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PUBLICATIONS

NBSIR 81-2350

Smoldering Combustion Hazards

of Thermal Insulation Materials

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U.S. DEPARTMENT OF COMMERCE

National Bureau of Standards

National Engineering Laboratory

Center for Fire Research

Washington, DC 20234

August 1981

Interim Report

Prepared for; Department of Energy

n ak Ridge National Laboratory

ak Ridge, Tennessee 37830

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NBSIR 81-2350 ^ ^ ^981

nc>^ a . SMOLDERING COMBUSTION HAZARDS

OF THERMAL INSULATION MATERIALS

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T. J. Ohlemiller

U.S. DEPARTMENT OF COMMERCE

National Bureau of Standards

National Engineering Laboratory

Center for Fire Research

Washington, DC 20234

August 1981

Interim Report

Prepared for; U.S. Department of Energy

Oak Ridge National Laboratory

Oak Ridge, Tennessee 37830

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

NATIONAL BUREAU OF STANDARDS. Ernest Ambler. Director

TABLE OF CONTENTS

Page

Abstract 1

1. Introduction 1

2. Smoldering Ignition 3 2.1 Nature of the Problem 3

2.2 Summary of Ignition Test Studies 4 2.3 Recommendations Concerning Smolder Ignition Test Development 5 2.4 Follow-on Work on Smolder Ignition 7

3. Smolder Propagation 7 3 . 1 Background 7 3.2 Objectives 8 3.3 System Studied and Procedures 8

3.4 Discussion of Typical Results 10

3.5 Limitations on Wave Structure Resolution 12

3.6 Factors Controlling Wave Shape 14

3.7 Some Effects of Boric Acid 17

3.8 Estimate of Heat Release Distribution 19

3.9 Summary of Wave Structure Findings 23

4. References 24

Figures 26-32

Appendix I Celluloslc Insulation Material III. Effects of Heat Flow Geometry on

Smolder Initiation 33

111
LIST OF FIGURES

Page

Figure 1. Attic mock-up adapted for smolder propagation studies 26

Figure 2. Typical profiles from floating thermocouples at different depths in the

insulation bed 27

Figure 3. Cross-sectional snapshot of steady-state smolder wave structure in

nominally unretarded insulation 28

Figure 4. Cross-sectional snapshot of steady-state smolder propagation in insulation

with 25% boric acid add-on 29

Figure 5. Sketch of oxygen inflow paths and heat release zones in spreading smolder

wave, inferred from analysis of wave structure profiles 30

Figure 6. Thermal analytical results for unretarded and boric acid-retarded

insulation 31

Figure 7. Heat effects (endothermic and exothermic) inferred from temperature-time

trace provided by shallow floating thermocouples 32

LIST OF TABLES

Table I. Lateral smolder propagation rates through 18 cm deep cellulosic

insulation layers 10

Table II. Summary of integrated heat effects along 1-D particle path through

smolder wave 21
Smoldering Combustion Hazards of Thermal Insulation Materials

Thomas J, Ohlemiller

National Bureau of Standards

Washington, D.C, 20234

The smoldering combustion hazards of cellulosic loose fill insulation

materials fall into three categories: smolder initiation, smolder propagation,

and transition from smoldering into flaming. Previous findings on the initiation problem are summarized briefly. They serve as the basis for recommendations on an improved smolder ignition test method which is designed to give

ignition temperatures comparable to those in practice. The proposed test

method requires checking against full-scale mock-up results before it can be

considered for implementation.

Smolder propagation, driven by buoyant convection, through a thick (18 cm)

layer of cellulosic insulation has been extensively examined. A heavy (25% addon) loading of boric acid (a widely used smolder retardant) cuts the propagation

rate in half (from approx. 0.3 to 0.15 cm/min) but does not come close to stopping this process. Analysis of experimental profiles for temperature, oxygen

level, and remaining organic fraction strongly indicates that the smolder wave

is oxygen-supply controlled and that it involves both first and second stages

of oxidative heat release from the insulation material. The balance of involvement of the tv;o stages varies with depth in the layer. It appears that efforts

to develop improved means of suppressing smolder propagation must be directed

at the entire oxidation process. However, since boric acid is fairly effective

at slowing the second stage of oxidation, most new efforts should be aimed at

the first stage of oxidation (which also is responsible for smolder initiation).

Smolder transition to flaming is the next problem to be studied; the above

results form the reference state that is perturbed into flaming when air is

forced to flow through or around the smoldering insulation.

1. Introduction

The work reported here is a continuation of that described in ORNL/Sub 7686/1 (ref. 1)

.

The work began at Princeton University and is continuing at the National Bureau of Standards

Center for Fire Research.

1
The focus of this work to date has been the smoldering combustion hazards of thermal

insulation materials, in particular, loose-fill cellulosic insulation. This type of insulation is a re-cycled paper product, typically made from newsprint which is mainly ground

wood. The normal combustibility of wood is accentuated when it is ground to a fibrous form

and utilized in low bulk density layers. The flaming aspects of this combustibility are relatively well-controlled by presently available additives such as borax. Smoldering com- bustibility is less well-controlled by present additives (e.g., boric acid). As a result,

smoldering may be initiated in state-of-the-art insulations that are improperly installed^

in contact with heat sources such as recessed light fixtures.

It should be noted that flaming and smoldering are fundamentally different kinds of

combustion so it is to be expected that they would respond differently to retardants.

Smoldering is not a well-characterized combustion process since it involves complex solid

phase oxidation chemistry coupled to mass and energy transport processes. One of the goals

of the present study is to shed as much light as possible on the nature of smoldering

phenomena in cellulosic insulations so that this form of combustion can be suppressed.

There are three aspects of smoldering combustibility that are pertinent to the

insulation usage hazard. First is the ignitability to smoldering. There are several questions about this aspect that one can ask. These concern the conditions that lead to smolder

initiation, the factors that influence this process, the feasibility of realistically

measuring the smolder ignitability tendency of an insulation, and the changes necessary to

lessen this tendency. Most of these questions have been addressed in the current work; the

status of the answers is discussed below.

The second aspect of smolder combustibility concerns the propagation process that occurs subsequent to ignition. The smolder reaction zone may spread outward from the ignition source indefinitely as long as there is a continuous layer of insulation. One could

ask the same questions as above about this process. It is currently under study and it is

apparent that the answers are substantially more complex than in the ignition situation.

The results to date are summarized below.

The third aspect of smolder combustibility is that of the transition from smoldering

to flaming combustion; this occurs, typically, if the smolder zone receives a sharply

increased air supply. Again the same questions apply. This subject has not yet been

examined but will be briefly scanned in the coming year. It is important because it is the

mechanism whereby the smoldering process leads to rapid development of threats to life

safety and property destruction.

In section 2 we briefly summarize the smolder ignition work. This work has been the

subject of three papers. The first two were included as appendices in last year's report

(ref. 1); they have been published in Combustion Science and Technology (ref. 2, 3). The

2
third paper is included here as Appendix I; it has been accepted for publication in the same journal. Recommendations for a revised ignition test method are discussed here in some detail. The next area of the smolder ignition problem in need of study is also

indicated.

The results to date on the process of smolder propagation are discussed in some detail

here (Section 3) since they have not been previously reported. Finally we conclude with a brief discussion of what problem areas need further study.

2. Smoldering Ignition

2.1 Nature of the Problem

The initiation of smoldering in cellulosic loose-fill insulations appears to be most

probably and most frequently a hazard when the insulation is in contact with large area

heat sources such as recessed light fixtures (1, 4, 5, 6, 7) . A light fixture is essentially a constant power heat source; the temperatures achieved within it at steady state

depend on the rate of heat loss. \iThen insulation is placed around the body of a fixture

recessed above the ceiling level, much of the pathway for heat loss is severely inhibited.

If an oversize lightbulb is then used, the steady state surface temperature of the fixture

may reach into the range from about 225-275°C where current insulations begin to react

appreciably with oxygen (1, 6). These reactions are exothermic so that they can become

self-accelerating. If the initiating source, i.e. the light fixture, is at or above a minimum critical temperature, a self-sustaining reaction process begins in the insulation.

It can spread away from the light fixture because the reactions themselves now produce

enough heat to ignite adjacent elements of the insulation.

The reactions that produce this heat are not in the gas phase but rather involve a direct attack of oxygen on the wood fibers that comprise the insulation. These heterogeneous reactions produce the unique characteristics of smoldering; their nature is not well

defined at present.

The minimum critical temperature for ignition of smoldering is not a unique property

of the insulation. Ignition occurs when the local rate of heat release continues to exceed

the rate of heat loss. Heat losses depend on the geometry of the heat source, the thickness of the Insulation, its thermal conductivity, etc. These factors, in addition to the

reactivity and exothermlcity of the Insulation, must enter into the determination of the

critical conditions that dictate the smolder ignition borderline. This complicates the

design of a test method that will provide a unique and meaningful measure of the tendency

of an insulation to ignite to smolder.

3
2.2 Summary of Ignition Test Studies

The total problem of a constant power, extended area heat source imbedded in a layer

of oxidizable insulation has never been treated analytically. Although this could be done

in principle, it would get quite involved and overly specialized if all of the realistic

details were incorporated. The behavior of the insulation is most directly responsive to

temperature; the light fixture can be viewed mainly as a device that imposes a temperature

on the boundary of the insulation and sets the general geometry for heat losses from the

insulation. In examining the behavior of the insulation then, it is simpler to replace the

light fixture with a constant temperature source and vary that temperature in separate

tests to find the minimum value that just produces ignition. In so doing we have somewhat

de-coupled the problem relative to what occurs in practice and this must be borne in mind

when the results are used to develop an ignition test method; a constant power source (e.g.,

a recessed light fixture) will not always behave the same as a constant temperature source although they are nearly equivalent in the almost steady-state tests used here.

We have stated that heat source geometry is important in that it affects the heat

losses from the initially reacting layer of the insulation. One of the goals of the current

work was to demonstrate this quantitatively; a second goal was to ascertain if this behavior

could be predicted quantitatively. If we understand the ignition behavior well enough to

make quantitative predictions about it, we have a firm basis for going to the next step,

i.e., developing a suitable test method for smolder ignition tendency.

The effect of heat source geometry on minimum source temperature to produce ignition

was studied for several configurations. These geometries were suggested by real heat sources such as those seen in residential attics but they were simplified to facilitate

comparisons with a numerical model of ignition. The various sources were all tested in an

"attic mock-up" so that the results would be reasonably transferable to the real world.

The details of the sources and the mock-up are given in Appendix I. Determination of the

borderline or minimum source temperature for smolder initiation required four or more separate tests for each configuration since the border can only be found by go/no-go testing.

(In such a test the source temperature is pre-set and held constant; one waits as much as

12 to 24 hours to see if ignition occurs.)

The minimum ignition temperature for the same insulation was found to vary from 235°C

to 385°C as heat source configuration was altered. The lowest value came from a source

that partially surrounds the insulation in a 90° corner. The highest value is for a hot

cable-like source traversing the 8 in. insulation bed at mid-depth.

These results were predicted with an accuracy of 10°C by a numerical model of the

ignition process. The model accounted for heat generation by a zero order reaction and

the transfer of that heat to ambient through the insulation. The "one reaction" responsible

for ignition in the insulation was shown to be the first overall reaction sequence exhibited

4
in themogravimetry (TG) and differential scanning calorimetry (DSC) studies of the

insulation. The source of the heat release in this first overall reaction (cellulose,

lignin, etc.) is a subject in need of further study.

Since the model was reasonably successful, it was used to infer a test configuration

that should satisfy two requirements: relatively short test time, and ignition temperatures

in the same range as those seen in practice with recessed light fixtures. These tend to be

mutually contradictory requirements (see Appendix I) but it appears that they can be approximated with, for example, a heat source that is shaped like a flat-bottomed bowl, having a diameter of about 15 cm and a depth of about half the diameter.

2.3 Recommendations Concerning Smolder Ignition Test Development

The above configuration is believed to be capable of producing smolder ignition in a nominal insulation at a borderline temperature of about 250°C in a one hour test time.

Yarborough, ^ al. (6) found ignition occurs with real light fixtures which achieve peak

temperatures of 225-275°C; such full-scale testing requires up to 24 hours per test.

The output of tests with the above configuration (after several go/no-go applications

of it at several temperatures) is the minimum source temperature that just produces ignition

(in this test) . A single insulation composition thus gives a single number (or a small

range, perhaps 5°C due to data scatter). By itself, this number is of limited meaning since

it is dependent not only on the insulation but also on the test configuration. It is put

into perspective by the previous work in this study (Appendix I) and the work of Yarborough

^ (ref. 6). Previous work establishes some correlation between ignition temperature

in this configuration and the ignition temperatures seen in idealized residential attic

heat sources. The work of ref. 6 provides approximate ignition temperatures for real light

fixture exposures. The combination implies that the small-scale test result will be in the

same range as that with real light fixtures.

This relation between the possible test method and the real hazard situation is fairly

well established but it needs to be strengthened. The problem is two-fold because of the

diversity of both light fixtures and insulations. It must be ascertained if:

1) the minimum ignition temperature ranking of insulation compositions depends on the

configuration of the light fixture producing ignition.

2) the ignition temperature-based ranking of insulations is substantially the same in

the small-scale test and in the real hazard situation.

The first concern is a product of the complex nature of the actual ignition chemistry

(represented by one reaction in the model discussed above) . Different light fixtures

5
comprise differing heat source geometries and so they will each produce differing ignition

temperatures for the same insulation . Since the way in which the real oxidation chemistry

occurs is somewhat temperature dependent and this dependency may be a function of insulation composition, it is possible that two different light fixtures, if used to rank a series of insulations for smolder ignition tendency, could yield differing ranking orders.

If this were so, there clearly would be an ambiguous correlation between a small-scale test

method and such realistic full-scale tests. It is necessary to determine if this occurs within some realistic range of both light fixtures and insulation compositions. This calls

for choosing a range of insulations and two or three distinctly different light fixtures

i

The fixtures would then be used in full-scale go/no-go tests where source temperature is

monitored in much the same way the small-scale tests are run; each fixture would be used to obtain the minimum source temperature for ignition of each insulation and the rankings from

each fixture would be compared. The non-constancy of temperature over the surface of each

fixture calls for careful consideration since it poses an ambiguity in the temperature to which the insulation is responding. Obtaining the source temperature distribution calls

for rather extensive instrumentation of the tests.

The second concern above arises from the same source, the temperature-variability of

the ignition chemistry. Even if the various fixtures above rank the series of insulations

in the same order, it must be established that the small-scale test will do the same. This

is a much briefer task since the small test is much faster. There would Inevitably be

quantitative differences (small to large-scale) in the ignition temperatures of the various

insulations. Even if this did not alter the ranking order (small to large-scale), these

differences should be carefully scrutinized for any implications they may have about

relative behavior on the two scales.

If the small-scale test method comes through the above two checks satisfactorily (as

expected), it is ready for round robin testing. Success here depends on careful constraints

on the design, manufacture and usage of the small-scale tester.

The result of all this provides a data base for assessing what constitutes acceptable

performance in the small-scale test. One could establish a series of performance classes

with each class encompassing a range of ignition temperatures in the test. The most preferred class would fall entirely above the range of test ignition temperatures that implies

ignition in the preceding full-scale light fixture tests. At present, it is not clear that

there are any commercial insulation compositions that would fall in this class if its lower

temperature bound is based on the most ignition prone light fixtures. Nevertheless, there

is incentive to set the levels of acceptability high both because repeated sub-ignition

exposures may drive smolder retardants such as boric acid out of the insulation, lowering

its ignition temperature, and because there are indications in the literature that wood

products may become more ignitable after prolonged (months or years) high temperature

exposure (ref. 8, 9, 10).

