NBSIR 77-1318

The Impact of A Room Fire on

A Corridor With Considerations

of Fuel Load, Ventilation and

Scaling

J. Quintiere, B. McCaffrey, T. Kashiwagi, K. Den Braven,

M. Harkleroad, J. Raines and W. Rinkinen

Center for Fire Research

Institute for Applied Research

National Bureau of Standards

Washington, D.C. 20234

November 1977

Final Report

U.S. DEPARTMENT OF COMMERCE

NATIONAL BUREAU OF STANDARDS

NBSIR 77-1318

THE IMPACT OF A ROOM FIRE ON

A CORRIDOR WITH CONSIDERATIONS

OF FUEL LOAD, VENTILATION AND

SCALING

J. Quintiere, B. McCaffrey, T. Kashiwagi, K. Den Braven

M. Harkleroad, J. Raines and W. Rinkinen

Center for Fire Research

Institute for Applied Research

National Bureau of Standards

Washington, D.C. 20234

November 1977

Final Report

U.S. DEPARTMENT OF COMMERCE, Juanita M. Kreps, Secretary

Dr. Sidney Harman, Under Secretary

Jordan J. Baruch, Assistant Secretary for Science and Technology

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Acting Director

PREFACE

This work was begun in the summer of 1974. It was motivated by

phenomena observed and recorded during previous tests of fire growth over

floor coverings in building corridors. The fire scenario in those tests was a room fire spreading into a corridor. The potential for fire spread

into the corridor would depend on the fire induced heating load in the

corridor. Also the resultant flow pattern in the corridor could influence

fire spread and smoke movement. The effect of room fire size, ventilation,

geometry and other variables on the movement of smoke and energy from a room fire to a corridor would broaden the understanding of this fire

scenario

To accomplish this objective a series of full-scale experiments

were conducted with the available room and corridor fire test facility

at the former site of the National Bureau of Standards in Washington, D.C.

(just before this site was abandoned). The scope of these experiments

was to investigate the effect of the room fire, natural ventilation effects

in the fire room, and the effect of an "obstruction" (floor fire or physical barrier) in the room doorway. Some of these data were used to

provide a basis for comparison with corresponding results derived from a scale model (l/7th geometric scale) room and corridor facility. In this

manner, the accuracy and limitations of scaling criteria could be evaluated, and the credibility of extrapolating results from scale model

experiments could be assessed.

This report has been prepared to document and make available the

results from the full-scale experiments and to present a complete set of

results for the scaling experiments.

iii

CONTENTS

Page

PREFACE iii

Abstract vi

INTRODUCTION vii

A SCALING STUDY OF A CORRIDOR SUBJECT TO A ROOM FIRE

(ASME Preprint) 1

COMPLETE SET OF SCALING RESULTS 14

DESCRIPTION OF THE FULL-SCALE EXPERIMENTS 46

CONCLUSIONS 49

v
Abstract

A study was conducted of heat transfer and temperature

field imparted to a corridor by a room fire in order to assess

the potential of fire spread along the corridor. Wood cribs were used as the fuel load which ranged from 20 to 120 kg.

Also the effects of ventilation and fuel load location were

examined. The results showed these effects to be significant.

A corresponding scale model study was conducted using

gas burners as a fuel supply. A scaling criteria was developed which emphasized the solid wall conductive and

gas phase convective transport processes. As a result

limitations were encountered at high temperature when radiation became more significant. In general the convective

processes appeared to scale well and radiation limitations

could be assessed through analysis.

Key words: corridors; fuel load; heat transfer; room fires;

scaling; ventilation.

vi
INTRODUCTION

This report will be based primarily on a presentation given at

the ASME/AIChE 17th National Heat Transfer conference. A preprint of

that presentation forms the main body of this text. This will be followed

by a complete set of curves for all the scaling experiments. These show

the comparison over time between the full-scale and model results. The

results supplement those given in the preprint.

In the last section of this report, a description of the full-scale

experiments will be given along with some general discussion. These data

have not been analyzed extensively. The complete set of data is on the

enclosed microfiche.

A SCALING STUDY OF A CORRIDOR SUBJECT TO A ROOM FIRE

(ASME Paper No. 77^HT-72)

(preprint follows)

vii

ited in 'll S A.

an ASME

publication

$3.00 PER COPY

$1.50 TO ASME MEMBERS

The Society shall not be responsible for statements or opinions advanced in papers or in discussion at meetings of the Society or of its Divisions or Sections, or printed in its publications. Discussion is printed only if the paper is published in an ASME journal or Proceedings

Released for general publication upon presentation.

Full credit should be given to ASME, the Technical Division, and the

author(s). A Scaling Study of a Corridor Subject to

a Room Fire

J. QUINTIERE

b. j. McCaffrey

Mechanical Engineers

Mems. ASME

T. KASH SWAGS

Materials Engineer

National Bureau of Standards,

Washington, D.C.

A study was made of the thermal and flow environment within a corridor subject

to a room fire of intensities of 300 to 1500 kW approximately. A corresponding model

study was done under l/7th geometric scale. Dimensionless groups were derived

from the conservation equations governing the gas and solid phases. A subset of dimensionless groups was identified as significant to establish Arteria to maintain

partial dynamic scaling between the model and prototype experiments. Good

results were achieved between the model and prototype for the convective process,

i.e., gas temperature, velocity, and convective heat transfer. Radiant heat transfer

did not scale, but an analysis of the data explains the lack of agreement in terms

of dimensionless groups that were not preserved in scaling. A secondary result yielded corridor convective heat transfer coefficients which could be correlated by a general relationship.

Contributed by the Heat Transfer Division of The American Society of Mechanical Engineers for presentation at the AIChE-ASME Heat Transfer Conference, Salt Lake City,

Utah, August 15-17, 1977. Manuscript received at ASME Headquarters April 20, 1977.

Copies will be available until May 1, 1978.

THt AMERICAN SOCIETY OF MECHANICAL ENGINEERS, UNITED ENGINEERING CENTER, 345 EAST 47th STREET, NEW YORK, N.Y. 10017

A Scaling Study of a Corridor Subject to

a Room Fire

J. QUINTIERE B. J. McCAFFREY T. KASHIWAGI

ABSTRACT k thermal conductivity of gas-phase

A study was made of the thermal and flow environment within a corridor subject to a room fire of intensities of 300 to 1500 kW approximately. A corresponding model study was done under l/7th geometric scale. Dimensionless groups were derived from the conservation equations governing the gas and solid phases. A

subset of dimensionless groups was identified as significant to establish criteria to maintain partial

dynamic scaling between the model and prototype experiments. Good results were achieved between the model

and prototype for the convective process, i.e., gas temperature, velocity, and convective heat transfer. Radiant heat transfer did not scale, but an analysis

of the data explains the lack of agreement in terms

of dimensionless groups that were not preserved in scaling. A secondary result yielded corridor convective heat transfer coefficients which could be correlated by a general relationship.

