Reference

NB3

Publi - cations

NBSIR 80-2120

A 1 1 10 S Dfl 75M 4

Fire Development in Residential

Basement Rooms

J. B. Fang and J. N. Breese

Center for Fire Research

National Engineering Laboratory

National Bureau of Standards

U S. Department of Commerce

Washington, DC 20234

Interim Report

October 1 980 — QC

100

,1156

30-2120

1980

Division of Energy, Building Technology and Standards

Office of Policy Development and Research

U.S. Department of Housing and Urban Development

Washington, DC 20410

NBSIR 80-2120

FIRE DEVELOPMENT IN RESIDENTIAL

BASEMENT ROOMS

J. B. Fang and J. N. Breese

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Center for Fire Research

National Engineering Laboratory

National Bureau of Standards

U.S. Department of Commerce

Washington, DC 20234

Interim Report

October 1980

Prepared for: Division of Energy, Building Technology and Standards

Office of Policy Development and Research

U.S. Department of Housing and Urban Development

Washington, DC 20410

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

Luther H. Hodges, Jr., Deputy Secretary

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

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director

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TABLE OF CONTENTS

Page

LIST OF TABLES iv

LIST OF FIGURES . • v Abstract 1

1. INTRODUCTION 2

2. EXPERIMENTAL DETAILS 4

2.1 Test Compartments 4 2.2 Fire Load Density 4 2.3 Furnishings 6 2.4 Interior Finish Materials 7 2.5 Forced Ventilation 9 2.6 Test Parameters 9

3. TEST MEASUREMENTS 10

3.1 Temperature 10

3.2 Heat Flux 11

3.3 Air Velocity at Doorway 12

3.4 Static Pressure 12

3.5 Optical Density of the Smoke 12

3.6 Gas Concentrations 12

3.7 Weight Loss 13

3.8 Flashover Indicators . . , 13

3.9 Data Acquisition 13

4. TEST PROCEDURE 13

5. RESULTS 14

5.1 Effect of Induced and Forced Ventilation 19

5.2 Effect of Fire Load Density 20

5.3 Effect of Lining Materials 21

5.4 Effect of Room Size 22

5.5 Effect of Ignition Source 22

6. DISCUSSION 22

6.1 The Experimentally Derived Temperature-Time Curve 22

6.2 Comparisons of Different Types and Measuring Locations

of Thermocouples 24

6.3 Doorway Flow Due to Ventilation Controlled Fires 26

6.4 Gas Temperatures Influenced by Type of Ceiling Materials . 27

6.5 Incident Heat Fluxes Calculated from Surface Temperatures . 28

7.. CONCLUSIONS 28

8. ACKNOWLEDGMENTS 29

9. REFERENCES 30

APPENDIX A - DERIVATION OF THE EQUATIONS USED FOR THE HEAT RELEASE

RATE CALCULATIONS ..... A-l

APPENDIX B - ESTIMATION OF THE INCIDENT HEAT FLUX B-l

iii
LIST OF TABLES

Page

Table 1. Types of interior finish materials found in basement

rooms 31

Table 2. List of combustible contents 32

Table 3. Experimental conditions of basement room fire tests ... 33

Table 4. Initial weights and total available heat of combustible

loads used for fire tests 34

Table 5. Test log during test 14 35

Table 6. Summary of critical times, heat release rates and gas

temperatures 38

Table 7. Summary of incident heat fluxes 39

Table 8. Summary of air flows, static pressures, smoke densities

and gas concentrations ' 40

Table 9. Average upper room gas temperature and heat release rate

at time of room flashover 41

Table 10. Average room gas temperature, heat release rate, flow

rates and calculated orifice coefficients for inflowing

and outflowing gases during active burning period .... 42

IV
LIST OF FIGURES

Page

Figure 1. Construction details of the test room 43

Figure 2. Layout of furniture and wall linings for the 3.3 x 3.3 m

test room 4 4 Figure 3. Layout of furniture and wall linings for the 3.3 x 4.9 m

test room 4 5 Figure 4. Frequency distribution of fire load data for residential

recreation rooms 46

Figure 5. Frequency distribution of fire load data for utility

rooms 4 7 Figure 6 . Frequency distribution of fire load data for other

basement rooms 48

Figure 7. Frequency distribution of the movable fire load density

in basement recreation rooms 49

Figure 8 a. Plan view of the 3.3 x 3.3 m test room showing

instrumentation layout 51

Figure 8 b. Plan view of the 3.3 x 4.9 m test room showing

instrumentation layout 52

Figure 9. Instrumentation layout - view of doorway and east

(front) wall showing gas thermocouples, velocity probes,

and smoke meters 5 3 Figure 10. Instrumentation layout - view of west (back) wall

showing wall thermocouples and heat flux meters 54

Figure 11a. Instrumentation layout - view of south wall of the

3.3

x 3.3 m test room showing wall thermocouples, heat

flux meter, and observation port 55

Figure lib. Instrumentation layout - view of south wall of the

3.3

x 4.9 m test room 56

Figure 12a. Instrumentation layout - view of north wall of the

3.3

x 3.3 m test room showing wall thermocouples .... 57

Figure 12b. Instrumentation layout - view of north wall of the

3.3

x 4.9 m test room 58

Figure 13a. Average gas temperatures as a function of time for tests

I, 2, and 4 59

Figure 13b. Average gas temperatures as a function of time for tests

5, 6 , and 7 Figure 13c. Average gas temperatures as a function of time for tests

8 , 9, and 10

Figure 13d. Average gas temperatures as a function of time for tests

II, 11A, and 12 62

Figure 13e. Average gas temperatures as a function of time for tests

13, 14, 15, and 16 63 v
LIST OF FIGURES (continued)

Page

Figure 14. Time variation of average gas temperatures as a function

of room ventilation 64

Figure 15. Time variation of total heat release rate as

a function of room ventilation 65

Figure 16. Time variation of average gas temperatures as a function

of fire load density 66

Figure 17. Relationship between heat released within fire room

during active burning period and total heat content of

combustible materials 67

Figure 18. Time variation of average gas temperatures as a function

of room lining materials 68

Figure 19. A comparison of heat release rates with various interior

finish materials 69

Figure 20. Time variation of average gas temperatures as a function

of room size 70

Figure 21. Time variation of heat flux incident at the center of

the floor as a function of room size 71

Figure 22. Time variation of average gas temperatures as a function

of ignition source 72

Figure 23. The range of variation in average gas temperatures for

residential room fires 73

Figure 24. A comparison of results obtained from ASTM E 119 and fast

response thermocouples for average gas temperatures ... 74

Figure 25. Comparison of fire exposure curve and the standard

ASTM E 119 fire exposure curve 75

Figure 26. Comparison of average gas temperatures as measured by

different types of thermocouples 76

Figure 27. A comparison of gas temperature averages calculated

using 27 and 3 locations . 77

Figure 28. The influence of the combustibility of the ceiling

materials on the vertical gas temperature profiles ... 78

Figure 29a. Comparison of the calculated and measured total heat

flux incident on the gypsum board ceiling 79

Figure 29b. Comparison of the calculated and measured total heat

flux incident on the gypsum board walls 80

Photo 1. Spread of fire over sofa cushions 3 minutes after

ignition in test 12 81

Photo 2. Full involvement of room contents 4 minutes after

ignition in test 12 81

Photo 3. Active burning stage 8 minutes after ignition in

test 12 82

vi
FIRE DEVELOPMENT IN RESIDENTIAL BASEMENT ROOMS

J. B. Fang and J. N. Breese

Abstract

A multi-phase study program has been established

to develop a rational test procedure for evaluating

the fire resistance of residential floor assemblies.

The first phase of this research program was aimed at

characterizing the severity of fires originating in

residential rooms and developing a specified set of

fire exposure conditions applicable for fire resistance

testing of floor constructions.

A total of 16 burnout tests were conducted to

investigate the fire behavior in typical residential

recreation rooms of single family houses. These fire

tests were usually run for one hour and were performed

in two instrumented test rooms, 3.3 x 3.3 x 2.4 m and

3.3 x 4.9 x 2.4 m in width, length, and height respectively, furnished with household furniture and lined

with interior finish materials typical of actual

occupancies. Measurements were made of the temperature,

heat flux, static pressure, smoke density, gas velocity,

species concentration, and oxygen consumption. The

effects of such parameters as the ventilation, fire

load density, initial item ignited, room size, and

thermal and flammable properties of the wall and ceiling

materials on the fire severity were evaluated quantitatively. A fire exposure temperature-time curve which

is different from the ASTM E 119 curve, has been developed

for testing the fire resistance of such building structures.

Key Words: Building fires; fire resistance; fire

tests; flow measurement; gas temperatures; heat release

rate; interior finishes; residential buildings; room

fires

.

1
1 . INTRODUCTION

The fire performance of construction materials and assemblies is

conventionally determined by subjecting the building components and structural

elements to a standardized exposure in a laboratory fire endurance test. Due

to extensive research on compartment fires and increased information available

on fire behavior, some doubts have been raised about the validity of the test

environments relative to actual fires.

The ASTM Standard E 119 time-temperature curve has been used widely in

the U.S. for over 60 years. The curve was developed in 1917 based on experience derived from all the known temperatures measured or inferred from building

fires and from fire tests made at various institutions such as the New York

Building Code Authority and Columbia University, and no major change has taken

place in the curve since then. The design of residential buildings has

changed considerably from that six decades ago, with the use of lightweight

construction instead of heavy masonry, and large windows versus small ones. Also, many new products including synthetic fabrics, finishes, laminates and

composites, have been introduced as the primary components for the furnishings

and lining materials of residential rooms. These materials tend to burn more

rapidly and may lose their integrity earlier and at lower temperatures than

woodbase materials. The real temperature development can be expected to be

quite different from the temperature history described by the standard curve.

To establish the fire resistance required for a structure or a partition

to withstand the action of the fully-developed fire without losing its structural integrity, insulating function or its load carrying capacity, it is

necessary to estimate the severity and duration of the expected fire. Several

years ago, an international cooperative research program [1]^ was carried out

under the auspices of the Conseil International du Batiment (CIB) to investigate the effects of various factors on the behavior of fully-developed fires

in single compartments. Numerous experiments were performed by burning standardized wood cribs in small-scale compartments under a wide range of conditions resulting from different combinations of such variables as compartment

size and shape, ventilation conditions, the amount and distribution of fuel

and type of wall materials. The results of this intensive experimental study

indicated that there was a substantial difference between fires in compartments

of different shapes, the influence of the compartment size was small, and the

effect of the thermal properties of the wall and ceiling materials was minor.

However, in real room fires the combustible contents such as the furnishings

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

this report.

2
in the form of fabrics, sheets, and foam padding may contain a certain amount

of highly combustible plastic materials, and it is not clear if the data

derived from using wood cribs are applicable.

In 1939, a total of five full-scale burnout tests involving typical

furnishings in a simulated three room residential occupancy building were

conducted at NBS to investigate the effect of different combustible loads on the intensity and duration of room fires. The fire load density studied

ranged from 25.4 to 58.6 kg/m2 (5.2 to 12 lb/ft 2

) of floor area. It was found that average spatial temperatures of individual rooms were generally

lower than the ASTM standard curve, except for fire load density greater than

41.5 kg/m2 where the average temperature exceeded the standard curve for

periods up to 22 minutes.

The severity of developing room fires depends upon the fire load, the

surface area and the thermal properties of the room boundaries, and the

ventilation conditions. A knowledge of the effect of these various parameters

on the development of fires in residential rooms is needed as a basis for

developing a rational procedure for evaluating the fire resistance of

construction materials and assemblies.

The overall objectives of the current fire endurance study sponsored by

the Office of Policy Development and Research within the U.S. Department of

Housing and Urban Development (HUD) at NBS are to develop a meaningful test

procedure for more realistically evaluating the fire resistance of floor/

ceiling assemblies in residential occupancies, and to suggest improved performance criteria for the fire-safe use of load-bearing structural components.

The first phase of this research program was specifically aimed at characterizing the behavior of fires originating in typically furnished basement

recreation rooms, providing technical data needed for the development of a mathematical model of fire growth and developing a rational set of fire

exposure conditions applicable to the fire resistance testing of residential

floor constructions.

This report presents the experimental results of a series of sixteen

full-scale room burnout tests. The test series was conducted to determine

the fire behavior over a range of typical fire load densities, room sizes,

ventilation conditions, and interior finish materials. The data include a description of the fire growth process in the fire room; the time to room

flashover, the time when flames emerged from the doorway, the time histories

of the upper gas temperature, the rate of heat released within the room; the

developed static pressure and heat flux levels at selected locations; and

the fire induced air flow at the door opening.

3
2. EXPERIMENTAL DETAILS

2.1 Test Compartments

Two burn rooms, one 3.3 x 3.3 m (10.7 x 10.7 ft) and the other 3.3 x 4.9 m (10.7 x 16 ft) both with 2.4 m (8 ft) ceiling height were constructed

consecutively to investigate the effects of room geometry on the fire

behavior. Both rooms had a 0.76 m (30 in) wide by 2.03 m (80 in) high

finished doorway opening located centrally in one of the 3.3 x 2.4 m (10.7 x

8 ft) walls to serve as the single source of natural ventilation for the rooms. Figures 1 to 3 show the construction details of the test rooms. The rooms also had a heat-resistant glass window, 0.41 m (16 in) wide by 0.2 m (8 in)

high situated symmetrically in one of the side walls and at a height of 1.4 m

(56 in) above the floor for observation purposes. The basic structure of the

burn rooms was 0.2 m (8 in) thick standard lightweight aggregate-concrete

block walls. The floor was concrete protected by a layer of 16 mm. Type X

gypsum wallboard. The ceiling was 13 mm (1/2 in) mineral board covered by a layer of 16 mm (5/8 in) Type X gypsum wallboard attached to W6 x 12 steel

I-beams resting on the top of the block walls through a layer of ceramic

fiber. The burn rooms were built within a large building to minimize the

exterior weather effects and the indoor air was controlled to provide an air

temperature of 21 + 4°C and a relative humidity of 42 + 8 percent. The

combustion products from the burn room were exhausted through a collection

hood to an afterburner for smoke abatement.

2.2 Fire Load Density

In order to select a typical range of loading density, material type and

geometric arrangement of the combustible contents for the present series of

full-scale room burnout tests, field data on fire loads in basement rooms for

residential occupancies were collected and analyzed. Figure 4 shows the

frequency distribution of the fire load density for the occupied recreation

rooms based on 70 surveyed homes situated in the Washington, D.C. metropolitan

area. These 70 homes were those for which information was available at the

time and represented a portion of the sample from an ongoing project to survey

fire and live loads in single family and mobile homes [2]. In figure 4, the

fire load density is expressed in terms of the total weight of combustible

materials per unit floor area. The numerical values shown at the top and

those at the bottom of each column represent the mean values of the total

and the movable fire loads, respectively, lying within a given interval.

The movable fire load included all of the combustible materials in the

furnishings and other items brought in for the service of the occupants after

4
the house was built. The total fire load consisted of the combustible

structural elements including the interior finish materials as well as the

movable fire load. The total weights of the combustible items were obtained

by comparing their dimensions measured during the survey with the manufacturers' data for items of similar size, materials and construction [2] . The

three levels of the total fire load density selected were 21, 28 and 42 kg/m 2

(4.3, 5.8 and 8.5 lb/ft 2) representing low, average and high ranges of the

combustible loads, which corresponded to the 28%, 50% and 83% fractiles based

on the normal distribution for approximating the total fire load observations.