6
If the author's experience with newsprint-based insulations holds true in general, the

ignition borders of most insulations in the small-scale test may fall in a rather narrow

range of temperatures (within 20 to 30°C of 250°C, for example). This would not mean the

test method is deficient, especially if the above checks of the test against the full-scale

hazard are satisfied. This outcome would instead reflect the rather weak effect present

smolder retardants have on smolder ignitability . This in turn implies that effective means

for fully eliminating the smolder ignition hazard are not yet in hand. The rather small

effect that retardants such as boric acid have on smolder initiation (ref. 1, also

Appendix I to this report) is, in a sense, implicit in the fact that smoldering remains a substantial combustion hazard. If really effective retardants were available (at reasonable

cost), the hazard would be eliminated. They are not.

2.4 Follow-on Uork on Smolder Ignition

The only smoldering analog that has been studied in depth is carbon oxidation. There

exists a substantial body of literature on the effectiveness of various additives in

inhibiting this oxidation. Unfortunately, this is likely to be of little or no use in

learning to inhibit smolder ignition in cellulosic insulation because the process starts not

with oxidation of the carbonaceous char but with much earlier, more complex reactions

(ref. 2, 3).

Most cellulosic loose-fill insulations are basically ground wood (newsprint). Wood

consists of several polymers and various inorganic contaminants. We do not presently know

which one (or more) of these components is responsible for the low temperature exotherm that

initiates smoldering. One of the major goals of the present work is to ascertain this and

then study the sensitivity of this exotherm to various additives. We are not assured of

finding the means to suppress smolder initiation but pertinent information on lessening its

probability will emerge.

3. Smolder Propagation

3.1 Background

After smolder is initiated in a cellulosic insulation by some heat source, it typically

becomes a self-sustaining process in a manner analogous to a flame. Some of the heat

released during ignition and subsequent combustion is transmitted to adjacent elements of

insulation, heating them to the point where they too ignite. This process can continue

indefinitely, spreading far away from the original heat source, but it need not do so. Propagation of smolder, like ignition, involves a balance between heat release rates and

heat loss rates. If the balance can be tilted in favor of the losses, smolder will extinguish. Unfortunately, for all of the practical materials and use conditions with which the

author is familiar, the balance favors continuing propagation. From the limited studies to

7
date, it appears that current materials will extinguish only in layer thicknesses less than

about 6-10 cm (ref. 3).

Smolder propagation is a complex process that has not been extensively studied. Nearly

all of the available experimental data are parametric in nature, mainly concerning the

effects of such variables as air supply on propagation rate. The most extensive parametric

study is that due to Palmer (11) ; however it includes only very limited information on the

internal structure of a propagating smolder wave, namely, part of the temperature distribution in a two-dimensional (2-D) smolder wave propagating through a thin layer (1-2 cm) of ' sawdust. A more detailed study of a simpler system, 1-D forward and reverse smolder propagation through cellulosic insulation, was completed by the present author and a co-worker;

a report on that work is in preparation.

3.2 Objectives

This limited amount of information on the details of smolder propagation, especially

in a realistic configuration, prompted the present work. There are two general objectives.

First, to gather evidence on what drives and controls smolder propagation in cellulosic

insulation; this clarifies the problem of inhibiting the propagation process. Second, to

establish the nature of the buoyantly-driven smolder case; this is a reference or base case

for the forced airflow situation where transition to flaming is much more probable.

3.3 System Studied and Procedures

The configuration is a thick (18 cm) horizontal layer of cellulosic loose-fill

insulation placed in the same attic mock-up as was used for the ignition tests discussed

above. Figure 1 shows this set-up with the igniter and instrumentation used in this propagation study. Since ignition behavior is not part of this study, smolder is initiated as quickly and uniformly as possible on one end face of the insulation bed. This is accom- plished with an electrical resistance heater that is formed into an up and down zig-zag

pattern with about 2 cm spacing (see Fig. 1); this is run from a constant temperature

controller, typically at 350-400°C.

Once smolder is initiated in this plane, it is the author's experience that it is

always self-sustaining and continues to spread in these thick (18 cm) beds, whether the

insulation is retarded or not. The oxygen supply to the spreading reaction zone is by diffusion and natural convection; the forced air flow case will be addressed in the future.

The smoke plume flow above the bed is inevitably somewhat disturbed by the necessity of

placing the whole mock-up in a laboratory fume hood; there is no direct cross flow over the

bed surface, only random eddies that do not appear to comprise a major disturbance.

8
Under these very weak (no forced flow) oxygen supply conditions, the propagation rate

of the smolder along the trough formed by the two boards (see Fig. 1) is extremely slow.

The smolder reaction zone may require 10-12 hours to reach the other end of the mock-up,

90 cm away, in the case of a retarded insulation. The rate of movement of this reaction

zone is followed by three sets of two thermocouples each, embedded at initially fixed depths

_2 in the insulation. These are 36 American wire gage (1.27 • 10 cm dia.) bare chromel/alumel

thermocouples arranged so as to "float" freely in the insulation. The bed depth decreases

as much as a factor of two during smolder wave passage so that these thermocouples must be

free to move downward with it to avoid disruption. The thermocouples are carefully put in

place during the packing of the insulation bed so that their spacing and initial height are well-defined. Given the thermocouple spacing and the temperature time histories at successive points through the bed, one readily calculates the lateral propagation velocity.

Typically it is based on the time necessary for the 250°C isotherm to pass between

successive thermocouple locations.

The structure of the smolder wave is a product of the factors which control propagation.

"Structure" means the spatial distribution of temperature, gas phase chemical species and

solid phase species. The details of the structure can suggest the dominant factors in the

propagation process and the relative roles of the first and second stages of heat generation

in driving smolder spread. The variables of primary interest are the temperature, oxygen

concentration and solid fuel distribution; it is these three on which we have focused.

Resolving a two-dimensional distribution is a much bigger problem than is resolving a 1-D distribution. The most feasible technique for the present problem involved a moveable

probe, inserted from above the fuel bed to various known depths. The probe always remained

at one longitudinal location (approx, at mid-length of the bed) and the smolder wave moved

past the point at a steady rate. Thus measurements (versus depth) made at successive times

could be converted, via the propagation velocity, to values at various lengths along the

smolder wave. Profiles of temperature and oxygen concentration were obtained in this manner. For temperature the probe was a 0.025 cm dia. sheath thermocouple (chromel/alumel)

which protruded about 1 cm from a 1 mm outside dia. stainless steel tube providing mechanical support. For oxygen the probe was a piece of 1 mm outside dia. tubing through which a local gas sample was drawn at 4 cm^/min. A Beckman 778 polarographic oxygen analyzer

measured the 0^ partial pressure in this sample stream. Both of these probes produced

inevitable disturbances, as discussed below.

The solid fuel distribution could not be measured during a test. Instead, after the

smolder wave had propagated sufficiently to provide other data, it was extinguished by being

smothered with nitrogen. This was carried out as quickly as possible to minimize deviations

from the actively propagating configuration. After extinction the bed was sectioned and

3 small samples (few cm ) were removed from fixed locations. These were ashed at 700-750“C

in air. From the ash content and knowledge of the nature of the ash precursor, one can

9
calculate the fraction of original organic material that remained in each sample. The

uncertainties here largely result from ambiguities in ash behavior as discussed below.

3.4 Discussion of Typical Results

About one dozen tests of this type have been carried out thus far using a commercial

insulation with a proprietary retardant blend, a nominally unretarded material, and that

material with 25% boric acid add-on. The smolder velocity results are summarized in Table I.

Table I Lateral Smolder Propagation Rates Through

18 cm Deep Cellulosic Insulation Layers

Insulation Composition

Bulk

2 Density (g/cm ) Smolder

Velocity (cm/min)

Commercial blend 4.45-10"^ 0.133 + .002^^^

(2) Nominal unretarded 3.55*10“^ 0.306 + .014^^^

Nominal unretarded +

25% boric acid

add-on(^) 4.43-10"^ 0.143 + .008^^^

Unretarded + 33%

boric acid add-on'' 4.73*10“^ 0.175

Average (+ avg. deviation) from three tests; two earlier tests with this material were

discarded because the thermocouple arrangement (vertical insertion) was found to be

disruptive.

( 2 ) From same company that made the commercial blend; ashing indicates this nominally

unretarded material actually contains 4-5% of the retardant mix that is in the commercial

blend

.

(3) Average (+ avg. deviation) of four tests; a fifth test with this material transitioned

spontaneously to flaming after 2-3/4 hours of smoldering.

(4) Boric acid was technical grade, through 200 mesh, milled into insulation for one hour in

a roller mill.

(5)

( 6 )

Average (+ avg. deviation) from two tests.

Material blended by a different manufacturer.

10
Note that the bulk densities of the insulation have been adjusted, as retardants were

used, to keep the density of the actual cellulosic material approximately constant at about

-2 3 3 the level in the unretarded case (3.55*10 g/cm or about 2.2 Ib/ft ).

In each test four values of smolder velocity were obtained from the time for the 250°C

isotherm to pass between successive pairs of the six floating thermocouples. While there

was typically a + 5% scatter in these values, it showed no trend with time or position,

indicating that the smolder wave was in a steady-propagation state in its journey along the

central 36 cm segment of the bed length (where all measurements were made) . The thermocouple

traces from the same depth in the wave (3 each for two depths) were virtually identical from

ambient conditions to more than 500°C, also confirming the steady state. In Table I, the

average of the four values per each test has itself been averaged among tests to give the

reported numbers.

It is apparent from Table I that the net effect of either a commercial blend of

retardants or a high loading of boric acid alone is merely to slow down the smolder spread

process (by about a factor of two); propagation is never brought to a halt. These results

confirm our earlier tentative conclusion (ref. 1) based on a smaller and somewhat less

realistic experimental set-up. We are interested in learning, however, just where and how

boric acid, one of the most effective smolder retardants available, has the effect that it

does. This could provide clues as to what is needed to find still better retardants.

The floating thermocouple traces from the two different depths in the wave (1.8 cm

from the top and 5.4 cm from the top) provide the first clues that the smolder wave structure

is not simply one-dimensional in nature. Fig. 2 shows typical temperature-time traces at

the two depths for both the nominally unretarded material and that material with 25% boric

acid added on. Note that the time scale is in hours; this is a very slow process. The

temperature- time history of a particle of insulation material is clearly different at the

two depths in the bed. A shallow particle undergoes a faster rate of heating from 100 to

450°C, quickly reaching a peak; the subsequent history then depends on the presence or absence of boric acid. A deep particle undergoes a less rapid initial heating but there is

further sustained heat-up leading ultimately to a higher peak temperature than in the

shallow position. These histories reflect the heat transfer and heat generation processes

in the moving smolder wave; we will return to this point below.

One cannot get a full picture of the wave structure from the above results. The

moveable probes provided much more data in this regard. Recall that a thermocouple and a gas withdrawal probe were inserted simultaneously from above (about 3 1/2 cm apart across

the width of the wave) to a series of fixed depths. This was done at intervals of ten or more minutes as the smolder wave moved steadily past the probe position. Temperature or oxygen mole fraction versus depth and time were then converted to purely spatial distribution data using the measured smolder velocity to convert time to distance. Finally, the

11
loci of various even values of the parameters (e.g., 50, 100, 200°C) were found by linear

interpolation of the recorded results and these points were connected by smooth lines. The

resulting profiles are indicators of the spatial structure of the wave. Fig. 3 shows typical results for the nominally unretarded material; Fig. 4 is for that material with 25%

boric acid added on. The top surface profile was obtained by simple measurements from a fixed reference point at successive times.

3.5 Limitations on Wave Structure Resolution

Before these profiles are discussed, it is important to note certain limitations in

them. First, insertion of a thermocouple from above gives results inherently less accurate

in this system than are the results from the floating thermocouples; nevertheless the

inserted thermocouple was necessary as a means of obtaining closely spaced data points since use of a large number of floating thermocouples is impractical. Both are comparably small

in thermal mass and this is not a problem. However, the probe inserted from above crosses

the isotherms at angles as high as 90°. Where these isotherms are closely spaced (region

of rapid temperature change) along the thermocouple leads, preferential conduction along

these leads will cause the thermocouple junction to be at a temperature different from the

local temperature it is supposed to be reading. This is a complex problem to assess quantitatively in the present system. It can be modeled plausibly by analogy to the classic

thermowell problem (ref. 12) but the result depends strongly on the effective conductivity

of the thermocouple lead and the heat transfer rate to it; neither of these values is known

for this system. The problem is minimized by using a very thin thermocouple; we have used

the smallest feasible size (0.025 cm dia.) consistent with penetrating the bed without

bending. The problem of temperature falsification due to this source is believed to be

minimal except perhaps in the one or two regions of most rapid temperature change along the

thermocouple lead. Any falsification results in shifting the apparent position of the

isotherms—upward when the thermocouple penetrates down into a gradient whose temperature

increases toward the junction and downward in the opposite gradient. The true peak

temperature in a small domain will also be underestimated.

The oxygen mole fraction profiles have a limited spatial resolution due to inevitable

sampling disturbances. Isokinetic sampling would be the ideal in this system but the local

gas flow velocity is so low here (estimated ~0.04 cm/sec) that the sample probe would have

2 to be about 1-2 cm in area to give the needed sample volume flow rate. The necessary

3 sample flow rate 4 cm /min) is set by system response time considerations and other limitations discussed below. To minimize thermal and mechanical disturbances upon probe

insertion we have used a gas probe that is a 1 mm OD by .75 mm ID stainless steel tube

3 drawing at about 4 cm /min. The gas velocity into the probe tip is then about 200-300 times

greater than the local flow velocity in the bed in the absence of the probe. The flow disturbance problem has been investigated somewhat (ref. 13) but not out in this very high

flow ratio domain. Rough extrapolation of those results would indicate a disturbance out

12
to about 10 probe radii (here this would equal about 5 mm) . In the absence of any normal

3 flow in the bed (infinite flow ratio), one can estimate that the 4 cm withdrawal in a typical one minute of sampling would come from a sphere of 1 cm in diameter around the probe

tip. We conclude that the oxygen sampling resolution is probably about + 1 cm (i.e. the

probe averages the sample over this domain). Clearly, steep gradients cannot be resolved;

the actual gradients in the bed are probably steeper than those shown in Fig. 3 and 4,

especially in the upper layers of the bed above the charred material. It was not possible

to vary the sampling flow rate appreciably to investigate this question. Greater flow rates

caused continual probe clogging problems; lesser flow rates extend the sampling time

excessively.

Sampling time is limited by another phenomenon. The bed shrinks as the smolder wave

passes through it but the local motions are not only downward, they are also lateral, up to

2 cm in the unretarded case. This lateral motion is locally inhibited by the probes while

they are temporarily in the bed for sampling purposes and the result can be an oblong hole

around the probe or even a total fracture across the width of the bed. When this happens,

the local conditions no longer are representative of an undisturbed smolder wave. This

problem was most severe toward the latter end of early tests, particularly with unretarded

insulation. A variety of small accommodations has been developed to minimize the problem

3 but short total immersion time for the probe is vital. One minute at 4 cm /min is needed to

purge the gas sampling line and allow the polarographic probe to come to equilibrium.

Figures 3 and 4 also show profiles of the remaining organic material at various

positions. As indicated previously, these data were obtained in the extinguished smolder wave by removing and ashing samples a few cubic centimeters in volume. This means that the

result obtained is an average over the sampled volume and we cannot get fine resolution of

rapidly changing domains. The resolution is comparable to the oxygen probe, about + 1 cm.