NOMENCLATURE

length

mean beam length of products

m mass flow rate

m, rate of burning

b

p pressure relative to initial static pressure

Poo C Pr Prandtl number,

k

q" rate of heat transfer per unit area

^ energy release per unit volume

Q heat flow rate

r stoichiometric mass air to fuel ratio

a

A

cP

c

s C

0

c

1 ,2 ,3

d

E

F

Gr

h

“d

H

c AH

absorption coefficient

area

specific heat of gas-phase

specific heat of solid-phase

flow coefficient

constants of the fluid

maximum doorway temperature ratio, T/T

energy flow rate

configuration factor

gravitational acceleration

p «

2g *

3 Grashof number,

2 P heat transfer coefficient

height of doorway

height of ceiling

net heat of combustion

flame height

P oo vJL Re Reynolds number, —-

t time

T temperature

u horizontal velocity component

v vertical velocity component

V velocity

W width

x horizontal coordinate

y vertical coordinate

Y concentration

a

^ absorbtivity of the gas products

<5 solid thickness

surface emissivity

< thermal diffusivity

P viscosity

dimensionless group

1
p density

a Stefan-Boltzmann constant

Subscripts

D room doorway

E corridor exit opening

FS full-scale

M model

r radiation

R reference

s solid

" initial

Superscripts

( )

' dimensionless

( ) average

INTRODUCTION

A common approach used to study the spread and growth of fire and its effect on the surroundings has been to examine a particular fire scenario by using a small-scale model experiment. The obvious advantages

of studying fire problems at a reduced size are less cost and ease of operation. However thq problem with

such an approach is the inability to maintain complete

dynamic similarity between the prototype full-scale

fire and the model. This would be evident by considering the paper of Williams (1) who lists no less than

28 dimensionless groups derived from the gas-phase

conservation equations alone. Thus novel techniques

and perceptive insight must be applied to fire problems, and the use of partial modeling must be employed.

Thus perfection is not likely achieved in the "art" of fire modeling. Nevertheless valuable results can be

achieved

Scale model studies have been applied to an assortment of fire problems. Gross and Robertson (2) and Thoirns (3) have used various size enclosures to study the rate of burning and other variables in fullydeveloped fires. These are generally fires which

"fill" the enclosure, with flames extended out of the openings. Heskescad has done similar studies as well

as modeling of sprinklered fires and the initial fire

environment ( 4 ,5 ) . The velocity and temperature

fields have been studied for large mass (10 5 m

2

) fires

by Lee (6) and the plume above a house fire by Yokoi

(7.) The technique of conducting model studies at elevated pressures (in order to preserve buoyancy,

inertia, and viscous effects) has been used by deRis

,

Kanury, and Yuen (8) for steady burning and by Alpert

(9) for modeling transient wood crib fires. Parker

and Lee ( 10 ) have used actual prototype room lining

material and furnishing items to investigate the fire

growth behavior in full and 1/4-scale rooms. All

of these studies have yielded generally good results,

but confidence in full-scale predictions from model

results remains uncertain.

The study reported here was motivated by a desire

to gain a greater understanding of factors influencing

corridor fire spread. Several years ago, experiments

were conducted to study fire spread over floor covering materials in building corridors ( 11 ) . The ignition source in this case was a room fire. These teste demonstrated that the fire produced conditions in the corridor had an ultimate effect on the nature of the fire spread over the floor covering material. In particular, radiant heat transfer to the floor was significant, but the significance of other factors,

such as corridor flow patterns, initial fire size, or geometry, were not fully appreciated.

In order to more expeditiously explore these factors a scale model study was initiated. To provide

a source of data for comparison with the model, a series of full-scale experiments were carried out to study the corridor flow and thermal conditions in terms of the room fire size. No floor covering flame spread was introduced in this work. These full-scale

results would then serve as a basis for validating

scale model results and thus could provide a measure

of the accuracy of model results and scaling criteria.

The scaling criteria were derived from the dominant

set of dimensionless groups generated by conservation

equations which best governed this problem. Ultimately

the results could be applied to the development of mathematical models ( 12 ) for the prediction of corridor heat transfer to a room fire under general geometric, material, and ventilation constraints. Fire

safety design analysis could then be applied to standards for selecting, for example, fLoor covering

materials in corridors.

DIMENSIONAL ANALYSIS

The dimensionless groups for scaling were developed from the governing differential equations.

Judgment was used in restricting the equations to include the significant thermal and fluid transport processes (e.g. viscous dissipation and pressure work

terms were neglected). The transport of multicomponent

gaseous species and the kinetics of combustion were

totally ignored in this formulation. The effect of combustion was included through a heat generation

term. This is justified since the field of interest

in this problem was essentially outside the combustion

zone. Also gas-phase radiation was neglected for simplicity in the gas-phase energy equation, but accounted for through a sur f ace-to-gas boundary condition for the solid-phase energy equation. All thermodynamic and transport properties were considered

constant except for density which follows an ideal gas

law. Since the scaling laws ultimately derived will

preserve temperature, and these properties are temperature dependent, the constant property assumption will

not affect the derived dimensionless groups. With

these restrictions the governing gas-phase dimensionless conservation equations follow. Without a loss in generality only the two-dimensional equations are presented

Gas-Phase Equations

Mass

(

& \ sp

. a (p 'u ') WJ 3t ' 3x

' 3 (p 'v') 3y' ( 1 )

2
Momentum the fire problem addresssed. Consequently from Eq . (3)

Vr

9 u'

9 t'

'-V

9 u'

9 x"

2

9 u"

9 x" 2

- 9 u"

7 s7 9 y

,2 „

.2

E -y/g* ( 8 )

where S. is the geometric length scale (for example, the height of the corridor) , and g is the gravitational

acceleration.

= T/T (9)

_1 d /3u' , 9 v"

\

3 3x' \3x' 3y'/

/ H \ 3y"

wJ 3t

'

, 9 v , -3v + U T—7- + V -7—

7

3x 3y

v)

(‘ ‘ °’) + fc

' 2

3 v'

,1

33y

( 2 )

where is the initial temperature.

p' = p/p vD

2

( 10 )

where p is the increase in pressure relative to the local initial hydrostatic pressure. An expression for the energy release rate per

unit volume, q,Vol ’ depends on the volume of the com- bustion zone. One representation is given as

“b

(t)AH

4Vol (11a)

Energy

, _1 3 / 3u^

3 3y L , - \ 3 x 3 y / _

(

! \ ar . . 3T' fe)

^ + " 77

+ v 3y'

where m^t) is a known time dependent fuel production

rate and AH is the net heat of combustion. This repre-

(3) sentation is based on assuming complete combustion which

occurs within a volume proportional to £

3

. This may be

reasonable for the application here since most of the visible flame zone is contained within the room, and

for large fires flames tend to fill at least the upper volume of the room. A more appropriate volume may be the flame volume

itself. For flame heights less than that of the room

height, this might be approximated by a cylinder of flame height, , with a base area equal to the initial

region where the fuel is introduced. Since turbulent

flame height can be represented as ( 13 )

(

P »

CPv) Cx'2 +

3y' 2 ) + (p-CpV.) "V01

= 16

Tl/5

(r4m)

p g(l-id)

. 2/5 “b

State

(A) « 16

2 . 2

2 P„ 8

Vs

T' = 1 for enclosures with openings. (5)

The dimensionless variables have been defined as follows

:

P' = p/p,

where p^ is the initial density.

