NIST Technical Note 1493

Aspects of the Fire Behavior of

Thermoplastic Materials

T.J.Ohlemiller

J.R. Shields

N~ National Institute of'Standards and Technology • U.S.Department of Commerce

NIST Technical Note 1493

Aspects of the Fire Behavior of 

Thermoplastic Materials

T.J. Ohlemiller 

J.R. Shields 

U.S. Department of Commerce 

Building and Fire Research Laboratory 

National Institute of Standards and Technology 

Gaithersburg, MD 20899

January 2008 

U.S. Department of Commerce 

Carlos M. Gutierrez, Secretary 

National Institute of Standards and Technology 

 James M. Turner, Acting Director 
Certain commercial entities, equipment, or materials may be identified in this 

document in order to describe an experimental procedure or concept adequately. Such 

identification is not intended to imply recommendation or endorsement by the 

National Institute of Standards and Technology, nor is it intended to imply that the 

entities, materials, or equipment are necessarily the best available for the purpose. 

National Institute of Standards and Technology Technical Note 1493 

Natl. Inst. Stand. Technol. Tech. Note 1493, 160 pages (January 2008) 

CODEN: NSPUE2 
1

Aspects of the Fire Behavior of Thermoplastic Materials 

T. J. Ohlemiller and J. R. Shields 

Abstract 

Thermoplastic polymers pose unique new challenges (resulting from the movement of burning 

melt material) for the understanding and control of fire growth on objects that incorporate them. 

Both full density thermoplastics such as polypropylene and polystyrene, as well as low density 

polyurethane foams are examined in this study, which has two overall goals: (1) assessment of 

potential test methods for quantitatively judging the flammability hazard of a thermoplastic 

(focused most specifically on polyurethane foams) and (2) fostering the development and 

validation of a model of fire growth over generic configurations of thermoplastic materials 

suggested by their end product use (particularly upholstered furniture). One such configuration is 

that used in the flammability test method, but early model development steps emphasize simpler 

configurations and materials. A critical aspect of modeling these materials is an adequate 

description of the viscosity of the polymer melt as a function of temperature; the viscosity can 

vary by several orders of magnitude. A procedure for deriving an empirical description of 

viscosity for full density thermoplastics, dependent only on temperature, is given but it requires 

extrapolation of melt viscosity out to burning temperatures. The procedure may be stymied by 

the complex behavior of polyurethane foam melts, indicating a need for further work. 

Gasification kinetics of the material, also needed in the fire growth model, are derived here from 

thermogravimetry for four full density thermoplastics. The more complex degradation behavior 

of polyurethane foam requires further work to derive these kinetics. The modeling process has 

been proceeding in stages of increasing complexity in conjunction with an outside contractor. 

The current stage focuses on two-dimensional, non-flaming melting plus gasification at heat 

fluxes comparable to those seen in fire growth. The present study has produced data on four full 

density thermoplastics and several polyurethane foams in this configuration which serves to test 

the developing model. Among the experimental results is the flux-dependent fraction of mass 

lost as melt flow from the heated sample surface. For polyurethane foams of varied composition, 

these results were found to vary over a wide range. This helped in understanding the relative fire 

growth behavior of these same foams when tested as roughly 30 cm by 60 cm slabs under the 

National Institute of Standards and Technology (NIST) Furniture Calorimeter. Similar scale tests 

were also done with thin sheets of polypropylene (the fire growth configuration which the model 

will attempt to predict first); these tests revealed complex flow dynamics in the melt pool fire and 

provided clues about the role of a pool fire in the overall fire growth process. As a result of the 

large scale foam tests we have proposed a tentative test configuration for polyurethane foams that 

are intended for use in upholstered chairs and a validation test series is being planned. The 

question of whether this test can be scaled down to use smaller amounts of foam is addressed; the 

scaling is difficult because of a mis-match in transient behaviors. 

Introduction. In this report the phrase “thermoplastic materials” is generalized to 

include any solid which produces a significant amount of a liquid phase during its 

burning. Thus we include not only traditional full-density thermoplastics like 

polypropylene or polystyrene but also such materials as polyurethane foam which, though 
2

it is a thermosetting polymer, degrades to yield, in part, a liquid phase when burning or 

when subjected to heat flux levels comparable to those seen during burning. 

Full density thermoplastics are commonly used in consumer electronics devices as, for 

example, the outer case or shell of a TV set or a computer monitor, keyboard, speakers, 

etc. They are increasingly common in office furniture and in automobiles, both under the 

hood and in the passenger cabin. Low density materials, in particular, polyurethane foam 

(which is actually an entire broad class of materials) are found as the soft cushioning in 

upholstered chairs, car seats and in mattresses. 

All of these thermoplastic materials yield extra complexity when they burn in the context 

of the products in which they are found. Under the influence of gravity, the liquid phase 

formed during thermal degradation flows downward. If, as is typical, this liquid is 

burning, then it extends the flaming zone on the solid downward onto whatever surfaces 

are available to catch the liquid (ultimately a horizontal floor or ground surface). This 

constitutes, at the least, a new form of flame spread on the object containing the 

thermoplastic (in addition to normal forms of flame spread over solid surfaces, which are 

typically fastest in the upward direction), extending the area of fuel that is burning and 

thereby increasing the overall heat release rate from the object. In many cases this 

downward flow of flaming liquid results in a pool fire under the object. If that pool fire 

is close enough to the object that its plume reaches the object, the result can be a selffeeding1

 pool fire that further enhances the rate of heat release from the burning system 

[1, 2, 3, 4]. 

In addition to this penchant for liquid-assisted flame spread, thermoplastic materials also 

tend to deform significantly as they burn. Thus large changes in the geometric shape of 

the burning object are common. These two aspects of thermoplastics make modeling fire 

growth on them, or on objects containing them, extraordinarily difficult. 

From the standpoint of U. S. residential fire statistics, the thermoplastic objects (or 

objects which are in large part thermoplastic) that are of most relevance to a goal of fire 

loss reduction are upholstered chairs and mattresses. The general problem posed by 

modeling of fire growth on such objects is essentially intractable due to the very wide 

variety of poorly characterized materials present as well as an almost infinite variation in 

construction details. A detailed fire growth model requires input parameters describing 

such things as the kinetics of degradation and gasification of each significant component, 

its melt viscosity-temperature relationship, its thermal properties, etc2

. This information 

is quite time-consuming to measure (and then it is only approximate; see below). The 

 

1

 Self-feeding here means that the pool fire plume increases the rate of melt generation on the surface of the 

thermoplastic. In turn this increases the size of the pool fire whose taller height, in turn, further increases 

the melt flow rate to the pool. This can potentially lead to exponential fire growth though this is frequently 

curtailed by the finite thickness of the materials involved. 