Translating the ash content of a solid sample to the local fraction of organic material

remaining is done by the following relation. It is made somewhat complicated because the

ash is a retardant residue that could have lost weight during the smolder process and during

the ashing of the sample.

e = (y6)

Fro

F sw

(1 - F /e) sw

(1 - F ) ro

Here c is the fraction of remaining organic in the sample, y is the weight of smolder

altered retardant per unit weight of original retardant, g is weight of ash produced per

unit weight of smolder-altered retardant, F^^ is the weight of original retardant per unit

weight of original insulation mix and F^^ is weight of final ash per unit weight of sample

from the extinguished wave. Only F is measured; F is set by the chosen mixture of

retardant and insulation. The quantities y and g require a knowledge of the behavior of

the retardant as a function of temperature. Furthermore, to apply them correctly to a given sample, one must know to what temperature it has been exposed.

13
The retardant mix in our commercial blend is known only in a qualitative sense, though

we can infer the probable weight fraction range of the components from published results on retardant effectiveness (ref. 14). From this we can in turn infer that the commercial

retardant would lose about half of its weight by 300-350°C and undergo no more weight loss

during the ashing process at 700-750°C. Thus we estimate that above 300-350°C for this

system y = 0.5 and 3 = 1.0. Since there is a small amount of the commercial retardant blend

in our nominally unretarded material, these values were used in converting ash values to

values of e for the nominally unretarded bed (Fig. 3). The behavior of boric acid alone is

much easier to characterize though still subject to some uncertainties. Boric acid undergoes

two successive water loss reactions to end up as ^20^ by about 300°C. The stoichiometry of

this is clear but the complication comes from the volatility of boric acid in a system (i.e.

degrading insulation) where both water and methanol are present. .We have ashed known mixtures of insulation plus boric acid and find that some small but significant loss of boron

does occur. The best values presently available for the boric acid treated insulations are essentially the same as those above, i.e., y = 0.5 and 3 = 1.0 for samples exposed to 300°C

or more in the smolder wave. Because of ambiguities and scatter in the boron loss aspect,

the value of y is probably uncertain within + 10%-15% and this puts an equal uncertainty in

the value of e. Because the retardant transformations may get more complex and time-dependent

below 300° C, we have not yet made any attempt to resolve the remaining-organic profiles

below 300°C in the smolder wave.

3.6 Factors Controlling Wave Shape

Despite this long list of caveats regarding the wave profiles in Fig. 3 and 4, they

contain a large amount of useful information. Interpretation of these results is still in

progress but several pertinent points can be made.

First, the length of the smolder wave should be noted; it is 30 to 60 cm long, about

2 to 3 times greater than the bed depth. This ratio of wavelength to layer thickness is

comparable to that found by Palmer (ref. 11) for sawdust layers a few centimeters deep,

suggesting that similar factors may control the rate of oxidation over this range of thicknesses. Essentially all realistic smolder configurations appear to be oxygen-limited (see

Appendix II), i.e., the overall rate of oxidation is limited by the rate of 0^ supply and

it increases when that supply increases. This appears to be the case here as well; note

the extensive domain in the high temperature region of the wave where 0^ is fully depleted.

In this horizontal layer configuration, O

2 can get to the reaction zone in two ways, by

natural convection flow and by diffusion from the surrounding atmosphere. Natural convec- tion'flow through the insulation is, of course, quite difficult due to the high specific

surface area of the fibers. We can estimate the magnitude of the convective flow velocity

by applying a balance between flow drag and buoyancy forces. This leads to the following

expression:

14
(?) (^)(?)

Here v is the desired estimate of the buoyant flow velocity in the insulation, g is the

gravitational acceleration, is the ambient air density, a^^ is the measured drag constant

for the insulation, T is an average temperature in the smolder region, is the ambient

temperature, and y are the gas viscosities at ambient temperature and at T, respectively.

The drag constant is measured on virgin insulation; it undoubtedly decreases due to mass

loss in some portions of the smolder wave. The "average" temperature in the wave is taken

as 350°C; the calculation is not very sensitive to the exact value. Putting this and other

typical numbers into the above equation gives v_ = 0.04 cm/sec; the actual values throughout

the wave could easily be + 50% different from this estimate but we are, for the moment, most

interested in the leading part of the wave where this value is reasonably appropriate.

Fig. 5 is a sketch of the two 0^ supply modes that we wish to compare. The sketch of the

convective flow is quite rough; no measurements inside the bed have been possible. The

buoyant plume above the hot reaction zone is shifted around considerably by the fume hood

flow in actual experiments so the average O

2 level at virtually all points above the smoldering layer is expected to remain essentially at 21% (even in less disturbed circumstances

this would be true over most of the smolder zone because the plume coming out of the porous

layer contracts sharply). Consider the convective inflow of O

2 from the left in Fig. 5 compared to the diffusive flow downward from above. The convective 0^^ flow is approximately

OC

The diffusive flow is approximately

These are mass flows in g/sec per unit width of the smolder layer; Y . is the ambient

oxygen mass fraction (above the bed) . Here we use i the layer depth, as a rough characteristic length both for the convective influx and for the length of the reaction zone affected

by these two O

2 fluxes. The O

2 diffusivity D is evaluated at the average layer temperature

above the reaction zone; AY (e Y ) is the change in 0„ mass fraction from the top of

the layer to center of the reaction zone. The average 0^ diffusion length is at most

equal to the distance from top of the bed to the edge of the O

2 depleted region in Fig. 3 or Fig. 4; because of probe volume averaging effects, the actual value of is probably less

than this. If typical numbers based on Fig. 4 are inserted in the above expressions, we

find

- 9 • 10 ^ g/sec

OD

10 ^ g/sec

15
This rough method of calculation probably overestimates the convective flux and

underestimates the diffusive flux The results suggest (but do not firmly establish)

that ©2 diffusion is the dominant 0

^ supply mode in these natural convection conditions

(This would change drastically if there was a forced flow through the insulation layer as can occur near the eaves in an attic. Such forced flow can induce a transition to flaming).

The conclusion that O2 diffusion probably dominates the O2 supply is reinforced by the

following rough argument concerning how O2 diffusion would dictate the shape of the reaction

zone. Suppose the actual reaction zone was thin compared to the insulation layer (assumption

of fast kinetics) ; then the reaction zone could be considered to be a surface whose shape we want to determine. Consider the O2 requirements and supply to a differential element of

that surface; following Gugan (ref. 19 )

Here the element of surface is of length ds and unit width; it is situated at depth y below

the top surface of the fuel layer. The layer is smoldering at a steady rate; we can thus assume ds is fixed and fresh fuel comes into this collapsed reaction zone at a rate set by

the smolder velocity v . The quantity entering ds per second is P^^v dy where p„ is the

bulk density of the fuel. Oxidation of this fuel requires n^ mass of O2 for mass of fuel

(stoichiometric coefficient). The required O2 must all come from above by diffusion through

the layer depth y. Then

n

0»SB''s'^y

This differential equation specifies the shape of the reaction surface by relating x and y on the surface.

With the origin (x = y

= 0 ) at the top of the layer where the reaction surface begins, this

specifies a parabolic reaction surface shape:

16
TOP

The qualitative similarity of this to the isotherms or organic fraction profiles of Fig. 3 and 4 is apparent. Even though the real reaction zone is thick and there is layer shrinkage,

ignored in the above analysis, the similarity of the two profiles lends credence to the

assertion that the general shape of the smolder wave is dictated by its dominant 0^ supply

mode, i.e. diffusion from above. One can even use the above simple expression to roughly

estimate the length of the smolder wave (by inserting y

= £) if n^ is known. It can be

shown that if n^ is approximated by summing estimates of the convective and diffusive O

2 supplies (over the whole wave) that go with a measured smolder rate, the predicted value of

the smolder wave length is of the correct order of magnitude (but may be off by a factor of

two) .

3.7 Some Effects of Boric Acid

The reaction zone in reality is not only thick (due to finite rate kinetics), it is

comprised of more than one reaction stage. Fig. 6, taken from ref. 1, shows how unretarded

and retarded (boric acid) insulation behaves during thermal analysis. There are two broad

stages of weight loss and heat release which we will refer to as char formation (1st stage)

and char oxidation. Recall that in Section II it was mentioned that smolder initiation is

largely due to the heat released in the char formation stage. It is of interest to determine to what extent these two overall reactions stages play a role in smolder propagation.

Inspection of Fig. 6 shows that boric acid does not affect these two stages equally. Rather

it is substantially more successful in slowing the heat release rate of char oxidation. The

actual oxidation chemistry underlying these two overall heat release/weight loss stages is

very likely quite different. The first stage probably involves 0

^ attack on molecules that

still bear substantial resemblance to the original cellulose, lignin, etc.; the second stage

probably more closely resembles carbon oxidation. Thus the influence of retardants and

catalysts on either stage is likely to be different; the effect of boric acid in Fig. 6 is

consistent with this. If we knew that only one stage dominated smolder propagation, efforts

at developing better retardants could be more sharply focused. However, as will be seen,

the evidence now points at both stages.

Fig. 6 serves to relate insulation weight loss (or fraction of original weight

remaining) with heat release stage. It should be noted that although the horizontal axis

in Fig. 6 is temperature, it alone is not an indicator of what stage of weight loss or heat

release is achieved. This is because both the heating rate and 0^ availability history are

different in the smolder wave than in thermal analysis.

17
The lowest sets of profiles in Fig. 3 and 4 are those of the fraction of remaining

organic material. Reference to Fig. 6 gives rough dividing lines between the two reaction

stages that can be used to help interpret the remaining organic profiles in Fig. 3 and 4. A plausible dividing line would be a vertical line dropped to the baseline from the lowest

point between the two reaction peaks. The dividing line between heat release stages for

the unretarded material occurs approximately at a remaining weight fraction of 0.2 to 0.25;

for the boric acid retarded material it is more diffuse but roughly 0.3 to 0.35 weight

fraction remaining. (Note that the 5°C/min heating rate in Fig. 6 is four to eight times

below the pertinent rates in the smolder wave; this is not believed to have a substantial

effect on the dividing lines assigned here.) Both the retarded and unretarded insulations

go unambiguously below their respective dividing lines indicating both involve first and

second stage heat release. The boric acid retarded insulation in Fig. 4 stops at a significantly higher fraction of remaining organic indicating that the boric acid cuts off the

last part of the char oxidation stage.

A further indication that boric acid effectively cuts off the final part of char

oxidation is shown in the isotherm profiles of Fig. 3 and 4. The "variable" portion of the

unretarded temperature profiles is caused by an "afterglow" phenomenon common to cellulosic

materials. Here it manifests itself as a dull orange-red glowing region, 1 cm or more in

size, that intermittently and randomly propagates around in the charred residue at the tail

end of the smolder wave. This process is particularly dangerous since, when it approaches

the partially quenched material that always exists at the boundaries of the smoldered insulation layer, it can generate, and ignite to flaming, pyrolysis gases. This happened spontaneously in one of the tests with our nominally unretarded material, i.e., transition to flaming in the absence of boric acid need not require a forced air flow. By lessening the

probability of transition to flaming, boric acid serves an important safety function even

if it is not very effective in retarding smolder initiation or propagation.

Still another indication of the ability of boric acid to retard the latter part of

char oxidation came in the sample ashing process used to determine the fraction of remaining

organic material. For samples containing boric acid, the final ashing temperature had to be raised from 700 to 750°C and the time at that temperature had to be more than doubled to assure virtually total burn-off of all organic material.

Returning to Fig. 3 and 4, we note again that smolder in both the retarded and

unretarded insulations involves both stages of heat release shown in Fig. 6. Furthermore,

there seems to be a somewhat differing distribution of the weight loss (and hence heat

release) processes. The second stage weight loss in the unretarded case is concentrated

in a thin region at the top of the insulation layer (at least in the first 20-30 cm of the

reaction zone). Below that there is a rather extensive region of nearly constant weight

fraction. This latter region coincides with the 0

^ depleted zone. Reaction is suppressed

here due to a lack of O

2

. The incoming ©2 from above is being totally consumed by the char

18
oxidation. The O

2 from the region ahead of the wave is evidently being totally consumed by

the first stage of oxidation, probably just inside the indicated line (actually a few millimeter zone) of color change in the insulation. It should be noted that the O

2 stoichiometry

of the two stages is almost certainly not equal. An estimate is available for the first

stage (ref. 15), 0.3 g 0

^

1 % fuel consumed; the value for the second stage is believed to be

higher in approximate proportion to its higher heat release (ref. 1), i.e., about six times

that of the first stage.

With boric acid there are similar features in the weight loss process but they are modified by the slower kinetics of char oxidation. There is now a rather thick region

(~2 cm) at the top of the insulation layer that is partially quenched, i.e., the weight

fraction remaining stays up around 0.25. This thicker quenched region is due to a shift in

the local rate of heat loss versus heat generation because of the slower char oxidation.

Below this the char oxidation process stops at about 0.16 weight fraction again because the

boric acid has retarded the oxidation kinetics to the point where the achievable heat release

rate, balanced against the inevitable (though here fairly slow) heat loss rate, cannot produce temperatures capable of consuming the remaining char in the time available, even though

0

^ is available. Below this kinetically retarded region is a layer of partial char oxidation and, apparently, some latter part of the first stage heat release as well. The initial

part of the first stage heat release is again occurring near the color change line but it

is terminated by consumption of the slow 0

^ supply from the unburned region of the material.

3.8 Estimate of Heat Release Distribution

If the available temperature data had a finer spatial resolution, one could back out

the heat release distribution quantitatively. This is an established procedure for simpler

systems such as laminar flames (ref. 16, ref. 17). Since it requires both first and second

derivatives of the temperature distribution, the data must be much more finely spaced than

that used in Fig. 3 and 4. We do, in fact, have very good data from the floating thermocouples placed in the bed but there is one substantial limitation on that data: it provides

a detailed temperature history only along two particle paths through the smolder wave, a shallow path and a deep path (see Fig. 2). It cannot provide data on the full 2-D heat

release distribution. However, the shallow thermocouple path through the wave happens to

follow a particle whose thermal history is nearly one-dimensional (1-D) because it moves perpendicular to the isotherms from about 100°C to nearly 500°C. For the particle to follow

a 1-D history, it must also not be greatly influenced by convective heat carried in the gas

since the gas crosses the particle path at varying (unknown) angles. It can be shown, using

the previous estimate of the gas velocity, that convective transport to a particle following

the shallow thermocouple path is only about 15-20% of the conductive heat flux it sees. This and other necessary approximations limit the accuracy of what follows but the result

is still informative.

19
\-Jhen the gas convective flux is neglected, a fuel particle proceeding on a path normal

to the isotherms of Fig. 3 or 4 obeys the following energy balance equation

ds s ds ) ‘^Sb'^s^s dTds + Q 0

Here the derivatives are taken along the direction of the particle path, v^ is the speed of

movement along that path which we take equal to the smolder spread velocity (somewhat in error near the high temperature end where bed shrinkage adds another velocity component)

;

3 is the local heat effect in cal/cm sec and it can be endothermic or exothermic. Note that

the equation applies to a particle moving through a stationary smolder wave at the smolder

velocity. The effective conductivity of the insulation undoubtedly increases somewhat

with temperature as the material degrades and a radiative transfer component (fairly small

in this material) adds in. Because of limited data, we are forced to take a constant value

of X^, using the data of ref. 18 to estimate a value at a mean temperature of about 200°C.

Finally we convert from length s to time t by using the measured smolder velocity and the

above equation, solved for Q becomes

-&)

dfr

dt^

The time derivatives are to be taken on the temperature-time histories for the shallow

thermocouples shown in Fig. 2.

Despite the very smooth appearance of the data in Fig. 2, it was obtained with a digital recorder that rounded the temperatures to the nearest degree. If one attempts to use the raw data to calculate Q, one finds the jitter in the second derivative causes wild

variations in Q (though the average behavior is consistent with that deduced below) . A

polynomial regression routine was used to fit smooth curves through the data and permit

smooth differentiation. The best fits (fraction of a degree) to the original digital

recorder data were obtained if the temperature history was broken into three or four segments and fitted piecemeal. However this invariably caused apparent discontinuities in Q versus temperature because the second derivatives, in particular were not continuous across

the juncture of any two fitted segments. Finally we used a one or two segment fit to the

whole range from about 75 to nearly 500°C. The fits are very good except at the ends of

the interval. The results are shown in Fig. 7 where the fitted temperature- time history

and the calculated Q vs. T are shown for the retarded and unretarded cases.

