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PUBLICATIONS

NBSIR 88-3753

9th Joint Panel Meeting of the

UJNR Panel on Fire Research and

Safety

Nora H. Jason

Barbara A. Houston

U.S. DEPARTMENT OF COMMERCE

National Bureau of Standards

National Engineering Laboratory

Center for Fire Research

Gaithersburg, MD 20899

April 1988

156 ATIONAL BUREAU OF STANDARDS

8803753

1988

C.2

Research Information Center

National Bureau of Stanuards

NBSIR 88-3753 Gaithersburg, Maryland 20099

9TH JOINT PANEL MEETING OF THE

UJNR PANEL ON FIRE RESEARCH AND

SAFETY

AlBsc

ClC/oo

/VO. ^3~7

Nora H. Jason

Barbara A. Houston

U.S. DEPARTMENT OF COMMERCE

National Bureau of Standards

National Engineering Laboratory

Center for Fire Research

Gaithersburg, MD 20899

April 1988

U.S. DEPARTMENT OF COMMERCE, C. William Verity, Secretary

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director

NINTH JOINT PANEL MEETING

OF THE UJNR PANEL ON

FIRE RESEARCH AND SAFETY

Joint with Combustion Toxicity and 6th Expert Meeting of the

U . S

. - Japan-Canada Cooperative Research Group on Toxicity of Combustion

Products from Building Materials and Interior Goods

May 4-8, 1987

Factory Mutual Research Corporation

Norwood, Massachusetts

TABLE OF CONTENTS

Page

Agenda vii

List of Members (Japan) xiv

List of Members (U.S.) xvi

List of Combustion Toxicity Participants xviii

Group Photograph xix

OPENING SESSION ! 1

FIRE AND SMOKE PHYSICS 3 Progress Reports:

Fire and Smoke Physics in Japan 4

Fire and Smoke Physics in the U.S 8 Forced Smolder Propagation in a Horizontal Fuel Layer and the

Transition to Flaming Combustion 13

Wood Ignition and Pyrolysis 24

Deterministic Properties of Turbulent Diffusion Flames from Low Q*

Fires 41

Transient Combustion in a Turbulent Fire 50

Numerical Simulation of Early Stage of a Compartment Fire by Field

Model 63

Critical Discussion on Fire Development 73

Radiation from Diffusion Flames-Smoke Point 83

Effect of Radiant Flux on Smoke Emission 100

The Distribution of Soot in Large Buoyant Diffusion Flames Ill

Predicting Capability of a Multiroom Fire Model 113

RISK, HAZARD, AND EVACUATION 130

Progress Reports:

Fire Risk Analysis Methods in the U.S 131

Risk, Hazard and Evacuation in Japan 136

Discussion 138

A Method of Fire Risk Analysis -- Estimation of Life Risk in

Buildings 139

Discussion 145

Response Surface Approximation of a Fire Risk Analysis Computer

Code 146

Discussion 155

An Engineering Fire Protection Design Assessment System 157

Discussion 170

Basic Structure of the Evacuation Safety Design Method 172

Discussion 188

Fire Risk Evaluation Method for Multi -Occupancy Building 189

Discussion 201

An Evacuation Model for the Use in Fire Safety Designing of

Buildings 203

Discussion 218

A Summary of the Assumptions and Limitations in Hazard I 219

Discussion 229

iii
Page

FIRE TOXICITY AND CHEMISTRY 231

Progress Reports

:

Fire and Toxicity Chemistry in Japan 232

Fire Toxicity and Chemistry Research in the U.S 235

Canadian Overview 238

Incapacitation of Mice Exposed to Gases in Full-Scale Fire Tests 240

Determination of the Kinetic Parameters and the Heat of Combustion

of the Volatiles for Wood 257

Correlation for the Generation of Carbon Monoxide in Smalland Large-Scale Fires 267

RISK, HAZARD, AND EVACUATION (CONT.) 279

EXITT - A Simulation Model of Occupant Decisions and Actions in

Residential Fires 280

Discussion 292

Emotional Instability in Fire Smoke -- Experiment of Human

Behavior in Smoked Filled Corridor 294

Discussion 304

Large Fires and Nuclear Winter 305

Discussion . ? 311

OPEN TECHNICAL SESSION 313

FMRC Sprinkler Research 314

Discussion 324

Progress in Modeling Sprinkler Performance 325

Combustion Behaviour of Building Materials Using a Model Box with

Air Supplied System 336

Fire Induced Flow in a Clean Room with Downward Vertical Laminar

Flow 357

FIRE TOXICITY AND CHEMISTRY (CONT.) 367

Toxicological Effects of Different Time Exposures to the Fire

Gases: CO or HCN or to CO Combined with HCN or C02 368

Discussion 384

Toxic Gas Test by the Several Pure and Mixture Gases 386

Discussion 396

Strategy of Fire Suppression by Using Fire Retardants 397

Discussion 404

Studies on the Toxicity' of Smoke Containing Hydrogen Chloride 405

Discussion 416

Toxic Gases from Combustion of Wood Treated with Retardants 417

Discussion 427

Bench- and Full-Scale Toxicity Tests of FRP Furniture 428

Discussion 441

Evolution of Toxic Gases from Experimental Fires in an Existing

Building 443

Discussion 454

Human Effects of Poisonous Gases Generated During Fire - Especially,

Presence of HCN and CO 455

iv
Page

A Proposed Method for the Simultaneous Analysis of HCN and N0X in

Fire Gases 461

Discussion 469

Evaluation of Combustion Toxicity from Various Materials Using a New Testing Apparatus 470

Discussion 488

A Review of the Five-Year Cooperative Program on Smoke Toxicity 490

Panel Discussion 493

OPEN TECHNICAL SESSION (CONT.) 495

Characterization of Flame Spread Over PMMA Using Holographic

Interferometry Sample Orientation Effects 496

Discussion 508

A Numerical Study of Periodic Alternating Flow of Fires 509

CLOSING SESSION 519

Discussion 520

Resolutions

Closing Addresses

Author's Index 52 7

v

AGENDA

9th UJNR PANEL ON FIRE RESEARCH AND SAFETY

MAY 4-8, 1987

NORWOOD, MASSACHUSETTS

Saturday. May 2

Japanese arrive; no activities.

Sunday . May 3

Early P.M.: Planning Meeting (Delegation Chairmen)

Selection of Session Chairmen

Plans for 10th UJNR

3:00 Tour of Boston and Quincy Market

Monday , May 4

9:00 Opening Ceremonies

Call to order by Dr. Jack E. Snell, U.S. Chairman

Welcoming Remarks by Mr. John A. Love, President of Factory Mutual

Engineering and Research

Remarks and Introductions of U.S. Panel Members by Dr. Jack E. Snell

Remarks and Introductions of Japanese Panel Members by Mr. Akira

Takahashi

9:40 Election of Officers

Approval of Proceedings of the 8th Panel Meeting

Approval of Agenda for the 9th Panel Meeting

Election of Session Chairmen

Appointment of Resolution Committee

Other Business

9:50 Group Photograph

FIRE AND SMOKE PHYSICS - (Chairman: John de Ris)

10:00 Y. Hasemi, Progress Report - Japan

10:15 E. Zukoski, Progress Report - U.S.

10:30 Refreshment Break

10:45 T.J. Ohlemiller, "Forced Smolder Propagation in a Horizontal Fuel

Layer and the Transition to Flaming Combustion" (delivered by

T. Kashiwagi)

11:30 A. Atreya and H. Emmons, "Wood Ignition and Pyrolysis"
12 : 15 Lunch

Chairman: Makoto Tsujimoto

1:30 Y. Hasemi and M. Nishihata, "Deterministic Properties of Turbulen

Diffusion Flames from Low Q* Fires"

2:15 H. Baum and R. Rehm, "Transient Combustion in a Turbulent Fire"

3:00 Refreshment Break

3:15 T. Yoshikawa, J. Mashige, T. Joh, and 0. Sugawa, "Numerical

Simulation of Early Stage of a Compartment Fire by Field Model"

4:00 T. Hirano, "Critical Discussion on Fire Development"

4:45 Adjourn

viii
Tuesday. May 5

8:00 Depart for FMRC West Gloucester Facility

9:30 Tour FMRC West Gloucester Facility

11:30 Lunch

12:30 Return to FMRC Norwood

PARALLEL SESSIONS

A. FIRE AND SMOKE PHYSICS (cont.) - (Chairman: Patrick J. Pagni)

1:30 J. de Ris , "Radiation from Diffusion Flames-Smoke Point"

2:15 G. Mulholland, V. Henzel, and V. Babrauskas , "Effect of Radiant

Flux on Smoke Emission"

3:00 Refreshment Break

3:15 E. Zukoski , "The Distribution of Soot in Large Buoyant Diffusion

Flames

4:00 K. Nakamura, "Predicting Capability of a Multiroom Fire Model"

4:45 Adj ourn

Evening: Cocktails and Formal Dinner at FMRC

B. RISK, HAZARD, AND EVACUATION - (Chairman: Harold E. Nelson)

1:30 J. Hall, Progress Report - U.S.

1:45 T. Tanaka, Progress Report - Japan

2:00 M. Tsujimoto, "A Method of Fire Risk Analysis -- Estimation of Life

Risk in Buildings"

2:45 Refreshment Break

3:00 M. Brandyberry and G. Apostolakis, "Response Surface Approximation

of a Fire Risk Analysis Computer Code"

3:45 H. Nelson, "An Engineering Fire Protection Design Assessment

System"

4:30 Adj ourn

Evening: Cocktails and Formal Dinner at FMRC

ix
Wednesday. May 6

-PARALLEL SESSIONS

B. RISK, HAZARD, AND EVACUATION (cont.) - (Chairman: James G

Quintiere

)

9:00 T. Tanaka, "Basic Structure of the Evacuation Safety Design

Method"

9:45 A. Sekizawa and T. Jin, "Fire Risk Evaluation Method for MultiOccupancy Buildings"

10:30 Refreshment Break

10:45 K. Takahashi and T. Tanaka, "An Evacuation Model for the Use in

Fire Safety Designing of Buildings"

11:30 R. Bukowski , "A Summary of the Assumptions and Limitations in

HAZARD 1"

12:15 Lunch

1:30 Tour of FMRC Laboratories

Evening: Light Supper

Fenway Park: Major League Baseball

Boston Red Sox vs. Oakland Athletics

A. FIRE TOXICITY AND CHEMISTRY - (Chairman: Toshisuke Hirano)

9:00 H. Suzuki, Progress Report - Japan

9:15 R. Gann, Progress Report - U.S.

9:30 T. Harmathy, Progress Report - Canada

9:45 T. Tanaka and M. Yoshida, "Incapacitation of Mice Exposed to Gases

in Full-Scale Fire Tests"

10:30 Refreshment Break

10:45 W. Parker, "Determination of the Kinetic Parameters and the Heat

of Combustion of the Volatiles for Wood"

11:30 A. Tewarson, "Correlation for the Generation of Carbon Monoxide in

Small- and Large-Scale Fires"

12:15 Lunch

1:30 Tour of FMRC Laboratories

Evening: Light Supper

Fenway Park: Major League Baseball

Boston Red Sox vs. Oakland Athletics

x
Thursday. May 7

PARALLEL SESSIONS

B. RISK, HAZARD, AND EVACUATION (cont.) - (Chairman: John R. Hall, Jr.)

9:00 B. M. Levin, "EXITT - .A Simulation Model of Occupant Decisions and

Actions in Residential Fires"

9:45 T. Jin and T. Yamada, "Emotional Instability in Fire Smoke--

Experiment of Human Behavior in Smoke Filled Corridor"

10:30 Refreshment Break

10:45 G. Carrier, "Large Fires and Nuclear Winter"

11:30 General Discussion

12:15 Lunch

B. OPEN TECHNICAL SESSION - (Chairman: Hiroaki Suzuki)

1:30 C. Yao , "FMRC Sprinkler Research"

2:15 D. Evans, "Progress in Modeling Sprinkler Performance"

3:00 Refreshment Break

3:15 F. Saito and M. Yoshida, "Combustion Behaviour of Building

Materials Using a Model Box with Air Supplied System"

4:00 0. Sugawa, Y. Oka and H. Hotta, "Fire Induced Flow in a Clean Room

with Downward Vertical Laminar Flow"

4:45 Adj ourn

Evening: Reception and Dinner at NFPA and Tour of Facilities

xx
Thursday, May 7 (cont .

)

A. FIRE TOXICITY AND CHEMISTRY (cont.) - (Chairman: Takeyoshi Tanaka)

9:00 B.C. Levin et al., "Toxicological Effects of Different Time

Exposures to the Fire Gases: CO or HCN or to CO Combined with HCN

or C02

"

9:45 T. Sakurai , "Toxic Gas Test by the Several Pure and Mixture Gases"

10:30 Refreshment Break

10:45 T. Hirano and I. Nakaya, "Strategy of Fire Suppression by Using

Fije Retardants"

11:30 G. Hartzell, A. Grand, and W. Switzer, "Studies on the Toxicity of

Smoke Containing Hydrogen Chloride"

12:15 Lunch

Chairman: Richard G. Gann

1:30 T. Hirata and S. Kawamoto, "Toxic Gases from Combustion of Wood

Treated with Retardants"

2:15 A. Kim and Y. Tsuchiya, "Bench- and Full-Scale Toxicity Tests of

FRP Furniture"

3:00 Refreshment Break

3:15 T. Morikawa, E. Yanai and T. Nishina, "Evolution of Toxic Gases

from Experimental Fires in an Existing Building"

4:00 Y. Tsuda and Y. Nishimaru, "Human Effects of Poisonous Gases

Generated During Fire - Especially, Presence of HCN and CO"

4:45 R. Curtis and Y. Tsuchiya, "A Proposed Method for the Simultaneous

Analysis of HCN and N0X in Fire Gases"

5:30 Adj ourn

Evening: Reception and Dinner at NFPA and Tour of Facilities

xii
Friday. May 8

PARALLEL SESSIONS

A. FIRE TOXICITY AND CHEMISTRY (cont.) - (Chairman: Raymond Friedman)

9:00 S. Yusa, "Evaluation of Combustion Toxicity from Various Materials

Using a New Testing Apparatus"

9:45 Panel Discussion: A Review of the Five-Year Cooperative Program on Smoke Toxicity; R. Gann, H. Suzuki, and Y. Tsiichiya

10:30 Adj ourn

B. OPEN TECHNICAL SESSION (cont.) - (Chairman: James G. Quintiere)

9:00 A. Ito and T. Kashiwagi , "Characterization of Flame Spread Over

PMMA Using Holographic Interferometry Sample Orientation Effects"

9:45 K. Satoh, "A Numerical Study of Periodic Alternating Flow of

Fires"

10:30 Adj ourn

10:45 CLOSING CEREMONIES

12:00 Lunch

Xlll
LIST OF MEMBERS (JAPAN)

Panel Members

Mr. Akira Takahashi (Japanese Chairman)

Director General

Building Research Institute

Ministry of Construction

Dr. Hiroaki Suzuki

Head, Smoke Control Division

Environment, Design and Fire Department

Building Research Institute

Ministry of Construction

Dr. Takeyoshi Tanaka

Chief Research Member

Testing and Evaluation Department

Building Research Institute

Ministry of Construction

Dr . Yuj i Hasemi

Chief Research Member

Smoke Control Division

Environment, Design and Fire Department

Building Research Institute

Ministry of Construction

Dr. Tadahisa Jin

Head, The Second Research Division

Fire Research Institute

Fire Defense Agency

Ministry of Home Affairs

Mr. Tokio Morikawa

Chief, Combustion Section

The First Research Division

Fire Research Institute

Fire Defense Agency

Ministry of Home Affairs

Dr. Toshimi Hirata

Chief Research Member

Forestry and Forest Products Research Institute

Ministry of Agriculture, Forestry and Fisheries

Prof. Makoto Tsuj imoto

Department of Architecture

Faculty of Engineering

University of Nagoya

xiv
Associate Members

Prof. Toshisuke Hirano

Department of Reaction Chemistry

Faculty of Engineering

University of Tokyo

Dr. Saburo Horiuchi

Chairman, Japanese Association of

Fire Science and Engineering

Dr. Yukio Tsuda

Department of Legal Medicine

School of Medicine

Yokohama City University

Dr. Osami Sugawa

Center for Science and Technology

Science University of Tokyo

Mr. Toshiro Sakurai

Japanese Association of Fire Science and Engineering

(Research Institute of Marine Engineering)

Mr. Kiyoshi Takahashi

Japanese Association of Fire Science and Engineering

(Technical Research Division, Fujita Corporation)

Mr. Kazuhito Nakamura

Japanese Association of Fire Science and Engineering

(Design Division, Shimizu Constructing Co., Ltd.)

Ms. Tomoko Joh

Japanese Association of Fire Science and Engineering

(Century Research Center)

Mr. Masashi Yoshida

Research Member

Testing and Evaluation Department

Building Research Institute

Ministry of Construction

Current NBS Visiting Scientist

Dr. Ai Sekizawa

Branch Chief Research Member

The Third Research Division

Fire Research Institute

Fire Defense Agency

Ministry of Home Affairs

(June 1986/1987, Visiting Scientist at NBS)

xv
LIST OF MEMBERS (U. S .

)

Par, el Members

Dr. Jack E. Snell (U.S. Chairman)

Director, Center for Fire Research

National Bureau of Standards

Dr. Richard G. Gann (CFR Coordinator)

Chief, Fire Measurement and Research Division

Center for Fire Research

National Bureau of Standards

Dr. Takashi Kashiwagi (Secretary)

Polymer Combustion Research

Fire Measurement and Research Division

Center for Fire Research

National Bureau of Standards

Dr. Raymond Friedman (retired)

Vice President of Research

Factory Mutual Research Corporation

Mr. Harold E. Nelson

Center for Fire Research

National Bureau of Standards

Prof. Patrick J. Pagni

Vice Chairman, Department of Mechanical

Engineering

University of California, Berkeley

Dr. John de Ris

Factory Mutual Research Corporation

Prof. Edward E. Zukoski

Division of Engineering and

Applied Sciences

California Institute of Technology

Dr. James G. Quintiere

Chief, Fire Science and Engineering Division

Center for Fire Research

National Bureau of Standards

Dr. John R. Hall, Jr.

Director, Fire Analysis Division

National Fire Protection Association

xvi
Honorary Members

Prof. Howard Emmons (retired)

Division of Applied Science

Harvard University

Prof. Hoyt C. Hottel (retired)

Department of Chemical Engineering

Massachusetts Institute of Technology

Associate Members

Prof. Arvind Atreya

Division of Engineering Research

Michigan State University

Dr. Gordon E. Hartzell

Head, Department of Fire Technology

Southwest Research Institute

Dr. Archibald Tewarson

Factory Mutual Research Corporation

Prof. George Apostolakis

Mechanical, Aerospace and Nuclear

Engineering Department

University of California, Los Angeles

Mr. Cheng Yao

Manager, Applied Research Department

Factory Mutual Research Corporation

Prof. George Carrier

Division of Applied Sciences

Harvard University

Dr. Howard R. Baum

Center for Fire Research

National Bureau of Standards

Dr. Barbara C. Levin

Center for Fire Research

National Bureau of Standards

Dr. Bernard M. Levin

Center for Fire Research

National Bureau of Standards

xvii
Combustion Toxicity Participant

Dr. Yoshi Tsuchiya

Division of Building Research

National Research Council of Canada

xviii
XIX

'

OPENING SESSION

1
Opening Session

Dr. Richard Gann, Coordinator for the 9th UJNR Panel Meeting, opened the joint

session by welcoming the delegation to the Factory Mutual Research

Corporation. He then introduced Mr. John Love, President of Factory Mutual

Engineering and Research for opening remarks. Dr. Jack Snell, Director,

Center for Fire Research, and Mr. Akira Takahashi, Director General, Building

Research Institute, introduced the United States and Japanese panel members.

The Combustion Toxicity and 6th Expert Meeting of the U . S

. -Japan- Canada

Cooperative Research Group on Toxicity of Combustion Products from Building

Materials and Interior Goods was held in conjunction with the 9th UJNR Panel

Meeting. The Combustion Toxicity meeting was called Fire - Toxicity and

Chemistry and took place on Wednesday, Thursday and Friday, May 6, 7 and 8,

1987 .

Selection of the Chairmen for the sessions was made. The Panel voted

unanimously to make the following individuals session chairmen:

Fire and Smoke Physics J. de Ris M. Tsuj imoto

P. Pagni

Risk

,

Hazard, and Evacuation H. Nelson J . Quintiere

J . Hall

Fire Toxicity and Chemistry T. Hirano T. Tanaka

R. Gann R. Friedman

Open Technical Session H. Suzuki J . Quintiere

The minutes of the 8th Joint Panel Meeting in Tsukuba were approved and will

be printed in the Proceedings to be published by the Japanese delegation.

The agenda for the week was reviewed and approved. Procedural details for the

sessions were discussed.

The opening session adjourned after the above-mentioned discussions were completed. The group photograph was taken.
FIRE AND SMOKE PHYSICS SESSION

3
PROGRESS REPORT ON FIRE AMD SMOKE PHYSICS IN JAPAN

Yuj i Hasemi

Building Research Institute

INTRODUCTION

Studies on fire and smoke physics reported after the previous UJNR. Joint

Meeting held in Tsukuba are reviewed. The following journals and preprints

are referred to for this purpose except for those essentially reported at

this or the previous meeting.

a) Transactions of Architectural Institute of Japan (AIJ)

.

b) Bulletine of Japanese Association of Fire Science and Engineering (JAFSE)

.

c) Proceedings of the Annual Meeting of AIJ.

d) Proceedings of the Annual Meeting of JAFSE.

e) Fire Science and Technology.

The studies introduced in this report may be classified into the following

research areas; the review will be made on the basis of this classification,

i) analysis or observation of overall behaviors of enclosure fires,

ii) dynamics of turbulent diffusion flames. • iii) flame spread.

iv) experiments and calculation of smoke layer formation process in a huge

enclosure

.

v) application of fire modeling to specific problems.

STUDIES ON OVERALL BEHAVIOR OF ENCLOSURE FIRES

Influence of the location of fuels on the behavior of an' enclosure fire was studied experimentally by Takeda(l,2) using 0.40m and 0.25m cube model

enclosures with PMMA slabs as the fuel on the floor, the wall and the

ceiling. He reported that burning rate is somewhat lower with the fuel on the

wall than on the floor, and for che fuel on the .ceiling, stable combustion

was obtained only for small openings (A /h < 0 . 006m “) . He also reported the

overall behavior of PMMA fires in a model enclosure with two openings, one window and one doorway, (3) and proposed a concept of "equivalent ventilation

parameter" to explain the burning rate based on the opening conditions. He

also compared "floor combustion" with "wall combustion" in this experiment,

and reported chat the behavior of the smoke flowing out of openings is very

different between these two conditions; the smoke flew out from only one opening at floor fire and it flew out from the two openings in the case of

wall fires.

While in these reports, he reported an oscillatory combustion in

enclosure fires, Nitta has cried to explain this oscillation by showing that an oscillation equation can be derived from the equations of energy and

momentum conservation for the fire enclosure(4) . In this theory, the

oscillation is attributed essentially to che interaction among che

ventilation rate, enclosure temperature and pressure difference on the

opening rather than the change of oxygen or fuel gas concentrations in the

combusting field.

4
DYNAMICS OF TURBULENT DIFFUSION FLAMES

Oscillatory behavior is also well known in turbulent diffusion flames.

Matsubara and Satoh of FRI(5,6) have tried to examine an idea of fluiddvnamxc

mechanism for this kind of oscillation. Matsubara assumed an interaction

between the buoyancy and the inductive power within the flame leading to the

generation of the oscillation; he made numerical calculations on the plumes

from a "ring" heat source using a field model, and obtained a mushroom—shaped

convection above the heat source. Satoh examined the influence of small

fluctuation on the behavior of a turbulent plume by using UNDSAFE, and

obtained a frequency—heat release rate relationship somewhat close to the

conventional experimental relationships obtained in enclosure fires.

Fire whirl is another characteristic fluiddynamic phenomenon in relation

with turbulent diffusion flames. Hasemi(7) has solved the basic equations

describing incompressible fire whirls, and obtained the distributions of

temperature and velocity as a function of external circulation. The

calculated velocity field is very close to the experiments of Emmons and

Mayle

.

FLAME SPREAD

Studies concerned with flame spread in fire situation have been reported

by Takeno, Kitaoka and Hirano(8) and Terai and Kato ( 9 , 10 , i 1 )

.

Takeno et al made measurements of flame spread and temperature

distribution using a 0.4m x 1 . 2m fuel tray containing glass beads and

kerosene to simulate a flame spread along a ground soaked with combustible

liquid(8). They reported that the flame spread velocity along the fuel

surface in this situatiqn is considerably smaller than for a combustion along

the liquid surface; this reduction suggests the dissipation of the Marangoni

effect in their experiments. They also reported that the temperature

distribution within the fuel bed is somewhat close to that generally observed

in a flame spread along a combustible solid.

Terai et al have reported a calculation model of the combustion on a combustible floor to predict flame spread from material properties of the

floor, assuming very stage of an enclosure fire(9)_. Their model is

essentially based on heat transfer taking account of radiation from flame to fuel surface and its surroundings as well as heat conduction within the

combusting solid, and for the two-dimensional calculation of the temperature

rise of the surrounding surface of a flame, they have applied the integral

equation method(lO). More recently, they have measured heat transfer within

and in the surroundings of a turbulent diffusion flame to obtain necessary

properties for this model, and finally compared some aspects between the

experiments and their model (11). While there is still considerable difference

between their calculation and experiments, their work is valuable to know

what component process should be studied to predict fire spread in relatively

early fires.

FIRE AND SMOKE IN LARGE ENCLOSURE

Studies on fire and smoke control in a large enclosure have been getting

popular in Japan, probably due to the growing interest on the construction of

air-supported structures. Katsuno et al(12,13), Tanaka et al(14), Nagasawa et

al( 15) and Akeno et al(16) reported full-scale and model experiments while

there are still many papers reporting smoke-layer calculations.

Katsuno et al reported an outline of full-scale fire experiments using

the former Kuramae Kokugikan Smo Arena conducted in 1984(12,13); the space

was approximately ' 6000m~ in area and 32m high and the • experiments using

methanol and clothings as fuels have provided basic data on the general

behavior of fire and smoke, and effectiveness of sprinklers and smoke venting

in a large structure. They reported that the observed smoke layer height is

5
close Co Che calculacion based on Thomas' encrainmenc model wich empirical

correccion of Tsujimoto.

Tanaka eC al developed an analycical calculacion model of smoke

formacion process in a large structure with/without forced vencing(14. Part

I) , and compared this theory wich experiments wich methanol as the fuel at

Che full scale fire test laboratory of 3RI , 27m x 20m in plan and 20m

hign(14, Part II). They reported that che theory is very close to the

experimental smoke layer height, if the lower bound of smoke layer in

experiment is defined as the height at che discontinuous change of

temperature and smoke density. They also made smoke layer measurements using

a few Tsukuba Expo, pavilions before their destruction, and compared che

results with che model (16). However, in some of the pressure measurement, che

reported interior pressure was always lower than the exterior one; this

suggests some difficulties in simultaneous pressure measurement around such a large structure.

Nagasawa ec al reported model experiments on smoke layer formation

process in a 2QIm long and 60m high air-supported structure using its 1/50

reduced model(15). They controlled heat release race at che anticipated fire source for the conservation of the Archimedes number, and reported chat the

resulted smoke layer height is close to a calculation based on Zukoski's

entrainment model.

SPECIFIC PROBLEMS

Hamada et al have reported some experimental aspects of flame from line

burner flowing along a surface composed of wall and ceiling(17), and

represented flame lengt£ and the distribution of temperature rise along a flame as functions of Q . for a line fire. Their result shows that che flame

height relationship is close to Hasemi's correlation for wall fires while

their experiment was made not only on a wall. They reported that their excess temperature along the ceiling is considerably higher than the experimental

relationship without wall reported hy Heskestad.

Mizuno et al(18) have observed burning behavior of upholstered

furnitures using BRI ' s calorimeter, and reported their calorific properties;

reported calorific potential of two different mattresses with urethanefoam

was L4.7MJ/kg and 12 . 8MJ/kg .

• Heat release rate measured by oxygen consumption

method was almost proportional to the mass loss rate measured with load cells

from very early stage to fully developed combustion.

Yamada et al have measured temperature, velocity, C0 ? concentration and

smoke particles of smoke in an approximately 20m long, “2. 5m high corridor

with 15 . 2kW propane line burner as the fuei(19). Their concern is to obtain

some criteria for the formation of two layers, since the formation of "clean"

layer in the lower part of a corridor is an important condition for

evacuation in fire. They suggest, from the vertical distribution of smoke

particles at several distance, the existence of some recirculation within the

ceiling layer in very far region from che fire source. Hasemi et at have derived a simple method to estimate material

properties concerned to ignitability from ignitability tests on che basis of

the concept ignition as a result of rise of surface temperature to ignition

temperature due to external radiation(20) . Estimated ignition temperature and

thermal inertia are very close to the values obtained by other independent

methods for industrial materials such as PMMA and particle board, however for

wood, the estimated properties using the data from many institutions are somewhat different from conventional method; the discrepancy is probably due

to the influence of moisture within the specimens.

Mima et al reported some experimental results on the movement of smoke

generated from smoke candles in a vertical laminar clean room(2I) . As

estimated naturally, the measured smoke density distribution suggests that

the diffusion of smoke within a laminar clean room is very insignificant.

6
REFERENCES

1 Takeda ,H.

, "Scale modeling of enclosure fire with PMMA as the fuel," (d) , 1986 . 2 Takeda, H.

, "Experimental investigation of PMMA compartment fire," (e)

,

No . 1 , 1985.

3 Takeda, H.

, "Wall fire behavior in the dual opening compartment," (e)

,

No. 1,2, 1986.

4 Nitta,K.,"A study on oscillatory fire," (c) , 1986.

5 Matsubara , Y.

, "A study on the oscillatory behavior in pool fires," (d) , 1986.

6 Satoh,K. , "Numerical study on the oscillatory behavior of fire plumes,"

(d) 1986.

7 Hasemi,Y. , "Deterministic modeling of flame-cored fire whirls," (a),

No. 356, 1985.

8 Takeno,K., Kitaoka.M. , and T . Hirano , "Flame spread along a ground soaked

with combustible liquid," (d)

,

1985.

9 Terai.T., and Kato , T ., "Calculation of the temperature rise at early stages

of enclosure fires," (a), 1985.

10 Terai.T., and Rato, T.

, "On calculation of floor temperature rise by

integral equations," (c)

,

1986.

11 Terai.T. , Private communication, 1987.

12 Katsuno,H., Torii.S., Kunimoto.Y., and Kawasaki, K.

, "Fire experiments using

the former Kuramae Kokugikan Smo Arena," (d)

,

1985.

13 Katsuno,H., Ishikawa,K., and Murakami, K.

, "Smoke behavior in a large

structure,", (d) , 1985.

14 Tanaka, T., and Yanama,T. , "Smoke control in large spaces, Part I, II," (c)

,

1985, or (e) , No . 1 , 1985..

15 Nagasawa.Y. , and Sakamoto ,K.

, "Scale model studies on smoke movement in an air-supported structures. Part I, II", (c)

,

1985.

16 Akeno,T. Tanaka, T. , Yamana.T., Takahashi,K. , and Nakamura ,K.

, "Smoke

venting experiments at Tsukuba Expo, pavilions. Part I-III," (c)

,

1986.

17 Hamada.T., Kato , T ., "Flow characteristics of flames along a surface

composed of vertical wall and ceiling," (d)

,

1985.

18 Yamanaka.K., Mizuno,T., and Kawagoe ,K.

, "A study on calorific property of

upholstered items," (c)

,

1986.

19 Yamada.T., and Jin,T. , "Experimental study of fire smoke mixing process in

a building," (c)

,

1986.

20 Hasemi.Y., and Yoshida ,M.

, "A simple estimation method of ignitability of

interior-lining materials," (c)

,

1986.

21 Mima,T., Ui,K. , and Ishii ,T.

, "Experimental study on smoke diffusion in

vertical laminar clean room," (c)

,

1986.

* The symbols (a)-(e) denote the journals or preprints as follows.

(a) : Transactions of Architectural Institute of Japan(AIJ), in Japanese.

(b) : Bulletine of Japanese Association of Fire Science and

Engineering (JAFSE) , in Japanese.

(c) : Proceedings of the Annual Meeting of AIJ, in Japanese.

(d) : Proceedings of the Annual Meeting of JAFSE, in Japanese.

(e) : Fire Science and Technology, in English.

7
SUMMARY REPORT

FIRE AND SMOKE PHYSICS SESSION

E. E. ZUKOSKI

Mechanical Engineering and Jet Propulsion

California Institute of Technology

Pasadena California 91125

ABSTRACT

A brief summary is given of some recent developments in the area of fire and

smoke physics.

1. INTRODUCTION

I have chosen a few topics which interest me and are relevant to fire and

smoke physics to review here. The selection of topics and papers is a rather personal one and certainly is not exhaustive.

2. FIRE-PLUME MODELING

The dependence of the length of diffusion flames on parameters such as fuel

flow rate and burner diameter indicate that there are three regimes of

behavior for these flames, see Zukoski (1985). These observations strongly

suggest that other properties of these flames will also differ in these

regimes and that the regimes should be used in discussing other data sets.

For example, when the the heat release per unit area of the fuel bed is very

small and the flame height is less than the burner diameter, regime I, the

flame is composed of a number of independent flame lets. The flame height is

less than the burner diameter. It appears to scale with the square of the

heat release rate per unit area and is independent of the total heat release

rate or burner diameter. For larger heat release rates per unit area,

regime III, the flamelets merge to form a single well-defined structure.

The flame height is larger than 3 burner diameters and scales with the total

heat release but is independent of the burner diameter. The important

parameter in fixing this transition appears to be the ratio of flame height

to burner diameter, Zf/D, and the transition occurs as this ratio increases

from 1 and 3.

In both of these regimes, the initial momentum flux of the fuel is

negligible compared with the change in momentum flux in the plume generated

by buoyancy forces. When the initial momentum flux is larger than the

buoyancy generated momentum flux in the flame, a third regime is observed in

which the flame length scales with the burner diameter and is independent of

the total heat release. Flame lengths in this regime are typically 50 to

200 burner diameters. The transition to this regime occurs when the flow

changes from a buoyancy controlled to a momentum controlled flow.

8
As the heat release increases, the transition between regimes I and' III is

completed when the heat release parameter Q* is near one. (Here Q* is

defined as [Q] / [ 0)0 Cpo To ( — /g D ) D~2 ]; and 0 ) 0 , To, Cdo are the

density, specific heat and temperature of the ambient fluid; D is the burner

diameter; g, the acceleration of gravity. The Q* parameter has been shown

to correlate flame length data for a wide range of fuels in regime III and

apparently also, in the transition regime II. Its usefulness as a correlation parameter for regime I has not been verified and it is not the

correct parameter to use in predicting the transition between regime III and

V.

The selection of the appropriate parameters to use in describing these

regimes and the clarification of the flow in regime I are important issues,

I believe.

3. INTERNAL STRUCTURE OF THE FIRE PLUME

Processes which occur within the fire plumes discussed above are very

complex and the high level of fluctuations and high temperatures make very

difficult the accurate measurement of velocity, composition, and

temperature. However, recent measurements have given us some additional

information concerning the flow in the flame itself. Interesting results

are reported by Miake-Lye and Toner (1986), and in a series of papers by

Gore and Faeth, e.g., Gore and Faeth (1986).

In the work of Miake-Lye and Toner, discussed in more detail later in this

meeting, a laser beam. was directed through a large diffusion flame and the

light scattered from soot was observed with photodiode array camera. The

diffusion flames studied covered the range 0.2 < Q* <3.0 and hence were in

regime II and the edge of III in the scheme discussed above. The presence

of soot was assumed to be a marker for a diffusion flame for this fuel-air

system. The most interesting result obtained in the study was that soot and

hence diffusion flames were observed to be present at points within the

outer boundary of the diffusion flame more than 40% of the time. Thus the

interior of the flame is a very complex region which contains geometrically

complex regions which are either fuel or air rich and which are separated by

diffusion flames. This result has interesting implications when the

radiation form flames is considered.

The experiments of Gore and Faeth clarify the influence of radiation on the

reactions in turbulent diffusion flames and suggest that some of the simple

models used to describe these flames are not applicable when radiation from

the soot removes a large part of the heat added by the combustion process.

4. CHEMICAL REACTIONS IN FLAMES

The chemical species formed in diffusion flames include the primary

constituents, water and carbon dioxide, but also a wide range of trace

species, some of which are toxic and hence of great interest in the study of

accidental fires. The most prevelant of these toxic species is carbon

monoxide which is found in the species produced in most accidental fires and

is toxic over short periods of exposure at levels greater than 0.1%.

Combust ion in Air : In a recent series of papers, Tewarson (1986) and DeRis

(1985) have published the results of a long series of tests on the

9
properties of a wide range of fuels. They report, among' other properties,

the species produced when the fuel is pyrolized in an inert atmosphere and

then burned in a laminar diffusion flame in air. The results show that the

fraction of the carbon in the fuel which becomes carbon monoxide or soot in

the combustion products varies very strongly with the chemistry of the fuel.

They propose a correlation for many of these products based on the soot

point for a laminar diffusion flame in which the products of pyrolysis are used as the fuel. These results can be used to predict the amounts of a wide range of species which are present in the products of diffusion flames

burning in air.

Combust ion in Two-Laver Svst ems : In many accidental fires, the supply of

oxygen can be greatly reduced in the upper part of the fire room. For

example, when a large diffusion flame is present in a room fire, a two-layer

situation may develop in which an upper layer contains the products of

combustion and the lower unvitiated air. As the fire develops, the upper

part of the flame may gradually extend into the upper v it iat ed- lay e r , due to

the growth of the fire and the thickness of the upper layer, while the

lower part of the flame will still be surrounded by unvitiated air. In

situations of this type, the concentration of oxygen in the upper layer may

be reduced to a small value and this almost inert gas will be strongly

entrained into the fire-plume which extends into the upper layer.

Under these circumstances, the chemical species formed by the fire may be

quite different from those found in the products of diffusion, flames burning

in unvitiated air. Studies of this situation were first made by Cetegen et

all (1984) for natural gas diffusion flames. A large hood was used to

collect the products of combustion and to model the upper layer. By varying

the spacing between the lower edge of the hood and the burner, the amount of

air entrained into the fire plume could be varied over a wide range and the

fuel-air ratio of the mixture entering the hood could easily be changed from

0.0 to 2.5 times the stoichiometric value.

Work with this apparatus was extended by L im (1984) and Miake Toner (1986),

and Beyler (1983) carried out similar studies of a wide range of fuels in a smaller hood. The interface in this apparatus was less well-defined and

mixing was not as complete.

In these experiments, several important results were observed. First , the

flame continues to bum in the vitiated gas contained within hood, even when

the separation between the burner and the lower edge of the hood was a small

fraction of the burner diameter. In this situation, no oxygen remained in

the upper layer and most of the flame was above the interface. The heat

released by the fire under these circumstances was determined from measured

values of the species present in the hood, and it was found to be no more

than ten percent less than the theoretical heat release. The theoretical

value was based on an equilibrium calculation for the flow of fuel and

entrained air into the fire plume and transported into upper layer. See

Toner ( 1986 ) .

Second . as long as there was more than 5% mole fraction of oxygen within the

hood, the products of combustion were close to their equilibrium values and were quite similar to those suggested by Tewarson. However, when the

mixture ratio for the gas entering the hood was above the stoichiometric

value, the oxygen mole fraction was reduced to less than 1% and

10
concentrations of the chemical species observed in the hood were far from

their equilibrium values for the mixture of elements involved and for the

gas temperatures observed.

The measured mole fractions of the species present in the products were well

correlated by the mixture ratio of fuel and air entering the upper layer in

the hood, see Beyler (1983), but were independent of the residence time of

the gas or the temperature of the gas within the upper layer. Observed

temperatures covered the range, 200 < T < 900 K, see Lim (1985) and Toner

(1986). These results suggest that the reaction in the flame are quenched

by the entrainment of inert gas into the flame at temperatures above 900 K.

The mole fractions of species observed in the upper layer are roughly

similar to those observed in samples withdrawn from points within the flame

zone of diffusion flames and averaged over long periods of time.

The observed mole fractions were found by Beyler to be a sensitive function

of the fuel chemistry. The concentration of carbon monoxide in the gas

within the upper layer were as high as 2 to 6Z when the mixture ratio

entering the hood was greater than 1.5 times the stoichiometric value.

These concentrations are high enough to be almost instantly lethal and hence

fires of the type discussed here are very potent sources of toxic gases.

Third, Beyler (1983) found that the products of reaction in the upper layer

would become flammable with air when the fuel-air ratios of the gas entering

the hood were high, e.g., around 2.2 times stoichiometric for methane. He was able to develop a model for prediction of the flammability limits for a' wide range of fuels.

As a final note, when the flammability limit was reached, burning started on

the interface and the products formed there were hotter than the gas within

the hood. This gravitationally unstable mass of hot gas was observed by Lim

to cause the gas within the hood to turn over rapidly. This process may be one of the causes for "f lashover .

Combust ion in V it iat ed Atmosphere : Products produced when the entire flame

is surrounded by a vitiated atmosphere are much closer to equilibrium values

and, because the flammability limit is reached when the concentration of

oxygen is reduced to values not much less than 10 mole percent, the

nonequil ibrium values of concentrations which were found in the two layer case are not observed here. A recent study of flammability limits, Lockwood

(1986), indicates that the flammability limit is better correlated by a limiting adiabatic flame temperature than- by the oxygen index method.

REFERENCES

Beyler, Craig Lynn, (1983) "Development and Burning of a Layer of Products

of Incomplete Combustion Generated by a Buoyant Diffusion Flame," PhD

thesis, Harvard University, Cambridge, Mass., Sept.

Cetegen, B. M. , Zukoski, E. E. , and Kubota, T., (1984) "Entrainment in the

Near and Far Field of Fire Plumes", Combust ion Sc ienc e and Technology . Vol.

39, pp. 305-331.

'

11
ae Ris , John, "A Scientific Approach to Flame Radiation and Material

Flammability," presented at the Eighth U.S. -Japan Panel on Fire Research and

Safety, Tsukuba, Japan, May 19-85.

Gore, J. P. and Faeth, G. M. , (1986) "Structure and Spectral Radiation

Properties of. Turbulent Ethylene/Air Diffusion Flames," The Twenty-First

Symposium (International) on Combustion . Munich, August.

Lim, Christopher, (1984) "I. Mixing in Doorway Flows, II. Entrainment in

Fire Plumes," AE Degree Thesis, CIT, June.

Miake-Lye, Richard C. and Toner, Stephen J., (1987) "Laser Soo t -Scat t er ing

Imaging of a Large Buoyant Diffusion Flame," Combustion and Flame 67: pp . 9-26 .

Tewarson, Archibald, (1986) "Prediction of Fire Properties of Materials Part

I: Aliphatic and Aromatic Hydrocarbons and Related Polymers," Technical

Report for MBS, Center for Fire Research, July.

Toner, Stephen J., (1986) "Entrainment, Chemistry and Structure of Fire

Plumes," PhD Degree Thesis, CIT, June.

Zukoski, E. E., (1985) "Fluid Dynamic Aspects of Room Fires,", Fire Safety

Science - Proceeding of the First International Symposium 1985 Howard W. Emmons Invited Lecture, pp . 1-30.

12
FORCED SMOLDER PROPAGATION IN A HORIZONTAL FUEL LAYER

AND THE TRANSITION TO FLAMING COMBUSTION*

Thomas J- Ohlemiller

Center for Fire Research

National Bureau of Standards

Gaithersburg, MD 20899

Abs tract

It is well known that a smoldering fuel responds to an increased oxygen supply

by becoming faster and hotter until, eventually, flames erupt. This sequence

is examined quantitatively for thick horizontal layers of a permeable fuel,

i.e., cellulosic insulation with air flowing over the top surface. Two

configurations are possible, forward and reverse smolder; both, are investigated experimentally. The influence of combustion retardants is also

examined; these include boric acid, a smolder retardant, and borax, a flaming

retardant. Both prevent the transition to flaming in the absence of adjacent

flammable material but are less effective in its presence. The overall

response of these various fuel mixtures and configurations suggests that both

kinetics and oxygen supply rate (not the latter alone) play substantial roles

in dictating smolder response to an air flow. Keywords: air flow, cellulosic

materials, flaming, insulation, retardants, smoldering.

1. INTRODUCTION

Cellulosic Insulation of the type investigated here is made from recycled

paper products, mostly newsprint. When properly ground to a finely fibrous

form, such products become an excellent thermal insulation. The inherent

combustibility of the base material (essentially wood) must be controlled,

however, by the addition of combustion retardants and by adherence to

installation practices that keep the insulation away from heat sources. When

heated sufficiently even retarded insulation may smolder (I).

Smoldering insulation becomes an immediate life threat when and if it

undergoes a transition into flaming combustion. For the usual attic type of

installation, this is most likely to be caused by air flowing over or through

the smoldering combustion zone as a result of winds impinging on the

structure. Such winds may directly penetrate various attic vents or induce

flow to/from rooms below the attic via ceiling holes around devices such as recessed light fixtures, vent fans, etc. It is clearly desirable to lessen

the tendency of a smoldering insulation material to transition into, flaming in

such a manner.

There have been two previous studies of smoldering horizontal fuel layers in

the presence of flowing air with some observations on the appearance of

flaming. Palmer [2] examined the behavior of wood and grass dusts in layers

up to 6 cm thick with air velocities up to 8 m/s; the configuration was quite

similar to that used here but the bulk densities were 2-3 times higher except

for the case of cork dust. The presence of flaming was noted but not examined

*Work sponsored by the Department of Energy, Building Thermal Envelope Systems

and Materials Program

13
in any decail. Leisch [3] examined Che behavior of grain kernel fragments,

grain and wood dust in 20 cm layers at air velocities up to 19 m/s. In this case the fuel layer was set into the bottom of the flow tunnel so that its top

surface was initially flush with the tunnel floor. Flames were noted, under

some conditions, to appear and to be trapped in the surface depression left by

the smoldering process.

The present study was undertaken to look at the impact of boch smolder and

flaming retardants on the forced flow smolder and flaming transition

process. Further details and discussion can be found in Ref. 11.

2. EXPERIMENTAL APPARATUS AND PROCEDURE

A flow tunnel, illustrated in Fig. 1, was constructed for the purpose of

examining the behavior of a smoldering insulation layer subjected to a flow of

air over its surface. The tunnel provides a straight test section 115 cm

long, 20.3 cm wide and 56 cm high; it is constructed from Marinite board. Air

is drawn through the tunnel by a centrifugal fan at the downstream end; the

speed of the fan can be varied to give flow velocities from 0.3 to 5.0 m/s.

The flow velocity is uniform over the central cross-section of the tunnel to

within ± 2%, as measured by an air velocity probe. Sidewall boundary layer

effects do not appear to influence the central 50% of the tunnel over the

majority of the length of the region where the insulation layer is placed.

The sample thus behaves, to a large degree, as though it is in a steady wind

blowing parallel to its top surface. The boundary layer on the top surface of

the sample grows from the tunnel inlet. From the peak flow velocity and the

longitudinal distance to the downstream end of the bed, one would infer that

the boundary layer is laminar in all cases. However, the fuel bed surface is

rough due to its particulate nature and to its uneven shrinkage during

smoldering; this introduces local separation in numerous places and, in

effect, causes a mildly turbulent boundary layer. This can be seen when smoke

is being generated by the smolder process.

The leading and trailing edges of the insulation bed are ramped at an angle of

35° to lessen flow separation effects.- The intersection of the upstream ramp

with the top surface is always smoothed and rounded before a test by pressing

down the insulation. The (Towns trearn intersection was not so treated and the

flow is always fully separated from there onward (downstream).

Flow velocity is measured continually during a test with a Sierra Instruments

Model 615 MHT air velocity meter*. The placement of the meter probe is ahead

of the sample bed where the tunnel cross-section has its full 56 cm height so

the flow velocity over the top of the bed is about 20% higher than that

recorded by the meter. The velocity values reported here are for the full

cross-section height. The local cross-section above the sample varies as it

smolders and shrinks; the tunnel was made 56 cm high to keep the corresponding

variation in flow velocity within 20%.

*Certain commercial equipment, instruments, or materials are identified in

this paper in order to adequately specify the experimental procedure. Such

identification does not imply recommendation or endorsement by the National

Bureau of Standards, nor does it imply that the materials or equipment

identified are necessarily the best available for the purpose.

14
Unrecarded insulation from a commercial manufacturer is fully conditioned to

50% relative humidity at 21-24°C. It is weighed out and then fluffed by hand

within a plastic bag to remove all clumps. It is packed in two stages, as uniformly as possible, into a pan that imposes the shape and size shown in

Fig. 1. The bulk density was the maximum achievable with hand packing,

0.08-0.10 g/cm , this spread of values results from a varying tendency of the

material to expand upward after it is removed from the packing pan. The same

range of bulk density for the organic portion of the insulation is used when

the material is retarded. The combustion retardants (boric acid and 10 mol

borax) were usually reagent grade but NF grade was used in a few tests. In

all cases the material is through 200 mesh. It is mixed with the insulation

in sufficient quantity to give a final product that is 15% by weight

retardant. Uniform dispersion is achieved by two hours of mixing on a roller

mill. Thereafter all procedures are the same as with the unrecarded

material. After placement in the flow tunnel, the vertical sides of the

insulation bed are in contact with either 1.27 cm Marinite board or 1.27 cm white pine board. The sample width is thus reduced to 17.8 cm. The bottom of

the sample is always in contact with 1.27 cm Marinite board. Thermocouples

(chromel/alumel in a 0.050 cm diameter stainless steel sheath) are inserted

laterally into the insulation bed so that the junction is at centerline middepth; three thermocouples are used, symmetrically placed about the sample

center with a nominal separation of 10 cm from one to the next.

The insulation layer can be ignited to smoldering on either end to produce

propagation in the same direction as the air flow (forward smolder) or in the

opposite direction (reverse smolder). The igniter is a 1.3 mm diameter

electrical heater formed into a zig-zag pattern 8 cm long x 13 cm wide so that

it uniformly -ignites the full width of the insulation bed. It is slipped

under the desired end of the bed, raised to 375°C and left in place for up to one hour. During ignition the air flow velocity through the flow tunnel is

0.15-0.30 m/s; it is raised to the desired test value (and subsequently held

constant) only after the smolder front on the top surface of the bed

approaches the top end of the ramp portion.

The progress of the smolder front is easily visible on the top of the bed

since it changes the insulation from a light gray to a dark black color. A

motor-driven, 35 mm camera, viewing the top of the bed through a window in the

top of the flow tunnel, monitors the smolder front progress by taking pictures

at precisely-timed intervals ([typically 10 min.). It also views ruled index

marks of 1.27 cm spacing on both sides of the bed. The photos are magnified

after a test and used to compute the smolder front velocity along the top of

the bed. Smolder velocity at mid-depth is inferred from the thermocouples via

their spacing and the time between 250°C isotherms.

3. RESULTS AND DISCUSSION

The smolder velocities given in the following figures are all derived from the

successive photos of the top surface of the fuel bed. The average value

derived from the three imbedded thermocouples generally agreed with this value

within 5 to 10% (and frequently closer). However, the two successive values

derived from the thermocouples almost always showed some acceleration of the

smolder front at the mid-depth of the bed. In contrast, the top surface

photos rarely showed any clear-cut acceleration or deceleration. Evidently

the smolder front was slowly changing shape as it moved through the bed; this

effect was not large enough to obscure any of the trends discussed here.

15
3.1 Behavior of Unretarded Insulation

Reverse smolder . Reverse smolder tests (propagation in direction opposite to

air flow) were run with two different batches of unretarded material. Figure

2 shows the measured smolder velocity as a function of air flow velocity.

There is no significant difference in smolder velocity behavior between these

two batches of material for this smolder mode. Both exhibit only the same

weak, acceleration of the smolder rate with increased air flow velocity; the

smolder velocity increases by 67% as the air velocity goes from 0 to 5 m/s.

The batch 1 material exhibited a type of localized flaming phenomenon also seen with this material in forward smolder; here it appeared at air flow

velocities greater Chan 4 m/s. At random intervals blue flames (CO burning)

appeared in residual char on top of the bed lasting up to 2 minutes and

spreading up to 10 cm through the residual char before extinguishing. They

never spread up to the leading edge of the smolder front on the top surface of

the bed where flaming based on pyrolysis products rather than CO could be

established. This flaming phenomenon was not viewed as a true transition from

smoldering to flaming.

Thus, no true transition occurred in reverse smolder over the range of air

velocities achievable with the present apparatus. For this reason, coupled

with the relatively weak, response of smolder velocity to air velocity in reverse smolder, it was concluded that the reverse smolder mode represents a distinctly lesser hazard than forward smolder (see below); it was not pursued

further with regard to combustion retardant effects.

Forward smolder . Figure 2 also shows the forward smolder behavior of the same

two batches of unretarded insulation material. Here the differing batches

exhibit a qualitatively similar smolder velocity response but there are significant quantitative differences. Most .of the quantitative differences in

forward smolder velocity between batch one and batch two material are probably

due to differences in bulk density; the batch two material was, in most cases,

packed to a bulk density about 20% greater than for batch one. From

stoichiometric considerations with an oxygen supply rate-limited process [1], one expects an increase in *bulk density to cause a comparable decrease in

smolder velocity. In addition, above a critical air flow velocity, batch 1 exhibited the CO flames in confined holes in the char, as discussed above;

batch 2 did not.

Flaming development from forward smolder . Both batches, at air flow

velocities between 2 and 2.5 m/s, exhibit a true transition to flaming combustion in that flames appear which move onto virgin fuel, consuming selfgenerated pyrolysis vapors. The CO flames in the char are not a necessary

precursor to the initiation of pyrolysis vapor flames since they did not occur with batch 2 material. The pyrolysis-vapor flames typically originated in a cavity-like depression or step formed just to the char side of the leading

edge of the smolder front on the top surface. Such a depression was formed

rather erratically, evidently by the combination of char shrinkage and

consumption. The interior surface of such cavities could be seen to glow just

prior to flaming ignition. Evidently the concave structure facilitated

locally higher temperatures due to decreased radiative losses; these structures may also have facilitated flaming ignition by allowing local build-up of

16
a flammable gas mixture and by providing a flame-holding region. The

complexity of these details implies that modeling of the transition process

will be difficult at best._

The pyrolysis-vapor flames originated at a local region (1-2 cm wide) along

the leading edge of the smolder front, spread laterally across the leading

edge quite quickly (f 1 sec), then spread downstream over the virgin portion

of the fuel bed. The spread is spatially erratic and transient. Since the

cellulosic insulation material is made ultimately from wood, it is a good char

former. Thus, a flame cannot be permanently stabilized on the exposed top

surface of the insulation bed; the build-up of only 2-3 mm of char slows the

evolution of pyrolysis vapors to the point where the flame blows off. The

flame thus at first spreads rapidly over a large area of the virgin fuel then

falls back to regions that provide a flame holder function. Slight depressions in the top surface of the virgin fuel area where flow separation occurs provide temporary anchor points until the char again becomes too thick (after

about lp2 minutes); the flame then dies but it frequently will have initiated

smoldering in the local region where it temporarily persisted. The longestlived flames (up to about 5 min.) are anchored at the original leading edge of

the smolder front where the smolder process provides pre-heated fuel and a supplementary source of pyrolysis vapors. Since the smolder process consumes fuel at a much slower rate than the flames, it too ultimately becomes inadequate as an anchor and the flames extinguish. The smolder process continues,

its geometry distorted by new initiations downstream; after a variable- time

interval, the whole process of flaming ignition, spread and extinction is

repeated with spread occurring only on areas of fuel not previously charred.

Mechanism of forward/reverse smolder differences. The differing responses of

forward and reverse smolder to increasing air flow pose interesting

mechanistic questions. It has not yet been possible to explore these in great

depth but some information has been obtained. It should be noted that Palmer

[2] also saw such differences though he did not pursue their cause.

Once' a smoldering process is past the initiation stage, it has generally been

accepted that its rate is controlled by the rate of oxygen supply [2,4,5,

6]. Here again there would seem to be room for such an interpretation. In

forward smolder the air flow impinges on the upward ramping surface of the

char; in reverse smolder the flow over the char is separated as is apparent

from the smoke behavior. One expects better oxygen transfer to the char in

the forward smolder case.

Mass transfer measurements . In view of the considerable surface roughness

here, it was decided that direct measurements of the local mass transfer rates

were desirable. The technique chosen, evaporation of a subliming solid

(camphor), has been used in the past mainly to infer local heat transfer rates

from local mass transfer rate measurements [7]. Unretarded cellulosic

insulation beds were smoldered in both forward and reverse configurations

( i . 9 5 m/s air flow velocity) and then quickly extinguished with nitrogen after

the top of the smolder wave progressed over about two-thirds of the length of

the fuel bed. Pre-weighed flakes of camphor (approximately 1 cm square,

1 1/2 - 2 V2 mm thick, with tape covering the bottom surface) were placed at

several lateral and longitudinal positions on the top surface of the char. A

given set of camphor flakes was exposed to a fixed air flow velocity over the

extinguished fuel bed surface for a period up to four hours; the weight loss

17
from each flake was Chen sufficient Co p,ermit a weighing accuracy of 1-2%.

From Che weight loss, Che flake area and Che vapor pressure of camphor, one can compute the -effective local mass transfer coefficient [7] .

The variation with position on the char surface is not very great (< 35% on

the forward smolder case from approximately 23 cm behind the leading edge of

the smolder front to approximately 6 cm in front of it). This variation was

not pursued extensively because it is small enough to be of secondary

importance. Of greatest interest is the behavior near the leading edge of the

smolder front; this has the greatest bearing on smolder propagation rate. The

difference there in mass transfer coefficients is only about 25-30% whereas

the smolder velocities can differ by a factor of two. In addition, smolder in

the two configurations shows- differing air velocity dependencies while the

dependency of the mass transfer coefficient is essentially independent of

configuration. Furthermore, despite the slightly larger coefficient in the

forward configuration, depletion of oxygen in the boundary layer over the

preceding char probably leads to a slower oxygen transport rate near the

leading edge in the forward case; the reverse case does not exhibit this

oxygen depletion. Thus, one is lead to conclude that oxygen transport alone

cannot account for the forward/reverse differences.

Probable source of the forward/reverse smolder differences . The leading edge

region of the smolder front is subject to the greatest heat losses. It loses

heat by conduction and radiation into surrounding virgin fuel; it also'

radiates from its top surface to the cold surroundings. In reverse smolder

this region also may suffer a net convective heat loss to the air flowing over

the surface, in spite of the flow separation. (Note that the convective heat

transfer coefficients will behave the same as the mass transfer

coefficients.) Certainly the virgin fuel just ahead of the leading edge of

the reverse smolder front (prior to flow separation) suffers a convective heat

loss because the air passing over it is at room temperature. In contrast,

forward smolder provides a configuration chat preheats the air flow as it

approaches the leading edge of the smolder front; one expects a net convective

heat* input to the char just upstream of the leading edge of the smolder front

as well as 'to the downstream virgin fuel not yet charred. In fact the top

surface temperature profile at 2 m/s shows thac in the forward smolder case,

the drying of water from the insulation is pushed several centimeters ahead of

the smolder reaction zone by the convective heat input. Drying is immediately

in front of the reaction zone for reverse smolder.

In view of these considerations and the observed smolder behavior, we hypothesize that the leading edge region of the smolder front is kineticallylimlted , particularly with regard to char oxidation, not oxvgen supply ratelimited (the remainder of the smolder front is oxygen supply rate-limited, as

is usually assumed). A kinetically-limi ted region will be highly responsive

to temperature (and additives, see below) and thus to the net heat flux; it

will be only weakly responsive to an enhanced oxygen flux.

Consider the forward smolder case first, in view of this hypothesis. An

increased air flow directly increases the heat transfer coefficient from the

char-heated air to the leading edge region causing some acceleration there.

Furthermore the increased air flow first impinges on the oxygen supply ratelimited portion of the char oxidation zone upstream of the leading edge

region. There it accelerates the char oxidation rate, raising the local

18
surface temperature and thereby heats the air further. This hotter air in

turn transfers more heat to the leading edge region enhancing its acceleration. This occurs in spite of the fact that the enhanced char oxidation

depletes more oxygen from the heated air in the boundary layer (which probably

overrides the increased mass transfer coefficient). If the leading edge of

the smolder front can thus move forward more and more rapidly due to the

forward transport of reaction heat by the gas, the remaining portion of the

smolder front (the oxygen supply limited portion) can keep up by adjusting its

shape (length) to assure that the oxygen supply rate per unit length of

smolder front is sufficient.

Accelerating char oxidation closer and closer to the leading edge of the

smolder zone puts a heat source of increasing strength and temperature in

closer proximity to the virgin fuel. The top surface temperature profile at

2 m/s shows that the forward smolder mode supports a temperature gradient

parallel to the surface that is 2 V2 x greater than in the reverse smolder

case. Ultimately the accelerated char oxidation at the leading edge leads to

the development of the concave structures, described above, which ignite

pyrolysis vapors to flaming.

The reverse smolder case might at first be expected to go slower with

increased air velocity on the basis of the preceding heat transfer considerations. In fact, Palmer [2] did find one material (cork dust) that smoldered

more slowly with increased air velocity above about 3 m/s in the reverse smolder configuration; many of his materials showed a constant reverse smolder

velocity with increased air flow velocity above about 1 to 2 m/s, though they

accelerated at lower air velocities. Here, Fig. 2 shows only a slow but

steady increase of reverse sraolde'r velocity. Evidently here the increased

convective heat loss from the leading edge region Is more than compensated by

other changes. The oxygen supply rate-limited region beyond the leading edge

(downstream) responds to the increased oxygen supply by getting hotter as in

the forward smolder case. The separated flow, with its attendant flow

reversal, helps move some of this heat up toward the leading edge. Conduction

and 'radiation through the fuel bed do so also. Furthermore, the leading edge

char zone is 50% thicker in the reverse case, lessening heat losses. The net

summation of the increased inputs of heat to the leading edge region is

evidently sufficient to outweigh the increased convective loss from the top

surface. The present arguments suggest that any model of the effects seen

must describe the gas phase above the fuel as well as the fuel bed itself,

since their coupling is a key "feature of the observed behavior.

It is interesting to note that the multi-dimensional nature of the smolder

propagation process studied here provides a new balance of processes

distinctly different than in one-dimensional smolder with similar materials

[8]. The overall results are considerably different: in one-dimensional

smolder, forward propagation is much slower than reverse propagation; here the

opposite is found to be true.

3.2 Behavior of Retarded Insulation

Both boric acid, a smolder inhibitor, and borax, a flaming inhibitor, were

tested separately for their influence on forced flow smolder (forward

configuration) and the transition process. Figure 3 summarizes the results.

It is apparent that they differ drastically in their influence on the smolder

19
rate. In the absence of any adjacent wood, they appear to be equally

effective in preventing flaming, i.e., neither permitted the appearance of

flaming at air velocities up to about 5m/s, the maximum value achievable with

the present apparatus. (In the presence of flammable wood sidewalls, much of

this effectiveness was lost; see Ref. 11.)

Thermal analysis (TG/DSC; 5°C/min in air) of the three material compositions

used here revealed that, as with pure cellulose (9), the retardants act mainly

on the char oxidation process. Boric acid substantially slows the rate of

heat release from this process but borax, a char oxidation promoter, appears

to have only weak, effect on this material, which already contains plenty of

catalytic metal ions (10, 11).

The striking differences in smolder velocities at equal air velocities in Fig.

3 are once again a demonstration that external oxygen supply rate alone does

not dictate smolder velocity. The preceding arguments about the importance of

the leading edge of the smolder zone again apply here. If it is kineticallylimited as hypothesized, it is sensitive to catalysts and inhibitors of the

char oxidation process. Thus boric acid slows the propagation by inhibiting

char oxidation in the leading edge region. By slowing down, the thermal wave

becomes thicker and the heat losses to the virgin fuel decrease. The char

oxidation process becomes less complete thereby insulating itself beneath the

char residue, cutting heat losses still further [1]. By covering itself with

such an insulating layer, the char oxidation zone interposes an additional

series resistance to oxygen transport that makes it less sensitive to changes

in external air flow velocity. Because of these interacting effects, quantitative prediction of the net effect of a given change in char oxidation rate,

for example, requires a detailed model of the entire two-dimensional system.

Borax, as noted, causes some acceleration of the char oxidation process.

The acceleration is small relative to the unretarded case and the smolder

‘velocities are only slightly greater. There is, however, no transition to

flaming here. The leading edge of the smolder front thus moves faster with

further air flow increases, becomes thinner and is able to tolerate the

increased heat losses that this implies. The remaining portion of the smolder

wave must adjust its shape to assure a sufficient oxygen supply rate so that

it can keep up. It has practically no "cover'* between the char reaction zone

and the external air flow (only a 1-2 mm quenched skin of char on the top of

the fuel bed) and so it is highly responsive to changes in the external air

flow velocity. .

The reaction zone length is stretched out in the borax case by about 40% to

about 37 cm from about 27 cm for boric acid. Thus not only does the borax

case assure its reaction zone a higher oxygen supply rate by remaining close

to the top surface of the fuel bed, it also stretches out the reaction zone at

a more shallow angle to facilitate a greater supply rate per unit length of

reaction front.

The failure of the transition to flaming to occur with either retardant is

presumably due to differing causes. For boric acid, the infrared TV viewing

the top surface makes it apparent that the surface is much cooler than the

unretarded (or borax) case. Thus it is probable that, for nearly all of the

air flow range, the temperature was insufficient to ignite a flammable mixture

of pyrolysis products even if it existed above the fuel bed surface. The

20
situation is evidently changing at flow velocities above 4.5 m/s since the

smolder velocity begins to accelerate; this is discussed somewhat in Ref. 11.

In view of the minor effect that borax has on the thermal analysis behavior,

its success in suppressing flaming is surprising. That is, it does not appear

to act by promoting char formation as it does with cellulose since the amount

of char in thermal analysis is about the same as for unretarded material. Its

mode of action is not presently known; it could be an indirect effect such as

a failure to facilitate formation of the Intermittent cavity-like depressions

at the leading edge of the smolder front (absence of these was noted with

borax present). Recall that these were the source of flaming ignition in the

unretarded case.

4. SUMMARY AND CONCLUSIONS

For a given insulation composition, there is a large difference in response

between. a smolder wave moving in the same direction as the accompanying air

flow (forward smolder) or in the opposite direction (reverse smolder).

Reverse smolder of an unretarded insulation showed only a weak increase in

velocity as the air flow increased; no transition to flaming was seen at air

velocities up to 5 m/s. Forward smolder of the same insulation was accelerated considerably more, and, at air velocities in the range 2 - 2 V2 tn/s

,

flames developed which spread over the uncharred material ahead of the smolder

zone. These differing responses are not due solely to differences in oxygen

transfer rate to the smoldering fuel. They are qualitatively consistent with

the idea that the leading edge of the smolder front is kinetically-limited and

is therefore sensitive to the differences in convective heat transfer implicit

in forward and reverse configurations. The overall smolder wave is thus

partially kinetically-controlled (leading edge region) and partially oxygen

transport-controlled (bulk of wave behind leading edge).

Borax and boric acid assert their influence, at least in part, through the

kinetically-controlled leading edge region by accelerating (borax) or slowing

(boric acid) the char oxidation kinetics. By slowing the char oxidation

kinetics to the point of incomplete char consumption, boric acid also manages

to slow the oxygen supply rate to the trailing portion of the smolder wave making it less responsive tb changes in air flow velocity. This partial

suppression of char oxidation is lost at sufficiently high air velocities and

a very hot smolder process is generated. Nevertheless, boric acid suppresses

flaming up to at least 5 m/s.' Borax does also, even though it somewhat

accelerates the smolder propagation process; Its mode of flaming suppression

in not clear.

Actual installations of cellulosic insulation almost invariably juxtapose

layers of the material with wooden structural supports. The wood, of course,

contains no retardants. When wooden sidewalls are Included in the forced-flow

smolder apparatus there is no significant impact on the critical velocity for

appearance of flaming with unretarded insulation. However, the benefits of

borax are completely lost; the glowing insulation smolder zone that does not

itself transition to flaming ignites the adjacent wood at essentially the same

air velocity as is seen with unretarded insulation. The benefits of boric

acid are reduced but not lost when wooden sidewalls are present; the critical

air flow velocity for appearance of flames drops from above 5 m/s to about

3-3/4 m/s. Since borax and boric acid are usually used together, it is

21
probable that the combination offers less margin of improvement than boric

acid alone. From this and the rather weak effect that boric acid has on increasing the minimum temperature for initiation of smoldering in ceilulosic

insulation, one sees that there is a continuing need for more effective

smolder retardants.

4. REFERENCES

[Ij Ohlemiller, T., "Smoldering Combustion Hazards of Thermal Insulation

Materials", National Bureau of Standards, NBSIR 81-2350, August 1981.

[2] Palmer, K. , "Smoldering Combustion in Dusts and Fibrous Materials",

Combust. Flame 1 (1957) 129.

[3] Leisch, S., "Smoldering Combustion in Horizontal Dust Layers", Ph.D.

Thesis, University of Michigan Dept, of Aerospace Engineering,

Nov. 1983.

[4] Moussa, N., Toong, T.-Y. and Garris, C., "Mechanism of Smoldering of

Ceilulosic Materials:, Sixteenth Symposium (International) on Combus tion , The Combustion Institute, Pittsburg (1977) p. 1447 .

[5] Summerfield, M. , Ohlemiller, T. and Sandusky, H., "A Thermophysical

Mathematical Model of Steady-Draw Smoking and Predictions of Overall

Cigarette Behavior", Combust. Flame 33 (1978) p. 263.

[6] Ohlemiller, T., Bellan, J. and Rogers, F., "A Model of Smoldering

Combustion Applied to Flexible Polyurethane Foams", Combust. Flame 36

(1979) p. 197.

[7] Eckert, E. and Goldstein, R., Measurements in Heat Transfer , 2nd ed.

,

McGraw-Hill (Hemisphere Publishing Corp.), Washington, DC (1976)

p. 407 ff.

[8] Ohlemiller, T. and Lucca, D., "An Experimental Comparison of Forward and

Reverse Smolder Propagation in Permeable Fuel Beds", Combust. Flame 5w

(1983) 131-147.

[9] Shafizadeh, F., Bradbury, A., De Groot, W. and Aanerud, T., “Role of

Inorganic Additives in the Smoldering Combustion of Cotton Cellulose",

Ind. Eng. Chem. , Prod. Res. Dev. 21 (1982) p. 97.

[10] McCarter, R., "Smoldering Combustion of Wood Fibers: Cause and

Prevention", J. Fire Flarnm. 9 (Jan. 1978) p. 119.

[11] Ohlemiller, T., "Forced Smolder Propagation and the Transition to

Flaming in Ceilulosic Insulation", National Bureau of Standards NBSIR

85-3212, Oct. 1985.

22
SMOLDER

VELOCITY

(cm/s)x10

IR TV

Air camera

(forward smolder) (reverse smolder)

Flgura 1 Schema c ic of flow tunne 1 apparatus for acudy of smolder propagation

with flow ov«r the fuel bad. Flow curmal cxoaa- section la 20.3 cm wide by 56 cm hign.

12

n 10

UNRETARDED INSULATION

More frequent

flames in char

\ n

'Pyrolysis flames'

V rfc'

/ A_~ZZ- 'Pyrolysis flames'

/ /w

More

frequent . flames

a, Forward smolder

'

' n char

o,e Reverse smolder

Open, dosed symbols = different batches

Subscript w = wood sidewalls

i i i i :

1 2 3 4 5

AIR FLOW VELOCITY (m/s)

Flgura 2 Measured ef fact of air vmlociey on aeolder velocity of air flowing over the fuel bed In both forward and reverse configuration*.

AIR FLOW VELOCITY (m/s)

Flgura 3 Measured effect of air velocity on smolder velocity of cvo

coaroosi Clone wich additives; all results are for forward smolder

mode. Flames appeared only when wood sidewalls were present.

23
Wood Ignition and Pyrolysis

ARVIND ATREYA

Mechanical Engineering Department

Michigan State University

and

HOWARD W. EMMONS

Division of Applied Sciences

Harvard University

ABSTRACT

In this study piloted ignition and pyrolysis of wood are investigated both

experimentally and theoretically with the intention of understanding the various

chemical and physical processes that are involved. Accurate measurements of

surface temperature and evolved mass flux were made under well controlled

external radiation conditions and time for sustained ignition was noted. It was found that even small quantities of adsorbed moisture; (4% dry weight) has a very noticeable effect on both evolved mass flux and heat flux. Good agreement

between the measured and calculated mass flux was obtained by using surface

temperature instead of incident heat flux as the boundary condition.

Discrepancies between the measured and calculated incident heat flux was found

to be due to exothermic gas phase reactions with oxygen. These exothermic gas

phase reactions eventually lead to spontaneous ignition. The effects of various

terms in the energy equation were also evaluated by comparison of "the

theoretical model with the experimental results. It was found that the

controversial heat of pyrolysis is small and can be neglected. Internal

convective heat transfer by volatiles and variation of thermal properties with

temperature were also found to be small effects. Piloted ignition was found to occur in the kinetically controlled pyrolysis regime. A simplied formula for

fuel mass flux was also developed.

1. INTRODUCTION

Wood and other cellulosic materials, such as cotton and paper products,

form the bulk of the fuels in many fires. These materials, when heated to a sufficiently high temperature, undergo thermal decomposition or pyrolysis to

produce combustible gases and char. These combustible gases mix with the

surrounding air to produce a flammable mixture which ignites either

spontaneously or with the help of an external flame or a spark. Once a flame is

established in the gas-phase, further thermal decomposition of the solid occurs

to produce the fuel which feeds the flame. The fuel production rate under

flaming conditions is often called the burning rate. Thus, study of pyrolysis

of cellulosic materials is of considerable importance for determining both the

critical conditions at ignition and the subsequent burning rate. This paper

presents an experimental and theoretical investigation into pyrolysis and

ignition of wood.

Pyrolysis of thick samples of wood involves both the physics of heat and

mass transfer and the kinetics of thermal decomposition. Studies of

decomposition chemistry have been performed on thermally thin samples [1-6] to

avoid complications caused by heat and mass transfer processes. .Kinetic

parameters for overall apparent decomposition reactions have been obtained from

these studies. For thermally thick samples of wood, heat and mass transfer

processes have also been studied with varying degrees of sophistication.

Theoretical models from purely thermal [7-8] (with infinite rate decomposition

kinetics) to fairly detailed numerical [9-12] (with finite rate decomposition

kinetics) are available in the literature. Experimentally, mass evolution rate

24
and temperatures at various locations in the thick solid have been measured for

different incident heat fluxes. Some investigators [13-15] have also measured

density profiles inside the solid using radiographic techniques.

Despite this substantial volume of information, large discrepancies in the

interpretation of data on pyrolysis of wood still exist. The best example of

these discrepancies is seen in the reported values of the heat of decomposition,

which is usually determined by comparing a theoretical model with the experimental results. The reported heats of thermal decomposition vary greatly from

significantly endothermic [-180 cal/gm; Kung's [10] Re-analysis of Bamford et

al's [9] data] to significantly exothermic [+444 cal/gm; Shivadev and

Emmons [3]. Some variation in this global property is to be expected because

the composition of the products of pyrolysis is not fixed. But, a variation as wide as reported seems due to our lack of understanding. The net result is that

the fuel production rate (required for both ignition and burning rate) as a function of heat input rate, cannoc, as yet, be reliably predicted.

2. EXPERIMENTS

In an attempt to understand the pyrolysis and ignition process, well

controlled small scale experiments were performed on horizontal samples of

various woods. The samples (8 cm x 8 cm x 1.9 cm) were pre-dried in an oven at

105 °C for 5 to 8 hours and then allowed to cool over anhydrous CaSO^ in a desiccator. These experiments were conducted in air using the apparatus shown in

Fig. 1. The water-cooled plate was suddenly removed, exposing the instrumented

sample to a constant incident heat flux. Incident heat flux was also

continuously monitored by using an off-center radiometer calibrated against the

heat flux at the center. The temperature of the electrical radiant heater, as measured by an infrared pyrometer, was about 1200K, which is representative of

fire radiation [16],

For the pyrolysis experiments, front and back surface temperatures and the

weight loss were measured as a function of time. The lb pilot flame and the 4

RPM motor were not used during these experiments. For the ignition experiments,

a small (2 mm dia.) pilot flame was periodically lowered to within a millimeter

of the sample surface and then quickly removed. The time required for sustained

ignition was noted and for some cases surface temperature was also measured.

The sample was slowly rotated (at 4 RPM) to avoid overheating a single spot on

the wood surface by repeated trials

.

In these experiments, weight loss of the sample was measured by a sensitive

weighing table (resolution + 0.001 gm) and surface temperature by a specially

fabricated film thermocuple in order to increase the contact area (see Ref. 17

for further details). Temperature measurements with the film thermocouple were

very repeatable (within + 3°C) and the physical contact remained excellent even

during charring. The data was collected by a microcomputer and the mass flux

was obtained by a polynomial spline fit to the weight loss data. Several

experiments, at different levels of incident heat flux (1.5 to 4.0 tf/cmy , were

conducted for each kind of wood for both ignition and pyrolysis. The results of

three pyrolysis experiments and corresponding numerical calculations (to be

described later) are shown in Figures 2a, 2b, and 2c for maple and Figures 3a,

3b, and 3c for vertical grain white pine[_VGPj. Results of other pyrolysis

experiments are summarized elsewhere [18]. The results of piloted ignition

experiments are shown in Figures 4 and 5.

The mass flux curves shown in Figures 2a and 3a do not start at the origin.

This is partly due to the disturbance in weight measurement caused by a sudden

removal of the cooling plate and primarily due to the desorption of absorbed

moisture, because no appreciable decomposition of wood is expected to start at

surface temperatures lower than 540K. The high heat flux tests [#8 in Fig. 2a

25
and #7 in Fig. 3a] had to be terminated due to the occurrence of spontaneous

ignition. These tests were also terminated at the first visible signs of

surface combustion or glowing. Often none or very little ash was observed on

the surface of the sample at the end of the test, although considerable cracking

of char had occurred.

The mass flux at ignition, showmi in Figure 5, was calculated using measured

surface temperature for a mo is ture - f ree slab. The model used for these

calculations is described below.

3 . THEORETICAL MODELS OF ^OOD PYROLYSIS

The data obtained in this study, provides a useful basis for developing a predictive model for pyrolysis of wood. The existing models of wood pyrolysis

[9-12] have numerically explored the effect of various parameters and terms of

the equations that were used, but only a few comparisons were made with the

experimental results. In the present study, data from at least eight different

experiments for each wood is compared with the model to evaluate the effect of

various parameters. An attempt is made to understand the effect of various

chemical and physical processes occurring during the pyrolysis of wood with the

help of experimental data and mathematical models.

3 . 1 Decomposition Kinetics

The experimental data of Figures 2a and 3a clearly shows that the mass flux

initially increases, reaches a maximum and then slowly decreases with time. A

qualitative comparison with purely thermal pyrolysis temperature models [7,8]

(which predict that for large times the mass flux decays as the square root of

time) suggest that diffusion of heat is the controlling mechanism for large

times. This is because heat has to diffuse through the thick char layer to

reach the reaction zone. However, in the initial stages of pyrolysis the

exposed surface of the solid produces most of the volatile mass flux. Thus,

surface boundary condition and decomposition kinetics are expected to play a dominant role for small times. Therefore, for a theoretical model to be valid

for both large and small times it must include both heat and mass transfer and

decomposition kinetics. A relevant question is: How much chemical detail needs

to be included to explain the observed behavior? Due to the complex

decomposition chemistry, most kinetic studies ( on both thermally thin and thick

samples of wood) have been limited to the determination of the activation energy

(E) and frequency (A) , in the following single overall decomposition reaction:

3pC O

- -(P

s (t) - p f

) A exp ( - E/RT ( t) ) , (1)

The values of E and A obtained by various authors from experiments on thermally thin samples of wood or cellulose fall into two distinct groups. Theg

first group can be approximately represented by E = 30 Kcals/gm mole and A s 10

/sec, and the second group by E - 54 Kcals/gm mole and A - 10^/sec. A closer

examination of the reported results show that experiments for the first group

were conducted in air and for the second groups in an inert atmosphere. This is

consistent with Shafizadeh [6] work which confirmed that pyrolysis proceeds

faster in air than in inert atmosphere.

Kinetic parameters obtained from thermally thick samples seem to agree with

those obtained from small samples in air, even though some experiments [15,19] were conducted in nitrogen flow. This apparent discrepancy is resolved by

Lewellen et al ' s [20] results for pyrolysis of a-cellulose in helium. They

found that reaction rates for a heating rate of 400°C/sec at 1 atm pressure were higher than those at 10,000°C/sec at 0.0005 atm pressure. Thus, greater the

26
residence time of the pyrolysis products in the cellulose matrix, higher is the

reaction rate. This suggests that increased residence time of the pyrolysis

products in thick samples .enhances the reaction rate in a manner perhaps similar

to that by oxygen in small samples. Lewellen et al [20] also showed that char

is not a primary product of the pyrolysis of cellulose. In their experiments,

there was not leftover char. Results of other investigators [6,21,22] show that

char yield depends on the temperature exposure history of the sample. It

increase with decrease in decomposition temperature and also with the duration

of sample exposure to lower temperatures. Hence, char yield would vary with

depth for thick sam-ples . To account for variation in char yield tiadorsky [23],

Broido [24] and Shafizadeh [6] have suggested a competing two reaction mechanism

(dehydration and depolymerization) for cellulose pyrolysis based on the results

of experiments on small samples.

All these results can be combined into a likely decomposition mechanism

shown in Fig. 6. In large samples, it is possible that levoglucosan pyrolyzes

as it travels through the hot char matrix [12,15], Since, the chemical structure and products of pyrolysis of levoglucosan are very similar to cellulose,

steps B and C of Fig. 6 are probably very similar. This would explain why large

samples showed a behavior similar to that of small samples pyrolyzing in air.

Thus for large samples, char- forming (or dehydration) reactions are expected to

dominate. The extent to which non-char- forming (or depolymerization) reactions

occur in real fire situations, however, is not negligible. This is evident from

the variable char yield obtained by Atreya [18] and Lee, et. al. [15] for wood.

Thus, for correctness, the likely overall decomposition scheme suggested in

Figure 6 should be used. However, due to the lack of appropriate kinetic data,

Equation (1) is used and char yield (or p is assumed to be a constant. As can

be seen from Table 1, for vertical grain pine (VG P) , two different char yields

had to be assumed to obtain agreement with the experimental results. Higher

char yield for lower temperature ' experiments is consistent with the work on

small samples [22].

3 .

Desorption of Moisture

Another noticeable effect in the mass flux data shown in Figures 2a and 3a

is that it quickly rises to about 0.2 mg/ cm sec and that the duration of its

stay at this level decreases with increased incident heat flux. It was found

that this is due to desorption of moisture.

Moisture is held in wood in three ways [25]: (i) Free or adsorbed water

present in the cell cavities. (ii) Adsorbed moisture, which can be up to 30% by

weight and is a few molecules thick layer held by hydrogen bonds. This layer is

monomolecular at about 5% adsorbed moisture. Wood under normal conditions

contains about 12% moisture. (iii) Water of constitution, which is a part of

the cellulose molecule. When wood is heated different types of moisture are released gradually with no distinction between them. The term "dry wood" is an arbitrarily defined standard (such as 24 hours at 105 °C over phosphorous pentaoxide) . Theoretically, wood may be considered "dry" when all the adsorbed

moisture and none of the water of constitution is extracted. Experimentally, it

is not possible to extract all the adsorbed moisture without thermally degrading

wood. 'Thus even though the samples used for the experiments were predried, some moisture flux is to be expected. This will significantly affect the ignition

process

.

In a theoretical formulation, moisture which is not a part of the molecule

(i.e., adsorbed moisture) must be considered as a separate specie because the

temperature dependence of its production rate is different from the rest of the

pyrolysis products. This is evident from Figures 2a and 3a. Thus, in the

27
present formulation, the rate of desorption is assumed to follow a first order

Arrhenius expression, viz.,

dp

~Z " -A p exp ( - E /RT) . (2)

a t mm m

The kinetic constants 'E ' and 'A ' along with the heat of desorption 'Q ' m m ° m were determined by best fit of this theoretical model to data from separate low

heat flux experiments (details are given in Ref. 18). Surface temperatures was always below 450 K to assure that only moisture desorption and no decomposition

occurred. Values of these constants, so obtained, are listed in Table 1. The

only remaining unknown in Equation (2) is the absolute value of the initial

concentration of adsorbed mo-isture. Since there is no way of determining this

value, it had to be assumed. Due to similar conditioning of all the samples

prior to the test, a fixed value was assumed for a kind of wood.

3 . 3 Conservation of Mass

If all the volatiles that are generated by thermal degradation and

desorption of moisture are convected out, then conservation of mass requires:

3M dp dp

g s m 3x TM 3t

+

3t (3)

were 'x' is the depth inside the solid, with x-0 as the top and x-L as the

bottom surface of the solid. M is the outward mass flux of volatiles, being

g

positive in the negative x-direction.

3 . 4 Conservation of Energy

Equations (1) to (3) are considered essential in order to explain the

experimental results shown in Figures 2a and 3a. The extent of detail required

for energy balance cannot, similarly, be determined by inspection. Thus,

ignoring kinetic energy of volatiles and work done by gravity and pressure,

energy balance is written as:

(pc

a pa

. 3T _d f _ _ >. 3T pa I + p c + p c )

- —

c pc m pm 3t 3x k —

( s3x J + l^J M

g

dp

s 3

• f Pf ] r , p

f ^ -Q + H - H — r

- H

< a c pr w J i p w J gj

(4)

3

( ^m

+ H V

The symbols are defined in the nomenclature.

In Equ. (4) , the left hand side is the increase in enthalpy of unpyrolyzed

active material, char, and moisture in the elemental volume. Terms on the right

hand side are: (i) influx of energy due to conduction, (ii) energy convected

out by volatile gases assumed to be in good thermal contact with the solid,

(iii) net heat required to produce a unit mass of pyrolysis gases (positive when

endothermic) , and ( iv) net heat required to evaporate a unit mass of moisture.

Here p and p are instantaneous densities of unpyro ly zed ac tive material and

char. At any instant, p (x,t)-p (x,t) + pc S3. (x,t), hence at t-0 , p - p and

* S ^

28 .
For the fixed char yield (~p^/p^> p

s ac t— is linearly related Co , p- p c

. U

hy '

f

s (t ' x) P a

(C-,x) + p f (5)

To simulate energetic effects due to cracking of char, convective heat

transfer between the volatiles and the solid matrix is not permitted for those

elemental slabs where p -p _ (i.e. o -0) and is completely retained for those

s

' t ‘a elemental slabs where p -p -p . For intermediate elements convective heat

s w a transfer is gradually turned off.

Following Tamanini [11], the thermal conductivity (k ) was assumed to be

linearly related to density as,

k

s p a w

( 6 )

where k and k were treated as linear functions of temperature [k - k

3 +

3. C T )1 and k was assumed to be a constant. Also, specific heat (C

ao m p assumed to linearlyJ varyJ with temperature as: C (T) - C°+ C (T-T ).

p p p

aO property values for maple and VGP are summarized in Table 1.

*k (T-

) was

The

3 . 5 Numerical Calculations and Boundary Conditions

The Equations (1), (2), (3) and (4) along with variable properties defined

by Eqs . (5) and (6) were numerically solved by using the Crank-Nicolson method

to express the finite difference form of the derivative's (details are given in

Ref. 18) Tamanini' s [11] computer code (SPYVAP) was modified to include

energetics, kinetics, and changes in physical properties due to desorption of

moisture, to simulate cracking, and to include the blowing correction in the

convective heat transfer coefficient.

Finally boundary conditions must be carefully selected for comparison with

experiments Previous investigators [10,11,12] have used the experimentally

prescribed incident heat flux as the boundary condition for the exposed surface.

This, however, results in interpretation of theoretical results due to: (i)

spectral absorptance of the solid surface relative to the spectral quality of

the incident radiation [27], (ii) changes in surface radiative properties during

decomposition, (iii) absorption of incident radiation by the decomposition

products [28], and (iv) exothermic surface reactions. The net effect is that

the boundary conditions no longer correspond to the experiments. These

uncertainties were eliminated in this work by using the measured surface

temperatures (Fig. 2c and 3c) instead of incident heat flux. Thus, only the

solid phase decomposition process was modeled. Incident heat flux required to

obtain the measure temperature was also calculated and compared with the

experimental value. Figures 2a and 3a show a comparison between measured and

calculated mass flux for three different experiment on maple and VGP and Figures

2b and 3b show the corresponding comparison for incident heat flux. In these

figures 'dots' are measured values and solid lines are the results of- numerical

calculations. The values of 'E' and 'A' (see Table 1) were chosen to obtain the

best fit between the theory and the data for all eight experiments. The best

29
fit of 'E' and 'A' are in good agreement with those reported in the literature

[ 2 ].

Similar calculations were performed to determine the critical fuel mass

flux at ignition. Measured surface temperatures (Fig. 4) were used as boundary

conditions and the time required for sustained flaming was plotted on the

calculated mass flux. The results of these calculations are shown in Figure 5. Note that the average fuel mass flux at ignition is in good agreement with the

literature [9] value of 0.25 mg/cm sec. Also note that moisture was not

included in these calculations for purposes of comparison with literature

values. Moisture content for these experiments was low (estimated - 4% by

weight). However, higher moisture content is expected to have a significant

effect on both time to ignition and mass flux at ignition.

4. DISCUSSION OF RESULTS

The agreement between the calculated and measured mass flux (Figures 2a and

3a) seems reasonable given the complexities of the process. The measured and

calculated incident heat fluxes (Figures 2b and 3b) do not, however, agree

equally well over the entire domain. This implies that the production rate of

fuel can be correctly predicted if the experimental surface temperature history

is used. In other words, the solid phase decomposition process is correctly

modeled. Hence, the effect of various terms in the energy equation (Equ. 4) can now be evaluated.

The third term on the RHS of Equ. 4 is the net heat required to produce a unit mass of pyrolysis gases. Ignoring the temperature dependence of the

specific heats for the purposes of discussion this term may be expressed as:

- Pr ]

o • f P -f 1 r p ? ^ c° - c°

L

- C 1-— »

— • 1

H H \

'

pw pc L p w J Pg l P w J

ao q Q

For representative values of the specific heats (C° - 0.33; C° -0.17 and C°

pw pc pg

- 0.26 cal/gm) , the quantity on. the left hand side is always positive and for a constant char yield of 21% it is 0.113 cal/gm°C. Hence, this term is the

difference between the heat of pyrolysis (Q) at the reference temperature (T ) and the stored energy released when a unit mass of pyrolysis gases are

generated. Figure 7 shows that the net energy required decreases with increase

in temperature. The numerical values in Fig. 7 were calculated by using a dynamic TGA curve from Browne and Tang [21] . Thus, for Q

- 30 cal/gm

(endothermic) pyrolysis at temperatures greater than 285°C would appear net

exothermic. For Q

- +60 cal/gm, the crossover from net endothermic to net

exothermic takes place at temperatures above 500°C. For the TGA curve shown,

pyrolysis is essentially complete at temperatures below 450°C, so that the

entire pyrolysis process would appear endothermic.

It is interesting to note (for Q

- +40 cal/gm, crossover at 372°C) that as the heating rate increases, a greater portion of decomposition would take place

at higher temperatures. This would tend to make the decomposition process net

exothermic. On the other hand, increasing the heating rate would reduce the

char yield, resulting in a lower slope of the energy release line. This would

shift the point crossover from endothermic to exothermic towards higher

temperatures. Hence, a compensating effect. From this discussion it is evident

that it is not possible to explain the large endothermic or exothermic effects

reported in the literature. Furthermore, it can be mathematically shown [12,18]

that a large exothermic heat of pyrolysis (>250 cal/gm) would lead to a thermal

runaway; a phenomena that has not been observed.

30 •
In an attempC Co resolve Che concroversy abouc Che magniCude and sign of

Che heaC of pyrolysis of wood, calculaCions for Q

- 30, 0.0 and -30 cal/gm were

compared wich che daCa for all eighc experimencs on maple and VGP. Ic was found

ChaC Che besC overall ag'reemenC was obcained by using 0.0 cal/gm. Figures 2a

and 2b show such a comparison for maple and Figures 3a and 3b for VGP. Thus ic seems ChaC che concrovers ial heac of pyrolysis is Coo small and can be ignored.

This considerably simplifies che energy equacion. The enCire Cerm mulciplying

3p /3t in Equ. 4 can be neglecced, because Che heac of pyrolysis 'Q' is small

and Che ocher Cerms are of comparable magnicudes and opposice sign (Fig. 7).

Concrary Co che heac of pyrolysis, desorpcion of even small quancicies of

moiscure has a very nociceable effecc. This is represenCed by che -las c Cerm in

Equ. 4. Figures 2a and 2b show che effecC of small amounts of moiscure for

ExperimenC #1 on maple. The dashed lines are Che resulc of calculacion for 0%

moiscure. The corresponding calculacion wich 4% moiscure is shown by che solid

line. For expc. #1 , using 0% moiscure resulced in poor agreemenc becween Che

calculaced and measured mass flux up Co 200 sec. and Che corresponding effecc on

2 Che heac flux was abouC 0.4W/cm .

The second Cerm on Che RHS of Equ. 4 is che heac carried away by volaciles

while Graveling chrough Che solid. For Che calculaCions shown by solid lines in

Figures 8a and 8b, Chis Cerm was sec equal Co zero. The resulcs indicaCe ChaC

Che heaC exchange becween Che volaciles and Che hoc solid macrix is noC a major

effecc and is furcher diminished by che loss of incimace conCacc due Co che

cracking of char.

The effecc of CemperaCure dependence of Che chermal conduccivicy and che

specific heac is shown by dashed lines in Figures 8a and 8b. The agreemenc wich

Che daCa was Che worse for large Cimes (or Cemperacures) . It is interescing Co

noCe that Che temperature dependence of thermal properties and the heat transfer

to volatiles have an opposite effecc on volatile mass flux, though the magnitude

of the former is somewhat greater. For the case when heat is not convected out

by the volatiles, more heat is retained in the interior to generate fuel gases.

For the case when thermal properties do not increase with temperature , less heat

is conducted into the solid thus reducing the evolved mass flux. The two

effects are therefore mutually compensating.

The next most noticeable result is that the agreement between calculated

and measured mass flux is quite resonable, but the heat flux required to obtain

the measured surface temperature (which was used as the boundary condition)

agrees well up to a certain time and then deviates considerably from the

measured values (Fig. 3b.). This suggest that the theoretical modeling has

ignored some very important phenomena which must occur at the surface. A closer

examination of the experimental data shows that (i) the deviation in the

required heat flux occurs slightly before the peak of the mass flux curves (Fig.

3a and 3b), (ii) the inflection points in the measured surface temperature

curves (Fig. 3c) correspond to the deviation in the required heat flux, (iii)

for high heat flux cases (expt. #8 of Fig. 2a and expt. #7 of Fig. 3a),

spontaneous ignition occurred at about this temperature. These observations

suggest that the gas phase reactions, that eventually lead to spontaneous

ignition, have started at this instant. For low heat flux cases, thermal

runaway to spontaneous ignition did not occur because of heat loss to the

surface and due to the attenuation in the fuel supply rate by the formation of

char. Thus, if exothermic gas phase reactions that supply additional heat to

the fuel surface do indeed take place, then the gas temperature slightly above

the surface must be higher. Figures 9 confirms this hypothesis. Other

investigators [29] have also reported similar inflection points in their surface

temperature measurements. To further confirm this hypothesis the experiments

need to be repeated in an inert atmosphere.

31
From these results it is very tempting to conclude that a large variation

(endo- exothermic) observed in the heat of pyrolysis of wood may be either due to

the desorption of moisture or to the exothermic gas phase reactions with oxygen

and that the' actual heat of pyrolysis is quite small and can be ignored. Note

that if incident heat flux rather than surface temperature was used as a boundary condition, then the only way to account for additional heat would be by

assuming exothermic heat of pyrolysis.

These results are also in agreement with the decomposition kinetics

suggested in Fig. 6. Depolymerization reaction A in Fig. 6 is due to scission

of CO bonds and is highly endothermic [23], Dehydration reactions B and C are not expected to be exothermic, except in the presence of oxygen. Hence, it is

likely that the observed high exo the rmic i ty is due to the oxidation of the

pyrolysis products. As described earlier (Fig. 7), physical arguments based on

temperature dependency of the enthalpy of the pyrolysis products can not account

for the large endo-exothermic effects reported in the literature.

Finally, from Figure 5 we note that piloted ignition occurs on the initial

rise of the mass flux curves -- a domain where surface boundary condition and

decomposition kinetics play a dominant role. The heat flux boundary condition

at the surface is given by:

k

^

3x x-0 <7 (T

4

-T

4

)

- h (T -T ) = f, (7)

5 co s aj

where f is the flux into the solid.

From Equations

density (i.e. p = ^w}

(1) and (3) assuming 0% moisture and small changes in

we obtain the fuel mass flux as:

rL

m" - A (P w

-P f

) exp(-E/RT) dx ( 8 )

Since, a thin surface layer produces this mass flux, T(x,t) may be approximated

by:

T (x , t) - T

s (t)

xf (t)

kT (9)

Substituting Equ. (9) in Equ. (8), integrating, and taking the lowest order term

of the asymptotic expansion for large E, we obtain:

m" - A (P w

-P f) (10)

Where, the upper limit of integration, L, was replaced by: (T -T^) k/f. Using

Equ. (10), volatile mass flux can be calculated from the knowledge of measured

surface temperature shown in Figure 4. Results of such calculations for

2 incident heat flux of 2.84 W/cm are shown in Figure 5 by the dashed curve. Given the approximations used in deriving Equ. (10), the agreement with a more

detailed calculation is excellent. Thus, Equ. (10) provides a useful

relationship between the critical surface temperature and critical fuel mass

flux at ignition.

32
5. CONCLUSIONS

In this study, pyrolysis and ignition of wood was investigated both

experimentally and theoretically. A phenomenological picture that emerges from

this work is shown in Fig. 10. The heat required for the decomposition reaction

was found to be small and can be ignored in the energy balance (Equ. 4) . On the

other hand, desorption of even small quantities of moisture was found to

significantly effect the energetics of the pyrolysis process. The experiments

also demonstrate the existence of exothermic gas phase reactions when pyrolysis

occurs in the presence of oxygen. These exothermic reactions eventually lead to

spontaneous ignition. However, more experiments under well controlled

atmosphere are needed to rule out the possibility of simultaneous exothermic

surface reactions.

The theoretical model developed for pyrolysis of wood with moisture,

successfully explains the experimental results. This model, however, is

incomplete because the char yield is apriori assumed rather than predicted. In

this model, using surface temperature rather than heat flux, as a boundary

condition, permits the separation of the phenomena occurring in the gas phase

from those occurring in the solid phase. A comparison of the experimental and

theoretical results shows that the internal convective heat transfer by

volatiles and variation of transport properties with temperature are small

effects

.

The results of piloted ignition are consistent with previous work and the

pyrolysis model. It was found that ignition occurs prior to significant

formation of char i.e. n the kinetically controlled pyrolysis regime. A

simplified formula for fuel mass flux derived for this regime agrees well with

more detailed calculations. This formula also provides a relationship between

the critical fuel mass flux and critical surface temperature at ignition.

ACKNOWLEDGEMENTS

The first author is especially indebted to Professor G.F. Carrier for

several helpful suggestions. This work was supported by the Center of Fire

Research, National Bureau of Standards.

A

C P k

h

L

E

€R

F

T

a

m"

tx

Pa

NOMENCLATURE

- Frequency factor (1/sec)

- Specific heat (cal/gm°C)

- Thermal conductivity (cal/gm°C sec)

- Heat transfer coefficient/e

- Thickness of the sample

- Activation Energy (kcal/gm mole/

- Emissivity (-absorptivity) of the surface

- Gas constant (Kcal/gm mole K)

- Prescribed incident heat flux

- Temperature (°K)

- Stefan -Boltzmann constant

2

- Outward mass flux (-M at x-0) mg/cm sec

§

- Time (sec)

- Distance; x-0, top surface; x-L, bottom surface of the sample

(cm)

3

- Density based on original volume (gm/cm ) 2

- Thermal diffusivity (cm /sec)

33
M

H

- Outward mass flux of volatile gases (water vapor and decomposition

products

)

- Thermal sensible specific enthalpy (-

i-T Cp (T) dT

Q

- Endothermic heat of decomposition of wood for a unit mass of volatiles

generated (cal/gm) , at ambient temperature.

- Heat of moisture desorption (cal/gm) at ambient temperature.

SUBSCRIPTS

® - Ambient

s - Solid (unpyrolyzed wood and char) , surface

w - Dry virgin wood before pyrolysis has started, i.e. , at t-0

c - Char (instantaneous value)

g

- Volatile gases (water vapor and decomposition products)

a - Unpyrolyzed active material (instantaneous value)

f - Final value after pyrolysis has completed, i.e., at t—

m - Adsorbed moisture

p

- Pyrolysis

SUPERSCRIPTS

a - Constant value

* - Temperature coefficient

REFERENCES

1. Tang, W.K. : Effect of Inorganic Salts of Pvrolvsis of Wood, AlphaCellulose, and Lignin, Determined bv Dynamic Thermo gravimetry . U.S. Forest

Service Research Paper, FPG71, Forest Products Laboratory, Madison,

Wisconsin, (1967)

.

2. Roberts

,

A.F.: Combustion and Flame . 14. 261 (1970) .

3. Shivadev

,

, U.K. and Emmons, H.W.: Combustion and Flame. 22, 223, (1974).

4. Min, K.Combustion and Flame. 30. 285 (1977).

5. Hileman, F.D. , Wojcik, L. J

. , Futrel, J.M., and Einhom, I.N. : Thermal Uses

and Properties of Carbohydrates and Lignins . Academic Press, New York, p.

49 (1976).

6. Shafizadeh, F.: Chemistry of Pvrolvsis and Combustion of Wood . Int'l

Conference on Residential Solid Fuels, Oregon Graduate Center, Portland,

Oregon (1981)

7. Kanury, M.A. : J. of Fire and Flammability . Vol 2, 191 (1971).

8. Delichatsios , M.A. , and de Ris , J.: Analytical model for the pvrolvsis of

charring materials . Fall Technical Meeting,. Eastern Section, 68 (1983).

9. Bamford, C.H., Crank, J., and Malan, D.H.: Proc . Cambridge Phil, Soc. 42,

166 (1946).

10. Rung, H.C.: Combustion an Flame . Vol 18, 185 (1972).

34
11 . Tamanini, F. : The third full-scale bedroom fire test of the Home Fire

Project Vol. II - - Analvs is of test Results , Technical Report, Factory Mutual

Research Corp . , Norwood, MA, 107 (1976).

12. Kansa, E.J., Perlee, H.E., and Chaiken, R.F.: Combustion and Flame 29, 311

(1977).

13. Kanury, A.M. : Thermal Decomposition Kinetics of Charring Polymeric Solids . FM Research Corporation Report 19721-4 (1971).

14. Nolan, P.F.

,

Brown, D.J.

,

and Rothwell, E. : 14th Symposium (Int'l j on Combustion . 1143, The Combustion Inst., (1973).

15. Lee, C

.

, Chaiken, R.F., and Singer, J.M.: 16 th Symposium (Int'l -

) on Combustion . 1459, The Combustion Inst., (1976).

16. Markstein, G.H." 18th Symposium (Int'l) on Combustion . The Combustion

Institute, 537, (1981).

17. Atreya, A. , Carpentier, C. , and Harkleroad, J. , "Effect of Sample

Orientation on Piloted Ignition and Flame Spread, " Accepted for publication

in The First International Symposium on Fire Safety Science , 1985.

18. Atreya, A., Pvrolvsis , ignition and fire spread on horizontal surfaces of

wood . Ph.D. Theses, Harvard University, Cambridge, (1983).

19. Roberts, A.F., and Clough, G.: 9th Symposium (Int'l) on Combustion . 158,

The Combustion Inst., (1963).

20. Lewellen, P.C., Peters, W.A. , and Howard, J.B.: 16th Symposium (Int'l) on Combustion . 1471, The Combustion Inst., (1976).

21. Browne, F.L., and Tang, W.K. : Fire Res. Abs . and Rev. . 4, 76 (1962).

22. Broido, A., and Nelson, M.A.: Combustion and Flame . 24, 263 (1975).

23. Madorsky, S.L.: Thermal Degradation of Organic Polymers

,

Robert E. Krieger

Publishing Company, Huntington, New York (1975) .

24. Broido, A.: Thermal Uses and Properties of Carbohydrates and Lignins . Academic Press, N.Y. 1976).

25. Stamm , A.J.: Wood and Cellulose Science . The Ronald Press Company, New

York (1964).

26. Hodgmann, C.D., Weast, R.C., and Selby, S.M.: Handbook of Chemistry and

Physics . Chemical Rubber Publishing Co., Cleveland (1960).

27. Wesson, H.R., Welker, J.R., and Sliepcevich, C.M.: Combustion and Flame . 16, 303 (1971).

28. Kashiwagi, T. : Fire Safety Journal . 3, 185, (1981).

29. Quintiere, J. , Harkleroad, M. and Walton, D.

,

Combustion Science

Technology . Vol. 32 67-89, (1983).

35
TABLE 1

THE FOLLOWING CONSTANTS WERE USED IN THE AHLYSIS OF PYROLYSIS

DATA FOR MAPLE AND VGP

Activation energy 'E'

Pre-exponential factor 'A'

Heat of pyrolyaia 'Q'

* Thermal conductivity 'k^'

* Thermal conductivity 'k ' c

Density of wood 'p ' w Char yield 'p./p '

Specific heac of wood 'C pw

Specific heat of wood 'C PC

Specific heat of gasses 'C Pg

Activation energy for adsorbed

moisture 'E

' m Pre-exponential factor for adsorbed moisture 'A

1 a Heat of moisture desorption

30 Kcals/gm mole (best fit parameter)

2

- 2.5 X 10 1/sec (best fit parameter)

• -30; 0.0; 30 cala.gm (variable under

study) (positive when endothermic)

- 4 x 10~ 4 0.29 x 10* 6 (T-T ) cal/

cm sec K ( Ref . 1

)

- 1.7 x 10 _4 0.29 x 10~ 6 (T-T ) cal/cm sec K (Ref. 16)

- 0.676 gm/cm^ (measured)

• 0.24 ( asumed constant)

• 0.35 + 0.36 x 10

4 (T-T ) cals/gm K (Ref. 16)

- 0.17 0.48 x 10 _4

(T-T ) cals/gm K (Ref. 24)

°

• 0.25 + 0.43 x 10

4 (T-T ) cals/gm K (Ref. 24)

0

• 10.5 Kcals/gm mole (Ref. 1)

- 4. 50E>3 1/sec (Ref. 1)

• 574 cals/gm (Ref. 1)

THE FOLLOWING CONSTANTS VZKE DIFFERENT FOR VCP

Activation energy 'E*

Pre-exponential factor 'A‘

Heat of pyrolysis 'Q'

* Thermal conductivity 'It

'

* Thermal conductivity 'k '

Density of wood 'p ' w Char yield 'p,/p ’ f w

Specific heat of wood 'C

’ pw

- 28 Kcals/gm mole (best fit parameter)

2 • 1.0 x 10 1 / lec (best fit parameter)

• 0.0 cals/gm (assumed)

- 6.30 x 10~ 4 + 0.29 x 1 0~ 6

( T-T ) cal/cm sec K (Ref. 1) 0

- 1.4 x 10 -4

+ 0.29 x 10~ 6 (T-T ) cal/cm sec K (assumed)

• 0.326 gm/cm^ (measured)

“ 0.32 (best fit far expt. numbers 1,2 4 3) 0.25 (best fit for expt. numbers 4, 5, 6,

7 4 8)

- 0.33 * 0.36 x 10' 4 (T-T ) cala/gm K (Smithsonian Tables ) °

*The temperature dependence was estimated by a simple resistance model with

intra-void radiation.

