NBSIR 80-2054

li The Burning of Wood

Cribs in an Enclosure

Volume I

and Plastic

James G. Quintiere, Bernard J. McCaffrey

Center for Fire Research

National Engineering Laboratory

National Bureau of Standards

U.S. Department of Commerce

Washington, DC 20234

November 1980

Final Report

, — Prepared for; QC oducts Research Committee

2QQ -ant No. RP-76-U-2: Assessing the Hazard of Cellular Plastic Fires in a Room

.U56

80-2054

1980

HATIOMAlt BURSAV or nAHBASM

UBHABT

APR 8 1981

THE BURNING OF WOOD AND PLASTIC

CRIBS IN AN ENCLOSURE:

VOLUME I

NBSIR 80-2054

no-L Clcc.-

OcfOo

U : no. : o f

James G. Quintiere, Bernard J. McCaffrey

Center for Fire Research

National Engineering Laboratory

National Bureau of Standards

U.S. Department of Commerce

Washington, DC 20234

November 1980

Final Report

Prepared for: Products Research Committee

Grant No. RP-76-U-2: Assessing the Hazard of Cellular Plastic Fires in a Room

*2

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

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

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director
.t«HtK7AW

MJUflMAnwi ‘‘iiiJ inihf>nj ? •< > '. -iJ

fS5f 8

» •'Al

'!_R

ifl''

ta I JkJKK -08 A»eev(

j. .'r^-

7''

j, ;

-

) ., .,. - ^ 01T3AJ*? anA.aoow ho oMivjjwia sht

;3HU20JaM^'ji^ Mtl afllflO -

"

I a MUJov

,. ,A

>7

*-—**-*r

•»

!_

ft

: w*j 'MrSk -

.V . ?y K"- ,fc?4 ?; . :.4 . 4.

** -ii m- .^A f.l

'?

P

ft..

* y' T ^**^.,.t’'/tW^;, •>C

l«r*OthlM

^

'. ' »; V'V

'

., '4KWwi^^^ ,'J

,?r

..‘'i’j -i^‘‘

; M.

1 j. 'l:^

- j

n«krP>fi *iw»<l t:

r< . er i

I

Ml,

.*’

,. ,

' **•

>v -ifl

n t! ....iyi'iii ^Y

iniij? V r-.i *ai<H 3i»wW ti»luU»3 'M^’ 9TM*«**^

»*

,

IKS'-

''

. .Hi. -V V

.!, ",?l » .t tywi^e J34«0uix!r« rjUtiW W •* '^

^Jr . *^J<kwA ^ '5iP 'j4A>i04

T

^

. ,' »*H

j ' rlV ...n #rt

.I’

* ; ^

’*n

«
TABLE OF CONTENTS

Page

LIST OF FIGURES V

LIST OF TABLES ix

TABLE OF NOMENCLATURE X

SUMMARY XV

Abstract 1

1. INTRODUCTION 2

2. EXPERIMENTAL DESIGN OF ENCLOSURE FIRES 4

2.1 Fire Test Room 4 2.2 Fuel 6

2.3 Instrumentation 10

2.3.1 Temperature 12

2.3.2 Heat Flux 12

2.3.3 Gas Concentrations 16

2.3.4 Mass Loss 16

2.3.5 Room Static Pressure 16

2.3.6 Doorway Velocity 17

2.3.7 Filter Paper Target 17

2.4 Data Recording 17

2.5 Experimental Procedure 19

2.6. Scope of the Experiments 22

3. RESULTS OF ROOM FIRE EXPERIMENTS 22

3.1 Observations 22

3.2 Average Peak Values 24

4. FREE BURN CRIB EXPERIMENTS 27

4.1 Instrumentation 30

4.1.1 Free Burn Basis 30

4.1.2 Oxygen Consumption 35

4.2 Results of Free Burns 35

5. ANALYSIS OF RESULTS 35

5.1 Rate of Mass Loss (m ) 41

5.2 Doorway Flow ...^ 45

5.3 Two Layer Concept 51

5.3.1 Neutral Plane (N) 51

5.3.2 Room Layer Height (D) 54

5.4 Room Static Pressure Differential (Ap) 58

5.5 Doorway Velocity and Flow Coefficient 63

5.6 Room Temperature (T) 67

5.7 Room Heat Flux 71

5.8 Combustion Product Composition 76

6. MATHEMATICAL MODEL 84

6.1 Conceptual Formulation 84

6.2 Steady-State Limitation 85

iii
TABLE OF CONTENTS (continued)

Page

6.3 Sub-Model Development and Verification 88

6.3.1 Effective Heat of Combustion and Air- Fuel Ratio .... 89

6.3.2 Fire Induced Flow 91

6.3. 2.1 Doorway Flows 93

6.3. 2.

Mixing Model 94

6. 3. 2.

3

Crib Fire Entrainment 102

6. 3. 2. 3.1 The Crib 102

6. 3. 2. 3. 1.1 Internal Channels . . 102

6. 3. 2. 3. 1.2 Exterior Surfaces of

Cribs 103

6. 3. 2. 3. 1.3 Contiguous Surfaces . 104

6. 3. 2. 3.

Fire Plume 10 5 6.3. 2.3.3 Results 107

6.3.3 Species Concentration . 109

6.3.4 Crib Fire Radiation 112

6.3.5 Room Heat Transfer 117

6. 3. 5.1 Upper Layer Emissivity 119

6. 3. 5.

Radiative Flux to Floor Target . . . .120

6. 3. 5.

3

Total Incident Heat Flux to a Ceiling

Target 125

6.3. 5.4 Conductive Ceiling Heat Flux 130

6.3.6 Crib Mass Loss Rate 131

6.4 Simulation Model 142

6.5 Comparison of Experimental and Theoretical Results 147

6.6 Empirical Results of Room Furniture Fires 164

7. CONCLUSIONS 167

8. ACKNOWLEDGEMENTS 168

9. REFERENCES 169

APPENDIX A - GASIFICATION RATE OF FREE BURNING CRIB FIRES 172

APPENDIX B - RADIATION EFFECT ON LOWER LAYER THERMOCOUPLES 177

Volume II, appendices A, B, and C are in microfiche form

in the back of this volume.

IV
LIST OF FIGURES

Page

Figure 2-1. Plan view of experimental facility 5

Figure 2-2. Plan view of test room 7

Figure 2-3. Crib configuration on the load platform for room fire

experiments 11

Figure 2-4. Room instrumentation - side view 14

Figure 2-5. Room instrumenation - thermocouple section with bead

diameter in mm 15

Figure 2-6. Experimental checklist 20

Figure 2-7. Reflecting globe for calibration of heat flux sensors . . 21

Figure 3-1. Cribs prior to ignition - PRC 7 25

Figure 3-2. Full involvement of cribs - PRC 7 25

Figure 3-3. Glowing char residue - PRC 7 25

Figure 3-4. Char residue of polyurethane cribs - PRC 14 25

Figure 3-5. Height of smoke layer interface above floor - PRC 7 ... 26

Figure 3-6. Smoke laden hot gases flowing from room into corridor - PRC 14 25

Figure 4-1. Basic layout of free burn instrumentation (dimensions

in m) 33

Figure 4-2. Configuration of cribs for free burn experiments .... 34

Figure 4-3. 0 ^ consumption apparatus (by Sensenig 19] ) 36

Figure 5-1. Symbol legend for figures in sections 5 and 6 4 0

Figure 5-2. Mass loss rate versus ventilation factor 42

Figure 5-3. Normalized mass loss rate versus ventilation factor ... 44

Figure 5-4. Room enhanced mass loss rate 46

Figure 5-5. Mass balance assessment of doorway flow rates 47

Figure 5-6. Airflow rate versus ventilation factor 49

Figure 5-7. Correlation for airflow rate 50

Figure 5-8a. Velocity neutral plane for PRC 8 52

Figure 5-8b. Pressure neutral plane for PRC 8 53

Figure 5-9. Doorway neutral plane versus upper layer temperature . . 55

Figure 5-10. Thermal and smoke layer heights 56

V
LIST OF FIGURES (continued)

Figure 5-11.

Figure 5-12.

Figure 5-13.

Figure 5-14.

Figure 5-15.

Figure 5-16 . Figure 5-17.

Figure 5-18.

Figure 5-19.

Figure 5-20.

Figure 5-21.

Figure 5-22.

Figure 5-23.

Figure 5-24.

Figure 5-25.

Figure 6-1.

Figure 6-2.

Figure 6-3a

.

Figure 6-3b.

Figure 6-4.

Figure 6-5.

Figure 6-6

.

Figure 6-7.

Figure 6-8.

Figure 6-9.

Page

Theirmal discontinuity versus upper layer gas temperature . 57

Room pressure difference at the ceiling 60

Average pressure distributions for each doorway size ... 61

Comparison of measured room pressure and calculated

hydrostatic pressure for PRC 8 62

Average normalized velocity distribution for each

doorway size 64

Local doorway flow coefficients 65

Doorway flow coefficients 68

Ceiling layer gas temperature versus fuel load 70

Correlation for upper layer gas temperature 72

Incident floor heat flux (K2) versus upper layer gas

temperature 73

A comparison of free burn to room floor flux 74

Ceiling incident heat flux versus upper layer gas

temperature 75

Oxygen and CO

2 concentrations with x = excess air (%) as a parameter 79

A consistency evaluation of the determination of excess,

air 81

Estimation of door CO concentrations 83

Conceptual model 86

Effective heat of combustion from room fire analysis ... 92

Comparison of calculated and measured doorway inflow

rates 95

Comparison of calculated and measured doorway outflow

rates 96

Comparison of calculated and measured doorway inflow

rates in kg/s C. =0.68 97

Entrainment concept 98

Analysis to determine and the mixing coefficient, . 100

Mixing correlation 101

Crib fire plume velocity and temperature correlations . . 106

Comparison of calculated and measured entrained flow by

the fire plume in the room experiments 108

VI
LIST OF FIGURES (continued)

Page

Figure 6-10. Conservation of species in upper and lower layers .... 110

Figure 6-11. Comparison of calculated and measured doorway CO

2 concentrations 113

Figure 6-12. Cylindrical flame model of Dayan and Tien [19] 115

Figure 6-13. Comparison of calculated and measured free burn radiative

incident target fluxes 118

Figure 6-14. Comparison of empirically calculated layer emissivity

with Modak's [22] method 121

Figure 6-15. Floor target flux by Modak and Matthews [23] analysis . . 122

Figure 6-16. Comparison of calculated and measured incident flux by

sensor H2 124

Figure 6-17. Derivation of radiative flux from the upper layer to the

ceiling element, dA 126

Figure 6-18. Ceiling convective heat flux analysis 127

Figure 6-19. Comparison of calculated and measured incident flux to

ceiling sensor Hi 129

Figure 6-20. Comparison of convective plus radiative heat flux into

the ceiling with conductive heat flux absorbed by the

ceiling 132

Figure 6-21. Correlation of free burn loss rate 135

Figure 6-22. Comparison of calculated and measured rate of mass loss

in the room fire experiments 141

Figure 6-23. Comparison of experimental and theoretical rates of

mass loss 148

Figure 6-24a. Comparison of experimental and theoretical airflow

rates - plastic fires 150

Figure 6-24b. Comparison of experimental and theoretical airflow

rates - wood fires 151

Figure 6-2 5a. Comparison of experimental and theoretical neutral

plane and thermal layer heights - plastic fires 152

Figure 6-2 5b. Comparison of experimental and theoretical neutral

plane and thermal layer heights - wood fires 153

Figure 6-2 6a. Comparison of experimental and theoretical upper layer

temperature - plastic fires 154

Figure 6-2 6b. Comparison of experimental and theoretical upper layer

temperature - wood fires 155

Figure 6-27a. Comparison of experimental and theoretical upper layer

oxygen concentration - plastic fires 156

vii
LIST OF FIGURES (continued)

Page

Figure 6-27b. Comparison of experimental and theoretical upper layer

oxygen concentration - wood fires 157

Figure 6-28. Comparison of experimental and theoretical lower layer

temperature 159

Figure 6-29a. Comparison of experimental and theoretical incident

floor flux - plastic fires 160

Figure 6-29b. Comparison of experimental and theoretical incident

floor flux - wood fires 161

Figure 6-30. Comparison of experimental and theoretical incident

floor heat flux as a function of wood and plastic rate

of mass loss 162

Figure 6-31. Calculated critical rates of mass loss for various size

room floor areas with 2.41 m high ceiling and a door 1.83 m

high and various doorway widths (also the input

parameters of table 6-11 apply) 163

Figure 6-32. Furniture fire distribution and flashover 165

Figure A-1. Correlation of crib burning rate data used for design . . 176

Figure B-1. Temperature correction versus incident radiant flux . . . 178

Figure B-2. Temperature correction versus thermocouple bead

diameter 17 9

Figure B-3. Comparison of gas temperatures measured by two

techniques 180

viii
LIST OF TABLES

Page

Table 2-1. Physical properties at 22®C 6

Table 2-2. Chemical and physical description of the fuel materials . . 8

Table 2-3. Geometry design of crib 10

Table 2-4. Type of instriomentation 13

Table 2-5. Correlation key for raw data 18

Table 2-6. Sequence of room fire experiments 23

Table 3-1. Average peak values for significant results 28

Table 4-1. Identification code for free burn instrumentation 31

Table 4-2. Change in probe positions for free burns with respect to

locations shown in figure 4-1 32

Table 4-3. Average peak values - free burn experiments 37

Table 4-4. Average peak values - O 2 consumption 39

Table 5-1. Effect of enclosure heat flux on filter paper target ... 77

Table 5-2. Gas concentrations for estimating room CO production ... 82

Table 6-1. Anticipated modeling parameters 87

Table 6-2. Air and O

2 - fuel ratio 90

Table 6-3. Effective heat of combustion results 91

Table 6-4. Crib channel flow 103

Table 6-5. Crib flow in exterior layer 103

Table 6-6. Flow between cribs 104

Table 6-7. Rate of air entrained over height of crib 104

Table 6-8. Absorption coefficient and flame temperature 116

Table 6-9. Pyrolysis due to radiation 137

Table 6-10. Evaluation of crib mass loss model 140

Table 6-11. Input parameters used in the theoretical calculations . . 146

Table A-1. Crib open burn results: rate of mass loss correlation . . 174

Table A-2. Estimated property values for crib material 175

XX
TABLE OF NOMENCLATURE

aA

A.

1

A

o ,

Jl o,S

0,T

o,u

A

W, Si w,u

A

V

V, s APV

b

b

B

c • V

w

D

e

E

E

f

f

F

1 ,

side dimension of ceiling quadrant, figures 6-15 and 6-17

crib exposed surface area, eq. 6-45

flame area (treated as a cylinder)

room floor area, WL

crib interior surface area, eq. 6-55

door area

door area below neutral plane

side projected area of crib assembly, eq. 6-51e

top projected area of crib assembly, eq. 6-57d

door area above neutral plane

room surface area below hot layer

room surface area above hot layer

2 cross-sectional area of crib vertical shafts, 4s

2 cross-sectional area of a single shaft, s average peak values corresponding to the burn period of 95% to

45% of crib initial mass

crib square stick thickness

side dimension of ceiling quadrant, figures 6-15 and 6-17

Spalding B-number, eq. 6-61b

specific heat at constant volume (gas-phase)

specific heat at constant pressure

specific heat of room "wall" material

constant parameter in eq. 6-50 and in eq. A-3

layer (thermal discontinuity) distance floor to hot layer

energy per unit mass of fluid

heat of combustion (appendix A)

rate of energy transport

fraction of total energy release radiated by flame

crib friction factor (appendix A)

radiation exchange shape factor between i and j

F^ - shape factor between cylindrical flame and elemental area

oriented normal to coordinate vectors i, j, and k respectively

x

I

I

1

i
TABLE OF NOMENCLATURE (continued)

g gravitational acceleration

=s'"o' number of side gaps for n^ cribs, eq. 6-57e

number of top gaps for n^ cribs, eq. 6-57d

h crib height (appendix A)

h convective heat transfer coefficient

H room height, or flame height including crib eq. 6-36

H

c crib height

- flame height measured from top of crib, eq. 6-44

H.

1

vertical distance between target element and layer, figure 6-15

«2

- vertical distance between target element and ceiling, figure 6-15

i enthalpy per unit mass

i/ j ,k - unit coordinate vectors

I Bessel function

1

M-l flame absorption coefficient

k

g

hot upper absorption coefficient

k

m

mixing coefficient, eq. 6-17 and 6-18

k

w

room "wall" material thermal conductivity

Z length of crib side

L room length, or distance between flame and target, figure 6-12

- main beam length, 4 (enclosed gas volume)

(surface area)

- Struve function

L

vap

effective heat of vaporization

A mass flow rate

mass flow rate per unit area

v,o

free burn pyrolysis rate per unit area

M molecular weight

M

c mass of crib assembly

n niimber of sticks per layer

n

c number of cribs

N number of stick layers, and neutral plane height

XI
TABLE OF NOMENCLATURE (continued)

P

P

s q

Q

Q*

r

r

rox

r

c

S

S(n^) -

t

t*

T

T

uu,v,w -

V

V

w

W

W( X

pressure

crib shaft perimeter (appendix A)

rate of heat transfer

rate of heat transfer per unit area (heat flux)

radiant heat flux

convective heat flux

conductive heat flux

rate of energy release by fire

dimensionless energy release, eq. 6-5 Oc

radial distance

stoichiometric air to fuel mass ratio

stoichiometric oxygen to fuel mass ratio

equivalent crib fire radius

control volume surface

number of crib sides facing the surroundings

time

nominal burning time, eq. 6-52b

temperature

adiabatic wall temperature

effective flame temperature

average of top three thermocouples in upper gas layer

internal energy per unit mass

direction cosiness

fluid velocity

control volume

control volume surface velocity

room width

doorway width

ratio of excess air after combustion to air required by

combustion

Y. mass concentration of specie i

1

xii
TABLE OF NOMENCLATURE (continued)

'1'

u

gc

"wall" conductance parameter, eq. 6-52b

^2 ~ cylindrical flame emissivity, eq. 6-35

mean flame intercept angle, figure 6-12

stoichiometric fuel to air ratio (appendix A)

mass production of specie i

w, Jl lower area ratio, f/A,

- upper area ratio, ^ normalizing velocity factor, eq. 6-20

" boundary- layer " factor, eq. 6-20

heat of vaporization (appendix A)

upper layer gas emissivity

emissivity of upper layer gas quadrant, eq. 6-59F

flame intercept angle, figure 6-12

radiation parameter, eq. 6-35d

stoichiometric coefficient

- 3.1416

density

Stefan- Boltzmann constant

dimensionless crib drag coefficient (appendix A)

mole fraction for specie i - 1/2 dimensionless crib pyrolysis rate, A^/Cb (appendix A)

ratio of char remaining to initial mass of fuel

SUBSCRIPTS

a - ambient air

b - burned within the enclosure

c - convective

c - crib

e - entrained

f - fuel

xiii
SUBSCRIPTS (continued)

f

F

g

i

in

k

£M

o

°2

0

r

R

s

S

T

uV

w

flame

room floor

gas-phase

specie

inflow

conductive

lower layer

mass transport

door opening

oxygen

outside surface of crib

radiative

energy release

heat flux sensor

side of crib

top of crib

upper layer

volatiles

room (wall) surfaces

xiv
SUMMARY

This study was motivated by the desire to achieve three objectives.

The first relates to the general objective of the Products Research Committee

(PRC) to provide a measure of the behavior of cellular plastics in burning

conditions related to real life. Specifically, the burning characteristics

of a structural (high density) polyurethane foam were compared with those of

wood. This comparison was made for uniformly ignited crib fires in a room.

For those experiments the relative fire risk of cellular structural plastics

in furniture items might be inferred.

The second objective was to experimentally determine the effects of fire

size, fuel type, and natural ventilation conditions on the resulting room

fire variables, such as temperature, heat flux, burning rate, and air flow

rate. This was accomplished by burning up to four cribs made of wood (sugar

pine) or polyurethane foam of nearly the same density as the wood. This

range of fire sizes was intended to simulate fires representative of small

furnishings to chairs of moderate size. In addition to variations in fuel

type and quantity, the room door width was varied.

The data from these experiments were analyzed in terms of quantities

averaged over the peak burning period defined as that time corresponding to

the interval from 5 to 55 percent crib mass loss. Thus, for each fuel, fire

size, and door opening, a set of "average peak values (APV) " could be defined

and recorded. These values tend to represent the most severe conditions

for that extent of fire involvement and door opening. By visualizing fire

growth as a succession of such "average peak values" together with tenability

criteria, a risk assessment can be made for various categories of fire hazard.

For example, flashover may be defined as having occurred when an incident

radiant flux of 2 W/cm is exceeded for a point on the floor. Through this

analysis, the data can yield the conditions for flashover in terms of fuel

type , amount , and doorway width

.

XV
The third objective was to develop a mathematical model to predict the

room fire conditions as a function of fuel type, amount, and room geometry.

The data collected were to serve as a basis for assessing the accuracy of the

model. The extent of this validation process would provide confidence limits

for model extrapolations to other conditions.

Experimental Results

The design of the experiments is discussed in section 2. A code was

used to label each run and its corresponding results. The code notation is:

12 3 4 PRC - where box 1 = experiment number

box 2 = W for wood or P for plastic crib

box 3 = number of cribs

box 4 = ventilation factor 1 = 0.79 m,

1/2 = 0.39 m, or

1/4 = 0.19 m, the width of the doorway

A complete set of results for the full period of each experiment is

tabulated in volume II along with graphical plots over time for many of the

measured variables. Also a set of the data are available on magnetic tape.

The average peak values for the room fire series and free burn experiments

are given in sections 3 and 4, respectively.

The average peak results were analyzed and are discussed in section 5.

Except for the plastic cribs burning at the 0.79 m doorway condition, the

mass loss rate (A^) was found to generally correlate as

A = function of A -4/ H (figure 5-3) v' o 0^0

for each fuel type.

where n

c is the number of cribs

A

0

is the area of the doorway

and H0

is the height of the doorway.

XVI
The 5/4-power dependence reflects the enhancement in pyrolysis due to

radiation from adjacent cribs. The ratio of crib mass loss rate in the

enclosure to its free burn value varies from 0.8 for small openings to 1.3

for large door openings (figure 5-4) . This displays the extent of the

enclosure radiative feedback and ventilation effects on the pyrolysis of the

cribs in this series of experiments. The enhancement by the room of incident

radiative heat flux to the floor ranged from 2 to 4 for the largest door

(0.79 m) (figure 5-21).

Two potentially useful correlations were developed but their generality

has not been determined. The air-flow rate (rti ) through the doorway was

Si found to behave as A

- V

(Ao^ o 1/2 ,in kg/s (figure 5-7)

where m^ is the pyrolysis rate in kg/s and is the ventilation factor

,5/2 in m

The temperature increase AT (°C) in the upper gas layer in the room due

to a fire energy release rate Q in kW for a room height H (m) and doorway

dimensions, H (m) and W was found to correlate as:

o o

AT = (5-19)

The first term in brackets was motivated by plume theory while the second

term in brackets represents the flow effect by the doorway. An alternative

dimensionless form of this correlation follows from an approximate form of

the equations in section 6.

T-T

a T

a

<x 1/2

where *b

AH

m

A

*w

is

is

is

is

the burning rate

an effective heat of combustion

the maximum possible doorway flow rate,

the heated internal surface area.

xvii
and h, is an effective heat conductance factor k V

ikpc

TTtj.* from equation 6-52.

The limit of applicability of this correlation has not yet been determined.

Doorway mass flow rates were determined from a limited number of velocity

and temperature measurements over a vertical centerline in the doorway. The

results yielded a mass balance to within + 20 percent for the largest opening and

agreed to within + 10 percent for the smaller doorways. A flow coefficient was

calculated for the inflow and outflow at the doorway and was found to be

and

where

C . = 0.68 out

C. = 0. 58 - 0. 31 W

in o 0.2m<W <0.8m. — o —

The position of flow reversal in the doorway, N^, (as measured from the floor)

was found to vary from N^/H^ = 0.5 to 0.7 for a corresponding mean upper gas

temperature T of 100®C to 800°C (figure 5-9)However, the elevation (N ) P

of zero pressure difference between the room and its ambience was always less

than or equal to the doorway neutral plane, (figures 5-8a and b) . This

suggested a downward sloping stagnation streamline from the doorway into the

room. The effective hot gas thermal layer (D^) was found to generally agree

with the height of the smoke layer (D^) (figure 5-10) . However, for large

fires with the smallest door, D was much less than and is indicative of

mixing between the upper and lower regions of the room. This phenomenon was

also confirmed by temperatures near the room floor as high as 87 °C (for

PRC-16-P-3-1/4) . Finally a hydrostatic pressure calculation, based on measured

room temperatures, reasonably predicted horizontal pressure differences between

the room and its ambience (figures 5-13 and 5-14) . This is significant because

these pressure differences are responsible for flow through room openings and a hydrostatic pressure assumption is used in current theoretical flow models.

Estimates made on measured chemical species, CO^ and 02/ indicated

good consistency between relevant measured variables and assumed chemical

reaction models for the sugar pine (CH^ 74^0 794 * ^ polyurethane

xviii
(PRC material GM-37, CH^ 27^0 0823^0 215^ going to complete combustion

(figures 5-23 and 5-24) . Also these calculations yielded a range of excess

air (i.e., the ratio of air flow rate into the room minus the rate of stoichiometric air required for combustion divided by the combustion air required) of

0.1 to 10. Although restricted ventilation effects for the smallest door

opening did reduce the fuel pyrolysis, there always appears to have been more

than sufficient air for complete combustion.