From figure 4, the low, average and high levels of the movable fire loads

corresponding to these total fire loads chosen for basement fire studies were

respectively 15, 23 and 37 kg/m2 (3.1, 4.7 and 7.6 lb/ft 2

) which were equivalent to the 26%, 50% and 88% fractiles on the basis of considering the spread

of combustible contents to be normally distributed. The frequency distributions of the total fire load data for utility rooms and other rooms such as basements or unfinished storage rooms are given in figures 5 and 6. In

general, the fire load density for utility rooms was slightly lower than that

for recreation rooms and its frequency seemed to follow a skew distribution,

whereas the frequency for other rooms decreased continuously with an increase

in the total fire load. The normal value of the movable fire load density

for an occupied basement utility room and that for other rooms in the basement were found to be approximately 13.7 kg/m2 (2.8 lb/ft 2

) and 15.6 kg/m2 (3.2 lb/ft 2

) of floor space, respectively. A summary of the type of interior

finish materials employed for the basement rooms of those 70 surveyed homes

located in the Washington, D.C. metropolitan area is given in table 1. As

illustrated in the table, most of the recreation rooms had gypsum wallboard

or plaster materials as the ceiling and plywood panels as the wall linings.

The majority of the utility and other rooms had exposed wood joists overhead

and concrete block or gypsum board finished walls.

The frequency distribution of the movable fire load in basement

recreation rooms based on 200 homes is shown in figure 7. These 200 homes

are the total number of single family detached homes included in the fire

load survey [2]

.

The average value of the movable fire load density was

24 kg/m2 (4.9 lb/ft 2 ). The 23 kg/m2 (4.7 lb/ft 2

) movable fire load density

used in these tests was based on a smaller sample of 70 homes which was

available at the initiation of this project. The floor area over which the

combustible loads were distributed in the 200 single family detached homes

surveyed ranged from 4.5 m 2 (49 ft 2

) to over 46 m 2 (500 ft 2

) with a mean

value of 32 m

2 (344 ft 2 ), which was approximately twice that of the larger

burn room used for the fire tests in this series.

5
The composition of the movable fire load employed in a typical test run was estimated based on the information supplied by the furniture manufacturers.

As given in table 2, these were found to be 11.1 kg/m2 of wood, 7.4 kg/m2 of

paper, 0.92 kg/m2 of fabric, 1.02 kg/m2 of plastic materials and 2.5 kg/m2 of

others for a 3.3 x 3.3 m recreation room furnished with 23 kg/m2 of combustible

contents. These material types and weights are comparable with those found

in the typical basement family rooms of single family detached homes in the

Washington, D.C. metropolitan area [2], where the average movable contents

consisted of 8.3 kg/m2 of wood, 9.4 kg/m2 of paper, 2.3 kg/m2 of plastics,

1.2 kg/m2 of fabric and 0.2 kg/m2 of others.

2.3 Furnishings

In order to minimize the possible effects due to the variations in

material type, weights and constructions of the movable combustibles on fire

behavior, and to duplicate the room contents for a number of fire tests,

identical pieces of household furniture were purchased in large lots directly

from either furniture manufacturers or large retail stores. For most of the

fire tests, the furnishings consisted of a sofa, an upholstered chair and

ottoman, an end table, a bookcase and a coffee table. For the tests involving

the large burn room, a loveseat, an end table and another bookcase were used

in addition to these furniture items. A description of the dimensions,

materials and weights of the combustible portions is given in table 2. These

varied slightly in size for the upholstered chair and the coffee and end

tables since the original styles used in the early stage of the test series

were subsequently discontinued. Figure 8(a) shows the floor plan of a furnished room illustrating the layout of the furniture and wall finish

materials for a typical test run. The large recreation room was furnished

similarly and is shown in figure 8 (b) . Included among the furnishings were

old record file papers (22 x 28 cm) , which were used to reach the required

fire load density. A total of 4.54 kg of the papers were placed on the

coffee table and 1.81 kg each were placed on the tops of the ottoman and

the end table, with the remainder in the bookcase (s).

The floor was covered with wall-to-wall carpeting except for two areas directly beneath the platforms of the weighing cells, where two identical - sized pieces of the carpet were placed atop the platforms. For all of the

fire tests, an olefin carpet with rubber foam cushion backing was used to

cover the concrete floor through a layer of 16 mm thick fire rated gypsum

wallboard

.

6
2.4 Interior Finish Materials

A summary of the test conditions and the interior finish materials used

in this study is presented in table 3. For each test, the interior walls of

the test room were finished with one of the following three types of wall

covering materials:

Gypsum wallboard . The wall framing consisted of nominal 25 x 76 mm

(1x3 in) wood furring strips spaced 0.40 m (16 in) on center and attached

to the concrete block walls with masonry nails. A single plate was used both

at the top and at the bottom of the furring strips. The 13 mm (1/2 in) thick

gypsum wallboard was applied with its long dimension parallel to the frame

with dry wall screws. All joints were taped and covered with joint compound.

The floor and ceiling seams were finished with wood trim. The walls were

painted with two coats of white interior flat latex paint.

Plywood paneling . The wall framing was constructed in a manner similar

to that of the gypsum wallboard. The 4 mm (5/32 in) thick prefinished and

printed 3-ply lauan plywood panels were nailed parallel to the framing with

interior hardened 25.4 mm (1 in) long paneling nails, 0.20 m (8 in) on center.

Prefinished wood shoe moulding (15.9 x 9.5 mm) was used to trim the edges of

the paneling walls.

Painted concrete block . The concrete block walls of the burn room were

painted with two coats of common masonry paint (calcium oxide 64%, silicon

oxide 22%, aluminum oxide, etc., 14%) to represent a typical finish for

basement rooms.

The following four types of ceiling materials were used:

Suspended ceiling . The installation included lay-in wood fiber base

ceiling tiles in a suspension grid. This system consisted of painted light

metal, wall angles 19 x 19 x 0.56 mm (3/4 x 3/4 in x 25 ga) , 38 x 25 x 1.12 mm

(1-1/2 x 1 in x double layer of 25 ga) main tee runners and 25 x 25 x 0.56 mm

(1 x 1 in x 25 ga) cross tees. The wall angles were nailed firmly to the

wood furring strips fastened to the block walls. The main tee runners were

installed at right angles to the steel studs, and the cross tees were hung

from the suspension wires fastened to the studs. The 0.61 m x 1.22 m x 12.7 mm

(2 ft x 4 ft x 1/2 in) thick, wood fiber base ceiling tiles (acoustical,

lay-in panel with washable finish, flame spread index class 200) were dropped

approximately 90 mm (3.5 in) below the steel studs to provide a ceiling

height of 2.26 m (89 in) measured above the carpet flooring.

7
Gypsum board . The ceiling framing members used were light gage,

galvanized C shape 92 x 33 x 0.56 mm (3-5/8 x 1-5/16 in x 25 ga) thick steel

stud, and spaced 0.4 m (16 in) apart. The 16 mm (5/8 in) thick, type X,

gypsum board sheet was applied with its long edges at right angles to the

steel stud, and fastened to its bottom flange with screws. All joints

between the adjoining pieces of gypsum board and screw heads were filled and

taped with joint compound. The gypsum wallboard surfaces were painted with

two coats of white color flat latex paint. The height of the finished gypsum

board ceiling above the olefin carpet floor was 2.33 m (92 in).

Exposed wood joists . The construction contained nominal 50.8 x 203.2 mm (2x8 in) wood joists spaced 0.41 m (16 in) on center, and a single layer of

15.9 mm (5/8 in) thick underlayment grade plywood subfloor. The joists were

resting on the upper sides of two nominal 51 x 203 mm (2x8 in) wood beams

supported by two nominal 51 x 203 mm wood posts. All of the wood beams and

posts were boxed within a double layer of fire rated 16 mm (5/8 in) thick

gypsum wallboard for eliminating fire penetration and fastened by nailing

into the block walls with masonry nails. Except for the joist spacing near

the center of the ceiling where heat flux meters were located, solid bridging

was installed at the midspan between the joists. Subflooring was laid perpendicular to the joists and fastened to each joist with nails. The ceiling

height measured from the top of the carpet floor to the bottom of the wood

joists was 2.14 m (84.2 in). In order to protect the mineral board sub-ceiling

from the intense fire, a layer of 19 mm to 25 mm thick plaster spray mixture

was applied to a metal lath hung approximately 13 mm (0.5 in) below the

sub-ceiling

Exposed steel joists . The framing used for the ceiling structure was

45 x 184 mm deep x 1.27 mm (1-3/4 x 7-1/4 in x 18 gage) thick, channel shaped

steel joists spaced at 0.61 m (24 in) on center. The joists were fastened to

nominal 51 x 203 mm (2 x 8 in ) wood rim joists. Each rim joist was attached

to the top side of an identical size of wood beam supported by two pieces of

nominal 51 x 203 mm wood posts. Wood rim joists, beams and posts were fastened

to the concrete block walls and protected from the fire by two layers of 16 mm

(5/8 in) thick, type X gypsum wallboard. No bridging was installed. A single

layer of 16 mm (5/8 in) thick underlayment grade plywood subfloor was installed

at right angles to the joists and secured to the steel framing with screws. The height of the bottom flanges of the joists above the carpet flooring was found to be 2.21 m.

8
2.5 Forced Ventilation

In three tests, air was supplied to the room to simulate operation of

a forced air central heating or cooling system.

For the fire tests involving forced draft, a constant rate of 170 m

3 /h

(100 CFM) air flow supplied by a blower was introduced into the test room via

a light gauge metal duct of 152 mm (6 in) in outside diameter and a 105 x 305 mm

(6 x 12 in) multi-louver register located at 1.83 m above the concrete floor

and 0.41 m from one of the room corners opposite of the doorway.

A 0.76 m wide by 1.99 m high metal door with one 0.20 m square, heat

resistant glass window for observation purposes, was employed for the two tests

with the door closed. The door in closed position provided a small opening,

0.76 m wide x 0.04 m high, between its bottom edge and the concrete floor for

venting the hot gases out of the fire room. Since the fire performance of the

door was not a research objective, the door was protected on the fire side

with 13 mm (1/2 in) thick noncombustible mineral board.

2.6 Test Parameters

The experimental conditions of the sixteen full-scale room burnout tests

conducted for determining the effects of various parameters on the behavior

of compartment fires are given in table 3. Several tests in the series were

designed to provide data related to specific materials and situations and

they would also be useful for development of mathematical models of fire

growth. Tests 1, 2 and 6 were intended to examine the influences of the

thermal and flammability properties of the wall materials on the fire buildup

process. The data from tests 1, 4, and 5 were expected to show the relationship between the heat release rate in the room and the fire load density.

An additional experiment (test 9) was performed to repeat test 1, which had

to be terminated early due to an unexpected fire involvement of the sur- rounding structures. Also, test 10 was a repeat of test 3 since much of the

temperature data was lost in the earlier test. Data from tests 3, 7, 9 and

11 were intended to establish the effects of open and closed doors and forced

ventilation. Test 9 was performed with a basic room configuration and average

level of combustible load to provide base line data. Data from tests 6 and

13 along with tests 8 and 14 were planned to compare the severity characteristics of room fires as a function of room size. In test 11A, the combustible

loads employed were those remaining from test 11, where no room flashover

and flame extinction were observed. In test 11A, the upholstered chair

instead of the sofa was used as the first ignition item to supply information

9
for establishing the effect of the ignition source on fire development. The

change in fire severity with gypsum board replacing plywood panel as wall

linings was determined in tests 12 and 13 for the larger burn room, and in

tests 6 and 9 for the smaller room. Tests 12 and 14 showed the influence

of the flammability properties of the ceiling finishes on the overall room

fire development. The data from tests 13, 15 and 16 provided the basis for

comparing the fire growth in rooms with unprotected floor joists with those

having finished ceilings.

The initial weights of the combustible materials used for all of the

fire tests are summarized in table 4. The wall finishes include the combustibles employed as the covering material, and the framing and moulding strips

for trimming the edges of the room walls. The estimate of total available

heat from the burning of the combustible materials for each run, also tabulated in table 4, was made using these weight data for interior finishes and

furnishings along with the information on weight percentages of the constituent

materials employed for constructing various furniture items as given in table

2. The calorific values used for these calculations were 18.6 MJ/kg (8000

Btu/lb) for wood, 46.5 MJ/kg (19900 Btu/lb) for olefin fabric, 16.7 MJ/kg

(7200 Btu/lb) for paper, and 26.5 MJ/kg (11400 Btu/lb) for polyurethane foam,

respectively. The estimated total available heat from the combustible loads

for all of the tests are also tabulated in table 4 and found to range from

approximately 5400 to 14000 MJ (5.12 x 10 6 to 13.3 x 10 6 Btu)

.

3. TEST MEASUREMENTS

3.1 Temperature

The temperatures within the room and on the exposed surfaces of the walls

and ceiling at selected locations were measured by bare beaded chromel-alumel

thermocouples constructed from 0,51 mm (No. 24 B. and S. gage) diameter wire

with a spherical junction of approximately 1.5 mm in diameter. After test 7,

the thermocouples were formed from the wires of two different cables in order

to prevent leakage of electric current between wires when it was found that

there was a reduction in the readings due to the charring of the cable insulation for gas temperatures above 800°C. Gas temperature measurements were

made at 10 stations including nine locations within the test room and one in

the doorway opening. A total of 42 thermocouples were distributed vertically

along nine thermocouple "trees" inside the room, located as shown in figures

8(a) and (b) , and 12 thermocouples were positioned at various heights along

the vertical centerline of the doorway as shown in figure 9. Indications of

the upper room gas temperatures were also determined with three commercial

10
metallic-sheathed mineral insulated fast response thermocouples and with three

ASTM E 119 standard protected furnace thermocouples. These sheathed thermocouples provided direct comparison with the ASTM E 119 curve and would be used

to control the gas temperature in the fire endurance furnace because they

required less frequent replacement than the bare beaded thermocouples due to

their greater mechanical strength. These thermocouples were placed centrally

0.58 m below the ceiling at the mid-point and the sixth point of the lengthwise span. The fast response thermocouples were made from chromel-alumel

wires with diameter of 1.02 mm (0.04 in ) and their hot junctions enclosed

in the sealed stainless steel sheaths, 6.4 mm (0.25 in ) in outside diameter

and 0.9 mm (0.036 in ) in wall thickness, and insulated with magnesium oxide.

The ASTM furnace thermocouples were of 1.02 mm (0.04 in ) chromel and alumel

wires with their hot junctions surrounded completely within sealed black

wrought iron pipes, nominal 13 mm (0.5 in ) in outside diameter and electrically

insulated with porcelain tubes.

The wall and ceiling surface temperatures were measured at 20 stations by

thermocouple beads mounted on the exposed and unexposed surfaces at the

selected locations. Five stations were located at the ceiling as shown in

figures 8 (a) and (b) ; one was at the front wall inside the test room for

monitoring full involvement of the combustible linings, shown in figure 9;

eight stations were distributed over the back wall near the ignition point on the sofa as illustrated in figure 10, and three each were placed at the left

and the right walls of the test rooms as shown in figures 11(a) to 12(b)

.

