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NBS

PUBLICATIONS

NBSIR 81-2453

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Quarter-Scale Modeling of Room

Fire Tests of Interior Finish

U S. DEPARTMENT OF COMMERCE

National Bureau of Standards

National Engineering Laboratory

Center for Fire Research

Washington, DC 20234

March 1982

Sponsored in part by:

Ship Damage Prevention and Control

Naval Sea Systems Command

artment of the Navy

lington, DC 20362

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NBSIR 81-2453 Ubbabt

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QUARTER-SCALE MODELING OF ROOM

FIRE TESTS OF INTERIOR FINISH

//-

B. T. Lee

U S. DEPARTMENT OF COMMERCE

National Bureau of Standards

National Engineering Laboratory

Center for Fire Research

Washington, DC 20234

March 1982

Sponsored in part by:

Ship Damage Prevention and Control

Naval Sea Systems Command

Department of the Navy

Washington, DC 20362

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

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director

TABLE OF CONTENTS

Page

LIST OF FIGURES iv

LIST OF TABLES vi

Abstract I

1. INTRODUCTION 2

1.1 General 2 1.2 Modeling Techniques 3

1.3 Objectives and Approach 8

2. EXPERIMENTAL TESTS 9

3. RESULTS AND DISCUSSION 13

3.1 Model Simulation of Room Fires 13

3.1.1 Nitrile Foam Rubber Insulation 13

3.1.2 Plywood Paneling 17

3.1.3 Fibrous Glass Insulation 19

3.1.4 Other Materials 23

3.2 Fire Location and Fire Buildup 24

3.3 Heat Production and Fire Buildup 25

4. SUMMARY AND CONCLUSIONS 32

5. ACKNOWLEDGMENTS 34

6. REFERENCES 34

iii
LIST OF FIGURES

Page

Figure 1. Room arrangements showing locations of fire source and instrumentation . 36

Figure 2A. Quarter-scale room fire test prior to ignition ... 37

Figure 2B. Quarter-scale room fire test at flashover 37

Figure 3. Top of doorway air temperature histories for several

full-scale fires and model tests using lintel III . . 38

Figure 4. Room air temperature profiles at time of peak doorway

air temperature 39

Figure 5. Doorway air temperature profiles at time of peak

doorway air temperature 39

Figure 6. Comparison of full-scale and model room fire test on plywood ' 40

Figure 7. Flashover time versus relative humidity for model room fire tests with pl3nyopd paneling 41

Figure 8. Room air temperature histories for full-scale fire

P2 and model fires M2I, M2II and M2III 42

Figure 9. Ceiling temperature histories for full-scale fire

P2 and model fires M2I, M2II and M2III 43

Figure 10. Room air temperature profiles at 240 S for full-scale

fire P2 and model fires M2I, M2II and M2III ..... 44

Figure 11. Doorway air temperature profiles at 240 S for fullscale fire P2 and model fires M2I, M2II and M2III . . 45

Figure 12. Doorway flow velocity profiles at 240 S for full-scale

fire P2 and model fires M2I, M2II and M2III 46

Figure 13. Ignition time of newspaper flashover ind3,ca^prs as a function of floor location ..... 47

iv
LIST OF FIGURES (Continued)

Page

Figure 14. Model room air temperature profiles at 240 S for

three burner locations 48

Figure 15. Full-scale room air temperature profiles at 240 S for three burner locations 49

Figure 16. Model doorway air temperature profiles at 240 S for three burner locations 50

Figure 17. Full-scale doorway air temperature profiles at

240 S for three burner locations 50

Figure 18. Model doorway flow velocity profiles at 240 S for

three burner locations 51

Figure 19. Full-scale doorway flow velocity profiles at 240 S for three burner locations 52

Figure 20. Model and full-scale room air temperature histories for

fire source at back corner 53

Figure 21. Model and full-scale ceiling temperature histories for

fire source at back corner 54

Figure 22. Model and full-scale doorway air temperature histories

for fire source at back corner 55

V
LIST OF TABLES

Page

Table 1. Interior finish materials used in room fire tests . . 56

Table 2. Model doorway dimensions 57

Table 3. Summary of full-scale tests and counterpart model

tests of nitrile foam rubber and pljwood paneling . . 58

Table 4. Summary of flashover times and temperatures in room

fire tests with exposed fibrous glass insulation

lining the walls and ceiling 59

Table 5. Summary of heat fluxes in room fire tests with fibrous

glass insulation 60

Table 6. Comparison of fire buildup in full-scale and

corresponding quarter-scale room fire tests of

nitrile foam rubber 61

Table 7. Flux measurements at time of flashover 62

Table 8. Comparison of room air temperatures for fire tests with

three different doorway openings 63

Table 9. Doorway flow velocities in fire tests with fibrous glass

insulation 64

Table 10. Comparison of flashover and flameover times (s) for

full-scale and quarter-scale tests 65

vi
QUARTER-SCALE MODELING OF ROOM FIRE TESTS OF INTERIOR FINISH

B. T. Lee

Abstract

A technique for modeling fire buildup in

rooms with combustible interior finish was refined

to achieve closer simulation of full-scale fire

development. Fire experiments were performed in

one-quarter scale model rooms and full-scale rooms

having a doorway opening. The interior finish test

materials were nitrile foam rubber, fibrous glass,

and plywood; a gas burner was employed as the fire

source in a rear corner of the room. It was necessary to lower the doorway opening in the model by as

much as 14 percent to obtain flashover with the same

equivalent heating rate that produced flashover in

the full-scale test. At the same time the width of

the doorway in the model was increased appropriately

to maintain the same volumetric air flow rate. The

effects of burner location and heating rate on flashover in a well-insulated room were also studied to

help select a suitable ignition source size and

placement for testing of interior finish materials.

The minimum heating rate needed to cause flashover

ina3x3x2.3m high room lined with fibrous glass

and having a 0.73 x 1.93 m high doorway opening would

1
entail placement of the heat source in a back corner

with the source having a heat release rate of 300 kW.

A corresponding rate for the quarter-scale room

would be 19 kW.

1 . INTRODUCTION

1.1 General

In many instances, full-scale room fire testing is the only means to

realistically assess the fire performance of materials. Laboratory fire

tests frequently do not predict the potential fire risk of interior

finish materials under actual room fire conditions. The problem is

partly with the test methods used and partly with the interpretation and

application of the results. There is presently no suitable analytical

prediction model for relating interior finish material fire test data to

compartment fire growth, although NBS and other organizations have ongoing programs to develop such a capability. Until such an analytical

prediction becomes feasible, adequate selection of fire safe materials

must often rely on ad hoc full-scale room fire testing. This latter

procedure is expensive, especially if many materials are to be evaluated.

A more economical and practical alternative is the employment of a

reduced-size model room fire test for the screening of materials.

2
Effective scale modeling methods can contribute significantly to an

improved understanding of room fire growth and, hence, help in the

development of a prediction model. Fire is intrinsically an accidental

occurrence, and the initiation and spread of fire in a room can occur in

a variety of ways. Employment of a reduced-scale physical model is the

only economical means of achieving sufficient parametric variation and

physical insight for a generalized understanding of the problem. Thus,

there is much incentive to develop a reduced-scale model which would

predict full-scale fire performance.

1.2 Modeling Techniques

Scaling techniques with small models have been successfully applied

in such diverse fire research areas as the temperature and air flow

fields in large area fires [1]^ and the modeling of burning rates and

flame spread by increasing ambient pressure [2,3]. Pressure modeling

has also been applied for simulating enclosure fires [4]. In pressure

modeling of fires, the ambient pressure is increased while the length

scales are reduced according to the two- thirds power of the pressure.

The two main drawbacks to this method are the problem of scaling the

radiation and the need for a large and costly pressure chamber capable

of handling operating pressures much above ambient conditions. Two

other reduced-scale modeling methods which have been used for simulating

room fire buildup cannot properly handle the contribution of combustible

walls and ceilings. One of these, developed at the Illinois Institute

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

end of this report.

3
of Technology Research Institute (IITRI) [5], requires a constant ratio

of heat release rate to the volumetric rate of air inflow in order to

maintain the same temperature in the room. When the horizontal dimensions

of the prototype compartment are reduced by a scale factor, the ITTRI

modeling criteria require that the vertical dimension should be proportional to the scale factor raised to the two-thirds power and the rate

of heat release be proportional to the square of the scale factor.

Thus, the rate of heat release is proportional to the floor area but not

to the wall area. If it is assumed that the mass burning rate of a

combustible wall is proportional to its area, then the heat release rate

of the wall material will exceed the requirement that it be proportional

to the floor area.

Factory Mutual Research Corporation approached the room fire scaling

problem from dimensional analysis considerations [6]. Their findings

indicate that the temperature and gas compositions in a room scale

reasonably well for geometrically similar enclosures where the heat

release rates are proportional to the 5/2 power of the scale factor.

This method assures that the ceiling of the model and prototype are at

homologous points of the convective column generated by the flame.

However, since the interior surface area of a room is proportional to

the square of the scale factor for geometric scaling, the fuel contribution from a combustible wall is too large for correct scaling using this

approach.

4
At the National Bureau of Standards a model having geometrically

scaled room dimensions has been effective in modeling fire buildup in

rooms with combustible walls and ceilings [7]. This modeling procedure

assumes that the severity of a room fire can best be described in terms

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

layer, as the term is used here, includes flaming and non-flaming gaseous

pyrolysis and combustion products. When this air temperature reaches

500°C there is rapid pyrolysis of the combustibles in the upper part of

the room. When this temperature reaches about 650 °C there is sufficient

radiation into the lower part of the compartment to ignite virtually

everything combustible. The maximum temperature which can be reached

with a given set of lining materials, furnishings, and ventilation can be

identified as the "fire buildup potential" of the compartment. The

estimation of the fire buildup potential of a room then is based on a

prediction of the maximum temperature rise. This prediction depends on

setting up an energy balance between the heat produced, the heat lost

through the lining materials, and the heat carried out the doorway. For

simplicity, the room is assumed to be divided into two uniform temperature regions with the higher air temperature, T, in the upper part of the

room and the ambient air temperature, T^, in the lower part of the space.