The relative risk of these fires can be considered with respect to their

flashover potential and their impact on the surroundings. The former aspect

was emphasized in this study, but some results can be applied to the latter

consideration. The thermal consumption of a filter-paper target sheet suspended in air (by either non-piloted ignition or thermal degradation under

radiant heating) was used as a criterion for flashover. Complete thermal

decomposition of the filter paper occurred under room conditions resulting

from the burning of three or more polyurethane cribs (i!i^ ^ 22 g/s) for all

door openings, and appeared to require more than four wood cribs (m^ > 27 g/s)

for complete destruction. If flashover by the remote ignition or pyrolysis

of other non-contiguous room furnishings is considered, then the filter paper

results suggest maximum tolerable burning rates for wood and plastic solid

fuels. A generalization of these results will be addressed later through

mathematical analysis. Measurements indicated that the filter paper was

completely consumed at an incident radiant flux of at least 1.77 W/cm .

An assessment of the relative risk of the surroundings with respect to

life safety might be considered based on an estimation of CO concentration

in the combustion product gases exiting from the room. A high of 0.8 percent CO

(by volxime) was estimated when three plastic cribs burned in a room with a

0.79 m door width (figure 5-25). CO was only directly measured in the

corridor adjoining the fire room for a limited number of experiments

(all at = 0.79 m) . In these cases, the CO/CO

2

ratios ranged from 0.70 to

0.95 for the polyurethane fires and 0.017 to 0.021 for the wood fires. These

ratios might be used to estimate CO levels for other conditions considered

in these experiments.
Theoretical Results

The basis of the mathematical model is a zone (control volume) analysis

applied to distinct homogeneous regions of the fire enclosure. The gas-phase

is divided into a lower layer (region) and an upper layer which includes the

fire plume. Conservation of mass, energy, and species equations are written

for each zone along with mass and energy equations governing processes at the

boundaries, i.e., wall, ceiling, and floor surfaces and the doorway. Transient

effects were not included in this analysis since it is reasonable to ignore

the time rate of change of mass and energy within a gas zone, and since the

model will address the average peak values derived from the experimental

results. Thus, a quasi-steady fire is being modeled. A weakness of this

approach is that the conduction into solid boundaries, a true transient, must

be approximated as constant. The inaccuracies attributable to this approximation limit the applicability of a steady-state model to a transient fire growth

process. However, it can easily be remedied by resorting to an accurate transient conduction calculation along with the other equations given in section 6.

This was not done in the present study since the emphasis is placed on "what

will happen" under given initial conditions of fuel burning in a room, as opposed to "when a critical event (such as flashover) will occur."

The equations are developed in section 6 and were solved numerically.

Complete convergence over all ranges of interest has not been achieved, but

the theoretical results do span the range of experimental conditions in this

study. Also in section 6, derived relationships for a particular phenomenon

are tested against the experimental results by using other experimental data

as input. This process provides some assessment of the accuracy of sub-model

equations used in the full fire growth simulation model. A summary of the

nature of the simulation model is given below. Its distinguishing features

include a lower layer that is allowed to be heated through a mixing process

between the upper and lower layer (see figures 6-1 and 6-10) . Oxygen

concentration in the lower layer can therefore be reduced, and this along

with radiation feedback from the enclosure affects the pyrolysis rate of the

XX
crib. Thus, a model for fuel pyrolysis in an enclosure contains the effect

of radiation for potential enhancement, and the counter-effect of reduced

oxygen. Indeed, results show that, as the rate of burning is increased and

the door opening is reduced, the mechanisms of radiant feedback and reduced

oxygen concentration due to mixing tend to cancel each other, resulting in

undramatic changes in crib mass loss rate. However, for other fuel

configurations these effects could be more pronounced.

The model can be reduced to seven nonlinear algebraic equations in

seven unknowns. Each equation constitutes a conservation law applied to a flow region (control volume) or interface between regions (boundary condition)

The input parameters, governing independent equations and dependent

relationships, are summarized below.

Model Summary

INPUT PARAMETERS

Physical properties:

g - gravitational acceleration

a - Stefan-Boltzmann constant

p ,T - reference (air) density, temperature

d. 3. c - specific heat (air)

Room Geometry

H, W, L - room height, width, length

H^, - doorway height, width

Crib Geometry:

n - number of cribs

c m" - free burn pyrolysis rate per unit area

v,o

£,b - stick length and thickness (square cross section)

n - number of sticks per layer

N - number of layers

Fuel Properties:

r - stoichiometric mass air to fuel ratio

AH - effective heat of combustion

xxi
f - fraction of energy radiated from flame

L - effective heat of vaporization

vap ^

0) - char fraction

V

kj - flame absorption coefficient

- effective flame temperature

M - mass of (single) crib

c

Model Parameters

C . , C . - in and out doorway flow coefficients

k^ - layer mixing coefficient (eq. 6-17)

C - heat transfer factor (eq. 6-50b)

x^ - excess air factor (eq. 6-41b)

a - wall conduction time factor (eq. 6-52)

h - lower layer surface heat transfer coefficient

O / ^

INDEPENDENT VARIABLES

N, D - hot upper layer height in doorway (neutral plane), in

room (thermal discontinuity)

Tg ^ , Tg

^ - upper, lower layer gas temperatures

T,, „ . - upper, lower layer surface temperatures W f u ^ W f - fuel pyrolysis rate

GOVERNING EQUATIONS

o

II

1

—

1

Conservation of mass - plume and upper layer

o

II

CM Conservation of mass - lower layer

f3=0: Conservation of energy - plume and upper layer

f^=0 = Conservation of energy - lower layer

o

II

in

iw Conservation of energy - upper solid/gas interface

o

II Conservation of energy - lower solid/gas interface

o

II Fuel pyrolysis model

DEPENDENT VARIABLES - Subroutines

A - air flow rate

a A^ - rate of entrainment between layers

e - upper layer gas emissivity

^ xxii
^

- radiant flux to upper surfaces

^

- radiant flux out of doorway

q^ ^

- radiant flux between upper and lower layers

q" - convective flux to upper surfaces

O f u

d" „ - radiant flux to lower surfaces ^r , a Yq ^

- oxygen mass concentration, upper layer

y_ I - oxygen mass concentration, lower layer

^

2

'

d" - incident radiant flux to crib sides

,s

d" - incident radiant flux to crib top ^r ,T ^

- incident radiant flux to floor target target ^

Results of the model are compared to experimental values in figures 6-23

through 6-30. The agreement may be considered fair to good. One problem in

making this type of comparison is that room gas properties do not behave as

upper and lower homogeneous zones as portrayed in the model. This fact is

underscored from measurements of lateral variations in gas temperatures between

the center of the room and doorway by as much as 25 percent (figures 6-26a and

b) . Finally, theoretical results are presented (figure 6-31) that predict the

minimum fuel mass loss required to achieve "flashover" (equivalent to an

2 incident floor flux of 2 W/cm ) as a function of sugar pine and rigid

polyurethane, room size, and doorway size. These results appear to be in

accord with results from furnishing item fires in rooms (figure 6-32).

Based on the theoretical and experimental results, an approximate

quantitative comparison of the relative risk of room flashover between these

two materials (sugar pine and structural polyurethane of comparable density)

is that roughly twice as much area involvement is required for wood over the

plastic. Although this conclusion may apply to other woods and plastics of

comparable properties as those examined in this study, it applies only to

non-spreading fires with a flashover criterion based on remote ignition.

xxiii
Jl

j I '^'V. < • • *(

§9r’ jE^/t**. -

Mt

/«-' !<U?% ~ Ip %9% X

i W"' ivj/oi T jcff'f *- — a

•I’ll » - »

•V*r' T-'jvc/ ^-7 xt-n

tsV;u^ l^iC/ w*4«

' _^r:|g :a

I ^ . -3 ‘jtf ’ n ’ /ffiT * t>* t ctfoblon t

- ' . ^ y.

'4R ^:-.n

i-.jT ' .’-ti jipf “ t

,,

.

.^- .• . . '*

, -•*> ;'# '»#r?y

4S.-0 -'•> iJJ .^^,^'s^div }W ^

; ,,r *'‘S M

‘ "^^v l^"v>^. . a; fr,^* .<. >,^ .>.. ifj* ^n'W«WWi»^^ rtpqoyli|j

^ ... . % '

I I iiiillMniiii' ri'tf't"t 111 in f . ftntf . i»adi; nu

/

' .‘H.

-,ii>^J •op

-<*£ !#<» ' V.fT ,.cf

-f» 1'<‘^ ?./.

ttarEM-imatiHiaiaw atoil t^;iop»'*^i>6fiaj^ * • :r

' .- 'aTI

•<^ ^irtjrioo<> IN«a woocfe %o "tap’Ufti ' ij

‘.^’'.*?>7f ?''a ''v

’ ;^.‘: - . '^Iri

Ufti »' fl j/'f!) '•ax^' ’liia-jiitnBWpi:*

"

,= ^ ,^’ IH/ '

- }y' -!*''*' ** , .'^'

*-‘jL nftLoJ -iiji e4S«*,

g

- • ' '

' Vx*'

.' ': Ji- ;^^' "

, ’(< 'tat» ml i'y%: -^91^ J-UXioq'^

,

''^ .^v '.W '•,.•' '

' .’lS' 'Tg

»^:JXw»aT( /«.^1W Mooo^-J

1

^ -'• 3

’ '*

oiUtAX K'7«^ « ,v-^ fix-*! U'-i-J ftfU fTO-'’t>^*3l _

. , >r V::r.i.^ -V

WWf’iHl * 1r- »‘VW !tO vXi|^^p«iU3p

(•y.* ' <tf‘I'

»v* *"i<i < *>'•'' .i . «./ ** «c>i44 t *4 '!

‘V. '

»

•k: ^ j ,-»A Of v|^'.f«. rfp»0£i4i,A

; ‘^^•^‘0^

* '

. .

^.>*1 -fSft f(,t .r"I>'- ':# «0‘

<• ,:, • 7<H«i...‘i 'wo4\r.*‘a^.37?' A , rt4iv p<iib*#uq‘ -‘toii!^

r-'%mrn> • ^#^, . .M

r
THE BURNING OF WOOD AND PLASTIC CRIBS IN AN ENCLOSURE:

VOLUME I James G. Quintiere and Bernard J. McCaffrey

Abstract

This study was designed to assess the fire hazard of

a cellular plastic material which has comparable structural

characteristics to wood. The study attempts to determine

the relative fire risk of such materials in furniture.

Rigid (high density) structural polyurethane foam and sugar

pine were selected for fuels and burned in the form of cribs

in a room. The crib loading and door width were parameters

experimentally varied. Twenty-one room fire experiments and

eight free burn experiments were conducted. Measurements

of mass loss, temperature, heat flux, CO

2

and O2 concentration were recorded. These data were analyzed and empirical

correlations were developed for air flow rate and upper gas

temperature. A theoretical fire simulation model was

developed and yielded results in fair to good agreement

with the data. An extrapolation with the theoretical model

was used to predict the critical (or minimum) fuel pyrolysis

rate to cause flashover (as implied by 2 W/cm of incident

radiation to the room floor) . This was done for various

size rooms and door openings. It appears that to cause

flashover, for a given room and door size, about twice as much wood must be involved in fire as the rigid polyurethane

material

.

Key words: Crib fires; experiments; mathematical model;

plastic; polyurethane; room fires; wood.

1
1 . INTRODUCTION

Understanding the developing fire in a room is a major objective of

fire research. The use of new materials in room furnishings and building

construction raises questions concerning their potential role in fire spread.

It is not a simple matter to assess the potential fire hazard of a room and

its contents. The growth of a fire within a room will depend on the nature

of its combustible contents, their arrangement, and the construction and

geometry of the room and its openings. Moreover, the concept of "fire hazard"

is subjective and multi-faceted. For example, the obscuration due to smoke,

the toxic effects of the combustion products, the rate of fire spread, and

their combination can all be indicators of hazard. In this study the concept

of hazard will be limited to the potential for fire growth.

In recent years mathematical models have been used to predict the dynamics

of a developing room fire [1-6]^. The basis of these models is an idealized

room fire concept which consists of a fire plume and upper and lower homogeneous gas zones (layers) . A rate of energy release expression is formulated

for the fire and the conservation equations are solved for the zones. In

reality, the term flashover is used by observers of fire to describe the

phenomena of rapid transition in a developing room fire from a small fire to

full room involvement. Although the occurrence of flashover is usually apparent

its cause may be attributed to one of several mechanisms. Moreover, these - , mechanisms have not been fully understood. Mathematical models currently I assess this transition by evaluating the rate of fire spread, the ignition rjj

of objects remote from the primary fire, or the excess of a room temperature

or heat flux. Thus, models can be very specific in adopting operational

definitions of flashover, but they may not be complete. The application of

modeling results to the prediction of flashover will be applied in this

context

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

of this report.

2
In this study the effects of flame spread over objects and combustible

room linings will not be considered. The simpler case of a discrete fire centrally located and burning at a nearly steady rate will be considered. This

relates to the real case of a furniture fire in a room with noncombustible

linings. Conditions at the peak fuel gasification rate will be determined

from the experimental results, and the potential for room flashover will be

evaluated based on the potential for ignition of a floor target. This

evaluation could lead to the determination of a critical (or minimum) fuel

mass loss rate that is just sufficient to cause remote ignition (or flashover)

in a room of given geometry and construction. Moreover, the results would apply

to the developing fire in which the increase in the rate of fuel gasification

is relatively slow. This is usually the case for furniture items before the

onset of flashover. Hence, this process yields a result that can determine

whether flashover will occur, but not when it will occur. Consequently, it

will be sufficient to use a steady-state mathematical model for this application.

In general, for slowly developing room fires in residential-size rooms, the

transient fluid dynamic effects are insignificant, but the solid conductive

heat transfer transient effects could be important. This latter effect will be

accounted for by approximately evaluating heat transfer into solid boundaries

at a time characteristic of the peak fire size; the time history of the solid

will not be computed. Ultimately, the intent is to use the mathematical model

to extrapolate to conditions beyond the range of the experiment. The validity

of this extrapolation can be judged from the accuracy of the theoretical results

relative to the experimental data. This summarizes the intent and nature of

this study.

In particular, the objectives of the investigation are listed below;

(1) To conduct room fire experiments in which fuel representative of a discrete furniture item will be burned; (In order to investigate the

performance of cellular plastics, a structural cellular plastic with

a density comparable to wood, and a wood material are selected as fuels. A series of experiments is conducted in which the quantity

3
and type of fuel are varied under natural ventilation conditions of

various doorway widths.)

(2) To determine and evaluate the peak average burning condition with

respect to a potential for remote ignition of a floor target;

(3) To develop mathematical models of relevant phenomena within the

framework of a two-layer room fire model, and attempt to verify

these models (or sub-models) by comparison with experimental results;

(4) To predict results using a steady-state two-layer mathematical room

fire model, and compare these predictions with the experimental

results; and

(5) To make a quantitative statement on the factors influencing flashover

for discrete room fires.

2. EXPERIMENTAL DESIGN OF ENCLOSURE FIRES

In this section the experimental facility will be described. The process

selecting the fuel configuration will be discussed. The type of instrumentatio

and measurements will be presented. And finally the experimental procedure

and scope of experiments will be listed.

2.1 Fire Test Room

The experiments were conducted in the NBS Fire Test Building (205) in

a test room which adjoined a corridor. A plan view of the facility is shown

in figure 2-1. The combustion products from the room fire would exit at the

open end of the corridor into an exhaust hood. To avoid the influence of the

corridor on burn room flow, the height of the room doorway was selected such

that the layer of hot combustion products in the corridor did not fall below

the top of the doorway. The doorway of the room was located at a corner and

had a channel-like door frame with a depth of 0.5 m approximately. These were

facility constraints and could not be easily modified. The width of the room

doorv/ay was varied over the series of experiments in three sizes: 0.79, 0.39,

and 0.19 + 0.01 m.

4
Not

to

scale

I

5

Figure

2-1.

Plan

view

of

experimental

facility
A plan view of the room is shown in figure 2-2. Its dimensions of

2.18 X 2.18 X 2.41 m high are relatively small by residential standards but not

uncommon for a minor bedroom or study in new housing. The room was lined with

2 a calcium silicate board ( Johns-Manville , Marinite XL ) with a thickness of

1.9 cm. This material has physical properties similar to gypsum board as shown in table 2-1. Also it should respond thermally as a semi-inf initely thick

wall over the anticipated burning duration of the experiment (10 minutes)

.

Table 2-1. Physical properties at 22®C

Density Thermal Specific Moisture

Conductivity Heat Content

(kg/m-^) (W/m-K) (kJ/kg-K) (%)

Marinite XL 700 0.11-0.14 1.12 <5

Gypsum board 960 0.17 1.1 1 <5

During the period of the experiments the ambient laboratory conditions '

were roughly 22 °C and 30-40 percent relative humidity.

2.2 Fuel

A wood and rigid polyurethane foam material v/ere selected. The source

of the polyurethane was the Products Research Committee (PRC) materials bank

and is GM-37 by their designation. Often throughout the text, the materials

will be referred to commonly as "wood" and "plastic". Both leave a char residue

when burned. A description of fuel materials is given in table 2-2.

2 Certain commercial equipment, instruments, or materials are identified ....in

this paper in order to adequately specify the experimental procedure. In no case does such identification imply recommendation or endorsement by the

National Bureau of Standards, nor does it imply that the material or

equipment identified is necessarily the best available for the purpose.

6
i

Figure 2-2 Plan view of test room
Table 2-2. Chemical and physical description of the fuel materials

' Density

kg/m^

Moisture

Content

Q.6

Element

C H

Q. Q. "O "O

composition (dry)

0 N

Q. Q. o '6

i Wood: Sugar pine,

clear 370 6.5 47.6 6.4 44.7 0.07

1

1 Plastic : Rigid

1 polyurethane,

GM-37

1 340 nil 66.5 7.0 19.0 6.4

A more difficult task than selecting the fuel materials was to select

their configuration. Perhaps the simplest fuel configuration is a horizontal

slab covering a portion of the room floor. This configuration burning in an

enclosure can sense a significant amount of energy feedback from the heated

enclosure. As a result of this feedback, its burning rate can be greatly

enhanced. Such a configuration is a relatively "simple" case to model. Unfortunately, both V700d and high density rigid polyurethane do not sustain combustion

in this configuration. Moreover, vertical slabs of these materials are also

difficult to burn unless some preheating or long duration ignition source is

applied. Consequently, the configuration of a "crib" (ordered array of sticks)

was selected.

The design of the crib configuration was subject to the following

specifications

;

(1) The burning rates should be over a relevant range. It was thought

that 5 to 20 g/s would be satisfactory. This represents small

wastebasket fires to moderate size chair fires. I

(2) The duration of the fires should be sufficiently long to provide

a realistic simulation of furniture-like fires, and should be

reasonably steady over several m.inutes of burning. A nominal fire

of 10 minutes was considered acceptable.

(3) The crib fires should be surface controlled and sufficiently open

in porosity to permit some room energy feedback.

(4) The ratio of the projected horizontal cross-sectional area of the

fuel to floor area should be small. Also the height of the fuel

8
element should be, if possible, below the thermal (or smoke) layer

in the room. These constraints should insure unrestricted ambient

air flow to the fuel.

The theory of free burning wood crib fires is well developed [7,8] and

was used to design the crib configuration subject to these specifications.

However, empirical parameters needed to be determined in order to apply the

theory polyurethane. A series of preliminary experiments burning wood and

plastic cribs of various sizes and arrangements was conducted. The theoretical and experimental results that led to a wood and plastic crib design are

given in appendix A. Insufficient data and uncertainty of empirical parameters

for polyurethane cribs do not fully insure that our design for polyurethane

is in the stick surface area-controlled burning regime. Also during burning the

polyurethane sticks swelled up to 50 percent of their original thickness, and

this phenomenon affects the certainty of a surface area-controlled crib design.

Observations showed that the plastic cribs would fissure and cause jets

of flame in spots whereas for wood the velocity of the volatiles leaving the

surface appears more uniform. For both materials there appears to be a distinctly greater fuel gasification rate per unit area for internal stick

surfaces than for outside surfaces of the crib. The results from appendix A

yield an average peak gasification rate per unit of exposed area, m^, for

the area-controlled regime as

ra" = C b"^/^ (2-1) V

1 5 where b is the thickness of a square stick in cm and C is 1.2 mg/cm ’ -s for

1 5 polyurethane and 0.88 mg/cm ' -s for sugar pine.

For our limited data over a range of crib sizes, equation (2-1) predicts

the mass loss rate to within + 20 percent for sugar pine and + 20 percent for

polyurethane. It should be emphasized that equation (2-1) corresponds to the

average peak gasification rate^, and that the reproducibility for our crib data

is about + 20 percent.

^The average peak gasification rate is defined in section 3, and is generally

the "best" slope of the fuel mass loss curve over the most active flaming period.

9
Based on these results, a crib was designed and it is presented in

table 2-3.

Table 2-3. Geometry design of crib

1

'

Square stick thickness b 3.5 cm

1

! Stick length 1 24.5 cm 1 i Spacing between sticks s 7.0 cm i Number of sticks per layer n 3

1

'

!

'

Number of layers N 10

, This design, for the wood and plastic materials, has a nominal peak average |

: mass loss rate of 5 g/s, a nominal burning duration of 12 minutes, and a |

'j duration during the peak burning period of 200-400 s. In order to vary the |

^ fire size, the cribs were burned in arrangements of one to four cribs. The ’

cribs were arranged on a platform, 0.18 m above the floor of the room, as : shown in figure 2-3. Heat flux sensor K2 is included for orientation. The :

combined height of the platform and crib was slightly under 30 percent of the

doorway height, and the area of the platform was about 10 percent of the room

floor area. Thus, the last specification on crib design appears to have been met

2.3 Instrumentation

The instrumentation was designed to characterize the thermal and flow

aspects of the compartment fire. The number of probes was chosen to avoid

an excessive amount of data to analyze and yet give a sufficient set of data

to, hopefully, fully characterize the phenomena. A vertical temperature profile was measured in the room outside the flame zone. Wall and ceiling surface

temperatures were measured, as well as the surface temperature of a filter

paper target at the floor level. The rate of fuel gasification was recorded.

10
n

nn

—

1

1

1

1

nH2

H2

Figure 2-3. Crib configuration on the load platform for room fire

experiments

11
along with the oxygen and carbon dioxide concentration of the combustion

product flow. Velocities and temperatures were measured in the doorway so that a mass flow rate could be determined. Total and radiative incident heat

fluxes were measured at significant locations in the room. (However, all of

our radiometer data is suspected to be invalid due to soot fouling of the

radiometer window. ) Also static pressure differences were measured between

the room and the surrounding ambience.

The type of instrumentation is described in table 2-4. Also given is a symbol and letter key that has been used to designate each sensor. The sensor

locations are shown in figures 2-2, 2-4, and 2-5.

2.3.1 Temperature

In figure 2-5 the number associated with the thermocouples gives the

bead diameter in mm. This set of thermocouples was installed after test PRC 8 and they were used through PRC 20 except for some individual replacement. Also

after PRC 8 temperature sensors T26-T30 were added in order to determine a "true",

gas temperature in the lower layer. The smaller bead thermocouples and the

aspirated probe would be less affected by radiant heating and thus yield a more

correct gas temperature. An assessment of radiation effects is given in

appendix B.

The surface monitoring thermocouples were held in place by their own

spring force, and half of their bead was pressed into the surface. The leads

from the bead were in contact with the surface for a short distance to minimize

wire conduction losses.

2.3.2 Heat Flux

The primary results of interest are the total incident heat flux measurement

at the ceiling (Hi) and at the floor (H2) . The floor sensor primarily represents the sum of the radiative heat flux from the flame and the enclosure, '

plus a small convective component. This is essentially the flux seen by the i

lower space thermocouples and the filter paper target. Unfortunately, the

12
00a

u

a

o

•H

e

03 •H

'O

cd

0)

0)oua

a

(d^9

0) M

4J <U

I *0

4) 4) •H S •H

CLO 4) S U

1-1 Cd

41 '4-1

Cd>

4)&4J ^

i a

iH 4) •H ^ CL 4J O *0 a^4)

T3

to

cd 40

cd iH

CL 'V

Cdo

»-» /-N aiH

4)

4) 4J

cj Cd C 2 cd 4J *» U 4) P u

fH CS

4) 41 M *HO

4) (0 tH Cd s ^ •H Q

a

>» H

1 CS

Cd H

<<:in 00^0 (

13

Velocity

pressure

P

(J

Pl-7

Bidirectional

velocity

probe,

capacitance

type

pressure

transducer

and

electronic

manometer

(Datametrics)
2.41 I

2.41

0> •H>0) •a•H

U)co

•H

+J

(0

4-> C

0)£

4J

(0C

•H£oo

'a*I CN

0)uD

t7> •H

Cm

14
0.30m

T25

0.51

Rl HI A PFILTER PAPER

T8

0.50

AT9

0.61

ATIO

0.62

Atii

0.60

ATI2

0.60

ATI3

0.43

ATI4

0.56

ATI5

0.50

ATI6

0.47

ATI7

0.54

AT2I

0.55

AT22

0.44

T23,T26,T27,T28

0.49,0.18,0.28, ASPIRATED

Figure 2-5. Room instrumentation - thermocouple section with bead

diameter in mm

15
windows on the radiometers could not be kept clean during a room fire

experiment so radiative and convective flux components could not be reliably

determined

.

2.3.3

Gas Concentrations

Oxygen and CO

2 volume concentrations were measured on a dry basis

(H

2

O was condensed out in the sampling line) in the exhaust products flowing

out of the doorway. Both measurements give information on the nature of the

combustion process. Also CO

2 concentration is needed to determine a gas

emissivity, and the O

2 concentration is required in order to estimate the rate

of energy release within the room.

