i*^ 0F o0j%

Vau of

*"

on NBS TECHNICAL NOTE 1128

U.S. DEPARTMENT OF COMMERCE/ National Bureau of Standards

An Oxygen Consumption Technique

for Determining the Contribution

of Interior Wall Finishes

to Room Fires

)'53

1128

SO
NATIONAL BUREAU OF STANDARDS

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'Headquarters and Laboratories at Gaithersburg, MD, unless otherwise noted; mailing address Washington, DC 20234. 'Some divisions within the center are located at Boulder, CO 80303.
An Oxygen Consumption Technique

for Determining the Contribution

of Interior Wall Finishes to Room Fires

°* STAKDABDt

tlBHABY

JUL 2 5 1980

Hot a_cc- -Ore

aaoo — /lb. //32

Darryl L. Sensenig

Center for Fire Research

National Engineering Laboratory

National Bureau of Standards

Washington, D.C. 20234

Sponsored in part by:

Armstrong Cork Company

Lancaster, Pennsylvania 17604

^

a

***eAu o* * i

/tJnr),ej>.\ mk~

U.S. DEPARTMENT OF COMMERCE, Philip M. Klutznick, Secretary

Luther H. Hodges, Jr., Deputy Secretary

Jordan J. Baruch, Assistant Secretary for Productivity, Technology and Innovation

5. NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director

Issued July 1980
National Bureau of Standards Technical Note 1128

Nat. Bur. Stand. (U.S.), Tech. Note 1 128, 87 pages (July 1980) CODEN: NBTNAE

U.S. GOVERNMENT PRINTING OFFICE

WASHINGTON: 1980

For sale by the Superintendent of Documents, U.S. Government Printing Office, Washington, D.C. 20402

Price $3.75

(Add 25 percent for other than U.S. mailing)
TABLE OF CONTENTS

Page

LIST OF FIGURES v

LIST OF TABLES vi

NOMENCLATURE vii

ABBREVIATIONS ix

Abstract . 1

1. INTRODUCTION 2

2. DETERMINATION OF HEAT RELEASE BY OXYGEN CONSUMPTION 5 2.1 Rate of Heat Release

2.2 Effective Heat of Combustion 8

3. APPLICATION OF OXYGEN CONSUMPTION TECHNIQUE 9 3.1 Laboratory Bench Test for Determining Rate of Heat Release and

Effective Heat of Combustion

3.1.1 Construction Details 9 3.1.2 Verification of Method 11

3.1.3 Illustration of Typical Results 12

3.2 Quarter-Scale Model Compartment 14

3.2.1 Construction Details 14

3.2.2 Apparatus to Measure Rate of Heat Release from Burning

Materials in Model Compartment 17

3.2.3 Convected Heat Through Doorway and Verification of

Method 19

3.2.4 Method of Obtaining Rate of Heat Release Per Unit Area

of Burning Surface in Model Compartment 20

3.3 Full-Scale Room Burns — Details of Room and Exhaust

Collection System 21

4. RESULTS 22

4.1 Quarter-Scale Model Compartment

4.1.1 Total Heat, Convected Heat, Upper Room Air Temperature,

Wall Ignition Time, and Flashover

4.1.2 Area and Rate of Heat Release Per Unit Area of Burning

Wall at Various Times Prior to Flashover 25

4.2 Full-Scale Room Burn 26

4.3 Laboratory Bench Tests . 29

4.3.1 Rate of Heat Release 29

4.3.2 Flame Spread 31

4.4 Comparison of Heat of Combustion With Effective Heat of

Combustion 33

5. RATIONALIZATION OF OBSERVED FLASHOVER TIMES IN MODEL COMPARTMENT . 34

6. SUMMARY 36

7. RECOMMENDATIONS 38

8. ACKNOWLEDGMENTS 39

9. REFERENCES 40

ill
TABLE OF CONTENTS (continued)

Page

APPENDIX A. Technical Data A-l

APPENDIX B. Scanning Technique Used to Determine Mean Velocity

Across the Exhaust Pipe Cross Section B-l

APPENDIX C. Crib Tests C-l

APPENDIX D. Errors in Calculation of the Rate of Heat Release .... D-l

IV
LIST OF FIGURES

Figure 1.

Figure 2.

Figure 3

Figure 4

Figure 5.

Figure 6.

Figure 7

Figure 8. Figure 9

Figure 10.

Figure 11,

Figure 12,

Figure 13.

Figure 14,

Figure 15,

Figure 16

Figure 17,

Figure 18

Figure 19

Figure 20

Page

Apparatus for measuring rate of heat release by oxygen

consumption 41

Comparison of the rate of heat release from burning

methane of known heat of combustion (HOC) at various flow

rates with rate of heat release by oxygen consumption ... 42

Comparison of the rate of heat release from burning

methane at a flow rate of 2.53 £/min (1.43 kW) with rate

of heat release by oxygen consumption 43

Rate of heat release -- Ohio state calorimeter and

oxygen consumption bench test comparison 44

Rate of heat release and effective heat of combustion of

16-mm particleboard ( ~ 50% RH) — bench test results ... 45

Rate of heat release and effective heat of combustion of

16-mm particleboard at 0% and 50% RH — bench test results 46

Model compartment — sketch 47

Model compartment — cross section 4 8

Quarter- scale model compartment 49

Full size room 49

Comparison of convected heat into hood with total rate of

heat release measured by oxygen consumption from burning

methane at various rates of heat release 50

Model compartment results -- 12-mm low density wood

fiberboard 51

Model compartment results -- 16-mm particleboard 52

Model compartment results — 2-ply 6-mm hardboard 53

Area of burning wall involved at various times to

flashover, quarter-scale model compartment — 12-mm

low density wood fiberboard 54

Area of burning wall involved at various times to

flashover, quarter-scale model compartment -- 16-mm

particleboard 55

Area of burning wall involved at various times to

flashover, quarter-scale model compartment -- 2-ply 6-mm

hardboard 56

Area of burning wall involved at various times, full-scale

room burn — 12-mm low density wood fiberboard 57

Area of burning wall involved at various times to

flashover, full-scale room burn — 2-ply 6-mm hardboard . . 58

Upper room- air temperature in center of room for full- and

quarter- scale enclosures lined with hardboard at 0% RH as a function of rate of heat release as determined by oxygen

consumption 59 v
Figure 21,

Figure 22,

Figure 23.

Figure 24,

Figure 25,

Figure 26,

Figure 1C.

LIST OF FIGURES (continued)

Page

Rate of heat release — low density wood f iberboard

bench test results 60

Rate of heat release — particleboard bench test

results , ... 61

Rate of heat release — hardboard bench test results ... 62

Flame spread distance vs time — low density f iberboard . . 63

Flame spread distance vs time — hardboard 64

Rate of flame spread as a function of surface

temperature 65

Rate of heat release from 3.4 kg cribs of polyurethane

and sugar pine under large hood c~2

LIST OF TABLES

Page

Table 1. Heat produced per volume of oxygen consumed for some common

materials 66

Table 2. Comparison of effective net heats of combustion with actual

gross and net heat of combustion (MJ/kg) 67

Table 3. Data summary for prediction of flashover times in model

compartment 67

VI
NOMENCLATURE

2 A = Cross sectional area of exhaust duct (m )

a = Number of carbon atoms in fuel molecule

2 b = Number of hydrogen atoms in one fuel molecule

c = Number of oxygen atoms in one fuel molecule

C = Heat capacity (kJ/kg K)

C. = Empirical constant defined by equation 13

C = Ratio of average velocity to centerline velocity in the exhaust duct

AP = Pitot-static tube differential pressure (Pa)

AT = T, - T,. (K) 1 A

3 D = Density of saturated water vapor in air (g/m )

E = Oxygen analyzer output during test (volts)

E = Oxygen analyzer output with normal air (volts)

2 F = Characteristic flame spread constant (m K /s)

h = Height of door (m)

2 H = Heat flux to surface (kW/m )

X = Thermal conductivity (kW/m K)

H = Length of preheating region for flame spread (m)

L = Length of boundary between burning and non-burning area (m)

m = Mass flow rate (kg/s)

ft = Molar flow rate of air into system

3.

ft. = Molar flow rate of argon into system

A„ . = Molar flow rate of water vapor into system H2°

fL. = Molar flow rate of nitrogen into system

w

2

ftn

= Molar flow rate of oxygen into system U

2

ft = Molar flow rate of combustion gases in the exhaust duct

Q = Total rate of heat release (kW)

Q = Rate of convected heat through the doorway (kW)

- 2 q = Mean rate of heat release per unit area (kW/m ) Q* = Critical rate of heat release at flashover in model compartment (kW)

R = Relative humidity (%)

P = Density (kg/m ) vii
NOMENCLATURE (continued)

3

P = Density of ambient air (kg/m )

p* = Density of ambient air (g/mole)

3

p = Density of stack gases (kg/m )

p* = Density of stack gases (g/mole)

t = Time (s)

t* = Time from ignition to flashover (s)

t

f = Flashover time (s)

t. = Ignition time (s) iy

T = Temperature of exhaust gas at point of pressure measurement (K)

T

1 = Temperature of exhaust gas entering hood (K)

T = Ambient temperature (K)

T = Standard temperature (298 K)

T = Pyrolysis temperature (K)

T = Material surface temperature (K)

9 = Oxygen depletion - the fraction of incoming oxygen molecules removed by

combustion

v_ = Flame spread velocity (m/s)

r V = Volume flow rate of air into the system referred to standard

a

3 conditions (m /s)

3 Vq = Volume flow rate of exhaust gas referred to standard conditions (m /s) w = Width of door (m)

X = Volume fraction of oxygen in the analyzer

X° = Volume fraction of oxygen in the analyzer prior to the test (oxygen

concentration of dry air)

X„ = Volume fraction of water vapor in normal air

X = Volume fraction of oxygen in normal air

X„ = Volume fraction of oxygen in exhaust duct

vxn
ABBREVIATIONS

CH - Convected heat

FB - Fiberboard

FO - Flashover

FSC - Flame spread classification

HB - Hardboard

HOC - Heat of combustion

MC - Model compartment

RH - Relative humidity

RHR - Rate of heat release

RWL - Rate of weight loss

URT - Upper room temperature

IX

AN OXYGEN CONSUMPTION TECHNIQUE FOR DETERMINING THE

CONTRIBUTION OF INTERIOR WALL FINISHES TO ROOM FIRES

Darryl L. Sensenig^

An oxygen consumption technique was developed for

determining the total rate of heat production in a room

fire. This was accomplished by measuring the volume flow

rate and the oxygen concentration of the exhaust gases

flowing through a collection hood. This method can be used

with unknown combinations of burning materials including

both interior finish and furnishings. By simultaneously

measuring the rate of oxygen consumption and the rate of mass loss the effective heat of combustion of the wall

linings were determined in a reduced- scale model room fire

test. The average heat release rate per unit area of the

wall linings was determined by recording the area of

involvement during the test and dividing this area into

the total rate of heat production at that time. The

enthalpy of the exhaust gases passing out of the doorway

was determined with the aid of an array of thermocouples

located at the entrance to the exhaust duct. By subtracting

the enthalpy flow through the doorway from the total rate

of heat production in the room, the heat losses through

the bounding surfaces were determined. Reduced-scale

and full-scale room fire tests and a bench test for heat

release rate using the oxygen consumption technique

This work was conducted from September 1976 to September 1977 while

Mr. Sensenig was a research associate at the National Bureau of Standards

for the Armstrong Cork Company.
are discussed in this report. Lateral flame spread

rates on vertical surfaces measured in the model room

fire tests and in a laboratory bench test are also

described.