- u , v

u = — and v = —

R R

( 6 )

(7)

where V

R

is a reference velocity that is determined by

the observation that buoyancy should be of equal order

of magnitude with the steady state inertia effect for

where m =

1 +

AH

rc T P-°

AH

rc T

p

»

~ 0 ( 0 . 1 )

this expression could be used to express a characteristic combustion volume as 2

2H^. Hence an alternate

expression to Eq. (11a) is

4Vol

3/5 , 3/5 (n^AH) ' (AH/-' 1 2 vV 5

16 V -m) (lib)

3
Time is normalized with a reference time t that

K should be characteristic of the fire behavior that is being studied. In the present study a characteristic

time will be associated with the duration of the fire

This was approximately 30 minutes. Hence

t D = burnout time ~ 30 minutes

R

( 12 )

The dimensionless groups that then result from

these equations are given below.

ir _ 1_ M ~ gas residence time (13)

1 t Y

8 burnout time

TT 2

TT 3

1/2 3/2

P mg 1

viscous force inertia force (14)

V»8

k

1/2j,3/2 thermal conduction

thermal convection (15)

TT 4 VH

P »

c pT„8

1/2 5/2

' i

combustion energy

thermal convection ,,, . (16a)

An alternate expression for tt^ from Eq . (lib) is

(rii^AH) (AH/r)

p Ji (=

pt.)

2

< 5/! (16b)

Thus scaling based on it over tt " differs by a factor

AH/ (rc T ) . 4 P

OO On examining the magnitude of these groups, it is found that for the full-scale corridor experiments of interest here

ifj ~ o(io

-3

),

*2 ~ "3 ~ 0(10" 6 ), \ ~ 0(10

_1

).

These estimates were based on l taken as the corridor

height. Consequently, it will be assumed that terms

in the equations can be neglected which have coefficients of 77 j ,

1t 2 > or T

3

• The physical significance

of these deletions follows: (1) The gas-phase is quasi-steady with burning rate contained in 77 ^ as the time-dependent "driving force." The time for fluid

to be transported through the corridor is considered

short compared with changes in the burning rate. Also

regular flow oscillations and turbulence are contained

in the unsteady terms, and thus are not considered

explicitly. (2) The viscous and thermal conduction

effects are ignored near solid boundaries. This will

have to be compensated for later when considering the transfer of heat at solid boundaries. It can be shown

that 77 2 an<5 ^3 may be redefined as

*

2 = 1/SGr (17)

and

77 3 = 1/Pr^Gr (18)

where Gr is a Grashof number and Pr is a Prandtl

number. In this scaling study the Gr will be greater

than 10 9 for both the model and full-scale system which

should insure turbulent flow. This would imply that turbulent mixing away from solid boundaries is independent of scale.

Solid-Phase Equations

In order to complete the derivation of the significant dimensionless groups, heat transfer at solid

boundaries needs to be considered. Also in the solidphase formulation, the radiative exchange with the layer of combustion products must be included. This

will be done by characterizing the gas at a single

temperature, T. For illustration, an inert solid of thickness 6 is considered heated by a hot_finite layer of combustion products at temperature T which is bounded below

by cold black surroundings at Tm

. This is representative of the corridor ceiling. The conduction equation

is given as

3T 8

2 T

s s 3t

< s 3x 2 s (19)

where is the solid thermal diffusivity. An internal

boundary condition has the form

3y _

x

s = 0: - ks = h (T ' V +<r

(20)

s where k is the thermal conductivity

s h is the convective heat transfer coefficient

and q" is the net radiant flux to the surface,

r The radiative flux can be represented as

= e (1 - a ) o(T

4 - T4

) s g

00 s + e a a

s g

—4 4 Or - t ) s

( 21 )

Here e is the surface emissivity and a is the

s g absorptivity of the gas layer which may Be represented

as

g

-at

1 - e g ( 22 )

where a is the absorption coefficient and is a characteristic mean beam length.

It is reasonable to expect that £~0(£). It should be noted that Eq . (21) does not contain geometric shape factors for surface to surface radiative change, and

in fact applies to radiative exchange between infinite

parallel plates. In general, the shape factors must

be regarded as additional dimensionless groups which

will automatically be satisfied if geometric similarity is maintained.

The convective heat transfer coefficient in

Eq. (20) must be a function of scale and therefore

must be represented by an appropriate relationship.

Since the nature of fire induced flow within an enclosure has been perceived as consisting of a cold floor

jet and hot ceiling jet, h is derived from a constant property turbulent boundary layer correlation for flow over a flat plate.

4
C l k

0 h = Re (23)

1 /3 where Cj is a constant for a given fluid, 0.036 Pr , and Re is the Reynolds number represented as

P«/g

/— . 3/2

Vo 2

' Re (24)

By combining Eqs. (19) through (24) and making terms dimensionless the following results:

3T

r _§

5 at'

3T'

x' = 0 : -

s ax'

a 2 T'

S 8x'2

s

-

1 (r-r>. v(l-.,)(l-r)

(25)

s

+ ^7^8 (T

' 4 -T

;4) - (26)

The dimensionless groups are

2

tt =

° ^

thermal penetration time . burnout time ^ ’ 5 K

s lR

IT =

6

0 . 2 , "

2 ^ convective heating

^ solid conductive heating

s

(28)

where C

2 is a fluid dependent parameter,

C k (—

)

1 u

0.8 0.4

e oT 6 , .. s ^

surface radiative heating

, solid conductive heating

a = ! _ e-(a2)(2 /2)

g

6

(29)

(30)

It should be pointed out that the above groups tt,. - tt^ hold for a solid boundary in which t > 6

2 /k . This

R s means that the thickness 6 is significant and the effects of heating are significantly felt throughout

the solid. Also the above groups apply to all gassolid interfaces since they all would have a boundary

condition similar to Eq . (20). However for thermally thick solids and for com- posite solids, an alternative set of dimensionless

groups apply. These are summarized below:

Thermally thick solid boundary, t << 6 /<

K S

, 2

TI /TT? (kPc)

'r ‘Vt. V

”6

0 2 C 2 2

•

e a T

s

a

( 31 )

Composite solid with layers A and B, t >_(&/<) for A and B

tt through tt for solid A

J O

IT S - /

TT

2 6 /k

( 32 )

10 (ksssA P c ) / (ksssB p c )

Composite solid with thick second layer, t << (6 /< ) R s B

t through tt for solid A

5 o (33)

10

Scaling Criteria

In order to achieve exact similarity between a model and prototype all the appropriate dimensionless

groups must be preserved in scaling. Indeed, the set

of groups presented here is only a partial set since

some restrictions have been made. Yet even this partial set is too numerous to achieve complete pre- servation. Thus a sub-set of the groups will be selected in order to design a model. It has already

been pointed out that for this application tTj, it

, and

TTj can be neglected. Also in designing a reduce!

size model it is impossible to preserve and tt^ similtaneously unless is manipulated in order to obtain

agreement between the model and the prototype. It would be difficult to control surface emissivity

appropriately especially since soot is likely to coat

the heated surfaces in these fire experiments. In fact e g was held the same by painting the walls and ceiling of the corridor with a high temperature resistant black paint. Thus, the rationale is to favor tt if convection is significant, and favor ir‘7 if radiation is important. Finally it is also impossible to accomplish similarity for flame or gas radiation, and

TTg will be ignored in designing a model.