3.2 Heat Flux

A total of nine water-cooled, Gardon type gauges were used for incident

total heat flux measurements. Three of these heat flux gauges were located

along the centerline of the ceiling, flush with its exposed surface: one flush with the bottom face of the mineral board sub-ceiling, two in the exposed

surface of the back wall, one each in the left wall and at the center of the

carpet floor viewing the ceiling as shown in figures 8, 10 and 11; and one on

a stand located outside the front of the room with its vertically oriented

sensing element positioned 3.05 m from the doorway opening and 1.02 m above

the concrete floor. One Gardon type, wide angle, water-cooled, radiometer

with an air-purged sapphire window attachment was mounted at the center of the

ceiling to measure the radiation from the hot gas layer and the heated

surrounding walls.

11
3.3

Air Velocity at Doorway

The vertical distribution of the gas velocity in the doorway was monitored with six bidirectional flow probes [3] connected to variable

reluctance differential pressure transducers and carrier demodulators. These

probes were located along the vertical centerline of the opening and positioned

at 0.13, 0.48, 0.84, 1.19, 1.55 and 1.91 m from the top of the doorway as shown in figure 9.

3.4

Static Pressure

Static pressure measurements of the gases in the fire room were made

with 64 mm (0.25 in ) steel pipes installed horizontally through the left

wall (south wall) with their open ends flush with the inner surfaces of the

wall lining panels. Five of these pressure probes were located along a vertical line 0.3 m from the front block wall at distances of 0.1, 0.3, 0.76,

1.32 and 2.13 m below the ceiling. Another probe was placed 0.3 m from the

rear block wall at 0.1 m below the ceiling. The probes were connected by

copper and tygon tubing to variable reluctance pressure transducers which

in turn were connected to carrier demodulators for continuous monitoring of

the static pressures in the room.

3.5

Optical Density of the Smoke

Optical density measurements of the smoke were made horizontally at five

heights, 0.08, 0.30, 0.53, 0.97 and 1.50 m from the top of the doorway, over

a 0.76 m long path along the doorway width as shown in figure 9, and over a 207 m path across the width of the exhaust hood, utilizing automobile spot

lights as the light sources and phototube detectors (PIN3DP) as the receivers.

Another measurement, with a path length of 2.07 m was made with a laser beam

collimated vertically along the doorway height. The receiver was a phototube

(1P39) located in a water-cooled, air purged insulated enclosure situated

0.30 m above the top of the doorway. The laser beam passed through a small

hole in the floor at the front of the doorway.

3.6

Gas Concentrations

The gas concentrations were monitored continuously at four locations.

Three sampling ports were situated in the doorway opening at heights of 0.08,

0.38 and 0.76 m down from the doorway top, and one was located in the smoke

stack, which was connected to the exhaust hood installed above the opening.

For the tests with the door closed, the open ends of three gas sampling probes

12
in the doorway opening were extended horizontally 0.31 m into the test room. After passing through a series of cold traps and glass wool filters, the gas

samples from the top and the middle ports were analyzed for 0

2

, CO

2 and CO,

and those from the bottom port were analyzed for C> 2 and CO. Samples withdrawn from combustion products flowing in the smoke stack were analyzed for

O

2

, C0

2 and CO. The oxygen concentrations were measured with electrochemical

type oxygen cells. The C0

2 and the CO were measured with non-dispersive

infrared analyzers.

3.7

Weight Loss

Continuous measurements were made of the weights of the sofa and the

upholstered chair during the tests. Each was located on a steel-frame platform suspended from a 227 kg capacity strain gage load cell. In order to

resist the intense heat from long duration fires, the platform was a monocoque

structure fabricated by bolting together two pieces of 3.6 mm (10 gage) thick

steel plate as the facings and two layers of 19 mm (3/4 in ) noncombustible

mineral board as the core material. The entire platform was suspended by a steel rod attached to the load cell transducer mounted on the I-beams resting

on the top of the concrete block walls.

3.8

Flashover Indicators

Two indicator specimens of crumpled ordinary newsprint and 0.78 mm thick

horizontally-oriented filter paper, each placed on the ring of a tripod

located at the center of the floor, were used to provide a direct indication

of the involvement of typical combustibles in room fires.

3.9

Data Acquisition

During the test, the data were collected with a digital data acquisition

system. The output signals from the thermocouples and the various transducers

were logged every 8 seconds on magnetic tape. The data were processed,

tabulated and plotted using a computer.

4 . TEST PROCEDURE

Since the test series aimed to quantify the severity characteristics of

basement room fires, which generally involved various stages of physical

processes from flaming fire spread to full room involvement and complete

burnout of combustible contents, a simulated newspaper ignition of the sofa

was employed. A section of the daily newspaper weighing 400 g was placed

13
along the backrest and the seat cushions in the middle of the sofa and

supported by a steel frame holder in order to provide a reproducible ignition

source. The paper was conditioned to equilibrium at an ambient temperature

of 23 + 3°C and a relative humidity of 50 +5 percent prior to the fire test.

The test was started by remotely igniting the newspaper with an electrically

heated book of paper matches. During the course of the test, visual observations were made of the progress of the room fire and recorded with an audio

tape cassette. Video tape and 35 mm slide coverage was provided on each test.

On some tests 16 mm motion picture film coverage was also provided. The

fire was allowed to burn completely for at least one hour, except for two

tests (tests 1 and 2) where the tests were shortened to avoid serious damage

to the surroundings and to the ceiling structures. A manually controlled,

fine spray sprinkler located near the center of the ceiling and manually

operated water sprays from hoses were used to extinguish the glowing ember

fires remaining after one hour.

5. RESULTS

The photographic records of the fire development in test 12 involving

a 3.3 x 4.9 m recreation room lined with lauan plywood panels on the walls

and a gypsum board ceiling, and furnished with combustible contents of 23

kg/m2 of the floor space, are shown in photos 1 to 3. The majority of the

test fires had a similar history of development. A log of both visual observations and records deduced from the video tape made through the doorway of

the fire room in a typical test (test 14) is given in table 5.

Selected results from this series of full-scale basement room burnout

tests are summarized in tables 6 and 7. Table 6 includes: time to appearance

of visible flames on the newspaper utilized as the ignition source; room

flashover times derived on the basis of spontaneous ignition of two indicator

specimens, on attainment of a level of 20 kW/m 2 for the incident heat flux

at the center of the floor; times for flames to emerge from the doorway; times

for spontaneous ignition of the olefin carpet flooring; the maximum rates of

heat released within the fire room calculated by two methods and the maximum

upper room gas temperature. The data show that the time required to obtain

sustained flaming of the newspaper ignition source varied from approximately

0.1 to 1.4 min with an average of 0.5 min after start of the test.

The times taken from the appearance of sustained flaming of the

newspaper to the spontaneous ignition of the flashover indicators and of the

carpet flooring, and to flame emergence from the doorway were obtained from

visual observations. As illustrated in table 6, no room flashover was observed

14
for fire tests with the door closed (tests 7 and 11) . Excluding the

supplementary test, where the fire starting from an upholstered chair took

slightly longer to reach flashover, the elapsed times ranged from 1.8 to 2.55

min with an average of 2.03 min for the ignition of the newsprint; and 1.8

to 2.72 min , with a mean of 2.13 min for the ignition of the filter paper.

The time at which the incident heat flux measured at the center of the floor

reached a level of 20 kW/m2 varied from 1.67 to 2.97 min with an average

value of 2.03 min. This was in good agreement with the times required for

onset of ignition of the indicators. The times taken for the flames to

issue out of the doorway and for the ignition of the carpet on the floor

ranged from 1.68 to 4.32 min (average 2.14 min ) and 1.87 to 3.50 min , (average 2.35 min ), respectively.

The mean temperature of the hot gas within the upper part of the fire

room for tests 1 to 7 was obtained by averaging the readings of thermocouples

placed at 27 stations; nine evenly-spaced locations each at 0.025, 0.584 and

1.17 m below the ceiling. It was found later that charring of the silicone

impregnation of the thermocouple wire insulation due to the fire, caused the

indicated temperatures above 800°C to be too low. This type thermocouple

with its insulation left on was also checked against an NBS standard type S

thermocouple in an electric furnace controlled at two selected temperature

levels. The accuracy of the thermocouple used was found to be 2 degrees high

at 500°C and ranged from 48 to 122 degrees low at 1000°C level maintained

for a period of one hour. This error was reduced by using a double-cable

thermocouple in which two cables ran in parallel and the bare measuring junction was formed by welding the positive wire (chromel) from one cable together

with the negative wire (alumel) from the other one. For tests 8 to 16, the

results from the average of the temperature measurements by the fast response

thermocouples at three locations were used to represent the room gas temperature since small thermocouple wires were easily broken and involved frequent

replacement. These metallic sheathed thermocouples will also be used for

monitoring and controlling the gas temperature of a gas-fired test furnace in

the latter part of this research project. As shown in table 6, the peak temperature of the upper room gas varied from approximately 190°C to 410°C for

the tests with closed door and 950°C to about 1080°C for the tests with the

open door.

Data on the rate of heat release in the fire room can provide useful

information for verification of the mathematical models of room fire growth.

The heat release rate can be computed using the oxygen consumption technique

since a great number of organic fuels under the presence of sufficient oxygen

burn completely and produce nearly the same amount of heat per unit mole, mass

or volume of oxygen consumed, which is almost independent of fuel type [4].

15
Recently, Parker [5] employed this principle to calculate the total rate of

heat production for a wide range of building materials tested in the ASTM

E 84 tunnel. Christian and Waterman [6] conducted full-scale fire tests to

characterize the fire hazard of various occupancies and used the data on oxygen depletion and air flow rate measurements to estimate the heat release

rate for various fires. Huggett [7] has recently reviewed the oxygen consumption technique for the measurement of heat release and found that incomplete

combustion and variation in the type of fuel have only a minor effect on the

result.

With the assumption that the combustion gas, which consists of the

combustion products and the depleted air, behaves as an ideal gas, the rate

of heat release from the complete combustion of the combustibles within the

fire room to carbon dioxide and water vapor can be derived based on the

conservation of mass and oxygen species as below

( a A v + 1 ) X

°2 a ^298.15/T

g j

U

g

A ( 1 )

where Q is the rate of heat release, in kW; K is heat of combustion of

combustible materials, expressed as the heat produced per unit volume of

oxygen consumed at the standard state of 25°C and 1 atm, in kJ/m 3

; X and

U ^ cl

*

. \ . . ^ Xq

^ are the concentrations of oxygen in the incoming air and m the combus-

^ * tion gas stream, respectively, in mole fractions; A v = A v/v^, the change in

the mole numbers per mole of combustible material or the fuel reacted;

4 » + v - ^0 + u co

2

'CO

v o. - v £ , in which v. is the stoichiometric coeffif D cient of gas species j; v = v /v f

, the ratio of stoichiometric coefficient

2 2 of oxygen to that of the fuel; a is the fraction of the fuel reacted, a = 1.0

for the fires under fuel controlled burning regime and 0 <_ a < 1.0 for ventilation limited fires; is the absolute temperature of the combustion gas,

in °K; is the component of the velocity of the combustion gas normal to the

plane of the doorway in m/s, and A is the cross-sectional area through which

it is exhausted, in m

2

. The detailed derivations of the equations utilized

for the heat release rate calculations are given in appendix A.

The total heat release rate in the fire room was computed based on the

data from the oxygen depletion, gas velocity and temperature measurements of

the combustion gas in the doorway, using equation 1 along with a constant K

value of 17010 KJ/m 3 [7] and assuming that the molecular weight of the

unburned fuel is equal to that of nitrogen gas during ventilation limited

burning. Since the major components of room combustibles were cellulosic

materials, wood with a chemical formula of CH. .,O n rc • 0.1H„0, in which

1 . 460.65 2 16
the attached 0.1 molecules of water account for 7 percent of moisture content

[8], was used to represent typical fire loads. The assumption that the

velocity, temperature and oxygen concentration can be approximated by a finite number of linear segments was employed for the energy production

calculation

.

In order to check the reliability of heat release rates derived based

on the oxygen consumption method, one gas burner fire experiment was made

with the 3.3 x 4.9 m burn room having concrete block on the walls and gypsum

board ceiling. Two gas burners, each 0.31 m square in size, were placed

side by side, 0.71 m away from the back wall and 1.3 m from the side walls

with a height of 0.31 m above the floor. These gas burners were used to serve as the fire source where propane gas was introduced into them at three

known flow rates ranging from 0.472 to 0.944 m

3 /min. The heat release rates

computed from the data on the oxygen concentration and gas flow rate measurements at the doorway opening were compared with the rates of heat output by

the gas burners calculated based on gas supply rate and found to agree

within + 12 percent.

The computed maximum heat release rate of an individual room fire during

its peak burning period is presented in table 6. As shown, the maximum rate

of heat produced inside the fire room based on the various measurements at

the doorway was found to range from 3530 to 6240 kW with an average of

5250 kW for the fires involving natural ventilation through an open door.

Significant quantities of gaseous fuel or excess pyrolyzates were observed to

burn outside of the fire room because of insufficient oxygen available for

combustion within the room. With the supply of air at a rate equal to that

of normal operating condition of the HVAC system in residential rooms, the

calculated heat output rate of the compartment fire increased significantly

as illustrated in table 6.

The rates of heat transfer to the selected locations on the room ceiling,

walls and floor, and in front of the doorway outside of the room are presented

in table 7. The peak level of heat flux incident at the ceiling varied from

approximately 90 to 180 kW/m 2 with a mean of about 130 kW/m 2 for the rooms

with an open door, and between 30 and 40 kW/m2 for the rooms with the door

closed. The data also show that with the fiberboard ceiling tile, higher

levels of surface heat flux were obtained due to the contribution from the

ceiling flames in addition to radiant and convective heating from hot gas

layer and the upper walls. The exposed surfaces of room walls were subject

to more intense heat flux than those at the ceiling, and the peak wall surface

flux produced from the fires in the well-ventilated room varied from about

120 to 230 kW/m2

. 17
The maximum heat flux incident at the center of the floor during the

fire test in the room with the open doorway, excluding test 2, which was stopped early, ranged from approximately 110 to 160 kW/m 2 with an average of

130 kW/m2. Also shown in the table, the peak rates of heat transfer, mainly

by means of radiative transport, to a point outside the room, 3.05 m from

the opening, were of the order of 33 kW/m 2

. This heat flux level was capable

of causing ignition of nearby combustible materials. An estimate of the

magnitude of radiative heat flux impinging on the vertically-oriented sensing

element of the heat flux gauge located outside of the room, due to the

emerging flames and the room fire was made. The assumptions employed in the

calculations are that (1) the flames emerge above the neutral plane located

at the lower one-third of the doorway height, (2) the flame axis has a con- stant angle of inclination of 45 degrees from the vertical plane; and (3) both

the flames and the room fire via the doorway opening can be considered as plane emitters at a constant temperature equal to that of the room gas. The

estimated radiant flux from the doorway alone was found to vary from 1.5 to

2.6 kW/m 2 with a mean of 2.1 kW/m 2

, for the six tests included in table 7.

The intensity of the thermal radiation emitted by the turbulent buoyant flames

issuing out from the room above the doorway was estimated to range from 26 to

46 kW/m2 with an average of 35 kW/m2

, using an absorption coefficient of

1.1 hi - ! for the flame gas. It is obvious that the radiation emitted by the

emerging flames constituted the major portion of the total energy received by

the heat flux gauge. Some difficulties were encountered with the radiometer

used for the direct measurement of the radiant flux inside the burning com- partment due to the deposition of soot and condensable matter on the window;

this caused a significant reduction in the total transmissivity throughout

the test.