Key words: Fire tests; flame spread; flashover; heat

release rate; ignition; interior finish; oxygen depletion;

room fires.

1. INTRODUCTION

The flame spread classification (FSC) obtained by the ASTM E-84 tunnel

2 test [1] is generally used by the building codes to regulate the surface

burning characteristics of interior finish materials. However, in order to

predict the actual contribution of these materials to fire growth in a room

it is necessary to perform expensive large-scale room fire tests. The number

of such tests may be reduced if a reliable analytical prediction model can be

developed which uses the fire properties of the interior finish materials as input data. A complete analytical description of the fire development in a room is extremely complex, considering the interactions of combustion

chemistry, heat transfer and fluid mechanics. Even the simplest fire remains

to be satisfactorily described. Much of the difficulty in the development of

a valid analytical model stems from the inability to properly interpret and use fire descriptors such as the ignitability, rate of heat release (RHR)

,

and rate of flame spread on materials as obtained from laboratory fire tests.

The influence of each of these "fire properties," as expressed by numerical

values from standardized laboratory tests, on room fire buildup has been

assessed by Fang [2]

.

2 Numbers in brackets refer to the literature references listed at the end of

this report.
A major effort is currently being devoted to the development of the

mathematical models. This requires reliable comparisons of the fire properties

of interior finish materials as they exist in room fires with those measured

in laboratory tests under conditions approximating those in the room. It

also requires reliable comparisons with important parameters in room fire

tests such as the total rate of heat production and the total heat loss rate

through the bounding surfaces of the room. There are no methods currently

available to determine these important quantities when combustible interior

finish is involved in the fire. There are heat release rate calorimeters

[3,4] which can be used to measure the heat release rate per unit area as a function of the external radiant flux. However, in order to establish the

proper exposure conditions in these calorimeters it is necessary to compare

these measured heat release rates with the heat release rate per unit area of

the burning material measured in a room fire. There has been no way of

obtaining the latter quantity.

A method based on oxygen consumption was investigated as a means of

measuring the instantaneous rate of heat production in the room fire. This

technique is based on the observation that for a given volume of oxygen

consumed there is a fixed amount of heat released, nearly independent of the

type of material burned [5] .

As combustible volatiles are being released during the fire, air

containing oxygen for their combustion enters the room through the lower part

of the doorway. The oxygen depleted gases are exhausted through the upper

part of the doorway serving as the single exit. Combustion can occur totally

within the compartment or flames can extend through the doorway as a result

of excessive volatilization. In either case, by analyzing the combustion

products collected in an exhaust system positioned above the doorway, the

amount of oxygen consumed in the burning process can be measured. Multiplying

the volume rate of oxygen consumed by the heat produced per unit volume of

oxygen provides a direct measure of the rate of heat released from the burning

materials within the compartment. 3
In order to verify the oxygen consumption technique for measuring heat

produced, it was first studied as a laboratory bench test, and then applied

to a quarter- scale model compartment fire and finally to a full size room

fire. The rate of heat release was measured by this method on small vertical

specimens in the bench test for various conditions of ignition, heat flux and

preheat. These conditions had a significant influence on the curves for the

rate of heat release per unit area versus time. In order to determine the

most suitable conditions for direct application to the room fire, model

compartment fire tests were conducted. A quarter—scale model ofa3x3x2.4

compartment containing combustible walls, a noncombustible ceiling and no furnishings was constructed for this purpose. The fire buildup created in

the model compartment by the combustible interior wall finishes was assessed

visually and by measurements of the total rate of heat release, area of

involvement, temperatures and heat fluxes. Accurate determination of the

total rate of heat release and the instantaneous burning area of the

compartment walls permitted comparisons with laboratory test measurements of

the rate of heat release per unit area. These comparisons then suggested the

most appropriate exposure conditions in the laboratory test for direct

application of these data to predict the fire buildup in the model compartment

The rate of heat produced obviously depends on the area burning, therefore,

another study was conducted in order to obtain basic information on rate of

flame spread. Many standardized "flame spread" tests exist; e.g., ASTM E 84,

ASTM E 162, etc., but for proper comparison, it is necessary to select

conditions of exposure and methods of computation that describe quantitatively

the flame spread along the wall surfaces within the room. In an attempt to

simulate the lateral flame spread occurring on the walls in the model room,

a bench test was set up to determine the rate of horizontal flame spread on

a vertical surface as a function of surface temperature.

The time to ignition of the wall is another important parameter that

must be considered in developing an analytical model describing fire buildup

in a room. This parameter was investigated with a gas diffusion burner,

4
representing a small quantity of burning contents or a small ignition source placed in the corner of the compartment with the flames impinging on the wall.

This study provides important information on the proper interpretation

and use of laboratory test data for their eventual application to an analytical approach for predicting fire buildup in a room, up to and including

flashover. Flashover is defined here as the condition when there is sufficient

radiation from the heated gases and upper room surfaces to ignite light

combustible materials in the lower part of the room.

2. DETERMINATION OF HEAT RELEASE BY OXYGEN CONSUMPTION

2.1 Rate of Heat Release

A method for measuring the heat release rate of burning materials was described by Parker [5] in connection with the ASTM E 84 tunnel test. This

method is based on the assumption that for a given volume of oxygen consumed

there is a fixed amount of heat released, nearly independent of the type of

material burned even though the heats of combustion may be considerably

different as seen in table 1. Similar computations were made from a table of

heats of combustion of organic compounds listed in the Handbook of Chemistry

and Physics [8] . The heat produced per volume of oxygen consumed was found

3 3 to be 16.7 MJ/m with a range of + 1 MJ/m at 25°C and one atmosphere

3 pressure. The product of this constant and the volume rate of oxygen

consumption provides a direct measure of the rate of heat release from a burning object. A small correction needs to be applied for carbon monoxide

3A more comprehensive review of the heat release per volume of oxygen consumed

done after the completion of this work suggests that a better value would be

17.1 MJ/m 3

, i.e., an increase of 3% [6], This difference arises partly

because the calculations on which table 1 were based assumed that all the

hydrogen went to water, whereas, for example, in PVC part of the hydrogen

goes to HC&; also the value for cellulose in table 1 was based on a value of

16.7 MJ/kg for the gross heat of combustion whereas a better value is

17.5 MJ/kg [7]
when it is formed in significant quantities since the heat release per volume

3 of oxygen consumed for combustion to CO (10.3 MJ/m ) is significantly less

than for combustion to C0

2

. Measurement of the CO/0,, ratio in the combustion

products should permit a suitable correction; however, it was not accounted

for in this study. The heat release rates discussed in this report are based

on the net heat of combustion (i.e., the water formed is assumed to be in the

gaseous state)

.

The volume rate of oxygen consumption was measured by collecting all of

the combustion products from a burning object along with an arbitrary amount

of surrounding makeup air. This mixture of air and combustion gases was collected in a hood and discharged at a measured volume flow rate and

temperature through an exhaust duct where it was analyzed for oxygen concentration. When there is no burning taking place, 20.8 percent of the volume

discharged is oxygen, assuming 293 K and 40% RH. For dry air the normal

oxygen concentration is 21 percent [5] . However, when burning occurs, the

percentage of oxygen in the flow discharged is lowered. The rate of heat

production, Q, is proportional to the rate of oxygen consumption which is

equal to the difference between the oxygen flowing into and out of the system

Q = 1.67 x 10 4 (XV - X_V_) (1) o a o o

where X is volume fraction of oxygen in normal air, V is the volume flow

rate of air into the system, Xq is the volume fraction of oxygen in the

exhaust duct and V„ is the total volume flow rate in the duct where V and

b a V refer to standard conditions. Note that Q in equation 1 is expressed in

3 kW while the heat produced per volume of oxygen consumed is in MJ/m thus the

3

" * factors of 10 in the numerical coefficient. It was assumed that V

a = Vg

and

that the oxygen concentration in the oxygen analyzer was the same as that in

the exhaust duct. Then

Q = 1.67 x 10 4 V

S

X

Q ( -^- ] = 1.67 x 10 4 V

g

X

o

6 (2)
where the oxygen depletion, 8, is defined within the limitations of these

X - X

g assumptions as 9 = —- . While these assumptions are not strictly

A.o correct, the approximation is adequate for organic solids at small oxygen

depletions. The extent of the error introduced is examined in appendix D.

From equation 2 it is seen that the effect of adding dilution air is simply

to increase V- and decrease 8 while keeping their product constant for the

same heat release rate.

The volume flow rate in the stack referred to standard conditions is

given by

/ Pn T \ 1/2 1/2

V

S = CV

A [jr

^

T^

where Cvis the ratio of the average velocity in the stack to the centerline

velocity (0.85 for the 5.08 cm diameter stack used in these experiments).

-3 2 A is the cross sectional area of the stack (2.03 x 10 m ) 3

p is the density of ambient air (1.18 kg/m ),

P is the standard pressure (1.01 x 10 Pa, 760 mm Hg)

,

P is the ambient pressure (Pa)

,

T is the standard temperature (298 K)

,

T is the temperature in the stack (K)

,

AP is the pressure difference across the pitot-static tube (Pa)

It was assumed that the ambient density of the stack gases is the same

as air. The error introduced is small and is discussed in appendix D. Then

a / , P T \ 1/2 " 1/2

Q = 1.67 x 10 4 X

Q

CV

A ( ^- p

2

f

2

) (AP) 6 (4)

P or substituting in the values for the above quantities assuming =— = 1

Q = 134 (AP/T) 1/2

8 (5)

Experimentally

,

E - E

-2= (6) E

o
where E and E are the oxygen analyzer output voltages for ambient and test

conditions respectively. It is assumed that the analyzer has a linear output

and that E = when X^ =0. It was assumed in equation 5 that X = 0.208

°2 ° which is the oxygen volume fraction of normal air at 293 K and 40% RH. The

calculation of the rate of heat release by this technique neglects the

increase in molar flow rate in the duct due to the generation of gases in

the fire which pass through the system uncombusted or produce more moles of

combustion products than the moles of oxygen consumed.