2.3.4

Mass Loss

A 45 kg capacity strain gage load platform, insensitive to eccentric

loads, was used to monitor the mass loss of the burning cribs. The platform

was contained in a stainless steel box surrounded by water-cooled fin tubing

as shown in figure 2-4. The top cover of the box was insulated and was

supported by the load platform. Sufficient clearance existed between the

top and bottom sides of the box. The accuracy of the system and fluctuations

under a static load were +50 g approximately; this suggests an error of up to

+ 3 percent during the peak average period in these experiments.

2.3.5

Room Static Pressure

Six pressure taps were located vertically in the corner of the room as shown in figures 2-2 and 2-4. After PRC 8 their positions were changed to a more optimum configuration for displaying the pressure distribution. These

pressure taps measured the static differential pressure across the wall of

the room. In general, it is this pressure difference which drives the flow

through any room openings during a fire. An assumption in developing a theory for fire induced flows is that the pressure distribution is hydrostatic

in a room. These measurements, along with the room temperature distribution,

would allow a check on this assumption.
2.3.6 Doorway Velocity

The velocity in the doorway was measured with seven bidirectional probes

as shown in figure 2-4. Their pressure differentials were sensed by seven

transducers whose output was sequenced by a multiplexer through a single

electronic manometer. Each transducer signal was integrated over 1 s intervals

(after PRC 8) and recorded every seventh data scan or every 35 seconds. This

arrangement proved to be satisfactory in determining the doorway mass flow

rate over the peak burning period.

2.3.7 Filter Paper Target

In order to assess the hazard of remote ignition of other objects, a

2 filter paper target was used. A 15 cm diameter sheet of Whatman Qualitative

No. 1 filter paper was suspended over the open face of a metal can. This was

located on the floor of the room as shown in figures 2-2 and 2-5. The paper

was not preconditioned, and a thermocouple (T24) monitored its tem.perature

.

The position of the target in the room did not enable us to view it during an

experiment, but the condition of the paper following the test v;as recorded.

2.4 Data Recording

Up to fifty channels of data were recorded by a high-speed data acquisition

system. The set of channels was scanned every 5 seconds at a maximum rate of

40 channels per second. The raw data were processed by a computer to yield

values in appropriate units for each channel number. A key for interpreting

the raw data (as given in volume II) is given in table 2-5. It should be

noted that channel 30 identifies the probe position corresponding to the

velocity recorded in channel 35. The number in channel 30 corresponds approximately to 10 times n where n is the digit on the bidirectional probe coding

label Pn. (See figure 2-4 or table 2-5)

.

Continuous color video recordings V7ere made for each experiment. The

television camera viewed through the lower half of the room doorvzay. The

visible portions of the room walls were marked by labeled horizontal lines

17
defining the height. This enabled the smoke layer height to be determined by

a study of the video tape for each test.

Table 2-5. Correlation key for raw data

Monitored Channel Units

Instrumentation Variable Number

LI Load Cell Platform Burning mass loss 31 g

H2 Floor Heat Flux Meter Total heat flux incident

on floor

32 W/cm^

T24 Thermocouple (TC) on Incidents ignition of 23 °C

Target Filter Paper target filter paper

T8- Room TC Rake 16 TC '

s

Representative room gas 7-22 °C

T23 including Ceiling gradient

T25 Wall TC Representative upper wall

temperature

24 °C

Tl- Doorway TC Rake Gas temperature distribution 0-6 °C

TV 7 TC ' s in doorway

PI- Bidirectional Velocity Gas outflow and inflow 35

P7 Probes (7) velocity (30 , probe ID) m/s

Gl, Doorway Gas Sampling 0„ and C0„ concentration 40

G2 Z 2. of exhaust gas

and

47

o/o

DPI- Static Pressure Room static pressure distri- 41-46 N/m^

DP 6 bution (differential-room

to ambient)

R1 Ceiling Radiant Energy Incident radiant heat flux 33 W/cm^

Meter to ceiling

HI Ceiling Heat Flux Meter Total incident heat flux

to ceiling

34 W/cm^

R2 Radiometer Heat flux from upper region

of room located 0.95 m above

i floor^

39 W/cm^

T26- Small Wire Thermocouples

\

T30 An Aspirated Thermocouple (T28)

Room gas temperature 25-29 °C

Added for PRC 13 - PRC 18

Added for PRC 9 - PRC 20

18
2.5 Experimental Procedure

Before each experiment the test room was cleaned of excess soot deposits,

and cracks in the lining were patched with a furnace cement. The probes were

cleaned and their positions checked and realigned if necessary. An instrumentation checklist (figure 2-6) was followed and calibrations of the load

platform and heat flux sensors were performed.

The heat flux sensors were calibrated in place by a reflecting sphere

apparatus shown in figure 2-7. A reference sensor was mounted flush to the

inside of the sphere. On the side opposite, a collar accommodated the sensor

to be checked so that its face was also flush with the inside of the sphere.

The heat source was a lamp mounted at the center of the sphere. The flux

was varied by varying the lamp voltage. A calibration constant had to be

applied to relate the test sensor to the standard reference sensor. This

factor depended on the orientation of the system and the radiometer window.

This procedure provided a cruick and convenient check of the heat flux sensors.

If a sensor deviated from its initial calibration it would then be removed and

recalibrated under laboratory conditions.

After the instrumentation was prepared, then the preparation for ignition

commenced. The wood cribs were kept in a conditioning room (20°C, 50% RH) for

several days prior to a test. Their moisture content ranged from 5 to 8 percent.

The cribs were arranged on the load platform (figure 2-3) and the spacing

between cribs was equal to the stick spacing, 7 cm. Each crib was placed on

a 30 cm square metal pan with a 1 cm high lip and was supported by two

Marinite strips 2 cm high. Forty ml of heptane was poured into each pan to

promote rapid and uniform ignition of the cribs. Ignition of all the pans

was virtually simultaneous. On ignition the data and video systems began

recording.

19
ODDQ

Pk

oE-*

>* I S§ ^

0) 0) o £

!

!CO ^ s

H O

<U -P

Ft <1) H

g s

0) r—

1

3 'O S 0)

CO Ft '3 Ft c o> a O

O -P

•H nH

T3 1

«

^ &

,o »o

0>

0) o

-P o

cd Fh

Fh O.

•H CO iH »H

CO a O -P

S'

S’

•H »H

CO Fh

CO ^

0) TJ

Ftp.g

^ CO

r-l C ' 0)

+>

to •H

(U x:o

iH

td +JC

0)eH

U

(Da

X

VO

I <No

3

C7> •Ha

20
7cm R

REFLECTING

SPHERE

pWx

QcS,->(r ujS^o

QujH<0

_ILU<Z

OLlIUJUJ

xtrxco

cruj <co

ds

OUJ

OlnomC

0)aX3

rH

4Jd

0) x:oco

•H

4J

COu

XI•H

cdouo

M-l

0) XO

rH

tJi C

•H

4JOO

rH

C4H0) caI CH

a)

}H3

O' •H

(X4

21
2.6 Scope of the Experiments

The scope of experiments was designed to range over fire sizes

representative of discret^ furniture fires, the biggest comparable to a chair fire. The ventilation conditions were to reflect typical door openings

(although our variable width doorway does not exactly simulate a door in various

open positions) including doorways small enough to be possibly ventilation

limited. A range of door sizes and crib loadings was selected in order to

span these conditions, and to provide a basis for comparison with theoretical

results and their extrapolation to other realistic conditions.

The sequence of experiments is given in table 2-6. The second through

fourth columns provide a descriptive code used for each experiment.

Test Number - Fuel Type - Number of Cribs - Door Factor

The code will be used throughout the text, and as a label on the figures

for identification purposes. Also given are some comments for each experiment

and the condition of the filter paper target following the test.

3. RESULTS OF ROOM FIRE EXPERIMENTS

In this section some observations made during the room fire experiments

will be presented along with the average peak burning data. A complete set

of the transient raw data, in the format given in table 2-5, is available on

magnetic tape, and a printed listing of the data is tabulated in volume II j of this report. Also included in volume II, are numerous computer generated

plots which display the behavior of each measured variable over time. |

3.1 Observations

Typical of room fire experiments, an immense amount of data was collected.,

By design, we have chosen to examine, for the most part, the quasi-steady i results during the peak burning period. This minimizes an analysis for each

experiment, and directs our analysis to the effects of the initial conditions '

on the "peak" fire conditions. Some observations made over the period of each

22
Table

2-6.

Sequence

of

room

fire

experiments

o

0)

d

oH

-P•H6,

•H

Fho

&o

I—

CO

cd

Fho

bo

Cl •H

rH•H

<UO

oa

tj

<uo

cd

I—

ft

0)

Fhbo

(h •H

iH•H

(1) O

o8

Td

CD -P

FhO

.C? CO rH cn

Cd CV & ft

oO E E E E O CD CD CD CD

ITN iH cH 1

1 rH

1 ,Q O O O O • U U Fh p

CD

Ph Ph Pi Pi

rH rH 1—

1—

-P rH rH 1—

rH bO

rH

0) a> CD P

O a O O

•H 'a O) Td rH Cd cd Cd Cd •H O o o O

<D rH iH 1—

rH O

<D 0) CD rH rH 1—

rH 1—

Cd rO rQ rQ •H •H •H •H •H

-P CO CO CO CO •H CO CO CO CO a O O O O M (H Ph Oh

ft \0

o

ft

fto

-p

cd

<uft

0) ft

o

ft

fto

p>

cd

0)ft

CD ft

-P0 Td, Fh h^ 0) Td CD

CD Fh CD ft CD P CD bOft Fh —I ft Fh CD Fh H cd Cd cd ft ft ft O ft

p! o o

1 I I O rH cSl

O O O O O O C\i C\i C\J C\i CM

e E

ft

p>•H^ Fh O O C7^ O I>- 0^^ ^ H-^ Cd o o O o

^ cd

Fh ft II II II o

O I—I Hloj h|4 Q

rH rH rH rH rH rH rlw hKh hKm rt|<> h(4 H<f nlm ny H|ni H|<f H<l

C^rHCMCMC^-s(--NtCMCMC^rH CMCM-sJ-C^-^C^^C^

CD

CDC

ft ft ft

CD

Fh ‘ cd

CO o

I ft & I ft

ft ft ft ft & ft ^ ft ft ft ft ft

os

• ft g,ft

W

CM C^ -sf ITN vO CO O O CMC»^s^^r^^DI>COC3^ O

CM

IN

t>

I £>- O

I

£>•

I

£>•

I cr\

I

£>

I CI I CM CM

CM CM

I I

I •sf CM

I C^CrvC^C^O^C^C^C^C^

oo 1 £>

CTl

1 rH

1 CM

1 O CM

1 CM1 CM1 CM

1 -d1 'd'

tI o

I •sf

I o

I ir\

£>

rI I>

I vtoI I lf^ ITN

!>

I sO vO O O

I cn

£>

I

I

£>-

£>•

I rH

I—

I ITN

£>

!>•

I vO

CM

I ITN

OI>

I vO

CM

I cr\

23
experiment are useful to consider here in order to give the reader some

impressions of these experiments. Complete color video records are available

for each test.

Some features of the experiments will be given: The wood cribs would

generally collapse at about 10 minutes after ignition whereas the plastic

cribs would swell and char but maintain a structure similar to its original

form. Figures 3-1 through 3-3 from PRC 7 (PRC 7-V7-4-1) display conditions

prior to ignition, full involvement of the cribs, and the glowing char remains

at the end of the test. The residue after combustion of the plastic cribs in

PRC 14 (PRC-14-P-2-1/4 ) is shown in figure 3-4. The char residue after the

flames extinguished was about 10 percent for wood and 13 percent for plastic.

Although no quantitative measure of the smoke was made, the general

character of the smoke was observed. The wood smoke tended to be light to

dark grey in color, while the smoke from the plastic cribs v/as black with

large visible flakes of soot. The smoke did tend to produce a distindt

interface within the room, but in the small doorv/ay cases the smoke did

stratify more and its interface did nearly fall to the floor. A plot of the

smoke layer interface with the room for PRC 7 is plotted in figure 3-5. This

was obtained by a study of the video record of the experiment. A photograph

of the smoke layer in the room doorway and corridor for PRC 14 is shown in

figure 3-6. It should be noted that the corridor smoke layer is above the

room door as desired in the experimental design.

3.2 Average Peak Values

Since a true steady burning period is not achievable / some arbitrary

determination of the peak values must be made. An average peak value (APV)

was determined for each measured variable and derived experimental parameter.

This was done in the following manner. The time interval was identified which

corresponded to the crib mass changing from 95 to 45 percent of its initial

mass. All parameters were numerically averaged over this time interval to

the average peak mass loss rate. This procedure, for the most part, yields

24
25

Figure

3-3.

Glowing

char

Figure

3-4.

Char

residue

of

polyFigure

3-6.

Smoke

laden

hot

gases

residue-PRC

7

urethane

cribs-PRC

14

flowing

from

room

Into

corridorPRC

14
r

.5

1

.

i a
CEILING

26

Figure

3-5.

Height

of

smoke

layer

interface

above

floor

-

PRC
£.4cWrt4i

eg

'>}.v.Ad

^tjoot
an average peak mass loss rate that is approximately + 15 percent of the rate

of mass loss in that time-averaging interval. The time-averaged peak values of

the raw data are listed for each run in volume II.

In table 3-1, significant average peak experimental results are given.

Listed are the time interval defined for peak "steady" burning and the

corresponding peak mass loss rate (m^) . Arbitrarily an average room upper

layer temperature (T) was defined as the arithmetic average of thermocouples

9, 10, and 11 near the ceiling. Also the ceiling (Tg) and upper wall temperature (T 25 ) 3-^® given. The calculated doorway flow rates, m^^ and ^^re

listed along with the doorway average velocities. The location of the zero

velocity (neutral plane height, N) position, the observed smoke height (D^)

,

and a defined thermal layer height (D^) are given in dimensionless variables

normalized with the doorway height (H^) . The thermal layer height was defined

2 as the height (z) at which d T is a maximum for the vertical gas temperature

^

2 dz

distribution in the room. This is generally the same as, or slightly below,

dT . the position where ^ is a maximum. The average volume concentrations

(dry basis) for CO

2 and O

2 are given for the upper doorway position. Other

values listed are the room static pressure differential (AP)

,

and the ceiling

(H2) and floor (Hi) total heat fluxes.

4. FREE BURN CRIB EXPERIMENTS

Following completion of the room fire experiments a series of free or

open burn experiments was conducted for each of the room crib arrangements.

Eight burns were performed - one through four wood and plastic cribs each.

Basically the same or similar instrumentation was used in these experiments

as used in the room. The experiments vzere conducted directly under the large

exhaust hood shown in figure 2-1. The purpose of this series of experiments

was to establish a basis for the rate of crib mass loss uninfluenced by a heated enclosure. Also, the direct radiation heat transfer from the crib

fires needed to be determined, and a means of computing the rate entrained

air into the crib fire had to be developed.

27
Table

3-1.

Average

peak

values

for

significant

results

C 0) O

T3 CU X

•H flj \ > 6-« Q

^ . rH (X

tN O rH E O £h o rtj <l) O w 5 ^ VIeh

> 0>

D ''C

1 o M

28
Table

3-1

(continued)

Cu

I lx°

0 *^0

29

PRC

20-P-3-1/4

5.07

4.24

2.90

-1.63

-1.10

-1.34

-1.33

11.57

9.38

3.60

-0.02

-1.63

-2.45
In addition to these experiments a single wood crib and a single plastic

crib were burned in the open in order to determine the effective heat of

combustion for each crib. This was done by measuring the rate of oxygen

consumption in the captured plume from the crib fire.

4.1 Instrumentation

The instrumentation for the set of eight experiments was more elaborate

and will be discussed first. Then the oxygen consumption experiments will be

described.

! 4.1.1 Free Burn Basis

The velocity and temperature fields were measured within, around, and

above the cribs. Thirteen bidirectional velocity probes and bare chromel-alumel

thermocouples were used. Seven of the bidirectional signals v/ere processed

through a single electronic manometer with each probe read every 35 seconds.

Two radiometers (with IRTRAN 2 windows) and two total heat flux sensors were

used to measure the radiation field. The same load platform as used in the

room fire experiment monitored the crib weight. Also O

2

and CO^ concentrations

were measured at the top of one of the crib channels. A list of the instruments|l

and their identification is shown in table 4-1.

. . li Figure 4-1 shows the distribution of the probes for the first experiment ]

with one wood crib (designated as PRC 21-W-l) . This instrumentation arrangemer^!

I was basically used for each subsequent experiment, except for the changes in J; probe positions as given in table 4-2. The crib configuration for each burn

is shown in figure 4-2 relative to sensor Hi.

30

j
Table 4-1. Identification code for free burn instrumentation

Channel

Niomber Symbol Sensor Function

0 Tl Temperature at bidirectional probe P6-V

1 T2 Temperature at bidirectional probe P5-V

2 T3 Temperature at bidirectional probe P6-D

3 T4 Temperature at bidirectional probe P2-V

4 T5 Temperature at bidirectional probe P4-V

5 T6 Temperature at bidirectional probe P3-V

6 T7 Temperature at bidirectional probe P7-D

7 T8 Temperature at bidirectional probe Pl-V

8 T9 Temperature at bidirectional probe P5-D

9 TIO Temperature at bidirectional probe P4-D

10 Til Temperature at bidirectional probe P3-D

11 T12 Temperature at bidirectional probe P2-D

12 T13 Temperature at bidirectional probe Pl-D

13 G1 Oxygen concentration (dry basis)

14 LI Mass

15 HI Floor heat :flux, horizontal element

16 H2 Flame plume heat flux, horizontal element

17 R1 Heat flux, vertical element

18 ID Sensor identification x 10 + 0.5

19 Pl-D to P7-D Bidirection probe velocity pressure

20 Pl-V Bidirection probe velocity pressure

21 P2-V Bidirection probe velocity pressure

22 P3-V Bidirection probe velocity pressure

23 P4-V Bidirection probe velocity pressure

24 P5-V Bidirection probe velocity pressure

25 P6-V Bidirection probe velocity pressure

26 R2 Heat flux, 'vertical element

27 G2 CO

2 concentration (dry basis)

31
Table 4-2. Change in probe positions for free burns

with respect to locations shown in figure 4-1

Distance from

vertical centerline Distance from Distance from

Exp. of load platform edge of crib edge of load platform

PRC 21-W-l - - -

PRC 22-W-2 P6-D to 31 cm P4-D to 1 cm — P7-D to 32 cm

PRC 23-W-3 I! !» -

PRC 24-W-4 !l r R2 to 57.5 cm

PRC 25-P-4 II R2 to 62 cm

PRC 26-P-3 >7 II R2 to 60 cm

PRC 27-P-2 P6-D to 21.5 cm n P.2 to 59.5 cm

P2-V to 9 cm P3-D to 3 cm HI to 54.5 cm

P7-D to 24.5 cm

PRC 28-P-l M P3-D to 5 cm R2 to 50 cm

Note: For PRC 22 through PRC 26 P3-D/T11 was placed at the top of a vertical

internal crib channel, and P1-V/T28 was centered in the gap between two

adjacent cribs.

32
UJ o^ O H O < K Ui t0. OD Z O HI

-I < z

«> ^ s 82

IS o a q (o 2 J W

<11.^ H O

V> < Q < ui S o I 5

<OOD ^

<3o O

33

Figure

4-1.

Basic

layout

of

free

burn

instrumentation

(dimensions

in

m)
CRIB

ARRANGEMENT

X

UJ

DD ni CVJ

I QX

CVJ

X

vjt Ql

X X

34

1

Figure

4-2.

Configuration

of

cribs

for

free

burn

experiments
4.1.2 Oxygen Consumption

A schematic of the oxygen consumption experiments is shown in figure 4-3.

It utilized the apparatus and procedure developed by Sensenig [9] for estimating

the energy release rate of burning items comparable in size to our cribs. A

single wood and plastic crib were studied. The significant results are the

rate of mass loss and the rate of oxygen mass flov7 in the exhaust gases.

4.2 Results of Free Burns

The results from these two sets of experiments will be summarized here.

The procedures for processing the data are similar to the room fire data

analysis, and only the average peak results will be presented and utilized in

this work. Table 4-3 gives the significant average peak values for the set of

eight free burn experiments. Also listed here are the average peak flame

heights based on observation of the video recordings of the experiments.

The oxygen consumption experimental results are summarized in table 4-4.

The single crib mass loss rates differ in tables 4-3 and 4-4, and this indicates

the degree of reproducibility as further discussed in appendix A.

5. ANALYSIS OF RESULTS

In this section the experimental average peak values will be analyzed in

order to evaluate experimental consistency, general trends and mathematical

correlations for these data, and to gain further phenomenological understanding

of fire phenomena. Separate sections will include the following topics:

mass rate, airflow rate, pressure, temperature, heat transfer, and finally,

a section on chemistry. In the following figures open symbols will signify -

a wood burn; filled symbols - a plastic burn; squares - 1 crib; triangles -

2 cribs; hexagons - 3 cribs; circles - 4 cribs; large symbols - widest door;

intermediate size symbols - intermediate size door/ small symbols - smallest

door (see figure 5-1)

.

35
consumption

apparatus

(by

Sensenig

[9]

)
Table

4-3.

Average

peak

values

-

free

bum

experimeni

> (0 00 I \ > rH g

I Q (A

<r* » \ > in g

0) .C 0) I O' >

fo -H o ‘H H 0) j5 M b X ITJ CJ

(A V) 4> (A CA (A > \

S3S-® s

37
•o

01oc

•HU

o

o

u

<n

I

§

H

- ^ >5^ > fp g

ro I \ rH rH g > a. *

> w

oj ( V. in > in g

0- in

id ^ id id id id

38
Table 4-4. Average peak values - O

2 consumption

Rate of Rate of

Peak Times Mass Loss 0^ Consumption

At m

V

Crib Material (s) (g/s) (g/s)

WOOD (Sugar Pine) 65-390 5.3 4.9

PLASTIC (Polyurethane) 60-255 7.0 10.3
LEGEND

NO. OF

CRIBS

AoA Im5/2|

1.96 0.98 0.49

1

2 A A A

3 0 O 0

4 O O o

OPEN SYMBOLS - WOOD

FILLED SYMBOLS - POLYURETHANE

Figure 5-1. Symbol legend for figures in sections 5 and 6

40
5.1 Rate of Mass Loss (iti ) V

The mass loss rates plotted against ventilation factor, A (A^, area

of opening; height of doorway) is shown in figure 5-2. Besides the data

for the three door widths the open or free burn values are also plotted on the

right-hand side and designated as A = The mass loss rates for the wood

in the open configuration range from 4.4 g/s for one crib to 25.6 g/s for four

cribs; and for the plastic cribs, the range is 6.05 g/s to 36.4 g/s. In general

the burning rate of plastic cribs is one and a half times that of wood. This

order of increase is preserved in the enclosure until the effects of the

enclosure (radiation feedback and insufficient airflow) begin to dominate.

The first effect of the enclosure can be seen in figure 5-2 by observing

the increase in mass loss rate as one moves from the open configuration,

= °° to A = 1.96, the largest door or the least ventilation constraining

of the enclosure runs. Except for one case, the mass loss rates increase over

that of the free burn cases. For one crib the increase is about 15% and for

two cribs the increase is about 30% for both wood and plastic. Increasing the

fuel load further to three and four cribs causes a net decrease in the enhancement which is gradual for the wood and quite dramatic for the plastic which at

four cribs exhibits a negative enhancement, i.e., the burning rate for four

plastic cribs in the enclosure is actually less than that of the free burn.

Radiation from the hot upper gas layer trapped in the enclosure and from

the ceiling and walls, absent in the free burns, is thought to be responsible

for the increased m^ seen in the enclosure. However, since the effect is

non-monotonic (enhancement decreasing after two cribs) , and even disappears in

the case of four plastic cribs, an additional mechanism appears to be operating.

The rate of combustion or burning rate within an enclosure is limited by the

airflov; rate through the enclosure opening, and ,it is empirically known that

the mass loss rate of wood cribs is limited by the airflow rate, i.e., a ventilation controlled fire. With the consideration that radiation feedback

and limiting airflow are operating in opposition, the behavior of the remaining

smaller door cases is readily explainable.

41
.04 r

.03

.01

J_

!

A

A

O

A

n

o

A

A

0

L2.0

Ao\/h^ (m5/2)

Figure 5-2. Mass loss rate versus ventilation factor

42
For the open crib burns there is a self-enhancement phenomena operating

distinctly from the enclosure enhancement. That is, two cribs burn at a higher m^ than the value obtained by multiplying the single rate by two. Three

and four cribs likewise follow this fuel loading dependence. A simple explanation involving radiation interaction between the cribs can be forwarded for

this self-enhancement. For one crib burning, the sticks making up the four

outer sides see relatively cool walls. When two cribs are assembled two of

the eight outer sides now see each other. Those mutually adjacent sides are

now going to burn more vigorously (like the internal sticks of a larger crib)

than those sides not receiving the high radiative fluxes.

For all four cribs utilized in this work the self-enhancement increases

with the number of cribs. The mass loss rate turns out to be a function of

the number of cribs (n^) to approximately the 5/4 power. Without self-enhancement m would be linear with the fuel loading (n ) . Since for all cribs the

enhancement is increasing with the number of cribs , it would appear that

inadequate airflow into the central regions of the large configurations is not

a major factor in these open burns. Hence any problems with airflow for the

enclosure burns are largely due to the enclosure restriction and are not due

to the configuration itself (this may not be necessarily true for more than

four cribs) . Thus the burning appears to be in the fuel surface area controlled

regime. Mass loss rate per unit "area" (or simply the number of cribs in this

case) is shown in figure 5-3 plotted against ventilation factor. Here the

self-enhancement correlation, the 5/4 power, is used for convenience in

collapsing the data. The behavior of both the wood and plastic fires described

previously is clearly differentiated in this plot. The behavior of the wood

fires may be characterized by a single function of the ventilation. Place the

cribs in a large-opening enclosure and m^ will increase, presumably due to

radiation feedback. Start closing the door and nothing very much happens, a further increase in radiation being balanced by lower airflows. Close the door

further and the restricted airflow will begin to dominate over radiation, thus

lowering the burning rate again (concave downward)

.