Table 8 summarizes the experimentally-determined flow rates of the

inflowing air and outflowing gas through the doorway opening; the static

pressure developed within the room; the neutral plane height above the floor

level at the peak of the fire; the maximum smoke densities of the exhausting

gases along the doorway width and across the length of the exhaust hood; and

the peak concentrations of CO and C0

2 measured at the doorway and in the

smoke stack. The data show that the maximum rate of air inflow induced by

the room fire ranged from approximately 1.0 to 1.8 kg/s and the hot gas outflow varied from about 2.2 to 3.4 kg/s. These variations in flow rates may

be attributed to differences in the characteristics of the room fires such as the temperatures of room gases.

The peak levels of the static pressure measured at 0.1 m below the

ceiling for all test fires were found to be relatively constant with an

18
average value of approximately 14 Pa for the rooms with an open door. The

pressure increased to 31 Pa for the tightly-closed compartment. The lowest

height of the neutral plane in the doorway ranged from 0.57 to 0.71 m for

all tests except for test 1 which was extinguished after only 7.5 minutes.

The optical density of the smoke and the concentrations of CO and C0

2 measured in the doorway of the ventilated room for most of the tests were of

the order of 3.5 OD/m, 9 percent and 20 percent, respectively.

The time variation of the spatially-averaged temperature of the hot gas

in the upper part of the room due to fires involving typical furniture and

interior finish materials for all of the tests is shown in figures 13 (a) to

(e) . These gas temperature versus time curves characterize the intensity

and duration of fires likely to arise in residential rooms under the specific

conditions of this test series and cover all stages of the fire development

including growth, flashover, active burning and decay. The gas temperature

data shown in the figures were obtained by bare bead, 0.5 mm wire diameter

thermocouples for tests 1 to 7 , where the temperatures above 800 °C were not

reliable, and the fast response furnace thermocouples were used for the

remaining tests. No radiation correction was made. As shown in the figures,

the rate of gas temperature rise and decay varies appreciably between the

tests, depending upon the air supply rate, the fire load density, the thermal

and combustible properties of the wall and ceiling materials, and the

dimensions of the room involved.

5.1 Effect of Induced and Forced Ventilation

The change with time of the measured average gas temperature and the

computed rate of heat release within the fire room as a function of room

ventilation are shown in figures 14 and 15, respectively. For the tests with

the open door, the fires were frequently controlled at different stages by

both regimes of combustion including fuel bed and ventilation control. During

the increase in gas temperature covering the period from ignition to flashover,

the rate of burning was determined by the exposed surface area of the burning

fuels and the rate of energy feedback from the overhead flames, hot gas layer

and the surrounding walls. During the ventilation controlled active burning

period which immediately followed room flashover, the increase in gas temperature followed the slowly rising temperature of the walls and ceiling enhanced

by an increase in the burning rate due to the increased rate of fire-induced

air flow through the doorway opening. During the decay period of the fire,

the gas temperature continuously fell because of the decline in the burning

rate, which was controlled by the diminishing fuel surface area.

19
The time histories of the computed heat release rate for the two closed

door tests (7 and 11) illustrated in figure 15 were derived based on the

calorific value multiplied by the measured weight loss rate of the burning

sofa involved. As shown in figure 15, the heat release rate increased with

an increase in the rate of air supply to the room either by mechanical means

or by natural ventilation (tests 9 and 10) . However, with the door open over

the range studied, forced ventilation had little influence on the rise in the

gas temperature since the heat gain arising from the increased rate of heat

released in the room was compensated by the increased rate of sensible heat

carried out by the hot gases. For the room with the door closed, no room

flashover occurred and extinction of the fire was observed, in spite of the

presence of a small ventilating opening located between the bottom of the

door and the room floor because the oxygen concentration in the room depleted

to levels approximately 9 to 11 percent, no longer supporting flaming combustion.

The lower heat release rates for the closed room with forced ventilation

(test 11) compared to the room with the door closed and no forced ventilation

(test 7) as illustrated in figure 15, was attributed to test variability.

5.2 Effect of Fire Load Density

A typical plot of the time variation of the average room gas temperature

as a function of the fire load density for fires developed in the 3.3 x 4.9 m residential room with an open door as the single source of induced ventilation

is shown in figure 16. There was shift in the time of the peak in the gas

temperature versus time curves and the duration of the vigorous burning

stage of the fire varied from approximately 25 to 55 minutes depending on fire load density, which ranged from 30 to 45 kg/m2

. This time shift was probably due in part to an additional sensible heat loss and gas dilution

resulting from the greater volatile generation rate during the early part of

the test for the higher fire load density thereby lowering the gas temperature

in the ventilation limited burning regime. The different thermal inertias

of the bounding surfaces in these tests also influenced the time of the peak

temperature. In general, the time duration increased with an increase in

fire load density because of more combustible materials available for the

fire to burn. If the duration of the active burning stage of a fire is

defined as the time interval between the rise in the heat release rate

immediately following room flashover and the start of the decrease in the

heat release rate from fuel combustion, this duration can be approximated by

LA (AH ) t = K — (2)

Q

20
where L is the fire load density, A

f

is floor area of the room, A is

calorific value of the combustibles, K is a proportionality constant and Q

is the heat release rate during the active burning period. Equation 2 assumes that a linear relationship exists between the total amount of heat

evolved during the. active burning period and the total heat content of the

combustible materials in the room as listed in table 4. The quantity of heat

given off by the room fire during its vigorous burning stage versus the total

available energy of the combustible loads, for all tests with the exception

of those involving the door closed or very short test durations, is shown in

figure 17. The method of least squares was used to fit a straight line to

the set of data points. The ratio of the amount of heat released during the

active burning period to the total available energy from the combustible

contents and interior finish for these tests ranged from 0.34 to 0.69 with

an average of 0.54.

5.3 Effect of Lining Materials

Figure 18 shows the influence of the room linings on the average gas

temperature developed in the fire room. For the tests with the combustible

walls there is a depression in the gas temperature, presumably due to dilution

by the pyrolysis products in a ventilation limited fire, and an extension of

the duration of the fire because of the greater total fire load. The burning

of the combustible pyrolysis products beyond the doorway caused a large

extension of the flames which resulted in intense radiation levels outside

of the room as seen in table 7.

The variation with time of the rates of heat produced by the fires in

the larger rooms lined with various interior finish materials on the walls

and ceiling are given in figure 19. The heat release rate generally

increased rapidly up to its peak with increasing rate of fuel volatilization

from the burning materials until the fire became ventilation limited, and

then remained relatively constant during the stage of full fire development.

The heat release rate then dropped sharply with the decline in fuel consumption and the return to the fuel bed controlled burning regime. These fires

had similar heat release rates since the rates of combustion of the volatile

products and oxidation of the carbonaceous residues were limited by the rate

of air entering through the ventilation opening or the size of the doorway

which was the same for all of the tests. The rates of heat release by these

fires under ventilation limited burning appeared to be independent of the

types of wall and ceiling materials used. However, the tests with the plywood walls exhibited vigorous burning of the excess pyrolyzate outside of the

room.

21
5.4 Effect of Room Size

The variation with time of the average room gas temperature measured in

two sizes of the rooms having the same movable fire load density and ventilation conditions is given in figure 20. Both rooms had comparable maximum

temperature values, but the fire in the larger room had a slightly longer

duration. This was attributed to more combustibles present in the large

room to support combustion.

Figure 21 displays the measured total heat flux incident at the center

of the floor as a function of time for two room sizes. As illustrated in the

figure and also shown in table 7, the larger room flux values are generally

higher than the corresponding values for the smaller room during the peak

steady burning period. The radiation path length through the flame gas or the mean beam length for the room filled with an absorbing grey gas became

larger with an increase in the room size, and thus the radiation flux density

to the floor from the hot gases increased with the increased gas emissivity.

5.5 Effect of Ignition Source

The average gas temperature versus time curve for the test with an upholstered chair as the ignition source is compared in figure 22 with that

for a test involving a sofa normally used as the first ignited item. It can

be seen from the figure that both room fires generally had similar peak gas

temperature levels and fire duration, except that it took a slightly longer

time for the chair fire to involve the plywood panels on the back wall and

then to reach full fire development. The burning rate of the sofa was sufficient to bring the room to flashover without any involvement of the plywood

panels

.

6. DISCUSSION

6.1 The Experimentally Derived Temperature-Time Curve

The range of average gas temperature-time curves for this series of tests

in which ordinary domestic furniture and interior finish materials were used

as the fuel and burned under a variety of test conditions is shown in figure

23. These temperatures were obtained by averaging the readings of three

metallic-sheathed mineral-insulated fast response thermocouples placed within

the fire room. The gas temperature rose rapidly as flashover was reached,

fell slightly and then increased up to peak levels ranging from approximately

950 to 1100°C followed by a gradual decrease. There were wide variations in

22
the times to arrive at the peak temperature levels and in the duration of

the active burning phase, depending on the quantity and the nature of the

combustible materials and the dimensions and ventilation conditions of the

room involved. However, the maximum gas temperatures attained in the fire

room are similar for most of the tests. Test 9 represents the standard test

cohditions (i.e., a 3.3 x 3.3 m recreation room finished with plywood paneling

walls and a gypsum board ceiling and furnished with a typical loading of 23

kg/m2 of household furniture) . The standard ASTM E 119 curve in which the

measurements of gas temperatures are made with the ASTM thermocouples, is

also plotted in the figure for comparison purposes. An approximate correction made to take into account the slow response of the ASTM thermocouple at

the early stage of fire buildup was estimated to be less than 150 degrees C.

It can be seen that these basement recreation room fires had a more rapid

temperature rise during their initial stages and a shorter period of time at

high temperatures than those represented by the standard ASTM curve. Consequently, the rate of development and the intensity of real fires involving

typical furniture and interior linings are significantly different from those

represented by the ASTM E 119 fire endurance test.

The temporal variation of the mean upper room gas temperature derived

from averaging the readings of three specified fast response thermocouples

for test 9 is compared in figure 24 with that obtained from three ASTM E 119

furnace thermocouples placed at the same locations. As expected, the heavy

cast-iron covered ASTM thermocouples gave lower readings than the fast

response thermocouples during the pre-f lashover period due to the heavier

mass of its thermocouple assembly. However, both types of thermocouples

showed roughly the same temperature levels in the time of post-f lashover

burning

The set of data on the average room gas temperature for test 9 as given in figure 23 was fitted with a polynomial equation using a trial and

error method based on seven selected experimental points. The fitted curve

along with the experimental curve is shown in figure 25 and can be expressed

by the equation:

a + a, t + a_t 2 + a_.t 3 o 1 2 3 1 + b-^t + t^t 2 + k^t 3 (3)

where T is the average upper room gas temperature, in decrees C, t is the

9

time variable, in seconds; and

23
for 0 < t < 90 sec : for 90 <_ t <_ 3600

a = o 0.250000 x 10 2 a =

o 0.1076 x 10 4

a

l = -0.786992 x 10°

-2

a

l = -0.1360 x 10

a

2 = 0.822586 x 10

-4

a

2 = 0.5987 x 10"

a

3 = -0.283153 x 10

-1

a

3 = -0.2996 x 10

ii

rH 13 -0.344181 x 10

-3

b

l = -0.1753 x 10

b

2 0.396824 x 10

-5

b

2 = 0.1786 x 10"

b

3 = -0.153081 x 10 b

3 = -0.1112 x 10

This experimentally derived gas temperature versus time curve, which

corresponded to an approximate average over the range of test conditions

studied in the test series, was considered to be a more realistic representation of room fire severity in residential buildings. It is recommended that

this newly-developed fire exposure be employed for evaluating the fire performance of floor-ceiling assemblies used in residential basement (or other

similarly furnished) rooms.

6.2 Comparisons of Different Types and Measuring Locations of Thermocouples

In order to provide some information on the influence of different types

of thermocouples on the measured gas temperature during a room burnout test,

five thermocouples of different construction were similarly exposed during

one of the tests. The thermocouples consisted of (a) regular 0.51 mm wire

diameter single cable, (b) 0.51 mm wire double cables (i.e., the junction was made from wires from 2 different cables) , (c) silica fiber insulated, (d) standard ASTM E 119, and (e) fast response mineral insulated sheathed thermocouples.

These were placed 0.58 m below the ceiling at three locations inside the test

rooms. The silica fiber insulated thermocouples were constructed from 24 B.

and S. gage (0.51 mm or 0.020 in diameter) chromel-alumel wire insulated by

a 0.38 mm thick, non-impregnated vitreous silica fiber covering. Figure 26

is a graph of the average gas temperature as measured by the various types of

thermocouples employed in test 15. The regular 24 B. and S. gage single and

double cables and silica fiber insulated thermocouples led the sheathed fast

response and ASTM thermocouples in response to the rapid rise in room gas

temperature during the early portion of the fire test. As the test progressed,

the temperature lags became less due to the slower rate of gas temperature

change. Apparently, these bare beaded thermocouples had relatively small

time constants compared with the furnace thermocouples due to the smaller

mass of the measuring junction directly exposed to the fire environment.

24
However, all of these thermocouples, except the regular single cable type,

had nearly the same temperature-time curves after 5 minutes. It can be noted

that the regular single cable thermocouples began to indicate lower values

after the gas temperature exceeded approximately 900°C resulting from the

charring of insulation materials impregnated with silicone modified resin.

The temperature above 800°C cannot be measured reliably with these thermocouples.

The average gas temperature derived from averaging the readings of three

regular double cable 0.51 mm thermocouples in the fire room is compared to

that obtained based on 27 thermocouples in figure 27. These three thermocouples were positioned along the centerline of the room at 0.58 m below the

ceiling where the fast response and ASTM E 119 furnace thermocouples were

also located. The 27 thermocouples used for the mean gas temperature were

distributed evenly in nine equally-spaced stations and placed 0.025, 0.58 and

1.17 m beneath the ceiling. The average gas temperature based on three

measuring points was slightly higher, but generally within 50 degrees C of

that with 27 locations in the early stage of the room fire, since an inhomogeneity of the local gas temperature existed in the hot upper layer resulting

from the mass addition of fuel volatiles from the decomposing ceiling materials.

However, there was no appreciable difference in the temperature levels as the

test progressed and a uniform temperature developed within the room volume.

In general, the average gas temperature obtained based on the thermocouples

placed at three locations was consistent with that derived from the 27 evenlyspaced stations.

The average upper room gas temperature and the heat release rate at the

time of room flashover is shown in table 9. Flashover corresponds to the

time at which the newsprint indicator placed at the center of the floor

ignited spontaneously. The average upper room gas temperature varied from

test to test and ranged from approximately 625 to 820°C. There were some

differences in the temperature levels registered by the different types of

thermocouples due to thermal inertia as noted before. The average upper room

gas temperature at flashover was found to be 706 + 56°C with a 68.3% confidence

level and 706 + 92°C with a 90% confidence level based on the regular double

cable 0.51 mm thermocouples. The fast response and the ASTM thermocouples

had reached 326 + 123°C and 78 + 56°C respectively at the time of flashover.