2.2 Effective Heat of Combustion

The heat of combustion is generally obtained by burning a specimen of

known weight in a pure oxygen atmosphere within a pressurized "bomb" and

measuring the total heat liberated. A material undergoes complete oxidation

under these conditions, which is often uncharacteristic of that which occurs

in a room fire. The potential heat of a material, defined as the difference

between its heat of combustion and the product of the mass fraction and the

heat of combustion of the residue remaining after exposure to a simulated

standard fire (750°C for 2 hours) , is a method of taking into account a more

realistic fire condition [9]. Most organic materials undergo complete

combustion to CO- and H„0 in the potential heat test. However, in the room

fire, flaming combustion of many materials is accompanied by a char forming

process which reduces the effective heat of combustion during the flaming

phase. The char structure formed has a much higher heat of combustion, which

is released during the glowing combustion stage of the fire. In order to

obtain the effective heat of combustion of materials under conditions

simulating the early stages of a room fire, a load cell was built into the

bench test for the rate of heat release so that the rate of weight loss and

rate of heat release could be obtained simultaneously.

The rate of mass loss was computed from the difference in mass between

15 seconds prior to and 15 seconds after' the time of the oxygen determination.

8
The ratio of the rate of heat release to the rate of mass loss is the

instantaneous effective heat of combustion which was measured as a function

of time throughout the test. If the specimen is completely consumed except

for its ash, the integral of the heat release rate over the duration of the

test divided by the total mass loss should yield the net heat of combustion.

This can be compared with the gross heat of combustion obtained in the oxygen

bomb calorimeter if the specimen is completely consumed and a suitable

correction has been made for the heat of condensation of the water formed

during the combustion.

3. APPLICATION OF OXYGEN CONSUMPTION TECHNIQUE

3.1 Laboratory Bench Test for Determining Rate of Heat Release

and Effective Heat of Combustion

3.1.1 Construction Details

A schematic of the apparatus is shown in figure 1. The specimen was placed in a holder designed so that burning was restricted to the front

surface, which was exposed to the flame and/or radiant flux, and could not occur along the specimen edges. The specimen was placed on a 0.5 kg

capacity water cooled strain gauge load cell calibrated for a sensitivity of

50 g/V and was positioned directly beneath a hood to collect the combustion

products. The combustion gases, as well as the surrounding makeup air, were

drawn into the hood and through a pipe of 50.8 mm diameter and 1.5 m length

3 by a squirrel cage blower operating at a nominal flow rate of 0.7 m /min. A

pitot-static tube and a thermocouple were placed in the center of the flow

stream at the mid length of the pipe to measure the velocity and temperature

needed for the volume flow calculations. The mean velocity across the pipe

cross section was found to be 0.85 of the maximum velocity in the center of

the pipe. This correction is due to the shape velocity profile which was

9
determined by scanning the cross section of the pipe with a pitot-static tube

using the averaging scheme found in the literature supplied by the manufacturer

of the pitot tubes (see appendix B)

.

The oxygen concentration in the exhaust stream was measured with an electrochemical oxygen cell which uses a KC1 solution with gold and zinc

electrodes. A tube for sampling the gas was placed in the exhaust pipe down

stream from the pitot-static tube and the thermocouple. The sample of exhaust

gas first passed through a dry ice cold trap and a glass fiber filter to

remove water and particulates and then through a vacuum pump. A valve in

line with a rotameter between the oxygen cell and the exhaust from the vacuum

3 pump was set to produce a flow rate of 16 cm /s through the cell. The cell

has an output voltage proportional to the oxygen concentration. A 300 kQ, 10

turn potentiometer placed across the output of the cell provided a means of

adjusting the output voltage to 100 mV when normal room air was being measured.

An output reading of mV was set to correspond to percent oxygen using pure

nitrogen. Millivolt deviations from 100 mV then are read directly as percent

oxygen depletions which are used in the calculation of rate of heat release.

Since the heat losses do not affect the calculated rate of heat release,

adiabatic chambers or constant temperatures within a confined space are not

necessary. In fact, the specimen is in an open environment without visual

restrictions except for the electric panel radiant heat source placed in

front of it for exposure to various heat flux levels. A study of the effect

of reduced oxygen levels on the rate of heat release would require enclosing

the specimen in a chamber. This was considered beyond the scope of the study,

but could be significant in cases where materials are burning in reduced

oxygen atmospheres . Ignition of the specimen was accomplished with a gas

pilot burner. A small pilot burner placed near the top or bottom of the

specimen may not make an appreciable contribution to the rate of heat release.

However, if the pilot is a line burner as shown in figure 1 and the gas flow

10
rate is large enough to actually bathe the specimen in flame a correction

must be made for the heat produced by the burner.

3.1.2 Verification of Method

The oxygen consumption method for determining the rate of heat release

was verified by burning a material of known heat output beneath the

collection hood. Methane (technical grade) was used since its heat of

combustion and flow rate were accurately known. It was burned at various

flow rates and the results of the comparison of this known RHR with the

measured RHR by 0^ consumption is shown in figure 2. This shows that the RHR

results are directly proportional under steady-state conditions with an accuracy of + 10%.

Shown in figure 3 are the results of a step increase in gas flow. The

broken line in figure 3 is the computed rate of heat release from the burning

of the methane gas at a volume flow rate of 2.53 Z/min with a net heat of

3 combustion of 34 MJ/m corresponding to 1.43 kW. The solid line is the

computed rate of heat release as a function of time determined by the oxygen

consumption method. Good agreement was obtained with the exception of a short period following ignition and extinguishment where the steady-state

value was reached relatively slowly due to the limited time response of the

measurement system for the oxygen concentration. This time response, which

was about 20 seconds for a step increase to reach 95 percent of its final

value, may be a source of dynamic error for cases where rapid changes occur,

but in most cases, e.g., for the solid materials examined in this report,

this time response is sufficiently fast. For materials exhibiting more rapid

changes in the rate of heat release, a computer correction can be utilized

similar to the numerical technique used to correct the Ohio State University

(OSU) heat release rate calorimeter data for apparatus delay [10]

.

11
Another means of verification of the oxygen consumption method is to

compare results with those obtained by other calorimetric techniques. The

rates of heat release of four materials were measured using the Ohio State

University calorimeter [4] and the oxygen consumption method. The OSU data was corrected for time response according to reference 10. The results are compared in figure 4. The specimen holder, specimen size 15.2 x 15.2 cm,

mounting technique, pilot ignition, and heat flux were identical for both

test methods in order to eliminate some of the possible sources of variation.

The results shown represent single runs on each apparatus rather than

averages, since only one test was conducted on each material by each method.

The agreement is quite good, perhaps better than would be expected,

since the differences are not much more than might be obtained from

repeatability checks on the same equipment. Slight discrepancies did exist

prior to the test termination for the fiberboard and polyurethane, but this was due to specimen fallout and combustion occurring on the back of the

specimen earlier in one test than in the other.

3.1.3 Illustration of Typical Results

Several rate of heat release test results on 16-mm thick particleboard

are shown in figures 5 and 6. These plots demonstrate typical results that

were obtained for an easily ignitable material with this equipment. In

figure 5 the three curves for the rate of heat release were obtained at

2 incident radiant exposures of 2, 2.5, and 3 W/cm . The exposure of the

specimen begins at time 0. There is a time delay before ignition and heat

release occurs; this depends upon the time required to attain the pyrolysis

temperature. During this time, the surface temperature increases until the

pilot flame at the bottom ignites the gases released from a small portion of

the specimen exposed to the pilot flame. The surface temperature of the

specimen continues to increase and the flame traverses the specimen surface

until the entire surface is aflame. Prior to total surface involvement, the

12
computed rate of heat release per unit area is disproportionately low, since

the results are obtained by dividing the rate of heat release by the exposed

surface area, even though it is not totally involved. Total surface

involvement occurs just prior to the peak heat release rate. The results

indicate that the higher the incident heat flux, the higher the peak rate of

heat release.

The effective heat of combustion obtained from the ratio of the rate of

heat release and the rate of mass loss is also shown in figures 5 and 6. In

the initial stages of the test as the temperature reached 100 °C, water was vaporized from the surface and the initial rate of mass loss was relatively

high compared to the rate of weight loss of the combustible volatiles.

Therefore, the effective heat of combustion curve, rather than being a step

function, gradually increased until a plateau was reached. The plateau

indicates that the effective heat of combustion was essentially constant.

Since the plateaus in the curves for the heat of combustion in figure 5 are identical for each exposure condition, it can be concluded that the effective

heat of combustion is independent of the radiant exposure over the range

tested.

The average effective heat of combustion is approximately 12.5 MJ/kg,

which is significantly lower than the net heat of combustion of 17.7 MJ/kg

for a specimen of this material conditioned to 50% RH at 23 °C as measured in

the oxygen bomb calorimeter.

These values are consistent with the fact that the water released in the

char forming process dilutes the combustible pyrolysis products and this

lowers the effective heat of combustion. Higher values might be expected

during the latter stages, but the tests were not carried far enough to obtain

a significant contribution from the char which was found to have a heat

of combustion of 33 MJ/kg. This argument is supported by figure 1C of

13
appendix C where an increase in the effective heat of combustion was obtained

during the latter stages of combustion of a wood crib.

The effect of the conditioning RH was also examined for its influence on

the rate of heat release and effective heat of combustion. The results shown

in figure 6 illustrate that predrying of particleboard not only increased

the rate of heat release, but also increased the average effective heat of

combustion by about 10 percent. The low absorbed water content minimized the

endothermic effect of water vaporization and thus permitted a higher rate of

temperature rise. This resulted in a higher pyrolysis rate and an increase

in the rate of heat release. The effective heat of combustion was increased

because the rate of weight loss of the specimens conditioned at 0% RH did not

include absorbed water loss.

3.2 Quarter-Scale Model Compartment

3.2.1 Construction Details

In order to examine the room fire phenomena in a laboratory size

enclosure a quarter-scale model of a 3x3x2. 4m room with a 0.76 m wide

and a 18.4 m high doorway was constructed utilizing the following scaling

rules [11,12]:

(1) Linear dimensions of the room are proportional to the scale factor.