36

38

Measured

surface

temperatures

for

piloted

ignition.
otPaTNajunoK

Fig. 6 Decompsition kinetics.

pOO»» JO I* IOU0IJO %

39

Fig.

7

Decomposition

energetics.
795

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on 3amva3dw3±

cuo

>

u

o

4-1

ui0) u

3

4J

cau

4) a

£

a) u

ui

caa£) T3C

<a

4J O

<a 4JU

3

onON

ao

—4

3 " '

3 3 ^

e e 3

—i a a — a. » “ u f>a

a a u O

4) U «4 « O ^ ti ® *

"r <a e u 3 £

3J C ^ -* *-»

a- jd a o —

m
FIRES

k DETERMINISTIC PROPERTIES OF TURBULENT DIFFUSION FLAMES FROM LOW Q

Yuji Hasemi

Building Research Institute

Ministry of Construction

Mitsuru Nishihata

Center for Fire Science and Technology

Science University of Tokyo

INTRODUCTION

Turbulent diffusion flames are generally a major element in the

development of a fire. In the early stages of a building fire, flames have a role to preheat the surfaces of the whole enclosure by radiation and to drive

hot gas to the ceiling layer. Also in urban or wildland fires, flame

radiation and confluence of individual flames into a huge flame have a significance not only for the development of fire but also the generation of

difficulties of fire suppression.

Practical properties concerning these aspects of fire safety have been

studied for various purposes ( 1-6) . Dimensional analysis of the flame geometry

has^^tablished the dependence of the ratio of flame height to fuel size on

Q/D , and its significance ^^s^^en often represented by a dimensionless

heat release rate Q = p C T g ^D or some other parameter representing an essentially same implication (4-6 , 10) ; temperature, velocity and entrainment

in the turbulent plume above a flame have been also found to be influenced by

the flame geometry, and its significance in the application of pc^int source

theory, location of virtual source, has been correlated with Q or flame

height (7-10) . While low Q fires are still probable both in building fires

and in mass fires(9-ll), in most of the experiments heat release rate has

been chosen to generate a fl^me height considerably greater than the

characteristic fuel size, i.e.Q » 0.1. Sand-wick experiments by Wood et al(12) provide only available data to have a peep of the practical properties

in the Q <0.1 regime. The dimensional analysis expects a dependence of flame

height only on the heat release rate per unit area, Q/D^, in the low Q

regime(l), and earlier experimental works seem to show this tendency ( 10 , 12)

however there is still considerable discrepancy and scattering among the

relevant experimental data. Also for the properties of the upper plume

regioij, the concept of virtual source is expected to fail in the limit of ow Q fires, since low Q fires must be essentially closer to a flat heat

-ource with finite area rather than a point sourc^.

The difficulty in the experiments on low Q fires is essentially the

necessity of relatively large fuel and therefore a huge experimental

facility. Since Wood et al's data still seem to show some instability

probably due to the use of sand wicks, accumulation of experimental data on low Q fires using various fuels is important as a basis for predicting their

practical properties.

T£ie purpose of this paper is to report some characteristic behaviors of

low Q fires obtained from experiments using a large porous refractory

burnej, which may allow a more controllable and stable combustion even in the

low Q regime than any other conventional fuels do.

41
EXPERIMENTAL DESCRIPTION

The experiments were conducted primarily using a porous square burner of

1.0m with propane as the fuel, and supplementary experiments were made using

0.20m and 0.50m square burners to study some source-size dependence of the

flame properties.

The burners were filled with ceramic beads of approximately 8mm

diameter, and the height of their surfaces was approximately 0.5m above the

floor of a laboratory, 27m by 20m in plan and 27m high without artificial

floor surround at the level of the burner exits. Anticipating that burner

exit velocity would become less uniform by reducing the fuel supply, plumbing

inside the 1.0m burner was carefully made; the burner surface appeared to be

covered with flames for every experiment reported in this paper.

Observations were made on the geometry of visible flames (general shape

and height) and centerline temperature.

Visible flame was observed by eyes and monitored by a video camera during each experiment. Since the general shape of flame was found to change

dramatically by the operation of fuel supply rate during a preparatory test

using the 1.0m burner, the shape of flame was carefully observed. Data of the

height of flame tips were obtained from observation of the videotape; the

reported values of flame height are the average of the data observed for more

than 3 minutes at the intervals of 1 s. Temperature measurements were made using 0.20m diameter chromel-alumel

thermocouples. The time constant of the temperature measurement system is

approximately 10 s. No corrections have been made for the radiation and

conduction losses from the junction of each thermocouple. To eliminate the

influence of accidental sway of the flame, temperature was monitored at each

height not only just above the center of each burner but also at 4 different

points 5cm apart from the center. The reported values of temperature are the

average of the centerline temperatures at the interval of 10 s over 3 minutes

while the centerline temperature was higher than those at any other points.

RESULTS

Flame Geometry

* The general shape of flame in the very low Q regime was apparently

quite different from tb^.t in the regimes of conventional studies (mostly Q > 0.1). For Q <43kW(Q < 0.039) of the 1.0m burner, fire was broken into

flamelets of approximately same height distributed over the burner surface.

At Q= 43kW(Q — 0.039), flamelets along the burner sides starts £o join

together to build "filmy" flames, and at approximately Q=130kW(Q =0.11),

small flametips starts to appear above the filmy flame. Figure 1 shows a

Figure 1 Shapes of flames

42
brief sketch of the characteristic flame shapes for these 3 regimes. The

filmy flames in the region II looked like a combusting boundary layer above a horizontal flat plate under forced convection. Similar transition of the

flame shape was also seen for the 0.50m burner, while the transition from

flamelets to filmy flames was somewhat ambiguous at the 0.20m burner. Figure

2 summarizes the Q-D relationships at these transitions; it is noteworthy

that the critical heat release rate for the transition from filmy fire to the

appearance of flametips is almost proportional to the square of the

characteristic fuel size, and practically represented as

(Q/D

2

) . = 140 (kW/m 2

) ( 1) crxt

The slope representing the transition from individual flamelets to filmy

flames seem to be somewhat less steeper than this while the slope itself is

not yet clear.

D (m)

Figure 2 Heat release rate vs. characteristic fuel size at the transition of

flame shape

Results of flame height measurements are summarized in Figure 3 and

compared with results from earlier works. While the flame height data for

L

f /D <0.5 or Q < 0.1 are generally scattered, it is evident at each

experiment that the slope tends to become sharper in the low L

f /D regime. For

Q < 43kW of the burner anf^^ 01" Q < 23kW of the 0.5Ckn burner, the

expression L^/D — §2 / 3 or ~ (Q/D ) may be adequate, while L../D seems to be

proportional to Q for Q >130kW of the 1.0m burner and for Q >35kW of the

0.50m burner, and the slope between these two regimes is somewhat less

steeper than the both regimes. This gentle slope in the very low Q regime is one significant result which has not been reported in • relevant previous

reports (10,12). Interestingly, these divisions of the regimes are corresponding to the division for general flame shape as shown in Figure 2.

43
10

_

——“I

1 1—I 1 I l i

I oa 1.0 m square c 3 h 8 o<> 0.5 m square c 3 h 3 m 0.2m square c 3 n a 7 0.6m square ch 4 , a 0.45 m square ch 4

I [

0.1

0.01

Cox a 3

<j 1.52 m dia. ch 3 coch3 ^ o 1.52 m dia. ch3oh

Wood 3 iia' ^

oi? . Of] *2, (3

<3 0°« 37

<j «C <*> o

0.01 0.1

Figure 3 Normalized flame height vs . Q

Characters depicting the regimes in Figure 1 (•rregime I,

c<>n: regime II , o<>a: regime III)

Centerline Temperature

Figure 4 shows the centerline excess temperatures for the 1 . Om^^mer as

a function of the height from the burner surface normalized by Q D; the

centerline excess temperatures are somewhat scattered, while they tend to

cluster around one curve obtained in previous experiments for around unity of

Q (2,3). Since the slopes of most of the data in the figure are evidently

less steeper than -5/3 while most of the temperature measurements were made

above the visible flames, the concept of virtual source was applied to adjust

the slope to -5/3 as shown in Figure 5. The adjusted data are divided into

two groups; one includes only the data from Q> 0.039, namely the regimes II

and III, and the other comes from those on Q <0.039 fires. The data on the

regimes II and III are scattering around the line obtained in the previous

paper(9) , whereas the other data are considerably lower than the previous

results; this may suggest that the concept of virtual source would fail in

the regime I.The centerline excels; ,£emperatures f or 0 .50m burner are also

shown as a function of both z/Q “ D and (z+Az)/ Q D in Figures 6 and 7. The difference of the regime I and the other in Figure 7 is less significant

than in Figure 5. However, the centerline temperature in the regime I is

still lower than those in the other regimes.

44
Figure 4

Figure 5

Centerline excess temperature

Characters depict the regimes

O : regime III) .

vs. normalized height (1.0m burner)

in Figure l(D:regime I, av: regime II,

Centerline excess temperature vs. normalized height from virtual

source (1.0m burner)

Characters depict the regimes in Figure 1( :regime I regime II,

O :regime III)

.

45
Figure 6 Centerline excess temperature vs. normalized height(0.50m burner)

Characters depict the regimes in Figure l(d:regime I, av: regime II,

O : regime III)

.

Figure 7 Centerline excess temperature vs. normalized height from virtual

source(0.50m burner)

Characters depict the regimes in Figure l(d:regime I,AV:regime II,

O :regime III)

.

46
The location of the virtual source thus obtained is* summarised in Figure

8; its depth tends to become lower with decrease of Q for Q > 0.039, and

apparently it turns to rise in the regime I. The virtual source locations for

Q >0.1 seems consistent with previous experimental works (7-9); however, for

the regime II, the location is considerably lower than the expected value

from the previous experiments, and in the regime I, the location is

scattered around 0.

0.5

Q

\N

<3 0

- 0.5

-1

- 1.5 0 0.1 0.2 0.3 0.4 0.5 Q*

* Figure 8 Location of virtual source vs. Q

Characters depict the regimes in Figure !(•: regime I, eo: regime II,

OO: regime III).

DISCUSSIONS

2 2 The present experiments seem to show that the L^cc (Q/D ) relationship

holds for very low Q fires, in the regime I of Figure" 1 or 2, as anticipated

from previous dimensional studies ( 1

, 10 , 1 1 ) . However, there is still

considerable discrepancy in the flame height between the 1.0m and 0.50m

burners, and *the transition from the regime I to II does not seem to be

governed by Q . Controlling mechanism of these transitions are important from

the viewpoint of fire safety; in a room fire, the difference of regimes

implies a significant difference of preheating of surrounding combustibles by

the flame, and, in a mass fire, this transition may provide some information

concerning the criticality for the confluence of individual flames into a gigantic flame which is attributed to one mechanism leading to a great loss

as seen at The Kanto Earthquake Fire(1923) . The following is an exploratory discussion to seek a responsible

mechanism for these criticalities based on the dimensional relationships

obtained in the present experiments, while it should be noted in advance that

different mechanisms may induce one identical dimensional relationship and

there is still some possibilities of other mechanisms leading to the

criticalities

.

One possible cause for these transitions is the criticality of oxygen

supply by entrainment. At the upper bound of the regime II for the 1.0m

burner, the filmy flame appeared to cover the whole surface of the burner,

and therefore the appearance of flametips is attributed to the lack of oxygen

supplied onto the burner surface. Actually the heat release rate at this

criticality is proportional to the burner surface area, D , as seen in Figure

2, and it is expected that there is a critical available oxygen or fresh air

rate per unit fuel surface area. Using the general value of reaction heat per

unit mass of air for usual natural fuels(2700kJ/kg, see Ref. 13), the critical

47
available entrained air, M, per unit area based on the above hypothesis is

estimated as

M/D 2 = 0.052 (kg/m 2

) ( 2)

The transition from individual flamelets to a filmy fire must be related

more to the entrainment through the circumference of the fuel surface, since

the nature of this criticality is somewhat "local” in the sense that a filmy

flame starts to develop along the margin of a fuel and gradually takes the

place of flamelets in the central part of the burner by the increase of Q. If

this criticality depends on the heat release rate per entrainment through the

circumference, Q/D may be a primary parameter characterizing this

criticality; this mechanism is worth studying since the power of D representing Q at this criticality in Figure 2 is somewhat close to 1.

CONCLUSIONS

Some experimental observations on low Q fires are reported. The

following conclusions can be drawn.

1 Flame behavior can be classified into 3 regimes according to the shape of

the flange.

^

2 For Q/D“ 140kW/m , flametips are observed above a continuous flame, and

flame height i| proportional to 2/3 power of heat release rate.

3 For very low Q , fire is broken in^o^individual flamelets, and flame height

is apparently proportional to (Q/D ) . In this regime, the concept of

virtual source fails evidently.

4 There is a transition regime between the above two regimes. In this regime,

fire looks like a combusting boundary layer under forced convection. The

dependence of flame height on heat input is less significant than in either

of the two regimes. The concept of virtual source seems to hold in this

region

.

5 Experiments using larger fuels is^encouraged to develop a model to predict

the practical properties of low Q fires.

ACKNOWLEDGEMENTS

The authors are indebted to Mr.Jun'ichi Izumi, research associate of BRI

from Mitsui Homes Co., for the assistance in the experiments. The authors are also grateful to Mr.Hideo Matsumoto of BRI who fabricated the major apparatus

for the experiments. His effort to build a large uniform burner is

gratefully appreciated.

TERMINOLOGY

C

Q*

Q

T

g

P

suffix

o ambient air

crit criticality

specific heat of air

characteristic fuel size

flame height

entrained air

heat release rate ^2 dimensionless heat release rate defined as Q/p T g D

absolute temperature

gravitational acceleration

density

48
REFERENCES

1 Thomas ,P .H. , Webster , C .T. , and Raf tery ,M.M. , "Some Experiments on Buoyant

Diffusion Flames," Combustion and Flame, Vol.5, 1961, pp359-367.

2 McCaffrey ,B.J. , "Purely Buoyant Diffusion Flames: Some Experimental

Results," NBSIR-79-1910 , 1979.

3 Cox,G., and Chitty,R.,"A Study of the Deterministic Properties of

Unbounded Fire Plumes," Combustion and Flame, Vol.39, 1980, pp. 191-209.

4 Zukoski , E . E. , Kubota,!., and Cetegen,B ., "Entrainment in Fire Plumes," Fire

Safety Journal, Vol.3, 1980/81, pp. 107-121.

5 Hasemi,Y., and Tokunaga ,T.

, "Deterministic Modeling of Unconfined Turbulent

Diffusion Flames," Transactions of ASME, Journal of Heat transfer, Vol.108,

1986, pp. 882-888.

6 Heskestad ,G.

, "Luminous Heights of Turbulent Diffusion Flames," Fire Safety

Journal, Vol.5, 1983, pp. 103-108.

7 Ce tegen ,B .M. , Zukoski ,E. E . , and Kubota , T ., "Entrainment and Flame Geometry

of Fire Plumes," NBS-GCR-82-402 , 1982.

8 Heskestad ,G.

, "Virtual Origins of Fire Plumes," Fire Safety Journal, Vol.5,

1983, pp. 109-114.

9 Hasemi,Y., and Tokunaga ,T.

, "Flame Geometry Effects on the Buoyant Plumes

from Turbulent Diffusion Flames," Fire Science and Technology, Vol.4, 1984,

pp . 15-26

.

10 Cox,G., and Chitty , R.

, "Some Source-dependent Effects of Unbounded Fires,"

Combustion and Flame, Vol.60, 1985, pp. 219-232.

11 Zukoski , E . E ., "Fluid Dynamic Aspects of Room Fires," Proc. First

International Symposium -on Fire Safety Science, 1985, pp.1-30.

12 Wood , B .D . , Blackshear , P . H . , and Eckert ,E.R.G. , "Mass Fire Model: An

Experimental Study of the Heat Transfer to Liquid Fuel Burning from a Sand-filled Pan Burner,"" Combustion Science and Technology, Vol.4, 1971,

pp. 113-129.

13 Huggett ,C.

, "Estimation of Heat release by Means of Oxygen Consumption

Measurements," Fire and Materials, Vol.4, No. 2, 1980, pp. 61-65.

49
TRANSIENT COMBUSTION IN A TURBULENT FIRE

Howard R. Baum

Ronald G. Rehm

Center for Fire Research

National Bureau of Standards

Gaithersburg, Maryland 20899, USA

ABSTRACT

A mathematical model of the local transient diffusion flame generated by a

- mixing in a turbulent eddy is presented. It is intended ultimately as a computational "molecule" to be imbedded in the authors existing numerical

simulation of large scale fire driven flows. The model consists of a three

dimensional strain-vortex field which exactly satisfies the constant property

Navier Stokes equations, a convection diffusion equation for the mixture

fraction which determines the thermodynamic quantities, and a potential flow

determined by the conservation of mass. The properties of the model are explained and its relevance to turbulent combustion are illustrated by sample

results

1. INTRODUCTIONThe purpose of this paper is the presentation of a mathematical model of gas

phase combustion appropriate for use in a field equation simulation of

enclosure fires. It is therefore necessary to summarize briefly how such

simulations are performed. Typically, the enclosure containing the fire is

divided into a large number of rectangular cells. Within each cell the laws

expressing the conservation of mass, momentum, and energy ip. each cell are represented in an approximate algebraic form so that a single value for the

velocity, temperature, and pressure is associated with each cell at any

instant of time. Given a set of values for these quantities at any instant of

time, the conservation laws then determine their values a small interval of

time later. Thus, for a prescribed initial state and room geometry, the

evolution of the gaseous properties in the room can be predicted.

Clearly, the accuracy with which such a simulation can be carried out depends

strongly upon the number of cells that can be employed in the calculation.

Using modern numerical techniques on current generation vector computers, it

is quite feasible to employ roughly 10 5 computational cells and several

thousand time intervals. Figure 1 shows a sample time frame from a calculation performed by the authors (1) . The motion is followed by seeding the flow

with thousands of point particles which originated in the fire. The (40) 3 grid used means that the minimum length scale in the computation is 7.5 cm. While this is adequate to resolve the largest scales of the turbulent fire

induced flow, the diffusion controlled combustion phenomena characteristic of

fires typically evolve on sub -millimeter scales. Indeed, the flow in figure 1 was calculated by replacing the fire with a prescribed heat source.

50
In order to study the combustion itself, it is necessary to abandon a grid

fixed in space. Instead, attention is focused on the fuel, since only in the

vicinity of the fuel can combustion occur. Each of the point particles is now taken to represent a fixed mass of fuel. The locatiQn of each blob of fuel at

each instant of time is determined by tracking the particles trajectory

through the large scale flow. If the rate of heat release of each fuel parcel

can be determined as a function of its local environment for each instant of

time; then the composite heat release rate of all the fuel parcels represents

the desired replacement for the prescribed heat release currently employed.

In order to determine the rate of heat release of an individual fuel blob, it

is necessary to change the scale of the analysis and consider in detail how it

burns. (See fig. 1).

2. MODEL FORMULATION

The combustion process viewed on the scale of an individual fuel parcel

contains three essential ingredients: First, the molecular mixing of fuel and

oxidizer and consequent release of heat; second, a local flow field on the

scale of the fuel parcel characteristic of turbulence, to enhance the

molecular mixing; finally, a mechanism whereby the flow is modified by the

combustion process.

To address the first point we consider a highly idealized reaction between a fuel F and oxygen leading to "perfect 11 combustion denoted by the reaction

"f F + u

o °2 -„ H2° + 1/1 c C02 + ^o®0 Q- C 1 )

Here, u i are the stoichiometric coefficients for the overall reaction and Q is

the energy released per unit mass of oxygen consumed. While such a model is

hopeless as a description of the real combustion chemistry, and says nothing

about the resulting carbon balance between soot, CO, and C02 it is a useful

starting point for the mixture fraction concept. This concept can be

generalized to permit real combustion chemistry to be incorporated (2). We

further assume that all species diffuse according to Ficks law with equal

diffusion coefficients D, and that the Lewis Number of the mixture is unity.

Then, in a frame of reference moving with the fuel parcel, the species mass fractions Yi and temperature T obey the following conservation equations

V- (pDVT) - q

f3Yi j— + u-VY^ - V • (pDVYi ) = m,

C2)

The quantities q and mi in equations (2) are the rate of heat released and

mass consumed for each species respectively per unit volume, while u is the

fluid velocity and p the density.

51
We are now in a position to introduce the mixture fraction variable. The

stoichiometry given by equation (1) implies the following relationships

between q, nq , the stoichiometric coefficients , and the molecular weight

mi

:

mF AF mF = m0 /vQ m0 = -q/Qi/

0 m0 = = mc /i/c mc These relationships allow all of eqs . (2) to be collapsed to a single equation

for the mixture fraction Z, defined as:

Z = (j8 - P°)/(P1

- F) (3)

Here /3 is any of the quantities conserved in the reaction

~ ^ F / u

F mF Y0 /^o mo

&2 = C

p T/Q + Yo (4)

03 = Yn

; 0* = YF /u7 mF + Yw /uw mw

2 0 5 = YwAA * YcAc mc = 0

The quantity 0a denotes the value of the conserved quantity in the ambient

environment seen by the individual fuel parcel; this value is ultimately

determined from the large scale flow calculation. The quantity 0

l is the initial value of f3 at the instant the fuel parcel is injected into the gas phase

from the burning object. The result is that all of equations (2) can be

replaced by a single equation and initial conditions for the mixture fraction

Z.

P

37. at

+ u-VZ V- (pDVZ)

Z ( o , r ) = 1

Z(o , r) = 0

in fuel

otherwise

(5)

The connection between Z and the physical variables is completed by adding an equation for the gas and the fast chemistry assumption to equations (3) and

(4)

The equation of state is assumed to have the form:

pT = p 9 Tffl (6)

The fast chemistry assumption means that fuel and oxidizer cannot coexist.

Thus, the quantity 0

,

represents fuel for Z > Z s

, the flame sheet value of the

mixture fraction, and oxidizer for Z < Z 3

. Requiring that fuel and oxidizer

vanish at the flame sheet readily yields the flame sheet temperature Ts and

mixture fraction Z s as:

52
z s = ( /vQ mQ ) { Yp /f}v ? mF + ^/u^)

( 7 )

T s = Tffl + {Yi + (i/ F mF /uQ mQ ) Y^ } ~ 1

p

The state relationships between Yi

, T, p, and Z are shown schematically in

fig. 2. While the sketch applies only to the idealized reaction given in eq.

(1), Bilger (2) has pointed out that laminar flame data can be used to

generate a realistic set of state relations that are applicable to turbulent

flows. Faeth (3) has exploited this idea to predict the structure and

radiation from turbulent buoyant jets. Pagni (4) has used variables

equivalent to Z in his analysis of excess pyrozylate

.

Next, we consider an appropriate description of the local flow field on the

scale of the fuel parcel. In a frame of reference moving with the large scale

fluid motion, it is always possible to express the local velocity u as a combination of a solenoidal (vortex- induced) velocity v and an irrotational

flow derived from a potential 4>

.

u = v + V<^

( 8 )

V • v = 0 ; V x v = w

The vortex induced velocity field is assumed to be a transient vortex

stretched by imposed strain field (5)

.

The geometry of this strain-vortex

field is illustrated in fig. 3. This geometry is known to be characteristic

of the velocity field in an individual turbulent eddy. In fact, it can be

shown that the smallest length scale in this model emerges naturally as Kolmogoroff scale (5).

The vortex induced velocity field is an exact solution of the constant

property Navier-Stokes equations. If v , v , and v9 are relatively the axial,

radial, and swirl components of the solenoidal velocity field, the NavierStobes equations have solutions of the form:

vx = ax; vr = -ar/2 ; v9 = Qro W(r,t)

r = r0 r ; t = (a) _1

t (9)

Here a is the imposed strain rate; Q the initial vorficity, and ro the initial

radius of the vortex. The intensification and collapse of the vortex induced

swirl are shown in figure 4 for a local Reynolds number aro

2 /V of 500.

Finally, we consider the feedback of the energy release on the flow field.

This consists of two parts. First, the heat released from the fuel parcels

acts as a distributed heat source for the large scale fluid motion; altering

the density field and thus changing the buoyancy forces which drive the large

scale flow. Second, the expansion due to the heat addition generates a

53
potential field which alters the local velocity on a small scale. Since

attention is focused on the local flow field on the scale of the individual

fuel parcel, we consider the second part.

To proceed, note that the conservation of mass can be written in the form:

pV • u = 0 (10)

Now recall that all thermodynamic variables can be related uniquely to the

mixture fraction variable Z. Using this fact together with the velocity

decomposition given by eq . (8) in eq. (10), an equation for the velocity

potential <p is obtained. The solution to this equation is:

= J + —- ( Y

1 + Y

a

) P.C 0 T, Snu ? 47T o ds n. dZ 1

k 3y,k lx-yl

Z= 2

J =

_ i/ m

vi _1_1

Pco C

p Ta

F ^ f p( Z)D(Z)dZ ; 0<Z<Z =

J = —r ~r- ^ f p(Z)D(Z)dZ ; Z S <Z<1

Pa,CpTa,

(ID

The physical content of eq. (11) is that the flow induced by the local heat

addition consists of two parts. The contribution denoted by J is the heat

flux conducted away from the flame sheet towards either the ambient gas or the

interior of the fuel parcel. The formulae differ on either side of the flame

sheet because the conduction heat flux vector is discontinuous at the flame

sheet itself. The discontinuity arises because in the fast chemistry limit,

the chemical heat release is confined to the flame sheet, a (possible multiply

connected) surface characterized by the mixture fraction taking on the value

Z s given by eq. (7). The remaining contribution to the portion <f> is the

direct effect of the heat addition. The integral involves the mixture fraction gradient evaluated at the flame sheet, integrated over the entire reac- ion surface. Although the two contributions give rise to comparable

velocities, the velocity gradients associated with the heat flux are much

larger, and these terms are much more important in coupling the velocity field

to the heat release

3. ANALYSIS AND DISCUSSION

The combustion model may now be summarized as follows: The gasified fuel is

to be represented by a large number of distinct parcels, which interact with

and modify their local environments. The rate of heat release q p for each

fuel parcel may be computed as:

54
-Q<- Yl + Y*)p d o ds n ( 12 ) k az_

dxk

In order to evaluate this formula, it is necessary to solve for Z. This

requires a solution to eq. (5), with the velocity field expressed by eqs . (8),

(9), and (11). Since the potential flow (eq. (11)) itself involves Z, and the

density p as well as the diffusivity D are both functions of Z, the resulting

system is highly non-linear. Moreover, the intense mixing induced by the

vortex field leads to an extremely convoluted interface across which the fuel

and oxidizer diffuse. The remainder of this paper discusses how these

difficulties may be approached and a solution for the mixture fraction

obtained

.

The starting point for the analysis is equation (5). Since p(Z) and D(Z) are non-differentiable functions of the mixture fraction, it is more convenient to

work with a modified mixture fraction F(Z) , defined by:

Z 1 F(Z) = / p(Z)D(Z)dZ/J* p (Z)D(Z)dZ (13)

0 0

The function F(Z) varies within the range 0<F<T, and is constant with value

F$

, say, along the flame sheet. The velocity potential can be written as:

4> = + K+ F(Z) ; 0<F<FS (14)

4> = <t> 5

- K_ F(Z) ; FS <F<1

Here, <j> s denotes the contribution of the surface integral over the flame sheet

in eq. (11), while the constants K+ and K_ are given by:

l / p(Z)D(Z)dZ

o

(15)

K. = / p(Z)D(Z)dZ

Pea

0 The mixture fraction equation can now be rewritten in terms of F in the form:

K. = — Y

1 P«Cp Tffl *

—H F -> « + (v+V0 s

) • VF + K+ (VF) 2 DV2 F

0<F<FS

OF —

— + (v+V<^ )

- VF - K (VF) 2 o z

5 DV2 F; Fs <F<1

(16)

55
The function F and its first derivative are continuous across the flame sheet

for times t>0 , while the initial conditions for F are identical with those for

Z (see equation (5)).

To proceed, it is necessary to approximate the diffusion coefficient D by a constant value, which will for the present taken to be the flame sheet value

D s

. We also note that at high Reynolds numbers the vortex induced velocity v

is both larger in magnitude and has much steeper gradients than the heat

release potential flow Vb 3

. This contribution to the potential flow is

henceforth neglected. Given these simplifications, it is possible to reduce

equations (16) to a single constant property linear equation without further

approximation. We introduce a new dependent variable, the pseudo -mixture

fraction G defined as follows:

G = a+ + b + exp (-K+ F/Ds

) ; 0<F<Fs

G = a. + b. exp (K_F/D S

) ; F S <F<1

( 17 )

The quantities a+ , a. , b + , and b_ are chosen so that the continuity of F and

VF at the flame sheet implies continuity of G and VG there. We also require

that G vanishes with F and that G=1 when F=l. This leads to a simple linear

system of algebraic equations which will not be displayed here. The more important result is that equations (16) collapse to the single linear equation

7T + (v-7) G = D V

2 G

at 3

G (0,r) = 1 in Fuel

( 18 )

G (0,r) = 0 otherwise.

At this point the issue of non-linearity has been dealt with; and we turn to

the effect of mixing by the vortex. This had been studied by Marble (6),

Karagozian and Marble (7), and Lundgren (8) within the context of an inviscid

flow field. Here, we indicate an alternative approach which allows for

viscous effects and can handle a variety of fuel geometries (5). The solution

for the stretched vortex flow represented by equations (9) can be used to

generate a set of Lagrangian cylindrical coordinates £, p, <j>; analogous to the

cylindrical coordinates x, r, 9 illustrated in figure 3. The transformation

is defined by the relations

£ - x exp (-t)

p - r exp (t/2)

- 9 - (fl/ae) i>(p,r)

T (p,r) = (p)' 1 f

V (p , r ) dr

0

( 19 )

56
r = € [ exp ( t) - 1]

w = exp ( t/2) V(p , r

)

€ = u/ar0 z

Note that the quantity £ is the reciprocal of the local Reynolds number which - sets the local viscous length scale in the vortex flow defined in equation

(9).

The introduction of Lagrangian coordinates transforms equation (18) into the

form:

3G

dr s

c j(l + r/O ' 3 3

2 G 1 d_ pdG

d£ 2 +

p 3 p dp

D dd 32G

p2 d<t> 2

- (Q/ae

)

r 1

3

3G dij> 3

Z G

) p dp d<t> dp 3 pd4>

1

( 20 )

Here, S

c is the Schmidt number u/D, and D^ is the effective azimuthal

diffusivity in Lagrangian cylindrical coordinates given by:

- +l ( n/ae ) 2 (|^)

2

. -( 21 )

Equations (20) and (21) are very important in understanding the role of

vorticity in turbulent combustion. Note that in Lagrangian coordinates there

is no advection by definition, but the diffusion is now both anisotropic and

inhomogeneous. In the absence of diffusion (S c -«o) G would not change from its

initial state in these coordinates. Thus, the mixing process is contained in

the transformation given by equations (19). As an example, the evolution of a cross section of an initially spherical parcel of fluid due to vortex mixing

is shown in figure 5. The shape in Lagrangian coordinates in the absence of

diffusion remains a circle displaced have a radius from the center of the

vortex

.

The diffusion itself is strongly enhanced by the vortex induced strain.

Figure 6 shows the quantity (Q/ae) plotted versus the Lagrangian radial

coordinate p for the velocity field (difr/dp) displayed in figure 4. Recall

that the local Reynolds number is 500 in this calculation. Even for this

relatively modest Reynolds number, figure 6 and equation (21) show that in the

region of most intense vorticity, the azimuthal diffusivity is enhanced by a factor of over 103 relative to its value in the absence of eddy motion. The

final stage of the analysis exploits this directionality to obtain a solution

for the pseudo -mixture fraction G.

57
The solution proceeds in two parts. Although a formal result can be obtained

for any initial shape of fuel parcel, we restrict attention at present to a sphere of radius R0 displaced from the center of the vortex as shown in figure

7. Then, in the absence of the vortex (i.e. strain only), the diffusion

equation can be readily solved in the form:

G =

2 T

- 0 r- s

= ., , i dA A J(A,£-a)exp - -o wc

r S erf (^)

H

(f -a + J R0

2 - A

2

)

r S + erf ^)

4

(a-e + J R0 2 - A

2

) ( 22 )

p = {

p

2 - 2b p cos 4> + b

2 )H

f ( r

)

= r (1 + r/2 e ) / ( 1+r /e

)

2

Here, a and b are parameters describing the displacement of the sphere from

the center of symmetry of the strain field. Denote this solution by G (Q=0)

Now this solution can always be expanded in a Fourier series in the form

G(Q = 0) = 2 exp(in^) Gn (p,£,r). (23) n—

0

Although the Fourier coefficients Gn can obviously be expressed as an integral, in practice it is undoubtably much more efficient to apply the Fast

Fourier Transform (FFT) algorithm to tabulated values of G(O=0) directly. In

any event, once values of G n have been obtained, a further analysis of

equation (20) shows that in the large Schmidt number large Reynolds number

limit (5), the solution to the full equation becomes:

G = 2 exp( in<£) Gn (p , £ , t) (p , r)

n=0

f n = exp{ -n2 (S e )' 1 (Q/ae) 2 /

, 3-0 N

(24)

Physically, the solution states that the vortex free solution is appropriate

in any region not yet caught up in the vortex. However, once the vortex

passes thru any region of space, the azimuthal diffusion dominates all else.

This effect is weakly modulated by the remaining anisotropy in the strain

induced (vortex- free) diffusion. The large Schmidt number limit is not

appropriate in a formal sense for gaseous mixtures. However, in the analogous

case of heat transfer, Lighthill (9) has shown that the large Prandtl number

limit, while even less appropriate for gases, in fact yields a very accurate

approximation for boundary layer heat transfer in gases. It is also worth

58
noting that equation (24) contains as special cases all previous results (6),

(7), (3) in the absence of viscous effects on the flow field.

4.

CONCLUDING REMARKS

The mathematical combustion model whose formulation and analysis is outlined

above, is but one ingredient required in a comprehensive simulation of fires

in enclosures. The roles of radiation and soot particulate formation have yet

to be addressed. These phenomena are also crucial to an understanding of fire

growth processes, since the radiative feedback to burning surfaces provides

the energy needed to gasify the fuel in the first place. However, it is

necessary to begin somewhere; and if the gas phase combustion can be adequately predicted by models like that described above, then a useful basis for more complete studies will be in place.

5

.

ACKNOWLEDGEMENT

This work was supported by the Air Force Office of Scientific Research under

Contract 4ISSA- 87 -0018

.

6

.

REFERENCES

.(1) Baum, H.R. and Rerun, R.G.: Combustion Science and Technology 40, p. 55,

1984.

(2) Bilger, R.W.: "Turbulent Reacting Flows" (P.A. Libby and F.A. Williams,

Eds.) p. 65, Springer -Verlag , 1980.

(3) Jeng, S-M., Lai, M-C., and Faeth, G.M.: Combustion Science and

Technology 40, p. 41, 1984.

(4) Pagni, P.J. and Shih, T.M.: "Sixteenth Symposium (International) on Combustion", p. 1329, The combustion Institute, 1976.

(5) Baum, H.R., Corley, D.M., and Rehm, R.G.: Twenty First International

Symposium on Combustion, The Combustion Institute (In Press).

(6) Marble, F.E.: "Recent Advances in the Aerospace ' Sciences " (C. Casci,

Ed.), p. 395, Plenum, 1985.

(7) Karagozian, A.R. and Marble, F.E.: Combustion Science and Technology

45, p. 65, 1985.

(8) Lundgren, T.S.; Phys . Fluids 40, p. 1641, 1985.

(9) Lighthill, M.J.: Proc. Roy. Soc. A 202, p. 359, 1950.

59
£=.002 Q= 2. TIME=O.IO f =.002 5= 2. TIME =0.50

£=.002 0= 2. TIME =1.0

Figure 5. Mixing of initially spherical fuel parcel by vortex displaced

one half radius from center. Local Reynolds number 500,

vorticity/strain ratio 2.

60

It
Buoyant Convection and

Combustion Models

Figure 1. Relationship of buoyant convection and combustion models

&££A T/O/VS

Figure 2. Sketch of state relationships between dependent variables and 3

61
D/SPLACPD SPAS'PE

Figure 6. Diffusivity enhancement

function

Figure 7. Spherical fuel parcel geometry

BB

62
NUMERICAL SIMULATION OF EARLY STAGE OF A COMPARTMENT FIRE BY FIELD MODEL

Tomoaki YOSHIKAWA a>

, Junji MASH I GE

b)

, Tomoko JO

b> ,and Osami SUGAWA

°

a) Japan Meteorological Agency, Meteorological Research Institute,

1-1, Nagamine Yatabe-cho, Tukuba, Ibaraki, 305, Japan

b) Century Research Center Corporation

3-6-2, Nihonbashi Honcho, Chuo-ku, Tokyo, 103, Japan

c) Department of Architecture, Faculty of Science and Technology,

Center for Fire Science and Technology,

Science University of Tokyo,

Yamazaki 2641, Noda City, Chiba, 278, Japan

ABSTRACT

We improve the field-equation model previously proposed in 1984 1

5

by

the inclusion of the mechanism of ignition by radiation, and the behavior

of smoke particles. We apply this extended field-equation model in analysis for the fire conditions, and compare those with experements. Our calculated values such as a temperature, velocities of gas and smoke concentration reasonably agree with the exprements.

1. THE BASIC EQUATIONS FOR THE FIELD EQUATION MODEL

Let us write the following basic equat i ons- and programs for the fluid

dynamics based on the techniques that has been developed from 1982 to 1984

.

I ) 2 > Equation of momentum,

3 ? f ( + 1

] _ s

at +

ax^ axj a^ l Jx

i

1 Equation_of continuity,

o

pi

ap a^’ot TT + a^

Equation of energy,

i! • JK Jpu;h

p + ^

at a^

dp a 3T , ap'V a(pu^yv

5

*-

5r'i7i cx 7ir ) - ( —+ Dt ax; ax;

' ' at ax

i Equation of transport and turbulent energy,

3Sq 3 pu

j.^ 3 35 3*5 3-5 3u, ,3ou, 3 ou ,

,

•iT + 1^-- 37{^^ +y ^

+ ^ ’ ** + ^

Equation of gas transport,

at ax^ axj Equation of state,

*here K - 4

1/1 ^ h “ Cp f

(1)

( 2 )

(3)

(4)

(5)

P-PRT

( 6 )

63
The basic idea of the chemical reaction and the radiation is the same

as in the previous study :>

. .However we develop the subroutines which

describe, accurately and in detail, the actual condition of a fire (such

as named of the behavior of smoke particle, the estimation of the combustion rate relating to the oxygen concentration, the ignition of the wall

and etc ).

2. THE EXTENDED F I ELD- EQUAT 1 ON MODEL

2.1. CALCULATION OF THE SMOKE CONCENTRATION

The amount of the smoke generation can be calculated by the extended

f i e I d -equat i on model with the combustion rate and smoke generation co- efficient. In the extended equation model, the movement of gas and smoke

is described by (5) and,

2 Cs , 35 Ti 2 x 1A -

it 2x. “ Jx, * (7)

Cs I smoke concentration in terms of extinction coefficient (1/m),

Cs = C/V

C I amount of smoky corpuscule per unit volume,

V : Volume of the space filled with smoke (m 3 ),

Cso ! concentration of smoke generation,

K I smoke generation coefficient,

AW 1 weight of combust i on(kg)

,

RB 1 combustion rate,

K RB! amount of smoke generation per unit time.

The definition of Cs is

Cs = 1/L LOGe (100/T), (8)

L ! beam length(m),

T ! transm i ss i v i ty (%)

Smoke generation coefficient C can be expressed as,

C = Cs * V . (9)

But because of the relations!

c

= k A W , (10)

K ! smoke generation coefficient,

Aw ! Weight of combustion (kg).

Therefore the generation rate of smoke, C can be represented as,

C = kAV,

where AW is combustion rate. As the output, Cs = C/V is more useful and

easier to understand the diffusion condition of smoke.

The amount of smoke generated from fire source is denoted by C which

is advecting and diffusing to other mesh cells. C(Cs V) should be deemed

as the number per unit volume. If the smoke inside the room does not leak,

the total amount of smoke produced by that time is written as I 0

l KAWdt

which is equal to the total amount of smoke ECi = ECsivi. The order of

magnitude of Cs concentration is about 0~30. The maximum value 30 cor- responds to the situation when the flash over occurs.

64
2.2. ESTIMATION OF GENERATION RATE OF ENERGY Ol'E THE OXYGEN CONCENTRATION

The oxygen concentration inside a flaming cell and the generation rate

of energy should be related.

Therefore, if the oxygen concentration inside the flaming cell is positive, the generation rate of energy of the calorie produced per unit

volume can be calculated.

When the oxygen concentration inside the flaming cell is negative, the

volume of calorie due to the amount of oxygen obtained from the flaming

object should be described as the generation rate of energy in the flaming

cell. Now negative concentration shows the amount of the lack of oxygen.

We show following basic equation.

(D When the oxygen concentration is positive,

HEATS = RB * QPER / (Ax) 3

, (11)

HEATS '. a generation rate of energy per unit hour and volume inside a flaming cell (Kca I /m 3 /sec)

,

RB '. speed of combustion (kg/sec)

QPER 1 a generation rate of energy per unit mass (Kca I /Kg),

AX : size of mesh (m).

(5) When the oxygen concentration is negative,

HEATS = (number of mol of oxygen in a flaming object)

% (generation rate of energy per 1 mol of oxygen)/ (Ax) 3

= RB*WP(3)/02M0L * HATL'02/( AX) 3

, (12)

HEATS T a generation rate of energy per unit hour and volume inside a flaming cell (kcal/m 3 /sec) , WP(3) I proportion of oxygen mass in a flaming object,

02M0L 1 mass of oxygen 1 mol (kg),

HATU021 released calorie volume when oxygen 1 mol is consumed.

2.3. IMPROVEMENT OF THE TEMPERATURE CALCULATION

ON THE SURFACE OF WALL

The temperature of wall surface by radiation is calculated by using

the sem i - i nf i n i te thick wall hypothesis.

The basic equation for this process is described as

T i ( t )

- T o = I o

t Q 1 (t)0(t- T )d t

n = EQi 0 n- i A t (13)

i -

1

Ti I temperature of surface of wall at the time t sec

T 0 1 temperature of outside surface of wall(from actual measurement)

Q i ( t) 1 radiation energy that the surface of wall obtains at the time t sec(Kcal/m 2 • sec : watt/cm 2

) 0 ! inditial response of surface temperature to heat input

when T=t sec, weight of getting calorie volume from t-At to t 0 i ( t - r i ) = 2*(v

r t-ri + Ati-V” t- r i )/vr n k p c (14)

65
Qi : the i th receiving calorie.

0n-i ! the i th weight function

, for the nth computation.

K ! thermal conductivity.

p density of material of wall.

C : specific heat of material of wall

2.4. IGNITION OF WALL

When the temperature of the wall surface reaches the pyrolycis temperature, the ignition is assumed to take place. Then heat flux to the

wall surface is assumed to be consumed in vaporization at the wall material.

The algorithm is described by the following 3 equations,

® SW’ = mv * HV (15)

SW’ : heat flux to the wall surface(Kca 1 /m 2 sec)

,

AHv 1 latent heat (Kcal/kg)( In case of Oak, AHv = 1341.5 Kcal/kg),

mv : total volume of generating gas at unit time per unit area (kg/m 2 sec)

.

(2) m0 2 = mv * wp(3) (16)

m0 2 • tota

I

' generat i ng volume of 0 2 per unit time (kg/m 2 sec),

WP(3) ! mass % of 0 2 in the generating gas from wall

(WP(3) depends on a substance).

© HEATW = m02 * Ax 2 /02M0L * HATL'02

^ (17)

HEATW I generating energy rate per unit time and unit volume at thecell bordering on the wall , 02M0L ! mass of 1 mol oxygen (kg/mol),

HATU02! releasing calorie when 1 mol of oxygen is consumed (Kcal/mol).

3. SIMULATION OF THE FULL-SCALE FIRE.

3.1 INPUT DATA FOR THE COMPUTATION

For the simulation of the full-scale fire in 1985, we calculate 2 ex- amples. In order to confirm the reliability of our model, we replaced the

subroutine which reasonably describes the combustion rate as shown in Fig2

-1 and F i §2 - 2 - The input data of the computation is shown in Table 1 and

Table 2.

tv Vail from jets average Qi t-At<T<r .

> y ^

At t t TIME

Fig 1 . Concept i on of Q

66
TABLE 1. CONDITION OF TWO COMPUTATION

Room Size (x,y,z)

CASE1

(2. 97m, 1.89m, 1.89m)

CASE2

(2 . 97m , 1 . 89m , 1 . 89m)

Door (width, height) (0.81m, 1 .35m) (0.81m, 1.35m)

Mesh size 0.27m 0 . 27m

Fi re source 3 mesh cells on the

floor of central

interior part

9 mesh cells on the

floor of left

interior part

Fue I Alcohol (CH 3 0H) Crib

Combustion rate

per 1 grid

RB=0. 00147/82. 5t 2 (0<t<82.5)

RB-0. 00147

(82. 5<t)

RB=0 - 14/82.

5

2 /*t 2 /15/9

(0<t<82.5sec)

RB=( (0.35/31 . 5 ) t

- 0 . 776

)/ 15/9

(82 - 5< t< 1 14sec)

RB- [ - (0 - 24/96/96)( t1 14) 2 +0/49] / 15/9

( 1 1 4<t)

X-axis

TIME (mi n)

Y-axis

RB per 15 sec (kg/15sec)

Energy release per

uni t mass 4600 kcal/kg 3600 kcal/kg

Energy per

0-) 1 mo 1

4 10KJ=97 .9kca

I

410KJ-97 . 9kca

Compos i t i on of

Combusting articles

CCD 37.52

(HO 12.62

COO 49.92

[not burning ratio]

02

CCO 43.52

(HO 5.22

COO 38.32

[not burning ratio]

132

Smoke generation 0.05

coeff i c i ent

Model

zflo

tC'

1 ' X Cl)

67
TABLE 2. Composition of the wall wood used in case 2

' fireproof board | Plywood

Thermal conductivity

(kcal /msecdeg)

(0 . 060+0 . 0009 9 )/3S00 2.667*10- 5

Density (kg/m 3 ) 400 588

Specific heat

(kcal /kgdeg) 0.2 0.38

T i me step of wall

temperature

computation (sec)

1 ~W~

Letent heat(kca 1 /kg) 1313.9

0 2

i n wal 1 wood (X) 38.3

Remark) d 1 temperature of wall wood(° C)

3.2. RESULT OF THE COMPUTATION.

CASE1

In Figure3, we show the distribution patterns of the temperature, the

atmospheric pressure variation and 0 2 concentration at the vertical section of y=0.945m and the horizontal section under the ceiling of z= 1.755m.

The patterns of every element shows results at steady state. And this is

not significantly different from the results ofour previous work. However,

as for the temperature and the 0 2 concentration, our present resluts show

that the temperature level increase and lower part of the 0 2 concentration

spread.

The atmospheric vectors are obtained for the region of the vertical

section of y=0.945m and x=1.485m and the horizohtal section of z=0.945m,

1.755m. Figure4 shows result at the steady state (90 sec). In the region

of the vertical section, the vertical stream forms the convective circulation which keeps the balance at the lower part of the center of the room. We observed the strongest stream 1.5m/s just above the fire source. Our

extended field equation model predicts the strong gush out 1.2m/s under

the handing wall at the entrance, and the inflow jet ( 1.2m/s ) in the

domain along the bottom of the region of the entrance. The distribution

patterns on the horizontal planes do not show any significant difference

comparing with our previous result.

Although the calculation of the time variation of each element was done at all meshes, we demonstrate the temperature, the atmospheric pressure, the turbulent energy and diffusion coefficient at 3 points from A to

C, and 0 2 concentration at the 4 points including the heat source D. Temperature (Figure5) becomes steady at 90sec and follows well the

rising of the combustion power. The temperature is 470° k under the ceiling and falls at the the middle point B between the ceiling and the ent- rance and further the temperature falls at the upper part of the door C. But the difference of the temperature between A and C is only 20° k. Since

the hot gas is thick and advection is fast, the decrease of the temperature

cased by diffusion is very small. The pressure deviation is also sensi68
tive to the change of the combustion and it becomes steady at about SOsec.

The level of the pressure is considerably low at the ceiling A, and it is

equivalent to the value at the middle point B and the upper part of the

entrance C. The pressure inversion at B and C (so-called a jet effect) did not

occur(Figure6). Also the change of the turbulent energy (Figure7) and the

diffusion coefficient (Figure8) are consistent with the tendency of the

temperature and the atmospheric pressure. At the ceiling A all of the two

quantities are larger than the ones at other places, and they become comparable at the entrance C and the middle B. Although, we do not observe the

inversion of turbulent energy and diffusion coefficient at steady state

our results of Figures suggests that the inversion would appear at the

time from 10 sec to 20sec. We expect the gush out associated with the

i nvers i on

.

The O 2 concentration decreases to about 18* at the heat source D but

the O 2 concentration decreases only 2 % near the door at A~(„ . ( F i gu re 9)

F i x3 Vertical and horizontal cross section of te*oera ture

,

pressure deviation and 0i coocen tn ti on In Case 1 (T=90.327sec)

H, 5 Tl«* t.H.U — a I [emm-TKlMrr U OkSEl

7Lvi;un.7i(i) P8isswe ormTicwoi/. 1 ) o, ca.-iCiAmrio.'K:)

Fi *4 vertical and horizontal cross setion of flow field in CLScl

1 unit vector jhovs t.0*/sec (Ta90.327sec)

!• f I t*< I I IX «* «

T ' li T 1 •« #/ trrrrurr Jr* » ClSTl

fi|7 Ti*« uriiiiM turtle*! mrn CJLST1 v.

r 10»

0 2 O C 3} K33

' (Sto

69
CASE2

Supposing that a wooden furniture (example, a desk) is placed at a cor- ner of a room, we calculate until the combustion peak.

The distribution of the temperature, the pressure deviation, the 0 2 concentration and the smoke concentration are described in a plammetric

map (vertical section of y=0.675m, 2=1. 755m). FigurelO shows the distribution of each elements at the combustion peak excluding the smoke con- centration. Comparing with the results of the easel, the temperature is

100° K higher under the ceiling and 50° K higher at the upper part of the

entrance. The pressure deviation shows almost the same level as in the

easel around 60 sec but both of the pressure deviation under the ceiling

and at the floor of the entrance increase about lN/m 2 at 90sec, which indicates the increase of the pressure. The 0 2 concentration of case2 is

similar to that of easel, but it is 1* higher approximately. The decrease

of concentration was less significant than the increase of the temperature. This is because of follows different quantities between easel and

case2, namely the place of the heat source, the mesh area for the heat source and furthermore combustion rate.

The distributions of the smoke concentration at 30, 60, 90 sec are shown in Figurell. We describe that i.t is all concentrated on the heat source inside the cell at the beginning, and it diffuse to the middle r of the entrance. Finally it gets higher than the one at the center of tut room.The movement of gas is shown for the vertical section of y=0.675m, the

vertical section of x=1.485m and the horizontal place of z=0.945m,

2 = 1 . 755m. Figurel2 describes the flow cond i t i oh of the combustion peak.

Since the fire source is placed at left side of the back cell, the unsymmetrical pattern appears. A right revolving vertex appears even on the

vertical section x= 1.485m. The velocity has the maximum of 2.0m/s above

the heat source and 1.8m/s at the upper part of the entrance.

We notice that a circulating flow (faster than lm/s) which forms a circulation from the lower part of. the entrance to the heat source and

heads to the upper part of the entrance just under the ceiling.

The time variaton of the temperature, the atmospheric pressure variation, the turbulent energy, the diffusion coefficient are the same as in

easel. We show their values at the point of A~C and the oxygen concentration at the point of A~~D. The temperature (Figurel3) rises hastily and

rises over 700° K just beneath the ceiling after 80 sec. The atmospheric

pressure variations (Figurel4) decrease at every points and get closer,

which means that the pressure of whole cell tends to be higher. The turbulent energy and the diffusion coefficient ( Figurel5, Figurel6) are rising hastily at the combustion peak. Also, at around 20 sec, the turbulence of the entrance slightly exceeds the under-ce i 1 i ng and the turbulence becomes bigger as it gets closer to the fire source and the condition of the local compression, so-called the entrance jet is not seen. The

0 2 concentration ( F i gu re 1 7) is 14* at the source, and 22* from the underceiling to the entrance.

70
I te*o<r»lure in CXSL2

H«I7 Ti»« nri»tion of Oi concentration

71
The temperature of the wall attaching to the furniture is higher than

600° k, and approximately 360° k at the opposite side, and 550° k at the

ceiling above the heat source and 760° k at the floor just under the furniture.

5. RESULT

We check the reliability of our model by using the same size room and

the same combustion rate as the ones of the experiments. As for the temperature under the ceiling, the gas velocities at the entrance and the 0 2 concentration under the ceiling, we obtained the reasonable values. But

we could not compare the time of the wall ignition, since the experimental of the time. 6 the wall ignition was not available at that moment.

We will calculate the time of wall ignition by our extended model with

the result of the new experiment. And we will add a new model which describes the situation when the door or window open suddenly by some accident.

REFERANCE

1) Takeshi Handa, K.Kawagoe, T.Yoshikawa, J.Mashige, T . Joh ! Numer i ca 1 si'

lation of early stage of a compartment fire, Fire Science and Techn;

gy, Vo I .4, No. 2, 91 , 1984.

2) Hasemi Y.! Numerical calculation of the natural convection in the

compartment, BRI Research Paper, No. 69, 1977.

72
CRITICAL DISCUSSION ON FIRE DEVELOPMENT

Toshisuke HIRANO

Department of Reaction Chemistry, Faculty of Engineering

The University of Tokyo

7-3-1 Hongo, Bunkyo-ku, Tokyo 113

ABSTRACT

A discussion has been made on the basic knowledge to be needed for the

understanding and prediction of fire development. It is pointed out that

for only limited cases, fire development can be predicted by presently

available data. Further, the prediction of fire development required for

the assessments of various fire hazards and the improvement of fire protection engineering is discussed, and the need of a stochastic expression is

emphasized.

1 . INTRODUCTION

Every fire hazard assessment is more or less based on the prediction

of fire development. For the evaluation of evacuation time, the rate of

fire development and the smoke behavior during fire development are needed.

For the recommendation of fireproof materials, the difference of

ignitabilitie s or flame spread rates under the conditions of fire

development should be interpreted. For other types of fire hazard

assessments, also the prediction of some aspects of fire development would

be required. Thus, a clear understanding of fire development seems to be

essential to reduction of the damages to be caused by fires.

In the prediction of fire development, a model is used, the quality of

which depends on the amount of knowledge of phenomena, and the reliability

of the predicted results apparently depends on the accuracies of available

data. The data usually used to predict fire development are concerning

various phenomena, such as ignition, flame spread, smoldering, burning zone behavior, fire plume, smoke behavior, heat release, radiation, convection,

ventilation, and so on. These data have been accumulated by a number of

previous studies [ 1 -9 ] • The data concerning some of these phenomena are

sufficient for the prediction of fire development, while those concerning

other phenomena are insufficient.

For the last decade, a number of excellent computer programs for the

prediction of fire behavior have been established. Consequently, an

accurate prediction of fire development seems possible if needed data are available and initial and boundary conditions are gi v en

[

1

0

- 1

4

]

.

Furthermore, even if the prediction were accurate enough, the predicted

result for a set of initial and boundary conditions would be only an

example. In general, it is doubtful that the predicted result is for the

most probable fire development, which is really required for fire hazard

assessments [ 1 5-17]

.

The objective of this paper is to indicate needed data for the prediction of fire development under arbitrary conditions, and to reveal the

necessity of appropriate prediction methods to explore the most probable

phenoemena. The dicussion presented in this paper will give us future

problems to be solved in fire research.

73
2. NEEDED KNOWLEDGE ON FIRE DEVELOPMENT

Figure 1 shows a block diagram representing the processes of fire

development. For the understanding or prediction of fire development,

knowledge of each phenomenon in Figure 1 is indispensable. Basically, our knowledge of a phenomenon should be based on presently available data

concerning it, so that to understand fire development to a further extent,

the accumulation of available data is necessary. In this section, the data

to accumulate urgently are discussed.

Ignition is the first process at fire occurrence. A large number of

studies have been carried out on this subject[ 1 -3,6,8] , and the mechanisms

of various types of ignition have been revealed. Also, quantities

representing ignition characteristics, such as ignition temperature,

minimum ignition energy, and ignition delay time, have been evaluated or measured. The results of theoretical and experimental studies on this

subject were summarized in several review papers [ 1 -3 >6,8 ] . A large number

of data are available.

However, if the initial and boundary conditions are actual ones, the

ignition characteristics will be not always predictable. It will be found

hard to predict the phenomenon that the ignition delay time at radiative

ignition depends on the orientation of the irradiated combustible material

surface as shown in Figure 2. Such a difficulty is attributable to lack of

data on the solid combustible ignition under natural convection. This

example seems to indicate that more data are needed for ignition under

various initial and boundary conditions. At present, the ignition process

under actual conditions is not always predictable. This may be the reason why the ignition process is not considered in the practical prediction of

fire development.

The next stage of fire development is flame spread. Flame spread

under various conditions have been examined and appropriate models have

been proposed[3,4]* The results of previous studies on this subject have

been summarized in several review papers [ 3 ,4, 6,7 ] . The mechanisms of

various types of flame spread have been explored and a large number of data

have been accumulated. The studies on this subject seem to be at the most

advanced stages among those on the subjects concerning fire development.

In spite of the accumulation of a large number of data and advanced

stages of studies, many ambiguities remain as yet. Flame spread under

particular conditions is not always predictable. In the case shown in

Figure 3, the flame spread rate depends on the both characteristics of

liquid and porous solid bed and no reasonable model for the prediction of

flame spread rate has been proposed. This example implies that the flame

spread rate over a material with complicated characteristics is hardly

predictable. The parameters to be considered in additiion to the material

characteristics for the prediction of the flame spread are the thickness,

convection, radiation, surface orientation, and thermal properties of the

material. Although the effects of these parameters on the flame spread

have been studied extensively, more data on particular types of flame

spread seem insufficient for the understanding or prediction of flame

spread under arbitrary conditions.

Smoldering and smoldering-flaming transition are important phenomena

to ubderstand early stages of fires as pointed out in the previous paper

[6]. However, available data concerning this subject are only a few. The

smoldering-flaming transition is a subject to be urged to study because the

processes of fire development following to it depend on it. Another

important problem is to examine the relation between the product composition and ambient conditions under which smoldering occurs. Very likely to

74
the flame spread, almost the same parmeters affect the smoldering, such as the thickness, convection, radiation orientation, and physical and chemical

properties of the material.

The rate of the combustion product generation must -be the knowledge

indispensable for fire modeling. However, little effort has been made to

predict it in a practical case when a piece of furniture, wall, ceiling, or floor is burning. Although needed data are on those various types of

burning zone behavior, most previous studies on the modeling of burning

zone behavior have concerned with pool burning!! 51-

Pool burning behavior is closely related to the fire intensity in a room and greatly affects the fire plume and smoke behavior. The rate of

burning depends not only on the burning zone size but also on its

configuration. However, very few data are available on the pool burning of

the configuration other than circular although in an actual fire the

burning zone configuration is not circular. Further, the burning rate in a small size pool should be distributed on the burning zone surface, so that

the burning rate at the center would be different from that near the

leading edge of the spreading flame.

The fire plume and gas flow inside a compartment are the subject

convenient for the computer simulation, so that many studies have been

performed on the prediction of these phenomena by using high-speed

computers [7] . Consequently, a number of excellent programs for the prediction of the fire plume behavior have been established. However, the

programs applicable to the prediction of fire induced gas flows in a large

size compartment are still very few. The effect of the mixing inside the

high temperature zone should be considered as pointed out in the previous

papers[6,7]. The gas flow behavior in a long corridor or tunnel is also

the problem to be solved because it is very important for the fire prevention and protection activities. This problem seems to be worth to investigate, although it is not easy one including the heat transfer between the

flowing gas and the wall.

3. IDEAL PREDICTION OF FIRE DEVELOPMENT

If the understanding of phenomena concerning fires were complete, a quantity H representing fire hazards would be evaluated under a given set

of initial and boundary conditions, i.e., H would be given by an equation

as follows:

H = f(t,x1 ,x2 ,x

3 ,y 1 ,y2

, z

1 ,z 2

, ), (l)

where t is the time, x^,x 2 ,Xo are the distances along the coordinates

representing each point in the space, y 1

, y 2

,... the initial conditions,

and z-pZ 2

,.... are the boundary conditions. In these conditions, the

values representing the material characteristics and space structure under

consideration must be included.

For an ideal case when f is given as a simple function of a variable

w, a critical value w

c of w can be found by assuming the value of f

c that

is a criterion of the safe and danger(Figure 4)- For a practical case, the

situation is different. The requirement for fire hazard assessments should

be not to find such a criterion but to find the optimum conditions to

minimize the total expense E for fires(Figure 5)- E is the sum of the loss

L to be caused by fires and the cost C of fire protection, and both L and C shuld be related to f,i.e.,

E = L + C = T

1 (f) + T

2

(f) , (2)

75
where T-j(f) and T

2

(-f) are equal to L and C, respectively, expressed using

the transformation equations relating L and C to f. The optimum value of a quantity w of the initial and boundary conditions(y.p y 2

» ...., , x 9 , ...) to minimize E can be found to solve the

following partial differential equation,

BE/aw = B(T 1 (f))/sw + a(T2 (f))/aw. (3)

The value w

Q

that is w to minimize E as shown in Figure 5 is eventually

required for fire protection engineering.

The effort to obtain w

Q

is equivalent to that to design for the

optimum fire protection. However, to do so is impossible because complete

understanding of phenomena concerning fires is unconceivable, so that the

details of the function f cannot be expected to determine. Further,

available data are not always appropriate to deduce E as mentioned in the

previous section. What we can do at present is to propose a model

represented by an approximate equation f instead of f and to estimate an approximate value w

aQ instead of w

Q on the basis of our knowledge on fire

development.

f to be deduced from-any one of presently available models is a function of time and space by which fire hazards can be predicted under a given set of initial and boundary conditions. This may be a reason why

the predicted results are not satisfactorily coincide with those expected

by other persons. The initial and boundary conditions assumed in a prediction seem to be different from those assumed in other prediction. In

actual cases, however, the initial and boundary conditions are variant, so that f based on a presently available model seem not suitable -to find w

aQ . The difficulty to find w aQ on the basis of the predicted results is

partly attributable to the uncertainty of selection of a set of initial and

boundary conditions. It is not easy to find a suitable set of conditions

among a large number of conceivable sets of conditions. For example, there

may not be generally acceptable reasons to decide the conditions of opening

and flammability. Both conditions obviously depend on the human behavior,

weather, and other known and unknown factors. In these situations, each

condition should be expressed stochastically[18,19l , and the concept of the

most probable condition is acceptable.

It may be recognized that there is a set of limiting conditions under

which the circumstances are kept to be safe against fires. These limiting

conditions are required for the fire protection engineering such as the

evacuation planning, detector development, building design, and development

of the technique for fighting fires. Each limiting condition may not be

determined uniquely and should be expressed stochastically, too.

For the hazard assessments, a specific value of f representing a criterion of the hazard and the mathematical expression of a function f

should be required. In the mathematical expression of f based on the

stochastic expression of the initial and boundary conditions, the relations

of these conditions to the fire development phenomena should be adopted.

Since the conditions are expressed stochastically, stochastic data of

phenomena concerning fire development should be required.

4. FUTURE RESEARCH ON FIRE DEVELOPMENT

In order to predict fire development in an appropriate accuracy, the

initial and boundary conditions to be assumed for the prediction of each

phenomenon should be given as accurate as necessary. Even if the initial

76
and boundary conditions were given accurately, the penomenon cannot be

predicted without an excellent model and a large number of data. Further,

knowledge of fire development is indispensable for the prediction. This

situation can be easily seen from the fact that only a few persons can predict the aspects from the time being under igni table conditions to

ignition although such aspects are very important information.

As discussed in Section 2, a large number of data under various conditions are needed. Important are the fundamental studies to accumulate

accurate and useful data. The test methods of material for fire protection

are also very important to understand the properties of materials during

fires. Also, it is desirable that the material properties under various

fire conditions can be inferred from the results of tests.

Other important discussion made in Section 3 is that by- the prediction

to be required for the fire protection engineering, the limiting values

should be derived from the data expressed stochastically. Thus, data should

be accumulate to stochastically express the physical or chemical

characteristics of fire development. Also, a model is required, based on which a stochastic expression of the function f is possible.

The probability density function g. of flame spread rate Vp under a certain set of conditions can be expressed as shown in Figure 6(a). If the

probability p^ of circumstances 'under the kth set of conditions is known as well as g v j£

» then g v

is given as follows:

where n is the conceivable nmber of the sets of conditions, and g v is

supposed to be expressed as shown in Figure 6(b). Further, the burning

zone area A^ can be calculated on the basis of g y and is supposed to be

expressed as shown in Figure 6(c).

The data needed in the above process to evaluate A-^ are those on flame

spread rate under various sets of conditions and p, , although both

requirements are not easy to satisfy. Although the data ho be required for

the prediction of other phenomena will be much more hard to obtain, we have

to make an effort to satisfy the requirements for the prediction really

needed for fire hazard assessments.

1. Kanury, A. M., "Ignition of Cellulosic Solids - A Review," Fire

Research Abstravts and Reviews, 1_4> PP* 24-52(1972).

2. N.F.P.A., "National Fire Codes," Vol. 1 3 » NFPA(1979).

3. Akita, K. and Hirano, T., "Ignition of PMMA and Flame Spread over Its

Surface," Saigai-no-Kenkyu, V1_, pp. 291-297(1980).

4- Fernandez-Pello, A. C. and Hirano, T., "Controlling Mechanisms of Flame

Spread," Combustion Science and Technology, 32 , pp. 1-31(1983).

5. Orloff, L., "Simplified Radiation Modeling of Pool Fires," Eighteenth

Symposium( International ) on Combustion, pp. 549-561, The Combustion

Institute, Pittsburgh ( 1 981 )

.

6. Hirano, T., "Combustion Phenomena at Early Stages of Fires," Proceedings

of the Eighth Joint Panel Meeting UJNR Panel on Fire Research and

Safety, pp. .214-227, Building Research Institute, Ministry of

Construction, Japanese Gov. (1985).

7. Zukoski, E. E., "Fluid Dynamic Aspects of Room Fires," Fire Safety

Science, Proceedings of the First International Symposium, pp. 1-30,

n

(4)

k=l

REFERENCES

77
Hemisphere, Washington ( 1 986)

.

8. Emmons, H. W. , "The Needed Fire Science," ibid, pp. 33—53

•

9. Friedman, R-, "Some Unsolved Fire Chemistry Problems," ibid, pp. 349-

359.

10. Emmons, H. W., "The Prediction of Fires in Building," Seventeenth

Symposium(International) on Combustion, pp. 1101-1111, The Combustion

Institute, Pittsburgh(l979).

11. Wakamatsu, T. and Hasemi, Y., "Progress in Fire Growth Modeling in

Japan," Proceedings of the Eighth Joint Panel Meeting UJNR Panel on Fire Research and Safety, pp. 125-135, Building Research Institute,

Ministry of Construction, Japanese Gov. (1985).

12. Rockett, J. A. and Morita, M., "The NBS/Harvard Mark VI Multi-Room Fire

Simulation," NBSIR 85-3281, U. S. Department of Commerce, National

Bureau of Standards ( 1 986)

.

13- Mitler, H. E. and Rockett, J. A., "How Accurate is Mathematical Fire

Modeling?" NBSIR 86-3459, U. S. Department of Commerce, National Bureau

of Standards (-1986).

14. Tanaka, T., "Present Status of 3uilding Fire Modeling and Review; Parts

1 & 2," Journal of the Japanese Association of Fire Science and

Engineering, 36, pp. 1-5(Vol. 5); pp. 1-9(Vol. 6) (1986).

15. Seeger, P. G. (Session Chair ), "Peopl e-Fire Interactions," Fire Safety

Science, Proceedings of the First International Symposium, pp. 495-590,

Hemisphere, Washington( 1 986)

.

16. Fitzgerald, R. W. (Session Chair ), "Trans lation of Research into

Practice," ibid, pp.^ 591-676.

17. Kersken-Bradley , M. (Session Chair) , "Statistic, Risk, and System

Analysis," ibid, pp._ 941-1056. - 18. Morishita, I., "A Stochastic Model on Fire Spread-A trial for

Evaluation of Fire Spread Risk," Proceedings of the Eighth Joint Panel

Meeting UJNR Panel on Fire Research and Safety, pp. 35-42, Building

Research Institute, Ministry of Construction, Japanese Gov. (1985).

19. Lin, W-C. T. and Williamson, R. B.,"The Use of Probabilistic Networks

for Analysis of Smoke Spread and the Egress of People in Buildings,"

Fire Safety Science, Proceedings of the First International Symposium,

pp 953-962, Hemisphere Washington(l986).

78
Evacuation

00c

o

cn

co

<u

co

oo

coo

u

.co

0)

cooo

79

Figure

1

. A

block

diagram

of

fire

development.
Figure 2. Ignition delay times at radiative heating ofhorizontal and vertical PMMA surfaces

.

Figure 3* Position-time (x-t) diagrams representing the

movements of leading flame edges. y s

: distance

from the bead layer surface to kerosene surface

d: bead diameter.

80
Figure 4. An illustrative example of the variation of f.

Figure 5- Illustrative examples of the variations of TMf),

T

2 (f) , and E.

81
(a)

(b)

(c)

Figure 6. Probability density functions(p.d.f.) of flame

spread rate V^. and burning area A^. (a) the p.d.f.

of *V r» under the kth set of conditions; (b) the overall

p.d.r. of V^.; (c) the overall p.d.f. of A^.

82
RADIATION FROM DIFFUSION FLAMES-SMOKE POINT

John de Rls

Factory Mutual Research Corporation

9th Joint Panel Meeting of the

UJNR Panel on Fire Research and Safety

May 4-8 , 1987

ABSTRACT

A wide-range of experiments provide extensive empirical evidence that the

laminar smoke-point value of a fuel provides a fundamental chemical measure of

a fuel's propensity to radiate and release hazardous products of combustion

from turbulent flames. It is also shown that the parameter 1/(BC +1) correlates the overall flame radiation from turbulent ethane wall fires. In addition wall-fire flame radiation measurements versus height show that the flame

radiation increases proportional to height in the lower fuel-rich portion of

the flames. This simple proportional ity is attributed to the absence of soot

oxidation in the lower flame region.

INTRODUCTION

The flammability of a material is strongly influenced by its flame radiation. At hazardous fire scales the radiative heat-feedback from the flames to

the pyrolyzing fuel surface is generally significantly greater than the con- vective heat-feedback. Fuels which release a greater fraction of their chemical energy in the form of thermal radiation (rather than convection) are generally more hazardous, because it is easier for radiation to return to the

pyrolyzing surface. At present, the two principal impediments to predicting

fire heat release rates and upward fire spread rates - namely solving the

flammability problem - are: 1) our lack of scientific knowledge of flame

radiation, and 2) the absence of some practical test method for evaluating the

flame radiation from materials of interest. In general, the relative impoi

tance of flame radiation increases with fire scale, so that it is impractical

to assess a material's flame radiation using conventional small-scale rate-ofheat-release test apparatii which, by imposing significant levels of external

radiation (to simulate large-scale combustion conditions), overpower any

direct flame radiation measurement . Instead, we shall have to become more

clever in our search for a practical method for measuring flame radiation.

83
This requires, however, a deeper scientific understanding of flame radiation

which is the subject of much of Factory Mutual's current flammability

research.

SMOKE-POINT CORRELATIONS

About four years ago Markstein 1 discovered that the radiant fraction from

buoyant turbulent fuel-jet flames burning in air is directly correlated by the

fuel's laminar smoke-point flame-height, as shown in Figure 1. The radiant

fraction itself from a buoyant turbulent fuel-jet is quite insensitive to the

overall flame size or (heat release rate) when the heat release rate is above

a f ew kilowatts. The smoke-point is defined as the maximum flame height for

which a laminar "candle-like" flame just does not emit soot from its flame

tip. The smoke-point of a fuel characterizes its tendency to produce soot.

Fuels with long smoke-point heights produce less soot than fuels with short

smoke-point heights. The smoke-point is a rather easy measurement to make for

liquid or gaseous fuels for which one can manually adjust the flame height by

controlling the rate of fuel supply to the flame whose height is proportional

to the chemical heat release rate. In retrospect, Markstein' s discovery is

not surprising because flame radiation comes predominantly from the soot

formed in the flame. For a given fuel burning in a small laminar buoyant

"candle-like" flame beneath its smoke-point height, the peak concentration of

soot is proportional to the flame residence time which, in turn, is proportional to the square root of the flame height. By changing the height of a laminar flame, one controls its flow residence time, x res » and its corresponding Damkohler number T re3 / T Ch ein which controls the influence of the flame

chemistry. As the laminar flame-height increases, it loses an increasing

fraction of its chemical heat-release in the form of radiation until the soot

at the flame tip is cooled below its critical oxidation temperature and un- burnt soot is released from the flame-tip as a result of radiative extinction.

Markstein 1 has shown that for hydrocarbon fuels burning in air at their smokepoint condition, about 301 of their chemical energy is released in the form of

radiation. For turbulent buoyant flames, as distinct from a laminar flames,

the turbulent microscale flow time (e.g. Kolmogorov time x - [vfi/u^] 1 // ^)

controls the final combustion processes and is relatively insensitive to the

overall fire heat release rate; so that those hydrocarbon fuels whose turbu84
lent buoyant flames (burning in air) release more than 30$ of their energy in

the form of radiation (such as propylene) also tend to release copious amounts

of smoke; whereas, those hydrocarbon fuels which release less than 30$ of

their energy in the form of radiation (such as propane) tend to release relatively little smoke from buoyant turbulent flames.

In a related study Tewarson measured the chemical energy fraction actually released, as well as the fraction of convecti ve heat release, xq> from

small turbulent pool fires for a variety of (noncharring ) solid and liquid

fuels indicated in Figure 2. From these two measurements he inferred the

fraction of radiative heat release, xr> difference. He also measured the

associated laminar smoke-point flame heights by vaporizing the various fuels

in an electrically-heated Erlenmeyer flask. As can be seen in Figure 2, his

correlation of radiative fraction for condensed fuels is almost identical to

2 Markstein's correlation for gaseous fuels. In addition Tewarson found that

the smoke-point also correlates the release of smoke, carbon monoxide and

unburned hydrocarbons for the same wide range of fuel types.

In another study on the blockage of radiative heat transfer to the base of

large (38 cm and 76 cm diameter) pool fires Markstein^ found that the peak

centerline soot concentration is also correlated by the smoke-point for gaseous fuels. Finally Kent-’ very recently has shown that the fuel smoke-point

correlates the peak soot concentration in pure momentum driven (e.g. forced

convective) turbulent fuel-jet diffusion flames for both aromatic and aliphatic fuels burning in air. Thus, we now have extensive empirical evidence that

the laminar smoke-point value of a fuel provides a fundamental chemical measure of a fuel's propensity to radiate and release hazardous products of

combustion from turbulent diffusion flames . This result can now guide the

development of a practical flame-radiation test method.

We note here that all the above empirical correlations involved burning

pure fuels in air with similar adiabatic flame temperatures and stoichiometric

mass ratios of ambient oxidant to supplied fuel. To obtain a more general and

deeper scientific understanding we need to extend our knowledge to; fuel

mixtures, fuels diluted with inerts such as N

2

, fuels diluted with products of

combustion such as C0 2 and H

2 0 which can chemically oxidize soot, and oxidants

having significantly enhanced or depleted oxygen concentrations as compared to

normal air. While such studies are of obvious intellectual interest in pro85
viding a more firmly based scientific understanding of flame-radiation, they can also yield important practical dividends. For example, we know that flame

radiation increases much more strongly than convective heat transfer with

ambient oxygen concentration. It is thus likely "that by performing standard

rate-of-heat release tests in different ambient oxygen concentrations, one may

be able to separately identify the radiant fraction of a burning material and

thus determine its expected flammability at large scale. Such a determination

will require a firmer basic scientific understanding of flame radiation.

Presently, FMRC researchers are exploring this possibility by measuring the

increase in flame radiation with ambient oxygen concentration for various

fuels in both laminar and turbulent flames. We will report on these results

at the next UJNR meeting.

WALL-FIRE COMBUSTION MODELING

Figure 3 shews the classical Shvab-Zeldovi ch conception of a diffusion

flame, where fuel diffuses toward the flame from the right and oxygen diffuses

i'n from the left. The flame position is determined by the stoichiometric

ratio, S, of the mass of oxidant mixture required to react with unit mass of

fuel mixture. It is presumed that both fuel and oxidant cannot exist instantaneously at the same position; because if this were not the case, the lesser

reactant (either fuel or oxidant) would be immediately consumed by the dominant reactant, thereby eliminating the lesser reactant. For diffusion flames

the chemical reaction rates are much faster than diffusion rates, so that the

overall rate of combustion is determined by diffusion of reactants into the

flame. For the Shvab-Zeldovi ch conception of a diffusion flame the rate of

heat release is proportional to the rate of reactant diffusion into the flame.

If one assumes that the diff usi viti es of the various species are the same and

also equal to the thermal diffusivity, then the peak flame temperature is

equal to the adiabatic stoichiometric equilibrium flame temperature for the

reactants provided they are completely consumed and there is no loss of energy

by radiation. One can mathematically show that the above theoretical conception holds for both steady and transient as well as laminar and turbulent

diffusion flames.

In reality, diffusion flame chemistry is considerably more complicated.

Fuel and oxygen molecules do not simply combine to form products with an

86
associated release of heat. Instead, numerous chemical intermediates are

£’ n formed and then consumed. Nevertheless, both theory 0 and experiment show

that the concentrations of all major species as well as lesser species such as carbon monoxide and molecular hydrogen (to a certain extent) can be correlated

by a mixture fraction variable such as

£

= S Y + 1 - F S + 1

Y /Y

o o°>

which is neither created nor destroyed by the chemical reactions. This type

of mixture fraction variable has the particular advantage of yielding straight

lines in diagrams for species concentrations in those regions where a particular species does not undergo chemical reaction. This allows one to use geometrical constructs for interpreting flame relationships

.

One can readily extend this simple Shvab-Zeldovich description to diffusion flames having incompleteness of combustion and release of heat by radiation, provided one assumes that: 1) a fixed fraction, 1-x A

, the supplied

fuel does not participate in the chemical reaction and acts like an inert

associated with the incompleteness of combustion, and 2) that a fixed fraction, Xr* °Y the theoretical ,(e.g. total) heat of combustion immediately

leaves the flame zone in the form of radiation. These assumptions lead to the

following definition of convective B-number

B = c q

c

S X, - C (T - T )] p S ® J

where q" is the convective heat transfer rate to the fuel surface and m" is

c the mass transfer of fuel from the surface. Here AH C is the theoretical net

heat of combustion, Tg is the surface temperature and T^ is the ambient reference temperature associated with the definition of AH C

. The dotted lines in Figure 3 show the temperature (or to be more precise,

the sensible enthalpy) profiles versus mixture fraction for fuel supplied from

a fuel-jet (B =® or <3"=0). If instead the same fuel is introduced to the fire u C from a pyrolyzing wall with a mass transfer rate m" and having a resultant

convective heat feedback rate <3" then the actual fuel concentration observed

c immediately adjacent to the fuel surface is reduced by the factor B_/(B +1) as c c indicated in Figure 3. This reduction comes from the back diffusion of pro87
ducts of combustion which, while not actually being absorbed by the wall,

nevertheless are present at the wall surface with a total concentration

1/(BC

+ 1) = 1-B /(B

c

+ 1.) adjacent to the fuel surface. Finally the introduction

of a cooler (or less hot) fuel surface can cause a reduction in the flame

temperature (or sensible enthalpy) as suggested in Figure 3*

These theoretical considerations suggest that the introduction of a pyrolyzing fuel surface providing a mass transfer, m" , and receiving a convective

heat feedback, <5” , should be equivalent (from a Shvab-Zeldovich viewpoint) to

supplying the flames from a fuel-jet supplying fuel at a concentrati on of

B„/(B

C +1) together with products of stoichiometric combustion with a total

concentration of 1/(B n

+ 1), all at a temperature equal to the surface temperature, Tg. By so doing the buoyant turbulent fuel-jet should replicate the

conditions at the fuel surface. By using this procedure one hypothetically

also replicates the full detailed fuel decomposition chemistry as the fuel

heats up while diffusing toward the flame surface. We are currently setting

up an experiment at FMRC with NBS sponsorship to test this hypothesis by

measuring the radiant fraction from fuel jets suitably diluted with same fuel's products of stoichiometric combustion generated in an auxiliary burner.

We anticipate that the addition of gaseous and CO 2 to the fuel supply will,

reduce the soot formation through both a reduction of the flame temperature as well as by chemically enhancing the soot oxidation through increased concen- trations of H

2 O and CO2 which are known to augment soot oxidation at high

temperatures

.

WALL-FIRE FLAME RADIATION MEASUREMENTS

Markstein^ has recently performed a series of flame radiation measurements

using a 38 cm wide by 1.83 n high wall burner producing two-dimensional flames

through use of 15 cm deep side-walls to prevent sideways inflow. Most of the

measurements involved use of a series of three 22 cm high water-cooled sintered-metal gas panels supplying ethane fuel-gas to the flames at a controlled

mass transfer rate, m" . A water-cooled heat transfer plate was present above

the sintered-metal gas burners. Measurements were also made for ethane supplied from a narrow slot burner at the base of the wall to simulate the case

of B ->® (or 6"-0). For each fire the total heat feedback rate to each wall

^ c section was measured by measuring the temperature rise (i.e. heat absorption

88
rate) of the cooling-water. Two types of flame radiation measurements were performed. One used a wide-angle radiometer to record the total flame radiation emitted by the fire; while the other used a scanning si it-radiometer to

record the flame emission versus height. Since the boundary layer flames were optically thin, it was assumed in all cases that half the flame radiation

returned to the wall and the other half radiated out into the laboratory away

from the wall. The convective heat transfer back to the wall was inferred for

each panel section by comparing the measured total heat feedback with the

associated radiative component by integrating the scanning radiometer measure- ments over the corr esponding panel section.

We are presently replacing the wall burner because it is only suitable for

using relatively non-sooty fuels such as ethane. We are also considerably

improving the spatial resolution of the scanning radiometer. Thus the mea- surements discussed here should be regarded as preliminary. Nevertheless,

they are rather interesting and can serve to guide the development of upward

fire spread models.

Figure 4 shows the total radiant fraction, Xr> from the wall flames plotted against 1/(B +1) which we showed in the previous section is the total

concentration of combustion products (chiefly ^0, CC^'and N

2

) at the wall

surface. One sees that the radiant fraction decreases linearly with increasing presence of combustion products. We also see that 1/(B

C +1) provides a reasonable correlating parameter which was anticipated by the theoretical

arguments of the previous section.

The radiant fraction for the slot burner was 19%, independent of the flame

height or fuel supply rate. This 19% fraction should be compared to the

radiant fraction of 21% measured for ethane buoyant turbulent fuel-jets burning in the open without an adjacent water-cooled wall. Thus the presence of

the water-cooled wall over the entire length of the wall flames results in

only a slight reduction in their radiant emission. Apparently, the overall

emission from the flames is much more sensitive to how the fuel is supplied to

the fire than any subsequent convective cooling by an adjacent wall surface.

Figure 9 shows the observed flame height as interpreted by the scanning

radiometer when the flame radiation dips below a certain arbitrarily chosen

threshold value. The flame heights seem to correlate reasonably well with the

heat release rate per unit burner width, as suggested by the theory

89
based on dimensional analysis for a turbulent line-fire. These threshold

flame height measurements do not appear to be sensitive to the pyrolysis zone heights

.

We obtain a rather different perspective when we examine the actual radiant flux distribution measurements shown in Figures 5 through 8. All these

flux distribution curves are very well correlated by: 1) normalizing the

'

i /p height, z, in the abscissa by Lj.=0.17(Q') where the flame radiation goes to

zero, and 2) normalizing the flux distribution, q"(z) in the ordinate by

• R

X d Q'/L„ so that the area under each curve is exactly equal to one-half; since

only one-half of the total flame radiation returns to the wall surface.

Alternative correlations of radiant flux distributions using the conventional

* 2/3

L - (Q r ) were much poorer. This surprised us. However, we note that the

1 * 2/3 prediction that L - (Q)~ is based on arguments relating to the rate of

chemi cal heat release. Here we are, instead, considering measurements of

flame radiation which is not expected to be locally proportional to the rate

Q of chemical heat release. In fact, from other studies , we know that most of

the chemical energy from turbulent diffusion flames is released considerably

earlier than the radiative energy. The soot is formed where the majority of

the chemical heat release occurs, and then continues to radiate until it is

completely oxidized in the upper flame region of soot oxidation.

DISCUSSION OF THE WALL RADIANT FLUX DISTRIBUTIONS

Examining Figure 5 for the slot burner more closely we see that

S (z)

dz

X rQ' m = (0.17)

where the initial slopes of the curves, m, are effectively the same in all cases. This initial slope seems to depend only on Xr and be independent of

both heat release per unit flame width, Q' and the number of burners. In this

region the flame is fuel-rich so that the rate of chemical heat release depends only on the rate of turbulent air entrainment into the flame as estabq lished by Delichatsios . It typically takes about fifteen mass units of air

to react stoichiometrically with unit mass of fuel. As each clump of air is

engulfed into the flame in this lower region, it is soon "eaten up" by the excess fuel. Because of the large stoichiometric mass ratio, S, both the

90
diffusion flame and fuel immediately penetrate into the air clump. The soot

which is formed cannot diffuse because its enormous molecular weight causes it

to have zero diffusivity. This soot is left behind in a hydrocarbon-rich gas

where it cannot be oxidized. The subsequent Lagrahgian evolution of this soot

volume cloud follows almost identically to the original entrained air clump

volume evolution. Thus, we expect the rate of soot formation to be proportional to the rate of combustion in the lower flame region, while the rate of

soot oxidation is probably negligible in this same lower region. The rate of

q combustion per unit area, as established by Delichatsios , is proportional to - 1/2 the upward velocity u-z , so that the accumulative upward flow of soot at a given height is proportional to z^ /2 or

> u Y

s

5 Z 3/2

where p is the mean gas density, Y is the mean soot mass concentration and 5 O

is the effective boundary layer thickness. Dividing by u we have the amount

of soot per unit wall area

P Y

S

6 ' 2

which is proportional to q"(z) in the lower flame region. In this lower

R region the local rate of combustion per unit area

<V’ U -u-z chem

1/2

so that the local radiant fraction

1/2

X R

(z)

<5{}(z) - z (z) chem

which starts at low values near the base of the flame and thus causes little

initial flame cooling by radiation in this lower flame region, and consequently, the effective flame radiation temperature T

R remains roughly constant, so that the flame radiation should initially be proportional to the amount of

soot per unit flame area. Thus, we have arrived at a plausible heuristic

explanation for why q"(z) - z in this lower region.

R

91
Continuing our attention on Figure 5, one might guess that most of the

original fuel is consumed around the radiation peak occurring at 30$ of L^.. At this point the combustion gases may be predominantly N

2

, CO, H

2

, H

2 0 and

C0 2 as well as soot. As the soot is depleted the radiant heat flux q"(z)

will decrease. Since ethane is a nonsmoky fuel, essentially all the soot will

be consumed because there is not enough soot to induce radiant extinction of

the flame with concomitant release of smoke and CO. Thus, the relatively

steep decline in radiation and the relatively small final decay tail of q"(z)

might have been anticipated. One might further speculate here that more sooty

fuels such as propylene will have considerably slower declines in radiation

and longer final decay tails.

Examining the remaining profiles in Figures 6, 7 and 8, one sees approximately the same initial rise in their fuel-rich regions near the base. Here

the rate of fuel supply is independent of z and the rate of combustion is

1 /2 again initially proportional to z , so that we again anticipate an initial

fuel-rich region. However, at least for the low mass transfer rates of Figure 8 with its three panels, the fuel-rich region should be much shorter and

the flame presumably soon becomes limited by the fuel supply at a lower

height. Once the initial fuel-rich region is eliminated, the soot that is

formed from further fuel supply from the wall is probably speedily oxidized

resulting in a local radiant fraction which is more constant with height.

Such speculations may be useful for: suggesting possible future experiments,

further ways of correlating the data or hopefully stimulating more rigorous

theoretical arguments.

Finally for completeness we show in Figure 10 a cross plot of the thres-

*

] /2 hold flame length measurements against the formula L^=0.17 (Q) which gave

the superior radiant flux distribution correlations. Clearly the threshold

which was used to determine the flame heights in Figure 9 is rather high. The

data seem to correlate on a parallel line which suggests that the possibility

that the present scanning radiometer data may lead to additional parallel

lines for other threshold values.

92
NOMENCLATURE

(4r

+ ^c }

AH

- C (T -T ) ] , [-]. L S p 3 ® — [ AHC

3

XA

'

- C (T - T ) ] , [-]

as *a p 3 *

M

C

AH C Fuel heat of combustion [E/M].

l 3 Smoke-point height, [ L ]

.

m" Mass flux, [M/L^t].

q"(z) Radiant heat flux to wall, [E/L“"t].

R q^(z) Convective heat flux to wall, [E/L^t].

q" Chemical heat release rate per unit wall area [E/L t]. chem

Q' Heat release rate per unit wall length, [E/Lt].

Q' XpQ’. CE/Lt].'

r a

5 Stoichioraetri c air-to-fuel mass ratio [-].

T

s Surface temperature, [8].

Ambient temperature, [9].

u Upward gas velocity, [L/t].

Y Soot mass fraction, [-].

s z Height, CL].

6 Boundary layer thickness, [L].

2 v Kinematic viscosity, [L /t].

p Gas density, [M/L^].

E Mixture fraction, [-].

X A Actual heat release rate fraction, [-].

X c Convective heat release rate fraction, [-].

X R Radiative heat release rate fraction, [-].

T res Microscale flow time, [t].

T chem Characteristic chemical time, [t].

93
REFERENCES

1. Markstein, G.H.: "Relationship Between Smoke-Point and Radiant Emission • from Buoyant Turbulent and Laminar Diffusion Flames," Twentieth Symposium

(International) on Combustion , The Combustion Institute, Pittsburgh, PA

0981).

2. Tewarson, A.: " Prediction of Fire Properties of Materials, Part I: Aliphatic and Aromatic Hydrocarbons and Related Polymers ," FMRC Tech. Rept. J.I.

0K3R3.RC-070( A) , Factory Mutual Research, Norwood, MA (July 1986).

3. Markstein, G.H.: "Radiation Emission from Wall Fires," Fall Technical

Meeting, Eastern Section, The Combustion Institute, Philadelphia, PA,

Nov. 9-6, 1985. Also private communication.

9. Markstein, G.H.: "Measurements on Gaseous-Fuel Pool Fires with a FiberOptic Absorption Probe," Comb. Sci . and Tech. , 39 , 215-33, (1989).

5. Kent, J.H.: "Turbulent Diffusion Flame Sooting- Relationship to SmokePoint Tests," Comb, and Flame , 67 , 223-233 (1987).

6. Miller, J.A., Kee, R.J., Smooke, M.D., and Grcar , J.R.: "The Computation

of the Structure and Extinction Limit of a Methane-Air Stagnation Point

Diffusion Flame," Western States Section, The Combustion Institute, Paper

WS5/CI 89-10, 1989.

7. Bilger, R.W.: "Reaction Rates in Diffusion Flames," Comb, and Flame , 30

277-289, (1977).

8. Tamanini, F.T.: "Direct Measurement of the Longitudinal Variation of

Burning Rate and Product Yield In Turbulent Diffusion Flames," Comb, and

Flame 5J_, 231-293 (1983).

9. Delichatsios , M.A.: "A Simple Algebraic Model for Turbulent Wall Fires,"

Twenty-First Symposium (International) on Ccmbustlon , The Combustion

Institute, Pittsburgh, PA, 1987.

94
0. 5r

0.4

? 1, 3~Sutadi ene

~ <b'lCx\Propy 1 ene

§0.3

° 0.2

"OO

cn

0. 1

0.0

0

Isobutene

^''.pthylene

^^Propc

Ethane

50 100 150 200 250

Smoke _ Point Flame Length, mm

Figure 1 : Correlation of radiative fraction, xp» for buoyant turbulent fueljet flames vs. laminar smoke-point flame height i for gaseous

fuels having similar flame temperatures . Ref. 1.

Figure 2 : Correlation of actual heat release fraction, x^» convective heat

release fraction, x^» and radiative heat release fraction, Xr» from

small turbulent pool fires vs. measured laminar smoke-point flame

height for various noncharring solid and liquid fuels. Ref. 2.

95
Radiative

Fraction

Diffusion Flame Combust/ov Model

Figure 3: Schvab-Zeldovich diffusion-flame combustion model for turbulent

wall fires showing how the presence of the wall reduces the

supplied fuel concentration and temperature in the gas-phase at the

wall. Here £ is the normalized conserved mixture fraction and 3

C is the convective mass transfer driving force.

4. 0 3. 0

B

o 3.5 2. 5 2. 0 1. 5 1. 0 0.5

Figure 4 : Corresponding correlation of the overall flame radiant

from two-dimensional turbulent ethane wall-fires vs. 1/(1 w ic

is the fraction of combustion products present at the fuel supply

surface. Ref. 3.

96
ETHANE WALL FIRES

z/L

f

F i gur e 5 : Correlation of the normalized local flame radiant heat release

q” vs. normalized height for ethane released from a slot at the

base of a 38 cm. wide wall-burner with s^^ walls. The radiant

distribution correlates best for - (Q’) . Ref. 3-

ETHANE WALL FIRES

*/W

Figure 6: Correlation of the normalized local flame radiant heat transfer,

vs, normalized height for ethane released from a single panel

(22 cm, high by 38 cm. wide) in the wall burner with side walls for

three different total B-numbers. Ref. 3 . 97
Figure

Figure 8:

ETHANE WALL FIRES

=1 r L

f/U

r

2 Panels

25.0 kV B Number 0.5

44.3 k* B Number 1.0

62.7 k* B Number 1.5

Beta F itT n = 2.. 2 , m

F(x) = X -a -

x)

9rl

,

L

F = 0.17(00

Q. 0 0. 1 0.2 0. 3 0. 4 0.5 0. 5 0. 7 0. 8 0. 9 1.0

z/L*

7: Correlation of the normalized local flame radiant heat transfer,

q" vs. normalized height for ethane released from two panels

($14 cm. high by 38 cm. wide) in a wall burner with side walls for

three different total B-numbers. Ref. 3-

ETHANE WALL FIRES

zAf Correlation of the normalized local flame radiant heat transfer

vs normalized height for ethane released from

C?6 cm. high by 38 cm. wide) In a wall burner with aide walla for

different total B-numbers. Ref. 3.

98
F i gur e 9 : Correlation of flame heights L

f interpreted from a fixed finite

threshold for versus traditional (Q'j . Note the slight data

curvature. Refl 3.

Figure 10 : Correlation of the flame heights L

f interpreted from a fixed finite

threshold for

qg versus (O')

7

. The line L

f

= 0.17 (Q'J ^ is

from the <5” = 0 values fitting the complete ( 2) profiles shown

in Figures p-^. Ref. 3. R

99
Effect of Radiant Flux on Smoke Emission

G. W. Mulholland, V. Henzel^, V. Babrauskas

National Bureau of Standards

Gaithersburg, Maryland 20899, USA

ABSTRACT

Smoke emission and the light extinction coefficient are measured for propane,

heptane, PMMA, wood, Prudhoe Bay crude oil, and rigid polyurethane in the Cone

Calorimecer for radiant fluxes up to 100 kW/m^ . 3y increasing the specific

burning rate in the Cone Calorimeter to match the burning rate in a large

scale test, a good correlation is obtained between the smoke properties

obtained with the Cone and the same properties obtained in large scale tests.

Time - resolved data are obtained for both smoke concentration and the optical

extinction coefficient. For wood, it was found that soot emission increases

during the later part of the burn while the optical properties remain

constant

.

1. INTRODUCTION

An Important element in assessing the hazard of a fire is the smoke emission.

Smoke from a developing fire will affect the visibility of escape routes. At

the present time there is a limited amount of information regarding emission

and optical properties of smoke as a function of material burning and fire

scale. while there have been some small scale studies on smoke emission

(Bankston e ~ a 1

.

[1], Tewarson [2]) for construe tion materials and materials

found in buildings and houses, there is no general correlation allowing one to

predict the smoke emission for a large scale fire based on small scale

results

.

The. study reported here was motivated by the observation [1,2] that the

burning rate of a fuel can have a pronounced effect on the smoke emission.

The hypothesis to be tested is that, for a given fuel, the smoke emission is

primarily controlled by the burning rate per unit area for the case of

overventilated burning. The dependence of smoke emission upon the burning

rate, however, may vary from fuel to fuel. In our study the smoke emission is

measured as a function of burning rate for a number of fuels and the results

compared with large scale free burn results. The burning rate is varied by

changing the radiant flux to the sample.

2. SMOKE MEASUREMENTS

The most rudimentary smoke quantity is the smoke yield, e , which is defined as the mass of smoke aerosol generated per mass of fuel consumed. This quantity

is determined by a mass flux method and by a carbon balance method. The key

measurements in the flux method are the measurement of the mass of smoke

•Nuclear Research Center Negev, Beer-Sheva 84 190, Israel.

100
collected on a filter, the mass loss of the fuel, the air flow rate to the

filter, and the total flow rate in the stack. The carbon balance method is

b'ased on the measurement of the carbon-containing combustion products

including C0. 2

, CO, and smoke aerosol. A standard 47 mm filter holder with

glass fiber filters and a tapered element oscillating microbalance (TEOM) were used for filter collection. The TEOM contains a removable filter element at

the end of an oscillating hollow rod, and the mass deposited on the filter is

determined from the change in the frequency of the rod. A nominal flow rate

of 2 1/min is maintained through the filter. The TEOM is virtually the only

commercial instrument capable of providing time-resolved information on smoke

emission for high concentrations of sooty smoke.

One of the principal optical properties of interest is the specific extinction

area, on fuel -pyrolysates basis, Of. For a flow system such as the Cone

Calorimeter, Of is obtained from the measurement of the extinction

coefficient, k, for a He-Ne laser beam (633 run) transmitted through the

flowing smoke

.

of = k/ (nif/V) , (1)

where m.e is the fuel mass loss rate and V is the volumetric flow rate through

the duct. As a heuristic example, a value of 1 m^/g for oc means chat if the

smoke produced by one gram of fuel were collected over aim- area, the light

incident on this area would be blocked (intensity reduced by a factor e).

Another optical propert3/ of incerest is the specific extinction area, on soot- mass basis, o s

, defined by

o 3 - k/ms , "

( 2 )

where m s is the mass concentration of smoke. The quantity o 3 is an intrinsic

property of the smoke depending on the wavelength of light, the opcical

properties of the smoke, and on the site distribution of the smoke. The

quantities Of and a s are related through the equation

<7f * € O s

. (3)

3. EXPERIMENTAL PROCEDURE

The smoke emission measurements were performed in the Cone Calorimeter

developed by 3abrauskas [3], The conical radiant source shown in Fig. 1 provides a uniform radiant flux over the 100 cm4

- sample for radiant fluxes up

to 100 kw/m- . After a warmup period for the radiant source to reach steady

state, the sample is inserted and ignited via a spark. The smoke aerosol and

combustion gases rise through the conical heater into the exhaust hood and are sampled from a horizontal section of pipe as indicated in Fig. *1. The fuel

burning rate is monitored with a load cell and, from monitoring the O 9 consumption in the exhaust gases, the heat release rate of the fuel is

determined [4]. Both the incident and transmitted laser beam are monitored

with photodiodes to achieve the high level of precision in the determination

of the percent transmission necessitated by the 10 cm path length. The gases

sampled include CO, CO 9 ,- O 2

, and HoO.

101
A square cross section diffusion burner (2.35 cm edge length) was used for

burning propane. Both the heptane and Prudhoe Bay crude oil were burned in a 8.5 cm diameter cylindrical vessel made of glass. The wood, PMMA, and rigid

polyurethane were in the form of square samples measuring 10 cm on an edge

with a thickness of about 2.5 cm.

Large scale fires up to 350 kW were also studied with a large scale version of

the Cone Calorimeter but without an external radiant flux. The apparatus has

a maximum flow of about 2 m^/s through a 2.4- m by 2.4 m collection hood

located 2.4 m above the floor. As indicated in Fig. 2, the smoke from the

fire is mixed by a tripper plate and then sampled 5 duct diameters downstream

by an isokinetic probe. The filter collection system allows sequential

collection of three filter samples over the course of a burn and is designed

to minimize particle loss by heating the sampling lines ana filters to the

stack temperature. As in the Cone, gas analysis and light extinction

measurements were performed.

4. RESULTS

The effect of the external radiant flux on the specific light extinction

coefficient, <7f, is shown in Fig. 3 for propane, PMMA, and wood. Propane is

unique, because the burning rate is not effected by the radiant flux. It is

set by the fuel flow rate, which has a value of about 3 1/min for all four

tests. The value of ac for propane combustion increases bv about a factor of

three as the radiant flux is increased from 30 to 100 kwT /m^ . With increasing

radiant flux, the smoke emission doubles as indicated in Table 1. The value

obtained for £ by the flux and carbon balance method agree to about 10 - 20 h

for all materials. Only the values obtained by the flux method are reported

in Table 1. Most entries in Table 1 correspond to an average of more than one test. Propane is the simplest case, since the external radiant flux affects

the flame processes but not the generation rate of the fuel vapor. For all

the other fuels, the external radiant flux affects both the generation rate of

fuel vapor and the flame processes.

The mass burning rate of PMMA is approximately tripled by increasing radiant

flux from 25 to 100 kw/m*-; however, in this case both the smoke yield and a

f

are relatively insensitive to the radiant flux. The results are also

insensitive to the sample orientation.

For the case of red oak, the smoke measurements are difficult ac low fluxes

because of the low smoke production. The uncertainty in ac is approximately

±0.005, which is comparable in magnitude to the observed value at the lowest

radiant flux (see Fig. 3). Both the smoke yield and the specific extinction

coefficient increase by almost a decade with increasing radiant flux. The

major peaks in the specific extinction coefficient mirror the same peaks seen

in the burning rate. The time dependence of £ shown in Fig. 4 also has a similar shape, though in this case there is a high level of noise. The

enhanced burning at the later stage apparently results from reduced conductive

heat loss when the thermal "wave" propagates through the entire sample. As

has been reported by 3ankston e t a

1

[1], the smoke generation is greater if

the sample is burning in a horizontal rather than vertical orientation.

102
Heptane was Che only fuel that exhibited a decrease in smoke emission wich an increase in the radiant flux. The mass burning rate of heptane increases more rapidly with increasing radiant flux than the other fuels because of the

higher vapor pressure of heptane. It was noted that for Che heptane fires

there was more pronounced boiling at the higher fluxes and also that the heat

release rate was several times larger than for the other fuels. It is not

known whether the decrease in smoke emission is a result of more efficient

combustion in the flame at the higher radiant fluxes or a result of the high

temperature in the exhaust system. For the other liquid fuel, Prudhce Bay

crude oil, the smoke emission was found to be relatively insensitive to the

burning race

.

As discussed above, the specific extinction coefficient, a s

, is an intensive

property of the smoke. The average value for each fuel is in the range 3 to

12 for the Cone data. The range in the average value of a s for the same fuels

in the large scale experiments, 8-9, i£ much less. The collection conditions

are more variable in the Cone experiments, with stack temperatures ranging

from 100°C to 400°C compared to 50°C to 175°C for the large scale experiments.

Also, the sampling system is isokinetic with the gas and wall temperature

matched for the large scale apparatus ; whereas , for the Cone the temperature

of the smoke decreases by as much as several hundred degrees before reaching

the filter element. The large difference between the gas temperature and the

temperature of sampling duct is expected to lead to significant particle

deposition resulting from thermophoresis.

The quantity a s is relatively constant as a^ function of time for burning wood

as indicated in Fig. 5, even though both e and jf increase significantly in

the later stage of the burn. This result suggests that the optical properties

of the soot remain relatively constant during the burn even though the soot

emission is increasing.

The results for the large scale burns are compared with the Cone results for

comparable burning rates in Table 2. ' The comparison for propane is somewhat

artificial, because the burning race is controlled by the gas flow race and

not by the external radiant flux. The lowest fuel flow race for the large

scale burner corresponds to the flame just barely covering the entire burner

surface. It is an interesting coincidence that there is an abrupt increase

in smoke emission with increasing fuel flow rate in the large scale

experiments and a similar size increase in smoke emission with increasing

radiant flux in the Cone.

The most reliably measured quantity for comparing large and small scale

results is <jf, since smoke deposition in the filter sampling line does not

effect this measurement. The large and small scale results in units of m“/g

respectively are: heptane 0.086 vs 0.078, Prudhoe Bay crude oil 0.96 vs

1.0,1, rigid polyurethane 0.79 vs 0.89, and wood 0.034 vs 0.028. The average

percent difference is about 12%, which is similar to the expected experimental

uncertainty. The largest difference is for wood. In this case, the large

scale measurements were made with sugar pine and the small scale with red oak.

In the other cases, identical materials were used at both scales. It is also

103
found that there is good agreement in the smoke emission values for the large

and small scale experiments for the same burning races.

One limitation of this method for intercomparing small and large scale tests

is the determination of the burning rate per area for complex structures such

as cribs. We roughly estimated the effective burning area of the crib to be

half of the total surface area of all the individual pieces. The factor of

two reduction cakes into account the limited burning of undersurfaces, the

physical overlap of the individual pieces, and radiation shielding of lower

pieces by upper pieces. Also, for the wood , we have taken an average of the

vertical and horizontal results for the cone measurements.

5. DISCUSSION

Experimental results support a good correlation between small and large scale

smoke emission results for the same specific burning rate. A demonstration of

the general validity/utility of this approach must await measurements on a wider range of materials including composites and a more reliable way to

estimate the burning area for complex structures.

Qualitatively, there are three ways in which large scale fires differ from

small fires: the burning rate, the flame radiation, and the fluid flow.

Applying an external radiant flux to a small sample enables one to match the

burning rates for small and large scale and to mimic to some extent the larger

fraction of radiant energy in the large fire, but not to reproduce the

vigorous turbulent mixing of a large scale fire. Data are needed at fire

scales up to at least 5 MW to determine the ,s ignif icance of the fluid flow in

the plume and radiant feedback in regard to the small scale correlation.

Data from studies by Tewarson [2] and by Bankston e t al

.

[1] are included in

Table 2. It is seen that there is good agreement in the values of e for ?MMA,

but that Tewarson' s value of e for heptane is a factor of two greater than our largest value. Also, the combustion efficiency reported by Tewarson, 0.93, is

less than observed in this study.

Bankston et al . f

1

1 report more than a two-fold decrease in smoke emission for r * o o wood as the radiant flux is increased from 25 kW/m** to 50 kW/mh , while we observe about a factor of two increase. The difference may result from our collecting smoke only after flaming combustion begins, while 3ankston e t al

.

[1] may have also collected "pyrolysis smoke" produced before ignition

occurred. It is known that the smoke yield from wood during pyrolysis is much

greater than during flaming combustion.

Bard and Pagni [5] have proposed a method for estimating the maximum

conversion of fuel carbon to soot based on the measurement of the maximum

volume fraction of soot within the flame. As a comparison, the maximum

yields obtained with the cone are 0.016 for PMMA and 0.013 for wood compared

to Bard and Pagnis' prediction of 0.024 for PMMA and about 0.018 for wood [5].

The values of a~ obtained with the Cone ranged from 8-12 m~/g and from

8-9 m

z /g for the large scale. Seader and Einhorn [6] obtained a mean value or 7.6 m“/g based on measurements on several plastics and wood in a chamber using

104
a polychromatic light source. Neuman and Steciak [7] obtain values of 10.3,

10.3, and 10.5 m~/g for a s based on flaming combustion of heptane, douglas

fir, and PMMA, respectively. The measurements were made at A=633 nm, which is

essentially identical to the wavelength used in our study. These results are consistent with the statement that a s for soot is independent of the fuel,

though systematic differences among laboratories remain to be resolved.