Also, the rate of heat release by the fire at flashover appeared to be independent of the room size, over the range studied, and varied from 1030 to

2420 kW probably depending upon the insulation properties of the room walls

involved. The heat release rate at the time of flashover was found to be

approximately 1480 kW for the rooms lined with plywood walls and wood fiber

tile ceiling, 1780 kW for gypsum board walls and ceiling, 1510 kW for plywood

25
walls /gypsum board ceiling and 2060 kW for concrete block walls/gypsum board

or exposed floor joist ceiling, respectively. It can be noted that the heat

release rates required to produce flashover also depend on how long they are maintained and on ' the ventilation conditions of the room involved.

6.3 Doorway Flow Due to Ventilation Controlled Fires

Information relative to the induced flow rates of air entering and

combustion gases leaving through an opening in the fire room is essential in

the development of the mathematical models for predicting fire behavior. The

hydraulics orifice flow analogy has been used reasonably successfully for the

prediction of the fire induced flow through the opening of an enclosure

[9-11] . For a fully-developed room fire, the hot gas layer often descends

to the floor level immediately after the occurrence of room flashover and

thus the height of thermal discontinuity may be neglected compared to the

neutral plane height during the active burning period. At this vigorous

combustion stage, the mass flow rates of cold air in and hot gas out of the

fire room under "choked" conditions [10] can be expressed in terms of the

average gas temperature and the neutral plane height [9] as

”1 = §VoA

o ^

2 ?Ho

2

< 4)

m = -=C_,p A

o 3 D

K o o 'M>)('->)('- k)

(5)

where nr and m

Q are the mass flow rates of the gases entering and leaving

the room, respectively, is the orifice coefficient, p Qis the density of

the ambient air, A

q

and H

q are the area and the height of the opening,

respectively, Z^ is the neutral plane height, and T

q

and T are the absolute

temperatures of the ambient air and the room gas, respectively.

Summary data on the average room gas temperature, the rate of heat

produced within the room, the neutral plane height, and the measured mass flow

rates of the hot and cold gases at the doorway during the active burning period

for the tests performed are tabulated in table 10. The height of neutral

plane above the floor was determined from the measured velocity reversal in

the doorway opening. The table also lists the air inflow rates estimated

from the average heat release rates along with a K value of 17010 kJ/m 3 of

oxygen consumed. As shown, the calculated rates of air flowing into the

room are normally higher than those obtained from direct measurement.

Gas velocity measurements were made along the vertical centerline of the

26
doorway opening. An experimentally determined coefficient, which is defined

as C v

= m/mL

, in which m and m

L are the mass flow rates based on average and

local velocities, respectively, is introduced to account for the variation

between the average and the centerline velocity. The ratios of the orifice

coefficient to the correction factor for velocity distribution across the

doorway width, C^/C^ for gas inflows and outflows are determined with the use

of equations 4 and 5 together with the measured or estimated flow rates, the

average room gas temperature and the neutral plane height, and are tabulated

in table 10. These doorway coefficients ranged from approximately 0.4 to

0.7 with a mean value of 0.54 for inflow and 1.0 to 1.6 with an average of

1.29 for outflow streams, respectively.

In the ventilation controlled burning regime, the amount of oxygen

contained in the cold air entering into the fire room was consumed completely

by combustion since the measured oxygen concentrations at the top part of the

doorway opening generally dropped to zero. From the heat release rate data

along with a calorific value of 18.6 MJ/kg for the room combustibles, the rate

of burning for the rooms with an open doorway can be related to air supply or the size of the opening and expressed by the approximate relation: m = (0.112

1/2

+ 0.006) A H , where m in kg/s, and A and H are the area, in m

2 and the — o o o o height, in m, of the opening, respectively. It can be seen that for these

1/2 test conditions, the ratio for m/A H has a constant coefficient of

5/2 ° ° 0.112 kg/m s, which is approximately 24 percent greater than the conventional

value of 0.09 kg/m^^s derived from the burning of wood-like materials [12,13].

6.4 Gas Temperatures Influenced by Types of Ceiling Materials

Typical vertical temperature profiles of the gas near the center of the

fire room for a test involving concrete block walls and wood joists/plywood

subfloor ceiling (test 15) and one test with gypsum board walls and ceiling

(test 13) for various times after full-involvement of the room are given in

figure 28. These profiles illustrate the nonuniformity of the temperature

distribution in the vicinity of the combustible ceiling attributed to the

cooling effect of the mass injection of gaseous volatiles from the decomposing

cellulosic materials. As shown for the gypsum board ceiling, the room gas

is well-mixed and exhibits no variation in gas temperature during the vigorous

burning stage of the fire.

27
6.5 Incident Heat Fluxes Calculated from Surface Temperatures

A method was sought for estimating the heat fluxes to the bounding walls

during the room fire and for assessing the accuracy and reliability of the

direct measurements by heat flux gauges. The total heat flux incident at the

enclosure walls can be calculated using the surface temperature histories and

the assumption that the ceiling or walls are semi-infinite solids. The

equations and procedures used for the computation are presented in appendix

B. Figures 29(a) and (b) show a comparison of the calculated and measured

surface heat fluxes incident at the center of the gypsum board ceiling and at

the mid-point of the upper half of the west gypsum board wall as a function

of time for test 6. As shown, the calculated peak heat fluxes are generally

in reasonably good agreement with the measured values. However, the measured

heat fluxes are to a water-cooled heat flux gauge and should be higher than

the calculated ones. Some of the differences between these curves are due to

the fact that the calcining of the gypsum board is not taken into account in

the calculation.

7. CONCLUSIONS

Based on the experimental results of 16 full-scale fire tests involving

furnished basement recreation rooms with prescribed variations in room size,

fire load densities, interior finish materials and ventilation conditions,

the following conclusions may be drawn:

1. The rate of development and the intensity of real fires involving

the burning of typical furniture and interior linings in a room

during the first 20 minutes may be significantly greater than

those defined by the ASTM E 119 standard time-temperature curve. A more realistic time temperature curve for residential

occupancies is presented in this report. This curve is considered

suitable for testing exposed floor constructions, floor-ceiling

assemblies, wall assemblies, columns or doors.

2. Room flashover did not occur for a closed room of the size tested,

even with normal forced air ventilation. However, for fires

involving a single open doorway, the rate of heat release within

the room increased with the addition of a forced air supply.

3. The addition of combustible interior finish materials on the

wall and ceiling did not increase the maximum temperature of a ventilation limited fire over that produced by the combustible

28
contents but did prolong its duration. Similarly, the fire

duration time increased with the fire load density. While the

intensity of the fire within the room was controlled by the

available air supply, the rate of burning outside of the room

was markedly increased due to the additional fuel produced by

the combustible finish and thus could be a major factor in fire

spread beyond the room of origin.

4. The effect of the room size on the fire intensity was small,

although the duration of the fire increased with an increase

in room dimensions for the same movable fire load density.

8 . ACKNOWLEDGMENTS

The room burnout test program was accomplished through the team efforts

of many members of the Fire Safety Engineering Division, especially

Messrs. Thomas Maher, Oscar Owens, Charles Veirtz, Ben Ramey, William Bailey,

Sam Steel, Daniel Debold, Roy Lindauer, Billy Lee and Douglas Walton, whose

valuable assistances are gratefully acknowledged.

The authors wish to thank Mr. Jack Bumgarner of National Archives and

Records Service, GSA, through Mr. Harold Nelson of the Center for Fire

Research, for providing disposable old record files used in the fire

experiments

.

This project is being sponsored by the U.S. Department of Housing and

Urban Development, under the program management of Mr. James C. McCollom of

the Division of Energy, Building Technology and Standards, a component of the

Department's Office of Policy Development and Research.

29
9. REFERENCES

[1] Thomas, P.H. and Heselden, A.J.M., Fully-Developed Fire in Single

Compartments, A Cooperative Research Programme of the Conseil

International du Batiment, Fire Research Note No. 923, Fire Research

Station, August 1972.

[2] Issen, L.A. , Single Family Residential Fire and Live Loads Survey,

Nat. Bur. Stand. (U.S.), NBSIR, in process.

[3] McCaffrey, B.J. and Heskestad, G., A Robust Bidirectional Low-Velocity

Probe for Flame and Fire Application, Combustion and Flame, 26 , 125 (1976)

[4] Thornton, W.M. , The Relation of Oxygen to the Heat of Combustion of

Organic Compounds, Philosophical Magazine, 3_3, 196 (1917) .

[5] Parker, W.J., An Investigation of the Fire Environment in the ASTM E 84

Tunnel Test, Nat. Bur. Stand. (U.S.), Tech Note 945 (August 1977).

[6] Christian, W.J. and Waterman, T.E., Characteristics of Full-Scale

Fires in Various Occupancies, Fire Technology, 7/3) , 205 (1971) .

[7] Huggett, C., Oxygen Consumption Calorimetry, a paper presented at the

1978 Technical Meeting of the Eastern Section, The Combustion Institute,

Nov. 29-Dec. 1, 1978.

[8] Harmathy, T.Z., A New Look at Compartment Fires, Part I, Fire Technology,

8(3)

,

196 (1972)

[9] Rockett, J.A., Fire Induced Gas Flow in an Enclosure, Combustion Science

and Technology, 1_2, 165 (1976).

[10] Prahl, J. and Emmons, H.W., Fire Induced Flow Through an Opening,

Combustion and Flame, 2_5, 369 (1975).

[11] Quintiere, J.G., The Growth of Fire in Building Compartments, ASTM-NBS

Symposium on Fire Standards and Safety, National Bureau of Standards,

Gaithersburg, Md. April 5-6, 1976.

[12] Heselden, A.J.M., Thomas, P.H. and Law, M. , Burning Rate of Ventilation

Controlled Fires in Compartments, Fire Technology, M2)

,

123 (1970)

[13] Bullen, M.L., A Combined Overall and Surface Energy Balance for FullyDeveloped Ventilation Controlled Liquid Fuel Fires in Compartments,

Fire Research, 1 , 171 (1977/78)

[14] Carslaw, H.S. and Jaeger, J.C., Conduction of Heat in Solids, Second

Edition (Oxford University Press, Oxford, 1959).

[15] Myers, G.E., Analytical Methods in Conduction Heat Transfer (McGraw-Hill

Book Co., New York, 1971).

30
Table

1.

Types

of

interior

finish

materials

found

in

basement

rooms oco

OOVO

CO oo

VO

CO oo

13

0) rC

in •HG

13P

03 £ 0 3 A W rH

Qd rH

>i fd o £

13ooItrH 0

fir 2

r- o>

o\ tN

CN

>nUc w o H

IT! rH

rO

m s s

'a* t

d) £

cn HCH

Pm

13P

03

g o Q) 3 £1 -P U

rH

W

Qj. <—

03 >1 m i—

o is CP

'O

!m - id -a o o A 9 ^ S <U X)

>. P o i—I -H O

ft Ip £

r-CM

O'. VO

VO

CM

Cr- G

•H

rH•H

0) U

13

<u

w

in -P o o w a O -H X g o W h) o

LO

LO

CM

CT> VO

CD

4-1 tT 0 03 (0 P P S 13 G

Q) 0 G a) n 0 (d u

Pi

p 3 CL) S a.

GO

•HP

03<

uo

a)

p

•H P

a) ,CPO

31 f
Table 2 List of combustible contents

Initial

Weight

Item (Kg)

Sofa

Upholstered

Chair

Ottoman

Loveseat

Coffee

Table

End

Table

Bookcase

Carpet

Material Size

Grade 1 100% olefin up- 2.15 m wide

holstery fabric, cotton 0.85 m deep

batting side arms, foamed 0.81 m high

polyurethane seat, back

and side cushions, wood

frame

frame-24

cover-6 .

4

foam-6 .

8

springs-7 .

3

cotton batting/

cardboard-8 .

5

Total 53

Grade 1 100% olefin (1) 0.78 m W (1) 24

cover fabric, urethane 0.90 m D

foam seat cushion. 0.91 m H

cotton batting, wood

frame (2) 0.79 m W frame-11 .

3

0.91 m D cover-1 .

8

0.93 m H foam-2 .

3

springs-1 .

4

cotton batting/

cardboard-6 .

2

Total 23

Grade 1 100% olefin 0.61 m M frame-4 .

5

cover fabric, urethane 0.46 m D cover-0 .

9

foam. wood frame 0.43 m H foam-0 .

9

cotton batting/

cardboard-1 .

4

Total 7.7

Grade 1 100% olefin

cover fabric, cotton

batting side arms,

urethane foam seat,

back and side cushions,

wood frame

Laminated veneer top

and sides, finished

hardwood

Laminated veneer top

and sides, finished

hardwood

Wood particle board

with melamine finished

surface

Level-loop olefin

pile with foam rubber

backing

1.52 m W

0.85 m D 40

0.81 m H

(1) 1.52 m W (1) 20

0.56 m D

0.43 m H

(2) 1.52 m W (2) 22

0.51 m D

0.41 m H

(1) 0.58 m W (1) 10.9

0.69 m D

0.51 m H

(2) 0.66 m W (2) 7.3

0.36 m D

0.56 m H

0.76 m W 42

0.36 m D

1.83 m H

3.2 x 3.2m 20

x 9 .

5

mm thick

32
Table

3.

Experimental

conditions

of

basement

room

fire

tests

ition

Condition |1

Forced 3

Air

(m

/s) o o

P-Oo o o o O o o

0.047

poo o o o o o o

m! ml

CD CD c p c g c G g G CD c c g CD c c c c C c

4> o CD CD CD CD CD CD 0 CD CD CD 0 CD CD CD 4) QJ 4) > o a ft ft ft ft ft i—

ft ft ft rH ft ft ft

a a, ft

p o o o o o O

u

O o o u

O

O o o o o H OJ t CO O'

CM CM P O LO P p 00 rH rH o CM vo ft P -H c p* ft O i—

VO 1—

VO vo P P O p' 0 pn ft on Os] on CM LO on on on on on on on on on on

CM Eh \

CD on o CM

O i—

rH rH CM CM CM CM CM on o rH CM a, • • • • • • • • • • • • • • • • P rH CM CM CM CM CM CM CM CM CM CM CM CM CM CM CM CM >i rd +j u

CD£ !h

W

CD 0 0; o •"i rH LO

O LO vc vo LO

CC an

O LO LO CO

*H & P CD o O CM CM i—

vo i—VO rH rH CM CM GO LO CM ft. ft

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fd -P £ 0 G -H

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fa CD rH CD rC -P •H £ ft on ft rH •—

ft O* on On cr. On ft ft on ft On O'. P CD • • • • • • • * • • • • • • • •

U3 4-> CM CM PJ LO

p CM CM CM CM CM CM CM CM CM CM CM CM 0 C CM CM CM i—

on CM CM CM CM CM CM CM CM CM CM CM CM

fa 0 E U 0U

a) CD rH rH

*Tj ml ra mi ml •H mi ml ml TJ mi mi •H CD P )4 u IH G G Eh G g G G >4 g Eh CD 4-J

Cr. fd fd fd fd fd fd fd fd fd fd fd fd fd 44 CD C 0 0 0 0 0 0 0 ml 0 0 0 0 0 0 ml CD •H

U

ffi ft

m ft ft

mm CD ffl ffi

mm ft ft CD •H 0

ml ml 0

fd •H

p pj pj E pj

gg

G

gg

g

g pj

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C P) •H CD p p P p P P CD p p P P p p CD rH p U CD CD CD CD CD CD CD ft CD CD CD CD CD CD ft ml CD

4) ft ft ft ft ft ft ft CD ft ft ft & ft ft CD 0 CD p >H s >i £ >i 2 >1 >i s 3 0 44

fd o o e>

u e>

u o ft

u o o o o o ft & on s

tr> G rM o a O £ 0 m3 0 o

CD rHG rH 1—

m fd fd ft ft 0 o

rd CD ffl ft CD CD !3 TS 4-> ml mi ml mi ml mi ml mi ml T3 m3 44 44 o CD o 0 0 g 0 o 0 0 o 0 0 E 0 CD CD 0 P 0 0 0 p 0 0 o 0 o 0 0 0 0 IH V4 3 O 3 3 3 CD 3 3 3 3 3 3 3 CD 3 O u c >i >i ft >i >1 >1 >i !>i I>1 G G

rH 0 •H rH l—

£ 1—

rH rH I—rH rH rH s 1 1 o 0

ftu ft ft ft e> ft ft ft ft ft ft ft

u ft

uU s CD — CM CM CM CM CM CM CM CM CM CM CM CM VO VD VD vO vO n N CM • • • • • • • • • • • • • • • • • 0 •rH g O o O o O o o o o O o o LO LO LT) LO on

ft ft i—

rH i—

rH rH rH rH rH rH rH i—

rH i—rH

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-p

CD

C <

CD 0 i—

CM CO 3' LO vDp CO ft

O i—

H CM on LO LO 2 rH rH rH H •—

rH rH rH

33
Table

4.