2

 It is apparent from the visual behavior of some fabric systems that prediction of their behavior in a fire 

also requires a description of their very thermally sensitive state of mechanical stress. Thermoplastic 

fabrics can split open and curl back abruptly as a result of tension forces. Chars from cotton fabrics shrink 

and then split in seemingly random ways. The behavior of fabric blends, of which there are unlimited 

numbers, is dependent on the nature of the constituent polymers and the details of their weaving. 
3

variability of the chemical composition of the components in these real objects, driven as 

it is by yearly fashion changes and economic considerations, is essentially unlimited. 

One could spend several years doing a detailed model of the fire behavior of a specific 

chair design only to find that it is irrelevant to what is currently on the market3

. The best 

one can eventually hope for in modeling fire growth on upholstered chairs and beds is 

generic representations of classes of such objects4

 that may be useful in such applications 

as design fires for performance-based fire codes. Successfully validated versions of such 

models are very likely a decade away. 

Such complex models are not the only path forward, if, by forward, we mean movement 

toward less flammable soft furnishings. Such movement has typically been 

predominantly empirical in nature and often occurs only when prompted by regulation. 

A nearer-term goal is a less complex model and a test method, used interactively as 

research tools, to assess and help improve the likely behavior of a given thermoplastic 

material if it were placed in the context of one of the types of objects mentioned above. 

One half of the goal is a test method whose results reliably predict that a thermoplastic 

material will be a major or minor hazard5

 if used in the context of a real object such as a 

chair or computer monitor. The other half of the goal is a model of the behavior of 

thermoplastic materials in the test method context and its purpose is to help gain a 

detailed understanding of how and why a material behaves as it does in the test context, 

thus providing clues as to how to improve the material’s behavior in real objects. 

For a test method to yield outputs that accurately reflect how a material will behave in the 

more complex context of a real object, such as a chair or a video monitor, which 

incorporates other materials and typically more complex geometries, it is necessary to 

capture the dominant factors that influence how the thermoplastic material of interest 

behaves.6 Work toward the goal of developing such a test is part of the study discussed 

here. 

 

3

 The cone calorimeter was developed in large part as a means of dealing with this daunting variety of 

materials combinations. The goal was to empirically characterize the heat release rate behavior per unit 

area of combinations of materials as a system and use that information in fire growth models. Because of 

significant differences in behavior in the cone versus that in the real objects, this goal has remained largely 

elusive, especially for soft furnishings [23, 24]. 4

 For example, a generic upholstered chair of intermediate mass covered by a thermoplastic fabric or a sofa 

covered by a charring fabric. 

5

 Hazard level in this context is to be primarily judged by the peak rate of heat release from the object and 

the time to that peak. These same metrics are carried over to the test method and to the model of the test 

method. 

6

 It is also necessary that the thermoplastic material that is the subject of the test be the dominant flammable 

material in the object. The approach being advocated here will not work if the fire growth of the real object 

is largely governed by factors other than the thermoplastic material that is the focus here. For many chairs 

and mattresses the heat release rate is essentially dominated by polyurethane foam behavior, especially 

when it is covered by a thermoplastic fabric. A charring fabric can significantly suppress some of the worst 

behavior of the foam but the test method would still seek to reveal the worst case potential of the foam, 

were it to be used with a thermoplastic fabric. Similarly, in objects such as a computer monitor or TV set, 

the fire behavior is dominated by the thermoplastic case and the test results on that material should give a 

good indication of the fire behavior of the real object. 
4

Modeling the growth of fire on thermoplastic materials or objects has long been 

recognized as a very difficult undertaking. In the present study we have approached this 

in a stepwise fashion with parallel experiments to test each model step. As a first step we 

have focused on non-burning behavior, looking just at the behavior of a simple vertical 

thermoplastic slab heated uniformly on one vertical face. The slab, in general, can 

respond to the heating by both melting and gasifying. In the next step, it can be burning 

on that face. In the following step it can be burning on its face and interacting with its 

own melt pool fire. In the final step, the slab is extended laterally, ignited locally and a 

melt pool-assisted flame can spread over and consume the entire slab7

. That, at least, is 

the plan. In practice, the first step has proven to be a major challenge. Commercial 

computational fluid dynamics (CFD) codes applied to this problem give impractically 

long solution times (many days per case) [5]. Currently, NIST has contracted to Reaction 

Engineering International (REI) of Salt Lake City for the development of a condensed 

phase code that can be coupled to the NIST Fire Dynamics Simulator (or similar gas 

phase flame model) as a description of the gas phase flame and its thermal feedback to 

the condensed phase. The outcome of this effort remains to be seen but significant 

progress is being made [25, 26]. K. Butler of NIST is also exploring the viability of 

particle methods, a relatively new class of numerical techniques for problems involving 

large geometry changes [27]. Much of the experimental work described in this report is 

in support of these modeling efforts, providing both model inputs and experimental 

results against which to test model predictions. In this report, modeling efforts are 

described only in relation to the experiments reported here. 

Emphasis in this study has generally been shifting from full-density thermoplastics 

toward polyurethane foams. The latter are part of the products involved in many more 

fatal fires in the U. S. and thus there is a stronger incentive to address this class of 

materials. These foams are, however, distinctly more complex than simple thermoplastic 

resins like polypropylene. This prompted an initial focus on the full density 

thermoplastics and the REI contract is aimed strictly at these. Thus many of the 

experiments described here (all those dealing with full density thermoplastics) have this 

focus. We have also begun exploring the behavior of polyurethane foams in certain 

situations, looking for viable configurations that could form the basis for a predictive 

flammability assessment. That work is also summarized here. 

EXPERIMENTAL MATERIALS AND METHODS. 

Materials. Three types of full-density thermoplastic polymer resin were looked at for 

this study; only two, polypropylene and polystyrene, proved to be suitable. Low 

molecular weight poly methylmethacrylate (PMMA, mean molecular weight of 120,000) 

was examined in a few radiative heating experiments (vertical slab, 25 kW/m2

; the 

experimental set-up for all such tests is described below). It was known that normal 

commercial PMMA (mean molecular weight of ca. 106

) does not exhibit melt flow 

 

7

 Numerous variations of this final configuration become possible, once the model is in hand. Of particular 

interest is a pair of thermoplastic slabs arranged symmetrically about a central plane so that the flame zones 

on each exchange radiation which intensifies the burning of each slab. Such an effect can be seen in the 

burning of upholstered furniture. A Vee configuration is used below in the test method development study. 
5

because its unzipping character during thermal degradation leaves its molecular weight 

largely unchanged, yielding a melt viscosity that is too high to permit significant flow at 

normal gravity. The lower MW PMMA has a melt viscosity that is at least 1000 × less 

and it did exhibit a tendency to flow downward. It also exhibited a behavioral aspect 

which the present study deemed too difficult to deal with – two phase flow. PMMA 

degrades exclusively to its monomer which is highly volatile at the conditions of interest. 

As a result it nucleates bubbles in depth, turning the surface region of the sample into a 

froth. Such a two-phase flow system cannot be effectively characterized by the available 

rheometric instrument. Thus it was decided not to pursue this polymer further. 