(2) Height of doorway is proportional to the scale factor.

(3) Width of doorway is proportional to the square root of the scale

factor.

(4) Thicknesses of the wall and ceiling lining materials are independent

of the scale factor.

(5) Heat release rate of exposure fire is proportional to the square of

the scale factor, and

(6) Times are independent of the scale factor.

14
Although exact scaling is impossible this procedure has been found to provide

a similar fire buildup scenario on the two scales which is adequate to

demonstrate the application of the oxygen consumption technique to the study

of room fires.

The model compartment constructed for this study (see sketch, figure 7) was designed with two novel features: (1) a simple means of relining the

walls with new materials without completely dismantling the instrumentation;

(2) the ability to obtain accurate weight losses of the burning walls without

errors introduced by attached wires and water lines to transducers and with

minimal correction for buoyancy forces on the floor and the ceiling. These

features were incorporated for expediency in conducting the tests and to

obtain accurate effective heats of combustion of the burning wall linings.

Since the test materials were placed only on the walls, the first design

feature was easily accomplished by the use of a removable, noncombustible

ceiling. In this way the wall specimens were easily inserted from the top.

The specimens were held in position by small clips at the top and bottom of

the wall. In order to obtain the weight loss of the walls with minimal

corrections from forces on the ceiling and floor, the ceiling and floor were

arranged to be completely independent from the walls. A cross section of the

model compartment is shown in figure 8. The interior dimensions were

0.75 x 0.75 x 0.60 m high, and the doorway dimensions were 0.38 m wide by

0.46 m high. There were two independent chambers. The outer chamber was the

structural housing forming airtight water seals with the inner chamber, at

the floor/wall and ceiling/wall intersections. The compartment ceiling was supported on the walls of the outer chamber, forming an airtight seal

preventing leaks of combustion gas, without contacting the inner chamber.

The inner chamber, used for holding the combustible walls in position, formed

an airtight enclosure except for the doorway and was suspended on three

equally spaced load cells located outside of the outer chamber while not

contacting the outer chamber. The inner chamber, including the combustible

15
walls, weighed approximately 75 kg and rested partially on load cells and

partially on springs. The load cells and springs were in parallel. The

total weight of the inner chamber was initially supported on the springs as shown in detail in figure 8. The adjustment of the load applied to the load

cells was accomplished by turning the thumb wheel supporting the spring and

inner chamber. This lowered the load cell attached to the support arm until

contact was made between the load cell and the pedestal. The thumb wheel was lowered farther until approximately 1 percent of the total 75 kg was supported

by each of the three load cells and the remainder supported by the springs.

The load cells had spring constants of 18.9 kg/mm, which were much higher

than the 0.6 kg/mm for the springs. Therefore, changes in weight of the

inner chamber resulted in minimal vertical displacement since most (97%) of

the weight change was seen by the load cell. The system was calibrated with

a set of weights. This design permitted high resolution of the weight change,

which was necessary since less than 1 percent of the total weight of the

inner chamber was lost during the burning of the compartment walls up to the

time of flashover.

With the requirement that the ceiling not be attached to the inner walls

because of buoyancy effects, a small gap was created along the wall/ceiling

intersection. The higher pressure in the upper compartment forced the flame

to lick back over the top of the inner wall and the flame was no longer

confined to traveling along the intersection. Because of this necessity for

a tight wall/ceiling intersection, a portion of the ceiling had to be attached

to the inner wall as shown in figure 8.

Pressures P, and P» above and below the ceiling attached to the walls

were essentially equal, thus no buoyancy correction was necessary. However,

for area A, pressures P, and P-. acted downward and upward respectively. Since

P-. was lower than P, a downward force was exerted on the inner chamber when

the upper room air temperature and pressure increased. A curve of air

pressure versus force correction was obtained by conducting burn tests in the

16
compartment containing noncombustible materials of constant weight. Thus an incremental force correction was applied to all weight loss data used for

calculations of the heat of combustion.

Other instrumentation within the compartment consisted of five vertical

assemblies "trees" of ten thermocouples each—one tree on each of the three

walls for measuring surface temperatures, one in the center of the compartment

to measure air temperatures from the floor to the ceiling, and one in the

doorway to measure the inflow air and exhaust gas temperatures. No

corrections were made for thermocouple radiation error. Radiation to the

floor was measured with a heat flux meter placed at floor level in the center

of the front left quadrant of the compartment.

A schematic of the model compartment is shown in figure 9. The doorway

in the model compartment was positioned off center on the front wall with one

side of the doorway 0.02 m from the corner, thus duplicating the physical

arrangement used in some previous compartment studies conducted at NBS . A

(0.08 x 0.08 m) gas diffusion burner having a flat surface of porous ceramic

fiber was used as an igniting source for the wall in the corner diagonally

across from the doorway.

The flow rate of the methane to the burner was adjusted to provide a heat input of 2.8 kW for all model compartment tests. This rate of gas flow

produced a flame height of approximately one half the wall height.

3.2.2 Apparatus to Measure Rate of Heat Release from

Burning Materials in Model Compartment

The quarter-scale model compartment was constructed in order to

determine the potential contribution of the interior wall finishes to room

fires. The oxygen consumption method was employed to determine the rate of

heat release. This method was applied to the compartment by placing an

17
exhaust hood directly above the doorway as shown in the sketch of the

apparatus in figure 7 so that the combustion products coming from the room

could be collected. The oxygen concentration in the combustion products was measured and the rate of heat release was calculated in the same manner as that used for the laboratory bench test method.

The difference between the exhaust collection system in the bench test

and the compartment was basically in the size and in the volume flow rate.

The combustion gases exhausted through the doorway of the model compartment,

as well as some of the surrounding air, were drawn into the hood by a squirrel

cage blower. The mixed gases were forced through a uniform pipe, 0.15 m

3 diameter and 4 m long, at a nominal flow rate of 7.1 m /min.

A pitot-static tube and four thermocouples connected in parallel were

placed near the center of the flow stream at the mid- length of the exhaust

pipe to measure the velocity and the temperature needed for the volume flow

calculations.

The mean velocity across the cross section of this pipe as measured by

the pitot-static tube was C = 0.95 of the maximum velocity measured in the

center of the pipe since the shape of the velocity profile was more uniform

in this larger diameter pipe. The measurement of the oxygen concentration

was identical to that described in section 3.1.1. Also mounted in the

exhaust pipe was a light obscuration meter for measuring the smoke generated

from the wall linings in the room. Smoke optical density measurements were

obtained but are not included in this report as their interpretation is

beyond the scope of this project.

18
3.2.3 Convected Heat Through Doorway and Verification of Method

Two important parameters in room fire tests are the total rate of heat

production and the total heat loss rate through the bounding surfaces of the

room. The total rate of heat production from burning materials within a room

can be determined by the oxygen consumption method. This heat is partly

absorbed by the interior surfaces of the room and partly radiated and

convected through the open doorway. The heat convected through the doorway,

Q , can be determined by measurements taken in the doorway, but more simply

by measurement of the temperature increase and volume flow rate of the gas

entering the hood.

Q

c = m C AT (7)

where m = mass flow rate, C = mean heat capacity of air at constant pressure

and AT = temperature difference between the air entering the exhaust hood and

the ambient air (T, - T ) (assuming that the values of the heat capacity and

_L a density of the exhausted combustion products are the same as those of air)

.

The mass flow rate through the exhaust pipe is equal to the product of

V„ and p measured at the pitot-static tube and is the same as that entering

the hood. The temperature of the mixture of surrounding makeup air and

heated gases from the burning compartment entering the hood was measured with

16 chromel-alumel thermocouples connected in parallel. Eight thermocouples

were equally spaced in the duct entrance, 0.08 m above the hood around a circle whose diameter was 0.5 times the diameter of the duct and the remaining

eight around a concentric circle whose diameter was 0.8 times that of the

duct. Air temperature measurements at the hood entrance are not affected by

heat losses in the exhaust system. Thus the enthalpy of the air entering the

hood closely approximates the heat convected through the doorway provided

there is no burning outside the doorway.

Equation (7) can be written in terms of measured quantities

19
1/2 / Ap \l/2

Q

c = CvAC (2p o

T

o

) (^j AT (8)

2

"} or substituting in the values for C„ = 0.95, A = 0.0182 m , p =1.18 kg/m , T

Q

= 298 K and C = 1.046 kJ/kg K,

Q„ = 0.48 (££) AT (9) T

)

Accuracy of the operation of the thermocouple array was checked by

burning methane gas with known flow rates directly beneath the exhaust

collection hood. Comparisons of the results in figure 11 exhibit good

agreement between the rates of heat production calculated from the air

temperature rise, the oxygen consumption, and the known flow rate of the

methane multiplied by its heat of combustion.

3.2.4 Method of Obtaining Rate of Heat Release Per Unit

Area of Burning Surface in Model Compartment

Another objective of this study was the measurement of the average rate

of heat release per unit area of the burning wall surface. within the

compartment. Comparison of these measurements with those obtained with the

laboratory bench test for the rate of heat release would thus establish the

most appropriate exposure conditions in the bench test for providing

equivalent room fire test data.

In order to determine the rate of heat release per unit area involved in

the compartment, simultaneous measurements of the total rate of heat release

and the area involved were made. The area involved was not measured

continuously while burning, but was obtained by conducting a number of tests

on different wall specimens of the same type of material and extinguishing

the flames at various times prior to flashover using a C0

2 extinguisher. The

three combustible walls were removed from the model compartment following

each fire test and the burned area was measured.

20
While the boundary between the areas involved and the areas lightly

charred but not burning was not well defined, the surfaces involved scratched

easily, whereas surfaces that had not begun to burn remained hard and did not

scratch. The areas involved were thus determined by lightly scraping the

surface with a sharp-edged tool. A tracing of the boundary was made and the

portion of the tracing paper representing the burning area was cut from the

overall tracing and weighed. The ratio of this weight to the weight per unit

area of the tracing paper was the method used for the area of involvement

determinations

.

The average heat release rates per unit area were obtained by dividing

the total rate of heat production at the termination of the test by the

measured area of involvement.

3.3 Full-Scale Room Burns — Details of Room

and Exhaust Collection System

The full-scale room burns conducted in this study were primarily

intended to demonstrate and confirm the application of the oxygen consumption

technique for measuring heat release in full-scale room fires.

To collect the combustion products from the room for the application of

the oxygen consumption method of determining heat release, a 1.2 x 1.5 m

hood, similar to that used in the model compartment, was constructed and

placed above the doorway. It was attached to a pipe 0.2 m in diameter

3 containing a high capacity blower (42 m /min) . The same procedure and

instrumentation described previously were used to determine the rate of heat

generated.