The pattern of Q vs. T is qualitatively the same for both the untreated and boric acid

treated insulation. That is, both show a low temperature endothermic region that fades out

to be replaced by a higher temperature exotherm which in turn tends to zero in the neighborhood of 500°C. These can be integrated to give the total heat effect as shown in Table II.

20
Table II Summary of Integrated Heat Effects Along 1-D

Particle Path Through Smolder Wave

Composition

Smolder

Velocity (cm/min)

3 Effect(cal/cm )

Integrated Heat

nominal unretarded 0.135 low temp, endo = -1.48

It It high temp, exo = +6.63

above with 25%

boric acid add-on 0.32 low temp, endo = -6.28

It II high temp, exo = +10.8

Direct comparison of this heat effect pattern with that in Fig. 6 is not appropriate; we

will relate the two below. The low temperature exotherm is initially due to water vaporization in both the retarded and unretarded cases. Note in Table II that the integrated endotherm is much larger in the retarded case. Because of the various approximations that went

into calculating these numbers they must be used with caution. However, it is plausible

that the endotherm be larger in the retarded case. The equilibrium moisture content of the

retarded insulation at room conditions is about 50% larger than the nominally unretarded

material. Furthermore, we noted previously that boric acid undergoes two successive dehydration reactions below 300°C. We estimate that these require nearly 200 calories per gram

of H^BO^ and the retarded insulation is 20% by weight boric acid. This makes the continuing

net endothermic ity up to about 200°C, seen in the upper half of Fig. 7, plausible in the

retarded insulation.

It should be noted that the moisture level at the leading edge of the smolder wave can

be well above the starting, room temperature equilibrium value. The wave drives water ahead

of it and it accumulates until the rate of deposition due to forward wave movement is

balanced by the outflow of water vapor due to buoyancy. There is a rather damp region at

the front of the wave that tends to stick to the temperature and gas flow probes.

Note that the endothermic ity of water vaporization and retardant degradation is

comparable in magnitude to the exothermic effects. This implies that these two heat sinks

have a substantial impact in slowing the rate of smolder spread. The greater endotherm in

the boric acid treated material could be a significant part of the reason for the slower

smolder spread in this case.

21
The net exotherms in the two cases (Table II) contain an unexpected feature—that for

the retarded case is larger than the unretarded case. This may be a false result of the

various approximations employed here; this point can be better studied in the simpler 1-D

smolder system which we have available. Note that the net heat effect (algebraic sum of

endo- and exothermic effects) is slightly larger for the unretarded material.

The net exotherms shown in Table II are not the total heat released by the particles

being followed. Inspection of the particle path shown in Fig. 4, for example, shows that

the temperature-time history for the retarded material in Fig. 7 ends when the remaining

organic weight fraction is still about 0.30 to 0.35; reference to Fig. 6 shows that this

corresponds approximately to the end of the first stage of heat release. This is roughly

true for the unretarded material as well. Part of the reason for the drop off in heat

release near 500°C is then kinetic—the next stage of oxidation is somewhat more difficult;

however, the major reason is almost certainly the lack of O

2 at that point in the smolder

wave. Inspection of Fig. 4, for example, shows that the particle path runs into the O

2 depleted zone as the particle approaches 500°C.

The portion of the smolder wave shown in Fig. 7 is a reverse smolder (or co-current)

part of the wave (see Appendix II for a description of this terminology) . The rest of the wave that follows, in which the char is oxidized by 0^ diffusing from above the insulation

layer, is a forward smolder (or counter-current) part of the wave. This dichotomy between

forward and reverse sections is inherent in the wave structure (which is in turn a result of

the 0^ supply modes, as discussed above) but there does not seem to be any fundamental reason

why the first and second stages of oxidation should fall completely within one section or the other.

It is of interest to cast the exothermic heat in Table II in a form more comparable

with the results of thermal analysis. By dividing by the bulk density of the insulation, we get the heat release in calories per gram of insulation. We use the initial bulk density

-2 3 of organic material (3.55*10 g/cm ); this ignores the slight shrinkage that occurs along

the particle path and refers the heat effect to the mass of starting material. When the DSC

results of Fig. 6 are integrated and similarly referred to the mass of starting material, we get about 1000 cal/g for the unretarded material and about 800 cal/g for the treated

material (first oxidation stage). On the other hand, the exotherms in Table II give about

190 cal/g for the unretarded material and 245 cal/g for the boric acid treated material.

Even allowing for the fact that the derived heat effects in Table II could be off by 2x due

to various approximations, it appears that less heat is released in the first oxidation

stage than is available according to the DSC results. This is not unexpected; the probable

explanation lies in an aspect of the cellulosic material's response to heating that is

discussed next.

22
As explained in ref. 1, the early stages of degradation of this material, like those of

many others, can be represented by two competing overall reactions

char + gases

char' + gases’

The upper pathway requires no 0

^ and is weakly endothermic; the lower pathway is exothermic.

The balance between the two is a function of the local O

2 concentration (and the rate of

heating). In the DSC results of Fig. 6, the O

2 level is constant at 21%. Fig. 3 and 4 show that it is decreasing from 21% to zero along the particle path we have been examining.

This shifts the above balance in favor of the non-oxidative pathway and cuts down the evolved

heat (mainly by removing fuel from the lower pathway)

.

Inspection of the O

2 and remaining weight fraction profiles in Fig. 3 a,nd 4 shows

another consequence of this 0

^ dependent competition. Particle paths lower in the layer

than the shallow one we have been considering undergo the first stage of degradation at

lower and lower levels of O

2 because they are farther in under the "overhanging" wave structure where O

2 diffuses at slower and slower rates (longer diffusion path) . We infer that

there is a continually decreasing contribution from first stage oxidative heat release with

increasing depth in the insulation layer. It is interesting to speculate that smolder in a low porosity material with the same chemical characteristics (e.g., solid wood or particle

board) would proceed with no practically significant contribution from first stage heat

release because O

2 could not penetrate the fuel from the front of the wave as it does here.

3.9 Summary of Wave Structure Findings

We have found a complex smolder wave structure that depends on a shifting balance of

heat release from both stages of oxidation of the cellulosic fuel. At the leading edge of

the smolder wave, which extends forward in the shallow part of the insulation layer, the

propagation process appears to be heavily dependent on the first stage of oxidation even

though it does not benefit from the full amount of heat available in this stage.

However, even here there is a dependency on heat fed forward from the char oxidation process;

this is indicated by the fact that a steady-state wave structure is quickly reached with

both stages of oxidation moving forward together. The further back one gets in the wave

(and the deeper in the fuel layer) the more the local heat release is dominated by char oxidation. Furthermore, the char oxidizing there (toward rear of wave) is predominantly a pyrolytic rather than oxidative char. (Differences between these chars have not yet been

explored .

)

23
Boric acid cuts the smolder wave speed by about a factor of two. It does this through

at least four effects. First it increases the equilibrium moisture content of the insulation. This, plus the second effect, the endothermicity of the boric acid dehydration process, puts a substantial heat sink on the front end of the smolder wave, slowing the rate at

which its heat generation can warm up new material. The two remaining effects are kinetic:

boric acid slows the rate of both the first and second oxidation stages. Its effect on the

first stage (presumably involving 0^ attack on polymers that still resemble the original

components of wood) is minimal. Its effect on the second oxidation stage (O

2 attack on the

char left by the first stage) is considerable but still insufficient to let heat losses \ quench the whole process. However, by suppressing the very end of the char oxidation stage,

the dangerous process of afterglow is prevented. This appears to preclude spontaneous transition to flaming combustion in buoyant air flow conditions (but not in forced air flow

conditions)

.

We now have further motivation to learn more about the nature of the first oxidation

stage (recall that it is also responsible for smolder initiation) . This is a current target

of study; the results should help in efforts to retard its occurrence. We also have a clear

picture of the buoyantly-driven smolder process which is the starting point for transition

to flaming as forced air flow is introduced. The complex structure of the wave implies that

it may have an equally complex response depending on the location and direction of the

forced air flow. We will explore some of these effects in our current work as well.

4. References

[1] Ohlerailler, T. and Rogers, F., "Smoldering Combustion Hazards of Thermal Insulation

Materials," Oak Ridge National Laboratory Report ORNL/Sub-7686/1, July 1980.

[2] Rogers, F. and Ohlemiller, T., "Cellulosic Insulation Material. I. Overall

Degradation Kinetics and Thermochemistry," Combust. Sci. Tech. , Vol. 24, No. 3 and 4,

1980.

[3] Ohlemiller, T. and Rogers, F., "Cellulosic Insulation Material. II. Effect of

Additives on Some Smolder Characteristics," Combust. Sci. Tech., Vol. 24, No. 3 and

4, 1980.

[4] Gross, D., "A Preliminary Study of the Fire Safety of Thermal Insulation for Use in

Attics or Enclosed Spaces in Residential Housing," National Bureau of Standards

Report NBSIR 78-1497, July 1978.

[5] Zicherman, J. and Fisher, F., "Fire Protection Problems Associated with Cellulose

Based Insulation Products," Society of Fire Protection Engineers Technology Report

78-7, May 1978.

[6] Yarborough, D. Yoo, K. and Konen, P., "Recessed Light Fixture Test Facility," Oak

Ridge National Laboratory Report ORNL/ Sub-7 504/3.

[7] Issen, L. , "Fire Performance of Loose Fill Cellulosic Insulation in Residential

Occupancies - A Progress Report," National Bureau of Standards Report NBSIR 80-2085,

August 1980.

24
[ 8 ] Schaffer, E., "On Smolder Initiation in Cellulosics Under Prolonged Low Level Heating,"

[9] Fire Technology .

McGuire, J., "Limiting Safe Surface Temperature of Combustible Materials," Fire

Technology , August 1969.

[10] U.S. Forest Products Laboratory, "Ignition and Charring Temperatures of Wood," Forest

Products Laboratory Report No. 1464 (revised), undated.

[11] Palmer, K. , "Smoldering Combustion in Dusts and Fibrous Materials," Combust. Flame I,

1957, p. 129.

[12] Eckert, E. and Drake, R., Analysis of Heat and Mass Transfer , McCraw Hill, New York

1972, p. 75 ff.

[13] Fristrom, R. and Westenberg, A., Flame Structure , McGraw-Hill, New York, 1965,

p. 190-191.

[14] Day, M. and Wiles, D., "Combustibility of Loose Fiber Fill Cellulose Insulation:

III. Evaluation of the Borax: Boric Acid: Aluminum Sulphate System," J. Cons. Prod.

Flamm. 6, June 1979, p. 105.

[15] Lucca, D., "A Comparative Experimental Study of Co-Current and Counter-Current Smolder

Propagation," MSE thesis. Mechanical and Aerospace Engineering Dept., Princeton

University, March, 1981.

[16] Fristrom, R. and Westenberg, A., op. cit, p. 85.

[17] Levy, A. and Weinberg, F., "Optical Flame Structure Studies: Examination of Reaction

Rate Laws in Lean Ethylene-Air Flames," Combust. Flame 3, 1959, p. 229.

[18] Tye, R. , "Heat Transmission in Cellulosic Fiber Insulation Materials," J. of Testing

and Evaluation 2, No. 3, p. 176.

[19] Gugan, K. , "Natural Smolder in Cigarettes" Combust. Flame 10 , June 1966, p. 161.

25
26

SIDE,

CROSS

SECTION

VIEW

Fig.

1

Attic

mock-up

adapted

for

smolder

propagation

studies.

Eighteen

centimeter

deep

in,

tion

layer

sits

in

trough

between

two

2x8

pine

boards.

Ignition

occurs

uniformly

left

end

and

propagation

achieves

a

steady

state

by

the

time

it

reaches

the

first

s

floating

thermocouples

(TC's).
m

0) 03

(U rO

>>

C r-i O iH

*rl Ct) JJ *H

rt -P

(U X!P OH

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CJ CO CCJ x: <u •M 6

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00

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cd

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(U X 03 P 0)U a d o <u P 3

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(34 cd X) X • H W -cf

CM

00

•H

(p

27
HEIGHT

ABOVE

BOTTOM

(cm)

HEIGHT

ABOVE

BOTTOM

(cm)

HEIGHT

ABOVE

BOTTOM

(cm)

RELATIVE DISTANCE (cm)

Fig. 3 Cross-sectional snapshot of steady-state smolder wave structure in

nominally unretarded insulation. Profiles of temperature, oxygen mole

fraction and fraction of remaining organic material superimposed on cross-section of insulation depth versus distance. Note that bottom

graph is shifted left about 15 cm relative to the upper two graphs.

28
HEIGHT

ABOVE

BOTTOM

|cm|

HEIGHT

ABOVE

BOTTOM

(cm)

HEIGHT

ABOVE

BOTTOM

(cm)

Fig. 4 Cross-sectional snapshot of steady-state smolder propagation in insula

tion with 25% boric acid add-on. Profiles of temperature, oxygen mole

fraction and fraction of remaining organic material superimposed on cross-section of insulation depth versus distance. Note that bottom

graph is shifted left about 20 cm relative to the two upper graphs.

29
u

<u Td

tH06M

bOC

•H

rt

0)

}-i 01

cn o

}-i ocu <uuS

4JO3

J-l 4-1 4J cn rt

CD <U

!2

C

cd

mo

03 ^ 03

4-1 ’H

cd 03 Oi 1—I ^ cd O c! 1—I cd G 6

•H

OU

G 4-1 03W3 Td

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4-3 <u

C/3

in

bO

•H

Pm I

30
(3K/1V3NI aSV313il 1V3H

(O CM

INBOtfad IM

31

Fig.

6

Thermal

analytical

results

for

unretarded

and

boric

acid-retarded

insulation.

Note

two

broad,

overlapping

stages

of

weight

loss

and

heat

release

with

both

compositions
TEMPERATURE

1

C)

TEMPERATURE

Td

RELATIVE TIME |min.|

+ 0.015

+ 0.010

+ 0.005

0

- 0.005

- 0.010

- 0.015

- 0.020

(/> E

o

o

Fig. 7 Heat effects (endothermic and exothermic) inferred from

temperature- time trace provided by shallow floating thermocouples. Temperature traces shown here are from the polynomial fits. Heat effects are inferred from these traces

via second equation on page 26.

32
APPENDIX I

Cellulosic Insulation Material

III. Effects of Heat Flow Geometry on Smolder Initiation

The variation with heat flow geometry of minimum heat source temperature

which causes smolder initiation in cellulosic insulation has been examined for

eight configurations. These configurations range from a wire-like source, to a flat planar source, to a corner formed by planar sources; they simulate such

heat sources as electrical conductors and light fixtures and span the geometry

range seen by insulation in practice. The ignition temperature for the same

20 cm deep insulation layer varies from 235°C (corner source) to 385°C (wire

source) . Other variables such as bulk density or the presence of smolder retardants have much less effect on ignition temperature. This behavior is rationalized in an approximate manner by simple heat generation/heat flow argximents; it

is predicted semi-quantitatively by a numerical model using reaction kinetics

derived from thermal analysis. A smolder ignitability test for cellulosic insulations is suggested that simulates exposures to recessed light fixtures. The

data here provide a correlation between results of such a test and ignition

temperatures in other configurations.