43
.009 r

.008 -

.007 -

u>

b4) .006 -

in

o .005 - >

•E

.004 -

.003 -

.002 -

.001 -

0

A

I

d

O

4d

o

A

1.0

L_

2.0

Ao^/H^ tm5/2

Figure 5-3. Normalized mass loss rate versus ventilation factor

44
The plastic follows this general behavior somewhat except that the

for smaller door openings is generally lower than that burn rate for large

openings. However, the different behavior at A = 1.96 m^'^^ is striking.

An explanation of these differences as seen in figure 5-3 is not at all evident

A further illustration of the com.plexity of enclosure effects on mass

loss rate is shown in figure 5-4. There, the ratio of the enclosure to free

burn mass loss rate is plotted against the incident heat flux to an element

on the floor. This flux reflects the degree of radiant feedback to the cribs.

A clear correlation in terms of heat flux and ventilation factor is not evident

But decreasing door width and increasing fuel load does appear to reduce the

enhancement.

A maximum enhancement of 1.3 for the largest door opening is nearly

identical with the results of Croce 110] for well ventilated (A Wh /m > 40 -

oyf o' V

m

5/2 -s/kg) wood crib fires in enclosures and for the same porosity value

1/2 1/2 As^ h

' =0.8 mm) used in this study.

(—

5 . 2 Doorway Flow

Figure 5-5 presents the mass balance - a comparison of the mass flow out,

i^out ' with the sum, m^^ plus the burning rate, m^ plus the burning rate,

(m^ accounts for up to 10 percent of the total for large fires in the smaller

door cases) . The doorway mass flow rates were determined by numerical

integration of

^irv - '^0 / (V/T ) dx (5-la)

and l4

= /all (V/T) it <5-lb)

where V and T are the velocity and temperature respectively , measured along

the vertical centerline (3-coordinate) of the doorway. The vertical distances

N and H -N are the regions over which the flow is in and out of the room

respectively. Implicit is the assumption of uniform velocity and temperature

in the horizontal direction over width W^. In general, as the door width

decreases the velocity increases, (even though mass flow decreases) especially

45
V

V

o

o

— CNI

CM

O

X

’5

< (D

O o

a

IT5 ir>

30Ji‘A Ul/Auj

)

46

I

I,

Figure

5-4.

Room

enhanced

mass

loss

rate
Mass

Balance

(s/3)|) ''ui+“!ui

47

Figure

5-5.

Mass

balance

assessment

of

doorway

flow

rates
the inflow; and hence, the capability of accurately measuring velocity improves.

There is less scatter in the 1/2-door result than in the one-door result and

less again in the 1/4-door runs.

Aside from the scatter there is a systematic bias in the results. About

10 to 15 percent more gas appears to be flowing out compared to the rate of air

going in plus the fuel gasification rate. This measurement imbalance may be due

to the use of centerline velocity values only. There is some evidence available

which indicates that the outflow velocities, being on the order of five times

target than the inflow velocities, may exhibit a more parabolically "shaped

velocity profile across the door as compared to that of the inflow. Measurement of the centerline velocity then would overestimate the outflow relative

to inflow. Until this point is resolved by future experiments, the best value

for flow rate will be taken as the numerical average of the two quantities,

m . and m. + ih , defined here as A , the airflow rate, out in V a

Figure 5-6 shows m^ plotted as a function of the ventilation factor. For

a given loading, m varies with ventilation factor to* approximately the 1/2

power. This compares with the classical result assumed for fully-developed

fires that m^ varies linearly with In every case the plastic cribs

have a higher airflow rate than the wood cribs, which is consistent with the

higher m^ for the plastic cribs resulting in higher room temperatures. For

the largest opening the m of all four plastic cribs are grouped more or less

3. together and the proper sequence is preserved, i.e., the four crib case draws

more air than three cribs, etc. They do not however exhibit the large variation

with fuel size as do the wood cribs in the large opening probably because

airflow is insensitive to average compartment temperature above 150 °C. Taking

advantage of the square root dependence of m on ventilation, figure 5-7 is

drawn to illustrate the dependence of airflow on burning rate or fuel loading.

A faired curve through the data yields approximately a 1/4-power dependence

3/4 -5/4 with the constant of proportionality being about 1.0 (kg/s) m . That is.

m

a

(5-2)

48
Figure 5-6. Airflow rate versus ventilation factor

49
I

I

(^/Q-^'i-S«3y|) V)/ ^

Ci

?

o:

ai]

! Zi

I

' a:

I V

i o:

I J

50

Fiqure

5-7

.

Correlation

for

airflow

rate
Looking solely at the plastic fires with the largest opening, the small power

dependence of m on m would sink further such that the flow would be independent of the burning rate. A complete correlation should probably include the

effect of compartment gas temperature as well, and this correlation may not

be generally applicable to other situations.

5.3 Two Layer Concept

Upon observation of compartment fires, a horizontal interfacial plane is

usually discerned between the smoky upper layer and relatively clear lower

layer. The smoke displays the primary flow pattern caused by the constraint

of the enclosure and buoyancy. Also for a compartment v;ith an open door, the

flow enters and leaves as illustrated in our experiments. Hence , the concept

of two layers has emerged in understanding and describing room fires. The

layers are separated by a surface extending from the flow reversal position at

the doorway into the room, separating the hot smoky region from the cool clear

region. Some optional definitions will be necessary to discuss this layer

height in our experiments.

5.3.1 Neutral Plane (N)

In the doorway the horizontal plane separating the hot outflowing

combustion products from the cool incoming fresh air is called the neutral

plane, a distance N above the floor. This is usually defined as the position

of zero pressure drop across the doorway such that the flow reverses direction

at that height. It can be identified as the position (N^) of zero velocity

interpolated from our data. Or it can be identified as the position (N^) of

zero pressure differential in the room. These experimentally derived heights

are plotted for PRC 8 in figure 5-8. In all cases, the velocity neutral plane

was higher than the pressure neutral plane N^. This may be a characteristic

of our long doorway for in the present setup the door frame was uncommonly

wide, i.e., in the order of 30 cm as opposed to a normal door frame of 10 cm. From visual observation the flow in appeared to fall as fluid moved from the

front edge of the door frame towards the rear part of the frame and into the room.

51
METERS

:d-

(M

PRC8-P-4-

VELOCITY NEUTROL PLPNE HEIGHT

OJ

OD

LO

OJ

<J)

<S?

<S

00

<S

<S> ©

150

I I

^

I 300 450 600

TIME (SEC)

750 900

Figure 5-8a. Velocity neutral plane for PRC 8

52
METERS

Figure 5-8b. Pressure neutral plane for PRC 8

53
This sinking phenomena has been noted previously but the effect is probably

exaggerated by this particular door frame. The interface or 'stagnation'

streamline along the distance of the door frame could be sketched from the

two measurements, and N^. Figure 5-9 presents the velocity-defined neutral

heights in the doorway normalized by the door height and plotted against the

representative upper gas temperature, T. This illustrates that the position

of the neutral plane is fairly stable over a large range of compartment

temperatures, dropping slightly with increasing temperature.

5.3.2 Room Layer Height (D)

Another parameter that is required for modeling is the height of the

thermal discontinuity in the room, D. This concept of a distinct layer height

comes from the idealized two-layer conceptual model of room fires. Unfortunately

nature doesn't conform to this idealization and the thermal discontinuity

instead of being sharp is rather a gradual, smoothly decreasing temperature

function from the more or less uniform upper gas temperature down to a uniform

ambient. Defining an effective thermal discontinuity then is somewhat arbitrary;

The one chosen here based on the APV temperatures involved finding the height

where the greatest change in the temperature gradient was occurring, i.e..

2 2 A T/Az (By differencing, the position defined to within at \

MAX'

a thermocouple location, therefore a second derivative condition was applied

in this region to uniquely locate the layer height) . Figure 5-10 compares this

definition denoted as D^/H^ , to the visually determined smoke layer height in

the room, D^/H^. The agreement in general appears to be good. The one and two

wood crib cases in the largest enclosure deviate to the right because of the

small amount of smoke generated from the small wood fires, i.e., one cannot see

the smoke. The three plastic crib cases in the smallest door deviate to the

left for the opposite reason - there is now too much smoke, it has diffused and

mixed into the cold layer.

The thermal discontinuity D^/H^ is shown in figure 5-11 plotted against

upper gas temperature. Also shown by the lines are the approximate location

54
0

55

Figure

5-9.

Doorway

neutral

plane

versus

upper

layer

temperature
i

U) 4JXi

O' •H

in

0)3 . tp ; •H I Pm

Lh

d

56

SMOKE

LAYER

HEIGHT,

Ds/Ho
APPROX.

Nv/Ho

II

CO OO O)

05 05

^

o

t

o o

4

°H/^a

57

Figure

5-11.

Thermal

discontinuity

versus

upper

layer

gas

temperature
of the doorway neutral plane taken from figure 5-9, (N-D is an important

parameter used in modeling door flow) . D^/H^ can be seen to be a strong function

of temperature and also of,„door width. Theory indicates that the maximum flow

is obtained when approaches the floor; this would be a fullyinvolved compartment fire. The trends here for lower with higher temperatures and smaller

door width are consistent with the flov; theory.

The airflow rate through the compartment is not only important from the

standpoint of limiting 0

^ to a large fire but also it can be responsible for

convecting away in the order of 50 percent of the total heat release of the fire.

Present modeling techniques start with an assumed vertical temperature distribution in the room. This distribution both above and below an assumed neutral

plane is integrated to yield the hydrostatic pressure distribution in the

enclosure and then the flow through the door is calculated using Bernoulli's

equation with an appropriate orifice coefficient. It has been previously shown

[11] that the room static pressure relative to ambient pressure is essential

hydrostatic. That is, it can be calculated directly from the room and ambient

vertical temperature distributions. The extent to v/hich that was true for the

current experiments, as well as implications to a two—layer flow model, will

be shown.

From hydrostatics, the room pressure differential is given as

for a uniform ambient air temperature T and density p . T is the room vertical,

0. 0

temperature distribution as a function of y measured (tupward) from the floor.

For a two -layer concept, the lower portion of the room is filled up to a height

D with ambient air at a temperature, T and at D the temperature rises abruptly

0

to T where it is uniform up to the ceiling. For this case the hydrostatic ^

5.4 Room Static Pressure Differential (Ap)

(5-3a)

58
pressure differential at height, y, is:

Ap = Pa g

^

(y-N) (5~3b)

Substituting for y the height of the ceiling and choosing a neutral plane, N,

which is assumed here to be independent of temperature, gives the curves in

figure 5-12. A neutral plane to door height ratio of 0.39 to 0.53, corresponding

to Np experimental values, yields Ap results that seem to bracket most of the

data.

Since the temperature factor, p g (1-T /T) appears to correlate the data

in figure 5-12, it can be used to normalize the vertical pressure distribution

data using the experimental T for each run. Figure 5-13 presents the APV Ap

normalized by p g (1-T /T) plotted against height, y in the enclosure for the

d. 3

three separate door widths. The symbols represent the numerical average of the

data at each height; the "error" bars indicate the high and low values. The

low side scatter seen for the largest opening could be reduced significantly

if the one wood crib case were not included. These were very small fires for

the large opening, resulting in very small signals. The decreased signal to

noise ratios are characteristic of the readings in the lower part of the room.

The signals there were approaching the sensitivity limit of these particular

instruments. Given a very noisy (acoustical not electrical) background like a fire situation, the scatter in the data will be a large fraction of the signal.

The APV averaging will filter out some of the larger fluctuations resulting in

the correct algebraic signs and order of magnitude of the results. (The more

sensitive pressure transducers available during this work were used exclusively

with the bidirectional velocity probes in the doorway, not for these room static

pressure measurements.)

The level of fluctuation in the pressure signal is shown in figure 5-14

for the set of pressure taps in PRC 8. Also shown is a calculation (using

eq . 5-2a) of hydrostatic pressure difference based on the time varying room

temperature distribution and neutral plane height shown in figure 5-8b.

59
o

o

0iu/Nl V

60

Figure

5-12.

Room

pressure

difference

at

the

ceiling
Ap/pag(l-^)|m)

(ui| A ‘dOOU WOUJ 33NV1SIQ

61

Figure

5-13,

Average

pressure

distributions

for

each

doorway

size
H0 l< M3

O' •H

62
In general it can be concluded that the room differential pressure field

is essentially hydrostatic and can be computed from the temperature. Also a two-layer model yields a good approximation for the pressure differential

provided the neutral plane height is knov;n.

5.5 Doorway Velocity and Flow Coefficient

The differential static pressure is responsible for the velocity profile

in the doorway. The velocity data were used previously to generate the airflow

rate. Figure 5-15 shows the door centerline velocity profile range for all

1/2 the runs normalized by (2g (T/T - 1)) and for each door size. The normalization factor comes directly from Bernoulli's equation and the two—layer model

for Ap. Like the pressure plot (figure 5-13) the symbols represent the average

of the data for that particular height and enclosure and the error bands

represent the high and low values. The lines are simply a faired representation

of the data to make the figure clearer. The neutral planes (N^) are higher than

those in the corresponding Ap plot in figure 5-13. Note also the correct order

in decreasing door width as well as the larger velocities with the smaller doors.

For a two -layer model, the upper curves would vary as (y-N)^/2 ^y^here y is the

height above the floor. The data does approximate this parabola somewhat.

However, for the incoming flov/ the two- layer model predicts a uniform velocity

below the hot upper layer with a curvature in the region between D and N. It

is probably this region where the model is breaking down owing to the lack of

a sharp thermal discontinuity in the data.

A measure of how effectively the pressure difference across the opening

is converted into velocity can be expressed by a flow coefficient. In an

understanding of conventional orifice flov7s, a flov; coefficient depends on

viscous and jet contraction effects, and is defined in terms of the entire

flow rate. For illustrative purposes, a local flow coefficient can be defined

which, on "averaging", is equivalent of a total flow coefficient. This

coefficient is defined as

C = ^measured (5-4a)

^theoretical

63
|ui| A ‘dOOU WOdd 33NV1SIQ

64

Figure

5-15.

Average

normalized

velocity

distribution

for

each
65

Figure

5-16.

Local

doorway

flow

coefficients
where

V

theoretical (5-4b)

In orifice flow, the actual velocity at the vena contracta v^ould he slightly

less, due to viscous effects, than the theoretical velocity predicted from

the imposed pressure drop Vp. Here it is of interest to determine how V7ell

the pressure field is responsible for the flow especially since the doorway

flow is counter-current. Note the actual pressure distribution is used from

figure 5-13 as opposed to one derived from temperature so the result is independent of any assumed model. If the pressure values at the measuring heights

are used then the actual measured velocity values can be found at the corresponding height from figure 5-15. Here, it should be noted that mean results

are being used from a given door width.

Figure 5-16 shows C spanning the door height for the three enclosures

configurations. The amount of averaging for both sets of pressure and

velocity data comprising the calculation of C will probably mask any variation

among the three door widths. What is clear, however, is that the outflow C is

considerably greater than that of the inflow. This is a result v^hich has been

obtained previously [12], and it is interesting to observe that the mismatch

is in the same direction as the lack of mass balance noted earlier. If instead

of using averages, a run by run calculation is performed, the scatter will be

greater but the conclusions will be the same - higher discharge coefficients

for the outflow. The calculation breaks down near the neutral plane

(cross hatched area on figure 5-16) where it was noted earlier that the point

of zero pressure difference inside the enclosure falls below the velocity

reversal point in the doorway. The familiar C = 0.68 for orifice flow is

sketched on figure 5-16 for reference. A Reynolds number based on door width

can be defined for these flows as: Re = m /W y where y is the viscosity and

a o is the doorway width. For both the inflow and outflow the above C's apply

4 for Re in the range 3 to 7 x 10 .

A more conventional or global flow coefficient can be defined for the

door flow. This definition is somewhat arbitrary and eventually needs to be

66
developed consistently with a flow model. The definition used in this analysis

was

^measured

^theoretical (5-5)

which applies to the in and out flow. The theoretical flow rate was defined

[12] as

H -N

.O V

^ = W p T

th, out o a a ^

I

' ^i)

and

m, ,

. = W p T

th, in o a a ^

I "i)

dz dz

"\

dz dz (5-6a)

:5~6b)

where in each equation z is taken as positive away from experimental neutral

plane N^, T^ are the interpolated corresponding room gas temperatures, and

T^ are the corresponding doorway gas temperatures at the measured positions.

The results of these calculations are shown in figure 5-17. Whereas the trend

differs from the "local" results of figure 5-15, the general magnitudes are

similar. In both cases the outflow coefficient is higher than the inflow

coefficient. While this is consistent with reference [12] it does not

necessarily agree with the constant value of 0.68 found by Prahl and Emmons

[13] . It appears that a more complete understanding is needed of factors on

which the C's depend.

5.6 Room Temperature (T)

It is the temperature level in the upper portion of the room which

promotes the spread and growth of the fire. The vertical and lateral variation

of these temperatures preclude assigning a definite unique value, but the

averages of T9, TIO and Til do tend to reflect an average maximum room

gas temperature T outside of the fire plume. The factors influencing T v/ill

be considered. Figure 5-18 presents T, plotted against fuel loading. Similar

to previous parameters the gas temperatures for the plastic fires are

considerably greater than those corresponding to the wood fires.

67
1N3I0IJJ303 M01J

o

68

Figure

5-17.

Doorway

flow

coefficients
The concept of "flashover" has often been applied to fires taking place

in enclosures. Localized burning continues until the upper gas temperatures

reach above about 500 to 700°C, resulting in certain critical values of radiant

flux to the floor, or more specifically to uninvolved fuel in the enclosure.

This is sufficient to "spontaneously" ignite most of the remaining fuel. This

can result in a fully involved fire sufficiently large enough to be deficient

with the room oxygen with consequent burning outside the enclosure. From

figure 5-18 it can be seen that four and three plastic cribs in all of the

enclosure appear to meet this qualitative description of flashover. (Limited

fuel in the present work precluded flaming out of the door.) Additional

confirmation comes from the fact that the cellulose filter paper tell-tabs

for these five cases (one repeat) did indeed ignite or were destroyed during

the fires. Those of the other large fires were discolored but remained intact

through the experiment. The four wood crib case does not exhibit the same

high temperatures in contrast to the three and four plastic crib cases.

The variation of T with ventilation can also be seen on figure 5-18.

There is a slight inverse relation of T with ventilation in all cases except

for the flashed-over plastic cribs where the variation, although slight, appears

to go directly with ventilation.

In a previous study [11] it was found that the upper gas tem.peratures

for smaller, "non-flashed over" fires could be correlated using plume theory.

Figure 5-19 shows AT, i.e., T - T , plotted against heat release rate, 0, to

d. the 2/3 pov;er divided by the height of the enclosure, H, to the 5/3 power, all

multiplied by an inverse ventilation parameter. The /H^^^ is a scaling

5/2 relationship from plume theory; is simply a normalizing factor required

in order for the ventilation factor, to scale properly, and 1/5 power was

chosen to best fit the data. The ventilation term, H ' /A J E , reduces

O 0^0

simply to H^/W^. Probably a room aspect ratio as well as the wall thermal

properties would also be required to generalize this correlation. But since

these parameters were not varied here, they were not included. In converting

mass loss rate to heat release rate the following heats of reaction were used:

69
Figure 5-18. Ceiling layer gas temperature versus fuel load

70
v7ood, 15.6 kJ/g and polyurethane, 22.2 kJ/g. The line shown on figure 5-19 is:

AT = 24.2

.2/3

0

h5/3

1/5

(5-6)

The result is within the scatter of the data from other enclosure fires given

in reference [11] . Like the results there, it is not obvious why the "flashedover" data are not following this correlation. It is not yet clear why a plume-like correlation should systematize the low temperature data, but it is

interesting to speculate that perhaps v/ith flashover the plume- like burning

behavior is lost and an area or volume of burning in the upper part of the room

is dominating. Also in figure 5-19 there is a gap between the low temperature

(T < 450°C) data and that of the high temperature (T > 600 °C) . This cap has

been observed previously [11] and may be a further clue for modelers trying to

understand flashover.

5.7 Room Heat Flux

The variation of incident heat transfer to the floor v;ith upper gas

temperature can be seen in figure 5-20, where the APV values of heat flux

meter H2 are plotted. This v;as a total heat flux meter more than a half

meter away from the load platform and not directly in the path of the doorway

air inflow. It may be assumed that all of the heat transfer v/ill be due to

radiation from above; however, the contribution from the crib is also significant. In fact, the heat flux to the floor target (Hi) in the free burns is

plotted against the corresponding floor flux (H2) in the room burns

(with = 0.79 m) in figure 5-21. The line labeled "no effect" represents

equal values of the floor flux; therefore, the horizontal distance from zero

to that line represents the flame contribution, and from the "no effect" line

to the data represents the enclosure contribution. The measured enclosure

flux is seen to be up to 3-1/2 times the free burn flux for the plastic cribs

and up to 2-1/2 times for the wood cribs.

The nominal incident floor heat flux value of 2 W/cm

2 may be used as an

indicator of the potential onset of flashover. The measured incident floor

71
72

Figure

5—19.

Correlation

for

upper

layer

gas

temperature
(jUio/M)

// •

73

5~20»

Incid6nt

floor

h©at

flux

(H2)

v©rsus

upp©r

lay©ir

t©mperatur©
NO

EFFECT

LO• O• lO O

(jiuo/Ml Ndna 33MJ

74

Figure

5-21

, A

comparison

of

free

burn

to

room

floor

flux
J I ir>

(jiuo/M) ^Nni30

|j

//•

75

Figure

5-22.

Ceiling

incident

heat

flux

versus

upper

layer

gas
flux and the disposition of the adjacent filter paper target is given in table

5-1. Hence, filter paper ignition or complete consumption occurred at a minimum of 1.77 W/cm

2 applied for roughly 200 s or more. Under more controlled

laboratory conditions, with radiant exposure to the same target configuration,

the paper charred black under 2.5 W/cm

2 and ripped at 120 s, but only decomposed

to a brown color under 1.5 W/cm

2

. Shivadev and Emmons [14] report on experimental ignition critical incident radiant flux for filter paper heated from below

as 2.4 W/cm

2

. Since the filter paper is only a representative room furnishing

material, it would seem sufficient to use the nom.inal value of 2 Vl/cm^ as a criterion for remote ignition or flashover potential.

Figure 5-22 presents the APV total incident heat flux (radiative + convective) to the ceiling plotted as a function of T. In this figure the extremely

high heat fluxes associated with flashover are evident and the gap between

the "plume burning behavior" and "flashover" can be seen to be dramatic. An

alternative criterion for flashover may be formulated in terms of this ceiling

heat flux based on the potential for significant pyrolysis and ignition of

ceiling lining material.

5.8 Combustion Product Composition 1

Dry O 2and CO

2

concentrations of the effluent gases were monitored

continuously by a single sample drawn from a point near the upper velocity i probe in the doorway. A correction could have been made to give the true

concentrations in the moisture laden products by estimating the H

2

O concentration'

This correction for the volume concentrations of specie i (O

2 or CO2

) would

be

^i ^i,DRY (1 - ^Hjo) (5-7a)

where Xt, r, is estimated as ^2°

= ^CO.

V 0

^2

(5-7b)

with = stoichiometric coefficient for the ith specie, The ratio
Table 5-1

Effect of Enclosure Heat Flux on Filter Paper Target

Experiments in

Decreasing Order

Disposition of

Filter Paper Average Peak

1 = charred

2 = consumed

Floor Flux

W/cm^

PRC 8 2 2.86

PRC 6 2 2.39

PRC 18 2 2.39

PRC 16, 20 2 1.77, 1.79

PRC 17 1 + 1.45

PRC 19 1 + 1.40

PRC 7 1 1.01

PRC 14 1 0.98

PRC 15 1 1.01

PRC 10 1- 1.00

77
is roughly 0.7 for the polyurethane and 1.2 for the wood (from section 6.3.1).

This correction would reduce the reported dry concentration values by no more

than 15 percent. A single sample is, only reliable if the gases in the upper

part of the room are well mixed and that the gas flow out the top of the door is

representative of that mixture. In fact, that measurement should be corrected

for profile shape to derive a true mixed-mean value. It is likely that simi*-

larity nearly exists for property profiles at the doorway such that (0.21 - Xn ) 2 and Xpn equivalent profiles. It can be shown that the profile shape factor CO2

relative to the maximum measurement is approximately 0.7 (section 6.3.1) e.g.,

mixed mean - 0.7 x^^q rmx

This correction and the "wet" correction factor will not be applied. Both

corrections taken together would yield mixed-mean values for CO

2 and 0^ in

the exhaust gases that are correspondingly lower and higher than those measured.

The following discussion will indicate the consistency and characteristics

of the O 2 and CO

2 measurements. (It should be noted that the chemical analysis

used will differ slightly from that in section 6. The consequences of the

differences are not significant so these analyses were not made consistent.)

Figure 5-23 gives the APV of O

2 and CO

2 for all the enclosure fires. As a guide the straight lines show an ideal complete combustion prediction for

assumed chemical composition of fuel with x being the excess amount of air:

fuel + (1+x) O

2 + 3.76(1+x)N

2 3.76(l+x)N2 (5-8)

for the wood experiment exceeds that of the urethane consistent with the

complete combustion model.