5. ACKNOWLEDGEMENTS

The authors gratefully acknowledge Nelson Bryner and William Twilley for

assisting in obtaining and reducing the smoke data.

6 . REFERENCES

[1] Bankston, C. P., Zinn, B. T., Browner, R. F. , and Powell, E. A., Aspects

of the Mechanisms of Smoke Generation by 3urning Materials, Combustion

and Flame . 41, 273-292 (1981).

[2] Tewarson, A., Prediction of Fire Properties of Fuels, Twe n tv -First S vmo

.

( Inti . ) on Combustion . The Combustion Institute, Pittsburg, 563-

570(1986).

[3] 3abrauskas , V. , Development of the Cone Calorimeter - A Bench-scale Heat

Release Apparatus Based on Oxygen Consumption (NBSIR 82-2611), [U.S.j

Nat. Bur. Stand. (1982).

[4] Huggett, C., Estimation of Race of Heat Release by Means of Oxygen

Consumption Measurements, Fire and Materials . 61-65 (1930).

[5] 3ard, S., and Pagni, ?. J., Spatial Variation of Soot Volume Fractions

in Pool Fire Diffusion Flames, First Svmr) . ('Inti.) on Fire Safety

Sc ience . C. E. Grant and P. J. Pagni, ed. New York, Hemisphere Pub. ,361-

369 (1986).

[6] Seader, J. D .

, and Einhorn, I. N., Some Physical, Chemical,

- Toxicological, and Physiological Aspects of Fire Smokes, Sixteenth Syrr?

Clntl.) on Combustion . The Combustion Institute, Pittsburg, 1423 -1444

(1977).

[7] Newman. J. S., and Steciak, J., Characterization of Particulates from

Diffusion Flames, Combustion and Flame . 67 . 55 - 64 (1987).

105
Table 1. Smoke Emission measured with the Cone Calorimeter

Fuel Radiant Burning Comb . Smoke Of ,m 2 /g

Flux, kW/m^ Rate , g/sm^ Eff

.

Yield

Propane 0 176 0.98 0.0057 0.019

30 176 - 0.0064 0.038

50 176 0.98 0.0087 0.057

100 176 - 0.0137 0.127

n-Heptane 0 10.0 0.99 0.0125 0.089

10 24.0 0.94 0.0131 0.083

20 31.9 0.97 0.0096 0.072

30 58.4 0.99 0.0062 0.050

a 0.93 0.027

Prudhoe 3 ay 0 4.5 40.

9

b 0.098 1.06

crude oil 25 10.7 38 .

1

0.096 1.01

40 18.1 36.9 0.083 1.01

50 23.9 35.5 0.084 0.98

PJiMA 25 16.4 0.96 0.0152 0.159

50 25.0 0 . 96 0.0137 0.169

75 37.3 0.95 0.0121 0.167

100 47.0 0 . 96 0.0160 0.164

a 0.95 • 0.016

Red Oak 25 , V 8.9 11.

8

b 0.0010 0.011

25 ,H 8.3 10.6 0.0020 0.022

50-7 11.7 11.3 0.0018 0.025

50 - H 11.8 10.9 0.0055 0.050

- 75 ,V 14.7 11.4 0.0033 0.035

- 75, H 14.6 10.8 0.0080 0.094

100, V 18.9 11.3 0.0039 0.060

100 -H 13 .

7

11.5 0.0129 0.120

Douglas fir c 25 0.025

50 <0.01

Rigid Poly- 50 5.3 0.85 0.080 0.89

urethane foam

V and H refer to the vertical and horizontal orientation of the

respec tively

.

aResults by Tewarson.

b Average heat of combustion, MJ/kg.

c Results by Bankston e t a 1

.

for sample 7.5 by 7.5 cm.

° S ’ m2 /S

6 .

4

7.3

9.5

7.7

6 .

6

7.9

8 . 3

11.7

10.3

12.5

11.7

10.9

12.5

11.3

10.3

11.2

13.4

9.0

11.0

14.6

9.3

9.8

9.4

s amp 1 e ,

106
Table 2. Comparison of Small and Large Scale Smoke Emission Results

Fuel/Conditions Q.kW Burning

Rate , g/sm2 Comb . Eff

.

Smoke

Yield

<7f ,m

2 /g .m2 /g

Prooane . 1 m 110 2.8 1 . 03 a 0.0077 0.049 - diff. burner 175 4.2 1. ll a 0.0168 0.154 8.1

2.35 cm burner

350 8.6 1 . 09 a 0.0199 0.161 7.8

0 kW/m2 4.2 176 0.98 0.0057 0.019 9.9

50 kW/m2 4.2 176 0.98 0.0087 0.051 7.3

100 kW/m2 4.2 176 - 0.0137 0.127 •9 .

5

n-Hentane

,

31 cm pool 70 24.9 • 0.89 0.0093 0.074 7.3

50 cm pool 240 28.3 0.94 0.0121 0.098 3 .

2

8.5 cm pool

10 kW/m“ 7.1 24.0 0.94 0.0131 0.083 6 .

6

20 kW/m2 10.0 31.9 0.97 0.0096 0.072 7.9

Ptudhoe 3av crude oil

40 cm pool 65 14.3 34.

5

b 0.090 0.96 9 .

5

60 cm pool

3.5 cm pool

185 (13.1) 0.085 3.7

25 kw/m2 2.3 10.7 33. l b 0.096 1.01 10.8

40 kW/m2

Wood

Sugar pine

3.3 13.1 36.9 0.083 1.00 12.5

1 crib 56 3 .

7

C 12.9° 0.0036 0.029 8 . 5

3 cribs 254 12.6 13.4 0.0042 0.040 9.4

Red oak, 10 cm

25 kW/m2 1.0 3.6 11.

2

b 0.0015 0.017 11.2

50 kW/'m2 1.3 11.3 11.1 0.0037 0.038 11.2

Polvure thane

.

rigid 1 crib 125 12.

0

C 0.68 0.085 0 . 74 9 .

1

2 cribs 310 14.6 0.68 0.101 0 .81 8 . 5

10 cm 50 kW/m2 2.5 5.3 0.85 0.080 0 . 39 9.4

a The erroneously large

the propane flow calibra

b The heat of combustion

combustion

tion.

in MJ/kg.

efficiency results frotn an uncertainty in

c The effective

area of all the

surface area for combustion is

individual sticks.

taken as half the total surfa<

107
<noe

cncccHocoe

Sri o

4-

Q

LLi o

_i

O

-J o • jj cn •H 0) a iri 03 <d

4-1 c.o

cd )ri f-ri Oas

'~-~-

a c

cn o

•rri 0) C/5 co cn

Sri -H

cc £ J 0)

Ol

CO

•ri^

10 2

_!

O

a: z

LU UJ > sO

a

cj

4-1

cn

CD •U

OJ

-ri 03

CJcn

03g

cn

4-1O3

as •H>

a.

CD

CJca

108 CO

til
PROPANE

RED OAK, HORIZONTAL

Fig. 3. Specific extinction area, Kf , versus time for radiant fluxes in

range 25-100 kW/m109
SMOKE

CONVERSION

FACTOR,

RED OAK, HORIZONTAL

TIME (s)

Fig. 4. Smoke conversion factor £ versus time

RED OAK

Fig. 5. Specific extinction coefficient Ks

versus time for radiant flux of 100 kW/m“

110
THE DISTRIBUTION OF SOOT IN LARGE BUOYANT DIFFUSION FLAMES

E. E. Zukoski

California Institute of Technology

Pasadena, California 91125

A study is being made of the internal structure of large buoyant fire plumes

with a new experimental technique which was developed by R. Miake-Lye and S.

J. Toner (1986) following a suggestion by P. Dimotakis. The technique

involves the measurement of the light scattered from a narrowly focused

argon-ion laser beam by soot which is generated by diffusion flames within

the fire plume itself. The beam is directed across the flame and study of

the scattered light allows us to measure the distribution of soot within the

fire plume. In contrast, photographs taken in light emitted by thermal

processes from soot particles allow only the exterior surface of the fire

plume to be studied.

The fire plumes used in these studies are methane-air diffusion flames

stabilized on a 50 cm diameter burner which utilizes an uncooled bed of

glass beads to produce a uniform flow of methane at the burner surface.

Flames, with total heat release rates in the range 20 to 200 kW, were chosen

for study because other characteristics of these flames have been studied

asnd reported elsewhere, e.g., Cetegen et al (1984). These flames are in

the transition regime between flames with height to diameter ratios which

are less than one and greater than three and thus are in' regime II defined

in Zukoski (1986).

In the original experiments, described in Miake-Lye and Toner (1987), the

laser beam was passed through the flame in a horizontal direction which

intersected a vertical line extending above the center of the burner. Data

was obtained for heights which ranged from the surface of the burner to

positions several burner diameters above the surface.

Light scattered from the laser beam was filtered to remove the background

radiation and then was focused on a 1024 pixel, linear solid-state, RETICON

array camera. The array was interrogated at a frequency of 586 hz and the

light intensity of each pixel was recorded as a function of the time. The

presence of scattered radiation with an amplitude above twice the background

radiation level was interpreted to mean that soot was present at that

position and time.

In preliminary experiments, we found that at positions where laser light was scattered, visible radiation from soot was present. Because of the rapid

cooling of soot particles in the absence of a diffusion flame, visible

radiation was assumed to be an indicator of the presence of a diffusion

flame. Eence, we expect and assume, in interpreting our data, that

scattered radiation -is also an indicator of the presence of a diffusion

flame.

Thus, if scattered laser light is present at a given point and at a given

time, we can reasonably assume that a diffusion flame, complete with fuel on

111
E. E. Zukoski UJNR-87

one side and oxygen on the other, was close to the point under observation

at the time of the measurement. In our picture, the fuel and oxygen may be

vitiated by combustion products, but both were assumed to be present.

The use of a sheet of laser light would have given us a much more complete

picture of the flow, but we were restricted to the beam because we did not

have a laser with the power required to produce a sheet with the requisite

intensity and by limitations on the data acquisition rate of the camera system used to record the scattered light intensity.

In ongoing experiments, the direction of the beam has been changed to

produce a vertical beam which passes through positions which correspond to

the centerline of the burner and 5 radial positions extending out to the

edge of the 50 cm diameter burner.

The most interesting result of this study has been the discovery that soot

is present in very complex and dense patterns within the flame and that it

is present between 40 to 60% of the time at all points within the exterior

envelope of the fire plume. The exterior envelope here is taken to be the

outer surface of the region illuminated by radiation from soot or chemiluminescence. This result has been observed to hold for elevations near

the surface of the burner and up to two burner diameters above the burner

surface. In our interpretation of the data, these results suggest that

diffusion flamelets, separating regions rich in fuel or oxygen, are densely

distributed within the exterior envelope of the fire plume. This result can only be true if transport of oxygen from the entrained air throughout the

fire plume is very rapid.

We believe that the large structures, shed from the burner in these flames

at a frequency of the order of several Hertz, are responsible for this rapid

transport and that secondary instabilities which develop on these structures

are responsible for the fine details of the diffusion flamelets observed

within the flames.

The data being obtained with the vertical orientation of the laser beam will

allow us to determine the geometry and celerity of the structures present in

the lower part of large diffusion flames and to continue our study of the

interior of the fire plume.

This work has been supported by the Gas Research Institute and the Fire

Research Center of the National Bureau of Standards.

REFERENCES

Miake-Lye, R. C. and Toner, S. J. (1987), "Laser Scot-Scattering Imaging of

a Large Buoyant Diffusion Flame", Combust ion and F lame . 67, 9-26 .

Cetegen, B.M., Zukoski, E. E. and Kubota, T. (1984), "Entrainmnet in the

Near and Far Field of Fire Plumes", Combust ion Sc ienc e and T echno logy , 39,

305-331.

Zukoski, E. E. ( 1985 ), "Fluid Dynamic Aspects of Room Fires", Proc . of F irst

Int e-mat iona I S vmp o s iun on Fare Safety Sc ienc e , Hemisphere Publishing Corp,

1986.

112
PREDICTING CAPABILITY OF A MULTIROOM FIRE MODEL

Kazuhito NAKAMURA

Design Division

Shimizu Construction Corporation

2-16-1 Kyobashi, Chuoh-ku, Tokyo 104

INTRODUCTION

For a performance based fire safety design method of

buildings, user friendly computer models to predict smoke

transport behavior in buildings is considered to be one of its

indispensable elements.

As such prediction models for smoke movement, chiefly two

types of zone models have been developed by now, that is, one

based on one layer concept and the other based on two layer

concept. The former will be convenient to predict smoke movement throughout large scale buildings, high-rise buildings and

so on, and the latter is thought to be beneficial when we need

to know some more detailed behavior of smoke around the room of

fire origin.

However, many of these models, which have been developed

primarily for research purpose, will more or less need to be

refined for the convenience of design practice as well as for

the improvement in accuracy to be used for design purposes. A

multiroom smoke movement model (Ref.l), which is based on two

layer concept, has been refined from this point of view. The

points of refinement include, a) the prediction of lower layer

properties, by which stack effect due to the temperature difference between inside and outside of building can be treated,

b) capability of dealing with mechanical smoke control, c) time

scheduled opening and closing of openings, by which the effects

of door opening due to evacuation or window breaking due to

fire on smoke movement can be assessed, d) capability of dealing with spaces of complex geometry such as a dome, and e) improvement of stability and speed up in numerical calculation.

Since most of the theoretical aspects of the model are

given in Ref. 2, this paper focus on the model of the opening

jet plume near opening region and some of the features for the

convenience of design practices.

OPENING JET PLUME NEAR OPENING

In this model, it is assumed that the plumes originated by

fire source and opening jets can penetrate through interfaces

of the upper and lower layers, due to which exchanges of mass

and heat take place. As a result the method to determine the

fraction of mass and heat penetration becomes one of the key

points.' In the region far away from the source of opening

plume, i.e., opening jet, the fraction can be calculated by

modeling opening jet plume as vertically rising plumes and assuming Gaussian profiles on the lateral profile of temperature

113
and vertical velocity in a plume and criterion temperature of

penetration (See Appendix 2.).

In the region near an opening, however, simple application

of Gaussian profile can cause a problem : i.e. the axis temperature of modeled plume zlTa can become higher than the temperature of the opening jet . This is illustrated by the

example shown in Fig. 2. In this example, an opening jet at

300® C enters a lower layer at 20® C and rises through the layer

entraining the air and penetrates into an upper layer at 300® C.

In this particular case, the excess temperature on the axis of

the modeled plume exceeds the temperature difference between

the opening jet and the ambient layer ZT > (=300-20 = 280 K) in

the region of Z < 0.72 m. So, if the criterion temperature for

plume penetration is assumed as

AT = = ( T a - T =o ) / 2 = 140 [ K]

the net heat gain of the upper layer becomes positive in the

region of Z < 0.25 m as is shown in Fig. 2, so that the upper

layer is heated while it has to be cooled in reality. Needless

to say, this is attributed to the sharp peak of Gaussian

profile of the temperature near axis, which is unrealistic in

the sense that the plume axis temperature jT a exceeds the opening jet temperature ATc.

It will be more reasonable to consider that an opening jet

initially has flat profiles of temperature and velocity, but

the profiles collapse as the distance Z increases and eventually turns into Gaussian profiles. Here an interim model is

proposed as follows to deal with the penetration in the region

near opening.

Let Za be the height where the excess temperature of the

axis of a plume with a Gaussian profile /ITj becomes equal to

the temperature difference between the opening jet and the ambient layer ZITj . The near region to an opening is defined as

the region where Z<Zd . It is assumed that the temperature

profile of the off central part of a plume follows Gaussian

profile with axis temperature dTo’

,

and inner part takes flat

temperature profile. Similar profile is assumed for the

velocity profile being WV as the axis velocity of plume. In

other words

,

iTd _ -& z vSJTV~ e (1)

Wd _ -yo 2 Wo’ K ( 2 ) where yd is the radius in Howarth transformation which corresponds to ZlTd on the Gaussian temperature profile.

Now, let mzt be the mass flow rate through near axis area

( 0 S y S y <i ), mx be the rate through off axis area ( yd^y ) and mz be the total mass flow rate, then

mc=m2i + m22 -(3) mn =2 7t / q7 J Wo

’ e 7 1 p =o b

2 y d y

= K Wo’ P CZ3 b ~ s -y •<- y .i 2 <4)

114
mz2 = 27r Wo’ e

-77&

p » b -ydy

• 7 d = 7c Wo ’ p ao b

2 e

771 (5)

Similarly, for the rate of heat flow through near and off axis

regions, we have

Q = Q i + Q 2 (6)

Qt = 27TCoJ‘^7d z]Ta’e ^ 77 1 wo’e 77 d p oo b

2 7 d 7

= k c 3 wo ’ p aa b

2 (JTb’ e ~^+ ® ^ 77 j2

7 d 2 ( 7 ) Q 2 = 27TCoJ!f dTo’e ^ 77 wo’e 77 pasb^d? 77 d

7T C 3 Wo ' P ® b ,m -(l+/? 2 )7d 2 - TT^ 4Ta e (a)

From Eqs.(l)-{8) and (A3), 7 <* can be obtained as follows:

1 1+ £* v Za

77 d 2 = 1-t-lAt ( -Z

7v

Z?

) 5 ' 3 -( 9 )

Using yd 2

, the rate of mass penetrating through a layer interface is given as follows:

iHomt - m 2 — 2 7c J" PIT. Wa

— -J7 2- e pccb 2

T7 dy

— m 2 — 7T W a p co b

2 e

-77

zlTc

1

( 77 d 2 — 77 2

) = mz{l--T

^~ e

4Tc

} (10)

where is the radius in Howarth transformation which corresponds to threshold temperature of penetration A T c . And the

penetration rate of heat is given as

q 0 = q -27r c d ff dTa’e ^ 77 w a

’ e

v pcab 2 v<iv

77 Ale

-(1 +/? 2 )y zlT s = Q-^ c jWo’pob 2

<d T a

’ ^ ^ gs

= Q {1 ~1 + (1+/? 2

) = Q k

(1+ /?

2

) ( 7 o 2 - v 2<f s ) 7

where

k = 1- 1 1 + (1+ /? 2

) 7 <1‘

(1 +/? 2)(7,2_ 7

2 t ^

•( 11 )

( 12 )

So, it follow that the penetration model near opening has been

obtained

.

It should be kept in mind, however, that this theory is not

perfect but of a compromised nature since the total mass flow

rate have to be calculated using Eqs .

( A 1 ) , ( A2 ) , i.e., the equations for the flow rate for far field region of plume with virtual point heat source as a adjusting parameter.

MODEL FEATURES FOR PRACTICAL APPLICATION

115
1.

Treatment of 'Mechanical Smoke Control System

For the convenience of evaluating efficiencies of smoke

control by means of air handling system, this computer model

has incorporated a feature to make it possible to deal with

smoke control system to some extent.

The method employed in this model is rather simple : The

air supply rate or smoke exhaust rate with time is given by the

user in terms of volume flow rate as input data and converted

into mass flow rate by the code. The mass rate of air given to

each layer as the result of air supply is determined at each

time step considering the height of the air inlet relative to

the height of the layer interface. Similar consideration is

takes to determine the mass rate of gases removed from each

layer by mechanical exhaustion.

2.

Opening and Closing of Openings

In the course of evacuation, condition of opening in a building may change. For example, doors in a building may be

opened or closed by evacuees or others, natural smoke vents may

be opened to let smoke go out, and fire doors may be closedautomatically or manually for confinement of fire. And these

changes of opening conditions, of course, greatly affect the

movement of smoke in the building. So a feature is incorporated. so that the caluculaion can be made under changing conditions of openings when the time schedules of openings are

specified by users as input data.

3.

Treatment of spaces of complex shape

Although

s imple shapes

,

shape cannot be

spaces or large

complex spaces

tangular space

most of building spaces have rectangular or other

the number of spaces which have somewhat complex

neglected in particular in case of high ceiling

scale spaces. This model can deal with such

by modeling a complex space as a stack of recuni ts

.

SAMPLE CALCULATIONS

In the followings, the results of three sample calculations

are shown to demonstrate the computation capability of the

present model.

1. Smoke Filling in Large Scale Space

A series of field experiments for smoke behavior were conducted using the pavilions of the international exposition held

in Tsukuba, 1985, taking advantage of the time lag between the

closing of the fair and the pulling down of the pavilions.

Of the pavilions in which fire experiments were run, the

pavilion that had been used for USSR’s exhibitions was a large

scale space with a saddle back roof as is shown in Fig. 5. Its

floor area was 2,000 m2 and the height of the roof ridge was

116
19.2 ra. As the smoke control measure, a mechanical smoke exhaustion system consisting of five smoke vents, a duct and a fan with the capacity of 1,000 m3/h was installed according to

the Building Standard Law of Japan.

Methanol in the six trays arranged together on the center

of the floor was used as the fire source. Its heat release

rate is estimated as about 2,400 kW.

The predicted and measured results of the temperature and

the interface height of the upper layer are shown in Fig. 6 for

the cases with and without mechanical smoke exhaustion. The

layer interface heights of the experiments were determined from

the thermocouple readings using the same method as is used by

Cooper et al (Ref. 3), and the space is modeled as the stack of

seven rectangular space unit for the calculation.

The predicted smoke layer height for no mechanical ventilation case agrees comparatively well with the measured results

although a trivial difference is observed at the initial stage.

On the other hand, while the predicted layer height for

mechanical smoke ventilation case is higher than for no

mechanical ventilation by one meter, no meaningful difference

can be observed between the experimental layer heights for the

two cases. The reasons why the effect of mechanical smoke exhaustion did not appear on the smoke layer heights may be attributed partly to the possible infiltration of outdoor air to

the smoke layer through invisibly small openings on the wall of

the pavilion and partly to the possible large mixing of the

smoke with air as it travels down to the edge of the space

after having impinged on the roof.

The predicted layer temperatures are consistently higher

than the measured. .The position of thermocouple trees may have

some relation on this difference, but the convective heat

transfer model may be responsible too. While the predicted

temperature difference is small between the cases with and

without mechanical exhaustion, the measured temperature difference is significant between the two cases. It is suspected

that because an exhaustion vent was located right above the

fire source, the gas in the plume, which is hotter than the

layer average, were exhausted effectively.

2. Mechanical Smoke Exhaustion at Office Space

Of the Tsukuba Expo. ’S3 pavilions used for the series of

the experiments, IBM pavilion was provided with a mechanical

smoke exhaustion system at its office area in compliance with

The Building Standard Law. The office area used for the experiments consisted of an office with 51 m. and a corridor of

1.75 m wide and 30 m long as is shown in Fig. 7. • The ceiling

height of this area was 2.7 m. One of the mechanical smoke

vents were located at a corner of the office and the other at

the end of the corridor and could exhaust smoke

0.9 m3/s and 1.3 m^/s, respectively.

The fire source was placed at the center of

office. The fuel was methanol and the size of

0.9 m x 0.9 m. The heat release rate is estimated as 320 kW.

The predicted and the experimentally observed behavior of

the smoke layers are shown in Fig. 9 for the case with mechaniat the rate of

floor of

the source

the

was

117
cal exhaustion, at the office and the corridor, and the case

without mechanical exhaustion. The figures schematically illustrates the section of area and the left hand, side and the

right hand side represents the office and the corridor,

respectively. The experimental temperature rise in the smoke

layers in Fig. 9 were obtained by thermocouple readings and are

shown by the shades of different 'thickness according to the

temperature

.

In the prediction, the corridor was divided into two parts

because it was considerably long and had a turn at the middle.

In the case of no mechanical ventilation, the predicted

layer temperatures are somewhat higher than those measured, but

generally speaking, the agreement of the predicted and experimentally obtained layer behavior can be said fair. In the

case when the mechanical ventilation operated, both the prediction and the experiment demonstrate that the smoke can be

effectively exhausted from the spaces by the smoke control

system

.

3. Prediction of Smoke Movement in a Model Building

Prediction was attempted for the smoke movement for a eight

story model building with an atrium. The building consists of

an atrium of which size is 20 m x 20 m x 32 mH and two rooms of

10 m x 20 m x 3 mH on .each floor. And on 3rd through 8th

floor, two fire shutters, 5 m x 3 mH each, are installed on

each opening facing to the atrium. On the top of the wall of

the atrium is a natural smoke vent of 12 m wide and 1 m high,

and on each floor an air supply fan with capacity 4 m^/s is

equipped

.

The calculation was run under the scenario described as

f ollows

:

time event

before 0 sec . the outdoor and indoor temperature were 30® C and

20® C respectively

0 sec . fire breaks out on the floor of the atrium

(heat release rate increase linearly from zero

to reach 5 MW at 300 sec. and then levels

off)

60 sec . detects fire, evacuation starts throughout the

building

a window is opened at the office on 7th floor

120 sec . the fire shutters on each floor close but on

5th and 6th floor one of the two failed to

close due to some trouble.

125 sec . mechanical air supply fan switches on 6th floor

270 sec . the natural vent opens at the top of the atrium

The results of the prediction are shown in Fig. 10. This

118
calculation for predicting the 600 seconds of fire took about

100 seconds of CPU time by BRI computer ( FACOM M-380Q )

.

CONCLUDING REMARKS

We acknowledge that the models on the entrainment of doorjet or the penetration of plumes are not based on very solid

understanding of the phenomena, so we may have to be careful

when we apply the model extensively to various problems.

However, perfect models are not necessarily available in many

other areas of engineering either, in particular when complex

problems have to be dealt with.

It is obvious that engineering tools to predict fire behavior can help rationalize fire safety measures of buildings

even if they are not perfect. In this sense, many of the existing models can be used in practical issues if sufficient attentions are paid on their limitations.

As for the present model, it is evaluated that with the introduction of partial penetration model of opening jets, the

model is refined so that predict smoke movement in building can

be predicted reasonably well without giving contradictory

results, although it does not seem so good in predicting convective heat transfer. Also, the model is equipped with

several features which, are thought to be beneficial for design

purposes

.

'The model still needs improvements on many respects, such

as convective heat transfer, lateral smoke propagation along

corridors, smoke behavior in narrow vertical shafts and so on, so we would like to continue its refinements absorbing better

understandings on these phenomena as they grow.

Re f erences

1. Tanaka, T., "A Model of Multiroom Fire Soread", NBSIR 83-

2718(1983), NBS

2. Nakamura, K. and Tanaka, T., "Refinement of A Multiroom Fire

Spread model", The 2nd ASME-JSME Meeting on Thermal Engineering

( 1987) , Hawai

i

3. Cetegen, B.M., Zukoski , E.E. and Kubota, T., "Entrainment

and Flame Geometry of Fire Plumes”, NBS-GCR-87-402 ( 1 982 ) , NBS

4. Tanaka, T., ’’A Proposed Model of the Behavior of Plumes in

Two Layer Models”, 8th UJNR Joint Panel Meeting on Fire Safety

(1985), BRI, Japan

5. Cooper, L.Y., Harkleroad, M., Quintiere, J. and Rinkinen,

W., "An Experimental Study of Upper Layer Stratification in

Full-Scale Multiroom Fire Scenarios”, Journal of Heat Transfer

Nov. 1982, Vol . 104

119
Fig.l Modeling of door jet plume penetration

«

note

)

a q • c » T« m •••% » c • ( T ® - Ti) m**-« -r k Q

Fig. 2 Effect of Modified Temperature and Flow Profile

Near Opening Region

120
Fig. 3 Limitation of Modeling of opening jet plume

by Gaussian Profile

Fig. 4 Model of Temperature profile Near Opening Region

121

A
Layer

Interface

Height

|

ta )

Temperature

l*C|

28

Bm

Fig. 6 Predicted and Experimented Results of Experiments

in USSR pavilion

122
1.750

0,250

,

5,000

Fig .-7 Area for the Experiments in IBM Pavilion
124

Fig.

8

Predicted

and

Experimental

Results

of

Smoke

Behavior

i

nI

DM

Pavilion
Office

Rooe C_L

XVoid' \

4- 10.000 20.000

10.000

PLAN

ll 8F Of flea Rooa^ >

L

7F

6F

L1—

L

3F T

lJL

SECTION

1 Ft ra Shuttar

IF Of flea Rooa

FI ra Sourca

Fig. 9 Outline of 8 story Model Building

T i me = 60 sec .

A thin smoke layer at

less than 30 *C begins to

fill the atrium.

i

T i me = 120 sec . The temperatures of smoke

layers are not yet very high

but smoke begins to enter the

rooms on each floors. On 7th

floor, the smoke layers expands by mixing with the air

flowing into the building

through the opened window.

Right after this point,

fire shutters are closed, but

one of the two shutters on

5th and 6th floors remain

open due to trouble.

At 125 sec., mechanical

air supply on 6th floor is

switched on.

Fig. 10 Smoke Layer Behavior in 8 story Model Building

125
240 sec.

T-rarng '-'•TT&rtv t -

- 5e?*J-vfc<S w»-vs ~ih-1- Jv«w3

.

* :*v. r „L

. - - '

: s^MlMNheaw^j&S^!:

i

.-v- nH ^^ jyv^ -vv>

.r

Time -

The smoke layer in the

atrium descends to near the

bottom and the layer temperature rises to more than 50

"C. The floors of which

shutters are closed are kept

comparatively safe. although

some infiltration of smoke

exists. On the floors of

which fire shutters are not

closed, significant amount of

smoke fills the rooms.

T i me = 300 s ec . The smoke is gradually

removed from the atrium

through the natural vent.

% ** poPi . * in*i| i

- , — .. aa

T i me = 600 sec

.

Hazardous condition

develops on 5th and 6th

floors, when the fire shutters failed to close, but the

temperature of 6th floor is

somewhat lower than that of

5th floor thanks to the

mechanical air supply.

u

Fig. 10 cont’d

126
Appendix 1 . Outline of the Model

The concept of the model is schematically illustrated

Fig.Al. This model assumes that opening jets behave just

fire plumes and partially penetrate through layer interfaces

Zone equations of this model are summarised as follow:

by

as

Fig.Al Concept of the Two layer Zone Model
Zone Equations

Upper Layer Temperature —Ti*= ^

( dt

1 c»p»iV*. dHmt 3

. t q

3 n»t . )

I C m* ' (T T .* ) -t-k* me (T9 -T. 3

) P a V 1

1

+ ( 1-k 13

) Mi* (Ti — T. 3) rMi' 1 (T. 3 — T. 3) -i-k’VMi 3 (Ti -T, 3

) + Z I ( l-k ss

) SS,. (TV-T. 3

) -i-SA'i . (T — T

.

3

) + k

sa SAji (TV-T. 3

) + ( l-k as

) ASu (TV-T. 3

) 4- AAV ( T

,

3 — T

.

3

) + k

aa AAi. (T, 3 -T. 3

) } ]

Lower Layer Temperature =

"c B pV"V. :

( . -t-q 3 n«t . )

> V C ( l~k« ) m» (Ts —

T

.*

)

+• Mi 3 ’ (T

.

3 —T. 3

) 4-k l3 .Vh 3 (T. -T. 3

) -r ( 1-k 13) Mi* (T> -T.*)

+ Z { SSV (T.’-T. 3

) +k**SSi. (Ti 3 — T. 3

) + ( 1 —

k

Sfl

) SA. • (Ti 3 -T.*

)

-i-AS’i. ( T 3 — T. 3

) f k

as

,\Sj . (Ti 3 — T. 3

) 4- ( 1 -kaa) .U,. (Ti 3 — T. 3

) ]

Mass Fraction of Species 1 within upper layer

TTV ’ i 3 = ^

1 w- [ Ti m» 3

. rmi ’ (Yu 3 — Y.

.

3

) — k* m* (Yu--Yu 3) at pi. v>i

+ ( 1 -k’VMi 3 (Ym -Yu 3

) ^Mi 3 ' ( Y i i " — Y i

.

3

) ^k 1

3

Mi 3 (Yu -Yu 3

) + 1 ! ( 1 -k ss

) SSi. (Yu 3 -Yu 3

) 4-SAV (Yu 3 -Yu 3

) ^k sa SA,. (Yu 3 -Yu 3

) '+ ( 1 -kas

) ASi. ( Y u* — Y u 3

) 4-AAV (Yu 3 -Yu 3

) ^k^AA,. (Yu 3 -Yu 3

) } ]

Mass Fraction of Species 1 within lower layer

j:Yu‘= 1 - Cr'm» 3

. - (1-k3

) m. (Yu -Yu 3

) dt p • . v*

.

-HMi 3 ’ (Yu 3 -Yu 3

) 4-k l3 Mi 3 (Yu -Yu 3

) +• (1-k 13) Mi* (Yu —Yu 3

) 4-Z !SS'i . (Yu 3 -Yu 3

) 4-k ss SS,. .(Yu 3 -Yu 3

) + ( l-k sa

) SA, , (Yh’-Yu 1

)

1 —AS ’j

. (Yu 3 -Yu 3

) -k as

AS,, (Yu 3 -Yu 3

) 4- ( 1 -kaa) AA,. (Yu 3 -Yu 3

) ) ]

Upper Layer Volume = -p"~v ~ ( A H m« 3

. — q

3

-<«>t . ) C 9 P 9 1 It

H W-7 [ k*maT» -r ( m«

‘ — k

3 m« ) T . 3 pi. 1 .

4- (1-k 13

) Mi 3 Ti — ( M

i

3

' — k

1

3

M

i

3

) T. 3 — k

13 Mi 3 T. -r ( Mi 3 ‘-k ,3 M. 3

) T. 3 -T. 3 Mz 3 rS { ( 1 -k ss

) SSi.Ti 3 - (SSV -k ss SS,, ) T. 3 + k

sa SA,.T, 3 + ( SA*, . -

k

sa SAi . ) T. 3

( 1 -kas

) AS.. Ti 3 - (AS’,. -k as ASi. ) T. 3 -rk aa AAJ ,Ti*-r (AA\ - k

aa

AA, . ) T. 3 - (SS.i+SA.i ) T. 3

) ]

Pressure Bal lance Equation

AH ( mb 3

; +mb‘ ) +• q

3 i«t. 4* q

3 i«t. rc« [T» mi t-Ti(Mi 3 — Mi 3

) — T.’M2 3 — T. 3 M2 3 + Z (Ti 3 (SSi.-rSAi.) rT, 3

( ASj . -t-AAj . ) -T . 3 (SS. ,+SA», ) -T . 3 (AS., -Mi,) ) ] = 0

12S
Appendix 2. Formation of Plumes (Ref. 2)

The flow rate of fire plume is calculated using the formulas developed by Zukoski et al(Ref.3). Opening jet plumes

are modeled as vertical thermal plumes and to calculate the

flow rate of the plumes at the height of Z, the equation of

plume flow rate for far field region by Zukoski et al(Ref.3) is

also applied.

Fictitious point heat source is assumed at a certain distance below an opening jet. The distance is given as

m i ^ -j — / •P'

! -<A1)

where

-(A2) Q = c o I Ti“T<a I m >i where Tj and T = are the temperature of opening jet and the

ambient layer, and rna is the mass flow rate of opening jet.

With the use of Z a , m: is also expressed as

-(A3) _ ( Z f Z 3 ms - m j

( ^ j

) —

Gaussian profile is assumed on the vertical velocity of

plume, then the mass penetration rate through a layer interface

is given as

- { ms ( 1- e ) ( A Tc< A Ta)

( ATcZ ZTa)

-(A4)

where 77 is plume radius in Howarth transformation. With

as the critical temperature for penetration,

ir - n ( AT- )

A

r 0 (A ~ K A T ? where A Ta and AT z are given as

is defined as

“(AS)

ATo= C (Ts-Ta)

<A6)

(A7)

In Eq.(A7), C is a constant to be determined empirically.

Similar consideration is applied on the penetration of heat

to yield:

q p«nt

C 0 T 03 m pent + k Q

0

where

k =

-(U/? 2 )y 2 T l-e

( Z T O < /j T a )

( ZTa)

(A3)

(A9)

129
RISK, HAZARD, AND EVACUATION SESSION

130
Progress in Fire Risk Analysis Methods

JOHN R. HALL, JR.

Fire Analysis Division

National Fire Protection Association

Quincy, Massachusetts 02269, USA

ABSTRACT

Progress in the development of fire risk analysis methods in the USA

since the Eighth UJNR Panel is described. There are several approaches

being pursued in parallel, involving different points of emphasis and

differing relationships with the rest of fire safety science.

1. INTRODUCTION

Fire risk analysis in the USA has progressed along several different

tracks since the Eighth UJNR Panel in 1 985. There are three principal

research approaches that may be defined by differences in emphasis. One

approach is favored by the nuclear power industry, the chemical industry,

and other economic sectors where the risks to be analyzed typically are defined by a single building or facility. The second approach is favored

by U.S. government regulatory agencies that must review regulatory

proposals affecting large parts of the general public. The third

approach is now being developed for use on fire risk evaluations of

products and furnishings in buildings.

2. FIRE RISK IN TECHNICALLY COMPLEX FACILITIES

Nuclear power plants in the USA are required to have probabilistic

risk assessments as part of the requirements for licensing. The model

typically used is a fault tree in which an undesirable outcome is

analyzed in terms of the logical and probabilistic relationships among

events that can produce or prevent it. Fault tree analysis is flexible

enough to include both mechanical and human factors. The measure of risk

is an overall probability of failure, where failure might be defined as a fire sufficiently severe to damage the safety components preventing core meltdown in a nuclear facility.

Because the emphasis is on measurement of the probability side of the

probabi 1 i ty/severi ty duality that defines risk, most research on this

approach also addresses aspects of probability estimation. The most

sophisticated treatment of uncertainty, for example, is found in this

track. Using Bayesian statistical methods, involving prior probability

distributions and rules for incorporating fire experience data, the

model -bui 1 der estimates not just a value for the possibility of each

event but a whole distribution for the values that that probability might

take.

This also involves close attention to the process of weighing

evidence. A recent paper by Ali Mosleh and George Apostolakis addressed

techniques for modeling levels of confidence in experts giving expert

estimates. Imprecise evidence on the probability of detection was addressed in a related paper by Nathan Siu and George Apostolakis.^

131
To make this approach useful for decision-makers, the concept of

acceptable risk is needed and reduces to an acceptable probability of

failure. Modelers have devoted considerable effort to establishing

conceptual bases for defining acceptable risk, generally in terms of such

natural reference points as unavoidable risk or the risk associated with

current practices or codes.

Professor Apostolakis of the University of California at Los Angeles

and his students are among the leading developers of theory in using this

approach. He will speak later in this session and provide more details.

Because this is the approach most firmly grounded in the US standards

system and most widely required by law, it is the most frequently applied

version of fire risk assessment in the USA today.

2. FIRE RISK IN REGULATORY AGENCIES

US government regulatory agencies increasingly are expected to

analyze the costs and benefits of proposed regulations. When the

anticipated benefits involve reductions in fire risk, the model they use involves a probability distribution over outcomes of measured severity.

This can involve the use of the decision tree format developed by Howard

Raiffa and others. Probability estimation is a major element of this

approach, but unlike the previous approach, this approach places equal

emphasis on techniques for measuring outcome severity. Outcomes are not

classified simply as failures or successes but are defined by the degree

of loss of life, health, or property. The overall risk measure is

therefore an expected loss.

This approach leans heavily on the methods of economics. It

therefore emphasizes such techniques as the reduction of several future

years of forecast costs and losses to equivalent present values and the

development of a dollar figure for the value of saving a statistical

life, that is, reducing the expected number of deaths by one. A recent

example in simple form is work done by the National Fire Protection

Association for the Cigarette Safety Act Technical Study Group. ^ In this study, a baseline of current fire losses involving cigarette

ignitions was created. Statistical models then were developed to project

those losses. Laboratory test results on the i gni teabil i ty of different

types of mattresses and upholstered furniture and the ignition

propensities of different types of cigarettes were incorporated into

probabilistic models, which produced parameters for use in the

statistical projections. Full-scale decision trees were not needed, but

the basic modeling format was of this second type.

Unlike the first approach, this second approach tends to address fire

problems in which a large quantity of data on real fires exists. This

reduces the uncertainty in the probability estimates.

There seems to be little development of new theory specific to fire

risk analysis now under way in this approach. However, this approach

represents the most corrmon approach to risk analysis generally in the USA

and therefore benefits from fundamental work being done on the general

models. Most of the current theoretical work addresses not the technical

points of estimating benefits but rather the related issues of

132
transforming expected outcomes into values embodying preferences. An

example of this is the ongoing research on factors influencing risk

aversion which, in simplified terms, is the tendency of most people to be

willing to pay more to avoid the chance of an unpleasant outcome than the

expected loss would seem to justify. For example, we might be willing to

pay $400 to avoid a one-third chance of having a fire that would damage

property worth $1,000.

3. FIRE RISK ASSESSMENT METHOD FOR PRODUCTS

The third approach is the subject of a major US research effort, now

in its first year, that is building a fire risk assessment method around

a core consisting of computerized models of the physics of fire

development, the biochemistry of toxic impact, and behavioral models of

human actions in fire. The other two modeling approaches use historical

fire experience, defined in statistical terms, as much as possible. They

address the physics and chemistry of fire primarily as sub-analyses that

permit the estimation of particular probability and severity

parameters. ’ This new third approach places the physics of fire at the

center of the model, using the HAZARD 1 method, which Richard Bukowski

will discuss later.

This approach works by subdividing the possible varieties of fire

into a set of scenario classes. Each scenario class is associated with a probability of occurrence, typically derived from historical fire data.

Each scenario class also is associated with a reference or benchmark

scenario. That is, a much more specifically defined class of fires is

taken from the larger scenario class, a sub-class with enough definition

to serve as a case for study by the hazard models and the human behavior

models. The latter models produce an expected number of deaths for that

case, and that is taken to be the expected number of deaths per fire for

any fire in that scenario class. Then the probability of the scenario

class provides the weighting of that death rate value in the overall risk measure.

This method involves a careful balance between the high level of

detail required by the hazard and human behavior models and the low level

of detail available in the data bases that provide probabilities. If the

scenario classes are too detailed, then the hazard model results clearly

will be appropriate to the entire class, but there may be very high

uncertainty in the estimates of the probability of the class. If the

scenario classes are not detailed enough, the probabilities may have high

confidence, but there may be low confidence that the benchmark scenario

truly represents the average scenario for the class.

Most of the theoretical work therefore involves the development of

crosswalks between data bases and models. For example, the many

varieties of items that can be involved in fire are grouped into a manageable number of classes defined by at most three or four values of

each of several key physical parameters--ra te of rise in rate of heat

release, ease of secondary ignition by a flaming heat source, peak heat

release rate. This permits the hazard model to be run on reference items

that represent classes defined by those physical parameters. Similar

techniques are used to address other needed crosswalks between data and

models.

133
The models can be tested by evaluating how well they reproduce

historical experience in terms of deaths per fire. The research is being

conducted by the National Bureau of Standards Center for Fire Research,

the National Fire Protection Association, and Ben jamin-Cl arke

Associates. A report on the model is due in August 1987.

4. THE FUTURE OF FIRE RISK MODELING

It is expected that the third approach will see a rapid rise in use once it becomes available. (The project to develop, test, and disseminate

it will be completed in August 1 989.) This new approach will satisfy the

demand for a model that permits explicit calibration in terms of real

fires, explicit modeling of new products, and extensive use of all other

branches of fire safety science. Tne rise in use of this model may be

constrained, however, in areas where legal requirements have already

created a presumption in favor of one of the other two approaches or in

situations where fast, inexpensive analysis is required, because the

third approach will not be fast or inexpensive, at least as it now appears. (By comparison, the first approach and particularly the second

approach can be done with different levels of detail, permitting

trade-offs of informational detail, precision, accuracy and completeness

against speed and affordability.) Over the longer term, all three

approaches are likely to borrow ideas and methods from each other,

because their different focuses are in many ways complementary.

134
REFERENCES

"The Assessment of Probability Distributions from Expert Opinions

with an Application to Seismic Fragility Curves," Ali Mosleh and George

Apostol aki s. Risk Analysis, Volume 6, Number 4 (December 1 986 ), pp.

447-461

.

"Modeling the Detection Rates of Fires in Nuclear Plants:

Development and Application of a Methodology for Treating Imprecise

Evidence," Nathan Siu and George Apostolakis, Risk Analysis , Volume 6,

Number 1 (March 1986), pp. 43-59.

3. "Final Report: Expected Changes in Fire Damages from Reducing

Cigarette Ignition Propensity," John R. Hall, Jr., NFPA Report to

Technical Study Group of- Cigarette Safety Act of 1984, August 29, 1 986.

^

.

Also appearing recently in the literature as an example of the first

approach: "Comparative Fire Risk Study of PCS Transformers ," Raymond F.

Boykin, Mardyros Kazarians, and Raymond A. Freeman, Risk Analysis , Volume

6, Number 4 (December 1 986), pp. 477-488.

135
PROGRESS REPORT ON RISK, HAZARD AND EVACUATION IN JAPAN

by

Takeyoshi TANAKA

Building Research Institute

Ministry of Construction

INTRODUCTION

In the area of research on the analyses of risk and hazards and on human

behavior concerning fire, it seems like the most important topics in Japan

are the completion of the two research projects carried out by Ministry of

Construction and Fire Defence Agency, respectively. So, it will be proper to

focus this progress report primarily on these two projects rather than

comprehensively covers the researches going on in Japan.

1. FIRE SAFETY DESIGN METHOD OF BUILDINGS

The research project: Fire Safety Design Method of Buildings was started by

Ministry of Construction in fiscal 1982 and carried out through the end of

fiscal 1986c The primary motivation of the project was to provide building

designers with a design method of fire safety for buildings usable as an alternative to the Building Standard Law and its associated orders, which are basically specification standards and whose structure has considerably lost

integrity since it has been basically unchanged for some 40 years.

In the Fire Safety Design Method, the objectives of fire safety for

buildings, which are abstructly and inclusively described in Article 35 of

the law, are expressed in more concrete form. Performance standards are introduced as long as possible and beneficial as the technical standards for

fire safety measures of buildings. The performance based technical standards

consist of Fire Safety Standards and Standard Fire and Occupants Conditions.

In principle, any design of building is supposed to be accepted as long as the design satisfy the fire safety standards under the standard fire and

occupants conditions.

Since the prediction methods for various aspects of fire behavior are indispensable to actually operate the fire safety design method, several

calculation methods and user oriented computer codes have been developed in

conjunction of the research projects, for example, one layer smoke movement

model, two layer smoke movement model, evacuation model, calculation method

for structural fire resistance, evaluation method for fire spread and so on.

The final report of this project will be published by the end of this fiscal

year. It is hoped chat the fire safety measures of buildings can be

significantly rationalized with the use of this design method.

136
2. FIRE RISK EVALUATION METHOD FOR MULTI-OCCUPANCY BUILDINGS

This research project was carried out from FY 1982 through FY 1986 by Fire

Safety Agency ,- Ministry of Home Affair. The objective of this project is to

provide fire officials with a practical and simple method to inspect and

evaluate the level of fire safety of existing multi-occupancy buildings.

In this method, the level of fire safety of an existing building is assessed

by the sum of the scores given to each of the primary fire protection

measures of the building. Scores of a fire protection measure is determined

by expert judgements or some other means according to its performance. The

system of the assessment consists of the two subsystems, i.e., the evaluation

method of potential fire risk and the evaluation method of evacuation

performance. Both of the evaluations are made on each floor basis.

3. SOME OF THE OTHER RESEARCH ACTIVITIES

Jin and Yamada conducted the experiments to investigate the effects of heat

radiation from smoke on evacuees as a continuation of their experimental

studies on emotional instability of evacuees in smoke. The effect of heat

radiation was assessed by the degradation of mental arithmetic ability of the

subjects

.

Tsujimoto proposed a method to evaluate life risk in an attempt to determine

the ' maximum allowable life risk due to fire in buildings. He applied the

method to existing hospitals and hotels and studied the relationship of the

life risk and the conditions of the buildings.

Field measurements were made by BRI for the egress behavior of the audience

of the pavilions of Tsukuba International Exposition, which was held in 1985.

Seven theater type pavilions were selected for the measurements and the

results were used for the validation of the evacuation model developed in the

5 year project of Ministry of Construction.

Field fire experiments were conducted by BRI also using the Tsukuba Expo. '85

pavilions to investigate the behavior of smoke and the effect of smoke

control measures. Eight pavilions with various geometries and smoke control measures were selected for the experiments and the results were used for the

validation of the two layer zone model for smoke movements.

REFERENCES

1) TANAKA, T. : Basic Structure of The Evacuation Safety Design Method, to be

presented at this session

2) SEKIZAWA, A. : Fire Risk Evaluation Method for Multi-occupancy Building,

to be presented at this session

3) JIN, T. and YAMADA, T. : The Experimental Studies on Emotional Instability

in Fire Smoke(Part 5 Radiative Heat Effects on Evacuees in a Fire Smoke) , Annual Meeting of Architectural Institute of Japan, 1986

4) TSUJIMOTO, M. : A Method of Fire Risk Analysis, to be presented at this

session.

5) TAKAHASHI , K. and TANAKA, T. : An Evacuation Model for The Use in Fire

Safety Designing of Buildings , to be presented at this session

6) NAKAMURA, K. : Predicting Capability of A Multiroom Fire Model, to be

presented at Fire and Smoke Physics session of this meeting*

137
Progress Report to the 9th UJNR Panel on Fire

RISK, HAZARD AND EVACUATION IN JAPAN

T. Tanaka, Building Research Institute, Japan

NELSON: Do you expect the fire safety design method to become part of a law

in Japan?

TANAKA: It is my personal understanding, and I will be able to explain it more in detail later, that the fire safety hazard law probably will be

administered in parallel for some time with the construction law. It also is

my understanding that the administrative agencies that have control of this to some extent understand, and are leaning towards, a revision of the law.

138
A METHOD OF FIRE RISK ANALYSIS

-Estimation of Life Risk in BuildingsMakoto TSUJIMOTO

Department of Architecture

Faculty of Engineering

Nagoya Uni vers i ty

Furo-cho, Chikusa-.ku, .Nagoya

464, JAPAN

ABSTRACT

As a way to determine the maximum admissible of life risk

caused by building 'Tire, a method -to calculate the life risk in

each floor of the -bui Idling is "proposed, and the distribution of

life risk covering the buildings (hospitals and hotels) in

Prefecture A in Japan is shown.

1 . Introduction

To protect a life of occupant is the minimum requirement

for the fire safety of the building. But it is generally

thought to be impossible to remove all the risk of a fire

from the building. So the maximum permissible of life risk must

be shown in some way in order to design the fire safety of a building. Until today, however, any method of risk analysis does

not enable us to estimate the life risk of an arbitrary space in a builbing quantitatively , and also the maximum admissable value

has never been proposed. Consequently the fire safety design has

been affected unreasonably by the fire regulations which always

had been established in haste after the severe fires.

In this paper, after the definition of the life risk, a method to estimate the life risk in each space in a building, and

the distribution of the values of the risk calculated by the

method for the existing buildings, are shown.

2. Definition of Life Risk

In this paper the life risk is defined as the expected

frequency that the occupants are trapped by the smoke generated

by a fire. The process of calculations is shown in Fig.l and as foil ows

.

First the size of the set corresponding the potential risk in

the space A, which is caused by the fire in the space I, is

given by the frequency of fires in the space I. When Fi stands

for the set, Fi is the life risk in the space A in case that

there is no measure against a fire.

139
Second if the sprinkler system is equipped or interior

finishing is restricted in the space I, Fi is reduced by

multipling the factors, Sp and/or In, which indicate the effect

of these coun termesures

.

Third the subset corresponding the case, where the evacuation

time is shorter than the smoke spread time, is deducted from the

set (Fi*Sp*In)

.

Forth the subset coresspondi ng the case, where there is

at least one staircase for evacuation free from smoke is deducted

also from the set. In third and forth processes it is assumed

that the smoke spreads at the fixed time (180 seconds) instantly

to every space through the doors left open, and that the smoke

does not spread across a closed door.

And the rest of the set is the life risk in the space A caused by the fire in the space F. Finally the life risk of space A .can be obtained by summa- rizing the risks caused by the fires in each space in the

building. In these calculations, it is assumed that the

situation to be trapped by the smoke.will happen only in the

spaces above the fire floor.

3. Calculation of Life Risk

The following equation is used to obtain the life risi: in

the space A.

n Ra= H, Fi *Sd i * I n i_*Plai *P2ai

i. Ra ; 1 i f e risk in the- space A

Fijfrecuency of fire in the space i Fi =f *Ai

flfire rate per unit floor area

. Aijfloor area of space i

Sp;factor indicating the effect of sprinkler system

Spi=0.1 sprinkler system is equipped with in the space i Sp i = 1 - is not equipped with

I n ; factor indicating the effect of interior finishing

In=0.5 interior finishing is restricted in the space i

I n= L

>• is not restricted

Plai ; probab i 1 i ty that the occupants in space A can finish evacuation by the smoke spread time from space i when Ts>Te then Plai=0

when Ts<Te then Plai=l

Ts;smoke spread time

Ts= 1 80

Te ; evacuat i on time from space A Te=Tl+T2 + T3+T4

Tljtime from an autbreak of fire to the detection

T2;time from the detection to the start of evacuation

T3; evacuation time from space A into staircase

T4; evacuat ion time in staircase

140
P2ai ;probabi 1 i ty that the occupants in space A can use

at least one staircase free from the smoke for

evacuation. It is derived from the probabilities

(shown in Table 1) that the each opening between

space i and staircase is open or closed.

n;the number of spaces, the fire in which causes the

life risk

4. Evaluation of Life Risk

To evaluate the life risk of a building by the method stated

above , it is necessary to survey the distribution of the life

risk of existing buildings. So 26 hotels and 23 hospitals are selected by random sampling from the buildings, whi-ch exist in a prefecture in Japan and each floor area of which is over 500 m

2 . And the life risks in each floor of these sampled buildings are calculated, and the distribution drawn from those values is shown

in Fi g. 2

.

By comparing the life risk of the floor intended to evaluate

with this distribution chart, ihe degree of the risk is quantitatively known on the whole.

By the way it is very difficult to show the propriety of the

assessment method -of this sort. Then to inquire into this

evaluation method, the Life risk in each floor shown in Fig. 2 is

compared with the characters of the floor, which are the time of

completion of the building, the number of the floors, and the

total floor area below the floor to be evaluated.

141
142
probability of a staircase door closed in case of fire

an interlocking fire door

vith a scoke detector a fire door with a fuse

a fire door closed

at ordinary tines

l fire door closed

aanually

ukok« detector /

heat fuse V

0 . s 0 . 4 0 . 7 0 . 1

an interlocking

fire sftutter vith

a smoke detector

jaota detector / i

a fire shutter closed

oanua-lly

"t

an opening

0 . 8 U . -1 0

Table 1. An example of the probability of a door closed in case of fire

The probability or' a -door with an interlocking closed was derived from

the result of testing the interlocking function and to the probaoility

of a door without an interlocking closed was applied the ratio of doors

kept close to all doors at the time of inquiry.

40i

2 30-

o

0

•3 20- U

VA

1 101

HOSPITALS

total nuaoer of floors 02

0

Ink(Ik. 200 4C0

life risk 6C0X r

t

)

40-1

” 30-

O

0

*5 20 - U

uA

1 I0 J

HOTELS

total nuader of floors 02

0

inn rfl

Figure 2. Distribution of

in a prefecture

2C0 4C0

life risk 600 x r

( tla«a/year

)

143

risk
Figure

200 -

100 -

1950 1960 1370 1980

a) relation between the life

xf

500

400

t 300

200 -

100

' " '

1 XT HOSPITAL HOTEL

500-

• • 400-

•

•

o

c 3 0 0^ •

• c

e

• 200- ® c

e 0 • e

. t • • * 100 • • *

• •

• « •

t .

i

1 ,

i 1 1 ; ' * ‘ ‘

1 1

1 '

' • •

• 23456789 10 23 456739 10 b ) relation between the life risk and the number of the floor

u

v—

xf

500

400-1

300

200 -

100 -

XT i HOSPITAL

500- • HOTEL |

e

•

•

•

400-

•

• 300^

•

•

•

e

o

200- . .

© c

« • • « • • • ° * *

. .*•.•**

100-

* €

« ® * « ©

•

t A

. --r - (

)

0 1000 2000 3000 4000 5000 0 1000 2000 3000 4000 5000

c) relation between the life risK. and the total floor area

below the floor evaluated

(ffl 2

)

3. Relation between the life risk and the characters of the floor

144
A METHOD OF FIRE RISK ANALYSIS

M. Tsujimoto, University of Nagoya, Japan

APOSTOLAKIS: You mentioned early in your talk that there was an acceptable

level of risk somewhere. How was that acceptable level determined? By whom?

TSUJIMOTO: This distribution is properly determined, likely, in a prefecture.

We decided the first 10 percent will be considered to be at risk. We decided

this is the risk factor. And, furthermore, for the local agencies, it's very

difficult to determine what is really the range of the fire risk.

In other words, among the officials, it's difficult to reach a consensus.

APOSTOLAKIS: Was a consensus reached?

TSUJIMOTO: As a committee we reached a consensus.

HALL: I noticed that some of the factors are dealt with as probability or probability type factors, and some take time effects and translate them into

probability

.

Are there any special problems involved in trying to mix these two things in one model?

TSUJIMOTO: As I explained, this is very difficult to get a rough estimate. I would like to explain the relationship between the two, and by utilizing the

variable and statistical information and so on, I'd like to really deal with

that.

145
Response Surface Approximation of a Fire Risk

Analysis Computer Code

Mark Brandyberry and George Apostolakis

Mechanical, Aerospace, and Nuclear Engineering Department

University of California

Los Angeles, CA 90024, USA

ABSTRACT

Fire Risk Analysis as currently used in nuclear plants is outlined along with a discussion of uncertainties in the methods. The use of the computer code COMPBRN is discussed for use in the fire growth analysis along with the use of response surface

methodology to quantify uncertainties in the code's use. Generalized response surfaces

are developed for the hot gas layer temperature and depth for a room of arbitrary

geometry as well as a surface for the time-to-damage in a scenario typical of those

analyzed in a nuclear plant fire risk analysis.

1.

INTRODUCTION

The evaluation of the risk associated with fires in nuclear power plants has attracted

considerable attention in the last few years. While it was believed earlier, on the basis of

crude estimates, that this risk was very small and, therefore, dominated by other risks,

more detailed analysis showed that the risks from fires could be. significant.. The

methodology is focused on the effects of fires on control and power cables and traces the

impact of their failure on the plant. Furthermore, the scenarios that are identified have

very low frequencies and special care must be taken to display the uncertainties. We have

found that the Bayesian interpretation of probability [1—2] provides a convenient

framework for the explicit and rigorous treatment of the uncertainties and the subjective

judgments that are necessary. Since the basic methodology has been presented elsewhere

[3 — 4] , its details will not be presented here.

The methodology is centered around the following three tasks:

1. The identification of the "critical" locations and the assessment of the frequency

of fires. 2. The estimation of fire growth times and the competing detection and suppression

times.

3. The response of the plant.

A location is declared as "critical", when the occurrence of a fire there has the potential

of creating an abnormal condition in the plant and damaging the engineered safety

functions (ESF). Thus, tasks 1 and 3 are related. As stated earlier, only fires that affect

control and power cables are considered. Figure l shows a situation, where the cable

trays are elevated. The plant analysis has already determined that trays A and B carry

cables, which if failed, would lead to undesirable consequences. The occurrence of a large

fire on the floor, as shown, could fail these cables, therefore, this location merits a detailed

146
analysis, and is "critical". In other situations the cable trays are closer to the floor and

one worries about a fire starting on one tray and propagating to the others.

In the second step the time Ta that it takes for the fire to propagate and damage the

cables of the trays is estimated using the computer code COMPBRN [5 — 6] . This code

uses empirical correlations for such quantities as the fire's mass burning rate and flame

height to characterize the fire as a heat source. It also employs heat and mass balances

to determine the extent and temperature of the layer of hot gases accumulating near the

room ceiling. The heat transferred to objects within the room and the room boundaries

is then calculated and the thermal response of these receivers determined. Damage is assumed to occur when the surface temperature of an object exceeds a predefined limit.

The purpose of this paper is to discuss the uncertainties associated with the use of

COMPBRN and to present analytical approximations ("response surfaces") to the code's

output.

2. UNCERTAINTIES

As stated earlier, the computer code COMPBRN is used to calculate the time to damage a specified component, TG . This time we call the deterministic reference model

(DRM) estimate. The question, now, is what kinds of uncertainties are associated with

this number.

. The uncertainties associated with modeling of physical processes can be thought of as being of two kinds: the statistical uncertainty and the state-of-knowledge uncertainty.

The statistical uncertainties stem from the random nature of fire; if we perform a particular experiment a large number of times under 'identical' conditions and measure

Tg , we will obtain a frequency distribution. Our physical model cannot predict the

environmental fluctuations which lead to this distribution. These uncertainties are inherent in the nature of fire. Even if our knowledge of fire were to increase dramatically,

we would not be able to reduce these uncertainties significantly. Such an increase in

knowledge, however, would markedly reduce our state-of-knowledge uncertainties.

There are two kinds of state-of-knowledge uncertainty: the parameter uncertainty and

the model uncertainty. The parameter uncertainty is due to insufficient knowledge about

what the input to the code should be. The model uncertainty is due to simplifying

assumptions and the fact that the models used may not accurately model the true physical

process.

The DRM is a collection of approximate models which are valid under certain

conditions. The modei uncertainty is due not only to our uncertainty in the accuracy of

each model's predictions under the conditions they were developed for, but also to our

uncertainty in the synthesis of these independent models. One of our primary concerns

is whether or not the synthesis contains enough component models (i.e., if all important

phenomena have been modeled).

The three types of uncertainty that we have discussed, i.e., statistical, parameter and

model uncertainty, must be expressed in terms of probability distributions, which will be

used in the risk assessment. The development of these distributions requires, again,

147
substantial judgment, since conventional statistical evidence is lacking. It is judged that

the variation of the estimated propagation time is dominated by state-of-knowledge

uncertainties. This means that the uncertainties in the models and the parameter values

are considered to overwhelm the statistical uncertainty, an assumption which is reasonable in the light of the current state of the art.

3. THE DISTRIBUTION OF Ta As shown above, the distribution of the scenario fire growth time is a necessary output

of the growth portion of the analysis. For a specific scenario to be analyzed, most of the

parameters that must be used as input to COMPBRN are either fixed by the problem,

e.g., its geometry, or are known to a satisfactory degree. On the other hand, significant

uncertainties (at least to the extent that they must be represented by probability

distributions) exist regarding the thermal and combustion properties of cable insulation

and jacket material as well as other combustibles. The principal sources of uncertainty

are the lack of information regarding the composition of these materials and of

experimental evidence.

There are several methods for propagating the parameter uncertainties through the

DRM, (Monte Carlo, Latin Hypercube Sampling, response surface methods) and each

has its advantages and disadvantages [7] . The method used in most fire risk analyses

so far is to form a response surface from a sampling of the DRM output and then to propagate the uncertainty distributions through the response surface by Monte Carlo

simulation. This technique has the advantage of using a minimal number of computer

runs while allowing the exploration of several input variables.

Techniques for setting up the computer runs and then generating the response

surface(s) are generally forms of statistical experimental designs (e.g. factorial designs)

and regression analysis [8] . The approach taken in this work is to use a form of a

factorial design to set up the computer runs and then to form the response surface from

polynomials. The parameter distributions can then be propagated through the surface(s)

using Monte Carlo simulation.

4. TYPICAL SCENARIOS IN NUCLEAR POWER PLANTS

From the results of nuclear plant fire risk analyses which have already been

performed, it is seen that the major source of fire risk to the plant is the vulnerability of

important cables to fire damage. Also, there are only a few general scenarios which occur

frequently in the various analyses. These scenarios are all combinations of an exposure

fire (usually postulated to be an oil pool fire or equivalent) and one or more important

cable trays in proximity to the fire. The three geometries which appear most often are shown in Figures 1-3. In these scenarios, two propagation times are usually found to be

important: the time for the exposure fire to ignite the first tray, and then the time for the

fire to propagate through the stack of trays to damage a second critical tray.

One of the simplest geometries, the exposure fire directly underneath a target cable

tray, forms a good starting point for analysis. The objective of this analysis is to develop

a simple procedure, through the use of response surface methodology and the

COMPBRN computer code, to enable an analyst to calculate the propagation time T0

148
without having to resort to running the computer code. Providing sufficient flexibility in

the procedure to enable it to be used in differing situations (room dimensions, tray

elevations) is important. In most use of response surface methods to date, the focus has

been to develop a surface for a specific geometry and fire situation to be used for

propagating uncertainties. The inclusion of geometrical variables complicates the

procedure tremendously, as we shall see.

The analysis begins by considering the factors which will affect the scenario and are known or thought to be important. Considerable experience has been gained with the

COMPBRN code at UCLA, and this has enabled the variables to be investigated to be

selected on the basis of known importance to the scenario results in most cases. For the

scenario in Figure 1., of major importance is the the characterization of the environment

which surrounds the cable tray. In the COMPBRN code, a target which is directly in the

fire plume is treated as being only subject to convective heat transfer from the plume. A

heat transfer coefficient is calculated on the basis of the fire strength and the tray height.

The effects of a hot gas layer (HGL), which may build up in a room, must also be

considered. COMPBRN treats the HGL as a single isothermal layer with a user-input

heat transfer coefficient (as in [9] ). Along with the plume or the HGL, a third possible

environment is when the tray is immersed in the flame itself. COMPBRN treats the

flame as a temperature-averaged isothermal cylinder with the heat transfer governed by

a user-input heat transfer coefficient. The code uses a simple "if-then" logic to calculate

the possible tray environments and in the presence of more than one (i.e., in the plume

and in the HGL), it will select the highest temperature environment.

The various temperatures and heat transfer coefficients are governed by a number of

different variables. The HGL temperature is governed by the fire size and the room

geometry, specifically,

rh0 = constant pyrolysis rate (kg/' m2 s)

Hf =net heat of combustion (kJ/'kg)

Rfl =fire radius (m)

Zc = ceiling height (m)

Ac = ceiling area ( m2

) Hd = doorway height (m)

WD = doorway width (m)

kc = ceiling thermal conductivity (W/'mK)

Since THOL is a continuous function of these parameters, to provide a means for

estimating THGL without using the computer code, a response surface has been developed.

Since COMPBRN is very inexpensive to run, an augmented full factorial design with

center and star points is utilized to=provide the data for the regression analysis. The

parameter ranges used are given in Table 1. Note that the response surfaces use variables

which are normalized to the difference from their means as given by the normalizing

equations presented in Table 1. As expected, the THCL turns out to be a complex function

of its regressor variables. The least complex acceptable surface includes 19 regressor

terms (see Table 2.) and reproduces the COMPBRN results within ±6K . A secondary

calculation with the same computer runs shows the depth of the HGL (from the ceiling)

to be well-defined by a simple linear combination of five of the regressors (Table 3.) which

reproduces COMPBRN's results within + 4 cm. It should be noted that the accuracy of

149
the response surfaces does not necessarily imply that COMPBRN's solutions are correct,

only that there is very little extra uncertainty introduced by using the response surface.

Characterizing the plume or flame temperature at the cable tray can be simply

accomplished by utilizing the same formulas that COMPBRN does. With several

variables either held constant and included in the coefficients or neglected due to insignificance to the model, a simple expression for the height of the flame above the floor

is [10] :

Hji = 30.64(*/-695(m/61

(1)

Depending upon whether the flame height is below or above the cable tray height Z

x

, (assumed known for an analysis) the environment temperature can be calculated from

[ 11 ]

:

„ Ktf/j ) 2,3

Tplume = 298( 1 + 0.00 1 3 1 ) (2)

m

Z < Hfl

' 1 flame Tf = ((- cPf R

7?

5.6697xl0

-8

( Rfl + 1Hfi

•)) 1/4

( 3 )

Thus, we now have reasonably simple equations to characterize the temperature of the

cable tray's environment.

The other environmental factor is the convective heat transfer coefficient at the tray.

For heat transfer within the flame or within the quiescent HGL, COMPBRN requires the user to input a heat transfer coefficient. In the flame. Ref. [12] suggests using a value or 20-23 W/ m2 K . We usually conservatively use 23 W/ m2 K . From experimental values

of the gas velocities in the HGL, a value of 10 W/ m2 K works well for heat transfer in the

HGL. Within the buoyant fire plume above the flame, COMPBRN internally calculates

a coefficient from:

‘plume = 4.97(

m0HjRfi

Z

1/3

(4)

This is adapted from [13]. We now have characterized the cable tray's environment.

Two other variables which control the cable damage time are the cable damage

temperature, TDam , and the cable thermal conductivity kcbl . Experience with the code

has shown that the thermal conductivity will account for 80-90 percent of the variation

in damage time with three other minor variables (cable density, specific heat, and

reflectivity) controlling the rest. For our purposes, the thermal conductivity is the

important variable.

Thus,, we see that:

Ta-f(Tem, ^emn ^ebb ^DAM)

Actually, it is not quite this simple and thus one of the problems with the response surface

approach arises: the dependent variable should be continuous over its range of

150
independent variables. When using a polynomial response surface to describe the

variation in a variable which has discontinuities, one must either judiciously choose the

ranges of the independent variables to avoid the discontinuities, of describe all deviations

from the fitted model to the user. A normal experimental design will vary each

independent variable over its entire range. In our case, as an example, this causes problems if, say, the plume and HGL temperatures are fairly low and the damage

temperature is high, since then the damage time will be infinite. We could leave these

points out of our model and still fit our surface, but then the surface will predict erroneous values of the damage time if a user attempts to evaluate the surface at one of

these discontinuous points. The discontinuities which arise when using COMPBRN and

an experimental design do not occur in regular patterns ( they do, however, follow the

physical models) and thus are very difficult to describe to the user. Therefore, it is necessary to carefully choose the independent variable ranges so as to essentially base our

analysis on : Time-to-damage given that there will be damage. Varying the damage

temperature also causes difficulties in continuity since a set of parameters which cause damage in a scenario with one damage temperature may not cause damage with another

(higher) damage temperature.

As a very simple example, we evaluate the time-to-damage surface for a cable tray

with a damage temperature set at a typical value of 550K. Since COMPBRN uses the

simple model of choosing the hottest environment to surround the cable tray, we may

investigate the cable damage time with respect to the local environment temperature and

heat transfer coefficient and the cable thermal conductivity without regard to the source

of the hot gases.

The response surface given in Table 4 characterizes the damage time. This surface

may be used for any situation in which a cable tray is immersed in an environment of

hot gases without significant external radiation, whether it is a flame, plume, or HGL.

The value of the heat transfer coefficient used will depend upon the specific situation.

Note also that this is really a "time-to-temperature" surface since COMPBRN uses the

simple criteria of an "ignition temperature". Thus, the TDAM calculated could be specified

by the user as time-to-ignition for scenarios to be propagated through a stack of trays

above the target tray or as time-to-damage for a cable short-circuit simulation. The

reader is cautioned that due to the discontinuous nature of the damage-time data (the

cable temperature response may be quite erratic when the environment is less than 100K

more than the damage temperature and the heat transfer coefficient is varied) the surface

presented here predicts the damage time within ± 60sec. near the center of the explored

region, but may become quite inaccurate with combinations of variables near the

boundaries, (see Figure 4.)

5. PARAMETER SELECTION

The ranges of values which were chosen for the various variable were selected to enable an analyst to 'simulate' a fire scenario using geometries and materials which could be

found in a nuclear plant. The ranges of rh0 and H;

- were selected to cover a wide range

of flammable liquids as shown in Table 6 [14]. The range of fire radii encompasses sizes

from that of a small container of fluid to a 1 meter diameter spill. Previous PRAs have

postulated fire diameters of 1 foot (.305 m) to 2 feet (.61 m) in diameter. The room

dimension limits and door sizes selected were judged to be representative of the smallest

151
room likely to be analyzed and of the largest for which hot gas layer effects are significant

for the fire sizes postulated. For the time-to-damage surface, the TDAM was chosen

arbitrarily as a representative value. The Tenv range was selected with the knowledge that

the code will not predict damage with an environment less than about 25-50K greater

than Tdam , and that values of Tgnv > 800 K produce damage within the first time step (15

sec here) for a TDAM = 550K and any kem in our range. The henv range was calculated

from the approximate maximum value of Eq. 6. and the HGL value of 10 W/'mK.

7. REFERENCES

1. Kaplan, S., and Garrick, B.J., "On the Quantitative Definition of Risk," Risk

Analysis, 1:1 1-37, 1981.

2. Apostolakis, G., "Data Analysis in Risk Assessment", Nuclear Engineering and

Design. 71:375-381, 1982.

3. Apostolakis, G., Kazarians, M., and Bley, D.C., "Methodology for Assessing the

Risk from Cable Fires," Nuclear Safety, 23: 391-407, 1982.

4. Kazarians, M., Siu, N.O., and Apostolakis, G., " Fire Risk Analysis for Nuclear

Power Plants: Methodological Developments and Applications," Risk Analysis. 5: 33-51, 1985.

5. Siu, N.O., and Apostolakis, G., "Probabilistic Models for Cable Tray Fires,"

Reliability Engineering, 3: 213-227, 1982.

6. Siu, N.O., "Physical Models for Compartment Fires," Reliability Engineering. : 229-252,1982.'

7. Iman, R.L., Helton J.C., "A Comparison of Uncertainty and Sensitivity Analysis

Techniques for Computer Models," Sandia National Laboratories,

NUREG/CR-3904, March 1985.

8. Box, G., Hunter, W., Hunter, J., Statistics For Experimenters. John Wiley and

Sons, Inc., 1978.

9. Rockett, J.A., "Fire Induced Gas Flows in an Enclosure,"

Combustion Science and Technology, 12(1976), 165.

10. Thomas, P.H., "The Size of Flames From Natural Fires," Tenth Symposium Unt'l)

on Combustion, 844-859 (1963).

11. Zukoski, E.E., Kubota, T., "Experimental Study of Environment and Heat

Transfer in a Room Fire. Mixing in Doorwav Flows and Entrainment in Fire

Plumes," NBS-GCR-85-487, 1985.

12. Neill, .DT., Welker, J.R., Sliepcevich, C.M., "Direct Contact Heat Transfer from

Buoyant Diffusion Flames,"

Journal of Fire and Flammability, 1, 289-301, 1970.

13. Veldman, C.C, Kubota, T., Zukoski, E.E, "An Experimental Investigation of the

Heat Transfer from a Buovant Gas Plume to a Horizontal Ceiling, Part I," NBS-GCR-77-97, June, 1975.

14. Babrauskas, V., "Estimating Large Pool Fire Burning Rates,"

Fire Technology, 4(19), 251-256, 1983.

152
Figure 1. Figure 2. Figure 3.

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- 2.00 .— — — — —— — — — — « - I - 7 *3 -I . 2

* -0.75 -0.25 0.25

r uV 1 M (J NMF N r ri)HF Figure 4. Contour Plot of T

G VS . T and h

153 env env

A

ncn-p^yM<«\

I

r,

3>u
PARAMETER RANGE NORMALIZED NORMALIZING EQUATION

VARIABLE

m (kg/m“K)

o .02 - .10 XI = 100m - 6.

0

H

f (MJ/kg) 20. - 47. X2 = . 2963H

f - 9.926

R

fl (m) .09 - .50 X3 = 19.512R

f - 5.756

Z (m)

r 2.4 - 6.1 X4 = 2.162Z - 9.189

C , ' A (nT)

c 9. - 900. X5 = . 00898A - 4.0808

c “n

(m) 1.8 - 3.1 X6 = 6.154H

d

- 15.077

W

n

(m) 1.2 - 3.7 X7 = 3.2W

n

- 7.84

k (W/m^K)

c .30 - 1.3 X8 = 8k - 6.4

c ^env (W/m^K) 10. - 300. X9 = . 0 1 16h - 1.7978

pnv

T (K) env

575 - 800 X10 = . 0 1495T - 10.2776 env

k , . (W/mK) cbl

.09 - .35 Xll = 10.8503k , . - 2.6583

cbl

Table 1. Parameter ranges and normalized variables.

Regressor/ Coefficient

XI 11.50 X1X2 1.04

] XI

2 -4.58

X^ 11.27 X1X3 9 . on

X3 30.24 X1X5 -. 707 I X3^ 1.48

xA -2.42 X2X3 2.58 hrrX5 -2 . 7 X3X5 -.754 I AJ

1 7

. / jU

Xh -1.77 X3X6 -.621 1 X8 .542

X8 1 Ln O X3X8 -1.27 , ! 5 X5 -. 130

Cons tant= 420.1 j Table 2. Response Surface for

Regressor/ Coefficient

XI .0149

X3 .0247

X4 .4629

X6 - . 0^28

X7 -.0528

Cons tant = 2.723

Table 3. Response Surface

for the HGL Depth (m)

.

Regressor/ Coefficient

X9 -.4279

X10 -.3929

Xll .4599

X9X10 .3689

9 X9 .3705

xio 2 .4085

constant = 3.4560

Table 4. Response Surface

for Ln(T ) (Ln(sec))

vj

Liquid (kg/m K) H £ (MJ/kg)

Heptane . 101 44 .

6

Gasoline .055 43.7

Kerosene .039 43.2

Transtormer Oil,

hydrocarbon

~'.7)W

Fuel Oil, heavy .035 iy.7

Methanol . (TT7 20.

d

Table 5. Some Typical Burning Properties

of Combustible Liquids.

154
RESPONSE SURFACE APPROXIMATION OF A FIRE RISK ANALYSIS COMPUTER CODE

M. Brandyberry and G. Apostolakis, University of California, Los Angeles, USA

NELSON: There are models much more sophisticated that COMPBURN, that can give more accurate answers for multiplicity of locations. Do you consider them too

large of an effort in the risk analysis to attempt to use them?

APOSTOLAKIS: Well, the uncertainties are always a major issue with us, and

the amount of effort that we can spend on a project like this has to be

balanced against the accuracy of the result. Given the parameter and model

uncertainties that I mentioned, most people would agree, at least in the

nuclear safety area, that it's not worth going to more sophisticated models.

Of course, the assessment of the model uncertainty may very well require the use of these more sophisticated models.

NELSON: So we've got to prove models with models?

APOSTOLAKIS: Or experiments could help.

HALL: A quick specific question on the first exercise of looking at models.

You had said that the exercise included with the code itself is quite

conservative. Was that because the uncertainty range was almost all in the

high side of one? For that E value? That from point 8 to 4 , and if you

assume one is saying that the model is bringing back the same value, so if

it's wrong it's almost certainly low...

APOSTOLAKIS: That's right.

HALL: And this is the parameter for high values and then safer conditions, it

that correct?

APOSTOLAKIS: That is correct.

QUINTIERE: I have a question on the response. You're seeking algebraic

representations for predicting quantities of heat. In seeking these

approximations you need to be sure that you identify all the attempted

parameters. But many empirical correlations are being developed that would be

useful in guiding this process

.

The part that I don't understand is the value of using the response surfaces.

Could you explain the value of using response surface in risk analysis?

APOSTOLAKIS: There are several questions that need to be answered. First,

the issue of important parameters. This is, of course, always an issue

because in essence you are building a new model. We are determining what

parameters are important by reading the literature, knowing what turned out to

be important in past applications, and asking other experts. If there are any

doubts about one or two parameters, then we include them in the model and we see how the beta coefficients turn out to be.

155
Regarding correlations, we only use response surfaces where we do not have

correlations. For example, we can get the temperature in the plume without

using the code by using simple correlations. Then we do not develop a response surface; in fact, the correlation is the response surface. But for

the hot gas "Layer, we were not able to find any correlation giving the depth

or the temperature, and this is where these methods are useful.

Now, how useful are they? You have all the advantages of a simple polynomial,

working with a simple polynomial. For example, you can propagate

uncertainties analytically. You can see immediately by looking at the

polynomial which parameters are important.

In general, it is more expensive to do a Monte Carlo simulation, but that is

not the case here. And another advantage that I see is that the response

surface may be used by people who are not really interested in becoming

experts in fire propagation, although that may be undesirable.

156
AN ENGINEERING FIRE PROTECTION DESIGN ASSESSMENT SYSTEM

Harold E. Nelson

Center for Fire Research

National Bureau of Standards

Gaithersburg, MD 20899

ABSTRACT

A fire protection engineering design assessment system to provide design

engineers with a tool to improve their ability to appraise the overall fire

safety in a facility and evaluate the impact of fire protection measures is

described. The system uses deterministic fire protection techniques and is

being implemented through an interactive computer program. The inputs to the

system require data on building layout, combustible contents, fire resistance,

detectors, sprinklers, smoke control, occupants, and fire safety objectives or criteria. The outputs include the predicted fire generated conditions within

the building as a function of time and an evaluation of the predicted

performance of the facility with user specified fire safety objectives.

Introduction

This paper discusses the plan for and status of an engineering fire protection

design assessment system. It also addresses the use of elements of this plan

in recent analysis of a major fire incident.

Both the U.S. and Japan are actively involved in several different approaches

to use recent developments in fire science and the modeling to analyze fire

safety in buildings. The emerging products vary widely in scope,

sophistication, and commitment demanded of the user. In general, as the

degree of sophistication and adherence to first principles of fire sciences

increases, the level of ’ complexity , user commitment required, cost of the

hardware involved, and CPU time increase.

The object of this engineering system is to produce a credible package of

coordinated engineering tools that can produce reasonably accurate

predictions, be operable on the current generation of personal computers, and

respond quickly to the user's commands. Simplifying assumptions will

necessarily be involved and appropriate safety factors required.

The development of the engineering approach is preceding through cooperative

efforts with various U.S. Federal agencies. In each case, the objective is to

support the fire protection engineering decisions that those agencies make for

federally owned buildings or facilities. Two recent publications represent

partial output. These are:

1. NBSIR 83-3308, "FIREFORM" - A Computerized Collection of Convenient Fire

Safety Computations [1]. This report describes a menu driven collection

of simple models and formulas felt to be useful in fire safety problem

analysis. These and similar simple models and formulas are at the heart

of the engineering system.

157
2. NBSIR 85-3298, "The Basic Structure of the Fire Protection Design

Assessment System. "[2] This report describes the basic structure and

details the flow chart of the assessment system. Parts but not all of

the detailed assessment system are now in operation.

The object of the project is to produce one or more systems that are compatible with building assessment and regulation, scientifically credible,

and easily used by the practicing engineer. It is intended to keep the

required engineering effort at a level that makes it economically feasible for

most designs and building evaluations.

Background

Historically, the relationship between expenditures on fire safety and the

actual impact of these expenditures has been hard to establish. Until the

past decade, virtually all fire safety knowledge was empirical. Fire safety

decisions were based on subjective opinions on how to prevent the recurrence- of prior disasters. Materials were generally judged on the basis of large

scale tests.

At the time current practice evolved, this empirical approach was sufficient.

Most buildings were inherently massive and highly compartmented . Wood and

paper were the prime combustibles of concern. The rate of change in building

technology was slow. The accumulative history of how buildings reacted when

exposed to fire was a good prediction of future expectations. It was in that

atmosphere that the system of consensus codes arose. These codes covered not

only fire but all other basic safety functions such as structural and plumbing

requirements

.

In the development of these codes credible technology was included wherever it

existed. When such was not available, committee consensus judgment was used.

The ultimate goal being to develop a complete system whereby a builder had

sufficient information to propose a design that an authority (typically a building commissioner) could review and concur in that effort. The code

approach evolved in the period between 1930 and the mid 1960s. It continues

with little or no advancement to dominate fire safety decisions to this day.

The result is a rigid set of requirements The objectives and expectations of

the committees and other consensus bodies in setting the requirements are generally not recorded. Requests to these bodies for interpretations often

very massively depending on the membership of the committee when the request

is received. Consequently, the value and intent of any requirement is seldom

apparent

.

Virtually every code has an equivalency clause. The equivalency clause

permits alternative approaches provided equal performance can be achieved. It

is, however, difficult to demonstrate performance when the factors that need

to be considered are established by consensus.

As a result, the collection of codes, standards, and criteria rather than fire

safety frequently become the objective. Expertise becomes entombed in

158
relacing fixed requirements to building materials and systems. Innovation,

rational design, and cost control are constrained and frustrated.

Over the past several decades, however, a relatively small but fortunately

persistent group of research scientists and engineers have made important

advances in understanding and quantifying fire and the responses of buildings,

peoples, and fire protection systems to fire. Their efforts have been

dedicated to determining the basic principles of unwanted fires and measuring

the variables involved. In recent years the scope has expanded to the

development of coordinated engineering approaches to predicting the course of

fire, the response of fire safety features, and the resulting impact on people, property and productive missions.

There is now an emerging fire protection engineering technology which can, in

many cases, be used to evaluate the fire safety performance of a building even

if it differs widely from the current perceptions of a code conforming

building. This same technology can be adapted to assess the impact of a proposed or existing safety requirement as applies to a specific inventory,

building, or set of circumstances. It is now possible to make a reasonable

engineering analysis of fire development and impact from the moment of

ignition to the final determination of the results of fire.

The design process starts with a given building condition. In fire safety

design this encompasses:

a. The building, its layout or shape, materials of construction, subsystems (utilities, electrical, fire alarm, sprinklers, etc.);

b. The intended use of the building, particularly in terms of the

combustible materials and hazard of operation involved in the activities

of the occupant;

c. The people in the building in terms of their areas of activity and

any consideration of special physical impairments (handicapped or disabilities) or other characteristics such as mission responsibilities

to stay in place as long as possible at time of emergency.

As a practical design problem, some of these qualities are reasonably fixed in

any building and some change from time to time. Many of these variables can affect the course or level of danger presented in a fire. To attempt to

analyze every conceivable variable and all the combinations of variables is

irrational

.

The engineering method, that is at the heart of the engineering assessment

system, follows the technique of design stress. The approach is similar to

that used in structural design. In this case, the stress is thermal and

environmental rather than a gravity or dynamic force load. Assessment is

based on fire or occupant situations that represent the most stressful

condition that the building is expected to endure. This design fire stress

starts with the expected fire load conditions, either typical for the

occupancy class housed or specific to a known case of concern. Other

conditions such as the position of doors (i.e., open or closed) and the

159
location and disposition of people in the building are evaluated on the basis

of

:

a. the most vulnerable situation,

b. a selected situation of particular interest, or

c. a series of selected design cases to determine the impact of one design

case versus another.

Fire is viewed as an energy driven stress on a building and its internal

environment. The building reacts to this stress by absorbing it, removing it, or undergoing change in response to the stress.

In recent years, the fire research community has made significant progress.

Fire modeling and similar activities can be used to discover, quantify, and

express in engineering terms the physics of fire and the manner that a building, its systems, and its occupants cope with fire stress.

The engineering assessment system is designed to bring current state-of-theart to the practicing fire safety professional in a form that is applicable to

the facilities involved and useful to and usable by fire safety professionals.

The Approach

The overall approach is shown in greater detail in NBS Report, "The Basic

Structure of the Fire Protection Design Assessment System. "[2] Figures 1 through 4 are flow charts from this report. These flow charts depict the

structure of the engineering assessment system. The specific approach

involved in the project is:

1, Assembling and as necessary advancing the current technoloev.

Much of the technology for the engineering approach exists. Some needs

refinement. In some cases new research, data development, and technology

development are' needed. During the past year, CFR has exercised many of the

elements in both in-house experiments and use of the engineering formulas to

mathematically reconstruct the conditions that occurred in serious accidental

fires .

Of critical importance to this project, a major part of the current effort is

being directed towards the enhancement of the Available Safe Egress Time model

(ASET.) [3] ASET is being revised to improve the manner in which heat

transfer to the walls is calculated in that model, extend the model from a single room with closed doors to one that can handle a series of rooms with

doors open, and provide a means for geometric changes (breaking of windows,

opening or closing of doors, and similar variations) during the course of the

computation. This will be the core model of the engineering assessment

method.

The improved core model and a considerable number of supporting computations

and formulas are being assembled into two packages. The first package will be

menu driven collection of all of the pertinent formulas and an accompanying

data base. This will place into the hands of the fire safety community readily

160
accessible tools aimed at fire protection engineers and other comparable

professional. A preliminary version of this is currently in broad circulation

under the title "FIREFORM. "

[ 1

]

2. Merging the elements into a single coherent system.

The second objective is to assemble the core model and accessory formulas

along with the appropriate data base into an interactive computerized system.

The system to be usable by persons who are familiar with building construction

materials and fire protection systems but not necessarily the detailed

engineering formulas. The user will be provided with data input screens at

his terminal. He will enter information such as the space dimensions,

materials of construction or finish, and type of occupancy (office, warehouse,

storage, trash room, etc.) The system will process the input and return and

appraisal of the conditions to be expected in case of fire and what safety

margin is likely to exist between the time needed for evacuation and the time

the space reaches intolerable conditions.

The output of the system will also be able to report the formulas and

variables internally used or developed by the system.

Current Uses

Currently, many of the elements of the engineering assessment system reside as individual programs on the author's computer and formulas available in various

publications. These have been used many times in responding to individual

fire protection engineering queries. Most recently, a number of these

formulas and procedures were collected and used to make an engineering

reconstruction of the fire phenomena that occurred at the Dupont Plaza Hotel

and Casino fire of December 31, 1986. In the analysis of the Dupont Plaza

fire, an attempt was made to accomplish the evaluation using methods that can

be quickly executed on personal computers or hand-held calculators. The

methods employed are, for the most part, elements destined for inclusion in

the engineering appraisal system. The objective of using these simpler

approaches was nearly accomplished. It was necessary, however, to resort to a more comprehensive model (emersion of the Harvard Model referred to as FIRST

[4] recently developed by Mitler and Rockett) to develop a credible estimate

of the development of the smoke layer in one of the rooms and the flow of fuel

from one room to another. As previously noted, the development of a rapid

response model to make reasonable estimates of smoke level development in

vented rooms and the flow of fuel from space to space are priority needs for

the development of the engineering analysis system.

A report of the analysis of the Dupont Plaza fire has been prepared by

Nelson. [ 6

]

Status and Direction

With the increasing use of the menu collection, "FIREFORM" and the exposure of

the concept resulting from its use on the Dupont Plaza fire, there has been a renewed interest in the engineering system. This has resulted in additional

support and the initiation of a three-year project. The object of this

161
project is to develop a working system, readily usable by practicing designers

and fire protection engineers. The scope of this initial version will be

limited to room and corridor arrangement in buildings that do not have

significant, unprotected vertical openings. The system will address smoke and

heat leakage from the floor of origin into stairways or shafts but will not,

at this stage, address the movement through the shafts or onto other floors.

The project plan calls for having an operating prototype system within one year and an implemented system in the hands of practicing engineers within

three years.

References

1. Nelson, H. E. ; "FIREFORM" - A Computerized Collection of Convenient Fire

Safety Computations; Report NBSIR 86-3308; National Bureau of Standards;

Gaithersburg, MD ; 1986

2. Nelson, H. E. and Walton, W. D,; The Basic Structure of the Fire Protection

Design Assessment System; Report NBSIR 85-3298; National Bureau of Standards;

Gaithersburg, MD ; 1986

3. Cooper, L. Y. , and Stroup, D. W. ; Calculating Available Safe Egress Time

from Fires; Report NBSIR 82-2587; National Bureau of Standards; Gaithersburg,

MG; 1982

4. Mitler, Henri E. and Rockett, John A.; Users' Guide to First, A

Comprehensive Single-Room Fire Model; NBS report to be published, National

Bureau of Standards, Gaithersburg, MD

5. Nelson, H. E.; An Engineering Analysis of the Early Stages of Fire

Development - The Fire at the Dupont Plaza Hotel and Casino, December 31,

1986; Report in preparation.

162
FIRE PROTECTION DESIGN ASSESSMENT SYSTEM

Figure 1. Fire Protection Deeign

General Model

Aaaeaement Syatam

163
Figure 2. Building Data Element

164
BUILDING DATA MODULES

ARCHITECTVFAl PATA

ROOM OF ORIGIN

DIMENSIONS

MATERIALS

OPENINGS

LEAKS

ADJOINING SPACE < S

)

DIMENSIONS

MATERIALS

OPENINGS

LEAKS

FUEL PACKAGE DATA

FIRST IGNITED

RATE OF:

ENERGY PRODUCTION

MATERIAL DESTRUCTION

PRODUCT (S) PRODUCTION

LOCATION

EXPOSED FUEL PACKAGES

BURNING RATES

IGNITION SUSCEPTIBILITY

LOCATION

TOTAL FUEL (ROOM OF ORIGIN)

MASS AVAILABLE

PYROLYSIS RATE

STRUCTURAL FIRE RESISTANCE

COMBUSTIBLE FUEL LOAD

RATED FIRE RESISTANCE OF BUILDING ASSEMBLIES

Figure 2. (continued) Building Data Element

165
BUILDING DATA MODULES

DETECTION EQUIPMENT

ROOM OF ORIGIN

SMOKE

LOCATION

HEAT

POSITION

RESPONSE TIME

ACTIVATION TEMPERATURE

SPRINKLERS

ROOM OF ORIGIN

LOCATION

RESPONSE TIME INDEX

ACTIVATION TEMPERATURE

SMOKE CONTROL

VENTS

SIZE (CAPACITY)

LOCATION

INITIATION METHOD

HV AC OR SMOKE CONTROL SYSTEM

PRESSURE DIFFERENTIAL

INITIATION METHOD

OCCUPANTS

NUMBER OF OCCUPANTS

CAPABILITIES

CONDITION (AWAKE, ASLEEP, ETC.)

MAXIMUM TRAVEL DISTANCE

DOORS

STAIRS

Figure 2. (continued) Building Data Element

166
Figura 3. Fira Safaty Qbjactivaa Elamant

167
CONDITIONS OR EVENTS FOR EXAMINATION

SMOKE LAYER TEMPERATURE

SMOKE LAYER LOCATION

SMOKE CONSTITUENCY

OPTICAL DENSITY

TOXIC CONCENTRATION

TIME TO FLASHOVER

FIRE DURATION

FIRE RESISTANCE OF A BUILDING ASSEMBLY

DETECTOR ACTIVATION

SPRINKLER ACTIVATION

SMOKE LEAKAGE

EVACUATION TIME

DESIGN LIMITS

HUMAN TOLERANCES

HEATED GAS TEMPERATURE

THERMAL RADIATION

SMOKE OPTICAL DENSITY

SMOKE TOXICITY

BUILDING ASSEMBLY TOLERANCES

FIRE RESISTANCE

SMOKE TRANSMISSION

CONTENTS TOLERANCES

HEATED GAS TEMPERATURE

THERMAL RADIATION

CHEMICAL EXPOSURE

SAFETY MARGIN

TIME AVAILABLE FOR EGRESS

TIME TO UNTENABLE CONDITIONS/TIME REQUIRED FOR EGRESS

Figure 3. (continued) Fire Safety Objectives Element

168
Figura 4. Raaulta Elament

169
AN ENGINEERING FIRE PROTECTION DESIGN ASSESSMENT SYSTEM

H. Nelson, National Bureau of Standards, USA

WATTS: You started by saying your objective was to find the least

sophisticated approach that would give you an adequate result?

NELSON: That's right.

WATTS: What is your criteria for adequacy in a result?

NELSON:- Technically, sufficient to make the judgment necessary with

professional confidence.

A.POSTOLAKIS: I wonder, since this was a real fire, whether you could get real

evidence to compare the model predictions with.

NELSON: I omitted one statement. I did not do this blind; I used evidence

and witnesses whenever there was question of a parameter, to attempt to choose

the right parameter.

JIN: Perhaps the answer will come out in the report that you alluded to

earlier, but my question is similar to the one that Dr. Apostolakis just

asked. Have you shown in the report how the calculations were done? I assume

that this was a real fire, and that you compared it with hypothetical

correlations. Therefore, will these be shown in your report, or is it

something that you might answer here in a different tone?

NELSON: I think the answer is "yes," and Dr. Jin will get one of the first

reports when they arrive.. If he gets a chance to look at it before the end of

the week, I would value his opinion as to whether I've achieved that or not.

TANAKA: If you're going to give report number 1 to Dr. Chin, I'd like to have

number 2

.

NELSON: I would be very happy, I'll sign it.

TANAKA: I am at a loss to understand why 85 people or so died in the casino.

Can you give some indication whether this was due to the lack of sufficient

exits, or was the fire too fast for them to make the exits?

NELSON: I think three items. The early stages of the fire were not

psychologically threatening and they were ignored. They had a maximum, by my

calculations which may not be that accurate, of about 40 seconds to egress

through a one meter wide door. There are disagreements about how many people

actually got out that door, but I believe in the neighborhood of 40 to 50.

Then, the smoke in the lobby was so intolerable -- this is a supposition -- this door was closed for protection purposes. At that point they attempted to

break windows to evaluate. Five or six succeeded in evacuating that way, but

then this window broke, and a gas flow with about 800 to 900 C flowed through

170
here, changing from gas to flame as it went in, I believe, about 20 seconds.

The deceased were found collected in front of that door, and near the areas where the windows were broken.

So, it was speed, egress capacity and delayed action.

CUSTER: In some of your diagrams there appear to be flames up in the ceiling

layer, yet you're reporting very low oxygen content?

NELSON: No, no. I skipped too quickly. The diagram I showed was 10 seconds

ahead of when the oxygen went virtually to zero, so I felt it was still

legitimate to show flame, through I'm not positive.

COTE: Since this was an existing fire in which there's been a fair amount of

reconstruction by virtue of being able to interview people, did you have any

surprises in the engineering method, in any of the pieces? Was there anything

that did not work as well as you thought it would, or did not come together as well as you hoped it might?

NELSON: The largest surprise to me was that a combustible material finish

that did not spread fire prior to flashover was a major contributor of fuel

after flashover. I think we need to pay more attention to a property that is

often referred to as heat of gasification. Also, I have learned how good

Drysdale's book is.

171
BASIC STRUCTURE OF THE EVACUATION SAFETY DESIGN METHOD

by

Takeyoshi TANAKA

Building Research Institute

Ministry of Construction

INTRODUCTION

Traditionally, fire safety design of buildings in Japan has been accomplished

through the specification standards described in The Building Standard Law,

The Fire Service Law, their enforcement orders and so on. While specification

standards are convenient to quickly decide the fire safety measures, their

drawbacks have long been recognized. They are usually inflexible to new type

of designs, new construction methods, and trivial differences from the

standards. Furthermore, literal application of specification standards can sometimes bring about demands which are not optimal but even ridiculous from

the viewpoint of fire safety performance.

In recognition of these -problems arising from designers, building officials

and others concerned, BRI started a research project called "Fire Safety

Design Method of Buildings" in an attempt to provide designers with a fire

safety design method alternative to the specification standards in The

Building Standard Law. This paper reports the basic structure of "The

Evacuation Safety Design Method" as a part of "Fire Safety Design Method of

Buildings"

RELATIONSHIP BETWEEN FIRE SAFETY DESIGN METHOD AND THE BUILDING STANDARD LAW

Since the primary motivation for developing Fire Safety Design Method of

Building was to establish a design method of fire safety for buildings that can be used as an alternative to the fire safety standards in The Building

Standard Law and its associated orders, it was thought to be important to

clarify the relationship of the two. At present, it is generally agreed that

the two should coexist for the time being as a dual design system for fire

safety measures, because the adequacy of Fire Safety Design Method needs to

be fully examined in the world of practice. Also, it is necessary that the

method have to be appreciated by many designers and building officials before

its full-scale application.

The relationship between Fire safety Design Method and the fire safety

standards in The Building Standard Law is illustrated by Fig.l. Since the two are supposed to function as a dual system of fire safety design, it follows

that the ultimate goal have to be the same for the two. In The Building

Standard Law, the principle for fire safety measures of buildings is given in

Article 35, but the description was thought to be too inclusive and abstract

to draw technical standards from it. So, in Fire Safety Design Method, the

principle was broken down into- the fundamental requirements for more concreteness as shown in Table 1.

172
MosC of the current technical standards in The Building Standard Law directly

specify materials, dimensions, construction methods etc. of various fire

safety measures of buildings, which is much of the source of complaints among

designers. So, in Fire Safety Design Method, attempts were made to write the

technical standards in terms of performance standards as long as possible and

beneficial. It should be kept in mind, however, that any technical standard

is more or less of temporary nature. They cannot help but to reflect the

current conditions of building construction and currently available

technologies and knowledges for fire safety. So, there is a good chance that

unexpected by the standards but still effective fire safety measures are invented by designers or others. Such a measure can be accepted in Fire

Safety Design Method as long as it meets the fundamental requirements.

Several fire tests are the almost only technical sources to supplement the

specification standards for fire safety in The Building Standard Law. On the

other hand, since predictions of various aspects of fire behavior have to

play a key role to run performance standards, Fire Safety Design Method is

provided with relevant calculation standards and computer codes to predict

fire behavior and evacuation conditions.

TECHNICAL STANDARDS FOR EVACUATION SAFETY

1. Basic Components of The Technical Standards for Evacuation Safety

On considering the basic structure of the performance based technical

standards for evacuation safety of buildings, the method being used in

structural design of buildings was taken as a model. In structural design,

standardized loading conditions regarding dead and live loads, earthquake and

wind are given and the stress which develops in each structural member is

required to remain within the allowable limit of the structural material

used. The loading conditions in structural design correspond to the fire

conditions in Evacuation Safety Design, which consist of conditions of fire

and occupants and fire scenarios. And allowable limit of stress corresponds

to human tolerance against the hazards associated with fire, that is; smoke,

heat, bruise and so on.

Basically, any design is can be accepted as long as it is proved that the

safety standards are satisfied under the standard fire conditions. The

concept of the procedure for the evacuation safety design using the

performance based technical standards is illustrated in Fig. 2.

2. Items to be Checked for Evacuation Safety

The items to be checked for evacuation safety of buildings are listed in

Table 2. Many of these have performance based safety standards which have to

be satisfied under relevant fire scenarios.

2.1. Restriction of Use of Building Materials Excessively Hazardous to Evacuation

This restriction reflects the toxicity tests currently required to the fire

proof materials by a Ministry of Construction Order. The technical standards

for evacuation safety has been constructed under the premise that extremely

hazardous materials are removed from buildings.

173
2.2. Adequacy of Evacuation Plan

In the current Building Standard Law, no description exists regarding how an evacuation plan has to be planned but the construction methods, materials,

dimensions, smoke control measures etc. of stairs, corridors, exits and other

building elements are directly specified without giving sufficient

explanation on the objectives. However, to make an evacuation plan of a building more effective in terms of performance and cost, designers have to

have clear strategies on how to evacuate the building occupants in the event

of fire and evacuation routes should be designed based on the strategies. The

intention of this item is to require that the designer of a building has to

frame up an evacuation plan which is realistic in the light of type of

occupancy, staff organization, fire detection and alarm system and other

concerning conditions of the building prior to the design of the evacuation

safety measures.

2.3. Security of Refuge

Final refuge of occupants in the event of fire has been traditionally

considered to be public ways or some sites of equivalent nature. Since in

such a place as a public way, people can get away freely whenever dangers due

to fire becomes imminent, public ways and such will continue to remain as the

refuges for most of the building in future as well. However, in case of

buildings from which occupants cannot be evacuated in reasonable time for some reasons, such as high-rise buildings and hospitals, there may be some cases where it is more rational to secure a refuge within the building or protect some parts of the building so as to be practically used as refuges in

the event of fire.

When a refuge is designed within a building or a site which is not connected

with public ways, the safety from various potential dangers due to fire has

to be carefully considered. The potential causes of danger for evacuees

include smoke, heat, air contamination, structural collapse and so on, as are shown in Fig. 3, so safety standards need to be considered on each of them.

Cl) Public ways and equivalent spaces

When a public way or such is taken as the refuge as is usual, virtually no check is required for the safety.

(2) Refuges on sites isolated from public ways

Although considered to be rare, refuges may have to be planned on building

sites which has no outlet to public ways in conjunction with the evacuation

routes within the building. This type of refuges need checks for the safety

against certain potential dangers.

(3) Refuges within buildings

We have little experience regarding refuges within buildings, yet sometimes

it is thought to be more reasonable to secure refuges within buildings rather

than to forcedly try to evacuate all the occupants out of a building. When

refuges are planned within buildings, however, sufficient level of safety has

to be secured for every conceivable hazards due to fire, which may attack

refuges through open air as well as through inside spaces.

174
2.4. Security of Safety of Evacuation Routes

It seems that evacuation safety means nothing but evacuation routes in The

Building Standard Law. Most of its standards concerning evacuation ' safety

primarily concern dimensions, materials, fire resistant rating and smoke

control measures of the parts of building which may play a role as evacuation routes in the event of fire. When we consider technical standards

for safety of evacuation based on performance, we have to analyze the

potential dangers associated with evacuation. As can be seen in Fig. 4,

obstruction on evacuation routes plays a key role in endangering evacuation

safety as well as the direct hazards such as smoke and heat.

The evacuation routes in a building are classified into the three for clarity

in design, that is, evacuation routes in the room of origin, those on the

floor of origin, and those on the floors of non origin.

3. Standard Fire Conditions

The standard fire conditions summarized in Table 3 correspond to the design

load standards in structural design and consist of fire conditions, occupants

conditions, fire scenarios ,. and other conditions affecting the evacuation

safety. A building which is designed according to Fire Safety Design Method

needs to satisfy the safety standards for evacuation safety under these

standard fire conditions. It should be kept in mind, however, that the

satisfaction of the safety standards does not necessarily mean to guarantee

sufficient level of safety but only secure a reasonable level of safety.

CONCLUDING REMARKS

’

The design method for evacuation safety incorporating performance based

technical standards for evacuation safety of buildings have been drafted.

They are given in terms of the safety standards and the standard fire

conditions. Basically, any design can be accepted as long as the safety

standards are satisfied under the standard fire conditions so that evacuation

safety measures can be rationalized in terms both of safety performance

itself and cost. Various prediction techniques for fire behavior and

evacuation will play important roles to operate the method.

A lot of excellent findings and engineering techniques do exist in fire

research area but there is little room to accommodate those in specification

standards. It is important to establish a performance standards for fire

safety so that those valuable results of researches can readily be transfered

to the world of practice.

175
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Table 1. Fundamental Requirements for Fire Safety of Buildings

I. REQUIREMENTS AS AN INDIVIDUAL BUILDING

1. Prevention of Occurrence and Rapid Propagation of Fire

2. Security of Evacuation Safety

2.1. Restriction of use of building materials excessively

hazardous to evacuees

2.2. Adequacy of evacuation plan

2.3. Security of safety of refuges

2.4. Security of evacuation routes

3. Prevention of Serious Social Troubles

3.1. Prevention of fife spread to others' space

3.2. Prevention of collapse or damage of building which

causes serious troubles of others or public

4. Security o‘f Fire Fighting Activities

4.1. Restriction of use of building materials excessively

hazardous to fire fighting activities

4.2. Security of activities for searching and rescuing occupants

4.3. Security of activities for extinguishing fire

II. REQUIREMENTS FOR CITY FIRE PREVENTION

1. Buildings in Fire Protection Districts

2. Buildings in Quasi-fire Protection Districts
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187
BASIC STRUCTURE OF THE EVACUATION SAFETY DESIGN METHOD

T. Tanaka, Building Research Institute, Japan

JIN: The report which Dr. Tanaka presented was the cumulative result of five

years of research by the entire staff of the concerned people in the

Construction Ministry, and for him to have condensed it into the very short

time slot that was allotted to him is a tremendous feat in itself.

HALL: I have two quick questions. Were the performance standards developed

using a group larger than fire scientists and other effected parties? The

second question, is the large number of independent standards an important

aspect in making the overall method conservative for regulation?

JIN: We in the Construction Ministry refer to this as the Overall Technology

Development Program. This program, which was explained, took five years. The

Construction Ministry worked in very close cooperation with both government,

private, and university people. I can't recall exactly how many subgroups

there were, but I believe there were more than 10 independent groups composed

of people from government, private and university levels. I do not believe

that we have arrived at a conservative standard.

Basically, what these standards ’call for is the establishment of minimum

safety requirements that are built into the materials themselves . Rather than

higher standards which provide the possibility of constructing better and

higher buildings at the same cost, which may be possible under this system,

that was not the major objective of our project. Our project was to call on

the expertise of the people involved with fire standards. As is the case in

the existing structural safety laws, which uses the same method in determining

their standards, the same materials may be used which would result in either a good building or a bad building. This is the expertise of the constructor.