1. Introduction

Cellulosic loose-fill insulation is essentially re-cycled newsprint, re-ground back to

a fibrous mass and intermixed with combustion-retarding additives. The original newsprint

is for the most part a mechanically ground wood (balsam fir, black spruce, southern pine,

depending on the mill location) shredded to the point of separating the individual wood

fibers (Britt, 1964). When this physical form is recovered during insulation manufacture,

the result is a high specific surface area, low bulk density material with a high resistance

to buoyant air flow. These properties make this material a very good thermal insulator but

they also accentuate the tendency of cellulosic materials to undergo self-sustaining heterogeneous oxidation, i.e., smoldering combustion. The combustion-retarding additives (borax

for flaming resistance and boric acid for smolder resistance are most popular) are reasonably

successful at suppressing flaming reactions but markedly less so for smoldering reactions.

Despite the presence of these additives, when this type of insulation, due to improper

installation, is exposed to possible attic heat sources (recessed light fixtures, chimneys,

overloaded electrical wires or junction boxes), self-heating can lead to smolder ignition

and later to flaming. There have been several residential fires that apparently began in

scenarios like this (Zicherman and Fisher, 1978; Gross, 1978).

33
The likelihood of heterogeneous self-heating yielding smolder initiation is greatest

when the cellulosic insulation remains in contact with a heat source for an extended period

of time (comparable to or greater than the time to yield a steady-state temperature distribution around the source if the insulation was chemically inert; we call this the thermal

relaxation tim^. For source geometries found in attics, typical relaxation times are from

two to twelve hours. A heat source that persists for this period will effectively preheat

the insulation, making it possible for a lesser self-heating rate to cause ignition.

If the exothermicity (heat release per gram reacted) of the reaction is sufficient, Che

solid fuel need not be excessively consumed during this heat-up interval and ignition can progress as the heat losses from the reaction zone diminish. There will be some minimum

heat source temperature for a given material (dependent on geometric configuration) just

sufficient to produce ignition after this thermal relaxation interval; this is the quantity

of interest here and is referred to as the minimtim ignition temperature borderline.

Problems such as this have been extensively examined (at least for simple geometries)

in the context of thermal ignition theory (Gray and Lee, 1967; Merzhanov, 1971). For the

present class of problem, this theory reduces to finding the highest heat source temperature

for which ignition does not occur. When reactant consumption is negligible, non-ignition

results in a steady state where heat from the source and from the self -heating near the source flows continually outward through the insulation to ambient. Theoretically and

experimentally, one looks for the limit on this steady state and identifies it with the

minimum ignition temperature.

In previous work (Rogers and Ohlemiller, 1980), quantitative expressions were given for

* the two overall stages of oxidative heat release from a cellulosic insulation ; these rate

expressions were successfully applied (Ohlemiller and Rogers, 1980a) to the prediction of

minimum smolder ignition temperatures for the simplest heat flow geometry (insulation layer

on a flat, heated plate). In that work, it was shown that it is the first overall oxidative

degradation stage of the insulation (peaking at approximately 320°C in a DSC scan at 5°C/min)

that is responsible for smolder initiation. The thermal ignition model of Bowes ^ al . (Thomas and Bowes, 1961, Bowes and Townshend, 1962) was used, together with the first stage

rate expression, to make a reasonably good estimate of minimum ignition temperature

dependency on insulation layer thickness.

Here those results are extended to more complex and realistic insulation/heat source configurations. The variability of minimum ignition temperature with heat sources analogous

to attic sources is explored together with the ability of thermal ignition theory to predict

this variability. We want to establish the variability of ignition hazard with differing

*The kinetic parameters of these are given in the Nomenclature Table; note that the Z values

there contain ambient oxygen concentration and bulk fuel density implicitly.

34
source types and also provide some "calibration" data for a possible test method. In

addition, we look briefly at what the theory implies is the necessary change in oxidation

kinetics of the insulation in order to raise ignition temperatures substantially. Finally,

implications for a possible improved smolder ignition test method are examined.

2. Experimental Ignition Measurements and Observations

2.1 Apparatus

Eight heat source configurations were examined; Figure 1 shows schematic sectional

views of them; two were one-dimensional flat plate cases (configuration F: 7 cm insulation

layer; configuration G: 5 cm layer) as in Ohlemiller and Rogers (1980a) . The six other

heat sources in Figure 1 were tested in an "attic mock-up" so that their thermal environment

would resemble the real-world hazard situation; this should facilitate a reasonably quantitative carry over of the present data to such scenarios. The sources are essentially

Idealized versions of an electrical wire (H) , a chimney face (D) and recessed light fixtures

(A, B, C, E). A schematic top view of the attic mock-up in which they were mounted is shown

in Fig. 2. Briefly, this consists of a trough 30.5 cm wide formed by two 2 by 8's (pine

floor joists; actual dimensions 4.8 cm thick by 18.3 cm high). Outside of the trough, on

the other side of each 2x8, is a 7.6 cm wide x 20 cm deep layer of mineral wool insulation

to help simulate a large Insulated attic expanse. The floor of the trough is a 6 mm thick

transite board. The floor is not in direct thermal contact with the bench on which the

mock-up sits so that, like an attic floor, it loses heat only by convection (and radiation)

to an open air space below it. This attention to the thermal environment of the heat source

is made necessary by the time scale of the tests (many hours) ; the heat source has time to

become influenced by these surroundings, especially in all the directions that fall in the

plane of Figure 1. Heat flow in the direction perpendicular to the plane of Figure 1 (toward the 2 by 8's at the sides of the trough) is less important (but still present) for

most of the configurations tested since they are designed to be fairly two-dimensional in

nature. These sources are, in all cases, at least 25.5 cm wide and centered laterally in

the trough. Under these circumstances, only configuration B in Figure 1 is significantly

influenced by conduction of heat in the direction perpendicular to the plane of the figure

(ratios of both source dimensions to the depth of the layer are about unity in this case)

.

This was confirmed by probing the steady-state temperature field in the various configurations.

The heat sources themselves are designed to be very uniform in temperature. The planar

sources in Figure 1 are made as follows. A long, thin heating element, shaped into a raster-scan-like pattern (with less than 2.5 cm spacing) is sandwiched between two brass

plates. The side away from the sample material is covered with 2.5 cm of ceramic felt insulation. The design provides very good temperature uniformity (+ 1°C except within about

2 cm of the edges) while still giving rapid heat-up (12-14 minutes required to reach typical

set point temperatures). The heat source in configuration H is a 4.8 mm diameter copper

35
tube that spans the trough and concentrically encloses a heating element. Here, as with the

planar sources, the temperature is controlled at the desired value (+ 1°C) by a controller

with a sensing thermocouple (.020 cm wire) in firm contact near the center of the source

(two sources and two controllers in configuration A). Because of an unavoidably large power

to mass ratio in configuration H (only), the temperature oscillated (period about 5 minutes)

-f 2-3° C about the set point. All heat sources were given a thin coating of colloidal

graphite paint to provide a reproducible surface . emissivity (radiation into a porous fuel

bed can be significant)

.

In configurations C and E, where the planar heat source is elevated above the floor of

the trough, the source forms the top of a sheet metal box (0.08 cm steel). The box penetrates the floor of the trough and is airtight at and above floor level. It is completely

open to the air space below the mock-up but does act as a trap for heated air as noted later.

Additional probes were used in various tests to help assess insulation behavior.

Typically these were 0.025 cm diameter sheath thermocouples (chromel-alumel) inserted parallel to the plane of the heat source. Junctions were placed both about 0.64 cm away from the source center and about 3.8 cm away; both were on the line normal to the source plane and

passing through the source center. The inner thermocouple (s) helps assess the proximity of

a given test condition to the minimum source temperature border for ignition (very near the

ignition border, the gradient near the heat source goes to zero or may become negative).

The outer thermocouple monitors the gradient away from the source on a larger scale and is a good indicator of the achievement of a non-igniting steady state (at which point the test can be terminated). In addition, in tests with several of the configurations, a long

syringe needle (1 mm OD) was mounted parallel and adjacent to the near-source thermocouple (s)

.

Small (1/2 cc) gas samples were periodically withdrawn (rate < 1/2 cc per minute) and

analyzed for CO, CO^, 0^ and by gas chromatography. The progress of oxygen consumption

and pre- ignition reactions could be followed in this manner.

The insulation used in most of the tests was a commercial blend containing a proprietary

mix of retardants sufficient to permit passage of current insulation smolder and flaming

tests . Unretarded (or nearly unretarded) material from the same manufacturer was used in

two tests with configuration D. All material was thoroughly conditioned to a relative

humidity of 50 + Prior to testing it was ball-milled as described previously (Ohlemiller

and Rogers, 1980a) in order to fluff it and permit bulk densities typical of commercial

application.

The CPSC Interim Standards for Cellulose Insulation (Federal Register, July 6, 1979)

require a critical radiant flux for sustained surface flaming of at least 0.12 w/cm^;

cigarette ignition of an insulation bed of a specific size must produce less than a 15%

weight loss.

36
The reader should bear in mind that the cellulosic insulation materials currently on the market vary in starting material and retardant blends. The results here apply to only

one manufacturers product but they are believed to be approximately valid for many others.

2.2 Procedure

The insulation was weighed out in four separate segments, each sufficient to provide

1/4 of the total layer depth (20.3 cm in all cases, i.e., the mock-up was always filled to

the top). Each segment was carefully distributed to minimize bulk density variations (a

potentially significant source of scatter in such tests even though average bulk density has

a fairly weak effect; see below). When the desired bulk density was the nominal value of

-2 3 4.4 X 10 g/cm , each layer required minimal compression (5 mm or less) to yield its 1/4

contribution to the total layer depth. When the target bulk density was 20% higher (5.3 x

-2 3 10 g/cm ), significantly greater compression was required.

In no case was the sample volume used sufficient to fill the full length of the mock-up

(91 cm); this would have taxed the sample preparation facilities excessively. Instead, only

the space around the heat source and out for a distance of at least 15 cm was occupied by

the sample. The remaining space was segmented-off by means of a steel wire screen (16 mesh)

3 and it was packed with mineral wool at a bulk density of approximately 0.10 g/cm . Flow

tests showed this packing density to provide a reasonable match to the sample in terms of

permeability. Handbook values of specific heat and thermal conductivity (Perry, 1950) imply

reasonable matches in these properties. In any event, the sample temperature at the

interface with mineral wool never exceeded 10-15°C above ambient, implying that this material

was well away from the more crucial parts of the temperature distribution.

For the two one-dimensional flat plate cases reported below, the apparatus was that

used previously (Ohlemiller and Rogers, 1980a). The insulation bulk density was the same nominal value as in the other tests but the thinner beds required packing in two layers

only

.

The tests themselves are necessarily trial and error attempts to locate the lowest source temperature yielding ignition regardless of exposure length. Several tests on successive samples are thus required to obtain the desired data point. In each test, after

packing, the source is brought quickly up to the desired temperature (12-14 minutes) and

held there. The thermal relaxation time of the sample is more difficult to analyze than in

the one-dimensional case previously treated (Ohlemiller and Rogers, 1980a) because of the

rather complex geometries. A range into which the relaxation time should fall can be estimated from the key dimensions of the particular configuration but the range is typically

2 quite wide. We use the simple 1-D formula given in the above reference, i.e., t = 0.4i /a,

where Z is the pertinent length and a is the thermal diffusivity of the insulation. For

example, from the two depths involved in configuration C, one estimates a thermal relaxation

37
time in the range from 2.2 to 9 hours; actual ignitions occurred in this case after about

3 hours.

The transient behavior of the sample was followed by the thermocouple and, in some

cases, the gas sampling probes. The probes themselves might somewhat alter the local

behavior but they could influence only a small fraction of the igniting material; there was

no indication that they altered the overall behavior. The use of the probes was purely an

optional exercise aimed at a better understanding of the ignition event. In principle, one could simply walk away from such a test and come back later to learn the result; ignitioii or no ignition. The probes were consistently useful, however, in indicating when no further

waiting was necessary for a non- ignition case. ^Then, in the neighborhood of the heat source,

temperature, CO and CO^ levels peaked and began a slow decline, the danger of ignition was

past. A few tests were continued up to 27 hours because there has been some concern that

cellulosic ignition is coupled uniquely to char build-up (Schaffer, 1979) . In no case did

this extended exposure have any noticeable effect or lead to any new exotherm. It should be

noted, however, that the present results cannot answer questions about long term changes in

ignitability resulting from months or years of exposure to lower temperatures.

The results of a given non-ignition test were used to judge what heat source temperature

to use in the next test (with a fresh sample) . If the thermocouple (s) near the source peaked within + 2°C of the source temperature, the next test temperature was about 5°C

higher. The goal was to determine the borderline (minimum) ignition temperature to within

about 5°C for each configuration. Generally, the position of the borderline was confirmed

by performing two tests just above it and two just below it. It probably is futile to

attempt to pin down the borderline to much better than 5°C since material and test condition

variability cause approximately this much scatter in results.

2.3 Observed Transient Behavior and Reactant Depletion in Tests

The qualitative trends noted with the gas sampling and temperature probes near the

heat source are similar to those shown previously for the one-dimensional geometry

(Ohlemiller and Rogers, 1980a). For conditions near the ignition borderline, the variations

of temperature and oxygen concentration with time are not greatly different for ignition and

non-ignition cases until late in the test. For non-ignition very near the borderline, the

temperature near the heat source (i.e., usually measured 6 mm away) gradually rises to 2-3°C

above the source temperature and then subsides back to a few degrees less than the source,

eventually approaching a steady state value. For the same case, the 0^ level drops to some rather large fraction of the ambient level, then slowly rises again as the local temperature

decreases slightly. Even in the case of configuration B with a high fuel bed density

-2 3 (5.3 X 10 g/cm ), where one would expect the greatest O

2 depletion, the O

2 level stays at

or above 80% of the ambient air level.

38
Behavior during a test where the source temperature is just above the borderline for

ignition is similar for an extended period of time. However, the near-source temperature

will continue rising gradually, typically accelerating appreciably when it is about 5°C

above the source (for deeply imbedded sources such as in B, the actual runaway may only be seen in the more distant thermocouple) . Only when the rate of temperature rise accelerates

does the 0^ level near the source drop below about 70% of the ambient air level. Subsequent

spread of the smolder zone is probably 02-liniited; it can require up to an additional 3 hours

(configuration B) to reach the visible surface of the insulation bed, where it is apparent

as a growing blackened zone that shrinks somewhat.

The effects of oxygen depletion on the ignition event are minimal since the depletion

itself is minimal. This is particularly apparent in light of the weak dependency of the

ignition reactions on oxygen concentration (0.43 power) (Rogers and Ohlemiller, 1980).

Consumption during ignition of the reactant for the first overall oxidative degradation

stage of the fuel cannot readily be measured directly. An estimate of depletion was made

previously (Ohlemiller and Rogers, 1980a) using reaction kinetics derived from thermal

analysis. It was shown there that depletion is less with deeper fuel beds such as used here

(with their lower ignition temperatures) but not negligible. One can estimate that solid

reactant depletion will slow the ignition reactions by something less than a factor of two;

this in turn translates into something less than a 10°C increase in ignition temperature.

(The actual effect may be half or less of this upper bound estimate)

.

Configuration H (hot wire-like source) is a definite exception with regard to the

magnitude of solid reactant depletion. The ignition temperature (see below) with this small source was so high that there were indications of substantial fuel consumption near the source.

With the exception of configuration H, we note that reactant consumption effects are

rather small. Furthermore, the behavior of a borderline ignition case is quite similar to a borderline non-ignition case until the former finally runs away. The experimental picture

thus reasonably resembles the physical picture implicit in the thermal ignition theoiry model

of Bowes and Townshend (1962). That is, the last non-ignition state (highest source temperature that does not produce an accelerated runaway to ignition) results in a nearly

steady-state temperature profile where heat flows from the source (and from exothermic

reactions), through the insulation to ambient. It is nearly steady because solid reactant

consumption takes substantially longer than does the primary thermal relaxation process due

to heat flow from the source. (It is reactant consumption that causes the peaking out and

reversal of O

2 and temperature trends in the non-ignition cases noted above; this implies

that reactant consumption time is not drastically less than the thermal relaxation time.)

