are assumed to be constant. Consistent with

information from the manufacturer, they are taken to be

thermal conductivity = k = 0.045 W/m®K

-7 2

thermal diffusivity = a = 2.67 (10 ) m /s

A mockup seat made up of cushions placed on a steel frame was placed in

the cabin directly in front of the open doorway which was exposed to the fuel

spill fire. The placement was likely to lead to the most rapid ignition of

the cushion material. In the present study, data from only eight tests,

designated as test numbers 104-111, are considered. The only parameter which

was varied from test to test was the seat cushion/fabric construction. One

test, test 111, is designated as the background test since it involved the

seat frame with no cushioning.

A schematic of the test set up is presented in figure 1.

- 3 -
During the tests the radiant heat flux near the doorway was measured with

fluxmeters facing outward toward the pan fire. Measurements were taken at two

elevations, 0.30 m and 0.91 m above the floor. At all times during each of

the tests the flux measured by the two gages were substantially similar. An

example of this can be seen in figure 2 where plots of the two fluxmeter

measurements, acquired during the background test, are presented. In the

analysis to follow it will be assumed that the radiant flux from the pan fire

is uniform and isotropic across the entire doorway. This flux will be taken

to be the flux measured by the lower of the two, near-doorway flxixmeters. The

measured flux history at this elevation for all eight tests are plotted

together in figure 3. This figure provides support for the assumption that

the exposure fire was closely reproduced in each of the tests of the series.

This study will consider near-ceiling temperature measurements which were

acquired from three thermocouples placed in the line which traverses the width

of the cabin and which is directly above the center of the doorway. As noted

in figure 1, one of these thermocouples, at designated position 1, was

directly above the center of the seat. The other two were 0.91 m (position 2)

and 1.83 m (position 3) from the first. The thermocouples were constructed

from 24 gage (0.000584 m diameter) chrome 1/alumel wire. The wire from each of

the three thermocouples was supported several centimeters from its bead, and

there was an attempt to position the bead close to the ceiling surface with

the hope that the bead temperatures would be substantially similar to the

respective, nearby, ceiling surface temperatures. Although there was no

attempt to measure them, bead-to-ceiling distances were probably of the order

of a millimeter. It is noteworthy that the placement of near-celling thermocouples in the present test series is different than in previous test series

-4-
where thermocouple bead- to-ceiling distances \^re of the order of a few

centimeters

.

Except for the background test, data for all tests ^re available only

until 120 s after ignition of the pan fire. The Test 111, background test

data were available for 240 s.

Up to 120 s after ignition, the measured temperatures for all eight tests

are presented in figures 4-6. Plots of measured temperature vs time at position 1, above the seat, are presented in figure 4. Plots of measured temperature histories at positions 2 and 3 are presented in figures 5 and 6,

respectively. As can be seen from these three figures, at each thermocouple

location the measured temperature histories are substantially similar for each

test. Based on this observation, it is reasonable to assume that, for the

threat scenario being simulated and up to the 120 s time mark, fire development in a single, raockup seat would not add significantly to the ceiling

surface material fire threat. For this reason, in the analysis of the present

test series it is assumed to be adequate to study the thermal response of the

celling only during the background test. Plots of the temperature histories

measured by the three, near-celling thermocouples during the background test

are presented in figure 7.

- 5-
2. AN ANALYSIS OF THE THERMAL RESPONSE OF THE CABIN CEILING MATERIAL

2.1 Components of the Surface Heat Flux

During post-crash fires of the type simulated in the test series there

are two major phenomena which can lead to relatively prompt lower surface

heating of the cabin ceiling.

The first of these phenomena Involves the thick flames and copious

products of combustion which engulf the exterior of the fuselage in the

vicinity of the threatened, open, cabin doorway. These lead to significant

components of both radiant and convective heat flux to the cabin ceiling.

The convective component comes about from the hot, buoyant gases of the

fuel spill fire which are captured by the open doorway. Upon entering the

cabin, these gases are driven upward toward the ceiling, forming an outward

(i.e., away from the doorway and toward the cabin interior) moving ceiling

jet. After spreading radially from the doorway, this ceiling jet is

redirected away from the general location of the doorway and toward the front

and rear of the cabin. Eventually the hot, captured, products of combustion

start to fill the cabin. They then participate in venting from the second

open doorway and in complicated entrainment processes which develop at the

fire-threatened/exposed open doorway itself. (It is noteworthy that a global

analysis of the external, fuel spill fire and the captured cabin flow under

rather general wind conditions have been presented previously in [2].)

-6 -
The second phenomenon which leads to heat flux to the lower ceiling

surface involves the growing fire which spreads in the cabin seating. As

mentioned earlier, the limited, single-seat scenario of the present tests

appears to result in only marginally important levels of ceiling heat flux.

Yet, as will be discussed later, fire spread in a fully outfitted cabin could

indeed lead to a significant additional threat to the cabin ceiling. The

seating fire leads to both radiation and convective heating of the ceiling.

The radiation would be primarily from the fire's combustion zone, and the

convection would result from the fire's plume-driven ceiling jet. This latter

ceiling jet would augment the previously mentioned, captured-gas-driven

ceiling jet.

Other components of heat flux to/ from both the upper and lower ceiling

surfaces are radiation from relatively cool, far-field surfaces and reradiation from the ceiling surfaces themselves. Finally, in an analysis of the

ceiling heating it is reasonable to account for natural convection cooling of

the upper ceiling surface.