It is further assumed that there is a continuous inflow of cool air

through a single open doorway into the lower portion of the room and hot

air carrying gaseous combustion products exhausting from the upper part

of the doorway. The mass of the pyrolysis products produced in the room

is neglected. The rate of energy storage term is generally small and is

neglected here. The heat balance can then be expressed as

5
Q + Z. qV A. = pCV(T-T ) + L(T-T ) ( 1 ) 1 X o o

where Q is the total heat release rate of all the combustibles in the

room exclusive of the interior finish materials, is the heat release

rate per unit area of the i-th finish material, A. is the fire involved

1

area of the i-th finish material, p, C, and V are the density, heat

capacity, and volumetric flow rate of the hot air and combustion products

exhausting from the room, and L is the ratio of the rate of heat loss (by

conduction through the linings of the room and by radiation through the

doorway) to the temperature rise. This leads to the following expression

for the temperature rise.

T - T

o

Q + Z.qVA ^ x^i i

pCV + L

( 2 )

When the terms in the numerator and denominator on the right hand side

are all divided by the floor area. A, the set of ratios which must remain

constant in order to achieve the same temperature rise in the model and

prototype becomes apparent, namely, Q/A, A^/A, L/A, and V/A. Q/A is

maintained constant by prescribing the strength of Q, which would be the

heat release rate of the gas burner used as the ignition source in the

full-scale and reduced-scale tests for testing interior finish materials.

A^/A and L/A are kept constant by geometric scaling, if the fires are

similar and the heat transfer coefficients are the same in full- and

reduced-scale. Unfortunately, scaling of L/A is only approximate as the

convective heat transfer is not scaled.

6
3H Since V « WH where W is the width of the doorway and H is its

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

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

factor. The wall above the doorway traps the hot combustion products

from the fire and is critical to the phenomena taking place in the room

so that this height was chosen to be scaled geometrically. For quarter

scale modeling, the doorway width then is half of its full scale value

while the other dimensions are only one quarter, except for the thickness

of the interior finish materials.

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

the prototype, to insure that the heat losses per unit area remain about

the same for the same interior air temperature. This is of great practical value since materials are tested in the thicknesses available in

the market place and composites do not pose an additional fabrication

problem.

However, the following problems are encountered with the scaling:

1. Since the lateral flame spread rate does not change with

scale, it is relatively too high in the model.

2. The flame heights are too high in the model.

7
3. The convective heat transfer coefficient is too low in

the model since air velocities are proportional to the

square root of the scale.

4. Radiation from the upper walls, ceiling, and hot air layer

is scale dependent when the hot air layer is semi-transparent

and a vertical temperature gradient exists.

5. The increased size of the doorway opening required to

scale the volumetric air flow rate permits slightly greater

heat losses and slightly less heat release from combustible

walls

.

1.3 Objectives and Approach

The principal goal of this study was to help develop a quarter-scale

room fire test for evaluating the potential fire hazard of interior

finish materials. The approach was to modify the existing NBS quarterscale working model [7], even at the risk of deviating from previously

existing scaling theory, such that the maximum temperatures reached and

the times to flashover and flameover* duplicate full scale tests closely

enough for screening purposes and for studying the fire buildup phenomena

on a much more economical scale. Two kinds of modifications were considered. One was to adjust the intensity of the heat source in the

*Flashover is defined here as the room fire condition where the thermal

radiation level becomes high enough to spontaneously ignite light com- bustible materials, such as newspaper, in the lower part of the room. Flameover is defined as the room fire condition where flames emerge from

the doorway.

8
model, which deviates from scaling theory. The other modification

involved an empirical adjustment of the doorway opening, which does not

3/2 violate scaling theory, as long as the quantity WH remains constant.

Changing the doorway in this manner would, however, affect the temperature

distribution and radiative environment in the model. Three significantly

different materials, fibrous glass, nitrile foam rubber, and plywood, were

employed in this investigation.

Secondary objectives were to determine the effects of heating rate and

location of heat release in the room on room flashover. This information

is essential for determining the size and placement of an ignition source

which should be sufficiently large to realistically appraise the fire

hazard potential of materials and, at the same time, would not be so

large as to overwhelm the material being evaluated.

2 . EXPERIMENTAL TESTS

Several full-scale room fire tests, where the walls and ceiling were

lined with nitrile foam rubber [8] and where the walls and celling were

lined with plywood paneling [9], were selected as prototype tests for

this modeling study. In addition, full-scale fire tests were performed

with fibrous glass Insulation fully covering the walls and ceiling to

provide further tests for modeling. These three interior finish materials

are described in table 1. The room used to test the foam and fibrous

glass had dimensions of 3.0 x 3.0 x 2.3 m high and a 0.73 x 1.93-m high

doorway offset to one side, figure 1. Since this room was used for a

9
study on fires in shipboard compartments, all four walls and the ceiling

had 6.4 mm thick aluminum alloy plating mounted on 51 x 102 mm steel

studs 0.41 m apart. The floor was covered with 3.2 mm thick aluminum

sheet. The test material was mounted directly on the room walls and

ceiling using an adhesive for the foam and screws with end washers for

the fibrous glass. The corresponding model test enclosure was a onequarter replica of this room except for the doorway dimensions and is

shown in figure 2. The test room with the plywood paneling was 3.0 x 3.0

X 2.4 m high with a 0.76 x 2.03-m high doorway. This room had concrete

block walls and a gypsum board-wood frame ceiling construction. Wood

studs, 25 X 75 mm, spaced at 0.41 m were used as spacers between the

plywood and the concrete walls and gypsum board ceiling. The model

version of this room used an enclosure lined with gypsum board. Wood

stud spacers and the plywood were then applied over this surface. In

both model enclosures three doorway lintel depths of one-quarter, 1.4

times one-quarter and 1.8 times one-quarter of the full-scale lintel

depth, labeled as lintels I, II, and III, respectively, were arbitrarily

chosen to evaluate the effect of doorway height. The dimensions for

these three openings are shown in table 2. Both full-scale and model

room fires were conducted within a large test building so that the effects

of wind and temperature extremes were minimized.

Location of all instrumentation in both size rooms is indicated in

figure 1. Measurements made in many of the tests to characterize the

thermal environment in the room included vertical temperature profiles

down from the ceiling and along the centerline of the doorway, vertical

10
distribution of air flow velocities in the doorway, thermal flux to the

ceiling and two walls, and radiation incident on the floor. Temperatures

were measured with thermocouples fabricated from 0.25 mm diameter chromel

and alumel wires. Ceiling and wall fluxes and the thermal radiation

Incident on the lower part of the room were monitored with water-cooled

Garden type of total heat flux gauges. Crumpled newspaper on the floor

was also used as flashover Indicators to show when the irradiance from

the heated air and upper room surfaces was sufficient to ignite light

combustible materials in the lower portion of the room. Pitot tubes were

located in the upper one-third and lowest one-third of the doorway height

in the large room. Bidirectional velocity probes [10] were employed in

the middle one-third of the doorway opening. These bi-directional probes

have the capability for measuring the velocity and the occurrence of any

flow reversal along the doorway. They were used to monitor the entire

height of the opening in the model.

In the tests with the nitrile foam rubber [8], a 0.305 x 0.305 m

porous plate diffusion flame burner, positioned 0.305 m above the floor

and in contact with both walls of one back corner of the room, served as

the fire exposure source in the full-scale tests. The burner used

methane gas and operated at steady heat release rates of 62 kW (for three

tests) and 94 kW (for one test). The full-scale room fire tests of the

plywood paneling used the same burner in the same location but with a

constant rate of 90 kW. The fire source in the model was a 76 x 76 mm

porous plate burner standing 76 mm high above the floor. The full-scale

tests and their corresponding model tests are outlined in table 3. For

11
the experiments using the fibrous glass material, eighteen fire tests

were conducted in the model, and six fire tests were performed in the

full-scale room. Conditions for these tests are outlined on table 4.

Several steady rates of heat release were employed using the above mentioned full-scale and model burners. Those rates ranged from a heat

output of 62 kW, which is smaller than that from the burning of some

small upholstered chairs, in the large room to rates which were sufficient

to produce flashover of the space. Propane was used for the first five

large scale tests shown in table 4. Test P5 was run at the highest flow

possible with the 100 psi propane. A switch to bottled methane was then

made in P12 to obtain a sufficient rate of heat release to cause flashover. In these tests with the fibrous glass insulation, the relationship

between the rate of heat release in the room and the time to achieve

flashover was also explored. As the heat release contribution from the

fibrous glass was difficult to predict, it was decided to first burn away

the combustible content of this insulation and then rely on the burner

for the sole source of heating in the actual testing. Moderate size

natural and propane gas fires were conducted for 900 s in the model and

large rooms, respectively, several hours before actual testing to burn

off any organic binder that might be present in the fibrous glass lining

the rooms. Natural gas was used for the first six model tests. Inadequate gas line pressure forced a switch to propane fires. Tests were

then run, using a propane and methane gas fire, respectively, with the

same rate of heat release as in test M2I to check whether any difference

would occur in the fire environment when a propane or methane fire was

12
used. Little difference, i.e., less than 5 percent variation, was

observed in the burner flame heights, room temperature profiles and flux

levels in the enclosure with the three gaseous fuels.

3. RESULTS AND DISCUSSION

3.1 Model Simulation of Room Fires

3.1.1 Nitrile Foam Rubber Insulation

Doorway and interior air temperatures and flashover times for the

four full-scale tests and their corresponding model tests are presented

in table 3 [8]. Flashover was assumed to have occurred at a time taken

to be the average of the times for the first two indicators (newspapers)

to ignite. Ignition of the newspaper, a specified flux level at some

location on the floor, or some designated minimum doorway or interior air

temperature can all be used to indicate the onset of flashover. However,

there are problems associated with the use of each method. Variation in

the thermal and physical properties of newspaper can result in the newspaper igniting over a range of fluxes between 17 and 25 kW/m^ [11].