Using the chemical model above, a consistency check of the data presented

can be made. The amount of excess air, x, can be determined directly from the

CO

2 and O2 measurements as well as from the burning rate and airflow rates

assuming some chemistry. From the CO

2

, O

2 measurements x = Xq /^qq wood

and X = 0.83 Xq /Xqq for the urethane. This measurement is designated x(02/ CO

2 J

78

1
CM

CO2 VOLUME CONCENTRATION |%]

Figure 5-23. Oxygen and CO

2

concentrations with x = excess air (%)

as a parameter

79
From the definition of x, another calculation of x, designated x(itj^, f can

be derived from the measurement of the burning rate, it, and airflow rate, it : 3.

x(it^^, it,) E ’^actual ^required _ *a ^’^b (5-9)

jD 3 • — • m . , rm, required b

where r is the stoichiometric air to fuel mass ratio. Using the above assumed

chemical composition of the fuel, wood = 4.64 gm air/gm volatile and rp^ =

9.42 gm air/gm volatile. Figure 5-25 shows the comparison of the two methods

for determining x. Over nearly two decades (10% to 1000% excess air) of

variation the agreement is probably good, the plastic cribs in the smallest

door showing the largest deviation which is probably more indicative of a defect!

in the assumed chemical model rather than in the measurements.

It should be pointed out that ordinary enclosures turn out to be designed

very poorly from a combustion chamber point of view. A significant amount of

air is swept into the room, entrained at the edge of the plume and the exiting

airstream and swept out again without ever taking part in the chemical process.

Furthermore, it appears that x does not necessarily have to go to zero for the

fire to be air limited and result in lov;er gasification rates. This is evident

from the previous discussion of the effect of radiation and ventilation factor

on mass loss rate.

Although doorway CO measurements were not made, an estimate of CO concentration in the room effluent could be made. In a companion toxicology experimen

[15] samples of the exhaust gases from the enclosure fires v/ere collected from

the adjoining corridor and brought through an animal exposure chamber where

measurements of CO and CO

2 were obtained. These measurements can be used to

estimate the concentration of CO being produced in the room burns. Provided

the location of the sampling line and sample gas flow rate remain approximately

fixed a dilution ratio may be obtained by comparing the doorway concentration

measurements of CO

2 with those of the animal exposure chamber. That is, a certain constant fraction of what is being exhausted from the room will arrive

in the chamber.

i

80
X(02,C02)

Figure 5-24. A consistency evaluation of the determination of

excess air

81
Table 5-2 Gas Concentrations for Estimating Room CO Production

Run

Nvimber

of

Cribs

Doorway

% 0 2

j %C 0 2

i

Animal

%C02

L Exposure

%C0

Chamber

%C0/%C02

Estimated

Doorway

%C0

WOOD

i

f PRC 5 2 18 .

’ 3.80 0.874 0 . 018 0 .021 0.078

PRC 2 3 16.7 ' 5.10 0.815 0.014 0 .017 0.088

PRC 7 4 15.6 1 6.94 1.23 0 . 021 0 .017 0.12

URETHANE j

1 PRC 1 1 18.6 2.52 0 . 201 0 .019 0 .095 0.24

PRC 4 2 15.1 ! 5.86 1.25 0 .122 0 . 098 0.57

PRC 6 3 10.7 00 I1— 1

—1.48 0 .104 0 . 070 0.83

82
1.0

EXCESS AIR, X |ma, nib)

Figure 5-25, Estimation of door CO concentrations

83
Table 5-2 contains the APV's of the concentrations in both the doorway

and animal exposure chamber for the widest door ,

' Aowo xfiT = 1.96 m

5/2 Looking

first at the exposure chamber gases, the ratio of CO and CO

2

is about 0.018

for the wood and about 0.087 for the urethane, a result consistent with the | greater air requirements for urethane burning. The temperatures at the sampling

location in the corridor were low enough to insure that further CO conversion

was frozen and hence the CO/CO

2 ratios will be preserved.

Figure 5-25 shows the estimated CO concentration in the doorway flow

plotted against the excess air, x (m , A, ) which is calculated using measured

airflow and burning rates. The CO concentration is obtained by multiplying

the inverse of doorway CO

2

value by the CO/CO

2

ratio for the animal exposure

chamber. Even at 800 percent excess air the amount of CO is significant and the

difference in the two fuel types is evident, i.e., 0.25 percent for urethane

and 0.08 percent for wood. As expected CO rises dramatically as x (m^,

decreases. This will result from decreasing ventilation, by decreasing the

doorway width for example, or by increasing the fire load for a fixed door width.

6 . MATHEMATICAL MODEL

In this section the mathematical fire simulation model will be developed,

and pieces of it will be checked against the average peak experimental results.

In this process, the assiamptions , limits of applicability, and a sense of

accuracy of the ingredients of the model should become apparent.

6.1 Conceptual Formulation

I

The model follows the basic two -layer approach currently being utilized

[1-6]. The enclosure is divided into an upper (u) and lower (£) volume, each

of spacially uniform properties with temperature (T ) and i^^ specie concentration (Y^)

Also a homogeneous combustion zone is identified in the lower

space. The corresponding upper and lower solid boundary surfaces are at uniform

temperature

due to pyrolysis (but not in our application since we neglect combustible

linings), pressure differences, and turbulent entrainment. Figure 6-1 depicts

the model for our case

T and T „ respectively. Mass can flow across zone boundaries

w, u w, £

^

84
The phenomena to be considered are illustrated also in figure 6-1.

Radiative and convective heat exchange is considered and the upper (T^ and

lower (T^ surfaces temperatures will be determined accordingly. Mass flow

will be calculated for a doorway configuration based on the hydrostatic pressure

difference model corrected for derived flow coefficients. In addition, mixing

between the layers due to shear effects of the incoming flow will be considered.

Consequently a lower layer oxygen (mass) concentration (Y_ ) and temperature

T different from the incoming air flow can be found. A new plume and crib

g ,

X, entrainment model will also be incorporated in the model.

The details of these sub-models will be discussed in the following

sections. It will be useful to preview that development by anticipating the

modeling parameters that would be required. These are listed in table 6-1.

6.2 Steady-State Limitation

Perhaps the most questionable aspect of a steady-state mathematical model

for our application is the approximate constant treatment of the transient

conduction wall loss. Unless the steady-state constraint is relaxed, the

effects of this approximation must be accepted in our model. However, the

general transient effects of fire growth are always present to some extent,

and it would be useful to illustrate the applicability and limitations of a steady-state model. This will be done by considering the conservation of

energy applied to the upper spatial zone (control volume) in figure 6-1.

The general conservation of energy (1st law) for a region fixed in space

is :

(just pressure work here) and e is the energy per unit mass in volume, V. In

application to the upper layer, the region is defined as bounded by the solid

upper surfaces, the plane of the doorway, and a lower surface at height D that

85
86

Mass

transfer

Heat

transfer
Table 6-1

Anticipated Modeling Parameters

Pertains to

Parameter

Fuel AH, effective heat of combustion

L , effective heat of vaporization

r, stoichiometric air to fuel m.ass ratio

f, fraction of energy radiated per total

(i) , fraction of char produced per unit mass

Crib

k^, flame absorption coefficient

T^, effective flame temperature

Geometry

Room

m" , free m.ass loss rate per unit area

v,o ^ Geometry

kpc, wall thermal inertia

t*, characteristic time

Flow C^, doorway flow coefficients

k^, mixing coefficient

Heat Transfer k , upper layer absorption coefficient

h , upper layer convective coefficient

9

87
has no relative motion of fluid across it (i.e., V, the velocity equals W

the surface velocity, except for the transport of the plume gases and

possible shear induced mixing across it. Neglecting kinetic and potential

energy of the fluid

e = u = c„ (T-T^) (6-2) V a.

equation (6-1) becomes

^ c (T-T ) ] + ^net, added dt v a

where m is the mass of fluid in the volume at any time

m is the rate of mass efflux

i is the enthalpy

The left hand side of (6-3) contains the transient conduction heat loss effect.

The first term of the right hand side of (6-3) is usually small, but is significant when the room is "filling" with hot gases or when the change in

temperature is rapid. A conventional size room would "fill" - layer falls and

gases begin to flow out of the door - in less than 30 seconds. Also a flashover transition would raise the gas temperature several hundred degrees Celsius

in less than 30 seconds. These two phenomena are not accounted for by a steady-state model. The second sum term of the right primarily contains the

fire energy release rate and the doorway rate of enthalpy efflux. The fire

energy release rate usually is slowly varying in time up to flashover. Also

the doorway flow rate, following the filling phenomena, (and mixing) is very

slowly varying in time. Hence the steady-state model can approximate very

well all terms except the first on the right, and thus should represent

conditions between "filling" and flashover with validity.

6.3 Sub-Model Development and Verification

This section precedes the development of the governing simulation model

equations, and establishes a basis for it. Property and necessary model

parameters will be determined. Necessary model relationships and subsidiary

relationships will be formulated and assessed. Although actual fire conditions

net, out flows (6-3)

88
[ do not perfectly fit a two-layer model, associated experimental parameters,

l!

I, such as N and D, have been derived from the data. These experimental values

I will provide sets of known independent variables for theoretically calculating

other variables, such as airflow rate, m . A theoretical value can then be

a compared to its experimentally derived value. The extent of agreement would

establish the acceptability of the theoretical approach.

6.3.1 Effective Heat of Combustion and Air-Fuel Ratio

A chemical equation, based on the analysis of table 2-2 and a char analysis

(except for sugar pine where a pure carbon char v;as assumed) was developed for

each fuel material. Char yields were based on the residue of the crib burns

after flaming ceased - approximately 10 percent for wood and 13 percent for the

plastic. The combustion reaction for each is then given as sugar pine (6.5% moisture)

CHf • O.O 966 H

2 O + 0.860 (O

2 + 3.77 N2

) (solid) (air)

0.777 CO

2 + 0.969 H2

O + 3.24 N

2 + 0.223 C

(char

)

Polyurethane

*^^1. 27^0. 0823°0. 215 + ^°2 ^2^

(solid) (air) — 0.844 CO

2 + 0.565 H2

O + 3.88 N

2 + 0.156 ^^0.0765 °0,0582

(char residue)

These assumed reactions will establish the stoichiometric relationships required

in subsequent modeling. The oxygen and air stoichiometric mass ratio, based

on volatile fuel since that relates directly to measured mass loss, is given

in table 6 - 2 .

89
Table 6-2

Air and 02“Fuel Ratio

rox

g 02 /g volatile fuel g air/g volatile fuel

Sugar Pine 1.14 4.92

Polyurethane 2.10 9.03

An effective heat of combustion, AH, was derived by considering two

analyses. First, from the free burn experiments

AH =

AH m»

ox 0. m (6-4)

Based on complete combustion of the fuels, AH^^ was estimated to be 13.7 to

15.3 kJ/g02 for the wood and 12.9 kJ/g02 for the plastic. In general, is

approximately 13 kJ/g O

2 for most fuels under various degrees of combustion

completeness. Using these values, substitution into equation (6-4) from table

2-2 yields AH = 12.7 - 14.1 kJ/g volatile fuel for sugar pine, and AH = 19.0

kJ/g volatile fuel for polyurethane.

A second approach was to analyze the room as a combustion calorimeter

utilizing the O

2 consumption method for detemnining AH. Based on the doorway

outflow it can be shown that:

H -N

o (0.23-Y ) pv dy (6-5) ox

where the property profiles are assumed similar as

•\ ( 6 - 6 ) H.-M j

with y measured up from the neutral plane.

Since the outflow rate is

m.A = WJ <«-’>

it follov/s from equations (6-5) - (607) that

0.23 - Xo\

0.23 - Yc Cpv) ma.1

90
Values for Y . (Gl) , m . and m are derived from table 3-1. The profile

ox, min out, v

^ factor, n+1 , was taken as 0.7 (v;hich corresponds to n between 1/2 and 1), and

2n+l

a nominal value of = 13 kJ/g O

2 was used. The results are shown in figure

6-2 along with theoretical (complete combustion) values of = 28 kJ/g and

AH^ = 17.7 to 19.7 kJ/g. Except for the data point of P-3-1/2, the room,

calorimeter results appear credible.

The heat of combustion results are tabulated in table 6-3. In anticipation

of the fact that the simulation model will yield an average upper layer temperature, which would be below actual room temperatures near the ceiling, somev/hat

maximum values were selected for AH in the modeling analysis. This vzould tend

to compensate for the inherent inability of the model to predict the maximum

room temperature.

Table 6-3

Effective Heat of Combustion Results

Theoretical From Free Burns From Room Exp. Mode 1

AH AH AH AH

(kJ/g) (kJ/g volatiles) (kJ/g volatiles) (kK/g)

Sugar Pine 17.7-19.7 12.7-14.1 11- 15.5 15.0

Polyurethane 28.0 19.0 14 . 5-26.0 23.0

6.3.2 Fire Induced Flow

Three phenomena of fire induced flows need be quantitatively described

in the mathematical model. The first is flow due to pressure differences

through enclosure openings. The theory of this is well developed; however,

the definition and determination of flow coefficients needs to be established.

This will be addressed. A second flow phenomena, not well quantified, is the

mixing that occurs in the vicinity of openings as jets of fluid entrain on

leaving and entering. Finally, flow induced by a fire plume can be predicted

by several buoyant plume theories, but none truly predicts the correct fluid

entrained near the base of the fire 116]. Moreover, no entrainment theory

exists for crib fires. These relationships will be developed.

91
m

CO

I

CO

lO

CO

CM

NOIlSnaNOO jo ivjh jaiiojjjj ojivniisj

92
6.3. 2.1 Doorway Flows

The doorway flow coefficients were determined in section 5. An average

analytic expression was used in the modeling to represent the variation in

coefficients with doorway width.

C ^ = 0.68 (6-9a) out

C . = 0.58 - 0.31 W (m) (6-9b) xn o

These coefficients were used in a two-layer flow model calculation.

For a two—layer model with a "warm" lower layer as shown in figure 6-1,

the doorway flow can be determined by integrating the general equations [12]

The results follow:

By expanding equation (6-lla) about b = 0 for small b (and N-D not small)

= I -of (N - P/2)

8 (I - (N-P)'

J

The second term in the bracket is of the order of 5 percent of the first term.

This second term was dropped in the following analysis, but equation (6-llb)

was used in the simulation later.

93
By using the experimental values of N, D, (=D^) (=T) from table 3-1,

, ^ and the C's from equation (6-9), the mass flow rates were calculated from

equations (6-10) and (6-lla) . The results are compared in figures 6-3a and 6-3b.

Some anomalous results are derived from eq. (6-lla) in a few cases because

measurements indicate that N was less than D. For these cases the computed

flow was set equal to zero; also, the same comparison is made for inflow results

in figure 6-4 with a constant value of C = 0.68 which is usually assumed. It

appears that the variable C values derived for the inflow provide better theoretical results. But a general conclusion here must await a more complete study.

6. 3. 2. 2 Mixing Model

Upon correction of the lower layer thermocouples (appendix B) or by direct

reading of the small diameter thermocouple or aspirated probes, it was clear

that mixing was occurring between the layers. This would increase the lower

layer gas temperature and decrease its oxygen concentration by dilution with

combustion products. The nature of this mixing is a turbulent shear entrainment that occurs as the cold air jet enters the doorway and diffuses dovmward.

It has been observed previously [17] , and a similar mixing phenomena also takes

place as a hot jet emerges from a doorway into another room impinging upon its

ceiling. These entrainment phenomena are illustrated by the sketch in figure

6-5.

A model for this mixing process is proposed as follows:

The net rate of mass entrained m^ is given as:

m

e Ae Ve

where is the density of the fluid entrained

A is an effective jet surface area

is the velocity normal to the jet surface.

( 6 - 12 )

The entrainment velocity is assumed proportional to an average jet velocity

(based on the mean doorway velocity) V:

V = k V (6-13)

e m

94
FLOW

RRTE

(THEORY)

.000

0-150

0.300

0.450

0.600

0.750

0.900

1.050

1.200

COMPARISON OF INDUCED FLOW RRTE CIN)

TWO LRYER MODEL

MEASURED AVERAGE PEAK VALUE

9

9

*

^ Aa A

A

A

*

0.000 0.100 0.200 0.300 0.400 0.500 0,600

FLOW RATE (EXP)

Figure 6-3a. Comparison of calculated and measured doorway inflow

rates

95
FLON

RATE

(THEORY)

.000

0.125

0.250

0.375

0.500

0.625

0.750

0.875

1.000

COMPfiRISON OF INDUCED FLON RPTE COUT)

TNO LPYER MODEL

"MEPSURED PVERPGE PEPK VPLUE

e

®

A

A

A

A

<S I 0 . 000

~I0.125 0.250 0.375 0.500

ELON ROTE (EXP)

“I0.625 0 . 750

Figure 6-3b, Comparison of calculated and measured doorway outflow

rates

96
FLOW

RATE

(THEORY)

.000

0,150

0.300

0.450

0.600

0,750

0,900

1.050

1.200

COMPPRISON OF INDUCED FLOW RPTE yN)

TWO LRYER MODEL

MEPSURED PVERPGE PEPK VPLUE

e

9

9

*

*

in

A

A

A

*

T

1 ® 1 1 1

1 0.000 0.100 0.200 0.300 0.400 0.500 0,600

FLOW RATE (EXP)

Figure 6 4. Comparison of calculated and measured doorway inflow rates

in kg/s = 0.68

97
98

Figure

6-5.

Entrainment

concept
Also the doorway mass flow rate is given as

A 4. = P V (K -N) out g,u out o o (6-14)

and

m. = p V. W N

in a in o (6-15)

Also the jet surface area is assumed to be

A = 6 W

e (v)

n-1

(6-16a)

where 6 = N - for the inflow (D^ and defined in fig. 6-5)

and 6 = D

2 - N for the outflow.

a)

n-1

(6-16b)

(6-16c)

is an empirical factor included to account for area changes in the

jet as it expands from the doorway into the corridor.

A 6 ;< 0 implies m^ = 0.

Hence for the flow in

and for the outflow

(6-17)

(6-18)

Results (unpublished) for a room-corridor model indicated k =2 and n = 1/4. m

Equation (6-17) was applied to the current room fire experiments.

In order to apply equation (6-17) to see if it fits the current data,

m^ had to be determined. It was not measured directly, but could be estimated

from an energy balance on the lower layer provided a mean lower (and upper)

layer temperature T was known. The corrected value T22 was selected for

g ,

X, T (and Tq , „ , was used for T ) . A conservation of energy equation

^ f ^ y^-LU^-LX 9 / was written for the control volume shown in figure 6-6. The results yield the

mixing ratio

'

T

<3f i

T

g,u

(6-19)

99
100

Figure

6-6.

Analysis

to

determine

A

and

the

mixing

coefficient
Figure 6-7. Mixing correlation

101
The accuracy of this estimate for could be of the order of + 50 percent.

Results correlated by equation (6-17) are plotted in figure 6-7 and data points

of questionable value are indicated. The was found to be 0.5 on average. It

could be as high as one, but is not two as derived in the small-scale results

also shown. It is not clear whether this discrepancy is due to inaccuracy

in the estimate for m^, the offset corner doorway of our room, or an inadequate theory. The simulation model will use the mixing model given by

equation (6-17) with = 0.5 and n = 1/4.

6. 3. 2. 3 Crib Fire Entrainment

The entrainment of air into the fire was viewed as being composed of two

components: the contribution from the crib(s) plus that from the fire above

the crib.

6. 3. 2. 3.1 The Crib

The flow of air and combustion products through and around the crib v/as

divided into three components: that through the four internal channels, the

flow up the exterior sides of the crib, and third, the flow between adjacent

cribs

6. 3. 2. 3. 1.1 Internal Channels

A pressure probe/thermocouple P1V/T8 or P3D/T11 sensed the flow coming

up through the channel. No systematic difference was noted for the temperature

and velocity for different numbers of cribs burning. In calculating mass flow,

it was assumed that the velocity is representative of an average across the

cross section of the channel exit and further, that the flow through the flow

monitored channel is the same as the flow through any of the four channels.

The results for mean temperature, velocity and mass flow for the four

experiments are found to be:

102
Table 6-4

Crib Channel Flow

T (°C) V (m/s) m/crib (kg/s)

Wood 874 3.15 0. 0191

Polyurethane 741 2.72 0. 0186

These mass flow rate results are consistent with the theory of Block [8]

(appendix A)

6. 3. 2. 3. 1.2 Exterior Surfaces of Cribs

The exterior sides are those not adjacent to sides of another crib. For

one crib there are four exterior sides, for two cribs there are three exterior

sides for each crib, etc. Sensors (P4D/T10 and P3D/T11) were located at the

top of the crib at various distances from the sides of the crib with a region

where the flow was expected to be upward. The shape of the profile for

those cases where more than one probe was present (one wood, one and two plastic)

indicated that a laminar-like distribution of flow velocity would be adequate

to describe the profile. That is, a two-constant parabolic fit was made for

these data. The constants (r and 6) were further refined using the data for

the remaining runs which had only a single probe in the exterior layer. The

distribution is given as

= I (i - 1)

'

with the constants r and 6 tabulated in table 6-5.

Table 6-5

Crib Flow in Exterior Layer

r (m/s) 6 (m) T m/exterior perimeter

(kg/s-m)

Wood 8.8 0 . 038 171 0.024

Polyurethane 18.1 0 . 074 504 0 .051

The mass flow per unit length of exterior side was obtained by integrating the

velocity profile assuming an average density from average temperature readings.

103
Tbe boundary layer thickness, 6, are of the same order as those calculated

natural convection along a vertical hot plate.

from

6. 3. 2. 3. 1.3 Contiguous Surfaces

In a manner similar to 6.3. 2. 3. 1.1, the average velocity and temperature

were measured (P1V/T8 when not used in the channel) and flow rate calculated

per intersection. One crib has no intersections, two have one, three have two

and four cribs have four intersections. The results are given in table 6-6.

Table 6-6

Flow Between Cribs

T (°C) V (m/ s

m/intersection (kg/s)

Wood 624 1.1 0.0088

Polyurethane 926 2.8 0.014

The calculated flows to the top edge of the crib obtained by summarizing the

three components

T+ni. . +m crib channel exterior contiguous

are shown in table 6-7.

Table 6-7

Rate of Air Entrained Over Height of Crib

iti (kg/s) crib '

Number of Cribs Wood Polyurethane

1 0.042 0.068

2 0.082 0.12

3 0.012 0.18

4 0.16 0.23

For the wood 45 to 50% of the flow comes through the interior channels, 30 to

55% from the exterior surfaces and the remaining 0 to 20% comes up through the

contiguous intersections. For the plastic only 25 to 30% comes from the interior channels and 45 to 75% comes from exterior surfaces while the contiguous

104
surfaces account for about the same percentage as the wood.

The flow was found to be proportional to burning rate

m . , = 8

crib ( 6 - 22 )

A more exact expression given in terms of potential heat release Q (using AH

of 15.6 kW/g/s for wood and 22 kW/g/s for plastic) is

m ., = 0.0022 Q crib

0.7 (6-23)

where Q is in kW and in kg/s.

6.3. 2. 3.

Fire Plume

Pressure probes and thermocouples were located on the centerline above

the crib up to a height of four times the crib height. Also, a rake of instruments was located at one crib height above the crib in order to obtain the

transverse distribution. Figure 6-8 presents results of centerline velocity

and temperature during the steady peak burning period. Velocity is normalized

by the nominal heat release rate raised to the 1/5 power and the height above

• 2/5 the top surface of the crib, z, is normalized by Q . This scaling comes

from a study 118] of diffusion flames above a gas burner and the results of

that study are shown by solid lines. Although exhibiting considerable scatter

the trend of the data in general is consistent with the gas burner results.

The temperatures are lower than those of the burner, but this will not significantly affect entrainment since its dependence on temperature is weak.

Included on the figure is a dimensionless parameter, V/ ^ 2g z AT/T^

which is the ratio of the measured velocity to a buoyancy generated velocity

due solely to a temperature difference. The line is the burner result and

again the crib data exhibit similar results to the burner. The fire plume

above the crib appears to be consistent with the results of the gas burner

even though the behavior at z = 0 is different in the two cases. The velocity

of the gases leaving the burner surface is negligible compared to the velocity

of the gases through the crib channels and sides.

105
AT(°C)

V/Q'^®(ni-s'-kw'^®

CENTERLINE CORRELATIONS

GAS BURNER CORRELATIONS

Z, measured from top of crib

Figure 6-8. Crib fire plume velocity and temperature correlations

106

V/V

2gzAT/To
Results in the radial direction (data not shown) indicate that the

transverse scaling developed in the gas burner study was also adequate for

predicting the behavior of the fire plume above the cribs. From that report

[18] the near field approximation for fire plume entrainment is

m. . , = 0.55z 0^/^ in kg/s (6-24) fxre plume * .

where z is the height above the top of the crib (in m) and Q is the nominal

heat release rate in kW.

6. 3. 2. 3.Results

Combining the two expressions for crib and plume entrainment yields

m =0.0022 +0.055 [D-0.53] (6-25a)

P

or approximately from eq. (6-22)

,

m = 8.m, + 0.055 [D-0.53] (6-25b)

p b

where D (m) is the height of the thermal discontinuity in the room. Also, the

crib is 0.35 m high and sits on weighing platform 0.18 m above the floor, hence

3 = D - 0.53 m. Entrainment by the fire plume is envisioned to occur over the

distance from the top of the crib to the thermal discontinuity D.