An ignition source, consisting of a horizontal gas diffusion burner

(0.3 x 0.3 m) , approximately 16 times the area of the one used in the model

compartment, was placed in one back corner of the room opposite the doorway.

21
The rate of gas flow to the burner was set 16 times higher than that used in

the model compartment to produce a flame height approximately one half the

height of the wall prior to wall ignition.

The dimensions and layout of the full size room are shown in figure 10,

and can be compared with those of the model compartment shown in figure 9.

This structure was not an exact 4X scale up of the model compartment. The

room had a doorway centered in one wall whereas the model compartment doorway

was off center. This difference did not appear to be critical to the overall

fire buildup. A significant difference was the unintentional departure from

proper scaling of the doorway height which was 1.88 m instead of the scaled

1.78 m height. This reduced the hot gas entrapment distance between the top

of the doorway and the ceiling as well as allowing too much airflow. The

consequence of the door opening is discussed in section 4.2. The second

full-scale room burn was conducted with the correctly scaled doorway height.

4 . RESULTS

4.1 Quarter-Scale Model Compartment

4.1.1 Total Heat, Convected Heat, Upper Room Air Temperature,

Wall Ignition Time, and Flashover

Three materials—low density wood fiberboard, particleboard, and

hardboard—were selected for this study on the basis of having different

rates of heat release, densities, and thermal properties, and should therefore

result in varying fire development in the model compartment. Each was essentially thermally thick for the preflashover period, and remained in

place when burned. The materials were readily available in convenient sheet

form. Specimens were installed on three walls of the model compartment; the

22
wall around the doorway was left unlined. The inner chamber was then

balanced on the load cells, and the ignition burner was placed in the corner

of the room, within 6 mm of the walls.

The rate of heat produced by the burning walls in the compartment plus

the heat from the ignition burner, 2.8 kW, are plotted as a function of time

in figures 12 to 14.

Three data points are labeled on the RHR curve indicating the time for

a particular occurrence in the compartment—wall ignition time, A; specific

area involvement, X; and flashover, 0. The burner was initially placed

arbitrarily within 6 mm of the walls. After obtaining significant variations

in the time of wall ignition, it was placed as close as possible without

contacting the walls. This resulted in shorter and more uniform ignition

times comparable to those obtained by the laboratory test for ease of ignition

[13] . See appendix A for comparison of test results of ease of ignition with

times for wall ignition.

2 The X on each RHR curve is the time required for the 0.093 m area of

the wall directly between the burner and the ceiling to become fully involved.

During this time the heat was transferred by direct flame impingement with

little additional external radiant heat flux. The flame spread direction was vertically upward during this period of fire development with the zone heated

by the flame extending far beyond the pyrolyzing area. To involve more wall

area, a transition occurred in the flame spread direction from vertically

upward to either lateral from the burning corner or vertically downward from

the ceiling. In this case the zone heated by the flame was small and the

preheating of the wall surfaces ahead of the flame by external thermal

radiation from other heated surfaces and by convective heating had a large

effect on the flame spread velocity. These distinctly different conditions

of flame spread would be characterized by different modes of heating if

comparable conditions were to be duplicated on a laboratory bench scale flame

spread test.
The symbol shows the time of flashover, defined in these experiments

as the time of ignition of crumpled newspaper on the floor of the compartment;

o this occurred at a measured mean radiant flux of 2.6 + 0.4 W/cm . The rate

of heat release at which room flashover occurred for the three materials

examined in the model compartment was 55+3 kW.

Several tests of a model compartment lined with each material were taken

to flashover, from which the mean RHR curves plotted in figures 12 to 14 were

obtained. Other tests were terminated by extinguishment of the fire with a COp extinguisher at selected times prior to flashover. The magnitude of the

rate of heat release at termination of each model compartment test is plotted

as a dot and labeled MC- ( ) on the mean RHR curve indicating the number of the

experiment. These values were used to determine the rate of heat release per

unit area as discussed in section 4.1.2.

Also plotted is the convected heat through the doorway. The ratio of the

convected heat (CH) to the rate of heat release (RHR) is approximately 0.5 for

each material, being slightly higher for the more thermally insulating

fiberboard. A higher ratio means less heat conducted to and through the

walls and ceiling and more heat loss through the doorway. The measurement of

the temperature rise of the gases entering the stack provides a simpler

approach to measuring convected heat through the doorway than the normal

method of dividing the doorway into sections and measuring the velocity and

temperature at each section. The use of the thermocouple grid in the hood

has the advantage of having a known mass flow rate and one whose spatial

distribution does not vary as greatly as the one in the doorway.

The third curve in figures 12 to 14 is the simultaneous plot of the

upper room air temperature (URT) measured 7.5 cm from the ceiling in the

center of the room.

24
4.1.2 Area and Rate of Heat Release Per Unit Area of Burning

Wall at Various Times Prior to Flashover

After termination of each model compartment test, at various times prior

to flashover, the three combustible walls were removed and analyzed for area involvement.

Tracings of the boundary between burned areas and non-burned areas for

each terminated test are shown in figures 15 to 17. Also included are the area involved and rate of heat release per unit area calculated at the

termination time for each test. Since the ignition time of the wall was not

constant for each test, due to inconsistent placement distance of the burner

from the wall, a corrected termination time is listed. The corrected

termination time is the time on the mean RHR vs time curve in figures 12 to

14 corresponding to the last RHR reading from each terminated test.

Also tabulated in the tables in figures 15 to 17 are test results on materials that had been conditioned at 01 RH prior to compartment testing.

Flashover for these tests occurred more rapidly, demonstrating that moisture

content can significantly alter the results of fire tests. All other

materials were stored in an area of near 50% RH prior to compartment testing.

As previously mentioned, one specific purpose of these compartment tests

was to determine the average rate of heat production per unit area involved.

This should provide insight into the proper use or a more meaningful

interpretation of the rate of heat release data from laboratory bench scale

tests. Within the compartment, each area element of the wall was exposed to

varying exposure conditions and was burning for different lengths of time.

This results in substantial differences in the rate of heat release for each

area element. The average rate of heat release per unit area was obtained

by dividing the rate of heat release by the area burning at the time of

measurement.

25
Significant differences in the rate of heat release per unit area were

obtained between the materials tested. The low density fiberboard had a

2 fairly constant and low rate of heat release per unit area of 6 W/cm . Nevertheless, flashover occurred in the shortest time—3 min 36 sec. The

compartments lined with particleboard and hardboard both required longer

times—8 min 45 sec and 8 min 53 sec, respectively, —to flashover, even though

their rates of heat release per unit area were higher. The particleboard had

2 a fairly constant rate of heat release per unit area of 9.5 W/cm from full

corner involvement to flashover, whereas the rate of heat release for the

2 2 hardboard is initially 11 W/cm , increasing to 16 W/cm at flashover. The

fact that the material with the lowest rate of heat release per unit area

caused flashover in the compartment in the shortest time emphasizes that the

time to ignition and the flame spread rate are also important fire parameters

for describing the fire buildup in a room.

4.2 Full-Scale Room Burn

Two full-scale room burns were run to establish a comparison of the

average rate of heat release per unit area in the quarter-scale and

full-scale room tests. Pictures and observations of the back wall marked

with grid lines were taken during the tests in order to determine the area of

involvement which is needed for calculating the rate of heat release per unit

area. The results from the model compartment and full-scale room burns were

in agreement. However, the fire growth in the first full-scale room test,

lined with low density fiberboard, was not a duplicate of the model compartment

performance. The room doorway was too high to satisfy the scaling rules, as was noted in section 3.3 and did not provide the properly scaled ventilation

(airflow) and trapping of the hot gases in the upper part of the room. Consequently, the room did not flash over and the severity of the fire was mild in comparison with the quarter-scale model compartment performance.

Details of room and model compartment interior dimensions and door sizes are shown in figures 9 and 10.

26
The boundary between the involved and non-involved areas for low density

fiberboard at various times during the test is plotted in figure 18. From

these plots, the burning area was determined. Time, room air temperature

10 cm below the ceiling, area involved, and rate of heat release per unit

area are included in the figure. Comparison with the mean rate of heat release

2 per unit area of 6 W/cm from the model compartment tabulated in figure 15

shows good agreement over the time period from 81 to 138 seconds.

Visual observations of the fire buildup in the room and the model

compartment were very similar during the first 2 min 15 sec. After this

time, areas which had been burning became extinguished with flaming occurring

only at the boundary between burned and unburned areas . Due to the higher

airflow and smaller distance from the doorway lintel to the ceiling in the

room, the temperatures in the upper part of the room remained lower than that

which occurred in the model compartment, providing insufficient heating of the

walls to sustain ignition after the char layer had been formed. Consequently,

the fire subsided and the test was terminated after 8 min by extinguishment

with water from a sprinkler.

Due to the limited availability of the full-scale test room, the

fiberboard was removed and the room was immediately relined with 2 -ply 6-mm

hardboard. The hardboard test was conducted on the following day. The

moisture content of the hardboard was probably higher than normal as a result

of the elevated relative humidity of the air from the water used to extinguish

the fiberboard test.

Properly scaled ventilation (airflow) was provided by lowering the height

of the doorway, see figure 10. The exhaust blower for the analysis system

was also modified for a higher volume flow to assure complete collection of

the combustion products over the whole fire growth period up to f lashover

.

27
The corner burner was ignited and after 2 min problems were encountered

with the exhaust system. The upper room air temperature had already reached

150 °C at which time the gas to the burner was shut off. The room was cooled

for one half hour to approximately 30 °C while new corner wall panels were

installed adjacent to the burner, and the blower motor was changed to prevent

another overload. Due to the 150°C temperature encountered in the upper gas

layer, the moisture content of the hardboard probably more closely approximate

the moisture content of the material conditioned to 0% RH that was tested in

the model compartment.

The full-scale hardboard test was restarted and ignition of the wall was

first observed at 1 min 30 sec. The boundary between the burning and non- involved areas at various times during the test is shown in figure 19. Also

included in this figure are the room air temperature 10 cm below the ceiling, area involved, and the computed rate of heat release per unit area. The first

computation of the RHR/area made at 2 min (approximately 1/2 min after the

first sign of wall ignition) is low in comparison with the model compartment

results, exhibiting a gradually increasing RHR/area as a result of higher

radiant flux on the burning surfaces.

During the latter stages of the test, the hood was not able to collect

all combustion products due to the limited blower capacity. Consequently,

the measured heat release rate after 5 min was low, causing incorrect RHR/area

results at flashover.