Initial

weights

and

total

available

heat

of

combustible

loads

used

for

fire

tests

Q) —

XI

rtf rH

•rH o O

O o o o o o o O o o o o o o < CP cO CP in o p o CM O o p CP CP 1—

CO lO

p

p CO co CO 00 o (P CP rH P cO o CD p CO p p~ p p> rH CP r—

rH

rH rH rH 1 rH

4-10H

c/i CD x:

c/i

tpc

-rH -—. CM

O

4t P- **

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rH fjl -H XI

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305. 152. .0

•H

l/l cO *^r

H1 in CO or CO in o o 00 o o

<D r—1 -—

o CO

p

O 00 P 00

p CM CM CO CO CM o o

CM CM rH cO rH rH 1—

•—

CM CM

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rH rH rH rH rH rH rH t—

rH

r

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4->

(1)

— rH

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m

a CP U M CP o o rH rH rH CO CO 'tf CM CM CO in o CO

rtf U

rH CM CM CM CM CM CM CM CM CM CM CM CO

u

0) ^ CP CO CM •'tf CP 00 00 rH CO in CM rH co o m in a CP CO IT) in in CM m CP CO CM rH CM p- CM

Pu

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p (M COrH

p

p P p

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rH

rH

rH i—rH

rHrHrHrH

4Jx:

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CP CP

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CM CM CM CM CM P

cy

-p

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CO CP

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CD X! 1 1 1 1 1 1 1 <P P- rH rH

O

CO CO

H1

PJ

CD CD rH .—

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H1 CP CP

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rtf

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p

p

p P P P CO pp CO P' p* ru

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CP CO CO cO co P- CD

p

rtf tJ x: x< u —

24. 23. CM

24. 23. 24. 1

24. 23. 23, 24, 23 23 23 23 22 23 22

rtf ^ CP CP CO CM

m CO in CM CO o CO CP CO CO in •^r in

44

D

CO in p*

p

m o p CO in o co CP co

CO

— in LD LD

m in in

m in in in in

Test No. r— CM on

H1 in co P CO CP 10 11 <

r-r— 12 13 14 15 9T

uooI—

44

XiP

(/)

T3Oo£r-H&G

rtf W

4-1

C/1 -Ho

TlOo2

c/i 4) T53

rHog

-K **

34

includes

plywood

subfloor
Table 5. Test log during test 14

Time

(min : s) Events

0 Ignition

0:04 Light smoke rising from the back of newspaper

0:15 Smoke over entire newspaper

0:19 Smoke almost to ceiling height

0:22 More smoke on the R.H.S. of paper than the L.H.S.

0:25 Smoke increasing more on the R.H.S. than the L.H.S.

0:35 Smoke hitting ceiling and spreading radially

0:38 Flame visible on left side of newspaper

0:50 Dark smoke halfway across ceiling

0:58 Back-rest cushion of sofa involved in burning

1:05 Flame height of about 0.91 m (3 ft) above sofa

1:11 Entire newspaper involved in burning

1:22 Hot smoke gas flowing out of the doorway opening

1:25 Central portion (1/3) of sofa involved in burning

1:33 Flames almost reaching to ceiling

1:40 Center half of sofa involved in fire

1:45 Smoke layer depth descending to 1/2 of room height

1:48 Fabric of arm cushions on sofa melting and splitting

1:55 Flames engulfing 2/3 to 3/4 of sofa

2:02 The flames halfway across left seat cushion, and still localized

2:05 Arm cushions burning and coffee table pyrolyzing

2:12 Wall paneling next to sofa involved in fire

2:19 Light smoke flames emerging from the top of doorway

2:20 Top of coffee table is burning

2:23 Newspaper indicator on floor ignites and is burning

2:26 Filter paper flashover indicator on floor catches fire

2:25 Heavy flames coming out from the doorway

2:30 Olefin carpet covering floor ignites

35
Table 5. Test log during test 14 (continued)

Time

(min: s) Events

2 : 35 Ottoman in front of bookcase is burning

2:38 Hot flaming zone extends down to 0.61 m (2 ft.) above floor

2:45 The whole front of the fire room appears to be in active

burning

3:39 Intense radiation emitted from the room through the doorway

3:49 Crackling sound occurs within the fire room

3:55 Mostly flames coming out of room down to 1.5 m below ceiling

in plane of doorway

4:20 A photoflood light exploding to produce a loud sound

5:35 Trace of smoke leaking out of room through cracks on the back

concrete block wall

5:55 Carpet is virtually consumed *

6:25 Curls from wall paneling fall onto floor

7:00 A piece of ceiling tile drops down on the left load cell

platform containing chair

8:05 Ceiling tiles almost burned away

10:05 The interface between the outgoing hot gas and the inflowing

cold air at the doorway is still at 1.5 m (third velocity

probe above floor) below ceiling

14:05 The upholstered chair starts collapsing

15:30 Smoky hot gas in the room begins lightening up

16:10 The bottom boundary of flames coming outside of room through

doorway still at level of third velocity probe

16:15 Chair collapsed

16:45 Crackling within the room is decreasing

17:10 Bright orange flames covering floor and dark orange flames

coming out of doorway

19:10 Flames in room become brighter

20:10 Bright orange flames filling the entire room volume

20:55 Most wall paneling are consumed; some furring strips still

burning

21:05 Flames out of room down to 1.14 m (at fourth velocity probe

above floor) below ceiling in plane of doorway

22:45 Crackling in the room is almost gone

36
Table 5. Test log during test 14 (continued)

Time

(min: s) Events

25:10 Piles of paper from bookcases still burning on floor

25:30 Very light intermittent flames coming out of doorway; no

black smoke visible in room

29:25 Little flaming is still going on some furring strips near

room corners; piles of glowing embers from bookcases remain

on floor

30:00 Stoppage of tape recording

37
Table

6.

Summary

of

critical

times,

heat

release

rates

and

gas

temperatures

0)00

cd

U • an t—

CM 00 LO cn <r CO on o • h ai a -— rH on 1 LO CO LO O no o ON rH LO 00 X 0) D cc e o co CO 1 On a> an o o o rH o on o o o ON

CO > a ct <U o rH rH rH rH rH rH rH s < o H A A A A A u cu

cO 4-1 0) cO X PS

Bo o o o o o o o o o o o o o B CU O 3 ON 1 co 0M no i on o i i rH On nO LO o 3 CO OS OS >—

LO 1 LO LO i LO nO i i CM CM rH o CO

B cd on CO LO LO LO no nO LO no nO LO

•H a) c X i—

cO cu S os co 4-J *H (U

m o 1^- 00 on CM LO

O 00 00 cn LO o 4-J 4H a CM 00 < 1 an rH o cn 1 1 o nO CM CM 00 rH LO

•H o S3 c cO CM CM CM rH CM cm Csl CM CM CM CM rH CO co

00

u

/“N

M cBSw/' 00c o CO •H cO CM CO LO CM

O CM LO CM LO CN oo CO CM

4-1 CU

w

B Z O CN On o 00 00 o On 00 CM i—

no On co

B V4

O i-t 0) cO (U H o CM CM rH CNJ 1 1 rH CM 1 1 1 1 cn Csl CH rH «H o 1 1 B M-l O c

cO pH cu

Q

H

cO NJ

0)

B

a o u LO o 00 o LO cn LO on CM a \ o CTn CTn no 00 nO O On O rH 00 00

<C 12

on lo \ 1 CM rH tH 1—

1 1 CN rH rH co rH CM CM 1—1 CM

a) o M—

B Csl cO rH

P-l

cu>c >4

B

X <u

a CM LO CO LO COO CM o CO 4.) at O r^- o CTn o rH 1 1 00 o o i—

rH 00 1 1 LO

*4 cO M—

in

a

M-l I—

o H cO CN CM CM rH CM CM CM rH CM Nt CM CM rH CM CM

pH

a

a

<U c B B o H o 4-J dS •H 1 a on LO r-* LO CO 00 o oo LO

O oo LO 00 CM c co qj CTn LO CTn oo On 00 O 00 00 00 rH o O LO

oc *H • • • • • • 8 • • • 8 • • • • • • 1—1 01 U X a

rH CM rH rH rH 1—1 CM 1—1 1—1 on CM CM rH CN CM

CU o o U CU C c-n B cd a> cO a c LO CM oo oo CM o Csl LO o CO LO

O Csl cn o O

B rH cO -H on CH CNJ on no LO nO on rH nO i—1 nO LO

•H pH QJ

B

H

a ^a o o o o o O o O o rH o O 1 1 o o o o

<0 iH

cd • •H B a

4-J o B

u <*• o 1 o rH on no nC o oo 00 CO co cn CM LO

•H o (D o CM CM i CM Csl CM CM CM cn CM CN CM CM CM CM CM CN c OS H "a Mco

*H 4-J LO U 4-1 CO C • cO o at -h rH o o o O

O o O o o O o O

O o P H B rH CO CO nO nO no nO nO nO nO nO nO nO nO nO nO 3Q

4-J

es No rH CM CO LO L0 r- 00 on

O rH 3 CM CO LO nO

H rH i—1 rH «H 1—1 t

i rH rH
Table 7. Summary of incident heat fluxes

Maximum Heat Flux to

Ceiling Walls Floor Outside of Room

Test Time Total (kw/m2

) Time Total (kW/m2

) Time Total Time Total

No. (min ) Max. Avg. (min ) Max

.

Avg

.

(min ) (kW/m2

) Max

.

(min ) (kW/m2

)

Max

.

1 2.3 106.0 99.9 2.1 116.0 90.7 3.1 133.0 — —

2 8.9 115.0 67.9 12.0 137.0 119.0 3.5 23.2 — —

3 12.9 107.0 75.5 11.5 192.0 150.0 11.3 120.0 — —

4 2.1 98.5 79.1 13.6 145.0 124.0 2.8 143.0 — —

5 12.6 124.0 73.8 15.1 180.0 146.0 12.9 138.0 — —

6 11.6 167.0 85.2 12.1 220.0 194.0 3.2 132.0 — —

7 1.9 32.7 26.0 1.6 36.7 19.5 2.0 2.7 — —

8 14.1 167.0 138.0 12.3 181.0 133.0 13.7 130.0 — —

9 13.3 87.0 47.4 15.3 178.0 154.0 13.6 124.0 — —

10 12.3 153.0 82.1 13.7 176.0 157.0 13.6 112.0 4.3 33.1

11 2.0 38.8 25.6 2.0 20.7 14.8 2.0 2.7 — —

11

A

13.6 121.0 82.1 15.3 162.0 132.0 5.1 119.0 5.6 33.2

12 19.1 130.0 72.2 24.4 136.0 112.0 4.0 134.0 6.3 27.4

13 16.4 138.0 101.0 18.3 229.0 175.0 13.6 162.0 — —

14 22.6 176.0 126.0 23.7 186.0 144.0 2.7 117.0 3.7 36.9

15 26.1 171.0 147.0 32.9 166.0 132.0 25.9 142.0 13.3 39.6

16 24.8 138.0 94.8 31.6 148.0 113.0 22.0 120.0 13.2 29.2

39
Table

8.

Summary

of

air

flows,

static

pressures,

smoke

densities

and

gas

concentrations

CN <3* 00 rH oi pH r- 00 1 I i i i i 1 i X 01 00 LO N< CN o CN ro u

(d •p

<A° i

• C\) r-H o r-H ro CO rH t" CO 00 N* rH CN rH N*

O CO 0 u >i • • • • • • • • • • • • • • r > id r- CO 01 01 01 1—

CO 01 o O 00 o CN

O 01 01 ' ' 5 i—

rH pH rH rH rH iH CN rH CM rH rH CM CN CM rH Po0O coH

-P N* r- N’ n* ro CO o rH ro CN

<d .x • • • • • • • • • 1 l i 1 i i i P o o o o o o o o o o O i i i 1 i i i -P (d c +j

cu CO oc o 0 u in 00 CN 10 ID ro o CM o r- ID 00 o CM

N1 u • (0 O', in oi O Cl LT) rH o Oi 01 o 00 r-' rH CM 00 • pH rH rH rH X P

fd o S oQ

CD in 00 ro 00

N1 lO CO

rH CO CO CO ID CM 01 rH l 10 N* N*

N1 CN £ CN c- lO 00 ro CN i t • 4 i • • • • • • \ faS • • • • • • i i N* CO CN CN ro CM CN ro • 0 co CM CN CM CN ro

<uQ o X fa 0 c e —

in >1

-P LO in N* 01 CO ID C" in CN 01 o 00

N1 • •H rd • r* rH ID CN CN i i c ID 00 i o r- N1 CM c O'. X IS3 5 Cn • • • • • i l • • • \ • • t • • • d C P > i—

in CM ro CO CN N* CM COm N*

N1 N* ro S 0) 0 <c c o

—•

Q

rH •P co CN 01 Oi 00 lO oi rH 00 iO 10

id <u fa r* ID i CO 10

m i m i0 in i 10

m r-'

m 10 10 c P C 0i — • • i • • • l • • • i • • • • • • •H -P rd •H £ o O o o o o o o o o o o o o a 3 r-l 0)

<D fa fa S

CD o p

•H

p • E -P in rd f" 10 CD 01 00 rH CO CN LO o 01 CD CO X O rd in fa l • i • • • i • 1 (0 O -P 0) — i ro i ro N* i CO 1 N* rH N* co N* CO 3* N1 a fa in p rH rH rH pH rH rH CO 1 1 rH pH rH rH rH

fa

fa U) \0

m CO

m CO 00 00 in oi

m

tm \ *—

rH 00 Oi CO 00 CO 10 in 3 3 Cn +J i • i • • • i • • • 1 • • • • • • 0 0 A 3 i eg i CN CN CN i CN CN ro 1 CN CN CN CN CN CN I—

P O

fa fa >,

Eh (0

X <D t 00 CO vO 00 i—1 m CN o 00 10 rH

(0 -P

O i 01 i in ro

O i rH CN CN 1 CN ro

m o ro

am o 3 i • i • • • l • • • 1 • • • • • • fa c H o rH rH rH rH rH rH rH i—

rH rH pH rH

4- # <

in 0 pH CN ro N' in ID 00 01 o rH

r

—

CN ro N* LO ID

QJ

a rH

r

—

i—

rH pH 1—

rH rH

tr

40
Table 9. Average upper room gas temperature and

heat release rate at time of room flashover

Test

No.