For the purpose of analysis the geometry of figure 8, which is somewhat

simpler than the actual cabin geometry, will be adopted.

Quantitative estimates for the above components of celling heat transfer

are developed below. Using these estimates, the boundary value problem for

the transient temperature distribution of the ceiling material is then

formulated and solved.

- 7 -
2.2 Radiation from the Doorway

The entire doorway opening is assumed to be a source of uniform diffuse

radiation. The radiant flux emitted from this doorway is assumed to have an

amplitude, q" , , (t), identical to that measured by the lower of the doorway flux calorimeters mentioned earlier.

Any line-of-sight interference by the seats of doorway-to-ceiling

radiation will be neglected.

Referring to figure 8, and using a well-known vlewfactor result given,

e.g., in [3], the vlewfactor, ^A-dA* between the area of the doorway opening,

A, and a horizontal element of lower celling surface, dA, lying on a line

normal to the plane of the doorway is given by

and where the dimensions a^ a

2

, b and c are defined in figure 9. Thus, the

thermal radiant flux from the outside, fuel-spill fire, through the door, and

to the celling element is taken to be

F ( 2 )

where

(3)

( 4 )

-8-
2.3 Captured External Fire Product Gases - An Equivalent Buoyant Source

The products of combustion of the external, simulated, fuel-spill fire

which are captured at the open cabin doorway rise in a plume to the ceiling of

the cabin, entraining far-field cabin air during their ascent. In the present

analysis the "captured gas" plume will be modeled by a nonradiating, equivalent, point source of buoyancy located at the center of the horizontal surface

of the mockup seat, and identified in figure 8. Note that all radiation

transfer from the exterior fuel spill fire to the cabin ceiling is assumed to

be taken account of in eq. (4). The strength of the equivalent source,

Q , , will be assumed to be directly proportional to q" . ^ . Thus ^equiv’ j r r ^rad-door

Q . = 3q" j , (3 in m ) equlv rad-door (5)

and 3 is an, as yet undetermined, constant.

As the captured-gas plume impinges on the ceiling near (thermocouple)

position 1, it forms an outward moving celling jet. As will be described

below, estimates of convective heat transfer from this jet to the ceiling

surface will make use of the results of [4].

2.4 Radiation and Convection for the Seating Fire

During the first 120 s of the post-crash fire, cabin ceiling heat

transfer contributions from the burning single mockup cabin seat did not

appear to be significant. However, in fully outfitted cabins, it is anticipated that this situation would be changed especially after the first minute

-9-
or two subsequent to ignition. This is the case since, by these times, fires

in FAA, multiple-seat, cabin test configurations have been observed to grow

and spread beyond single seat involvement. Since the present analysis will be

extended to fully outfitted cabin scenarios, ceiling heat transfer with an

important contribution from the seating fire will be included here at the

outset

.

The seating fire is simulated by a time-dependent point source of energy

release rate, Q , which is assumed to be located, together with the nonS6 3.C radiating source, Q . , at the center of the horizontal surface of the equiv

outer, external-fire-exposed, doorway seat. In contrast to ^ nonzero

fraction, X , of Q ^ is assumed to be radiated uniformly over a sphere

surrounding the combustion zone, and to the far field. The remaining rate-ofenergy release, fl - X ] Q , drives the buoyant fire plume upward.

Thus, the radiation from the seating fire which is incident on the ceiling is

assumed to be

rad-seat

Xr ,seat (1 2/u2 r /H >3/2i

( 6 )

where, as defined in figure 8, H is the source-to-ceiling distance and r is

the radial distance from the source as measured along the ceiling.

• • Q ^ would vary from one seat cushion construction to another. Q ^seat ^ seat

would typically have to be estimated from test data, and then specified in the

present analysis. X would also vary somewhat from one construction to

IT I S63

L

another, although it is reasonable to choose the value X = 0.35, a value

IT I S 6 ct L which characterizes the radiation from flaming combustion zones of many

- 10 -
practical fuel assemblies [5]. In all calculations to be presented here, the

0.35 value will be adopted.

2.5 Convective Heat Transfer From a Combined, Equivalent Source of Buoyancy

A single point source of buoyancy, Q, is used to describe the total

generation of upward moving combustion gases which are driven by a combination

of the equivalent, captured-gas-buoyancy source, seating-firebuoyancy ^ source, fl-X ^ ^)0 Thus

r,seat^ seat

Q=Q .+[l-X )q equiv ^ r,seat'^ seat

(7)

All convective heat transfer to the cabin ceiling will be modeled by

a Q-generated, plume-driven, ceiling jet. This heat transfer will be estimated with the use of the algorithm, reproduced below, which was developed in

[4], and used in [6]. Thus, the convective rate of heat transfer to the

ceiling surface will be estimated from

• M *^conv ,L ad - ( 8 )

In the above, T_ is the instantaneous, absolute temperature of the lower

S y Li ceiling surface, T^^ is the absolute temperature that would be attained by the

lower surface of an adiabatic ceiling.

^amb^

T

amb

10.22 exp (-1.77 r/H) , 0 ^ r/H < 0.75

2.10 (r/H)“°’®^, 0.75 ^ r/H

\

Q* is a dimensionless buoyant source strength.