All we're seeking in this regulatory situation is to provide a law which would

mandate a minimum safety requirement within each material this is used.

QUINTIERE: Since this is the work of many people, is there some documentation

upon which this work was established? How were the criteria established, by

underlying physics or just judgment?

JIN: At the present time there is no information broken down to that extent.

However, as I explained, this is the culmination of a 5-year program, and we hope to have a full report published by the year end which should contain the

information you're seeking.

QUINTIERE: I must have missed it.

JIN: You did not miss it, I purposely did not refer to it. As you probably

are aware, Japan, the United States and Canada are currently engaged in a joint study on toxicity, and I had hoped to have a chance to refer to that

before I made any conclusive statements. Actually in terms of technical

standards themselves, there are certain portions which are really not

conclusive as yet.

188
Fire Risk Evaluation Method for Multi -occupancy Building

A Report of Five Year Research Project

conducted by Fire Defence Agency

by

Ai SEKIZAWA

and

Tadahisa JIN

Fire Research Institute

Fire Defence Agency

Ministry of Home Affairs

1 . Introduction

This report intends to present a summary of the results of five year research

project for developing a "Fire Risk Evaluation Method for Multi - occupancy

Building" sponsored by Fire Defence Agency, Ministry of Home Affairs. This

research work was performed by the technical committee organized by Fire

Defence Agency, and ended in 1986 fiscal year.

The main purpose of this project was to provide a practical, manageable and

concise method for building fire safety evaluation, which will be mainly used

by fire officials in fire inspection.

Therefore, the scheme and selected factors for estimating fire risk and

evacuation performance in this method are focussed on those items that address

substantial fire protection measures such as sprinklers, fire alarm system,

fire walls/doors, compartmentation and so forth.

Additionally, this method is actually more like a scoring system than a scientifically-based risk prediction model, and also this method is designed

by using many assumptions and expert judgments. In this sense, this method

will need further verification and improvement, based on a lot of case

studies. But, still we consider this method to be useful as a practical tool

for fire inspection and also for self-diagnosis of building fire safety by

building owners.

2. Framework of the Fire Risk Evaluation Method

The framework of the Fire risk evaluation method is shown in FIg.l on page 2.

One of the significant features of our method is to evaluate each floor's fire

risk independently, because each floor in a multi-occupancy building should

have different potential fire risk corresponding to its variety of occupancy,

occupant load/characteristics, building configuration and features of fire

protection measures etc.

Also, the fire risk of a whole building can be estimated, using each floor's

fire risk and an additional modification coefficient in terms of vertical fire

spread risk both inside and outside of the building.

The other important feature of this method is that the evaluation system is

composed of two major subsystems ; the evaluation method of potential fire

risk and the evaluation method of evacuation performance, as shown in Fig.l .

189
The meaning of the former one is actually the estimate of potential property

damage, which we call simply potential fire risk hereafter in this report.

The final estimate of each of potential fire risk and evacuation performance

is expressed by using grading scores corresponding to the numerical estimation

results which are derived from each evaluation method described later.

In this report, we omitted the explanation about the treatment of such factors

as vertical fire spread risk, fire fighters' activity and fire prevention

management which are used for modifying the final estimate, simply owing to

limited space. But, we think it will not affect a readers' understanding of

this outline of our fire risk evaluation method.

Fire risk diagnosis of a building concerned

t

Fig. 1 Framework of "Fire Risk Evaluation Method"

190
3. Evaluation Method of Potential Fire Risk in Each Floor ( S )

3.1 Basic idea

Potential fire risk in each floor ( S ) is evaluated by considering potential

relative damage of fire in unit spaces, which a floor can be divided into.

The basic premises in the estimation process of potential fire risk are as follows ;

1) Each floor in a building is considered to be an aggregation of such unit

spaces as defined later. Each unit space could have different occupancy or use and also different size of floor area.

2) The potential fire damage in an unit space ( D ) is estimated by the sum of

the relative damage caused by a fire inside the unit space itself and the

relative damage suffering from adjoining unit spaces. So, such two factors

as the potential possibility of fire outbreak / fire spread in each unit

space, and the fire performance of partition components between adjoining

unit spaces are the main bases for estimation of potential fire damage ( D )_

3) Potential fire risk in a floor ( S ) is represented by a grading score corresponding to max[ Di ] in that floor.

4) Potential fire risk decreases in proportion to the number of compartments

in the floor, if the size of floor area is held constant.

Because compartmentation should work for limiting the spread of fire, and

also potential fire damage in an unit space should be smaller as the size

of unit space decreases

.

3.2 Estimation method of potential fire damage in each unit space ( D )

Potential fire damage in each unit space ( D i ) consists of roughly two components such as the

relative damage caused by a fire inside itself and

the relative damage suffering from the fire spread

beyond adjoining unit spaces. In. the illustration

of a simple floor plan, the fire damage of #3 unit

space ( D3 ) for example is considered to be the sum of following fire damages as ;

^3 ^3 + ^2*^2 3 + ^1*^13 + ^4**^4 3 3-1

where generally ;

P : Relative fire damage in case that a fire occurs inside "i" unit space.

Cj

i

: Fire performance of partition components

between "j " unit space and "i" unit space

Pj *Ch

i

: Relative fire damage of "i" unit space

suffering from the fire spread beyond

adjoining "j " unit apace when a fire

starts in " j" unit space.

© p i

( Network )

191
3.3 Relative fire damage In unit space ( P )

3.3.1 Definition of unit space

Unit space is basically a compartment room or space enclosed with fire rated

walls and fire rated doors. Usually, each unit space has its own occupancy.

However, in case that some adjoining rooms are divided with non- rated walls

and openings which do not work for limiting fire spread between those rooms

,

we should consider a series of those rooms as a single unit space which might

have several different occupancies.. ( See the explanatory illustration below )

Fire rated

wall

Non- rated

wall

3.3.2 Estimation method of relative fire damage in unit space

Relative fire damage in unit space ( P, ) is estimated by using following

expression [ 3-2 ]

.

In this expression, "P0ik " most: important basis for estimation which

stands for the potential relative fire risk per unit floor area in terms of

both outbreak and internal spread of fire in each occupancy space (k) of each

unit space (i). This value was determined by expert judgment based on fire

statistics data .. ( See Tab . 1 on page 6 )

"a ik

" is a modification coefficient for "P0ik ", which reflects the condition of

fire protection measures in each occupancy space (k) of each unit space (i).

"a" works to reduce the value of "PQ

" according to the condition of fire

protection equipment in case that the unit/occupancy space concerned is

installed with automatic sprinklers, automatic alarm system, or wall hydrant.

The value of "a” is also given by expert judgment. ( See Tab . 2 on page 7 )

Pi - S ( P0lk * alk * Alk ) [ 3-2 ]

•

k = 1

where ;

n : The number of occupancies involved in "i" unit space

P 0ik : Basic relative value of fire risk per unit floor area for given "k"

occupancy in "i" unit space ( /m2

) a ik : Modification coefficient in terms of fire protection equipment for

"k" occupancy space in "i" unit space

A, k

: The floor area of "k" occupancy space in "i" unit space ( m

2

)

"A" unit space ( Occupancy = corridor )

"B"unit space

( Occupancy

= Office )

( Occupancy 1 = Office )

unit space

( occupancy 2 = Meeting

room )

>

( Occupancy 3 = Storage )

192
3.4 Fire performance of partition components between unit spaces ( C )

Fire performance of partition components ( i ) stands for the relative

intensity of impact by the fire spread from j

unit space to "i" unit space;

that is, the fire performance of walls and/or doors connecting unit spaces in

terms of potential fire spread risk. ( See Tab. 3 on page 7 )

In some case, "C^" is not equivalent to "C^" because there are such walls

that have non-combustible finish on one side and combustible finish on the

other side.

"C" does not mean the substantial fire performance of partition components,

but means the relative index of it. The fire performance of partition

components between a couple of particular unit spaces is represented by the

maximum value ( the value of the weakest part ) among those "C" of walls

and/or doors between these two unit spaces

.

The value of "C" shown in Tab . 3 was determined by expert judgment.

3.5 Determination of potential fire risk in each floor

As described before, potential fire damage in each unit space ( D ) is

evaluated by using "P" and "C" in the expression [ 3-1 ]. Therefore, in each

floor, there are the same number of "D"s as the number of unit spaces which

the floor is divided into. Among those "D i

11 ( i=l,2,...,n ), the max[ D i ] is adopted for the representative estimate score [ S ] of potential fire risk

in the floor concerned.

The grading score of "S" is given a value of "1" through "10", using the table

below which shows the corresponding score for the range of "D"

.

The lower

score indicates a safer case.

Range of D i Corresponding grading score ( S )

0

31

61

30

60

90

1

2

3

91 -- 140

141 -- 190

191 -- 240

4

56

241 -- 310

311 -- 380

381 -- 450

451 --

7

89 10

193
Tab. 1 Basic relative value of fire risk by occupancy of an unit/occupancy

space in multi -occupancy building ( PQ )

Relative

of fire

value

risk

1

1

1

1

Occupancy / Use of an unit space in a building

t

1

1.00

1

1

1 Industrial occupancies treating fires, and/or hazardous

materials such as laboratory, boiler room etc.

1 0.90 1 Parking garage

1 0.85 1 Kitchen in restaurants

Higher

0.75 1 «-•

1

Stores in the presence of sizable fire load such as clothes store, book store etc.

0.70 1 Office room in the presence of sizable fire load

0.65 1 *"• Residential unit

0.60 1

1

Stores well arranged and/or not in the presence of

sizable fire load

Lower 0.55 1 4“ # School room, Guest seat space of theater, playhouse etc

0.50 «-• Sick room / consulting room in health care occupancy

0.40 1 *

1 Office room well arranged and/or not in the presence of

sizable fire load

1 0.30 Bathroom etc.

1 0.10 1

1

1

Common use space such as corridors, rest room, entrance

hall and elevator hall etc.

*1 Relative fire risk values in this table were determined by expert

judgment based on fire statistics. This value stands for the

potential fire risk per unit floor area in terms of both outbreak

and internal fire spread in an unit/occupancy space characterized

by given particular occupancy/use

.

*2 Relative fire risk value in this table is denoted only for typical

conditions of main occupancies. Therefore, for other occupancies

or uses, the relative value needs to be determined by comparing its

actual conditions with those of main occupancies shown in the table.

194
Tab. 2 Modification coefficient in terms of fire protection equipments

installed in an unit/occupancy space ( a )

Modification coefficient

( a ) The condition of fire protection equipment

installed in an unit/occupancy space (*1)

1.0

0.5

0.1

-*-• With no sprinklers

«-• With sprinklers not so well maintained (*2)

*-• With sprinklers well maintained

1.0

0.8

0.6

«-• With neither automatic alarm nor wall hydrant

*-• With automatic alarm but no wall hydrant

<-• With both automatic alarm and wall hydrant

*1 As for determination of "a" in a particular unit/occupancy space, the

minimum value of "a" is adopted among those categories that come under

the particular condition of fire protection equipments installed in

the unit/occupancy space concerned.

*2 If there is some obstruction of water discharge and/or the presence of

non-covered area by sprinklers, it is considered the case as not so well maintained.

Tab. 3 Relative value of fire performance of partition components

in terms of fire spread risk beyond an unit space ( C )

Fire performance of partition components | C

Walls ; | 1. Fire resistant wall ranked as "A" ( 1 hour or more ) | 0.0

2. Fire resistant wall ranked as "B" ( 30 min.- 1 hour ) | 0.3

3. Fire retardant wall ranked as "A" | 0.6

( Non-combustible materials in both backing and finish )

|

4. Fire retardant wall ranked as "B" | 0.8

( Non-combustible materials only in finish ) |

Doors ; | 1. Self-closing fire door ranked as "A" | 0.1

2. Automatic closing fire door ranked as "A" (smoke & heat)'

|

0.2

3

.

Elevator door | 0.3

4. Self-closing fire door ranked as "B" | 0.5

5. Automatic closing fire door ranked as "B" (smoke & heat) | 0.6

6. Non-combustible door but not deemed as fire door I 0.8

* Ranking definition of "A" and "B" in this table is derived from

Building Standard Law in Japan.

195
4. Evaluation Method of Evacuation Performance in Each Floor ( E ) 4.1 Basic idea

Evacuation performance in each floor concerned is evaluated by comparing

the "time" needed for evacuation with the "criterion time" created in

advance. Now, the "time" needed for evacuation can be divided into

following two parts as ;

T0

: Time needed to alert occupants and initiate their evacuation.

This means the lapse of time from fire initiation or fire

detection until starting evacuation of occupants. So, we call

it "T0

" or "evacuation starting time" hereafter. T0 should be

taken into account for an evacuation estimate, because in some fires a lot of fatalities are caused by delay in alerting

occupants. T0 is calculated in the slowest case and is used for

the whole building concerned.

Tx

: Time needed for evacuation action itself. This means the lapse of

time from starting evacuation of occupants on the floor concerned

until completion by reaching a safe place. So, we call it "T 1 " or "evacuation action time" hereafter. And, T x is calculated in each

floor .

.

V

Ti

o V

T 1 V

| Evacuation starting time | Evacuation action time

1 1 1 Fire initiation Starting of Completion of

or Fire detection evacuation evacuation

For evaluation of evacuation performance, firstly we estimate TQ and T., separately in each floor by using estimation methods mentioned later.

Second , we make a grading score corresponding to evacuation performance of

each floor, comparing T0 and T, respectively with each criterion time which

we call "CT0

" and "CT 1 " hereafter.

The advantage of this method is that we can diagnose where the problem is in

each floor, i.e., whether the problem exists in evacuation starting time or

in evacuation action time. Also, we can estimate the sum of T0 and T x

.

4.2 Estimation method of "Evacuation starting time" ( T0 )

Evacuation starting time ( T0 ) consists of the following four parts of the

process for confirming a fire through alerting occupants about the fire as mentioned below. So, T0 can be estimated by summing up T0

x

, T0

2

, T Q3 and

T0

4

like an expression [ 4-1 ].

T0 = T0 ]_ + T0 2 + T0 3 + T0 4

[ 4-1 ]

( sec )

196
where ;

T0

x

: Time needed for detection of a fire incident. The length of this

time depends on the condition of fire detection and alarm system

in a building concerned. So, we adopted the appropriate value

for T0

x

corresponding to that condition by expert judgement.

TQ2 : Time needed for confirmation of a fire. In the usual case,

personnel who work at a security office in the building go to the

origin of an incident and confirm it as a fire.

T0

2

is estimated by calculating both the horizontal and vertical

travel time from the security office room to the fire origin at

the assumed furthest location.

T0

3

: Time needed to give evacuation directions to occupants from the

security office. This is considered to depend on the condition

of the emergency communication system like emergency telephone

or speaker system in the building. The appropriate value is given

for T0

3

corresponding to that condition by expert judgement.

T0

4

: Time needed for occupants to prepare for starting evacuation

after they get notice of fire incident. The appropriate value is

also given for T0 4 corresponding to the type of building by

expert judgement, because characteristics of occupants are different by the type of building.

Each part of evacuation starting time above -pientioned is estimated as follows ;

1) T01 * Fire detection and alarm system is fully equipped .... T0

l

= 0 * Else •

. T0 1

- 2/ A0 2 ^ ^0 2 * Emergency elevator is equipped .... T0

2

= L1 / +( h/ + 30 )+ 2*L2 /.

V 1 V « V

* Equipped -with both emergency telephone system

and emergency speaker system T0

3

= 30 * Equipped with emergency speaker system only T0

3

=60

* Equipped with emergency telephone system only ....... T0 3 =90

* Else TQ3 =120

4 ) t04 * If the building type is one of accommodation types

like hotels , dormitories and etc . . T0 4 =60

* Else •. T04 = 0

Footnotes : "o

L1 ’ L2 h

V,

V„

Floor area of the floor of fire origin (m2

) Horizontal travel distance (m)

Height from the floor of security office to the

floor of fire origin (m)

Horizontal rapid walk velocity ( 2.0 m/sec ) Velocity of emergency elevator ( 2.5 m/sec ) Vertical velocity of man ( 0.25 m/sec )

197
4.3 Estimation method of "Evacuation action time" ( Tx ) Evacuation action time ( T 1 ) in each floor is estimated by calculating the

lapse of time needed for all occupants on the floor to reach a safe place.

As for determination of a safe place, there are the following three cases according to the condition of fire stairs in the building concerned. stairs”* First . if the building has "special fire * 1 with smokeproof enclosure

or outside fire stairs, the fire stairs can be deemed as a safe place.

Second , if the building does not have such special fire stairs but has

"simple fire stairs"* 2 with self-closing fire door or automatic closing fire

door, a safe place is a street floor which means any floor level accessible

from the street or from outside the building at ground level.

Thirdly . if the building does not have any of such fire stairs as mentioned

above, a safe place should be outside of a street floor.

Therefore, estimation method of T x in each case is different according to

the travel distance of evacuation as follows ;

1) The first case ( The safe place is fire stairs on same floor ) The most dominant part of horizontal evacuation to the fire stairs on the same floor is considered to be the time needed for all occupants to pass

the exit which leads to the fire stairs. So, T x in each floor is

calculated by using expression [ 4-2 ]

.

- »/( B * 1.5 )

' 4 - 2 1

2) The second case ( The safe place is a street floor ) Tj_ is estimated by the sum of the time needed for passing the exit to fire

stairs described in the first case and the vertical travel time through

fire stairs to reach a street floor. So, the calculation method of T, is

expressed as follows.

T l ‘ % B * 1.5 ) + H/

V, t 4 ' 3 1

3) The third case ( The safe place is outside of a street floor ) In this case, horizontal travel time to reach outside of the building

from fire stairs at the street floor, is added to T x in the second case.

Tl - N

/( B * 1.5 ) + H/V3 + L>\ [ 4 ' 4 1

Footnotes ;

*1 : Fire stairs with an attached smokeproof enclosure enclosed in

fire rated construction and with self-closing fire doors

*2 : Fire stairs enclosed in fire rated construction and with selfclosing or automatic closing fire door

N : Occupant load on the floor concerned ( # of persons ) B : Total width of exits to fire stairs on the floor ( m ) 1.5 : Flow coefficient at exit ( # of persons/ m,sec ) H : Height from a street floor to the' floor concerned ( m ) L3

: Horizontal travel distance from fire stairs

at a street floor to outside of the building ( m ) V4

: horizontal walk velocity ( 1.0 m/sec )

198
4.4 Estimation of evacuation performance in each floor

4.4.1

Criterion time for evacuation

Currently, there is neither a well-grounded data base nor standard methods for

the determination of criterion time for evacuation.

However, in the current guide book for building fire safety design in Japan,

there appear recommendations for the maximum allowable criterion time for all

occupants on the floor to evacuate into fire stairs on the same floor level.

This criterion time is now broadly used for building safety design in Japan,

and is expressed by " 8/ Ai " ( Ai

: floor area of "i" floor ), which was given by expert judgment in an authorized committee.

But, this criterion time contains both "evacuation starting time" and

"evacuation action time" in its meaning. Therefore, we decided to take the

half of this criterion time length ( 4/ Ai ) for our criterion time of

"evacuation action time" ( CT, ) , which is considered to depend on size of

each floor area. On the other hand, we should treat the criterion for "evacuation starting

time" somewhat differently, because this criterion time does not seem to

depend on size of each floor area, but should have a common standard time for

all kinds of building. So, at present, we adopted 240 (sec) as the criterion

of "evacuation starting time" ( CT0 ) , which is supposed to be a standard

value from the results of case studies.

4.4.2

Determination of grading score for evacuation performance estimate .

The estimate of evacuation performance in each floor ( E ) is expressed by

using grading score, which is "1" ( the best rank ) through "10" ( the worst

rank ) . This grading score is determined, based on TQ , T L comparing with

CT0 , CT 1

. In practice, we plot a point for each floor's estimate on the coordinate plane

by using "T0 value" on X axis and "T 1 value" on Y axis. And we make a grading

score for each point by zoning on the coordinate plane with lines of criterion

time and total time limitation of T0 and T l as illustrated in Fig. 2 on page 12

By the way, in case that -the sum of T0 and T : is a little bit over the sum of

CTq and CT X

, we should think there is some problem for evacuation on the floor

concerned. However, we consider this case is better than such a case that

one of T0 and T x is far over its criterion time, even though the sum of TQ and

T : does not exceed the sum of CT0 and CT X. Therefore, we make a lower score as for the former case ( score area "5" in

Fig. 2 ) than the latter case ( score area "6" and "7" in Fig. 2 ).

4.4.3

Modification in terms of the conditions of occupants and building

The estimate of evacuation performance ( E ) above-mentioned is determined,

assuming usual occupants who will be capable of evacuation without assistance

under the average conditions of the building configuration.

Therefore, if the characteristics of occupants and the building configuration

in the floor concerned involves one or more of the following cases , the

estimate should be modified by shifting up/down the grading score for each of

199
Che corresponding cases, provided that Che final modified score shall noc be

less Chan "l" and also noc be more chan "10".

1) 10% -30% of occupanC load are considered Co be incapable of

self -movemenc in evacuacion. ( + 1 ) 2) More Chan 30% of occupanC load are considered Co be

incapable of self -movemenc in evacuacion. ( + 2 ) 3) There is one or more obscruccion(s) on a pach of cravel

Co an exic. ( 3 or more curvings / gap of floor level ) ( + 1 ) 4) There is a dead-end hallway which is more chan

20 m long . • * ( +2 ) 5) There is a dead-end hallway which has an exCerior balcony

wich a fixed ladder ac Che dead-end. ( + '1 ) 6) There is an exCerior balcony wich a fixed ladder

accessible direccly from main rooms. (

- 1 ) 7) There is a conCinuous exCerior balcony moscly all around

Che exCernal walls . (-2)

CT0 ( sec )

Fig. 2 Grading score of evacuation performance

based on zoning by T0 value and Tt value

* Ai

: Size of floor area of "i" floor

( m

2

)

200
FIRE RISK EVALUATION METHOD FOR MULTI -OCCUPANCY BUILDINGS

A. Sekizawa, Fire Research Institute, Japan

APOSTOLAKIS: There must be a lack of uncertainty calculated in these

findings. Are you making any allowances?

SEKIZAWA: Well, I believe there are a lot of certainties in characterization.

In regard to uncertainty, there are two points. One is the calculation itself

does not reflect the real accuracy, and, secondly, there are many variations

of the opinions. Therefore, different people may have different results. I can report to you that all 11 of the largest cities in Japan, under the

direction of the Fire Agency, adopted this result. Very desirable results

were expected out of this study. Also, this method is better understood

today, so there is not much difference between now and then.

This method is quite different from the scientifically based prediction model,

so it has a lot of uncertainty, so we need verification based on a lot of case studies in the future.

APOSTOLAKIS: But if you know the method that will make the allowance for

these uncertainties, have you tried to be conservative, or have you tried to

be realistic?

SEKIZAWA: What do you mean allowance?

APOSTOLAKIS: These time studies that you were estimating for the

investigative damage potential seems to me not to be worthwhile. The number

that comes out must be very uncertain, so whatever you do with that number

later, must somehow leave no uncertainties. Are you trying to get a conservative number or what? That's what I mean by making an allowance

for the uncertainty.

HALL: You mentioned that verification will be done using case studies. How

will that work? What will you measure to confirm the accuracy of the method?

SEKIZAWA: It's a very hard question, but we have to build a database. Then

this method will be used because we cannot verify this method directly right

now.

HALL: So, would you be using some indirect plausibility checks?

SEKIZAWA: Yes.

HALL: In other words, if you can't measure the parameters directly, you

simply try to work consequences out of the method that, can be verified? Is

that how it's worked? If you can't measure some of these key parameters

directly, which now we can't then you'd be measuring at some derived

consequence. Is there an example of how that might work? What kind of data

might you collect?

201
SEKIZAWA: The database is derived from this method. The file shows, and

some experts can understand the style of the buildings or if they are dangerous buildings or safer buildings by looking at the distribution, whether

it is at the threshold point, or in the risky range.

HALL: Let me see if I understand you. You're saying experts can look and try

to say these are safe buildings, or these are dangerous buildings? Does the

method also classify them in that way?

SEKIZAWA: Yes.

202
AN EVACUATION MODEL FOR THE USE

IN FIRE SAFETY DESIGNING OF BUILDINGS

by

Kiyoshi TAKAHASHI

Fujita Corporation

4-6-15 Sendagaya, Sibuya-ku, Tokyo 151

and

Takeyoshi TANAKA

Building Research Institute

Ministry of Construction

Oho-machi, Tukuba-gun, Ibaraki 305

SYNOPSIS

The theories and the principal features of the computer model developed

in a research project of Ministry of Construction: "Development of Fire

Safety Design Method of buildings" as a tool to predict and evaluate

evacuation time of people in fire are outlined. Also, the predicted results

using the model are compared with the field measurements of the egress time

of the audience of Tukuba Expo' 85 pavilions and of the evacuation drill in a high-rise building in Tokyo.

INTRODUCTION

To date, fire safety measures of buildings in Japan have been determined

almost always according to the specification standards of The Building

Standard Low of Japan. However, it is necessary to replace these

specification standards with performance ones as most as possible for the

rationalization both in fire safety performance and cost of the measures and

for more flexibility of design of architectures. In recognition of this

necessity. Building Research Institute (BRI) have carried out a research

project "Development of Fire Safety Design Method of Building" with the

cooperation of universities, private cons true tion^companies and designers.

To introduce performance standards into a fire safety design method,

many kinds of prediction methods are indispensable for evaluating various

aspects of fire behaviors. And since a prediction method of fire behavior is one of the most important ones, an attempt was made to develop a computer

evacuation model. This paper outlines the theories of the model and also

demonstrates its predicting capability by means of the comparisons with the

field measurements conducted during Tukuba Int'l Expo *85 and an evacuation

drill at a high-rise building in Shinjuku, Tokyo.

EVACUATION MODEL

1. Scope of The Model

Many evacuation models have been developed by now to evaluate human

behaviors in fire. And one of the directions of the developments is to try

203
to predict or explain human behaviors in really hazardous conditions in fire

by means of introducing models concerning human reactions to fire or smoke.

Thos’e models may be interesting as a basic tool or an attempt to interpret

human behaviors in fire, but it will be too ambitious to use this type of

models for fire safety design purposes. The reasons are: human behavior when

encountering a fire will vary from one individual to another, there is

virtually no reasonable means to validate human reaction models to fire

conditions at present, and it will require too much time or too many runs until the fire safety design of a building is decided because this type of

models will be sensitive to fire conditions, so give different answers under

different fire scenarios, including location of the room of origin, behavior

of smoke movement

.

On the other hand, if the scope of the application of a model is

restricted to only design purposes, the model does not necessarily have to be

very sophisticated. Let's take an example: suppose that the kinetic

characteristics of a structural material are known only for the elastic

domain, then it is possible to design a structure of a building so that the

stresses of the structural elements do not exceed the limit of its elasticity

thus avoiding the risk of collapse due to the uncertain knowledge in its

plastic domain. Similarly, in an evacuation design, if prediction of

evacuation is possible only when the level of hazard to evacuees is not too

high, a building can be so designed, by means of compartmentation , smoke

control and so on, that the necessary level of safety can be attained during

the period of evacuation.

The present model was developed according to the consideration briefed

in the above and adopts the basic premises as follows:

a) The necessary level of safety of evacuation route is maintained

throughout the period of evacuation so that orderly evacuation can be carried

out.

b) The evacuation plan is adequate enough to be carried out in actual fire

situations as is planned.

2. Modeling of Building Spaces

3uilding spaces are modeled as an evacuation space system composed of

six kinds of space elements i)-vi) and two imaginary spaces vii) and viii) as

is shown in Fig.l. The latter two are introduced to make the concept of the

calculation procedure clear.

i) room : A space with exits in which occupants exist at the onset of fire

and make egress movement. A room with multiple exits is further

divided into room units.

ii) path : A space on a floor with an entrance and an exit in which

evacuees do not exist at first but enters from another space element

and move toward the exit. A path is further divided into path units

for convenience of calculation.

iii) stairs : A space which connects floors at different levels and has

multiple entrances and exits. A stair functions like paths joined one after another.

iv) vestibule : A space with an entrance and an exit to a stair. A

vestibule function like a path. The number of evacuees entering a vestibule is counted to judge if the evacuation on the floor has been

completed

.

v) hall : A space with multiple entrances and exits which are connected by

multiple paths.

vi) refuge : A final destination of evacuees.

204
vii) link : An imaginary space without capacity which regulates the flow of evacuees at a joint part of space elements,

viii) crowding : An imaginary overflow space at the exit of each space

element, which accommodates evacuees who cannot make immediate egress

to the next space.

Fig.l An Example of Modeling of Building Spaces

3. Modeling of Evacuees' Movements

1) Evacuees

It is assumed that the evacuees are homogeneous in evacuation ability

and move like a fluid.

2) Initial distribution of evacuees

it is assumed that the evacuees are uniformly distributed in rooms.

3) Starting time of evacuation

In actual fires, it is not unusual that the evacuation starts

significantly after the onset of fire, and this delay time can never be

neglected in the total evacuation time. It is not easy to assess the

starting time of evacuation because this depends on various conditions such

as nature of occupants, fire detection and alarming systems, guiding method

of evacuees and building spaces, but this model treats the time as input

data.

4) Movements of evacuees in the space elements

i) Movement in rooms

As was stated before, when a room has multiple exits, the room is

further divided into the room units each of which corresponds to an exit as

is shown in Fig. 2.

i — r

> A ? (l) (2)

* /

1 \ . \ / A \ ! O +

\ ‘ y p*. \

i Fig. 2 Division of a room into room units

The evacuees in a room unit, which are distributed uniformly at the initial

point of time, starts to move toward the exit simultaneously. Two types of course of approach to an exit are considered as follows:

a-1) L- shape approach to exit

In such rooms as large offices, where many office furniture usually

exist, it will be difficult for evacuees to approach to an exit directly.

Hence, in such a case, it is assumed that evacuees take L-shape course to

205
approach to an exit.

Let's consider first the case where the exit locates at a comer of a room. It will be known that the evacuees who arrive at the exit from the

start by a certain elapsed time t are those who were initially within the

L-shape distance v*t, in other words, those in the shaded area in Fig. 3.

-a+b-v-t—

f

Fig. 3 The Number of Evacuees Arriving at the exit Within Time t

In mathematical form, the number is given as follows:

P =

( vt r

2

, a 2

P

where

1 2

{ ab

ab

tv

p

P

( 0 ^ c ^ ) + a ( v t — a )

}

( a+ b — vt Y

} <7< tS

( —< t £ V

V

b_v

a + b

)

( 1 )

the length of shorter side of the room (m)

the length of longer side of the room (m) a certain elapsed time from the start (sec)

walk speed (m/sec)

density of evacuees (person/m2

) number of evacuees arriving at the exit by time t (person)

Incidentally, if ever evacuees do not take a literal L-shape course with

one turn but take a course with several turns, it does not make difference in

the distance to an exit as long as they do not go backward.

a-2) Centripetal approach to exits

In rooms with little obstacles for walk, such as assembly rooms without

fixed seats, party rooms and gymnasiums, it will be reasonable to assume that

evacuees make centripetal approach to the exit. In this case, referring to

Fig. 4, it can be seen that the number of evacuees arriving at an exit is

given as follows:

ff(vt) 2

( 0 ^ t )

P = {-

( avt ) SIN 9 ,

( avt ) SIN 0,

+ TT ( vt )

+ n ( vt ) 2

(T -9,)

2 n

(-f- -0, -

} (4-<ts-4-) (2)

2 Tt

9i

)

^

( b vt ) S IN d7

9

= COS H

(“77") 9 2 = C0S H

(“7“ < t

206
where a : the length of shorter side of the room (m)

b : the length of longer side of the room (m)

t : a certain elapsed time from the start (sec) v : walk speed (m/sec)

p : density of evacuees (person/m2 ) P : number of evacuees arriving at the exit by time t (person)

*

Fig.

4

In general, the exit of a room unit does not necessarily locate at a corner, but as can be seen in Fig. 5, a room unit can be regarded as a combination of the two parts with an exit at a corner. So, the number of the evacuees arriving at the exit can be calculated as the sum of the numbers of

the evacuees of the two parts.

o

b) Movement between crowdings

Since evacuees try to escape from fire hazards as fast as possible, it

will be reasonable to assume that in a room with multiple exits, they choose

the exit through which they are likely to be able to make fastest egress from

the room. As the result, the numbers of the crowdings at the exits of a room

are redistributed to minimize the expected time of the dissolution of all the

crowdings of the exits. Because the expected minimum time for -'the

dissolution of all the crowdings is

Tmin = ZCr(i)

EPout(i)

when the egress at each exit continues at the

redistribution of each crowding is given as rate of Pout(i), the optimal

C

fr(i) = Pout ( i) * Tmin

where Pout(i) : egress rate at exit i (person/sec)

Cr(i) : number of crowding at exit i before redistribution (person)

Cr(i) : number of crowding at exit i after redistribution (person)

Note that Pout(i) may change with time, as the condition of the space outside

exit i changes.

207
ii) Movement in paths, stairs and vestibules

In this model, evacuees entering a space element are assumed to proceed

orderly toward the exit without passing over and going backward. Since

paths, stairs and vestibules are divided into the units with the length that can be cleared at the time increment, that is v*Jt, it follows that the

evacuees in a space unit move to the forward unit at every time step.

the direction of movement

time t n n1

5 4 3 2 i next space

\

N\

N\

N\

N\

N\

time t+dt n n -

1

5 4 3 2 1 next space

| v« A tl v A t(

Fig. 6

iii) Movement in halls

In a hall with multiple exits, the evacuees entering through an entrance

are assumed to choose the exit which allows the shortest egress time from the

hall taking into account the distances to the exits, the amount of crowdings

and the rates of egress at the exits,

a) Exit selection in an empty hall

Let's consider first the case where the evacuees enters in an empty hall

where a crowding does not exist at any exit. Let P be the number of evacuees entering through an entrance during the period of time t. If the numbers of

evacuees are allotted optimally so that the egress time from the hall becomes

minimum, the egress times at all the exits must become the same, that is

„ , , . PI „ . P2 ^ , Pi Tmin =tl + -Ir = c2+-5T =~-=ti+— = •• ••

where ti : time to arrive at exit i (sec)

Ni : egress flow capacity of exit i (person/sec)

Pi : number of evacuees allotted to exit i (person)

i : exit number in order of distance

( 3 )

"rom Eqn.(3), we obtain

£ ( Niti + Pi ) P + £( Niti ) Tmin = " - " ~ _ £ Ni £ Ni

So, it follows that the optimal allotment of evacuees is given as Pi = Ni (Train - ti ) (i=l, 2. 3,”~,n)

( 4 )

( 5 ) where n is the number of exit. However, this problem is not so simple, that

is, the optimal solution is not necessarily given when all the exits are

used. Let's take the simplest example, if all the evacuees choose the

nearest exit, the egress time will be

T = t i + ( 6 )

So, if T < 1 2 » the evacuees can go out of the hall faster when they choose

only the nearest exit rather than when they choose more.

208
Fig. 7 Exit selection in am empty hall

After all, it follows that the optimal time Tmin and allotment Pi have to

be calculated according to the procedure described in Fig. 8a.

Fig. 8a

b) Exit selection in crowded halls

As the time elapses, a hall will be crowded with evacuees and crowdings

may be formed at exits. The evacuees entering a hall at such conditions are assumed to pick up the exit so as to minimize the egress time taking into

account of the crowdings at the exits and the numbers of the evacuees

approaching to the exits as well as the distances and the egress rates. In

this case, since the expected egress times of the evacuees entering the hall

during a time increment dt are the same for all the exits chosen if the

allotment is optimal, the following equation holds:

C'1+PI (T2 + P2 C*l+Pi (7) Tmin = tld jT[

" — t2H ^

— ti + "

209
where Ri is the rate of egress at exit i at that time, and Ci is given as

( Ci+Wi- RiCi ( Ci+Wi > Riti)

c i — < ( 8 )

( 0 (Ci+Wi ^Riti)

where Ci : number of the evacuees in the crowding at exit i (person)

Wi : total number of evacuees arriving at exit i from the entrance

within time ti (person)

Ri : rate of egress at exit i (person/sec)

From Eqn.(3), Tmin is given as follows

P + E (C'i + Riti)

Tmin = ( 9 ) E Ri

and optimal allotment is given as

Pi= Ri { Tmin - ( ti + -|^)} (10)

However, exits have to be examined in order of the magnitude of ti+ Ci/Ri

from the smallest to larger as is shown in Fig. 8b. Note that the procedure

to yield Tmin described in Fig. 8b is general and include the case of empty

halls.

Fig. 8b

210
c) Movement from an entrance to an exit

In this model, each entrance of a hall is connected with every exit with

imaginary paths, so the evacuees entering the hall through the entrances move

toward the exits just as they move in a path.

iv) Movement through a link

The movement of evacuees from a space element to another is treated by

the concept of movement through a link. Since the number of evacuees which

move through interface of space elements cannot be simply determined because

it depends on the number of the evacuees ready to move, the available

capacity of the space to move into, the widths of the openings and the number

of space elements combined. Links are • introduced to deal with this

complex movement around the joint parts of the spaces,

a) Total number of evacuees moving through a link

Let's consider a general case where multiple space elements from which

evacuees come out are combined with multiple space elements to which they

move into. A link is imagined between the farmer and the latter as is shown

in Fig. 9. Suppose that during a time step Jt, movement of evacuees took

(i) IlK

•Ni *

link N'jm« | mi A'jf

(j)

Fig. 9 A link combining multiple space elements

place in each space element and evacuees have been added to the crowdings of

the space elements in the upstream of the link and empty spaces have appeared

in the downstream spaces. In this case, the number of evacuees which can be

allowed in a downstream spaces is determined either by the available space or flow capacity of the entrance, that is

Pmax.j = MIN ( N*j jt , PmaxA'j ) ( 11 ) where N*j : flow capacity of entrance of space j (person/sec)

A'j : space available In path j (person/ sec)

P max : maximum density of evacuees (person/m2

)

211
So, the total number of evacuees allowed in the downstream spaces E Pmax,j is

E Pmax , j =E MIN ( N'j At , o maxA'j ) ( 12 )

On the other hand, since the maximum number of evacuees which can egress from an upstream space element is determined either by the number of evacuees waiting to egress, or the egress capacity of the exit,

Pmax,i = MIN(Nidt,Ci) (13)

where Ci is the number of evacuees in the crowding of space i, and the total

for all the upstream space elements E Pmax,i is

EPmax.i =E MIN ( Nidt ,Ci)

( U ) Finally, it follows that actual number of evacuees PL which move through the

link is given as

:

PL = MIN {EMIN( N i dt , C i ) , E MIN ( N'j At . p maxA'j )} ( 15 ) b) Allotment of evacuees to each path

Now that the total number of the evacuees moving from one side to the

other, the number allotted to each space element, in other words, what

portion of the total number comes out of a certain space element i in the

upstream spaces and what portion goes into a certain space element j in the

downstream space element, has to be determined. Suppose that they have been

determined, the number of evacuees coming out from space element itPout(i),

and the number going into space element j:Pin(j) are expressed as follows:

Pout(i)=/3i PL and Pin(j) = (A'

j

PL ( 16 ) where and j are the ratio of the evacuees allotted to upstream space

element i and downstream space element j respectively. Needless to say,

E /3i = 1 and E /3'j = 1 . And here, 0± and /3'j are assumed to be

proportional to Pmax,i and Pfoax.j respectively. Hence,

Pmax ,

i

/3 i= —

E Pmax

,

i

Pm a x ,

j

E Pmax ,

j

MIN ( N i J t ,Ci)

E MIN ( Ni At . Ci)

MIN( N'j At, P maxA'j ) E MIN ( N'j At, P maxA'j )

( 17 )

COMPARISONS BETWEEN FIELD MEASUREMENTS AND PREDICTION

1. Egress of Audience from Tukuba Expo '85 Pavilions

1,1 Field measurement of the egress of the audience from Tukuba Expo' 85

Pavilions

Field measurements were made for the egress behavior of the audience of seven pavilions of Tukuba International Exposition held in 1985. The plans

and the sections of the seating area of the selected pavilions are shown in

212
Fig. 10. These theaters were vacated after each films or performance for next

turn.

The positions of VTR. cameras for the measurements are shown also in

Fig. 10. Several turns of measurements were made in each theater and the

results of the time of the egress completions are given in Table 1. The

typical tendency of. egress is shown by the example shown in Fig. 11. The

number of audience remaining in the theater counted from the video recordings

is plotted as a function of time but the counting was only possible for the

initial and the final regions of egress. For the primary region in between,

it 'is too difficult to count the number of audience, but a simple

interpolation as shown in Fig. 11 is thought to be sufficient to estimate the

number.

In the final region, a tailing appeared in every turn due to a part of

the audience egressed leisurely and the time of egress completion was lengthen by the tailing. In a truely emergent condition, this tailing would

disappear. If the tailing is cut off by extrapolating the slope at primary

region as shown in Fig. 11, the egress time is shortened by 10~30sec. These

results are given in Table 1.

1.2 Prediction of egress time

The egress time of the audience from the theaters were predicted using

the evacuation model described. in the above. The two cases were considered

for each theater, i.e: case(a) in which simply L-shape approach to exits is

assumed in the theater space, and case(b) in which the theater is composed of

space units such as rooms and paths. In the latter, the walk speeds at

horizontal paths and stairs are taken as 1.0 m/s and 0.5 m/s, respectively,

and in the former, 0.7 m/s was employed as the average walk speed considering

that the space consists of horizontal paths and stairs. In both cases, the

egress coefficients of exits on horizontal floor and on stairs, are taken as

1.5 person/ (m*s) and 1.3 person/ (m« s) , respectively.

The results of the predictions are given in Table 1. As can be seen,

the predicted egress times of case(b) are consistently longer than case(a)

.

This means that in the prediction for case(b), crowdings are formed in some parts within seating area of the theater. In fact, crowdings were observed

along the alleys in many of the theaters in the field measurement.

1.3 Comparison between the predictions and the measurements

It can be seen in Table 1 that the predicted egress times are generally

shorter than the measured ones and the predictions assuming case(b) give the

better agreement with the measurement except FR. pavilion. So, it may be said

that it is more appropriate to model a theater space as a combination of

space units than to simply consider a room with L-shape approach in most of

the theaters, in which crowdings are usually formed along alleys. On the

other hand, in FR. pavilion, since the seats were low bench type ones so that

people can more freely over the seats, the calculation assuming case (a) gave

a better prediction.

There are some other reasons why the measured egress times are longer

than the predictions. Needless to say, the theaters are not in emergent

conditions, so the people egressed more or less leisurely, in every case. In case of SH and TD pavilions, it was rainy on the day of measurements, so the

people stopped outside the exits to prepare rain outfits and delayed the

egress of the remaining people. In NT pavilion, people on the upper layer

213
went down through the sajsa stairs besides the theater and merged with the

people from the theater.

Taking these factors into consideration, the present model can be said

reasonable as a prediction tool to evaluate evacuation time in fire safety

design of buildings.

2. Evacuation drill in a high-rise building

2.1 Field measurement of the evacuation drill

A field measurement for egress behavior was made taking advantage of an evacuation drill carried out in a 50 story high-rise building in 1985. All

the participants of the drill started the evacuation movements by the order

delivered through loud-speaker simultaneously throughout the building and

went down to stairs to outside the building. The outline of the building and

the number of the drill participants on each floor who used the east

stairwell are shown in Fig. 12 and Table 2, respectively. The measurements

were made on the east stairwell using the VTR cameras placed on a couple of

points in the stairs and by the portable VTR cameras carried by the observers

who went down the stairs with the drill participants. The evacuation took

960 sec to complete.

2.2 Prediction and comparison

The prediction were made for the east stairwell. The initial number of

participants on each floor is shown in Table 2, but the number in ( ) were estimated from the total number of the participants and the results of

measurements since they were not exactly known.

As the speed of movement for the horizontal paths and for the stair, 1.2

m/s and 0.6 m/s (horizontal component of velocity) are employed respectively

and as the egress coefficients through doorways, 1.5 person/ (m * s) and 1.3

person/ (m • s) are used for the horizontal paths and for the stairs,

respectively.

The time to complete evacuation was predicted as 885 sec. The predicted

time is shorter than the measured by 75 sec. This may be partly because a certain time was used for the preparation, gathering or roll calls by

leaders. However, even if these factors are not taken into account 75sec of

difference will be evaluated well within the allowable limit of accuracy

compared with the evacuation time.

CONCLUDING REMARKS

An evacuation model has been developed to evaluate evacuation time,

which is one of the important components in performance based Fire Safety

Design Method of Buildings. The capability for prediction of the model were examined in comparison of the predictions with the field measurements of the

egress behaviors in Expo '85 pavilions and a high-rise building. As is easily

understood, a perfect prediction of people behaviors cannot be expected,

however, the present evacuation model can be evaluated as reasonable as a tool to be used for design purpose.

214
Table 1

Results

of

measurements

and

predictions

of

egress

completion

time

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215

front

*1:

Tape

start

delayed

by

back

*2:

Opening

of

exit

door6

right
u

X— <u -a a ns U aj CQ U 4)

G M»m w aj O

a <-* w as c

« *H c O O MO « C V -n G U

<q «-» « a

k U 4) H 4) 0) M U OJ Us U

C X ** O CJ a <u <** & a *q

< a u a u t

cn

3300

eo

*H

Cfa
Fig. 11 Typical tendency of egress from theaters

section ssrSa i : 3.000

stairwell

Fig 0 12 Outline of the high-rise building

Table 2 The number of participants to

the evacuation drill who used

the east stairwell

F (person) F (person) F (person)

l — 13 — 35 —

2 — 19 — 38 (2 1) 3 —

20 4 S 37 —

4 — 21 — 33 (2 1) S —

22 4 S 39 —

8 — 23 — 40 —

7 1 8 24 —

41 —

8 — 25 7 0 42 (2 1) 9 5 0 28 4 0 43 —

10 — 27 — 44 (2 1)

11 5 0 23 7 0 45 —

12 — 29 — 46 (2 1) 13 5 0 20 (2 1) 47 —

14 — 31 — 43 (2 2)

IS s 0 32 (2 1) 49 —

18 — 33 — 50 (1 5)

17 — 34 (2 l) If 8 9 1 217
AN EVACUATION MODEL FOR THE USE IN FIRE SAFETY DESIGNING OF BUILDINGS

K. Takahashi, Japanese Association of Fire Science and Engineering,

Japan

NELSON: You took a lot of data with TV cameras. Are you familiar with the

assumptions on flow versus density that Predeshinsky and Fruin and Pauls are using? Did your data confirm or challenge those approaches?

First, are you familiar with the work that I'm referring to?

TAKAHASHI: I'm familiar with the data from Predeshinsky.

NELSON: His theory is that the speed of the evacuees is a function of the

density of the population. Did your observations confirm this, or were they

not appropriate?

TAKAHASHI: I haven't really concluded anything about the evacuation speed.

NELSON: Are you familiar with the effective width theory?

TAKAHASHI: Yes, I'm very familiar with it.

NELSON: Did you observe it in your videos? Did you observe people staying

away from the edges in your videos?

TAKAHASHI: I haven't really quite observed that to reach any kind of

conclusion from your viewpoint; however, we focused on the time to completion.

I would like to explore speed and density in the future.

HALL: Your method seems to be flexible enough to be adapted for cases where

the individuals do not know the best route out, or where exits are blocked or no longer available due to the fire. Do you have plans to examine cases like

this?

TAKAHASHI: We choose the best exits for this work; however, another path to

the exit is along the evacuation route.

TANAKA: We have taken into consideration during the calculation process the

possibility of the exit being blocked. However, we have not taken into

consideration the possibility that the entrance or exit may be blocked after

the fire starts.

218
A SUMMARY OF THE ASSUMPTIONS

AND LIMITATIONS IN HAZARD I

Richard W. Bukowski P.E., Head

Hazard Analysis

Center for Fire Research

National Bureau of Standards

Gaithersburg, MD. 20899

INTRODUCTION

The first version of the Hazard Assessment Method - HAZARD I consists of a three volume report and a set of 5 1/4 in. disks containing

the software necessary to conduct hazard analyses of products used in

residential occupancies. All of the software provided will operate on any

IBM1 PC (XT or AT) or compatible MS-DOS computer with the following minimum

hardware configuration:

• 512k MEMORY

• GRAPHICS CARD (IBM CGA OR EGA,. OR HERCULES COMPATIBLE)

• HARD DISK DRIVE

• MATH CO-PROCESSOR (8087 OR 80287)

• PRINTER

• DBASEIII+ (REQUIRED ONLY FOR THE DATA BASE PROGRAM)

A block diagram of the important phenonema which must be included in a hazard analysis and their inter-dependencies is given in figure 1. The

organization of the HAZARD I software package is shown in figure 2. It

includes a scenario development guidance utility (PRODUCT . ONE) ; an interactive program for inputing data to the fire model (FINPUT) ; a data

base program (FIREDATA) and files of thermophysical, thermochemical, and

reference toxicity data; the FAST model for multi-compartment energy and

mass transport; a graphics utility for plotting data (FASTPLOT) ; a detector/sprinkler activation model (DETACT) ; an evacuation model which

includes human decision/behavior (EXITT) ; and a tenability model (TENAB)

which evaluates the impact of the predicted exposure of the occupants in

1 The use of company names or trade names within this report is made

only for the purpose of identifying those computer hardware or software

products with which the compatibility of the programs of HAZARD I has been

tested. Such use does not constitute any endorsement of those products by

the National Bureau of Standards.

219
terms of incapacitation or lethality from temperature or toxic gases or incapacitation by second degree burns from radiant flux exposure. A stepby-step procedure for performing a hazard analysis using this software

package is presented in chapter 4, and user guides for each of the programs

is contained in chapter 6.

Also included in the report is the set of eight representative example

cases of typical residential fires which were established by two panels of

outside experts. These consist of fire scenarios, selected by a panel

composed of representatives of the major fire services organizations, in

each of three single - family residences which were considered to represent

typical homes by a panel from the model building codes, NFPA, and the

architectural community. A description of the process of developing these cases and the results obtained for each is included in chapter 5, and the

complete documentation (input data file listings, program outputs and

graphs of selected, predicted variables) is in Volume 2.

While the scope of this first hazard assessment method is limited to

residential occupancies, the work necessary to extend it to other occupancy

classes is underway. This first version is being introduced now to allow

the fire protection community to become familiar with the method and its use. Feedback from users is crucial to refinement of the method and its

ultimate utility.

GENERAL

:

HAZARD I is a first, prototype version of a methodology for

making quantitative predictions of the hazard to the occupants of a building from a fire within that building. It consists of a collection of

data, procedures, and computer programs which are used to s imulate the

important time -dependant phenomena, generally as depicted in figure 1 (although this first version does not account for all of the phenomena

listed). The major functions provided include;

1. the production of energy and mass (smoke and gases) by one or more burning objects in one room, based on small- or large-scale

measurements

,

2. the buoyancy- driven transport of this energy and mass throughout a series of user specified rooms and connections (doors, windows,

cracks , etc . )

,

3. the resulting temperatures, smoke densities, and gas

concentrations after accounting for heat transfer to surfaces and

dilution by mixing with clean air,

4. the evacuation process of a user specified set of occupants

accounting for delays in notification, decision making, behavioral

interactions, and inherent capabilities,

5. and the impact of the exposure of these occupants to the predicted

room environments as they move through the building, in terms of the

expected fatalities, and the time, location, and cause of each.

220
As can be seen from this list, the method involves a mix of the "hard"

sciences (physics, chemistry, fluid mechanics and heat transfer) and the

"soft" sciences (biology, toxicology, and human behavior). Less obvious

are the areas where fundamental laws (conservation of mass, energy, and

momentum) can be used, versus where empirical correlations or even "educated guesses" must be employed to bridge gaps in existing knowledge.

The latter, along with compromises required by considerations of

operational practicality result in the introduction of uncertainties in the

results which must be quantified by experimental verification and

validation. Until such is completed, the user should understand the

inherent assumptions and limitations of the procedures and programs in

order to make estimates of these uncertainties.

SCOPE LIMITATION: The scope of HAZARD I has been limited to one- and twofamily residential structures, as a result of limitations in both technical

knowledge and resources. The technical limitations involve insufficient

understanding or capability to model such large -building phenomena as the

distribution of heat, smoke, and gases through the HVAC system, and the

complex flows in stairways. While these should have a negligeable

influence on hazard development in small buildings, the effects in large

buildings can be great. Also, the behavioral rules first implemented in

the EXITT model for small building evacuation have not yet been

incorporated into the large building evacuation models.

In addition to these technical considerations, available resources required

that the first data base be limited in the number of data entries supplied.

The obvious solution was to concentrate on materials and products typical

in residential occupancies. If one considers that the majority of US fire

losses are in these areas, this does not represent a bad starting point for

the first system; and the scope will be broadened in HAZARD II by

improvements in the technology and expansion of the data base.

PROGRf\MS AND PROCEDURES

Figure 2 presents the HAZARD I software package

developed to implement the assessment of hazard as shown in figure 1. Of

the nine programs shown, only five perform calculations and so have

assumptions associated with them. The others ( PRODUCT . ONE , FIREDATA,

FINPUT, and FASTPLOT) perform utility and user interface functions only.

Of this latter group, the only point which should be made in the context of

assumptions deals with the database files searched by the programs in

FIREDATA. The data provided in the cone and furniture calorimeter files

are measured values from individual samples tested in these devices under a specified set of conditions. While the materials are identified

generically, it should be understood that such data are not necessarily

representative of the behavior of that generic material. Some variation

would be expected, even on a set of samples from the same lot; and no attempt was made to obtain representative samples for test. Also, data in

the thermophysical properties file were taken either from manufacturer's

data or from literature sources with no attempt to verify values or to

determine their representativeness. Finally, the data in the toxicity file

are published values from the sources indicated. Replicate tests were sometimes done, and some of the sources provide confidence intervals for

these data. The material identifications are those provided in the sources

and cannot be verified.

221
SPECIFIED FIRE: The greatest weakness of HAZARD I is in the specification

of the rates of energy and mass release by the burning item(s). The

procedures available to the user and their associated limitations are:

1. For the Furniture Calorimeter, a product (chair, table, bookcase,

etc.) is placed under a large collection hood and ignited by a 50 kW

gas burner (simulating a wastebasket) placed adjacent to the item for

120 sec. The combustion process then proceeds under assumed "freeburning" conditions, and the release rate data are measured.

Potential sources of uncertainty here include measurement errors related to the instruments and the degree to which " free -burning"

conditions are not achieved (radiation from the gases under the hood

or the hood itself, and restrictions in the air entrained by the

object causing locally reduced oxygen concentrations affecting the

combustion chemistry) in the apparatus. There are limited

experimental data for upholstered furniture which suggest that prior

to the onset of flashover, the influence of the compartment on the

burning behavior of the item is small, where it is the only item

burning and it is ignited in a similar fashion. The differences

obtained from the use of different ignition sources or locations of

application have not been explored.

2. Where small-scale calorimeter data are used, procedures are provided to extrapolate to the behavior of a full size item. These

procedures are based on empirical correlations to data which exhibit

significant scatter, thus limiting their accuracy. For example, the

peak heat release rate estimated by the "triangular approximation"

method for upholstered furniture averages 91% (range 46% to 103%) of

the measured value for a group of chairs with non-combustible frames,

but only 63% (range 46% to 83%) for other chairs with combustible

frames. Also, the triangle does not include the "tail" of the curve produced primarily by the burning of the frame after the fabric and

filler are consumed.

3. The data and procedures provided relate directly only to contents

items under "free burning" conditions, initiated by relatively large

flaming sources. Almost no data exist for release rates under

smoldering combustion, nor for the high external flux and low oxygen

conditions post- flashover which strongly influence the combustion

chemistry. While the program MLTFUEL allows multiple items burning

simultaneously to be converted to a single "equivalent" specified

fire, it does not account for the energy interchange of such items.

Thus, for other ignition scenarios, multiple items burning

simultaneously (which exchange energy by radiation and convection)

,

combustible interior finish, and post- flashover conditions, the

procedures provided can only give estimates which are often non- conservative (the actual release rates would be greater than

estimated) . The only way to account for all of these complex

phenonema is to conduct a full-scale room burn and input the release

rates to the transport model. Alternatively, a detailed combustion

model such as HARVARD V can be used as the source of the specified

fire data, although it too has its own assumptions and limitations.

222
TRANSPORT

The distribution of energy and mass throughout the rooms included in the simulation is done in the model FAST. This is a zone (control volume) model with all of the attendant assumptions and

limitations of such models. The basic assumption is that each room can be

divided into two or more zones each of which is internally uniform in

temperature and composition. In FAST, all rooms have two zones except the

fire room, which has an additional zone for the fire plume (see figure 3 for a representative diagram) . The boundary between the two layers in a room is called the interface.

Other than in very large rooms, the zone assumption does not generally

represent a large source of error since, at least in the spaces close to

the fire, experiments show that buoyantly-stratified layers clearly form.

While in an experiment the temperature can be seen to vary within a given

layer, these variations are small compared to the temperature difference

between the layers.

Beyond the basic zone assumptions, the implementation of any model requires

a mixture of established theory (e.g. conservation equations), empirical

correlations where there are data but no theory (e.g. flow and entrainment

coefficients), and approximations where there are neither (e.g. postflashover combustion chemistry) or where their effect is considered

secondary compared to the "cost” of inclusion. An example of the latter is

the fact that none of the models account for the variation of the thermal

properties of structural materials with temperature. While this would be

fairly simple to add to the computer code, data are scarce over a broad

range of temperature even for the most common materials, and the estimated

error from this assumption is less than for others which should be

addressed first.

With a highly complex model such as FAST, the only reasonable method of

assessing the assumptions and limitations is through the verification and

validation process, which is ongoing at CFR. Until the results of this

process are available, the following list should provide some insight.

1. The "specified fire" input by the user is not subject to the

influences of the room as discussed previously. If a large mass loss

rate is entered, the model will proceed, regardless of the fact that

the available oxygen will limit the portion of that mass which will

burn in the room. This will result in extremely high predicted

temperatures; exceeding the adiabatic flame temperature.

2. With no extant chemistry, the species yields specified will be

used regardless of the oxygen concentration. It is known that the

chemistry changes in low oxygen, with an attendant increase in the

yields of products of incomplete combustion such as CO. Room fire

test data shows that under post- flashover conditions, the production

rate (yield) of CO increases (and C02 decreases) in the room, but much

of the excess CO burns (to C02

) in the door flame. The resulting

concentrations may not differ substantially from the pre - flashover

condition

223
3. The entrainment coefficients are empirically-determined values.

Small errors in this value will have a small effect on the fire plume

or the flow in a single door plume. In a multi -compartment model such

as FAST however, small errors in each door plume are multiplicative as the flow proceeds through many compartments, possibly resulting in a significant error in the furthest rooms.

4. The introduction of gases into the lower layer of each room occurs primarily due to mixing at connections and from the downward flows

along walls (where contact with the wall cools the gas and reduces its

buoyancy) . Since for the former an appropriately- documented mixing

coefficient is not known, and for the latter the associated theory is

only now being developed, neither is included in FAST. Thus, the

lower layer cannot receive smoke or gases and can heat only by

radiation from the upper part of the room heating the lower walls and

floor, which then convect to the lower layer.

5. Plumes are the only mechanism to move energy and mass into the

upper layer of a room, and are formed only by the burning item(s) in

the fire room and by a door jet of upper layer gases flowing through

an opening. When warm, lower layer gases move through a low opening

(e.g. the undercut of a door) by expansion, they are "trapped" in the

lower layer of the room into which they flowed until the upper layer

in the source room drops to the level of the undercut and the door jet

forms. Thus, for a time the receiving room will show a lower layer

temperature which exceeds the upper (a physically- impossible

condition) . Since the temperatures are low and no gases flow until

the upper layer drops to the undercut, this should not present a problem for hazard analyses

.

6. FAST is the only zone model which does not assume that room pressures remain constant. This allows FAST to track pressure

fluctuations caused by transient phenonema and to work on room(s)

which are not vented to the outside. Under some conditions, the- door

flows can become imbalanced (inflow exceeds outflow) for a significant

time, causing the room pressures to rise and errors in species

concentrations

.

OCCUPANT BEHAVIOR AND EVACUATION: The EXITT model is a fairlystraightforward "node and arc" evacuation model to which an extensive

series of behavioral rules has been added. The assumptions of interest are thus inherent in these rules, and the limitations are mostly associated

with behavior not yet included. For example, the model does not have

people re-entering the building, as they sometimes do. In addition, the

current model is completely deterministic - a specific set of circumstances

always results in a specific action. The data on which the rules were based sometimes identifies several potential actions (e.g. under this

condition, 60% of the time they do A and 40% of the time they do B) . This

represents probabilistic branching, which will be implemented in the

future

.

224
Some of the rules are qualitative (e.g. a man's first action is to

investigate) and some are quantitative (e.g. a woman between the ages of x and y walks at z meters per minute) . The assumed values in quantitative

rules are called parameter values, and the documentation for the model

identifies each, the reason for assigning that value, and how the user can change it.

ACTIVATION OF THERMAL DEVICES: The activation of heat detectors or sprinklers is handled in the program DETACT. The report (included in an appendix to the HAZARD I report) describes the underlying theory and

assumptions used. The basic assumption is one of quasi-steady ceiling

layer gas flow under an unconfined ceiling (no walls) . It is based on the

experimental study done by Factory Mutual Research Corp . for the Fire

Detection Institute and on which the NFPA 72E, Appendix C methods were

developed.

TENABILITY CRITERIA: The evaluation of the impact of exposure on the

occupants is done in the program TENAB . Individual determinations are done

for both incapacitation and lethality from temperature and toxicity (both

with a concentration- time product and a Fractional Effective Dose) along

with potential incapacitation from burns due to flux exposure. No

interactions are currently included (e.g. temperature exposure does not

affect rate of uptake of toxic species) . The basis for the threshold

values used and the derivation of the equations in the FED calculation are provided in a chapter on Tenability Limits, which contains an extensive

list of references. For all cases except flux exposure, the user can easily change the limit values used (and is encouraged to do so as a sensitivity test). Also, the method of presentation of the output of TENAB

facilitates the observation of the sensitivity of the result to the value

selected

The limiting values for temperature are based on the general literature,

which includes some human data. The flux criterion comes from work done

with pig skin, which is generally held to be very similar to human skin (it

is used for temporary grafts for burn patients) . The toxicity data

however, are from the body of combustion toxicology literature which is

based entirely on animal exposures (primarily rodents). Thus, it must be

assumed that humans will exhibit a similar physiological response to the

exposure. Due to the difficulties of assessing the onset of incapacitation

in an animal, the data used- are primarily those where death was the end

point. The values used for incapacitation are assumed to be half of the

lethal level.

The Ct parameter is an attempt to include the toxic impact of the material

without differentiating the constituent gases. This represents a broad

assumption, and will not include the effect of diminished oxygen. The FED

parameter is the first version of the N-Gas model, including the effects

and interactions of the gases CO, C02

, and HCN. Again, the effect of

diminished oxygen is not included at this time. The N-Gas model is under

continuing development, and additional gases will be added as the data are obtained. It is expected that oxygen and the first irritant gas (HC2) will

be included in the next version.

225
ficuri;

l

INTERRELATIONSHIPS

OF

MAJOR

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OF

A

FIRE

HAZARD

MODEL

226

Irdlants

Intoructions
FIGURE 2 HAZARD I SOFTWARE

227
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228
A SUMMARY OF THE ASSUMPTION AND LIMITATIONS OF HAZARD ONE

R. Bukowski , given by H. Nelson, National Bureau of Standards, USA

WATTS: This model seems to have the same objective as the NFPA building fire

simulation model that was written 10 years ago. How do the results of these

two compare?

NELSON: I will answer your question obtusely. I believe the building fire

simulation model is as good a frame for analysis as any other I have seen. However, I also feel that it didn't have scientifically sound methods for

making its calculations; it was a frame awaiting methodology. As it was published, it used a regression analysis from a series of tests and/or

experienced data. It was as valid as that data. Unfortunately, the data was extracted from a narrow series of residential and model home fires.

Therefore, I felt it was extraordinarily limited.

If that's wrong, I invite anybody else to make their own comparison.

EMMONS: This is an excellent model which is still in its early stages, and it

has obvious limitations, so we'd like to note that there are already models

that could be used in place of the specified fire, and should be, I think, put

in as selectable by the user, so that they may have a specified fire, or they

select one that will predict the fire, and read them properly.

A final observation is somewhat more in the distant future; it is the fact

that unless the improved fire growth model is included, one cannot begin to

obtain real data based on rules. There are thousands of fires every year, in some of which persons die, and in a few of those we would know enough to

develop buildings, the source of ignition, the growth of the fire, to be able

to use the observed position about it to reduce the processes.

NELSON: And to test the human decisions too.

HALL: Your invitation on the comparison for BFSM, I think the contrast I would draw is that BFSM used a probabilistic format throughout, and it forced

everything into that even if it was actually deterministic. This used a deterministic format throughout and forces everything into it even if it's

probabilistic. It is as difficult to compare those as to compare Dr.

Sekizawa's point score approach to this kind of modeling. The format

differences are so great.

NELSON: I think that's an excellent picture.

SEKIZAWA: The question I have is, rather than determining the building itself

and the hazards to the building, are there any thoughts which lean towards the

content safety?

NELSON: Do you mean the impact of the contents on the safety of people?

SEKIZAWA: Yes.

NELSON: You cannot model without considering what is burning, and what is

burning is contents

229
SEKIZAWA: Then, in chat case, who are you targeting to be the most prolific

user for this model?

NELSON: I'll have to give my opinion. I believe the people who make

furniture or wall linings, particularly those that make synthetic materials

will wish or be forced to use this type of model to determine whether the

hazard produced by their product is acceptable. Let me continue a moment. I believe that this system, HAZARD 1, will be applied to prototypical buildings

to establish a baseline of hazard that is normally accepted in society, and

then variables such as a new style of plastic or a new use of a material, will

be entered in the system to find out the change in the overall hazard.

230
FIRE TOXICITY AND CHEMISTRY

231
PROGRESS REPORT

on FIRE AND TOXICITY CHEMISTRY IN JAPAN

by

Hiroaki SUZUKI

Head, Smoke Control Division

Environment, Design & Fire Department

Building Research Institute

Ministry of Construction

1 Tatehara, Oho-Machi

Tsukuba-Gun, Ibaraki-Ken, 305

JAPAN

Studies of fire and toxicity chemistry have been pursued from the point

of production of toxic gases from building materials, behavior of animal

exposed to laboratory environment and to smoke from full scale fires in order

to get lethal conditions of animal in toxic gases.

Researchers concerned with toxicity work in Japan are about 20 who

belong to national research organizations, universities and laboratories

under juridical foundations, etc. Some of them study independently and

others cooperated each other under a project.

The project of combustion toxicity from building materials and interior

goods in fires had been conducted for these 5 years since April 1982. This

project has been supported by the. Science and Technology Agency, Japanese

Government and has covered the following subjects in corporation of research

organizations and universities.

1. Establishment of combustion conditions

2. Burning tests in full-scale model rooms and in Laboratory test apparatus

3. Numerical analysis of burning procedures

4. Development of testing apparatus

5. Evaluation method of toxicity by animal exposure

The items of the results of the projects which have been conducted in

these 2 years, are mentioned below.

a) Burning behavior of building materials under fire conditions (1)

b) Prediction methods of compartment fires in their early stages (2)

c) Production of toxic gases in a semi full-scale compartment under forced

air-control conditions (3)

d) Application of burning conditions obtained from full-scale fire

experiments to a test apparatus (4)

e) Evaluation of toxicity of combustion products from textile (5)

f) Toxicity test using mixed gases (6)

g) Evaluation of combustion toxicity of building materials using a new testing apparatus (7)

h) Toxicity effect of pure and mixed gases (8)

i) Toxicity effects of combustion products to human bodies (9)

j) Analysis of combustion products from fire treated wood (10)

232
k) Combustion toxicity from foamed plastics (11)

l) Incapacitation of mice exposed to combustion products generated in

full-scale fire experiments (12)

m) Evaluation of toxic gases from experimental fires in an existing

building (13)

Some of these were reported in annual meetings of Architectural

Institute of Japan and of Japan Association for Fire Science and Engineering

as well as reported at the 5th Expert Meeting of the US-Japan-Canada

Cooperative Research Group on Toxicity of Combustion Products from Building

Materials and Interior Goods. The last 7 subjects are to be reported in the

9th Joint Panel Meeting UJNR Panel on Fire Research & Safety jointed with the

6th Expert Meeting of the US-Japan-Canada Cooperative Research Group on Toxicity of Combustion Products from Building Materials and Interior Goods.

Some papers which were not envolved in the UJNR Joint Meeting and the

Expert Meeting are summarized as follows.

K. Kishitani (14) et al exposed mice in the combustion products which

generated in full-scale two storied wooden houses. Fire initiated in the

living room on the first floor and spread rooms in the same floor and the

upper along the staircase. Mice in a corridor on the second floor were incapacitated at 18 minutes and 43 seconds when the concentration of CO

became suddenly high. The CO was 1.4% and temperature was higher than 98°C

in the atmosphere. At another site, mice were incapacitated at 34 minutes

and 33 seconds when CO became 0.7% and temperature became higher than. 12Q°C.

In other case, mice* were incapacitated at 39 minutes and 5 seconds when

CO was 0.65% and temperature was higher than 60°C. In these full scale

experiments mice seemed to be incapacitated not only caused by CO

concentration but also high temperatures.

k. Kishitani et al (15) exposed mice, which were trained with electric

shock and non-trained, to experimental conditions with various concentration

of CO gas. In results, trained mice lost their knowledge. Heavier groups of

mice in weight died in earlier time in the same concentration of CO (0.6% and

1.3%) than the lighter ones. This seemed that heavier group of mice needed

more oxygen than the lighter.

M. Hirota (16) predicted the concentration of CO along with movement of

smoke in the central staircase Using mathematical simulation in a 5-storymodel building.

(References)

(1) T. Tanaka, M. Yoshida and I. Nakaya: Burning Characteristics of Building

Materials under Fire Conditions, (*1)

(2) T. Yoshikawa, K. Kawagoe and T. Joh: Numerical Simulation of Early Stage

of a Compartment Fire (*1)

(3) T. Morikawa, E. Yanai: Toxic Gases Evolution from Air-Controlled Fires

in A Semi-Full Scale Room (*1)

(4) S. Yusa: Application of Burning Conditions Obtained from Full Scale

Experiments to Labo-Scale Test Apparatus (*1)

(5) M. Furuva: Evaluation of Toxicity on Combustion Gas from Fiber Materials

(* 1 )

233
(6) T. Sakurai: Toxic Gas Test by the Several Pure and the Mixed Gas (*1)

(7) S. Yusa: Evaluations of Combustion Toxic from Various Materials by Newly

Developed Apparatus (*2)

(8) T. Sakurai: Toxic Gas Test by the Several Pure and Mixture Gases (*2)

(9) Y. Tsuda and Y. Nishimaru: Human Effects of Poisoneous Gases Generated

during Fire Especially, Presence of HCN and CO (*2)

(10) T. Hirata and Sumire Kawamoto: Toxic Gases from Combustion of Wood

Treated with Retardants (*2)

(11) F. Saito and M. Yoshida: Combustion Behavior of Building Materials Using

a Model Box with Air Supplied System (*2)

(12) T. Tanaka and M. Yoshida: Incapacitation of Mice Exposed to Gases in

Full-Scale Fire Tests (*2)

(13) T. Morikawa and E. Yanai: Evaluation of Toxic Gases from Experimental

Fires in an Existing Building (*2)

(14) K. Kishitani, F. Kon, S. Sugahara, H. Sato, S. Yusa and S. Uesugi: Fullscale Fire Experiments Using Traditional Wooden Houses, Summaries of

Technical Papers of Annual Meeting of Architectural Institute of Japan

(1986)

(15) K. Kishitani, M. Yun, S. Sugahara and K. Nishine: Experimental Studies

on Physiological Effects by Combustion Products, Summaries of Technical

Papers of Annual Meeting of Architectural Institute of Japan (1986)

(16) M. Hirota: Application of Smoke Movement Model to Fire Safety Plan,

Summaries of Technical Papers of Annual Meeting of Architectural

Institute of Japan (1986)

(*1) : The 5th Expert Meeting of the US-Japan-Canada Cooperative Research

Group on Toxicity of Combustion Products from Building Materials and

Interior Goods

(*2) : The 9th Joint Panel Meeting UJNR Panel on Fire Research & Safety

234
Progress Report on

Fire Toxicity and Chemistry Research

in the United States

RICHARD G. GANN

Center for Fire Research

National Bureau of Standards

Gaithersburg, MD 20899

FIRE TOXICITY RESEARCH

The research directions discussed at the 8th Meeting of the U.S. Panel on Fire Research and Safety have been vigorously pursued during the past two

years. These include generating the framework in which smoke toxicity

measurements should be evaluated, developing the appropriate measurements

to make, and reducing the dependence on animal testing for that purpose. A

subsequent progress report was presented at the 5th Expert Meeting of the

Tripartite Cooperative Study Group on the Toxicity of Combustion Products

fourteen months ago

Most recently, Researchers at the National Bureau of Standards have

generated a large set of rat lethality data for CO, CO 2

, and HCN for a range of exposure times. Staff at the Southwest Research Institute have

completed studies on the toxicity of CO and HC( in air, as well as comparison with smoke from polyvinylchloride. A simplified version of the

N-gas model has been incorporated in the prototype NBS fire hazard

assessment method, HAZARD 1 , with an upgrade to be included in the next

version.

A quick series of experiments coupling the Cone Calorimeter with the

University of Pittsburgh animal exposure system was an engineering success

Further work is underway to determine the value of this approach to

generating data for fire hazard models.

The Vinyl Institute, a division of the Society of the Plastics Industry,

has completed a comprehensive study on the generation and decay of HC( in room fires. Two publications are in progress, and an integration of the

results into the fire hazard models is in progress.

Forensic scientists at St. Johns University of New York, with the

collaboration of NBS staff, are further refining the protocols and analytic

methodology for determining HCN in blood.

235
NBS, with the support of the Consumer Product Safety Commission, has

completed a first cut at developing the bases for comparing toxicity data

from bench-scale tests with data from full-scale experiments. These

preliminary results point the way for a second series of tests to be

performed this summer.

The National Research Council's Committee on Fire Toxicology has issued its

final report: Fire & Smoke: Understanding the Hazards . A second NRC panel,

the Committee on Toxicity Hazards of Materials Used in Rail Vehicles, is

dormant. A third NRC panel, Toxicity of Complex Mixtures, has incorporated

results from toxicity studies of combustion atmospheres. That report is to

be released soon.

While the research emphasis has been on developing input data for fire

hazard assessment methods, the legislative concentration has been on standalone test methods. The State of New York has passed a regulation stating

that all building products sold in the State must be tested for toxic

potency using the University of Pittsburgh test method. Compliance is to

be phased in over a four-year period. Since there may be as many as 10^

different products to be tested, this action has started a sizable effort

aimed at creating classes of products for representative testing. Since

the first listing of products must be completed by December, 1987, new information on the progress in this area will be developing rapidly.

FIRE CHEMISTRY

Research in this area has been directed at understanding how polymeric

materials burn and how their combustion leads to the formation of specific

combustion products, e.g., soot and toxic gases.

Late in 1986, Gann, Drews and Dipert published a review chapter on polymer

flammability. This discusses the fundamental concepts of materials

burning, degradation and fire retardancy mechanisms, and current test

methods

.

Significant progress has been made in understanding the mechanisms of

polymer gasification. Kashiwagi and co-workers at NBS have published

several papers on the chemical degradation of and thermal profiles in

burning poly (methylmethacrylate) . Parker (NBS), Ohlemiller (NBS) and

Atreya (Michigan State University) have studied the mechanisms of wood

decomposition. Orloff and de Ris at Factory Mutual Research Corporation

have studied correlations between concentrations of gaseous species

produced by fires of varying size.

236
Studies of the effects of fire retardants (FR) on the thermal decomposition

of polymers are also underway. Wilkie (Marquette University) is studying

the chemistry of FR-promoted crosslinking in poly(ethyleneterephthalate)

Starnes (Polytechnic University of New York) is investigating similar

effects of copper compounds in poly (vinylchloride)

Soot formation is a central issue. There is a large effort in the U.S. to

determine the chemistry involved in the nucleation of soot particles, the

growth from C]_ molecules to species that inevitably lead to particles . Several complex chemical kinetic models have been proposed, mainly based on premixed flame data. A sizable data base on premixed flames has been

obtained by Howard and co-workers at M.I.T. Recent research at NBS by

Smyth and Miller has probed diffusion flames, allowing extension to the

types of combustion prevalent in fires. Research at the United

Technologies Research Center has investigated the effects of additives on soot yields.

Mulholland and co-workers have been studying the agglomerate structure of

soot using transmission electron microscopy and computer simulation.

Babrauskas and Mulholland have begun studying possible correlations between

smoke yields and obscuration at bench- and full-scale.

Understanding the yields of carbon monoxide and other principal toxicants

in fires is important for accurate fire hazard modeling. Tewarson at

Factory Mutual Research Corporation and research staff at NBS are developing correlations for CO formation from the wealth of existing smalland large-scale data.

237
Canadian Overview

by

T.Z. Harmathy

Institute for Research in Construction, National Research Council of Canada

Montreal Road, Ottawa, K1A 0R6 , Canada

Prepared for the 6th Canada-Japan-USA Trilateral Meeting

The staff of the Fire Research Section of the Institute for Research in

Construction, National Research Council of Canada, consists of

11 professionals, 12 technical officers, 1 laboratory attendant, and

1 secretary. Of the 11 professionals 10 are research officers, and 1 is a guest worker.

The objective of the Fire Research Section is to provide research

support to the effort of improving fire safety in buildings and reducing the

cost of fire protection. This objective is achieved by 1) responding to the

research needs of building product manufacturers, building designers,

regulatory officials, and the fire services, 2) Darticipating in the

development of building codes and fire test standards, 3) building up basic

information necessary for converting from the present regulation-based fire

safety system to a knowledge-based system.

A signficant portion of the Section's work is carried out in the

National Fire Laboratory located in Almonte, Ontario, about 60 km from the

NRCC campus. It is a national facility for conducting full scale studies of

the characteristics of large fires, smoke control measures in tall

buildings, and fire fighting techniques and equipment. The facilities are available on a contract basis to the private sector, government agencies and

educational institutions for the development of new products, systems, and

procedures

.

The program of the Section for the 1987-88 fiscal year has not yet been

approved. In its tentative form it consists of four programs:

F3 - Fire Resistance of Building Elements and Assemblies

F7 - Fire Performance of Materials and Products

F8 - Hazard Assessment of Building Fires

F9 - Fire Spread Via Exterior Walls

The four programs are subdivided into 18 projects. Two of the 18

projects are related directly to fire gas toxicity:

F7.1 - Air quality in houses following a fire

F7.2 - Rate of heat release and smoke

One other related project was completed in the 1986-87 fiscal year:

FI. 7 - Gaseous decomposition products of nitrogen-containing polymers

The following progress has been reported:

FI. 7 - A TGA/APCI/MS/MS system for the study of pyrolysis and combustion

products was developed. Using that system, the pyrolysis of

polyacrylonitrile was studied.

238
F7.1 - Literature study was conducted on methods of odour control and on the

composition of fumigants. Preliminary experiments were carried out.

The purpose of the work is to establish criteria for the habitability

of houses after fire incidents.

F7.2 - The existing ASTM E906 heat release rate apparatus was modified to

make it suitable for the test proposed by the Federal Aviation

Administration (U.S.A.). Further changes in the FAA test method are expected. The test will yield information on the rate of heat

release, and smoke and toxicity generated during the burning of

materials used in forming the interior surfaces of aircrafts.

Several other projects of the Section overlap on the subject of smoke

and toxicity: Volatile organic components in indoor air, Fully developed

fires of plastics. Gas-phase combustion in fully-developed fires, and

Development of facilities for a standard room-fire test.

239
INCAPACITATION OF MICE EXPOSED TO GASES IN FULL SCALE FIRE TESTS

by

Masashi YOSHIDA

and

Takeyoshi TANAKA

Building Research Institute

Ministry of Construction

1 Tatehara, Oho-machi, Tsukuba-gun

Ibaraki-ken 305, Japan

1. INTRODUCTION

Attempts are being made to develop the method to rationally evaluate the

potentially toxic hazards of fire effluent gases due to burning of various

materials in fire. The test methods to be incorporated into the evaluation

methods of toxic hazards need to be laboratory scale ones in view of

reproducibility, cost and labor. However, since the conditions of burning of

materials can be different in many respects between small scale tests and

real fires, the validity of the test and evaluation methods to be proposed

need to be examined at least by some full scale fire test results. A series

of experiments in which mice were exposed to the gases produced in full scale

fire were conducted to contribute the data to the verification of toxicity

test and evaluation methods.

2. TOXIC HAZARDS AND FIRE SCENARIOS

Once a fire breaks out, certain kinds of combustion products, which are more

or less always harmful, are inevitably produced. Therefore, obviously, it is

impossible to remove all the materials around us on the ground that they are potentially hazardous in the event of fire. If there is any material that

needs to be rejected from the viewpoint of fire safety, it will be one which

causes extremely toxic hazards. So, for clarity, we classify the hazards due

to the toxicity of fire effluents into the two; (a) ordinary hazards and

(b) extraordinary hazards, as follows:

(a) ordinary hazards : the hazards caused by usual combustion products

in fire, i.e. heat, oxygen depletion, carbon

dioxide and, in a sense, carbon monoxide.

(b) extraordinary hazards : the hazards caused by unusual combustion

products

.

As for ordinary hazards, we cannot help but suffer them, otherwise we have to sacrifice tremendous convenience we are enjoying from various kinds of

organic materials. Only when the toxicity of the combustion products or the

amount of the production of toxic products of a material is unusually high,

we need to make the decision whether to allow or reject it considering the

potential hazards in fire and the benefits in every day lives.

240
Another issue concerning toxic hazards due to fire effluents is the

definition of the fire scenarios to be considered, i.e., on what situations

of fire we should discuss the hazards of toxicity. As such scenarios, we

assume that the following two are generally acceptable.

(fire condition) (potential victims)

(i) Scenario(l)

(ii) Scenario(2)

a part of contents in the room of origin is burning.

considerable part of the room of origin is involved

in fire.

residents in the room of

origin(can be disabled,

sleeping, etc.)

people outside the room of

origin(can be evacuees,

fire fighters, etc.)

In either scenario, we have to consider significantly cooled down gases,

otherwise people will be killed almost in an instant by heat, but it is also

important to note that the concentrations of ordinary gases in combustion

product gases, -e.g., oxygen and carbon dioxide, and the temperature are not

independent each other in real fire since mixing with fresh air is thought to

be one of the dominant mode of cooling of fire gases. One of the purpose of

this series of experiments is to investigate che hazards of fire effluents

under the existence of the heat of fire, which can never be disregarded in

real fires.

3. EXPERIMENTS FOR SCENARIO (1)

3.1. Experimentals

3.1.1. Full scale fire test apparatus

BRI full scale fire test apparatus was used for the series of the experiments

for Scenario (1). As shown in Fig.l, this apparatus has two rooms connected

by a doorway. The materials were burned in che bum room. If the scenario (1) was literally interpreted the mice should have been exposed in the bum room,

but they were exposed to the gases in the adjacent room because it was difficult to bum the building materials and still obtain che temperature

acceptably low for the exposure in the burn room despite that the loading of

material was not so large.

3.1.2. Fire source

The kinds of fuel burned in this series of experiments are as follows:

a) Propane gas

b) Yezo spruce (crib)

c) Hard board (crib)

d) Plywood (non-treated , crib))

e) Plywood treated with 5% of DAP (crib)

f) Plywood created with 30% of DAP (crib)

g) Wool carpet

h) Nylon carpet

i) Aclylo-Nitrile carpet

The propane gas was used as one of the fuels in an attempt to determine the

reference level of toxic hazards since any unusual toxicant is not considered

to exist in its combustion products, in other words, only usual products are supposed to be responsible for che hazards. On the other hand, if any

unusually toxic gas is produced from the burning of a realistic material, it

will cause shorter incapacitation time of mice.

241
3.1.3. Method of exposure

The mice are exposed Co Che combuscion produce gases ac chree posicions of

different heighes as shown in Fig.l. The Cemperacure and che concencracions

of 0-, C0

? and CO were measured ac Che posicions of exposure. Typical

examples of Che measured cemperacure and che gas concencracions ac a poinc

are shown in Fig. 3 for che cases of propane gas source and a realiscic fuel.

As can be seen from che example in chis figure, while che cemperacure and gas

concencracion hiscories of propane fire exhibiC almosc sceady scace condicions ac che exposure poincs afeer cercain Cime elapsed, chose for a realiscic fire could hardly gee such condicions.

So, monicoring che cemperacure of Che exposure posicions and deciding che

appropriace Cime, che mice in che rocary cages as shown in Fig. 2. were inserced inco che chree posicions.

3.2. Experimental Resulcs

3.2.1. Resulcs of experimencs

The incapacicacion cimes of mice ac each posicion in each experimenc are shown in Table 1.. The cesc condicions and che average cemperacure and gas

concencracions are also given in Table 1.

3.2.2. Reference hazard level

The incapacicacion cimes of mice exposed co che gases of propane fire are plocced versus che average cemperacures of exposure in Fig. 4. As can be seen

in chis figure, some extent of dispersion is observed among che daca, so ic was choughc chac che incapacicacion cime cannoe be fully explained in cerms

of cemperacure alone alchough mosc influential. So, assuming an equacion

similar with Harber's law but slightly modified by adding the factor on CO^

concentration as follows:

(a x (T - To) + b x (C - Co)) x (t -co) - Const (1)

where T : temperature

C : concentration

c : cime

a regression technique was applied co yields che coefficients in Eqn.(l),

i.e„, a, b. To, Co, co and Const, so as Co best fit the test resulcs for

propane fire. Only che CO., concentration daca were used in obtaining che

coefficiencs as the gas concencracion data since che 0 o deplecion and che CO,, increase are so closely coupled chac eicher gives similar resulcs and CO

concencracion was negligibly low. The result of che best fit regression is

shown by the solid line in Fig. 5.

3.2.3. Incapacitacion cime by che gases from realiscic materials

The results of incapacitation cimes by fire effluents of realistic materials

are shown in Figs. 6(a), (b), (c) . It follows that a material is more hazardous than propane gas if che plots for che material lie below che

reference line, but unexpectedly we could noc find a material clearly more hazardous than che reference level. This may be partly because the mice were

exposed co che acmosphere more or less at declining stage in the case of

realistic fuels, but some ocher factor such as che concentrations of H.,0

might have been involved in che results. If we compare che resulcs for

carpets, which are synthetic materials, with chose for Yezo spruce, a natural

wood, any notable difference cannot be observed. This seems Co indicate that

242
Che incapacitation of mice are determined almost by the usual causes of

hazard i.e. heat, C0

2

, 0

? depletion and such. The same is true with the other

kinds of wood products except that the plywood treated with 30 % DAP may be

somewhat more dangerous than the others.

4. EXPERIMENTS FOR SCENARIO (2)

4.1. Experimentals

4.1.1. Full scale fire test apparatus

For the series of experiments for scenario (2) , the full scale fire test

facility was modified as shown in Fig. 7. The opening of the bum room was narrowed to simulate ventilation controlled fires. The fire effluents flowing

out through the opening was led to the flow channel set up outside the

facility. The flow channel has three weirs implemented to change the

temperature and the concentrations of gases along the channel by entraining

fresh air at each weir.

4.1.2. Fire source

The materials burned as the fuels in this series of experiments are as follows

a) Propane gas

b) Yezo spruce (crib)

c) Plywood (non-treated , crib)

d) Insulation board (qrib)

e) Hard board (crib)

f) Particle board (crib)

g) Phenol foam (crib)

h) Isocyanurate foam(crib)

These materials were burned in the form of cribs except propane gas. Since

the opening of the bum room was limited, most of the experiments were run under ventilation controlled condition.

4.1.3. Method of exposure

The mice were exposed to the gases at three positions on the flow channel at

certain time after ignition, which was decided to be appropriate for the

exposure judging from the temperature. As in the experiments for the scenario

(1) , the temperature and the concentrations of 0.,, C0 ? and CO are measured at

the positions of exposure and the movements of the mice were monitored by the

rotation of mouse cages.

4.2. Experimental Results

4.2.1. Results of the experiments

Examples of the history of the temperature and the concentration of gases at

a position of exposure are shown in Fig. 3. It can be seen in the examples

that while the CO concentration for propane source is insignificant, that for

a realistic material is considerably high. The results of the incapacitation

time of the mice are shown in Table 2 together with the test conditions and

the average temperature T and concentrations C of the gases, which were

calculated as:

243
where T : temperature at the position of exposure

C : concentration of a gas at the position of exposure

t^ : time at which exposure starts

t

?

: average time at which the mice ceased to rotate cages

4.2.1. Reference level of hazard

The incapacitation times of mice for propane fire source are plotted versus

the average exposure temperatures in Fig. 9. The dispersion of the data does

not appear so significant. The solid line indicates the best fit regression

curve of the incapacitation time as a function of temperature alone.

4.2.2. The incapacitation time by fire effluents from realistic materials

The incapacitaion times of the mice exposed to the effluents of the realistic

materials are shown in Figs. 10(a), (b) , (c) . It can be seen that the times

of incapacitation for the realistic materials are consistently shorter than

those for propane fire. On the other hand, as can be seen in Fig. 10, the

difference of the incapacitation times among most of the realistic materials

are not remarkable although phenol foam and isocianurate foam may be

significantly more toxic than the others. In interpreting this result, it

should be kept in mind that this series of experiments were mostly conducted

under ventilation controlled condition and by this reason, not negligible

amount of CO was yielded because of the incomplete combustion in the case of

the realistic materials, while in the case of propane fire CO yield was not

so significant. The effects of CO is suspected to have been important in

shortening the incapacitation times for realistic materials although we have

not analyzed the data in detail.

5. CONCLUDING REMARKS

Some data of the incapacitation times of mice exposed to the fire effluents

of full scale fire tests were obtained. With the limitation of the

apparatuses for gas analysis, the analyses of complex gases could not be made

so the data may not be sufficient for constructing a method to evaluate toxic

hazard of materials. However, it is felt that the hazards due to usual

products of combustion can never be neglected, in particular, heat may play

very important role in threatening human life in usual fire conditions. As

long as this series of experiments are concerned, heat seemed to be most

dominant cause of the incapacitation of the mice, although very unusual

materials were not tested. More thorough studies are necessary before a solid

conclusion is drawn.

244
Incapacitation

time

and

exposure

condition

of

the

mice

^ecensrio(l)

1 ^

g

I - y O C5 ^5 — o O o —

O

; Cl Cl Cl Cl c-3 c-3 C-3 C-3 C-3 C-3 C! C-3 C-3 Cl o |

j

e s ui 00 O HI

cd «<;

1 i 1 o o o o O o o

- g s e je

— * 00 oo \ \ \ 30

30 -XlO r—00 -XP~00 4000 -XCM

00 -X

00 -X

0-3

00 -xCl00 -XCO00 -XCM

CM 40

H5P '-0° H3° C-3 c-3 C-3 CM CM

1

Materials

<0c

38a*o ®

C 33 Oe O

oN

<U >-

spruce P

1

y

wood

PUT

- 5

X

**om

1 H

3

cu 5 _ CM Cl H3» ul 40 P** CO ai o —

CM C-3 -^D 2 — — ~

245

House

Position

II

—1

. 5

M

from

the

floor

2)

S

. D“S

t

and

ard

Deviation

3)L"Life

H-0.9H

from

the

floor

D“0.3ll

from

the

floor
Table

246
Incapacitation

time

and

exposure

condition

of

the

mice

^scenarlo(2)

CNV

-3

43 H

0P4'--0C4P40 OOOOOOOO— —oooo 04 o —

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Incapacitation

time

of

mice

CD>C

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| d^dd - —- <o o 04 03 o o o g4 03 r-^ o o- ‘ r*u*or-— o-^roo'-s-oo — —

—

|

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’ p^op-—

030-040300 ‘ o*

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—• co — cm — —

Average

Temp

(a) mcso^csj

| N-o^au:rp-aoinooNOCu5a ooooooooo — r— r— — C3 03 — 04 OO OOOO or— 030004 0 000003 OOO^CM — o— o — oo P— ''O’ O ^ OO^CSOOO''S*000-3-00"3rOO'T r— O^C— 00 0-T0 0 004

Average

gas

concentration

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| o — ssn , cr tr-csurr-ca aerrma—csaocaMC

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(

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1

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oo «» «. : u oe

**« -3

I C ^ « 2 5 = 5 = 5 = 5 = 5 - 5 - 5 = 5 = 5 = 5 = 5 = 5 =

« 3 = o o o o o o c5 o o o o o o o o o o o o o o oo oo

O =£ =

« 06 s 2" s 5" IT 2 \ \ \ \ \ \ \ \

so oX -X

3£ -X

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. 3

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4 2 &

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PL. >•

247

1)K

.

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position
Table

248
Table

249
f

3

„ 4i5

iii

Sobols

a Sucriou. "irarxcoupLa

® Sara Clarxocouple

* Gas saxolirr

Plan

a Seat flux uacar ax she oallu

* Eaac flux ra:ar ax TMa caf I f

a

g

* Bldiracriaxai prassura probe

£<J yfcusa rrrsr*/ cures

Fig.l Experimental setup for scenario(l) experiments

Fig. 2 Mouse Rocary Cage

250
a) Propane gas

_ 5

4 . i

j .

CMa

Z

(M

CJ

n

Fig

0 LO 2Q 20 10 50 SO 70 30 90 100 L 10 120 TIMS MIN

b) Yezo spruce

3 ExampLey of the temperature and gas concentration histories

251
,065*(Xco

2“

.943)+

,0071*

(temp4

1.2)

Temperature

80

70

o9- 50

5Q

40

30

10 20 30 10 Incapacitation time (min)

Fig. 4 Incapacition time V.S. temperature for propane gas (scenario (1)

)

<=»

Propane gas

18 L/min

x

x

X TDrCt

4- 12 L/min

O '6 L/min

a x

Q) 3

x ** XX X ° ® O OD cO

Fig. 5 Best fit regression of the incapacition time for propane gas

252
.

Q65*(Xco 2

-

.

943) +

.

0071*

(TempA

1

.2)

.

065*

(Xco 2

-

.

9«) +

.

007

1*

(Teinp-41

. 2)

.

065*

(Xco2

-

.

943)+.

0071*

(Temp41

. 2)

0.4

0.3

0.2

Oj

0

10 20 _ 30 10 Incapacitation time (min)

I’** x Yezo spruce

O Hard board

\o a ® * a -5,

•

standard hazard

*—*— at

line

a) Yezo spruce and Hard board

>

\ Standard

\ hazard line

i x Plywood

i a ?1wood treated with

| 57. of DAP

V

<x2^\°

rh^]aa a

O PIwood treated with

30% of DAP

J 0*0 \

a

X X X XX

X a

XX— wx *

L0 20 30 40 Incapacitation time (min)

b) Plywood , PWT- 5% and PWT~30%

\ Standard 'r

>

lazard line

i x Wool carpet

D> Nylon carpet

- Acivlo—Nitrile came t

V. x \

o*»< SOB]

>®t»0 r> c

X 2 X5< rsn ^ -*r-—^

** X => x*>

10 20 30 40 Incapacitation time (min)

c) Wool carpet, Nylon carpet and Aclylo-Nitrile carpet

fig. 6 Incapacition times for realistic materials

253
txhausc duee

txiiause duce

>

_ 5

t

_ 5 j ^ 2 2.5

?

'. 5 unie (a) & posision at exposure

Fig. 7 Experimental setup for scenario (2) experiments
a) Propane gas

b) Yezo spruce

Fig. 8 Exampley of Che Cemperacure and gas concentration histories

80

70

09 SQ

0)

1 50

<Ua.sH

d0

30

i0 20 30 io Incapacitation tine (min)

Fig. 9 Incapacition time V.S. temperature for propane gas

255

HR

1* *£* 'OCkss ,

1 Prooane gas

1

\xm

x WX

*"* \

X "H X

*

' x -cm **

<* X*

X * hazard Line - S tanderd

(scenario (2)

)
Temperature

*C

Temperature

a(]

Temperature

'

1

\ S tandai"d hazard line

x Hard board

l a Insulation board

0 Particle board

a"* ^

x /X X

* x qx

aa

-yy ^XQ *

« .. Oy * * *

X X 0 0

oo J

iO 20 20 40 Incapacitation time (min)

b) Hard borad, Insulation borad and Particle board

l S t and a rd hazard lin

x phenol foam

e 0 Isocyanurate foam

\

X x* x«

X #

cfx

0

cT3 ^ X XX x XX *

X X oo o c 0=

to

J 20 30 Incapacitation time (min)

c) Phenol foam and Isocyanurate foam

F-fo-.lO Incapacitation time of mice for realistic materials

256

40
DETERMINATION OF THE KINETIC PARAMETERS AND THE

HEAT OF COMBUSTION OF THE VOLATILES FOR WOOD

William J. Parker 1 Center for Fire Research

National Bureau of Standards, USA

ABSTRACT

An apparatus for measuring the heat of combustion of the volatiles and the

kinetic parameters for the mass loss rate of the various chemical components

of wood as a function of their degree of char is described. The results of some of the measurements on cellulose made with this apparatus are presented

1. INTRODUCTION

The research described in this paper is part of a project in the Center for

Fire Research (CFR) at the National Bureau of Standards (NBS) to provide a heat release rate model for wood that can be used as a sub-model for the

flame spread and room fire growth models. A heat release rate model has

been described by Parker (1) which takes char shrinkage and multiple

chemical components into account along with the change in the thermophysical

properties with temperature and mass retention fraction. The mass retention

fraction is defined as the mass at any given time divided by the original

mass and is designated in this paper by the symbol, Z. The crucial factor

in validating the model is to have reliable input data on the thermophysical

and thermochemical properties of the wood and its chemical components . These

components include cellulose, lignin, mannan, xylan and galactan This paper

describes the method used for the determination of the heat of combustion of

the volatiles and the kinetic parameters which describe the rate of mass

loss during pyrolysis. The heat of combustion of the volatiles released

during the pyrolysis of char forming materials is less than the heat of

combustion of the same material as measured with an oxygen bomb calorimeter

and can vary during the course of the pyrolysis.

Ordinarily, the kinetic parameters (i.e., the effective activation energy

and pre - exponential factor) are measured by thermogravimetric analysis

(TGA) . However, the heating rates in TGA are typically less than 20 degrees

per minute. Consequently, a considerable fraction of the material has

pyrolyzed by the time it reaches the temperature range of rapid pyrolysis

that would be typical in the interior of a burning piece of wood. Thus, the

calculated activation energy is characteristic of a lower temperature

reaction, perhaps more typical of smoldering. While 100 °C/min heating

rates are available in the TGA apparatus they result in unacceptably large

temperature gradients in the specimen. Measurements are needed in the

temperature range at which most of the pyrolysis occurs during flaming wood

combustion. This temperature range will depend on the component. It will

Part of this work was performed during the author's tenure as a visiting scientist at the Forest Products Laboratory during the summer of 1985.

257
be relatively high for lignin, low for hemicellulose , and in between for

cellulose (2). Generally, however, most of the pyrolysis occurs at

temperatures between 300° C and 400° C even though the pyrolysis process is

initiated below 300° C and the char continues to degrade even at temperatures

above 600° C.

The heat of combustion of the volatiles might be deduced from measurements

in an oxygen bomb calorimeter on virgin material and on chars of the same material at a series of mass retention fractions. However, the net heat of

combustion is required for modeling, while the oxygen bomb calorimeter measures the gross heat of combustion. The quantity of liquid water formed

must be known at each mass retention fraction in order to make the

correction. Also the determination requires many specimens, each of which

must have a mass of a gram or more of the charred material. This means

that several grams of the original material are needed. The separation of

pure unaltered components of wood is very tedious and so the quantities

available, except for cellulose, are limited.

Susott (3) measured the heat of combustion of the volatiles from forest

fuels using a combination of TGA and evolved gas analysis (EGA) . The EGA uses a catalytic converter to completely oxidize the volatile pyrolysis

products to carbon dioxide and water. The amount of oxygen consumed in this

process is measured and used to calculate the potential heat release rate.

This rate is divided by the mass loss rate, determined by TGA, to yield the

heat of combustion of the volatiles. This technique has the advantage of

requiring a very small specimen size, on the order of 5 mg. However, it

does have the disadvantage of depending on two specimens having the same identical mass loss rate history in two different instruments. Furthermore,

the rate of temperature rise of the specimen is limited to around 20 °C per

minute, which limits the applicability to the lower temperature reactions.

2 . EXPERIMENTAL

Because of the above difficulties, the PYROCAT (combination pyrolyzer and

catalytic converter) was developed to measure the kinetic parameters and the

heat of combustion of the volatiles for this project. In its present state

this method is applicable to materials containing only carbon, hydrogen and

oxygen. By heating the specimen rapidly to the temperature region where the

measurements are being made, steep temperature gradients in the specimen are confined to the early period where less than 10% of the mass is lost. The

specimen is pyrolyzed in a stream of nitrogen and the volatiles are swept

into a catalytic converter where they are converted to carbon dioxide and

water as in Susott' s method (3). In the PYROCAT, however, the converted

gases are passed through carbon dioxide and water vapor analyzers as well as the oxygen analyzer. From the concentrations, Xq 2 , XC02 and XH20 , of

oxygen, carbon dioxide and water vapor measured in the analyzers and the mass flow rate, mN

2

, of nitrogen flowing into the system, the molar flows of

carbon, hydrogen and oxygen from the specimen can be calculated from the

formulas. (4)

:

258
n*c 3 mN

2 X^. 0 2 / ( 1 "Xq 2

2

) /7 (1)

mH mN2^H2o/^'^H20 '^C02

niQ — 4mN 2 [ (XH 2 o / ( ^ ‘^H 2 0 ) / '^O 2 *^C 0 2

) + 2 ( Xq 2 +Xg g 2 )/(l‘Xo2'Xco2)

-2XO2

o /( 1-X02o )]/7 (3)

where Xg 2

0 is the measured oxygen concentration prior to the run. The

calculation does not depend on nor provide any information on how these

elements are combined in the hundreds of' different chemical species making

up the volatile pyrolysis products. However, the total mass flow from the

specimen is equal to the sum of the masses of these three elements.

Furthermore, the external oxygen supply rate required for complete

combustion of these volatiles is given by

mg, = (8mg/3 + 8mH

- mg) . (4)

Huggett (5) found that, to within an engineering approximation of ±5%, the

heat released per unit mass of oxygen consumed is 13.1 MJ/kg for a large

variety of polymeric solids. The potential rate of heat release can be

calculated by multiplying the oxygen mass requirement evaluated in equation

4 by this constant. The heat of combustion is simply the ratio of the rate

of heat release to the rate of mass loss. Thus both the heat of combustion

of the volatiles and the mass loss rate for the calculation of the kinetic

parameters can be determined on a single specimen. Data from several

specimens pyrolyzed in different temperature ranges need to be accumulated

to determine the kinetic parameters.

The PYROCAT was first used in a project at the Forest Products Laboratory

during the summer of 1985 to obtain the heat of combustion of the volatiles

and the kinetic parameters for the components of Douglas fir. Figure 1 shows a schematic of the apparatus. A 2.5 cm diameter quartz tube passes

through two 30 cm long tube furnaces. A flow of 26 mg/sec of preheated

nitrogen comes in from the right. It passes over the specimen which is

located in an aluminum pan on top of a chrome plated copper block heat sink.

The cellulose specimens were in the form of 0.75 mm (30 mil) thick sheets

with masses between 275 and 325 mg. The other materials were in the form of

powders with masses between 50 and 300 mg. Oxygen flows in from the left

through a 12.5 mm diameter quartz tube mixing with the nitrogen and the

pyrolysis products. The short ceramic rod located in the outer quartz tube

in the zone where the two tube furnaces are in contact serves to increase

the velocity of the nitrogen at that location. This was intended to block

the diffusion of oxygen into the region of the specimen. No confirming

measurements of the oxygen concentration in this region were made. However,

there was no visual indication of char oxidation in any of the specimens

tested. The nickel - chromium heating wire was added to raise the temperature

of the pyrolysis products and to induce some cracking so that there would be

no condensation on the quartz tube in this region. The wire was found to be

sufficiently heated by the furnace that it was not necessary to apply

electrical power to it. There was never any evidence of condensation. As

the mixture passes out through the platinum gauze catalyst at 900° C the

259
pyrolysis products are converted to carbon dioxide and water. The nitrogen

and oxygen flows into the system were measured with rotameters. The oxygen

flow was adjusted to provide a 35% oxygen concentration in the catalytic

converter in the absence of pyrolysis. This initial concentration was high

enough to insure that the concentration never dropped below 21%. This

insured that the conversion was complete as indicated by the absence of a measurable concentration of carbon monoxide. After passing through a heated

filter to eliminate the possibility of particulates reaching the analyzers,

the exhaust stream divides. One part goes through an infrared water vapor

analyzer and the rest through a cold trap and desiccant before going through

the oxygen and carbon dioxide analyzers. There are rotameters in the

exhaust of each of the analyzers to insure that the flows are in the range

specified by the manufacturers. This is around 10 ml/sec. The CO meter is

present simply to indicate if there is detectable CO in the exhaust.

Prior to the test several flow rates-'xof methane were added to the nitrogen

and converted to carbon dioxide and water vapor in the catalytic converter

in order to calibrate the analyzers. In addition, the carbon dioxide

analyzer and the oxygen analyzer were checked with bottled calibration gas

.

There is a solenoid valve, not shown in figure 1, which cuts off the flow of

oxygen into the tube when the end is open for specimen insertion. Otherwise

oxygen would flow out over the specimen. Nitrogen flowing out of the end of

the tube during specimen insertion prevents the diffusion of atmospheric

oxygen into the system. Chromel - alumel thermocouples, 0.25 mm (10 mil) in

diameter, are attached to the copper block and to the specimen. In the case

of powdered specimens the junction is buried in the powder while for

cellulose it is threaded through the sheet. Since there is a positive

pressure in the line, any leakage in the system beyond the quartz tube where

the combustion gases are well mixed would not affect the concentrations

measured in the analyzers

.

The uniformity of the specimen exposure over its lateral extent was estimated by cutting up a cellulose specimen into nine equal parts and

measuring the total mass loss of each piece with a balance after a mass loss

of approximately 30% was obtained. There was a 25% difference in mass loss

per unit area across the length of the specimen which could not be

eliminated. Considering that the rate of mass loss of cellulose essentially

doubles every 10 degrees this represents a temperature difference of about 3 degrees .

The output voltages from the analyzers and thermocouples were recorded at 5 second intervals on a floppy disk. Capacitors were placed across the inputs

to the computer in order to reduce the noise. The data were analyzed later

using an electronic spread sheet.

3. RESULTS FOR CELLULOSE

Figure 2 is a plot of the millimoles per second of carbon, hydrogen and

oxygen flowing out of the specimen as a function of time for one of the

cellulose runs. The specimen temperature increased from 350 to 370° C over

the course of the run. The upswing of the hydrogen at the beginning was due

to the release of water. Most of this water peak was lost because of a

260
transient in the oxygen concentration caused by cutting off the supply

during specimen insertion. Although the shapes of the curves depend on the

temperature history of the specimen they do demonstrate that the carbon,

hydrogen and oxygen content of the volatiles can be followed during the run. Figure 3 shows the fractional mass loss rate as a function of the mass retention fraction (Z) for a cellulose specimen. The fractional mass loss

rate is the sum of the mass flow rates hydrogen, carbon and oxygen from the

specimen divided by its original mass. The right hand edge of the plot where

Z=1 corresponds to the virgin cellulose. The mass loss rate rose to a maximum as the pyrolysis proceeded because the specimen temperature

continued to rise as seen in the figure 4. In principle the specimen

temperature does not need to be constant. It is only necessary that it gets

into the proper range before the specimen has lost more than a small portion

(say 10 percent) of its original mass. However, a correction has to be made

for the temperature change that occurs over the gas transit time from the

thermocouple to the analyzers. The greater the change the greater the

potential error in the correction. The gas transit time was 15 seconds. No

correction was applied for flow dispersion.