With these considerations in mind, we adapt the Bowes and Townshend model to the present

geometries below. First, however, the measured ignition data will be given.

39
2.4 Measured Minimum Ignition Temperatures

The results for eight experimental configurations are summarized in Figure 3. Two of

these are one-dimensional flat plate configurations (two different layer depths) similar to

those reported previously (Ohlemiller and Rogers, 1980b). They are included so as to put

such results into the context of the present more realistic configurations.

It is apparent that the configurations vary widely in their smolder ignition tendency

as measured by minimum ignition temperature; there is a 150°C range between the corner (A)

and hot wire (H) configurations. As will be seen below, the important factor that varies

across this range of configurations, causing their ignition temperatures to differ, is the

rate of heat loss and heat generation per unit volume from the hottest layers of insulation.

The range covered here probably encompasses most practical ignition hazard situations.

On the basis of relative minimum ignition temperatures it appears that an overheated

electrical wire is a substantially lesser smolder ignition hazard than the other larger area

sources. (There are, of course, other elements that may affect hazard assessment.) It is

probable that the overload needed in a copper conductor to produce the observed ignition

temperature would open a properly rated circuit breaker before it could be achieved (Gross,

1978). On the other hand, the work of Yarborough ^ (1979) and of Gross (1978) clearly

demonstrates that recessed light fixtures covered with Insulation and containing light

bulbs in excess of their rated power can achieve peak temperatures in the range seen for

configurations A to E; some of these cases were found to ignite.

Yarborough found that the peak temperatures achieved in insulated recessed light

fixtures correlated reasonably well with the electrical power per unit ceiling-exposed

area. This area on the ceiling side of the fixture is where 85-90% of the electrical heat

is dissipated; only 10-15% goes through the insulation surrounding the fixture in the attic.

This led Yarborough to the reasonable conclusion that increasing the insulation thickness

on top of the light fixture would not substantially increase its peak temperature. (The

greatest thickness on top of the fixture in those studies was 7.6 cm). Even so, the present

work predicts an increased hazard from increased thickness; compare configurations F and G

in Figure 3, for example.

Note that configuration B has a lower ignition temperature than configuration C. The

source dimensions are the same in both cases; only the insulation depth varies.

Configuration B can be viewed in two ways: (1) as a simulation of a recessed light fixture

of the same dimensions as C that is covered on top with twice the thickness of insulation

or, (2) as a simulation of a fixture mounted flush on the underside of the attic ceiling.

As a simulation of role (1) one finds, as expected, that B has a lower minimum ignition

temperature than C. The role (2) simulation is deceptive, however. It is wrong to conclude

that the results in Fig. 3 imply that the fixture mounted on the underside of the ceiling

40
is more dangerous than the same size fixture recessed as in C. The fixture on the underside

of the ceiling will dissipate virtually all of its electrical heat to the room below the

attic and, being essentially uninsulated, would require much more electrical power input to achieve the minimum ignition temperature shown in Fig. 3. (Note that role (2) also neglects

the heat flow resistance of the ceiling itself and, in effect says that the indicated area

of the upper surface of the ceiling itself is at the temperature of source B.) One should

bear in mind that the electrical power input to an immersed heat source is a separate factor

in hazard assessment that could not be Included here.

Configuration A is probably a worst case, not likely to be found in practice. It is

worth noting, however, that the complex shape of some light fixtures creates concave corners

that could increase their hazard somewhat.

2.5 Bulk Density and Retardant Effects

These factors were examined only briefly here to see if their effects were consistent

with those seen previously (Ohlemiller and Rogers, 1980a and Ohlemiller and Rogers, 1980b).

A bulk density increase of 50% was previously found to lower the minimum ignition temperature

by 10° C or less for a 7 cm layer of the particular insulation used in the one-dimensional

-2 3 flat plate configuration. Here the bulk density was increased by 20% (to 5.3 x 10 g/cm ) in configuration D (vertical flat plate heat source) and in configuration B (same heat source flush with bottom floor of mock-up); the effect in both cases was a lowering of the

ignition temperature by 5°C or less.

Bulk density changes affect several parameters in the ignition problem: thermal

conductivity of the insulation, bed resistance to air Inflow and solid fuel concentration.

Changes in thermal conductivity are quite small (Brookhaven National Laboratory, 1978) and

are such as to counter the effects of changes in fuel concentration. This last variable is

shown by the model below or that of Bowes and Townshend to have a rather weak effect on ignition temperature. The effect of bulk density on bed flow resistance to air is large

but, as was seen above, oxygen depletion is not a serious problem even in the higher density

tests. This type of test is thus fairly insensitive to bulk density variations although

caution is needed as noted below. This is reassuring since this material is rather difficult

to pack in a highly reproducible fashion.

Unretarded insulation was examined only in two tests with configuration D (vertical

3 flat plate source). The bulk density was 0.042 g/cm so that the actual concentration of

solid fuel was somewhat higher here than in the retarded cases (where roughly 25% of the

bulk density is accounted for by the retardants). The results, though sketchy, confirm the

rather small change in ignition temperature induced by present-day retardants (Ohlemiller

and Rogers, 1980a). Here the unretarded borderline lies only 15-20°C below that for the

commercially retarded material as it did in the previous tests. The implications of this

41
for the insulation ignition problem were discussed previously (Ohlemiller and Rogers,

1980a).

2.6 Material Variability

All material used in these tests was obtained from a single manufacturer and was made

on the same day. Nevertheless, when an attempt was made to reproduce the original ignition

borderline for configuration B using insulation from a second bag, the new borderline was

found to be reproducibly 10°C higher (see Table II below). This prompted a series of '

further tests to discern if changes in experimental conditions (fume hood cross-flow rate;

mock-up mechanical support, etc.) had caused the change; none of these had any effect. The new bag of insulation was found, however, to have a 30% higher equilibrium moisture content

(~13% vs 10%) and a 15% lower ash content (1 hour at 650°C; 17% vs 20%). Clearly there was

a change in retardant mix and possibly loading level; the paper stock may also have varied.

The nature of the chemical change has not been identified but these results provide a measure

of variability in these materials on any given day.

3. MODELING OF IGNITION BEHAVIOR

3.1 Phenomenological Model of Configuration Effects on Ignition

The above results coupled with our previous work indicate that for cellulosic

insulation the heat flow configuration has more impact on smolder ignition temperature than

any other variable. It is instructive to formulate some crude models of the heat flow/heat

generation for a few cases to see why this is so. In a subsequent section we will solve the

problem more exactly but the simple heat balance effects that influence ignition temperature

are somewhat obscured. The goal here is limited to demonstrating that the shifting balance

of heat generation rate versus heat loss rate leads quite naturally to an ignition temperature that varies with geometry. The models posed here for three geometries are intended

only to demonstrate the source of this variation; they are not intended as formulae for

routine use in calculating ignition temperatures since substantial approximations are made

at certain points. The reader who needs no convincing as to the source of the ignition

temperature variability may wish to proceed to the next section.

When ignition does not occur for a given source temperature, it is because the heat

from the source and from the reacting insulation near the source is lost to the ambient

surroundings. Thus we have in general;

Here the first integral is the total inflow of heat to the insulation as a result of

conduction from the heated source to the cooler elements of adjacent insulation; the second

Integral sums the heat release contributions from the entire volume of the insulation (only

( 1 )

42
that near the heat source is significant, as will be seen) ; the third integral sums all the

losses from the surfaces of the insulation where it is hotter than the adjacent material

(ambient surroundings) . Because of the Arrhenius temperature dependence of the chemical

heat generation term, (Rogers and Ohlemiller, 1980) it and its integral are negligible

until the heat source temperature approaches the borderline for ignition. As the borderline

is approached there is an abrupt change in the relative importance of the two terms on the

left hand side of Equation (1); the second term grows and the first shrinks. This latter

effect occurs because the reaction, again due to its Arrhenius temperature dependence,

proceeds fastest right where the insulation is hottest, i.e., at the interface with the heat

source. Heat evolution there raises the local temperature in the insulation just enough to

drive the temperature gradient at the interface approximately to zero (Bowes and Townshend,

1962) just as the ignition borderline is reached. (This is only locally true for certain

geometries; see below). Further increase in the source temperature causes the chemical

reaction term to grow so much that the steady state balance of Equation (1) is no longer

possible and ignition occurs. We will use the idea that, at the borderline, the second and

third terms of Equation (1) are essentially in balance; this becomes our criterion for

achievement of the minimum ignition temperature. This will be done for three geometries by

estimation of the values of these two terms for each case. It will be seen that different

geometries change the minimum ignition temperature chiefly because they imply changes in the

volume of material generating heat relative to the surface area losing heat.

First consider a one-dimensional flat geometry such as that in configurations F and G. Define an effective reaction layer depth, £ , adjacent to the hot plate as follows

R

Unit area of the hot plate is implicit here so the volumetric integration becomes an integration over the insulation layer depth, £. Subscript 1 means that we expect the first

overall oxidation stage of the cellulosic insulation (recall that two overall stages were

found, as described earlier) to be responsible for ignition in this case; this was shown to

be true previously (Ohlemiller and Rogers, 1980a). The kinetic parameters (Q^^ the heat

release, the pre-exponential factor and E^^, the effective activation energy) obtained

from thermal analysis were shown in this reference to give good estimates of ignition

temperatures when used in the Bowes and Townshend model. Here we are constructing a more

approximate model to estimate the effect of changing geometry. We estimate the value of

the middle integral above by inserting the linear, steady state temperature profile for the

non-reactive case (Carslaw and Jaeger, 1959a). In view of the above discussion about the

gradient at the interface with the source, this is clearly approximate; it works best a few

degrees below the ignition border rather than at it. When this is done, we can solve for

£ after performing the integration (the Frank-Kamenetskii approximation to the exponential

is used as in Bowes and Townshend, 1962).

exp ( 2 )

43
( 3 )

where ^ (h/A) ^ (hA/A) )

is the negative of the slope of the linear temperature profile. The other symbols can be

found in the Nomenclature Table. When evaluated, £ , which measures the reacting volume,

K

is found to be a small fraction of the layer depth (see below)

.

Now it is convenient to focus attention on this reaction layer and apply Equation (1)

to it rather than to the whole insulation bed. The same statements made above about the

balance of terms at the ignition border apply here as well. Therefore we still estimate the

ignition temperature at the border by equating the chemical heat generation in the reaction

layer to the rate of heat loss from the layer. Thus, for this one-dimensional geometry we

get

exp(-Ei/RTp) = a(j^) (5)

Here in estimating the heat loss from the reaction zone, we used ambient temperature T^

rather than the more accurate outer surface temperature available from the inert solution;

for the present purpose of comparing geometric effects under one set of conditions this is

adequate. Equation (5) can be solved, by trial and error for example, to obtain the

predicted border ignition temperature T^ for the one-dimensional flat plate case; we report

the results below.

Consider next the case of a hot wire at mid-depth in an insulation layer

(configuration H) . Here we do not have an exact temperature distribution for the inert case. For present purposes, it is sufficient to use the solution for a concentric cylindrical bed

of insulation surrounding the wire (Carslaw and Jaeger, 1959b); we take the radius of the

cylinder as the half-depth of the bed. Use of the radial geometry means we can only roughly

approximate configuration H; all that is needed here is an approximation to the reacting

volume very near the wire, obtained by linearizing the temperature distribution there.

(Alternatively, we could treat the concentric cylindrical bed instead, as an example of an extreme convex heat source geometry; the numerical estimates would be very similar.) The

linearized profile is

T(r) = T

P

(h£_/X)(T - TJ

£ p A

1 + (hSi^/X) in (r/£s 1 ) ( 6)

The slope here is about an order of magnitude steeper than in the one-dimensional flat case due to the higher surface to volume ratio here; this implies a substantially thinner reaction

zone. Furthermore, the radial geometry near the source implies a higher heat loss rate per

unit of reaction volume. These two changes will be found to imply a substantially higher

44
ignition temperature than in the flat plate case. Proceeding in a manner exactly analogous

to the one-dimensional case, we find

‘r =

‘s

1 [b’(E/RT'

^[l + b'(E/HT2)]j

1/2

where b' (h£_/X)(T - T.

I £ ^ + (h£2/^)

(7)

( 8 )

and the ignition condition, equating heat produced and lost in the thin reaction zone around

the wire is:

(1/2)QZ(£^^ - £g) exp (-E/RT^) = X (2Tr£p

(

^^ 75 /) (9)

Here £' is the outer radius of the cylindrical reaction zone so that the actual reaction

R

layer is (£' - £ ); £„ is the radius of the hot wire. The heat loss estimate on the right

is rough; the denominator is the mean of the two pertinent length scales £ and £

2

. Again,

the predicted minimum ignition temperature can be found by solving for T^. This time we

have not specified which reaction stage is pertinent since we will look at both below.

Finally, consider the concave corner, configuration A. Here there is a new complication.

The reaction zone does not extend equally over the interface between the bed and the source. The near-corner region is concave and has a higher surface to volume ratio, now working

oppositely to the hot wire situation, so that there is a favored zone here where reaction

occurs most extensively. It is this zone to which we want to apply our ignition criterion.

By means of flux plotting (Eckert and Drake, 1972) or by use of the exact numerical solutions

described below, one can show that the igniting volume is nearly isothermal near the border

and is roughly triangular (with the converging hot walls forming two sides of the triangle)

.

The isothermal lines in the corner are roughly twice as far from the nearest point on the

heat source as they are when one considers the same isotherm well away from the corner

region. In the latter case, the isotherms approach the spacing (from the source or from

each other) they would have in the planar 1-D case. The reaction volume of interest here is

thus roughly given by the following

V = (1/2) (4/2 Z^)(2/2 £^) = 8 £^ ^ (10)

where £ is the one-dimensional reaction layer thickness as expressed by Equations (3) and

(4); the actual value of £ is greater here because the pertinent layer depth in Equation (4) R

is greater. '

Estimating the heat loss rate from this volume is equally difficult in this case. Again one must resort to flux plotting or to the exact numerical solution to make a reasonable estimate. The final result, analogous to Equation (5) or (9), is:

45
8£^ exp(-E^/RT ) = A (4/2 ( 11 ) (\ - "a)

r 2 ^

2.1/2

(ai + a^)

Here a^ and are the lengths of the heated walls and we have found that the effective

length for conductive loss from the favored reaction volume is approximately equal in length

to a diagonal joining the ends of these walls. (The heat loss flux vector is roughly

oriented out from the corner along the opposite diagonal.) Equations (11) and (5) are

similar; comparing them, one finds that the igniting volume suffers a substantially lesser

loss rate in the case of the corner configuration than in the case of the flat plate source.

When numbers are put into the above equations, the results in Table I are obtained.

Table I Estimated Dependence of Minimum Ignition Temperature on Configuration

Conf iguration

Reaction Layer Ignition Temperature

Thickness (cm) Estimate (°C)

1-D flat plate, 7 cm insul.

1st stage kinetics

0.4 292

corner w. 20 cm insul.

1st stage kinetics

1.2 255

0.24 cm radius wire, 20 cm insul.

1st stage kinetics

0.05 350

0.24 cm radius wire, 20 cm insul.

2nd stage kinetics

0.05 480

The approximations are least in the one-dimensional flat plate case and here our estimate applied to this specific case is only 7°C higher than the experimental result in

Figure 3. The approximations are clearly less good in the other cases but we have in fact

captured the trend of change with configuration reasonably well. The goal here was not to

make quantitative predictions of ignition behavior but rather to illustrate that the variable

ignition temperature behavior can be rationalized on the basis of simple heat balance

arguments. We have demonstrated that geometry has a big impact on reaction volume (and thus

total heat generation rate at a given temperature) and on heat losses from that volume;

these differences are made up by changes in heat source temperature that finally destroy the

heat balance and cause ignition.