Since the area of the walls or fuel source of the full-scale room was 16

times that of the quarter-scale model compartment, then 16 times the rate of

heat release would be expected to produce the same upper air temperature at

identical times if properly scaled. The total rate of heat release from

hardboard conditioned at 0% RH and tested in the model compartment and the

total rate of heat release from the full-scale test scaled down by a factor

of 16 are plotted in figure 20 as a function of the upper room air

28
temperatures, 2.5 cm from ceiling in model compartment and 10 cm from ceiling

in full-scale room. The agreement is striking.

4.3 Laboratory Bench Tests

4.3.1 Rate of Heat Release

Various modes of operation of the laboratory bench test for rate of heat

release by 0» consumption were investigated, each significantly influencing

the rate at which heat was produced from the burning materials . These were conducted in order to establish the exposure conditions in the bench test

resulting in a mean rate of heat release per unit area the same as that

obtained from the model compartment. In order to apply rate of heat release

data from the laboratory bench test to describe the buildup in a room, it was desirable that the data be obtained at exposure conditions simulating those

in the room as closely as possible. These conditions included the preheat of

the walls from the hot upper gas layer and radiation from the surrounding

heated surfaces, in addition to direct flame heating and external radiation

during the burning period.

These conditions were closely simulated in the bench test by preheating

2 the specimen to a temperature of 290 °C using a radiant exposure of 2 W/cm . 2 An external radiant flux was then applied at levels of 0, 1, 2 and 3 W/cm

and the simultaneous application of a bathing flame from a line pilot burner

2 that produced an additional heat flux of 3.5 W/cm . The results of these

tests are shown in figures 21 to 23 along with the mean rate of heat release

per unit area (broken line) obtained in the model compartment tests.

To properly interpret the results from the bench test as they relate to

compartment results, the conditions with respect to time within the

compartment must be considered. Comparison on an equal time basis is not

valid, since the exposure conditions in the bench test were adjusted so that

29
ignition occurred at or near the time of zero minutes , whereas in the model

compartment, ignition occurred at later times. During the early stages of

fire growth the only involvement is the corner, directly above the burner.

This area is exposed to the bathed flame with little additional heat flux

from the surrounding surfaces. The rate of heat release with no external

radiation in the heat release rate calorimeter adequately describes the heat

released in the room fire during this time for all three materials. Later,

the upper room air and surfaces became heated thus exposing burning areas to

higher heat fluxes. These higher heat fluxes cause higher rates of heat

release unless the time for which the surface has been burning is on the

decreasing portion of the curve for rate of heat release. Both the fiberboard

and the particleboard exhibit a peak rate of heat release followed by a gradually decreasing rate. Areas that have been burning for some time at

low additional heat fluxes show decreasing rates of heat release. It is

speculated that burning areas that continue to receive gradually increasing

heat fluxes produce nearly constant heat outputs, while other areas just

becoming involved produce heat outputs higher than the mean. The mean heat

release rate per unit area from all of the burning areas remains nearly

2 2 constant at 6 W/cm for fiberboard and 9 . 5 W/cm for particleboard throughout

the model compartment burn duration.

The curves for the rate of heat release for hardboard exposed to a constant flux do not exhibit a peak followed by a gradual decrease with time,

but more nearly a plateau. The rate of heat release per unit area from

2 hardboard in the model compartment is initially 11 W/cm and increases to

2 16 W/cm .

The rate of heat release curves from the bench test are not extended to

the time to flashover because none of the area involved in the model

compartment has burned for that duration. In fact the maximum duration of

burn of any area is from ignition to flashover with the major portion of the

involved area at flashover becoming ignited after full corner involvement.

30
The time between full corner involvement and flashover, labeled X and

respectively in figures 12 to 14, is the basis used for defining the mean RHR

curves from the model compartment shown in figures 21 to 23.

A comparison of the mean heat release rate at each exposure condition in

figures 21 to 23 indicates that the rate of heat release for all three

materials tested in the model compartment may be approximated on the bench

test by exposure to a bathed flame with no external radiation initially,

2 increasing to 1 W/cm at a time comparable to the time duration from full

corner involvement to flashover.

4.3.2 Flame Spread

The results of section 4.1.2 demonstrate no direct relationship between

rate of heat release per unit area and the time to flashover. Other fire

parameters must also be important to adequately describe the severity and

rate of buildup of a room fire. One such parameter is the flame spread rate

which determines the rate of area involvement.

In the model compartment the flames from the fire source impinge upon

and then involve a portion of the wall after heating it to its pyrolysis

temperature. The additional fuel generation causes the flames to extend and

then the pyrolyzing area increases causing further flame extension until the

entire area directly between the burner and the ceiling is burning and the

flames are extending along the wall/ceiling intersections. Beyond this time,

the exposure conditions for further wall area involvement become quite

different. To involve more area the flame must propagate in directions

either perpendicular to or opposite to the fire plume flow, that is, lateral

or downward. The additional wall surfaces that become involved are convectively preheated by the hot, upper gas layer and radiatively preheated

by the other high temperature surfaces rather than by the direct fire plume

impingement. Flame contact, however, still supplies the heat needed to bring

31
the surface from the preheat to the pyrolysis temperature. If there is little

heat trapped in the upper room further spread of the flame will be minimal

and the conditions of an open corner fire test would be closely simulated.

However, if a sufficiently high upper room air temperature is obtained, the

flames will spread downward and laterally involving additional surface area until the rate of heat release becomes sufficient to flash over the room.

To simulate the flame spread on the model compartment walls during the

latter stages of fire growth resulting in flashover, lateral flame spread

tests on a vertical surface were conducted. Vertical specimens 7.5 cm high

x 30 cm long were exposed to various levels of radiant heat flux with the

simultaneous application of a vertical line pilot burner at one end of the

specimen.

The flame propagation distance was measured as a function of time

(figures 24 and 25) . The average surface temperatures of the non-burning

area ahead of the flame front were measured with surface applied thermocouples

and are recorded on each curve. The slope of these curves or rate of flame

spread was computed and plotted in figure 26 as a function of the surface

temperature.

For a semi-infinite solid exposed to a constant heat flux, H, the rise

in surface temperature is given by classical heat conduction theory [14] by

T - T = -2- -*£ (10) ° /fr /ApC

If the preheating region ahead of the pyrolysis zone of a thermally thick

combustible solid has an effective length £ and an average rate of heat

transfer H, the time to heat the surface from T at the leading edge of this

region to the pyrolysis temperature T is given by

TT_ ApC

4 2 H ( t

p

- ».r

32
and the flame spreading velocity is

2 v

f =

t = ¥I^c

= - — (11)

(t -t r h - t v

where A, p, and C are the thermal conductivity, density, and heat capacity of

the solid, and F is assumed to be characteristic of the material.

It was verified that for hardboard and fiberboard (particleboard was not

tested) the rate of flame spread could be expressed by this equation. This

was accomplished by conducting a nonlinear regression analysis where the

solution represented by the X's in figure 26 indicates a good fit. The T

for both materials was determined to be nearly identical at 350°C, which is

a reasonable approximation of the pyrolysis temperature of cellulose. The

proportionality constant F was found to be 3310 for fiberboard and 800 for

hardboard when the flame spread rate is expressed in meters per min or 55

and 13.3, respectively, when the flame spread is expressed in m/s. The ratio

of these proportionality constants indicates that the relative flame spread

rate for fiberboard is approximately four times that of hardboard at

equivalent surface temperatures. The F can be considered to be a characteristic of the lateral flame spread rate of a material on the wall and might be

used as one of the parameters needed to predict fire buildup in a compartment,

4.4 Comparison of Heat of Combustion With Effective Heat of Combustion

The effective heat of combustion of many solid materials differs

significantly from their actual net heat of combustion. To demonstrate these

differences, effective heats of combustion were obtained in the model

compartment and in the bench test for rate of heat release and compared to

the actual net heat of combustion obtained with the oxygen bomb calorimeter.

The results are shown in table 2

.

33
Formation of char in a pyrolyzing solid which was not burned in these

tests reduces the heat of combustion during the flaming phase. This does not

occur in the oxygen bomb calorimeter which consumes the char. This is at

least part of the reason for the 20 to 40 percent difference between the

effective net and actual net heats of combustion.

The differences in the bench test and model compartment results are partly attributable to moisture in the model compartment being vaporized from

heated surfaces that are not burning. This increases the rate of weight loss

which in turn lowers the apparent effective net heat of combustion.

5. RATIONALIZATION OF OBSERVED FLASHOVER TIMES IN MODEL COMPARTMENT

Ignition, flame spread rate, and heat release rate of combustible wall

linings all influence the rate of fire buildup in a room, ultimately leading

to a condition necessary for flashover. Each of these fire phenomena has

been quantified by measurements taken on laboratory bench tests, at exposures

approximating the conditions attained in the model compartment. Using each

of these properties is essential to evaluate a material's performance in the

model compartment.

The rate of heat release from the burning walls of the compartment at

any time is equal to the product of the area of involvement at that time and

the average heat release rate per unit area. This can be expressed as

follows.

Q = qLv

p

t

- 2 where q = mean rate of heat release per unit area, (kW/m )

,

L = length of the boundary between burning and non-burning area, (m)

,

v_ = flame spread rate, (m/s) , and

r t = time (s)

34
Note that L and vp are average values over time t.

The time difference t* from wall ignition t. to flashover t c is the lg ro time necessary to produce a critical rate of heat release and upper room air

temperature necessary for flashover. The critical rate of heat release, Q*,

at which the compartment used in this study flashed over was measured to be

55 kW for the three materials examined (see section 4.1.1). In terms of the

critical rate of heat release,

Combining equations 11 and 12 it is possible to express the flashover

time tf in terms of the time to ignition, the average heat release rate, and

a characteristic flame spread constant,

t. . t . ^Kl<£ (i . i

- ,.. . = t. + -A- (13) ig

q F

Actually T is a function of time but equation 13 is based on the

assumption that there is a characteristic surface temperature just ahead of

the flame as it moves across the surface. Substitution of the experimental

data of t . and tf from the model compartment and q and F from the bench

tests into this expression results in an empirical value for C. which is

nearly the same for both materials as seen in table 3

.

It can be seen, however, from the expansion of C. in equation 13 that it

would be expected to be strongly scale dependent through L and Q* , and material

dependent through T . It is also somewhat insulation dependent through Q*

,

(since a reduction in heat losses through the walls results in a lower

required heat release rate to produce flashover)

.

Even though only two materials were evaluated, it appears plausible that

this same constant C. may be applicable to other materials having the same T

35
if evaluated in like manner. This rationalization suggests a conceivable

approach for applying bench test results to determine a fire hazard rating

for interior finish materials.