The fractional mass loss rate and the temperature at mass retention

fractions of 0.9, 0.8, ..., 0.3 were picked off the fractional mass loss

rate and temperature curves for each of seven runs on cellulose, each in

different temperature ranges. These were used to produce a set of Arrhenius

plots, one for each mass retention fraction, like the one in figure 5 for

Z=0 . 6 . From these plots the effective activation energy and preexponential factor were determined as a function of the mass retention

fraction. The average value of the effective activation energy over the mass retention fraction range of 0.9 to 0.3 was determined to be 45

kcal/mole with a range of ± 3 kcal/mole. Using this average value of the

effective activation energy, the pre - exponential factor in units of

fractional mass loss per second could be expressed by

A = 2.5 x 10 13 (Z-0.06) . (5)

Figure 6 shows the average heat of combustion of the volatiles as a function

of the mass retention fraction over all of the runs on cellulose. The

vertical bars indicate the range of the data and the points represent the

averages. The average of all runs over the complete pyrolysis period is

13.8 MJ/kg.

4. DISCUSSION

The PYROCAT appears to be a promising tool for determining the instantaneous

carbon, hydrogen and oxygen content of the volatiles during pyrolysis. This

information was used here to determine the instantaneous heat of combustion

of the volatiles and the mass loss rate needed for the calculation of the

kinetic parameters. The accuracy of the results is limited by (1)

perturbations in the oxygen concentration at the beginning of the test due

to specimen insertion, (2) a small temperature variation across the

specimen, (3) changes in specimen temperature with time, (4) small changes

261
in the flow rates of nitrogen and oxygen into the system during the test and

(4) system noise

.

The mass loss rates for cellulose could be described quite well by an Arrhenius equation. However, because of the complexity of the reactions

involved it would be difficult to assign any physical significance to an activation energy. Hence the term, "effective activation energy," is used

here simply as a parameter to calculate the mass loss rate. There are a wide range of activation energies reported in the literature for cellulose

in nitrogen with a number of investigators reporting values around 54

kcal/mole (6) primarily using TGA. Although it varied some with the mass retention fraction an average value of 45.1 kcal/mole was obtained on this

project. This is in reasonable agreement with the 42 kcal/mole reported by

Lipska and Wodley (7) between 276 C and 360 C using isothermal pyrolysis.

The net heat of combustion of cellulose as reported by Huggett (5) is 16.1

MJ/kg. The average net heat of combustion of its volatiles measured on this

project was 13.8 MJ/kg. The lower value was due to dilution by the water

released in the char forming process. This dilution was greatest during the

beginning of the pyrolysis period where the heat of combustion of the

volatiles was only projected to be about 9 MJ/kg. Shafizadeh (8) found a gross heat of combustion of 15.2 MJ/kg for the volatiles from cellulose

based on the difference in the heat of combustion of cellulose and its char

using the oxygen bomb as described earlier in this paper. Assuming that his

measured char yield of 14.9% was pure carbon, a correction was made here for

the heat of vaporization of the water. On this basis the net heat of

combustion of the volatiles was calculated to be 13.7 MJ/kg which is in

agreement with the average value of 13.8 MJ/kg found in this project.

4. ACKNOWLEDGMENT

I am grateful to the Forest Products Laboratory for providing me with the

opportunity to carry out the experimental part of this research as a visiting scientist in their laboratory.

5. REFERENCES

1. Parker, W.J., "Prediction of the Heat Release Rate of Wood," Fire

Safety Science - Proceedings of the First International Symposium,

pp 207-216, (1985)

2. Beall, F.C. and Eickner, H.W., "Thermal Degradation of Wood

Components: a review of the literature," U.S.D.A Forest Service Research

Paper FPL 130 (May 1970).

3. Susott, R.A., "Characterization of the Thermal Properties of Forest Fuels

by Combustible Gas Analysis," Forest Sci. Vol.28, No. 2, pp . 404-420

(1982)

4. Parker, W.J., "Development of a Model for the Heat Release Rate of

Wood- A Status Report , "NBSIR 85-3163, Nat. Bur. Stand. (U.S.), May 1985.

262
5. Huggett, C., "Estimation of the Rate of Heat Release by Oxygen

Consumption Measurements," Fire and Materials, Vol.4, pp . 61-65 (1980).

6. Roberts, A.F., "A Review of Kinetics Data for the pyrolysis of Wood and

Related Substances," Combustion and Flame, 14, pp . 261-272 (1970)

7. Lipska A.E. and Wodley, F.A., J. App . Polymer Sci.,13 pp . 851 -865 (1969).

8. Shafizadeh, F., "Combustion, Combustibility, and Heat Release Rate of

Forest Fuels," AlChE SYMPOSIUM SERIES No. 177, Vol. 74, pp . 765-82

(1978) .

FLOW

urrui

FLOW

W€TF« FLOWWTTE* &

CN 4 TiM

*1

T AM 1 01

| FLOW mem» flow wr»m

FLOW

IffTO

1 covmjxrmiJL OAJ MTTZ* 02

AMALTTCT

CO. COj

1. Schematic of PYROCAT (Pyrolyzer and Catalytic Converter)

263
RATE

OF

MASS

LOSS

(

1

/•)

i

tO'

2. Molar Flow Rates of Carbon, Hydrogen and Oxygen

from Cellulose Specimen in the 350°C to 370°C

Temperature Range

3. Fractional Mass Loss Rate of Cellulose Specimen

264
MASS

LOSS

RATE

4 . Temperature of Cellulose Specimen

RECIPROCAL OF ABSOLUTE TEMPERATURE

(1/K) x 10 5

5. Arrhenius Plot for Cellulose at a Mass Retention Fraction of 0.6

265
MASS RETENTION FRACTION (2)

6. Heat of Combustion of Volatiles from Cellulose

266
CORRELATION FOR THE GENERATION OF CARBON MONOXIDE

IN SMALL- AND LARGE-SCALE FIRES

Archibald Tewarson

Factory Mutual Research Corporation

Norwood, MA 02062 USA

Ninth Meeting of the United States/Japan Panel

on Fire Research and Safety, May 1987

ABSTRACT

Generation of carbon monoxide in small- and large-scale fires, under variable

ventilation conditions, is discussed. The ratio of the mass generation rate

of CO to mass generation rate of material vapors, defined as the yield of CO, was found to have a weak dependency on fire scale, but a strong dependency on

the chemical structure of the materials and additives and fire ventilation; it was highest for the highly halogenated materials and decreased as the chemical

structure changed from aromatic to aliphatic.

The yield of CO was found to increase with decrease in fire ventilation,

reaching values as high as 0.30 or about 20? of the maximum possible theoretical. yield of CO.

From the application of the National Bureau of Standards' N-gas model for fire

toxicity, it was found that hazardous conditions could be created if mass aij

to-material vapor ratio was about 1.4 to 2.0 times the mass air-to-fuel stoichiometric ratio for diffusion flames of propane, propylene and wood cribs,

assuming CO and CO 2 to be the only products of concern. However, if the fire

products were diluted by about fifteen to twenty times their volume by air,

hazardous conditions were found to be absent even when fires were highly

under- ventilated and the yield of CO was very high.

1 . INTRODUCTION

Accurate data on the causes of deaths associated with fires are difficult to

obtain. However, it is generally accepted that 70 to 80? of fire deaths re- sult from smoke inhalation [1]. Smoke is a mixture of various gaseous, liquid

and solid compounds such as CO, C0

2

, depleted 0

2

, HC£, HCN, saturated and

unsaturated hydrocarbons , alcohols, aldehydes, etc. Some of the compounds

present in smoke, individually or in combination, are considered to be responsible for not only immediate or delayed deaths, but also for impeding escape

and causing injuries.

CO, one of the major toxicants present in smoke, binds the red blood cells and

forms carboxyhemoglobin (COHb) , which interferes with oxygen transport in the

body. COHb concentrations of 50 to 60? are generally accepted as fatal [1].

Human physiological response to CO has been investigated by several researchers , some of their results are summarized in Figure 1, taken from Reference

2. CO is the cause or a contributing cause in 80? of fire deaths based on the

data collected in Maryland [1]. The effects of CO exposure to rats and mice

have also been investigated [1,3,4]. For example, CO concentrations of 2667

and 8636 ppm for 15 minutes of exposure are found to cause incapacitation and

death in 50? of exposed rats, respectively [ 3 ]. Factors that are considered

267
to be important for CO hazard in fires are [1-4]: 1) concentration of CO; 2)

exposure time; 3) health of the people expected to be exposed; and 4) concen- trations and generic nature of other compounds generated in the fire, which

may affect the CO response of exposed people.

This paper presents an analysis of the generation of CO in fires to show cor- relations between small- and large-scale fires.

2. GENERATION OF CO

CO is generated in flaming and nonflaming fires and is associated with incompleteness of combustion. CO concentration in fires is governed by several

factors: 1) nature of chemical bonds and additives in materials; 2) fire ven- tilation, which includes mixing and flow of fire products and air; and 3) fire

size

.

The mass generation rate of CO is directly proportional to the mass generation

rate of material vapors, where the propor tional ity constant is defined as the

yield of CO. Thus at any time t, the yield of CO can be expressed as:

Y co(c) °C0 (t)/G

f (C) ’ ( 1 )

and the overall average yield of CO as:

t . t . CO / G£

q

(t) dt / /

G’’ (t) dt,

or, Y co - W

Co /w

f,

( 2 )

(3)

where Yqq is the yield of CO (g/g); G" is the mass generation rate of CO per

unit surface area of the material (g/rrrs); G" is the mass generation rate of

material vapors per unit surface area of the material (g/m 2 s); Wnn is the

total mass of CO generated in the fire

rial vapors generated in the fire (g).

,

. 'CO (g); and Wj. is the total mass of mateIn fires, CO is generated as a result of the conversion of the carbon atoms

(and oxygen atoms if present in the structure). The- maximum possible theoretical yield of CO can be calculated from the elemental composition of the

material

,

k CQ = n (MW) C0 /(MW) f

, (4)

where k^g is the maximum possible theoretical yield of CO (g/g); n is the number of carbon atoms in the structure of the material; (MW)gg is the molecular

weight of CO, which is equal to 28 g/mole; and (MW)^ is the molecular weight

of the material calculated from its elemental composition (g/mole). The generation efficiency of CO thus can be expressed as:

f C0

= YC0 /kC0 * ^ 5)

where f co

is the generation efficiency of CO.

268
2.1 EFFECT OF NATURE OF CHEMICAL BONDS ON THE GENERATION OF CO

i Materials consist of combinations of atoms, such as' carbon, hydrogen, oxygen,

nitrogen, sulfur, halogen, etc. The generation of CO depends on the nature of

chemical bonds between these atoms as shown in Figures 2 and 3 • In the

Figures, Y C q

is plotted against the molecular weight of saturated and unsaturated hydrocarbons . Aliphatic hydrocarbons with single, double and triple

bonds are defined as alkanes, alkenes and alkynes, such as ethane, ethylene,

and acetylene, respectively. Benzene is an aromatic hydrocarbon. Hydrocarbons with chemical structures consisting of both aliphatic and aromatic units

are defined as arenes such as styrene and toluene.

Data in Figure 2 show that Y P q

increases with increase in the degree of unsaturation in the car bon-car bon bonds and presence of aromatic rings in the

structure. For example, Y rp increases from normal heptane (CyH^), heptene

(CyH-j jj) , and heptyne (CyH-jg) vo toluene (CyHg). Data in Figure 3 show that

Yqq also increases when nydrogen atoms are substituted with larger groups such

as a methyl radical, CH^ (substituted alkanes) and when structures are cyclic

with loss of hydrogen atoms (cyclic alkanes or alkenes). For example, Y r g

increases from normal pentane (CgH^g). methyl butane (C^H^g) to cyclopentane

(C' 5 H 10 ^‘ The data in Figures 2 and 3 show that the molecular weight also effects Yqq.

Table 1 Yield of CO for Well-Ventilated and

Under -Ventilated Fires of Materials

Materi al Well-Ventilated Under-Venti.

F i r es

a Fires b

Methane (gas) NR 0.175

Propane (gas) 0.005 0.229

Propylene (gas) 0.020 0.200

Hexane (liquid) 0.01 1 0.195

Toluene (liquid) 0.066 0.107

Methanol (liquid) 0.001 0.236

Ethanol (liquid) 0.002 0.21 9 Isopropanol (liquid) 0.002 0. 168

Acetone (liquid) 0.003 0.304

Pine (solid) 0.0011 0. 138

Polymethylmethacrylate (granular

)

0.010 0.189

Silicone (solid) 0.021 NR

Polyethylene foams 0.023 NR

Polyethylene (granular) 0.024 0.180

Polypropylene (granular) 0.024 NR

Nylon (granular) 0.038 NR

Flexible polyurethane foams 0. 01 0 to 0. 042 NR

Rigid polyurethane foams 0.025 to 0. 051 NR

Polystyrene foams 0.060 NR

Polystyrene (granular) 0.060 NR

Polyvinyl chloride (granular) 0.063 NR

Fluoropolymers (granular) 0. 120 NR

a: Data taken from Reference 6.

b: Data taken from Reference 7.

NR: Not reported.

269
The effect of chemical structure on the generation of CO is also noted in

materials other than hydrocarbons . The data are listed in Table 1 for wellventilated flaming fires. The data in Table 1 show that Y P q

increases from

aliphatic to aromatic to halogenated structures.

2.2 EFFECT OF FIRE VENTILATION ON THE GENERATION OF CO

The fire ventilation can be defined in terms of mass air-to-f uel-stoichiometri c f racti on (<f> )

,

<+> = 6 M /G" A k a f a

( 6 )

where 5 is the fraction of entrained air which reacts with material vapors; M

a is the mass flow rate of air (g/s); A is the surface area of material (m^);

and k a is the mass air-to-fuel stoichiometric ratio. In the calculations, 6 is usually considered as unity. When <f> < 1.0, fires are defined as underventilated, and when <j> > 1.0, fires are defined as well-ventilated

.

Data for Y^q for under- ventil ated fires (<j> < 0.83), are included in Table 1,

which are significantly higher than the data for well-venti lated fires. Depending on the material, Y-q values can be as high as 0.30 for under-ventilated fires. High Y^q values have also been found in other studies on propane

diffusion flames [8]". In a small-scale flow-through furnace heated to 800°C,

very high Yqq values have been measured for highly under-ventilated conditions, as can be noted in Table 2 [9].

Table 2

’ Maximum Yield of CO Under Highly

Under-Ventilated Combustion3

Material Y c°

(g/g)

kc°

(g/g)

oo Methanol 0.88 0.88 1 .00

Ethanol 0.80 1 .22 0.66

Acetone 0. 91 1 . 45 0.63

n-Hexane 1 .01 1.95 0.52

Benzene 1 . 24 2.15 0.58

Benzene-Methanol (1:1) 1 .05 NA NA

a: Data taken from Reference 9. Fuel flow rate = 0.005 g/s;

Dimensions of the tube furnace are not given in the paper.

NA : Not applicable.

Data for Y

Cq, calculated from the ratio of the mass generation rate of CO to mass generation rate of material vapors, are available for small- to largescale fires, which are plotted in Figure 4 against air-to-fuel stoichiometric

fraction. The data in Figure 4 show that as <p decreases, Y

Cq

increases, and

for <j> < 2.0, Y Cq

increases very rapidly. For propane and propylene diffusion

flames, the maximum value of Yqq is about 0.27, occurring at a 4

) value of

about 0.80, which is similar to Beyler's data included in Table 1 for <j> < 0.83

[7]. For small- to large- size wood crib fires, and larger-scale wood-lined

tunnel fires, the maximum value of Y Cqis about 0.24, occurring at a <p value

of about 0.50. The Yqq profiles for wood cribs burning in our apparatus and

various sizes of enclosures and in larger-scale wood-lined tunnel fires are very similar.

270
The data in Figure 4 also show that for similar <p values, there is a good cor- relation between Y co data for fires of small- and large-size cribs and wood

lined on the walls of a larger-scale tunnel.

2.3 EFFECT OF FIRE SCALE ON THE GENERATION OF CO

From Eq . ( 1 )

,

G" = Y G"

CO CO f

and

Gf " ^n

/L ’

( 7 )

( 8 )

where q" is the net heat flux received by material, equal to the sum of external heat? flux and flame heat minus the surface reradiation loss (kW/m"); and L is the heat of gasification of the material (kj/g).

If fire products and air are well mixed and V is the total

rate of fire product-air mixture, then the concentration of CO

ed as:c co

c co A/V r

volumetric flow can be expresses

where C^q is the concentration of CO (g/m-3)

From Eqs . ( 7 ) and ( 9)

,

cco

y co 0? A/ h

• (10)

Eq. (10) suggests that the concentration of CO in various fire scales would be

governed by Y^q and by G” A/V (which can be defined as the effective concen- tration of material vapors).

Data presented in Figure 4 suggest that Y r Qhas a weak dependency on the fire

scale for wood crib and wood-lined tunnel fires. Data presented in Table 3 also suggest that the yields of CO as well as CO

2 and HCN have weak dependencies on the fire scale. Data presented in previous sections, however, show

that Y co

is strongly dependent on fire ventilation; thus the concentration of

CO is expected to depend more strongly on fire ventilation, flow and mixing of

fire products and air rather than on the fire scale.

With increase in fire scale the following will increase: 1) q" , and from Eq

(8), G£; 2) A; and V . From Eq. (10), increase in G” _ and A wi^l result in an increase in CO concentration, whereas an increase in will result in a de- crease in CO concentration. The net affect of the increase of fire scale on

the concentration of CO thus is not as obvious as it appears to be.

2.11 CO FIRE TOXICITY

For the assessment of smoke toxicity without using animals, National Bureau of

Standards has proposed a N-gas model [11]. The N-gas model assumes that toxicity of fire smoke is determined mainly by a small number of gases; these may

act additively, synergistically , or antagonistically [11]. The model suggests

that all the animals exposed to fire products are expected to die when

271
raE C . / ( LC-- ) . > 1.0,

j-1 J ' J where C. is the concentration of the compound j (ppm); m is the number of

chemical compounds generated in the fire, and (LC^q)j is the concentr ation of

compound j producing death in 50? of the animals for a defined exposure period

(ppm). The suggested relationship has been found to be valid for mixtures of

CO and HCN with concentrations in the range of 25, 50 and 75? of their respective OC^q values [11].

Table 3 Average Yields of CO 2

. CO, and HCN

for Well-Ventilated Turbulent Flaming Fires of Materials 3

Material Surface Area Yj (g/g)

(nr) co 2 CO HCN

Methanol 9.68 1 .29 < 0.001 NA

2. 32 1 .30 < 0.001 NA

0.008 1 .32 < 0.001 NA

Rigid Polyurethane Foam 7 1 .50 0.027 0.01 0

0.008 1 .51 0.036 0.01 2 Polymethylmethacrylate 2.37 2. 1 1 0.008 NA

0.073 2.10 0.010 NA

0.008 2.15 0.01 1 NA

High Temperature Hydrocarbon 2.37 ND ND ND

Flui ds 0.008 2.69 0.019 NA

Heptane 0.920 2.83 0.0150 NA

0.059 2.92 0.0090 NA

0.910 2.82 0.0081 NA

0.008 2.89 0.0091 NA

Corrugated Paper Boxes with 53 1 .25 0.0085 NA

Shredded Paper (19? by Weight)

on Wood Pallets ( 38 ? by Weight)

Corrugated Paper Box with 0.060 1.12 0.0082 NA

Shredded Paper (25? by Weight)

Corrugated Paper Box with 0.060 1.25 0.01 9 NA

Shredded Paper. (62? by Weight)

Pine Wood Crib 0.062 1.33 0.005 NA

Douglas Fir 0.007 1.31 0.009 NA

Red Oak 0.007 1.27 0.009 NA

Cellulose Powder 0.007 1 .29 0.002 NA

a: Data taken from References 6. ND: Not determined.

NA : Not applicable.

In this paper, we have used the N-gas model for simplified application to CO

and CO

2 mixtures for diffusion flames of propane and propylene and wood cribs.

From Eq. (10), expressing Cqq in ppm,

c co

(ppnl) ‘W’co’ ( VV x 106 00

272
and for C0 2

,

cco

2

(ppm) (Y co

2

/p co

2

) <VV x 106 (,2)

Recently Levin et al [12] have developed a relationship for the toxicological

synergism between CO and C0 2 for hazardous environment

m c co /(c

co 2

- b) ^ ’-o • (13)

where m is the slope and b is the intercept of the curve showing toxicological

synergistic relationship between the concentration of C0 2 and CO. For C^q <_ 50,000 ppm, m = -28 and b = 117,000 ppm and for C rQ > 50,000 ppm, m = 1 50 ind

b = -

31 3,000 ppm [1 2] . 2

From Eqs (11), (12) and (13), condition for hazardous environment can be ex- pressed as follows

m (Yco

/(, co> Wx 106

> 1.0

(i co

2

/p co

2

) VV x 10 ‘ - 6

(IK)

Using p CQ = 1165 g/m^ and p co = 1842 g/m^ at 293°K and expressing left hand

side of Eq . (14) as toxicity parameter, T , the following relationships can be

derived:

1 ) For c co 2 1

50,000 ppm

- 0.2054 Y rn

T = ^

(15)

x 0.0046 Y cq

- (V

T

/G

f

)

2) For C CQ > 50,000 ppm

T __ °' ,n< Y C0

X 0.0017 Y cq

+ (V

T

/G

f

)

( 16 )

Figure 5 shows a plot of the toxicity parameter plotted against air-to-fuel

stoichiometric fraction for propane, propylene and wood crib, where data for

CO and C0 2 were measured in our apparatus [6]. The data in Figure 5 show that

if fire products are not diluted, hazardous conditions, where T

x

_> 1.0, are achieved for <f> < 1 . 8 for propane and propylene and <p < 0.8 for pine wood

cribs. If fire products are diluted, hazardous conditions change such as shown in Figure 6 for propane and pine wood cribs. The data show that hazardous conditions are not present in the fires of wood cribs and propane when

fire products are diluted by air by about ten and fifteen times their volume,

respectively, even if the yield of CO may be very high and fires may be highly

under ventilated.

273
2.5 CONCLUSION

From the analysis of the data presented in this paper, the following can be

concluded: 1) the yield of CO is weakly dependent on the fire scale for turbulent fires; 2) the yield of CO is strongly dependent on the chemical structure

of the materials; it increases as the structure changes from aliphatic to aro- matic to highly halogenated; 3) the yield of CO is strongly dependent on the

fire ventilation; it increases with decrease in ventilation. Values as high

as 0.30 have been found for the yield of CO in highly under-ventil ated fires;

4) combination of the data from National Bureau of Standards' fire toxicity

study on CO and C0 9 and our flammability study suggest that if CO and CO 2 were

the only gases of concern, the hazardous environments may not be present if

fire products are diluted by about ten and fifteen times their volume by air

for wood cribs and propane, r especti vely , even if fires are highly unde;

ventilated and Y^q values may be very. high. It is interesting to note that

fire products from other materials can also be diluted to eliminate hazardous

conditions, but the extent of dilution may depend on the generic nature of the

material. It may be possible to rank materials on the basis of extent of

dilution required to eliminate hazardous conditions.

ACKNOWLEDGEMENT

Financial Support for the study by the National Bureau of Standards, Center

for Fire Research under Grant No. 60NANB4D-0043 is deeply appreciated.

NOMENCLATURE

A Surface area of the material (m ) Cqq Concentration of CO (g/m^ or ppm)

f^Q Generation efficiency of CO (-)

G

'r

k C0

L

MW

n

Q"

Mass generation rate of material

vapors per unit surface area of

the material (g/m^s)

Mass generation rate of CO per

unit surface area of the material

(g/m 2s

)

Mass air-to-fuel stoichiometric

rate (g/g)

Maximum possible theoretical yield

of CO (g/g)

Heat of gasification of material

(kJ/g)

Molecular weight (g/mole)

Number of carbon atoms (-)

Heat flux per unit surface area of

the material (kW/m^)

Toxicity parameter (-)

t Time (s)

V„, Total volumetric flow rate of fire

T R product-air mixture (m-’/s)

W^q Total mass of CO generated (g)

Wf Total mass of material vapors

generated (g)

Y co Yield of CO (g/g)

Greek Symbols

6 Fraction of entrained air which

reacts with material vapors

p Density of compound j (g/m^)

j

Subscripts

f Material vapors

n Net

Superscr i pts

• Per unit of time (1/s)

" Per unit surface area ( 1 /m )
REFERENCES

1. Fire and Smoke: Understanding the Hazards. Committee on Fire Toxicology.

Board on Environmental Studies and Toxicology Commission on Life Sciences,

National Research Council, National Academy Press, Washington, DC 1986.

2. Tewarson, A., "The Effects of Fire-Exposed Electrical Wiring Systems on Escape Potential from Buildings. Part I - A Literature Review of Pyrolysis/Combustion Products and Toxicitus - Poly (Vinyl Chloride)," Factory

Mutual Research Corporation, Norwood, MA, Technical Report 22991, RC75-T97 , December ( 1 975 )

3. Hartzell, G.E., Priest, D.N., and Switzer, W.G., "Modeling of Toxicological Effects of Fire Gases: II. Mathematical Modeling of Intoxication of

Rats by Carbon Monoxide and Hydrogen Cyanide," J. Fire Sciences 3_, 115

(1985).

9. Kaplan, H.L., Grand, A.F., and Hartzell, G.E., "A Critical Review of the

State-of-the-Art of Combustion Toxicology," Southwest Research Institute,

San Antonio, TX, Final Report SWRI Project No. 01-6862, June 1982.

5. Tewarson, A., "Prediction of Fire Properties of Materials, Part I. Aliphatic and Aromatic Hydrocarbons and Related Polymers," National Bureau of

Standards, Gaithersburg , MD , Technical Report NBS-GCR-86-521 , December

1 986.

6. Tewarson, A., "Prediction of Fire Properties of Materials," National

Bureau of Standards, Gaithersburg , MD, Technical Report (to be issued).

7. Beyler, C.L., "Major Species Production by Solid Fuels in a Two-Layer Compartment Fire Environment , " Fire Safety Science, Proceedings of the First

International Symposium, Hemisphere Publishing Corporation, New York, NY,

p. 931 , 1 986.

8. Lomax, S. , and Simmons, R.F., "The Formation of Carbon Monoxide from Diffusion Flames," Fire Safety Science, Proceedings of the First International Symposium, Hemisphere Publishing Corporation, New York, NY, p. 991,

1986 .

9.

Morikawa, T. , "Effect of Supply Rate and Concentration of Oxygen and Fuel

Location on CO Evolution in Combustion," J. Fire Sciences 1 , 369, 1983.

10. Tewarson, A., "Fully Developed Enclosure Fires of Wood Cribs," Twentieth

Symposium (International) on Combustion , The Combustion Institute, Pittsburg, PA, p. 155, 1989.

11. Babrauskas, V., Levin, C., and Gann, R.G., "A New approach to Fire Toxicity Data for Hazard Evaluation," ASTM Standardization News, September

1986 . 12. Levin, B.C., Paabo, M., Gurman, J.L., Harris, S.E., and Braun, E., "Evidence of Toxicological Synergism Between Carbon Monoxide and Carbon Dioxide," Toxi cology (to be published).

275
Human Ptty*iotog*ai R#*pon&s to CO

10,000

B1~ 1,000 o

o

100

30

<R*f 243) (FW 95) (Ref 242) (Rsf 238) —Fens In —

- —tmmedlatg 1 mm —Fetal In « daoth tolhr

“

- —FcfttS In < ! hr

• Feta* in hr. Tofe-onc* —jr— (mm) | Mcrtskd poisoning ——Tblercnc*

• —Oeefh Inlhr IDangerous Ini hr. — —Daatti St2hr*. ~'f~ Unoteoaont hi hr (i hr) —Difficulty — NFPA,

i Slioht effect In 1 hr

"foiaronca of om&uMtion donasrous for 3 (lto2hr) C mm — Hdlucinoilon* m

'/t to 2 hr* *T~ No effect In i he Po(*ontnq in 1 to 2 hr* —— Hallucination m ^

to 2 hr*

(Rffif. 260) -j

• Pot*onn3

In 2 to 3 hr*. —Minimum farpc<*ontng

1

- —HaoOoctv* in 2 to 3 hr* —No tymptom* J —No pononmq —AJIowOBV* for ——Tbfaronc* (8hr».) -J

:

Saverol hr* •OSHA limit for j-J

" 8 hr txpotur* J (Ref. 261)

o

o

Figure 1. Summary of Results of Some Selected Studies on Human Physiological

Response to CO, Taken from Reference 2. Reference Numbers Identified in the Figure Appear in Reference 2.

Figure 2. Yield of CO as a Function of Molecular Weight for Well-Ventilated

Fires of Normal Alkanes, Alkenes and Alkynes and Arenea. Data Taken fro® Reference 5.

27 6
Maisculer Wsighi [g/mola]

Figure 3. Yield of CO as a Function of the Molecular Weight for WellVentilated Firea of Normal, Substituted and Cyclic Alkanes,

Alkenes, Dienes and Alkynes. Data Taken from Reference 5.

Alr-fo—Fu@l Stoichiometric Fraction

Figure 4. Yield of CO as a Function of Alr-to-Fuei Stoichioisetric Fraction. Data Taken frota References 6 and 10.

277
Toxicity

Parameter

[-]

Figure 5. Toxicological Effects Due to CO and CO2 Generated in Diffusion

Flames. No Downstream Dilution of Fire Products.

Figure 6. Toxicological Effects Due to CO and CO2 Generated in Diffusion

Flames. Varying Downstream Dilution of Fire Products.

278
RISK, HAZARD, AND EVACUATION (CONT.)

279
EXITT - A Simulation Model of Occupant

Decisions and Actions in Residential Fires

Bernard M. Levin

Center for Fire Research

National Bureau of Standards

Gaithersburg, Maryland 20899, USA

ABSTRACT

EXITT is a discrete event simulation of occupant decisions and actions in a simulated fire. Before the simulation starts, the characteristics of a residence, a fire in that residence, and the occupants of the residence are entered into the computer. Based on a large set of decision rules, the

occupants "make" decisions which are a function of the smoke conditions in

the building, the characteristics and status of the occupants (including

their capabilities), and the available travel routes. The occupants

investigate the fire, alert and assist others, and evacuate the building.

The simulation ends when all the occupants are either out of the building or are trapped by the smoke

.

1. INTRODUCTION

The EXITT model simulates occupant decisions and actions in fire emergencies

in -small residential buildings. In assigning decisions to an occupant, the

computer considers such factors as: age of occupant; sex; whether occupant

is asleep; smoke conditions; whether smoke detector is sounding and how loud

it sounds; whether occupant needs help in moving; and location,

capabilities, condition and status of the occupants. This version of the

model does not consider the heat or toxic components of the smoke. The

permitted actions include: investigate the fire; alert others; awake others; rescue others; and evacuate/escape. Actions not permitted in the current

version include: telephoning fire department from within the building;

fighting the fire; and re-entering the building to make a second rescue.

This paper describes the EXITT model, with emphasis on the rules used in

determining the simulated decisions and actions of the occupants. The

simulation is currently being developed as a series of progressively more sophisticated models. This paper describes an early interactive version of

the model.

A more extensive description of the model, including a users 's guide is in

preparation and should be published during the summer of 1987 [Levin, 1987],

Upon its publication, the computer program will be available from the author

and from the Center for Fire Research.

Imbedded in the decisions rules are parameters that can be easily changed.

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Developing improved values for these parameters is a major part of the

future development of the model.

This simulation program is designed to run on the IBM- PC computer (and

similar microcomputers using the MS-DOS operating system) using the BASICA

programming language. 1

The user controls the model through the keyboard by answering simple

questions that appear on the screen. (A "batch" version of the model, that

does not require or permit user interaction during the running of the model,

is also available in the FORTRAN programming language: the decision rules

in this paper apply also to the batch version.) The decisions of the

occupants are reported on the screen and on the printer. The movements of

the occupants are displayed graphically on the screen, reported on the

printer, and stored in a data file. The user can suppress outputs on the

printer and on the screen.

All the decision rules programmed in the computer model are designed to make

the decisions as similar as possible to those that building occupants would

make. The decision rules are based on: 1. a limited number of controlled

experiments; 2. case studies of occupant actions in residential fires; and

3. the judgment of the author. Whenever the rules are based directly on data in the literature or specific case studies, reference is made to such

data. Otherwise, the rules are based on the author's judgment.

The fires, buildings and occupants that are modelled can come from two

sources: 1. three buildings -- each with two or more fires and two or more

sets of occupants -- are stored in the computer programs and associated data

files; and 2. the user can substitute a new building, fire and/or set of

occupants by answering a large number of simple questions appearing on the screen.

2_ INPUT VARIABLES AND PARAMETERS

Building . The building is represented within the computer by nodes that

represent rooms, exits and secondary locations within rooms; and by links or distances between adjacent nodes. The major data used to define a building

are: the number of rooms, nodes and exits; the height of each room; the room location of each node; nature of each exit (door or window); and the

distances between adjacent nodes. Windows that cannot or would not be used

in a fire are not entered into the computer as exits, e.g., those with a window air conditioner installed.

Smoke . The program is designed to use the output of the FAST model- -or any

other similar model- -for distributing smoke throughout the building over

time [Jones, 1984], EXITT assumes a two layer smoke model. However, it is

1 The use of company names or trade names within this paper is made only

for the purpose of identifying those computer hardware or software products

with which the. compatibility of the programs of EXITT has been tested. Such

use does not constitute any endorsement of those products by the National

Bureau of Standards.

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assumed that a small proportion of the smoke in the upper layer gets into

the lower layer so that there is an odor of smoke in the lower layer. Exitt

accepts as input the smoke density in the upper layer and the depth or height of the two layers in each room at the beginning of each time period,

e.g., every five seconds.

The measure of optical density for which the model is being calibrated is

the one used by Jin in his studies of human behavior in smoke.

OD = In (L0 /L)

where L0 is the initial light intensity which reduced to a value of 1 over a path of one meter. This measure is consistent with the well recognized fact

that when people perceive a physical stimulus, the perceived intensity tends

to vary directly with the log of the physical intensity of the stimulus,

One important factor in making action choices in a residential fire is the

properties of the smoke in the occupant's room. A measure of the

psychological impact of smoke is determined as follows;

S-2*0D*D/H where,

S is the psychological impact of the smoke.

OD is the optical density of the smoke in the upper layer,

D is the depth of the upper layer, and

H is the height of the room.

This expression is based on the assumption that the impact varies directly

with the optical density (i.e., the log of the amount of smoke in the upper

layer) and with the depth of the upper layer relative to the height of the room. The formula is an arbitrary representation of this assumption.

Some of the decision rules and definitions that involve S include:

Occupants will not move to a node where S>0 . 5 unless the depth of the

lower layer (H-D) is at least 1.2 meters, i.e., unless the occupant can crawl under the smoke.

Occupants will not move lower layer is at least

Occupants will increase room where S>0 . 1

.

to a room where S>0 .

4

1.2 meters.

their travel speed by

unless the depth of the

30% after encountering a

Occupants will terminate an investigation if they are in a room where

S>0.05. They will terminate their investigation before entering a room where S>0 . 1

.

Once an occupant is in a room where S>0 . 1 , he will respond more quickly:

the changes are described below as consequences of believing the fire to

be serious.

When S>0 . 4 there are prohibitions and penalties: these are described

below as consequences of encountering "bad smoke.”

Each of the above mentioned thresholds is an input parameter and can be

easily changed as we continually improve the calibration of the model, i.e.,

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modify the model to better correspond to behavior in real fire emergencies.

Although the values selected are consistent with a conservative

interpretation of Jin's data [1976], these values will be reconsidered as part of the further development of the model.

Noise and Alarm . The background noise level in a room affects the ability

of an occupant to hear the alarm, both in real fires and in the model. The

background noise is preset at 35 decibels. The user can easily override this

value, on a room by room basis, by entering a larger value when defining the

fire scenario

Another input, related to a specific fire scenario, is the loudness of each

smoke detector in each room, including the room in which it is located- -it

is a function of distances and of which doors are open. The impact of the

alarm is a function of the difference between the volume of the alarm and

the background noise.

Characteristics of the Occupants . The user has an opportunity to either

define or modify the characteristics of the occupants through the keyboard.

The characteristics are: age, sex, normal travel speed, whether or not the

occupant needs help in evacuation, whether or not the person is awake, room

location, and, if the occupant is asleep, a measure of how difficult it is

for the occupant to awaken.

There are a number of additional parameters imbedded in the decision rules

which are described below. These include: the age below which a child is

considered as a baby unable to initiate any action; and the times required

to perform various actions, such as waking a sleeping adult occupant when

the fire does not appear to be serious. Some of these, additional parameters

are described in Section 5, Delays, Pauses and Action Times.

3. DECISION RULES

Introduction . There are two types of occupants: those who are fully capable

when awake and those who need assistance in moving. The decision rules

apply only to those who are capable when awake. Those who need assistance

moving make no decisions and their movements are determined by their

"rescuer"

At the beginning of the simulation, all occupants are unaware of the fire

and the potential danger. Actions and decisions are assigned, in part,

based. on the smoke conditions in each room at the beginning of the

appropriate time period.

The following paragraphs describe the sequential steps the computer follows

in determining the decisions and actions of one occupant for one time

period. The computer goes through these steps for each capable occupant for

the first time period and then repeats the process for each subsequent time

period, in turn, until all the occupants are either out of the building or trapped by the fire. (For each step, the computer considers all occupants

before proceeding to the next step.)

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Aware of Fire . The first step in determining the actions of an occupant is

to determine if and when an occupant is sufficiently aware of the fire cues

(i.e., smoke, sound of alarm and visible flame) to undertake an action. If

the occupant became aware of the fire cues in a previous time period, he

will remain aware of the fire cues for this and all subsequent time periods.

An occupant becomes aware of the fire when the fire cues are sufficiently

strong. Obviously, stronger cues are needed to awaken and alert a sleeping

occupant than to alert an awake occupant. The fire cues are: the sound of

the smoke detector; the odor of smoke; and, for awake occupants, visible

smoke and visible flame. If the weighted sum of the intensities of the cues reach a prescribed threshold, the occupant will be flagged as being aware of

the fire cues. (If the fire cues are of borderline intensity, the occupant

will become awake and aware after an assigned delay.) If the fire cues are not sufficiently strong for the occupant to become aware of the fire during

the current time period, the consideration of this occupant for the time - period is completed.

The following basic equation, for determining if and when an occupant will

start to respond to the fire cues, was suggested by the empirical results of

Nober et al.[1981]. While Nober studied only the response of the smoke

detector alarm, his results were generalized for the odor of smoke, and the

sight of smoke

.

T = 70 - 4(C-20) and

C = (A-N)+X1+X2+X3+X4 where,

T is the delay time, in seconds, before the occupant will start his first

action

;

C is the sum of the sensory impacts on the occupant;

A is the sound intensity of the smoke detector as heard by the occupant

N is the background noise;

XI is impact of an awake occupant seeing flame. It is set sufficiently

high to assure a rapid response whenever an awake occupant is in ' the same room as a visible flame and set equal to zero if the occupant is asleep;

X2 is impact of an occupant smelling smoke- -it is a function of the smoke

density and smoke depth and applies to both sleeping and awake occupants.

It varies directly with S, the psychological impact of smoke when the

smoke remains above 1.2 meters. However, its value dramatically

increases when the upper smoke level gets down to the height of a person

in a bed;

X3 is impact of an awake occupant seeing smoke- -it is a function of the

product of the smoke density and smoke depth in the upper layer and also

varies directly with S; and

X4=0 if the (typical) occupant is asleep and X4=15 if the occupant is

awake. This reflects the fact that more stimuli are require to awake

than to alert an occupant. The value of 15 is based on the data in Nober

[Nober et al , 1981], Occupants who have difficulty waking could be

assigned negative values of X4

.

Subject to the restrictions:

A-N cannot be less than zero. If N>A let A-N = 0;

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If C < 20 then T = infinity (i.e., 99999 in the computer). This

restriction is based on Nober's data where occupants usually either

responded within 70 seconds or remained asleep for the remainder of the

test period.

XI and X3 equal zero if the occupant is asleep.

The model as described above assumes that the response is a function of the sum of the impacts of different sensory cues . This assumes that the

relevant aspects of the perceptual processing of olfactory, visual and

auditory cues are similar. There does seem to be a surprisingly consistent

perceptual rule (Fechner's Law) that the intensity of a perception varies

directly with the log of the physical stimulus. (While this rule has broad

applicability, it is not universal and only approximate [Boring, 1950].)

Since our measure of the psychological impact of all the cues are based on

the log of a physical measure, the impacts to the three types of cues can be

assumed to be roughly comparable. The decision to sum the impacts of the

three cues is based on the assumption that simple behavioral rules are better than complicated ones when there is no technical reason to select a complicated one. Furthermore, simple summing is consistent with the results

of Fletcher and Munson who found that a tone heard binaurally seems twice as loud as the same tone heard monaurally [Licklider, 1951].

The physical measure of each fire cue is measured in different units and

they must be converted to a single measure of sensory impact. The cue most

easy to quantify, and the one for which we have the most data, is the sound

of the smoke alarm which is measured in decibels. It was decided to use "equivalent decibels" as the single measure of sensory impact. The impacts

of the other sensory cues are "converted" to equivalent decibels, i.e., the

values of X2 and X3 are transformed to the number of decibels that would

approximate an equivalent impact in alerting occupants . (When the flame is

visible to an awake occupant, the value of XI is set sufficiently high to ensure a rapid response.) The transformation factors are input parameters:

their values will be the subject of future research and analysis.

While occupants respond more quickly to strong fire cues, there is a minimum

duration of time required to awaken or become aware of the fire cues , select

an action, and perform preparatory actions. These minimum times range from

1 to 10 seconds depending on whether the occupant is asleep and the amount

of smoke

For each occupant, a time to start his actions is computed independently at

each time period until the occupant starts his first action: that is, a different time to start his actions will be computed each time period. He

will start his action at the earliest time among those computed.

Assigning Actions to Occupants Who Are Aware of Fire . If an occupant has

been assigned an action in a previous time period, he will be given an opportunity to complete that action before any consideration is made

regarding additional actions.

Investigation Top Priority . The normal first action is to investigate the

fire cues to determine the nature of the hazard. However, there are a number of exceptions, i.e., situations that would make investigation either

285
a' lower priority or an unreasonable choice. [Levin, 1985] If the computer

determines that none of these exceptions applies, the computer assigns, the room with the most smoke as the GOAL, labels the occupant as investigating,

and assigns him the task of going to the room with the most smoke.

One special situation that would cause investigation to be a low priority

arises when an adult female occupant is in the same building as a baby- -in a case study provided by Keating and Loftus a mother rescued her baby before

determining if it was necessary. [Keating and Loftus, 1984]

Investigation is not permitted: if the occupant has already completed an investigation; if the occupant has been in a room with moderate or bad

smoke; or if the occupant has been awakened or alerted by an occupant for

whom investigation is not permitted.

If the exceptions do apply, the computer considers the following alternative

actions in the order given below.

Help Occupant in Same Room The computer determines if there is another

occupant in the same room who needs help. If that occupant is fully

capable but asleep, he will be awakened. If he needs assistance moving,

he will get that assistance. (If more than one occupant qualifies for

help, the sleeping occupant is given priority.)

Help Occupant in Different Room . If there is one or more persons in a different room(s) who needs to be alerted, rescued or awakened, the

computer will make two assignments: tentatively assign the fully capable

occupant a person to alert, rescue or awaken; and assign the capable

occupant the action of going to the room of that person . (Once he

arrives at that room, a new action will be assigned based on fire

situation at that time and the capabilities of the persons in the room. Exception: if he is going to the room of an awake and capable person who

needs to be alerted, he will automatically alert that occupant.) The

priority order of these tentative assignments for helping persons in

• different rooms is: alert capable adult; rescue other occupant; wake

other occupant; and alert child.

Investigate . If investigation is still a permitted choice, the occupant

is assigned the task of investigating. The computer assigns the room with the most smoke as the GOAL, labels the occupant as investigating,

and assigns him the task of going to the room with the most smoke.

Egress . If none of the above alternative assignments apply, the occupant

is assigned the action of evacuating.

Every capable occupant is considered for helping an occupant in the same room before any occupant is considered for helping an occupant in a different room.

In this version of the model, an occupant over the age of ten functions as an adult, that is, they follow the priority list presented above for adults.

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A child who is 8,9, or 10 will rescue any occupant in the same room and will

go to another room to awaken or alert another occupant but will not go to

another room for the purpose of rescuing an occupant. Children 7 and

younger do not assist others out of the building but will wake or alert

other occupants who are older. A child is considered to be a baby if his

age is equal to or less than the parameter BABY: babies do not initiate any

actions. The tentative value of BABY is 3.

The general rationale for the above priority order is to determine if there

is a need for positive action, to assist those known to require help, and

then to assist those who might require help. (When it is unknown if a sleeping occupant has awakened, he can be considered as someone who might

require help.) 'If two people are known to require help, provide help to the one needing more limited help, that is, a sleeping but otherwise capable

occupant. The rule that supercedes all others is to help someone in the same room before helping an occupant in different room.

4. TRAVEL WITHIN THE BUILDING

Occupants move within the building from node to node. The path assigned is

determined by finding the shortest path to an exit based on a shortest path

algorithm which contains an option for assigning penalties for going through

bad smoke and for leaving through windows

The route to be taken to a designated room, the best exit, or another

designated node is determined by a shortest path algorithm. Normally when

the occupant is investigating or going to assist another occupant, a straightforward shortest distance is determined.

If the occupant wishes to egress or evacuate the building, or if the

occupant had encountered too much smoke when going to assist another

occupant, then the path with the lowest number of demerits is selected.

Each meter of travel is assigned one demerit, leaving by a window is

assigned 100 demerits, and going to a node through bad smoke is assigned 200

demerits. If the smoke at a node is intolerable, that node cannot be part

of a route. If smoke is blocking all routes to the designated node, the

occupant will decide to escape. If the occupant is escaping, the route out

of the building with the least demerits is selected. If all escape routes

are also blocked, he will be considered trapped. (See section on Smoke in

Section 2 for decision rules regarding moving or not moving through bad

smoke .

)

As an occupant attempts to move, whenever he encounters an intolerable

amount of smoke based on the criteria in Section 2, he will stop moving and

the computer will redetermine his best route to his destination, i.e., the

route with the fewest demerits. If the shortest path algorithm fails to

find an acceptable path, the computer will look for the best route out of

the building. If all exit routes are blocked by smoke, the occupant is

considered trapped.

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5. DELAYS, PAUSES AND ACTION TIMES

Introduction - The time consuming activities of an occupant can be

classified into three categories.

Movement from one node to another. He travels the shortest path at the

speed defined below.

Delays and pauses. These activities include time to awaken, time to make

decisions, and time to prepare for action.

Assisting actions, i.e., waking another occupant and preparing another

occupant for egress.

Soeed . The travel speed of each occupant is set at:

1.3 meters per second for normal conditions;

1.69 m/s (30% faster than normal) if an occupant should consider the fire

to be serious (e.g., he has been in a room with heavy smoke);

0.65 m/s (50% of normal) if the occupant is assisting another occupant,

or 0.845 m/s if the occupant also considers the fire to be serious;

0.78 m/s (60% of normal) if the smoke is bad (i.e., S>0.4) and if the

depth of the lower layer is less than 1.5 meters, i.e. if the occupant

has to "crawl 1 ' under the smoke. (0.52 m/s if the occupant is also

assisting another occupant.)

The normal travel speed and all the modification factors are parameters that can be set by the user. While these values will be reconsidered as part of

further model development, they are similar to those reported by Jin [1976],

Whenever an occupant moves, his actions are printed on the screen and on the

printer, recorded in a data file, and graphically represented on the screen. (Printing on the screen or on the printer can be suppressed.)

Delav Times . The delay time, the decision time, and the time to perform

assisting actions (hereafter, collectively called Delay Times) depend on the

occupant characteristics, the fire characteristics and the impact of the

fire cues on the occupant. The length of these Delay Times are determined

by a set of decision rules as described below. These Delay Times can be

changed rather easily and all assigned values should be considered as tentative

.

Minimum Response Time . The normal (i.e., smoke is not bad) minimum response

(delay) time is 6 seconds for awake occupants: this includes decision and

preparation time. The normal minimum response time for sleeping occupants

is 10 seconds: this also includes decision and response times. These

values are based on the work of Nober [Nober et al , 1981] . The status of

sleeping occupants is changed to awake status whenever the remaining

response time is 6 seconds or less.

TPAUSE . An occupant is assigned normally a delay time of TPAUSE seconds

whenever: he completes his investigation or terminates his movement along a route because of intolerable smoke; or changes his mind about helping

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another occupant. This delay includes the time required to choose a new

action. TPAUSE is tentatively set equal to 3 seconds.

Decrease in Preparation Time Due to Heavy Smoke . When an occupant is

subjected to normal fire stimuli, a ten second response Delay Time is

assigned to a sleeping occupant and six seconds to an awake occupant. (Note

the response time will be greater if the fire stimuli are not sufficiently

strong for an immediate response.) However, if the occupant believes the

fire to be serious, the maximum Delay Time for the occupant becomes four

seconds and if the smoke in the room is bad, the maximum Delay Time becomes

1 second.

Hesitation Due to Not Being Alone . Research by Latane and Darley [Latane

and Darley, 1968] has shown that when the fire cues are noticed but not

immediately compelling, adults will hesitate in their responses if other

capable adults are in the same room. A simple explanation is that there is

a failure to respond due to a feeling of shared responsibility. The

computer program accounts for this by delaying responses by one time period

for each time period where: there is no one that needs to be rescued,

alerted, or roused; there is a second capable occupant in the room; the

smoke detectors are not sounding; and the sum of the psychological impacts

of the fire cues is less than 30, i.e.

,

if C<30 then T=99999 , where C and T are defined in Section 2. (The threshold for this hesitancy is 30 rather

than 20 to reflect that more stimuli or cues than the "minimum for reponse"

are required to prevent the "hesitation due to not being alone.")

Time Required to Alert. Wake or Prepare for Evacuation . Whenever one occupant assists another, time for providing or receiving the service must

be assigned. The following times are assigned:

If Occupant J is alerting a fully capable and awake adult, he moves to

the node of the other occupant. Once he arrives at that node, Occupant J

starts his next action with no delay or decision time charged. The

occupant being alerted is assigned a Delay Time of five seconds or 2.5

seconds depending on whether the alerting occupant believes the fire to

be serious

There are two types of assistance that an occupant may be flagged as needing: waking; and help moving. If an occupant is asleep and does not

need help moving, the delay is five seconds for the occupant doing the

waking. For the occupant who is being awakened the delay is ten seconds - -5 seconds for waking plus 5 seconds decision and preparation time.

However, if the assisting occupant believes the fire to be serious, then

his time devoted to waking would be only 2.5 seconds and the total Delay

Time for the previously sleeping occupant would be 5 seconds.

If an occupant needs help moving, the Delay Time (at the time the

assisting occupant arrives at the location of the other occupant) is

usually 10 seconds if the disabled occupant is awake and 12 seconds if he

is asleep. However, if the disabled occupant is a baby, the Delay Time,

in seconds, is the baby's age plus 4. (It does not take long to pick up

a baby and wrap him or her in a blanket.) In addition, if the capable

occupant believes the fire to be serious, the previously determined Delay

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Time is halved. For example, if the fire is believed to be serious, the

Delay Time for helping a 2 year old baby would be 3 seconds ((2+4)/2).

5 . SMOKE DETECTORS

The building may have up to three smoke detectors. These smoke detectors

are independent and are not interconnected in any way. It is necessary to

provide the locations of the smoke detectors and how loud each detector

would sound in each room of the building. A smoke detector will sound if

the smoke density of the upper layer smoke is at least .015, and the depth

of the upper layer is .15 meters or greater.

An option in the program is to consider the smoke detectors as broken. It

is, therefore, easy to determine the effect of smoke detectors by running

the program twice with the same fire and occupants, once with smoke

detectors working and once with them broken.

6. CURRENT LIMITATIONS AND FUTURE DEVELOPMENT

The model as described in this paper is a preliminary version of a model

under development. The development, improvement and expansion of the model

is a continuing activity. The user should be aware of the limitations of

the current model which will be distributed in the summer of 1987. These

limitations include:

1. The model is deterministic. Only typical behavior is modelled:

aberrant behavior is not permitted.

2. Improved calibration is required to upgrade its validity.

Calibration means changing parameter values in the model based on: a. analyses of data in the technical literature; b. judgments of a panel of

experts; c. analyses of in depth interviews of survivors of residential

fires; and, d. attempts to simulate behavior in real fire emergencies for

which we have information. An intensive effort to improve the

calibration of the model, is scheduled for the next year.

3 Some typical actions are not included, e.g, fighting the fire.

4. Occupants respond to smoke conditions but not to heat conditions.

The current development program is designed to overcome and eliminate these

limitations. The program can be made probabilistic, and heat will be added

as a factor in decision making and route selection.

The model permits the user at the keyboard to override a very limited set of

the occupant decisions assigned by the computer. Within the next few

months, the user will be given the opportunity to override most, if not all,

of these assigned decisions. This will permit the model to be used in

studying the effect on safety of alternative decisions by the occupants.

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REFERENCES

Boring, E.G., A History of Experimental Psychology, Appleton- Century- Crofts

,

Inc. New York, 1950, p.289.

Jin, T., Visibility through Fire Smoke (Part 5 Allowable Smoke Density for

Escape from Fire, Report of Fire Research Institute of Japan, No. 42,

September 1976.

Jones, W.W. , A Model for the Transport of Fire, Smoke and Toxic Gases

(FAST), NBSIR 84-2934, National Bureau of Standards, Gaithersburg, MD , 1984.

Keating, J.P. and Loftus , E.F., Post Fire Interviews : Development and Field

Validation of the Behavioral Sequence Interview Technique, National Bureau

of Standards, Report Number GCR-84-477, Gaithersburg, Md. , 1984. Interview

records provided by the authors in private communication.

Levin, B.M. , Design as a Function of Responses to Fire Cues, General

Proceedings of Conference Research and Design 85: Architectural Application

of Design and Technology Research held at Los Angeles, California, March 14-

18, 1985, American Institute of Architects Foundation, Washington, D.C.,

1985, pp. 289-294.

Levin, B.M., EXITT--A Simulation Model of Occupant Decisions and Actions in

Residential Fires: Users Guide and Program Description, NBSIR 87-3591,

National Bureau of Standards, Gaithersburg, MD , 1987.

Latane , B., and Darley, J.M., Group Inhibition of Bystander Intervention in

Emergencies, Journal of Personality Psychology, Vol.10, No . 3

,

Nov. 1968,

pp. 215-221.

Licklider, J.C.R., Basic Correlates of the Auditory Stimulus, in S.S.

Stevens, Handbook of Experimental Psychology, John Wiley & Sons, Inc, New

York, 1951, p.1031. '

Nober, E.H., et al , Waking Effectiveness of Household Smoke and Fire

Detection Devices, Fire Journal, Vol.75, No . 4

,

July 1981, pp . 86 - 91 , 130

291
EXITT - A SIMULATION MODEL OF OCCUPANT DECISIONS AND ACTIONS IN RESIDENTIAL

FIRES

B.M. Levin, National Bureau of Standards, USA

HALL: As was mentioned earlier in the week, this is not only an important

model in its's own right, but this is an important aspect of HAZARD 1.

JIN: I have two questions. You said that any person, not necessarily an expert on fire, could be a user of this model. Could you be more specific on that? Could just an ordinary individual use it? My question is when you said

ordinary people, can you define whether it would apply, for example, mostly to

designers of new buildings? Would they find this useful in determining

whether to build a building a certain way because of certain circumstances of

people, or would it be for existing buildings where people who are concerned

with fire safety could use it as a means of analyzing the safety level of the

building? Is it mostly intended for the ordinary public, as a public

education system method concerning fire evacuation, or would it be for the use of firemen to educate the public in evacuating their own homes?

These are all possibilities, but where is the focus of your building this

model?

LEVIN: First, it's a simple enough model so that it can be used by the

general public. My son took it upon himself to write instructions so that a person who knows nothing about a computer can use it, even to the point that once you put a disc in the machine you must close the computer door.

In terms of the professional use of the model, I do not visualize it being

used very much by people who are interested in design. It is being used now

in the HAZARD model that Bud Nelson described in Mr. Bukowski's paper, but I think of it being used a lot in public education. However, I think it may be more useful for helping the people design the programs for the public. For

example, when I first joined the Center for Fire Research, I was told that it was safe to sleep only with the doors to the bedrooms closed. However, most

Americans will not sleep with their bedroom door closed if they have an infant

child in another room. I think my model can be used to analyze this situation

in a more professional manner than has been done.

I think the model has another use in addition to your list, in helping us determine where we should put our efforts in learning more about human

behavior and fire. The gaps in our knowledge become very apparent when

working with a model.

JIN: In this particular model you talk primarily about the action of the

people. What I would like to point out is that the action of people is

variable. People tend to do things differently under similar circumstances.

Professor Keating's data must have shown through an accumulation of the

tremendous volume of data, an average action of a person. This is not to say

that everybody that he interviewed acted the same way under the given

circumstances. In actuality, given the same s ituation ,_ perhaps a group of

people A would act 60 percent one way, B group would act 20 percent another

way, and C group would act 10 percent, and so on. I have long studied this

problem of human behavior under this type of stress, and I believe that

292
statistical input also is very necessary. As Mr. Nelson said yesterday,

statistical input is very valuable. Perhaps you may input this into your

modeling in the future.

LEVIN: Of course. I do plan to put in the probabilities as you suggested,

recognizing that it is a difficult thing to do in terms of developing what the

probabilities should be. On the other hand, I found something very

interesting in analyzing Professor Keating's raw data. I've looked at the

data of Dr. Woods from England, from Professor Bryan here in the United

States, from Keating, and they always point out that the first action, a percentage of the time it's investigation, but then they list a lot of other

actions that occur as a first action.

Professor Keating's raw data gave me an opportunity to look at this phenomena

in more depth. His raw data are in-depth interviews where he asked people not

only what they did, but what they tried to accomplish. From this data, I was able to infer how much people knew about the fire when they first became aware

that there was something wrong.

When I looked at the first response of people who knew something was wrong,

but did not know the specifics of the fire, I found they almost invariably,

over 90 percent of the time, investigated. One of the exceptions was a mother

who went to rescue her baby, not knowing whether there was a need to rescue her or not. In my model, mothers rescue their infant babies before they

investigate. In Professor Keating's data all those who woke up in a room with

flames, of course, did not investigate, but if we were analyzing the data the

way the data had previously been analyzed, they would just have been

classified as somebody who did not investigate.

You're, of course, right that we must worry about probabilities, but I feel

less uncomfortable due to my analysis of Professor Keating's data.

293
EMOTIONAL INSTABILITY IN FIRE SMOKE

( Experiment of human behavior in smoke filled corridor )

Tadahisa JIN. Tokiyoshi YAM ADA

Fire Research Institute

Fire Defense Agency

Ministry of Home Affairs

3-14-1 Nakahara , Mi taka-sh i , Tokyo 1S1

ABSTRACT

An experimental study was conducted to estimate smoke and

heat effects on an emotional instability under smoke condition.

Mental arithmetic and walking speed are used as index of the

instability and moreover a feeling on smoke and heat was

investigated by a questionnaire. From the experiment, high

adaptability of humane to smoke condition was recognized in

thinner smoke, and the heat of smoke affects a thinking ability

down over about 2,000 (kcal/m2 h) was found.

1 . BACKGROUND

A smoke obstruction against safety evacuation has been

pointed out through many fires in the past. A density of CO gas

and lack of 0 2 in fire smoke are physiological limited factors

to decide life or death. However physical and psychological

effects of thinner smoke, a decrease of visibility through smoke

and/or smoke irritation for instance, tend to give evacuees

panics and indirectly lead to death.

Many experimental studies on such smoke effects have been

conducted in our laboratory with the smoke of normal

temperature. In these experiments, hot smoke effect on human

were not considered, so an additional experiment was conducted

to estimate a complex effect of heat and smoke with electric

heat radiator instead of hot smoke.

About 30 subjects were exposed to a smoke environment and

in each experiment, the physiological and psychological effect

of smoke and heat radiation was investigated by mental

arithmetics and a questionnaire.

2. METHOD AND CONDITION OF EXPERIMENT

31 subjects participated in this experiment. The

experiment was conducted in the smoke filled corridor. Each

subject was imposed to walk forward in this corridor answering

the arithmetic questions, simple additional and subtractive

calculations, as show in Fig. 1.

294
2.1. EXPERIMENTAL CONDITION OF THE CORRIDOR

The experiment was conducted in the straight corridor of

about II m depth, 2.5 m height, 1.2 m width. On its floor,

five mats, in which an on-off switch was equipped, were set to

know the stand position of subjects in a smoke environment. At

each of these five points, subjects stopped to answer arithmetic

questions as mentioned later.

The corridor was filled with white smoke generated by

smoldering japanese cedar crib chips in electric furnace before

starting each experiment. At the beginning of the experiment,

a smoke density expressed by the extinction coefficient was

adjusted to 1.2 (1/m) and weaken slowly to the 0.1 (1/m) during

30 minutes. In this 30 minutes duration, two or three subjects

individually entered into the corridor by turns under some kinds

of smoke density levels.

A temperature inside the corridor was about 20 °C. At the

inner position of A and B, the heat was radiated to subjects

with electric radiator ( at A -->12 Kw, 3 -->3 Kw) set at a ceiling.

A maximum heat flux density was 2,030 (Kcal/m h) at A

point and 1,370 (Kcal/m2 h) at B-point of the 1.5 m height from

the floor level. Mean Radiant Temperature measured with glove

thermometer levels about 82 °C -at A-point and 75 °C at B-point.

Inside the corridor was illuminated by fluorescent lamps at

some points, and ‘illumination levels in the range of 30 to 100

(lx) on the floor.

2.2 MENTAL ARITHMETIC

A mental arithmetic method was adopted to estimate a degree

of emotional instability under smoke condition in this

experiment. A rate of correct answer to simple arithmetic

questions was expected to decrease as Increasing the human

emotional instability in certain degree. So this rate was

selected as an index of emotional instability. The arithmetic

questions were informed to each subject by a loud speaker to

hear at everywhere in the corridor.

The mental arithmetic questions were transcribed with an endless tape and 10 questions were continuously given at each of

five answer positions (A to E). The three kinds of question

interval speed, 1.5 (sec/question) for male and 2.0 to 2.5

(sec/question) for female were selected to correct the

individual difference.

This set of 10 questions was repeatedly given at intervals

of 10 sec. Subject was imposed to answer at each five point and

to move forward at next deeper position for this 10 sec.

interval and again stop to answer the questions at new point.

The answers were voice recorded with a handy type tape recorder

and was rated the answers later.

These data are normalized by the correct rate of 90

questions under normal condition, (no smoke and heat radiation),

for there are self differences of calculation ability.

295
2.3 SUBJECTS

31 adults aged 20 to 51 (male 14, female 17) participated

in this experiment. These females were mainly house wives and

males were personnels of F.R.I.

Before the experiment, no detail informations of the

corridor were not given and the corridor state without smoke was

not displayed. Only fundamental instructions of experiment and

notices for safety experiment were given,i.e.,a corridor filled

with smoke is strait and has an end , subjects are allowed to

return back whenever they want to break.

The subjects were protected from irritating smoke with 15

times piled towel on their mouth and nose. About 90 % of soot

are removed by this towel as investigated in the past.il]

3. RESULT AND CONSIDERATION

The Fig. 3 shows how the correct rate of answer to

arithmetic questions changes as the subjects go forward in the

corridor. The rate at A-point decreases at less than 10 % of

those of other position even in case of no smoke. It may

indicate the effect of heat radiation to emotional instability,

however 'when the heat radiation is small, such as 3-point no decrease is not found.

The walking speed between each answer position is measured

in each experiment .with mat switches as reference data

represents the emotional instability. The followings are

further considerations of this walking speed and the correct

rates from the view point of the emotional instability.

3.1. INTERRUPTION OF EXPERIMENT

At the beginning of each experiment, a smoke density was

not necessarily of the same level. The average values were in

the rage of Cs = 0.92 ± 0.21 (1/m). At first time experiment

trial, 17 subjects (male 5, female 11) could reach the deepest A

-point, and other 14 subjects (male 8, female 6) returned back

before reaching to the A-point.

Fig. 3 shows a relation between attainable distance and the

normalized correct answer rate to arithmetic questions of the

latter 14 subjects. Seven subjects abandoned to walk forward in

corridor at the nearest point of E (2.8 m from entrance) and the

correct rate drops to about seven percent of normal correct

rate. The six subjects of these seven subjects broke at the

second trial under condition of Cs = 0.66 (1/m) and moreover two

of them did at the third trial (Cs = 0.62).

At the first trial in thick smoke, three subjects returned

back at D-point (5.3 m distance from entrance), other three at C

-point (6.9 m), one at B-point (8.6 m), and the each correct

answer rate was about 35 %, 55 % and 75 %. The correct rate

decreases as the subjects returned back earlier, so this result

indicated the emotional instability is larger when the

subjects returned back earlier.

296
3.2.

CORRECT ANSWER RATE AND EMOTIONAL INSTABILITY

Some subjects could reach the end of corridor in the first

trial under thicker smoke. Fig. 3 shows the relation between

the attainable distance and the correct answer rate. This

average rate of 17 subjects levels at the lowest just after

entering the corridor (at E-point) and it tends to rise up at

deeper position. However, in the inquiries conducted after

experiment, the subjects indicated that they felt more unrest as going f o rward

These results indicate that the low correct answer rate

does not necessarily explain the instability for this type of

subjects. The cause of discrepant results seems to be led by

an observatory fact, subjects assimilate the smoke condition.

From the view point of physiological aspect, an abrupt

encounter against irritating smoke attacks subjects’ eyes and/or

throats, and disturb to open their eyes. This physiological

discomfort lets the correct rates fall down, however the

discomfort was lighten with the lapse of time.

The assimilation factor compensates the correct rate

physiologically but can not recover the emotional instability

as the same degree in psychological aspects. So the rate rises

up as seen in Fig. 3

3.3.

HEAT FLUX EFFECT ON CORRECT ANSWER RATE

In Fig. 3, included the mental arithmetic results under no smoke condition with dot line. The data at A-point drop 10 %

lower than those of after points. This decrease is considered

to be caused by the heat radiation from an, electric heater set

at the ceiling.

The same drops at A-point was also apparently recognized

under smoke condition presented with normal line. Nevertheless,

no drops are not found at B-?oint.

The maximum heat flux intensity was 2,030 (Xcal/m'h) at 1.5

m height from floor of A-point and 1,370 (Kcal/mh) at 3-point.

So the limit value of heat flux effect on a thinking ability is

within the range of between 2,030 and 1,370 (Kcal/ ra

J h).

3.4.

SMOKE DENSITY EFFECT ON CORRECT ANSWER RATE

Fig. 4 shows how the smoke density affects the correct

answer rate. It indicates that the correct rate decreases as increasing the smoke density almost linearly. This relation has

an agreement with the other experimental results conducted in

the pas t . [2]

The assimilation effect is recognized in this figure

clearly. The dot line shows the second trial results of the

same subjects. The correct rates get about 30 % better than

those of the first trial. It concluded that the assimilation

depend on the number of experiment trial and the exposed time

duration in smoke condition.

297
3.5. WALKING SPEED IN FIRE SMOKE

The walking speed between each point was measured in this

experiment. Fig. 5 shows the results. The walking speed goes

down in proportion to the distance from the entrance. The

variance between male and female data are not distinct in

absolute walking speed, however the declining rate of female in

the deeper point is higher.

The smoke density effect on the walking speed was shown in

Fig. 6. It was almost the same result of the correct answer

rate, that is, the walking speed goes down as increasing smoke

density. The drop rate of walking speed is lower than the other

former experimental result. C 3 1 This may be caused by

experimental conditions and methods' difference, especially the

use of towel to protect fire smoke are important. About 90 % of

soot in smoke is filtered out with this towel and it lessens the

smoke irritation effect of smoke.

4. INQUIRY RESULT

Some questions were given to each subject after

experiments. One of them is question on discomfortable factor

of smoke: "What are the discomfortable factors of smoke ? Select

three factors concerned with fire smoke from following items in

order !” The Table* 1 shows the result.

Some differences were found between male and female. For

the male, the physiological factors are mainly selected. The

irritation to eyes and/or throats is the first or second

selected item, "The obstruction of inhalation" is selected as second or third discomfortable factor. As, compared with the

above male's answer, females tend to feel discomfort due to

psychological factor.

As the first cause of discomfort, the obstruction of

visibility is selected as well as. the irritation and inhalation

problem. And as the second factor, this obstruction of

visibility is aiso cited. This experimental result indicates

that the emotional instability in thick fire smoke is not

necessarily caused by the same factor for male and female.

Under thin smoke condition, without distinction of sex, the

smoke irritation and heat flux are selected as first or second

factor.

The table 2 shows the inquiry result of the feeling on radiation heat flux. At the B-point, the power of electric

heater (about 3 Kw ) is so small to effect on subjects

discomfort. The answer was mainly "just fell heat a little",

the other point (point-A with 12 Kw heater), 44

answered "hot but not intolerable" under smoke

however this rate increase to 50 % under no smoke

Moreover the 7 % subjects answered "intolerably hot"

condition, and the 29 % under no smoke condition.

The intensity of heat flux on subjects is almost the same

between under smoke of Cs = 1.0(l/m) and no smoke condition in

this experiment. So the existence of smoke itself may

psychologically affect the felling on heat flux.

% subjects

condi t i o n

,

condi t i o n

.

under smoke

298
5. CONCLUSION REMARKS

From this experiment in the corridor, the following remarks

are ob tai ned

1. The emotional instability caused by physiological

factor, such as smoke irritation, is more important

than that by psychological unrest under high irritant

smoke

2. The degree of suffering from fire smoke decreases as time goes by. The assimilation to smoke condition is

recognized in thin smoke, so condition of the moment,

when evacuee first encounter aga i ns t fire smoke , i s very

smoke

important. However, the

occurs human unrests.

long time s t ay 1 n the

o

o . The walking speed in fire smoke deer eases as increasing smoke density and walking d i s t ance from

entrance in the corridor.

4 . for this exper i men t

,

the discomfort of s u b j ec t s against the- fire smoke was caused ma i n 1 y by the

i rr i tat i on of physiological factor fo r male, on the

o ther hand the psychological effect, the d ec r e as e o f visibility for instance, is import ant factor for

f ema 1 e

5 . The thirfking ability is affected by the h eat flux o f over 2 , 000 ( k c a 1 / rri h ) in fire smoke

The emotional instability in fire smoke i s composed o f psychological and physiological aspects

The degree 0 f contribution of each factor is not clear up in t h i s exp eriment

Further i nves t i gat i on s in this task are expected i n f u ture

References

1) T.JIN: KASAI, Japanese Assoc, of Fire Sci. & Eng. Vo 1.22,

No . 5 , p . 26 , ( 1 981 ) 2) T.JIN: 3ul 1 . of Japanese Assoc. of Fire Sci. 8 Eng.

Vo 1.32, No . 2 , p . 1 , ( 1982 ) 3) T.JIN: J. Fire & Flammability : 9, p. 135, (April 19 78)

299

A
le n trance!

ic o r r 1 a oil electric heater

A

'

Fig. 1 ; The outline of experiment.

Fig. 2 , The relation between the distance from entrance and the

correct answer rate.

(For the 14 subjects (8 males, 6 females) returned back

with in terrup t i on

)

300
Fig. 3 ; The relation between the distance from entrance and the

correct answer rate.

(For the 17 subjects (5 males, 11 females) reached at

the end of corridor in first trial)

EXTINCTION COEFFICIENT C (1/m)

Fig. 4 ; The relation between the smoke extinction coefficient and

the correct answer rate.

( At the E- point and in the first and second trial )

301
Q j

I

1 i

I l_ 0 2 . 4 6 3 10 12

DISTANCE From ENTRANCE (m)

Fig. 3 ; The relation between the distance from entrance and the

walking sjeed.

j ! 1 ! ! L

2 .4 .6 .8 1.0 1.2 EXTINCTION COEFFICIENT C

$

(1/m)

Fig. 6 The relation between the smoke extinction coefficient anc the walking speed .

302
Table 1; The discomfortable factor of the fire smoke

Q. ” Select three factors concerned with fire smoke

from foil wing items in order !”

selecti on items

male ( %) fe male :%)

I

s

'- 9 nd 3” 1 51 2

ns • Qro O

a 1 fou 1 smel

1( 72) 2(142) 3(212) 1( 62) 0( 02) ; i( 72)

b 1 smoke i rr i tat i on 8(57%) 4(292) 1( 72) 7(412) 6(352) : 2(132)

£

c ! obstruction of inhalaion 5(36%) 4(292) 5(362) 4(242) 3(182) 1 6(402)

d 1 obstruction of visi bi

ty 0( 02) 3(212) 3(212) 5(292)

r>{ or'V \

1 5(332)

e I heat 0( 02) 1( 72) 0( 02) 0( 02) 1( 62) : o( 02 ) = f feel i ng of isolation 0( 02) 0( 02) 1( 72) 1 0( 02) 1( 62) : !( 72)

‘

1 others 0( 02) 0( 02) 1( 72) 0( 02) 0( 02) 1 0( 02) a 1 fou 1 smel

• 2(142) 2(172) 3(302) 1 2(132) 7(502) : 2(152)

b smoke i rn tat ion 6(432) 0( 02) 2(202) 3(192) 4(292) : 2(152)

O c ! obstruction of inhalaion 3(212) 3(252) 4(402) 2(132) 1( 72) : 3(232)

d obstruction of visibi 1 ty 0( 02) 1( 82) 0( 02) 0( 02) 0( 02) : 1( 82)

e ! heat * 3(212) 5(422) 1(102) 8(502) Q( 02) : 2(152)

HZ f feel ing of isolation 0( 02) 0( 02) 0( 02) 1( 62) 1

<1 N zv 1 y : 2(152)

c

o others 0( 02) 1( 82) 0( 02) 0( 02) 0( 02) : K 82)

Table 2 ; The feeling for the

no smoke conditions

heat radiation under smoke and

Q. How do you reel for the neat flux on you at the deepest point- A ?

”

( for a! I subjects)

se 1 ected i terns wi th smoke wi th

i out smoke

1 . intolerably hot 2 ( 7%) I 8 (2 9%)

2 * hot but not intolerable 1 3 (44%) 1 1 4 (5 0%)

3 . comfortably warm 7 (23%) ! 4 ( 1 4%)

4. just feel heat a little 4 ( 1 3%) 1 2 ( 7%)

57 no fee 1 i ng 4 ( 1 3%) 1 0 ( 0%)

303
EMOTIONAL INSTABILITY IN FIRE AND SMOKE -- EXPERIMENT OF HUMAN BEHAVIOR IN

SMOKE FILLED CORRIDOR

T, Jin and T. Yaraada, Fire Research Institute, Japan

EMMONS: In your last chart there were a number of people who felt that the

heat was intolerable, and there were 13 who had no feeling. That seems odd

that some found it intolerable and some didn't fee-1 anything?

JIN: Ordinarily in the general public, I believe, there couldn't be such a situation, but this particular group may have had some individuals who were impervious to heat. I'd like to ask Mr. Sekizawa to give an answer on that as perhaps he studied it more.

SEKIZAWA: I'm one of those who did not particularly feel any effects of the

heat, so I can vouch for that. My attention was focused on being able to

withstand the smoke rather than feeling the heat. I did notice the heat, but

I didn't find it intolerable ... it was just hot.

HOTTEL: I'm concerned about whether the sample of the subjects were typical

because of thinking about three characteristics of subjects that might make

them act differently. One, speed of addition under normal conditions, which

might be high in engineers if you have an abnormal proportion of them there,

giving a weighted answer. Another is a natural tendency to emotional

instability as against the characteristic of getting a j ob done that you're

supposed to do. Those are very different characteristics, and which of those

drove a person is hard to visualize. Was it emotional instability, or that

characteristic of "I want to do a job well."

In your first set of tests, for example, how different people got hot would be

related to those things. It seems to me that these characteristics would

cause one to worry about whether a sample high in engineers was typical of

humanity

.

JIN: Before answering that I would just like to indicate that in using these

test individuals, we used each one only once. And, of course, since we used

housewives for the females we used them only once. I indicated employees of

the Institute; however, other than Sekizawa, who was involved in the test, the

test individuals were do-called part-time employees, student employees, who

are not particularly involved in smoke activities. Even then, I was very

concerned that there was a non- representational cross section among the males,

and I was prepared to throw out all the data relevant to the male responses,

but the responses were very close to the females, so I used them.

304
LARGE FIRES AND NUCLEAR WINTER

George F. Carrier

Harvard University

Division of Applied Sciences

Cambridge, MA 02138

1 . INTRODUCTION

In their attempts to estimate the effects of a major nuclear exchange on the subsequent dynamic and thermodynamic state of the atmosphere most

(maybe all?) of the investigators believe that the smoke from large

urban fires might be the most substantial contributor to the evolution

of the state of that atmosphere. Accordingly, many of the questions that

must be answered deal with issues related to large fires. In this brief

discussion before an audience drawn from the "fire community” I think it

will be useful to distinguish three aspects of the composite

phenomenology. The first, chronologically, is that which deals with a characterization of the hypothesized weapons usage scenario and must

provide as its output a characterization of the conditions that would

prevail shortly after the exchange has occurred. I will use the general

and imprecise phrase, "a characterization of", several times because, as nearly as I can tell, no-one has compiled carefully and credibly a list

of all of the specifications that play significant roles in the analysis

of the phenomena that could occur. In this particular instance I do have

in mind the need for a characterization of the redistribution of

structural combustibles (largely wood and synthetic materials?) and the

redistribution of non-combustible material (structural concrete, stone,

plaster, reinforcing and any other sources of rubble) the degree of

exposure of highly volatile fuels (e.g. devastated filling stations,

broken gas lines, the contents of commercial warehouses where solvents,

polymers, etc. are stored), and perhaps many others. I admit immediately

that I do not know, for example, the kind of detail needed in the

characterization of the rubble distribution but I will return to that

topic shortly.

I do suggest that the foregoing subdivision of the composite problem is

not an issue to be submitted to the fire community.

The second subdivision of the estimation process is that which addresses

the phenomonology that starts with the foregoing "redistribution inputs"

and has as its output a quantitative description of the characterizing

features of the deposition of smoke in the atmosphere at, perhaps, a day

or so after the exchange.* This subdivision is. the business of the fire

community and I will return to it

The third, highly intricate, aspect of the phenomenology requires the

unravelling of the evolution of the state of the atmosphere that would

*This rather fuzzy timing statement avoids lots of alternative conflict

durations that stem from alternate options related to the timing of

weapons deliveries. There may be many options (and classes of such

options) that can be studied if and when they seem to deserve attention.

305
accompany the interaction of solar radiation, the smoke laden atmosphere

("initially" in the state at the end of subdivision #2), and the land

masses and ocean of the planet. This also is not the business of the

fire community and I will return now to aspect #2.*

2. LARGE FIRES

It seems highly plausible and it is widely believed that the mechanical

devastation caused by the use of a multi-kilo-ton weapon on a target in an urban or near-urban location would be accompanied by the exposure of so much highly volatile "fuel" that fires, very heterogeneously

distributed over an area of a few to several km^' would be initiated and

sustained. The distribution in space and time of pvrolizate production,

of flame location, of smoldering reaction, of air access, and of oxygen

depletion, would also be very heterogeneous. It follows that, depending

on the evolution of the distributions so listed, such macroscopic

quantities as total heat release per unit time, air entrainment rate as

a function of altitude and time (which also depends on the ambient

dynamic state of the atmosphere) fraction of "fuel" consumed, total -rate

of smoke generation, character of the smoke generated (e.g. fraction

formed in flaming combustion vs. fraction formed in smoldering) and many

others are not entities for which there is a well documented body of

knowledge and understanding. To the contrary, in fact, there has been, even for the much smaller fires that have received scientific scrutiny,

almost no attention whatever to the dependence of burning rate on any of

the parameters characterizing fuel type, fuel distribution geometry,

fuel wetness, and external or self-induced wind field. Without a quantification of the foregoing dependencies, one cannot hope to

estimate in a credible way the distribution of smoke deposition

altitudes. Furthermore, without analogous correlations of (1) the

amount of smoke generated and (2) the variation of optical properties of

that smoke with combustion type (again e.g. smoldering vs. flaming) with

the foregoing parameters, one cannot tighten up the estimates of the

amounts and properties of the smoke that is distributed over those

altitudes. And, of course, the frustration that stems from the absence

of the information to which we have just alluded is amplified greatly by

the knowledge that a very broad variety of rubble types and of possible

rubble distributions must inevitably imply that the number of

experiments and the number of mechanisms whose effects must be

incorporated into the predictive models seem to be ever increasing

functions of the time during which we have been examining this problem.

3 . STATUS OF THE FIRE PROBLEM

Several investigators have modelled the dynamics of large fire

plumes . I know of three approaches

:

(1) Extension of the classical Morton, Taylor, Turner plume

theory. All of these adopt a postulated fuel consumption rate based

largely on anecdotal reports of the duration of large historical fires

(e.g. the Hamburg fire of World War II). Ordinarily, the analysis is

based on time independent conservation laws

.

* All readers will realize, of course, that many other questions such as, "what is the effect of all this on the biota (especially agricultural

biota) and the human population in various parts of the planet?" may be

of serious concern to all of us.

306
(2) Large hydro-code calculations. These also postulate the fuel

consumption history but, in this case, the consumption rate varies with

time and the fire evolves with time. I have not seen many attempts to

validate the hydrocode results by comparing known results in other

parameter ranges with code results.

(3) Initial attempts to devise models that distinguish the effects

of heterogeneity. These are at an early stage of development but without

this approach it is doubtful that models will be developed within which

one can treat the fuel consumption rate as an output of the analysis

rather than a postulated input.

In none of the foregoing families of estimation procedures has

there been any consideration of the "sheltering of fuel" by noncombustible rubble. Two aspects of that sheltering require attention.

One would speak to its role in altering the low level velocity field and

its implications for the availability of oxygen; the other would

influence the rate at which heat transfer to, and hence pryrolization

of, the fuel would be altered by the field of obstacles. Both of these

may have important implications for the net fuel consumption rate, for

the amount of smoke produced and possibly for the properties of that

smoke. The role of rubble deserves attention.

When a plume emerges from an intense fire in a very humid lower

atmosphere it can rise to altitudes at which, despite the entrainment

aloft of drier air, some of the moisture condenses . Observations suggest

that, in many such plumes, the water droplets accumulate large amounts

of smoke (I know of no quantification associated with these events) and

the rain that sometimes ensues carries significant amounts . of

particulate matter to the ground. Other observations suggest that, in

many such fires, capping clouds of condensed moisture form above the

darker plume but rain does not ensue. In these cases it is unclear as to

whether the presence of the condensate has any significant effect on the

morphology of the smoke

If one is to acquire a predictive capability regarding the amounts

of smoke and the optical properties of the smoke that might be deposited

in the upper troposphere as a result of nuclear weapon usage one must

find a far better quantification of the processes that occur in the

upper reaches of the plume than has yet been attempted. It seems clear

that the degree of supersaturation achieved by the plume gases is an important parameter in the determination of the droplet formation and

growth; and the local dynamics is important in the determination of the

degree of supersaturation. Since the amount of smoke that survives the

early scavenging and the optical properties of that smoke are crucially

important in the hypothesized alteration of the atmospheric state, it is

clear that any ability to quantify those alterations has as a prerequisite a quantified characterization of the dynamics and cloud

physics of the upper part of the plume.

4. NATURAL VARIABILITY

At this time, in addition to very large uncertainties in the

possible weapons usage and to those in the atmospheric response models

there are, as described in the foregoing, very large uncertainties in

the amounts, properties, and distributions of smoke that would ensue from a given weapons usage. At this time, there is in place no ongoing

307
organized experimental program with goals that would provide a body of

knowledge from which one could hope to build an understanding which

would allow him to scale the behavior of the necessarily smaller

experimental fires to those of interest in the Nuclear Winter context.

Unless one addresses the questions discussed in the earlier sections of

this paper, one will not have, at this time or at any other time, any

credible means of reducing the current enormous uncertainty regarding

the potential severity of the Nuclear Winter phenomenology.

There is, of course, no guarantee that a large and successful

research effort directed to these and other questions could reduce those

uncertainties. It is very possible that the sensitivity of the burning

rate--and therefore the plume height--and therefore the moisture

dependent scavenging processes--and therefore the surviving optical

blockage —to the atmospheric stability and humidity (they are coupled,

of course), to the distribution of rubble, to the variability of the

fuels contained in urban areas, and to the variability of other

parameters (e.g. weapons usage scenario) would leave us with a possible

range of smoke depositions that is as broad as that that characterizes

our present state (lack) of knowledge.

I favor an effort to improve our understanding of large fires and

their atmospheric depositions but (to repeat myself) it is not clear

that such an effort will remove the current uncertainties concerning the

atmospheric modifications at issue.

308
Large Fires and Nuclear Winter George F. Carrier

Bibliography

Albini, F . A . 1984 Wildland fires. American Scientist 72 , 590-597.

Bohren, C.F., and Huffman, D.R. 1983 Absorption and...S.ca.t.tegingi..o £ . L ight

by Small Particles . New York, N.Y: John Wiley.

Bohren, C.F., and Brown, G.M. 1981 Once in a blue moon. Weatherwise

11(3), 129-130.

Briggs, G.A. 1984. Plume rise and buoyancy effects. Atmospheric Science

and Power Production, pp . 327-366. Oak Ridge, TN : U.S. Department

of Energy, Technical Information Center.

Brunswig, H. 1982. Feuerstrum fiber Hamburg. Stuttgart: Motorbuch Verlag.

Carrier, G.F., Fendell, F.E., and Feldman, P.S. 1985 Firestorms. Journal

of Heat Transfer 107 , 19-27.

Carrier, G.F., Fendell, F.E., and Feldman, P.S. 1984. Big fires.

Combustion Science and Technology H, 135-162.

Colbeck, I., and Harrison, R.M. 1986 The atmospheric effects of nuclear

war —a review. Atmospheric Environment H, 1673-1681.

Committee for the Compilation of Materials on Damage Caused by Atomic

Bombs in Hiroshima and Nagasaki 1981. Hiroshima and Nacrasaki--The

Physical, Medical, and Social Effects of the Atomic Bombings. New

York, N.Y.: Basic Books.

Crutzen, P.J. 1984 Darkness after a nuclear war. Ambio JJ1, 52-54.

Crutzen, P.J., Galbally, I.E., and Bruhl, C. 1984. Atmospheric effects

from post-nuclear fires. Climatic Change £, 323-364.

Crutzen, P.J., and Birks, J.W. 1982 The atmosphere after a nuclear war:

twilight at noon. Ambio H, 114-125.

Dotto, L. 1986 Planet Earth in Jeopardy: Environmental Consequences of

Nuclear War . New York, N.Y.: John Wiley.

Levi, B.G., and Rothman, T. 1985 Nuclear winter: a matter of degrees.

Physics Today 2.2.(9), 58-65.

Luke, R.H., and McArthur, A . G . 1978 Bus hf ires in Australia. Canberra,

Australia: Australian Government Publishing Service.

Lyons, J.W. 1985 Fire . New York, N.Y.: Scientific American Books.

Malone, R.C., Auer, L.H., Glatzmaier, G.A., and Wood, M.C. 1986. Nuclear

winter: three-dimensional simulations including interactive

transport, scavenging, and solar heating of smoke. Journal of

Geophysical Research 9 ID

,

1039-1053.

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Manins, P.C. 1985 Cloud heights and stratospheric injections resulting

from a thermonuclear war. Atmospheric Environment .Li, 1245-1255.

Morton, B.R. 1957 Buoyant plumes in a moist atmosphere. Journal of

Fluid Mechanics Z, 127-144.

Mullholland, G.W., ed. 1987 Report on DNA/NBS Sponsored Workshop on Smoke Emission and Properties, November 13-14, 1986. Gaithersburg,

MD : National Bureau of Standards, Center for Fire Research.

.National Academy of Sciences, Committee on the Atmospheric Effects of

Nuclear Explosions 1985 The Effects on the Atmosphere of a Nuclear

Exchange . Washington, D.C.: National Academy Press.

Pittock, A.B., Ackerman, T.P., Crutzen, P.J., MacCracken, M.C., Shapiro,

C.S. and Turco, R.P. 1986 Environmental Consequences of Nuclear

War. Volume 1: Physical and Atmospheric Effects. New York, N.Y.:

John Wiley (SCOPE 28)

.

Pyne, S.J. 1984 Introduction to Wildland Fire--Fire Management in the

United States. New York, N.Y.: John Wiley.

Pyne, S.J. 1982 Fire in America — a Cultural History of Wildland and

Rural Fire . Princeton, N.J.: Princeton University,

Radke, L.F. 1983 Preliminary measurement of the size distribution of

cloud interstitial aerosol. Preciptation Scavenging , pp . 71-78,

New York, N.Y.: Elsevier.

Radke, L.F., Stith, J.L., Hegg, D.A., and Hobbs, P.V. 1978 Airborne

studies of particles and gases from forest fires. Journal of the

Air Pollution Control Association 2 8

.

30-34.

Rodhe, H. 1987 Book reviews, "Environmental Consequences of Nuclear

War" and "Planet Earth in Jeopardy: Environmental Consequences of

Nuclear War." Bulletin of the American Meteorological Society

£3.(2), 143-144.

Rogers, R.R. 1979 A Short Course in Cloud Physics, 2nd ed . New York,

N . Y . : Pergamon

.

Schneider, S.H. 1987 Climate modeling. Scientific American 256 (5) , 72-

. 80 .

Slinn, W.G.N. 1984 Preciptation scavenging. Atmospheric Science and

Power Prediction , pp . 466-532. Oak Ridge, TN : U.S. Department of

Energy, Technical Information Center.

Turco, R.P., Toon, O.B., Ackerman, T., Pollock, J.B., and Sagan, C. 1983

Nuclear winter: global consequences of multiple nuclear

explosions. Science 222., 1283-1292.

Twomey, S. 1977 Atmospheric Aerosols

.

New York, N.Y.: Elsevier.

310
LARGE FIRES AND NUCLEAR WINTER

G. Carrier, Harvard University, USA

ZUKOSKI : I understand that you're interested in getting this stuff up to

30,000 feet. It seems to me that even the old Turner Taylor Model is going to

likely fail us, because in a place like Los Angeles, the size of the fire

could be 30,000 feet.

CARRIER: I agree that one can expect sort of marginal results at best from

that kind of model, but what I was talking about were extensions of the model,

in which, for example, the point source is long gone, and in which the actual

state of the atmosphere, it's degree of stratification, is taken into account.

There even have been some investigations in which we try to take account of

the ambient circulation, but it only can be expected to give us gross

estimates of the height of the column, and cannot begin to give us the kind of

detail we need in connection with the condensation removal questions in

particular

I should add that the model has been very useful on some occasions in showing

how inappropriate some other kinds of calculations are by virtue of the total

unacceptability of the results when viewed against the Turner Taylor Model as

a comparison base.

EMMONS: First, relative to the general circulation problem, is it possible

that one could study some other planet to find a model that would be valid? I think, for example, Venus always would be covered with a cloud. Would that

provide a possibility for developing a model appropriate in this more general

situation?

CARRIER: There have been some attempts to use the Martian dust cloud

phenomena as a validation base for the modified GCM's, I suppose

with some success. However, our knowledge of what goes on, on a scale that

would allow us to talk about precipitation patterns on the earth, would be

very important to those who wish to indulge in agriculture after such an event. It doesn't permit that small of scale of phenomena to be resolved, so it's of limited validity.

Venus, of course, has such a different rotation rate, and such enormous

stability that I think any suggestions that might provide useful information

of that sort have been put aside rather quickly. There are natural phenomena

which provide severe tests in one way or another of GCM's, and people are actively searching for cases in point that might help with the validation or refutation of the things that are not being found in this context.

EMMONS: My second question is a very detailed one. The presence of a water

cloud above the smoke cloud implies separation of the smoke cloud from the

gases. Has gravity slipped from that?

CARRIER: The optical interaction of sunlight with the moisture cloud is such,

I think, that it gives an altered appearance to the region where there are water droplets. I was not suggesting, and indeed it's known not to be true,

that there are no smoke particles in that region. There are, but one does not

have a clear picture, or even a consistent picture, from observation to

311
observation

droplets as separation

,

as to how much there is

,

opposed to separate from

and I'm quite sure there

and how much of it is in or on the

the droplets. I didn't mean to suggest

can' t be one

.
OPEN TECHNICAL SESSION

313
FMRC Sprinkler Research

CHENG YAO

Manager, Research Division

Factory Mutual Research Corporation

Norwood, Massachusetts 02062

U.S.A.

Ninth Meeting of the United States/Japan Panel on Fire Research and Safety, May 1987

1. INTRODUCTION

Traditionally automatic sprinkler systems have been designed to control the

fire spread and to wet and cool the building structure, thereby preventing

collapse and complete wreckage. Sprinkler protection requirements of different storage occupancies in terms of water density and area of water demands

were developed primarily from large-scale fire test programs which used a defensive strategy or conservative concept of "fire control." This concept permits lower densities and assumes large numbers of sprinklers operating (in the

range of 20 to 30). Fires in today's high-piled storage occupancies have

challenged the control concept, pushing the sprinklers of current design to

the limit of their effectiveness. A new approach to provide a cost-effective

solution to the high challenge storage fire protection was urgently needed.

Factory Mutual has supported a substantial sprinkler technology research program to study the principles of sprinkler fire protection. Knowledge developed has been used as a basis to reexamine the sprinkler protection concept

and to develop new and effective sprinklers for the occupancies which cannot

be protected with standard (spray) sprinklers. This work led to the development of the large-drop sprinkler (1971-1980) to control high-challenge storage

fires, fast-response residential sprinklers (1 976-1979) to maintain a survivable environment in residential areas, and the early suppression fast response

(ESFR) sprinkler system (1984-1986) to suppress severe storage hazard fires

that are beyond the protection capability of even large-drop sprinklers without supplementation with expensive in-rack sprinklers.

It is the intention of this paper to present an overview of our sprinkler

technology program and to give examples of some recent and future advances in

the application of this technology.

2. SPRINKLER TECHNOLOGY RESEARCH

The objectives of the sprinkler technology research program are:

1) to quantify the thermal responsiveness of automatic sprinklers, which can

be used to determine the fire challenge at the time of sprinkler operation;

2) to quantify the fire challenge in terms of its combustibility and extinguishability , which can be used to predict the density of water required

to be delivered at the top of the burning storage array for early fire

suppression;

314
3) to quantify the sprinkler discharge characteristi cs , which can be used to

predict the density of water actually penetrating the fire plume and delivered onto the top of the burning array;

4) to develop a phenomenological model of a sprinklered fire in a storage

warehouse including mathematical simulation of various components of a comprehensive sprinkler performance model, which can be used to predict

performance of sprinkler systems against different types cf fire challenges .

2.1 SPRINKLER RESPONSE PREDICTION

For a given fire situation, sprinkler response time is determined by the thermal responsiveness of the sprinkler and the gas temperature and velocity at

the sprinkler location. Assuming the heat responsive element cf a sprinkler

is heated purely by forced convection until actuation, the thermal responsiveness of the sprinkler heat sensing element can be quantified by a Response

Time Index (RTI). RTI can be measured in Factory Mutual's plunge test tunnel.

Ten different brands of sprinklers responded with fairly constant RTI over the

velocity (1.5 to 5.2 m/s) and temperature ranges examined^ 1

. RTI and sprinkler temperature rating had been demonstrated to be sufficient to predict re- sponse of conventional (slow-response ) sprinklers in full-scale rack storage

fire tests ^

.

he protection cf

that the heat

The introduction of low RTI fast-response sprinklers for t slow growing fires in residential situations has demonstrated

loss by conduction from the heat sensing element to the sprinkler fitting can become

time.

a significant factor in the

An attempt was, jt hen made to

thermal response model

determination of the sprinkler response

incorporate conductive heat loss in the

Adding a conductive heat loss terra, the heat balance on the heat sensing element becomes:

me (dT

e /dt ) = hA (T -Te )-C' (Te -T Q

)

where m

ch

A

C

T

T

T

t

eg

o

mass of element,

specific heat of element,

convective heat transfer coefficient,

surface area of element,

conductance of sprinkler,

temperature of element,

gas temperature,

ambient temperature,

time

Substituting

RTI = u 1/2 (mc/hA),

C = C' ( RTI/mc )

(1 )

The solution of Eq (1) for a plunge test can be written:

315
RTI

( 2 ) i+c/u

1/2 m [

i

-at (

i +C/U

1 /2

)/at -] ea g where ATfia is the actuation temperature of the heat responsive element above

ambient.

The conduction parameter, C, can be measured in the plunge test tunnel using a "prolonged exposure technique"^ . The C value measured among 11 sprinkler

models range from 0.5 to 1.56 m

1 s'! '/ ^. Figure 1 shows the temperature rise

AT 0 as a function of time, calculated for C=0 (no conductive heat loss) and

for C=1 (a realistic value). The element with C=1 actuates much later than

that for C = 0. A series of room fire tests has been conducted using both pool

fires and crib fires. Actual sprinkler response times correlate well with

predictions based on the model using the measured gas temperature and velocity

values and laboratory values of RTI and C ^ '

.

It is also demonstrated that

conduction correction is not necessary to predict response time for fastresponse sprinklers in industrial applications such as ESFR sprinkler systems.

2.2 QUANTIFICATION OF FIRE CHALLENGE AT SPRINKLER RESPONSE

The principal challenge a storage fire presents to a sprinkler system at the

time of sprinkler response is the combustibility of the ignition stacks of

storage commodities. This combustibility may be expressed in terms of convective heat release rate at the time of sprinkler actuation, the associated upward thrust of the fire plume and the ceiling flow char acteristi cs of the fire

products. Empirical correlations of ceiling flow velocity and temperature

have been established in terms of convective heat release rate and ceiling

clearance for cases where a quasi-steady approximation of the ceiling flow is

valid^ . In order to predict the lew RTI sprinkler response accurately,

generalized correlations of the gas temperature and velocity in the initial

transient ceiling flow has to be established.

A series of storage fire tests has been conducted under the Fire Products Collector (FPC), a large calorimeter which collects all fire products and determines the fire's heat release rate. The thrust of the plume, derived from the

measured initial transient plume temperature and velocity profiles, has been

correlated with the fire growth history. The same tests will be repeated

under a 30-ft (9 m) high ceiling to develop generalized correlations for initial transient ceiling flow and the fire growth history (heat release rate

history) and ceiling clearance.

A computer program for predicting the ESFR sprinkler response will be developed using the correlations for transient ceiling flows developed in this

work. The accuracy of the computer prediction of ESFR sprinkler response and

the fire plume thrust force will be compared with the measured values in the

fire tests.

2.3 QUANTIFICATION OF STORAGE FIRE SUPPRESSION - RDD

Over the last decade, attempts have been made to quantify the fire suppression

requirement for different stored commodities at the time of sprinkler actuation while the fire is still confined within the ignition stacks of storage

array. This fire suppression parameter is denoted as the required delivered

density of the storage array or RDD. The RDD is the water density which must

316
be delivered to the top surface of a burning storage array in order to achieve

early suppression.

A large-scale fire suppression test apparatus, as shewn in Figure 2, was developed. It consists of a special water applicator that is capable of spreading a fairly uniform distribution of water directly over the top surface of

the fuel, thus simulating water that has penetrated the fire plume. This apparatus is positioned below the FPC. The water applicator is actuated at a specific heat release rate (as measured by the FPC), which corresponds to that

of an ESFR sprinkler actuating in a 30-ft (9 m) high building (see Section

2 . 2 )

Figure 3 shows the results of three HDD tests of identical rack storage of

plastic commodity in a 8-ft x 8-ft x 20-ft high (2.4-m x 2.4-m x 6.1-m) configuration. Water was applied at the same heat release rate. The fire was suppressed at a delivered density of 0.97 gpm/ft“ (19.2 mm/min); at 0.92 gpm/ft"

(17.2 mm/min) fire redeveloped, and at 0.50 gpm/ft 2 (20.9 mm/min) a faster

suppression was achieved.

Numerous combinations of storage commodity and configuration and ignition con- dition were examined. The storage commodities tested were:

1) Polystyrene jars in ccmpar tmented cartons;

2) Metal lined double triwall cartons;

3) Aerosol products of different hazard levels;

4) Roll paper of different forms, including polyethylene film coated and

chemically treated.

2.4 COMMODITY HAZARD EVALUATION FOR FIRE CONTROL

Sprinkler protection requirements of storage occupancies for fire control systems were developed from large-scale fire tests using different standardized

test commodities. The qualitative test procedure presently used to evaluate

or classify the hazard of a storage commodity relies to a large degree on subjective judgement of trained observers. The objectives of this new project is

to develop an improved commodity evaluation test procedure based on the com- bustibility and controllability of the storage commodity as quantified by the

FPC. The test procedure is similar to the RDD test, except that the water

applied is not intended to suppress the fire, but merely keep the fire from

becoming out of control.

Eased on both freeburn and fire control test results, five parameters were selected for indexing the fire hazard of storage commodities. They are fire

growth rate, maximum convective and radiative heat release rates, convective

energy release and maximum automatic sprinkler link temperature response. All

the fire control parameters were determined as functions of water delivered

density. Classification criteria based on these parameters will be established upon the successful validation with the full-scale fire test results.

2.5 QUANTIFICATION OF SPRINKLER PERFORMANCE - ADD

Under a no-fire condition, the water distribution over a specific area of con- cern (for example, the top cross-section area of the 'ignition stacks) is a function of sprinkler design, sprinkler location, number of operating sprink317
lers and discharge pressure. In the presence of a fire, however, the density

actually delivered at the top of the ignition stacks, the Actual Delivered

Density (ADD), i s ' inf luenced by the updraft of the fire plume. The upward

draft in the plume presents a barrier to the penetration of drops; only a fraction of the discharge density can be delivered at the base of the fire

plume

.

As shown in Figure 4, the ADD apparatus consists of four simulated storage

commodities and a heptane-spray fire source. In each of the simulated commo- dities, there are four water collection pans. Under each pan there is an automated water metering system. There are 20 water collection pans, 16 at

the top and M at the base of the flue spaces separating the simulated commodities. This apparatus is placed under a large suspended ceiling which is mov- able to provide different clearances between the test array and sprinklers.

In addition to ADD tests at specific test conditions, a statistical technique

is being used to design, collect, analyze and map the ADD performance of the

sprinkler over a broad range of operational variables. Figure 5 shows a performance map for one of the prototype ESFR sprinklers.

2.6 COMPUTER MODELING OF SPRINKLER SPRAY AND FIRE PLUME INTERACTION

The drops may be generally expected to penetrate if the terminal drop velocity

is greater than the gas velocities in the fire plume, or the drop size (d) is

larger than a critical maximum drop size (d ) that can just be stopped by the

centerline plume flow. This type of penetration is ref erred to as gravitational penetration. A second type of penetration, momentum penetration, oc- curs if the total thrust of the water spray (M

s

) overpowers the upward thrust

of the plume (M

p

).

Numerical computation to model sprinkler spray interaction^ with a fire

plume gas has been run on the computer for a wide range of sprinkler discharge

conditions and fire intensities to predict spray penetration ratio, Pen.

Pen = ADD/D.d

where D_ is the rate at which water is applied to a unit area without fire.

CL Analysis of these solutions has led to the identification of a parameter con- trolling the penetration of droplets to the base of the fire plume. Penetration ratios calculated are shown in Figure 6 as a function of a thrust ratio,

M

s /Mp, a drop-size ratio, d/d

c

, and a velocity ratio, U inj/Up0, evaluated at

the elevation of the spray source. Here U inj

is the droplet effective injection velocity and U p0 is the plume centerline velocity. These controlling

parameters involves Doth characterise cs of the momentum and gravitational

penetration mechanisms. When the combined controlling parameters are greater

than a critical value, mere than 80% of the water discharge arrives at the

plume base.

The water flux, drop size distribution and thrust force of the ESFR sprinkler

will be measured. These data will be used as an input to the model to check

the ability of this model to predict spray-plume interation in terms of ADD.

318
2.7 1/12 SCALE SPRINKLER SYSTEM MODELING

In addition to the simple density/area rule, sprinkler performance is effected

by at least six principal variables: ceiling clearance; sprinkler temperature

rating; sprinkler responsiveness (RTI); water pressure; sprinkler type and

orifice size; and burning and extinguishment characteristi cs of the combustible. Large numbers of parametric tests have been conducted to investigate

the interrelations between the sprinkler variables in a reduced-scale model

(1/12.5 of full scale). According to the principle of Froude number modeling'"^ , the fire heat release rate is scaled to L^ , where L is a representative length dimension. Drop diameter and the initial drop velocity and the

diameter of the heat sensing element are all scaled to L W~. A model building

1.44-m high was constructed to represent a full-scale building 18 -m high.

The first series of modeling tests were designed to simulate full-scale fire

tests, using the standard 1/2 in. (13 mm) sprinklers to control the fire in

the rack storage of ordinary combustibles in cardboard cartons. A total of

164 fire tests -was performed with various combinations of clearance, temperature rating, RTI and water pressure. Where comparisons were possible, model

results generally agreed with full-scale test experience.

Another series of modeling tests to simulate full-scale early suppression fire

tests using the ESFR sprinkler system is underway.

2.8 PHENOMENOLOGICAL MODELING OF SPRINKLERED WAREHOUSE FIRES

Work is about to begin on the development of a phenomenological sprinkler performance model intended to. simulate a high-challenge sprinklered warehouse

fire. The model will be used in situations for which there are no specific

full-scale test data, but the relevant commodity combustibility and extinguishability data described in Sections 2.3 and 2.4, and the sprinkler characterizations described in Sections 2.1 and 2.5 are available. Phenomena to be

incorporated in the fire simulations include: 1) freeburn fire growth in

terms of heat release rate and flame spread; 2) plume and ceiling layer flows

as described in Section 2.2; 3) sprinkler actuation times and spray discharge

characteristics; 4) spray-plume interactions in terms of the parameters described in Section 2.6; and 5) delivered water suppression capability as characterized by data/correlations described in Sections 2.3 and 2.4 including

effects on fire heat release rates (and possibly flame spread) so as to allow

calculations to proceed until the fire is suppressed and/or flame spread

ceases. The model will be initially formulated as a deterministic calculation, but uncertainti es in simulating these phenomena will be quantified and

eventually incorporated into a probabilistic formulation that can be exercised

along with probabilistic descriptions of the fire scenario (ignition location,

storage configuration details, etc.).

3. THE DEVELOPMENT OF THE ESFR SPRINKLER SYSTEM

ESFR stands for Early Suppression Fast Response. The objective of the ESFR

program was to develop an ESFR sprinkler system which will provide "early suppression" performance on high challenge storage fires represented by up to

25-ft (7.6 m) high rack storage of standard plastic commodity in a 30-ft

(9.1 m) high building. Such a severe fire challenge is beyond the protection

319
capability of standard and even the large-drop sprinklers without supplementary and expensive in-rack sprinklers.

The basic ESFR concept is that if a sufficient quantity of water can be delivered onto the burning storage commodities at the early stage of development to

suppress the fire quickly before a severe challenge situation is reached,

fewer number of sprinklers will operate and less fire and water damage will

resul t

.

In principle, early suppression performance is determined by three factors

which can be independently measured: 1) Response Time Index (RTI); 2) Actual

Delivered Density (ADD); 3) Required Delivered Density (RDD). The RTI and the

time-dependant nature of ADD and RDD determines the early suppression performance of the ESFR sprinkler system. The premise of the ESFR system is to ensure ADD in excess of RDD, regardless of the ignition location.

After two years of extensive research and development, a prototype, de

as CPF sprinkler, was chosen as representati ve of the ESFR sprinkler

is of pendent design and has a K factor of 1 A gpm/psi 1/2 (20.1

kPa 1/2

), 0.7 in. (17.8 mm) orifice diamecer. The sDrinkler link used

at 1 65° F (79°C) with a RTI of 50 ft 172

s 1/2 (28 m

1/2

s 1/2

).