( 9 )

- 11 -
g ( 10 ) 1/2 ^

3/2

)

g is the acceleration of gravity, and density,

absolute temperature and specific heat, respectively, of the ambient air.

Also, h is the heat transfer coefficient

7.75 Re"^*^ [l - (5.0 - 0.390 Re® ) (r/H) ]

,

0 < r/H < 0.2

0.213 Re"^*^ (r/H)"^*^^, 0.2 < r/H < 1.03 (11)

0.217 Re"^*^ (r/H)”^*^, 1.03 < r/H

where

r „l/2 -*1/3

'’amb %

* "

Re = gl/2 ^3/2 ^*1/3^^

( 12 )

and V is the kinematic viscosity of the ambient air.

The above algorithm will be used in the present calculations. In doing

so, two major assumptions are made; namely 1) effects of the inevitable,

growing, upper-cabin smoke layer are relatively weak during the early times of

interest, and 2) the Interactions of the ceiling jet and lateral cabin wall

surfaces, especially surfaces immediately to the doorway side of the plumeceiling Impingement point will not lead to total, instantaneous, heat transfer

flux amplitudes which are significantly larger than peak, instantaneous values

that will be estimated with their neglect.

- 12-
2.6 Radiation Between the Lower Ceiling Surface and the Far-Field Cabin

Surfaces

At all times subsequent to ignition, the lower ceiling surface at temperature T„ T is assumed to diffusely radiate to the initially ambient temperature , illuminated surfaces of the cabin and its furnishings. In response to

this radiation, the temperatures of those lower surfaces outside of the

seating combustion zone also increase with time. However, for times of

interest here, it is assumed that these latter temperature increases above

Tamb always relatively small compared to the characteristic increases of

T_ j. Accordingly, at a given radial position of the lower ceiling surface,

S ^Lj the net radiation exchange between the ceiling and the nonburning surfaces

below can be approximated by a net reradiation flux

^rerad ,L (13)

where a is the Stephan-Boltzmann constant, and e is the emlttance/absorptance

Lj of the assumed grey, lower ceiling surface.

2.7 Heat Transfer from the Upper Ceiling Surface

Heat is transferred through the ceiling, and eventually the temperature

of its upper surface, which is also assumed to be exposed to a constant T^^^

environment, begins to rise.

The rate of heat transfer from the upper ceiling surface has convective

and radiative components, q" ,, and q" , respectively. These can be conv,U rerad,

U

estimated from

- 13-
q’* „ = (t XT “ T , ] conVjU TJ ^ s,U arab^

4 4

^rerad,U ^"^s,U "^amb^

where ^s,U the instantaneous upper surface temperature, hy is an effective

heat transfer coefficient, and is the emittance/absorptance of the assumed

grey, upper ceiling surface. The value for h^ to be used in the present

calculations will be [7]

h^ = 1.675 It „ - T , W/m^ (T in °K) U I s , U amb

(15)

2.8 The Boundary Value Problem for the Ceiling, and the Method of Its Solution

The absolute temperature field of the celling material is assumed to be

governed by the Fourier heat conduction equation. Initially, the ceiling is

taken to be of uniform temperature, T^^^. The net rates of heat transfer to

the lower and upper surfaces, q" and q" , respectively, are given by

Li U

Q = n** + n** fl” — n** Moor-celling ^rad-seat ^conv,L '^rerad,!

n" = — n" — n" ^conv,U ^rerad,U

(16)

For times of interest here, radial gradients of variables of the problem

are assumed to be small enough so that conduction in the ceiling is quasi-one

dimensional in space, l.e., T = T (Z, t; r)

,

where Z is the indepth celling

coordinate

.

An illustration of the final, idealized, post-crash fire scenario to be

analyzed here is presented in figure 10.

- 14-
A computer program for solving the above problem was developed. The

solution to the heat conduction equation for the ceiling at every radial

position of interest is by finite differences. The algorithm for this was

taken from [8] and [9]. For a given calculation, N ^ 20 equally spaced points

are positioned at the surfaces and through the thickness of the ceiling. The

spacing of these, 6Z, is selected to be large enough (based on a maximum time

step) to insure stability of the calculation. Throughout a calculation, the

change in time for all time steps is made small enough so that, at a given

lower surface node, the temperature increase from time step to time step never

exceeds one percent of the current value of T at that node.

3. CALCULATION OF THE RESPONSE OF THE CEILING IN THE POST-CRASH TEST SIMULATON

3.1 Calculations for the Test 111 Ceiling Response

The algorithm described in the last section was used to predict the '

ceiling response during the first 240 s of the Test 111, post-crash simulation .

Eq. (1) values of k and a were used for the 0.0127 m thick Kaowool®

ceiling. The ceiling surfaces were assumed to absorb and radiate as black

bodies [e = e = l). Consistent with previous discussion, for this test

U ij

Q of eq. (6) was taken to be identically zero, and q" , , (t) of eq. (5)

was taken to be identical to the Test 111, underseat fluxmeter measurements.

Ceiling temperatures at positions 1, 2 and 3 were computed for different 6's

in the range 0 —< 6 —< 6.0 m . (This range of 6 leads to the approximate* <)••qulv

range 0 < Q . < 300 kW.) — equiv —

- 15-
o The computed lower celling histories for 3 " 0. and 3.0 m are plotted In

figure 11 together with the corresponding, measured, near-celling temperatures

of figure 7.