Figure 6-9 shows the derived entrainment expression, labeled ' against ^ the measured flow through the door, m The abscissa is the ^ measured

mean of the two flows m , and m. + A. . The ordinate requires the heat out mb ^ release rate and the height of the thermal discontinuity. A 45° line is

provided for comparing the two results. The agreement is generally within + 0.25

percent. For those data with D j< 0.53 m for large fires and small doors, the

agreement as expected is poor. The thermal discontinuity is approaching and

falling below the top of the crib thereby cutting off the entrainment of the

fire plume. Those points then will show only what the crib contributes to

entrainment about one half to somewhat more than one half of the measured

flow. In reality the fire pliime will entrain from a region around the fire

107
r

lil PREDICTED = .0022

6**'^ +.055 0®'® (0—.53)

lil MEASURED (kg/s)

Figure 6-9. Comparison of calculated and measured entrained flow by

the fire plume in the room experiments

108
up beyond the thermal discontinuity and an account of this can be incorporated

into the model.

6.3.3 Species Concentration

In both the upper and lower layer it is important, for many reasons, to

knov7 the concentration of various gaseous and particulate (soot) concentrations. Carbon dioxide, H

2 O, and soot are needed to predict the absorptionemission properties of the gases. Oxygen and CO are significant with respect

to toxicity, and "smoke" particulate concentration determines obscuration.

Some expressions for calculating these concentrations in the upper and lower

layer will be derived and applied in the forthcoming analyses.

A conservation of species mass will be made for the upper layer including

the combustion plume (CV^) , and the lower layer (CV^^) . This is shown in

figure 6-10. The mass concentrations of species i are assumed uniform in the

upper (Y^ and lower (Y^ layers.

In general, the conservation equation is

V ^

- V:

(6-26a)

where is the net rate of production of specie i.

For steady-state and Y^ uniform in the control volume.

E lY, tvi)- ' VSpecific application to CV^ (fig. 6-10) yields:

TVl„ W\w -V- W

• \ \y /• ’ \ V / Vi K, fV\b - c-i) V, lA, (i-w)

(6-26b)

( 6 - 27 ) 'a

' '"v

where is the stoichiometric coefficient of i (m = 1 , i = reactant^* m = -2,

i = a product)

Vj is the stoichiometric coefficient of the fuel (solid)

is the molecular weight of i

is the molecular weight of the fuel (solid)

m, is the rate of volatile fuel reacted within CV^. b I 1

- 0) is the ratio of volatile fuel produced to solid fuel burned

109
110

Figure

6-10.

Conservation

of

species

in

upper

and

lower

layers
and (6-28) *b

=

in ' r < A /in V _ a' V

A— , r > m /m

r a' V

For the lower layer, where no generation is considered,

(m + m ) Y. - m y. - m Y. -AY. =0 (6-29)

e a ±,l e i,u a i,u a i,a

Also the volume or mole fraction (X^) can be estimated from

29 Y.

X. = 1 ^ M.

1

(6-30)

where a mean mixture molecular weight of 29 is used. Some special results

will be considered.

For i = CO., and H„0, Y. =0 and subtracting equation (6-29) from

A 2. / d.

Y. =( —I It ]/ Im + m / (6-31) 1,U \v^

J \ a V/

Substituting the stoichiometric coefficients from the chemical equations in

6.3.1 and the values in table 3-1 into equation (6-31) for CO

2 yielded the

comparison shown in figure 6-11. The corrections discussed in 5.8 were not

accounted for in the doorway CO

2 values.

For 1 = «2' = 0.23 and similarly

*. -M) (fe-) (6-32a)

Y.i,u m + m

a V

It should be noted that the stoichiometric air to volatile fuel mass ratio

r =

(1-w)

Thus

0.23 j^l - r Aj^/A^ (6-32b)

2,u 1 + (A /A ) v' a

111
For the lower layer, from equation (6-29)

2,1

0.23

1 + (lllg/*^)

(6-33)

Similar analysis can be applied to other components of interest for which

chemical information exists. The level of accuracy depends on the assumed

stoichiometry, the model, and, of course, the interpretation of the measured

value. The extent of agreement shown in figure 6-11 is a good measure of the

anticipated accuracy for this (sub)model.

6.3.4 Crib Fire Radiation

A comparison of the free burn and room burn radiation field as shown

in figure 5-20 suggests that a significant amount of radiant energy is transferred directly from the crib fire in the room experiments. Hence, a model

is needed to predict this flux field. Ultimately in the room fire simulation

model the incident floor flux will be calculated as the super position of the

room upper layer contribution plus the crib fire flux. The crib fire model

will not account for distortions in flame shape due to the room geometry. The

flame heights did not vary greatly with crib loading and did not impinge on

the ceiling in this study.

The mathematical model selected is based on the analysis of Dayan and Tienj

[19] for a cylindrical grey flame (figure 6-12) . The crib and flames above

will be represented by an equivalent cylinder of radius

r =

c * TT (6-34a)

where ^ is the projected top surface area of the cribs

A

o , T

= n +

c ^g,T ^T ^^c^ (6-34b)

where G (n ) are the number of gaps for n cribs and a, the gap area =

X C C f X 2 0.245 x 0.07 m . The height of the cylinder is the sum of the crib height

plus flame length above it. The analysis by Dayan and Tien yielded the

112
C02 VOLUME FRACTION, DOORWAY EXIT (%|

Figure 6-11. Comparison of calculated and measured doorway CO,^

concentrations

i

113
incident radiant flame flux to a target element dA with a unit normal vector

n = ui + vj + wk for h/r^ ^ 3

q" = a ^“l^l 2 2 ’3 3 (6-35a)

where

= (1 - -

2

) + 2e ^ ( - + ^ ) = 1-e 0 1 Vi 1 1.2 (6-35b)

a „ -

0^3 = IT [I^(y) -L^(y)] = 1 -1.9y

0 = y < 0.6

y =

2 r kc f

sin 3

F. =

F^ =

1 r 2

-j— (—^) (it - 20 + sin 20 ) .u

4ir L o o

1 ""c TT- (t ) (tt - 20 + sin 20 ) .V

2tt L o o

-

1 (— 2 ^

) cos 0 . w TT L O

(6-35c)

(6-35d)

(6-35e)

(6-35f

(6-35g)

with k^ and an effective absorption coefficient and flame temperature

respectively. Since the crib fire is a hybrid of hot sticks and flame, k^

and are not truly representative of a homogeneous flame. A somewhat

arbitrary but suitable process was selected to determine these values. The

kj was selected from the available and appropriate literature results, and

the "best" to make the model results agree with measurements Rl for one

and four cribs was determined. This is summarized in table 6-8.

I

I

114
115

Figure

6-12.

Cylindrical

flame

model

of

Dayan

and

Tien

[19]
Table 6-8

Absorption Coefficient and Flame Temperature

Literature Value

Source Material

^f

(m (K)

Markstein [20] Polyurethane foam,

pool fire

1.28 + 0.17 1408 + 12

Hagglund and

Persson [21]

Flames above wood

crib fires of

thickness 0.2 to

2 m

0.75 - 1.3 873 - 1303

Modeling Values for Cribs

k _f ^f

Crib Material (m (K)

Rigid Polyurethane 1.3 1200

Sugar Pine 0.8 1100

116
Equation (6-35) was applied to the free burn experiments to calculate the

incident flux at the sensor locations shown in figure 4-1. The theoretical

results are compared to the measured values in figure 6-13. These results

indicate the degree of accuracy of this cylindrical flame model.

Another parameter needed in the simulation model is the fraction of total

energy release from the fire plume that is radiation. In a room, this energy

will be transported to all portions of the room and intermediary absorbing gases.

For the model considered, the total radiation energy rate would be

= A^E^a T^ (6-36)

where A^ is the total surface area, 2tt r (H + r ) f c c

— L k is the emissivity, 1 - e m f

. , , ^ 3.6 Volume , o 1 with L = = 1.8 T—;

; =—777 m Area 1/r + 1/H

and

a is the Stephen-Boltzmann constant.

For the values selected previously for and T^ (table 6-8) and H (table 4-3)

,

it was found that the fraction, f

where f

= b

is 0.16 to 0.20 with an average of 0.18 for wood, and 0.20 to 0.27 with an

average of 0.22 for the polyurethane.

6.3.5 Room Heat Transfer

In this section the room heat transfer processes will be formulated. The

radiative transport and upper layer emissivity will be defined, convection heat

transfer to the ceiling will be based on existing correlations, and an approximate

wall conductive model will be utilized. These relationships will be verified

by comparison with the experimental results.

117
CALCULATED

FLUX

(W/cm^j

FREE BURNS

RADIATION TO REMOTE SURFACE ELEMENTS

CYLINDRICAL FLAME MODEL

3.0

2.5

2.0

1.5

1.0

0.5

0

k(m'| Tf(K) SYMBOL

CRIBS 1 2 3 4 WOOD 0.80 1100 OPEN

R1 6 A 6 6 PU 1.3 1200 FILLED R2 D- A- o- o

HI 9 9

H2 - •A -o -o

0 0.5 1.0 1.5 2.0 2.5 3.0

MEASURED FLUX |W/cm^|

Figure 6-13. Comparison of calculated and measured free burn radiative

incident target fluxes

118
6.3. 5.1 Upper Layer Emissivity

Fundamentally the emissivity (absorptivity) of the layer gases can be

calculated provided ^ their composition, and Y , are known. The ^ soot,

soot concentration was not measured in these experiments. However, it can be

estimated that the fraction of soot produced per mass of fuel decomposed (g)

is roughly 0.01 for wood and 0.05 for most plastics. With this estimate and

the calculation code developed by Modak 120] , the upper layer emissivity was

determined. But since this calculation relies on an estimate for g, an

empirical approach was taken, for computational convenience, that yielded

comparable values.

Modak 's calculation is given functionally below

^MODAK ^MODAK ^CO^, ^soot, ^m ' "^g^ (6-38)

The volume fractions were determined from equation (6-31) for H

2

O and CO

2

/ and

X 4. = 3 soot (^-3 V!lx

\^soot/\ a ^ *v/

(6-39)

where is the density of the upper layer = p^T^/T

3 and PgQQ^ is the density of soot, estimated as 2 g/cm . The path length

was taken for this calculation as 4 (layer volume/surface area) or

_ 2 WL (H-D)

m (H-D) (W+L) + V7L (6-40)

and Tg the bulk upper layer temperature was taken as T from table 3-1.

The empirical method is based on the assumption of a grey gas with a bulk representative absorbing-emitting specie concentration proportional to

m^/ (m^ + A^) - And since the spectral absorption coefficient is proportional

to concentration, it is assumed that

[m / (m + m ) ]

, _ V V a ^g “ [m_/ (m_ + m^ )

]

V V flame

(6-41a)

119
where the flame specie concentration factor is

(

fiz \

\ V a /flame

(6-41b)

with the fraction of stoichiometric air required by the flame. Literature

results suggest that x may be as high as 20. It seemed that x =1, i.e.,a

3. 3

stoichiometric flame, yielded satisfactory agreement between

^APPROX (6-42)

This comparison is shown in figure 6-14. Although a better fit could have

been achieved by a slightly larger value of x , it was not done since the

3

certainty of B is questionable in the alternative analysis. Equation (6-41)

and (6-42) will be used in the simulation model.

6.3. 5. 2 Radiative Flux to Floor Target

The incident radiative flux to the floor sensor H2 is the siim of the

flame and upper layer contribution.

d" = d" + d" ^r,2 ^r, layer -r, flame

From section 6.3.4 the flame flux is given as

(6-43)

*^r, flame “ °"3 ^3 °

”^f (6-44a)

with L = 0.906 m and H = is the crib height, 0.35 m, and is

an average flame height above the crib based on the free burns.

!

0.63 m, wood

(6-44b)

0.87 m, polyurethane

The layer contribution is based on the analysis of Modak and Matthews

[23]

It is given as the sum of contributions from four layer sectors shown

in figure 6-15.

4 ^r, layer E

i=l

^1 (6-45a)

120
X0dddV3

Figure 6-14. Comparison of empirically calculated layer emissivity

with Modak's [22] method

121

MODAK
L=2.18

Figure 6-15. Floor target flux by Modak and Matthews [23] analysis '

f

122

SIDE

VIEW
e -k L

g m T

g

+ -k L

e g m

oT 4 -| c

tan (6-45b)

where

2ab (H„ - H ) T — — m (H

2 - H^) (a + b) + ab

Tg is the upper layer gas temperature

is the upper layer surface temperature

and and are the lower layer surface temperatures. This model assumes

a non-absorbing, non-emitting lower layer gas.

Values from table 3-1 were used as input data to evaluate q" These

results were then compared to the measured values from sensor H2 . The thermal

height D was used, T^ = T^, T^ = (T8 + T25)/2, and ~ “ 22°C. Figure 6-16

shows the computed flux, (Q2) plotted against the experimental values

r , z (QEXP) . Also tabulated on the figure are the flame flux, ^ (QF)

,

and

the layer flux, layer' computed flux overestimates the measured

values at high flux which corresponds to the hot, smoky fires. The reason

for this is that the floor sensor is responding to the flux emitted from the

lower portion of the layer which is at a lower temperature than T. The

thermocouples near the ceiling representative of T cannot be "seen" through

the smoke by the sensor. This is a deficiency of the model and may necessitate the inclusion of a stratified temperature to cope with radiation transfer.

The level of improvement required here will depend on the application.

123
TOTRL RRDIRTIVE FLUX TO FLOOR SENSOR H2

U..

<X) TEST ID Q2EXP Q2 QF QR QR8

U/Cf1>x>k2 U/CMHot:2 U)/CM>k*2

1 0.153 0. 134 0.080 0.054 0.035

2 0.143 0.164 0.080 0.084 0 . 063

3 0.248 0.290 0. 164 0.125 0.113 Q

G? 4 0.543 0.484 0.238 0.246 0.256

5 1.010 0.789 0.313 0.471 0.489

6 0.381 0.305 0. 196 0. 108 0.096

7 1.040 0.805 0.381 0.423 0.455

8 2.390 2.945 0.531 2.414 2.596

9 2.860 5.718 0.686 5.032 5.236

10 0.420 0.312 0. 164 0.147 0.146

11 0.580 0.521 0.238 0.284 0.301

12 1.918 9.924 9.919 9.996 8.6§1

CD 13 1.000 0.944 0.381 0.563 0.651

14 2.390 4.328 0.531 3.798 3.915

15 0.430 0.403 0. 164 0.239 0.259

16 1.450 1.484 0.318 1.166 1.263

17 1.400 1 . 403 0.3 lit 1 . 032 1 . 195

G?

G? 18 0.390 0.367 0. 196 0.171 0. 180

19 0.980 1.155 0.381 0.774 0.855

L/5 20 1.770 3.549 0.531 3.318 3. 177

21 1.790 3.721 0.531 3.190 3 . 35b

G?

:± “

®

<S? <S

00

Q

CU

®

cs? G? ®

T—

G? a®

e w

G?

G? 1 1 1 1 1 0.00 0.50 1.00 1.50 2.00 2.50 3.00

Q2EXP

Figure 6-16. Comparison of calculated and measured incident flux by

sensor H2

124
6. 3. 5. 3 Total Incident Heat Flux to a Ceiling Target

The heat flux to ceiling sensor Hi has two components: a convective

flux, q" ; and a radiative flux q .

qV = q" + q" (6-45) 'll ^r

The radiative component will be derived using the methodology described

by Seigel and Howell [24] . Hi is located as shown in figure 6-17 as a surface

element dA, at T^, the cooling water temperature. The surfaces bounding the

upper layer, labeled 2, are the walls at a uniform temperature measured by

thermocouple T25. The interfacial plane of the thermal discontinuity, labeled

3, is considered to be at a temperature T = 22°C which approximates the

3. global temperature of the lower region. The absorbed radiative flux by sensor

HI is then given as

<4; = ('12 qa) + ''12 (1 - ^g)

+ ^13 (1 - “ ^3^ - o (6-46)

The configuration factors are

^12 ^13 , ^ and

13 E F. (a.,b.:1 1 1

i=l

^ (^a) ^ ik).

i

2 2 a + (H-D)

b^ + (H-D)^

Thus

,

q'' = e oT + ^r ,1 g g

+ F

(1 - f'13

) f

- ^g)

13

" ^g)

a T

w

4 4 aT - aT

a s

(6-47)

(6-48a)

(6-48b)

(6-48c)

(6-48d)

(6-49a)

125
dAi

IHIIiTs

126

I

Figure

6-17.

Derivation

of

radiative

flux

from

the

upper

layer

to

the

ceiling

element,

dA
CO

CM

csi

II

127

Figure

6~18.

Ceiling

convective

heat

flux

analysis
with

e = 1 - e (6-49b)

g

and L = m

4 X Volume

Area

(2WL (H-D)

(H-D) (W+L) + WL (6-49c)

The convective flux to HI is given by Zukoski and Kubota [23] as

h (T , - T ) c ad w

(6-50a)

where

is the adiabatic wall temperature

T is the wall temperature w

and h is the convective coefficient

c which

h

c Q*V3 C (r/H) (6-50b)

(1-f)

p c T

a p a ^

^ (6-SOc)

(see figure 6-18) . The convective coefficient falls off as the radial position

j

I r increases from the stagnation point where the plume strikes the ceiling. i

I The parameter C corresponding to the sensor location (r/H) = 0.4 was taken j

'I as 0.025. Also the adiabatic wall temperature is taken to be T here (from

[ table 3-1), and T^ is the sensor temperature, T^. I

I

Unfortunately, the ceiling radiometer did not function satisfactorily in

j these tests, and therefore the convective and radiative components could not •

be resolved separately. However, the theoretical components (QR and QC) are

j computed and the total measured flux (QlEXP) is compared to the computed flux |

I (Q1 in figure 6-19) . Overall the computed flux is slightly higher than

I the measured values. It is significant to observe that for the cases of poor

agreement in figure 6-16, the computed ceiling flux is in good agreement with j HI values. The same radiative error does not occur here because the gas j temperature used (T) is the primary source of the radiative ceiling flux in '

128
Q

CO

FLUX TO CEILING TPRGET HI

TEST ID QlEXP

LJ/MC>XH<2 Q1

u/cm>K2

QC

U/Cri*!t<2 QR

U/CM>loX2

1 0 . 000 0.363 0.357 0.006

2 0.336 0.390 0.341 0.048

<s 3 0.774 0.833 0.723 0.111

4 1.210 1.608 1.291 0.317

mJ M 5 2.540 2.497 1.829 0.668

6 0.655 0.745 0.632 0.114

7 2.060 2.245 1.637 0.607

8 7.780 7^376 3.644 3.732

<s? 9 10.400 10.927 4.406 6.521

Q 10 0.810 1.137 0.942 0.195

CM _ 11 1.300 1.649 1.289 0.360

12 2.250 2.595 1.795 5.551

13 1.900 2.241 1.532 0.709

14 6.650 7.509 3.276 4.233

15 0.900 1.275 0.973 0.303

<S? 16 3. 120 3.324 2.003 1.321

17 1.950 3 . 057 l.!155 1.202

G? _

18 0.710 0.807 0.629 0.178

19 2.110 2.489 1.567 0.922

20 5.700 6.348 2.886 3.463

21 5.940 6.747 3.087 3.660

e

O

00

Q

CO

9

9

G?

ziQQ

CM

9

9

® A Q O

<S 0 . 00

“1

2.00

~i

4.00

__i I I I e .00 8.00 10.00 12.00

QlEXP

Figure 6-19. Comparison of calculated and measured incident flux to

ceiling sensor HI

129
these hot smoky fires. Although the comparison in figure 6-19 does not

directly confirm the accuracy of either the convective or the radiant model,

the agreement over the range of flux levels recorded does imply good accuracy

for both models.

6. 3. 5. 4 Conductive Ceiling Heat Flux

A model is needed to calculate the conductive heat flux into the heated

enclosure surfaces. Since a quasi-steady simulation model is being considered

for the average peak burning period, an average conductive flux must be

determined that reflects the past history of wall heating. This will be

approximated, although the transient conduction problem can be solved exactly

with probably the greatest accuracy of all the phenomena being considered.

This approximation is traded against solving the full transient problem

(which essentially would require a transient burning rate input along with the

transient conduction model)

For a constant surface temperature T^ for a semi-infinite solid (applicable

to our walls) Carslaw and Jaeger [26] gives the surface heat flux as

(kpc

6"

^x=o

- T^) W 3 (6-51)

yfrit

where k,p,c are the wall material conductivity, density, and specific heat,

respectively. The average heat flux over the time interval t* is

2 q" = —

^x=o

Ttt

_V.(kpc)

(T

w

(6-52a)

where t* is taken as the characteristic burn time

t* =

(1-oj ) M a c m

V

(6-52b)

where M is the mass of the crib

c and a is a constant selected as 0.5 such that t* is half of the nominal

burn time. This gives a t* of roughly 300 s, comparable to the duration of

the steady peak burn period.

130
The properties of Marinite XL, the lining material are given as:

c

k

P

w

w

w

=696 kg/m^

= (1.12 kJ/kg-K) [1 + 6.7 x lO""^ (T^ - T^)]

= (0.125 X 10^ kW/m-K) [1 + 6.7 x

10“'^ k“^ (T - T ). w a

The value of p and k [27] were measured, c was taken from the literature, w w w

and the temperature variation from Taminini [28].

A measure of the accuracy of this model can be assessed by considering

the heat flux balance at the ceiling position v/here thermocouple T8 measured

the surface temperature. At that position the conductive flux (q" or QK) must

equal the sum (QIN) of the convective (q^ or QC) and radiative (q^ or QR)

fluxes. Since these can be evaluated separately, it can be determined how

well

(or QK = QC + QR = QIN as shown in figure 6-20) at T8 . qj^ is determined from

equation (6-52) where T^ = T8. q^ is found from, equation (6-50) and from

equation (6-49) where T^ = T25 and T^ = T8 . The results are shown in figure

6-20. The results are not in good agreement; and since the models for q^

and q^ appear good, the convective flux model is weak. It could be fixed by

decreasing the parameter a to roughly, 0.1, but a lot of scatter would still

remain. The accuracy of the temperature measurements could have a bearing

on this result since the temperature differences are small compared to their

absolute values.

Perhaps the most significant variable in achieving accuracy in a mathematical fire model is the rate of energy release. This is implied by the

correlation given in equation (5-6) which displays the strong dependence of

upper layer temperature on energy release rate. For discrete room furnishing

items the rate of energy release could be adequately determined from an

apparatus similar to that shown in figure 4-3. However, that technique would

yield the free burn value, which does not account for the room effects.

(6-53)

6.3.6 Crib Mass Loss Rate

131
QIN

FLUX BPLRNCE PT POINT (TG) ON CEILING

CD

IJQ

CU

«S

<S? 9

<SCO 9

9

9 TEST ID QK QIN QC QR

U/Cri>tolc2 U/Cm>k2

G? 1 0. 152 0. 151 0. 193 -0.043

CD 9 2 0.031 -0.030 0.038 -0.068

<S “ 3 0,203 0.272 0.303 -0.031

® ^9 4 0.308 0.547 0.497 0.050

9 5 0.280 0.485 0.463 0.021

6 0.243 0.322 0.302 0.020

7 0.287 0.481 0.383 0.098

Q S 0.332 0.806 0.423 0.383 9 9 0. 170 0.492 0.231 0.261

G? " 10 0.341 0.576 0.510 0.066

11 0.321 0.570 0.518 0.052

9^ 12 0.315 0.709 0.523 B.156

13 0.432 0.969 0.617 0.352

9

14 0. 141 0.466 0.185 0.281

G? 15 0.318 0.577 0.469 0. 108

C\J 16 0.205 0.440 0.331 0. 109

G? 17 0.233 0.581 0.392 0.189

18 0.220 0.325 0.276 0.049

9 19 0.246 0.594 0.339 0.255

20 0. 199 0.851 0.276 0.575

21 0. 155 0.621 0.210 0.411

G?

G?

G? 1 1 1 1 1 1 0 .®0 ro 0.40 0.60 00 1 .00 1 .20

QK

Figure 6-20, Comparison of convective plus radiative heat flux into

the ceiling with conductive heat flux absorbed by the

ceiling

132
Recently Thomas et al. 129] has shown that flashover, when viewed as a burning

instability (not a remote ignition phenomenon )

,

requires an energy release

rate of no more than 50 percent above the free burn value at the onset of this

instability. In any case, the free burn value could serve as a basis for

describing the room burning rate.

For (wood) crib fires fundamental correlations exist (appendix A) which

can predict the free burn rate for a crib structure. This approach was not

adopted here because of the inability of these correlations to account for

burning rate enhancement in a burning assembly of cribs. This self-enhancement

discussed in section 5.1 was due to radiative interaction between cribs in

an assembly. Existing correlation could predict the burning rate of a single

crib, but not an assembly of cribs. Therefore an empirical analysis was

developed to account for the enhancement present in an assembly of like cribs.

If one observes a burning crib it is clear that the intensity of flaming

and pyrolysis is greatq,r for the internal sticks than the outer stick surfaces.

In fact at times the outer stick surfaces may have limited involvement.

Consequently an internal crib surface area (A^) was defined which could also

account for the crib faces adjacent to each other. The total exposed surface

area (A) of a single crib is

2 2 A = n N (4bP + 2b ) - 2b (N - 1) (6-54)

= 0.897 m^

where n, N, b, and £ are crib configuration parameters defined in section 2.

The internal area of a configuration of cribs is defined as

A. = 0.897 n - (Nb £) S (n ) - £^n (6-55)

1 c c c

where S (n^) is the number of crib sides exposed to open surroundings. In fact

A^ is the total area A minus the projected surface area of the sides exposed

to open surroundings; so A^ includes the exposed surfaces within a crib and

the adjacent surfaces between cribs. The free burn rate is then given as

133
ih (6-56a) v,o

ih"v,o

A.

1

where m" is the mass loss rate per unit area and is constant for each

material. In general,

^ depends on crib configuration and material, and

can be readily predicted for wood m^aterials by the analysis in appendix A.