- - ...... ... ... . ... f Average Upper Room Gas

Temperature (°C)

Heat

Release

Rate

Regular (kW)

Fast

Response

ASTM

E .119

1 624 1030

2 822 — 1620

4 694 — — 1190

5 710 — — 1880

6 757 — — 2420

8 726 264 52 1560

9 705 214 49 1610

10 763 325 54 1850

11A 755 398 86 1470

12 699 254 50 1560

13 632 256 45 1130

14 685 231 36 1390

15 666 394 118 2360

16 640 602 212 2210

41
Table

10.

Average

room

gas

temperature,

heat

release

rate,

flow

rates

and

calculated

orifice

coefficients

for

inflowing

and

outflowing

gases

during

active

burning

period

/-SQ /—N a 13 v— o CD OS 43 o rH

m o CN CN sO 00 rH Oh LO

rH 44 Os as o CN CN •vt LO LO co LO co co CN o 44 3

id 4-J 03

O o 1—1 rH rH rH rH rH i—1 1—1 rH rH rH

a) 3 cd

•rl

O CD o E

•H

4-1

4-4CDo • CJ 1 13 ad •H G S • O 00 1—1 i—1 CN .<fr Os o as CN CO

4-) 4J o ad sO as rH CN LO CN 00 CN i—1 o G £ CO G 4-i S o w 3 44 o o o *H rH rH rH 1 1 rH o 1 1 rH i—1 u 1—1 o 44 6 G Q M 13 1 G sO CO o SO rH CO CN CO CN rH rH CO G ! -A LO sO LO SO LO LO SO <!* LO G 3 E 03 o o o o O o o O o O o o o

/-N !3 /'"S 03 o 13

V. lH

G o 04 sO LO LO Os 00 sO as CN as

oo CO 4-1

G >4 00 so 00 CO CN 00 LO "d- CN CN rH

dd cd 4-1

G 3 o 3 E m rH «H rH CN 04 CN CN C\1 CN CN CN CN CN O v—

GuG • ad & 1 13 ad o •H G B • o Ol CN SO LO CN 1—1

LO rH CO sO 5 4J 4-1 o ad o o CO sO SO 00 00 00 SO <r o 44 m G 44 r—1 G w B 44 dd 1—1

rH rH 1 1 i—irH rH 1—1 1—1 i—l i—l

rH rH

P-i

M

13

03 1 G ON as 00 rH sO 00 rH as o 00 00 LO o

03 •H

G ! -4 00 as 00 00 sO LO LO rH LO sO

cd <d

G 3 s E 03 o o o o O o o o o rH o o o G • -U 03 G

ao cd

G 4-J o o CO o O o o o o o o o o > CD G G E 1—1

CN sO so CN 00 OS rH < X » 1 Pi Pi o o as as m as sO LO rH rH OS <3* G CO CO m <t LO m LO LO LO

Pi

rHG

G 4-J • >4 C 3d CO LO as rH LO CO as 00 CN co

ao 4-1

G ao sO sO SO so 00 > G 1—1 •h a • < CD Pu (D W o o o O o o o o o o o o o E X B { ao o 03 r^- sO CO CO CN o i—1 as as > o G G O LO O o rH as LO 00 as CO o sO o i—1 <d ad O E-H 'w' 00 00 oo 00 00 00 oo 00 Os 00 00 00

/^Na

•H 00 CN rH o <r 00 a o • ai ao H CN LO 00 CN o CO o LO o > G rH 1—1 rH rH CN t—1 rH i—1 CN CN CN CO CO

•H •H TJ U G O u M •H

a < 3 G o sO as o os LO o o o CN CN as rH

CQG

P-i P4 04 rH CN rH CN CN CN CN CN rH CN CN LJ • 03 O 04 "d" m sO 00 as O CN CO LO sO G E H

rH rH

rH rH rH «—i

rH 1 1 1

42
STEEL

I-BEAMS

0.41 m

O.C. £

LJ>Zo5

LU

LU

LJ><

_l

CL eooP

•P

tfl CD -p

<D x:

-P

M-lO

cn

fd -p

cuCo

•H

-popp

-pwcou

<DPH

(L

43
3200

mm

44

Figure

2.

Layout

of

furniture

and

wall

linings

for

the

3.3x

3.3m

test

room
Figure 3. Layout of furniture and wall linings for the 3.3 x 4.9 m test room

45

4826

mm
FREQUENCY

(%)

(psf

)

0 I 2 34 567 8 9 10 II 12/over

0 9.8 19.5 29.3 39.0 48.8 58.6/over

TOTAL FIRE LOAD DENSITY

( kg / m

2

)

Figure 4. Frequency distribution of fire load data for

residential recreation rooms

46
(psf)

TOTAL FIRE LOAD DENSITY

(kg/m2

)

Figure 5. Frequency distribution of fire load data for utility rooms

47
FREQUENCY

(%)

(psf)

01 2 3 45 6 78 9 10 II 12 /over

TOTAL FIRE LOAD DENSITY

(kg/m2

)

Figure 6. Frequency distribution of fire load data for other basement rooms

48
(psf)

( kg / m2

)

Figure 7. Frequency distribution of the movable fire load density

in basement recreation rooms

49
Key to figures 8a and 8b

Symbol

X

®

z

m

o

O

A

A

Type of Transducers

3 Thermocouples

7 Thermocouples

5 Thermocouples

5 Thermocouples

11 Thermocouples

1 ASTM E 119 and

1 Fast Response Furnace

Thermocouple

1 Total Heat Flux Meter

1 Total Heat Flux Meter

1 Total Heat Flux Meter

1 Radiometer

5 Static Pressure Probes

1 Static Pressure Probe

3 Gas Sample Ports

1 Load Cell

Sprinkler

Distance Below the Ceiling (m)

0.025, 0.58, 1.17

0, 0.025, 0.15, 0.30, 0.58, 1.17, 1.73

0, 0.025, 0.30, 0.58, 1.17

0, 0.025, 0.58, 1.17, 1.73

0, 0.025, 0.15, 0.30, 0.58, 0.89, 1.17,

1.45, 1.73, 1.93, 2.31

0.58

0

-0.09 (above the ceiling)

2.34

0

0.10, 0.30, 0.76, 1.32, 2.13

0.10

0.38, 0.69, 1.07

50
SOUTH

WEST

T

o

rom

o

N-o

E

E

o

N

O

8

in

EAST

Figure 8a. Plan view of the 3.3 x 3.3 in test room

showing instrumentation layout

51

NORTH
SOUTH

WEST

Figure 8b. Plan view of the 3.3 x 4.9 m test room

showing instrumentation layout

52

NORTH
DISTANCE

FROM

TOP

OF

DOORWAY

(m) IOW IO o — o o

roo

10<Md

00 U)

<0 o d

roto *

00 (MO

<t> — ro

m

IO

roN 0>

N

o>o

CD

0>

00o >X > X X >X >X

X >X

>.

x:

-8SQ- •ui QZ+-

UJOOO5

cc

IlJ X

\- o

h*8

gcQ

CD

a:

LlJ [_ q:

LjJ iu 2 u

uj 2 o u* 5 O

(O 2

CO 2 *"*

O § g uj N |— £= a: a: H O UJ 7 x > ^ X > Q D 1

53 0P0

-P0e0 ,—. *

-P o g E 0 0 p

TO '— G0

-P00 0 0 0

X!

T3

O

G

p 0 a

>i >i 0 -p £ -H P O o o o 1—1 0>

4MO ». 0 5 0 0H a > d0 i u0

-p E d p o 0

fd

p

t—i 0 g 0 o Pd •H

-P Cn

fd

G

-P •H G £ 0 o e & d 0 p

-p r—1 m i—1 G 0 H £

CTt0pd

in "H

Pm
< UJ Ofr l 6S0 —

\*" 80 0 T 5 i ZzUl

£E £L <o

UiZ8>

t— < r—1

UJ tc. rtf (IX Q_ 5 X

ac Ul QJ < £ X o u. a) 1 o g / xx X -— o

r—\ X X

(IX CD

X 5 fo0)

4-1

x 0

XS 5 c

<D re •H> 02

CD 1 »—I a

4-1

D 3 O 0 o

>1 o

rfl E

r—( 2-4 CD C X o x XXc

a)gpxX

(IX

as5

CPc

•Ho C X M 0)

a)xo

tn *H

U-4

54
55

Figure

11a.

Instrumentation

layout

-

view

of

south

wall

of

the

3.3

x

3.3

m

test

room

showing

wall

thermocouples,

heat

flux

meter,

and

observation

port
2.41 m

f2.41

56

Figure

lib.

Instrumentation

layout -

view

of

south

wall

of

the

3.3x

4.9m

test

room
57

Figure

12a.

Instrumentation

layout

-

view

of

north

wall

of

the

3.3

x

3.3

m

test

room

showing

wall

thermocouples
58

Figure

12b.

Instrumentation

layout

-

view

of

north

wall

of

the

3.3

x

4.9

m

test

room
TIME

*««

SEC

600

1200

1800

2400

3000

3600

o

( 4 a

8

o

o

CD 8 S §

r* to us 8 8 8 8 ° m cm —

0 030 » 3«niVH3dK31 SVO

59

Figure

13a.

Average

gas

temperatures

as

a

function

of

time

for

tests

1,

2,

and

4

(temperatures

above

800°C

are

not

reliable)
TIME

SEC

600

1200

1800

2400

3000

3600

o

o

c* oO

o

o

o

a

o

o

a a

o a

o

o

o

o

o 07 00 r* (0 10 «n 0 030 •• 3«niVM3dh31 svo

8

CM 8

60

Figure

13b.

Average

gas

temperatures

as

a

function

of

time

for

tests

5,

6,

and

7

(temperatures

above

800°C

are

not

reliable)
TIME

**•

SEC

600

1200

1800

2400

3000

3600

3 030 •• aanivkGdflu svo

61

Figure

13c.

Average

gas

temperatures

as

a

function

of

time

for

tests

8,

9,

and

10
TIME

««•

SEC

600

1200

1600

2400

3000

3600

o

o

O o

o

a

8

o

a

CD o

o

r* 8 8(C UJ 8 8

«n 8

04 8

0 030 *• 3«niVH3dW3i SV0

62

Figure

13d.

Average

gas

temperatures

as

a

function

of

time

for

tests

11,

11A,

and

12
TIME

***

SEC

600

1200

1800

2400

3000

3600 IIlIi 63

Figure

13e.

Average

gas

temperatures

asa

function

of

time

for

tests

13,

14,

15,

and

16
COo

10

<N

COQ)

54P

4-1 rd

54

CDa,g

<u 4->

COrdao

to

rd

S-t

rd

O

5-1 44O 44O GO C H 0

4-1 -H

rd 4-1 •H u

54

C

rd P > 44

(U rd g

•H CO E4 rd

QJ

54P

to •H

64

ventilation
600

1200

1800

2400

3000

3600 i 65

Figure

15.

Time

variation

of

total

heat

release

rate

asa

function

of

room

ventilation
BOO

1200

1800

2400

3000

3600

66

Figure

16.

Time

variation

of

average

gas

temperatures

as

a

function

of

fire

load

density
00001

o

o

oX

<

hLU

O

O

5

LUX

00 CO CJ

PIAI-10 ‘Q3SV313U 1V3H

67

Figure

17.

Relationship

between

heat

released

within

fire

room

during

active

burning

period

and

total

heat

content

of

combustible

materials
UJ2

F

0 030 •• 3«f\lVb3dh3i SVO

fd

68

Figure

18.

Time

variation

of

average

gas

temperatures

as

function

of

room

lining

materials
Z

2

1

•

Ui2

<=

M* ••• ILVy 3SV333M 1V3H

69

Figure

19.

A

comparison

of

heat

release

rates

with

various

interior

finish

materials
COo

ino

eno

ega Za!

Ulac o 030 3yniva3dn3i svo

in

(DUP

-P

fdU0ae

0) +Jm

fdtn

0)

a)

n

ch-h

fd

m

p

fd op

4-1O 44O co c

•H o

-P -H

fd -P •H O P c

fd P > 4-1 0 td e

-h cn Eh fd O

CM0pp

tn •H

Cm

70
600

1200

1800

2400

3000

3600

71

Figure

21.

Time

variation

of

heat

flux

incident

at

the

center

of

the

floor

asa

function

of

room

size
72

Figure

22.

Time

variation

of

average

gas

temperatures
0 030 •• 3dniva3dW3i svo

73

Figure

23.

The

range

of

variation

in

average

gas

temperatures

for

residential

room

fires
TIME

SEC

600

1200

1800

2400

3000

3600

0 030 •• 3«niVH3dfl31 SVO

74

Figure

24.

A

comparison

of

results

obtained

from

ASTM

E

119

and

fast

response

thermocouples

for

average

gas

temperatures
TIME

SEC

600

1200

1800

2400

3000

3600

CM»o

J 030 3aniVM3dH3i svo

CM

CM

en

CO

o

in

z

a °

!

m •

Uia

P

CM

a

o >

>-l u m g

o V

75

Figure

25.

Comparison

of

fire

exposure

curve

derive^

room

fire

test

9

and

ASTM

E

119

standard
TIME

•••

SEC

600 1

ZOG

1800

2400

3000

3600 3 030 •• 3aniVM3dh3i svo

TS0upw

to0e0) to

CO0uP

-u

to

5-1 0

0460

-P

cn

(0CP0

Cp

(0

5-1 0>

fd

M-loco0H

5-i fOOU

r\i0

5-i P

CP •H

76

by

different

types

of

thermocouples
TIME

SEC

600

1200

1800

2400

3000

3600

0 030 •• 3MfUVH3dh3i SVO

I

I 77

Figure

27.

A

comparison

of

gas

temperature

averages

calculated

using

27

and

3

locations
05 05

o CM CM

IN 0NHI30 INOiH 33NV1SIQ

78

Figure

28.

The

influence

of

the

combustibility

of

the

ceiling

materials

on

the

vertical

gas

temperature

profiles
TIME

»**

SEC

600

1200

1600

2400

3000

3600

79

Figure

29a.

Comparison

of

the

calculated

and

measured

total

heat

flux

incident

on

the

gypsum

board

ceiling
TIME

**•

SEC

600

1200

1800

2400

3000

3800

xmj 1V3M

80

Figure

29b.

Comparison

of

the

calculated

and

measured

total

heat

flux

incident

on

the

gypsum

board

walls
Photo 1. Spread of fire over sofa

cushions 3 min after ignition in

test 12

Photo 2. Full involvement of room

contents 4 min after ignition in

test 12

81
Photo 3. Active burning stage

8 min after ignition in test 12

82
APPENDIX A DERIVATION OF THE EQUATIONS USED FOR THE

HEAT RELEASE RATE CALCULATIONS

Consider a -burning compartment where the combustible materials being

heated by the fire, the hot gas layer and the compartment walls degrade into

volatile decomposition products and carbonaceous residue or char. The

volatile components and the char react with the air surrounding the materials

,

and the mixture of combustion products and depleted air subsequently moves

upward to fill the upper space. The heat is liberated during the combustion

of the volatiles in the gas-phase and char oxidation on the solid surfaces.