Polypropylene and polystyrene resins were also explored and they showed a minimal 

tendency to bubble in our melt flow experiments. Both of these resins tend to undergo 

random scission when heated sufficiently (above the ca. 200 oC to 250 oC range where 

they are molded commercially), lowering the mean molecular weight. This, along with 

heat-induced increases in polymer chain mobility, lowers the melt viscosity, yielding a 

fluid melt. These two resin types thus became the exclusive focus for the full density 

thermoplastic aspects of the study8

. Samples of commercial resins were obtained in bead 

form. Extremes (low and high) of initial melt viscosity were sought, as indicated in the 

manufacturer’s reported melt flow index; this proved to cover only about one order of 

magnitude (or less), presumably to keep the moldability of the resins within bounds set 

by commercial machines. Table 1 lists the properties of the four thermoplastic resins 

used in the two-dimensional radiative heating and burning studies described below. 

Higher melt index indicates lower melt viscosity (at the measurement conditions). 

The test samples were molded from beaded commercial resins in a heated press. The 

available mold system could not remove all air bubbles from the mid-depth of the 2.5 cm 

deep samples. This resulted in a sort of sandwich structure in which the two faces of a 

sample, to a depth of roughly 7 mm to 8 mm was bubble-free; in the middle, between 

these two clear areas, the sample had myriad fine bubbles, more or less evenly dispersed. 

This had no visible effect on the experiments and the bubbles did not noticeably affect 

the melt flow process. 

A commercial polypropylene, generally similar in its melt and burning behavior to those 

in Table 1, was obtained in the form of large thin sheets (3 mm and 6 mm thick); this was 

used in the exploratory lateral fire spread studies described below. No further data on 

this polymer are available. 

As noted above, polyurethane foam is a class of materials, based on the reaction of an 

isocyanate with a polyol. Commercial formulations may contain as many as eight 

different ingredients, including these two plus catalysts, blowing agents, surfactants and 

fire retardants. The nature of the polyol can be widely varied but typically, for the 

 

8

 It is worth noting that there is a variety of other widely-used thermoplastics that do not exhibit the 

relatively simple behavior seen here with polypropylene and polystyrene. In earlier work we found 

polyethylene to be prone to form a seemingly cross-linked skin on the sample face that captured melting 

polymer behind it. In other studies ABS was prone to form a bubbly, partially charring, viscous mass, as 

was polycarbonate. 
6

applications of interest, it consists of a polyether based on ethylene and/or propylene 

oxide. The isocyanate is typically toluene diisocyanate (TDI) but may also be methylene 

diphenyl diisocyanate (MDI). The physical properties can be widely varied by adjusting 

the nature and proportions of these many ingredients. All of these in turn affect the 

details of the fire behavior of polyurethane foams. For this study we obtained samples of 

foams of widely varying fire behavior. Table 2 lists the limited information available on 

them. The fire retardant content ranges from zero to high; the graphited foam passes the 

FAA kerosene burner test in the context of an aircraft seat [28]. Two viscoelastic foams 

were included; these are becoming increasingly popular in mattress applications.

There have been several types of experiments involved in developing input data for 

modeling, in obtaining comparison data for the models and in exploring the threedimensional spread of pool fire-assisted flames. These are described briefly here along 

with some limitations where appropriate. 

Thermogravimetry (TG). This is a common laboratory technique for characterizing the 

weight loss characteristics (via gasification) of a condensed phase material. A 

representative sample of a few milligrams in size is programmed upward in temperature 

at a fixed rate in an inert atmosphere (nitrogen) while it is being weighed on a very 

accurate balance. The weight is accurate to approximately ± 1 % and the temperature to 

approximately ± 2 oC. Here we are interested in using this technique to estimate the 

Arrhenius kinetic parameters that provide a simple description of the gasification process 

of our thermoplastic materials. The technique has thus far been applied to two 

polypropylene resins and two polystyrene resins. All four of these can be reasonably 

well described by a single gasification reaction that leaves no residue. The technique can 

also yield information on the relative stability of differing materials. We have applied it 

for this purpose to the polyurethane foams in this study. 

To obtain the kinetic parameters of a gasification reaction, it is necessary to obtain data at 

two or more heating rates. Those heating rates must be low enough that the sample, as it 

is programmed upward in temperature, is spatially uniform in temperature. This 

necessitates quite low heating rates. Values of (0.5, 2.0 and 5) oC/min were used here. 

As the heating rate is increased, the reaction is shifted to higher temperatures (the result 

of a simple competition between the rate of reaction and the rate of heating). This shift is 

a measure of the effective activation energy of the gasification reaction. There have been 

numerous methods proposed in the thermal analysis literature for obtaining the kinetic 

parameters from these types of experiments. One of the oldest and simplest is that due to 

Kissinger [6], which was used here. It uses the shift in the peak reaction rate to obtain the 

kinetic parameters. This leads to some deficiency in the exact description of the overall 

weight loss curve as will be seen below9

. 

There are two other concerns with this mode of obtaining kinetic parameters that apply to 

the fire-level heat fluxes used in the experiments below. The first concern is that these 

 

9

 Appendix 3 describes the exploration of a different TG data reduction technique which uses all of the 

data. This technique pointed to more complex behavior for the one polypropylene resin to which it was 

applied. 
7

fire level heat fluxes heat the resin at a substantially higher rate than is the case in the 

thermogravimetric experiments. This higher heating rate pushes the reaction to higher 

temperatures where the controlling mechanism could, in principle, be changed. One can 

only assume that no such change takes place. With simple resins like polystyrene and 

polypropylene this is more likely to be the case10 than for a complex material like 

polyurethane, but direct evidence of change (or not) is very difficult to obtain. 

The other concern has to do with the atmosphere. In the thermogravimetric experiments 

it is nitrogen; in the radiative melt/gasification experiments (described below) it is usually 

air. The degradation of all polymers of interest is potentially affected by oxygen11. 

When a polymer is burning, oxygen typically cannot reach its surface, having been 

consumed in adjacent flames. We are ultimately interested only in flaming but some of 

the modeling and experiments do not involve flaming. Rather than build an elaborate 

enclosed apparatus for these non-flaming experiments, we have conducted them in room 

air. We have made some attempt (described below) to assess whether excluding most 

oxygen influences these results, but this issue is not entirely resolved and should be kept 

in mind when model and experiment for non-flaming conditions are compared. 

Polymer Melt Viscosity. This is a fundamental issue which must be dealt with in any 

modeling effort aimed at thermoplastic materials; one must have an appropriate 

description of the flow properties of the liquid phase of the polymer of interest. In 

general, and especially at lower temperatures, polymer melts exhibit non-Newtonian flow 

behavior, exhibiting such features as a shear rate dependence to their effective melt 

viscosity. Fortunately, at higher temperatures, comparable to those near the surface of a 

burning object, most polymer melts become more nearly Newtonian. Our first major 

simplification then is to assume that the melt behavior is always Newtonian and that all 

one needs to describe the polymer flow is its melt viscosity. 