3.2 The Importance of

If convective celling heating from doorway-captured products of combustion Is equivalent to a near-seat buoyancy source of the order of a few

hundred kW, then the calculated results plotted In figure 11 Indicate that

such heating Is not significant compared to doorway radiation. (Note that

3=0. and 3.0 m of figure 11 leads to a

Qgq^j^^

of zero and approximately

150 kW, respectively. Also, except for the very earliest few seconds subsequent to ignition, convection from the relatively weak source associated

2 with 3 = 3.0 m is seen to lead to net cooling of the strongly irradiated

ceiling surface.) This result Is consistent with earlier observations where

variations in single seat cushion construction (peak energy release rates

likely never exceeding the few hundred kW level) did not lead to significant

differences in near-celling temperature measurements (see figures 4-6).

3.3 Comparisons Between Computed and Measured Temperatures

As can be seen in figure 11 the peak computed values of celling temperature compare favorably with the corresponding peak temperatures measured by

the near-ceiling thermocouples. However, the basic qualitative characteristics of the computed and measured transient thermal responses are significantly different. In particular, the measured temperatures of the thermocouples

do not have the same type of rapid response which the numerical solution

- 16-
properly predicts for the lower ceiling surface temperatures. Also, the close

tracking of the position 2 and 3 thermocouples at early times does not compare

favorably with a like tracking of the computed, lower ceiling temperatures at

these two positions.

Two conclusions result from these observations; namely, the thermocouples

are not measuring the temperature of the near-by ceiling surface, and, therefore, data to support the validity of the analysis are not evident. As a

result of these conclusions an analysis of the response of the thermocouples

was carried out. This analysis, to be reported in the next section, had a

two-fold purpose; first, to explain the measured thermocouple responses, and

second, to bring a measure of experimental validation, albeit indirect, for

the predicted ceiling response.

4. AN ANALYSIS OF THE THEEIMAL RESPONSE OF THE NEAR-CEILING THERMOCOUPLES

4.1 A Model Equation for the Temperature of the Thermocouples

The objective of the present analysis is to predict the thermal response

of the thermocouples when appropriately placed within the figure 10 scenario,

near, but not touching the ceiling. The procedure for positioning these

devices prior to testing was such that the thermocouple wires were essentially

parallel to the lower ceiling surface and at a distance, d, of the order of

0.001 m. The actual orientation of the 24 gage (diameter = 0.000584 m) wire

relative to the doorway plane is unknown. As depicted in figure 12, the

analysis will consider two extreme configurations for the wire, viz., normal

and parallel to the doorway.

- 17 -
The characteristic time for' conductive heat transfer through the wire

thickness Is relatively small, of the order of tenths of a second. The

analysis will therefore assume that throughout the experiment the wire Is

spatially uniform in temperature.

«

The density and specific heat of the chromel/alumel wire (a 95 percent

Nickel alloy) will be taken to be identical to that of Nickel, viz.,

3 density = p = 8800 kg/m

specific heat = = 460 Ws/(kg®K)

From references [4] and [10] or from reference [11] it is possible to

conclude that the thickness of the ceiling jet within which the thermocouples

are submerged are of the order of several centimeters. With a characteristic

thermocouple-to-ceillng separation distance, d, of the order of 0.001 m, it is

therefore reasonable to assume that gas velocities local to the thermocouple

wire are so small that forced convection heating compared to radiative heating

of the wire is negligible. On account of the fineness of the wire, the

characteristic Grashoff numbers would be relatively small, and it is also

reasonable to neglect natural convection heating of the thermocouples.

The thermal analysis which emerges from the above discussion leads to the

following equation for the temperature, T^, of the thermocouple wire

dT

J pc D

4 p dt

= 6' +6' +6' - O' ^door-wire ^ceiling-wire ^amb-wire ^wlre (17)

where
O' = reradiation from the wire, per unit length of wire ( 18 ) wire , = ttDoT^w

O'amb-wire

radiation from lower ambient temperature surfaces to

the wire , per unit length of wire

(19)

2 = ^ a T

amb

6

'ceiling-wire

radiation from the ceiling to the wire,

per unit length of wire

Tm

2 a Ts ,L

( 20 )

•, _ radiation from the doorway to the wire,

^door-wire per unit length of wire

= oO q”

door- ceiling

( 21 )

In eq.^ (21), q" is given in eq. ^door-ceiling (4), and, depending on the wire

configuration

,

a = 1 for wire normal to the door plane, i.e., configuration 2

of figure 12.

( 22 )

a. “ l/sin0 (see figure 9 for definition of 9) for wire parallel

to the door plane, i.e., configuration 1 of figure 12.

In the above, all surfaces are assumed to radiate and absorb as black

surfaces

.

To obtain a solution for the of a thermocouple at a given position

along the cabin ceiling one would specify a and T .(t) at that position, use

S I u the measured values of q" , , to obtain q” , and solve eq. (17) ^rad-door ^door-celling ^

subject to the initial condition T (t = 0) = T , . w amb

- 19 -
4.2 Solutions for T„ in the Test 111 Scenario w

The above procedure for predicting T^(t) was applied to the Test 111

scenario. The analysis was carried out numerically for a thermocouple in

position 1, 2 or 3 and in configuration 1 or 2. In each case, the history for

Ts required in the analysis, was taken from the ceiling

temperature calculations described earlier.