The fire induces an inflow of cold air through the lower part and an outflow

of combustion gases via the upper part of the ventilation opening.

In order to derive the relevant equations for calculating the rate of

heat release in the compartment, the following simplifying assumptions were

made:

1. The fire compartment can be divided into two separate regions

with a layer of hot gas above and a cold air layer below. In

each region the gases are perfectly mixed with no variations

in their composition and temperature.

2. Following the pyrolysis of the combustible materials, the fuel

volatiles and the char react instantaneously with the surrounding

air

.

The continuity equations describing the conservation of the total mass

and the gas species around the upper region of the compartment can be written

as

d K) .

dt

d (vp X . /W ) \ g jg/ g/

= F W + F_W_ - F W

a a f f g g

= F X . -FX. - v ,r cV' a ja g ]g j f

(A-l)

(A- 2)

where V is the volume occupied by the combustion gases or the hot gases in

the upper part of the compartment; p is the density of combustion gases;

t is time; F^ and are the molar flow rate and the average molecular

weight of the gas in stream i, respectively; the subscripts a, f and g

denote the air, fuel and combustion gas streams, respectively; X^ is the

concentration of the gas species j in the stream i expressed in mole fraction;

the gas species include oxygen, fuel, nitrogen, CO

2

, CO and H

2 0; r^ is the

A-l
reaction rate or the rate of disappearance of the fuel in moles per unit volume

* of the combustion gas, and is the ratio of the stoichiometric coefficient

of the gas species j to that of the fuel.

These mass balance equations contain terms to account for the change

in the mass of the gases in the compartment, the flow in of the cold air and

the fuel, the flow out of the effluent gas, and the consumption of the fuel

by combustion. In general, the rate of change of the total mass of the gases

in the compartment is negligible compared with the inlet and outlet mass flow

rates

.

An expression for the rate of oxygen consumption in the compartment

can be obtained from equation (A-2) by neglecting the rate of mass accumulation term. Thus,

F X /F - X_

a 0

2 a g C^g (A- 3

)

Using this equation, the rate of heat release in the fire compartment

is expressed by

Q = r

f

V (-c)

= (AH /v. F X. F - X_ F

C

l °2 / \ a °2/ g °2 g / g " K ' FaX

°2 a/

F

g

- X0

2 g) ( 238 - 15 /Tg) V

g

( A-4

)

where AH is the heat of combustion of the fuel, in kJ/g-mole; K = AH / ( v v)

,

c c U

2 the heat of combustion of the fuel per unit volume of oxygen consumed at the

standard state of 25 °C and 1 atm; v is the molar volume of oxygen at 25 °C

and 1 atm, e.g., v is equal to 22.414 x 10 ^ x (298.15/273.15) or 0.02446

3 m /g-mole for an ideal gas; T^ and are the absolute temperature and the

volumetric flow rate of the combustion gases leaving the compartment,

respectively

.

With the assumption that both the combustion gas and the air behave as perfect gases, equation (A-4) can be rewritten as

Q = K X_ V T / V T - X 00 a a g/ \ g a) C^gj ^298 . 15/Tg^ (A- 5

)

where V and T are the volumetric flow rate and the absolute temperature,

cl d

respectively, of the air flowing into the compartment.

A-2
The fuel is assumed to react irreversibly with the air to form carbon

dioxide, carbon monoxide and water according to the reaction,

v f [Fuel] + v 0

^

1

0

2 + <3N 2

] = V C0; C0

2 + v^CO + v^H.,0 + dv^Nj (A-6)

where v,, v__ , v.n , and are the stoichiometric coefficients of the

il U

2 u id 2 U

fuel, C0

2

, CO and H

2 0, respectively, and d is the nitrogen/oxygen mole ratio

of air.

The. stoichiometric combustion equation for a fuel of composition

C^H^O with air can be expressed in the following form:

C

a

Hb°c +

(a + b'T2

-*

) [°2 + 3 - 76 N

2

= (a-x)C0

2 + H2 0 + xCO + 3.76 ^a +

b~ 2 ^~ 2x

. j

^ (A-7)

For the overall reaction, the molar rate of consumption or production

of a gas species can be related to its stoichiometric coefficients. Thus

r

f = ^2 = _^2

f 0

2 co

2

CO

CO

^2 °'

u h

2 o

(A-8

)

Using equation (A-2) an overall mole balance of all of the component

gases in the upper hot gas layer can be written as

F (X_ + X. + X„_ + X_,_ + X„

g \ Fg 0

2 g CC> 2 g COg H

= F. + F + r,V (Av*) ( A-9

)

r a r

where Av = Av/v^, the change in the mole numbers per mole of the fuel

reacted, and Av = + v^ + v CQ - - v f.

During the fire growth period, the rate of burning is generally

controlled by the fuel supply rate, and all of the fuel volatiles and char

produced convert completely into the combustion products. For fires in the

fuel limited burning regime the molar flow rate of combustion gas venting

through the opening can be obtained by simplifying equation (A-9) based on

the mole fraction constraint for the effluent stream and r^V = F^ . Thus,

F = F + (Av* + 1)F- (A-10)

cf a -L

,°g + x

N

2 g

A3
For post-f lashover fires, the burning rate is limited by the compartment

ventilation or the rate of air inflow. Since there is no oxygen present in

the combustion gas, its molar flow rate can be derived in a similar manner

with the exception that r^V = aF^, where a is the fraction of the fuel

reacted

.

F = F + (a Av* + 1) F_

g a f (A-ll)

Combining equations (A-3) and (A-10) to solve for F /F and substituting

a g

the resulting expression into equation (A-4) gives

Q = K/X

°2 a

*

(A v + 298.15/T V

g

(A-12)

For the fires under ventilation controlled burning, another expression

for the heat release rate can be obtained from equations (A-3), (A-4) and

(A-ll) as

Q = K/X1 - X_ 0-a 0-g (298.15/T

g) V

g

(A-13

)

The analysis of the combustion gas leaving the compartment is usually

made on a dry basis where the water vapor in the sampling line is condensed

and removed prior to the determination of the gas composition . The moisturefree combustion gas produced from the fires under fuel limited burning consists

of CC>

2

, CO, O

2 and ^ • The concentrations of these gases as measured with

the gas analyzers can be expressed by

X^L = F X„ /P

co

2 g co2 g

XC0

= F

g

XCOg/P

X“ = F X /P

°2 g °

2 g

where P = F [X__ + X__ + X_ \ + F X.. g^ C0

2 g COg °2 g/ a N

2 a

(A-14

)

These simultaneous equations are solved for the concentrations of CO

2

, CO and O

2 in the combustion gas stream. Thus,

XC0

2 g =

(

F

aXN

2 a/F

g)(

XC0

2 / XS2 )

XCOg =

(

F

a

X

N

2 a/F

g)(

X

C0

/XS

2 )

X

0

2 g =(F

a

X

N

2 a/F

g)(

XS

2

/<

2 )

A-4
v. VA _ VA VA VA where x

n

2 1 x o

2 xco2 xco

The concentration of water vapor contained in the combustion gases can

be estimated based on the ratio of the moles of H

2 0 produced to the moles

of' CC>

2 and CO formed from the burning of the fuel as given in the general

combustion equation. Thus,

H

2 0g Vr2 / \

C°2

+ VC°)

XC0

2 g

+ XCOg

v h

2 o/ (v co2 + V C0,

XC0

2 + XCo)[F

a X

N

2 a/(F

aX

N

2/J

(A-16

)

The mole fraction of the nitrogen, which lumps all of the components

of air except oxygen and is considered as an inert species throughout the

fire process, can be expressed by

X.T = F /F N„g a N a/ g

(A- 17

)

Since the summation of all of the mole fractions of 0

2

, C0

2

, CO, H

2 0

and N

2 gases in the combustion gas stream must be equal to unity, this

results in

F X. T a N

2 a

j 1/(1 + Y) (Vn2 )

=

where Y = + ^

(A-18

)

Substituting this relation into equations (A-15) to (A-17) yields

XC0

2 g “ XC0

2 j^ + Y '

X

COg

“ XC0 /

(1 + Y)

X0

2 g

= X

0

2 /

(1 + Y)

xM = xf; /(i + y) N

2 g n

2//

v

X

H

2 0g

= Y/ (1 + Y ) (A-19)

The average molecular weight of the combustion gases can be determined

from the equation,

W = 32X + 44X + 28X + 18X„ . + 28XM (A-20)

g 0

2 g C0

2 g COg H

2 0g N2 g

A5
For the fires under ventilation controlled burning, the combustion gas

consists of CC>

2

, CO, H

2 0, N

2 and the unburned fuel. The molar concentrations

of these gases can be derived in a similar manner to those for fuel controlled

fires. Thus,

XC0

2 g

X

C°2 /

(1 + Y

' )

XCOg XC0 /

(1 + Y,)

x

f g

= 4 n1 + Y '»

,B

N

2 g n2

H

2 0g

= XN^(1 + Y ' )

= Y'/ (1 + Y'

(A-21)

,B . where is the concentration of the moisture-free gas j in the combustion

gas stream during the ventilation limited burning regime;

Y' = v H„0

^

XC0

2 + XC0^/(^V C02 + V C0^

-j VB _ VB „B B and xn

2 “

1 XC0

2 “ xco

" x

f

The average molecular weight of the combustion gas can be estimated from

the following equation:

W = 44X„_

g C00 g + 28X_, + 18X _ + W-X^. COg H

2 0g f fg + 28X

N

2 g (A-22

The combination of equation (A-l), in which the mass accumulation term

is neglected, and equation (A-ll) gives the ratios of the molar flow rates

of the air inflow or fuel supply to that of the exhausting combustion , gas

Thus

F /F =

a g

* / W (a Av + 1) - g f /

W (aAv +1) - Wr

a f

VF

g =

(

W

a

- W

g)

:

W ( a A v + 1 ) - a f

(A-23)

(A-24

An expression for the fraction of the fuel converted can be obtained by

the rearrangement of equation (A-23). Thus,

W£ -W + F ( W - W,

f g a \ a f

Av W - F W /F ' g a a/ g(A-25)

where F /F = V T / (V T ) and V is the volumetric flow rate of inflowing air.

cl CJ 3. CJ CJ cl cl

A-6
The total rate of heat release within the fire compartment can be

calculated by substituting these derived and measured quantities into either

equation (A-5) or equations (A-12) and (A-13)

.

.

Al I’UNDIX B ESTIMATION OF THE INCIDENT HEAT FLUX

The temperature distribution in a semi-infinite slab initially at a uniform temperature and subjected to a constant heat flux on the exposed

surface is given by [14]

2 cr" T( X ,t) = qa k

(B-l)

where q" is the net heat flux entering the slab at its exposed surface; k

ci and a are the thermal conductivity and thermal diffusivity, respectively,

of the slab; t is the time; x is the distance into the slab from the exposed

surface and T^ is its initial temperature.

From Duhamel ' s theorem [15], the unsteady temperature at the surface of

a slab exposed to a time varying heat flux, q" (t) , is given by

cl

<°' t > = I V? /

^

(T) dT + a

0 -^t-T

(B-2

)

The integral in the above equation can be evaluated numerically. Thus,

the surface temperature at any time can be calculated from the following

equation, with a constant duration 0 for each time step.

T ( o , t) = JL. v

^~

£

qa

(n)

£

Vt_ ( n_1 ) 0 “Vt-n0 + T

±

s/tT

k n=l

(B-3

)

where N = t/0.

An expression for the net heat flux into the slab can be obtained by

rearranging equation (B-3) to give

q" (N) = 0 1/2

a. T ( 0 , t

) -T

.

s 1

N-l

E

n=l

q"(n) d ^[N- (n-l)] 0- (N-n) 0 (B-4 )

where C

The time variation of the net heat flux absorbed by the walls and ceiling

of the recreation room through their exposed surfaces was computed using the

measured surface temperature history data. The heat fluxes incident at the

room walls and ceiling, which were assumed to be semi-infinite solids with

temperature-independent thermal properties, were estimated from the calculated

net heat flux and the experimentally determined surface temperatures using

the following expressions for the net flux entering into the material, q",

d

B-l
(B— 5

)

and the incident heat flux acting on the exposed surface, q V .

b" = e (H-aT 4

) + h (T - T ) ^a s s g s

= e H + h (T - T ) (B-6)

s g s

where e g is total emissivity of the surface of the slab, H is the incident

radiative heat flux, h is the convective heat transfer coefficient, a is

the Stefan-Boltzmann constant, and T is the bulk temperature of the sur- rounding hot gas. It is assumed that the total absorptivity of the surface

for the incident radiation is equal to its total emissivity at the temperature,

T , as it would be for a "grey" body.

B-2
IREV. 2-»0) U.S. dept, of comm. BIBLIOGRAPHIC DATA

SHEET (See'/nstruct/ons)

1. PUBLICATION OR

REPORT NO. NBSIR 80-2120

2. Performing Organ. Report No. S. Publication bate

October 1980

4. TITLE AND SUBTITLE

FIRE DEVELOPMENT IN RESIDENTIAL BASEMENT ROOMS

5. AUTHOR(S)

J. B. Fang and J. N. Breese

6. PERFORMING ORGANIZATION (If joint or other than NBS. see instructions) 7. Contract/Grant No. NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE 1. Type of Report & Period Covered

WASHINGTON, D.C. 20234 Interim

^F^fyllZATION NAME AND COMPLETE ADDRESS (Street, City, State, ZIP)

Division of Energy, Building Technology and Standards

Office of Policy Development and Research

U.S. Department of Housing and Urban Development

.Washington, D.C. 2D4.1.Q

10.

SUPPLEMENTARY NOTES

| | 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

bi bl iography or literature survey, mention it here) A multi-phase study program has been established

to develop a rational test procedure for evaluating

the fire resistance of residential floor assemblies.

The first phase of this research program was aimed at characterizing the severity of fires originating in residential rooms and developing a specified set of

fire exposure conditions applicable for fire resistance

testing of floor constructions.

A total of 16 burnout tests were conducted to investigate the fire behavior in typical residential

recreation rooms of single family houses. These fire

tests were usually run for one hour and were performed

in two instrumented test rooms, 3.3 x 3.3 x 2.4 m and

3.3 x 4.9 x 2.4 m in width, length, and height respectively, furnished with household furniture and lined

with interior finish materials typical of actual

occupancies. Measurements were made cf the temperature,

/ heat flux, static pressure, smoke density, gas velocity.

12.

KEY WORDS (Six to twelve entries; alphabetical order: cabitalize only proper names; and separate key words by semicolons)

Building fires; fire resistance; fire

tests; flow measurement; gas temperatures; heat release

rate; interior finishes; residential buildings; room

fires.

13. AVAILABILITY 14. NO. OF

Unlimited

PRINTED PAGES

1 | For Official Distribution. Do Not Release to NTIS

n Order From Superintendent of Documents, U.S. Government Printing Office, Washington, D.C. 97

20402. 15. Price

£X] Order From National Technical Information Service (NTIS), Springfield, VA. 22161 $ 8.00

USCOMM-DC 8043-P80

:

species concentration, and oxygen consumption. The effects of such parameters as the ventilation, fire load density, initial item ignited, room size, and thermal and flammable properties of the wall and ceiling

materials on the fire severity were evaluated quantitatively. A fire exposure temperature- time curve which

is different from the ASTM E 119 curve, has been developed

for testing the fire resistance of such building structures.
*

• , .