Non-uniform distribution of thermal fluxes throughout the room also poses

a problem for determining flashover with both flux meters and newspaper

flashover Indicators. Non-uniform distribution of temperatures throughout the room and thermal radiation errors associated with thermocouple

measurement of temperature affect the accuracy of interior and doorway

13
air temperature measurements. All of the above techniques for determining

flashover can, however, help support each other. Moreover, where an ignition of the newsprint indicator had occurred, room and doorway temperatures

either had attained or continued to increase to at least 650°C and 550®C,

respectively. The latter temperature levels are above those usually

prescribed for the onset of flashover.

The full-scale tests FS-1, FS-2, and FS-3 all attained flashover

within the test period of 600 s. However, only one of the corresponding

tests with the model having the lintel I doorway (test 6) reached flashover. Tests 4 and 10 reached peak doorway temperatures of 215 °C or less.

When the doorway height was lowered by 7 percent, as was the case with

the lintel II opening, two of the three model tests, numbers 22 and 23,

attained flashover. However, in test 24 which was the counterpart of

test FS-3, the temperatures in the doorway peaked at only 288 °C. Lowering

the doorway further to 14 percent as in test 34, corresponding to the

lintel III model version of test FS-3, resulted in flashover. In general,

whenever flashover occurred in the model, it developed slower than in the

corresponding full-scale test. For the full-scale test FS-4, the peak

doorway temperature was about 300®C. Lowering the lintel from II to III

in the model did not increase model doorway temperatures much beyond what

may be regarded as experimental variation expected between similar tests.

Again, the model fire developed slower, but the peak fire buildup, as

evidenced by the maximum air temperature in the doorway, simulated the

full-scale behavior. Figure 3 shows the doorway air temperature variation

with time for the model tests having the lintel III doorway and for their

14
counterpart full size fires. A condensed version of table 3, showing a

comparison of the fire buildup in the full- and quarter-scale compartment

fires of the nitrile foam rubber, is given in table 5.

Flux measurements at the time of flashover for the full-scale and

corresponding quarter-scale tests are indicated in table 6. The average

irradiance on the floor at which flashover occurred was 22 kW/m^ for the

full-scale fires FS-1, FS-2, and FS-3, and 24 kW/m^ for the model fires

23, 33, and 34. The irradiance level on the floor for test FS-4 was also

lower than that for its counterpart model test. This was also observed

in tests where the interior finish was an inert fibrous glass Insulation

(refer to section 3.1.3). The higher flux probably resulted from the

relatively taller burner flames in the model and from the relatively

larger heated surface area in the model which occurred as a consequence

of the lowered lintel.

The vertical air temperature profiles inside the room and along the

doorway for three full-scale fires, at their respective times of peak air

temperature as measured at a location 102 mm below the doorway lintel,

are shown in figures 4 and 5. Superimposed on the figures are the corresponding test data from the model tests 33, 34, and 51 with the lintel

III doorway. The times to reach the peak doorway air temperature for

tests FS-2, 33, and 34 also corresponded to the times for flashover,

whereas flashover occurred seconds after the peak doorway air temperature

was attained in test FS-3. In the model tests 33 and 34, temperatures

were lower than in the full-scale test and, consequently, the model fires

15
required a longer fire exposure of the insulation in order to reach

flashover. For model test 51 and its corresponding full-scale fire FS-4,

the fire was confined principally to the zone in contact with the flames

from the ignition source, producing only a moderate air temperature rise.

The vertical distribution of temperatures below the center of the ceiling

indicated much higher temperatures in the model than those in the fullscale fire. However, these temperature differences in the upper part of

the compartment did not show up in the doorway temperature profiles.

This may be explained by the relatively taller flame heights in the

model. The higher flames could have resulted in a more intense localized

heating of the thermocouples in the upper region of the compartment.

However, the heated air in the compartment space became well stirred by

the time it reached the doorway. This is evident from the similarity in

the doorway temperature profiles for the two tests.

Repeatability of these model tests and the feasibility of adjusting

the intensity of the ignition source to achieve satisfactory modeling

were also studied with the results indicated in table 7. Tests 5 and 17

were repeat tests of tests 4 and 16, respectively, and show good repeatability between tests in the doorway air temperatures but less agreement

in the interior temperatures. Tests 16 and 17 showed that an increase in

the heat release rate of the ignition source by as much as 50 percent can

sometimes have little effect on room fire buildup. For the C2 insulation

used in that test, the lowering of the doorway by 7 percent as in test 22

had a much greater effect on the fire growth than a significant increase

in the size of the fire initiation source. For the B2 insulation, a

16
50 percent increase in the intensity of the ignition source led to flashover, but test 34 demonstrated that small decreases in doorway height

could also lead to equally large increases in room interior and doorway

air temperatures.

3.1.2 Plywood Paneling

Data from the two full scale room fire tests and the corresponding

model tests 1 and 2, conducted in ^ previous study [9], are outlined in

table 3. Data from model tests 3, 4 and 5, which were performed for this

present study, have been included in table 3. Model tests 1 and 2 used a

doorway height which was 0.90 as high as the quarter-scaled doorway.

Model test 3 used a quarter-scaled doorway while tests 4 and 5 had doorway

heights 0.93 and 0.86 as high, respectively, as that in test 3. The

plywood used for all of these tests appeared to be of the same material.

However, full-scale test 1 did not reach flashover, while full-scale test

2 conducted one week later experienced flashover in 158 s. Model test 1,

conducted immediately after the full-scale test 1, also did not reach

flashover, while model test 2, performed in the same day as was fullscale test 2, had one of its newspaper flashover indicators ignite at

185 s. Figure 6 shows good agreement in the chronological development of

the fire between model test 2 and its counterpart full-scale fire test 2.

Model tests 3, 4, and 5 were conducted over the range of relative humidities from 35 to 45 percent. Room flashover times for these three tests

occurred between 310 and 350 s. The different degree of fire buildup

between the full scale tests 1 and 2 and the differences among model

17
tests 1, 2, and the last three model tests could be ascribed to a combination of varying properties for the plywood paneling and to differences

in humidity conditions during each test.

Unfortunately, humidity conditions were unknown for the full-scale

tests and for the model tests 1 and 2, and no more of the same paneling

was available for a study on the effect of humidity on room fire growth

using this material. A substitute plywood paneling was employed for this

latter purpose. This plywood was tested in the model having a 14 percent

lowered doorway opening under several humidity conditions. The data

shown in figure 7 indicated that the times required to reach flashover

varied randomly between 223 and 323 s over the range of relative humidities from 22 to 76 percent. Evidently, the variation in material

properties and its effect on the room fire development overwhelmed the

differences due to humidity. It is possible that the plywood used in the

full-scale tests and in the model tests 1 to 5 was more sensitive to

humidity and, thus, could have accounted for the differences between

tests shown in table 3. On the other hand, different batches of the

plywood paneling having different material properties could have been

used for each of the following sets of tests: (1) full-scale test 1 and

model test 1, (2) full-scale test 2 and model test 2, and (3) model tests

3, 4, and 5. Care was taken to assure that material from a single batch

was used in the last three model tests and that these tests were performed

under a narrow range of humidity conditions. Results in table 3 showed

that only small differences occurred as a consequence of modifying the

height of the doorway when this plywood material was used. This was

18
believed to be the result of rapid flame development over the surfaces

of the wall and ceiling materials.

3.1.3 Fibrous Glass Insulation

Results of the fire tests M2I, M2II, and M2III in the quarter-size

models having the geometrically scaled, 7 percent and 14 percent lowered

doorway heights, respectively, are compared with the data from the corresponding full-size room fire test P2 in figures 8, 9, 10, 11, and 12

and in table 4. Temperature histories near and on the ceiling for these

tests are shown in figures 8 and 9. Vertical temperature profiles inside

the room and at the doorway opening are presented in figures 10 and 11.

These temperature histories and profiles show that the model having the

geometrically scaled doorway height gives a better simulation of the

temperatures in the full-scale room fire.

velocities along the centerline of the doorway opening are indicated in figure 12. The three models have about the same velocity

distribution along the height of the doorway. Doorway flow velocities

for all of the fire tests on the fibrous glass Insulation are given in

table 9. The scaling criteria suggest that velocity should scale as the

square root of the scale factor, meaning that the full-scale values

should be twice as large as the model velocities. Inspection of the test

data indicates that this is roughly the case. The calculated volumetric

flow rate determined from these profiles was uncertain by as much as

30 percent as the calculated outflow was about 20 to 30 percent larger

19
than the calculated inflow. There is evidence that volumetric flow rates

calculated from just centerline velocities could give outflow rates which

are 20 to 30 percent higher than the calculated inflow rates, and that

this apparent excess outflow results from an insufficient mapping of the

flow across the entire doorway area [12,13]. Nevertheless, the agreement

between the temperature profiles between M2I and P2 suggested that the

volumetric inflow rates to the model room properly scaled the Inflow to

the full-scale fire.

Although geometric scaling of the doorway height leads to adequate

reproduction of the temperature profiles, provided the rate of heat

production is properly scaled, the heat fluxes to the room surface are

not scaled properly. This is due to: (1) the flame height being relatively higher in the model, (2) the velocity of the flow of hot air and

combustion gases in the model room being lower, and (3) the thickness of

the layer of hot air and combustion gases being less in the model. A

greater flame impingement area from the burner in the model results in a

larger area of high surface temperature. This results in a general increase in thermal radiation levels in the room. On the other hand, the

convective and radiative heat transfer from the hot air and combustion

gases to the model room surfaces would be lower as these quantities decrease with decreasing velocity and decreasing thickness of this hot

layer, respectively. Convective heating is the dominant mode of heat

transfer near and within the flame zone, which extends upwards from the

ignition source. With combustible materials, the flame spread away from

this zone would be slower in the model. This has been the case with the

20
fires discussed in sections 3.1.1 and 3.1.2 where the times to attain

room flashover were slower in the model than those for the full-scale

fires. For the fibrous glass insulation, which did not contribute to

the fire, thermal flux measurements for tests M2I, M2II, M2III, and P2

(presented in table 8) indicate that the total heat flux measured at the

heat flux meter locations was greater in the model fires that in the fullsize counterpart test. The data indicate that the heat flux values along

the walls and on the floot were highest in the model with the lintel III

doorway. This results from a thicker heated layer of air and combustion

gases and more surface exposed to the elevated temperature caused by the

lowering of the doorway in the model.