Figure 6-21 shows this relationship plotted with the free burn data. The

slope for the correlation yields:

1

2 9.5 g/m -s, wood

(6-56b)

2 14.0 g/m -s, polyurethane

This establishes the free burn basis; the room effects remain to be considered.

Two effects of the room on the burning of an item will be the extra heat

flux received and the diminished oxygen available for combustion. The

following model is proposed:

m = i!i" A. + + V v,o 1 R

AitiOX

where

and

d" T

= —A m + 4"

o ,T ^r ,

L

vap

A

o , s

AltoxIt"v,oA.

1

(1

Y

ox , £ > 0.23 ’

(6-57a)

(6-57b)

(6-57c)

Alt represents the gasification rate due to heat flux from the room

R

to the top surface (A „) of the crib, (5" m; and the exposed sides (A ) of

0,1 ^^1 0,0

the crib,

^o,T

" ^T^^c^(6-57d)

A = bN£ S (n ) + £s G (n ) (6-57e)

o , s c s c

where G^(n,c) is the number of top gaps between adjacent cribs

and G^(n^) is the number of side gaps between adjacent cribs. The projected

areas facing the surroundings are used here.

134
(s/3| 3iVd SS01 SSVW Ndna 33dd

0) +J

u

3

•H

Em

135

CRIB

CONFIGURATION

INTERNAL

AREA,

Aj

(m^j
An effective heat of vaporization was derived for each fuel material.

This was done using an irradiation source to pyrolyze wood and polyurethane

in air while measuring the mass loss. The rate of mass loss was surprisingly

constant over a wide range of mass fraction remaining. The results are

shown in table 6-9. Nominal values of

L

vap

5 kJ/g, wood

\S kJ/g, polyurethane

were used in the computations presented here.

(6-58)

These values are nominal averages of the results shown in table 6-9 which

neglected fuel surface heat loss effects. Neglection of the heat loss in the

determination of L was an error discovered during the preparation of this

report. Assuming a linear relationship between m^ and qV (as given in table

6-a) would yield

qV - q"

vap m'V

where q^ is the surface heat loss. For the vertical samples pyrolyzed in air,

the data suggest L =1.9 kJ/g for wood (sugar pine), and L =2.9 kJ/g

V3.p V3.p

for the rigid polyurethane (GM-37). The value q^ was the same for both

materials, and corresponds to a radiative loss from a surface at 500 °C. This

compares v/ith Tewar son's 130] values of 1.82 kJ/g for Douglas fir and 4.52

kJ/g for polyurethane (GM-37) pyrolyzed in nitrogen.

Furthermore, in the analyses to follow (figures 6-22 and 6-23)

these

"more correct" values of produced changes of within 10 percent of the

current calculated theoretical values.

To complete this model the radiant fluxes to the crib surfaces must

be presented. The methodology of Modak and Matthews [23] were used. For

simplicity the area elements were considered centered with respect to the

four upper quadrants (i.e., ~ “ (2.18/2)m for i = 1,4). The top element

was at the crib height (H^ + 0.18 m) and the side element was at the mid

position (H^/2 + 0.18 m) . The sum of the flux from the four quadrants is

136
Table 6-9

Pyrolysis Due to Radiation

Fuel Orientation Incident

flux

m" L V vap1 **

W/cm^ g/cm -s J/g

Wood vertical 3.5 0.80x10"^ 4387

Wood vertical 3.0 0.527x10"^ 5688

Wood horizontal 2.5 0.54x10"^ 4610

Wood horizontal 2.5 0.544x10"^

-3

4595

Polyurethane vertical 3.5 0.50x10

-3

7030

Polyurethane vertical 3.0 0.343x10 8755

Polyurethane horizontal 2.0 (ignited after

4 minutes)

*Data by W. Twilley 131] .

** This evaluation neglected heat loss at the fuel surface.

137
I

given as

,T = 4 i FuT

4 , 4

E oT + (1 - e ) oT

gc g,u gc w,u

where

uT —1 - a . , -1 / a (^) tan — a \ a

£T

1_

2 TT

H. V

— - 2 (l-> tan (j^) a a J

2 2 + a

gc

me

D - (H + 0.18 m) > 0 c —

, -k Lmc

1 - e g

2a (H-D)

2 (H-D) + a

(6-59a)

(6-59b)

(6-59c)

(6-59d)

(6-59e)

(6-59f)

(6-59g)

and H„ = H - (H + 0.18 m)

A" = F

, s uc

4 4

e aT + (1-e ) aT

gc g,u gc w,u

(6-59h)

+ + V2)aT;_^ (6-60a)

where this is the sum of the upper and lower half-space contributions, and

H.

, ^

1 uc IT

ZC TT tan-1 (g-

(6-60b)

(6-60c)

with L^, H, and e^^ defined as in equation (6-59) .

Finally the rationale for the formulation of the oxygen effect will be

discussed. In the combustion of simple systems, the burning rate due to

convective effects can be described in terms of the Spalding mass transfer

number B , i . e .

K = (y It + 1>

where h is the convective heat transfer coefficient and

Cg is the specific heat of the gases.

(6-61a)

and B =

Y _ AH/r - c (T - T ) o2 ox g s a L

vap

(6-61b)

138

1
Hence, for a given fuel B is directly proportional to Y- - in this case the

oxygen concentration in the lower space, . Kanury [32] demonstrated that

^2

for charring materials, equation (6-61) is not sufficient to describe this

burning phenomena unless B is defined to account for additional effects such

as stick thickness and radiant intensity. Hence, it would be reaching to

utilize equation (6-61) for the crib burning rate.

However, it is desirable to keep the analysis simple until a more definite

approach develops. Moreover a study of crib burning in diluted atmospheres

[33] indicated that, in oxygen concentrations down to 17 percent (by volume) , the

rate of burning was proportional to oxygen concentration. Since for small

B-number, the Spalding relationship in equation (6-61) would also suggest

direct proportionality to Y„ , then is assumed that

^2

Y

Am = (m ) free D (6-62)

ox V 0.23

The accuracy of this model was examined by substituting experimental

values of m , m , D, T = T and T = (T8 + T25)/2 (with T . = T . = T ) v a g,u w,u ' g , Z w,2- a

for table 3-1 into formulae (6-57) through (6-62).

4 The results are given in table 6-10 and the calculated and measured

values of m^ are compared in figure 6-22. The results in table 6-10 show that

the enhanced radiation effect appears to be comparably counteracted by the

oxygen effect. For these data however, both effects are not more than 20

percent of the mass loss rate. Hence, the room effects are not great. This

is consistent with the discussion of experimental results in section 5.1.

139
Table

6-10.

Evaluation

of

crib

mss

loss

model

cr» U\ n to CTl ru in VO LO r^ VO t'J ro in CO in CD

ro UT o ro ro (M CD VO ro V CM O 00 rw r^ in (M 'T Q O ro VO o CM VO ro CO ro r- CM in •V G> V CM ro

CD O (D G3 o S) G) — G} o O G) CM GO G) CO CM (M

•E

<\

X

VO CM ro VO cn VO G) ro CD rv- ^D r- cn CM CD CM ro in VO IV

G) G! CM CM CM G) hw CO •tr CD ro CM in CD •sr G) 0 X CO G1 G) G) CM GO CM CM IM. CD CM In. cn in vr CO ro ro CM G> VO

CD \ « C3 G) G) G) GD G) G) G) G) G> CD GD GD G» —

*

CO ro ro GD CM <T

I I I I I I I I ) I I I I I I i I I I

_l in CM W ‘A in CD in rv VO CO ro tv ro «»« ro un in CD CM CD a CD CD r- ro CM ro o *-4 CD CM VO CM CD in VO CM in CM OD

CJ CO CD cn VO ro CO CM CD r- IV CO VO VO r^ ro «' c IV VO ro iv CM > \ • IZ (J •vf •V •—« CD rv VO VO rv GD 00 VO in VO LO •V VO V to ro

^-4 CM CM CM -r-l CM w^ CM CM CM CM

I

I

I

I

I

I

•E

(U m i/) > r U

CD GD G G) G) G) CD CD G) CD CD CD GD GD GD (D GD CD G CD G)

CO CD VC •vj- CO cn CM cn •V to 00 CM C»J tv CD Sj in tv ro

CD CO VO cn CD CM cn tv CM v GD CM •vf 'S' m in cn cn

in in GD CO VO GD ro in CD OD LO in vr G> in vf CM

«—

t

CM CM to ro CM w < CM r«i CM CM CM

CM

Q- « CD vr vr to cn cn CO CO CD tv tv to VO in VO tv in VO cn VO vO vj in GD VO to VO ro 00 vr CM cn 00 o ViD GD 00 CM w4 CM CM

GD 1 to VO in M- in to CD VO to ro to CM CM cn 'vr VO

- ct: \ 0 3 CD GD GD CD GD CD GD to VO CD CD ID Q •vr Q - GD CD to to

cn ui CM 'O * tv VO CM to to VO CM CM tn GD IV in cn CD 'vr CM ro CD IV in

•vr in tv tvj CD VO cn tv tv CO ro CD cn -vr CM cn rv cn CD in

= to r CD G CD CM CD tn CM O CM CM cn *-• in in CD vr in VO

•tr O' \ • C3 3 (S3 G G CD GD GD CD G CM CD CD GD G GD GD GD CD CD

GD CD cn tv cn cn tv 00 VO CD tv to |v VO GD 4—4 03 rv VT cn

_J CD to to CM CM CM CM CM CM CM to CM CM m 'H CM CD CD »-< tn tn CO X \ CM CM CM CM CM CM CM CM CM CM CM CM CM CM CM CM CM CM O CD • GD GD GD GD GD GD GD GD GD GD GD GD GD CD GD CD GD GD GD CD 1

CJ CM (VI TT •vr vf 'vr

i

^ 1 CJ \ \ \ \ \ \ s \ s s s

\ w-i 4-4 —44 4—4 •-H 4-4 *-• y4

1 1 1 1 1 1 1 1 1 1 1

1 1 1 1 1 1 1 1 1 to •Vf CM to CM •vr vj- CM to ro 1 O ^-4 CM to Tf CM to vr CM 1 1 1 1 1

1 1 1 1 1 1

1 1 1 1 1 1 1 1 3 3 CL CL 3 3 3 CL CL CL Q. 3 3 3 3 3 CL CL Cl Q. 3 1 1 1 1 1 1 1 1 1 1 1 : H- 1 1 1 1 1 1 1 1 1 in GD CO to tv cn CM vr VO CD I in CD to in CM tv •VJ VO 00 cn 4—1 « 4 CM 1 U tJ t_D U tJ C_D CD U u CJ u u U u U u CJ CJ tJ CJ 0 : a: QC a: oc DC O' (^ O' O'. O' O' ni D^ oc Qt OiC O' CL ct: 0. Q. Q. CL Q. CL CL U. CL CL CL CL CL CL (L (L CL CL CL (X. a.

1

140
MV

(CRLC)

G/S

20.0

25.0

30.0

35.0

40

.

ENCLOSURE EFFECT ON BURNING RPTE

6

e

e 9

<s?

LO _

<s? ® _

Q

LD

<S? “I 1 1 10.0 15.0 20.0

MV (EXP) G/S

0.0 5.0 25.0 30 )

Figure 6-22. Comparison of calculated and measured rate of mass loss

in the room fire experiments

141
6.4 Simulation Model

The formulation of the governing equations essentially has been made in

' the previous sections. However, some modification must be made in the

radiative exchanges since the analysis of section 6.3.5 dealt with an exchange

with respect to a target elemental area, not a finite surface area. Following

the methodology given in Siegel and Howell [24] and the approach illustrated

in section 6. 3. 5. 3 yields the necessary radiative exchange. The details of

this will be omitted here; only the formulation will be listed. Basically

it will consist of seven nonlinear algebraic equations with subsidiary

relationships. The reduced set of unknowns (x^) are:

x^ = N, the height of the neutral plane

X

2 = D, the height of the thermal discontinuity

x^ = ^

^ '*^he upper layer gas temperature

x^ “ u' upper layer surface temperature

x^ = Tg the lower layer gas temperature

X,- = T . , the lower layer surface temperature

D w, £ and x^ = m^, the rate of mass loss.

The equations have the formi

f . (x. ) = 0

1 1 (6-63)

A mass balance on the plume and upper layer yields:

f, = m ,-m - in = 0

1 out a V

(6-64)

where (the doorway flow out) is given by equation (6-10) and m^ (the

air flow) by equation (6-llb)

.

A mass balance on the lower layer yields

:

f^=ih - ih - m = 0

2 p a e (6-65)

where m (the crib fire entrainment rate) is given by equation (6-25b) and

P

(the entrainment rate from the upper to the lower layer) is taken from

142
equation (6-17) with n = 1/4 and = 0.5.

The energy balance on the upper layer and flame plume is the algebraic

sum of the energy release (E ) , the rate of energy losses by mass transport

(Ej^) , and heat transport (E^) . q

= “

-

2

>

( 6 - 66 )

(6-67)

where half the energy radiated from the flame is assumed to be absorbed by

the upper layer. Also, the rate of fuel reacted within the enclosure is

m , m > rm V a — V

m / m < rm a/r , a v

( 6 - 68 )

and the net rate of energy transport by fluid flow from the enclosure is

E- = (m + m + m ) c (T - T J - m c (T* - T J (6-69) M a V e p g,u g,Ji v p

where the last term, due to pyrolysis gases emerging from the surface at

temperature T* , will be neglected.

The net rate of heat transfer from the upper layer and flame plume is

E^ = (q" + q" ) A + q" A + q" + df" ^ A^ L ^r,u ^c,u w,u ^r,o w,u ^o,u -"r,D F

(6-70)

The upper surface area is

A = WL + 2 (W + L) (H - D) - A

w,u o,u

(6-71)

The upper portion of the doorway opening is

A = W (H - D) o,u o o

(6-72)

The floor or layer surface area is

Ap = WL (6-73)

143
The radiative flux to the upper surface is

4" = E a + Y (1 - e ) - [1 - (l-y )(l-e ) ] oT'^ ^r,u g g,u 'u q' \7 ,

1

' ’u q'

w,u

where Y = A„/A

u F w,u

(6-74a)

(6-74b)

and Eg is given by equation (6-49b) with from equation (6-41) . The

radiative flux lost through the opening is

4 4 4 d" = e oT + (1-e ) e T - aT ^r,o gg^u g w,u a (6-75)

The net flux through the layer interface from the upper to the lower space

is

A A A (6-76)

4 4 4

c^'' ^r,D = e oT + (1-e ) oT^ - T „ g g,u g w,u w,£

The convective flux to the upper surfaces is

(5" = h (T - T ) ^c,u c,u g,u w,u

(6-77)

where h^ ^

is given by equation (6-50) with C = 0.01 based on an average

convective coefficient. The convective flux to the side walls has not been

adequately quantified, so C is based on an "extended" ceiling. An energy

balance on the lower layer is

A

^4 "^g, £~'^a (a + a

't - T \ (T

g,u *'a ) (A^ + A^) ^ q , !i) = 0 (6-78) ^a e/\^' / P'a e where the gas is not considered to be absorbing or emitting, and the convective

(last) term was dropped in the final calculation.

Also energy balances at the upper and lower gas-surface interfaces yield

two equations.

fr = qi - 4

" - 4

" =0 5 ^k,u ^r,u ^c,u

where d," is given by equation (6-52) with T = T = T,

•^k,u www,u

and

f^ = d" - n ~ g" n 0

6 ^k,£ ^r , £ ^c ,

£

(6-79)

(6-80a)

144
where q" is given by equation (6-52) with T = T

jC ^3c W W f 6". = h „ (T „ - T „) with h „ estimated as 0.01 KW/m‘‘ ^c,l c,l g,z w,Jl c,l ' - K

d" = Y

4 4 4

e a + (1 - e ) oT - aT „ g g,u g w,u w,£ + 7=r AH

2 A

(6-80b)

(6-80c)

w, £

^ £ ^F^'^w , £ (6-80d)

and A „=2(W+L)D+WL w, £

Finally, the seventh equation is developed from equation (6-57)

Y,

(6-80e)

f_ = A - m" A.

7 V v,o 1

02 , £ +d" A „ ^r ,T o , T ^r , s o , S _ Q

^vap 0.23

The necessary subsidiary equations (6-57 through 6-60) also apply,

(6-81)

The resultant set of nonlinear algebraic equations to be solved are

^1 (x^. X2, X3, ^

4

' X5, ^6' x^)

^2 (Xj^, X2, X3, X4, X7)

^3 (Xj^, X2, X3, X4, X5, ^6' X,^)

£4 (x^. X2, X3, X4)

^5 (x^. X2, X3, X4, X5, ^6' X^)

*6 (x^. X2, X3, X4, X5, ^6'

^7 (x^. X2, X3, X4, X5, ^6'

= 0

= 0

(6-82) = 0

= 0

= 0

= 0

= 0

This set of equations was solved by a Newton-Raphson method. Problems

of convergence do occur with this current algorithm so that a complete set of

calculational results were not achieved for all conditions desired. A summary

of the input parameters used in the basic theoretical calculations is shown

in table 6-11.

145
Table 6-11

Input Parameters Used in the Theoretical Calculations

Parameter Plastic Wood

r (g air/g volatile fuel) 9.03 4.92

AH (kJ/g volatile fuel) 23. 15.

f (kJ radiation/kJ total energy 0.22 0.18

L (kJ incident flux/g volatile fuel) 8. 5.

vap

to (g char/g solid fuel) 0.13 0.10

, -1

k^ (m ) 1.3 0.80

T^ (K) 1200. 1100.

(kg crib) 3.1 3.4

C . 0.58 1

—

o cn 1 1 W (m) in o

C ^ out 0.68

k 0.5

m

C 0.01

X 1. a

a 0.5

^c,£ (kW/m^-K) 0.01

146
!|

I

t 6.5 Comparison of Experimental and Theoretical Results

The results of the theoretical simulation model were compared to the

average peak results of the enclosure fire experiments. The theoretical

i results were calculated over a range of doorway widths and crib loadings of

I up to four or more. The combination of initial conditions for which results

i are available depend, at this time, on the program's ability to converge to a

I

I solution. The same symbol legend used in section 5 for the data points applies

I here as well. The format of the comparison will be the variable for comparison

j versus door width with the number of cribs as a parameter.

!

I The rate of mass loss for both the wood and plastic cribs is presented

'l

I in figure 6-23. The constancy of the theoretical results is in contrast to

I the actual curvature as the data points imply. This discrepancy is probably

I due, on one hand, to an inadequate prediction of the radiative enhancement

effect for large door sizes; and, on the other hand, an inadequate treatment

of the diminution of mass loss rate with oxygen depletion in the lower layer.

j Also calculations for small door sizes, which would lead to results with

I

I downward curvature, have not been successfully run. Despite this inability

to match the curvature of the experimental results, the theoretical results

are within 25 percent of the measured mass loss rates.

The results for the rate of airflow into the enclosure are shown in

figures 6-24a and 6-24b for the plastic and wood respectively. Here both

the inflow rate and airflow rate derived from the outflow measurement were

plotted. The scatter indicates the accuracy of these data. The theoretical

j result gives an impression of order to the data and may indeed be more accurate,

i Of course, until better data is achieved a conclusion statement of the

theoretical accuracy cannot be made. It should be noted that a reduction in

the outflow coefficient reduced the computed rate of airflow by a proportional

magnitude

i

1

I 147

i
WOOD

and

PLASTIC

Figure

6-23.

Comparison

of

experimental

and

theoretical

rates

of
The (velocity) neutral plane heights and thermal discontinuity heights

are shown in figures (6-25a and b) . Realizing, especially, that the thermal

discontinuity is a nebulous quantity, the theory appears reasonably adequate.

The next two figures (6-20a and b) examine upper layer gas temperature.

The lack of a unique layer temperature or an imposed procedure to derive a bulk temperature for the gases leaves a somewhat arbitrary choice of comparing

experimental results to the theory. To confront this issue, the average

ceiling layer temperature T from table 3-1 was plotted along with a bar tab

indicating the temperature of the gases leaving at the top of the doorway.

These differences are discussed in volume II and are clearly shown here. With

these considerations for room temperature, the theoretical result, T^ may be

judged as in fair agreement with the data. It should be noted that for the

polyurethane case a reduction in the wall conductive parameter a from 0.5 to

0.1 reduced the computed T by 10 percent change in AH from kJ/g to 19 kJ/g,

g , u reduced T^ ^ by 15 percent, and a change in both a and AH yielded about 25

percent reduction in T^ All of these changes would only influence the

computed airflow rate by 3 percent at most.

In figures 6-27a and b the volumetric oxygen concentrations are plotted

(dry basis) for the upper doorway probe. If corrected for water vapor, the

data would be lower; if corrected for the vertical distribution in the doorway,

the data would be higher than the average value which more correctly corresponds

to the predicted value; hence, these are compensating corrections. Thus, the

experimental values shown are probably close to a mixed-mean value, and are

consistent with the theoretical quantity, Xq The theoretical results

appear to be in fair agreement with these data. An adjustment in the stoichiometric air-fuel ratio (r) could lead to improved results. A reduction in r of

about 25 percent was suggested by analysis of the experimental results but this

was not made in the simulation model computations.

149
PLASTIC

oo CO ir> ^

CO

C3 o o

(s/3)i| aivti M01J m

CM

150

Figure

6-24a.

Comparison

of

experimental

and

theoretical

airflow

rates

-

plastic

fires
WOOD

J_

CO ir» CO

|s/3)|| 31VU AA01J illV

151

Figure

6-24b.

Comparison

of

experimental

and

theoretical

airflow

rates

~

wood

fires
TAG

UP

-

Neutral

plane

TAG

DOWN

-

Thermal

discontinuity

L to

|ui) 3AOaV 33NV1SIQ

152

Figure

6-25a.

Comparison

of

experimental

and

theoretical

neutral

plane

and

thermal

layer

heights

-

plastic

fires
TAG

UP

-

Neutral

plane

TAG

DOWN

-

Thermal

discontinuity

L

cq

|UJ| U001J 3A09V 33NV1SIQ

JL

CVI

153

Figure

6-25b.

Comparison

of

experimental

and

theoretical

neutral

plane

and

thermal

layer

heights

-

wood

fires
PLASTIC

o

GO

i

Ool 3dniVd3dN3i SV9 d3AV1 d3ddn

I

154

Figure

6-26a,

Comparison

of

experimental

and

theoretical

upper

layer

temperature

-

plastic

fires
WOOD

155

Figure

6-26b.

Comparison

of

experimental

and

theoretical

upper

layer

temperature

-

wood

fires
PLASTIC

156
WOOD

l%l N0liVtliN33N03 3WmOA N39AX0 ttSAVI U3ddn

157

Figure

6-27b.

Comparison

of

experimental

and

theoretical

upper

layer

oxygen

concentration

-

wood

fires
A measure of the accuracy of the calculated lower layer gas temperature

is given in figure 6-28. Again, the existence of a non-unique experimental

lower layer gas temperature makes a comparison with the theoretical variable

T a bit tenuous. Despite some ambiguity in this interpretaion , however,

g ,

Jc the theory appears to be close to the mark.

The incident flux to the floor target (H2) was also derived from the

model. For this calculation, the analysis of section 6. 3. 5. 2 applies with

g g,u' c w,u' a b w,Z. Because of the strong dependence of

flux on room temperature, the variation of temperature in the room, and the

smoke blockage effects in the lower layer, this prediction lacks some necessary

refinements. Nevertheless its results are acceptable. The comparisons are

shown in figures 6-29a and b.

Another viewpoint is to consider the incident floor target flux a function

of the rate of mass loss. A comparison on this plot relaxes the burden on

the theoretical model of predicting the mass loss rate. This is shown in

figure 6-30.

The results presented in figure 6-30 have practical utility. They can

serve to indicate whether flashover by remote ignition is likely to occur for

2 a given fire. A nominal ignition criterion of 2 W/cm will be used to make

this assessment. Furthermore, additional computations were done for larger

room sizes than the experimental test room. The critical mass loss rate at

2 which a flux of 2 W/cm would be reached at a floor position (given by the

distance of sensor H2 from the fire source as in the experiments) was determined as a function of room size and doorway width. These results are shown

in figure 6-31.

These results suggest that, with respect to the hazard of remote ignition

of adjacent items in a room, the rigid polyurethane foam material is more

hazardous than the sugar pine wood. Figure 6-31 indicates that for equal room

and doorway sizes roughly 50 percent more wood than polyurethane must gasify

to cause remote ignition in a room. Since the burning rate per unit area is

158
WOOD

and

PLASTIC

DATA:

Aspirated

probe

(corrected)

or

the

mean

of

T22

and

T23

corrected

05

CO

oo

CO

CD

CD

UD

Ool 3dniVd3dN31 SV9 U3AV1 d3AA01

CO CN

159

Figure

6-28,

Comparison

of

experimental

and

theoretical

lower

layer

temperature
PLASTIC

Figure

6-29a.

Comparison

of

experimental

and

theoretical

incident

floor

fluxplastic

fires
WOOD

CO

I;iii3/M| 1N3W313 U0013 01 XflU 1N30I3NI

161

Figure

6-29b.

Comparison

of

experimental

and

theoretical

incident

floor

flux

-

wood

fires
PLASTIC WOOD

r

in

CO CO

ip

csi inCM

in

3UI3/MI 1N3IM313 UOOU 01 XflU 1N30I3NI

Figure

6-30,

Comparison

of

experimental

and

theoretical

incident

floor

heat

flux

as

a

function

of

wood

and

plastic

rate

of

mass

loss
(s/3| 31V8 SS01 SSVW 1V3I1IU3 a3lV1fl31V3

(U

163

Figure

6-31.