?^ated

dm ^ /min/

is rated

The most important performance requirement of the ESFR system is to provide

ADD in excess of RDD under a worst-case scenario of up to 25-ft (7.6 ra) high

rack storage of standard plastic commodity in a 30-ft (9.1 m) high building,

regardless of the ignition location. Figure 7 shows the ADD value of the ESFR

prototype sprinkler at different clearances between the sprinkler and the top

of the storage for two different spri nkl er-f ire source conf igurations ; i.e., one and four sprinklers centered above the fire source. In comparison with

the RDD values measured for different storage heights, the prototype ESFR

meets performance requirements for different ignition (fire source) locations.

It also shows that with one sprinkler discharging centered over the fire

source, high clearance is critical; with four sprinklers, low clearance is

critical

.

( q

\ Eleven large-scale fire tests v ' of double- and multiple-row rack storage fire

tests were conducted. Results of these fire tests not only confirm that the

prototype ESFR sprinkler would perform well in critical conditions described

above, they also demonstrate that even under the worst failure mode condition

tested, ignition between two sprinklers with one of them plugged, the fire was still controlled with 11 sprinklers.

Factory Mutual has just released a draft installation standard for public re- view. The ESFR sprinkler system protects palletized, solid-piled and open

frame single-row, double-row, multiple-row and portable rack storage of most common materials up to 25-ft (7.6 m) high, in buildings up to 30-ft (9.1 m)

high. Specific commodities protectable by ESFR sprinklers include Class I through IV commodities, cartoned unexpanded plastics, and cartoned foamed-inplace polyurethane packaging material. Also, any lesser hazard commodities,

lower storage heights, or lower ceiling heights are all protected with the same water demand: 12 sprinklers at a minimum of 50 psi (395 kPa).

320
4. THE DEVELOPMENT OF THE RESIDENTIAL SPRINKLER SYSTEM

The quick-response concept in sprinkler technology as applied to residential

occupancies has been brought to fruition by FM through the support of the

United States Fire Administration (USFA)/Federal Emergency Management Agency

(FEMA). From 1976 to 1 982, research results demonstrated the need of fastresponse (low RTI) sprinklers with different discharge patterns for providing

adequate life-safety and property protection in residential fires. Based on

the information gained

1 ^

1

, new- generation , quick-response (low RTI) residential sprinklers have been developed by the U.S. Sprinkler Industry. In

1980, a new version of the NFPA-13D Standard on Installation of Residential

Sprinkler Systems was adopted by the National Fire Protection Association

(NFPA). In 1983, research results indicated that there is a need of sidewall

sprinkler installation rules with regard to system water demand and sprinkler

location

1

.

Currently, USFA/FEMA is sponsoring a two-phase research program at FMRC to

investigate the Effect of Beams and Ceiling Slopes on Residential Sprinkler

Performance, and to evaluate the Performance of Quick-Response Sprinklers in

Office and Hotel Occupancies.

REFERENCES

1) Heskestad, G. and Smith, H.F., "Investigation of a New Sprinkler Sensitivity Approval Test: The Plunge Test," FMRC 22485, Factory Mutual

Research Corporation, Norwood, Massachusetts, December 1976.

2) Kung, H.C., Spaulding, R.D. and You, H.Z., "Response to Sprinkler Links

to Rack Storage Fires" FMRC J.I. 0G2E7.RA, November 1984.

3) Pepi , J.S., "Relationship Between Thermal Response and Fire Suppression

Characteristics of Automatic Sprinkler in a West-Pipe System" Presentation at Fire Detection and Suppression Symposium, SFPE, March 9-11, 1987.

4) Heskestad, G. and Bill, R.G., "Quantification of Thermal Responsiveness

of Automatic Sprinklers - RTI, ETC." Presentation at Fire Detection and

Suppression Symposium, SFPE, Maritime Institute of Technology, Linthicum

Heights, Maryland, March 9-11, 1987.

5) You, H.Z. and Kung, H.C., ""Strong Buoyant Plumes of Growing Rack Storage

Fires," Twentieth Symposium on Combustion, The Combustion Institute,

1 984.

6) Alpert, R.L. and Friedman, R., "Computer Modeling of Sprinkler Spray

Interaction with a Fire Plume" Presentation to the Scientific Section of

the International Meeting at Avila, Spain, October 7-9, 1986.

7) Heskestad, G., "Parametric Study of Sprinkler Fire in Reduced-Scale

Model" FMRC J.I. 0G0E5.RU, June 1986.

8) Fire Detection and Suppression Symposium, SFPE, Maritime Institute of

Technology, Linthicum Heights, Maryland, March 9-11, 1987.

9) Chicarello, P.J., Troup, J.M.A., Dean, R.K., "Large-Scale Fire Tests

Evaluation of Early Suppression Fast Response (ESFR) Automatic Sprinklers," Prepared for National Fire Protection Research Foundation, and

Factory Mutual Research Corporation, FMRC 0M2R5.RR and 0M0J7.RR,

May 1986.

10)

Kung, H.C., Spaulding, R.D. and Hill, E.E., "Sprinkler Performance in

Residential Fire Tests," Technical Report FMRC Serial No. 22574, Factory

Mutual Research Corporation, Norwood, MA, July 1980.

321
11) Kung, H.C., Spaulding, R.D., Hill, E.E., and Symonds , A.P., "Field Evaluation of Residential Prototype Sprinkler, Lcs Angeles Fire Test Program,"

Technical Report, FMRC J.I. 0E0R3.R(1), Factory Mutual Research Corporation, Norwood, MA, February 1982.

12) Kung, H.C., Spaulding, R.D., and Hill, E.E., "Residential Fire Tests with

Sidewall Sprinklers," Technical Report, FMRC J.I. 0J1N0.RA, Factory

Mutual Research Corporation, Norwood, MA, December 1985.

ig. 1 Temperature rise of a hypothetical,

RTI = 200 (m/s) ^ sprinkler in

environment of constant temperature

(AT = 110°C) and velocity (1 m/s) for

two

6different values of the conduction

parameter

Fig. 2 Test setup for the determination of

heat release rate and required

delivered density of storage

commodities within the ignition stack

of storage arrangement

Fig. 3 Results of three RDD tests of 20-ft

(6.1 m) high rack storage of plastic

commodities

322
ig. 4

SPRINKLES DISCHAACE PRESSURE

Palg (kSa)

Fig. 5 ADD performance map for prototype ESFR

sprinklers with the fire source centered

between two discharging sprinklers, 10 ft

(3 m) clearance, 10 ft (3 m) spacing

ADD test apparatus for measuring the between sprinklers

ADD performances of candidate ESFR

spr inklers

o^eCMyWp^+o^zCd/dj^+o.ozyCu^/u^)’-6

ig.6 Optimized correlation for heat release

and ceiling height effects (one sprinkler

operated directly over the fire- source)

Clearance to Sprinkler ( m )

Storage Height ( m )

Fig. 7 Comparison of ADD value measured at

different clearances with RDD values

measured for rack storage of plastic

commodity at different storage height

(CPF sprinkler, 3 m x 3 m spacing,

345 kPa, 1034 kW convective heat

release rate)

323
FMRC SPRINKLER RESEARCH

C. Yao , Factory Mutual Research Corporation, USA

HAS EMI : Did you consider the radiation effect?

YAO: At the moment, we ignore the radiation effect. In this on-going

program, we're going to use a 9 -meter high ceiling and will be looking

carefully into the radiation effect on sprinkler response . In a month or two

,

we will have a report addressing the radiation effect and sprinkler response

in an industrial fire situation.

If it becomes important, we may have to modify the model to include radiation,

but that would not effect the RTI measurement or conductive term measurement.

We just measure them separately.

ZUKOSKI : You had some measurements of transient plume motion?

YAO: Yes.

QUINTIERE: I noticed on the slide for the B ceiling that it appeared to be a combustible ceiling. If that is the case, what is the effect of a combustible

ceiling?

YAO: The slide I showed was borrowed from a large drop sprinkler program that

we ran ten years ago. The ceiling was involved about three or four minutes

after ignition. Also, as soon as you suppress a fire, even a fire like that,

it disappears within ten seconds. Without a fire plume to feed it, the

ceiling flow doesn't propagate itself without a fire pushing it. In the

model,' we use a non- combus tible ceiling. We assume, at this moment, we're

talking about a sprinkler operating within thirty or forty seconds; I don't

think the ceiling has a chance to get involved. It may char a little, but if

it suppresses the fire as fast as we want, we don't have to worry about it.

But, when we validate the model in Rhode Island, we will use a combustible

ceiling to validate the results.

324
Progress in Modeling Sprinkler Performance

DAVID D. EVANS, WILLIAM D. WALTON AND

BERNARD MCCAFFREY

Center for Fire Research

National Bureau of Standards

Gaithersburg, MD 20899, USA

ABSTRACT

The Center for Fire Research has begun the construction of a fire suppression

submodel to be included in one of the current zone fire models. Initial

studies will concentrate on the extinction of liquid pool fires by water

sprays, but may be generalized to other fuels within the same submodel

framework. Studies related to the overall performance of sprinkler systems

that are detailed in this report are: Sprinkler response time predictions

within enclosures, sprinkler spray characteristics, droplet evaporation time

predictions, and wood crib fire suppression.

1. INTRODUCTION

Water spray sprinkler systems are a common and effective means of providing

automatic fire protection for facilities. For decades, these systems have

been put in place in industrial facilities to provide for prevention of spread

and possible extinction of accidental fires. During the past decade,

considerable research has been performed to extend the use of sprinklers to

ordinary residential living spaces in which new technology systems may provide

both property and life safety protection in the event of accidental fires.

Even though there is considerable qualitative information about the various

physical and combustion phenomena associated with fire suppression using water

sprays, there is relatively little quantitative information available from

measurements in studies of room fire extinguishment. In the common cases of a bedroom or living room fires, which have been studies for over a decade to

quantify fire growth processes, there exists plenty of debate as to the

relative importance of several mechanisms that operate simultaneously during

the suppression of a furniture fire. In each case, the possible physical and

chemical mechanisms involved may be enumerated. In most cases, quantitative

studies under idealized research conditions can be found in the literature to

quantify the individual physical and energetic mechanisms contributing to the

fire suppression action of the water spray in a compartment fire. The

energetics associated with the extinction of fires by a) cooling of a solid

liquid fuel below some critical temperature, b) dilution of a miscible liquid

fuel, c) separation of the fuel vapors from a means of oxidation, d) reducing

the critical level of radiation flux from the flame and surroundings back to

the fuel, e) dilution of the oxidizer concentration by steam are well known.

Additionally, there are data available, albeit less than is desirable, which

325
propose to give critical water spray fluxes needed to extinguish selected fuel

packages

.

Except for the recent work of Pietrzak e_£ jit [1], there has not been any

efforts to produce a single predictive model for analysis of extinguishment of

accidental fires in enclosures using water sprays. The Center for Fire (CFR)

at the National Bureau of Standards (NBS) has underway a research program to

construct modules of an extinction model which may be implemented into one of

the available fire growth room fire codes, such as FIRST [2] or FAST [3],

Although the technical content 0 f the suppression subroutines mav change as more information is obtained through the next decade of research necessary to

perfect the modeling, the framevork should remain unchanged. For example, the

position of the automatic spriniler within the enclosure, the initial

trajectory and drop sizes in the spray, and the volume flow rate of water will

always be required as input to the model. As the spray mav interact directly

with the burning surface, or be blocked by obstructions, the addition of a suppression module to the current room fire models will require more input

into the model about the fuel and room geometry than is currently used in

routine execution of the existing codes. The initial framework for the

suppression submodel and required changes to the existing room fire models is

being assembled this year. During the next year, predictive methods for

liquid pool fire extinction within ventilated enclosures will be formulated

and partially validated. As part of the support for the fire suppression

submodel limited studies were completed 'in 1986 to quantify individual

physical phenomena involved with the prediction of sprinkler performance and

the collection of full scale test data. The results of these studies

conducted as part of the in-house CFR research program, the extensive

cooperative research studies with U.S. and foreign government agencies, and

the CFR university grants program are summarized below.

2. SPRINKLER RESPONSE TIME

Of primary importance to the modeling of sprinkler system performance is the

prediction of the time of activation of the sprinkler which determines at what

time during the growth of the fire suppression by water spray will begin. Two

activities were completed. First, the formulation and coding of the

compartment fire sprinkler activation model that considers the effects of an accumulated warm gas layer in the upper portion of the room was completed and

added to the popular ASET code [4], As part of this overall analysis, studies

of two-layer plume flow characteristics were completed at laboratory scale by

Evans and Morehart [5], Figure 1 shows an example case of the agreement

between predicted and measured plume flow temperatures above and below the

warm-cool layer interface in an enclosure. This analysis forms the basic

analysis that allows the substantial research work and experimental

information available for tests using very large or unconfined ceilings to be

used in the modeling of sprinkler activation in small rooms typical of

residential living spaces.

Second, the extensive analytical information for thermally activated devices

installed below large unobstructed ceilings was organized into both tabular

form and a user friendly computer codes by Evans and Stroup [6,7], These

works provide practicing fire protection engineers with a needed tool for

326
TEMPERATURE

FUSE

ADOVE

AMBIENT

(K)

Fig. 1. Measured and predicted temperatures for

burner located 0.77 m below the ceiling.

327
analysis of expected sprinkler system response time that can easily be used as

a guide for selection of appropriate equipment and installation spacing.

3. SPRINKLER SPRAYS

Sprinkler spray drop size and trajectories are important input information to

the fire suppression subroutines under development at CFR. Wendt and Prahl at

Case Western Reserve University in Cleveland Ohio have reported on studies of

spray trajectories produced from idealized and actual sprinklers [8], One

objective of their study was to understand the dynamics of the water sheet

break-up that determines initial dropsizes and trajectories at the sprinkler.

Another was to formulate a model for the spray-gas interactions that was capable of predicting water spray flux distributions in the absence of fire.

Figure 2 shows a comparison of the calculated and measured radial distribution

of water spray flux per unit area at floor level from their tests of idealized

sprinklers. Using radial spray distributions that are easily measured, and

the available physical models, adequate information for initial drop size

distributions and trajectories may be calculated that potentially avoid the

difficult and tedious task of physical drop size measurement for the purpose

of suppression modeling. Further studies are required in which both spray

distribution and drop size (unavailable to these investigators) information is

used to validate the model.

Extensive drop size distribution measurements are being collected in

cooperative studies funded in part by the Swedish Fire Research Board which

have brought together extinguishment specialists in CFR, the British Fire

Research Station, and the Polytechnic of the South Bank, London to examine the

interactions of sprinkler sprays with hot gas layers. To provide initial

spray drop size and trajectory information from actual sprinkler hardware, the

spray characteristics are being measured at CFR. Figure 3 shows a histogram

of the spray droplet distribution and droplet velocity distribution from a pendent sprinkler at a sample location near the discharge from the deflector.

Average velocity of droplets in the spray is indicated to be 5.65 m/s an angle

of -86.2° to the plane of the deflector. This is one of over a hundred

positions measured in this single spray.

Detailed initial trajectory, drop size and radial spray flux distributions

from sprinklers will be used to further evaluate the predictive capabilities

developed by Wendt and Prahl in future studies.

4 . DROPLET EVAPORATION

The Center for Fire Research through cooperative studies with the University

of Maryland is quantifying the heat transfer processes involved in the

evaporation of water droplets on hot surfaces. Investigators diMarzo and

Evans hope to eventually construct a thermal model for solid fuel extinction

based on the detailed understanding of both single and multiple droplet

evaporation processes. Initial studies have been performed on the simplified

problem of single droplet evaporation from a heated non-porous aluminum

surface. Later studies will examine the effects of porosity, material thermal

conductivity, and external heat flux.

328
NONDIMENSIONAL

RAIN

RATE

q'-(Q

R

/Q)/(2r'Ar’>

NONDIMENSIONAL RADIAL POSITION, r'-r/R

Figure 2 Calculated and measured radial distributions

of water spray flux per unit area for an idealized sprinkler.

329
DROPLET

COUNT

DROP

VELOCITY,

m/e

Figure 3 Droplet size distribution and velocity distribution in

the spray from a pendent sprinkler

330
Initial studies [9] examined the case of high thermal conductivity solids for

which the surface temperature of the solid under the evaporating droplet is

nearly constant. Setting this constant temperature equal to the available

analytic solution for the contact temperature between semi - inf ini te bodies of

differing thermal properties essentially decouples the analysis of the solid

substrate and droplet evaporation processes. This permits a rather simple

overall analysis of the evaporation process that yield predictions of

evaporation time which are compared with measured values in Figure 4 for three

different initial drop diameters. Figure 5 shows the comparison of calculated

and measured time dependent variations of droplet volume for a 30 microliter

droplet evaporating on an aluminum surface initially at 98°C.

Further studies underway quantify the evaporation process for low thermal

conductivity solids, and multiple droplet exposures characteristic of sprays

effects during fire suppression.

5. LARGE SCALE TESTING

Water spray extinguishment phenomena associated with both actual furniture

fires and research fuel loads such as cribs have been examined by Walton at

CFR. This research has two objectives, one is to provide practical

information on the performance of actual sprinkler system. The other is to

examine the range of water sprays and activation time that are effective in

controlling fires in easily characterized fuels for aiding the development of

sprinkler suppression models.

In an applications based study for the U.S. General Services Administration

and the U.S. Fire Administration it was shown that new technology residential

fast response sprinklers could be effective protection for office space fires

that presently require the use of greater flow rate standard sprinklers [10],

A series of tests underway at this time is designed to examine the effect of

water spray flux on the extinguishment of various fuel packages. Initial

tests have focused on fires in wood cribs constructed of fir with a total

weight of approximately 21 kg. The cribs consist of 8 layers with 6 sticks

per layer where each stick is 38 by 44 by 609 mm. A sprinkler head is located

1.75 m above and 2.02 m away from the crib. Water spray flux to the area of

the crib is measured in a separate test.

Water is applied to the crib when the maximum and nearly steady burning rate

of the crib has been obtained. Rate of heat release measurements of the crib

fire are taken using oxygen consumption calorimetry. Figure 6 shows the heat

release rate for three tests where time has been shifted so the sprinkler

operation occurs at the same time (approximately 450 s) for each of the tests.

The spray flux per unit area for test 511 was 0.058 mm/s , for test 514 was 0.054 mm/s and for test 516 was 0.041 mm/s. Although these tests cannot be

compared directly because of the different ignition sequences, it appears the

initial reduction in heat release rate due to the application of the three

different spray fluxes of water was nearly the same for all tests. Further

work will be performed in an effort to understand this phenomena as it may be an important finding that could simplify modeling of the suppression process.

331
EVAPORATION

TIME

(s)

75 80 85 90 95 100

INITIAL SOLID SURFACE TEMPERATURE (°C)

Figure 4 Predicted and measured droplet evaporation

times for three droplet diameters

332
NON

DIMENSIONAL

VOLUME

(V/V

0

)

Figure 5 Calculated and measured time dependent variation

of droplet volume for a 30 microliter droplet

333
TESTS

SP511,

SP514

S

SP516

(MW) aiey aseaisy }eaH

334

Figure

6

Heat

release

rate

history

for

three

crib

fires

before

and

after

sprinkler

response.
6. FUTURE WORK

Studies will continue to quantify the effects of sprinkler discharges on fire

growth. The thrust of new programs will be on the detailed studies of the

effects of low water volume discharges on fire spread rates. Work will

continue on the formulation of the suppression submodel to be accommodated in one of the zone fire models under development at CFR.

7 . REFERENCES

[1] Pietrzak, L.M., Johanson, G.A., and Ball, J.A., A Physically Based Fire

Suppression Computer Simulation for Pos t - Flashover Compartment Fires, Vol.

Ill, Simulation Formulation, Software Documentation and Experimental

Requirements, Mission Research Corporation Report MRS-R-846, 735 State Street,

P.0. Drawer 719, Santa Barbara, CA 93102, March 1986.

[2] Mitler, H.E. and Rockett, J.A., User's Guide to FIRST, A Comprehensive

Single-Room Fire Model, in preparation as an NBSIR.

[3] Walton, W.D., Baer, S.R., Jones, W.W., User's Guide for FAST, NBSIR 85-

3284, U.S. Dept, of Commerce, National Bureau of Standards, Gaithersburg, MD

20899, December 1985.

[4] Walton, W.D., ASET-B, A Room Fire Program for Personal Computers, NBSIR

85-3144-1, U.S. Dept, of Commerce, National Bureau of Standards, Gaithersburg,

MD 20899, December 1985.

[5] Evans, D. and Morehart, J., Investigation of the Effects of a Stratified

Two Layer Environment on Fire Plume Temperatures, Proceedings of the 1987

ASME-JSME Thermal Engineering Joint Conference, Honolulu, Hawaii, March 22-27,

1987, Volume I, pp . 381-386, 1987.

[6] Stroup, D. , Evans, D. , and Martin, P. , Evaluating Thermal Fire Detection

Systems, NBS Special Publication 712 (English Units) and 713 (SI Units), U.S.

Dept, of Commerce, National Bureau of Standards, Gaithersburg, MD 20899,

April 1986.

[7] Evans, D.D. and Stroup, D.W., Methods to Calculate the Response Time of

Heat and Smoke Detectors Installed Below Large Unobstructed Ceilings. Fire

Technology . Volume 22, 1, p. 54-65, February 1986.

[8] Wendt, B., and Prahl, J.M., Discharge Distribution Performance for an Axisymetric Model of a Fire Sprinkler Head, NBS -GCR- 86 - 517 , U.S. Dept, of

Commerce, National Bureau of Standards, Gaithersburg, MD 20899, October 1986.

[9] DiMarzo, M., Evans, D.D. and Trehan, A.K., The Cooling Effect of a Single

Evaporating Droplet on a Hot Semi - inf inite Metal Body, Proceedings of the 1987

ASME-JSME Thermal Engineering Joint Conference, Honolulu, Hawaii, March 22-27,

1987, Volume 1, pp . 409-416, 1987.

[10] Walton, W.D. and Budnick, E.K., Office Occupancy Sprinkler Tests, in

preparation as a NBSIR.

335
COMBUSTION BEHAVIOUR OF BUILDING MATERIALS USING

A MODEL BOX WITH AIR SUPPLIED SYSTEM

Fumiharu SAITO

Testing Lab of Center for Better Living

2 Tatehara Oho-cho, Thukuba-gun,

Ibaraki Pref, Japan

Masashi YOSHIDA

Building Research Institute, Ministrg of Construction

1 Tatehare Oho-cho, Tsukuba-gun,

Ibaraki Pref, Japan.

PREFACE

Fire behabvior of compartments depends on the relation between

the building conditions such as geometrical configuration of the

compartments and size of openings including doors and windows,

and the material conditions such as interlining materials,

furniture etc. (hereaf ter. referred to as materials),

existing in the compartments.

Recently, research on mathematical modeling of fire behavior of

compartments has been developed remarkabley, and it becomes

possible to estimate it accurately.

However, it is still difficult to estimate burning behavior on the stage of flashover which is a characteristic phenomenon of

changing of burning behavior during a steady state of compartment

fire, because the combustion mechanism of various materials at

that stage is quite complicated.

Many combustibility test methods have been developed in order to estimate the fire risk of materials. In most cases, plate shape

specimens of materials have been tested in an open space with

sufficient air supply for burning. Since these tests have not

been conducted in any limited ventilating conditions such like

those at real compartment fire, it is difficult to utilize these

test results for estimating fire behavior of compartments.

For the purpose of such estimation, it is essential to evaluate,

quantitatively, the burning behav-ior of materials in the

condition in which major factors of compartment fire are took

into consideration as parameters.

While many kind of factors which might govern compartment fire

have been pointed out. it is possible to extract major factors

from them by focusing the target which we want to clarify in a fire scenario.

336
The aim of this study is to explain the burning behavior of

materials in compartment fire.

Burning of materials in a compartment fire is a oxidizing

reaction in the limited space; therefore, the fire behavior is

controlled by the relationship between the rate of release of

combustibles from the materials and rate of air supply to the

compartment. The rate of release of combustibles is in proportion

to exposed area of the materials, On the other hand, rate of air

supply depends on the temperature on the compartment and shape of

open i ngs of it.

Temperature in a compartment under fire, fundamentally, depends

on calorific value and rate of pyrolysis of the materials

involved. However, it should be noticed that those two values

which we consider are the special values in the condition of

compartment fire, and are not the values which are described in a table of constants of materials. Furthermore, the rate of

pyrolysis is accelerated by feedback of heat within the

compartment

.

From the above point of view, a combustibility test apparatus

with a forced ventilation system has been developed, and several

materials have been tested with the variation of exposed area of

the material and rate of* air supply.

The aim of the study is to clarify the burning characteristics of

materials with relation to rate of air supply; and in addition,

toxicity of smoke generation from the burning has been evaluated

by vecording the behaviour of mice using a rotary caze installated the a ajusting switch which were exposed to the

smoke.

1. Combustion mechanism of materials in compartment fire.

When burning behavior of materials in compartment fire is to be

quantified experimentally, it is necessary to design a test

apparatus and to develop a test method suitable for the purpose

of quantification.

Once compartment fire occurs and develops, the burning behavior

changes, in general, from a stage of fuel rich combustion to a stage of ventilation control fire where the rate of pyrolysis of

materials is balanced with rate of air supply. It may be

usually considered that compartment fire immediately turns into a developed stage of fire from flashover. However, on condition of

materials involved and shape of openings, there is another stage

of fire in which the compartment is filled with smoke and only

smoke but no flame spouts from the openings. It may be called as smoky fire. And in proportion as the area of openings decrease,

the duration of the smoky fire lengthens. By investigations into

many cases of recent building fire, it has been recognized that

the smoke release period is very long because the openings of

building are rather small, it is also well recognized that rate

of release of smoke and CO gas in fire greatly affects the

337
safety of lives of residents in the building. Therefore, this

release rate is one of the most important characteristics of

materials in case of consideration into the stage of smoky fire.

Combustion of materials can be classify, in general, into

smoldering and flaming. Since the smoldering is considered as a burning behavior which occurs at the lower limit of combustion on the surface of burning materials, application of the theory of

smoldering to solve the combustion mechanism of smoky fire brings

a lot of problems with it. Namely, the factors controlling the

oxidizing reaction in a closed space filled with pyrolytic

decomposition products and with limited air supply differs from

the factors controlling the oxidizing reaction in a open space

with sufficient air supply and lower limit of combustion of

pyrolytic decomposition products.

The oxidizing reaction of combutsibles in a compartment can be

defined as the primary oxidizing reaction, and that of

combutibles released out from the compartment into surrounding

air can be defined as the second oxidizing reaction. In this

second oxidinzing reaction, the aspect of the reaction varies

widely depending on the phase of mixing of the combustibles with

surrounding air. In other words, it should be noticed that, even if the concentration of combustibles released out from the

compartment of fire is. thick enough, there is a difference on degree of oxidizing reaction between the case that the

combustibles are released into an open space with sufficient air

and the case that they are released into a limited space such as

a corridor in the building.

In general, ratio of oxidizing reaction of combustible materials

in a compartment against that at the outside of the compartment

varies depending on Ratio of Air SupplyCratio of rate of air

supply against the area of exposed surface of the material

involved in the compartment) That is; When the RAS (ratio of air

supply) is small, the compartment is filled with lots of smoke

because of shortage of air. In this case, if the temperature of

the smoke is relatively low, only the smoke but no flame runs

out from the openings; but if the temperature of the smoke is

rather high, flaming occurs at the outside of the openings. When

the ratio of air supply is large, combustion occurs much more in

the compartment than at the outside of the compartment. When a fire load of heat, smoke etc. resulting from a compartment fire

is to be quantified from the point of view of the risk of upward

fire spread in buildings, oxidizing reaction of pyrolytic

decomposition products spreading in an open space should be

investigated. On the other hand, when the fire load which flows

and spreads in a building is to be quantified, oxidizing reaction

in a limited space imaginated as a corridor etc. should be

investigated. Therefore, it is necessary to design a fire

scenario suitable for the purpose of each study.

Figure-1,2 shows an example of results of fire tests in a model

inner dimension of which is 0.84m wide, 1.68m deep and 0.84m high

with a opening adjacent to

- unlimited open space. Even if the

338
rate of decomposition of the interlining material in the

compartment model is almost constant, rate of discharge of smoke

and CO gas through the opening varies depending on the oxidizing

condition of the smoke and gas at the outside of the

comnpar tmen t . When oxidation of the smoke and gas does not

occur, rate of discharge of smoke and CO gas becomes large.

However, when the oxidation occurs, the rate of heat release

becomes large, but rate of discharge of CO gas extremely

decreases. In addtion, in case that the compartment has a large

opening or depth of it is short, the degree of oxidizing reaction

in the compartment is relatively large, because of increase of

air supply or well mixing of the pyrolytic products with

ai r. (model C)

It is presumed that ratio of degree of oxidizing reaction inside

and outside of a compartment depends on the pyrolytic

decomposition mechanism of materials involved in the compartment,

the arrangement of the materials, configuration of the

compartment and the openings etc. Accordingly, there remains a lot of problems to be solved in order to estimate the ratio of

oxidizing reaction theoretically.

From the above mentioned point of view, limitation has been made

to consider the oxidizing reaction in a compar tmen t ( pr i mary

oxidizing reaction) as a first step to elucidate the combustion

reaction of materials on compartment fire, and accordingly, some apparatus have been conductd'for the purpose of understanding the

combustion characteristics of materials with the parameter of

rate of air supply. This may cause both an evaluation, from

safetry side, of toxicity of smoke spreading in a building, and

an estimation of fire load to division construction wi'th regard

to the heat release.

2 Experiments

2. Specimen

The tested materials are shown in Table-1. Three dimensional

specimen was constructed with three side of walls and a ceiling,

and was set on the back side of the combustion chamber. There

were two types of specimens the exposed area of which was 0.154m"

and 0 . 207(11* .

2.2 Test Apparatus and Test Method

The test apparatus consists of a heat source, a combustion

chamber, an air supplying system, a mice exposure chamber and a gas analyzing systesm as shown in Figure-2.

Ceramic burners of 10cm by 10cm were used as the heat source and were set in front of the wall surface of the specimen with

spacing of 1.0cm. The mixture of 3.0 1/min. of propane gas and

14.0 1/min. of air (primary air)was suppluied to the heat

339
source. The rate of heat release of the heat source was 5„lKJ/sec. Another amount of air -supply to the combustion

chamber ( secondary air) was decided taking into acount the full

scale test data excuted by the Building Research Institute,

Ministry of Construction, Japan.

3 Test Resu 1 ts

From the masured values of gas flow rate and gas concentration

through the chimney of the apparotus, rate of heat re 1 ease ( RHR)

,

total heat release, rate of CO release(RCO) and the amount of

generation of CO were calculated by the oxygen consumption

technique. Heat release per a unit mass loss(KQ) and generation

of CO gas per a unit mass loss(KCO) were calculated by the weight

difference of the specimen between that before test and that

after test, and by the total heat release and the total CO gas

re 1 ease

.

In order to evaluate the toxicity of generataed smoke and gas,

the generated gas was taken into the mice exposure chamber at a rate of 100 1/min. ( this value was lead from the number of change

of air of 50 times/hour). And incapacitation time of the exposed

mice was measured by means of recording the movement of rotary

cages of the mice.

4. Discussion

1) Test with air supplied system

i) It has been recognized from the test results that RHRCrate of

heat release)and RCCKrate of CO release) both from materials were

in proportion to the exposed surface area of the mater i al s . (Fi

g

3,4) It has been also recognized that there was a close

relationship among the heat release, the characteristics of CO

gas release and ASRCrate of air supply per unit area of exposed

surface of the material). In other words, it has been observed

that RHR is in proportion to ASR, and that RCO is in inverse

proportion to ASR. A similar tendency has been observed in an effective heat release of materi als ( the amount of heat release

per unit pyrolytic decomposition amount) and the amount of CO

gas release, respect iveiy. (Fig 5)

ii) In the process of the oxidizing reaction of pyrolytic

decomposition products on the compartment, it is considered that

the pyrolytic decomposition mechanism of wooden materials is

relatively simpler than that of plastic materials which may

generate many kind of intermediate products. This idea seems to be supported by the test results which shows the- face that the

characteristics of heat release and CO gas release of wooden

materials has close relationship with ASR, but that the

characteristics of plastic materials has no clear relationship

wi th ASR. (Fig 5.2 , 5.3)

340
The following experimental equation has been obtained with

regard to the characteristics of heat release and CO gas

release of wooden materials:

There is a limitation on application of the equation, because KQ can not exceed the amount of heat release in complete combustion

while KQ increases in proportion to the rate of air supply.

iii) KQ is in proportion to KCO. and it's proportional

coefficient depends on the material.

2) Comparison between the air supply system and natural air

infiltration

Air supply into the compartment in real fire is made by the

buoyancy of heated gas and air in the compartment. And mixing of

air with the pyrolytic decomposition products in the compartment

is done depending on the folw of heated gas in it. On the other

hand, with a forced air supply system, the pyrolytic

decomposition products is mixed with air better than in condition

of natural air infiltration, and oxidizing reaction is done more effectively. Thus, it is necessary to underatand the

relationship between effectiveness of forced air supply system

and that of natiral air infiltration in order to apply the test

results taken under the condition of forced air supply to evaruattion of compartment fire.

J. Rockett and K. Steckler have proposed an equation to estimate

the rate of air flow into a compartment using temperature of hot

layer and opening factor under the steady state burning condition

taking into account the real compartment fire. And it is also

possible to estimate the rate of air flow into the compartment

using measured information of RHR.

Accordingly, rate of air flow into two types of model box of

0.84m x 1.68m x 0.84 and 0.84 x 0.84m 0.84m, which have openings

adjacent to open space, was measured by bi-directional pitot

tubes during fire tests. Figure-6 shows the relationship

between the measured air flow and the air flow estimated by

temperature, opening factor and RHR.

In case that burning also occurs at the outside of the

compartment in addition to inner combustiun, air flow tends to be

over-estimated by the portion of RHR made at the out side of it.

Experimental Formula (Air Supplied System)

RHR/m (KJ/s.m ) ROO/m (g/s.m ) KQ(KJ/g)

KCO ( g/g)

2.0 J6

2.0-0.014 0

0.4 0

0.35-0.004 0

0

: Air Supply Ratio = Rate of Air

Supp 1 y ( 1 /s )/Sur f ace area(m“)

341
It is impossible to evaluate the combustion reaction of

compartment fire at the inside of the compartment independent

from that at the outside. However, incase that the combustion

reaction made at the outside of compartment is very little, it is

still possible to estimate rate of air flow into it by RHR.

Moreover, according to the obsevation on the air flow at the

openi ngs during the f i re test, i t seems the most real i s t i

c

approach that to estimate the rate of ai r f 1 ow i n to the

compartment by RHR, and the estimation has been used for

discussion of the test results.

1) Rate of pyrolytic decmposition

Pyrolytic decomposition is promoted by the feedback heat made in

the compartment because the heat release in it is in proportion

to ASR. (Fig 8)

2 Character i s t i cs of heat release

Figure-10 shows the relationship between ASR and KQ. that is; KQ

increases in proportion to increase of ASR, and reaches to the

level of 80% of theoretical complete combustion when ASR exceeds

40. The line on the Figure shows the calculated results by the

experimental formula of the case of forced air supply. It is

considered that effective mixng of pyrolytic decompsition

products with air and over-estimation of rate of air flow cause increaes of the value of heat release(KQ) in case of forced air

supply in comparison with that in case of natural air

infiltration at the same ASR condition.

(3) CO gas release

When opening area is small, oxidizing reaction is made only in

the compartment because temperature of the smoke is low even if

the condition of mixture of gas and air which runs out from the

opening reaches in the flammable condition. However, when the

opening area is large, oxidizing reaction with flaming occurs at

both inside and outside of the compartment, because temmperature

and concentration of smake satisfies the energitic and component

condition for combustion. Rate of CO gas release in the latter case is smaller than that in the former case.

When the opening area is small and flaming is not observed at the

outside of the compartment, there is a good agreement between the

measured characteristics of CO release and that calculated by the

experimental formula of the case of forced air supply.

(4) Ratio of KQ and KCO

It has been recognized that there is a close relationship between

the heat release and CO gas reelase on the combustion of

materials, and that value of KQ/KCO is in proportion to ASR.

342
CONCLUSION

In the early stage of compartment fire, smoky fire, which is an oxidizing reaction in the fuel rich condition, exists, because

the temperature rise is low and the amount of air supply to the

compartment is small. And it is known that the continuation of a fire causes the collaspe of glass of windows and openings, and

that due to increase of opening area, the rate of air supply

increases and the fire becomes vigorous. The smoky fire in the

early stage of fire is one of the serious problems for the safety

of residents in buildings which should be solved as soon as possible. Therefore, this experimental- study was conducted for

the purpose of evaluating combustion of materials in the early

stage of fire. It has baeen recognized thata the evaluation of

the burning characteristics of materials utilizing the method of

the forced air supply is usseful in some conditions; that is, the case of small opening. However, there still remains many

problems in understanding all of the burning behavior of

materials as a function of ASR.

(end)

343
T<'ib!c 1 Test Malerials

AUjreviaLion Material lliiokicss(rnn) Oxygen lixJex W

IB

I'Ve isf

PF<N)

IM yvwxxJ

Pi Bor Insulation Board

PVC Poan

Po! y isosyam laic Poan

Hiemlic Poan (Resole Type)

Ilicnolic Pan (Novolac Type)

4.0

9.0

25.0

25.0

25.0

25.0

21.0

29.0

23.0

38.0

26.5

Table 2-1 Size of Specimen

tvne K ( rm \ exposed surface a. d \ Cul

)

u \ cm i c (an)

area (d) r

s'

A 30.0 25.0 11.0 0. 154 b

| D 21.5 25.0 21.5 0.20 7

Table 2-2 Specification of Test Apparatus

inner dimension 0.25X0.50X0.25 (m)

volume of chamber 0.0312 (m!)

1 inings wa 1

1 : 3 cel 1 IngP

exposed surface area 0.154, 0.207 id)

rate of air supply 230 345 460 (4 / min)

propane 3 4 / min ( 5. 1 KJ/s ) pre-mix air 14. 4 / min

test duration 10 min

Table 2-3 'Air-Supply-Ratio

exposed surface rate of air supply air supply ratio

area id) (4 /min) (4 /s • d)

Type 460 49.8

0. 154 345 37.3

A 230 24.9

Type 460 37.0

0.207 345 27.8

D 230 18.5

Table 3 Experimental Formula (Air Supplied System)'

RlIR/m 2 (KJ/s-m2

) .2.00

R00/m 2 (g/s-m 2

) 2.0-0.0140

KQ(KJ/gj 0.40

KCO(g/g) 0.35 - 0.001^

i>: Air Supply Ratio - Rate of Airsupply(2/s)/Sorface area(m 2

)

344
SAMPLE I E392Aa-05*9*l8 SAMPLE 1 B9-Ab-05*9* 1

9

345

RHR (KJ/s)

m (3/s)

t emp (*C)
Fig

1-2

Temperature

Distribution

(Mode!

Box)

Fig 3

Rate

of

Hest

Release

with

Air

Supplied

System
(H)Jd

o

CO. O

rf, o

CMo

u.

ca ^ cn

o!

c_o

f?a ro

of

CO

Gas

Generation

with

Air

Suool

ieri

System
RHR/

(

KJ/Svi)

S

o

o

KQ(

KJ/g

)

Fig 5 Relation Between the RHR, RCO, KQ, KCO vs ASR ($)

350
I

a. i a. I ^ I t.

| * ,

pjeog uoi'iGinsui poo^id WBOJ 3Ad 351

Fig

^-2

Rearranging

the

Test

Result

of

Various

Materials

with

Air

Supplied

System(l)

Td:

Time

to

Death
Ql.

V.’ >-l

F

jg

5.3

Rearranging

the

Test

Result

of

Various

Materials

with

Air

Supplied

System

Td:

Time

to

Death
o»o

"S

CO N,

ocO «=*

s*

OCM

crz CO

<af O

8

o

o

COO

Q.X

i

<

i o

Q.

oQ

Q.X

*S

os

3:

as

(s/no

(S/W m

o>

o

=3

<0

ZDCJ3

(MIDI) 'snieA poiP|non|i?o : o

353

o

*o

co

x ZT- COo 2„

a>

TZ3 o

o

o

O'c

a>c

1

o

o

COo
m

354
Air

Suppl

ied

System

G>

TsT

CO \

-e- - T3

U“3 <o

O) \C7> - 8

355

Fig

12

KQ/KCO

vs

Air

Supply

Ratio

(IB)
Air

Suppl

ied

System i 1 i_ 8

TS oU-3

Fig

12

KQ/KCO

vs

Air

Supply

Ratio

(IBJ
FIRE INDUCED FLOW IN A CLEAN ROOM WITH

DOWNWARD VERTICAL LAMINAR FLOW

BY

Osami SUGAWA*. Yasushi OKA**, and Hirofumi HOTTA***

*) Center for Fire Science and Technology. Science University

of Tokyo. 2641 Yamasaki. Noda, Chiba 278

**) Department of Architecture. Faculty of Science and

Technology, Science University of Tokyo

***) R&D Division. Nohmi Bosai Kogyo Co., Ltd. 3-53-7 Honmachi

.

Shibuya-ku 151 Tokyo

ABSTRACT

A clean room was used for this experiment which had downward

vertical flow of about 22-cm/sec. A line fire source was set

near to the center. Heat flux of the fire source was about 6KW

.

Temperature. downward and upward velocities were measured. The

equilibrium zones where dynamic pressure given by upward flow

balanced against the one given by downward flow were pursued

experimentally. Flow Visualization was also conducted to measure the location of eddies. The fire induced flow was treated

as a viscous incompressible. 2-dimensional flow. Nav i er-Stokes

equations were applied to the flow with the Boussinesq approximation. Assumptions of adiabatic condition and non-slip flow

condition at boundary were adopted. Implicit finite difference

scheme method was employed for the numerical calculation.

Distributions of temperature, velocity, and locations of the e- quilibrium zones were almost coincided with the calculated

ones

.

1 . I ntroduct i on

The interaction and behavior between fire produced flow and

ventilation flow in air conditioned or clean rooms are matters

of some concern in the area of fire safety and fire detection.

Generally. the buoyant flow from an early stage of a fire is

very weak compared with the forced ventilation air flow. The

early stages of a fire produced air flow would generally be

diluted with and included within the forced ventilation flow.

Consequently the detection of a fire is drastically delayed. To

ensure adequate fire safety, it is important to detect a fire

as early as possible. The buoyant flow from a very early stage

of fire is very weak to get to the sensor mounted on the

ceiling. The Japanese fire code requires all fire detectors to

be placed on the ceiling of rooms. However if we consider the

effects of the flow in clean room, it is likely that such

detectors would be inadequate for the efficient detection of

the early stages of. a fire. As the first step of the study on

an interaction between ventilation flow and fire induced flow,

we began with the 2

,-dimensional counter flows of upward and

357
downward flow in a clean room. Laminar downward flow is given

in the clean room, therefore, if we employ the line fire

source, the flows in the clean room could be treated as a two

dimensional flow.

2 . Exper i mental

2-1. Room: Figure 1 shows the vertical cross section of the

experimental clean room with a downward vertical flow. Clean

air supplies into the room through HEPA (High Efficiency

Particulate Air) filters set in the ceiling, and returns to the

ducts located both sides of the room. Table 1 shows the

specifications of the experimental clean room.

Table 1 Specification of the Clean Room

Size

Vent i 1 at i on

system

Down flow

veloci ty

HEPA f i 1 ters

6.4 m(W) x 9 m(L) x 3.2 m(H)

Downward Vertical Laminar Flow . . . Suppl

y

200 times/h

0.30 m/sec (nominal value)

0.22 m/sec (observed value in this time)

570 mm x 1170 mm x 42 sets, separator type

Chemical benches were

set on both sides of

the room. These were

different sizes and

shapes. There was no

symmetrical interior

arrangement in the

room.

2-2. Fire Source:

The line fire source

of 1 cm(W) x 85 cm

(L) was employed to

get a two-dimensional

fire induced flow.

Methanol was

a fuel which

plied

server

rate

ml /sec

from

at a of

An

Figure 1 The cross section of the

clean room for experiment. unit:mm

used as was supa fuel

constant

0 . 37-0 .

4

important consideration regarding the choice of the

fire source is its characteristic form of two-dimensional upward flow and also its stationary heat release rate. The line

fire source was positioned just off center, of the room approximately 297 cm apart from the right side wail. The height of

the source was about 78 cm above the free access floor. Combustion of the fuel had been kept constant for about 1 hour

before the run started.

2-3. Measurement: In order to measure the temperature in the

clean room. thirty-three 0.32 mm diameter al umel -chromel

thermocouples with 10 cm spacing between them were placed on a about 3 m movable wire. The set of the thermocouple line was

moved horizontally and vertically to cover the all regions of

358
the air flow. Figure 2 shows the positions in the room where

temperatures were measured. Another three thermocouples were

set on the HEPA filter and on both sides of wails.

A bidirectional tube was used to measure the dynamic pressure

for downward and upward flow. The tube was moved systematically

for vertically and/or horizontally. as shown in Figure 3, to

find out the locations where the dynamic pressure measured zero

indicating a balance between upward and downward flows. A

smoke-wire system was utilized to visualize the positions

where the velocity indicated almost zero, as well as indicate

the size and positions of eddies and flow patterns.

Figure 2

The positions

where temperatures were

measured by K

type themocoupl es

.

.Figure 3

The positions

where velocities

were measured.

fire source

2-4. Simulation Model

A field model was employed to describe the flow charact er i s t i cs

in the clean room. The following assumptions were made:

1) The fire induced flow and downward flow were treated as viscous incompressible fluid.

2) The driving force for convection was governed by density

differences as calculated from temperature differences.

359
3 ) Buoyant force as calculated from temperature difference

was included into Nav i er-Stokes equation.

4) Adiabatic condition for wall and floor was employed.

5) Two-dimensional flow was induced in the clean room even the

the fire source and fire induced flow were given.

6) Free access floor gave no effect to the air flow.

7) Effective thermal di f f us i bi 1 i ty and eddy viscosity were

kept constant.

8) Fire source was given at constant heat release rate of 6KW

.

The calculation was started under the initial conditions of a constant downward air flow formed in advance to the start of

the. model fire.

In this study two dimensional rectangular coordinates (x,y)

were employed with the cell size of 10 cm x 10 cm. Calculations

were based on a differential field model under the preceding

assumptions. The purpose of the present study is not to develop

a technique or method for numerical calculations, but is a computer experiment which is able to verify the full scale

experiment carried out in a clean room. The basic concept to

construct the numerical equations can be found in reference 1,2

although they deal with the fire induced flow in free boundary.

Physical parameter values used in the calculation are

demonstrated in Table 2.

Table 2

i )

l i )

l l i )

i v)

v

)

I ncompress i bl

e

Implicit

Equation of Heat Flow. : Finite difference method:

Boundary Condition

Temperature of Heat Source

Temperature Atmosphere

Down flow Velocity

Physical Properties

Effective Thermometric D i f f us i b i 1 i ty

Effective Eddy Viscosity : Finite Difference scheme

viscous fluid

450

20

0.2

C

°c

m/sec

2.2

.515

T i me I nterval : 0.1 sec

Mesh I nterval X--d i rec.t i on : 0.1 m Mesh I nterval Y--d l r ect i on : 0.1 m Solving Simultaneous Method

Maximum Iterate Counts

Criterion of Convergence

Solving Method

Type of Data

250 times

1.0 x 10

Success i ve

-5

1 0

1 0

-4

- 2 m /sec

m /sec

Over Relaxation

Double Precision

3. Results

3-1. Experimental results

a) heat Flux

The heat flux was calculated based on the consumption rate of

the fuel and its heat of combustion at 25 C to gaseous C02 and

H20. The estimated heat flux was about 6KW. The free board of

the aluminum vessel was about 2-3 mm and which was kept

constant during the experiment.

360
b) velocity 2.0 Velocities of the down

flow were measured firstly without the line

fire source. They were

Plotted in Figure 4 with

the mark of a. Down flow

velocities indicated 0.2

- 0.22 m/sec independent

of height in region of

0.55 - 1.65 m above the

free access floor. The

vertical velocities of

the down flow began to

reduce as they approaches

to the height of 0.4 m above the free access

floor. Visualization of

the down flow was determined with the sinoke-wire

and generated smoke. The

direction of the down

flow shifted clearly from

a vertical to a horizontal direction in the re- gion of 0.4m - 0.5m above

the free access floor.

This phenomena was ceused

by the pulling of the

down flow by the return

air intake for through

the ducts which were set

on the concrete floor.

These ducts were located

along both walls and were 0.6 in below the free access floor.

The center line upward flow without down flow indicated about

1.3 -1.4 m/sec independent of height in the measuring region

except just above the fire source. The velocities without down

flow are demonstrated in Figure 4 with the mark of A. Figure 4

also shows center line upward velocities with down flow by the

mark of o. The behavior of the upward velocities with and

without down flow were very similar with each other. However,

the difference in the velocities between both upward flows was

about 0.2m/sec. and which was attributed to the pushing caused

by the down flow.

c) temperature

Typical results on temperature distribution in the vertical

direction are shown in Figure 5. The center-line temperature

distribution is not included in this figure. As hot air went up

toward the ceiling, it spread out beneath the ceiling as it was

pushed down by down flow. The location, therefore, where the

maximum excess' temperature appeared dropped slightly with increasing of the horizontal distance from the center line. In

the case of without down flow, the maximum excess temperatures

in the vertical distribution at various points were observed

just under the ceiling.

1.5

Cl '-0

Oi X

0.5 -

0.0

DOWNWARD UPWARO

< 1

- 1.0

a

a

a* o

•a0.0 l.O Velocity (m/sec)

2.0

a wtthout fire source and with down flow

a with fire source and without down flow

o with fire source and with down flow

Figure 4 Variation of upward

velocity along the center line.

361
d) dynamic pressure ceiling

equ i 1 i br i um

A bidirectional tube

was moved systematically in a ver t i cal

and/or horizontal

direction. Dynamic

pressure as well as temperature at the

tube point were

examined. Probe and

thermocouple were

ma i nta i ned

pos i t i on

,

spacing

each, for 2

D i f f er ences

sure were

on an

recorder

at each

1 0 cm

between

- 3 min.

in presr ecorded

analog pen

and showed

slight deviations

from the mean value.

In the marginal zone

where both upward

and down flow were

present . the equ i 1 i

brium zone in the

dynamic pressure*

were provided where

dynamic pressure

given by upward flow

balanced against the

one given by downward flows. The

equilibrium zone are

shown in Figure 6.

i i i i i i 0.0 0.5 1.0 1.5 2.0 Distance from fire source ( m ) Figure 5 Vertical distribution of temperature in the clean room with down flow.

F : fire source point

Figure 6 Equilibrium zones where dynamic

pressure balanced against the one given by

_ _ _ , , _ , , downward flow. - :experiment. o icalculated

3-2. Computed Results

Calculation were performed for the same configurations as the

clean room where the experiments carried out. Down flow from

the ceiling was taken into account in the calculation. The heat

source position was chosen at just off-center center

corresponding to the experimental position of the fire source,

a) heat source

Temperature of the heat source for the calculation was chosen

from results of experimental measurement. Calculated heat flux

increased with time from t= 0 to 10 sec. and reached the stationary state after 10 sec showing about 6KW.

b) velocity

Figure 7 shows the calculated and experimental center line

upward velocities. Calculated upward velocities was about 0.85

m/sec in the computational section just above the fire source.

362
The closest measuring position corresponding to

velocities was about 2 cm above the fire source

the calculated

and indicated

0.5 - crease

source

0.3 - source

0.8 m/sec. Calculated velocities increase with an inin height within the region 0 - 0.3 m above the fire

In the region of

0.9 m above the fire

they appeared to be

maintained almost constant

of about 1.1 - 1.2 m/sec.

In the upper region . however. over 1 m above the

fire source calculated velocities decreased with

height

.

In the lower regions beneath the fire source. upward velocity decreased

towards the free access

floor. At 0.4 m below the

fire source. calculated

velocities were about 0.3

m/sec and measured ones

were 0.3 - 0.7 m/sec. The

results of calculated velocities seem correspond

well with the measured one.

c) temperature

Figure 8 shows the vertical

temperature distributions

of calculated temperatures.

These temperatures are

slightly higher than the

measured ones. Highest

excess temperature measured

at a distance of 0.4m from

the center line was observed

Compare with this, calculated

red 30 - center

.

t i on

both

Velocity (m/s)

Figure 7 Vartat

velocl ties

cal cul ated

on of the center 1 I ne

of exper l mena t a 1 data and ones

.

at around 10cm under the ceiling,

ones, highest temperatures occur40 cm under the ceiling at 0.4, 0.6, and 0.8m from the

The hot layer under the ceiling obtained from calculais deeper than the measured ones. The vertical profiles of

sets of temperatures, however. are very similar with each

other. We adopted adiabatic condition, therefore, no heat loss

was taken into account in the calculation. The discrepancy in

temperature between calculation and measurement elucidates that

the heat loss must be taken into account to get well concordant

results with the experiment.

d) dynamic pressure

The equilibrium zone in dynamic pressure given by the balance

between upward and downward flows were estimated depending on

the calculated velocities and densities. The flow density was

given as a function of temperature. We employed the equilibrium

zone where the ratio between the upward and downward dynamic

pressure indicate 0.9 - 1.1. Estimated equilibrium zone are

Plotted in Figure 6. There coincide with locations of measured

ones

.

363
4 . Discussion

Figure 9 shows average temperatures and velocities along the

center line from the fire source with and without down flow.

ceiling

0.0 0.5 10 1.5 2.0 Dfstance from fire source (m)

Figure 8 Vertical distribution of

calculated temperatures in the clean room.

_ 500

u

a>CD£u

>>u0

$ 100

XDaa 50 3

Go

01Da

a*CL 10

E

Qj

1 5 10 50 100 200

Height from fire source (cm)

Figure 9 Variations of temperatures and velocities

along the center line of the upward flow with and

without down flow. Yokoi’s data are also plotted.

—

1

t 1 * a c — $

KZH

A

1 A

c 1

1

A

.*

1 1 V n it n i IT p m n

• ;

I : Yokoi ( Vel. ) Temp.

ourc)ata

Vel.

Vel. without downward flow

i 1 11 HIM L-l-J

o :

:

364
The result clearly describes that excess temperature decreased

with increase of height in the measuring region of 0.05m to

0.8m above the fire source. Average velocity was approximately

constant along the center line in the region of 0 - 2m - 1.0m.

However. within the near field to the fire source, it seems to

increase with height. We also plotted Yokoi's results, reference 3 - 5. in Figure 10 who carried out the experiments without

any ventilated flows. Yokoi utilized a line fire source of

about 1 m length and 1 cm width using MeOH as a fuel. The size

and shape of the fire source he used is almost same as ours.

The behavior of temperature and velocity along the center line

are compared with Yokoi's data. It was found that our data were

in fair agreement with Yokoi's. Therefore, the upward flow we

made can be taken as a line fire source. Despite the effect of

the downward flow. our data clearly demonstrated that the

temperature and velocity distribution of the upward flow from

the fire source was conserved along the center line. The downward flow only began to effect the upward flow with increasing

distance from the fire source. In the far region from the

center line, the horizontal distribution of upward velocity and

temperature spread out widely. and their width are wider than

those obtained by Yokoi. Convection in the clean room was

clearly visualized by the -smoke-w i re method. Figure 10 shows

the calculated stream lines calculated. Even in the ventilated

enclosure with down flow, convection seems a governing flow.

Entrainment to the upward flow, therefore. is thought to be a driving force of the upward flow in the far region from the

center line of the plume.

Figure 10 Calculated stream lines, "F" is a position

of fire source

.

An experimental study on the flow behavior with ventilated flow

in a clean room has just started. We have still discrepancies

between measured values of temperature and velocity and calculated ones. It is necessary to obtain physical parameters,

such as diffusion coefficient and thermal d i f f us i b i 1 i ty in the

counter flows. to enable the calculated values to be closer

approximations to the measured conditions.

5. References

1) Hasemi Y. "Numerical Calculation of the Natural Convection in

Fire Compartment", BRI Research Paper No. 69, Feb. 1977

365
2) Morita M., Nishimoto T.. Handa T.. and Hayashi K..

"Feasibility of Numerical Computational Methods of Heat Flow

in Fire Compartment." Proceedings 8th UNJR Meeting, Tsukuba,

1985

3) Yokoi S. . "Upward Current from an Infinite Line Fire Source"

(in Japanease), Bull, of the Fire Prevention Society of Japan

Vol . 10, No. 1 ( 1961

)

4) Yokoi S. , "Kasai" Vol. 3, No. 1 (1953)

5) Yokoi S. , "Study of the Prevention of Fire-Spread Caused by

Hot Upward Flow", BRI Research Paper No. 34, Jun. 1960

366
FIRE TOXICITY AND CHEMISTRY (CONT.)

367
U.S. -Japan Panel on Fire Research and Safety, May, 1987

TOXICOLOGICAL EFFECTS OF DIFFERENT TIME EXPOSURES

TO THE FIRE GASES: CARBON MONOXIDE OR HYDROGEN CYANIDE

OR TO CAPRON MONOXIDE COMBINED WITH HYDROGEN CYANIDE OR CARBON DIOXIDE

Barbara C. Levin, Joshua L. Gurrnan1

, Maya Paabo , Laura Baier2, and Trudi Holt2 National Bureau of Standards, Gaithersburg, MD 20899

ABSTRACT

The toxicity of single and multiple fire gases is being studied to determine

whether the toxic effects of a material's combustion products can be explained

by the toxicological interactions (as indicated by lethality) of the primary

fire gases or if minor, more obscure gases need to be considered. LC 50 values

for Fischer 344 rats have been calculated for carbon monoxide (CO) and

hydrogen cyanide (HCN) (as individual gases in air) for 1, 2, 5, 10, 20, 30,

and 60 minute exposures plus relevant pos t - exposure periods using the NBS

Toxicity Test Method apparatus. The concentration- time products for the

various HCN exposures plus 24 hour pos t - exposure times are constant, but the

concentration- time products for CO decreased for the 1, 2, and 5 minute

exposures and then increased for the longer times. In exposures ranging from

5 to 60 minutes, the toxic effects of CO plus HCN were additive and, except

for the 5 minute exposures, the presence of 5% carbon dioxide (C02

) decreased

the LC 50 values of CO. These results will be used in the computer model

currently under development in the Center for Fire Research to predict the

hazard that people will experience under various fire scenarios.

1. INTRODUCTION

Experimental research is being conducted at the National Bureau of Standards

to characterize the toxicological effects of the individual and combined

primary gases produced in fires. The three main purposes of these studies

are: (1) to accumulate baseline data to determine the extent to which the

lethality of a material's combustion products can be explained by the major

gases produced in fires; (2) to develop a bioanalytical test procedure, which

minimizes the use of animals, to predict the approximate LC 50 values of

polymeric materials and to screen for extremely potent or unusually toxic

thermal degradation products; and (3) to provide input data for computer fire

hazard models designed to predict the human tenability of fire gas

atmospheres

.

In prior research [ 1 , 2

]

3 in the Center for Fire Research at the National

Bureau of Standards, LC 5

0

values for Fischer 344 rats were calculated for the

individual gases, carbon monoxide (CO), hydrogen cyanide^ (HCN), or decreased

oxygen (02

) , for 30 minute exposures plus relevant post-exposure periods using

the NBS Toxicity Test Method apparatus. Thirty minute combination experiments

1 American Iron and Steel Institute, Washington, DC

2 University of Pittsburgh School of Public Health, Pittsburgh, PA

3 Numbers in brackets refer to the literature references listed at end.

368
UJNR, May, 1987

with CO and HCN indicated that they act in an additive manner. Synergistic

effects were found when the animals were exposed to certain combinations of CO

and carbon dioxide (C02 ). A comparison of the concentrations of the major

combustion products generated from a number of polymeric materials at their

LC 5Q (30 minute exposure plus 14 day post-exposure) values with the combined

pure gas results indicated that, in most cases, the observed toxicity could be

explained by the toxicological interactions of the examined primary fire gases

- CO, CO,, and HCN. These results have provided necessary information for the

computer model currently being developed in the Center for Fire Research to

predict the hazard that people may experience under various fire scenarios.

The research to be described here was sponsored by the Society of the Plastics

Industry, Inc., the International Isocyanate Institute, the U.S. Army Medical

Research Institute of Chemical Defense, and the National Bureau of Standards.

These experiments were designed to extend the 30 minute results to other

exposure times. LC 50 values were determined for CO and HCN (as individual

gases in air) for 1, 2, 5, 10, 20, and 60 minute exposures plus relevant postexposure periods using the NBS Toxicity Test Method apparatus. The additivity

of the toxic effects of CO and HCN observed during 30 minute exposures and the

effect of 5% C02 on the LC 50 values of CO were examined for the other time

periods

.

2 . METHODS

2 . 1 Materials

:

The gases (CO, C07 , HCN, and N2

) utilized in this study were commercially supplied in various concentrations in air or nitrogen. Since

certain distributors add SO, to their HCN tanks as stabilizers, one of the HCN

gas cylinders was analyzed by both infrared analysis and BaS0 4 titration. The

results of these tests, which were confirmed by our distributor, indicated

that less than 1 ppm S02 was present in our HCN cylinders.

2 . 2 Animals

:

Fischer 344 male rats , weighing 200-300 grams, were obtained

from the Harlan Sprague - Dawley Company4 (Valkersville , MD) , Hilltop Lab

Animals (Scottdale, PA) or from Taconic Farms (Germantown, NY). They were allowed to acclimate to our laboratory conditions for at least 10 days prior

to experimentation. Animal care and maintenance were performed in accordance

with the procedures outlined in the National Institutes of Health's "Guide for

the Care and Use of Laboratory Animals." Each rat was housed individually in

suspended stainless steel cages and provided with food (Ralston Purina Rat

Chow 5012) and water ad libitum . Twelve hours of fluorescent lighting per day were provided using an automatic timer.

2.3 Animal Exposure System: The acute inhalation toxicity of the primary fire

gases alone and in combination was assessed using the chemical analysis system

4 Certain commercial equipment, instruments, materials or companies are identified in this paper in order to specify the experimental procedure

adequately. In no case does such identification imply recommendation or endorsement by the National Bureau of Standards, nor does it imply that the

equipment or material identified is necessarily the best available for the

purpose

.

369
UJNR, May 1987

and, with the exception of the one minute HCN experiments, the 200 liter

animal exposure chamber that was designed for the NBS Toxicity Test Method[3],

The animal exposure system is a closed design in which all the gases are kept

in the 200 liter rectangular chamber for the duration of the experiment. In

the one minute HCN experiments, a 4.5 liter flow-through exposure system was used in place of the 200 liter chamber. This change was deemed necessary to

reduce the amount of HCN gas used in the one minute experiments and to perform

the experiments under negative pressure to prevent an inadvertent HCN exposure

of the personnel.

2.4

Gas Analysis; Carbon monoxide and C02 were measured continuously by non- dispersive infrared analyzers. Oxygen concentrations were measured

continuously with a paramagnetic analyzer. The CO, C02 , and 02 data were recorded by an on-line computer every 15 seconds. All gases that were removed

for chemical analysis (with the exception of HCN) were returned to the

chamber. Syringe samples (100 jii) of the chamber atmosphere were analyzed for

HCN approximately every three minutes with a gas chromatograph equipped with a thermionic detector [4]. All given gas concentrations are the average

exposure values which were calculated by integrating the area under the

instrument response cur/e and dividing by the exposure time. The variability

in these square-wave exposures increased as the exposure time increased. For

example , the average percent change from the average one minute CO values was

0.4%; whereas that for the 60 minute CO values was 8.5%. For HCN, the average

percent change from the 5 minute values was 1.3%; while that from the 60

minute values was 15.1%.

2.5 Blood Analyses : Animals designated for blood analysis were cannulated.

This procedure involved anesthetizing the animals with pentobarbital sodium

(0.09 g/kg) and surgically inserting a cannula into the animal's femoral

artery. Sixteen to 24 hours later, blood samples (0.5 mi or 0.8 mi, if blood

cyanide was also measured) were taken during and following the exposures from

unanesthetized cannulated rats with syringes flushed with heparin. The

carboxyhemoglobin (COHb) and pH were measured with a Co-Oximeter 282

(Instrumentation Laboratory Inc., Lexington, MA) and a Corning 168 pH Blood

Gas System, respectively. Hydrogen cyanide in the blood was determined by a head space technique using a gas chromatograph equipped with a nitrogen/phosphorus detector.

2.6 Test Procedure : The various gases were added to the chamber until the

desired concentrations (monitored by the analytical instrumentation) were reached. A fan in the chamber assured adequate mixing of the gases prior to

exposure of the animals. Six rats were exposed in each experiment except the one minute HCN experiments, in which 5 rats were exposed per test. Animals

were placed in restrainers which were then inserted into the portholes located

along the front of the exposure chamber such that only the heads of the

animals were exposed. To achieve square-wave exposures, these portholes were fitted with rubber stoppers during the time that the gas concentrations were reaching equilibrium. Insertion of the restrained animals caused the rubber

stoppers to fall into the chamber and simultaneously exposed the animals to

the chamber atmosphere. Exposures were for set exposure times of 1 , 2, 5, 10,

20, 30, and 60 minutes. Blood was taken from the cannulated animals at zero

time, at various times throughout the exposure and just before the end of the

370
UJNR , May, 1987

exposure (one or two animals were cannulated in each experiment) . The

toxicological endpoint was death. The percentage of animals dying at each gas

concentration was plotted to produce a concentration- response curve from which

LC 50 values were calculated for each exposure time and for the exposure plus

post-exposure period. The L

C

50 in these cases is defined as the concentration

of test gas in the chamber (ppm or %, where 1% - 10000 ppm) which caused 50%

of the animals to die during the exposure only or during the exposure plus the

post - exposure observation period. The LC 50 values and their 95% confidence

limits were calculated by the statistical method of Litchfield and Wilcoxon

[5], All animals (including the controls) were weighed daily from the day of

arrival until the end of the 14 day post - exposure observation period except

for the cannulated ones which were sacrificed following exposure. In some

cases where the animals were still losing weight at the end of the 14 day

post - exposure observation period, they were kept until they either recovered

(as indicated by three days of successive weight gain) or died.

2.7' Experimental Design for Individual and Combined Gas Studies : The LC 5

0

values and their 95% confidence limits were determined for the individual

gases - CO and HCN for 1, 2, 5, 10, 20, 30 and 60 minutes. Although 30 minute

data have been published previously [2], some of these values were reexamined.

In the case of the combined CO and HCN experiments, the additivity of the

toxic effects of the two gases was examined by exposing the animals to

combinations of approximately 50% of the LC 5

Q

values of each gas for each time

period. In the case of the combined CO and CO., experiments, the concentration

of CO, was kept constant at roughly 5%, while the concentration of CO was varied to determine the effect on the LC eo for each exposure period.

2.8 Kinetic Blood Uptake Studies : Some experiments were designed to determine

the effect of the presence of other gases on the rate of formation and final

equilibrium level of COHb , blood cyanide and arterial pH. In these

experiments, all six rats were cannulated and exposed to pre-set

concentrations of the gases. Blood (0.5 m2) was sampled at approximately one minute intervals sequentially from different rats such that an excessive

amount of blood was not taken from any one rat.

3 . RESULTS

3,1 Carbon Monoxide: Two sets of LC 50 values and 95% confidence limits for CO

in air were calculated for each of the different exposure times ranging from

one minute to 60 minutes. These LC 50 values were obtained with animals

supplied by the Harlan Sprague - Dawley Company (pre-1986 30 minute values),

Hilltop Lab Animal Company (all other pre-1986 values) or from Taconic Farms

(1986 and later values) and are shown in Table 1. The two sets of values are plotted in Figure 1 along with the plot of the CO concentration- time products,

which are calculated by multiplying the LC 5

0

values by their respective

exposure times. The COHb equilibrium levels and the rate of formation of COHb

depends upon the test concentration to which the animals are exposed (Fig. 2).

3.2 Hvdrogen Cyanide : The LC 50 values and their 95% confidence limits for HCN

in air for the exposure times ranging from one to thirty minutes plus a 24

hour post- exposure observation period are shown in Figure 3 and Table 2. The

371
UJNR , May, 1987

Table 1. LC 50 Values® for 1Carbon Monoxide Exposures for Various Times

Exposure

time C0b oo

o (min) (ppm) (ppm)

1 107000 (102000-112000

2 36000 (34000-38000) 42500 (42000-43000)

5 12600 (12400-12800) 14000 d (12000-16000)

10 7600 (7300-7900) 9800 (9200-10500)

20 5600 (5400-5800) 7400 (7100-7700)

30 4600 (4400-4800) 6600 (6100-7300)

60 3400 (3100-3600) 4900 (4600-5200)

a. Calculated using deaths within exposure plus 24 hours

b. Data prior to 1986

c. 1986 data

d. Calculated via the Litchfield and Wilcoxon procedure for significantly

heterogeneous data [5].

60 minute LC 50 value does not have 95% confidence limits because it was not

statistically calculated, but rather was estimated from the results of six

experiments. In these exposures (78, 84, 89, 90, 100, and 100 ppm), 25% of

the animals died at the three lower concentrations . and 100% of the animals

died at the three higher concentrations. Since the concentration- response

curve was so steep, it was deemed unnecessary to try to obtain further data

points between the 25% deaths and the 100% deaths.

Table 2. Acute Inhalation Exposures to HCN

Exposure Lethality Incapacitation

time LC 5 0

S EC 50

b (min) (ppm) (ppm)

1 3000 (2500- 3600) c 1700 (1200-2400)

2 1600 (1400-1800) 1100 (800-1300)

5 570 (460-710) 390 (310-490)

10 290 (250-340) 170 (150-190)

20 170 (160-180) ND

30 110 (95-130) ND

60 90 ND

a. LC 50 based cn deaths occuring within exposure plus 24 hours postexposure . b. EC 50 based on lack of righting reflex of animals immediately following

exposure

.

c. 95% confidence limits.

ND . Not determined.

372
UJNR , May, 1987

The values which had been determined prior to 1986 were checked and had not

changed. Also shown in Table 2 are the concentrations of HCN which caused

incapacitation (as indicated by the lack of the righting reflex) in 50% of the

animals following exposures of 1 to 10 minutes. Time-weighted average blood

cyanide levels from representative exposures are shown in Figure 4.

3.3 Carbon Monoxide and Hydrogen Cvanide : Prior thirty minute experiments on

the lethal effects of combinations of CO and HCN in air indicated that these

two gases act in an additive manner such that if:

[COJ

LC 5 o C0 ( 3 0 min)

thcni

lc 50 hcn ( 30 min)

* 1+15% ( 1 )

some of the exposed animals will die (the values in brackets are the test

concentrations of the gases) (Table 3). If the left hand side of the above

equation is greater or less than this value, all the animals will die or live,

respectively. When this equation is equal to 1, 50% of the test animals would

be expected to die. However, as the concentration- response curves of CO and

HCN as individual gases are very steep, the deaths of 1 to 5 animals during

these 30 minute experiments were considered indicative that the concentrations

of the combined gases were close to the LC 50 values. The estimated + 15%

variability is based on the observation that the number of deaths within the

range of approximately 0.98 and 1.10 varied from 1 to 5 and the scarcity of

information for the range between 0.86 and 0.98 (Table 3).

Table 3 . Carbon Monoxide and Hydrogen Cyanide - 30 Minute Exposures

Exp

.

Fraction of lc 50 No. died COHb

No. CO HCN CO + HCN Six animals (%)

1 1.00 0 1.00 3 84*

*

2 1.00 0 1.00 5 83

3 0.75 0.27 1.02 4 79*

4 0.75 0.38 1.13 6 67*

5 0.52 0.58 1.10 3

- 6 0.50 0.53 1.03 5 67*

7 0.49 0.49 0.98 5 71*

8 0.38 0.48 0.86 0 - 9 0.28 0.78 1.06 3 45*

10 0.27 0.79 1.06 3 43*

11 0 0.98 0.98 2

- 12 0 1.00 1.00 1

- 13 0 1.08 1.08 4 -

LC50 of CO - 4600 ppm

LC 50 of HCN - 160 ppm

* Animal died during exposure.

LC 5q values used in 'these calculations were based on deaths within the 30 min

exposures only.

373
UJNR , May, 1987

When approximately 50% of the LC 50 values of each of the different time

exposures (10, 20 and 60 minutes) of CO and HCN were combined, the number of

animals that died out of the six exposed was between 2 and 4 (Table 4). In

the 5 minute experiments, the HCN concentrations were kept constant between 47

- 53% of the 5 minute LC 50 value and the CO was varied to determine the effect

of HCN on the CO LC 50 value. This new CO (with HCN) LC 50 value was 7600 ppm,

which, if doubled, would fall within the 95% confidence limits (12000 - 16000

ppm) of the CO without HCN LC 5

0

five minute value of 14000 ppm.

Table 4. Additivity of Carbon Monoxide and Hydrogen Cyanide Toxicity

Exposure

t ime

(min) CO

Percent of LC 5

0

a HCN Total

No . an

died/ti

10 52 47 99 3/6

52 47 99 2/6

20 51 60 111 4/6

50 57 107 4/6

60 51 56 107 2/6

53 56 109 4/6

a. 1986 data

Table 3 also shows that as [CO] decreases and [HCN] increases; the animals die

at lower concentrations of CO and, consequently, COHb levels lower than if the

CO were given alone. A standard curve developed in our laboratory of

equilibrium levels of COHb generated at various atmospheric CO concentrations

showed that the COHb values obtained in the combined CO and HCN experiments

are systematically lower than those expected at equilibrium from comparable CO

exposures (data not shown) . The degree of reduction of COHb depends on the

HCN concentration. Polymeric materials showed similar results, in that

animals exposed to the thermal decomposition products from polymeric materials

which produce relatively high HCN concentrations have lower levels of COHb

than those predicted by the atmospheric concentrations of CO.

This lower - than- expected level of COHb in the presence of both HCN and CO was examined to determine if the HCN was affecting the initial rate of uptake and

formation of COHb or the final equilibrium levels. Figure 5 shows that at the

concentrations tested (1470 ppm of CO alone or 1450 ppm CO plus 100 ppm HCN),

the initial rate of formation of COHb was the same in the presence and absence

of HCN. The final COHb equilibrium, however, was lower when the animals were exposed to both gases.

3.4 Carbon Monoxide and Carbon Dioxide : In 30 minute experiments, sublethal

374
UJNR , May, 1987

concentrations of C02 (1 to 17. 3% 5

) in combination with sublethal levels of CO

(2500 to 4000 ppm) caused deaths of some of the rats both during exposures and

the following 24 hours (Fig. 6) [1]. The solid line in Figure 6 was drawn to

separate the experiments in which some animals died from those with no deaths.

Based on this line, the most toxic combination of these two gases appears to

be approximately 5% C02 and 2500 ppm CO. Above and below 5% C02 (1 to 17.3%), more CO (but still concentrations below the 95% confidence limits of the CO

LC 5

0

value) was necessary to produce the deaths.

The LC 50 values of CO for the time exposures of 5 , 10, 20 and 60 minutes were determined in the presence and absence of approximately 5% C02 and, with the

exception of the 5 minute case, the LC 50 values were statistically

significantly lower (based on the 95% confidence limits) in the presence of 5%

C02 than in its absence (Table 5). In the 5 minute exposures, the CO LC 50

value in the presence of 5% C0-, was lower than in the absence of C02 , but the

95% confidence limits overlapped.

Table 5. LC 50 Values for CO Alone and CO Plus C02 Exposure

time

(min)

C0a

( DDro ) CO + 5% C02

a (rum)

5 14000 (12000-16000) 11700 (10800-12700)

10 9800 (9200-10500) 7800 (7500-8200)

20 7400 (7100-7700) 5600 (5500-5800)

30 6600 (6100-7300) 3900 (3400-4500)

60 4900 (4600-5200) 3100 (2800-3400)

a. 1986 data

4. DISCUSSION

The atmosphere produced by a fire is a dynamic, rapidly changing combination

of toxic gases, particulates, reduced oxygen and high temperatures. Each

component of this combination is capable of producing conditions which are incompatible with life or will act to prevent safe egress from a burning

building. As statistics have shown that the majority of fire fatalities

result from smoke inhalation and not from burns [29], the importance of the

identification of the responsible combustion products produced from materials

under different fire conditions becomes apparent. Of equal importance is

information on the toxic concentrations of these gases for various exposure

times

.

The toxicologist is faced with the problems of determining the toxicity of the

individual combustion products and their combinations and how these toxic

effects are affected by changing exposure times. Many studies have been

5 Even though 17.3% C02 was not tested by itself, it is considered non- lethal because 17.7% C02 in the presence of 3200 ppm of CO did not produce any

deaths (Fig . 6 )

.

375
UJNR, May, 1987

performed on the toxicity of single common fire gases, e.g., see review by

Kimmerle [6] plus papers on CO [1,2,6,7,8,9,10,11,12]; C02 [1,2,9, 12]; HCN

[2,6,8,10,11,13]; and 02 [2,7,10,11,12]. Some studies have examined the

toxicological interactions of various combinations of these fire gases or the

additional effects of decreased oxygen atmospheres, e.g., CO and HCN

[2,6,10,13,14,15,16,17,18,19,20,21,22,23] CO and C02 [1,12,24,25,26,27, 28];

C02 and 02 [2,12]; CO and 02 [7,10,12] ; HCN and 02 [10]; CO, C02 , and 02 [2,10,12,27]; and CO, C02 , HCN, and 02 [10].

Prior research at N3S has examined an initial set of primary fire gases - CO,

C02 , HCN, and low 02

- as both individual gases and in various combinations.

Our toxicological results on 30 minute exposures of rats to these single gases

showed good agreement with values found in the literature (Table 6) . In our determination of the LC^ 0 values for CO for the different exposure times, we found that the values shifted when the animal supplier changed even though

both sets were designated Fischer 344 male rats. Therefore, it is important

in any set of experiments to determine the LC 5

0

values for the specific

animals being used and to check periodically the values to ensure that they

have not changed.

Table 6. Thirty Minute Lethal Exposures to Individual Gases

Animal

Gas Species LCc Q

Post

Exposure

Time Reference

CO (ppm) rats 5000-6000“ 0 9 rats 5500 NI 6 rats 4600b 0

f 2 rats 5000= 0

£ 2 rats 5000d 0 20

rats 5275® 24 hrs 8 19

rats 6600b 0

f this pape

mice 3570 NI 31

mice 3500 10 min 11

HCN (ppm) rats 142 NI 6 rats 200 NI 6 rats 160b 0 2 rats 110b 24 hrs 2 rats 157® 0 23

rats 116® 24 hrs 8 19

mice 166 10 min 11

a

.

estimated from data provided

b. square-wave exposure • c

.

gradual exposure

d. amount needed to kill in 30 min

e . calculated from CT product

f

.

animals kept 14 days, but all deaths occurred during the 30 min expo

g- animals kept 24 hours , but no change s were observed after one hour

.

NI

.

not indicated.

376
UJNR, May, 1987

Haber's rule states that the toxic concentration multiplied by the exposure

time is equal to a constant. If this rule held true for all cases, one could

examine one exposure time and extrapolate to all other times. There are very

few cases in which Haber's rule has been shown to hold true, and the reason is

that living organisms have the ability to metabolize and excrete most drugs

and toxicants before they accumulate to fatal levels. An interesting finding

from the current study, however, is that the concentration- time product for

the HCN LC 50 s is a constant if the 24 hour post-exposure deaths are included

in the LC 50 calculation. On the other hand, as expected, the CO LC 50

concentration- time product increases for the exposure periods ranging from 5 minutes to 60 minutes. An unforseen result was the initial decrease in the CO

concentration- time products for the 1, 2, and 5 minute exposure times. This

probably means that the animals need more than 5 minutes to eliminate these

high levels of CO.

We had found previously that the toxicities of CO and HCN were additive if

calculated as ratios of their 30 minute lethal concentrations. Our research

reported here shows that the additivity observed during the 30 minute

experiments is also seen in experiments in which the exposure times are 5, 10,

20, and 60 minutes. Although different experimental designs were used by

Hartzell et al . [18], Kishitani and Nakamura [10], Lynch [19], Smith et al

.

[22], Pitt et al. [21], and Yamamoto and Kuwahara [23], these investigators

also concluded that the toxicological effects of CO and HCN were additive.

Kimmerle [6] and Moss et al . [20], on the other hand, claimed that CO and HCN

showed synergistic effects. They hypothesized that the synergism they

observed in their experiments could arise from an HCN- induced increase in

respiration rate which would cause a more rapid uptake of carbon monoxide.

Results reported by Matij ak- Schaper and Alarie [11], however, show that HCN

causes a significant decrease in the respiration rate of mice. This decrease

became apparent three minutes after the start of an exposure to 100 ppm of

HCN. As little as 20 ppm of HCN caused a 20% decrease in the respiratory rate

and 63 ppm caused a 50% decrease. These investigators also showed that CO

produced a respiratory rate depression, although the onset was later than that seen with HCN. Therefore, the two gases together are unlikely to increase the

CO uptake rate. Our examination of this issue showed that as the CO

concentration decreases and the HCN concentration increases, the COHb level

decreases (Table 3) . This decrease is actually lower than the COHb

equilibrium level expected from the comparable CO concentration. Our

investigation of the kinetic uptake of CO and the formation of COHb in the

presence and absence of HCN showed that the initial rate of formation of COHb

is the same; however, the final COHb equilibrium level is reduced when HCN is

present (Fig. 5). In other words, the animals are not receiving more CO in

the presence of HCN, but less. This depressive effect of HCN on CO uptake may

help explain why some people dying in fires have low COHb levels (<50%

)

[14,15,16,17,29].

Higgins et al . [13], however, did not observe any statistically significant

changes in the 5 minute LC 50 value of HCN when rats were also exposed to

enough CO (2100 ppm) to generate 25% COHb. Their 5 minute HCN LC 50 value

(503 ppm with 95% confidence limits of 403 to 626 ppm ) agrees well with our

five minute LC50 value of 570 ppm with 95% confidence limits of 460 to 710

ppm. When they exposed rats to 2100 ppm of CO plus HCN, the LC 5

0

value of HCN

377
UJNR , May, 1987

became 467 ppm with 95% confidence limits of 395 to 553 ppm. Calculating

equation (1) using exposure concentrations of HCN and CO of 503 and 2100 ppm,

respectively and 5 minute LC5 0 values of HCN and CO of 503 ppm and 14000 ppm,

respectively results in a value of 1.15, which is within the experimental

error of the calculation.

There is little human data that one can compare with the animal results. One

such set of data comes from Anderson and his coworkers [14,15,16,17] and their

studies of fire deaths in the Glasgow area and the United Kingdom. They found

that 78% of the fire deaths had blood cyanide levels greater than normal.

Although they considered HCN a contributory factor, they did not find any

evidence of the additivity of the toxicities of CO and HCN. Their conclusions

are based on the assumption that 50% COHb is the lethal concentration and

since high levels of blood cyanide were usually found in conjunction with high

levels of CO, they concluded that these deaths (>50% COHb) were due to CO.

If, however, as found in our animal studies, 50% COHb is not sufficient to cause death, then reevaluation of their data may show the additive effects of

CO and HCN.

A detailed study has been made on the interaction of CO and C02 [1]. Pryor et

al. [12] found that 40% (400,000 ppm) C02 was necessary before one rat died in

a 30 minute exposure, an indication that C02 is relatively non- toxic.

However, except for the 5 minute exposures, our experiments indicated that the

presence of 5% C0 2 decreased the CO LC 5

0

values for the exposure times tested.

This synergistic interaction appears to be due to two mechanisms: 1. C02 , at

low concentrations, acts as a respiratory stimulant, which increases the rate

of uptake of CO (Fig. 7); and 2. CO and CO, together produce an increased and

prolonged respiratory and metabolic acidosis (Fig. 8). Acidosis is a disturbance of the acid-base balance of the body and, in this case, results

from an increase in acidic chemicals (e.g., carbonic acid from the reaction of

C02 with H2 0 and lactic acid as a result'of anaerobic metabolism). Depression

of the central nervous system is the major effect of acidosis. pH values

below 7.0 will cause disorientation and coma. If such severe acidosis is not

treated, death will occur [30].

5. SUMMARY

1. Carbon monoxide LC 50 values (the concentration of CO necessary to kill 50%

of the Fischer 344 rats during the exposure) were determined for exposure

times ranging from 1 to 60 minutes and found to be 107000 ppm for 1 minute,

42500 ppm for 2 minutes, 14000 ppm for 5 minutes, 9800 ppm for 10 minutes,

7400 ppm for 20 minutes, 6600 ppm for 30 minutes, and 4900 ppm for 60 minutes.

These values were higher than values determined previously with rats from a different supplier.

2. Hydrogen cyanide LC 5

0

values (calculated based on deaths during the

exposure plus a 24 hours post - exposure observation time) were. 3000 ppm for 1 minute, 1600 ppm for 2 minutes, 570 ppm for 5 minutes, 290 ppm for 10 minutes,

170 ppm for 20 minutes, 110 ppm for 30 minutes, and 90 ppm for 60 minutes.

378
UJNR , May, 1987

3. The CO concentration- time products decreased as the time increased from 1 to 5 minutes and increased from 5 to 60 minutes; whereas, the HCN

concentration- time products were constant throughout all the tested times.

4. The toxicity (as indicated by deaths of the rats) of the combined CO and

HCN gas exposures were additive for the tested times ranging from 5 to 60

minutes

.

5. The combined CO and C02 gas exposures indicated that 5% C02 decreased the

LC 50 values (i.e., increased the toxicity) of the CO for the tested times

ranging from 10 to 60 minutes. The decrease in the 5 minute LC 50 exposures

were not statistically significant.

6.

REFERENCES

1. Levin, B. C., Paabo , M., Gurman, J. L. , Harris, S.E., and Braun, E. (1987) . Toxicological interaction between carbon monoxide and carbon

dioxide. Toxicology (In Press).

2. Levin, B.C., Paabo, M., Gurman, J.L., and Harris, S.E. (1987). Effects

of exposure to single or multiple combinations of the predominant toxic

gases and low oxygen atmospheres produced in fires. Fundam. Appl.

Toxicol. (In Press).

3. Levin, B. C., Fowell, A. J., Birky, M. M. , Paabo, M., Stolte, A., and

Malek, D. (1982). Further development of a test method for the

assessment of the acute inhalation toxicity of combustion products

.

NBSIR 82-2532, National Bureau of Standards, Gaithersburg, MD

.

•4, Paabo, M., Birky, M.M., and Womble, S.E. (1979). Analysis of hydrogen

cyanide in fire environments. J. Comb. Tox. 6 , 99-108.

5. Litchfield, J. T. , and Wilcoxon, F. (1949). A simplified method of

evaluating dose-effect experiments. J. Pharmacol. & Exp. Therapeut. 96,

99-113.

6. Kimmerle, M. G. (1974). Aspects and methodology for the evaluation of

toxicological parameters during fire exposure. J. Fire & Flamm . / Comb.

Tox. 1, 4-51.

7. Cagliostro, D. E. and Islas, A. (1982). The effects of reduced oxygen

and of carbon monoxide on performance in a mouse pole-jump apparatus.

J. Comb. Tox. 9, 187-193.

8. Hartzell, G. E. , Priest, D. N. , and Switzer, W. G. (1985). Modeling of

toxicological effects of fire gases: II. Mathematical modeling of

intoxication of rats by carbon monoxide and hydrogen cyanide. J. Fire

Sci. 3, 115-128.

379
UJNR, May, 1987

9. Herpol, C., Minne , R, , and Van Outryve E. (1976). Biological evaluation

of the toxicity of gases produced under fire conditions by synthetic

materials. Part 1: Methods and preliminary experiments concerning the

reaction of animals to simple mixtures of air and carbon dioxide or carbon monoxide. Comb, Sci. Technol. 12, 217-228.

10. Kishitani, K. , and Nakamura, K. (1979). Research on evaluation of

toxicities of combustion gases generated during fires. In "Fire

Research and Safety" (Ed. by M.A. Sherald) , National Bureau of

Standards Special Publication 540; Proceedings of the Third Joint Panel

Conference of the U.S.- Japan Cooperative Program in Natural Resources

held at the National Bureau of Standards, Gaithersburg, MD (March, 1978)

11. Matij ak- Schaper , M. and Alarie, Y. (1982). Toxicity of carbon monoxide,

hydrogen cyanide and low oxygen. J. Comb. Tox. 9 , 21-61.

12. Pryor, A. J., Fear, F. A., and Wheeler, R. J., 1974: Mass Life Fire

Hazard: Experimental Study of the Life Hazard of Combustion Products in

Structural Fires. J.Fire & Flam. /Comb. Tox., 1, 191-235.

13. Higgins, E. A., Fiorca, V., Thomas, A. A., and Davis, H. V. (1972).

Acute toxicity of brief exposures to HF, HCi, N02

, and HCN with and

without CO. Fire Tech. 8, 120-130.

14. Anderson, R. A., Thomson, I., and Harland, W. A. (1979). The importance

of cyanide and organic nitriles in fire fatalities. Fire & Matls . 3,

91-99. . . . ;

15. Anderson, R. A., Watson, A. A., and Harland, W. A. (1981a). Fire deaths

in the Glasgow area: I. General considerations and pathology. Med. Sci.

Law. 21, 175-183.

16. Anderson, R. A., Watson, A. A.,, and Harland, W. A. (1981b). Fire deaths

in the Glasgow area: II. The role of carbon monoxide. Med. Sci. Law 21,

288-294.

17. Anderson, R. A., Willetts, K. N., Cheng, K. N., and Harland, W. A.

(1983). Fire deaths in the United Kingdom 1976 - 82. Fire & Matls. 7 , 67-72.

18. Hartzell, G. E., Switzer, W. G., and Priest, D. N. (1985). Modeling of

toxicological effects of fire gases: V. Mathematical modeling of

intoxication of rats by combined carbon monoxide and hydrogen cyanide

atmospheres. J. Fire Sci. 3, 330-342.

19. Lynch, R. D. (1975). On the non-existence of synergism between inhaled

hydrogen cyanide and carbon monoxide. Fire Research Note No. 1035, Fire

Research Station, Borehamwood, Hertfordshire, WD6 2BL, England.

20. Moss, R. H. , Jackson, C. F. , and Seiberlich, J. (1951). Toxicity of

Carbon monoxide and hydrogen cyanide gas mixtures. Arch. Indust.. Hyg.

Occup. Med. 4 , 53-64.

380
UJNR May, 1987

21. Pitt, B. R. , Radford, E. P., Gurtner, G. H. , and Traystman, R. J.

(1979) . Interaction of carbon monoxide and cyanide on cerebral

circulation and metabolism. Arch. Environ. Hlth. 34, 354-359.

22. Smith, P. W. , Crane, C. R. , Sanders, D. C., Abbott, J. K. , and Endecott,

B. (1976). Effects of exposure to carbon monoxide and hydrogen cyanide.

In the Proceedings of an International Symposium on The Physiological

and Toxicological Aspects of Combustion Products held at The University

of Utah, March, 1974. Published by the National Academy of Sciences,

Washington, DC, pp . 75-88.

23. Yamamoto, K. , and Kuwahara, C. (1981). A study on the combined action of

CO and HCN in terms of concentration- time products. Z. Rechtsmed. 86,

287-294.

24. Crane, C. R., Sanders, D. C., Endecott, B. R. , Abbott, J. K. , and Smith,

P. W. (1977). Inhalation toxicology: 1. Design of a small animal test

system. II. Determination of the relative toxic hazards of 75 aircraft

cabin materials. Report No. FAA-AM-77-9, Dept, of Transportation,

Federal Aviation Administration, Washington, DC.

25. Edginton, J. A. G., and Lynch, R. D. (1975). The acute inhalation

toxicity of CO from burning wood. Fire Research Note No. 1040. Fire

Research Station, Salisbury, Wiltshire, U. K.

26. Gaume , J. G., Bartek, P., and Rostami, H. J. (1971). Experimental

results on time of useful function (TUF) after exposure to mixtures of

serious contaminants. Aerospace Med. p. 987-990.

27. Rodkey, F. L. and Collison, H. A. (1979). Effects of oxygen and carbon

dioxide on carbon monoxide toxicity. J. Comb. Tox. 6, 208-212.

28. Nelson, G. L. , Hixon, E. J., and Denine , E. P. (1978). Combustion

product toxicity studies of engineering plastics. J. Comb. Tox. 5, 222-

238.

29. Birky, M. M. , Halpin, B. M. , Caplan, Y. H., Fisher, R. S., McAllister,

J. M. , and Dixon, A. M. (1979). Fire fatality study. Fire 6: Matls. 3,

211-217.

30. Guyton, A.C. (1976). Textbook of Medical Physiology, Fifth Edition,

W.B. Saunders Co., Philadelphia, London, and Toronto.

31. Hilado, C.J. and Cumming, H.J. (1977). Effect of carbon monoxide in

Swiss Albino Mice. J. Comb. Tox. 4, 216-230.

381
HCM

(LC50

VALUES

(ppml

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100

*0

I

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<

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o

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200

ISO

120

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I

i

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vu

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Figure 1. ^50 v *lues and cone entr at ion- time products for CO for various exposure times. Figure 2 Rate of formation of COEb from exposures to various

atmospheric concentrations of CO (ppm « CO concentration).

1 2 5 10 20 50 100

TIME (min)

I

\

J

IUa

z

<

>u

a

o

o

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bU><

AVERAGE HCN CONCENTRATION (ppm)

Figure 3. HCN LCjq values for different exposure

tunes plus a 24 hour post- exposure period 382

Figure 4 . Time -weighted average blood cyanide levels in animals

exposed to LCjq values of HCN for various exposure times
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TOXICOLOGICAL EFFECTS OF DIFFERENT TIME EXPOSURES TO FIRE GASES; CO OR HCN OR

CO COMBINED WITH HCN OR C02 B.C. Levin, National Bureau of Standards, USA

VOICE; This may be somewhat related to the next topic, but what's the

concentration level of combination of CO and C02 ?

LEVIN: They can vary. Based on the curve seen in Fig. 6, you can see the

variation. For example, if the CO were 3500 ppm and the C02 was 4 percent,

you should have deaths

.

TSUCHIYA: What is the oxygen concentration?

LEVIN: In all of these pure gas cases, we have kept our oxygen concentration

at about normal, ambient level. Sometimes in our material decomposition,

because it is a closed system and the animals use some oxygen and the

combustion uses some oxygen, it may drop to about 18 percent in some cases.

VOICE: Could you clarify that? When you add five percent C02

, what is the

actual molar percentage of oxygen in that test? When you add just C0 2

, let's

say?

LEVIN: Well, what we're doing here is using cylinders of C02 in oxygen, so that we're changing the nitrogen concentration in the system.

VOICE: You're substituting C02 for nitrogen. Thank you.

TSUCHIYA: I also appreciate this study. I found that these show synergism

of CO and C02

. This is the only case so far statistically confirmed with fire

gases. I appreciate this result. Second point, I'd like to see your slide on

CT product (Fig. 1).

LEVIN: This one?

TSUCHIYA: Yes, that's correct. You mentioned that in the case of CO, that it

does deviate from Haber's rule?

LEVIN: Right.

TSUCHIYA: Is this finding correct?

LEVIN: That's correct.

TSUCHIYA: And the second point, about HCN, you mentioned that CT is constant.

LEVIN: It appears to be when the LC 50 values include the deaths that occur within the first 24 hours following the exposures.

384
TSUCHIYA: But rather, I think, this is again normal phenomena. I studied

many sets of animal data presented by other people, and I found for Haber's

Rule it may be written: ACT =1. A is a proportionality constant, C is the

concentration, T is the time , okay?

So we know that actually when experimental data deviate from this rule at both

ends, the high concentration or very low concentration, and we modified

Haber's Rule into this form, then we found experimental data can be better

represented and we found values of C o and T0 in many cases, with many

different gases. But proof of these, biologists like you, need to give

physical meaning to these constants.

This case (HCN) is a special case of difference, a special case where To is 0

.

I see the data that way, because I studied other sets of animal data, and

those data fit well in this type of an equation. It is normal for the case of

C Q and other gases, that Co and To are positive values, but in the case of

hydrogen cyanide, To is quite often zero, so this is one of those cases.

LEVIN: Thank you very much.

TEWARSON: On that slide, the primary equation that you wrote, it says mCO

divided by C02 minus b, where b is the intercept. Now if I look at it here,

is that intercept the LC 50 of C02 ?

LEVIN: No. m is the slope and b is the intercept of these two lines, and

those numbers change depending on whether C02 is above or below 5 percent. I have that in a paper I'll send to you. The reference is Levin et al .

,

Fundamental and Applied Toxicology . Vol. 9, pp . 236-250, 1987.

385
TOXIC GAS TEST BY THE SEVERAL PURE AND MIXTURE GASES

Toshirou Sakurai

Research Institute of Marine Engineering

5-12, Fujimicho 1-chome, Higashimurayama city

Tokyo 189 Japan

Abstract

The relationship between incapacitation time and apnoea time of mice and gas

concentration in case of single toxic gas condition and mixed gas condition with 6

types of toxic gases ( CO, COz, HC 1 , HCN, NH

3

, NO 2) and N2 gas was investigated for the

purpose of establishing an estimating method for toxicity of combustion gases generated from interior materials of building fire.

1. Test gases and mice

Six type gases were used; those were CO, CO 2 , HC1, HCN, NH 3 and N0 Z . Those gases except HCN were obtained as pressurized or liquified highly pure gases.

HCN gas was generated from KCN. At first, the tests were carried out with several concentration of the single

gas. After those tests, the toxicity of some combinations of the toxic gases was examined. Tow levels of concentration of each gas in the mixture were chosen so that

by the concentration at the single gas condition, incapacitation time of mice came

to 10 minutes or 20 minutes. However, the concentration of CO 2 was set at 2,5 and 8%

Also, the relationship between incapacitation time of mice and ambient temperature

was investigated.

The mice used in the study were female of ICR type of Jcl mice; the weight was

wi thin 18(gf) to 22(gf)

.

2. Test method and test apparatus -

Toxic gases are mixed in the gas controller, and the mixture is led to mice

exposing chamber of the combustion gas toxicity test apparatus (The Notification No,

1231 of the Ministry of Construction). Eight mice are put in rotating cage individually and exposed to the toxic gases in the exposing chamber. Incapacitation time of

each mouse is recorded. Apnoea time of each mouse is judged by eye. And concentration of each mouse is measured by gas analyzers.

The concentration of each toxic gas is adjusted roughly by control ing the

intake flow rate and intake duration of gas. The toxic gases are mixed with the

carrier gas of air of flow rate 10 (1/min) in a mixing chamber before lead to the

exposing chamber so as not to expose mice directly with the thick toxic gases.

In case of tests used mixture including CO 2 and N 2 ; CO 2 and N z are used as the

carrier gas. And for in case of test single C0 Z , 02 concentration was kept on about

21 (%) by supplying 0 2 into the test chamber so as to eliminate the effect of hypoxia.

Concentration of CO, CO 2 , HC1, N0 Z , NH 3 and O 2 are measured during each test by

cotinuous gas concentration analyzers; the maxmum concentration (or minimum concentration of 02 ) are recorded as the representative gas concentration of the test. The concentration of HCN is measured by the gas detecting tubes after one minute

starting of intaking the gas and for three minutes.

Each rotating cage of mouse generates a electrical signal for half a rotation

of the cage. Incapacitation time of the mouse is caliculated from the record of the

signals. The time after that a mouse dose not move for more than five minutes is defined as the incapacitation time of the mouse, because great diviation of incapacitation time of mice occasionally observed when the time after that a mouse dose not

386
move completly is defined as the incapacitation time.

Ambient temperature in the exposing chamber is controled within 30±2(*C) and

40±2(*C) except the case of tests for relationship between ambient temperature and

incapaci tationtime of mice.

3. Tes t resul ts

Test results of single gas, the relationship between gas concentration and incapacitation time of mice and the relationship between incapacitation time and apnoea

time are given an example in Figure - 1 to Figure - 8 . Experimental curves of the relationship shown in those figures are described as equation (1). Table - 1 shows the cofficients ‘a’ and ‘ b

’ in the equation for each

type of gas and the gas concentration on the experimental curve at the incapacitation time of mice of 5, 10, 20 and 30 minutes.

C (or T) = a / t + b ( 1 )

Where, C

T

t a, b

concentration of gas ( X or ppm ) temperature ( *C ) incapacitation time of mice ( minutes ) cofficient depending on the type of gas

By the equation ( 1 ), gas concentration relating to five minutes of incapacitation time of mice was calculated for each type of toxic gas. And measured gas con- centration was yield to a non-dimensional gas concentration index ( Ig ) by the

following equation:

Ig = ( Cm — b-) / ( C 5 - b )

• •

• ( 2 )

Where, Cm : measured gas concentration

C 5 : gas concentration corresponding to five minutes of

incapacitation of mice on the experimental curve

The relationship between incapci tation time of mice and gas concentration index

derived equation ( 2 ) are given an example in Figure - 9 to Figure - 10. The curves

in those figures are ideal curves; Ig = 5 / t .

Results of the tests with mixture gases, the relationship between incapacitation time of mice and algebraical sum of gas concentration index of the component

toxic gases and the relationship between incapacitation time of mice and apnoea time

of mice are given an example in Figure-11 to Figure-24. However, concentration of CO 2 and 0z were not yield to gas concentration index because the concentration was low.

4. Consideration

Results of tests with single gas condition, except COz, the test results are in

accord with the ideal relationship of Ig = 5 / t ; and those shows that the product

of gas concentration index and incapacitation time of mice is constant.

Table-1 show autline of consideration of the relationship between gas cocentration

and incapacitation time and the relationship between incapacitation time and apnoea

time.

Results of tests with mixture gas condition, Table-2 to Table-6 show autline of

consideration of the relationship between gas cocentration and incapacitation time

and the relationship between incapacitation time and apnoea time.

387
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393
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395

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ffl)& —
TOXIC GAS TEST BY THE SEVERAL PURE AND MIXED GASES

T. Sakurai, Japanese Association of Fire Science and Engineering, Japan

LEVIN: During your talk, you mentioned that your results on CO and C02 differ

from my results. I believe the reason for this difference is from differences

in the experimental procedures. I am not looking at lethal levels of CO,

adding C02 and looking for decreases in the time of death because at those

high concentrations of CO, the additional C0 2 may not make a big difference.

I am looking at lower levels of CO, adding the C02 , and finding that deaths

occur. Sometimes these deaths occur 24 hours later, so the time to death

within the exposure are not as important as the occurrence of death. The

insult that the C02 adds may take time for the body to recognize or for the

effect to take place. So, in your case, you're looking at the incapacitation

level and looking for changes in the time. I think this may be the

difference

.

SAKURAI: In our case too, we do not experiment with C0 2 toxicity as such. As

I explained at the beginning, when C02 concentration is increased, in 30

minutes, you have to have more than 20 percent; otherwise, incapacitation

wouldn't occur, but the maximum was only 8 percent. We are interested only in

effect of C0 2

, that is the main objective of our experiments.

And, another thing, we are looking at the animals' incapacitations within 30

minutes or one hour, a rather short time. Your case is like 24 hours, so it

is possible to get some different results because of the difference of time.

TSUDA: You have mentioned the additivity of C02

. In the case of the patienc

sent to the hospital because of the CO poisoning, they would have such a test.

In that case, not only 0

2

, we add the C0 2 also, mixed C02

, in order to

encourage the breathing of the patient, and then take out CO from the patient,

so that the patient will be away from the toxicity of CO.

What do you think of this kind of phenomena?

LEVIN: I'd like to say that the C0 2 does stimulate respiration. If you're in

the presence of CO, it will stimulate the uptake of the CO. If you are in the

presence of oxygen, it will increase the rate of oxygen uptake into the body.

Under these latter conditions, CO is reduced in the body (it's breathing it

out). However, C0 2 also increases the acidosis of blood, and this may add a complication

.

396
Strategy of Fire Suppression by Using Fire Retardants.

Toshisuke Hirano

Uni vers i ty o f Tokyo

7-3-1 Hongo, Bunkyo-ku, Tokyo 173.

and

Ichiroh Nakaya

Building Research Institute

1 Tatehara, Oho-cho, Tsukuba-gun,

Ibaraki 305, Japan.

ABSTRACT

The demand for the building materials treated with fire retardant chemicals

is increasing. To assure healthy and safe human life, these materials should

not pollute the air in the building even when they are involved in fires.

The chemicals used for fire retardation are desirable to be inexpensive and

easy to make treatment in addition to having no adverse effects on the

physical properties of the materials. Unfortunately, none of the currently

available chemicals meets all of these requirements. To develop the ideal

chemical, a cooperative study by many researchers or specialists of various

research fields, such as combustion science, chemical engineering,

mechanical engineering, chemistry, toxicology and economy, is necessary. The

development of the ideal chemical is the current world-wide concern. For the

great success in this work, we proposed the initiation of an international

cooperative study on this, topic.

l.The role of fire retardants 'in fire suppression.

In many countries, some fire protections are required for materials used

in a building. If ail of them were incombustible, there would be no building

fire but the number of furnishings and furnitures would be deadly limited

and our lives would become restricted, unhealthy and wearisome. As we can't

do without combustible materials, this is not a realistic method to overcome

fire hazards.

There are two ways to improve the quality of materials against fires.

One way is to combine a combustible material and an incombustible one. The

other way is to treat a material with fire retardant(s) . In the former way,

the material is usually concealed behind the incombustible material.

Furthermore, the total amount of combustibles, which may be important for

the evaluation of fire hazards under fully developed fire conditions, is not

decreased. So the latter way seems to be preferable in real situations.

The facility added to the materials by the treatment with fire

retardants, is increase in the ignition delay time, slowing down of the

flame spread speed or decrease in the pyrolysis or heat release rate.

Usually these facilities don't work individually but together in a complex

way. Such as, flame spread speed will be slowed down when ignition delay

time become long or heat release rate are decreased.

2. Current fire retardants

In Japan, the number of chemicals permitted to use for fire retardations

were about 250 in 1980(1). They are used as additive fire retardants,

reactive fire retardants or surface coatings.

The additive fire retardants are generally used for cellulosic materials

397
or thermoplastic resins with few exceptions (2) . The chemicals used as the

additive fire retardants are injected by putting pressure or penetrated into

cellulosic materials, or added to the polymer during processing plastic

res i ns

.

These

phosphorus

i norgani

c

eas 1 ness

chemicals belong to the halogenated compounds, the organocompounds, the halogenated organo-phosphorus compounds and/or the

compounds (Tab 1 e 1). The advantages of these fire retardants are

in processing and this lowers the manufacture cost. At the same

time, however,

materials.

the .chemicals may be.easy to migrate and to ooze from the

Table 1. Typical chemicals used as additive fire retardants.

Types Typical Chemicals

Halogenated Compounds Po 1 ych 1 oroparaff i

n

Chrolo-polyethylen

Perchl oro-pentacyci odecane

Tetra- broraoethane

Hexabrorao-cyc 1 odecane

Organopnospnorous Compounds Tri methyl -phosphate

Tri ethy 1 -phosphate

Tri butyl -phosphate

Tri octyl -phosphate

Tributoxyethyl -phosphate

Octy 1 - d i pheny 1 phosphate

Triphenyl -phosphate

Halogenated Organophosphorous

Compounds

Tns-chloroethy 1 -phosphate

Tris-2-chloropropy I -phosphate

Tris-2.3-dichl oropropy 1 -phosphate

Tris-2.3-broraopropy 1 -phosphate

Tris-bromochl oropropy 1 -phosphate

2.

3-

d i bromopropy 1 -phospnate

2. 3ch 1 oropropy 1 -phospnate

Inorganic Compounds Antimony Oxide

Boron Compounds

Aluminium Compounds

Molybdenum Compounds

The reactive fire retardants are reactive monomers or crosslinking

agents and are primarily used in unsaturated polyesters, epoxy resins and

polyurethane foams(2). The chemicals used as the reactive fire retardants

are chemically reacted with the polymer structure at some processing stage.

These chemicals belong to the acid compounds and/or the phospheric

polyol compounds (Tab 1 e 2)-. The main advantage of these chemicals is the

permanence of the fire retardance imparted. In most cases chemically

reacting the fire retardant into the polymer essentially eliminates longterm migration of the fire retardant. But these materials may be more expensive in the long run because of reacting the materials with the

polymer. In some case the processing techniques are very di f f i cul t (2 )

.