The effect of lowering the doorway in the model on fire buildup was

also investigated for the case where the heating rate from the gas burner

was raised sufficiently for rapid flashover of the room. Results of the

model fire tests M12I, M12II, and M12III, having the lintel I, II and III

openings corresponding to the geometrically scaled, 7 percent, and 14

percent lowered doorway heights, respectively, are compared with data

from prototype test P12 in tables 4 and 8. When the burner heating rate

was raised sufficiently for rapid flashover in the model and full-scale

room fires, flames covered much of the ceiling in both cases and the

relative areas covered by the flames did not differ appreciably between

the model and full-scale tests. However, the layer of heated air and

combustion gases was thicker in the full-scale fire than was the hot

layer in the model fire which led to a more intense radiative environment

in the full-scale room. The data showed that the gap in the flashover

21
times resulting from these differences in the model and full-scale fires

could be bridged considerably by lowering the doorway, which consequently

increased the thickness of the layer of hot air and combustion gases in

the model. The time required to reach flashover decreased from 138 to

90 s as the doorway opening was lowered from the lintel I to the lintel

III height. The flashover time of 90 s for the model with the lintel III

opening compared well with the time of 78 s required for flashover in its

counterpart full-scale test P12.

In many of the tests, the incident thermal flux value on the floor

at the time of flashover was lower than the 17 to 25 kW/m^ range expected

to ignite the newspaper flashover indicators. A simple experiment was

performed to observe the effect of room flux distribution on the ignition

times of crumpled newspaper under one room flashover condition. The same

full-scale test room with the fibrous glass lining and corner gas burner

was used. The floor of the test room was divided into nine equal sections with a piece of crumpled newspaper placed in the center of each

section. A heat release rate of about 500 kW was used for the burner.

Ignition times for the newspaper flashover indicators were recorded and

are given in figure 13. The results showed newspaper ignition times

varying between 39 and 92 s, with the longest times occurring at the

three corners farthest from the burner. Figure 13 also indicated that

the placement of the floor flux meter for the tests in table 8 was at a

position where the flux was too low, resulting in values significantly

lower than the 17 to 25 kW/m^ range anticipated at the time of flashover.

22
3.1.4 Other Materials

Several other materials have been evaluated with both the model

having the 14 percent lowered doorway and full-scale room fire testing in

a separate study [9]. In that study, three types of polyurethane, a

foil-faced and an unfaced polyisocyanurate, fiberboard, and unfaced

fibrous glass were used. Further description of these materials along

with some of the test results from that study are outlined in table 10.

Included in table 10 is a summary of the room flameover and flashover

times for all of the tests with the nitrile foam rubber, plywood paneling

and fibrous glass insulation where there were both full-scale and quarter

scale counterpart tests. The results indicated that materials which

resulted in rapid flashover in the full-scale tests also did so in the

model tests; while those materials which did not contribute significantly

to the full-scale room fire development also did not contribute much to

the model room fire growth. In general, fire buildup times (e.g., times

to reach flameover and flashover), took longer in the model, and this

can be seen in the data for the three polyurethanes. This could result

in situations where the fire development just barely reaches flameover

and/or flashover conditions in the full scale test, but would not do so

in the model test. This situation is well Illustrated with the polyisocyanurate test where the time to reach flashover was 368 s and no

flashover happened in the corresponding model test. However, in borderline flameover or flashover situations such as this, other indicators

such as the interior and doorway air temperatures in the model could be

used instead to help predict full scale behavior.

23
3.2 Fire Location and Fire Buildup

Three different burner positions along the floor were considered in

the room fire tests to help assess the effect of burner location on fire

buildup and as a further check on the model’s ability to follow the fullscale fire development. These locations, as indicated in figure 1, were

at a back corner, the center of the back wall, and the center of the

floor. The effect of placement of the fire source on the temperature and

the flux levels in the room and on the temperatures and flow velocities

along the doorway opening for both the model and full-scale fires are

given in figures 14-19 and in tables 4 and 8. The figures show that as

the burner was moved from the floor center to the back wall, and from

there to the corner of the room, the heated air in the room and the hot

exhaust through the doorway became more and more concentrated in the

upper regions of the room and near the top of the doorway. In general,

the model fires behaved much like the full-scale tests except for somewhat higher air temperatures resulting from the lowered doorway in the

model. The flow through the doorway did not change appreciably when the

burner was relocated from the back wall to the corner. Thermal flux

levels for the three burner locations in both of the quarter- and fullsize fires are presented in table 8. Ceiling fluxes for the corner

burner placement were higher than those for the fire source positioned

against the back wall. This resulted from the longer flames when the

fire was in the corner. The shortest flames occurred when the burner was

at the room center. However, at the latter location the flame was directly

under the flux meter and, consequently, the instrument saw a higher flux

24
than when the fire was placed at the back wall. Thermal fluxes, in

general, were lower for the full-scale fires. This was a direct consequence of the relatively shorter flames in the full-size tests.

The highest ceiling temperature and air temperatures in the room and

at the doorway occurred when the burner was used in the corner as in test

M2II instead of being used against the back wall or at the center of the

floor as in tests M3II and M4ll, respectively. Locating the burner or

another fire initiation source at the center of the floor is not suitable

as the flames from any reasonable size fire can not directly contact the

wall material. Placement of the ignition source against the back wall or

in the corner are two practical alternatives. Corner placement of the

burner or another fire initiation source appears to be more desirable as

the room interior finish would be evaluated under the more demanding but

yet realistic fire exposure for any given rate of heat release for the

ignition source.

3.3 Heat Production Rate and Fire Buildup

A wide range of heat release rates in the room fire tests was employed to determine combinations of heating rates and exposure times

required for room flashover and to check the performance of the model

under conditions of modest heating all the way up to flashover. Table 4

arranges the various quarter- and full-size fire tests in the order of

increasing rates of heat generation in the room. Figures 20-22 show the

temperature buildup near the ceiling, on the ceiling, and along the top

25
of the doorway opening for corner placement of the fire source in the

model and full-scale tests. For the two lower heating rates shown, the

model tests Mill and M2II or M2III exhibited higher air and surface

temperatures than those in the full size tests PI and P2. This was

expected as a consequence of the lowered doorway height in the quarter

size enclosure. For model tests Mill and M2II the ceiling flux was also

higher than that in their corresponding tests Pi and P2 due to the taller

flames in the model fires. However, at the full-scale equivalent heating

rate of 460 kW, the temperatures in test P12 exceeded those for the

corresponding model test M12II or M12III. Table 8 indicates that at the

latter heating rate the ceiling flux level in the full-scale fire was

also higher than that in the counterpart model test at their respective

times of flashover. There are two explanations as to why air temperatures and the ceiling flux were higher in this full-scale test. The

first reason is that the ceiling in both P12 and M12II or M12III were

fully covered by the burner flames, meaning that the longer flame in the

model merely spills over and out of the opening and would not contribute

significantly more to the fire. Secondly, the higher convective heating

and thicker hot layer of air and combustion gases in the full-scale fire

apparently overshadowed the additional trapping of heat in the room when

the lowered lintel height was used in the corresponding model test,

resulting in higher ceiling flux and higher air temperatures in the fullscale room.

The degree of flame coverage of the ceiling plays a dominant role in

the determination of the thermal flux environment in the fire room. The

26
full-size test P5 was conducted with a 50 percent higher scaled rate of

heat release than those in the model tests M2II and M2III, but it did

have the same degree of flame coverage of the ceiling as those in the two

small-scale fires. Except for the wall fluxes in test M2II, the flux

measurements in table 8 Indicate that the flux levels at the floor, wall,

and ceiling were about the same for the three tests.

Heating rates leading to flashover of the space for the three burner

locations in the room are indicated in table 4. As the interior finish

in the room fire experiments was an inert fibrous glass having good

insulating properties, these rates of heating could represent minimum

rates needed in practice. The lowest full-scale equivalent rate of heat

generation found to lead to flashover in the model room was 300 kW,

maintained for almost 420 s, at the corner of the room. At a heating

rate of 460 kW in the corner, flashover times of 80 and 100 s were found

for the full-scale and model tests, respectively. Increasing the rate to

645 kW resulted in flashover of the model room in only 45 s. When the

fire source was placed against the back wall, a heating rate of 375 and

410 kW led to flashover at 240 s in test MIOII and at 126 s in test

Mini, respectively. With the burner at the floor center, a rate of

460 kW could not ignite the newsprint flashover indicators over an exposure time of 480 s, at which time the heating rate was increased to

494 kW. After another 45 s, the room flashed over. When a rate of 475 kW

was employed with the burner at the floor center, flashover occurred in

138 s.

27
Other flashover experiments have also been conducted in reducedscale models [14,15]. Fire tests in a similar sized quarter-size enclosure lined v;i th gypsum board [14] demonstrated that a full-scale equivalent

rate of heat production of 670 kW, maintained over a 300 s period, at the

floor center, could result in flashover of the room. Waterman [15]

performed similar experiments with an one-eighth scale model of a 3.7 x

3.7 X 2.4-m high room having an interior lining of cement-asbestos

board. Instead of a doorway opening, his enclosure had a scaled down

version of a 1.22 m wide and 1.37 m high window. However, the ventilation

3/2 factor, WH , discussed in section 1.2, was the same as that for our

model doorway. He found that about 650 kW along one side wall was needed

to flashover the space. Both of these previous studies employed models

lined with materials having much higher values of thermal conductivity,

k, as well as thermal inertia, kpc, (the product of the thermal conductivity, density, and heat capacity of the material). The thermal inertia

of a material determines the rate of the temperature rise of a surface

exposed to the fire. If the thermal inertia is low, the surface temperature then rises rapidly, and the irradiance from the heated surfaces

increases even faster. Thermal conductivity, on the other hand, determines the ultimate extent of the temperature rise. Gypsum and asbestos

boards have k values of about 0.17 and 0.11 W/mC, respectively, as compared

with 0.035 W/mC for fibrous glass. Furthermore, the room which was lined

with the cement asbestos board also had a 40 percent greater interior

surface area for thermal losses. Consequently, the heating rates for

room flashover in those studies were considerably higher than the rates

found with exclosures which were lined with the fibrous glass.