Calculated

critical

rates

of

mass

loss

for

various

siz

room

floor

areas

with

2.41

m

high

ceiling

and

a

door

1.83

m

high

and

various

doorway

widths

(also

the

input

parameters

of

table

6-11

apply)
I

about 50 percent higher for the plastic cribs than the wood cribs, about twice asj

much wood as plastic must be burning to produce an equivalent hazard. That

•'1 is, about twice the surface area of fire involvement is required by wood to

| cause remote ignition than is required by the polyurethane. This follows since

j I (^v,o) PU ^

: (m" ) . .i V , o wood 'i

i-'i and

^*v,crit^ wood/' (m . ) = 1.5 v,crit PU 1

Therefore, the fire area, , has the ratio

^f,wood 2.25

A ~

f ,PU

to cause equivalent room heat flux. Thus more fire involvement is required

for wood furnishings than for plastic to cause flashover in the same room.

i

I

-.'I

6.6 Empirical Results of Room Furniture Fires

The present study applies to discrete fires in a room without the

consideration of combustible lining materials and their effect on fire growth.

Although the fuel element examined was cribs, it does bear some similarity

to furniture assemblies. Moreover the range of burning rates considered and

the extrapolation by analysis to other conditions does produce a relevant

ensemble of results. This will be put into perspective by considering the

results from a series of furniture fire studies found in the literature [34-41]

1

I

i

1

In a previous review of room fires [1] a tabulation was prepared of mass

loss rates at the onset of perceived room flashover or the peak mass loss

rates without flashover for a series of furniture items. It should be realized that those results were derived from experimental fires of rooms perhaps

slightly larger than that of the present study and usually with one conventional size opening. In this tabulation no segregation due to room geometry

or fuel composition was made. Moreover, since flashover did not always occur

at any particular mass loss rate, a frequency of flashover was determined.

These results are shown in figure 6-32.

164
DISTRIBUTION

OF

MAXIMUM

MASS

LOSS

RATES

AND

FLASHOVER

FREQUENCY

IN

ROOM

FIRE

EXPERIMENTS

d3A0HSV1d JO AQNJnDJdJ

GO CO C^J

(S/s) SN31I ONINUnS JO A1ISN3Q Ainiavsoud

u

•H

Pm

165

MAXIMUM

MASS

LOSS

RATES

DURING

BURNING

(g/sj

7. CONCLUSIONS

The primary goal of this study has been to make an assessment of the

potential fire hazard of high density polyurethane foam with respect to its

use in room furnishings. The study contrasted the characteristics of wood

and plastic fires in a room environment. Based on these results and their

analyses the following conclusions were drawn.

1. With respect to the potential for flashover by remote ignition, the

plastic material (rigid polyurethane, GM37) is roughly twice as prone to

cause flashover as wood (sugar pine) . This is based on the critical fuel

surface area burning in a given room; it does not account for the rate of

spread over the fuel. By analysis this conclusion can be traced to the

difference in heats of combustion and the rates of gasification of burning

between the two materials.

2. For these crib experiments, the effects of the room environment on the

fire were of secondary significance. Enhancement by radiation feedback and

reduction in burning due to oxygen deficiency appeared to be of the order

of +25 percent at most. This is likely to be true for some furniture items,

but especially not the case in which flame spread on horizontal surfaces

is a factor.

3. Although the stoichiometric air requirements are distinctly different

for the two fuels, room ventilation was not a significant factor in controlling

room heating over the range of doorway widths considered. The empirical

correlation of equation (56) clearly shows that room temperature is a strong function of energy release rate, but a weak function of the ventilation factor. In fact, for this study, the theoretical calculations indicate

a significant diminution of the mass loss rate only below doorway widths of

0.2 m. Thus, it is likely that ventilation will have a significant effect

on reducing the fuel gasification rate only for very small room openings.

While this ventilation controlled (small fires) fire regime may not be a hazard from a fire growth point of view, it will tend to increase the

production of CO. Igy
4. The development of the theoretical model and selective comparisons

with the experimental results provide an assessment of the model's accuracy.

The two-layer model is an approximation of reality and hence will. never be

perfectly consistent with experimental results. It is likely, however,

that the imperfections of the model are becoming comparable with the

uncertainty of some experimental measurements.

8 . ACKNOWLEDGEMENTS

The authors are especially indebted to Margaret Harkleroad and William

Rinkinen for their assistance. Mrs. Harkleroad prepared computer programs

for most of the data analyses, developed the computer plotting routines

used in this report, and maintained a library of programs used in this

study. Mr. Rinkinen coordinated the design, construction, and allocation

of instr\imentation in the experimental study. He designed a reflecting

globe heat flux calibration device that enabled a quick check of the room

mounted sensors. He maintained drawings, plans, and was responsible for

the video records of each experiment.

Also, we wish to express our appreciation to the Products Research

Committee (PRC) for their financial support which enabled this study to be

carried out.

Finally, we would like to note the contributions of Robbin Annunzi,

who typed the manuscript, and Betty Oravec, who edited the manuscript for

publication

.

168
9 . REFERENCES

[1] Quintiere, J. G. , Growth of fire in building compartments. Fire

Standards and Safety, ASTM STP 614, 131-167 (1977).

[2] Quintiere, J. G. , McCaffrey, B. J. and Den Braven, K. , Experimental and

theoretical analysis of quasi-steady small-scale enclosure fires,

Nat. Bur. Stand. (U.S.), NBSIR 78-1511 (October 1978).

[3] Emmons, H. W. , Mitler, H. E. and Trefethen, L. N. , Computer Fire Code

III, Home Fire Project Tech. Rept. No. 25, Harvard (1978).

[4] Pape, R. and Waterman, T. , Modifications to the RFIRES pre-flashover room

fire computer model. IITRI Project J6400, IIT Research Institute,

Chicago, 111. (March 1977).

[5] MacArthur, C. D. and Myers, J. F. , Dayton aircraft cabin fire model

validation. Phase 1, Rept. No. FAA-RD-78-57 , University of Dayton

Research Institute, Dayton, Ohio (March 1978).

[6] Smith, E. E. and Heibel, J. T. , Evaluation performance of cellular

plastics in fire systems. PRC Proj. No. RP-75-1-36, Ohio State

University, Columbus, Ohio (1977?).

[7] Gross, D. , Experiments of the burning of cross piles of woods. Journal

of Research, National Bureau of Standards, Vol. 66 C, No. 2, 99-105

(April-June 1962)

[8] Block, J. A. , A theoretical and experimental study of non-propagating

free-burning fires, 13th Symposium (International) on Combustion, The

Combustion Institute, 971-978 (1971).

[9] Sensenig, D. L. , New concept for determining rate of heat release by

oxygen consiomption , presented at the Fall Technical Meeting, Eastern

Section of the Combustion Institute, Miami Beach, Fla. (Nov. 1978) (To

be published)

[10] Croce, P. A., Modeling vented enclosure fire Part 1. Quasi-steady woodcrib source fires, FMRC J. I. 7AOR5.GU, Factory Mutual Research Corp.

,

Norwood, MA, (July 1978).

[11] McCaffrey, B. J. and Rockett, J. A., Nat. Bur. Stand. (U.S.), J. Res.,

Vol. 82, No. 2, 107-117 (1977).

[12] Quintiere, J. G. and Den Braven, K. , Some theoretical aspects of fire

induced flows through doorways in a room-corridor scale model, NBSIR

78-1512, Nat. Bur. Stan., Washington, D.C. (Oct. 1978).

[13] Prahl, J. and Emmons, H. W. , Combustion and Flame, Vol. 25, No. 3.,

369-385, (Dec. 1975).

[14] Shivadev, U. K. and Emmons, H. W. , Combustion a.id Flame, Vol. 22,

223-236 (1974).

[15] Birky, M. , Private communication, 1978.

[16] Smith, R. K. , Morton, B. R. and Leslie, L. M. , J. Fluid Mech. , Vol. 68,

Part 1, 1-19 (1975)

169
[17] Quintiere, J. G. , McCaffrey, B. J. and Rinkinen, W. J. , Fire and Material

Vol. 2, No. 1, 18-24 (1978).

[18] McCaffrey, B. J. , Purely buoyant diffusion flames: some experimental

results, Nat. Bur. Stand., (U.S.,) NBSIR 79-1910 (Oct. 1979).

[19] Dayan, A. and Tien, C. L. , Combustion Science and Technology, Vol. 9.

41-47 (1974).

[20] Markstein, G. , Radiative properties of plastic pool fires, FMRC J.I.

7AOR3.GU, Factory Mutual Research Corp. , Norwood, Mass. (July 1978).

[21] Hagglund, B. and Persson, L. E., An experimental study of radiation from

wood flames, FOA 4 Rapport, C 4589-D6 (A3), Forsvarets Forskningsanstalt,

Stockholm, Sweden (June 1974).

[22] Modak, A. T., Radiation from products of combustion, FMRC J.I. OAOE6.BU-1

Factory Mutual Research Corp. Norwood, Mass. (Oct. 1978).

[23] Modak, A. T. and Matthews, M. K. Radiation augmented fires with enclosures, Tech. Rept. 22361-8, Factory Mutual Research Corp., Norwood, Mass.

(Feb. 1977).

[24] Siegel, R. and Howell, J. R. , Thermal radiation heat transfer, Vol. Ill,

NASA SP-164, National Aeronautics and Space Administration, Washington,

D.C. (1971).

[25] Zukoski, E. E. and Kubota, T. , An experimental investigation of the heat

transfer from a buoyant gas plume to a horizontal ceiling - part 2.

Effects of ceiling layer, Nat. Bur. Stand. (U.S.), NBS-GCR-77-98

,

NBS Grant No. 5-9004 (1975).

[26] Carslaw, H. S. and Jaeger, J. C., Conduction of heat in solids, 2nd Ed.

Oxford University Press, London (1959).

[27] Siu, C. I., Personal communication (1978).

[28] Tamanini, F. , Personal communication (1977).

[29] Thomas, P. H., Bullen, M. L., Quintiere, J. G. and McCaffrey, B. J.,

Flashover and instabilities in fire behavior. Combustion and Flame,

Vol. 38, 159-171 (1980) .

[30] Tewarson, A., Experimental evaluation of flammabilty parameters of

polymeric materials, FMRC J.I. 1A6R1. PRC Grant No. RP-75-1-33A,

Factory Mutual Research Corp., Norwood, Mass. (Feb. 1979).

[31] Twilley, W. , Personal communications (1978).

[32] Kanury, A. M. , Rate of charring combustion in a fire, 14th Symposiiom

(International) on Combustion, The Combustion Institute, 1131-1142

(1973)

.

[33] Durak, 0. T. , Combustion of wood cribs in diluted atmosphere, NBS Grant

No. G5-9009, Northwestern University, Evanston, 111. (June 1976).

[34] Waterman, T. E., Determination of fire conditions supporting room

flashover. Final Rept. IITRI Project M6131, DASA 1886, Defense Atomic

Support Agency, Washington, D.C. (Sept. 1966).

[35] Croce, P. A. and Emmons, H. W. , The large-scale bedroom fire test,

July 11, 1973, FMRC Ser. No. 21011.4, Factory Mutual Research Corp.,

Norwood, Mass. (July 1974).

170
[36] Croce, P. A., A study of room fire development: the second full-scale

bedroom fire test of the home fire project (July 24, 1974), Factory

Mutual Research Corp. , Norwood, Mass. (June 1975).

[37] Hagglund, B. , Jansson, R. and Onnermark, B. , Fire development in residential rooms after ignition from nuclear explosions, FOA Rep. C 20016-D6

(A3), Forsvarets Forskningsanstalt , Stockholm, Sweden (Nov. 1974).

[38] Gross, D. and Fang, J. B. , The definition of a low intensity fire,

Nat. Bur. Stand. (U.S.), Spec. Publ. 361, Vol. 1 (March 1972).

[39] Fang, J. B. , Measurement of the behavior of incidental fires in a compartment, Nat. Bur. Stand. (U.S.), NBSIR 75-679 (March 1975).

[40] Theobald, C. R. , The critical distance for ignition from some items

of furniture, FR Note 736, Fire Research Station, Borehamwood, England

(Dec. 1968).

[41] Christian, W. J. and Waterman, T. E., Fire Technology, Vol. 7, No. 3,

205-217, (Aug. 1971).

171
APPENDIX A. GASIFICATION RATE OF FREE BURNING CRIB FIRES

A series of wood and plastic crib fire experiments were conducted in

order to design a crib for the room fire experiments. The burning rate of

wood crib fires has been studied by Gross [7] and Block [8]

,

and the theory

of Block was used as a basis for correlating our experimental results.

Basically crib fires burn in two regimes. The porosity controlled

regime is where the airflow through the crib controls its rate of mass

loss. Cribs with close stick spacing fall into this regime. For the

porosity controlled regime, Block derived the mass loss rate per exposed

I surface area to be

If

where G

and c() =

= 1/2 f p

1 1 1— 1

-e- 1

1 (1) 11— r 1

-e- 1

0

(N r< "^O

CL Q. 1

11— 1

[^( A

1/2

P hf

2 ^ is a dimensionless drag coefficient

vs

G/(j) (A-1)

Also B

(B + 1) + {2/[l + p/Pq) x"^] }

and B = [Ey - c^ (T^ - T^)]/e.

The symbols conform to those used by Block [8] and are listed below:

f = friction factor = 0.13

g = gravitational acceleration

h = height of crib

p = density of fluid in a crib shaft

p = density of ambient air

P^ = perimeter of a shaft

A = cross-sectional area of a shaft

vs

E = heat of combustion

e = heat of volatilization

Cp = specific heat of solid fuel

172
= surface temperature of burning solid

= ambient air temperature

Furthermore, Block showed that for densely packed cribs and most cribs

in general, most of the air flow enters from the bottom of crib rather than

from the sides. This vertical flow rate per unit shaft area A can be ^ vs

estimated as

(A-2a)

or for (J) ^ 1

^

s <=0 vs f O \ O /

2 For our design crib calculated as 0.89 kg/m -s or 17.5 g/s per

crib.

In the open or surface area controlled regime the mass loss rate per

unit exposed area is given as

a" = Cb“^ (A-3) V

where n = 1/2 according to Block, and C is an empirical constant, determined

for the polyurethane and taken from Block [8] for wood specie.

173
Table A-1. Crib open burn results: rate of mass loss correlation

Material

Density

P g/cm^

bcm

scm

n N C

mg/cm^ ‘ ^-s

¥ m

V

g/s

PINE, Ponderosa 0.54 1 5 3 10 0.520 1.03''’ 1.89 1.70 1.14

PINE, Ponderosa 0.42 4 6 3 10 1.73 1.03’’’ 2.84 6.5 1.26

PINE, Eastern 0.36 1 5 3 10 0.520 1. 1.94 1.6 1.10

PINE, Eastern 0.36 4 6 3 10 1.73 1. 2.94 4.0 0.81

PINE , Sugar + Ponderosa

0.39 - 0.43 3.5 7 3 10 1.30 1. 3.25 5.0 1.04

PINE, Sugar 0. 37 3.5 7 3 10 1.30 0.88"'’ 3.69 4.4 1.04

PINE, Sugar 0.37 3.5 7 3 10 1.30 0.88’*’ 3.69 5.2 1.21

Rigid Polyurethane 0.46 1 5 3 10 0.520 1.2 1.22 1.60 0.92

Rigid Polyurethane 0.46 1.5 7 2 4 0.223 1.2 1.78 0.^6 0.86

Rigid Polyurethane GM-37 0.34 1 5 3 10 0.52 1.2 1.23 1.78 1.02

Rigid Polyurethane GM-37 0. 34 4 6 3 10 1.73 1.2 1.85 4.0 0.66

Rigid Polyurethane GM-37 0.34 1 2 3 10 1.30 1.2 0.62 0.38 0.43

Rigid Polyurethane GM-37 0. 34 3.5 7 3 10 1.30 1.2 2.18 7.2 1.24

Rigid Polyurethane GM-37 0. 34 3.5 7 3 10 1.30 1.2 2.18 8.0 1.39

Rigid Polyurethane GM-37 0.34 3.5 7 3 10 1.30 1.2 2.18 6.05 0.94

^from Block

1/2

174
Table A

Estimated property values for crib material

Effective Heat of

Stoichiometric Specific Heat of surface volatiliMaterial fuel to air ratio heat combustion temp

zation B

A c E T e P s g/g J/g-°c kJ/g °C kJ/g

Wood* 0.27 1.26 18.6 370 1.78 2.56

Polyurethane0.13 1.26 23.4 325 1.52 1.80

*from Block [8]

^from literature

The data are tabulated in table A-1. The correlation by Block [8] is

based on 'i' defined as the ratio of equation (A-1) to equation (A- 3) . The

II predicted specific mass loss rate is the smaller of the values calculated

from equation (A-1) or equation (A- 3) . The parameters required in equation

(A-1) are listed in table A- 2. They represent available values determined

at the time this analysis was made, and are not necessarily the same as the

values used in section 6. A thorough analysis of the polyurethane cribs

has not been made, so those may not be the "best" values listed in table

A-2. Also in the calculations, the density ratio, P/P^/ '^as estimated as

0.2. The experimental results are plotted in figure A-1. The plastic and

wood crib designs are identical in geometry but differ in '1' values as shown. The design crib has f = 2.18 for the polyurethane, and T = 3.69 for

the sugar pine. Clearly there is uncertainty in selecting the "best"

property values for the polyurethane from limited data, and since its

sticks swell on burning. The swelling effect could cause the design value

for 'F to decrease shifting toward the porosity controlled burning tegime.

For the plastic crib design value of 'F = 2.18, an anticipated maximum swelling

of 50 percent in stick thickness reduced >F to 1.09 —still in the surface area

regime of burning.

175
176
APPENDIX B. RADIATION EFFECT ON LOWER LAYER THERMOCOUPLES

It was observed early in the experimental results that the gas temperatures

in the lower portion of the room were significantly above the incoming air

temperature. The lowest thermocouple (T23) read as high as 118°C (PRC 8-P-4-1)

It was unclear whether this was due to radiation or heated gas in the lower

space due to mixing. To resolve these results, thermocouple bead sizes were

recorded for experiments subsequent to PRC 8, and small wire thermocouples

and an aspirated probe were used (see figure 2-5)

In order to assess the radiation effect on a thermocouple under normal

atmosphere temperature, thermocouples were irradiated with a known radiant

flux from a quartz lamp. The irradiated thermocouple was in "still" air

but probably was subjected to a low natural convection induced flow. The

schematic of this experiment and its results are shown in figure B-1.

(Data points have been omitted for the aspirated probe.) The results yield

2 2 about a -5° C/W/cm correction for the aspirated probe and a -20° C/W/cm

correction for the normal size thermocouples used (0.5 mm bead diameter).

Figure B-2 is a plot of the same results for the "backed" beads (the reflective

metal backing shown in the schematic of figure B-1) in terms of bead diameter.

These results can be used to correct thermocouples due to thermal radiation

effects for the case of irradiation over one bead hemisphere and thermocouple

temperature above the true surrounding gas temperature.

A comparison was made between the bare thermocouple T23 and the aspirated

probe T28 for the average peak burning periods. Also, the radiation corrections

applied to both are also shown in figure B-3. In all but one case, the corrected

data appear to be in general agreement. Also, a heated lower layer was clearly

occurring in these fires.

177
RADIANT INCIDENT FLUX (W/cm^)

Figure B-1. Temperature correction versus incident radiant flux

178
3W/cm

179

BEAD

DIAMETER

(mm)
Figure B-3. Comparison of gas ^temperatures measured by two

techniques

180
NBS«n4A (REV, 2 «ec) U.S. DEPT. OF COMM.

BIBLIOGRAPHIC DATA

SHEET (See instructions)

PUBLICATION OR

REPORT NO.

NBSIR 80-2054

2. Performing Organ. Report No

3. Publication Date November 1980

4.

TITLE AND SUBTITLE

THE BURNING OF WOOD AND PLASTIC CRIBS IN AN ENCLOSURE

VOLUME I

5.

AUTHOR(S)

James G. Quintiere and Bernard J. McCaffrey

6. PERFORMING ORGANIZATION (If joint or other than NBS, see instructions)

national bureau of standards

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

7. Contract/Grant No. Grant No. RP-76-U-2

8. Type of Report & Period Covered

Final

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

Prepared for

Products Research Committee

c/o Dr. John Lyons National Bureau of Standards

Washington^ D.C. 2Q234

10. SUPPLEMENTARY NOTES

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

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

bibliography or literature survey, mention it here)

This study was designed to assess the fire hazard of a cellular plastic material

which has comparable structural characteristics to wood. The study attempts to determine

the relative fire risk of such materials in furniture. Rigid (high density) structural

polyurethane foam and supar pine were selected for fuels and burned in the form of cribs

in a room. The crib loading and door width were parameters experimentally varied.

Twenty-one room fire experiments and eight free burn experiments were conducted. Measuriments of mass loss, temperature, heat flux, CO2 and O2 concentration were recorded. The?e

data were analyzed and empirical correlations were developed for air flow rate and upper

gas temperature. A theoretical fire simulation model was developed and yielded results 4

-u fair to good agreement with the data. An extrapolation with the theoretical model was

used to predict the critical (or minimum) fuel pyrolysis rate to cause flashover

(as implied by 2 W/cm^ of incident radiation to the room floor). This was done for variejus

size rooms and door openings. It appears that to cause flashover, for a given room and

door size, about twice as much wood must be Involved in fire as the rigid polyurethane

material.

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

Crib fires; experiments; mathematical model; plastic;

polyurethane; room fires; wood.

13. AVAILABILITY

I Unlimited

I I For Official Distribution. Do Not Release to NTIS

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

20402. [^Order From National Technical Information Service (NTIS), Springfield, VA. 22161

14. NO. OF

PRINTED PAGES

202

15. Price

$17.00

USCOMM-DC 6043-P80
' ^ »fei ta:5},^(!/^ vi ^ 19 ,t^4>0

V*-< >VtiiH AKt>«jJ>

.OM T^M3;»

^?os--oa >aaaif

a^H<yU?90U819

««i\T33H?

3^rrr»u2 owa 3jt*t >,

’ a«o30ao«3 fiA wx aaw:> oka ciootf *§0 anr. ,

, X mv'ciov

' m . ... o ^ t^^llOKTUA

.oi! ias*iD

li^MWW^-^*WWWW^>i^ P III WIP f j l9 *«JtT .1

^tv^'iAcrt^ bnii ii' ?>cu>t.

Oi.i -jg,

I .

' i

'

I i1i%i« t mi » > ' ^' » ‘

' *»• 'K-W *0 lMo\ |fl H*C»rt AiiHAilWi*’ Of 'HWO^HaT.t

c- - ^ ' ^

'

all

<0Hlk«HAT« oO UAJMVUe J«r?d4T«ilii:

*tuf ^,0 ,0OTfimi«Afi

t .»;wi>ci3»6M ifHU = ’ »05 OKA

z

OvM>}« C 4

JL

lUi^fibrte^a

*» ikU :~ 7d \V\ Z

30l i>flii^qax^

*aaa ,«^a'jbo:

<ulot.. .3C o\i tAdOiM

‘AiwiNfe# • »A0* 'i'oit'i*

<Wt# ^giftOtr^ •O'ji Z' 1 ^ H ,t*^n'** *»j#»***it< »* «44ai«4IMl

><1lMVng8

<p ]

r. j

T4^iX90 B ^o,^biMia/i *»aBf>* oJ «*w rhu'i* aXa^k'j?' •.. jiojjgfr. . lo a<JT *bex?>t oJ^fcJi«l'S33oel*4A %I<tr*'' aq»so^ i-WHalifc

biglii ol rfoat la i«»±T nil! uvAS; »r*3 «itjL

lo Axfi *ft> h»l?* t«i »i*w tnlq btf» caoI ** ¥ll<»S45»«il%iKpc» «te»W •JOAb iusM (U35 auT »«s;Q7 » otl

^TUi»^ , Qt9v Bioawilifq^k irrvd **3^ litgX* bn« eia«lT»<pi<» aiH ««oo'» Bqo yii

aai;>»?jaao» :0 br» >00 a»Ad . TO3«r4>qw»d i«®oI

‘ #^*1 '*^11 il» lol l^olA^b 919V nAollfcXvT^oa Xaolii^oa tsa» fcax^lBQ*

ft, . .1- * >"# fcelilg^ bpiioS&>4b «AV Uboer noli^aXMAie ayJO

Jtt- .>!«« j£*ilJ4^o;rd^ ya rtoXsAiaqeiJxJ* aA

^rvtti^eAO 94»1 *rli»Y'^'^"*X? /9i'3 (?Hj«U£iX4iriwM^Xfe>^i^ pl^!

i«v »irtr .CiOvil Aooi *ill o3 pol3fc$b«t

toi^ * 3M»vi>d»AXl 3CtlBl '^3 1W<3 >30J:41SS»V<! biX6 BlCa^^sX^

iriJ*»; i’ U*^ iti-l'A tt» »Tjtl fiti bpvXovnl: »<J 3ti^ l>acw i{oiJft a# i»od*

Jy . I

'

; -•

. w»i. ! ii-. II • " » i ' ivAj

a#i: *!. \ r 4C i)0 f0$ 120

'YMot(M44«ii V 1^* l^NUrtl^- *^1 c* • 1) i:n4&'

ini^«jelq XAoi5/«cisddfAtn Souil di'xO

,hCtoxi/ T»/taAieiirYXDq

Ai > ',*1 *4

1 i<;A» *0#» ' ’*«

.»f 't .!( L

vu.yts

pm^ vtu©AJIAV:A Cl’ IQ '4 pAitfiiur. of-9A,>

ii i¥vi ift»wii'iiirf*<i6 A-it .*ni»i<‘*^M»Ci/#. i'^ •**>*^ GJ

%' 4

f-
TrjitI*j\i