Configuration H was examined for both stages of decomposition because the ignition

temperature implied by the first stage is so high. It was shown previously (Ohlemiller and

Rogers, 1980a) that in circumstances where the ignition temperature is high there can be

extensive consumption of the first stage reactant in the early transient part of an ignition

test. Here the ignition temperature is so high that there could be total consumption of

first stage reactant within roughly the thickness shown in Table I during the transient

46
(without causing ignition) . Then as the transient gradient near the source diminishes and

conditions for ignition become more favorable, ignition will have to be as a result of either

further stage one reaction outside the consumed annular region or stage two reactions within

the annular region (or both) . Neither case is well-described by the present models but the

results in Table I do imply minimal assistance from reaction stage 2 (char oxidation)

;

comparing the second stage prediction with the actual value in Figure 3 shows that the former

is 100°C higher. It is thus more likely that the first stage produces ignition but requires

a higher source temperature to overcome the insulating effect of the depleted layer. The

results below provide some further support for this.

3.2 Numerical Model of Ignition Behavior

The above arguments are informative but too imprecise for further purposes here. We

want to make a more quantitative test of the ability of thermal ignition theory to predict

the results in Figure 2 and also to put the theory to use in suggesting test configurations.

The model here simply generalizes that of Bowes and Townshend to the various

two-dimensional configurations of Figure 1. Thus we solve the steady-state two-dimensional

heat conduction problem incorporating the possibility of a zeroth order, Arrhenius-type

exothermic chemical reaction within the insulation. We have noted above that oxygen

depletion is minimal and can be neglected without significant error. Except in the case of

configuration H, fuel depletion is also small. Our previous results (Ohlemiller and Rogers,

1980a) imply that for configurations A to G, reactant depletion at most results in an upward

shift of ignition temperature by 5 to 10°C and probably less (shift is greatest for highest

ignition temperature). As in the experiments, we are interested in finding the upper limit on on the steady-state; this is the minimum ignition border.

The model here includes only conductive heat transfer through the insulation material

itself. It was noted briefly in a previous section that radiation can become a significant

heat transfer mode in porous materials and it is not zero here. However, an estimate of the

infrared radiative path-length for this particular material (few tenths of a millimeter,

ref. 27) implies its inclusion would be a minor correction here. Buoyant convection of air

through the permeable fuel bed is more significant in some cases. One can estimate the

order of magnitude of the flow velocity from a force balance of drag versus buoyancy

(ref. 16); the result is a few tenths of a millimeter per second. This is liable to be

approached only in configurations with extended hot vertical surfaces such as A and D. In

such a case, the convective flux near the source can amount to 30% of the conductive flux

from the source but they are at 90° with respect to each other. In view of the considerable

complexity that inclusion of internal buoyant convection would entail, coupled with the

results below that Indicate changes in effective thermal conductivity have a very minor

effect on ignition temperature, there is little Incentive to include this convective heat

transfer mode.

47
In contrast to the rough model of the previous section, here we treat the entire

insulation bed, allowing the possibility of reaction throughout; the boundary conditions

dictate the locus and extent of reaction. The governing equation is thus:

The reaction parameters Q, Z, and E are derived from thermal analysis as indicated above.

The parameter Z is a composite number containing oxygen and fuel concentrations (bulk

density), both of which are taken here as constant. The thermal conductivity, X, is constant

here. This parameter was initially taken as temperature-dependent using the data of Tye

(1974). However, this proved to have little effect (<3°C) on the minimum ignition border

and was costly in terms of computer memory; all cases reported here are for X constant.

Solutions to this steady-state equation are sought subject to the types of boundary conditions implicit in the various configurations of Figure 1. Thus, in general, there can be up

to six sides on the domain of interest. For example, configurations C and E are treated by

solving one symmetrical half of the domain, splitting it along the mid-line of the heat

source; the half-domain then has six sides with two due to the indentation into the fuel

bed by the heat source. Each side, or portion of a side, is subject to one of the following

types of conditions:

Here h can be zero if the border in question is adiabatic (such as the midline just

mentioned) ; it can be due to the usual convective heat transfer boundary layer for a surface

with the appropriate orientation. Finally, in the case of a surface corresponding to the

bottom of the attic mock-up, it can be the series equivalent due to the transite board and

the convective boundary layer below it.

The assignment of the appropriate boundary condition to a border segment is, for the

most part, obvious from Figure 1 with two exceptions. Since the domain of solution had to

be truncated at a finite insulation bed length, some compromise was necessary, just as in

the experiments. The typical bed length used was 40 cm and an adiabatic boundary condition

was imposed at the cold end. With this length, this condition, which is a simplification,

has little effect. The other non-obvious boundary condition is pertinent to configurations C

and E where one part of the indented boundary corresponds to the surface of the sheet metal

box holding the heat source above the floor of the mock-up at mid-bed depth. During the

experiments with these configurations, measurements were made of the temperature of the air

trapped within the sheet metal box; a mean value of 147“C was used as the ambient temperature

together with Equation (14) for this boundary.

( 12 )

T = T = constant

P

(13) or

(14)

48
The derivatives on the left side of Equation (12) were approximated by central

differences; the exponential term on the right side was quasi-linearized to facilitate

solution. The iterative process that quasi-linearization implies was combined with the

iteration inherent in the solution algorithms. Because we are looking for an inherent

stability limit on the solution of Equation (12) and this could be confused with numerical

instabilities, two separate solution algorithms were employed. All of the cases were solved

using successive overrelaxation, called SOR (Ames, 1965); a few cases were also solved as a check using an alternating direction implicit method, called ADI (Carnahan, et al., 1969).

The SOR program was first tested against an analytical solution derived for a chemically

inert case similar to (though somewhat less complex than) configuration A. This defined the

iteration convergence tolerance necessary to assure temperature distributions within 1°C or less of the exact solution. The program was then checked against the model of Bowes and

Townshend as to its predictions of the minimum ignition border for one-dimensional flat

plate configurations (e.g., G and F in Figure 2). The SOR program gave results 2 or 3°C

higher, depending on layer thickness; recall that the present solution does not use the

Frank-Kamenetskii approximation to the Arrhenius factor whereas Bowes and Townshend did.

An ADI solution method is generally considered to be faster and more efficient than the

SOR method (McDonald, 1979). For the rather unique requirements of the present problem,

this was not found to be the case. Here we are interested in solutions very near where the

reaction term is changing rapidly and becoming dominant. This forced the use of large values

of the ADI iteration parameter, corresponding to small values of a pseudo-time step, to avoid

numerical instability. The ADI method was significantly more difficult and expensive to use

in this case. It was used only to check the borderline for configurations E, H and a 10 cm analog of configuration F. The results were the same as with the SOR method except for an

analog to configuration H where ADI gave a result 2-3 °C higher; this latter configuration

posed special problems as discussed below. It is interesting to note that this method,

which essentially integrates the time dependent problem, does giye virtually the same ignition border as that found from the upper limit on the steady state solution.

An overall energy balance check was made on each solution and typically was satisfied

to 10% or better. This could have been improved by means of smaller space step sizes or further iterations but this was found to be unnecessary; highly accurate temperature distributions were not the objective, rather the ignition borderline was the goal and it was not

significantly affected by this level of solution inaccuracy.

It is interesting to note that the numerical problem was rather forgiving with regard

to space step size. The reaction layer thicknesses in Table I somewhat underestimate the

scale of the extended reaction zone but they are indicative of its small magnitude. The

step sizes used typically only provided a few points across the width of the reaction zone. Nevertheless, the minimum ignition border shifts only 1-2°C when the step size is halved.

49
3.3 Comparison of Numerical Model with Experimeitt

The temperature profiles for the non- igniting cases are generally consistent with what

one would expect from the preceding arguments about heat flow and heat generation. Heat

generation increases rapidly as one approaches the ignition border and is strongly sensitive

to small changes in the local temperature. In a case such as configuration F or G, heat

evolution occurs in a layer very near the source that extends equally in the lateral direction; in cases B, C, E and D, it is concentrated near the center of the source because of

the lateral decrease in temperature near the source edges. As expected, in configuration A,

heat generation is concentrated in the hot corner.

Here we are interested not so much in the temperature profiles as in the upper limit

solutions to Equation (12), their variation with configuration and how that variation

compares with that seen experimentally. The results are shown in Table II. (Numerical

values of model parameters used can be found in the Nomenclature Table.) Except as noted

for configuration H, first stage kinetic parameters were used in the model calculations.

Table II

Comparison of Numerical Model and Experimental Results

for Minimum Ignition Temperature

Configuration T„ihCC) .Model

A: 90° corner w. two plates 234-235 243

B: 20 cm plate, 20 cm deep 255-256

(266)*

265

C: 20 cm plate, 10 cm deep 272 272

D; 20 cm plate, vertical 276-277 269

E: 10 cm plate, 10 cm deep 282 284

F: 1-D plate, 7 cm layer 285 284

G: 1-D plate, 5 cm layer 298-301 297

H: 0.48 cm dia. wire, mid-depth 380-385 /320\ ** \435y

due to material variability; see text.

**upper value for first stage kinetic parameters, lower value for second stage; both are somewhat different heat source than in expt; see text.

In general, the ability of the model to account for both the magnitude and the

variation of the minimum ignition temperature is quite good. This is a confirmation of the

utility of a thermal ignition theory model and of kinetic data derived from thermal analysis.

The exact quantitative agreement for certain cases must be viewed as somewhat fortuitous, a consequence of mutual cancellation of small errors. In view of several approximations—neglect

50
of solid fuel consumption, constant thermal properties, the less than perfect match between

the kinetic rate law and the thermal analysis data—one should not really expect less than a 5-10°C differential between model and experiment.

Configuration H which mimics an overloaded electrical conductor, clearly is not as well

modeled as the other cases. The agreement is poor for at least two basic reasons. First,

the actual configuration used in the model has a different geometry than the hot wire. A

small round source that produces very high temperature gradients is not easily handled

without refinements to the present computer program which has only fixed, equal space step

sizes. Since this is a lower hazard configuration, no attempt was made to model it exactly.

To get reasonably small space steps, the model focuses on one quarter of the insulation bed

as representative, ignoring the differences in heat loss between top and bottom fuel bed

surfaces. The model heat source is at one corner of the quarter and is two perpendicular

0.40 cm segments at fixed temperature enveloping the corner. The source is thus convex

rather than concave (the latter posed numerical stability problems) and is significantly

larger than the experimental source; this lowers the minimum ignition temperature for either

set of reaction kinetics. The 0.2 cm space step used is comparable to the reaction layer

depth; this very poor resolution probably accounts for the somewhat differing predictions of

the two solution algorithms.

The second reason for the discrepancy between model and experiment was mentioned

previously: the experiment probably involves extensive or total first reactant consumption

in the layer next to the source. The results in Table II reinforce the previous indication

that, when ignition does occur experimentally, it is due to further first reactant exo- thermicity rather than to the oxidation of the second reactant (char); as before, the

prediction based on second stage kinetics is quite a bit higher than the experimental value.

3.4 Some Parameter Effects in the Numerical Model

In our previous work (Ohlemiller and Rogers, 1980a) we noted that the ignition

temperature is rather insensitive to the thermal and physical properties of the insulation.

This is because these factors, when varied in the rather small realistic range, tend to have

roughly linear effects on the heat flow balance that determines ignition; the exponential

Arrhenius factor requires very little temperature change to compensate for such balance

shifts

.

In the model, as in the experiments, the effect of insulation bulk density is minimal

(recall that the parameter Z in Equation (12) is proportional to bulk density) . In experiments with configurations B and D, a 20% increase in bulk density caused less than a 5°C

decrease in minimum ignition temperature. The model predicts a 3“C decrease with

configuration D and a 2“C decrease with configuration B.

51
Ambient temperature (i.e., of the surroundings and of the insulation far from the heat

source) is again predicted to be a variable of secondary importance. A change from 20°C to

80°C with configuration A decreases the predicted minimum ignition temperature only about

2°C. There is no experimental data for comparison. Yarborough (1979) measured a substantially larger effect of ambient temperature on the maximum temperature achieved in recessed

light fixtures; this, however, was observed in non- igniting conditions where rapid

compensation by the Arrhenius factor did not come into play.

The predicted effect of retardant removal can be examined as was done previously for

one-dimensional configurations (Ohlemiller and Rogers, 1980a); reaction kinetic parameters

were determined from thermal analysis by the same techniques (Ohlemiller and Rogers, 1980b)

.

The model predicts that for configuration D, the unretarded insulation will have a minimum

ignition temperature about 14°C lower than the retarded case (255°C vs. 269°C) . The

experiments for this configuration with no retardant (limited in number) gave a decrease of

about 15-20°C (255-260°C vs. 276°C). As with the one-dimensional cases previously studied,

model and experiment agree that the smolder retardants here are not very effective in

improving smolder ignition resistance.

3.5 Implications for Needed Retardant Effectiveness

In view of the rather weak effect of these retardants, it is of interest to ask just

how much the exothermic behavior of the insulation must be diminished in order to make it

safely resistant to smolder initiation under typical exposure conditions. This is really a two-part question. One must first estimate roughly how much the ignition temperature has to

be raised above temperatures realized in practice in order to preclude nearly all insulation

ignitions. Secondly, one must determine how much the chemical parameters of the insulation

have to be changed in order to raise the ignition temperature this much.

If one assumes that recessed light fixture ignitions comprise the bulk of the smolder

initiation hazard, the data of Yarborough et al. (1979), combined with the present results,

suggest that the overlap in minimum smolder ignition temperatures and achievable maximum

temperatures in recessed light fixtures is 50°C or less. Then an approximate answer to the

first part of the hazard reduction problem is that the minlmvim ignition temperature of the

insulation needs to be raised by 50°C. (Again, this does not address the effects of long

term exposure to elevated temperatures.)

In previous work (Rogers and Ohlemiller, 1980a), it was shown that boric acid, the most

effective cellulosic smolder inhibitor available, acts rather weakly on the first overall

oxidation stage and does not change its effective activation energy. Here we therefore hold

activation energy constant and ask what reduction in pre-exponential factor (Q*Z in

Equation (12)) is needed to obtain the desired 50°C increase in minimum ignition temperature

The answer depends somewhat on configuration. For configuration F (one-dimensional flat

52
source, 7 cm insulation layer), the product Q*Z must be reduced by a factor of 18; for

configuration A (90° corner source immersed in 20 cm layer of Insulation), the product must

be reduced by a factor of 25. These are drastic changes; boric acid achieves only about a factor of 2 or 3 reduction at practical loading levels (10-15% add-on by wgt.). The 50°C

increase appears to call for virtual suppression of exothermicity from the insulation up to

300°C or more (as measured by DSC at l-5°C/min) . In short, this is not a goal that appears

achievable in the near future.

The nature of the exothermic reactions responsible for this heat release is unknown.

It is not even known which component(s) of this wood product—cellulose, hemi-celluloses,

lignin—are chiefly responsible. Shafizadeh (1979) has studied the low temperature oxidation

of cellulose. There have been numerous other studies of wood component behavior (Beall and

Eickner, 1970). The fact which emerges most clearly is that such studies must be ultimately

done on the insulation itself; its complex history as re-cycled newsprint implies

contamination that could substantially affect its behavior.

3.6 Implications for Smoldering Ignition Test Methods

Some general points should be emphasized first. It should be clear from this work and

the work of many others (see, e.g.. Gray and Lee, 1967; Merzhanov and Averson, 1971) that

there is no unique ignition temperature characteristic of a given insulation material.

Ignition results from an imbalance of heat generation and heat losses so the temperature at

which it occurs depends on all factors affecting these processes. Another point worth

noting is that there is no indication in any work to date that a cellulosic insulation has

been made smolder proof; it can only be made more resistant to smolder as measured by

various tests. As the test becomes more severe, a previously non-igniting material will

ignite.