28
In general, the heat release rate needed for room flashover depends

on room construction and configuration; the type, quantity, and distribution of combustible materials or fuel; and, on the ventilation conditions.

Neglect of one or more of these factors could lead to serious error in

predicting the potential of a room for flashover. For example, one such

simplification method assumes that only the fuel heat release rate, q,

and the available air supply expressed in terms of the room ventilation

3/2 factor WH may be adequate for estimating the room flashover potential

[16] . The expression derived in that study can be given as

q

= 600 WH3/2

(kW)

3/2 5/2 where WH is in units of m .In the present test series with the

5/2 fibrous glass lining, a ventilation factor of 0.123 m was used in the

quarter-scale tests. Using the above formula, this ventilation factor

corresponded to 73 kW, which when multiplied by the scaling factor of 16,

corresponded to a full-scale counterpart of 1170 kW. However, the results

outlined in table 4 showed that a rate of heat release of only 300 kW

would be required for flashover of the room. In all practicality, this

300 kW represented a minimum value needed for flashover in the fire test

room. For more typical room lining materials having a higher thermal inertia, the heat release rates needed for flashover would be higher. For

room fires involving combustible interior finish and furnishings, instead

of just the methane or propane fuel source, the flames and combustion

products would have a higher optical density. The latter could help obscure radiation from the heated room surfaces and, hence, also result in

29
requiring a higher rate of heat release for flashover. Nevertheless,

even though the simplified formula was satisfactory for many situations,

it was entirely inadequate for the room fire tests performed here.

A procedure, which takes into account the physical properties of the

interior finish, room size, and doorway openings, is also available for

estimating room temperature as a function of heat production rate [17],

This method assumes that the room is at a uniform temperature, e.g.,

equal to the air temperature measured near the celling, and that the

thermal loss per unit surface area is uniform over the entire surface of

the room, including the floor and door areas. The expression given in

that study is as follows:

=

j

•{

1/2

• IhA A S wo o

where the heat transfer coefficient h is given as

and where, for the case of the full-scale room lined with fibrous glass

insulation.

A = Opening area, W x H (1.4 m^)

o ^ o o

’

A = Effective surface area for heat transfer including door area w

= 2(L xH+LxW+HxW) or (45.6 m^)

c - Specific heat of ceiling/wall material (0.8 kj*kg ^*K ^),

30
Cp = Specific heat of air (1.1 kj*kg ^*K ^)

,

g = Gravitational constant (9.8 m*s ^)

,

h = Effective heat transfer coefficient through ceiling/walls

(kW»m ^) ,

H = Height of room (2.3 m) ,

= Height of room opening (1.93 m)

,

k = Thermal conductivity of ceiling/wall material (0.035 kW*m ^*K

q = Heat release rate of fire (kW)

,

0 = Density of ceiling/wall material (60 kg*m ,

= Ambient air density (approximately 1.2 kg*m ,

t = Time of exposure (s)

,

T = Temperature (°C or K)

,

T^ = Ambient temperature (approximately 20°C or 293 K)

,

AT = Temperature rise relative to ambient

= T - T (“C),

o

’

W = Width of room (3.0 m) , and

W^ = Width of room opening (0.73 m)

.

For model test M14II in table 4, an air temperature of about 600°C was achieved at a time of 138 s. If the corresponding full-scale test

31
behaved in the same manner, then a fuel heat release rate, q, of about

450 kW would be obtained for the test using the above procedure. This

coir.pared well with the full-scale equivalent value of 475 kW actually

used. In model test M7II, an air temperature of 544®C was reached at

402 s. The calculated full-scale equivalent for this test would be 295 kw

compared with the 300 kW used.

4. SUMMARY AND CONCLUSIONS

1. The interior finish materials used for this study were nitrile foam

rubber, plywood paneling, and fibrous glass insulation. In addition,

data from another study [9] involving polyurethane, polyisocyanurate,

fiber board, and fibrous glass were used to support this work.

2. Empirical adjustment of the scaling criteria for the doorway with

regard to the quarter-scale model room fire test developed at NBS

resulted in closer agreement with full-scale room fire behavior. It

was necessary to lower the doorway opening in the model by 14 percent

to obtain flashover with the materials considered in this study and

with the model equivalent of the gas burner heat release rate that

produced flashover in the full-scale test. However, a theoretical

basis is still needed to justify the general use of this modified

doorway

.

3. The modified quarter-scale model room fire test can also duplicate

the peak fire buildup, as evidenced by the peak air temperatures

32
reached inside the room and at the doorway, in full-scale room fires

not having the potential for flashover. The duplication of the fullscale test results is close enough for screening purposes.

4. The modified model test still takes longer to reach the peak fire

buildup than does its counterpart full-scale test. Consequently, in

situations where the fire development in the full-scale test just

barely reaches flashover, the model fire may only approach flashover

conditions. In these cases, other indicators such as the peak

interior and doorway air temperatures in the model could be used to

help predict the flashover potential of the full-scale tests.

5. Corner placement of the ignition source appears desirable as the

room interior finish would be evaluated under the more demanding

fire exposure location than, e.g., at the center of the floor or

against the back wall, for any given rate of heat release for the

ignition source.

6. A minimum heating rate of 300 kW is required to cause flashover in a

3 X 3 X 2.3 m room lined with fibrous glass and having a 0.76 x 2.03 m

high doorway opening when the fire is located in one back corner of

the room. The heating rate required to cause flashover would depend

to some extent on the characteristics of the fire and, in most

cases, be higher for other wall finish materials.

33
7. The severity of the ignition source in room fire experiments should

be expressed in terms of some fraction of the minimum heating rate

required to cause room flashover. The source should be large enough

to adequately assess the fire hazard potential of materials, but

should not be so large as to overwhelm the material being evaluated.

5 . ACKNOWLEDGMENTS

Appreciation is expressed to Mr. Charles Veirtz, who built the test

room and performed the actual testing, and to Mr. W.J. Parker for his

helpful comments throughout the project.

6 . REFERENCES

[1] Lee, B.T., Laboratory Scaling of the Fluid Mechanical Aspects of

Large Fires, Combustion Science and Technology, 1972, Vol. 4,

pp. 233-239.

[2] de Ris, J. , Kanury, A.M. , and Yuen, M.C., Pressure Modeling of

Fires, Fourteenth Symposium (International) on Combustion,

pg. 1033, The Combustion Institute, Pittsburgh, PA, 1973.

[3] Kanury, A.M. , Pressure Modeling of Pan Fires with a Variety of

Plastics, FMRC Tech. Report 19721-9, 1973.

[4] Alpert, R.L., FMRC, Personal Communications.

[5] Waterman, T.E., Room Flashover-Scaling of Fire Conditions, Fire

Technology, Vol. 5, No. 1, Feb. 1969, pp. 52-58.

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

(International) on Combustion, pp. 1021-1030, The Combustion

Institute, Pittsburgh, PA, 1973.

34
[7] Parker, W.J. and Lee, B.T., A Small-Scale Enclosure for Characterizing the Fire Buildup Potential of a Room, NBSIR 75-710, Nat.

Bur. Stds. (U.S.) Wash., D.C. June 1975.

[8] Lee, B.T. and Breese, J.N., Submarine Compartment Fire Study-Fire

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

Bur. Stds. (U.S.) Wash., D.C. May 1979.

[9] Parker, W.J., Comparison of Selected Room Fire Tests with ASTM E84

and Other Laboratory Fire Tests - A Cooperative Program with the

National Research Council of Canada, NBSIR in preparation.

[10] McCaffrey, B.J. and Heskestad, G., A Robust Bidirectional LowVelocity Probe for Flame and Fire Application, Combustion and

Flame, Vol. 26, pp. 125-127 (1976).

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

Materials, NBS Technical Note 879, Nat. Bur. Stds. (U.S.) Wash.,

D.C. June 1975.

[12] Quintiere, J. and McCaffrey, B., The Burning of Wood and Plastic

Cribs in an Enclosure, Vol. I, NBSIR 80-2054, Nat. Bur. Stds.

(U.S.) Wash., D.C. Nov. 1980.

[13] Tu, K.M. and Babrauskas, V., The Calibrajiion of a Burn Room for

Fire Tests on Furnishings, NBS Technical Note 981, Nat. Bur. Stds.

(U.S.) Wash., D.C. December 1978.

[14] Lee, B.T. and Parker, W.J., Naval Shipboard Fire Risk Criteria - Berthing Compartment Fire Study and Fire Performance Guidelines,

NBSIR 76-1052, Nat. Bur. Stds. (U.S.) Wash., D.C. July 1976.

[15] Waterman, T.E., Scaled Room Flashover, Final Technical Report,

Contract DAHC 20-70-C-0308 , OCD Work Unit 2534G, (April 1971).

[16] Babrauskas, V., Estimating Room Flashover Potential, Fire Technology,

Vol. 16, No. 2, pp. 94-104, May 1980.

[17] McCaffrey, B.J., Quintiere, J.G., and Harkleroad, M.F., Estimating

Room Temperatures and the Likelihood of Flashover Using Fire Data

Correlations, Fire Technology, pp. 98-119, May 1981.

35
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36
Figure 2B. Quarter-scale room fire test at flashover

L

37
FULL

SCALE

DATA

TAKEN

102

mm

BELOW

LINTEL

MODEL

DATA

TAKEN

2^j

mm

BELOW

LINTEL

o

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3o '3dniVd3dM31

38

Figure

3.

Top

of

doorway

air

temperature

histories

for

several

full-scale

fires

and

model

tests

using

lintel

III
TEMPERATURE. °C

Figure 4. Room air temperature profiles at time of peak doorway air temperature

TEMPERATURE, °C

Figure 5. Doorway air temperature profiles at time of peak doorway air temperature

39
Figure 6. Comparison of full-scale and model room fire test on plywood

40
Tests conducted over

temperature range

from 24 to 26°C

0 l ^ J ^ \ \ I L.^ ^ 0 20 40 60 80 100

RELATIVE HUMIDITY, PERCENT

Figure 7. Flashover time versus relative humidity for

model room fire tests with plywood paneling

41
3o 'BdnivdadHBi

A2

Figure

8.