398
Table 2. Typical chemicals used as reactive fire retardants.

Types Typical Chemicals

Acids Compounds Tetrachlorophthal ic anhydride

Tetrabroraoohthal ic anhydride

Chlorendic acid

Chlorendic anhydride

D i al 1 y 1 ch 1 orendate

Polyols Po 1 yo 1 s

' vhc i ch contain

u ,

1 Halogen groups

Phosphorous groups

Halogen and phosphorous groups

The use of surface coatings has been mainly concentrated to increase the

fire retardance of wooden products or textiles by painting. The surface

coatings can be non- i n tumescen t or intumescent. The non- 1 n tumescen t coatings

have a long history and are more widely used than the intumescent in

addition that they are more economical. On the other hand, intumescent

coatings are undoubtedly more effective in providing a high degree of

protection on the underlying material(3). In any way, this type of fire

retardant is easy to process but sometimes lacks the permanence of the fire

retardance.

The chemicals used in the non- i n tumescen t coatings belong to halogenated

alkyd, fire retarded polyurethanes, epoxy resins with a reasonable degree of

fire retardance, merami ne- formaldehyde resins or phenol-formaldehyde resins.

On the other hand, those used in the intumescent coatings belong to carbonific compounds, catalysts which promote decomposition of the

carbomfic compounds, spumific compounds or resin binders(3).

3. Suitable fire retardants

Most of the fire retardants belong to the halogenated compounds or the

halogenated organocompounds . It is well known that the amount of smoke

evolved during the burning of halogenated polymers are significantly greater

than those derived from the corresponding non-hal ogena ted polymers.

Furthermore, halogenic acids, which are known to be highly toxic, are emitted as the final product.

These flame retardants which act chemically in the gas phase are of course generally react readily with active flame propagating species such as

a hydroxyl radical. Then these active species are no longer available for

the oxidative reaction of the gaseous fuel products. As cracking of the

carbon-rich compounds tends to occur through the oxidation process, the lack

of the active species cause eventually the formation of large amount of the

carbon-rich solids which constitute smoke or carbon monoxide, which is also

highly toxic, by incomplete combustion. The evolution of the smoke or toxic

gas species is undesirable because it impedes egress and fire brigade

activities which enlarge damages to be caused by fires. Thus the additives

introduced to reduce the flammability of materials may increase the fire

hazard at the same time.

The other flame retardant chemicals belong to inorganic or organo399
phosphorus compounds. These chemicals take effects mainly in solid phase and are usually encourage the formation of a char layer and/or water. The char

layer is well known to act as the good thermal barrier. And water will

increases the total latent heat of the material’s pylorysis. As the emission

of the carbon-rich volatiles are suppressed, they are seldom to promote the

formation of toxic gas species or smoke during the combustion of materials.

But the cost to get satisfactory fire retardance using this kind of

chemicals is more expensive than using the halogenated compounds.

The ideal chemicals used as additive fire retardant should be

inexpensive,-- colorless," easily incorporated into the polymer composition,

stable to heat and light, efficient in its fire retardant properties,

nonmigrating, and have no adverse effects on the physical properties of the

polymer. The toxicology of the chemicals and the gases produced by

combustion are also of great concern(2). The same items should be required

to the chemicals used as the reactive fire retardant or the surface coating.

Unfortunately, most currently available chemicals seldom meet all these

requirements

.

4. The expected study subjects

Fire retardant chemicals take effect in gas phase and/or solid phase. In gas phase they quench active species or slow down the oxidative reaction

speed. In solid phase, they suppress the production of combustible

volatiles.

It is known that the hydroxyl radical plays great role in oxidative

reactions of hydrogen and. carbon. To quench the hydroxyl radicals, halogen

atoms are usually used. These halogen atoms finally become halogenic acids.

These acids are highly toxic. Furthermore, a great amount of smoke is

produced as described previously.

On the contrary, the amount of smoke generation should become small if

the concentration of the hydroxy radical is increased. This might be

verified by the empirical knowledge that the generation of smoke is

dramatically decreased when water spray is introduced into the flame region

and that metal oxide with some crystal water act as smoke suppresser.

To control combustion reactions simply by dilution, the amount of inert

gas required may be very large. If the adequate amount of inert gas can be

introduced, this lowers the gas temperature and slow down the reaction speed

of oxidation of combustible compounds because it help removing of active

species by promoting their recombination reactions, which may increase the

amount of smoke.

If carbons can be fixed into the solid phase and no carbon rich

volatiles are not produced, the amounts of smoke generation and heat release

should be reduced. Phosphorus compounds are known to promote the dehydration

reaction in solid phase and to form a char layer at a fuel surface. Though

there seems to be no report about the smoke suppression by phosphorus

compounds, the amount of heat release is known to decrease by the treatment

with the phosphorus compounds. This might be because these compounds also

work in gas phase as the trap of active species(4).

Covering fuel surface by thermal barrier like a char layer or some coatings are effective to keep the fuel pylorysis at low rate. The char

layer is known to have poor thermal conductivity and high pylorysis

temperature. Phosphorus compounds are usually used for this purpose. This

reaction is said to be dehydration but the details of its mechanism has not

been fully revealed yet because it proceeds in the solid phase.

Intumescent or non- i n tumescen t coatings, are also used as the thermal

400
barrier. Sometimes, the spumific agents are added into the coatings or fire

retardants to form thick thermal insulating layer over the fuel surface. The

spumific mechanism has fairly well revealed. As the coated chemicals exist

on the fuel surface, they are easy to migrated and to lose the fire

retardance.

As described previously, the ideal fire retardant chemicals should be

inexpensive, colorless, easy to process, stable to heat and light, harmless

even when the materials treated by them were burned, but the chemical, that

meets all of these requirements, has not been produced yet. To resolve all

of the problems remained, the new ideal fire retardant chemicals or retarded

materials should be developed using the knowledge obtained through studies

in various research fields (Figure 1).

Figure 1. A Diagram to Develope New fire Retarded Materisis.

401
5. Currently required research activities

In last July, a new group has founded in Japan to study the controlling

mechanisms of combustion reactions. This group is composed of the

specialists of various research fields and they belong to various

organizations. For example, there are chemists, mechanical engineers,

chemical engineers and combustion scientists who belong to universities,

governmental institutes or private companies. In this group, discussion to understand controlling mechanisms of combustion reactions has mainly been

done. Through the discussion, all of the members understood what is known

or unknown.

The controlling mechanisms of combustion reactions by currently used

chemicals have almost revealed but further studies to find the other

possible mechanisms are still needed for the development of the ideal

combustion control technique. To promote this study, the fundamental study

by combustion scientists with the supports of chemical and mechanical

engineers may be important. Furthermore, the discussions in the aspects of

chemistry, toxicology and economy are indispensable because newly developed

materials treated by the new chemical should be easy to produce, safe and

cheap

.

So the development of new types of fire retardants and extinguishers is desirable to be commenced by combustion scientists, chemical engineers,

mechanical engineers, chemists, toxicologists and economists (Figure 2).

Figure_2. Tentative Plan for The Project Study.

As the development of new type of fire retardant chemicals or retardation techniques is world-wide concern, it will be wonderful if this

project study can be promoted internationally. We proposed the initiation of

a cooperative study to Dr . Gann (NBS) , Dr . Harmathy (NRCC) , Dr . Woo 1 1 ey(FRS) and

402
have a panel discussion on the theme of "Now and Future Aspects on Combustion Control in Fire" in the Second International Symposium on Fire

Safety Science to be held in Tokyo on June 13-17, 1988.

Now we would like to propose the initiation of an international

cooperative study to all of attendants at this conference again.

References

.

1.

Japan Fire Retardant Association: Sen'i To Bouen, San Print (1978).

2.

A.H.Landrock: Handbook of Plastics Flammability and Combustion Toxicology,

Noyes Publication (1983).

3.

C.F.Cullis and M.M.Hirschler: The Combustion of Organic Polymers, Oxford

University Press (1981).

4.

H.B. Palmer and D.J.Seery: Combust. Flame 4, 213 (1960).

403
STRATEGY OF FIRE SUPPRESSION BY USING FIRE RETARDANTS

I. Nakaya and T. Hirano , University of Tokyo, Japan

TEWARSON: In the United States, there is a great interest to produce non- originated low smoke additives. In my individual research, I find that for

most incorporation that is introduced, mostly into the final product, it does

not only reduce smoke corrosive compounds, but also has reduced the level of

emission. I would like to hear what is being done in Japan.

HIRANO: We are in the process of putting all this information, state and city

posts together, so upon the completing of this report, we will provide you

with a copy in the new form. When you receive the report, I would appreciate

your comments.

As far as the flame suppression deterrent is concerned, they also have any

number of activities in Japan. As far as practical applications are concerned, traditionally the Japanese have imported the kind of fire

suppression deterrent technique from the United States, not vice versa.

I will try to expand upon the initial stage. Dr. Nakaya came up with the idea

that he continue with the initial stage. As far as the basic research

activities are concerned, there are a number of basic research activities in

different countries all over the world, and on different topics. In this

instance, we appreciate the fact that there is information and many documents

available on various subjects.

Of course, each country is interested in developing excellent fire retardants.

We too are interested in developing excellent fire retardants. In order to

support this activity, we need basic research.

One more unsettled problem remains. This is due to the fact that our policy

is not clear. As far as my personal policy is concerned, it's clear in my

mind. However, the collective policy, various people with different

disciplines, must come up with an integrated policy. We have not yet come up

with an integrated policy and that's the reason why this is still fuzzy.

In addition to that, we also would appreciate Dr. Gann's participation on the

thermal decomposition issue. As far as the cable fire is concerned, Dr.

Tewarson can contribute to our organization. At the present, we have started

to study fiber optics.

404
STUDIES ON THE TOXICITY OF SMOKE CONTAINING HYDROGEN CHLORIDE

G. E. Hartzell , A. F. Grand and W. G. Switzer

Department of Fire Technology

Southwest Research Institute

6220 Culebra Road

San Antonio, Texas 78238

ABSTRACT

Studies of rodent lethality due to exposure to HC1 , as well as to mixtures of HC1 and CO, have shown different apparent toxicological effects at low and at high concentrations of HC1 . At low concentrations of HC1 , sensory

irritation causes a decrease in respiratory minute volume, with somewhat

slower loading of CO and a delay in incapacitation. This effect is observable

only at low concentrations of CO. At much higher HC1 concentrations, pulmonary irritant effects are observed leading to postexposure lethality. An

empirical analysis of the data for mixtures of HC1 and CO suggests that

exposure doses leading to lethality may be additive.

The lethal toxic potency ( L C

Q

) of PVC smoke may be largely, but not

entirely, accounted for by the HCT produced. However, PVC smoke exhibits a greater incidence of early postexposure deaths. The early deaths, which may

be partially attributable to a combined effect of CO and H Cl , may also be

linked to the pattern of respiratory penetration by the smoke. There is

evidence that components other than HC 1 are present which cause PVC smoke to

be more irritating than H Cl alone.

INTRODUCTION

Hydrogen chloride (HC1 ) is a relatively common component of fire

effluents about which there has been much speculation and controversy. The

studies reported here represent a continuation of our efforts to clarify the

role of HC1 in combustion toxicology.

In order to place new work in perspective, findings from selected

previous studies will also be discussed. An attempt will also be made to

develop the practical application of this knowledge to exposure of human

subjects

.

EXPERIMENTAL METHODS AND MATERIALS

EXPOSURE OF ANIMALS

Apparatus

The exposure apparatus used in these studies, illustrated in Figure 1, is

similar in size and shape to that commonly used in the NBS smoke toxicity test

[1]. However, it was modified to create a "flow-through" system. The air

flow through the exposure chamber ranged from about 30 L/min up to about

405
200 L/min. This wide range of air flows was due to modification of the ex- haust system during the course of the program in order to provide greater

flexibility and control in generating and maintaining specific gas atmospheres. The higher flow rates permitted more rapid equilibration of the test

atmospheres and more stability in the atmosphere once established.

SODA LIME CO

TUBE SAMPLING

AIR

Figure 1. Flow-through exposure apparatus

The animal restraints and isolation system, used to protect the animals

from the test atmosphere until equilibrium was established, was the same as that used in a previous study [2]. Incapacitation was determined using con- ventional leg-flexion shock avoidance apparatus [3].

Exposure Protocol

Each animal exposure test involved the exposure of six, young adult, male

Sprague-Oawl ey rats positioned in tubular restrainers to provide for head-only

exposure. Following insertion of the tubed animals into position through the

wall of the chamber, the animal isolation system was closed and breathing air

provided during the establishment of the desired test atmosphere. Thirtyminute exposures were initiated upon opening of the isolation system and termination of supplemental breathing air to the animals. End points recorded were incapacitation and lethality. Surviving animals were observed until

either postexposure death or 14 days. Animals surviving 14 days were subsequently sacrifi ced

.

Values for LC

5 Q

' s and 95-percent confidence limits were calculated from

the data using a standard probit program.

For experiments during which it was desired to follow the time course of

blood pH or COHb saturation, blood samples were withdrawn sequentially from

several rats implanted with femoral artery cannulas. No cannulated rats were

used in the calculation of IC

5Q

values.

406
GENERATION OF TEST ATMOSPHERES

Pure Gas Atmospheres

"Pure gas" atmospheres were generated by the metering of the appropriate

gases from cylinder sources into the exposure chamber, with continuous

analytical monitoring of concentrations. Suppl emental 0

2 was provided as necessary to maintain the 0, concentration at 20.5 ± 1.0%.

Sources were:

Carbon Monoxide (99.9%): Scientific Gas Products

Hydrogen Chloride (99.5%): Scientific Gas Products

Polyvinyl Chloride (PVC) Fire Effluent Atmospheres

A continuous combustion apparatus, developed at SwRI , was used for the

studies reported. Shown in Figure 2, it consists of a quartz tube 1 m long

and 7.6 cm in diameter, with two radiant heating devices positioned exterior

to the tube. One of several stainless steel combustion boats (ranging from

50 to 110 cm in length), containing a weighed quantity of PVC, was pulled at a constant rate of travel inside the tube past the radiant heater. The heat

flux, sample weight and rate of travel of the ccmbustion boat were predetermi ned to achieve the desired concentration of HC1 in the fire effluent.

QUARTZ LAMP

DILUTION HEATER

Figure 2. Continuous combustion apparatus

A constant flow of dry air (20 L/min) was metered into the combustion

tube, with additional room air (from 30 L/min to approximately 200 L/min)

being drawn into the system by an exhaust blower. The total flow of air

through the apparatus was measured either by an orifice meter in the exhaust

line or by calculation based on dilution of a carefully metered pure gas prior

to the combustion experiments.

407
Different radiant heat furnaces were used for flaming and for nonflaming

conditions. Nonflaming combustion was conducted at 3 W/cm 2

, using a cone heater (the same as used in the NBS cone cal or imeter ) . However, the cone heater was unable to produce the higher fluxes (6 W/cm 2

) necessary for flaming

PVC experiments. Therefore, a tungsten-quartz lamp heating device was used

for the flaming mode. A small coil of resistance wire, connected to a variac

and mounted inside the combustion tube, was the ignition source used to sus- tain flaming combustion.

Shown in Figure 3 is the relationship found between the HC1 concentration

produced and the quantity of PVC burned, expressed as mg*L _1

. The latter

values were estimated from the mass of PVC charged, sample travel and air flow

rates .

=&

uu>

<

8000

7000

6000

5000

4000

3000

2000

1000

0

0 5 10 15 20 25

FOAMING

a NONFLAMING

mg PVC A AJRFLOV

Figure 3. Plot of HC1 concentration produced as a function

of mass of PVC charged per liter of airflow

The polyvinyl chloride (PVC) used was natural, unplasticized material in

pellet form supplied by the B. F. Goodrich Chemical Company. It contained two

parts lubricant and two parts stabilizer per hundred parts resin.

ANALYSIS OF TEST ATMOSPHERES

Asphyxiant Toxicants

Analyses of exposure atmospheres for CO, CO, and 0, were continuously

conducted using a closed loop sampling system and tne following instruments:

Aq8
Beckman 865

Beckman 865

Beckman OM-11

Carbon Monoxide:

Carbon Dioxide:

Oxygen

:

Hydrogen Chloride

The method used for continuous monitoring of hydrogen chloride (HC1) gas,

both in a "pure gas" atmosphere and in smoke, was based on that reported in a French standard test method [4], with instrumentation and procedures modified

in our 1 aboratori es . Complete verification of this method, including measurement of i nterf erences , has not been done. Therefore, for most experiments,

"back-up" analyses of HC1 were also performed by soda-lime tube (SLT) sampling

followed by aqueous extraction and titration [5].

A schematic drawing of the continuous HC1 method is illustrated in

Figure 4. A sample of an atmosphere containing HC1 is drawn continuously by a pump into a gas-solution impinger containing 0.1 N HNO, aqueous solution. A

si 1 ver/ si 1 ver chloride electrode (or electrode pair) and a pipette for introduction of silver nitrate ( AgN 0 3

) titrant are positioned in the cell. The

basic principle is to maintain a predetermi ned electrical potential (emf) in

the test solution by autcmati cal ly metering in AgN0

3 to react with chloride

ion. The rate of addition of titrant is directly proportional to the concen- tration of HC1 in the sample gas stream.

CRT DISPLAY

CELL

Figure 4. Continuous analysis of HC1

In operation, any increase in HC1 is detected as a change in the emf of

the solution. This value (mv) is monitored by a digital pH/mv meter connected

409
to a "stat titrator" control module, which controls the dispensing of silver

nitrate titrant solution through an automatic buret. When the emf of the

solution deviates from the set point due to addition of chloride, the buret

dispenses AgN0

3 in order to return the emf to the set point. This process

continues as needed to maintain the solution at the set point, which is near

the equivalence point of a sil ver-chl ori de titration.

The higher the concentration of HC1 in the test atmosphere (or the higher

the gas flow rate), the faster the silver nitrate is run into the titration

cell. The output of the automatic burette (ml of titrant) is linked directly

to an Apple lie mi crocomputer for computation of H C 1 concentration. A mass flowmeter is used to monitor (and control) the gas flow, its output also going

to the computer. Calculations are performed every 10 seconds in order to

time-average the quantity of AgNO, being used. Due -to fluctuations in the

quantity of HC1 introduced into the impinger during any given 10-second

increment, an additional time average is performed over three successive readings. This final averaging "smoothes out" the curve of HC1 vs time. The increment of time for each calculation and the interval for determining the

"average" HC1 are both under the control of the operator. The following data

are displayed on the comDuter monitor every 10 seconds (or any other increment

sel ected )

:

1) emf of the solution (mV);

2) total volume of AgNO, consumed (ml);

3) sample gas flow rate (cc/min);

4) concentration of HC1 (ppm);

5) "average" HC1 concentration (ppm) over the selected

time interval; and

6) current run time (minutes and seconds).

Immediately following each experimental run, a plot of HC1 concentration

vs time may be constructed from the stored data.

RESULTS

A total of eight LC0 (30-minute exposure plus 14-day observation) values were determined. Data oocained are summarized in Tables 1 and 2.

TABLE 1. CARBON MONOXIDE IC

5Q

VALUES*

(30-MINUTE EXPOSURE + 14 DAYS)

Atmosphere LC

5Q (ppm)

CO 6400

CO (600 ppm HC1) 5700

CO (1000 ppm HC1

)

5300

CO (1000 ppm HC1/PVC smoke) 7100

*Va1ues are rounded to the nearest 100 ppm

410
TABLE 2. HYDROGEN CHLORIDE LC

5Q

VALUES*

(30-MINUTE EXPOSURE + 14 DAYS)

Atmosphere LC

50 (ppm)

HC1 3800 HC1/PVC (Nonflaming) 2900

HC1/PVC (Flaming) 2100

HC1 (3000 ppm CO) 2100

*Values are rounded to the nearest 100 ppm

DISCUSSION

Although hydrogen chloride can exert toxicological effects in a variety

of complex ways, two are relevant to this study: 1) as a sensory irritant and

2) as a pulmonary irr.itant. At low concentrations, it is primarily a sensory

irritant resulting in depression of a rodent's respiratory rate and respiratory minute volume (RMV) . The effect, which is concentrati on- rel ated rather

than dose-related, is often quantified as that concentration causing a 50-percent decrease in respiratory rate ( RD

5 0

) - The RD

5Q

for the SpragueDawley rat is about 600 ppm [2].

Sensory irritation is an effect of the inhalation of HC1 which, although

generally not leading to lethality, does have a perturbing effect on the inhalation and resulting toxicity of CO. Earlier work (time- to-eff ect studies)

reported that in the presence of HC1 , r-espiratory rate and minute volume

depression increase somewhat the observed times to the toxicological effects

of CO, particularly if the CO concentration is relatively low, i.e., less than

about 4000 ppm [2]. Above approximately 4000 ppm CO, the presence of the irritant has increasingly less influence on times to physiological effects due

to the very rapid loading of CO. Concentrations of HC1 above about 1000 ppm

have little additional effect over that observed with 1000 ppm since

increasingly less additional depression of RMV occurs above that concentration. The overall effect of the presence of HC1 on intoxication of rats by CO

is relatively minor and is mostly observed in a "window" of perhaps 400 to

1000 ppm HC1 and with CO concentrations up to about 4000 ppm.

It must be emphasized that the effects of a sensory irritant on CO

toxicity may be limited to rodents. The significance and applicability to

human exposure remain to be determined from primate studies. There is evidence that, with primates, exposure to a sensory irritant results in decreased

arterial blood p0 7 values which may increase its sensitivity to CO [6].

A comparison of the sensory irritant effect of smoke produced frcm the

burning of PVC with that of simple gaseous HC1 is of particular interest. An

equivalent concentration of HC1 contained in PVC smoke (nonflaming decomposition) increased observed times-to-i ncapacitati on over most of the concentrations of CO studied, not just at the lower concentrations, thus giving

indication of sensory irritation somewhat greater than that produced by HC1 , alone. Furthermore, the 30-minute LC.n of CO in the presence of 1000 ppm HC1

contained- in PVC smoke (nonflaming) .17100 ppm with 95-percent confidence of

4300 to 6500 ppm) was greater than that for CO in the presence of simple HC1

411
(5300 ppm with 95-percent confidence of 6300 to 8000 ppm). This "lower toxicity" of CO in the presence of smoke produced from the nonflaming decomposition of PVC is another indication that irritants more potent than HC1 may be

present which act to depress further the RMV and slow the loading of CO. (The

greater irritancy of PVC smoke than would be expected from the HC1 content has

previously been reported based on experimental measurements [7]). It must be

emphasized that these effects are reported for exposure of rodent; primates,

including humans, have not been studied.

Hydrogen chloride, when present at high concentrations, is thought to

penetrate into the lower respiratory tract and exerts its lethal toxic effect

as a pulmonary irritant [3]. Pulmonary edema and subsequent death often

follow exposure to HC1 at high concentrations. Postexposure lethality results

from exposure doses much lower than those required for either incapacitation

or wi t hi n-exposure lethality [8].

Concentrati on-response relationships for postexposure lethality of rats

exposed to HC1 have been studied over a range of exposure times from 5 minutes

to 60 minutes [9]. Exposure doses (LCt

5Q 's) ranged from about 80,000 ppm-mi n to about 170,000 ppm-min, respectively. Higher exposure doses can be tolerated if the HC1 concentration is relatively low. Conversely, progress! vely

higher concentrations of HC1 result in tolerance to decreasing exposure

doses. The 30-minute LC

5Q (postexposure lethality) for HC1 was determined to

be about 3700 ppm within a static system [9] and about 3800 ppm for a flowthrough system. Considering the 95-percent confidence limits for postexposure

lethality due to pulmonary irritants, the difference between these two values

is not significant.

Very limited experience with baboons has indicated that 150,000 ppm-min

(5- to 15-minute exposures) is probably on the threshold of postexposure

lethality [9]. Although there has been no experience with human exposures

under controlled conditions, somewhat conservative estimates based on the

nonhuman primate experience would indicate that perhaps 100,000 ppm-min would

likely result in very severe ccmpl i cati ons and even death.

The 30-minute (plus 14-day observation) LC

5Q

of HC1 as contained in smoke

from the nonfl ami ng decomposition of PVC (little or no CO present) was found

to be 2900 ppm (95-percent confidence of 2200 to 3700 ppm). Although somewhat

lower than the value of 3800 ppm (95-percent confidence of 3100 to 4800 ppm)

for HC1 , itself, it would be most reasonable to conclude that the postexposure

lethal toxicity (l_C 5Q ) of nonflaming PVC smoke can largely be accounted for by

the HC1 content. Further support for this statement can be obtained from the

data plotted in Figure 3, in which an HC1 concentration of 2900 ppm corresponds to a PVC concentration (mass charged) of about 10 rng'L

-1

. Although this

value is somewhat lower than that reported from the NBS cup furnace test [1],

a greater observed toxicity of PVC smoke would be expected from the combustion

method used in this study due to the absence of any HC1 generation delay time

and also the absence of HC1 decay.

The mechanism of action of HC1 is so very different physiologically from

that of CO, that it would be anticipated that these two toxicants should be

considered separately in predicting hazardous exposure conditions through the use of the Fractional Effective Dose model [10]. Fractional doses would not

be expected to be additive, as in the case of CO and HCN. However, the LC 50

412
data on mixtures of CO _ and HC1 tempt one to consider additivity on an empirical basis. In Table 3 are shown data for seme limited mixtures of CO and HC1

in which the summation of fractional lethal doses approximates unity for

50-percent lethality with 30-minute exposures. In Table 4 is shown a com- parable treatment of data using PVC as the source of HC1 . It would appear

that the increased lethal toxicity of HC1 as contained in smoke frcm the

flaming combustion of PVC may be attributed only in part to the contribution

of CO produced. The full explanation may be quite complex, however, since the

gas/aerosol /parti cul ate nature of real smoke may result in a different pattern

of respiratory penetration. For example, particulate deposition in the upper

respiratory tract may obstruct breathing of rodents and further stress an already compromised oxygen transport system.

TABLE 3 SUMMATION OF FRACTIONAL EFFECTIVE (LETHAL) OOSES FOR 30-MINUTE EXPOSURE

OF RATS TO MIXTURES OF CO ANO HC1

CO HC1 . Fractional Lethal Fractional Lethal £ FED

ppm Pose ppm Pose (Lethal ity )

6400 1.0 — — 1.0 5700 0.39 600 0.16 1.05 5300 0.33 1000 0.26 1.09 3000 0.49 2100 0.55 1.04 — — 3800 1.0 1.0

TABLE 4 SUMMATION OF FRACTIONAL EFFECTIVE (LETHAL) OOSES FOR 30-MINUTE

OF RATS TO MIXTURES OF CO AND HC1 AS PRODUCED FROM PVC

CO HC1

EXPOSURE

l FED (Lethal ity

)

ppm

Fractional Lethal Oose ppm

Fractional Lethal

Oose

ca. 700 ca. 0.11 2100 ( fl ami ng) 0.55 ca. 0.66

7100 1.11 1000 ( nonf 1 ami ng ) 0.26 1.37 — — 2900 ( nonfl ami ng) 0. 76 0.76

As discussed earlier, smoke from the nonflaming decomposition of PVC

would appear to be more irritating that expected frcm the HC1 content. Thus,

in the case of CO being the major toxicant, the summation of FED's was found

to be considerably greater than unity when HC1 from nonflaming PVC smoke was present at the sensory irritation level. Although smoke from the nonflaming

combustion of PVC would appear, based on summation of H C 1 FED's, to be more

toxic (as a pulmonary irritant) than expected, the significance of the difference is questionable.

Particularly striking, both in the case of the flaming combustion of PVC

(with which up to 1680 ppm CO was produced) and in the case of HC1 in the

presence of 3000 ppm CO supplied from a cylinder source, was the incidence of

postexposure deaths the same day as the exposure, usually within an hour.

413
This is rare, except at high concentrations, for HC1 , alone. Data showing the

incidence of these early postexposure deaths are shown in Table 5.

TABLE 5 COMPOSITE LETHALITY OF RATS EXPOSED TO ATMOSPHERES

CONTAINING HYDROGEN CHLORIDE*

Atmosphere

Animals

Exp os ed Exposure

Number of Oeaths

Day 0 Days 1-4 Days 5-14

LC

50 HC1 in ppm

H Cl 60 0 0 21 9 3800

PVC Smoke-Nonflaming 36 0 0 10 3 2900

PVC Smok e-Fl ami ng 30 1 4 5 19 2100

HC1 + 3000 ppm CO 30 I 2 8 17 2100

* At nominal approximate HC1 concentrations of 1800, 2600, 3500, 4300, 5500 and 6500 ppm.

Attempts to elucidate the effect of CO on the postexposure lethality of

HC1 were inconclusive, but did suggest s anew hat impaired ability of rats to

return to homeostasis following exposure. Exposure of rats to a mixture of

5300 ppm HC1 and 3000 ppm CO was character!' zed by a relatively rapid initial

drop in blood pH and slower than normal postexposure unloading of COHb. Even

with these data (compared with those for CO alone), it is difficult to speculate as to the cause of the incidence of same day postexposure lethality. The

observations reported with rodents may pose seme significance in the case of

immediate postexposure comDl i cati ons following human exposures to mixtures of

CO and HC1 (e.g., prolonged hypoxemic conditions after rescue or escape).

In summary, it would seem that the lethal toxicity of combinations of CO

and HC1 can be predicted from an FED model. However, only limited success was obtained when applied to real smoke from PVC. This may be due to complications in respiratory penetration patterns involved with the complex gas/

aerosol /parti cul ate mixtures found in smoke.

ACKNOWLEDGEMENT

This work was supported under U. S. National Bureau of Standards Grant

No. 60NANB6D0635. Development of the analytical and combustion methodology

described was supported by the Vinyl Institute of The Society of the Plastics

Industry, I ncorporated

.

REFERENCES

1. Levin, B. C., A. J. Fowell, M. M. Birky, M. Paabo, A. Stolte and D. Malek,

"Further Development of a Test Method for the Assessment of the Acute

Inhalation Toxicity of Combustion Products," NSSIR 82-2532, National

Bureau of Standards, Gai thersburg , Maryland (1982).

2. Hartzell , G. E., H. W. Stacy, W. G. Switzer, D. N. Priest and S. C. Packham, "Modeling of Toxicological Effects of Fire Gases: IV. Intoxication

of Rats by Carbon Monoxide in the Presence of an Irritant," J. Fire Sciences , Vol . 3, pp. 263-279 '(July/August 1985).

414
3. Klimisch, H. J., J. E. Doe, G. E. Hartzell, S. C. Packham, J. Paulunn and

D. A. Purser, "Bioassay Procedures for Fire Effluents: Basic Principles,

Criteria and Methodology," J . Fire Sci ences (In Press).

4. French Standard (AFNOR) FD X 70101 December 1982, "Analysis of Combustion

and Pyrolysis Gases - Smoke Chamber Method."

5. Grand, A. F., H. L. Kaplan, J. J. Beitel , W. G. Switzer and G. E. Hartzell, "An Evaluation of Toxic Hazards from Full-Scale Furnished Rocm

Fire Studies," Fire Safety: Science and Engineering, ASTM STP 882 , T. Z. Harmathy, Ed., American Society for Testing and Materials, Philadelphia,

Pennsylvania, pp . 330-353 (1985).

6. Kaplan, H. 1., A. Anzueto, W. G. Switzer and R. K. Hinderer, "Effects of

Hydrogen Chloride on Respiratory Response and Pulmonary Function of the

Baboon," J. Toxicology and Environmental Health, submitted for publication

(March 19877:

7. Barrow, C. S., Y. Alarie and M. F. Stock, "Sensory Irritation and Incapacitation Evoked by Thermal Decomposition Products of Polymers and

Comparisons with Known Sensory Irritants," Arch. Env. Health, Vol . 33,

p. 79 (1978)

.

8. Kaplan, H. L., A. F. Grand, W. G. Switzer, D. S. Mitchell, W. R. Rogers

and G. E. Hartzell, "Effects of Combustion Gases on Escape Performance of

the Baboon and the Rat," J. Fire Sciences, Vol. 3, pp . 228-244 (July/August 1985)

.

9. Hartzell, G. E., S. C. Packham, A. F. Grand and W. G. Switzer, "Modeling

of Toxicological Effects of Fire Gases: III. Quantification of PostExposure Lethality of Rats from Exposure >to HC1 Atmospheres," J. Fire

Sci ences , Vol. 3, pp. 195-207 (May /June 1985).

10. Hartzell, G. E., D. N. Priest and W. G. Switzer, "Modeling of Toxicological Effects of Fire Gases: II. Mathematical Modeling of Intoxication of

Rats by Carbon Monoxide and Hydrogen Cyanide," J. Fire Sciences, Vol. 3,

pp. 115-128 (March/April 1985).

“ '

415
STUDIES ON THE TOXICITY OF SMOKE CONTAINING HYDROGEN CHLORIDE

G. Hartzeil, A. Grand, and W. Switzer, Southwest Research Institute, USA

TEWARSON: Would you discuss the impact of the low HC 2. concentrations?

HARTZELL: Low concentrations of HC.2 , such as a few hundred parts per million,

are extremely irritating, and may very well cause a person to not want to

escape. Physiologically, he could escape a much higher concentration. And

the concentrations required to kill are very high. In terms of toxicity,

hydrogen chloride is really not all that toxic. On a scale of irritants,

hydrogen chloride is not very irritating. There are worse irritants. We know

very little about behavioral effects, not physical but behavioral effects, of

people exposed to these very strong irritants

.

TSUCHIYA: Can you describe the this difference with PVC flaming smoke and PVC non- flaming smoke?

HARTZELL: The non- flaming is not particularly visible, very light smoke,

whereas the flaming is quite noticeable. To some extent, we're able to get

the same results of HCJ? and CO, as we obtained from flaming PVC. However,

flaming PVC also had soot particles and water, so there are other

complications

.

GANN: You mentioned a factor of 30 between the mortality data and some escape

data. Is this hundred versus 30?

HARTZELL: At the last Tripartite Committee meeting, results from Dr. Alarie's

lab on running speed and potential for escape, remembering that the LC 50 's were widespread, something like several thousand parts per million, they

measured the onset of slowing of burning of a couple hundred parts per

million, and significantly determined well below 1,000. Dr. Alarie announced

that the difference between the two would be a factor of 7.

416
TOXIC GASES FROM COMBUSTION OF WOOD

TREATED WITH RETARDANTS

Toshimi HIRATA and Sumire KAWAMOTO

Forestry & Forest Products Res, Inst.

P. 0, Box 16, Tsukuba Norin Kenkyu Dancni-Nai,

Ibaraki, 305 Japan

ABSTRACT

Combustion toxicity of wood has been attributed to CO. That for wood

treated with retardants, however, could not satisfactorily be explained

exclusively by the CO evolution. Therefore, toxicants other than CO were researched. First, evolutions of N0.c and SO-? were determined from retardant

treated wood during JIS A 1321 tests by means of IR spectroanalyzers

.

Secondly HCN evolution from the wood heated in an imaging furnace was measured by ion chromatography wxth an electrochemical detector. The ECN

results are discussed in terms of the dependences on air flow rate and

temperature. The data are useful for the reasonable explanation of the

combustion toxicity of the retardant treated wood,

1.. INTRODUCTION

Combustion toxicity of wood treated with several retardants has been

reported (1), which is determined in incapacitation and death of mice

according to the Ministry* of Construction Noticef ication 1231. The reported

data have been rearranged and reviewed with respect to the incapacitation

time against the CO concentration in the mouse-exposed chamber and the

survived mouse percentage against the retardant contents. The results

obviously show that the combustion toxicity cannot be explained exclusively

by the CO evolution, though it has been accepted to be attributed to the

CO production.

In the rearrangement a relation between the CO concentration and

incapacitation time was approximated by a regression function

Y = 6.43 X

-0 ’ 46

, CD

where Y is the incapacitation time and X the maximum CO concentration, as shown in Fig. 1. It is noticeable that two samples treated with guanidine

sulfamate and ammonium sulfamate yielded a lot of plots downward deviated

from the regression line, and that on the contrary, boron compound treated

samples gave many plots deviated upward. On the other hand, the plots about

the death of mice indicate interesting tendencies that the survived mice

increased with the content of boron compounds, but that with ammonium

phosphate and ammonium sulfamate they decreased after the increases to the

maxima, as shown in Fig. 2. Furthermore, for the guanidine sulfamate

treated samples, the survived mouse percentage at four chemical content

levels are 0 except 4.2 % at a 15 % chemical content level. Since the CO

evolution decreased with the increase in the chemical content, these

observations suggest that combustion toxicity for retardant treated wood

depends not only on CO but also other toxic gases. Therefore, the present

study has been carried out to search such toxicants.

2 . EXPERIMENTAL

2-1. Samples

Wood used for the JIS A 1321 Surface Flammability Test is 22 x 22 x 1

417
cm red lauan (Shores, rubroshorea section) of the air-dried specific gravity

0.43 to 0.51. The surface of specimens which was finished by a planer is

approximately parallel to the fiber direction. By means of a full-cell

process they were impregnated with diammonium hydrogen phosphate ammonium

bromide, guanidine sulfamate, ammonium sulfamate, boric acid, sodium

tetraborate decahydrate, sodium chloride, or sodium borate plus boric acid.

All of the chemicals were CP grade. Before and after the impregnation the

specimens were conditioned at constant temperature 22° C and relative

humidity 65 %. Therefore, the chemical concents in the specimens could be

determined from the weight differences. Before the tests they were dried at

40° C for 72 h and then cured with silica gel for 24 h in order to preclude

effects of water in the wood on the combustion.

Another sample for isothermal heating is pulverized Japanese cedar

(Chriptomeria japonica) wood which passed through a 16 mesh but not through

a 40 mesh screen. The sample was immersed in diammonium hydrogen phosphate,

ammonium dihydrogen phosphate, guanidine sulfamate, or ammonium sulfamate

solution in the different concentrations. The sample recovered from the

solution was freeze-dried and the chemical contencs were calculated from the

weight increase. The water-extracted amount from the wood was found to be

negligible. Before the heating the samples were dried longer than 48 h at a temperature 60° C. Resin samples of urea-formaldehyde and melamine-formaldehyde, tested

partly, were prepared in the “Laboratory. The former resin was symthesized

from 1 mole of urea and 1.8 mole of formaldehyde in catalysts, hexamethylenetetramine and sodium hydroxide by being hardened with a hardener

ammonium chloride at a temperature 105° C. The latter was prepared from

pure oligomer soak MA-50 (from Honen Food Oil Inc.) by heating at 120° C for

2 h. Both the samples were pulverized so as to pass a 200 mesh screen. Before the heating they were also dried at 60° C.

2-

2. Heating and gas analysis

For measurement of N0.

c and SO 2 the 22 x 22 x 1 cm specimens were heated in a surface flammability test apparatus in accordance with JIS A

1321-1975. A part of combustion gas in a flow rate 3 L/min was introduced

from the smoke chamber into Fuji Electric Infrared Gas Analyzers Model ZAL

and ZBYB to analyze SO? and N0X , respectively.

On the other hand, the pulverized wood about 1000 mg or resin about 30

mg was heated by an imaging furnace in a thermal balance Sinku Riko TG

1500RH-J-S at a temperature 600°, 800°, or 1000° C in an air flow rate 30,

100, 300, or 600 mL/min. The sample temperature reached a desired value

from the room temperature within 1 min. Pyrolysis gas was introduced into

a series of two spiral traps through a pass heated at 300° C, as shown in

Fig. 3. The traps were chilled by dry ice—methanol in Dewar flasks but the

first one was heated at the gas inlet portion by the mantle heater of 300°

C. Therefore, it condensed products in a wide temperature range. The

products in the traps were resolved in 0.1 N sodium hydroxide solution and

analyzed for CN~ by Yokogawa-Hokusin Ion Chromatoanalyzer Model IC-100 with

an electrochemical decector Model ED-11.

3. RESULTS AND DISCUSSION

3-

1. Evolutions of N0X and SO 2 Concentration of NO^ in the smoke chamber increased with the heating

time, as shown by examples in Fig. 4. The gas accumulated in the smoke

chamber reached the maximum concentration in a time range near 10 min. It is

noted that the concentration started to increase earlier with specimens

treated with' nitrogen containing retardants.

418
The maximum N0X concencra cion in 10 min increased with Che concencs of

guanidine sulfamace, ammonium phosphate, and ammonium sulfamate, but was reduced wich boron comDounds and sodium chloride, as shown in Fig. 5. That

for ammonium bromide seems Co increase wich Che samll certain concent but

Co be retained constant at chis increased value in spice of che concent

increase more chan che certain point around 10 kg/m3. Similar relations were

found for the maximum NCk, concentration normalized for che weight loss of

specimens, namely, che maximum NC^ concentration per unit weighc loss. N0X

is accepced to be originated from both of fuel and air, and called fuel N0X

and thermal N0X , respectively. The increase in N0X

from che specimens

treated with che nicrogen containing retardanCs is obviously accri'oucable

Co che fuel NOx production, as shown in Fig. 6, where che N0X

per unit

weighc loss is shown co increase wich che nicrogen content of che created

specimens. It may be said on che basis of a regression line in che figure

Chat guanidine sulfamate and ammonium sulfamace produced more N0X than

ammonium bromide and ammonium phosphate ac an equvalenc nicrogen concent.

It should be noced chac che latter two retardants contain nicrogen in

ammonium groups while guanidine sulfamace has it in amino and imino groups,

and that ammonium sulfamate at a middle place of Che NOx production has both

ammonium and amino groups. Therefore, amino and imino groups mighc produce

more NOx chan an ammonium group.

The decrease in NCt^- with* che treatments by the chemicals other chan

che above retardants is inferred to be caused by che reduction in thermal

NOx , since NOx increased wich ca9, as shown in Fig. 7. The cd9 is assumed

to be an indicator for heac relaease, which decreased with che retardant

treatment. By the way, NO^ from che untreated wood is attributable co not

only Chermal one but also co fuel NOx, because wood usually contains

nitrogen less chan 1 % of che weight (2-7)

„

Also SOg was generated from che combustion and accumulated in che

smoke chamber, as shown in Fig. 8. The generation is especially high from

the specimens treated with sulfur containing retardants./ The maximum SOp

concentration per unit weight loss of specimens increased with Che ammonium

sulfamate and guanidine sulfamate contents, and slightly with ammonium

bromide, as shown in Fig. 9. However, chose for ammonium phosphate and

boron compounds decreased with che concents, and chac for sodium chloride

is constant independently of che concent. Since che sulfur content of

ammonium sulfamate (28.1 % in che weight) is higher chan that of guanidine

sulfamate (20.5 %) , che higher S02 concentration was yielded from che

ammonium sulfamace created wood chan from che guanidine sulfamace created

one. The SO 2 production per unit weighc loss seems almost equally co increase with che sulfur content between the guanidine sulfamace and

ammonium sulfamate created samples. Chough che latter samples gave a linear

regression line with a slighcly steeper slope and more scattered plocs, as shown in Fig. 10. Since wood usually contains sulfur less chan 0.1 % of Che

weight(5,3), the untreated specimens seem to produce low SOg concentrations.

Sulfur as well as nitrogen is believed co present in proteins in wood. It

is considered chat Che proteins is first pyrolyzed and chen S02 is formed.

The pyrolysis and che SO 2 formation would be affected by che recardancs.

3-2. Evolution of HCN

All HCN production values were converted into Che molar concentration

in che sodium hydroxide .solucion per one gram of burnc samples.

The HCN productions from che wood samples increased with che recardanc

concents, as shown in Fig. 11 where resulcs from che heacing ac 800° C are given. Ic is shown thac che HCN production is che highesc from che guanidine
sulfamate treated wood and the lowest from the ammonium phosphate treated

samples. No difference is discernible between the samples treated with

diammonium hydrogen phosphate and ammonium dihydrogen phosphate. The HCN

increase with the retardant content is linear in a low air supply rate as

30 mL/min, but in higher air flow rates the increase rate decreasingly

changed with the content and was inclined to be leveled off. The HCN

production was reduced with the heating temperature lowered to 600° C. Also

at this temperature in 30 mL/min of air flow rates the production increased

linearly with the retardant contents, as shown in Fig. 12. From the linear

plots the HCN production per unit mass of retardants was determined and

listed together with that at 800° C in Table 1. The table clearly shows the

decrease in the HCN production with the reduced temperature.

The effects of air are more definitely seen in Fig. 13. Here the

differences of the HCN production among the retardant contents are shown to be noticeably large in the lowest air flow rate but gradually to decrease

with the increase in the air flow. The HCN production in the highest air

flow 600 mL/min, especially from the ammonium phosphate treated wood is

almost equal irrespectively of the different content of the retardant.

The nitrogen contents in the weight of retardants are 35.9 % for

guanidine sulfamace, 24.6 % for ammonium sulfamate, 21.2 % for diammonium

hydrogen phosphate, and 12.2 % for ammonium dihydrogen phosphate,

respectively. Therefore, the HCN production appeared to depend on the

nitrogen content. From plots "of the HCN production against the nitrogen

content in Fig. 14, however, it is seen that again guanidine sulfamate

produced the highest HCN and ammonium phosphate gave the lowest HCN

production in an equivalent nitrogen content. This suggests that HCN

production is different depending on conformation of nitrogen compounds and

that amino and imino groups yield more HCN than an ammonium group. Furthermore, it should be noted that • guanidine sulfamate owned C-N bonds by itself.

The HCN productions from urea and melamine resins decreased with the

increase in the air supply 30 to 100 mL/min, and slightly decreased or was retained constant in spite of the air supply more than 100 mL/min, as shown

in Fig. 15. Tne temperature dependences of the HCN production in 30, 300,

and 600 mL/min of the air flows are given in Fig. 16, where the temperature

dependence is shown to decrease with the air flow rate.

Tne monomers of the two resins, urea and melamine have two and three

amino groups bonded to carbon, respectively, as shown in Fig, 17. Furthermore, melamine combines three nitrogens in its heterocycle. Then both of

the monomers own two nitrogens for every- carbon. From the elemental analysis,

the urea and melamine resins are estimated to contain nitrogen in an approximately equal quantity, namely about 31 and 28 Z of the weight,

respectively. Nevertheless, the latter resin usually gave HCN more than two

times of that from the former.

The urea resin suddenly starts to lose the weight in a high rate at

ahout 230° C in both air and helium atmospheres of the thermo gravimetry (TG)

C9). On the other hand, it is found that the melamine resin decomposes

slowly at first and rapidly loses the weight at about 360° C in air and

405° C in helium flow of the TGG-0) . The rapid weight loss probably reveals

the heterocyclic ring cleavage. It is interesting that the thermally stable

resin produced more HCN than the unstable resin. It might occur in the

pyrolyses that the less stable resin caused a lot of scission of C-N bonds

to form more oxidized products, while the melamine resin retained more C-N

bonds for a longer time . and formed more HCN, In addition it has to be taken

into consideration that the heterocycles of the melamine resin, which

imparted the high thermal stability, have three C-N bonds for every carbon.

420
and chat the urea resin has for the most part two C-N covalent bonds for

every carbon except side groups such as methylol, methylene, and so forth.

These factors may be the cause of the different HCN productions between the

two resins.

If the whole of nitrogen in the samples is converted into HCN, though

the reaction mechanism in fact may be different depending on the nitrogen

compounds, one gram of the retardants or the resins should produce HCN

ranging from 8700 ppm in mole for ammonium dihydrogen phosphate to 25600 ppm

for guanidine sulfamate. At the present study, however, the HCN production

is far low, namely, which ranges from 0,07 % of the above potential value

for diammonium hydrogen phosphate to 0.91 % for the melamine resin.

Guanidine sulfamate yielded the highest percentage 0.21 % of the treated

wood samples, which compares one half of the highest value of ones for the urea resin.

4 , CONCLUSIONS

1) Evolution of NO.^ from combustion of wood increased with nitrogen

containing retardants and decreased with, other retardants due to reduction

in thermal NO... Amino and imino grouns in the retardants produced more NO.,

than an ammonium group,

2) Wood treated with sulfur containing retardants yielded more S0 ? than other retardants.

3) HCN production from- wood treated with, nitrogen containing retardants

increased with the contents , -and decreased with. the increase in the air

flow rate and the decrease in temperature 800° to 600° C. Amino and imino

groups in the retardants^ produced more HCN than an ammonium group.

4) Urea resin produced less HCN than melamine resin, probably because

the urea resin was easier to 'be decomposed and oxidized at the C-N bonds

than the melamine resin.

5) It can more satisfactorily explain results previously reported of

combustion toxicity (.1) that toxic gases other than CO, especially HCN 'were observed from combustion of wood treated with guanidine' sulfamate , ammonium

sulfamate, and ammonium phosphate.

Acknowledgement

The authors wish to express their thanks to Inst., for the advice and cooperation in the HCN

Y. Fukui, Forestry 4 Forest Products Res. Inst.,

measurement

.

Dr. So Yusa, Building Res.

measurements, and to Mr.

for the N0.r and SO?

1)

2 )

3)

4)

5)

6 )

7)

8 ) 9)

References

Uesugi, S., Fukui, Y. , and Abe, H. : J. Jaoan Wood Res. Soc., 2_7’ ^8

0.981).

Nikitin, N. I.: "The Chemistry of Cellulose and Wood", Israel Program

for Scientific Translation, Jerusalem, p. 34(1962).

Furukawa, T. : J. Japan Forestry Soc., 45_, 99(1963)

Kato, Y., Doi, K. , and Furukawa, T. : J. Japan Fores-try Soc., 45_, 191 0963)

Tilman, D. A.: "Wood as an Energy Resources", Academic Press, N, Y. , p. 71, 73, and 84(1978)

Bechl, J. S. and ochers: ReDt. Office Technol. Assessment Congress

U. So, OTA-C-78-339. , p. 2460979)

Karchesky, J. and Koch, ?. : U. S. Forest Service General Technical

Repto, S 0-240979) . Sakurai, K. : Bull. Forestry Forest Prods. Res, Inst, , No, 335, 141(1986)

Hirata, T. , Oguro, A., and Inoue, A.: Wood Preservation, 12 , 279 0986)

421
Perenlogo

of

iurvIveU

mice

10) Hiraca , T. : Not published.

Maximum CO concenrration ( %

Fig. 1 Changes in capacitation tine with CO concencracion

oat- (NM4)jHP04 —-• — * MadOa

- —O— : nh4SQjn*»

Chemical concentration canned fa wood ( % )

Fig. 3 Heating apparatus

and HCN traps

Pig. 2 Changes in percentage of

survived nice with retardant

concentration applied to wood soecimens

422
(fi/lirtM)

01X

SSOI

|l|fi|.nv\

|

HIM

Incl

•

qN

Time ( min )

Fig* 4 N0X increase wich

hearing cime

Fig. 6 NOx increase wich nitrogen

concent

tig, o - i

Amount cnemicaia ;n wood ( '<q/m‘ )

z 4 rad * !0‘* CC min )

Fig. 7 NOx increase wich cd9

423
SO;

per

uni I

wclylil

loss

(

ppm/g

)

Tiffit ( mm j

SO 2 increase with

heating time

Fig. 9-1

dmounr of cnemtcais in -*000 ( <q/m‘ )

Fig. 9 Changes in SO 7 productions

with retardant contents

Retardant contents ot wood ( % )

Fig. 11 1 Air flow rate:

30 mL/min

424
& : (

• HSOaNM* —

x xyusoifi*** O « ( NM* txt^O* • t N»-UHrP04 r

/

4 X

/ o

Q t untr«al*^

Retardant contents of wood ( * )

Fig. 11-2 Air flow rate:

300 mL/min

HCN increase with

-retardant contents at

temperature 800° C

Fig. 12 HCN increase with

retardant contents at

temperature 600° C

Table 1 HCN production per one gram of retardants or nitrogen (ppm)

Retardants 800° C Retardant Nitrogen 600° C Retarddnt Nitrogen

NH4 H2 PO4 and

(NH4 ) 2 HPO4 6.4 34.6 3.1 14.6

NH4 SO3 NH2 27.4 112.0 19.5 79.3

CN 3 H 5 .HSO 3 NH 9 52.6 146.5 42.6 118.7

Fig. 13-1 Fig. 13-2

425
I

ICN

concentration

x

IQ

'1(

ppm )

400

f 300*C . 00/r</

A.r ft av* cz:e ( )

13 Changes in HCN production

with air flow race Changes in HCN

production with

nicrogen concents

Fig. 15 HCN production

TemoercTure ( *C ) Fig. 16 Temperacure deoer.dencas of

HCN production from urea

and melamine resins

from urea and melamine resins

HzN — C C —NH z

ii i N N C

i NH z Melamine

Fig. 17 Chemical formulae of urea and melamine

NH ZX

CO

/

NH z

Urea

426
TOXIC GASES FROM COMBUSTION OF WOOD TREATED WITH RETARDANTS

T. HIrata, Forestry and Forest Products Research Institute, Japan

GANN: If these same experiments were done in large scale, and the compartment

were allowed to proceed to flashover, is there a reasonable possibility that

all the nitrogen that is released would be converted to N0.

c ?

HIRATA: I don't think so.

427
BENCH- AND FULL-SCALE TOXICITY TESTS OF FRP FURNITURE

by Andrew Kim and Yoshio Tsuchiya

IRC, NRC , Ottawa, Canada, K1A 0R6

prepared for presentation before 6th Canada- Japan-US

A

Trilateral Cooperative Study on Fire Gas Toxicity

to be held at Factory Mutual Research Corporation, May 1987

Introduction

Dormitory-type buildings, such as university residences or penal

institutions, are compartmentalized to minimize fire spread from one room to

another. A severe fire, however, will still spread smoke and toxic gases to

other parts of the building, posing danger to the residents. The severity

of the fire and the production and spread of the smoke and toxic gases

depends on the contents of the room, interior lining materials, and

ventilation. In many dormitories the rooms contain simple furnishings,

often made from fire-retardant materials.

Bench-scale tests and full-scale tests were conducted at the Fire

Research Section, Institute for Research in Construction, National Research

Council of Canada (NRCC) to study the toxicity of fire gases produced by

three types of fibreglass reinforced plastic (FRP) furnishings. The first

part of the tests consisted of pyrolyzing a milligram quantity of the

specimens and, using a gas chromatograph/ mass spectrometer with a pyrolysis

unit ( Pyrolysis/GC/MS) , the pyrolysis products were identified and semiquantified. Then a gram quantity of the specimen was burned in a horizontal

quartz tube, the combustion products were collected by scrubbing with water,

and the water soluble components of the combustible gases were analyzed by

ion chromatography (IC).

In the second part of the investigations, full-scale room burn tests were carried out using the room-burn facility at the National Fire

Laboratory, NRCC. Prior to the burns, preliminary tests were conducted to

determine the type and severity of a fire involving combustible materials

normally found in a dormitory room. That fire was then simulated using a gas burner with a nominal 300 mm x 300 mm porous top surface (propane sand

burner) [ 1

]

.

Finally, the FRP furnishings made from the three different

types of fire-retardant material were burned in the test room, using the

simulated fire as ignition source.

Materials

Furniture made from three types of FRP (with different types and

amounts of fire retardants) were tested. They were identified by their

colours as orange, white and blue furniture.

Bench-Scale Tests

Pyrolysis/ GC/MS

A milligram quantity of the specimen was pyrolyzed using a Chemical

428
Data System Inc. Model #100 Pyroprobe* at a nominal temperature of 1000°C.

The products were directly injected into a Hewlett Packard Model 5996 GC/MS.

A methyl-silicone bonded phase capillary column, 50 m x 0.2 mm i.d. , was used for the GC separation. The oven temperature was programmed to increase

from -50°C to 250°C at a rate of 4°C/min. The components of the GC effluent

were identified by MS, using a spectra library and a search program, both of

which were supplied by the manufacturer of the instrument.

IC Analysis of Combustion Products

In a 1 m long, 17 mm i.d. quar

was heated in a tube furnace, a 0.5

in an air current. The temperature

The combustion gases were scrubbed

analyzed by IC.

All chemicals used in IC analy

received from the commercial source

a 4.6 mm i.d 25 cm long PRP-X100 co material was poly(s tyrene-divinylbe

p-hydroxybenzoic acid aqueous solut

was delivered to the column at the

SP8700 pump. The samples were intr

a 7001 Auto Injector fitted with a with a Vydac 6000CD conductivity de

checked with cyanide specific elect

Results

tz t ube , a 2 70 mm long se ction of which

g s pec imen was burnt on a ceramic boat

of the ove n was cont roll ed at 800°

C

• with wa ter; the disso Ived components wen

sis wer e of reagent g rade and used a s

s

.

The IC s eparation was carried ou t on lumn ( Hamil

t

on Co. )

.

The packing

nzene)

.

The carrier was a 4 millimo lar

ion

,

CO ntrol led at pH 8.6 with NaOH. It rate of 2 mL/min by a Spe ctra Physic s oduced through a Rheo dyne 7010 valve and

200 U L loop

.

The det ecti on was perf orme*

tect or. The concentr ations of CN- were

rode se parat ely

.

Approximately 60 organic compounds were generated from the pyrolysis of

FRP samples. Most of the compounds were common for all three samples. The

pyrolysis products of orange and blue furniture were especially similar.

The most significant products were styrene for all three, methyl

methacrylate for orange and blue furniture, and benzof urandione for white

furniture. The quantities of these compounds were roughly a few percent by

weight of original materials.

Hydrogen chloride was the main toxic gas detected by

quantity of hydrogen bromide was also detected. Quantitat

shown in Table 1. Based on the measured quantities of HC1

products, it was estimated that both the orange and blue f about the same amount of fire retardant, whereas the white

contained only 20% of the amount in the other two.

IC. A trace

ive results are

in the pyrolysis

urniture contained

furniture

Full-Scale Tests

Test Room

The test room was built according to the ASTM "Proposed

FIRE TEST OF WALL AND CEILING MATERIALS AND ASSEMBLIES" [l].

Method for ROOM

The interior

*Some commercial equipment is identified in this paper to adequately

describe the experimental procedure. Such identification does not imply

recommendation or endorsement by the National Research Council, nor does it

imply that the equipment identified is the best for the purpose.

429
floor dimensions were 2.44 m x 3.66 m, and Che ceiling height was 2.44 m. The Cest room was a concreCe-block structure with a concrete floor. The

interior surfaces were covered with calcium-silicate boards having a density

of 736 kg/m 3 and a nominal thickness of 12 mm.

The room had one 0.

2.44 m x 2.44 m walls,

insulation on the inside

0.8 m x 1.0 m portion of

allow the viewing of the

approximately 30 mm was

76 m x 2.03m doo

A plywood door (c

surface) was ins

the door was cov interior when th

left between the

rway in the centre of on overed with ceramic fibr

tailed in the doorway,

ered with polycarbonate

e door was closed. A ga

bottom of the door and t

e of

eThe

glaz

p o f he s

the

central

ing to

ill.

A schematic drawing of the experimental set-up is shown in Fig. 1. At

the top of the doorway, immediately adjacent to the exterior wall, a hood

was installed to collect the combustion products leaving the room. The hood

measured 2.4 m x 2.4 m, with a depth of 1.1 m. It was connected through a 0.92 m cubic plenum to a 0.4 m diamater horizontal exhaust duct with a circular aperture of 0.3 m at the entrance. The downstream end of the

exhaust duct contained a fan with a capacity of 2.4 m

3 /s. The length of the

horizontal exhaust duct was approximately 6 m.

Instrumentation

A thermocouple and a bi-directional flow-probe were centre of the exhaust duct, approximately 3.6 m from the

and a laser smoke meter was placed in the duct approximat

downstream from the thermocouple location.

installed

centre of

ely 0.3 m

in the

the hood,

Two gas-sampling ports we approximately 1.2 m above the

approximately 4.3 ra downstream

concentrations were measured u concentrations, a paramagnetic

taken in evacuated flasks from

analysis.

Description of Tests

The tests were carried ou three preparatory tests (Tests

the burning characteristics of

dormitory room, so that these

propane sand burner. The simu

fire in the second phase tests

The combustible materials

4 towels, 4 bed sheets, 2 pill

2 pairs of jeans, 2 jean jacke

of writing paper, 1 roll of to

2 parkas, 1 plastic model and on all combustible materials p severe fire.

r

f

s

e provided, one

loor, the other

from the centre

ing non-dispersi

oxygen analyzer,

the room and the

at the side of- the

in the exhaust due

of the hood. The

ve inf rared analyz

Additional gas s exhaust duct for

room

,

t,

CO and C02

ers, and 02

ample s were

IC

in two phas es. The first phase consisted of

#1, n, and # 3) . The:y were aimed at studying

com bust ible sol ids typically found in a typical fires could then be simulated by a ated fi re would then be used as the ignition

in the tests consiste d of 2 mattresses , ow cases, 2 bedspreads, 2 blankets, 2 pillows,

ts, 8 books, 1 newspaper, 12 magazines, 1 pad

ilet paper, 2 T-shirts, 1 package of matches,

1 can of Lysol. The can of Lysol was sprayed

rior to ignition to create intentionally a

In Test #1, the combustible materials

as severe a fire as possible (for example, were torn and crumpled, and the sheets and

were purposely arranged to yield

all books, magazines, and papers

towels were torn and spread out).

430
In Test // 2, these materials were piled in the test room without giving

consideration to the arrangement. Test // 3 was devised to simulate Tests //I and #2 using a propane sand burner (300 mm x 300 mm size) with no solid

combustibles in the room.

The second phase testing consisted of three full-scale fire tests (A,

B, and C) with three types of FRP furniture. The furniture consisted of a bed, a-clothes locker, a desk, and a stool, arranged in the test room to resemble the furniture arrangement of a typical dormitory room (Figs. 2 and

3). These were the only combustibles in the room. The propane sand burner

(providing the ignition fire) was placed at the junction between the bed and

the clothes locker to give the furniture maximum fire exposure. The tests were started by spark ignition of the sand burner. The propane flow rate to

the burner was controlled to correspond to the heat release rate and

duration of the fire created by burning combustibles in the preparatory

tests.

It was hypothesized that in a real fire situation, e.g., one created

intentionally by an occupant in a prison, the door would be kept closed at

the beginning, and it would take approximately 3 min for people outside of

the room to notice the fire and open the door. Accordingly, for the initial

3 min of each test, the door was closed, and the temperature and the

concentrations of CO, C02 , and 02 in the room were measured. At 3 min, the

door was opened and left open for the remainder of the test.

Gas temperatures, velocities, and concentrations of CO, C02 , and 02,

and optical density in the exhaust duct were continuously measured,

following the opening of the door. Gas samples were also collected in

evacuated flasks, using the sampling port on the side of the room, at 1 min,

and 2 min 30 sec. A few more gas samples were collected after that time

from the exhaust duct at irregular intervals.

Results

Figure 4 shows the concentration of oxygen in the room vs time for

Test #2. The oxygen concentration decreased sharply during the initial

3 min of the test. When the door was opened at 3 min, there was a sudden

influx of fresh air into the room, while hot, smoky fire gases exited

through the upper half of the doorway. The oxygen concentration in the room suddenly increased and the combustibles reignited. The subsequent fire

growth was very rapid.

Figure 5 shows the concentration of CO in the room vs time for Test //2. The maximum concentration occurred while the door was closed. The peak

value, 0.7%, was dangerously high, considering that the danger level is 0.4%

for short-term exposure to CO [ 2]

.

Test #3 was a simulation test of Tests #1 and // 2, using a propane sand

burner. After several tries, the propane flow rates were determined to

approximate the room temperature and the release rate in Tests #1 and # 2.

The variation of the propane flow rate used in Test #3 is shown in Table 2.

The rate of heat release for the tests with the three types of FRP

furniture is shown in Fig. 7. The information was derived using the "Oxygen

depletion method" [ l] . The heat-release rate in Test A (orange funiture)

was small compared to that in Tests B and C. In Test A it was close to that

431
of Che ignition fire (Test #3), indicating that the orange furniture did not

contribute much to the severity of the fire- In contrast, the white

furniture ignited quickly and, once ignited, contributed significantly to

the room fire.

All three types of furniture produced heavy smoke with

gases. Figure 8 shows for Tests A, B, and C, the variation

density of the fire gases measured with a laser smoke meter

diameter exhaust duct. The gas flow rate was approximately

included some entrained fresh air.

The results of the toxicity test for the fire gases in Tests A, B, and

C are shown in Table 3. The primary toxic components were HC1 and HBr. Gas

samples taken at 1 min and 2.5 min represent gas concentrations in the test room with the door closed. The remaining concentrations relate to the

samples collected in the exhaust duct. The latter values were converted

into rate of generation of these gases in the room, using information on the

gas flow rate in the exhaust duct.

The concentrations of HC1 given in Table 3 were not very high compared

to the danger level, which in a 30-60 min exposure is 1500-2000 ppm | 2

! . HBr has a threshold limit value (TLV) of 3 ppm ! 2_] , but no data are available on the concentration of HBr causing acute toxicity. Even though

there is not necessarily a direct relationship between the TLV and the

danger level, knowing that the danger levels for HBr and HC1 are in the same range, one can estimate that the probable danger level of HBr, based on the

TLV for HBr and HC1 (5 ppm), is approximately 1000 ppm in a 30-60 min

exposure.

The white furniture, which contained benzof urandione and styrene,

produced gases of particularly high concentration in HBr for the first

minute during Test B. A possible explanation for the finding is that the

furniture was burning fiercely at that time and, with the door closed, all

the fire gases accumulated in the room. In this test the locker ignited

within 30 sec of the test, resulting in rapid fire development in the room. The intense fire lasted for approximately 30 sec, after which the flames

started to decrease. No flame was visible at 1 min 10 sec, probably because

of the lack of oxygen in the room. Among the three types of furniture, only

the white furniture contained Br.

Comparing the HC1 and HBr concentrations in Test B, it was noted that

the concentration of HBr in the first two measurements (taken from the room

at 1 min and 2.5 min) was much higher than that of HC1. However, the HBr

concentration in the remaining three measurements (taken from the exhaust

duct at 4 min, 6 min, and 10 min) was approximately the same or even lower

than that of HC1. A probable explanation is that HBr has a higher boiling

point than HC1, and thus has a higher tendency to condense on surfaces.

Therefore, the HBr concentration in the exhaust duct is expected to be lower

than in the test room . The true HBr generation rate in the room might have

been substantially larger than the rate calculated from HBr concentration

measurements in the exhaust duct.

While the door was closed, concentrations of CO and C02 in the room

were dangerously high in all tests. The maximum concentration of CO was between 0.4% and 1.0%, higher than the 0.4% danger level for short- term

exposure of CO. The average concentration of C02 was between 5% and 7%,

some toxic

of the optical

across the 0.4 m

0.7 m^/s and

432
slightly lower than the 10% danger limit for short-term exposure | 2 J «

Even with a sufficiently large ignition fire, the orange furniture did

not burn well, owing probably to the fire retardant contained in the

material. However, the blue furniture with the same amount of fire

retardant burned reasonably well. This observation can be attributed to the

difference in geometry of the two lockers, which were made by two different

manufacturers. The orange locker was wide and shallow and had four shelves,

whereas the white and blue lockers had only one shelf at the top and were narrower and deeper. In Tests B and C, the flame impinged directly on the

bottom of the locker's top shelf and on its inside back surface. With the

orange locker (Test A), the flames impinged only superficially on the front

surface of the locker.

Summary

Full-scale room burn tests were carried out to study the toxicity of

the fire gases produced by three different types of FRP furniture, referred

to as orange, white, and blue. Combustibles typically found in a dormitory

room were used to produce an ignition fire, which was sufficiently large to

nearly cause flashover in the room.

According to the hypothetical scenario (door closed for the first

3 min) , the occupant would have been in danger because of high CO and C02

and very low 02 concentrations in the room during the first 3 min, even

without any FRP furniture present.

References

1. "Proposed Method for ROOM FIRE TEST OF WALL AND CEILING MATERIALS AND

ASSEMBLIES", pp. 1618-1638, 1982 Annual Book of ASTM Standards,

Part 18.

2. Sumi , K. and Tsuchiya, Y. , "Assessment of Relative Toxicity of Materials

Toxicity Index", paper presented at the International Symposium on Toxicity and Physiology of Combustion Products, University of Utah,

March 22-26, 1976, NRCC 15367.

433
TABLE 1. Determination of HC1 and HBr

in % by weight of original samples

Sample HC1 (%) HBr (%)

Orange 2.76 trace

Blue 2.93 undetectable

White 0.58 0.1

TABLE 2. Propane Flow Rate

Time (min) Propane flow rate (L/min)

0 - 1.5 375

1.5 - 3.0 0

3.0 - 3.5 300

3.5 - 4.0 375

4.0 - 5.0 350

5.0 - 6.0 320

6.0 - 7.0 275

7.0 - 8.0 250

8.0 - 10.0 225

10.0 - 12.0 175

12.0 - 16.0 125

16.0 - 24.0 105

24.0 -'30.0 85

434
TABLE 3. Results of Analysis using Ion Chromatography

(A) TEST A (orange furniture)

HC1 generation rate

Time

(min

)

HC1 concentration

(ppm

)

in the room

(g/min

)

1.0 175.8

2.5 0.0 - 4.0 16.5 1.22

6.0 135.8 9.33

10.0 164.3 1 1.65

15.0 88.9 7.62

25.0 34.4 3.14

(B) TEST B (white furniture)

HC1 generation rate HBr generation rate

Time HC1 cone. in the room HBr cone. in the room

(min) (ppm) (g/min) ( ppm) (g/min)

1.0 14.7 - 1132.1 - 2.5 1.4 - 79.3 - 4.0 23.8 1.64 29.8 4.55

6.0 98.9 6. 10 29.2 4.00

10.0 14.8 0.83 10.2 1.27

(C) TEST C (blue furniture

)

HC1 generation rate

Time HC1 concentration in the room

(min

)

( ppm ) (g/min )

1.0 360.9 - 2.5 17.6 - 4.0 436.9 31.48

6.0 1184.4 75.87

10.0 874.7 50.78

15.0 114.5 9.59

435
FIGURE 1 SCHEMATIC OIAGRAM OF TEST ROOM

436

GAS

SAMPLING
LU

Q_<

on cz. < O

O Qo

LU

Q_C

LTl

lo c£ < o

O Q437

FIGURE

3

FlGURE2

EXPERIMENTAL

SET-UP

FOR

TEST

B

AND

EXPERIMENTAL

SET-UP

FOR

TEST

A

(WHITE

AND

BLUE

FURNITURE)

(ORANGE

FURNITURE)
CO

CONCENTRATION.

%

0

9

CONCENTRATION.

FIGURE 4 CONCENTRATION OF OXYGEN IN THE ROOM

vs TIME IN TEST No 2

TIME, ti i n

FIGURE 5 CONCENTRATION OF CO IN THE ROOM vs TIME IN TEST No 2

438
RATE

OF

HEAT

RELEASE,

MW

TEMP,

300

200

0 5 10 15 20 25 30 TIME, min

FIGURE 6 AVERAGE ROOM TEMPERATURE vs TIME IN THE

PREPARATORY. TESTS

FIGURE 7 RATE OF HEAT RELEASE vs TIME FOR THREE TESTS

WITH FRP FURNITURE AND FOR THE REFERENCE

TEST (TEST No 3)

439
OPTICAL

DENSITY

3.0

2.5 -

2.0 -

1.5 -

L.O -

TEST A TEST B TES T C TES T 3

10 15 20 25 30 TIME, min

FIGURE 8

* OPTICAL DENSITY vs TIME FOR THE THREE TESTS

WITH FRP FURNITURE AND FOR THE REFERENCE

TEST (TEST No 3)

440
BENCH AND FULL SCALE TOXICITY TESTS OF FRP FURNITURE

A. Kim and Y. Tsuchiya, National Research Council of Canada

TEWARSON: You were measuring temperature, and on the basis of that, you set

the propane flow. After that, you were measuring rate of heat release, and

that, I suppose, was in the duct. Is that rate of heat release reaction,

chemical heat release rate, or convective heat release rate?

KIM: Convective heat release.

TEWARSON: Okay, so that is a convective heat release. Therefore, if one is

using gas temperatures in the room, is it in any way proportional to

convective heat release rate?

KIM: No, temperature use is after that.

TEWARSON: HCN stays in the air in the room, is that in the room or in the

duct?

KIM: In the duct.

TEWARSON: But your Table 3 says HCN generation rate in the room.

KIM: HCN rate is determined at the duct. PPM means determined at the room,

1.2 meters from the floor.

TEWARSON: And also the temperature in the Figure 7 is the temperature in the

due t ?

KIM: I don't think so, the temperature is in the room.

TEWARSON: I was confused as to whether it was only in the room.

KIM: The temperature in the duct is much lower than in the room.

TEWARSON: Is there an actual simulation using the room? How many

thermocouples were used?

KIM: One to three thermocouples were used.

TEWARSON: At one place?

KIM: At one place.

TEWARSON: These were measured in the duct?

KIM: Yes, everything mixed at the duct.

GANN: The rates of heat release that you show, are there enough to flashover

the room?

KIM: Yes, but in this case, my calculations are scenario -based

.

441
GANN: Did the room flash over? The reason I ask this, if this sample is

totally burned, and if the room is very hot, I would expect near quantitative

yields of HCi . In a very crude visual integration of the numbers table, it

doesn' t seem to give that'.

KIM: It seemed valid.

GANN: CO levels were very low?

KIM: We have a complete record.

MORIKAWA: This is a full-scale test. The temperature seems not to be going

up much, and the CO density was low when we were testing. Maybe you should

have had more things so that you would have had the conditions of flashover.

KIM: Of course I mentioned that. However, we have only one set for each

material

.

442
EVOLUTION OF TOXIC GASES FROM EXPERIMENTAL

FIRES IN AN EXISTING BUILDING

T. Morikawa and E. Yanai

Fire Research Institute, Ministry of Home Affairs,

14-1, Nakahara 3 Chome, Mitaka, Tokyo 181

T. Nishina

Shizuoka Prefectural Institute of Public Health

and Environmental Science,

4-27-2 Kitaando, Shizuoka 420

ABSTRACT

Two compartment fire experiments were conducted in a two story ferroconcrete building to investigate evolution of toxic gases and toxicity of

atmospheres in the burn room and its surrounding area. One was the

combustion of natural polymer contents, and the other was for synthetic

polymer-containing contents. Major toxicants evolved such as CO, HCN , HC1

S0 2 , NOx, acrolein, formaldehyde were determined and toxic effects were evaluated by means of toxicity index, z(Ci/Cfi). Mice and rabbits were

used as test animals and exposed to fire effluent gases in the burn room

and exposure boxes. For rabbits, blood analysis and other biological

examinations were carried out to find out the cause of the death or incapac itation.

INTRODUCTION

Recently there has been a growing number of fires which have caused

death by toxic gases poisoning.

Synthetic materials have' often been blamed for great evolution of

toxic gases, although as much toxic gases could be evolved from natural

materials such as wood and wool. Toxic gases evolution depends not only

on the material burning but also on such factors as fire atmospheric

temperature, oxygen concentration, and perhaps room size.

For the present study, fire experiments were conducted in a dormitory

that was to be dismantled, to investigate the evolution of fire toxic

gases in a home environment using chemical and biological analyses.

EXPERIMENTAL

Fire experiments

The building offered for fire experiments was a two story ferroconcrete building which had been a nurses dormitory. The plan view of the

first floor is shown in Fig. 1, and that of the second floor was almost

the same. Two rooms, the third and seventh rooms on the first floor were

used as burn rooms. They are the same in shape and space, and have a floor space of 16m2 and the ceiling height of 25Ccm each. The fixed

combustibles of each room are 350kg (22kg/m2

) in weight including built-in

closets and cabinets, tatami (straw mat), window frames and pillows. But

the combustibles below the tatami -covered floor were excluded, because

they are not usually involved in fire.

Furniture, carpets, curtains, apparel and other house contents were

carried into each room. Such mobile combustibles were 450kg for each room The list of the mobile combustibles is presented in Tab. 1. The fixed and

mobile combustibles in "Room 3 were all natural polymers such as wood, wool

cotton, and leather, while those in Room 7 were 27% synthetic and 7 3 %

natural polymers by weight.

The layout of the combustibles for room 3 is shown in Fig. 2, and

that for room 7 is nearly the same. A sketch of the burn room, corridors

443
and a stairway hall is shown in Fig. 3. All the doors and windows except

a window in the landing of the stairway hall were closed. The opening of

the window in the landing was 73cm x 100cm (height). The first floor

corridor was blocked with fire resistive boards at the far end of the

stairway hall, so that smoke and gases could not go into the corridor

beyond the boards.

A cotton cushion (50cm x 50cm, 200g) which was smoldering over the

entire surface was used as a fire source; It was inserted between another

cushion and a blanket on the floor with a bedding layer over all.

Measurements and analyses

Temperature and optical smoke density were measured at various points

in the burn room, the corridor, and the stiarway hall, using C.A. t.nermo- couples of 0.32mm dia. for temperature measurement and smoke meters with a tengstem lamp as a light source and SE cell as a detector for smoke measurement.

Fire effluent gases were sampled into 200ml syringes and 5 % teflon

bags at two points; one just below the ceiling in the center of the burn room and the other at the landing of the stairway hall, as shown in Fig. 3. Gas samples were taken at appropriate time intervals during the fire

experiment. Sampled gases were subjected to gas chromotography , ion

chromatography and colorimetric analyses.

Animal exposure tests were conducted using mice (dd strain, female

2 Og ) and rabbits (Japanese white, female, 2kg) as test animals. Four mice were placed in the center of the burn room, two just below the ceiling and

two on the floor, and 3 mice were placed just below the ceiling of the

second floor of the stairway hall. Another 3 mice were placed in each of

2 animal exposure boxes, ‘into which fire effluent gases were introduced.

One box received gas from a point just below the ceiling in the burn room

and second from the second floor stairway hall. The fire effluent gases were cooled by ice water before being introduced into the animal exposure

boxes. Each mouse was contained in a free rotary wheel.

The box has a room for 3 rabbits accommodation in addition to an animal exposure room where also 3 mice were accommodated. The head of each

rabbit was inserted into the animal exposure room. Time to incapaciation , or, time to wheel rotation stop was measured

for all the mice. A behavioral observation was made for mice and rabbits

in animal exposure boxes. Rabbits were subjected to electrocardiograDhic

analysis, body temperature measurement and urine tests during the exposure.

Blood analysis and patholo-nistological examinations were conducted for

rabbits after they stopped breathing.

RESULTS AND D ISUCSS ION

Behavior of fires

Temperature in the burn room with time for Exp. 1 and Exp. 2 are shown

in Fig. 4. The maximum temperature was recorded at nearly 1000°C for both

experiments. Optical smoke density failed to be measured a little after

flaming because of fire heat. During the smoldering stage, the maximum

optical smoke desnity in the burn room was recorded at 6m -1 for Exp. 1 and

0.2m -1 for Exp. 2. In Exp. 1, after a 39 min. -long period of smoldering, flaming occurred,

then soon ceased after spreading to a cardboard box containing paper waste.

At 49 min., when one of the windows (140cm high and 270cm wide) facing

outside was opened to 15cm wide, flaming occurred again, and developed to some extent. At 81 min., when the window was closed, flames disappeared

again apparently because of suffocation. At 86 min., when the window was opened again to one fourth of the width, falming came back, and developed

444
to flashover at 108 min. Then smoke and flames came out of the window.

After 120 min. the fire was in its decay stage.

In Exp. 2, after smoldering continued for 30 min., flaming occurred and

developed nearly to flashover. However, it ceased suddenly when fire

almost spread to the ceiling, probably because of oxygen deficiency. At 50

min., when one of the windows was opened by 15cm, flaming occurred again,

and the fire developed to flashover at 60 min. with very thick yellow

brown smoke coming out of the window. Fire became more violent after 72

min. when the window was opened by one fourth of the width. A few minutes

later, some of the window panes were broken by fire heat, and fire developed even more violently. Then fire entered into the decay stage.

Evolution of toxic gases

For Exp. 1 toxic gas concentrations in the burn room with time are

shown in Fig. 5. During the smoldering combustion, almost no change was

seen in concentration of all the gases determined. Around 40 min. after

ignition when flaming occurred, 0 2 concentration dropped to 18% and C0 2 concentration rose to 3%, CO concentration also rose to 0.3%. However,

other toxicants such as HCN were not recognizably detected. Substantative

changes in concentration of gases did not occur until flashover at 108 min..

After flashover most toxic gas concentrations rapidly increased, reached

their peaks, and then decreased. The concentrations at 110 min. were: CO

5%, C0 9 16%, HCN 530ppm and acrolein 16ppm. The concentration of 0 2 was

2 . 8 %. Toxic gas concentrations in the stairway hall were not determined

until flaming occurred, because closed windows and a door on the corridor

side of the burn room were kept intact before that.

After the flashover, most toxic gas concentrations increased rapidly,

though much less than in the burn room, as shown in Fig. 6. For Exp. 2, toxic gas concentrations in the burn room with time are

shown in Fig. 7. As was the case with Exp. 1, most toxic gas concentrations increased just after the occurrence of flaming, then decreased after

cessation of the flaming, and again increased drastically to the maximum

after flashover. The concentrations just after the flashover were: CO

6.5%, C0 2 13%, HCN 2500ppm, HC1 800ppm and acrolein 150ppm, and were much

higher than those in Exp. 1. The concentration of 0 2 was 3%. Toxic gas

concentrations in the stairway hall after the occurrence of flaming are

shown in Fig. 8. Most toxic gas concentrations hit their peaks around

at 80 min.. The maximum concentrations were: CO 2%, C0 2 7%, HCN 400ppm

and acrolein lOppm. The minimum concentration of 0 2 was 13%. The maximum

concentration of HC1 , which occurred around at 60 min., was 187ppm.

The oxygen concentration as a function of the C0 2 /C0 ratio in the

burn room is shown in Fig. 9 for both Exp. 1 and Exp. 2. It is noted that

in the fully developed stage of fire, the C0 2 /C0 ratio is lower in Exp. 2 than in Exp. 1, although the oxygen level was almost the same in both

experiments. This indicates that the relative ventilation was lower in

the former experiment because of the former's large burning rate.

Biological analysis

1) Behavior of animals

Times to incapacitation of mice in free rotary wheels after ignition

are presented in Tab. 2 for both Exp. 1 and Exp. 2. Incapacitation did

not occur in any case of the mice during the smoldering combustion which

lasted until 39 min.- in Exp. 1 and 32 min. in Exp. 2. In the burn room,

incapacitation occurred shortly after the start of flaming, and in the

stairway hall it had to wait until flashover occurred. The mice in the

animal exposure boxes placed outside the building took a little longer

445
time to incapacitation, probably because fire effluent gases were cooled

before introduced into the boxes.

The results of the observation of rabbits are presented in Tab. 3. As

all the rabbits were exposed to fire gases which were cooled, their body

temperature did not rise during the experiment. Snivel was observed during

the smoldering stage of fire in which CO was almost negligibly low in

concentration. Sniveling must have been caused by irritant gases at

instrumentally undetectable concentration, especially acrolein which could

irritate even at below 1 ppm in concentration. Paralysis of limb appeared

as well as chain strokes 1-5 min. before breathing stop occurred. Another

noted behavior was occasional closing of eyes during the exposure.

2) Blood analyses

Levels of CO-Hb and cyanide in the blood of rabbits just after their

breathing stop almost equivalent to incapacitation are shown in Fig. 10.

Cyanide levels are higher for Exp. 2 than for Exp. 1, while CO-Hb levels

are almost the same for both experiments. The lethal equivalence curve

is based on CO-Hb lethal level of 50% and cyanide level of 3ug/ml

.

In these experiments, CO seemed to be a larger contributer to incapacitation than HCN. Erythrocyte and leucocyte counts, and hematocrit

values in the blood of rabbits are presented in Tab. 4. Both erythrocyte

and leucocyte counts decreased a little, but hematocrite values decreased

as much as 22.4% on the average.

3) Other analyses

Electrocardiographic analysis, urine tests, histological examinations

were conducted for rabbits.

It was found in electrocardiographic analysis that there were a reverse of T wave and a large fluctuations of beat rates. In the urine

tests, blood and protein were detected in urine, which indicates that

kidney trouble was caused by the gases inhaled during the exposure. The

organs on which pathl o-histol ogical examinations were conducted were heart,

kidney, lung and spleen. Abnormalities were recognized in some of the

organs including cardiac muscle necrosis, probably due to oxygen deficiency.

Toxicity evaluation

Interactions of toxic effects of different gases have not yet been

established. Total toxicity inaex (Ci/Cfi, where Ci is the concentration

of gas component i, and Cfi is its lethal concentration) was used in the

present study in evaluating the toxicity of fire effluent gases, since

toxic effects of every toxicant are treated equally in this method.

Lethal concentrations for 5-10 minutes exposure are 5000ppm for CO, 350ppm

for HCN, SOOppm for HC1 and SO 2 , and 30 ppm for acrolein [1], The lethal

concentration of formaldehyde was estimated to be 180ppm, because acrolein

is reported to be 6 times stronger than formaldehyde in toxicity to rats

C2].

The total toxicity index in the burn room with time is shown in Fig.

11 for Exp. 1 and Exp. 2. In both experiments, the total toxicity index was far below 1 in the

burn room during the smoldering combustion stage when no animals collupsed

At the flashover, the total toxicity index increased rapidly, reached its

maximum in a few minutes, and then soon began to decline. Generally after

flashover, the total toxicity index was higher in Exp. 2 than Exp. 1. The

maximum total toxicity index was 13 for Exp. 1 compared with 26 for Exp. 2. The toxicity index of CO was compared with the sum of toxicity indexes

of other toxicants, as shown in Fig. 12. It can be seen from this that

the former is apparently larger than the latter.

446
In the stairway hall, the total toxicity index increased rapidly

after flashover, and the maximum value was 3.6 for Exp. 1 and 5.9 for Exp. 2.which endorsed the incapacitation caused to the animals in the exposure'

boxes

.

REFERENCES

1 Terrill, J. B. et al. Science, 200, pp. 1343-1347 (1978)

2 Alexeeff, G. V. , J. Fire Sciences, 2, pp. 362-379 (1984)

447
Tab. 1 List of “.coil a conoustibl as

Exp. 1 ixc. 2

Chests of drawers 120 kc Chests of drawers 63 kc Oesk 13 Oesx 41 Chair 2 Cha ir d 3ook case 10 Sock case 23 root warmer d fzzz wcrr.sr n

Shoes case 7 Shoes case 20 Car see (wool) 20 Television set 27 Certains (cotton) d 7 el eonone

<* 2 ieccir.es (cotter.) ::_2 Caroet (PAN) C Pillows (cotton) 3 Certains (PAN, Polyester) 11.1 ilanxets (wcei

)

Ciotr.es (cotton, sslk. 21.3

5cccir.es ( PAN , PA, poiysscsr)

23.4

and wool) Mattress (flexible S d RuO C2T urstnane foam) Shoes 3 Pillows (svntr.etsc) 2.2

L sacner ilankets (PAN) 7 . d Carcooac coxes ii Clctr.es (PAM, PA

,

i r r News caters 10 :c 1 vaster

)

icoxs 1 Cj ?VC SC55CS 2 Woos s*i:cx3 11 Artificial leather d . s Wail caters d .News :eoe-s 10 Plywcac i7 .

3

5‘cc<s ICO Wccc scicxs n

r

’ VWCCC d. . d P.icic uretr.ar.e fear. 7 PVC well cecers 4

7 etc i —o Tccdi d ; <c (Synthetic 27. S'.)

Tao. 2 T ime to incaoaci tation of mica

Location Tescsc

Time to incatacitation (men) mica

CJCC. 1 cxo. 2

Earn -ter. .No. 1 40 Pail ore

(Ceil inc;

2. 33 3Z

5err: reen i cu 22 [r-.zzr;

Z 50 22

Sea irway i 1C9 c'4 nail

z 10= 63

2 109 S2

Exrcsura box 1 dc 34 (Sum room)

2 19 34

3 43 34

Excosure box 1 111 62 (stairway hall

)

2 111 oZ

3 112 5Z

448
nesuics ct cssarvasicn cr rasrcs in axscsra :oxes

(

= C)

sxsssura e.xscsurs

Observed z:~a arsar isr, Irion (min)

Paralysis amvei ms craiS.O'; nc sou:

j

Z.O ZA

i. 1

23. SO ICS 29.70 ICS 29.90 1C3

110 nono

rwa y na i

1

113

I I c

1 U

:~.~s was nou c;aa r

Tdc. i Eryurrcc y-.a end leu:cscy-a ecu•-4 a.-.u herauccr oca values of racicc in 5.x ccs

u

r*5 -CX5S

— r- - - ( I - --r - - y 3

‘

•

1 C r — U/

. =i --- ' *•“ P22T ZZ exccsura exccsura Lxccsura exccsura exccsura axe: sura

Nc. 1 511 T 2T“ 3 1 533 9.223 -Z. 3 3 . 3

z 4-1

T

ii i 10. 2-22 z t r L. J L - . I d. - 3 2 233 3,-00 d ^ r n 53.3 Z-.u

For susirwav hall

Nc.l •r3 -55 3,200 c ^ -n -3 . 3 3G.3

2

d.3 £ -1

1

4,000 a,7GC 37.0 23.0

3 :co 302 2,700 4,200 22.0 21.0

E.xo.2 For hum roan

No . 1 £C2 IP c 3,200 1 ,200 41.0 22.0

2 5 5 , £30 2 ,533 *"L - U

"

1 Q

3 322 -c7 3,200 3,200 41.0 24.0 For soairwav hall NO . 1 51

1

533 d 3,000 41.0 55.0

2 470 429 9,100 4,000 41.0 23.0

3

' 433 LA q *nn l; 24.0 L4

LA: Uccracsry accident

449
Fig. I A plan viev of the first floor of the building

offered for fire experiments

Curtains Viiiii ow Shoes case

A I s n i lion source

E.T P . 1

Fig. 2 Layout of the room contents for Exo.l

Fig. 3 A sketch of experimental set-uo

450

i\M
'

(»'l>l()7.

X)

upivJ^V*

4

ra,l,l

UU<'. X)

IIJII

Fig. 4 Variations or ta=:erat’jr

calling in the csr.tsr cf

in ; tire just the son rocs

be i ov the

Exp . 1

S ta 1 rvay ha I I

Burn rooa

;aco

o

Burn rooo

ho

no. —• SOi HCHO

20 40 50 30 100 120 Time (o i n)

Fig. 5 Variations of concentrations

vitli tine in burn rooo for

of gases

Exp1

S

T i ae (a i n)

Sta i rvay ha I I

Fig.S Variations of concentrations of gases

with tine in stairvay hall for E.tP-1

451
PiS- 7 Variations of concentrations

vi th tiae in burn roo« for

of ?ases

ELxb.2

Tine (am)

Pig. 3 Variations of concsn tra ;i ons of gases

with tiae in stairway ball for Lxc.Z

• La ;-2

O'0 1C 20

Oa (S)

Pig. 9 COa/CO ratio vs. Oa concentration

Pig 10 CO - lib of

vs. cyanide concentrations

rat> its just after brea tiling

n b I ood

s too

452
Exp. I 8urn room

o 20 *10 60 60 100 120 Time (min)

- 10-

g co a HCN

10 HCL

Q Acrolein 3 HCHO

NO.

n

ClI ... 31 ^

si s

: 31

5l 55

i

01 gl

0 g

fj ^

hi si

AG SO Tine Cain)

glgl

gig!

?'e)| gjl&i

aa

Fig. 11 Variations of total toxicity index with tine

in burn roo« for Exo. 1 ana Exp.

2

Fig. 12 Toxicity index of CO vs. sura of toxicity index

of gases excluding CO

453
EVOLUTION OF TOXIC GASES FROM EXPERIMENTAL FIRES IN AN EXISTING BUILDING

T. Morikawa, E. Yanai , and T. Nishina, Fire Research Institute, Japan

RARTZELL: In reading over the text of the behavior of fires, it seems to me that the way in which the fires were manipulated was different in the two cases because of difficulty in keeping the fire going. In fire number two,

for example, the fire apparently burst into flame which would have gone out

had the window not been open to cause the flame to continue. Somehow, it seems to me that because of the difference in the way these two fires were handled, it's difficult to compare not? only the times of the events, but also

what happened because they were not treated identically.

MORIKAWA: As far as the condition of the equipment is concerned, there were common denominators between the two. One, the smoldering was the same; also,

the window was shut before and also opened. That kind of manipulation was the same. However, the time was different and then the flashover manipulation was

the same. So-

, we feel that we obtained satisfactory data out of this

experiment by focussing on the flashover phenomena and the smoldering.

HARTZELL: On the animal data, regarding heartbeat, for example, and other

things, how many of the animals were involved in that data? One animal or many animals?

MORIKAWA: As far as the electrocardiogram was concerned, we measured only one animal per experiment. However, we used three animals in the cage.

LEVIN: When did you gather the blood data?

MORIKAWA: Right after the breathing stopped, we immediately took the animal

out and then examined the red cells and the white cells. Therefore, breathing

had stopped, but the heart was still beating.

454
HUMAN EFFECTS OF POISONOUS GASES GENERATED DURING FIRE

ESPECIALLY, PRESENCE OF HCN AND CO

Yukio TSUDA and Yoichi NISHIMARU

Department of Legal Medicine, Yokohama City University

School of Medicine

3-9 Fukuura, Kanazawa-ku, Yokohama 236, Japan

ABSTRACT

In Japan it is well documented that there are abou

fire fatalities for a year. 'W Their cause of death had competed with

temperature, asphyxiation for oxygenmonoxide poisoning. But there has been a

2 000

burn for high

n and carbon

significant increase

in the number of fire victims recently, not only due to burns

but also because of poisonous gases produced from fires.

Structural probrems buildings as well as construction

materials and living commodities contribute to the production

of such gases. It is known that the burning of various

lead to oxygen depletion and release

HC1 , NH 3 and other gases.

Many reports and references have already been

of CO- CO , HCN

made regarding

0

2 ' CO 2 , CO. But the emphasis shall be on which is of the

greatest interest to the author. Therefore, the author

discusses human effects of poisonous gases generated during

fire- especially, presence of HCN and CO.

INTRODUCTION

In Japan, it is well documented that ’ther' are about 2000

fire fatalities for a year ( Table 1 ) . There categories are death due to burning, death by carbon

monoxide poisoning and asphyxia, death by contusion and

fructure, human torch suicide, and others.

Detailed autopsy studies of 67 fire fatalities in Department

of Legal Medicine, Yokohama City University School of Medicine

for four years from 1982 to 1985.

A medicolegal analysis was made of 67 cases in which COHb and

CN in blood levels were determined.

According to the study for the author many fire fatalities

are due to CO and HCN.

Findings from Fatality Study

Table 1 shows a number of fire fatalities for eight years

from 1978 to 1985 in Japan, Kanagawa Prefecture and Yokohama

City . Fig.l is shown a number of fire fatalities who were autopsied

in Department of Legal Medicine , Yokohama City University

School of Medicine.

According to the author's study approximately 30%-40% of the

fire fatalities are due to human torch suicide.

455
Burns have been classified into four degrees:

First degree - reddening of the skin ( erythema ) Second degree - blistering of the skin ( blister ) Third degree - loss of epidermis ( crust ) Fourth degree - charring of subcutaneus tissue ( carbonaz a t ion ) Burns degree of self-ignition ( human torch suicide ) are

third degree and fourth degree but their carbon monoxide level

in blood is normal or low level, and poorly diagnonosed

include soot in the trachea in addition to burn of respiratory

tract and they are not pugilistic attitude. Blood color is red

On the other hand, the cases of the fire fatalities are shown

the first degree, the second degree, the third degree and

fourth degree. Many fatalities are found dead in houses after

a fire. Note the soot present in the trachea, and the edema of

the larynx. There were 30%-8Q% saturation of the blood with

carbon monoxide. Blood color is crimson. Its presence in blood

together with soot i n the air passages which are usually

acutely engorged, constitutes proof that the victim was alive

at the time of fire and, usually, that the death was due to

the fire. In this manner it is possible to explain deaths

which occur in a fire when no burns or other signs of violence

are found on the body and there is no evidence of natural

disease

.

In these circumstances it is usual to find that the blood of

adult victims, especially the elderly, is about ha 1 f - s a tu r a t ed

with carbon monoxide; the results range from case to case

between 30 and 80 per cent. Althoug other factors, for example

shock, may play a part, this degree of saturation in an elderly person is of itself capable of causing death.

However, the number of the lower degree fatalities and burns

caused after death have increased recently; they are hard and

yellowish in color.

Blood color is not crimson often and a thin, friable,

chocolate colored, sickleshaped hematoma is not present in the

epidural space in the anterior and middle cranial fossae

inferiorly.

According to the authors' study, they are due to poisonous

gases generated during fire.

Those gases are C 0 ^

, CO, HCN, HC1, NH 3 and other gases.

Therefore, the authors have examined these poisonous gasesfoctors in blood of the fire fatalities, especially CO and

CN" . RESULTS OF BLOOD EXAMINATION

Inhalation of fumes had caused death from carbon monoxide and

anoxia . According to the authors' study, they are due to some poisonous gases generated during fire - CO^, CO, HCN, HC1, NH 3 and

othergases.

The authors have examined these poisonous gases factors in

blood of the fire f a t a 1 i t i e s , e s pe c ial 1 y CO and CN~.

i) Concentration of COHb ( Fig. 2 ) Under 10% : 22 cases, the level of 10% : 6 cases

the level of 20% : 7 cases, the level of 30% : 6 ta s e s

,

456
the level o f 4 0% 3 cases

,

the level o f 5 0% 3 cases

the level o f 60% 2 cases

,

the level o f 7 0% 3 cases

the level o f 30% 9 cases

,

the level o f 90% 1 cases

ii) of CN in blood ( Fig. 3 ) Under 1 ppm 3 cases

,

Under 10 ppm : 9 cases

,

the 1 evel o f 10 ppm 3 cases, the level o f 2 0 ppm : 6 cases

the level o f 30 ppm 5 cases , the level o f 40 ppm : 5 cases

the level o f 5 0 ppm 12 cases, the level o f 60 ppm : 4 cases

the 1 evel o f 70 ppm 2 cases, the level o f 3 0 ppm : 2 cases

the level o f 90 ppm 2 cases, above 100 ppm 14 cas 0 s . i i i ) pH of blood ( Fig . 4 )

59

39

09

2 9

4 9

2 cases

4 11

7

cases

,

cases , cases

,

cases

,

cases , cases , cases . 65 cases were acidosis,

relation of concentration

which are the

30 - 5.39

50-6 . 59

70-6 . 79

90-6.99

10-7.19

30-7.39

cases

10

8

5

2

1

cases

cases

cases

cases

cases

Under 6.29

6.40-6.49

6.60-6

30-6

00-7

20-7

40-7

Except

iv ) Mutua 1 The cases

and the

recently . v) %C> 2 Hb ( Normal value : 3 5 -98% ) Under 10% : 34 cases, the level of 10% : 16

the level of 20% : 11 cases, the. level of 30%

the level of 60% : 1 case, the level of 90%

The low letel of %02Hb results from anoxia.

CO Hb and CN ( Fig

of high level of

5 cm'

ion of low level of COHb have increased

cases ,

: 3 cas

2 cases.

SUMMURY

In old day, the cause of death had competed with a burn for

high temperature, anoxia for oxygen-starvation and carbon

monoxide poisoning.

But there has been a significant increase in the number of

fire fatilities recently, not only due to burns but also

because of poisonous gases produced from fires.

These facts were demonstrated by f-indings from inspection

and autopsy of fatalities, and by results of blood examination

o f f a t al i t i e s . The concentration of COHb in blood of fire

fatalities is low level and they have contained CN in blood.

I HCN gas is absorbed through the mucose the tr.cnea and lungs

by inhalation . When cyanide is absorbed and enters the blood

stream, functions of the oxiginate ferment are suppressed in

internal -organs and tissue cells; and the oxygen in blood is

prevented from being consumed. CN absorbed in the blood stream

in other words, immediately unites with ferments such as cy to chrome -oxyda s e , which are widely dispersed in cells and

interferes with their functions.

As a condition of recovering victims who were helped from a fire, how much of the damages done by CN , which is more

poisonous than CO is reversible?

This is a question that draws much interest and needs further

in the future

.

457
90-ager

458

Fig.
459
.

40

-

7.49
A Proposed Method for the Simultaneous Analysis of

HCN and N0„ in Fire Gases

by

R. D. Curtis* and Yoshio Tsuchiya

IRC, NRC , Ottawa, Canada, K1A 0R6

Prepared for presentation before

6th Canada-Japan-USA Trilaterial Cooperative Study on Fire Gas Toxicity

to be held at Factory Mutual Research Corporation, May 1987

Abstract

In fires involving the burning of nitrogen-containing construction

materials and building contents (e.g. , carpets made of nylon or polyacrylonitrile), various toxic inorganic gases are generated. These

gases can be trapped and absorbed in aqueous or alkaline media and the

anions resulting from their dissociation may be studied by Ion

Chromatography (IC). In particular, an ion chromatographic method based on

the method of DuVal, Fritz and Gjerde (1) was studied for the simultaneous

analysis of HCN and N0

X

. The method proved to be satisfactory for the

simultaneous quantification of cyanide, nitrate, and nitrite ions. A linear

range for CN

- analysis was established from 0.5 ppm to 20 ppm in solution.

*presently a graduate student at Dalhousie University, N.S.
INTRODUCTION

During building fires, various quantities of inorganic and organic

gases are produced. Samples of these gases can be trapped and then scrubbed

with an alkaline solution. Dissolved gases in the solution dissociate, at

least partially, to their respective anions. The resulting anion mixture is

then prepared for analysis and subjected to an IC separation. The main

advantage of the IC method is its ability to separate and quantify a variety

of anions whose predecessors are commonly found in fire gases (e.g. , HCN,

HC1 , HBr , HI, N0

X

).

Due to their relatively high toxicity, the analysis of HCN and NO is

important. Many methods have been developed for the analysis of cyanide

ion. A commonly used method is the absorption of HCN gas in strongly

alkaline media and then determination of free cyanide using the specific ion

electrode. Various IC methods also exist (2, 3), however, there has been no effective method for the simultaneous analysis of HCN and other fire gases

such as NO . The problem has been that cyanide chromatographs very poorly

at low pH (less than 8) and interferes with other common anions at higher pH

(8 to 12). Tsuchiya (4) has reported a simultaneous method using both a conductivity detector and an amperometric detector for monitoring the IC

effluents. The high sensitivity of the amperometric detector to cyanide (5)

has made its application advantageous. However, the pH used was 4.5 which

is generally considered unacceptable for the analysis of cyanide. At this

PH the cyanide is present mostly as HCN molecules with little cyanide ions.

The chromatographic separation of cyanide is poor at low pH. In fact it

passes through the column with little or no retention and elutes at or near

the solvent front. Thus the method is susceptible to interference from

other nonretained species which are easily oxidized at the electrode

surface. In particular, sulfide (another weak acid present as an HS~/S - /HzS

mixture) may present severe interferences.

The

cyanogen

reaction

proposed method involves the conversion of hydrogen cyanide to

iodide and hydroiodic acid using an ethanolic iodine solution,

which, has been known for years (6) is shown below (Eq. 1).

The

HCN + 1 2

; r ICN + I + H

+ K

gq

= .7 3 ( 1

)

The reaction, as described by Beran and Bruckenstein (7), is pH dependent

and quantitative over the pH range 2 to 7. Thus the alkaline solution

containing free cyanide must be adjusted to pH using a noninterfering buffer

just prior to the reaction with iodine. As well, for the reaction to be

quantitative, an excess of iodine must be added. This excess iodine must be

removed prior to the IC analysis. Assuming a quantitative reaction, the

amount of free iodide generated should be proportional to the original

amount of cyanide in the sample. Thus cyanide is determined indirectly by

chromatographical ly separating and quantifying iodide ion.

EXPERIMENTAL

Chemicals

All chemicals were of reagent grade and used as supplied from the

commercial source. Water for all aqueous solutions was prepared using a

462
Corning Mega-Pure System* which was composed of an ion exchange column,

adsorbent column and borosilicate glass still. It is essential to have ion

free wat-er to prevent erroneous results from being obtained on the ion

chromatograms.

Calibration and testing solutions were prepared fresh daily from sodium

cyanide and sodium hydroxide (pH = prepared and solutions for analysis

prior to the analysis.

A 0.05 N sodium acetate buffer

acetic acid and concentrated sodium

1) . A 1000 ppm NaCN stock solution was were prepared from this immediately

(pH = 5.0) was prepared from glacial

hyd roxide

.

The iodine solution was prepared daily by dissolving sufficient solid

iodine in 95% ethanol to make a 0.1 N solution. To prevent the

decomposition of the solution, the container was wrapped in tin foil and

stored in the dark.

A column containing the Amberlite XAI>-4 resin was made in a 12 cm glass

tube fitted with a glass frit. The resin is a polymeric adsorbent which can

be used to remove hydrophobic solutes from relatively polar solvents. It is

a crosslinked polystyrene type polymer with a very high surface area and

excellent physical durability. Samples were injected on to the top of the

column and allowed to elute by gravity so that sufficient time to absorb the excess iodine was possible.

A 0.05 N nitric acid solution was prepared in acetone for the purpose

of removing the absorbed iodine from the XAD-4 resin.

Standard solutions of other common anions (such as Cl , NO 2 , S04

-

) were prepared by dissolving their corresponding sodium salts

NO 3 , and

in

water.

Apparatus

Ion chromatographic separation was carried out on a 4.1 mm i.d. , 250 mm

long ion exchange column (Hamilton, high resolution PRP-X100 column ( 8 )).

The packing was a rigid macroporous po ly ( s tyrene-divinylbenzene ) co-polymer,

chemically bound with trimethyl ammonium groups to provide exchange

capabilities for anion IC. The eluent used was o-phthalic acid (aqueous , 2.5 mmol) buffered to pH 6.0 with sodium hydroxide. The eluent was delivered to the column at a flow rate of 1.50 mL/min using a Spectra

Physics SP8700 pump. Samples were introduced using a Rheodyne 7010 valve

and 7001 Auto injector fitted with a 200 uL loop. The separated anions

leaving the column were detected using a Vydac 6000CD conductivity detector.

A Hewlett Packard Reporting Integrator Model 3390A was connected to the

detector and was used to record the ion chromatographs.

*Some commercial equipment is identified in this paper to adequately specify

the experimental procedure. Such identification does not imply

recommendation or endorsement by the National Research Council, nor does it

imply that the equipment identified is the best for the purpose.

463
PROCEDUkE

Liquid Samples

Samples for analysis should be prepared just prior to the analysis from

the stock solution of sodium cyanide. This prevents the loss of cyanide by

hydrolysis of cyanide ion to hydrogen cyanide.

Pipette accurately 40 mL of the cyanide sample (in pH 11 NaOH) into a 50 mL volumetric flask. To the flask add 1 mL of the 0.05 N sodium acetate

buffer and shake well. Using a microsyringe add 0.6 mL* of the ethanolic

iodine solution to the flask and dilute to volume with water. Shake well

and then allow the solution to stand for at least 10 to 15 minutes. From

the 50 mL flask take a 10 mL aliquot and pass it through the XAD-4

precolumn. This process not only removes the excess iodine but also removes

the water soluble organic gases. The effluent from the precolumn containing

the anions of N02~, NO’j -

, I - and others is then chromatographically

separated for quantification. A calibration plot was prepared by analyzing

several standards and plotting the area of the respective peak as a function

of the cyanide concentration.

In between samples the excess iodine was removed from the precolumn

using 10 mL of the nitric acid acetone mixture. Sufficient quantities

(50 mL) of water were passed through the column to ensure that all the

nitrate ions are removed from the packing. This is an important step since

nitrate is one of the ions that is of prime interest in the analysis of fire

gases by IC.

Fire Gas Samples

A 5 L glass volumetric flask was evacuated. For obtaining a fire gas

sample a piece of Orion cloth (a polyacrylonitrile based fabric) was burned

under controlled conditions in an oxygen index apparatus. The fire gases

evolved on combustion of the material were collected in the flask by

releasing the vacuum and drawing them into the flask. To the flask inject

with a syringe through a rubber fitting 100 mL of 0.001 N sodium hydroxide.

Shake the flask well to absorb all the fire gases in the alkaline media.

Release the vacuum and then shake the flask again. Let the flask sit for at

least 10 minutes and then pour the solution from the flask into a clean

beaker. Pipette 40 mL of the sample solution into a 50 mL volumetric flask

and process and analyze in the same fashion as the calibration solutions.

From the calibration plot determine the level of cyanide in the sample and

(correcting for dilutions) the amount of HCN in the original fire gas.