The results of the T^ calculations are presented in figures 13, 14, and

2 2 2

15 for 3 values of 2.0 m , 3.0 m , and 4.0 m , respectively. Also included in

these figures are plots of the measured T^ values previously presented in

figures 7 and 8.

4.3 Comparison Between Computed and Measured Temperatures - An Optimum

Choice for the Value of 3

As can be seen in figures 13-15, the computed and measured thermocouple

temperatures compare as favorably as one could reasonably hope for in an

analysis of the kind of experimental fire scenario under investigation.

Perhaps of greatest significance is the early-time thermocouple temperature predictions, which were of particular concern in the previous ceiling

temperature - thermocouple temperature comparisons of figure 11. Here, the

simulations of the early, near- linear responses of the thermocouples are

noteworthy.

Of further significance is the fact that the results of the calculations

reveal a possible explanation for the previously noted, close tracking of the

-20-
response of the thermocouples at positions 2 and 3. As can be observed in

each of the three figures, the present analysis predicts such behavior if the

thermocouple wire at position 2 was normal to the door plane (configuration

2), and the thermocouple wire at position 3 was parallel to the door plane

(configuration 1).

Figures 13-15 provide a basis for selecting the "best" value for 3 with

which to carry out the ceiling temperature calculation in the context of the

post-crash cabin fire scenario under investigation. In this regard, the 3

predicting a ceiling response which, in turn, yields the most favorable

comparisons between calculated and measured values of would be the obvious

choice. As is evident from the figures, the predictions are not very

2 sensitive to 3 variations in the appropriate range 2. 0-4.0 m . Furthermore,

2 2 2 of the values 3 = 2.0 m , 3.0 m , and 4.0 m , all yield reasonable predictions, and no one of these values clearly yields more favorable predictions

2 than the others. 3 = 3.0 m will be chosen as the "best" value.

5. PREDICTING THE POST-CRASH TIME-TO-IGNITION OF CEILING CONSTRUCTIONS

IN A FULLY SEATED CABIN

5.1 Using the Algorithm to Estimate a Measure of Post-Crash Cabin Life Safety

The results of the previous section provide some confidence in the postcrash ceiling thermal response algorithm. To use the algorithm in a manner

that would simulate the post-crash fire exposure of a fully seated cabin, It

is necessary to include the effects of fire spread in an array of seating

adjacent to the exposed, open doorway. As was discussed earlier, this would

be done by inputing appropriate, nonzero, X and Q terms in eqns . (5) r,seat seat

-21 -
and (7), Then, using the k and a of a candidate ceiling material, tthe algorithm would calculate the ceiling's time-dependent, post-crash, thermal

response

.

In the most likely case of a combustible celling material, one could, for

example, predict the time for the lower surface to reach a characteristic

ignition temperature . Since cabin flashover would not likely occur prior to

ceiling ignition, such a tlme-to-ignitlon would provide a reasonable measure

of post-crash cabin fire safety, viz., the minimum time available for

f passengers to evacuate the cabin or the Available Safe Egress Time (ASET)

[12]. Hopefully, evaluations of practical cabin ceiling material candidates

would lead to associated Ignition times, or minimum ASET's, which exceed the

time required for cabin evacuation. In any event, the greater the time-toignition of a material the better.

In the case of a noncombustible ceiling, time-to-ignltion in the above

discussion would be replaced by time to reach some agreed upon celling temperature (e.g., 600°C), which would typically be associated with cabin flashover.

5.2 Estimates of Post-Crash Fire Growth in Arrays of Cabin Seats

The objective of reference [13] was to obtain estimates of the energy

release rate of post-crash fires growing through arrays of cabin seats. Based

on FAA, full-scale, 21 seat tests which were similar to Tests 104-111, estimates of fire growth in two types of seat construction were obtained. The

first type of seats, designated as "regular" seats v«re made of fire retarded

polyurethane foam covered with wool-nylon fabric. The second seat construc-

-22-
tion was similar to the first, except that it included a blocking layer

constructed of a 0.0048 m thick sheet of neoprene with a polyester scrim.

The estimates of Q (t) for the two types of seats are plotted in

seat

figure 16. The plots terminate at times 140 s and 185 s, at which times

video-tape recordings of the tests indicated the initiation of either flashover (the 140 s time) or of rapid development of total obscuration (the 185 s

time)

.

The Q estimates of figure 16 will be used below to evaluate the postS0 3 tl crash response of a specific, honeycomb celling material.

5.3 Post-Crash Response of a Honeycomb Ceiling Material - Estimates of

Time- to-Igni tion

The algorithm developed here was used to estimate the post-crash thermal

response of a 0.0254 m thick, honeycomb composite, aircraft lining material

with an epoxy fiberite covering. The effective thermal properties of the

composite were measured, and found to be [14]

k = 5.9 dO"^) kW/(m°K)

a = 4.8 do"^) m^/s (23)

P = 110. kg/m^

Cp = 1.11 U/(kg°K)

X of eq. (7) was taken to be 0.35 and Q was simulated by the r,seat seat

plots of figure 16. The computer program written to exercise the algorithm

was actually designed to accept pairs of [t, Q (t)j data points as input, seat

-23-
and then to linearly interpolate between these to obtain Q at any

seat

arbitrary time during the calculation, 3 of eq. (5) was taken to be 3.0 m ,

e and e were taken to be 1

.