The problem is still complex. At these high temperatures the polymer is degrading and 

its mean molecular weight is decreasing as a result. Both increased molecular mobility 

 

10 The degradation of such polymers proceeds by a free-radical mechanism. The initial, randomly-placed 

break of the polymer backbone, forming a pair of highly reactive free radicals, is probably the controlling 

step. It is followed by a series of rapid, free-radical chain propagation, chain transfer and chain termination 

reactions occurring in series and parallel that serve to lower the mean molecular weight of the broken chain 

and others in its neighborhood. As this sequence progresses (given a high enough temperature), species are 

generated which are volatile enough and concentrated enough to be able to nucleate vapor bubbles. These 

bubbles grow by diffusion of low molecular weight species from the surrounding melt. The pressure in the 

bubbles builds, causing them to swell toward the nearest free surface (open to the atmosphere) and burst. 

The degradation of polyurethanes is probably not greatly dissimilar but the larger number of types of bonds 

between a larger variety of types of atoms opens the way to more competing paths and the possibility that 

this competition will be more likely to shift the dominant mechanism as conditions change. 

11 The extent to which polymer degradation can be affected by oxygen tends to be limited by the very slow 

diffusion of oxygen into the condensed phase and the time available for this diffusion. Running the 

thermogravimetric experiments in an air atmosphere would tend to give an exaggerated oxygen effect 

because the sample exposure time is much longer than in the non-flaming experiments. One would have to 

develop a specialized model, applicable only to the non-flaming experiments, to fully account for the 

possible impact of oxygen on the observed behavior. This would be a major diversion from the present 

goals. 
8

and decreased molecular weight contribute strongly to the decreased viscosity of a 

polymer melt in the temperature range of interest (roughly 300 oC to 450 oC). The 

decrease in molecular weight is a consequence of a complex series of chemical reactions 

that are shortening the average length of the polymer backbone. If this kinetic aspect of 

the polymer behavior had to be included in a model, the thermoplastic fire growth 

modeling problem would essentially be intractable. Appropriate kinetic constants for the 

various degradation steps are not available for commercial polymers and, in any event, 

solving for the shifting mean molecular weight would add too much complexity to the 

model to make it practical. Instead, for the full density thermoplastics, we have 

developed an approximate, empirical approach that yields a viscosity description that 

appears to be a function of only one variable – temperature. This single variable 

description is obtained piecemeal. The bulk of it (at lower to intermediate temperatures) 

comes from straightforward rheometry on the virgin polymer resin. The rheometer is a 

sophisticated instrument (Paar-Physica Model UDS 200)12 that, in this case, measures the 

force needed to sustain steady, constant shear rate rotation of a circular (typically 25 mm 

dia) flat plate on top of a 1 mm thick layer of molten resin which, in turn, rests atop a 

fixed plate. An attached computer converts the force measurements into a melt viscosity 

via a standard mathematical solution for this shear flow. This device can yield viscosity 

results for standard materials at a fixed temperature within 10 % of their specified value. 

With successive measurements on melt samples from a given experiment, the results for 

viscosity as a function of temperature may vary by ± 10 % or somewhat more if the melt 

is contaminated with particles. The entire measurement assembly is enclosed in a 

nitrogen purged oven that is programmed upward in temperature at 1 oC/min until the 

sample degradation process has reached such an extent that gas bubbles from the sample 

disrupt the force measurement. The materials examined typically do show a mild shear 

rate dependence of melt viscosity (shear thinning is a non-Newtonian effect) so the 

applied shear rate (set by the rotation rate) is kept at a rate (typically 0.1 s-1) estimated to 

apply to the falling films seen in the polymer melt experiments described below. In this 

manner, with the virgin resin one might get melt viscosity data from under 200 oC to 

somewhat over 300 oC for a polypropylene or polystyrene. 

Simply extrapolating these data out to 400 oC or higher would be highly risky since the 

viscosity is changing over a couple orders of magnitude and the temperature dependence, 

on a semi-log plot, is not a straight line once thermal degradation sets in (as it does well 

before bubbling is seen). Instead, we have obtained two (or three) approximate “data 

points” on this extrapolated line by capturing melts generated at higher temperatures in 

radiation-driven melt flow experiments (described below). The surface temperatures (at 

two different heat fluxes) that are associated with these melts are measured in the 

experiments. The captured melts are then subjected to rheometry in the same manner as 

the virgin resin. Their entire viscosity curve is shifted downward by the fact that they 

have been thermally degraded to a lower mean molecular weight. This viscosity, again, 

 

12 Certain trade names and company products are mentioned in the text or identified in an illustration in 

order to specify adequately the experimental procedure and equipment used. In no case does such 

identification imply recommendation or endorsement by the National Institute of Standards and 

Technology, nor does it imply that the products are necessarily the best available for the purpose. 
9

can only be measured up to temperatures at which further degradation sets in. These 

results are then extrapolated out to the measured surface temperature at which the melt 

was generated to get an estimate of the point on the viscosity curve represented by the 

degraded melt on the heated sample surface. There are further reservations/limitations 

one must consider with these results, as discussed below. However, one has, from this 

procedure, a first approximation to a temperature-only dependence for polymer melt 

viscosity. Ultimately this dependence is fitted with a polynomial description of the 

behavior on a semi-log plot and this becomes the input for a model. 

It should be noted that the viscosity-temperature behavior for the polymer melts is nonreversible because of the chemical degradation involved. That is, when the melt cools to 

a lower temperature after it escapes from the hot polymer surface, its viscosity does not 

go back up along the extrapolated curve described above. Rather it follows the 

temperature behavior measured for the melt obtained from the highest incident heat flux 

seen by the polymer surface. 

Two-Dimensional Radiation-Driven Melt/Gasification/Flaming Experiments. These 

are the counterparts (and test cases) to the early stages of fire growth model development 

described above. The cone heater in the cone calorimeter is used as the radiation heat 

source. Incident radiant heat fluxes on the samples are measured with a water-cooled 

Schmidt-Boelter gage, previously calibrated to an accuracy of approximately ± 3 %. 

There have been a few variants of the exact way in which these experiments were done; 

here we describe the most recent since it is the source of the data presented later. Figure 

1 shows a schematic side view of the layout of the experiment. The cone heater is 

oriented vertically on the left. The irradiated sample face is also oriented vertically and it 

is contained within a square, stainless steel sheath that constitutes a feed tube. The front 

of this tube is 7.5 cm from the front (wider) face of the cone heater. This greater than 

normal spacing between sample face and cone heater permits better visual access needed 

for thermocouple measurements of the sample surface temperature. It reduces the 

maximum available heat flux but does little to diminish the uniformity of the heat flux 

falling on the sample face. It also permits easier access for insertion of pans or bottles for 

capturing the melt flowing downward off of the sample face. As the sample face recedes, 

the sample is fed forward in the tube to keep its face very nearly fixed at the tube exit 

plane. This feeding process is stepwise and manually driven in the case of full density 

polymer samples; it is continuous and motor-driven in the case of polyurethane foam 

samples13. 