Room

air

temperature

histories

for

full-scale

fire

P2

and

model

fires

M2I,

M2II,

and

M2III
43

Figure

9.

Ceiling

temperature

histories

for

full-scale

fire

P2

and

model

fires

M2I,

M2II,

and

M2III
Figure 10. Room air temperature profiles at 240 s for

full-scale fire P2 and model fires M21, M2II, and M2II1

44
100

Figure 11. Doorway air temperature profiles at 240 s fo

full-scale fire P2 and model fires M2I, M2II, and M2III

45
PERCENT

DOORWAY

HEIGHT

50

100- Gi

' »: • a

90- CA A

•

80- ^ A

•

70- £> A

•

60-

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- -40

% -30

^ 4k - -20 • M2 1 O M2 II ^ - .in A M3 III A P2

A

1 1

() 50 100

VELOCITY, m/min

150

Figure 12. Doorway flow velocity profiles at 240 s for

full-scale fire P2 and model fires M2I, M2II, and M2III

46
(56)

-r

(74)

' Gas burner

I L

1391

[ |56|

I

I

I

I

(52)

I (43)

I

I 1

I Fluxmeter

( ®

(45) I (69)

I

I

Burner rate of heat release - 500 kW

Times in parenthesis given in seconds

Figure 13. Ignition time of newspaper flashover

indicators as a function of floor location

47
Figure 14. Model room air temperature profiles at 240 s for three burner locations

48
Figure 15. Full-scale room air temperature profiles at 240 s for three burner locations

49
1H9I3H AVMMOOQ lN33d3d

QJ^ M

3 C

4J 0

ro -H U +j

<U (0

D-1 U

e O

QJ 1—

I

-P

OoUJ

oe

LU

CL

1H9I3H AVMdOOQ lN33d3d

50

Figure

16.

Model

doorway

air

temperature

Figure

17.

Full-scale

doorway

air

profiles

at

240

s

for

three

burner

locations

profiles

at

240

s

for

three

burner
PERCENT

DOORWAY

HEIGHT

-100

-90

-80 ®

-70

A- -o#

- -50

AO - -40

ap - -30

£S» -20

10

O •

om ^

o M2 It, corner • M3 II, back ^ M4 II, center

-50 0 50 100 150

VELOCITY, M/min

Figure 18. Model doorway flow velocity profiles at 240 s for three burner locations

51
PERCENT

DOORWAY

HEIGHT

tIOO

-90

O A

O A •

--80 o

^

-70

A

A

AO

mo

£mo

60'4- O

50

40

430

20

4 10

-50

A »

o P2, corner • P3, back

A P4, center

—j I i_

50 100 150

VELOCITY, m/min

-X

200

Figure 19. Full-scale doorway flow velocity profiles at 240 s for three burner locations

52
900

800

700

600

500

400

300

200

100

0

20. Model and full-scale room air temperature

histories for fire source at back corner

53
800

700

600

500

400

300

200

100

0

Lgure . Model

f and full-scale ceilingor fire source at back

54

temperature

corner histories
Figure 22. Model and full-scale doorway air temperature histories

for fire source at back corner

55
Table 1. Interior finish materials used in room fire tests

Thickness Density

Material (mm) (kg/m^)

Nitrile Rubber Foam* B2 27 90

Nitrile Rubber Foam* C2 27 120

Fibrous Glass Insulation with

Glass Cloth Facing

25 60

Printed Lauan Plywood Paneling 4 510

Intumescent Paint 0-987 0.25 900

* Polyvinyl chloride acrylonitrile butadiene closed cell foam

56
P4

CO

4-» P4

cO

3) •K

H CD'

mO VD

M 1 m fa CD 00 M CO t—

'O’ •

31 o O

00

fa

Co

•H

4J

CO tH3 ^ N COC CO M •p

3i

CO cO

CO 3)

a

CO

M O'

m CJ^ vO CO I—

M 1 Cs| 'O' O C3^

CL5 CO 1—

vt •O’ • 31 o w CO O C 3) o o

»-l 'O’

m • C •H rH

(U fa 6

•H 13

T3

C

cO

cd

B & cO

u

O <0 o fa u

a o p *

13 !-i cO a-\ CO vO o

Q)

M 31 00 00 vO o

iH ,J3 o 'O’ CO • (U 43 1 1—!

13

d CO

O fa 3l S c5

Q) • •H

CVJ

4-1 (U •H

!3 -Q CO cO 4-1 tH CO (0 H CO cO N N

CU

O •H cO o CvJ •H

H C/1 C/1 LO CO CO C/1 1 1 m C' t O r—

fa fa fa

4-) fa fa !—

3 3 3J 13 fa fa fa

(U

CO 33r4

(U 13 P O K 43 S 43 60 P g ^—\

•H

fa

g

g CO

CO

gg 33

Q

g '—

'w' P Po

T—

43 43 cd

CO 43 60 P |S P P •H 13 U

ti fa (0 *H o

•H CO 33 !3 o

fa

Q

n

kO fa 13 1—

cO cd 13

CO CO CO & S CO

4-1 fa

p P p fa

(3 cO 31O o cd *H o o o CJ hJ CO fa Q Q CO

It

>-l (1) 4J

cd •K

3) fa O' CT. r*H CNI

m

fa 1 r-.m <T\ 00

fa CO <T • 31 o oNw^

00

(UP

cd •K

31

K

fa O' fa CN vD CO

fa 1 CO fa (N cr,

(U fa 'O’ 'O’ • 31 o

60

O

3l •H

fa 'O’ CU • 31 !—

cd fa 60 13 31 O •W O fa & fa

P^ P CU fa CU C3 fa p •K p a L|-l P cd cs 00 1—

o o

CO fa 3 o o 00 o

CU o fa in CO # CU H fa o

CO cd

O c ip P o P

31 13 CO

CU

CO o

33p 1—

fn

(U CU CU (U cd 13 1—

t—

I—

5 O cd cd cd P S o o vO CM CM o o

C/d CO o CO v£) c/1 o 1 1 'O’ o fa 1 13

fa fa CM fa fa 1—

fa CU 3 3l 31 43

fa fa fa P

<+-l O Vj

0)

CL, 4J O t-i P p cd

a 43 33 43 / V 60 <U W P g •pH 43 1 fa /''-S g g (U P CU

CU V-^

g 33 31 o P fa B o o 1—1 43 43 cd p CO

CU 43 60 P & fa CU p P •H 13 P E

31 fa CU •pH o (U ‘p •H CU 33 12 o CJ P

hJ

Q

fa

33 (3cd CU I—1 13 fa cd cd 13 p d

0) CU CU S & cU CO fa

4J fa P P P fa •pH cd d cd 31 o o cd a >

•H o •H o o o •K -sc hJ C/l Q Q C/l 57
Table

3.

Summary

of

full-scale

tests

and

counterpart

model

tests

of

nitrile

rubber

foam

and

plywood

paneling

c cu o O

•H O O CTJ U U

•H 0) CO tH C e T3

OO -H nH G

j-l 4-1 U-l -H

o o fo r>* vD fO-vT-xi <»

r«-

CTn

K

•K vo iT> moo m CO *<r iH m

iH iH CO CO ro

0) E-^ 6 ^ •H O CO H 4-> w

o o 00 CO I O no CO

<r

CO r«» O CO \D r>. 00 o CM CO vo m

th <r CO m

o o o o

I CO m I-* m

iH cn CO cn

•X

CC3

CC3 S

y>o

o

T3

H

CM 00

r- I a^ m m

o m m

CO .H 00

vo CM m

^ m

m <y\ <r MO

MO 00 O

0^ 00 o

iH CM m

00 I—

I

0> 00 .H CM CM CO

CO m vo CO

v£i 0^ >—I CO

'4' \0 '>0 \D

OJ H

e ^ •H O CO H w ^

O CM vO O O CO rx. CO <r m CO o "<r MT MOCMCMO CMCMCM COO

co-Mj-vocr* Ovoco rx.00

r-i <r Mf <r iH iH O ^ O 00 vD

CM SJ- CO 0> iH

r-H rH «-H f~l CM

0> CX^ ago a 0) o

Mt pxx CO

px. <r 00

px. vX> v£)

CM 00 px. O CM

CO CO xa-

\D O

0^

Px. px.

CO;|cM iH m x3- OvlcO px. ON xa

px. px. >X> sO

CUeH

H

3O

OJg

CO

M-i

COe

o ^O CO XJ 'x^

px. m x3- CM CO x4- CM cMx^-m co8co*4- xj

00

CO 8 V o 00 00 CO m o 00 ON CO mm CM iH m

rH iH rH rH CM CM CM

(Uc

M

3

cn

cO C3

C

o

•H CU AJ O

•H H /-X C 3 S

00 O ^

•H CO

CMCM CMCMCM CMCMCM

\0 \0 \Q yQ yO

xd- -xT o O O O O

ON ON O' ON ON ON

CM CM

v£> rH rH sO rH

CO (0 fO o o o

CO CO CO

'O ’O 'O

(U CU CM O) —I rH vD rH

CJ o o

CO CO CO

CU <u xa- «u QJ rH rH On rH rH

CO cO CO CO

CJ O CJ o

CO CO CO CO

T3 TJ O CU O CU ON rH ON rH

CO CO a u

CO CO

"S 'S *s

u 3 3

CO CO CO

px.px.rx. O O O CM CM CM

O O O O O O O O

c

o

•H

3

COc

CO 4-» g -S) O -H O 0) Q

<U

COo

CO

CM CM CM U U OCM CM CM CM o c_> o u

CM CM CM CM

p::) P3 CO CO

Kil

CM CM CM

CO CO CO

T3

(U <U rH N CO •H O

cn TU CO

I 0) rH rH m

rH CO ON 3 U • lx, CO O

T3 T3

(U <u

0) rH rH

CM cO cO •H O O

Cfl T3 CO CO

I 0) rH rH m vO> rH CO ON C» 3 a • • Ph CO o o

*3 TJ *3 TJ (DO) <u 3 3 rH rH 3 rH rH N 3 3 N 3 3

•H CJ O *H O CJ W T3 CO CO W CO CO

I 3 I rH rH fO NO rH m nO

rH 3 ON CO rH ON CO 3 U • • 3 • • Pm CO O O [i, O O

rH xa- x3- <r 3 'Xx, -Xx Xx-. Cm *-H rH rH

CO m

Cm NO CM

CO CM ro

Cm xd- CM CO

cn xaI 1 CO O xd- x3- c/3 xa- rH

Cm rH CM ro Cm xo- m

Oh P.H

*33

3 rH N 3

•H CJ W CO

I rH O

rH ON 3 • Cm O

3 Xxx Cm rH

CO I & §

Pm cu

^ I s §

cu CU

"o 'o o

o o o III

cu cu cu

T33

rH m m

3 ON 00 o • • CO o o

xd- xd- «d- Xx. Xx^ XXx.