L U

The predicted history of the lower surface of the honeycomb ceiling

directly above the doorway seat is plotted in figure 17 for both "regular”

seating and "blocked" seating.

The ignition temperature, Tj[gn> honeycomb material had been

measured previously, and was found to be [15]

T, = 536°C (24) ign

As can be seen in figure 17 the calculation predicts onset of celling ignition

at 148 and 204 s for "regular" and "blocked" seating, respectively. For

cabins outfitted with ceilings of the honeycomb material, the use of the

blocked rather than the unblocked seating construction would lead to a 56 s

advantage in time available for post-crash cabin evacuation,

6. SUMMARY AND CONCLUSIONS

An algorithm was developed to predict the thermal response of aircraft

ceiling materials during a specific, post-crash, external fuel-spill, fire

scenario. Experimental measurements of near-ceiling temperatures in a series

of eight, full-scale, post-crash, single seat, test simulations provided

indirect validation of the algorithm. Two other full-scale tests, each of

which Involved fire spread through arrays of seating with different types of

-24-
seat construction (fire retarded polyurethane foam covered with a wool-nylon

fabric, with and without a neoprene sheet-polyester scrim blocking layer)

provided the input data required to exercise the algorithm.

The post-crash, time-to-ignition of a ceiling construction, which can be

associated with the time available for passengers to safely evacuate an aircraft, was recommended as one possible measure of the fire safety of the cabin

ceiling-seat construction system (i.e., changing either the ceiling or the

seat-construction has an impact on cabin fire safety) .

Relative to the post-crash fire scenario considered here, the algorithm

was used to predict the response of a candidate honeycomb ceiling material

when used in a wide-body aircraft, with and without seat blocking. Times- toceiling ignition were estimated to be 148 and 204 s with and without seat

blocking, respectively.

In a similar way, the algorithm developed here could be used to estimate

the time-to-ignition of any other candidate ceiling material which would be

used in cabins having either blocked or unblocked seating. Required inputs

are the thickness and the effective thermal conductivity and dlffusivlty of

the material.

7 . ACKNOWLEDGMENTS

This work was supported by the Federal Aviation Administration of the

U.S. Department of Transportation.

-25 -
8. REFERENCES

[1] Hill, R.G., Johnson, G.R. and Sarkos, C.P., Postcrash Fuel Fire Hazard

Measurements in a Wlde-Body Aircraft Cabin, FAA-NA-79-42, Federal

Aviation Admin., NAFEC, Atlantic City, NJ, 1979.

[2] Emmons, H.W. , The Ingestion of Flames and Fire Gases Into a Hole in an Aircraft Cabin for Arbitrary Tilt Angles and Wind Speed, Home Fire ProJ

.

Rpt. 52, Harvard Unlv. Div. Appl. Sciences, Cambridge, MA, 1982.

[3] Eckert, E.R.G. and Drake, R.M., Heat and Mass Transfer, McGraw-Hill, 2nd

Ed., 1959.

[4] Cooper, L.Y., Heat Transfer from a Buoyant Plume to an Unconflned

Ceiling, J. Heat Transfer, 104, p. 446, 1982.

[5] Cooper, L.Y., A Mathematical Model for Estimating Available Safe Egress

Time in Fires, Fire and Materials, 6, p. 135, 1982.

[6] Cooper, L.Y., Thermal Response of Unconfined Ceilings Above Growing

Fires and the Importance of Convective Heat Transfer, 22nd National Heat

Transfer Conf., ASME Paper 84-HT-105, 1984 and NBSIR 84-2856, National

Bureau of Standards, Washington, DC, 1984.

[7] Yousef, W.W. , Tarasuk, J.D. and McKeen, W.J., Free Convection Heat

Transfer from Upward-Facing, Isothermal, Horizontal Surfaces, J. Heat

Transfer, 104, p. 493, 1982.

[8] Emmons, H.W. , The Prediction of Fires in Buildings, 17th Symp. (Inter.)

on Combustion, p. 1101, 1979.

[9] Mitler, H.E. and Emmons, H.W. , Documentation for the Fifth Harvard

Computer Fire Code, Home Fire Project Tech. Rpt. 45, Harvard Univ.

,

Cambridge, MA, 1981.

[10] Poreh, M., Tsuei, Y.G. and Cermak, J.E., Investigation of a Turbulent

Radial Wall Jet, ASME Journal of Applied Mechanics, p. 457, 1967.

[11] Alpert, R.L., Turbulent Ceiling-Jet Induced by Large-Scale Fires,

Combustion Science and Technology, Vol. 11, p. 197, 1975.

[12] Cooper, L.Y., A Concept of Estimating Safe Available Egress Time, Fire

Safety Journal, Vol. 5, p. 135, 1983.

[13] Steckler, K. , National Bureau of Standards, Chapter 1 of the Role of

Aircraft Panel Materials in Cabin Fires and Their Properties, Quintiere

et al, to appear as NBSIR, National Bureau of Standards, Washington, DC.

[14] Parker, W., National Bureau of Standards, private communication.

[15] Harkleroad, M., Quintiere, J. and Walton, W. , Radiative Ignition and

Opposed Flame Spread Measurements on Materials, DOT/FAA/CT-83/28

(National Bureau of Standards Rpt. to) Federal Aviation Admin., Atlantic

City, NJ, 1983.