The sheath tube cross-section is 10 cm square and this is the face size of the full density 

thermoplastic samples. Typically they were 5 cm thick, formed from back-to-back 

2.5 cm thick samples. As noted above, the 2.5 cm thick samples had a central layer of 

fine bubbles from the molding process that formed them. Again, any effect of this bubble 

layer did not visually manifest itself in the experiments. However, they must have had 

 

13 The same slide mechanism is used for both types of samples but the available motor drive could not 

handle the very slow feeding rates seen with the solid polymer samples. Foam, with its much lower 

density, requires a much faster feed rate to keep the regressing sample face fixed at the tube exit. 
10

some effect on the thermal conductivity of the samples, presumably reducing it locally, 

and this affects, to some extent, any comparison with model results. 

The polyurethane foam samples posed a different problem. If they were the same crosssectional size as the full-density resins, they behaved very poorly, receding unevenly over 

their face, especially at their outer edges. They appeared to be highly sensitive to heat 

loses around their periphery.14 In order to minimize this effect for the foams, the sample 

cross-section was reduced to 5 cm square. The lateral sides and top of this smaller sample 

were additionally insulated by 2.5 cm thick layers of polyisocyanurate foam15 fixed 

within the stainless steel sheath tube (the sample bottom could not be similarly insulated 

because the feed mechanism came through the bottom of the tube). This essentially 

eliminated the problem. The foam samples were 40 cm long and, as noted above, were 

fed forward by a motor drive at a rate (set by the test operator) that matched their 

regression rate. 

As shown in Fig. 1, the sample assembly (sample, sheath tube, motor-driven slide, base 

plates) sits on top of a high capacity scale with a 0.1 g resolution. As noted, when the 

sample was a full-density thermoplastic, the sample was manually fed forward at 

intervals by turning a crank on the slide mechanism. This momentarily interrupted the 

weighing process but this had essentially no effect on the ability to follow the decreasing 

sample weight. The steady-state rates of weight loss reported here are obtained by fitting 

a straight line to the weight vs time data; they are accurate to approximately ± 3 %. 

The material flowing from the sample face was captured either in an aluminum foil pan 

or in a bottle. The weight of this was followed by a scale with a resolution better than 

0.1 g. In some experiments, the sample face was burning. In such cases it was desirable 

that the captured melt not continue to burn. This prompted the use of deep bottles as the 

capture vessel having the top opening partially covered with aluminum foil (and in some 

cases the base of the capture bottle was water cooled). This caused burning to halt as a 

result of a lack of oxygen. 

Some of the polyurethane foams tended to accumulate a denser char-like “skin” on the 

sample face as the bulk of the foam both gasified and melted, flowing downward. 

Ultimately it was noted that this skin probably served to block oxygen from participating 

in the melt process immediately behind it. The skin was therefore allowed to stay in place 

during most of each foam experiment, being removed (with tweezers) only when it 

threatened to slide off into the melt capture bottle. The skin typically began re-forming 

within 10 s to15 s of removal, but it required removal only every several minutes. 

The sample surface temperature was measured in these experiments with a 0.05 cm dia 

stainless steel sheathed thermocouple (chromel/alumel) that was moved forward at 

 

14 The stainless steel tube sheath was insulated on its outer surface with 1.3 cm of ceramic fiber insulation. 

This left the tube wall itself (ca. 0.6 mm thick) as a weak heat sink which had no effect on the full density 

thermoplastics but a large effect on the foam. 

15 The ends of the isocyanurate foam blocks were baked at the highest flux to be used. This charred the 

ends of the blocks and eliminated them as a spurious source of weight loss from the sample assembly. 
11

intervals through the central hole in the cone heater; typically three measurements were 

made per test, after the sample was in a steady regression state. The lead length of the 

thermocouple was supported on a 3 mm dia ceramic tube (kept small to minimize 

shadowing of the cone heater during surface temperature measurements)16. The last 6 

mm of the sheath before the end with the junction was bent downward at 90o so that this 

section of the thermocouple contacted the surface in a parallel manner, essentially 

eliminating conductive lead losses (a length equal to12 diameters of the sheath was in the 

surface melt). When a surface temperature measurement was to be made, the 

thermocouple support was fed forward manually on a mechanical slide. When it 

contacted the surface, a mark was placed in the computer file. The thermocouple was 

typically oscillated back and forth a mm or so within the surface melt layer to assure that 

it was thoroughly wetted and to seek a maximum value. (In the case of the polyurethane 

foams that formed a skin, the thermocouple stayed to the outside of the skin; when, in a 

few cases, the measurement was made right after skin removal, the results were not 

consistently different.) The thermocouple was kept in contact for ca. 60 s and then 

removed, with a second mark being made in the data file. 

The accurate measurement of surface temperature in this manner is not trivial (and there 

are no easy alternative methods given the strong infrared flux from the cone heater). 

There are two significant sources of error arising from (1) the finite size of the 

thermocouple in comparison to the thermal wave in the solid and (2) radiative heating of 

the thermocouple by cone heater radiation that penetrates the outer layers of the sample. 

The first source of error can be estimated using the steady-state temperature profile based 

on neglect of the melt flow layer (see Appendix 1 for a steady-state model of the 

melt/gasification process). The characteristic depth of this profile is 

 ( / ) c = λs mss Cs l

where λs is the thermal conductivity of the condensed phase, mss is the steady-state mass 

loss rate from unit area of the sample and Cs is the sample heat capacity. This length, 

coupled with the measured surface temperature values, allows one to estimate the 

temperature gradient at the sample surface. The thermocouple spans a non-negligible 

segment of this gradient. Assuming that the thermocouple in such a situation would tend 

to read the temperature value at the middle of the segment, one finds that, for this reason, 

it will tend to read a value less than the actual surface temperature and this error varies 

from about 7 oC at 20 kW/m2

 to 20 oC at 45 kW/m2. This estimate assumes that the 

sheath of the thermocouple is immersed in the sample surface, and flush with it. 

The second source of error has the opposite effect – it tends to give a higher reading than 

the actual value at the sample surface. It is more difficult to estimate and the estimate 

requires information on radiation penetration that is currently only available for 

polypropylene. Measurements of cone heater radiation penetration in a sheet of 

commercial polypropylene were made several years ago using the NIST Building and 

 

16 It is estimated that the shadow of this support blocked less than 5% of the cone heater radiation in the 

vicinity of the thermocouple junction. 
12

Fire Research Laboratory (BFRL) gasification apparatus (which uses the same type of 

cone heater). We assume that they apply to the two polypropylene resins used here. We 

cannot assume that they apply to the polystyrene resins or the polyurethane foams since 

the differing compositions of these materials result in differing infrared spectra. The 

black, char-like “skin” that forms on some of the polyurethanes probably absorbs the 

cone radiation in a very short depth, but this needs to be checked. Available infrared 

spectra for polystyrenes (Sadtler Standard Spectra) are ambiguous as to whether they 

differ appreciably from the polypropylene17. 