In summary, the following groups will be preserved

in scaling (within the limits of material constraints) :

tit (t) AH

1/2 5/2 ’ P coC pT

oo g *

5
6

7

C

2

«.°* 2 6 e oT 6 s “

and Trg and ir 1Q if necessary. It is implicit that geo- metric similarity must also be preserved. As a result

of these scaling criteria, the dependent variables will

scale as follows:

(temperature) = Tp

S

(velocity) FS

(34)

(35)

(floor incident radiant flux)

= ^FS

(if *7 & (36)

at the same time and corresponding coordinates. Since

approximations introduced in arriving at this scaling

criteria have been based on judgment and physical

constraints, the results will indicate the accuracy

of relationships (34)-(36). The internal dimensions of the room-corridor

model were selected to be l/7th of full-scale. This

geometric scale was selected for its convenient

overall size relative to our laboratory space yet

large enough in height to give a Gr > 10 9 to insure

turbulent flow. The construction and lining materials

were selected according to the dimensionless groups

and available materials. The extent to which the dimensionless groups were preserved is shown in Table

1. Table 2 lists the materials and some of their

properties. It is apparent from Table 1 that perfect

agreement of the dimensionless groups could not be

achieved. This was due to the unlikely prospect of finding available materials to meet the exact model

requirements in size and properties. For the walls

and ceiling linings an attempt was made to match it 5 and tt : thus, ir_ was uncontrolled. The outer material

6 7

on the room walls and ceiling was not considered very important within the duration of these fires. Therefore if 9 would be irrelevant for the thermally thick outer layer, but 71 jq would be significant. For the floor, which was expected to be a cold thermal sink

for radiation, agreement among the dimensionless

groups was considered rather loosely. In fact, as opposed to the results in Table 1, the floor was considered to be a semi-infinite layer of asbestos

cement board (which has properties similar to brick)

Under this approximation the group 71 5/

77 7 is preserved.

EXPERIMENTAL DESIGN AND PROCEDURE

Full-Scale

In this series, nine experiments were conducted.

The effects of room fire fuel loading, location, and ventilation were examined in the series. Only the five experiments in which room fuel loading was varied

will be considered in this report. The room was 2.6 m x 2.4 m x 2.8 m high with a single opening, a doorway 2.0 m high x 0.79 m wide, leading to the corridor. The corridor was 2.44 m x 2.44 m x 9.14 m

long with a window at one end which was 0.66 m above

the floor and 1.7 m high x 1.0 m wide. The window

was oriented at a corner and was connected to the outside by a 1. m long duct. The instrumentation

consisted of 53 Chromel-Alumel 0.25 mm thermocouples

(T) , five total heat flux (H) and radiant heat flux

(R) sensors, eight pitot-static tubes (P) monitored

by variable reluctance pressure transducers, and five

electrolytic oxygen cells (0) . The overall layout of these probes is shown in Figure 1. The heat flux sensors were water-cooled thermopile sensing elements

coated with an 0.96 emissivity paint. The radiometers

had their sensing element shielded by a CaF2 window

scrubbed continously by small air jets. These heat

flux sensors were calibrated at the start and end of

the test series. A shift in calibration was generally

within 5% except for H5 (20%) which was cleaned and repainted after the first test and R5 was ignored

since its window broke . The radiometers were not

Table 1. - Dimensionless groups

2

6 c2 n°-

2 «

3 e o t o

TT S

°° - ( 4;

\

(kpc)

s, INNER

5 K

s K k

s ks

I OUTER LAYER

(kpC)

s, OUTER

Corridor Full-scale 0.85 3.3 0.13

Walls and Ceiling Model 1.1 4. 3 0.24 ~ —

Room Full-scale 0.93 9.2 0.35 —38.0 0.021 Walls and Ceiling Model 1.60 11.0 0.62 1.1 0.087

Floor Full-scale 0.046 0.21 0.0081 oo 2.0 Model 0.73 0.58 0.032 — —

* Notes: (1) t c = 0.5 hr, (2) e = 1, (3) T = 295 K.

K S

00

6
Table 2. - Lining materials and their properties

Thickness k

s

c

s

P K

S (kpc)s

CORRIDOR WALLS AND CEILING - cm W/ (m-K) kJ / (kg-K) kg/m 3 -6 2 . 10 m / s

2 4 2 kJ / (m -K -s)

Full-scale: Gypsum board 1.6 0.17 1.1 960 0.17 0.18

Model: Calcium silicate board 1.9 0.11 1.0 580 0.18 0.066

ROOM WALLS AND CEILING -

Inner Lining:

Full-scale: Spray coated insulation 2.5 0.099 0.84 300 0.39 0.025

Model: Alumina-silica block 2.5 0.056 1.0 240 0.22 0.014

Outer Lining: Full-scale: Concrete block 18 0.73 0.84 1900 0.46 1.2 Model: Plywood 1.3 0.12 2.5 540 0.085 0.16

FLOOR -

Inner Lining:

Full-scale: Asbestos-cement board 0.64 1.1 . 1.0 2100 0.49 2.3 Model: Asbestos-cement board 2.5 1.1 1.0 2100 0.49 2-3

Outer Lining: Full-scale: Brick 6.4 0.73 0.84 1900 0.46 1.2 Model: None — — —

1 — —

Fig. 1 Corridor geometry and instrumentation for full- scale experiments (distances in metres) T- thermocouples , R- radiometer, H- total heat

flux, P- pitot static tube, 0- oxygen con- centration.

generally useful since the CaF2 has a transmissivity

which falls off to zero at 12pm which leads to a 25% error or greater for source temperatures below 400°C. The oxygen cells and pressure transducers were calibrated before each run. The fuel burned was cribs of Hemlock made of 3.8 cm x 3.8 cm square sticks spaced 7.6 cm apart in 16 layers with 4 sticks per layer. Each crib weighed

approximately 19 kg and had a moisture content of 10

to 13%. The mass loss of the fuel during combustion

was monitored by means of a platform suspended from a strain-gage load cell transducer which was in a cooled container above the room. In the five experi- ments presented here, the cribs were centered on the platform which was approximated 15 cm above the floor. The fuel loading varied from one to five

cribs through the five tests. The cribs were ignited

by pans of heptane and generally burned to a less than 10% ash residue in 30 minutes.

The series of five experiments presented here were run over a two week period during evenings in the spring. The same gypsum board lining material

was used throughout the tests, however, an earlier

experiment damaged a section of the ceiling near the room doorway. This section was covered with a 2 m x

1 m x 1 cm sheet of asbestos mill board (930 kg/m^)

for added protection. It should be pointed out that this earlier test produced ceiling temperatures of 200°C to 500°C which should have been sufficient to remove the water of hydration (20% by mass) of the gypsum board ceiling. Thereafter only absorbed

moisture of up to 5% may have been present in the gypsum board. This moisture content is comparable to that of the asbestos cement board used in the model.

Model

The model was constructed of the materials des- cribed in Table 2 with internal dimensions equal to l/7th full-scale. In order to validate scaling, it was

felt that only selected data would need to be recorded.

Four 0.12 mm Chromel-Alumel thermocouples each recorded the vertical temperature distribution in the room

doorway, corridor, and exit window. Three additional

thermocouples monitored corridor ceiling temperature.

Flush mounted water-cooled total heat flux sensors measured incident flux at three positions along the

7
floor and one location at the ceiling. Two pitotstatic probes were used to derive the exit flow velocity at the window. All of these sensors were located

at positions corresponding to their counterpart in the full-scale experiment.

In both the model and full-scale, data were

recorded on a high-speed multi-channel analog-digital

acquisition system at a record rate of every 10 seconds

for 30 minutes.