-26-
9 . NOMENCLATURE

^1» ^2* ^ dimensions. Fig. 9

specific heat

D wire diameter

^A-dA view factor, Eq. (2)

g acceleration of gravity

H seat fire-to-ceiling distance

lower/upper surface heat transfer coefficient

h dimensionless heat transfer coefficient, Eq. (12)

k thermal conductivity

N number of grid points in ceiling analysis

•Q enthalpy flux in plume, Eq. (7)

Q*

• dimensionless Q, Eq. (10)

Q . equiv equivalent fire strength

•Qseat

•O'amb-wire

strength of seat fire

radiant heat transfer to wire per unit length

radiation from ambient to wire

Q'ceiling-wire radiation from ceiling to wire

^door-wire

^ire

radiation from doorway to wire

radiation from wire

^i

rates of heat transfer per unit area

• ti q q ^conVjU, conv,L

convection to upper/ lower ceiling

• II ^door- ceiling radiation from doorway to ceiling

• II *^rad-door

radiation from doorway

• ti ^rad-seat radiation from seat fire to celling

a" a"

^reradjU* ^rerad,L

radiation from upper/ lower celling

- 27 -
A" A" net heat transfer to upper/lower celling

Re Reynold’s number, Eq. (12)

r distance from plume Impingement point

Tad adiabatic celling temperature, Eq. (9)

Tamb ambient temperature

Tign Ignition temperature

Ts,U» Tg^L upper/ lower surface celling temperature

Tw thermocouple wire temperature

t time

a thermal dlffuslvlty/wlre configuration constant, Eq. (22)

8 a constant

r r 1’ 2

constants, Eq. (3)

6Z Indepth spacing of celling grid points

"u’ \

lower/upper celling emlsslvlty

e configuration angle. Fig. 9

Xr jseat

• fraction of Q ^ radiated

seat

V kinematic viscosity

P density

pamb density of ambient

a Stef an-Boltzraann constant

-28-
29

Figure

1.

A

schematic

of

the

test

setup
(gUjo/M) xnu 1V3I-I

Figure

2.

Measured

doorway

heat

flux

vs

time

(lower

fluxmeter

upper

fluxmeter

):

Test

111

(background)
(aUJO/M) XnU 1V3H

31

TIME

(s)

Figure

3.

Measured

doorway

heat

flux

vs

time

(lower

fluxmeter)*

Tests

104-111
280.0

(>+.) 1N3I9IAIV 3A09V 3SId 3aniVa3dlAI3i

32

Figure

4.

Measured

temperature

vs

time

at

position

Tests

104-111
200.0 OJ CO

o

o

OJ

o• o

o

o

00

o• o

o

o

OJ

o

e^i

(Mo) iN3iaHv 3Aoav 3siy 3yniva3di^3i

33

TIME

(s)

Figure

5.

Measured

temperature

vs

time

at

position

Tests

104-111
160.0

:>i) lN3iaiAIV 3A09V 3Sia 3aniva3dl/^3i

fO

34

Figure

6*

Measured

temperature

vs

time

at

position

Tests

104-111
(>io) iN3ignv 3Aoav 3sia 3aniva3dN3i

35

TIME

(s)

Figure

7.

Measured

temperature

vs

time

at

positions
1.83

(0

k-(DOEX3

u.

36

Figure

8.A

simplified

version

of

the

post-crash

fire

scenario
Ch

I C

CO

I<N

37

Figure

9.

The

vlewf

actor

between

the

doorway

and

a

ceiling

element
rerad,

U

^conv,U

38

Figure

10.

The

Idealized

post—

crash

fire

scenario
300.0 o o o o o

lO O LO O LO

OJ OJ

o

o

OJ

o• o

o

OJ

o

d

CO

o

o

Cvj

o

o

00

o

o

o

K

(Mo) iNaai^v 3Aoav 3Sid 3aniva3d^N3i

39

Figure

11.

Computed

Test

111

lower

ceiling

surface

temperatures

(+: B

»

0.;

x: B

*

3.0

m"^)

and

corresponding,

measured

near-ceiling

temperatures

at

positions

1,

2,

and

3
Thermocouple configuration

Q<

LUX

(T

LU>o 5> $m>O

cr

u.

CVJ

fo \

<N

a o 2 ^ o 3 E

0) c £ o

I- o

0) J=

4J •wocI

0)y

(0 o

MH

CO 00 C V O fH

•H O.

•M fl <d o h u ® 2 o u u

bO

0) ti

.uX

0) I

CMg3

bO

>»

cC

V— o

40
o o o o o o O ID O LO O UO

OJ OJ tH tH

o

o

OJ

o• o

o

OJ

o

o

CO

o

o

OJ

o

o

00

o

o

o

u

o.

Z3o

o

o

a; — rg C C ft O O H

LiJ

U 4J

CO (0 0)

Ui W

« • 3 3 3

0) ^ bO 00 CO Hcm .. «, B u-i 0) TO CCS 0) O O O

P • O CJ .. 3 CN

03 «

(0 U ft ft •

OJ H H

a cn

TO TO

0) <u •

*J u

V o

>-i -H ^

3 TO T3

TO3 m

TO

(U

l-l 0^

0) 0) ^ u a o.