For polypropylene, the available radiation penetration data show that about half of the 

incident radiation reaches a depth of 4 mm. Then the thermocouple, just immersed in the 

surface, as assumed above, sees a largely undiminished radiant flux on the side facing the 

heat source. It is immersed in a slow downward flow of melt around it that removes the 

absorbed radiant heat. As a first approximation, we do a steady-state heat balance on this 

cylindrical object which has a flow, parallel to the cylinder axis, over its entire periphery 

while absorbing radiation on half of its surface. That flux is taken as the value at 1⁄4 of the 

thermocouple sheath diameter, which, for polypropylene, is about 87 % of the flux 

incident on the sample surface. This heat balance yields an estimate that, at an incident 

flux of 20 kW/m2

 the thermocouple reads about 20 oC higher than the surrounding melt 

and at a flux of 45 kW/m2

 it reads 45 oC higher than the melt. This estimate has 

significant uncertainties due the fact that the heat flow situation around the thermocouple 

is substantially more complex than assumed. The melt flowing past the thermocouple in 

the area closest to the sample surface is thin and thus less effective in cooling the 

thermocouple. On the other hand, the melt behind the thermocouple is in its shadow and 

receives less radiant heat, is cooler, and thus is more effective at cooling the 

thermocouple. The “front” half of the thermocouple does not actually receive the full 

radiant flux over its area; a diminishing view factor with respect to the cone heater lowers 

it on the sides parallel to the longitudinal axis of the sample. Although these effects are 

beyond any simple estimate, the overall implication appears to be that the preceding 

estimate of the radiation-induced error in the thermocouple reading is probably high. 

When combined with the above opposite effect that lowers the thermocouple reading, it is 

likely that the overall effect is that the reading is still somewhat high, but probably not 

more than 20 oC at the high flux end and substantially less high at the low flux end. 

Much of the data reported here is for non-flaming experiments. In those experiments the 

cone heater was allowed to come to a steady temperature (corresponding to a measured 

heat flux level at the sample face) while the sample was protected from its radiation. At 

time zero, a heat shield was removed from the sample face to give a square-wave start to 

the radiation exposure. As the sample heated up and began to gasify/melt, its surface 

would begin to recede. The test operator would use the manual or motor-driven feed 

assembly to feed the sample forward to keep it approximately in the same plane18. The 

 

17 See Appendix 4 for a further discussion. That work was completed only after the bulk of this report was 

written. 

18 The radiant flux varied by about 2.5 %/mm in the neighborhood of the end of the square tube sheath 

enclosing the sample. The full density thermoplastics, whose feeding speed was adjusted manually at 
13

sample and melt capture weights were displayed continually during each test. When it 

was apparent that the melt capture weight was increasing linearly with time, indicating a 

quasi-steady state was achieved, surface temperature measurements were begun. The 

polyurethane foam samples tended to develop a non-planar face with the upper 1⁄4 to 1/3 

of the surface sloping somewhat away from the heat source; this did not appear to have 

any major effect on the results obtained. For the full density thermoplastic tests, the 

experiments were carried out for 10 min to 13 min and the regressing surface was less 

than midway into the front, 2.5 cm thick sample (typically into the bubble layer there). 

For the polyurethane foams, about 2/3 of the 40 cm length of the foam sample was 

consumed. 

For analogous burning experiments, the test was started and generally conducted in the 

same manner except that the radiant flux levels were lower and the sample face was 

ignited as the radiation exposure began, using a multiple jet methane flame waved over 

the face. 

Three-Dimensional, Melt Pool Fire-Assisted, Fire Growth Tests. These intermediate 

scale tests were carried out under the NIST 3 meter calorimeter hood. These were 

exploratory tests, aimed mainly at examining the behavior of several compositions of 

polyurethane foam but also included some tests on a commercial polypropylene (as 

noted, it was different from the two resins used in the cone heater tests described above). 

Figure 2 shows a photo of the test apparatus. The sample, in this case, is a sheet of this 

commercial polypropylene (30 cm tall and 60 cm long). The sample has a set of grid 

lines forming squares 5 cm on a side which facilitate measuring the rate of flame 

spread19. The sample is held by two clamps, one along its top edge and one along it right 

vertical end. The sample is to be ignited on its left vertical end. Here the sample is 

suspended with its base 77 cm above a catch surface made of calcium silicate board 

(Marinite P, chosen because its thermal inertia20 is generally similar to that of wood). 

This height is the one of two used in these tests. The other was approximately 4 cm. The 

two heights were chosen to minimize the interaction (large height) between the sample 

and any pool fire on the catch surface or maximize that interaction (small height). When 

the sample was in its lower location, it rested on a fine nichrome wire “grid” formed by 

winding the wire back and forth between the two rods visible in the figure above the 

Marinite surface.21 It was still supported by a clamp on the right end, as well. When the 

sample was a slab (typically 7.5 cm to 10 cm thick) of polyurethane foam, it was mainly 

supported by impaling the right end of the slab on a set of spikes inserted into the right 

end face at mid-depth. Many of the foam tests involved a Vee configuration with two 

foam slabs held at an approximately 30o angle, each impaled on its own set of spikes (see 

 intervals, were generally held within about ± 1 1⁄2 mm; the polyurethane foam samples varied by about 

twice this due to the shape changes they tended to exhibit. 

19 Post-test measurements of flame front position, typically along the top edge of a foam slab were made 

using video tapes of the test and were accurate to better than ± 1 cm. 

20 Thermal inertia is the product of thermal conductivity, heat capacity and density. It is a measure of the 

amount of heat the material will absorb when brought into contact with another material, in this case, the 

melt from the test sample. 

21 The average spacing between successive wires was 10 cm or more. They posed no noticeable barrier to 

flame propagation on the base of any of the test samples. 
14

Fig. 3). It proved desirable to add additional stiff steel wires between the two foam slabs 

in two or more locations near the opposite (open) end of the Vee configuration to keep 

the samples in place for as long as possible as they burned away. 

The sample and the large framework supporting it all ultimately rest on a scale placed on 

the floor; this scale has a better than 0.1 g resolution. The catch surface and the pan 

containing it rest on a separate scale22; this scale has a 0.1 g resolution. In this manner, 

the mass of the sample and of the melt pool it generates can be separately measured. 

In initial tests, several configurations of fire growth were examined with one of the foam 

types (Z0 FR Foam), mostly with the close spacing above the Marinite catch surface. 