The fire source in the model was four watercooled brass diffusion flame burners centered above

the floor of the room. They were sized and located

according to l/7th geometric scaling relative to the height and cross-sectional area of a wood crib. The burner height was 5.5 cm above the floor and had a coarse screen covered port diameter of 6 cm. The fuel was methane (CH4 ) metered through a high and low capacity rotometer. The criterion for establishing the correlating flow rate was given by Eq . (16a). That

is

,

m

CH4 ,M '"WOOD CRIB

where (AH) W^0D = 16.3 MJ/kg for Hemlock with 12% moisture, and (Ah) = 50.0 MJ/kg. The burning

L.ri4 rate was determined by taking a five-point derivative

of the crib mass loss data and averaging this over 100 second intervals. Then the CH^ flow rate was manually

controlled in the model experiments over corresponding

100 second intervals. The number of burners was equal

to the number of cribs burned in a given experiment

except four burners were used to simulate the five

crib case. From the relationship for free burning

turbulent flame height ( 13 ) it follows that

H

H

f ,FS

f ,M (38)

Thus, geometric scaling of flame heights will not be

achieved. This distortion would even be greater for flames that impinge and extend under a ceiling.

RESULTS

General Observations

More smoke appeared to be produced in the full- scale crib fires than in the model experiments. In fact for 3 or more cribs dark grey smoke obscured

vision along the corridor. Also fire brands from the wood cribs were generated in all fuel loadings and their trajectory followed the jet of combustion pro- ducts which issued from the room. The corresponding

flame lengths were longer in the model fires. In the

5 crib full-scale burns, flames scrubbed the corridor

wall opposite the room doorway. However, in the cor- responding model case, flames extended to thermocouple

T23. The pressure transducers used in the full-scale

experiments were later found to have had problems with

zero drift and did not yield consistently reliable

results. Despite this problem, it was deduced that

the mass flow rates were 0.75 + 0.25 kg/s for the doorway and 1.0 + 0.2 kg/s for the window; generally,

independent of fuel loading. The lowest oxygen concentration (dry gas basis) for the room fire products

occurred for 5 cribs and was recorded at 3 per cent (sensor 01). The average peak burning rates ranged

from 18 to 92 g/s, roughly 18 g/s per crib. None of these fires were ventilation (air) Limited.

Typical Scaling Comparisons

Selected channels of data were compared directly

for the model and full-scale results. The two sets of data were processed and plotted by a computer. As a representative example, the 3 crib fuel loading case will be presented. The "input function" for scaling

was the rate of burning for the cribs. The resultant

mass loss for 3 cribs is shown in Figure 2 along with

TIME | s

Fig. 2 Mass and mass loss rate for the 3 crib full- scale experiment

the rate function prescribed in the model. The erratic

rate behavior at the end of burning was probably caused

by cribs collapsing. Temperature comparisons are shown

in Figures 3 through 6 for the room doorway, corridor

vertical traverse at 1.5 m (5 ft.), the exit window,

and the ceiling, respectively. In general, the results were poorest for the ceiling temperature. As

a consequence, the radiation heat transfer to the floor, shown in Figure 7, is also lacking in agreement. The full-scale window velocity data was the most reliable and therefore that was compared to the model results. These data compare favorably (Fig.

8) except perhaps for a "drift" problem in full- scale V7. Also shown in Figure 8 is a ceiling heat transfer coefficient (HTC A) determined at 1.5 m. This was defined as

where q 4 is the measured flux by sensor HA, T is the

gas temperature measured by T23, and T is 22®°C, the temperature of the water-cooled sensor. The ceiling

flux data was not corrected for radiation, and con- sequently an overestimate in h of up to 30 per cent in full-scale could result. It follows from Eq . (23) and

8
CORRIDOR FIRE - 3 CRIBS

ROOM DClORUO i CORRIDOR FJRE - 3 CRIBS

EXIT WINDOW

o

Fig. 3 Room doorway temperatures for model and full- scale experiments (3 cribs).

Fig. 5 Exit window temperatures for model and full- scale experiments (3 cribs).

Fig. 4 Corridor vertical temperature distribution Fig. 6 Corridor ceiling temperatures for model and

for model and full-scale and experiments full-scale experiments (3 cribs)

.

(3 cribs)

.

9
CORRIDOR FIRE - 3 CRIB^

RODIPTION __ full. ecntE

1X7 8CQLE

Fig. 7 Corridor floor irradiance for model and full-scale experiments (3 cribs).

CORRIDOR FIRE - 3 CRIES

HERT TRANSFER COEFFICIENT

VELOCITY

Fig. 8 Corridor ceiling heat transfer coefficient and exit window velocities for model and full- scale experiments (3 cribs). Note: h (HTC 4) for model was scaled Jjo 2 full-scale by multiplicative factor of 7

' .

(24) that the scaling factor for h was Jt

' which was applied to the model values for h in Figure 8. Overall results for the fire scaling experiments

are compared in Figures 9 through 11. Figure 9 displays corridor vertical temperature distributions

tures for various room fire sizes and for model

and full-scale experiments.

Fig. 10 Comparison of full-scale and scaled model

ceiling heat transfer coefficients as a function of full-scale burning rate.
Fig. 11 Comparison of full-scale and model floor irradiance at maximum steady burning.

where Che configuration factor was ignored and e was taken as 1. These two equations can be solved for q"

ir terms of T' for values of ir & and n given in Table

1 , and^fcaken as 0.11 for the model an^ 0.25 for full- scale based on no smoke and characteristic CO 2 and HjO concentrations of 10 per cent each and a beam

length of 1 m. Since smoke was more apparent in fullscale, Tig was estimated as 0.50 to account for smoke.

The correction factor or correlating curve was then determined by solving for

(q ) FULL-SCALE

^ MODEL

constant (42)

where T' was based on T23 (from Figure 9). This

factor was applied to the model data of H2 and H3 to yield the curves in Figure 11. In analyzing the ceiling heat transfer coefficient

results, it appeared that the turbulent flat plate cor- relation was valid, and that room or doorway tem- perature affected the results. It is believed that

the maximum doorway jet velocity controls the convective corridor ceiling heating at least near the doorway (perhaps for distances within a door height, H^) . This velocity can be ideally expressed as ( 14 )

for the T23 thermocouple traverse during peak burning

rates. The higher gas temperatures in the model for the 5 crib case were the result of flames extending to those thermocouples. This was a distortion in flame length between model and full-scale. Figure 10 displays the convective heat transfer coefficient scaling

results in terms of the peak crib burning rates. The

bars on the full-scale results indicate the range during maximum "steady" burning conditions. Finally,

Figure 11 compares the incident floor radiative heat flux for the model and full-scale at average peak

burning as derived from the nearly flat portion of the flux curves in Figure 7. As shown, the model flux values are generally lower than the full-scale corres- ponding values. The reason for this can be traced back to the lack of preservation in the radiation

groups JTy and ir.. This can be quantitatively demonstrated by the following analysis.

ANALYSIS

Better agreement may be achieved for the floor

flux scaling results, or their differences may be explained by applying a correction factor to the data which accounts for differences in rr 7 and it (as well as

it.). This was done by assuming no conduction loss to the ceiling during peak burning, and considering a ceiling heat balance from Eq . (26)

0 = tt 6 (T' - T') + tt 7 (1 - ir 8 )(l - T

' 4

)

+ it 7

it g (T

' 4 - T

' 4 ). (40)

The dimensionless radiative flux to the floor can be expressed as

1 -d d(l+d

l/3

)

1/2

where d is the temperature ratio T/T ,

and

«D

C

0

is the door height,

is a flow coefficient (0.7) .