4J«

>-< a;a

^ •• ••

<U /\ o 4J V/

0)Ui3

bO

1-t

Cl,

;>^o) iN3iai^V 3A09V 3Sia 3aniva3dt^3i

41
300.0 o o o o o

lO O ID O LO

OJ OJ tH

o

o

OJ

o• o

o

OJ

o

o

CO

o• o

cvj

o

o

00

o

o

o

(Mo) iN3Ia^NV 3A09V 3Siy 3aniVa3d^^3i

a

3

O

o

o

0) J=

4J

4J

to •

fll Rcm

8

•o

0) o

M • 3 CO

as(0 I

X-/ C OB

(0 0)h

^ CN S £ O O H

•H -H

*J JJ •«

(0 <0 (U P P h

a a s

bo bO 09 •H *H rt «w V

c c e o o

o o

ap

«op

019 T3 O.

0) 8

T3

4) PIO

.V

0) 4) • 4J P

O V

•H x-l •O 03

(V 01 P P

a a

p

0^ /V o V

tvp

3

bO

fa

42
I

0• q 0 0 0 0 0

0 d d d d

• 0 0 LO 0 ID 0 LO

ro OJ OJ T—

1

o

o

OJ

o• o

o

OJ

o

o

CO

o

o

OJ

o

o

00

o

o

o

a

3o

a

o

bJ

0) X

05 • QJ Hcm

e

T3

05 O l-l • 3 '3- 00

03 N

05e 00. T3C 0) a 05p

•3 3

05 4J

4-5 oa C5 U

t4 05 "3 D.

05 e

Ui 05

0-1 4-5

CM g

•H

4J 4J 05 3 3 t-i l-i t-i 3

3 3b bO C50 3

•H tH 3

>4.1 (4-1 e § § 3 O

hV>3 T3 3 05 •

4-5 4-5 3 3

•wi 1-5 T3 "3 3 3 3 ^

CX O.

/V O

V- m

s

3

t50

(H5 ) iN3ianv 3Aoav 3sia 3aniva3di^3i

43
1400

1200

1000

800

600

400

200

0

Regular seats

Blocking layer seats

40 80 120 160 200

TIME, t (s)

Figure 16. Estimate for Q for arrays of polyurethane seats

with and without®Blocking layers [13] ( extrapolated

from curves of [13])

44
o o o o o

o o o o o

CO lO CO CVi

(Xo) iNBiai^v 3A08V 38IU BunivuadtMai

45

TIME

(s)

Figure

17.

Predicted

lower

surface,

post-crash,

temperature

of

the

honeycomb

celling

material

in

a

cabin

outfitted

with

polyurethane

seat

cushions;

with

(

. )

and

without

(

)

blocking

layers
NBS.114A iREv. 2 -80 ) U.S. DEPT. OF COMM.

BIBLIOGRAPHIC DATA

SHEET (See Instructions)

1. PUBLICATION OR

REPORT NO.

2. Performing Organ. Report No. 3. Publication Date

NBSIR 84-2912 October 1984

4

.

TITLE AND SUBTITLE

The Thermal Response of Aircraft Cabin Ceiling Materials During a Post-Crash,

External Fuel-Spill, Fire Scenario

5.

AUTHOR(S)

Leonard Y. Cooper

6.

PERFORMING ORGANIZATION (If joint or other than NBS, see Instructions) 7. Contract/Grant No. national bureau of STANDARDS

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

I. Type of Report & Period Covered

9.

SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (Street, City, State, ZIP)

Federal Aviation Administration

Technical Center

Atlantic City, N.J. 08405

10.

SUPPLEMENTARY NOTES

I I Document describes a computer program; SF-185, FIPS Software Summary, Is attached.

11.

ABSTRACT (A 200-word or less factual summary of most significant Information, If document Includes a significant

bibliography or literature survey, mention It here)

An algorithm is developed to predict the thermal response of aircraft ceiling materials

during a post-crash fire scenario. The scenario Involves an aircraft's emergency exit

doorway which opens directly onto the flames of an external, fuel-spill fire which

engulf a large portion of the fuselage. Data of near-celling temperatures acquired

during a series of eight, full-scale, wlde-body aircraft cabin, post-crash test

simulations provide indirect validation of the algorithm. These tests involved cabins

outfitted with single, mockup seats. Two other full-scale cabin tests Involving

fire spread through twenty-one seat arrays with different types of seat construction

provide the input data required to exercise the algorithm in evaluations of fully

outfitted cabins.

Relative to the post—crash scenario, a measure of cabin fire safety is proposed, viz.,

the post-crash time—to-ceiling—ignition. This measure would be used as a surrogate for

the post-crash time available for passengers to safely evacuate the cabin. In this

sense, the algorithm is exercised in an example evaluation of the fire safety of a candidate honeycomb ceiling material used together in cabin systems Involving polyurethane ceiling seating, with or without a blocking layer.

12.

KEY WORDS (S/x to twelve entries; alphabetical order; capitalize only proper names; and separate key words by semicolons)

aircraft compartments; aircraft fires; algorithms; ceilings; egress; exits;

fire safety; fire spread; upholstered furniture

13.

AVAILABILITY 14. NO. OF

PRINTED PAGES

Unlimited

03] Pot Official Distribution. Do Not Release to NTIS

Order From Superintendent of Documents, U.S. Government Printinf Office, Washington, D.C.

20402.

52

15. Price

Q3] Order From National Technical Information Service (NTIS), Springfield, VA. 22161 $10.00

uscomM'DC e04S>pao