Pm Pm Cu

POC

S g u

CJ O 3

> m m o

« w

3 3 rH 3 3 C s s o

i!i

i

58
Table

4.

Summary

of

flashover

times

and

temperatures

in

room

fire

tests

with

exposed

fibrous

glass

insulation

lining

the

walls

and

ceiling

•K 111 111 111 111 111 111 •K 1 1 s 1 1 1 >1 B 1 1 1 1 1 1 fd 0) -- 1 1 1 1 1 1 S Eh U 00 CN ro r- ro KD 1 ro (Ti 00 in 1 o in o 00 1 r- o 1 ro o <T 1 <T 1 ro

Vh o ON r- 00 (M I-H r- 1 rH in CO tN 1 r- (Ti r- ID 00 1 o r- 1 <D ro 1 CM CM 1 00 0 U ^ I-H rH CM ro ro CM 1 ro CM CM CN 1 ro ro ro ro 1 1 in in 1 in 1 in 0 •H 1 1 1 1 1 1 Q C 111 111 111 111 111 111 K 11 11 11 11 11 11 K 1 1 1 1 1 1 Cn • 1 1 1 1 1 1 C a.— 1 1 1 1 1 1 •H B u (N o ro in 1 <y» vn m 1 00 CM 00 I-H 1 CM CM 1 in in VD in 1 in 1 CM 1—

tUo 1—

<y\ I-H 00 '<D ro 1 ro I-H 1 1 CM rH VD ro 1 CM 1 r-H o VD iH 1 1 (T •H Eh ^ CM I-H m ro ro ro 1 ro CM ro 1 ro in 1 in 1 VD in r- 1 in 1 in

tu 1 1 1 1 1 1 u 111 111 111 111 111 111

ft O. 111 111 111 111 111 111 B 1 1 1 1 1 1 B (U U in a\ m 00 ro 1 o CM 00 in 1 CM in (T in 1 1—1 CM 1 o a^ rH o 1 CM rH 1 00 0 Eh o IT) o CM I-H 1 00 o I-H c» 1 00 ro ro VD 1 (T r- 1 ro I-H CM CM 1 I-H 00 1 rH o CM CM ro ro 1 ro ro ro CM 1 in in <D in 1 in VD 1 VD r- r- 00 1 VD in 1 ra; 1 1 1 1 1 1 •H 1 1 1 1 1 1 <c 111 11 111 111 111 111

>H 11 11 11 11 11 11 1 1 1 1 1 1 > 1 1 1 1 1 1 0 0) -- <u Q) <U (U CU Q) 1 (U 0) 0) OJ 1 (D 0 CM CM 00 1 o VD 1 00 CM o 00 1 O 00 1 m

s:

B to c c C c c C 1 c c C

C 1 c C o 00 I-H 1 CM 1 ro o (T r- 1 00 ro 1 to H 0 0 o 0 0 0 1 0 0 0 0 1 0 0 ^ CN ro 1 CM rH 1 rH r-H 1 I-H 1

(t) Eh 2 2 2 2 2 1 :s 2 2 2 1 2 2 1 1 1 A 1 rH 1 1 1 1 1 1 111 111 l11 111 11 111

CM 11 s1 11 11 11 11

.. s 1 1 1 1 1 1 o 1 1 1 1 1 1 4-1 r-| 1 1 1 1 1 1

(t) td 1 1 1 1 1 1 0) o 1 1 1 1 1 1 K to 1 1 1 1 1 1 — , 1 1 1 1 1 1 4H S (N (N o o o o 1 o o o o 1 o o o O o 1 in O 1 O o O o 1 O in 1 in 0 X VD 1 1 o o "ir 1 r- rH 1 VD VD VD VO 1 VD r*- 1 I-H I-H r-H f-H 1 r-H fH 1—

rH 1 CN CN

m ro ro 1 ro 1 1 1 VD

(U (U 1 1 1 1 1 1 4-1 to 1 1 1 1 1 1 m Id 1 1 1 1 1 1 a ti) 1 1 s 1 1 1 1—

1 1 1 1 1 1 (U 1 1 1 1 1 1 pi, 11 11 11 11 11 11

in 111 11 11 11 11 11 (U 4C 1 1 1 1 1 1 C to o o o u 1 u PU o 04 1 04 04 Oh 04 CU 1 Ph Ch 1 CU CU p M 1 CU CU 1 CU

>H fd 2 2 2 2 1 2 1 1 1 1 1 3 U 1 1 1 1 1 1 m 11111 Sh Sh 11111 11111 11111 11111 Jh 11111 1 r-H I-H <U <U 1 1 I-H I-H 1 1 (V <D 1 c 1 I-H I-H 4-1 4-1 1 1 I-H r-H 1 1 1 Jh 0 u >H IH >4 >H Jh 1 cd fd C C 1 n Sh Sh S4 Jh 1 (d fd 1 Jh S4 Sh 1 a C 1 Jh (U -H (U 0) 0) 0) tu 0) 1 s s Q) Q) 1 0) (U (U tu (V 1 S 2 1 <v (U CJ 01 1 0) 0) 1 (V C 4-1 c c c c c C 1 U U 1 c C C c a 1 1 c c c C 1 u u 1 C M 'H )H >H V4 >H Sh Jh 1 1 Sh Sh Sh Sh Jh 1 44 44 1 Jh Sh Sh 1 1 H 3 to 0 0 o 0 0 0 1 O o Sh Sh 1 0 0 0 0 0 1 u o 1 0 0 0 0 1 Jh H 1 0

CQ 0 u

u

uuuU 1 fd fd 0 0 1 uuUUU 1 fd fd 1 uuUU 1 0 0 1 U 1 cQ m 0 0 1 1 CO OQ 1 1 0 0 1 1 1—

rH 1 1 1 1 rH rH 1 111 b 111 111 11 11 Ph 11 1 (U 111 111 111 11 11 11 i-H f—

1 1 1 1 CM 1 1 1—1 Id I-H CM 1 ro 'S' 1 m 1 1 rH 1 1 3 U

ti in

Oi 11 04 11 Oi 11 11 CU 11 11 4-1 1 1 1 1 1 1 to 1 1 1 1 1 1 (U 1 1 1 1 H 1 1 Eh >

1 H 1 1 1 H H 1 H H 1 H H 1 H

(U

HH

H 1 I-H

H 1 H

HHH 1 MH 1 H

M

H 1 H

H 1 H

13

H

H

HM 1 H

H 1 HHHH 1 O r-H 1 CM CM CM 1 ro 1 m 0 1—

CM CM (N 1 ro 1 ID r- 00 <T 1 I-H rH 1 rH I-H rH 1 rH rH 1 rH 2 2 2 2 2 111 S 2 111 2 2 2 2 111 2 2 111 2 2 2 111 2 s: 111 2

59

*NG,

P,

and M

refer

to

natural

gas,

propane,

and

methane,

respectively.

**At

time

of

flashover,

otherwise

at

240

s.

For

test

M13II,

temperatures

given

at

240
Table

5.

Comparison

of

fire

buildup

in

full-scale

and

corresponding

quarter-scale

room

fire

tests

of

nitrile

foam

rubber 1

-Kcu3

XI

CM t—1 u S-i S-i S-I S-I S-I S-I S-I o "H 0) d d d d d d d 3 > > > > U > > > CJ CP > CP CP CP

01 pq o o o o o o o O o 0 O o o 0

0) 4P pp P3 pp LO pp pp pp VO 00 PP r-. 00 CTi d d d d d d t—1 d d d (Ti 00 d a\ 00 a\

00

dd

d

d CMddd 1—1

CM

d CM CM CM

O *H r—i I—1 1—i 1—1

.H 1—1

I—1 1—1 O [Ml Pm Pm Pm Pm Pm Pm Pm Pm do 00d

Ml •H

Ml pp PP pp

d Ml 13 2: 13 :3 13 13 13 ;3 13 ;3 00 00 00

00 0)

OO

OOOOO

O o O

O •H •H

1—1

cn i-J hP hP i-P i-P hP i-P hP t-p I-P I-P PP PP PP

0)u

XI XI XI Xl X) XI

pj J

dd

ddddd

0) 1—1 d 1—1

1—1 d 1—1 1—1 d 1—! 1—1 N d N d d N d d N d d Q) 'H d •H d d •H d d •H d d Pj

d Ml ca Xl cq CO X3 CO CO CO X3 CO CO CO CO CO CS

5; x: 1 d 1 d 1 d 1 fSQ) d 00 rH 1—I m 1—1 1—1

CO MP 1—1 I—1cn vD r—l cn VO o -H I—1 d CT^ rM d o^ 00 r—1 d CJV 00 1—1

<3V 00 o 01 3 d • 3 d • • 3 d • • 3 . • Q Pd Pm c/n o Pm CO o o Pm CO o o Pm o o •H

od

d 1—1 :§ 01

H

M

M o u Ml M H M M M M M o o d 1 M M 1 M M H 1 M M M 1 M M oo

'H

C

03

4-1 1—1

CM on

CO • 1 1 1 1 0) o CO CO CO CM ro CO o CO 1—1 Eh S Pm vo csi Pm <3- CM CO Pm I—1

CNI CO IT)

v-i o

•u

njo

•H

X)d

•HU

q;>o

00 4=: 3 (fi 'iP 1^ cO Ml o d d 00 d CM 3 3 ov U-i o 1 o o 1 CP <d la 2 O U-lo do

•H

J-J d

I—I 3wdM

CNo O

CM CM

pq PQ

do

•H

4-J •Hd

bO

•Hdo

XI

OJ

cfi d

pq

•K 60

If
Table 6. Flux measurements at time of flashover

Test

Time

(s)

Floor

(kW/m^)

Upper Left Wall

(kW/m^)

Upper Right Wall

(kW/m^)

Ceiling

(kW/m^)

FS-1 30 17.3

23 64 20.0 49 55 50

FS-2 30 29.5 —

33 47 25.3 48 76 54

FS-3 46 20.0 — —

24 186 2.8* 7.4* 6.7* 6.6*

34 97 25.2 56 67 52

FS-4 468 1.9* — 44 360 4.6* 7.5* - - 51 552 5.6* 15* 13* 14*

*Flashover did not occur. Values given at the time of peak doorway

temperature.