(43)

Substituting this expression for V into Eq . (23) and taking l = H^, it follows that

Q 4

h = C

3 C

p (4gVc 0

4 9) • Hp * 1 -d d(l+d 1/3

)

(44)

where C 3 is a constant for the fluid. Taking T as doorway temperature T1 and evaluating fluid properties

at normal room temperature the results of this corre- lation are presented in Figure 12. The data were

derived from quasi-steady periods during the scaling

experiments. Also plotted there are results for other model tests which had various width exit door- ways instead of the window used in the scaling series.

Finally an approximate energy balance was done

for the full-scale 3 crib fire condition. This analysis

assumed a nominal air flow of 1 kg/s, and considered

the average peak burning period only. A sketch of the analysis follows:

E (by mass loss)

L J COMBUSTION

rflb AHWood

= (55 g/s) (16 kJ/g) = 880 kW

4 e a?

s

‘

<!-*«> t ;

4 + IT. — 4 T' (41)

11
Q RADIATION FROM HOT LAYER TO FLOOR

Fig. 12 A correlation for ceiling heat transfer

coefficient

M

H

COMBUSTION (by °2 de P letlon ) ~ m

a irAY0

2 AH

0

2

= (1 kg/s)(0.21-0.14) (l3240 = 927 kW

kJ DOORWAY ENTHALPY

AairVT = (1 kg/8)(1 - 05 ki=K>

(520 K) = 520 kW

DOORWAY RADIATION

oT

4 A = (5.7 x 10_U kW

n2 -K4

(810 K)

4 (1 m

2

) = 25 kW

CORRIDOR CONVECTION hA(T T

s

) TO CEILING & WALLS

0.045 ) (50 m

2

) (370-315) m -K /

= 124 kW

- -

![" —b __ 1

Q CORRIDOR RADIATION

= (1-a )aT - a a(T-T ) g s g s' A

_ _ FROM CEILING &

L _l WALLS TO GAS

= 219 kW for a =0.25

g

= 97 kW for a = 0.5

g

WINDOW

ENTHALPY

= i £ q" . A.

F i=l ^ 1

/ 1 \/2.4 x 9.1 \0T5/y 4 m2J

kW . 7+5+2. 2+1. 8 -^ = 175 kW

m

2>

m . c AT = (1 kg/s) (l.05 +-+

air p V kg-K

(250 K) = 250 kW

Thus about two-thirds of the energy released in the room reaches the corridor, and more than half of that energy

is lost in the corridor. Admittedly these results are very approximate, and only serve to give an order of magnitude of the flow of energy in the corridor.

DISCUSSION AND CONCLUSIONS

The results should illustrate the limitations of modeling due to the inability to preserve all the significant dimensionless groups. This handicap will

always be present in fire problems due to the con- straint of the physical and chemical processes as well

as the constraint imposed by material properties. These difficulties will likely increase with the complexity of the problem studied.

Apparent in this study were several troublesome

points. There was an ambiguity in selecting the burning rate in the model since or -n^' could have

been used. Moreover the effective heat of combustion

of solid fuels burned in fire experiments may not be equal to their theoretical values due to incomplete

combustion. Also the radiation dimensionless groups

tt 7 and Tt g had to be ignored which resulted in unfavorable, but consistent, results for ceiling temperature

and floor irradiance. However, by analysis, it was demonstrated that correlation of floor irradiance

could be achieved. Distortion of flame lengths were anticipated and contributed to good agreement for floor heat flux H2 in the five crib case. This is one example of "good" agreement achieved by compensating effects associated with lack of preservation in flame length and radiation.

In general, the gas-phase convective results for temperature, velocity, and heat transfer coefficient were good as anticipated. In fact the corridor

thermal and velocity fields have been studied very extensively in the model by McCaffrey and Quintiere

( 15,16 ) under various experimental conditions with

good correlation to full-scale behavior. An example

from this work (16) is shown in Figure 13 for corridor mid-plane vertical temperature and velocity

distribution under nearly geometric and energy input

similarity conditions. The flow reversals in the corridor were found to result from the restriction of the corridor exit, and were also observed for the model configuration given in Figure 1. Finally, the correlation for heat transfer

coefficient given in Figure 11 appears to have merit

despite the way in which it was derived. A more

careful investigation under steady state conditions

with correction for radiative heating should improve

12
SOFFIT a DOOR

Qfs = 300 kW, H FS = 2 4m

q model = 2 0 hmooel = 0 35 m

TCC)

Fig. 13 A comparison of corridor vertical temperature

and velocity profiles between model and full- scale under approximate similarity conditions

(16).

the accuracy of the results and perhaps decrease the heat transfer coefficient (by about 30 per cent at most) .It would be useful if actual floor covering fire

spread could be simulated in the model since this was the subject of the original full-scale experiments

(11 ) . Although this may be possible, the design of such model experiments would be subject to assumptions,

and the accuracy and interpretation of results would

be subject to uncertainties. The results of the relatively simpler experiment modeled here should convey the difficulties of attempting to model the corridor flame spread problem. From the results of this study it may be con- cluded that:

1. convective processes associated with fires

in enclosures can be scaled sucessfully;

2. radiative transport is not scaled successfully but results can be improved through

analysis or increasing the size of the model;

3. the accuracy of small-scale fire growth

studies will depend on the importance of radiation in the prototype problem and how successful it is modeled.

ACKNOWLEDGMENTS

The study could not have been effectively done without the support of several persons. The authors would like to express their appreciation to William

Rinkinen, who designed and built many of the compo- nents in these experiments, Jim Raines, who operated

the data acquisition system, Margaret Harkleroad , who

programmed the computer for data retrieval and com- parison, and Karen Den Braven who assisted in the model experiments and analysis.

REFERENCES

1 Williams, F. A., "Scaling Mass Fires," Fire

Research Abstract Review, Vol. 11, 1969, pp. 1-23.

2 Gross, D. and Robertson, A. F., "Experimental

Fires in Enclosures," Tenth Symposium (International)

on Combustion, The Combustion Institute, 1965,

pp. 931-942.

3 Thomas, P. H., "Some Studies of Building Fires

Using Models," in The Use of Models in Fire Research , W. G. Berl , ed . , National Academy of Science - National Research Council, Washington, D.C., 1961,

pp. 150-185.

4 Heskestad, G., "Similarity Relations for the Initial Convective Flow Generated by Fire,” ASME paper 72-WA/HT-17 , Nov. 1972.

5 Heskestad, G., "Physical Modeling of Fire,"

J. Fire and Flammability, Vol. 6, 1975, pp. 253-273.

6 Lee, B. T. , "Laboratory Scaling of the Fluid

Mechanical Aspects of Large Fires," Combustion Science

Technology, Vol. 5, 1972, pp. 233-239.

7 Yokoi, S., "Upward Convection Current from a Burning Wooden House," in The Use of Models in Fire

Research , W. G. Berl, ed . , National Academy of Science

- National Research Council, Washington, D.C., 1961,

pp. 186-218.

8 de Ris , J., Kanury , A. M. , and Yuen, M. C., "Pressure Modeling of Fires," Fourteenth Symposium

(International) on Combustion, The Combustion

Institute, 1973, pp . 1033-1044.

9 Alpert , R. L., "Pressure Modeling Transient

Crib Fires," ASME Paper No. 75-HT-6, August 1975.

10 Parker, W. J. and Lee, B. T. , "A Small-Scale

Enclosure for Characterizing the Fire Buildup

Potential in a Room," NBSIR 75-710, National Bureau

of Standards, Washington, D.C., June 1975.