If after addition of iodine to the sample a yellow colour is not

observed, the sample may require a dilution to fit the linear range of

quantitative analysis (see discussion).

*0.6 mL of the iodine solution is enough to give a two-fold excess at a cyanide concentration of 20 ppm. For samples less than 10 ppm CN , 0.3 mL

of the iodine solution is sufficient for at least a two-fold excess. The

amount of iodine added also affects the sensitivity of the method, as discussed later.

464
RESULTS AND DISCUSSION

The conductometric detection of iodide ion is much more sensitive than

that of cyanide ion. Iodide chromatographs well under the conditions used

and has a retention time of about 14 minutes. A linear range for iodide

quantitation was established from 1 ppm to greater than 100 ppm. A possible

expansion of the method would be to use an amperometric detector for low

level cyanide solutions. This type of detector would broadly increase the

dynamic range for analysis as it is sensitive to iodide at the sub-ppb

level.

Iodide appears to chromatograph well in the presence of other common

anions such as nitrite, nitrate and sulphate. A chromatogram of a fire gas

sample from burning Orion cloth is shown in Fig. 1. The retention times were for sulphate about 8 minutes; for nitrate 5.5 minutes; for nitrite 3.8

minutes and for chloride 3.1 minutes. Note that retention times vary

depending on the IC conditions and also vary slightly with a change in

sample concentration. The initial peak, which is always present in single

column IC, results from the difference in conductances of the eluent and the

injected sample. The large negative dip after the initial peak results from

the acetate ions introduced just prior to the reaction of cyanide with

iodine. Initially this negative peak caused problems in the simultaneous

analysis of nitrite, nitrate and cyanide. That is, the peak corresponding

to the nitrite ions was too early in the chromatograph and thus not observed

(Compare with the small peak for Cl

- in Fig. 1. This level of Cl

- is about

15 ppm and the area response is not representative of the correct amount of

Cl

- (see Fig. 2 at about 4 mins)). By lowering the eluent concentration and

decreasing the pH, the retention times of all anions increase. For example,

a chromatogram of Cl

-

, NO2

-

, NO3

- and SO4

- was obtained using a 1 mmol

o-phthalic acid (pH 6.0), as shown in Fig. 2. On this chromatogram chloride

elutes at greater than 4 minutes, which should be sufficient to avoid the

negative acetate peak. Further experimentation is required to establish the

best conditions for the analysis. Note that the eluent concentration should

not be lowered too much since the dynamic range for that analysis would be

decreased substantially (i.e. , the maximum concentration that can be

quantitated is Towered). The pH should be kept as close to 6 as possible.

According to DuVal et al (1), the iodide peak height remains relatively

constant between pH 4 to pH 6. At pH values greater than 6 the peak height

and thus area increases rapidly. As well in alkaline media the potential

exists for iodine to oxidize other compounds and produce further errors in

iodide quantitation. At lower PH nitrite can react with the released

nitrite according to the reaction shown below (Eq. 2):

21" + 2N02 + 4H" 2H20 + 2N0 + I2 ( 2 )

Despite the possible interferences, the present experimental procedure is

suitable for the simultaneous analysis of N0X and HCN in fire gases.

The reaction between cyanide and iodine appeared to be quantitative

over the range of 1 ppm to 20 ppm in solution. An experimental calibration

curve is shown in Fig. 3. It appears from the calibration curve that the

background level of iodide ion is quite high (about 15000 area units). When

only 0.3 mL of iodine was added to the reacting solution the background

level was much lower (about 1200 area units). The sensitivity was much

better under these conditions, but the linear range decreased to a maximum

465
of just greater than 10 ppm cyanide in solution (see Fig. 4). The

calibration curve tailed off at 12 to 15 ppm for the latter case and at

about 25 ppm for the former case.

In fire gases the gas phase concentration of HCN is usually of the

order of 10 ppm to 300 ppm. With the present experimental procedure the

resulting solution of the dissolved gases would contain cyanide in solution

in the range of 0.6 ppm to 19 ppm. This range of concentrations fits the

linear dynamic range of the method quite well. Another advantage of the

method is that the nitrite and nitrate ions do not present any observable

interferences. Both N02~ and NO3 - can be quantified over the 1 ppm to

100 ppm range while simultaneously cyanide is quantified as iodide.

CONCLUSION

Ion chromatography is not a continuous monitoring method for fire gases

but it has the distinct advantage of being able to quantify a variety of

inorganic fire gases simultaneously. In particular, a method suitable for

the simultaneous analysis of HCN and N0

X was developed. Future research

about IC conditions and possible interferences should help to further

strengthen the method.

ACKNOWLEDGEMENTS

One of the authors (R. Curtis) extends his thanks to the IRC for giving

him the opportunity to pursue this research project. The authors would like

to thank Brian Stewart and Mike Clermont for the help along the way.

REFERENCES

1. DuVal, D.L. , Fritz, J.S., and Gjerde, D.T., Indirect Determination of

Cyanide by Single Column Ion Chromatography, Analytical Chemistry, 54 ,

' 830, 1982.

2. Bond, A.M. , Heritage, I.D., and Wallace, G.G., Simultaneous

Determination of Free Sulfide and Cyanide by Ion Chromatography with

Electrochemical Detection, Analytical Chemistry, _54_, 582, 1982.

3. Okada, T. , and Kuwamoto, T. , Potassium Hydroxide Eluent for

Nonsuppressed Anion Chromatography of Cyanide, Sulfide, Arsenite and

Other Weak Acids, Analytical Chemistry, 5_7_> 829, 1985.

4. Tsuchiya, Y. , Ion Chromatography for the Analysis of Fire Gases,

Prepared for submission to the 4th expert meeting of Canada-Japan-US

A

Trilateral Cooperative Study on Fire Gas Toxicity, Division of Building

Research, National Research Council of Canada, Montreal Road Campus,

Ottawa, Ontario, April 1985.

5. Koch, W.F., The Determination of Trace Levels of Cyanide by Ion

Chromatography with Electrochemical Detection, Journal of Research of

the National Bureau of Standards, 88, No. 3, May-June 1983.

6. Treadwall, F.P. , and Hall, W.T., Analytical Chemistry , Volume 1, 6th

ed. , John Wiley and Sons Inc. , p. 330-337, 1930.

7. Beran, P. , and Bruckenstein , S. , Pneumatoamperometric Determination of

Nanogram Amounts of Cyanide, Analytical Chemistry, _52_, 1183, 1980.

8. Lee, D.P. , A New Anion Exchange Phase for Ion Chromatography, Journal of

Chromatographic Science, 22, 327, August 1984.

466
COCN

time, min

FIGURE 1 ION CHROMATOGRAM OF FIRE GASES

RESULTING FROM 8URNING OF ORION CLOTH

TIME, min

FIGURE 2 ION CHROMATOGRAM USING A DILUTED

ELUENT - 1 mmol 0- PH THA L I C ACID (PH 6.0)

467
FIGURE 3 STANDARD CALIBRATION CURVE

FROM 1 TO 20 ppm CYANIDE (r - 0.986)

x 10

o

Q_oo

Q£O>—

<cq;o

FIGURE 4 STANDARD CALIBRATION CURVE

US ING ONLY 0. 3 ml OF THE I0D INE • S OLUTI ON ( r - 0. 986)

468
A PROPOSED METHOD FOR THE SIMULTANEOUS ANALYSIS OF HCN AND NO:< IN FIRE GASES

R. Curtis and Y. Tsuchiya, National Research Council of Canada

GANN:. What is the cycle time for running samples?

TSUCHIYA: Running time is about this, but before that, there is treatment of

the samples each time. In the first stage we collected for combustion gas in

glass flasks. We found that NOX goes through the wall of the bags, so we now

use glass bottles for collecting NOX gas, mixed with oxygen and oxidized to

N03

. N0 ? doesn't dissolve in water. We have to combat N02 through N0 3. Then

we dissolve the gas in an aqueous solution and filter that solution to remove organic material. It takes about 2 hours.

GANN: Including the iodine treatment?

TSUCHIYA: I don't know the time for the iodine treatment.

469
EVALUATION OF COMBUSTION TOXICITY FROM

VARIOUS MATERIALS USING A NEW TESTING APPARATUS

by

Shyuitsu YUSA

Building- Research Institute, Ministry of Constraction

1 Tatehara, Oho-Machi, Tsukuba-Gun

Ibaraki Prefecture, JAPAN

(For Ninth U.S. -Japan Natural Resources Panel Meeting

On Fire Research And Safety, Factory Mutual Research Corporation

Norwood, Massachusetts 02062 U.S. A.)

May 1987

470
EVAUATION OF COMBUSTION TOXICITY FROM VARIOUS MATERIALS

USING A NEW TESTING APPARATUS

by

Shyuitsu YUSA

Building Research Institute, Ministry of Construction

1 Tatehara, Oho-Machi, Tsukuba-Gun

Ibaraki Prefecture, JAPAN

ABSTRACT

The Building Research Institute is developing an apparatus for toxicity

test of building materials. The toxicity of combustion products of typical

materials are evaluated by animal, exposure tests using the new apparatus. The

weights of the materials when generating combustion products sufficient to cause incapacitation of mice under given exposure conditions (steady - state) are obtained and combined with the time to incapacitation. It is shown that

the relative degrees of toxicities of combustion products of materials may be

expressed by those values (ECT).

1. INTRODUCTION

The new apparatus for testing combustion toxicity was developed at Building

Research Institute, Ministry of Constraction in Japan around 1983. It has been

developed in the research program of the United States - Canada - Japan

Trilateral Research Group on Toxicity of Combustion Products from Building

Materials and Iuterior Goods (from 1982 FY to 1986 FY). The apparatus is

designed to supply air or N

2 mixture in furnace, and to simulate the condition

to which the materials are exposed in real fire.

2. METHODS

The materials used for testing were those indicated Table 1. They were made

into specimens 10cm in diameter, dried in an oven at temperatures of 45 to 55

471
°C for at least 24 hours, then left in a desiccator containing silica gel for more than 24 hours, upon which they were testes.

The testing apparatus comprises combustion device and an exposure chamber

as shown in Figure 1. The cone guartz combustion tube is held inside a cone radiative electric furnace. The mass loss of the samples during the tests is

measured continuously by the load cell. From the bottom of quartz tube and

around the sample, air is supplied to continue the combustion. At the top the cone quartz tube, the gasket is set up to make the time to introduce the

products arbitrary so that a selected part of combustion could be introduced

to the exposure chamber. The exposure chamber is designed to produce a square wave (Steady - State) exposure condition to the animals.

Specifications of the apparatus are described in the following:

(1)

The combustion system

:heating system electric radiation heater (cone furnace

of ISO type

)

: furnace cone shaped quartz tube (40 mm in inner

diameter at the upper part, 164 mm in

inner diameter at the lower part, 3 mm

thick)

:air supply system mechanical supply (air, air + ^

mixture), air temperature control system

:test sample commercial products, horizontal

orientation

(2)

The exposure system

: exposure method static, whole body exposure, 15-30

minutes of exposure, square wave exposure

: animal species : mouse (dd strain weight 20 +2g)

number : 6

( 3

)

Measurements

:mass loss continuous measurement during the test

: chamber atmosphere — continuous measurement of the concentrations of C^, CO

2

, C0» and temperature

: concentrations of HCN and HC1 at a fixed time interval

: effect to animals time to incapacitation, etc.
The combustion products generated from flaming combustion or non flaming

combustion are introduced into the exposure chamber for the several duration

to make various concentrations of products. The animal cover is pulled up following the constant concentration of the products, and the exposure to 6 animals in revolving cages starts. The conditions of mine causing the cages to

revolve are recorded as electrical pulses by proximity switches. The incapacitation time to incapacitation is defined as the length of time from

start of exposure until the point that a mouse could no longer make its cage

revolve. External radiation levels were mainly 2.5 w/cm , and partially 3.5

and 4.5 w/cra“. The air supply rates are selected by which the flaming mode is

sustained in the minimum rates. However the non flaming mode appeared

partially.

The gas concentration inside the exposure chamber is measured by a magnetic

type analyzer for C^, and by an infrared ray analyzer for C07 and CO. To

determine the concentration of hydrogen cyanide in exposure chamber 100 ml of

sample gas was collected every two minutes into a syringe barrel and absorbed

in an 0.1 N - NaOH solution, and then analyzed by lOn cneomatography . To

determine the concentration of hydrogen chloride in combustion gases 200 ml of

sample gas was collected every two or three minutes into a syringe barrel with

approximately 20 ml of alkaline absorption (0.1N NaOH) and analyzed by an ion

chromatography. The CO, CO

2

, and 0 o data and temperature measurements at the

animal exposure position were collected by an on - line computer every 6 seconds as were the weight of the sample. The COHb value in the blood of mouse

is determined by the IL282 CO-OXIMETER, although not for all the case.

3. Results and Considerations

Though the combustion of tested materials was almost near steady state,

there were some materials which had both flaming and non flaming mode in one test run. These were MRIS, Acrylic, and Modacrylic. In the case of PVC, there was some scatter in formation of flame. The test results of the various

materials are indicated in Table 2.

The exposure method used in this study is, as mentioned above, intended the

square wave exposure. Howevere, although the steady state was obtained in the case of O

2

, CO

2

, and CO, the difficulty in maintaining. HC1 concentraiton con473
stant was shown at the last UJNR meeting-^. In the case of HCN, Fig. 2 shows

that the concentration falls by degrees but not so much as is the case of HC1.

With regard to the test results of various materials, the relationship between mass concentrations (mass loss/chamber volume ;mg/l) and average times to

incapacitation of mice, and also average concentraitons of gases indicated in

Figs 3 through 9. In the Fig. 3 the result when oxygen partial pressure is

varied is also indicated. It might be learned that low oxygen, concentration

increased CO generation from the material and resulted in shortening the times

to incapacitaiton of mice. However, the difference between the two is not so much that there needs more detailed study for clear conclusion.

Wood based materials such as Lauan, INSLBD, and PATCBD combusted relatively

stabilized so that correlation between the mass concentration and gas con- centration is good. And it can explain the the variation of the time to

incapacitation. Compared with those materials, synthetic materials and wool

show some scatter in combustion. Among these, FPU shows relatively clear

results as indicated in Fig. 6. With regard to the PTFE, it did not decompose

at the 2.7 w/cm“ heating while decomposed at 3.3 w/cm" (NF) and 4.8 w/cm^ (F).

In any case there need further consideration on combustion.

In the combustion toxicity field the LC^q is commonly used and LCT^q

( mg/1 'min) is used to compare materials. In this study, using the concept, the

ECT which is obtained by multiplying mass concentration by time to incapacitation is shown in Table 2 and indicated in Fig. 10. As a matter of course, it

may not be possible to compare with LCT directly. However, it is thought the

method in this present study, although approximate, is fairly effective in

comparison of materials.

References

1) S. Yusa, Toxicity testing of fire effluents. Proceedings of the 8th UJNR

Panel on Fire Research and Safety, 13-21 Hay, 1985

474
Table 1. Materials Tested

Material Form Thickness

(mm)

Density

(g/cnr1

)

Lauan Board 9.3 0.50 0.53

Insulation board Board 9.0 0.26

Particle board Board 12.5 0.61 0.68

Melamine resin-impregnated Rigid sheet 1.5 1.12

sheet

Acrylic Carpet 10

Polyamide ( Nylon

)

Carpet 10

Wool Felt 10 0.20

Flexible Polyurethane Flexible foam 10 0.02

Polyvinyl chloride Board 3 1.38

Modacrylic Fiber

PTFE Flexible sheet

475
Comb.

Time

Mass

Ave.

concent

of

gase's

CO

Ave.

time

to

ECT

Material

Air

mode

to

Ig

concent

CO

HCN

HC1

/

incapacitation

COHb

Mortality

(1/min)

(sec)

(mg/1)

(%)

(ppm)

(ppm)

CO^

(min)

(mg/l*min)

(%)

VO vo VO vo VO VO vO vO vO VO vO vO VO VO vQ VO VO VO vO VQ vO VO VO VO vO vO \ \ \ \ \ \ \ \ \ \ \ \ \ \ N \ \ \ \ \ \ \ \ \ \ \ ro VO VO rH VO vO ro vO in 04 vO ro vO VO vO VO vO VC

1

VO iD O VO VQ vO VO

vO 03 ro O o 04 o 04 O o o

04 VO o VO rH vO o o 33 03 03

i i 1 1 1 1 i 1 I 1 1 1 1 1 i • • • • • « ro 03 'T O m n O' 04 TT

'33 O' VO O' in O' n vQ VO vO vQ

r* rH o* ro O' a 'O* O o vO rH CD O o O rr VQ VO ro O' O' o or ro n 03 ro 03 vo in LD 04 rH o o in 04 rH in CD ro VO 04 o* o VO 03 03 04 <

1

a> 04 rH O' T 03 04 rH o O' vO rr rr rH rH ro rH rH n CD ro VO rH 04

r* cn IT) O 04 O' o 33 o r' 03 04 ro 04 03 04 VO n ro o o CD ro 04 ro LD 04 04 in ro ro in O' ro 04 ro ro rH rH i—

1

rH vD rH 04 rH rH rH

03 04 O'! vO ro VO 04 rH 'T o rr O' rH rH O' O' ro in O' vO ro T3* ro n CD

04 in rH ro O' <n O' o in CD in ro o TT rH rH CD rH 33 CD VO ro 04 O'

rH o rH TT o o rr o rH ro rH vO o o o o o O ro O' 04 o rH o ro o

-H -H +i -H +1 -H -H -H -h -H -H +1 +1 -H -H -H -H -H +i +1 -H -H +) -H -H -H ro 04 vO ro o 03 O rH rH ro CO O' CD ro in 03 33 O TT 03 ro 04 n

VO o 03 O O' TT IT) 03 in vO vO CD O' rH 03 rH o 03 o CD O' rr O' rH 04 rH

04 O' rH Tp 04 ro in 04 VO in 04 ro 04 rH rH 04 r*' O O' 04 n 04 rH

rH rH rH rH rH rH o O' 03 ro vO LD 03 O' 03 CD CO o T in 03 o O' vO O' CD vO ro n

r—1 04 rH rH ro 04 04 ro ro ro O o 04 04 04 ro 04 ro 04 04 04 04 04 04 n ID o o o o o o o o o o O o o o o o o o o o o o o rH o o

1 1 1 1 1 i i 1 1 1 ! 1 1 i i 1 i 1 I i 1

rH

1

n

1

n

i

03

i

vo

1

n

i i i i 1 1 i 1 1 1 1 I 1 i 1 1 1 1 i i -t 03 o n ro

rH rH rH rH

04 TT 03 LD vO TT O' o in 04 in TT in vO rH rH rr vO in n ro rH ID

o* 04 VO TT 04 vO rr o vO T ro 33 03 o 04 o in ro 04 04 O' 03 ro n o rH O o rH o o rH o o o o O o rH rH rH o o o o o o o o o

o VO VO o o o o vO vO 04 o o o 04 04 vO VO o o o o o

03 04 O' 04 CO VO 04 CD o 03 03 04 vO CD r' 03 O' O' O' o 04 n 03

rH 04 vO LD 04 uo vO vO vo ID 33 04 rH in CO rH 03 in ro n o o 04 ro

03 O' LD rH vO T 03 in rr o CO O' o rH o CD ’T rH rH ro vD n vQ

rH rH rH rH rH rH

in in n LD o O o n o o o in in in o o o o in o O o n n

ro ro ro ro ro rH rH rH rH rH rH rH i in vO in in n n vO i

Cn

Cu ti Cn Cn Cn Uj Cu Cn Cn Ou Cn Cn Cn Cn Cn Cn Cn Cn Cn Cn Cn Cn Cn + Cij Cn

Ci

4

2

2

04 + 2

o O VO O O O O • « • • • • • 03 VO LD vO o 33 O vD

rH rH O

in rr 03 ro 04 03 O in 04 04 n O' O* 03 03

« vO in VO in in VO VO O' in O' O' n 04 n CD

04 • • « • • • • 04 04 04 04 04 04 04 04 04 04 04 04 04 04 04

Q

c 33 a

(TJ 2 CD cn a 03 u M

(0 2 6- CC >4 H CL 2

476
Comb.

Time

Mass

Ave.

concent

of

gases

CO

Ave

.

time

to

ECT

Material

q'

Air

mode

to

Ig

concent

CO

HCN

HC1

/

incapacitation

COHb

Mortality

(

1

/min )

(sec)

(mg/1)

(%)

(ppm)

(ppm)

C02

(min)

(mg/1*

min)

%

co c

0

co CO CD CD CD CD CD CD CD CD KD CD CD CD CD cD cD CD CD CD CD N \ \ \ \ \ s \ \ \ \ \ \ \ \ \ N. \ \ \ \ \

co CD <o VO CD rH rH CD CM CD CD o CD cD CD cD CD CD 04 *cD cO CD o O o O

o o o O o o o o O o O O o O O rH o O' a

cD CD CM LO CD LO CM ro o in in rH rH rr • • • • • • 1

1 1 1 I O' rH o o

O' IT) 03 O' 03 o* CD o ro CM O' ID rH

rH CM CM CM P0 CM rH rr CD CD

co m rr O' o o o o o o o o o o O o O O O O o o o o

rr CM 03 LD CO rH ro ro LO 03 o r- cD o CM (—

1

ro CD 1 1 03 n O' Hp

LT) rH IT) CM m ID O' CD o- ro CD O' CD CM in nr in in ro O' o 03 O' ro CM CM CM o o LO Tp LD CD CD CM C^ o in 03 rH o o* nr CD rH rH CM

rH rH rH CM rH rH in ro n

O' CD i—

1

03 LT) TT CM in CD 03 03 rr CD 03 o 03 CD CD C3 CM rr rH

rH CM rH CM 03 o IT) o ro O' CD o o CD G3 in O' m ro O' 03 CM CM rH

<-H o o o o CM o o rH o rH rH rH rH o ro rH rH O' rH o o o

H +i -H -H -H -H -H +i +i -H -H -H +1 +i -H -H -H -H +i -H +i +i +) +i LD CM IT) HT CD 03 P0 ro 03 o LO 03 CD O' rr CD in CM CM or rH CM

• TT LT) CD m O' O' o CM o o CD CD 03 ro CD 03 CM CM CD CM CM CM

PO rH rH rH CM ro rH CM CD CM LT) CM rH •o* m CM O in n n ro rH rH rH rH rH rH

ro Q LD CD O' ro 03 03 03 CD CD o r- rH r~ CM o> n rH cD CD CO ro

CM CM CM CM CM CM rH rH rH - O O rH o CM ro ro CM CM ro HT O O o CM o o o o o o o o o O O o o o o o o o o o O O o o

o o o o

o n n o

1 ! Ill 1 III CD o CM CO

* cn CO CM in

h (N h mrr-rcom

m rr m rr

cn ro ro cm pi h cn cn h

(N in TT id rH o rp

rH CO O O <0 rH O'

ro rH ro ro rH rH

O' ( cm to

rH G*) CD CM

rH ro ro m

cD O' 03 CO CD 03 03 CD ro cn HT O' rH o rH o rH n ro CD ro ro ro CM O O O o CM rH rH rH rH o o o rH CM HT cD ro ro ro rH o o o o

O O o o o o o o o o o o o o o O o o o o o o o o

o o o o o o o o o o o o o cD o CM o O O o CM T o o

rr cD c CM o CD CD 03 CD CM n O O' o 03 o O CM o O' o o

o HT o CM n O' CD ro ro cD ro ro CD ro cD n 03 cD rH CD CD n cD o

rH rH rH rH T CM CM CM CM ro rH rH n rH 04 ro o 03 O' ro rH rH rH CM

o o o o o o CD n n o o O' n o o o o n o n n o o rH rH rH rH CD CD n O' O' rH 03 CD 03 O' CM T N* i

Cu Cu

Cl, 2 2

2 Cl, Cu CU Cl, Ch Cr.

j

Cl4 CU Cl, Cl, Cl, Cl, Cl, Cu Cl, Cu Cu Cu Ctj Cu + + Cu 4- Cu Cu 2

CUO o o O o o • • • • • « cD cD o CO n CD

rH rH

CM o CN H? ^p CM O 03 o O' n 03 CM ^P n T or O o O' CN O' cD CD CD O' O' CD n 0- O' cD CD 03 CD CD O' ro HP n CD CD cD n n

CN CN CN CN CN CN CM CN <N CN CN CN CN CM CM CM CM CM CM CM CN CM CM CM O

a >

•H u

rH c u > 0 rH <a u rH 0 2 o 2 o 0 CU > 0 < 2 3 Cu a, 2

477

PTEE

2.74

6.0

-

-

~

~

°*

00

°/

6

4.77

F

30

5.60

0.01

-

24.26

±2.86

135.90

-

0/6

3

84

NF

-

6.67

0.02

-

-

9.72

±2.39

64.80

0.70

5/6
Exposure

Chamber

Made

of

PMMA

r- > 0 u

-H

a

4->

u

03 -u

S-4 54 4J jQ CP 03 c U c n

0) < •I—1 c o o *i-4 c 01 a 0 4J 03 C 0 — cj 0 •H >1

4-) 3 54 r— 1 c c 03 4-1 01 a) ‘-a 0 c 3

43 -u 1-4 •H 0 3 oi 04 —. 3

r; 4-1 40 01 3 a 0 <u 0 1-i 4-1 •i—1 3 o 0 c i-4 4-1 a 03 ra 3

"a aj 03 4-1 .—. C

G

O

aj e 0 0 O ra

a '—i G 54 03 i-4 i-1 CJ 4-1 0) _Q E 03 i-4 01 a 3 01 E E 3 03 01

CL. 44 4-4 ra ra 0 4-1

O

— aj CJ 40

O --1 ra ~i

n 0 CJ 'T 03 < E 4-1 <u — 03 — 01 2 >-o 43 03 c OJ 01 .2. j3 <u 4J 1-1 03 ra 03 4J ra 04 0 r- 33 -Q > cu z

•f“4

C 01 3 1-4 * i—4 03 £

4-1 (U £h •H 0 £-> 03 4-1 44 re 0) Q 5-4 Q G ra 3 44 .—. 3 a> 'V X N G -1—4 4-1 "a 3

S-i a; cn 03 03

—J- 3 'G

X 01

0) 03 0 4J 01 44 ra 44 -H c ra 4J ra G 33 03 X! OJ cc 0 • H 3 2 X 03 — •—

33 43 4-1 z

<—1

(U si 4-1

a -e 0 •r-4 OvJ

—

1-4 ra 0) •—4

C

a 0 c 03 c 2 a c 03 -u) 1-4 CP 4-1 '~0 > —4 —4 4—1 oj a 4—1 1—4 + o '—

0 0 CO * G a <—4 rH

cj ai 4-1 01 ra 1-4 03 o u

—

,—i

CP aj 4-1 4J 43. 03 i—i •*—1 03

ra i-l 01 4—1 01 0 cn 4—i oa 44 < 0) CJ

32 ra ra <u 33 33 03 4-1 i—-4 5 CP ra 03 03 E 03 o - a 03

44 -H 01 43 0 3 <a 3 03

u

—

a

cr 4-1 XI

X

U 0 -"4 0 r—i -H ra 0 < < 0 O 03 G CJ 03 CJ 03 < CO G

478

Figure 1.

Teat

apparatus
o o o o CO 04 <— Q

(rndd) N3H

CN

Cm

479

Time

(min)
o

03 U0T1B11U30U03

cn s>

1-0

o

o

oO

o

0) X

e

03

co

CO O o cn ,

co

CO

03 2

c

Q

in

cn

( UTUI ) UOpBipedBOUT 00 3UJTX

Q

480

Fig.

J

Result

on

animal

exposure

test

for

lauan.
(%) 00

inO ©

ID

D

•o ,—

|

o

r_ Oi

<De

a

r~Ho>

s-l <yg

rO a

o w ^ <no

tnw

<TJ

S> Q

(utui) uotq.Hq.T:3Ed^3ai oq. auiTj,

O' •H

481
500

(uidd) NDH

o o O O o o o o o OO CV i

482
24.8

min

PTFE(F)

(uidd) NOH

o

o

LD O

o*3

" T

O

n

O

o

CNJ o

o

% ) 03 LDO Q

ID

o

o

o

LD

a)5

r—

I

o

>u

cu A

ra o

cn

cn0

cncn

fC 2

Q

ID o id

(uxui) uot^B^Toedeoux 04- auixj,

Q

<T> •H

kl

483
MOD (

NF

)

( mdd ) N3 H

0 O O O

O

i_n =r CO C\J Q

(UTUI) UOI^E^TOBdBOUI Oq. SUITJ,

484

p-^g

7

Mass

loss/Chamber

volume

(mg/1)
(uidd) MOH

485

Mass

loss/Chamber

volume

(mg/1)
(%) TOH uoi^ei^uaouoo

486
487

Fig. 10
EVALUATION OF COMBUSTION TOXICITY FROM VARIOUS MATERIALS BY NEWLY DEVELOPED

APPARATUS

S. Yusa, Building Research Institute, Japan

TEWARSON : In the charts that you are showing, grams per liter, this is the

concentration?

SUSUKI: Yes.

TEWARSON: Basically you are measuring this potential yield?

SUZUKI: Yes.

TEWARSON: Because you must have a density per grams of CO per gram of the mass loss rate

.

SUZUKI: Yes, it is the yield of the products. Yield was divided by the size

of the room.

TEWARSON: But the interesting part of the data is that as you change the air

flow, you are changing the fire ventilation which, in turn, is changing the

constraint, which, in turn, is effecting the targeted capacitation

possibility

.

SUZUKI: Yes.

TEWARSON: Therefore, this is the first time that the toxicity method is

taking into consideration the importance of fire ventilation.

SUZUKI: I appreciate your comment. Mr. Yusa will be happy to learn of it.

TEWARSON: However, you also may have to do some other plots where you could

put the fire ventilation versus the effect on animals.

SUZUKI: At this moment, we do not have enough data, we only have enough data

to see the tendencies, but I think time will solve these shortcomings. When

we were developing this new apparatus we referenced Mr. Tewarson's apparatus a lot, so I'd like to acknowledge it here.

LEVIN: In your Figure 9, I don't understand why, as the mass increases, the

HCJ? is accurate.

SUZUKI: Regarding the result of the CO and the C02

, we have quite accurate

data. However, regarding HC1 , the data seems varied and not the steady result

we can get here. At this moment, we can say we lack good data for HC1 . We

still have to gather more RC2 data so that we will have more exactness. One more thing, at BRI , we have been measuring HCJ? by certain time intervals.

However, at Factory Mutual, when we visited, they seemed to have the measuring

apparatus constantly checking HCi . Perhaps we should adopt that type of

instrument in Japan too for this kind of continuous experiment.

TSUCHIYA: About this HCi concentration, I suspect that HCi in the sample is

mainly consumed at the maximum point of the concentration curve.

488
SUZUKI: May I ask you a question? What do you think, is it that things are burnt in a short time, or that a certain amount of burning things is taking a long time to burn, and so HCI can be absorbed by the wall or some other

things ?

TSUCHIYA: I don't know the answer to your question. Is your question related

to the pyrolysis reaction of PVC? After the first stage of pyrolysis, HCI

goes up first; then carbon further decompose. You measure these two stages of

reaction. I don't think this is caused by absorption.

489
A REVIEW OF THE FIVE-YEAR COOPERATIVE PROGRAM ON SMOKE TOXICITY

by R. Gann, H. Suzuki, and Y. Tsuchiya

Presentation bv Richard Gann

The five-year period is drawing to a close and we have just finished the 6th

Expert Panel Meeting. In reviewing the accomplishments of this group, I am very encouraged by the amount of work that has been performed. And yet we still are not quite at the end point. Let me briefly summarize what I see as some of the principal accomplishments

.

Over this five-year period, we have apparently reached a consensus that the use of mixed gas data is an excellent way to represent the toxicity of fire

smokes. To this end a sizable amount of new data and analysis of data have

been performed in all countries. We are now well on the way to using this as

a basis for reducing the dependence on animal testing. At the same time, this

data is highly usable as input for fire hazard and risk modeling. The second

observation that I would make is that we are finally beginning to do the

proper experiments at real scale to allow us to determine the relevance of the

various bench-scale toxicity measurement methods. I don't think that the fact

that we are just doing this now is a mistake on our part. Much of the

knowledge that allows us to do relevant testing today has come from the prior

five years of research. The third area I would like to comment on is the

animal end point that we are using. While most of the data that we are taking

is still neither on animal deaths or on very serious impairment of the

animals, we are just starting to get some data on the threshold for abnormal

behavior as a result of smoke inhalation. If we understand the concentrations

of smoke that barely produce effects and, at the other end of the spectrum,

the concentrations that produce animal deaths . we then are in a position to assess realistically the kinds of safety factors that need to be used. I think we need to do far more work on this threshold measurement, such has been

performed by Professor Alarie and reported at the last Expert Meeting.

Those are the major accomplishments I see. I now would like to spend a moment

on where we have not yet been. Five years ago we set a very ambitious goal

for ourselves; that was the development and testing of a single test method

for smoke toxicity. This was, in fact, a quite proper goal. However, it was

far too ambitious for the amount of resources we have been able to develop. I feel meeting this goal is, in fact, possible and that if we don't set

aggressive goals for ourselves, we would not have achieved what we have over

the last five years.

I will conclude by stating that I hope this cooperative effort will continue,

that significant effort will continue to be expended on the multiple gas

approach, that both experimental and analytical efforts will be spent on validating small-scale test methods, and that we will indeed be able to define

the proper criteria for a bench-scale data generator.

490
Presentation by Hiroaki Suzuki

I appreciate the fact that Dr. Gann and Dr. Tsuchiya attended the first

Tripartite Meeting on toxicity and contributed to the progress of the study on toxicity. One of the difficulties in studying the toxicity issue is that we cannot find the definitive fixed criteria for the exposure conditions

.

Therefore, we had to conduct bench- scale tests and then we had to find out the

relevance of these bench-scale tests to large scale. We had to find a correlation between the two experiments. Our next task was to implement the

apparatus design and gather data. Then we had to come up with some sort of a criterion for prediction. So these are the tasks for the scientists. On the

slide, I have listed five major objectives. We have achieved quite a bit on

the objectives. Also, there are certain cases where, even though an excellent

test method was developed, this test method was not necessarily applicable to

the situation. For instance, if we were to use certain measurements for a test method, we might come up with the kind of data that indicates low level

or high level toxicity. However, when you report the results and conditions

in the actual situation, as often is the case, we tend to forget about the

conditions. For instance, the experimental data we gathered under the certain

conditions cannot necessarily predict toxicity of the certain toxicants or predict the toxicity level. As scientists we should be able to evaluate the

toxicity of the materials even if the evaluation of the combustible material

is not necessarily in harmony with commercial bias. As a scientist I feel

that it is important to find the basic conditions and also the basic

principles for explaining toxicity. We have so far experimented with the

toxicity level by using animals such as mice; however, if we are to experiment

with that kind of toxicity on human beings, that seems to be rather cruel.

Yesterday the medical doctor showed slides where the dead, grotesque body,

caused by fire, was really a horrible sight. Therefore, we must prevent that

type of casualty at all costs by finding out the scientific facts. Since we cannot experiment on a human being at the toxicity level, but, after the fact, we can examine the dead body and gather data from that, then we can use that

data for preventing that kind of casualty. Even though we have almost

completed the five-year project, our research activity must continue. In the

future we should continue the trilateral research cooperation. We may not be

able to have this kind of meeting on a regular basis but, with the cooperation

of the UJNR, we should have a Expert Meeting such as this as often as we can. We also need help from the research organizations throughout the world.

491
Presentation by Yoshi Tsuchiva

I would like to review our activities in this Tripartite cooperation on fire

gas toxicity. Incidentally, I am the only person who has witnessed the

development from the beginning. This cooperation officially started in 1982.

In June 1982, there was a director's meeting in Ottawa, At that meeting the

three countries agreed to cooperate in the fire gas toxicity study. Prior to

the Tripartite cooperation there was a bilateral cooperation between Canada

and Japan, starting in 1980. This cooperation was a little bit different from

the present Tripartite cooperation. The cooperation was on a much narrower

subject and there was much closer cooperation. Discussions started with Dr.

Saito of BRI and myself. At that time BRI already had expertise in animal

testing. We had no experience, even today, in animal testing but we had

expertise in analysis of fire gases. We helped the Japanese group by giving

our expertise in chemical analysis and we were helped by the Japanese group

with their knowledge of animal testing. We exchanged only information, but we ran experiments together. I proposed to use the Ohio State University heat

release rate apparatus for the generation of combustion gases. We had one at

that time and BRI built one. We studied the components of combustion gases

by burning various materials in this apparatus at the same conditions of

combustion and in Japan they burned the same materials and determined the

effect on animals. That was the start of bilateral cooperation. Before that,

about 18 years ago, I proposed to run animal tests at our institution. At

that time not many research institutions were running animal tests. I proposed a study of fire gases using animals, but our director did not give us approval for that. The possible reason was that we were not biologists or physiologists, but chemists. The positive reason was that we had some expertise in analysis at that time. After the cooperation started, we had a good reason to develop our analytical expertise. We have developed analytical

facilities since this cooperation started. Now we are one of the best

equipped laboratories as far as fire gas analysis is concerned. In these

cooperative activities we worked in three areas. One is the development of

analytical techniques . . We have GCMS with which we can analyze organic

components of fire gases. We have APCIMS and with it we can analyze fire

gases in real time. Also, we have ICHPLC with which we can analyze inorganic

gases and aldehydes. In regard to analytical information, our function in the

cooperative program is disseminating information to other countries. The

second area is mathematical analysis of fire gas toxicity. In this area we developed mathematical equations in cooperation with members of BRI. For

necessary data to input into this equation, we used animal data built by many

groups, for example, Dr. Saito, Mr. Sakurai of Japan and Dr. Hartzell and many

other groups in the United States. At this moment we can predict

incapacitation time or lethal time in a given toxic gas atmosphere. In this

area, I believe this cooperation has had very fruitful results. Our future area is full-scale tests. This is just my review of the cooperation. In the

future even this official cooperation may terminate, but actual research will

proceed

.

492
PANEL DISCUSSION: A REVIEW OF THE FIVE YEAR COOPERATIVE PROGRAM ON SMOKE

TOXICITY

R. Gann, H. Suzuki, and Y. Tsuchiya

HARTZELL: I would like to go along with those remarks made by Dr. Gann and

Dr. Tsuchiya, in that I feel amazing progress has been made in the last five

years for being able to predict toxicological effects. Certainly, as far as rodents are concerned, and with some additional information, I think we can go

a long way towards predicting those effects for humans . In some ways, it's a shame that so many of our resources have been spent in the last few years on test method development, and not into this type of thing. Unfortunately, in

this country this is still true that many of our resources are being spent, I think, in the wrong direction.

The second comment, in the nature of test methodology, I'd like to say that I am more impressed all the time with the test method under development by Dr.

Yusa. I've been involved for a number of years with the work of ISO TC92 FC3

,

and I must say that I feel that Dr. Yusa's method probably satisfies more of

the criteria being developed in ISO than any of the other test methods

currently in use or being considered. I commend the Japanese and Dr. Yusa.

TEWARSON : Where would we find the data from the Ohio State University?

TSUCHIYA: Unfortunately, these results will not be published. We found that,

in most cases, animals are not necessary. In this study, a very narrow range

was used to determine very accurate information to determine the effect on living animals. In most cases, under normal fire conditions, all of the

animals died or did not die. Therefore, in most cases, after fire conditions,

animals aren't necessary.

TEWARSON: If I requested the data, will I be able to get it?

TSUCHIYA: No, I don't think so; the analysis is not complete.

TANAKA: This is not a question, but I would like to express my greatest

appreciation for the American and Canadian scholars' cooperation in achieving

the results

.

Since the analysis we are all experts involved with the toxicity study, some

of you might have questions about the Tripartite Cooperative in this effort.

To be honest with you, when we initiated this topic, we asked for help from

Canada and the United States. For this cooperative joint project, we received

a rather larger sum of funds from the Japanese Science and Technology Agency.

Because of this large sum of funds, we could study effectively the toxicity of

the materials and could provide the necessary data.

The major benefits obtained from- this project to the Tripartite Cooperative

were that the American and Canadian researchers contributed to Japan, in terms

of philosophy and methodology and data on toxicity. In Japan, as you know,

the Japanese building code specifies the toxicity test method. Therefore, we need to upgrade the existing toxicity test method. In this sense, this

meeting was very meaningful in terms of learning about upgrading the test

methods. Because of the funds received from the fire and technology agencies,

493
we could establish many research institutes, laboratories, facilities, and

what not. In addition to that, we could establish the research group.

Therefore, we want to continue this kind of research activity from now on, and

we also will need your help in the future

.

494
OPEN TECHNICAL SESSION (CONT.)

495
Characterization Of Flame Spread Over PMMA Using Holographic Interferometry

Sample Orientation Effects

Akihiko Ito 1 and Takashi Kashiwagi

Center For Fire Research

National Bureau Of Standard

Gaithersburg, MD 20899

ABSTRACT

Flame spread over a thermally thick slab of PMMA at several angles of

sample orientation from -90° (vertically downward flame spread ) to 9 = +90 ° (vertically upward flame spread) in air was investigated by measuring

temperature dis tributions within the PMMA sample in the vicinity of the

leading edge of the flame front using holographic interferometry. Samples

with widths of 0.32, 0.47, 1.0, and 2.5 cm, a thickness of 2.5 cm and a length of 30 cm were used. The measured net heat flux from the gas phase to

the sample surface at the vaporization front of the sample is about 7 U/cm2 for downward flame spread (9 < 0°), 6.5 W/cm2 at 9 = +10° and 2.8 W/cm2 at 9 = +90°. However, the total net heat transfer rate increases with an increase in the angle of sample orientation, because the characteristic

heating length, defined as the distance from the adiabatic point on the

sample surface to the vaporization point, increases with an increase in the

orientation angle of the sample. The total net heat transfer rate into the

sample from the gas phase is about 56% of the total net heat transfer input

to the sample at 0 = -90°

,

78% at 9 - 0°

,

87% at 9 - +10° , and 99% at fl = +90°. Therefore, heat transfer from the gas phase into the unburnt fuel

ahead of the vaporization point is the dominant heat transfer path for all

angles of orientation. This was clearly demonstrated by the net heat flow

vector patterns within the sample. The conductive heat transfer rate

through the sample decreases with an increase in flame spread rate (increase

in the sampTe orientation angle) due to insufficient time being available

for the slbw thermal wave to travel through the sample.

1 . INTRODUCTION

In the initial stages of building fires, flame spread over the surface

of combustible solid materials such as a floor covering, a wall or interior

furnishings is the key determinant of the rate of the fire growth. The

controlling mechanism of flame spread appears to differ with the surrounding

conditions, such as the ambient oxygen concentration or the gas flow

velocity . (

1

• 2

)

The flame spread rate in air is controlled by the rate of

thermal energy feedback from the flame to the unburned fuel surface . (

3

4

)

This heat feedback rate differs significantly with the direction of flame

spread relative to the direction of the gas flow/ 55 Since upward flame

spread rate over a PMMA surface is one or two orders of magnitude faster

than downward flame spread/ 8

' 85 it is expected that the heat feedback rate

differs significantly with the sample orientation angle. In practical

^uest worker from Oita University, Oita Japan.

Current address: Oita University, 700 Dannoharu, Oita Japan.

496
application, flame spread over various furnishings occurs at many different

orientation angles

.

In order to determine heat feedback rates from the flame to the unburnt

fuel surface, one needs the detailed temperature profiles in the gas phase

and also in the condensed phase. For this reason, Fernandez - Pello e t al

,

( 7

5

Hirano et_al, {3) and Ray e t al ( 9

5

measured the temperature distributions

using thermocouples in the gas and the condensed phases for downward

vertical or horizontal flame spread on PMMA. Saito et al ( 1 0 5 measured the

surface temperature histories for upward flame spread over a PMMA surface.

However, when using thermocouples, it is difficult to measure the

temperature distribution with sufficient spatial resolution within the

critical thin thermal layer in the vicinity of the traveling flame front.

Therefore, heat feedback rates calculated from the measured temperature

distribution using thermocouples do not have high accuracy, particularly at

the sample surface due to the steep temperature gradient at the surface.

The present authors recently applied a holographic interferometry

technique for determining the temperature distribution within a transparent

solid such as PMMA by measuring changes in refractive index of the

material/ 115 A high spatial resolution, 30 pm, in the temperature

distribution measurement is obtained by this technique. Therefore, the heat

flux distribution can be determined very accurately. In our previous

paper/ 125 the temperature distributions in a thick slab of PMMA during

vertically downward flame spread were measured by this technique. In this

study, the same technique was applied to flame spread over a PMMA, sample in

air at several angles of orientation from 9 = -90° (vertically downward

flame spread) to + 90° (vertically upward flame spread) . Trie results of

temperature and net heat flux distributions in the PMMA sample are presented, and the effects of the sample orientation angle on the net heat

feedback rates and the heat transfer process in the sample are discussed in

this paper.

2. EXPERIMENTAL PROCEDURE AND APPARATUS

Since the detailed description of the relationship between temperature

and the fringe number and also the detailed discussion of the technique were described elsewhere/ 115 only a brief description is presented in this

paper. The relationship between the temperature and the fringe number, N, can be expressed as (115

B /I (dn

p /dT)dT + qB( T - T„ ) [ £ (dru/dT)dT + n „

- na<B ] - A(2N - l)/2

CO -**03 where B is the width of the sample. The term dr^/dT is the thermooptic

coefficient of the material, a is the linear thermal expansion coefficient,

is the refractive index of the material at the ambient temperature, na<a is the refractive index of air at the ambient temperature, and A is the

wavelength of the light. The thermooptic coefficient and the linear thermal

expansion coefficient of PMMA have been recorded up to a little over the

glass transition temperature (about 105°C)/ 13) Above the glass transition

temperature, accurate data for these coefficients are not currently

available. Therefore, the surface temperature history measured with a fine

thermocouple was used as the reference temperature to analyze interference

fringes above the glass transition temperature.

497
Since the refractive index change of PMMA due to temperature change is

much larger than those of gases, the fringe number is more sensitive to

temperature for PMMA than for gases. Therefore, a short path length through

the sample is required to be able to resolve each fringe. However, a wide

sample is needed to insure that temperature distributions in the gas phase

and in the condensed phase are two dimensional. This condition is necessary

because the technique used in this study is based on the integral of the

change in refractive index along the sample width. The experimental

configuration was arranged to satisfy the above two conflicting

requirements/ 125 Two thin sheets of quartz with a 0.09 cm thickness were installed at both sides of the PMMA sample with a small clearance. This

clearance is close to a quenching distance, about 0.15-0.17 cm, which

prevents flames passing over both sides of the sample and insures twodimensional flame spread. These sheets of quartz hardly affect the fringe

shift because their thermooptic coefficient and thermal expansion

coefficient are much less than those of PMMA and also their thickness is

much less than that of the sample. A PMMA sample was supported by an aluminum holder to form many different orientation angles with respect to

the horizontal plane as shown in Fig. 1. For the recording of holograms,

the double exposure method (14) was employed for the experiments in which

steady or near steady flame spread was observed. However, the real time

method(14) was used for the experiment with unsteady accelerating flame

spread which was observed in the vertically upward mode. The optical setup

and the holographic procedure were described in our previous paper/ 125

A 25 pm wire diameter chromel-alumel thermocouple was used to measure

the history of the surface temperature of the sample during the flame

approaching period. To assure good contact between the thermocouple and the

PMMA surface, the thermocouple was heated electrically and pressed onto the

surface prior to the test. The surface temperature measurement by the

thermocouple was conducted simultaneously with the interferometric

measurement. Cast slabs of Lucite L (E.I. du Pont de Nemours <S Co.) 2 with

widths of 0.32, 0.47, 1.0 and 2.5 cm, a thickness of 2.5 cm and a length of

30 cm were used as samples in this study. The selection of Lucite L was based on its cleaner burning than previously studied Plexiglas G. (15)

3. RESULTS AND DISCUSSION

3.1 FLAME SPREAD RATE

The obser/ed steady flame spread rate at 9 - -90° (vertically downward)

was 5 . 8x10" 3 cm/s for a sample width, 3, of 0.32 cm, 5. 5x10' 3 cm/s for B = 0.47 cm and 5.0x10' 3 cm/s for B > 1.0 cm. The results on flame spread rate

for the sample 0.47 cm wide, 2.5 cm thick and 30 cm long at various angles

of the orientation are shown in Fig. 2; they are compared with the results

reported in previous studies 1

7

' 8) using different PMMA samples. The

comparison indicates that flame spread rates measured in this study are slightly higher than those previously reported at -90°= 0 < 0° but they

become slightly slower than those reported at J > +10°. Since it was reported that flame spread rate over Plexiglas G was slightly slower than

2 Certain materials are identified in this paper in order to adequately

specify the experimental procedure. Such identification does not imply that

the material is necessarily the best available for the purpose.

498
that over Lucite L at 8 - -90°

,

( 15

1

this slight difference in flame spread

rate shown in Fig. 2 might be due to the difference in material. The

vaporization front was a straight line across the sample surface

perpendicular to the direction of downward flame spread. Therefore, it is

assumed here that the specific experimental configuration shown in Fig.l and

described in the previous section did provide the two-dimensional flame

spread comparable to a large width sample at 9 < 0°

.

For vertically upward flame spread (

9

- +90°'), the initial short flame

is laminar and blue. After the flame travels about 10 cm from the ignition

line, the blue flame changes to a yellow flame with some fluctuation. The

temperature distribution was measured at 20 cm from the ignition line

.

Although the flame is yellow and unstable, it appears to be in the

transition regime and may not be a fully developed turbulent flame. Flame

spread is accelerating and it is difficult to measure its rate accurately.

Although the measured flame spread rate and the temperature distribution

which will be described in the next section are not universal values at 9 = +90° , the observed trends are very clear compared with those at other

sample orientation angles

.

3.2 TEMPERATURE AND HEAT FLUX DISTRIBUTIONS

Typical interferograms are shown in Fig. 3. Since a condensed fuel must

supply evolved combustible products to the gas phase to form a flame, the

vaporization front of the sample surface is defined at x = 0 in this study

where x is the coordinate parallel to the sample surface and it increases

with downstream distance. The visible flame front extends further forward

from the vaporization front. This overhang distance of the flame increases

with an increase in the angle of the sample orientation.

In these interf erograms , it is noted that the fringe pattern sometimes

behaves abnormally. This is caused by the glass transition in the PMMA

sample. The thermooptic coefficient of PMMA decreases with an increase in

temperature (negative value), but the thermal expansion coefficient

increases with an increase in temperature . ( 1 1 5 Near the glass transition

temperature, about 105°C, the term including these coefficients becomes

close to zero. Therefore', the fringe pattern appears to be at a singular

point at the glass transition temperature of PMMA. Except for the glass

transition layer, however, the fringes shown in Fig. 3 correspond to

isotherms

.

Fig. 3(a) shows that the surface temperature at 9 - -90° starts to

increase from about x =* -1 cm, and the adiabatic condition at the surface is

reached at about -0.34 cm , where the fringe is normal to the surface. The

sample surface slightly loses heat to the gas phase at x < -0.34 cm because

the interior temperature of the sample is slightly higher than the surface

temperature. The number of fringes gradually increases until the glass

transition layer and it increases significantly within the distance of 0.2 cm from the glass transition layer to the vaporization front indicating a steep temperature increase. In the distance between the adiabatic point and

the vaporization front, net heat is transferred from the gas phase into the

sample. A "characteristic length for flame spread", 2, can be defined as this distance. This characteristic length increases with the increase in

the angle of the sample orientation. For the vertically upward flame spread

shown in Fig. 3(b), the adiabatic condition is not found in this figure

499
because this point is located well ahead of the vaporization front, about at

x - - 5 cm

.

Surface temperature histories measured by a thermocouple at various

sample orientation angles are shown in Fig. 4. The results show that the

surface temperature at the vaporization front (x- 0) is about 380°C at the

angles of -90°, -30° and 0°. However, this vaporization temperature tends

to decrease with an increase in the angle of the sample orientation for

upward flame spread, for example, about 375° C at 6 - +10° and about 365° C at

8 - +90° . The difference in vaporization temperature between upward and

downward flame spread must be due to the difference in heating rate. For

vertically downward flame spread, the sample is rapidly heated from ambient

temperature to the vaporization temperature within a short time. On the

other hand, since the flame front in upward flame spread extends further

ahead of the vaporization front, the sample is slowly heated over a long

time .

The temperature distributions within the condensed phase were calculated

from the interf erograms using the surface temperatures measured by a thermocouple above the glass transition temperature region. By taking the

spatial derivative of the temperature distributions, conductive heat fluxes

in the direction normal to the surface, q are calculated. The results for

the conductive heat flux at the surface, q yQ , are plotted in Fig. 5. Ahead

of the vaporization front, the meaning of qy0 can be defined from the energy

balance at the surface which can be expressed as

9y o - A

s (3T/3y) y __ 0 - A

g (3T/3y)

y

_ + 0 + Qr ,

- £

s a(Ts

4

- Tffl 4

) at y

- 0 (2)

where A

s and A

g are the thermal conductivities of the sample and gas,

respectively, Q r £ is the radiative flux from flame, £

s is the emissivity of

the sample, a is the Stefan-Boltzman constant, T s is the surface temperature

of the sample and T0 is an ambient temperature. The first term of the right

hand side of Eq.(2) represents the convective/conductive heat flux from the

gas phase. The third term represents the radiative heat loss from the

surface to surroundings. Therefore, qy

0

is the net heat flux from the gas

phase at the sample surface. The results shown in Fig. 7 indicate that the

value of q y Q at 9 - -90° increases sharply from 0 to about 7 W/cm2 within a short characteristic distance of about 0.34 cm. The value of q yQ at x - 0

for downward flame spread (9 < 0°) does not change significantly with the

sample orientation angle and it remains at around 7 W/cm2. However, it

tends to decrease with an increase in the angle of the sample orientation

for upward flame spread. Especially for vertically upward flame spread

(i9 = +90°), the value of qy0 increases gradually from about 5 cm ahead of

the vaporization front, and its value does not change significantly within a distance of 2 cm from the vaporization front. The value of q y0 reaches to

about 2.8 W/cm2 at the vaporization front. The main reason for this

difference between 2.8 W/cm2 at x - 0 for vertically upward flame spread and

7 W/cm2 for vertically downward flame spread is considered to be accounted

for by the observation that the upward spreading flame is far away from the

sample surface; therefore, the conductive/convective heat flux from the

flame is much smaller than that of the downward spreading flame. Although

the experimental condition is significantly different from this study, a value of 2.5 W/cm2 was reported for a steady wall burning study. (17)

500
Conductive heat fluxes in the direction parallel to the sample surface,

q x , were calculated from the measured temperature distributions similarly to

qy

. The results for heat flux conducted through the sample at x - 0 , Qx 0 , are plotted in Fig. 6. This figure indicates that the value of qx Q decreases significantly with an increase in the sample orientation angle,

for example, from about 3.2 W/cm2 at 9 - -90° to about 0.1 W/cm2 at

9 - +90° . Since the value of qxQ decreases in spite of increasing in flame

spread rate with an increase in 9 , the contribution of conductive heat

transfer through the sample to the overall energy balance for flame spread

decreases with the increase in the sample orientation angle. Further

detailed discussion of the overall energy balance will be given later.

3.3 FIDUE SPREAD MECHANISM

Isotherm contours and heat flux vectors (determined from the values of

qy and q x ) in the sample in the vicinity of the flame front were obtained to

examine the heat flow path in the sample. The typical results are shown in

Fig. 7(a) at 9 - -90° and 7(b) at 9 - +90°. Broken lines in these figures

are isotherms and arrows are heat flux vectors which indicate the magnitude

of the heat flux and the direction of heat flow. For vertically downward

flame spread, heat flows into the sample at relatively steep angles and

large magnitudes especially near the vicinity of the vaporization front.

The magnitude of the heat flux decreases rapidly with the distance away from

the vaporization front. At about 0.4 cm ahead of the vaporization front, the

direction of the heat flow is nearly parallel to the sample surface and

beyond 0.4 cm the heat starts to flow toward the surface from the inside of

the sample. The overall heat flow pattern at 9 - 0° is similar to that at

9 = -90°. On the other hand, for vertically upward flame spread shown in

Fig. 7(b), the direction of the heat flow over the entire figure is almost

normal to the sample surface . This indicates that heat conduction through

the sample in the direction parallel to the surface is negligible for the

vertical upward flame spread mode.

The temperature and heat flux distributions were used to construct an energy balance in the condensed phase . The control volume is the layer from

x - 0 to -® and y

= 0 to -L (L is the depth of the sample). Then, the net

heat transfer rate into the control volume from the gas phase is

Q

g

o o

*

f* 3T

A dx -

s dy

- CO i o

dx

,

(3)

and the amount of heat conducted into the control volume through the plane

of x = 0 is

Q

c

o 3T

A —

s dx

J -L

dy x-o

o

q dy xo

-L

( 4 )

501
The rate of heat loss from the rear sample surface is

Q

L

3T

3y y- -L

dx ( 5 )

Then, the net heat accumulated in the control volume is

Qt = + Qc

* Ql

This must be equal to the net flow of thermal enthalpy

"T

\

Q

= P v dy

s -L

c dT

s x=o

i®

( 6 )

(7)

The following thermal properties were used to calculate the above

integrals; thermal conductivity of the sample A

s - 2.1 x 10' 3 J/cm s K, (18)

density p s = 1.19 g/cm3

, and values of specific heat C

s based on the results

of ref. (19). The calculated values of,Q , Qc

, QT

, and Q for each angle of

the sample orientation are shown in Table 1. The results show that the net

heat transfer rate from the gas phase into the control volume, Q increases

with an increase in the sample orientation angle, but the conductive heat

transfer rate through the sample across the plane of x - 0, Qc , decreases

with an increase in the angle. The net heat accumulated in the control

volume, QT , also increases with an increase in the angle and its change with

the angle is roughly proportional to a change in flame spread rate. The

values of QT are about 3-7% larger than those of the net flow of thermal

enthalpy, Q. This difference is mainly due to heat loss by surface

reradiation from the sides of the sample, which is not counted in the energy

balance. The fraction of Q ? in the total net heat transfer rate, QT

, as described by Eq.(6) is about 56% at# - -90°, 70% at 9 - -30°, 78% at 9 = 0° , 87% at 9 - +10° and 99% at 9 - +90°. Therefore, the heat transfer from the

gas phase into the sample surface ahead of the vaporization front is the

dominant heat transfer path for all angles of orientation of thermally thick

PMMA samples . This does not necessarily mean that heat conduction through

the sample in the direction parallel to the surface is negligible. This is

particularly true for flame spread at from 9 - -90° to 9 -0°.

The isotherm contours and heat flow vectors shown in Figs. 7 indicate

that the high temperature region is located only in the vicinity of the

flame front, i.e., there is a "critical zone" defined by the two length

scales 1 and S in the x and y directions, respectively. The characteristic

length i can be defined as the distance between the vaporization front and

the location where the adiabatic condition is attained at the surface . Then

the characteristic time for the non- vaporizing sample to be exposed to the

heat from the gas phase , r , can be expressed as

r - i / V

The thermal wave penetration depth,

period can be expressed as

5 - J ar

6 , during this characteristic time

502
where a is the thermal diffusivity of the sample material. The values of 2 were determined from the temperature distributions and r can be calculated

from the values of 2 and V. then, using the previously described thermal

properties of the material, the values of 5 were calculated. The results

are listed in Table 2. It indicates that r and 5 do not change

significantly with change in the sample orientation angle, although 1 changes significantly as discussed previously. Therefore, the thermal wave velocity, 5 /r , of the sample material is also independent of the angle. The

thermal wave has two components; one is in the direction normal to the

sample surface, normal heat conduction, and the other is in the direction

parallel to the surface, streamwise heat conduction. It appears that the

normal heat conduction is nearly independent with a change in the sample

orientation angle due to a nearly constant value of r as shown in Table 2. However, since the flame spread velocity becomes much larger than the

thermal wave velocity, S/r, for upward flame spread, the effects of the

streamwise heat conduction become less significant with an increase in the

angle

.

4 . Cone lus ions

(1) The surface temperature of PMMA at the vaporization front is about 380°C

for downward flame spread. However, it tends to become lower with an increase in the sample orientation angle, such as 375°C at 9 = +10° and

365°C at 6 = +90°. This is due to the difference in heating rate.

(2) The net heat flux from the gas phase at the vaporization front is about

7 W/cm2 for the downward flame spread (9 < 0°). It decreases with an increase in the sample orientation angle and reaches about 2.8 W/cm2 for

vertically upward flame spread.

(3) The total net heat transfer rate from the gas phase into the sample

ahead of the vaporization front increases with an increase in the sample

orientation angle and its fraction of the total net heat transfer rate into

the control volume increases from 56% at 9 - -90°, 78% at 9 * 0° to 99% at

9 = +90°. Therefore, the heat transfer from the gas phase into the sample

ahead of the vaporization front is the dominant heat transfer path at all

angles of sample orientation. The role of streamwise heat conduction

through the sample decreases with an increase in flame spread rate due to

insufficient time being available for the slow thermal wave to conduct

sufficient energy through the sample.

(4) The characteristic length, 2, defined as the distance between the

vaporization front and the location where the adiabatic condition is

attained at the surface, increases significantly with an increase in the

sample orientation angle. However, the thermal wave penetration depth, 5,

and characteristic time, r, are nearly independent of the sample orientation

angle

.

5 . REFERENCES

1. Magee, R. S. and McAlevy III, R.F., J. Fire and Flammability. 2, 217

(1971).

2. Ray, S. R. and Glassman, I., Combust. Sci, Tech. 32,33 (1983).

3. Williams, F. A., Sixteenth Symposium (International) on Combustion, The

Combustion Institute, 1281 (1977).

503
4. Fernandez - Pello , A. C. and Hirano , T., Combust. Sci Tech. 32, 1 (1983).

5. I to, A. , " Flame Spread Over Thin Combustible Solid", Ph . D Thesis, Tokyo

Institute of Technology, Tokyo, Japan (1979).

6. Orloff.L., de Ris , J., and Markstein, G.H., Fifteenth Symposium

(International) on Combustion, The Combustion Institute, 183 (1975).

7. Fernandez - Pello , A. C. , and Williams, F. A. , Fifteenth Symposium

(International) on Combustion, The Combustion Institute, 217 (1975).

8. Hirano, T. , Koshida, T. , and Akita, K. , 3ullet. Japan Assoc. Fire Sci.

Eng. 27, 33 (1977) . 9. Ray, S. R. , " Flame Spread over Solid Fuels", Ph . D. Thesis, Princeton

Univ. (1981).

10. Saito, K. , Quintiere, J. and Williams, F. A., Fire Safety ScienceProceedings of the First International Symposium, Hemisphere Publishing Co.,

75 (1986).

11. Ito,A and Kashiwagi , T., " Measurement Technique for Determining the

Temperature Distribution in a Transparent Solid Using Holographic

Interferometry", (in press) Aon lied Optics . 12. Ito, A. and Kashiwagi, T., "Temperature Measurements in PMMA During

Downward Flame Spread in Air Using Holographic Interferometry" , Twenty First

Symposium (International) on Combustion, (in press).

13. Brandrup J., and Immergut, E.H. Ed. " Polymer Handbook”, John Wiley &

Sons, New York, v-56, 1975.

14. Vest.C.M. " Holographic Interferometry ", Wiley, New York, chapt. 6,

1979.

15. Kashiwagi, T. , Inaba, A., and Brown, J. E. , Fire Safety Science - Proceedings of the First International Sympos ium , Hemisphere Pub. Co., 483

(1986) . 16. Waxier, R. M., Horowitz, D., and Feldman, A., Applied Ootics 18, 101

(1979) . 17. Quintiere, J., Harkleroad, M. , and Hasemi, Y. , Combust. Sci, Tech, . 48,

191 (1986)

.

18. Hand, S.D. and Horsfall, F., J. Phvsics E, Scientific Instruments 8, 687

(1975) . 19. Bares, V. and Wunderlich, B .

.

J . Polymer Sci. 11, 861 (1973).

Table 1. Net heat transfer rates and net flow of thermal enthalpy.

9 V (cm/s) Q. (W/cm) Q„ (W/cm) QT (W/cm) Q (W/cm)

g C I

-90° 5.5 x 10' 3 0.52 0.42 0.93 0.89

-30° 6.0 x 10* 3 0.71 0.33 1.02 0.94

0° 7.4 x 10' 3 1.0 0.29 1.3 1.2

+10° 9.5 x 10' 3 1.3 0.22 1.5 1.5

+90° 8.3 x 10' 2 (8.2) (0.13) (8.3)

( ) means not the steady state value.

504
&/r (cm/s)

Table 2. Characteristic lengths and characteristic time

9 V (cm/s) 2 (cm) r (s) 5 (cm) at x-0

-90° 5.5 x 10' 3 0.34 62 0 . 24 3.9 x 10" 3

-30° 6.0 x 10' 3 0.36 60 0.23 3.8 x 10' 3

0° 7.4 x 10' 3 0.44 60 0.23 3.9 x 10' 3

+10° 9.5 x 10' 3 0.48 51 0.21 4.1 x 10' 3

+ 90° 8.3 x 10' 2 5.0 60 0.22 3.7 x 10' 3

Fig.l. Schematic illustration of the

experimental apparatus

.

1 ANGLE OF INCLINATION, 9 decrees

Ve,,lcal, V vemcai. y oo^nwafd upwafc.

Fig. 2. Relation between flame spread

rate and the sample orientation

angl-e

.

505
Direction of flame spread >

(

Surface

0

= - 90 °

Fig . 3 . Typical

various

(a) 9 -

interferogram pictures at

sample orientation angles,

-90° , and (b) 9 - +90°

.

x : err. i

Fig. 4. Surface temperature distributions

ahead of the vaporization front

(x - 0 at the vaporization front)

at various sample orientation angles.

x l rr

'

0 -0 1 -0 2 - 0.3 - 0.4 -5 -6

Fig. 5. Distributions of normal heat flux

at the surface along the distance

ahead of the vaporization front.

506
Flame

Spread

Rate

4

o

cr x

3

<

LUI

y( cm)

Fig. 6. Distributions of streamwise heat

flux at x - 0 plane with the sample

depth

.

(Nel heal llui scale) 2W/cm^

i i 1 1 1 1 r

-0 4 -0 3 - 0.2 -0 1 0

y cm

o C

Q c ® Q 2 w/cm 2 (Net heat Muj scale)

y cm

Fig. 7. Isotherm contours (dashed lines)

and heat flow vectors (arrows) at

various sample orientation angles,

(a) d - -90°, and (b) $ - + 90°.

507
CHARACTERIZATION OF FLAME SPREAD OVER PMMA USING HOLOGRAPHIC INTERFEROMETRY

SAMPLE ORIENTATION EFFECTS

A. Ito and T. Kashiwagi , National Bureau of Standards, USA

JIN: He's asking a question of whether the heat loss from or to sides of the

sample was important or not, or whether the thing was taken only on a twodimensional plane.

KASHIWAGI: There was a barrier (a fused silicon plate) spaced between the

samples so there would be no heat overflow to the side. Although the flame

was against the sample, there was a certain amount of heat loss to the sides.

But they were able to measure the difference of heat loss around the sample

and the barrier. However, it was so insignificant that they did not put it on

the chart. In order to get a more accurate reading, they would have to

install sensors on an order of magnitude much greater than what they did in

order to get an accurate reading, otherwise, it's just an academic exercise in

indicating that there had been some heat loss

.

The barriers which prevent the flame from going to the side were not directly

attached to the samples per se, because by doing that you have a misreading

because there's a heat loss through both the barrier and the sample. It was close but not touching.

QUINTIERE: Do you have any explanation as to why the characteristics of time

is apparently a constant for all those orientations?

KASHIWAGI: In this case, it comes down to applications. If you heat so, a kind of ghostly thermal pressure conduction of the normal side, is about the same because ventilation comes out the same . Then as you go into the

sampling, it's about the same. So, whatever is happening, total energy is

about the same. It could be a very long time for a low spot, or a high spot

in a shorter time.

QUINTIERE: Vaporization temperature is the same for all cases?

KASHIWAGI: Very constant.

QUINTIERE: So, if I interpret this, due to the particular distribution of

heat flux, the distance somehow the integral of that distance --

KASHIWAGI: Right, it doesn't matter how you do it.

QUINTIERE: It's the same?

KASHIWAGI: Right.

508
A NUMERICAL STUDY OF PERIODIC

ALTERNATING FLOW OF FIRES

Kohyu Satoh

Fire Research Institute of Japan

3-14, Nakanara, Mitaka, Tokyo 131

ABSTRACT

It is pointed out that large-scale turbulence can accelerate combustion

due to surface and bulk mechanism. This suggestes that fluctuations of

fires based on turbulence may play a significant role in fire growth. The

author et al. have numerically studied regular oscillatory flows caused

by turbulence. It was found in those studies that the frequencies of

oscillatory flows depend upon the l/3rd power of heating rate, similarly

to experiments. Recently it has been revealed that numerical calculations

can reproduce turbulent flows extremely similar to the actual dynamic

fluid flows (particularly in case of the wake behind a circular cylinder) . So-called computer fluid dynamics are generally being applied in designing

aircrafts. In the future numerical fire simulations corresponding to

a "numerical wind tunnel" may be available in designing fire safety

systems. Currently the author has investigated the dynamic behavior of

both a plume above a heat source and a turbulent jet from a nozzle,

relevant to a "dynamic numerical fire". The objective of the present

report is to discuss the periodic oscillatory behavior of fire flows due

to vortices, based on the numerical results calculated by the author and

by the other investigators. In these results calculated, numerous vortices are periodically created near the stream and cause the periodic

fluctuations of flows, growing up in the downstream. The present twodimensional calculations showed the existence of vortices appearing

alternately on opposite sides of both a plume and a turbulent jet axis.

These periodic fluctuations are essentially the same as the Karman- vortex

street in the wake behind a circular cylinder.

1. INTRODUCTION

A characteristic feature of actual fires is the dynamic motion

accompanying combustion. Oscillatory flows due to turbulence may play a significant role in fire growth. Frank- Kamenetskii 1 an a Williamson et al. 2 pointed out that large-scale turbulence can accelerate comoustion,

that is, the energy release from the flames may be accelerated. Two

possible mechanisms, surface and bulk mechanism, are suggested. Thus the

investigation of turbulent flow of fires should be important to analyse

fire growth. The author et al.

3

* 1- 0 have numerically investigated

oscillatory flows due to turbulence, using a finite difference computer

code UNDSAFE 11. The flows in some cases calculated were very regularly

oscillating. In particular the venting flow from floor to ceiling caused

by a central heat source on the floor in an enclosure was highly periodic.

The oscillatory frequencies of flows depended upon the l/3rd power of

the heating rate. The relatioship was quite similar to the one in

experiments as shown in Fig.l-A. The two-dimensional oscillating flows

depicted by velocity vectors and isotherms are shown in Fig. 1-8, where the

air outflow rate is maximum at t=36.50 sec (t:time after heating) and

minimum at t=37.Q0 to 37.25 sec. The same phase as at t=36.5 sec appears

509
at t=37.75 sec, i.e, the frequency is 0.8 Hz for the heating rate 10 kV.

It has recently been revealed that numerical calculations can reproduce turbulent flows extremely similar to the actual dynamic fluid

flows. In particular the turbulence near the wake of a circular cylinder

or a rectangular block located in a uniform flow has very precisely

been calculated 1 "

’ 1 4

, It is well-known that the Karman vortex street

near the wake behind a block located in a uniform flow shows a periodic

oscillation. Ihe so-called computer fluid dynamics (or "numerical wind

tunnel") is now being applied in designing aircrafts. Hence in the future

a "numerical fire room" corresponding to a "numerical wind tunnel”

may be available in designing fire safety systems. Currently the author

has investigated the dynamic behavior of both a plume above a heat source

and a turbulent jet from a nozzle, relevant to a "dynamic numerical fire".

Time-dependent oscillatory flows due to a localized heat source have

recently been calculated by some other researchers, too. Ihese oscillatory

flows must be based on the same origin. The objective of the present

report is to discuss the periodic oscillatory benavior of fire flows

due to vortices, based on the numerical results calculated by the author

and by the researchers in different fields of fluid mechanics.

2. NUMERICAL RESULTS OF PERIODIC MOTION OF A PLUME

ABOVE A HEAT SOURCE AND A TURBULENT JET

Two oscillatory flows calculated here are first represented, since

they are typical of oscillatory flow benaviors. Calculations are conducted

using the two-dimensional finite difference UNDSAFE computer code which

is based on a micro contro 1-voluae method. Detailed procedures for

calculations are not jhown here, but are described in the previous

reports 1 s 16. The original code has, however, been largely refined.

2.1 Swaying motion of a plume

Forstrom et al. 17 first measured the natural swaying motion of a plume

above a line heat source, although the flow was very slow and laminar.

Noto et al. 13 experimentally investigated the relationship between

swaying' frequency and heating rate, using a fine electrically-heated

wire as a heat source in an enclosure (0.3 x 0.8 x 1.0 m (height) ). The

author has studied the flow behavior by two-dimensional calculations,

since the motion of flows is reported to be in a two-dimensional plane and

to be quite regular. For calculations the flow domain is mainly divided

into about 10000 grids, e.g. nonuniform 103 x 110 and uniform 30 x 100

grid.

A line heater is located at 0.4 m from the left side wall and 0.4 m above the floor. The heating rate varies from 10 to 120 watts/m. A small

disturbance (by dislocating the heat source to an adjacent grid briefly)

is numerically added to a plume to accerelate the commencement of swaying

motion. The plume above the line heat source .however, began to sway

spontaneously , even without the disturbance to the plume. And a longer

time was required for the swaying motion to begin spontaneously in the

calculations. The swaying period was almost constant and the frequency

depends upon the heating rate of a line heater. Fig.2-A shows the relationship between the frequency and the heating rate, from experimental

measurements by Noto et al. and present numerical calculations. The

dynamic motion of a plume calculated is shown in Fig. 2-3, in which the

swaying motion of a plume and vortices along a plume can be seen. The

thermal conditions . at the walls had a quite large effect upon the

510
swaying frequency, that is, cooling of the ceiling reduces the frequency

of swaying motion.

The relationship in the experiments was expressed as Re = 0.33 Ra 1 -^ , where Re is the Reynolds number defined by Re=f-L 2 / v and the Rayleigh

number Ra = L

3 *g* j9 *Q/ X / v

2 (f : frequency, L: reference length, v : kinematic viscosity, g: gravitational acceleration, Q: heating rate, X:

heat conductivity). A close relationship was given by the present

numerical calculations as shown in Fig.2~A.

At the initial stage before swaying motion began, nearly symmetric

pressure distribution was shown at both sides of the plume. The pressure

imbalance at the both corners of enclosure was, however, made

accompanying the swaying motion of plume. A small vortex was alternatingly created at each side of the plume and grows moving upwards

with swaying motion.

2.2 Turbulent free jet

Antonia et al. 13 studied a flag-like flapping motion of a twodimensional turbulent free jet. It is suggested that the motion can be

represented by a two-dimensional von Karman vortex-street pattern.

Shimizu 20 and Hasegawa et al. 21 studied the jet numerically by the so- called discrete-vortex method, but the maximum time after the ejection of

a jet and the maximum distance from a nozzle investigated were quite

limited. The present study investigates the jet motion using a twodimensional finite difference method. The regiems studied by the three

researchers are shown as follows in dimensionless and dimensional

forms

:

Maximum distance from nozzle

Maximum time after ejection

Reynolds number

Reference speed

R eference length

Shimizu 20 Hasegawa et al. 21 30(15 cm) 10 ( 10cm)

160 ( .01s) 60 ( . 06 s)

26700 6667

(30 m/s) (10 m/s)

(0.5 cm) (lcm)

Present study

400(400 cm)

1 50 ( . 1 5 s

)

6667

(10 m/s

)

( 1 cm

)

Temperature and density of the flow are fixed constant in the present

calculations. Figs.3-A and 3 show the behavior of vortices and isooaric

lines along the jet, calculated by dividing the numerical domain

into nonuniform 111 x 117 grid. In this calculation a short-time disturbance was added to the jet (by dislocating the nozzle to one adjacent

grid briefly), in order to accerelate the initiation of the flapping.

Even without the disturbance the flapping of a jet began spontaneously in

this calculation. The vortices appeared alternately on opposite sides

of the jet. The bending locations of a flapping jet coincided with the

locations of vortices. The frequency of the periodic flapping motion

of a jet depends upon the initial speed of the jet from the nozzle, i.e.

Reynolds number. Due to the flapping motion of the jet, the half-width

of the jet increases and the maximum velocity decreases along the jet axis

as shown in Figs. 3- A and B, based on the time-averaged velocity

distributions

.

3. Examples of other oscillatory flows calculated numerically

The two flows (mentioned above) calculated by the author definitely

display periodic oscillations. Recently time-dependent flows have

widely been investigated by numerical calculations. Examples of other

oscillatory flows calculated by several investigators are:

511
3.1 Unsteady wake behind a circular cylinder or a rectangular block

Numerous numerical studies solving the unsteady Navier-Stokes equations

have been conscerned with the unsteady features of the wake behind a circular cylinder or a rectangular block placed in a uniform flow. It is

considered as a very convenient test case for fluid-dynamic computations.

Hence some numerical results calculated are extremely precise.

For example, Braza et al. 12 studied the initiation mechanism for

vortex shedding in the range of the Reynolds number of 100 - 1000.

Fig.5~A shows the alternating vortices calculated, where the loss of

symmetry was initiated by a numerical perturbation (by adding a brief

rotation to the cylinder artificially). The periodic flow pattern was extremely similar to that in experiments, although the numerical

perturbation was not always the same as the physically existing

purtubations in the experiments. Further, Ogawa et al. 13 investigated

the vortex shedding characteristics, too. The unsteady flow pattern

in the wake, with the Reynolds numoer of 400000, was extremely precise

as shown in Fig . 5-8 , displaying a periodic alternating behavior. A number

of small vortices regularly leaving the surface of a circular cylinder

are observed. In this calculation no particular perturbation was reported

as having been used to initiate the vortex shedding.

These alternating flows in the wake behind a block must be common with

the oscillatory flows of_ turbulent jet mentioned above. Cantwell et

al. 14 reported contours for mean Reynolds stress in the wake behind a circular cylinder as shown in Fig.5-C, which displays a similar structure

shown in Fig. 3-3 for a turbulent jet.

3.2 Unsteady thermal coavection caused by a localized heat source

Kuwahara et al. 22 attempted to investigate periodic motions of thermal

convection due to a ring- type heat source, with the intent of eventually

studying firewhirls. The size of the ring was 0.5 km for the inner radius

and 2.5 km for the outer one. Numerical calculations were based on axisymmetric cylindrical co-ordinates with the 3oussi'nesq approximation.

The isotherms above a hear source were highly time- dependent as snown

in Fig. 6. Profiles at 6600, 10800 ana 15500 seconds belong to the same phase. A vortex is periodically created near the ring-type heat source. Three-dimensional calculations will additionally give three-dimensional

twisting flows around the plume axis. This author 23 studied the thermal

convection between two circular plates, using axisymmetric cylindrical

coordinates, as shown in Fig. 7. The flow was weakly oscillatory.

Matubara 24 studied the merging and oscillatory flows of buoyant plumes

above two line heat sources as shown in Fig. 8. Kou et al. 2S calculated periodic variations of smoke concentration

around an aircraft fuselage in a crosswind near the ground, as shown

in Fig. 9. Kuwahara et al. 25 numerically studied the time-dependent

flow in a channel with a partially heated floor. The flow pattern

is highly oscillatory as shown in Fig. 10. These oscillatory flows must

be based on vortices created periodically around a heat source. Nakamura et al. 27 attempted to computationally determine the coupling

between flame size, the speed of the induced (combustion) air, and the

burning rate. Fig. 11 represents a steady state flame of a 3unzen

burner. A weakly oscillting temperature profile around the top of a flame

is suggested. Baum et al. 28 studied the plume pulsations in a cubic

enclosure, as shown in Fig. 12 where an artificial disturbance was added.

In these oscillatory flows periodically created vortices are suggested.

512
4. DISCUSSION

[Numerous vortices around flows are commonly created in buoyant flows

shown in Figs. 2 and 5 and forced cool flows shown in Fig. 6. In twodimensional flows, vortices are alternatively created in each side of

main stream, where pressure imbalances are determined by the von Karman

vortex-street. In three-dimensional calculations such as by Baum et al.

,

too, numerous vortices should be created and grow to cause pulsations.

The thermal plume will be stable if there are perfectly no disturbance

around the plume. However, even very careful experiments can not be free

from physically existing disturbances. Likewise, if there are perfectly

no truncation errors in calculations, the plume above a heat source will

not oscillate. The oscillations in the wake calculated by Ogawa et al.

must be initiated by small truncation errors accompanying calculations.

Thus, an increase in numerical precision will delay the commencement

of oscillations. Purturbations added to the plume accerelate the

commencement of oscillations. However, these effects in calculations

will not greatly affect the natural oscillatory frequency itself, but

affect the time oscillations commence. An artificial purturbation and

small truncation errors in calculations are only a trigger in initiating

the oscillatory flows. Hence the natural oscillatory flows calculated as above reflect the physical phenomena.

Vortices created around buoyant flows and turbulent jets play a significant role in the turoulent flows. Thus ccmoustions , or fires, must

be greatly affected by the oscillatory flows.

5. CONCLUSION

The present two-dimensional calculations showed the existence of

vortices appearing alternately on opposite sides of both a plume and a turbulent jet axis. In these results calculated, numerous vortices are periodically created near the stream and cause the periodic fluctuations

of flows. These periodic fluctuations are essentially the same as the

Karman- vortex street in the wake behind a circular cylinder, which

has been simulated well by many investigators.

The origin initiating oscillatory flows between numerical calculations

and experiments differs each other, but they only play as a trigger to initiate the periodic flows, that is, they only affect the time

oscillations commence. Oscillatory behavior of flows in some numerical

examples mentioned above are quite common and are very similar to the

actual fluid flows. Thus the natural oscillatory flows calculated as above reflect the physical phenomena.

Vortices created around buoyant flows and turbulent jets play a significant role in the turbulent flows. Thus combustions , or fires, must

be greatly affected by the oscillatory flows.

REFERENCES

1

)

Frank-Kamenetskii , D . A . : Diffusion and Heat Transfer in Chemical Kinetics , 2nd edition Tr. by J. P. Appleton, Plenum Press, New York, p. 560(1969)

2)

Williamson , R . B . et al .: "Observations of Large Scale Turbulence in Corner

Wall Experiments", Combustion Science and Technology , Vol.41, p. 83 (1984)

3)

Satoh,K. et al. : "A Numerical and Expermental Study of the Relationship

between Periodic Temperature Variations of Outflowing Air and Heating

513
Rate in an Enclosure", Proc. 16th Conf. Fluid Mech- (Japan) , p. 154 ( 1934)

4)

Satoh,K. et al.:"A Numerical Finite-Difference Study of the Oscillatory

Behavior of Vertically Vented Compartments", Numerical Properties and

Methodologies in Heat Transfer , Hemisphere Publishing Co.,?. 5 17(1931)

5)

Satoh,K., et al. :"Flow and Temperature Oscillations of Fire in a Cubic

Room vith Ceiling and Floor Vents (Part 3 Relationship between Frequency

and Heat Source Strength), R e o . Fire Res. Inst ., No. 57, p. 79 ( 1984)

6)

Saton,K. et al.:"Flov and Temperature Oscillations of Fire in a Cubic

Enclosure with Ceiling and Floor Vents (Part 1 Oscillatory Behaviors & Heat Source Strength), ibid. , No. 55, p. 31 (1983)

7)

Satoh,K. :"A Numerical Study of Swaying Motion of Fire Plume(in

Japanese)", Proc. 23rd Japanese Symposium of Heat Transfer , p. 424(1986)

3) Sat oh , X .

: "A Numerical Finite Difference Stuay of Natural Swaying Motion

of Plume above a Heated Line Source", Proc. the 2nd ASME-JSME Thermal

Engineering Joint Conference , Honolulu, Vol.4, p. 41(1987)

9)

Saton , X .

: "A Numerical Study of Two-Dimensional Free-Jet"(in Japanese),

Proc. 1 3 1 h Conference of 7i u i d M e c h a n i c s 'Japan) . ? . 4 2 ( 1 9 8 6

)

10)

Satoh, ec al. : "Three-Dimensional Numerical Simulation of Aircraft

Cabin Fires", Proc. 24th Aircraft Symposium (Japan) , p. 326(1986)

11)

Liu,V. et ai. : "UNDSAFE-II , A Computer Code for BuoYant Turbulent Flow

in an Enclosure vith Thermal Radiation", Tech. Reo. of the Department of

Aerospace and Mech. Eng, of Univ. of Notre Dame . TR — 79002-73 — 3 (1973)

12)

3raza,M. et ai .: "Numerical Study of the Pressure and Velocity Fields in

the Near Wake of a Circular Cylinder", J. Fluid w ecn. , Vo 1 . 1 65 , p . 79 (1986)

13)

0gava,S. et al.:”Fiov around a Circular Cyiinder(in Japanese)", Proc

.

17th Conference of Fluid Mechanics(Jaoan) , a . 22(1985)

14)

Cant veil , 3 . ex al.:"An Experimental Stuay of Transport in the Turouient

near Wake of a Circulac Cylinder", J . Fluid Mech

.

, Vol. 136, p, 321(1983)

15)

Ku,A.C. et al. : "Numerical Modeling of Unsteady Buoyant Flows Generated

by Fire in a Corridor", 16th Symo. (Int.) on Comb, , p. 1373 (1977)

15) Patankar ,S . W . ^Numerical Heat Transfer and Fluid Flow, McGraw-Hill ( 1980)

17) Forstroo , S . J . et al. : "Experiments on the Buoyant Plume above a Heated

Horizontal Wire", Int. J Heat Mass Transfer , Vol. 10, p. 321 (1967)

I3)Notc at ai. :"Svaying Motion in Thermal Plume above a Heat Source",

Bull. Japan Soc. Mech. Eng. (Ser. B) , Vol. 50, No. 457, p. 2179(1934)

19)

Antonia,R.A. et a 1 . : " 0 n the Organized Motion of a Turbulent Plane Jet",

1. Fluid Mech. , Vol. 134, p. 49 (1983)

20

)

Shimizu , S

.: "Numerica Analysis of Two-Dimensional Jet by a DiscreteVortex Method”, Bull. Japan See. Mech. Eng. (Ser. 3) , No. 4 72, p. 3852(1 985)

21

)

Hasegava , T . ec al .: "Numerical Simulation of Two-Dimensional Jet by a Discrete-Vortex Method", ibid. , Vol. 52, No. 475, ?.!4oi(1986)

22)

Kuwahara , K . et al. ^"Thermal Convection caused by Ring~:ype Heat

Source", J. Physical Soc. Japan , Vol. 51, No. 11, p. 3711(1982)

23

)

Satoh , X .

: "A Numerical Study and Experimental Verifications of Air

Entrainment by Fire Plume", Fire Scl. A Tech ., Vol. 2, No. 2, p. 127 (1982)

24)

Matubara , "A Consideration of the Oscillatory Behavior of Flame of Pool

Fire", Proc. 1986 Conf. Japan Soc. Fire Sci. and Eng. ", p. 111(1986)

25)

Kou,H.S. et al. .’"Turbulent Buoyant Flow and Pressure Variations around

an Aircraft Fuselage in a Cross Wind near the Ground”, Fire Saiety

Science-Proceedings of the First International Symposium , p. 173 ( 1985)

26)

Kuwahara , K .

: "Computation of Thermal Convection with a Large Temperature

Difference", Proc. 4th Int. Conf. Applied Num. Modeling , p.565v!984)

27)

Nakamura , S . :"A Computational Model for Fire Research", Proceedings of

International Symposium on Computational Fluid Dynamics , p. 678 (1985)

2S)Baum, H . "Calculations of Three Dimensional Buoyant Plumes in Enclosure"

Combustion Science and Technology , Vol.4., p. 55 (1984)

514
Fig. 1-A Oscillatory frequency of buoyant plume vs. heating rate

of central volumetric heat source in a vertically vented room.

Fig. 1-3 Velocity vectors and temperature contours.

Heating rate Q=10 kW, enclosure dimensions : 51 x 61 ca.

( 1

) -(4 ): Present calculation

( 5 ): Experiment by floto at al. Fig.2-A Reynolds number on frequency

of swaying plume vs. Rayleigh number.

Enclosure dimensions : SO x 100 ca. Lines (1) to (4) depend on boundary

conditions (details are found

in Ref. (8) )

.

515
Fig. 2-9 (right)

Swaying notion

of plume depicted by

velocity vectors (upper)

and isotherms (lower)

.

Heating rate Q= 1 20 ktf/m. Calculated by nonuniform

103 x 110 grid.

Fig. 2- A (middle left)

Flapping motion of jet

deoictea by velocities.

Re-5667.

Fig. 3-3 (niadele right)

Isooaric contours.

U

1 — ——

,

1

1

i • i |

i

i i

1 1

1 1 III i

1 i : mu

-Li

f • ' JJ1J2

1 II II —' —

L r

i

i 1 1

i

1

i Mini II

1

• II 1 Mill

|_ _ _ _!_|j

j_|||j_

i ij i m

Jill.

>oo

x 1000

r ; i H

in 'i i tin it |

| 1

1

7

*

:

:

rrt :

:

:

'

i i i i i i i i 3 , I' l

1

1

i

1 II 1 I J

. 1 II II 1 1 1 1 i i 5

i t 1 Ill

III

1

! 1 _Li J_U1

: 1 Mil

_ _! J . _ Tl

Fig.4-B (bottom right)

Maximum velocity

along jet axis.

Fig. 4- A (bottom left) Half width of velocity distribution

along jet axis (x)

.

516
k Mumm

y^:fS^.c' -;

BSSTo'-

•

\

'

. 'W/s'/ ///•' ', mmmr

Fig. 5-3 Vortex shedaing

for Re=4Q000.

(By Ogava et al.

13

)

Fig.5-A Oscillatory notion of wake

behind a circular cylinder.

(By Braza et al.

12

)

-J OOUM— - o.OCUOJ

-C Vs.

.... jprr

A >'

0.0694 J -*3 0204— \ \

0 0204 —* 0.0I22 -*

Fig.5-C Contours for mean Reynolds shearing stress.

(By Cantwell et al.

14

)

i o

/SOBARlC CONTOURS

.-cm

"'I 1

j \ X?,/

v. k.

.a : ico -un

!

! i i‘f ; i r

_r^ :reo 1*00

: ! 'i!i

: V

=

.

_ ,-Np^

-.-CS ::ts f

'

. • *i

’.13 iM UCC ' f

- ^ :::

^4

. ^

-3S -CO ..CO v •

1 r~ r ; !

1 ! OS.

13 •r»

i ii V7 -rr~s

.-3 -3 jrco

!

j 1 :•/ -(

\ k2

A\

Fig. 5 Theraal convection

caused by ring-type heat source. (By Kuvahara et al. 22

)

517
Fig. 7 Buoyant plume above a circular

plate- type heat source. Weak oscillation

vas obserbed. Cylindrical coordinates

vere employed. (By the present author 23

)

5_2 x v\

j j

-ifmec ) J i

/

•

f> ?! j: 1 j

3 c

l

| ^

l

i

!i

! |

(.9.5a*c

ii

II

1

;

/ \

J u

• 1 —i*C V- [

Fig. 8 Merging of plumes above tvo line

heat sources. (By Matubara )

Fig. 9 Oscillatory buoyant plume due to fire

located closely to an aircraft fuselage.

(By Kou et al.

25

)

Fig. 11 Slov diffusion flame. The fuel gas

mixture was ignited just after the

opening of a gas valve.

(By Nakamura et al.

27

)

Fig. 10 Thermal convective flow in- a channel

with partially heateci^ floor

.

(By Kuvanara et al. "

)

Fig. 12 Three dimensional buoyant plume in

a cubic enclosure.

(By Baum et al.

2

)

518
CLOSING CEREMONIES

519
CLOSING CEREMONIES

JIN: We would like to begin our formal closing ceremonies. When the Japanese

delegation arrived in Boston, our impression was that this was the end of the

winter season. In the one week that we have been privileged to be here, we have seen the magnolia bushes begin to blossom and the maple trees begin to

put on their red leaves as they start to flower. We have the definite

impression that spring has finally arrived. They have told us that the New

England spring is a very short lived one. We have found that this also

applies to this particular conference. It has already begun and ended in a matter of moments that we have been conscious of its existence. It has been

customary, and I believe I, as the visitor, have the privilege of beginning

the closing ceremonies.

There are many very important items in the resolution. Four individuals, Dr.

Pagni as the lead, participated in the formulation of the final resolution.

Specifically, I must say that they began their work last night and finalized

it this morning. There is a reason for Dr. Pagni' s passion in assisting us in

this regard. The reason being that if we get too involved in arguments over these resolutions, Dr. Pagni will miss his airplane, so I would like to

request everybody's cooperation in passing the resolutions post haste.

I would like to request that Dr. Pagni start off by reading the English

version of the resolution.

PAGNI: We left the fate of the fat mice out of the resolutions.

The resolutions follow Mr. Chairman:

The members of the United States -Japan Natural Resources Board on Fire

Research and Safety are quite pleased with the results of the 9th Joint

Meeting, held in Boston, Massachusetts , May 3rd through 8th, 1987. We wish to thank the National Bureau of Standards, the National Fire Protection

Association, and especially Factory Mutual Research Corporation, for the New

England hospitality we have all enjoyed. We also are grateful for the

insights provided by Professor Hottel from MIT and Professor Emmons from

Harvard University. We note again that each time we meet, the quality and

quantity of our technical communication improves . The four workshops on special topics proved informative and stimulating for all participants.

The toxicology discussions held jointly with Canadian representatives showed

good progress. We hope this Tripartite Cooperative will continue.

The following resolutions summarize the consensus reached:

It is hereby resolved that:

Resolution 1: The objectives of the meetings of this Panel remain to:

A. Exchange the latest technical information.

B. Promote cooperative research on fire safety science.

C. Encourage the innovation necessary for the development of new testing

methods, designs and standards.

520
D. Form a multinational consensus in the field of computer-based fire

modeling

.

E. Explore the possibility of developing performance fire codes.

Resolution 2: The 10th Meeting of the UJNR Panel on Fire Research and Safety

will be held in June of 1988 in Tokyo or Tsukuba, Japan.

Resolution 3: The same format as the 9th Meeting is proposed for the 10th

Meeting, with shorter sessions to accommodate attendance of panel members and

associates at the Second International Symposium on Fire Safety Science.

Resolution 3A: Agenda Topics:

1 . Fire and smoke physics.

2. Fire and toxicity chemistry

3 . Risk

,

hazard and evacuation

4. Open technical session.

Resolution 3B: The discussion of each topic should begin with an outline of

progress made in each field between meetings by both Japan and the United

S tates

.

Resolution 4: Cooperative fire research programs should be continued. These

are especially valuable where the facilities are somewhat different in our two

countries. It is suggested that joint research programs be organized on:

1. The structure of large, buoyant diffusion flame plumes.

2. Smoke movement and fire growth computer modeling.

3. Fire retardant and combustion product chemistry.

4. The proper definition of risk and risk analysis.

5. Other topics at the discretion of the Chairman.

In addition, the exchange of working personnel should continue to be pursued

vigorously. During the past two years, Dr. D. Evans, an NBS employee, spent a valuable month in Japan. Dr. A. Sekizawa of FRI , Mr. Y. Yamauchi of Hochici

Corporation, Professor A. Ito of Ohita University and Mr. H. Otani of IHI have

conducted useful research projects in the United States.

Resolution 5: Panel members are encouraged to exchange information of

interest through the respective Chairmen between meetings . Research reports

issued in each country should be exchanged every three months. The titles of

these reports will be sent to the Panel members in each country who can identify the ones they would like to receive from their Chairman.

It is important that all the relevant reports for the next meeting be sent in

time to be received in the other country at least one month before the 10th

Meeting

.

Mr. Chairman, those are the resolutions.

TAKAHASHI : I was .told that both resolutions would be read in their respective

languages, so Mr. Jin will read the Japanese version of the resolutions.

521
JIN: Prior to the reading of the resolutions, I'd like to make a commtent to

just the Japanese delegation. The resolutions were made in such haste that

perhaps the translation may not be as accurate as desired. Therefore, I'm

asking the Japanese delegation to rely -on the English version as final.

TAKAHASHI : I believe, after reading over the resolutions, that there were

three main points of quite high importance in my mind. One important issue is

the statement that refers to the toxicity studies, and the recommendations

that these studies be conducted under a Tripartite Cooperative System. The

second is that there is a need to decide at this Session that the 10th Panel

Meeting to be held either at Tokyo or at Scuba. I would prefer that the

decision on this particular site be delayed if that is possible.

The third important point is, I believe, the addition of sub -items 3 and 4 in

the 4th resolution, which called for additional studies. I believe it is a very good thing for this type of Panel to have their sights set on as wide and

as broad a spectrum as possible.

Are there any questions or discussions that anybody might have concerning this

resolution?

SNELL: For my part, I'd simply like to say I agree with your three points and

observations: that the Tripartite continue, that the decision on location be

deferred based on your situation in Japan, and thirdly, I too think of some other topics that we might wish to pursue on cooperative projects and look

forward to discussing those with you later.

JIN: Thank you very much for your comments. Then would it be all right to assume that we have accepted the resolutions?

In deference to Dr. Pagni's very tight schedule, we seem to all have agreed

that the resolutions are acceptable.

I'd like to proceed on to the next item on the agenda and call on Dr. Gann to

summarize his views on the current meeting.

GANN: Good morning again. It is a pleasure to be asked to summarize the

value of this meeting. Unfortunately, I was unable to attend the Risk, Hazard

and Evacuation session because we were holding parallel meetings. However, I have read those papers, and I plan to include some discussion of them in my

remarks

.

In my opinion, this is easily the best of the UJNR meetings that I've

attended. There were some research areas in which remarkable, innovative

progress has been made. There are some other areas in which the progress can best be described as incremental. I think the mixture of these two is

healthy, but I think that in the future, we should reconsider the spacing of

the meetings so that we can have major progress to report in virtually all of

the areas

.

There is an aspect of maturity about the technical nature of this Panel.

We've all now known each other for many years, and the exchange of ideas seems

to be far more fluid. In many of the projects that were reported here during

the last week, it's obvious that ideas that have contributed to those projects

522
have 1 come from both countries. Since this is the essence of holding these

kinds of meetings, I think we should applaud ourselves for taking advantage of

it

.

In thinking about the sessions and the papers on fire and smoke physics, I consider that this is by far the most mature, the most well-developed of the

disciplines discussed this week. Because of this large technical base and

experience, it is not surprising that there are remarkable advances each time

we meet, both in the understanding of the physical phenomena of fire, and in

the mathematical modeling. We've just spent the last hour or so discussing

the accomplishments of the Toxicity Tripartite Group over the last five years.

Suffice it to say that this year, as in each of the prior years, the amount of

progress, the evolution of our thinking, and the wholesale generation of data

is extremely encouraging. The most gratifying and most rapid progress has

been made in the area of risk and hazard evaluation. Perhaps the youngest of

the fire disciplines is the true fire chemistry field. For decades, we have

been making analytical measurements of the products of combustion and the

history of how materials fall apart during fires. However, there are still

very few papers on the true chemistry of how those processes occur.

First of all, as the hosts of the prior UJNR meetings in the United States, I and my colleagues from the National Bureau of Standards greatly appreciate the

excellent hosting job provided by Factory Mutual Research. In particular, Ms.

Madeline Anderson has been an excellent resource.

In addition, the opportunity to tour laboratories that many of us have not seen before, or certainly not recently, allows us to visualize the kinds of

apparatus and where the results come from. I also think we should provide a round of applause for our interpreters who have made a fairly significant

language barrier disappear this week.

I would now like to make a few brief announcements on procedure. As in the

past, we will be preparing proceedings from this meeting. We hopefully have

tape recordings of the discussions from the various papers and will be

transcribing it. We have the original copies, I believe, of virtually all the

papers. As in the past, we allow about a one-month period after the meeting

for those of you who would like to submit revised versions. If you send the

revised version to NBS within the next four to six weeks, we will be happy to use it in place of the original. If not, we'll just go ahead and use the

paper you provided prior to the meeting.

Next, I'm going to preempt Dr. Snell, who normally has the pleasure of

presenting some sort of parting souvenir for the attendees at the meeting. As

I recall, approximately two years ago in Tsukuba, we were presented small

models of a traditional Japanese fire standard from years past. In fact, I still have mine sitting on my desk. This year, we have another totally unique

parting gift for all participants. The National Fire Projection Association

has published a special commemorative facsimile version of the very first Fire

Protection Handbook. For those of you who are familiar with the current

version, it is extremely large and thick and contains almost everything you

need to know about fire. This is the same version of that book 90 years ago.

In the carton in the second row, there are copies for everybody. For those

people whom already have left the meeting, we ask that people please take

copies back to them.

523
And last, after the closing ceremonies are over, we would like to take a photograph of all the people who are still here. Please go out the back door

and find a suitable location. This should only take a moment or two and you

will still be able to catch your flight.

SNELL: My remarks will be very brief because the efficient Dr. Gann has

stated most of them already and has presented them very well. I too am impressed with the quality of this meeting, as distinct from our previous

meetings that I've observed. Even though I missed the important heat transfer

experiments Wednesday night, I am pleased the Red Sox not only won the game

but saw to it that nobody became ill.

More seriously, after reading the papers and hearing your discussions of them,

I think that the time has come for us to begin working on a new type of

cooperative project. A number of years ago, Professor Emmons challenged us to develop the pieces for predicting fire hazard and safety. Reflecting on the

comments of Professors Hottel and Emmons this week, it seems that we now are

in a position to confront how to bring these pieces together. It's a problem

bigger than either of us can solve separately, so I propose we go at it

collectively. Dr. Tanaka has noted the success the Japanese have had in

obtaining resources for cooperative efforts. I think the time has come for us

to try to work the same game the other way, so you might expect us coming to

you to ask for similar help in the near future.

I am pleased that the experiment of holding this meeting outside of Washington

has been so successful, and warmly support Dr. Gann's comments concerning the

hospitality of our New England friends. In sitting right here, one observes

the youthful energetic appearance of the current Japanese delegation. Dr.

Gann called it maturity; maybe there are other words that come to your mind

when you observe the other side. Be that as it may, we shall continue to do our best effort to stay the course.

Now, and on behalf of the U.S. Panel, and our hosts at NFPA and Factory Mutual

Research Corporation, we wish you all a safe journey back home.

TAKAHASHI : We have finally reached the end and I too would like to make a few

words of comment. Many people have already stated this, but I would like to

repeat and add my thanks to the very kind and helpful reception given to us by

the NBS , and particularly by FMRC and NFPA. Judging from your warm reception,

we were able to gather from the warmness of the feelings that were addressed

to us from the American side. And one more word of appreciation. I would

like to comment on the Tripartite Cooperation between Japan, the United States

and Canada. We talked of success of these joint meetings but joint meetings

cannot succeed unless there is a meeting of minds of very important and

respected participants. Therefore, I hope that the 9th Meeting will lead to

the 10th, 11th, and infinitum, and that as many participants as possible will

continue for as long as possible. I also would like to express my thanks to

the people who were backstage in this particular conference. We cannot

disregard the tremendous part that was played by Ms. Madeline Anderson in

making this conference both memorable and good in feeling.

This is not part of the agenda but I would like to present her with a memorial

.

524
ANDERSON: Thank you very much.

TAKAHASHI: She is not here today but there is one other very important person

in these meetings and that is Ms. Debbie Cramer. She was the one responsible

for wading through the voluminous papers that was exchanged prior to the

meetings. I would like to ask Dr. Snell to present this to her in my place.

SNELL: I would be honored to do so.

TAKAHASHI: I don't believe there was any more difficult situation and I'd

also like to thank the interpreters.

We have now reached the end and I'd like to say that this will close the

meeting. I hope to see you in June 1988 in Japan.

525
'
AUTHOR INDEX

A

Apostolakis, G., 146

Atreya, A. , 24

B

Babrauskas , V. , 100

Baler , L. , 368

Baum , H . , 50

Brandyberry, M. , 146

Bukowski , R. , 219

C

Carrier , G . , 305

Curtis , R. , 461

D

de Ris , J . , 83

E

Emmons , H . , 24

Evans , D . , 325

G

Gann, R. , 235 , 490

Grand, A. , 405

Gurman, J . , 368

H

Hall, J. , 131

Harmathy, T. , 238

Harttell , G . , 405

Hasemi , Y. , 4 , 41

Henzel , V . , 100

Hirano, T. , 73, 397

Hirata, T., 417

Holt, T. , 368

Hotta, H. , 357

I

Ito , A. , 496

J

Jin, T. , 189 , 294

Joh, T. , 63

K

Kashiwagi , T . , 496

Kawamoto , S . 417

Kim, A. , 428

L

Levin, B . C . , 368

Levin, B .M. , 280

M

Mashige , J . , 63

Morikawa, T. , 443

Mulholland, G. , 100

N

Nakamura, K. , 113

Nakaya, I . , 397

Nelson, H . , 157

Nishihata, M . , 41

Nishimaru, Y. , 455

Nishina, T. , 443

0

Ohlemiller , T. J .

,

13

Oka, Y. , 357

527
p

Paabo , M . , 368

Parker , W. , 257

R

Rehm , R . , 50

S

Saito , F . , 336

Sakurai , T . , 386

Satoh, K. , 509

Sekizawa , A . , 189

Sugawa , 0 . , 63 , 357

Suzuki, H. , 232, 491

Switzer , W . , 405

T

Takahashi , K. , 203

Tanaka, T., 136, 172, 203, 240

Tewarson , A . , 267

Tsuchiya, Y.-, 428, 461, 492

Tsuda, Y. , 455

Tsuj imoto , M . , 139

Y

Yamada , T . , 294

Yanai , E . , 443

Yao , C. , 314

Yoshida , M . , 336

Yoshikawa , T . , 63

Yusa , S . , 470

Z

Zukoski , E . , 8 , 111

528
N BS-1 1 4A (rev. 2 -ac) C.5.S OGPT, DE. OF COMM. BIBLIOGRAPHIC DATA

1. PUBLICATION OR

REPORT NO.

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

SHEET (See instructions) NBSIR 88-3753 April 1988

4. TITLE AND SUBTITLE

9TH JOINT PANEL MEETING OF THE UJNR PANEL ON FIRE RESEARCH AND SAFETY PROCEEDINGS

5. AUTHOR(S) , Nora H. Jason and Barbara A. Houston, Editors

6. PE R FORMING ORGAN I ZATION (If joint or other than N BS. see in stru ction s) 7. Contract/Grant No. NATIONAL BUREAU OF STANDARDS

U.S. DEPARTMENT OF COMMERCE

GAITHERSBURG, MD 20899

8. Type of Report & Period Covere

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

)

d

10. SUPPLEMENTARY NOTES

|

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

11. ABSTRACT (A 200-word or less factual summary of most significant in formation . If document includes a significant

bibliography or literature survey, mention it here)

The 9th Joint Panel Meeting of the United States-Japan Panel on Fire

Research and Safety was held jointly with the Combustion Toxicity and

6th Expert Meeting of the U . S

. -Japan-Canada Cooperative Research Group

on Toxicity of Combustion Products from Building Materials and Interior

Goods at the Factory Mutual Research Corporation, Norwood, MA. Technical

sessions were in the areas of: Fire and -Smoke Physics; Risk, Hazard and

Evacuation; and Fire Toxicity and Evacuation. Progress reports were presented in each area, in addition to state-of-the-art papers. The

next conference will be held in Japan in June 1988.

12. KEY WORDS (Six to twelve entries; alphabetical order; capitalize only proper names; and separate key words by semicolons)

building fires; -evacuation; fire hazards; fire models; human behavior;

risk analysis; smoke; smoldering combustion; sprinklers; toxicity;

wood

.

13. AVAILABILITY 14. NO. OF

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PRINTED PAGES

| | For Official Distribution. Do Not Release to NTIS 547

| Order From Superintendent of Documents, U.S. Government Printing Office, Washington, D.C. 20402. 15. Price pq Order From National Technical Information Service (NTIS), Springfield, VA. 22161 $42.95

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A

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