These included a single slab burned in the orientation shown in Fig. 2, a single slab laid 

flat on the wire support, ignited on one end and a single slab oriented with its long axis 

vertical, ignited on the bottom of one face. A few foam experiments were also run in the 

Fig. 2 orientation but close to the Marinite and with a radiant flux incident on one large 

face. The two-slab experiments were all in the Vee configuration described above, 

mostly with the sample close to the Marinite. 

Single slabs (both full density samples and foams) were ignited with a propane-fed Tee 

burner, visible to the left end of the sample in Fig. 2. These were fire growth studies, not 

ignition studies so for tests with the sample near the Marinite catch surface, the igniter 

was held on until a pool fire was clearly present with a plume that could begin to interact 

with the lower parts of the sample. The same time was used when the sample was raised 

above the Marinite, too high for there to be an interaction with any pool fire. In the Vee 

configuration tests (done only with foams) the igniter (of a different design) was at the 

narrow end of the Vee (see Fig. 3) and ignited adjacent ends of the two foam slabs 

simultaneously. 

All tests were recorded with two video cameras. One camera gave an overall view and 

allowed following the flame spread over the 5 cm grid system on the sample face(s). The 

other was used to get close-up pictures of the melt coagulation and flow process for 

mechanistic insights. Heat release rate23 and sample/pool weights were also recorded as a 

function of time. 

 

22 The pan was insulated from the scale with a 2.5 cm thick slab of ceramic fiber insulation. Two 

thermocouples placed on top of the scale, beneath this insulation, never showed more than 2 oC temperature 

rise during a test. We infer that the scale readings were unaffected by thermal transients from the fires 

above. 

23 There is no complete analysis of the accuracy of heat release rate measurements in this calorimeter 

available as yet. However, there is a detailed analysis of this issue for the NIST 6 meter calorimeter which 

operates in a very similar manner [23]. The heat release rate has an expanded uncertainty (95 % confidence 

interval) of 11 %. To a first approximation this same uncertainty can be applied here to HRR 

measurements in the NIST furniture calorimeter and in the cone calorimeter. 
15

RESULTS AND DISCUSSION 

Gasification Kinetics from Thermogravimetry. As noted previously, one step 

Arrhenius gasification kinetic descriptions were obtained by the Kissinger method for 

two polypropylene and two polystyrene resins. No thermogravimetry-based kinetics 

have as yet been obtained on any of the polyurethanes used in this study. However, it is 

known from previous work [7, 8] that polyurethanes are substantially more complex, 

requiring two or more reaction steps to capture their behavior (especially given the fact 

that they start out the heating process as a cross-linked polymer). 

Figures 4a , 4b, and 4c show the experimental and predicted thermogravimetric weight 

loss curves at the three experimental heating rates for PD 702N polypropylene. The 

predicted curves come from plugging the Kissinger-Method kinetic constants into the 

following rate of weight loss equation 

 ( / )/ ( / ) exp( / ) d w w0 dt A w w0 Etga R T n β = tga −

where β is the constant heating rate, ) ( / w w0 is the fractional weight of sample remaining 

(also called the conversion), t is time, Atga is the pre-exponential factor, n is the reaction 

order, Etga is the effective activation energy of the thermogravimetric weight loss 

process, R is the gas constant and T is the absolute temperature of the sample. The above 

equation has been solved using a fourth-order Runge-Kutta numerical scheme to produce 

the “predicted” curves in Figure 4. 

 

In the calculated curves for Figure 4, we have taken the reaction to be first order in the 

quantity ) ( / w w0 . A reaction order other than unity does not make much physical sense 

in the present context. In fact, in the actual experiments, the effect of reaction order 

should be indiscernible since the “reactant”, i.e., the polymer, appears to remain at full 

concentration (full density) as it regresses. However, inspection of Fig. 4 shows that, 

with a reaction order of unity, the match is best for the lowest heating rate and looks 

increasingly poor at higher heating rates as the conversion progresses past about 0.2. The 

results do still accurately catch the upward shift of the onset of the reaction with heating 

rate. 

The match between the predicted curves and the thermogravimetric curves can, on 

average, be improved by taking into account a non-unity value of reaction order. In 

effect, the values of reaction order, n, and the pre-exponential factor, Atga , interact - as n

goes down, Atga goes up to compensate; the value of Etga is not affected at all by this 

trade-off. By lowering the value of n from 1.0 to 0.85 for this particular polypropylene, 

the slope of the mid-section of the reaction rate curve is made to come substantially 

closer to the experimental behavior. This mid-range agreement tends to come at the 

expense of the early reaction rate behavior – the predicted curve with the lower value of n 

tends to be delayed in the onset of the reaction. The experimental curves cannot be 
16

matched perfectly with this single reaction model but this description should be generally 

adequate for describing the gasification reaction as it competes with other processes in 

the experiments described below24. 

Table 3, which summarizes the kinetic parameters for the four polymer resins, includes 

two sets of values for PP702N polypropylene based on the above discussion of the effect 

of reaction order. The value of Etga shifts slightly in the second set of parameters, not 

because of the change in reaction order, but because a somewhat different fit was used to 

the available data from replicate tests. The two sets thus represent two plausible versions 

of the gasification rate parameters for this polymer resin. 

Fig. 5 shows the prediction of how that reaction rate (in this case, for PS 666D 

polystyrene) shifts still further upward as the heating rates are raised up to the sort of 

level seen in the radiative heating experiments. The upward shift in significant onset of 

the reaction is substantial. The amount of the shift depends on the magnitude of the 

effective activation energy, Etga ; the larger that value, the less is the upward shift. Table 

3 shows that all of the activation energies are roughly comparable so all should shift by 

an amount comparable to that shown in Fig. 5, given equal heating rates. However, the 

effective heating rate in a radiative gasification experiment, or in more complex burning 

configurations, depends on the overall rate of mass loss from the sample. This can vary 

substantially among polymer resins so the heating rate can also vary substantially and 

thus also can the absolute upward shift of the onset of the gasification reaction. 

As noted previously, it must be assumed here that this upward temperature shift in the 

gasification reaction does not change the controlling mechanism and so the kinetic 

parameters in Table 3 are valid under fire-like conditions. 

Thermogravimetric Scan of Polyurethane Foam Behavior. Single heating rate tests 

(5 oC/min in nitrogen) were run on the various polyurethane foams used in the present 

study. These foams tended to behave quite differently in the fire tests described below. 

The present data provide one set of clues as to why this is so. 

Figures 6 and 7 are examples of the thermogravimetric curves for two of the polyurethane 

foams, NFR and P, respectively. The other foams gave qualitatively very similar results. 

There are two principal peaks of weight loss, rather than one as for the full density 

thermoplastics discussed above. As discussed further below, the first peak is due to the 

loss of the isocyanate-related moieties and the second is due to the gasification of the 

polyol. (The source of the minor intermediate peak for the P foam is unknown.) 