11 Quintiere, J., "Some Observations on Building

Corridor Fires," Fifteenth Symposium (International)

on Combustion, The Combustion Institute, 1974,

pp. 163-174.

12 Quintiere, J. , "The Application and Interpretation of a Test Method to Determine the Hazard of Floor Covering Fire Spread in Building Corridors," International Symposium: Fire Safety of Combustible Materials,

Univ. Edinburgh, Oct. 1975, pp . 355-366a.

13 Steward, F. R. , "Prediction of the Height of Turbulent Diffusion Buoyant Flames," Combustion Science

Technology, Vol. 2, 1970, pp . 203-212.

14 Kawagoe, K. , "Fire Behavior in Rooms," Rep.

No. 27, The Building Research Institute, Japan,

Sept. 1958.

15 McCaffrey, B. J. , and Quintiere, J. G., "Fire-Induced Corridor Flow in a Scale Model Study,"

Symposium on the Control of Smoke Movement in Building

Fires, Proceedings, Conseil International du Batiment

(CIB), Garston, U.K. Nov. 4-5, 1975, pp . 33-47.

16 McCaffrey, B. J., and Quintiere, J. G., "Buoyancy Driven Countercurrent Flows Generated by a Fire Source," Turbulent Buoyant Convection , Hemisphere

Publ. Corp., (to be published), pp . 457-472.

13
COMPLETE SET OF SCALING RESULTS

The complete set of scaling results follow and supplement

Figures 3 - 8 of the previous section. Also Figure 1 is repeated at

a larger size for clarity.
T30

15

Figure

1.

Corridor

geometry

and

instrumentation

for

the

full-scale

experiments

(distance

in

meters)
300

200

0

300

600

900

CORRIDOR FIRE - 1 CRIB

ROOM DOORWPY

>i i »n lii i)<wnr‘^iV%i4A'

<3 T T 0 600 1200

TIME CS)

Figure 14a

Figures 14a-e: TEMPERATURES AT ROOM DOORWAY

16
009

00ti

002

0

002

001

CORRIDOR FIRE - 2 CRIBS

ROOM DOORWPY

Figure 14b

17
100

200

0

200

400

600

Figure 14c

18
T4

CC)

CORRIDOR FIRE - 4 CRIES

ROOM D00RWP7

Figure 14d

19
006

009

008

0

000

001

CORRIDOR FIRE - 5 CRIBS

ROOM DOORWRY

Figure 14e

20
T25

CC)

T23

CC)

200

400

0

200

400

600

CORRIDOR FIRE - 1 CRIB

5 FT. CORRIDOR

.=! O

00

ru

<s

<\i

®

0 —I

1 600 1200

TIME (S)

Figure 15a

Figures 15a-e: TEMPERATURES IN CORRIDOR AT 1.5 m (5 ft) from rear wall

21
T25

(C)

T23

CC)

100

200

0

100

200

300

CORRIDOR FIRE - 2 CRIBS

5 FT. CORRIDOR

Figure 15b

22
T25

(C)

T23

CC)

100

200

0

100

200

300

CORRIDOR FIRE - 3 CRIBS

5 FT. CORRIDOR

Figure 15c

23
T25

CC)

T23

(C)

200

400

0

300

900

900

CORRIDOR FIRE - 4 CRIBS

5 FT. CORRIDOR

Figure 15d

24
T25

CC)

T23

CC)

200

400

0

300

600

900

CORRIDOR FIRE - 5 CRIBS

5 FT. CORRIDOR

Figure 15e
T52

CC)

T50

(C)

100

200

0

200

400

600

CORRIDOR FIRE - 1 CRIB

CEILING

FULL SCSLE

1X7 SCOLE

Figure 16a

Figures 16a-e: CORRIDOR CEILING TEMPERATURES

26
T52

(C)

TS0

CC)

100

200

0

100

200

300

CORRIDOR FIRE - 2 CRIBS

CEILING

FULL SCRLC

Figure 16b

27
T52

(C)

T50

(C)

100

200

0

100

200

300

3 CRIBS

Figure 16c

28
TS2

(C)

T50

(C)

100

200

0

200

400

600

CORRIDOR FIRE - 4 CRIBS

CEILING

___ FULL SCBLE

\/7 SCBLE

Figure 16d
T52

CC)

T50

CC)

100

200

0

200

400

600

CORRIDOR FIRE - 5 CRIBS

CEILING _ FULL SCBLE

Figure 16e

30
HFM

3

CW/CM*)

HFM

2

CW/CM

8)

.00

0.05

0.10

0.00

0.05

0.10

CORRIDOR FIRE - 1 CRIB

ROD I OT 1 ON

FULL. SC3LE

1/7 SCBLE

U)n

o

Figure 17a

Figures 17a-e: INCIDENT RADIATION TO CORRIDOR FLOOR

31
HFM

3

(

W/CM

a)

HFM

2

(W/CM«)

0

0.20

0.00

0.15

0.30

0.45

CORRIDOR FIRE - 2 CRIBS

RQDIQTION

FULL SCPLE

1 /7 SCPLE

Figure 17b

32
CORRIDOR FIRE - 3 CRIBS

ROD 1 AT I ON

FULL. SCALE

1X7 SCALE

OL OC LU UJ

Figure 17c

33
05

'l

00

't

05

'0

00

'0

00'

I

0S

’

0

00'

(a

WO/M

)

5

WdH

(a

WO/M)

£

WdH

CORRIDOR FIRE - 4 CRIBS

RADIATION

FULL SCALE

1X7 SCALE

Figure 17d
HFM

3

C

kl/CM

*)

HFM

2

CW/CM

a)

.00

0.50

1.00

0.00

0.50

1.00

1.50

CORRIDOR FIRE - 5 CRIBS

ROD 1 PT I ON

Figure 17e

35
200

400

0

200

400

600

CORRIDOR FIRE - 1 CRIB

EXIT WINDOW

Figure 18a

Figures 18a-e: TEMPERATURES AT THE CORRIDOR EXIT WINDOW

36
T9

(C)

CORRIDOR FIRE - 2 CRIBS

EXIT WINDOW

TIME CS) TIME (S)

0 600 1200

TIME (S) TIME (S)

Figure 18b

37
00£

003

00

I

0

001

(0)

91

(0)

61

CORRIDOR FIRE - 3 CRIBS

EXIT WINDOW

<3in

<3 ' I 0 600

TIME (S)

~i1200

Figure 18c

38
T9

(C)

T6

CC)

200

400

0

200

400

600

CORRIDOR FIRE - 4 CRIBS

EXIT WINDOW

Figure 18d

39
100

200

0

100

200

300

CORRIDOR FIRE - 5 CRIBS

EXIT WINDOL-i

Figure 18e

40
VELOCITY

S

(M/SEC)

HTC

4

CW/M*

K)

CORRIDOR FIRE - 1 CRIB

HEAT TRANSFER COEFFICIENT

VELOCITY

FULL. SCfiLE

1X7 SCnLE

Figure 19a

Figures 19a-e: CEILING HEAT TRANSFER COEFFICIENT

AND

EXIT WINDOW VELOCITIES

41
VELOCITY

S

(M/SEC)

HTC

4

(W/M*

K)

CORRIDOR FIRE - 2 CRIBS

HEfiT TRANSFER COEFFICIENT

VELOCITY

FUI-U SCRLE

1/7 9CPLE

Figure 19b

42
VELOCITY

6

(M/SEC)

HTC

4

CW/M

2

K)

CORRIDOR FIRE - 3 CRIES

HEAT TRANSFER COEFFICIENT

VELOCITY_ FULL SCALE

\/T SCALE

Figure 19c

43
''ELoerrr

§

(H/sec)

htc

4

cw/m*

k)

CORRIDOR FIRE - 4 CRIBS

HEAT TRANSFER COEFFICIENT

VELOCITY__ FULL SCRLE

1X7 SCRLE

Figure 19d

44
VELOCITY

G

(M/SEC)

HTC

4

(

W/M

*

K)

CORRIDOR FIRE - 5 CRIBS

HEfiT TRRNSFER COEFFICIENT

VELOCITY—„ FULL SCRLE

1/7 SCRLE

Figure 19e

45
DESCRIPTION OF THE FULL-SCALE EXPERIMENTS

A total of nine experiments were run during a two-week period in

June of 1974. All were run in the evening, and two runs were conducted

on the same evening on occasion. The instrumentation for the corridor

was described in the previous section and is shown in figure 1.