6 . SUMMARY

1. A technique was developed to determine the rate of heat produced

from a fire by measuring the rate of oxygen consumed. The principle utilized

for this determination is that for a given volume of oxygen consumed there is

3 associated a fixed amount of heat produced of 16.7 MJ/m . It was verified

on a laboratory bench scale test using a known flow rate of methane that this

technique has an accuracy of + 10 percent when the composition of the burning

material is known. Taking into account the differences in heat released per

volume of oxygen consumed the accuracy is + 15 percent for unknown materials

of concern in the room fire environment. The rates of heat release are comparable when determined by oxygen consumption and by calorimetric methods.

The technique was demonstrated to be versatile in application to various

burning situations including a laboratory bench test for heat release rate,

a reduced- scale model compartment fire test, and full-scale room burns.

2. A quarter-scale model compartment was designed to evaluate the

contribution of wall linings to room fires in which the walls could be

continuously weighed during the fire with a sensitivity of a few grams. The

rate of weight loss in conjunction with the rate of heat release from the

compartment walls permits computation of the effective heat of combustion in

a simulated fire situation.

3. The laboratory bench test for determining the rate of heat release

was equipped with a load cell to simultaneously measure heat release and

weight loss rates in order to obtain the effective heat of combustion. The

effective heats of combustion measured in the model compartment were similar

36
to those obtained in the bench RHR tests, but both were significantly lower

(20-40%) than the net heat of combustion determined by the oxygen bomb

calorimeter.

4. Utilization of the model compartment permitted accurate computation

of the rate of heat release per unit area involved in a dynamic fire

situation. Comparison with the laboratory bench test for the rate of heat

release operated in various modes established that the heat release rates are similar when the bench test exposure conditions include a bathed flame

2 exposure with additional thermal radiation of 1 W/cm .

5. A thermocouple grid at the entrance to the stack along with mass

flow rate measurements with a thermocouple and pitot- static tube downstream

of the grid was set up to measure the heat convected through the doorway of

the model compartment. Subtracting this heat from the total heat produced

gives the heat lost through the walls, ceiling, and floor of the compartment.

This method should prove to be valuable in checking the energy balance when

an analytical model is developed to predict fire growth in a room.

6. It was verified that for hardboard and fiberboard the lateral flame

spread rate on a vertical surface could be expressed by the equation

F F

v„ = — ;r . The proportionality constant F is defined as the

(T

o

- T

s)'

characteristic flame spread constant for a material. The constant values of

F for fiberboard and hardboard are 55 and 13.3, respectively, when the flame

spread rate is expressed in meters per second. This indicates a flame spread

ratio of approximately 4 for these two materials at equivalent surface

temperatures.

7. The heat release rate of interior finish materials when considered

separately is not a suitable indicator for estimating flashover times in a room fire. It was shown that wall ignition time is related to the time to

37
ignition in the ease of ignition test and that the time from wall ignition to

flashover in a compartment lined with combustible walls is inversely

proportional to the product of the rate of heat release and the characteristic

flame spread constant F.

7 . RECOMMENDATIONS

1. Research should be continued in the further development of a bench

test for the rate of heat release and the effective heat of combustion using

the oxygen consumption technique along with a continuous rate of weight loss

measurement using a load cell. This technique is accurate, versatile, and is

less expensive than other methods. In addition to its predictive power for

room fires it could serve as an economical method for evaluating flame

retardant treatments.

2. A flame spread test should be further developed in which the

horizontal flame spread rate along a vertical surface is measured as a function of the temperature of the surface ahead of the flame. The data from

such a test along with the data from suitable laboratory tests for heat

release rate and time to ignition can potentially be used to predict the time

to flashover in a room.

3. The oxygen consumption method should be used as a standard technique

for measuring the rate of heat release in large-scale room fire tests for

interior finish and furnishing materials. The conventional method of

measuring the rate of weight loss and multiplying by the heat of combustion

has some difficulties. It is usually not practical to weigh the walls of a full-scale room, and, in the case of burning furniture, there are usually

several different materials burning with different and unknown heats of

combustion. Furthermore, the heat of combustion which is measured with the

oxygen bomb calorimeter can, at least for cellulosic fuels, be considerably

higher than the effective heat of combustion which should be applied during

large-scale fires.
4. Equipment utilizing the oxygen consumption concept for measuring the

rate of heat production from burning materials could be applied to a variety

of other tests, such as the flooring radiant panel, the fire endurance test

for wall and floor/ceiling assemblies involving combustible constructions,

and the ASTM E 84 tunnel test.

5. The calorimetric methods presently used to determine the heat release

rate of materials depend on the temperature rise of the stack gases. If the

flames from the specimens are highly luminous an appreciable part of the heat

may be radiated away before it reaches the measuring point in the stack. If

the walls are well insulated most of this heat will eventually be picked up

and measured but the shape of the heat release rate versus time curve will be

distorted. There has been no systematic study of the magnitude of this error

for interior finish materials. The oxygen consumption method of measuring

the heat release rate (which is not affected by heat losses) could be used to

compare the results of these two methods to examine the effect of heat losses

in the existing calorimeters.

8 . ACKNOWLEDGMENTS

The author is indebted to those at Armstrong Cork Company responsible

for making possible my year as a Research Associate at the National Bureau of

Standards where this research was conducted. My gratitude goes to Hank Roux,

Armstrong, for his guidance and to individuals of the Center for Fire Research

who expressed interest in this work by assisting, by supplying needed

instrumentation, and by the encouragement and guidance through helpful

discussions. In particular, I wish to thank Bill Parker who was most

instrumental in the success of this research effort. His expertise and

devotion to fire research were an inspiration. Critical to this research

effort was data reduction. For this I am greatly indebted to Jim Jackson,

Armstrong, who wrote the necessary programming and processed the data.

39
I would like to express my gratitude to the CFR readers, D. Gross,

W. Parker, B. Lee, C. Huggett, A. Robertson, and V. Babrauskas, who reviewed

this report and for their very useful comments.

9 . REFERENCES

[1] Standard method of test for surface burning characteristics of building

materials, ASTM E 84-79a, American Society for Testing and Materials,

Philadelphia, Pa. (1979).

[2] Fang, J. B. , Fire buildup in a room and the role of interior finish

materials, Nat. Bur. Stand. (U.S.), Tech. Note 879 (June 1975).

[3] 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,

135-151 (1972)

.

[4] Smith, E. E., Heat release rate of building materials — Ignition, heat

release, and noncombustibility of materials, ASTM STP 502, American

Society for Testing and Materials, 119-134 (1972)

.

[5] Parker, W. J., An investigation of the fire environment in the ASTM E 84

tunnel test, Nat. Bur. Stand. (U.S.), Tech. Note 945 (August 1977).

[6] Huggett, C, Oxygen Consumption Calorimetry, Proceedings Eastern Section

of the Combustion Institute, 1978 Fall Technical Meeting.

[7] Jessup, R. S. and Prosen, E. J., Heats of combustion and formation of

cellulose and nitrocellulose (cellulose nitrate), Nat. Bur. Stand. (U.S.),

J. 44, 387-393 (1950)

.

[8] Handbook of Chemistry and Physics, 57th Edition, CRC Press (1976-1977) .

[9] Loftus, J. J., Gross, D. and Robertson, A. F., Potential heat — A method

of measuring the heat release of materials in building fires, Proceedings

of the ASTM, Philadelphia, Pa., Vol. 61 (1961).

[10] Evans, D. D. and Breden, L. H., Time delay correction for heat release

rate data, Fire Technology, Vol. 14, No. 2, 85-96 (May 1978).

[11] Parker, W. J. and Lee, B. T. , Fire buildup in reduced size enclosures,

Nat. Bur. Stand. (U.S.), Special Publication 411 (November 1974).

[12] Waterman, T. E., Room flashover — Model studies, Fire Technology,

316-325 (November 1972)

.

[13] Lawson, J. R. and Parker, W. J., Development of an ease of ignition test

using flame impingement, Nat. Bur. Stand. (U.S.), NBSIR to be published.

[14] Carslaw, H. S, and Jaeger, J. C. , Conduction of Heat in Solids, Second

Edition, Oxford University Press, p. 75 (1959)

.

[15] Baumeister, T. , Editor, Mark's Standard for Mechanical Engineers, Seventh

Edition, 4-72 (1967)

.

[16] Fang, J. B. and Breese, J. N. , Fire development in residential basement

rooms, Nat. Bur. Stand. (U.S.), NBSIR to be published.

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66
Table 2. Comparison of effective net heats of combustion with actual gross

and net heat of combustion (MJ/kg)

Material Bench RHR test Model compartment Oxygen

Gross

bomb calorimeter

Net

Fiberboard 13.5 12.5 18.8 17.2

Particleboard 12.5 10.0 19.4 17.7

Hardboard 12.5 11.0 19.7 18.0

Table 3. Data summary for prediction of flashover

times in model compartment

Material

Fiberboard

Hardboard

Wall ignition

time, (s)

27

147

Flashover

time, (s)

216

534

kW/m"

60

130

K

2 -m/s

55

13.3

K • kW/m

6.3 x 10'

6.7 x 10

67

APPENDIX A

Technical Data

Thermophysical Properties of Materials Employed in This Experimental Study

Density,

kg/m 3

307

Specific

at 23 e kJ/ (kg

heat

C

K)

Thermal conductivity

at 23°C

kW/(m K)

Fiberboard 1.43 55 x 10" 6

Particleboard 674 1.97 119 x 10" 6

Hardboard 900 1.49 126 x 10" 6

Ignition Times Obtained on the NBS Test for Ease of Ignition

Compared With Time to Ignite Wall in Model Compartment

Ignition time, seconds

Ease of ignition test Model compartment

Fiberboard 20 27

Particleboard 96 108

Hardboard 103 147

Measured at Armstrong Cork Company

A-l

APPENDIX B

Scanning Technique Used to Determine Mean

Velocity Across the Exhaust Pipe Cross Section

A traverse of the pipe was made with the pitot-static tube, averaging

the readings taken at positions marked "+" in the pipe according to the

sketch. The factor, C , was obtained by dividing the average velocity by the

center velocity over a range of flow rates.

B-l

APPENDIX C

Crib Tests

An adjunct to this study included the burning of two cribs beneath the

large room collection hood. This was done in order to demonstrate the

applicability of this method of determining rate of heat release and effective

heat of combustion from other large fires. The cribs each weighing 3.4 kg

3 were constructed of polyurethane (320 kg/m ) and sugar pine. Placement of

the cribs on a load cell during burning permitted calculations for the rate

of weight loss to determine effective heats of combustion. The rate of heat

release as a function of time (RHR) was determined by oxygen consumption and

is plotted in figure 27. The broken lines labeled A, A , B, B , etc. are plotted at the mean RHR associated with the computed rate of weight loss (RWL) over the same period.