Under these circumstances, the safest way to assess an Insulation's smolder potential

in application is to subject it to the heat sources and conditions it will see in its normal

end use. This has been done to a limited extent (Yarborough 1979; Gross, 1978) but

it clearly is not practicable for production line testing; the source variations are too numerous and the tests can take 12 hours or more. Small-scale testing is unavoidable. In

addition to improved convenience, such tests should provide conditions that are pertinent to

the real hazard situation and, if possible, are simple enough to be quantitatively analyzed.

Existing or suggested smolder ignition test methods include the use of such heat sources as

a cigarette (Gross, 1978), a 0.95 cm dia. electrical cartridge heater (Day » 1979) and

a smoldering bed of unretarded cellulosic Insulation (Issen, 1980) . These all offer the

advantages of simplicity and speed but each has certain shortcomings as well. A cigarette

as a heat source is deceptively simple in appearance but is actually quite complex

(Summerfield ^ aJ . , 1978); it does not resemble sources typically encountered in practice

and its behavior could be modified by additives in the insulation under test. There are

53
some data indicating a partial correlation between cigarette tests and recessed light

fixture tests (Gross, 1978) but it is impossible to know how general this may be since the

cigarette test is too complex for detailed analysis.

The electrical cartridge heater used by Day ^ is a simpler and more invariant

heat source than a cigarette. However, it is a small source; the present results indicate

that such small sources must be raised to higher temperatures than occur with recessed light

fixtures. By getting away from the appropriate temperature range, one always runs the risk

that the relative ignitability of two materials will change (Bowes and Townshend, 1962)

.

Furthermore, the mode of usage suggested by Day al. is dynamic in nature rather than

static (limit on steady state); one measures an ignition delay time as an ignitability

ranking parameter. In view of these facts, the test as proposed seems more suited to

measuring insulation ignition resistance to localized transient heat sources such as a short

circuiting electrical conductor.

The use of a smoldering mass of unretarded insulation as an "ignition" source (Issen,

1980) is, in a sense, a modification of the cigarette test in which the smoldering heat source has been enlarged. There are again some data showing a partial correlation between

this test and recessed light fixture tests; again the same difficulty pertains in generalizing the results due to the complexity of the test phenomena. This test appears to be more

an examination of continued propagation than of ignition.

The present work can be applied to the design of a test procedure which sacrifices

some convenience and simplicity but gains comparability with the recessed light fixture

hazard. We have two conflicting design goals: (1) short test time; (2) ignition temperatures in the same range as with immersed light fixtures. The first calls for a thin

insulation layer and the second for a thick layer. It was noted previously that test time

with a one-dimensional flat insulation layer is approximately given by (0.4 £ /a), where £

is the layer thickness and a is its thermal diffusivity. In a test, if ignition is not

occurring by this time, one can stop waiting since the layer is as pre-heated as it will

get. (Any extended heat-up time for the source is essentially additive.) Since this implies

that test time increases rapidly with layer thickness, we are forced to consider thin layers.

A layer of the order of 7 cm or more is desirable to assure reproducibility in packing and

permeability characteristics; thinner layers may have increasing problems in this respect.

The preceding expression, (0.4 £ /a) evaluated using thermal conductivity data of Tye (1974)

and representative numbers for density and heat capacity, gives a maximum test time of about

one hour for a 7 cm layer. On this basis it appears undesirable to go any thicker.

However, Figure 2 indicates that, even at 7 cm, a one-dimensional layer has a minimum

ignition temperature of 285°C. Inspection of the results of Yarborough e^ al^. (1979) implies

that the ignition problem with assorted light fixtures begins at about 240-250°C; thus we

want a test that places the minimum ignition border of a nominal commercial material in this

range. This can be achieved without increasing the layer thickness or test time by changing

54
from a flat plate heat source to one that is more "concave," as in configuration A. For

example, the present model indicates that the material used in these tests would ignite at

257°C if placed in a two-dimensional "trough" 15 cm wide and 7 1/2 cm deep with bottom and

sides held at this temperature. A square or circular container with the same dimensions and

heating would give a lower ignition temperature. This could be adjusted further downward by

decreasing the width somewhat and the maximum test time would remain about one hour. Thus,

by manipulation of the heat source geometry, the minimum ignition temperature for this

nominal material can be brought down to 240-250°C.

In view of the variability of attic heat sources and the inevitable small scatter

5°C) in test results of the present type, such a test would best be used to establish

performance classes corresponding to some range of ignition temperatures. For example, the

data of Yarborough suggest that recessed light fixture ignitions occur in the range from

225°C to 275°C; this range could be broken into a series of performance classes having

ignition temperature ranges from 15 to 20°C. These borders and the suggested ignition temperature goal in the heat source design are based on the limited data presently available; more work with a greater variety of real heat sources and commercial insulations is needed

before they could be firmly established.

With a test such as this, the performance class of a new material could be established

in a short time by using four identical heat sources set at the lower temperature limit of

each of the above classes. One merely looks for ignition or non-ignition in each case. If a material falls near a borderline, as indicated by the fact that ignition requires a large

fraction of the one hour test time, the test at that temperature should be repeated. Routine

production checks would require only the checking of the lower temperature limit of the

performance class.

In a go/no-go test such as this, one should realize that variation in a parameter such

as bulk density can cause a borderline material to go from Ignition to no ignition or vice

versa. This is in spite of the fact that bulk density was shown above to have only a weak

effect on ignition temperature. This is inherent in any go/no-go test and implies a need

for care in preparation.

This test has the advantage of closely resembling the recessed light fixture hazard in

terms of both type of exposure (low intensity, long term) and temperature. Furthermore, as was shown here it can be mathematically modeled. The results in the present study provide

some calibration of the meaning of the test results in terms of other sizes and shapes of

heat sources as well as other depths of insulation. This puts the results in a context which

adds a perspective on the degree of hazard represented by a given ignition behavior in this

test

.

55
4. CONCLUSIONS

Heat source/insulation configuration influences smolder ignition temperature more than

any other parameter of this problem. The minimum heat source temperature necessary to cause

ignition can be varied by at least 150°C as a result of variations in heat source geometry

ranging from extended "concave" sources to small "convex" sources. The various potential

ignition sources in an attic space thus pose widely varying degrees of hazard. The impact

of other parameters and even of smolder retardants is much less than this configuration

effect. The reasonableness of this behavior has been established both by crude phenomeno-'

logical models of heat flow/heat generation and by more accurate computer modeling. The

model of this behavior implies that quite large reductions in heat generation rate below

300°C will be needed in order to eliminate the bulk of the smolder ignition hazard in cellulosic loose-fill insulations. The model also permits one to design a smolder test which

provides a reasonably accurate measure of ignition tendency in realistic usage circumstances

5 . ACKNOWLEDGMENTS

This work was sponsored by the Department of Energy under ORNL Sub Task IA-6113-4. The

author would like to acknowledge frequent useful conversations with Dr. F. E. Rogers during

the early stages of this work.

56
6 . NOMENCLATURE

E - effective activation energy

E^ = 41,500 cal/mol

= 39,300 cal/mol

h - convective heat transfer coefficient

-4 2 h=^ 2.2 -3.0 10 cal/cm sec°C; upward facing horizontal surface

-5 2 h - 5.0 • 10 cal/cm sec°C; downward facing horizontal surface plus 8 mm transite

board

£ - insulation layer thickness

^2 - one-half of insulation layer thickness

£ - effective reaction layer thickness

R

£g - radius of cylindrical heat source

Q

- heat release/g of material reacting

q

- flux of heat/unit area

q' - heat release/unit volume

r - radial distance coordinate

R - universal gas constant

S - area

T - temperature

- ambient temperature

Tp - temperature of heat source

V - volume

X - distance coordinate

y

- distance coordinate

57
Z - composite pre-exponential factor

11 3 = 1.66 • 10 g/cm sec

Z^ = 5.67 • 10^ g/cm^sec

X - thermal conductivity of insulation

X - 1.27 • 10 ^ cal/cm sec°C (Tye, 1974)

58
7 . REFERENCES

[1] Ames, W. (1965), Nonlinear Partial Differential Equations in Engineering , Academic

Press, New York, p. 378.

[2] Beall, F. and Eickner, H. (1970), Thermal Degradation of Wood Components: A Review of

the Literature, USDA Forest Products Laboratory Research Paper FPL 130.

[3] Bowes, P. and Townshend, S. (1962), Ignition of Combustible Dusts on Hot Surfaces,

Brit. J. Appl. Phys. 13, p. 105.

[4] Britt, K. W. (1964), Handbook of Pulp and Paper Technology , Reinhold, New York,

Chap. 2 and 9.

[5] Brookhaven National Laboratory (1978), An Assessment of Thermal Insulation Materials

and Systems for Building Applications, Brookhaven Rept. BNL-50862.

[6] Carnahan, B. , Luther, H. and Wilkes, J. (1969), Applied Numerical Methods , Wiley,

New York, p. 508.

[7] Carslaw, H. and Jaeger, J. (1959a), Heat Conduction in Solids , Oxford University

Press, London, p. 126.

[8] Carslaw, H. and Jaeger, J. (1959b), op cit, p. 189.

[9] Day, M. , Suprunchuk, T. and Wiles, D. (1979), A Technique for Evaluating the Smolder

Resistance of Loose Fiber Fill Cellulose Insulation, J. Consumer Prod. Flamm. , Vol. 6, p. 233.

[10] Eckert, E. and Drake, R. (1972), Analysis of Heat and Mass Transfer , p. 110.

[11] Gray, P. and Lee, P. (1967), Thermal Explosion Theory, in Oxidation and Combustion

Reviews (C. F. Tipper, Ed.), Elsevier, New York.

[12] Gross, D. (1978), A Preliminary Study of the Fire Safety of Thermal Insulation for

Use in Attics or Enclosed Spaces in Residential Housing, National Bureau of Standards

NBSIR 78-1497, Washington, D.C.

[13] Issen, L. (1980), Fire Performance of Loose Fill Cellulosic Insulation in Residential

Occupancies - A Progress Report, National Bureau of Standards NBSIR 80-2085,

Washington, D.C.

[14] McDonald, H. (1979), Combustion Modeling in Two and Three Dimensions—Some Numerical

Considerations, Prog. Energy Combust. Sci. , Vol. 5, p. 97-122.

[15] Merzhanov, A. and Averson, A. (1971), The Present State of the Thermal Ignition

Theory; An Invited Review, Combustion and Flame , Vol. 16, pp. 89-124.

[16] Ohlemiller, T. and Rogers, F. (1980a), Cellulosic Insulation Material II. Effect of

Additives on Some Smolder Characteristics, Combustion Science and Technology, in press.

[17] Ohlemiller, T. and Rogers, F. (1980b), Smoldering Combustion Hazards of Thermal

Insulation Materials, final report on DOE Contract ORNL/76086, Princeton University,

Princeton, N.J.

[18] Perry, J. (ed.) (1950), Chemical Engineers Handbook , Third Edition, McGraw-Hill,

New York.

[19] Rogers, F. and Ohlemiller, T. (1980), Cellulosic Insulation Material I. Overall

Degradation Kinetics and Thermochemistry, Combustion Science and Technology, in press.

[20] Schaffer, E. (1979), On Smolder Initiation in Cellulosics Under Prolonged Low Level

Heating, to be published in Fire Technology.

59
[21] Shafizadeh, F. and Bradbury, A. (1979), Thermal Degradation of Cellulose in Nitrogen

and Air at Low Temperature, J. Appl» Poly. Sci. , Vol. 23, p. 1431-1442.

[22] Suramerf ield, M. , Ohlemiller, T. and Sandusky, H. (1978), Thermophysical Mathematical

Model of Steady-Draw Smoking and Predictions of Overall Cigarette Behavior, Combustion

and Flame , Vol. 33, pp. 263-279.

[23] Thomas, P. and Bowes, P. (1961), Thermal Ignition in a Slab with One Face at a Constant

High Temperature, Trans. Farad. Soc. 57 Part II, p. 2087.

[24] Tye, R. (1974), Heat Transmission in Cellulosic Fiber Insulation Materials, J . of

Testing and Evaluation , Vol. 2, No. 3, p. 176.

[25] Yarborough, D. , Yoo, K. and Konen, P. (1979), Recessed Light Fixture Test Facility,

Oak Ridge National Laboratory ORNL/SUB-7504/3

.

[26] Zicherman, J. and Fisher, F. (1978), Fire Protection Problems Associated with

Cellulose-Based Insulation Products, Society of Fire Protection Engineering Technology

Report #78-7, Boston, Mass.

[27] Lucca, D. (1981), "An Investigation of Co-Current and Counter-Current Smoldering

Combustion in Particulated Fuel Beds" MSE Thesis, Dept, of Mechanical and Aerospace

Engineering, Princeton University, Princeton, N.J.

60
61

Figure

1.

Schematic

cross-sectional

view

of

the

eight

configurations

tested.

The

double-line

border

segment

represents

the

fixed

temperature

heat

source;

elsewhere

the

insulation

bed

rests

on 6

mm

asbestos

board.

The

dimensions

shown

are

in

centimeters

and

they

apply

to

all

similarly

sized

parts.

Configurations F

and G

provide

width

to

depth

ratios

that

yield

essentially

1-D

heat

flow

at

the

sample

center.
91.4cm

62

SECTION

IS

6mm

ASBESTOS

BOARD)
Q

I I I I ^ I I I I I ABCDEFGH

CONFIGURATION

Figure 3. Results of ignition tests on the

eight heat source configurations. F and G are

the 1-D layers (7 cm and 5 cm, respectively)

.

63
U.S. DEPT. OF COMM.

BIBLIOGRAPHIC DATA

1. PUBLICATION OR

REPORT NO.

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

SHEET (See instructions) NBSIR 81-2350 August 1981

4. TITLE AND SUBTITLE

Smoldering Combustion Hazards of Thermal Insulation Materials

5. AUTHOR(S)

T. J. Ohlemiller

6. PERFORMING ORGANIZATION (If joint or other than NBS, see instructions)

national bureau of standards

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

7. Contract/Grant No. ORNL Sub Task IA-6113-4

8. Type of Report & Period Covered

Interim Report

Oct. 1, 1979-Apr. 30, 198

9.

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

U.S. Department of Energy

Oak Ridge National Laboratory

Oak Ridge, Tennessee 37830

10.

SUPPLEMENTARY NOTES

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

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

bibliography or literature survey, mention it here)

The smoldering combustion hazards of cellulosic loose fill insulation materials

fall into three categories; smolder initiation, smolder propagation and transition from

smoldering into flaming. Our previous findings on the initiation problem are summarized

briefly. They serve as the basis for recommendations on an improved smolder ignition , test method which is designed to give ignition temperatures comparable to those in practice. The proposed test method requires checking against full-scale mock-up results

before it can be considered for implementation.

Smolder propagation, driven by buoyant convection, through a thick (18 cm) layer of

cellulosic insulation has been extensively examined. A heavy (25% add-on) loading of

boric acid cuts the propagation rate in half (from approx. 0.3 to 0,15 cm/min) but does

not come close to stopping this process. Analysis of experimental profiles for temperature, oxygen level and remaining organic fraction strongly indicate that the smolder

wave is oxygen-supply controlled and that it involves both first and second stages of

oxidative heat release from the insulation material. The balance of involvement of the

two stages varies with depth in the layer. It appears that efforts to develop improved

means of suppressing smolder propagation must be directed at the entire oxidation process. However, since boric acid is fairly effective at slowing the second stage of oxidation, most new efforts should be aimed at the first stage of oxidation (which also is

responsible for smolder initiation).

12. KEY WORDS (Six to twelve entries; alphabetical order; capitalize only proper names; and separate key words by semicolons)

Boric acid; cellulosic insulation; ignition; propagation; smoldering combustion;

test methods; wave structure

13.

AVAILABILITY

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