61
Table 7. Comparison of room air temperatures for fire

tests with three different doorway openings

Test

Doorway

height Insulation

Source

setting

(kW)

Max. interior

upper air temp.

(°C)

Max. doorway

air temp.

(°C)

Flashover

times

(s)

4 Scaled C2 3.9 216 215 00 5 * Scaled C2 3.9 293 213 00 16 Scaled C2 5.9 410 240 00 17** Scaled C2 5.9 477 240 00 22 0.93 Scaled C2 3.9 707 585 43

33 0.86 Scaled C2 3.9 604 451 47

10 Scaled B2 3.9 221 196 00 19 Scaled B2 5.9 756 579 70

24 0.93 Scaled B2 3.9 410 288 00 34 0.86 Scaled B2 3.9 646 500 97

*Repeat of test 4 **Repeat of test 16

62
Table

8.

Summary

of

heat

fluxes

in

room

fire

tests

with

fibrous

glass

insulation

u XCM CO m CJN vX5 00 00 CO vD o^ CM in <r CO O CM m O O CO O 1 OS CM o

p o < m CO <r rH CO in in CO OS 00 CO CO OS O 00

(OH rH rH rH rH rH rH rH rH rH rH rH CM rH

1

1

CMU O CO X o • • • 0 CO 1 1 1 1 o rH <r CO 1 1 1 1 1 c CM CM CM CM

QJ (N CO E

<T1 :s s rH

4J }-< o CM 00 00 rH CO l—i CO in 00 CM CO 00 in CO 00 'd’ 1—

!

rH <3- CO O rH o • W) CO CM CO 00 o o o r*- o rH CM CO vo as os O CM OS CM CM

•H C CM CM CM CM CM CM CM in CO CO CO CM m

11 CO

CM

}< u 00 in in rH X o • • • • • • 3 CO 1 I 1 1 rH CO 1 in rH \ 1—

1

c rH rH rH rH rH CM

cu CM CO

I—I E

rH

XJ u rH r**. •vT rH <r CO n o rH O in CM CM rH r>. CO 'd- O O rH CO

4-1 o • Q) CO CO CM 00 00 CM a^ rv cr. CM <r <r rH m v£> o o vD CO CM sD

H-l c rH CM CM CM CM CM CM CM CO CO CM CO CO

cu CO

00 /-~s c •3C CM •H X E rH ON Mt r>* CM CM v£> (T» in vD in rH Os 00 'd- Os in CO rH

f—

1

a • • •H rH S LO o CM o in O CO O m rH o o **d‘ o <r CM CM rH O CM OS 00

Q) Uu A! rH rH rH rH rH rH CM CM CM CM CM CO CM CO CO in in CO CO U

U

0) 1 > O CU CU CU (U (U <u CU (U (U cu (U CU CM CM 00 o vD 00 CM O 00 o 00 in o E / ^ c c C C d d d d d d d d O 00 rH MJ- CM CO o ON 00 CO 1 •H CO o o O o o o o o o o O o •«d’ CM CO CM rH rH 'd* 1 w H '«—^ IS :a a a a A 1 cd 1^ 1 1

1

1

1

1

CM 1

1

1

1 /-~s I 1 0) 1 1 4J rH 1 1 1 cd Cd 1 I 0) O 1 1 1 « 1 1 1 N-X 1 1 1 13 CM CM o o o o o o o o O o o O O n O 1 O o O o 1 o m 1 in 0 X vO 'd* Mt MT O r*^ o <3* rH 1 vD O nD 1 1 rH rH rH rH rH rH CM CM CO CO CO CO 1 Mf 'd- 1 'd- 'd* 1 NO

0) cu 1 1 1 4J CO 1 1 1 cd cd 1 1 1 oi 0) 1 1 1 rH 1 1 1 0) 1 1 1

u Pi

1

1

1

1

1

1

1

1

1

1

1

1

1

1 Pi u

1

1

1

1

1

1

1 rH rH cu <U rH rH 1 1 CU cu 1 c rH rH 4J 4J rH 1 1 4-1 4J 1 u o U u u u u u 01 Cd d d Pi Pi Pi U Pi Cd Cd 1 Pi Pi u u 1 d d 1 u

CD •H CU 0) 0) cu cu cu 5 a cu CU CU <U (U <U CU a a 1 (U <U <u <u 1 lU (U 1 cu C 4-> C c d c d d u CJ d d d d d 1 d d d d 1 CJ CJ 1 d

u •H U M U u u u Pi Pi Pi u .ii 1 Pi Pi Pi Pi 1 ( u

;3 CO O o 0 O o 0 o u Pi U o O O O o O O \ O O O O 1 Pi Pi 1 n

PU O CJ u CJ CJ u CJ cd cd O O CJ CJ U CJ u Cd Cd 1 U U CJ CJ 1 O O 1 u

CP CP 0 0 PP PP 1 1 o O \

\ 1 \ Pm a 1

1

1

1

1

1

PP 1

1

1

1 CU

1

1

1

1

1

1

1

1

1 1—

1

rH 1 CM 1 1 rH Cd rH CM CO vt in 1

1—

1

1 1 r CJ Pm PH PH 1 Pm 1 1 U-i CO 1 1 1 4-J

1 1 1 (0 \ 1 I <u I 1 1 H 1 M 1 1 rH M M M 1 hH M 1 M M 1 PH

(U h- M M H M M M M M M M i H M l-H 1 PH PH 1 PH

Td M f- M M M M M HH M O rH 1 CM CM CM 1 CO <r 1 in c r“ CM CM CM CO CO Os 1 1 rH 1 rH s a a a a a a a a a a a a 1 a a a 1 a a 1 a

63

*Data

at

time

of

flashover,

otherwise

at

240

s.

For

test

M13II,

data

given

at

240

s.

Sensor

1

at

3/4

room

height.

Sensor

2

at

5/8

room

height

.
Table

9.

Doorway

flow

velocities

in

fire

tests

with

fibrous

glass

insulation

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rC

bO

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pC4 r—1

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CN 1 1 1 1 1 <44 > CJ 4-) 1—1

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dd 1 *H •H 1 P4 1 o QJ

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d J-i P4 1 & S 1 P4 1 Po Po Po 1 Po 1 QJ ( P4 CJ rH

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d •H •H •H •H •H 1 rH rH 1 *H 1 O o o 1 O o 1 -H 1 •H Pm o H 3 1 Pm Pm 1 Pm 1 PM PM Pm 1 PM PM 1 Pm 1 Pm •K -sc •SC 1 1 1 1 1 1 SC 1 1 1 1 1 1 •Sc 65
NBS-114A (REV. 2-QG) U.S. DEPT. OF COMM. BIBLIOGRAPHIC DATA

SHEET (See instructions)

1. PUBLICATION OR

REPORT NO.

NBSIR 81-2453

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

March 1982

4.

TITLE AND SUBTITLE

Quarter-Scale Modeling of Room Fire Tests of Interior Finish 5.

AUTHOR(S)

B. T. Lee

6.

PERFORMING ORGANIZATION (If joint or other than MBS. see instructions)

national bureau of standards

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

7. Contract/Grant No.

8. Type of Report & Period Covered

9.

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

Sponsored in part by;

Ship Damage Prevention and Control

Naval Sea Systems Command

Department of the Navy

Washington, DC 20362

10.

SUPPLEMENTARY NOTES

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

11.

ABSTRACT fA 200-word or less factual summary of most significant information. If document includes a significant

bi bl iography or literature survey, mention it here)

A technique for modeling fire build-up in rooms with combustible interior finish was refined to achieve closer simulation of full-scale fire development. Fire experiments

were performed in one-quarter scale model rooms and full-scale rooms having a doorway

opening. The interior finish test materials were nitrile foam rubber, fibrous glass,

and plywood; a gas burner was employed as the fire source in a rear corner of the room.

It was necessary to lower the doorway opening in the model by as much as 14 percent to

obtain flashover with the same equivalent heating rate that produced flashover in the

full-scale test. At the same time, the width of the doorway in the model was increased

appropriately to maintain the same volumetric air flow rate. The effects of burner

location and heating rate on flashover in a well-insulated room were also studied to

help select a suitable ignition source size and placement for testing of interior

finish materials. The minimum heating rate needed to cause flashover ina 3x3x2. 3m

high room lined with fibrous glass and having a 0.73 x 1.93 m high doorway opening

would entail placement of the heat source in a back corner with the source having a heat release rate of 300 kW. A corresponding rate for the quarter-scale room would be

19 kW.

12.

KEY WORDS (Six to twelve entries; alphabetical order; capitalize only proper names; and separate key words by semicolon s) Fire growth; flashover; heat release rate; physical modeling; room fires; scale

models

13. AVAILABILITY

Unlimited

I I For Official Distribution. Do Not Release to NTIS

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

71

15. Price

$9.00

USCOMM-DC 6043-P80