The adjoining room in which the primary fires were contained was

2.8 m high and 2.4 m x 2.6 m in floor area. Two small vents (18 cm x

35 cm) to permit additional inlet airflow into the burn room were

located about 30 cm above the floor on opposite walls. These vents

were either open or closed.

The motivation of the experiments was to determine the effect of

the room vents, the effect of the fuel location, and the effect of an obstacle or floor fire at the room doorway.

The wood cribs were hemlock, and made of 3.8 cm square sticks, with

7.6 cm between the sticks, in 16 layers, having four sticks in each

layer (except only two in the bottom layer) . All the cribs were placed

on a platform which covered most of the room floor and the platform was suspended a short distance above it. The platform was hung by a cable

from a cooled load cell. The cribs were ignited with 250 g of heptane

centered in a pan below each crib.

Table 3 gives a summary of experiments. The arrangement and

quantity of fuel are given along with ventilation conditions. Experiments 8 and 9 were motivated by a behavior in the flow pattern that

appeared to occur in the floor covering corridor fire experiments [11].

This was a decrease in the gas flow rate exchange at the room doorway

which was speculated to have been caused by the advancing (opposing the

airflow) floor fire. That is, as the floor covering fire increased in

size, it would tend to become less affected by the air flowing along the

floor. The consequences of this would be a taller floor flame, more

forward heat transfer, and an increase in fire spread - a possible

mechanism of rapid fire growth (flashover) . Unfortunately because of

a draft problem with some of the pressure transducers for the velocity

46
Table

3

Summary

of

Experimental

Conditions

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47
probes, an illumination of this phenomena was not accomplished by the

results of Experiments 8 and 9.

Except for the scaling analysis, the results of the full-scale

experiments were not extensively analyzed. They will simply be documented here for the benefit of others who are concerned about particular

features of this type of fire scenario.

The complete set of data are enclosed on microfiche with this

report. Some features of the results will be discussed.

Figures 20, 21 and 22 show the temperature distribution on the

corridor during the peak of the room fire for four cribs under three

different conditions. Some observations are apparent. First, near the room doorway the temperature field is skewed by the effect of the

doorway jet impinging on the corridor wall and turning down the corridor.

Eventually, the temperature distribution in the hot gas flow becomes

more uniform across the corridor width. The maximum corridor heating

occurs for the condition of centered cribs with room vents closed. The

increase in air supply with the vents open, lowers the temperatures overall. For the case of the corner location, the cribs burned at a slower

rate, thus resulting in the lowest corridor temperature. This latter

case could have had the opposite effect (over the centered, closed vent

case) had the burning rate been the same. This would follow since it

is expected that less air would be entrained into corner fires than room

centered fires. Figure 23 shows these same effects in terms of heat

flux to the corridor ceiling and floor.

48
CONCLUSIONS

It has been shown that some scale model results could be accurately

extrapolated to full-scale counterpart systems. However there can never

be complete dynamic similarity in fire problems and this fact leads to

uncertainty in the reliability of the scale model results. For the

design conditions of these experiments, i.e. slowly varying fires, it was found that convective processes were modeled well. Since radiation

effects were ignored in the scaling criteria, it was anticipated that

results would be inaccurate when radiation becomes significant.

However, the degree of disagreement in radiant flux could be estimated.

The full-scale experiments demonstrated that additional room ventilation and the location of the fuel load in the room, did affect the

heating conditions of the corridor. The temperature field in the

corridor appeared to be affected locally by geometry; but by 7 meters

downstream of the room, the field became two-dimensional and invariant

across the corridor width.

49
50

Figure

20.

Peak

average

temperatures

in

°C

for

Ruri

1:

4

cribs,

centered,

vents

open
51

Figure

21.

Peak

average

temperatures

in

°C

for

Run

6:

4

cribs,

centered,

verts

closed
52

Figure

22.

Peak

average

temperatures

in

°C

for

Run

7:

4

cribs,

corner

location,

vents

closed
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53
NBS-1 14A (REV. 7-73)

U.S. DEPT. OF COMM. BIBLIOGRAPHIC DATA

SHEET

1. PUBLICATION OR REPORT NO. fclBSIR 77- 13182. Gov’t Accession

No.

3. Recipient’s Accession No.

4. TITLE AND SUBTITLE

The Impact of a Room Fire on a Corridor with Considerations

of Fuel Load, Ventilation, and Scaling

5. Publication Date November 1977

6. Performing Organization Code

7. author(S) j ^ Quintiere, E. McCaffrey, T. Kashiwagi, K. Den Bravek?’ Performin s 0re an - Re P

ort n°. Harkleroad, J. Raines and W. Rmkin&n

9. PERFORMING ORGANIZATION NAME AND ADDRESS

NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

10. Project/Task/Work Unit No, 491

11. Contract/Grant No.

12. Sponsoring Organization Name and Complete Address (Street, City, State, ZIP)

Same as No. 9

13. Type of Report & Period Covered

Final Report

14. Sponsoring Agency Code

15. SUPPLEMENTARY NOTES

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

bibliography or literature survey, mention it here.)

A study was conducted of heat transfer and temperature field imparted to a corridor by a room fire in order to assess the potential of fire spread along the

corridor. Wood cribs were used as the fuel load which ranged from 20 to 120 kg.

Also the effects of ventilation and fuel load location were examined. The results

showed these effects to be significant.

A corresponding scale model study was conducted using gas burners as a fuel

supply. A scaling criteria was developed which emphasized the solid wall conductive

and gas phase convective transport processes. As a result limitations were encountered at high temperature when radiation became more significant. In

general the convective processes appeared to scale well and radiation limitations

could be assessed through analysis.

17. KEY WORDS (six to twelve entries,' alphabetical order; capitalize only the first letter of the first key word unless a proper name; separated by semicolons)

Corridors; fuel load; heat transfer; room fires; scaling; ventilation

18. AVAILABILITY [X Unlimited

r For Official Distribution. Do Not Release to NTIS

d! Order From Sup. of Doc., U.S. Government Printing Office Washington, D.C. 20402, SD Cat. No. CU

k ! Order From National Technical Information Service (NTIS)

Springfield, Virginia 22151

19. SECURITY CLASS

(THIS REPORT)

UNCL ASSIFIED

21. NO. OF PAGES

60

20. SECURITY CLASS

(THIS PAGE)

UNCLASSIFIED

22. Price

$ 5-25

USCOMM-DC 29042- P 74 I