The heat of combustion (HOC) was determined during each of these time

intervals by dividing the RHR by the RWL. The increase in the HOC at latter

times during intervals D and D probably reflects the combustion of the char.

C-l
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C-2
APPENDIX D

Errors in Calculation of the Rate of Heat Release

In section 2.1 it was assumed that (1) the molar flow rate of the

combustion gases in the exhaust duct was equal to the molar flow rate of the

air into the system, (2) the oxygen concentration in the analyzer was equal

to that in the exhaust duct, (3) the ambient density of the stack gases was

equal to that of air, (4) the concentration of CO in the exhaust gases was negligible, and (5) the incoming air was dry.

The errors due to the assumptions in section 2.1 will be examined for

fuels of the form C H, which undergo the following complete reaction in air

to CO- and H_0:

0no C H.O + a + £ "^ n^ 0-> + n„ N + n n A + ft„ ~H-0

2 abc I 4 2/0-2 No 2 A H2 * 9A

2 aC°2 + ! -f) [v

6ft 2

l

H2° +

(a + I

" f)

[

(1 "

9>fto

2 °2 + A

N

2

N

2 + V + AH

2 H

2°] (14)

where Aq , L , 1, and n„ Q are the molar flow rates of oxygen, nitrogen,

argon, and water vapor into the system. The oxygen depletion, 0, is defined

as the fraction of the incoming oxygen molecules that are removed by

combustion.

Equation 2 can be more accurately written

Q = 1.67 x 10 4 X

Q Va

9 = 1.67 x 10 4 X

Q Vg

I ^fJ 6 (15)

Equation 14 can be used to determine i— and 9.

V

S

D-l
The ratio of the molar flow rate of the gases leaving the system, A , to

s the molar flow rate of the air entering the system A is found from equation 14

to be

*_s *B , X

Q

e (l + l

A .". 7 b c

a V

a

a +

4 "2

(16)

where X = y o ft + A„ + A. + A„ _. °2 . N

2

A H 2°

is the oxygen concentration of the ambient

air and Vc and V are the volume flow rates of the gas in the exhaust duct

and of the air entering the system respectively. These volume flow rates are referred to standard conditions of temperature and pressure.

The oxygen concentration X in the analyzer for the case where the water

is trapped but CO- is allowed to pass is also found from equation 14,

+ I - f )

(l " *) Xo

x

o

=

I1 - X

h) (' !-!)-• !-!)*o

(17)

where XTT =

H2°

H AT + A7I + A_ + A„

~

2 N2 A HjO

ambient air.

is the volume fraction of water vapor in the

Equation 17 can be solved for 9

e =

x

l - x.

l

- b c v

4 "2 c a + & - £

4 2

-1

(18)

Prior to the test 9=0 and X_ = X* =

c c 1 - X

because of the trapping of the

H

moisture in the ambient air. Hence

s^u-d -t) vi-!-f) v '

D-2

(19)
Combining equations 15, 16, and 19

1.67 x 10

4 X^ V c (e„ - e)/E,

O S V O / c Q

= P

; Mt*fV *!-! 'I- 1 v(»*!-f

(2o:

E - E X° - X

where —= = —=5—— is the fractional change in the voltage output of

o c the analyzer assuming it is linear.

From equation 20 it can be seen that the only effect of the ambient

humidity on the calculations is to reduce the value of X which is given by

the relation

X

q = 0.2095/(l + X

R

^ _1

(21)

where 0.2095 is the normal concentration of oxygen in dry air at sea level.

However, if X is to be determined by the oxygen analyzer it must be done

without the water being trapped. The volume fraction of water, X , is given

by

X

H = 0.0224 (D/18) yIq (22

where D is the density of saturated water vapor in air in g/m . It is given

as a function of temperature in the Handbook of Chemistry and Physics [8]

.

R is the relative humidity (RH) in percent. If X is taken to 0.208, its

value for 20°C and 40% RH, its variation under normal conditions is less than

1%.

The factors in the denominator of equation 20 account for the difference

in molar flow rates of the combustion gases in the stack and the incoming air

and the difference between the oxygen concentration in the analyzer and in the

stack. The numerator in equation 20 is essentially equation 2 which was used

for the calculations during this project. The errors incurred by this

procedure can be examined by writing equation 20 in terms of some common fuels

D-3
For cellulose, C^H, -0 C 6 10 5

1.67 x 10 4 X Vc /E - E

Q = i Z c B v 7§

S Mi? I (23) 1 + 56X76 I E

o \ o

For methane, CH.

1.67 x 10 4 X Ve /E - E o S / o (i + e X

Q/2) (1 - x

c/2)

For -CH_- polymers

Eo

(24)

1.67 x 10 4 X

Q Vg

/E - E

Q

=

71 j. fl v 7TN /I Z~v 7TV I w I (25) (i + e V3

) C 1 - V3

)" Eo

For cellulose the error is negligible at low oxygen depletions but

reaches 9% at a depletion of 0.5. For methane even at 9 = there is a 10.5%

error which reduces to a 0.3% error at 9 = 0.5. For -CH~- polymers the error

at 6 = is 7% and less than 0.2% at 9 = 0.5.

Other sources of error considered to be small were gases generated in

the fire and not converted to CO? and H_0. Also neglected was the

consideration of CO present in the stack and in the cell. Since the heat

release per volume of

2 consumed is substantially greater for CO than for

other fuels the presence of CO leads to another error in heat release rate

calculation. The heat of combustion of CO is 11.76 MJ/m at 20°C [15]. Then

3 the heat released per volume of 0_ consumed is 23.2 MJ/m at 25°C. The direct i 3 combustion of carbon to CO_ yields 16.7 MJ/m at 25°C. Then the heat produced

3

per volume of Q~ consumed in oxidizing the fuel to CO is 8.8 MJ/m (i.e.,

8.8 = 2 (16.0) -23.2) . After flashover in a room fire CO levels can reach as high as 11% in the exhaust gases at the doorway when the oxygen concentration

— has gone nearly to zero [16] . In this extreme case, approximately 25% of the

oxygen consumed has gone into CO production. Then the total rate of heat

3 release will be 13% less than that calculated based on 16.7 MJ/m for all of

the fuel being completely oxidized. Again a correction can be made if the CO

D-4
concentration is monitored in the stack gases. Also due to burning outside

of the doorway much of the CO will be converted to C02 prior to the gas

sampling point in the exhaust stack.

In calculating the volume flow rate in the stack, it was assumed that

the ambient density of the stack gases was that of air. In order to account

for the effect of the ambient density of the stack gases, p , on the

calculation of the volume flow rate using the pitot-static tube it is

necessary to replace p by p in equation 3 and rewrite in the form,

**-<A%r(k>>) <- 1/2 (26)

<©'Al/2

Then[—-j becomes the correction factor. Assuming dry air, the stack gas

density in g/mole, p*, deduced from equation 14 is given by

s

0.216 (44a + 9b) + a + £ - £ (29.0 - 6.729)

P* = r h

* (27)

3 0.219 a+| + a+|-| (1- 0.219)

For cellulose this becomes

r>* 174 + 346

p s 6 + 1.059 (28)

For 6=0 p* = p* = 29.0

s o For 8=1 p* - 29.5

s

/P*\l/2 /p \

For complete depletion of the oxygen the correction factor 1 rr ) =

\ o~

)

0.991 or a maximum error of 1% for cellulosic materials.

For methane

p* = 58 4- 3.366

p s 2 + 0.219 y '*'

D-5
(»

1/2

= 1.022 for 6 - 1

or a maximum error of 2%.

For -CH-- polymers

n * 43.5 + 2.946

P " 1.5 + 0.1050 (30)

(if = 1.001 for 6 = 1

or a maximum error of 0.1%.

Since the oxygen depletion factor, 9, is normally much less than unity

the effect of the chemical composition of the combustion gases on the ambient

density of the stack gases can be neglected.

D-6
NBS-114A (REV. 9-78)

U.S. DEPT. OF COMM. BIBLIOGRAPHIC DATA

SHEET

1. PUBLICATION OR REPORT NO. NBS Tech Note 1128

2.Gov*t Accession No. 3. Recipient's Accession Wo.

4. TITLE AND SUBTITLE

AN OXYGEN CONSUMPTION TECHNIQUE FOR DETERMINING THE

CONTRIBUTION OF INTERIOR WALL FINISHES TO ROOM FIRES

5. Publication Date July 19b\)

$. Performing"Organization Code

7. AUTHOR(S) 8. Performing Organ. Report No. Darryl L. Sensenig

9. PERFORMING ORGANIZATION NAME AND ADDRESS

NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE

WASHINGTON, DC 20234

10. Project/Task/Work Unit No.

11. Contract/Grant No.

12. SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (Street, City, state, zip)

Sponsored in part by: Armstrong Cork Company

Lancaster, Pennsylvania 17604

13. Type of Report & Period Covered Final

14. Sponsoring Agency Code

15. SUPPLEMENTARY NOTES

| | Document describes a computer program; SF-185, F I PS Software Summary, is attached.

16. 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 oxygen consumption technique was developed for determining the total rate of

heat production in a room fire. This was accomplished by measuring the volume flow

rate and the oxygen concentration of the exhaust gases flowing through a collection

hood. This method can be used with unknown combinations of burning materials including

both interior finish and furnishings. By simultaneously measuring the rate of oxygen

consumption and the rate of mass loss the effective heat of combustion of the wall

linings were determined in a reduced-scale model room fire test. The average heat

release rate per unit area of the wall linings was determined by recording the area of

involvement during the test and dividing this area into the total rate of heat

production at that time. The enthalpy of the exhaust gases passing out of the doorway

was determined with the aid of an array of thermocouples located at the entrance to the

exhaust duct. By subtracting the enthalpy flow through the doorway from the total rate

of heat production in the room, the heat losses through the bounding surfaces were

determined. Reduced-scale and full-scale room fire tests and a bench test for heat

release rate using the oxygen consumption technique are discussed in this report.

Lateral flame spread rates on vertical surfaces measured in the model room fire tests

and in a laboratory bench test are also described.

17. KEY WORDS (six to twelve entries; alphabetical order; capitalize only the first letter of the first key word unless a proper name;

separated by semicolons) Fire tests; flame spread; flashover; heat release rate; ignition; interior finish;

oxygen depletion; room fires.

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