The relative thermal stability of the foams is of interest in interpreting their fire behavior 

results below. Actually, there are several measures of foam stability which one might 

apply. Table 4 lists four such measures. The first three come directly from the 

thermogravimetric curves: temperature at which 2 % of the sample mass is lost (as a 

 

24 See Appendix 3 for further TG kinetic analysis, using a different data reduction technique, performed 

after the bulk of this report was written. 
17

measure of the onset of foam gasification), the temperature of the first major weight loss 

peak and the temperature of the second major weight loss peak. Thermogravimetry does 

not allow observation of the sample as it is heated and one other desirable measure of 

foam stability is the temperature at which its physical structure collapses (essentially the 

temperature at which the foam appears to melt). Here this was measured by placing a 

small (few mg) sample of foam on the heated stage of the rheometer as it was 

programmed upward at 5 oC/min. The stage is enclosed by an insulated, nitrogen-purged 

oven with windows. The sample was observed during its temperature rise and the onset 

of physical collapse was noted. This extended over a range of temperatures that is 

probably a result of both kinetic and heat transfer limitations; the low end of the range 

should suffer least from the heat transfer effects. The melt initiation temperatures are 

accurate to approximately ± 5 oC. 

The first point to notice from Table 4 is that the onset of the physical collapse of the foam 

is associated fairly closely with the first weight loss peak. This makes sense – this peak 

involves the loss of the moieties which hold the polyol chains together in a rubber-like 

structure. What is somewhat surprising is that these cross-links appear to have to be so 

extensively destroyed before the foam actually collapses.25 In any event, it is noteworthy, 

in light of the fire behavior results below, that Foam P tends to collapse at the lowest 

temperature (though not by a great margin).26

Another point to note is that the first two columns in Table 4 do not give a consistent 

message about which foam is most/least prone to gasification. Foam Z” reaches 2 % 

weight loss at the lowest temperature but the graphited foam has the lowest temperature 

for the first mass loss peak. This mixed story is, at least in part, a result of the fact that 

these foams are complex mixtures of ingredients of varying volatility/stability. The early 

weight loss (2 %) could be due to a “secondary” component (in Z”, most likely one of the 

flame retardants) that is more volatile than the foam itself. Thus the first weight loss 

peak is probably a more relevant indicator of overall thermal stability and, surprisingly, 

the graphited foam is thus judged least stable. Foam P is next. 

Inspection of the placement of the first and second weight loss peaks (and the onset of 

foam collapse) for NFR, Z0 and Z” foams suggests they are all of similar composition. It 

is noteworthy that they all came from the same manufacturer and have nominally the 

same density. 

The K and L viscoelastic foams do not seem to stand out strongly in any particular way 

here (aside from having the highest temperatures for 2 % weight loss). Their fire 

behavior, as seen below, was unique, however. 

 

25 This need to reach near the peak of the first reaction may be, to some extent, a result of the way the 

sample was heated mainly from below in the rheometer experiments 

26 The graphited foam did not collapse. The recorded temperature range corresponds to that over which the 

graphite granules could be seen expanding. There was no liquid phase discernible at any point with this 

foam. 
18

Finally, it is necessary to add a caveat deriving from the low heating rate used in these 

thermogravimetry experiments. It is related to points made above – the low heating rate 

here and the high heating rate in actual burning could, in principal, lead to a re-ordering 

(or an accentuation of differences) of the various measures in Table 4. All should shift to 

higher temperatures at higher heating rates but some may shift more than others. 

Polymer Melt Viscosity. The procedure for obtaining a temperature-only depiction of 

polymer melt viscosity was described above. Figures 8, 9, 10 and 11 show the results for 

the four full density thermoplastics used in this study. Again, in all cases, the viscositytemperature relation is that passing through the indicated data for the undegraded resin 

and then through the two (or three) points inferred at higher temperatures using the 

extrapolated data from the melt samples obtained in radiative heating experiments. 

There are significant shortcomings in the process used here to get an approximate 

viscosity-temperature (only) relationship. The extrapolation of the data for the captured 

melt is linear and, in some cases, it is a long extrapolation. The linear extrapolation 

essentially assumes that the slope of the melt data being used for this purpose is a result 

of enhanced polymer chain mobility, not decreasing mean molecular weight. The fact 

that the slope of the melt data generally parallels that of the lower temperature portion of 

the undegraded polymer tends to confirm this assumption. In effect, the melt sample 

appears to have undergone as much degradation as it is going to undergo until it reaches 

temperatures comparable to or greater than those at the surface of the sample which 

generated that melt. There are degradation kinetics issues underlying this that need 

further discussion; we address this further below. Another concern in the extrapolation is 

that the temperatures we extrapolate to are themselves subject to the measurement 

uncertainties described above. Furthermore, all of this is being done on a semi-log plot. 

Given this, the melt viscosity values that are inferred, especially at the highest 

temperatures that are ultimately of most relevance to a burning polymer sample, are 

uncertain to at least a factor of two to three. This uncertainty may not seem too bad when 

one looks at the enormous ranges over which melt viscosity is changing for all of the 

polymer resins considered here. Note, however, that the part of the curve which counts 

the most in explaining/predicting fire behavior of thermoplastics is at the right side in 

Figs. 8-11 and this is where the uncertainty is greatest. 

The melt samples used for the above extrapolations were obtained from 10 cm high 

samples. The polymer melt/gasification model that uses these data as input is to be 

compared with data from samples of this same dimension so this is not a concern in that 

context. However, the same viscosity data will eventually be used to test fire growth 

models on larger samples. The reason that this could be a concern relates to the 

degradation kinetics issue, just mentioned above. The melt samples whose behavior was 

used in Figs. 8 to 11 are a product of a certain exposure time to the elevated temperatures. 

The “average” element of the melt sample originated at mid-height on the sample and 

then flowed down off the sample and cooled to room temperature. The time required for 

this depends on the viscosity of the melt at the sample surface temperature. The polymer 

degradation process may have been proceeding at a finite rate during this flow process 
19

and so its extent could depend on the length of time spent at the surface temperature. A 

taller sample, with its longer flow time could then yield a melt sample that is further 

degraded, a fact not reflected in the viscosity-temperature relations derived from Figs. 8 

to 11. 

To check for the possibility of this effect, experiments were run with both 10 cm and 20 

cm tall samples. Here it was necessary to switch from the cone heater to a gas-fired 

radiant panel to obtain a uniform flux over samples of both heights. With this set-up the 

only useable radiant flux level was 20 kW/m2

; at higher fluxes the samples could not be 

prevented from igniting to flaming rather than simply undergoing the desired process of 

passively melting/gasifying. Replicate samples of PD702N polypropylene and PS 666D 

polystyrene were tested at this flux using the two sample heights. Viscosity 

measurements were then made on the captured melts. For the polypropylene, the 

variability between replicate melt samples (a factor of about 1.8 in this case) was equal to 

the variation seen with melt samples from the two different sample heights. For the 

polystyrene, the results were basically the same – there was no clear cut effect of the 

increased sample height on the melt viscosity of the captured melts. 

In addition, for the polystyrene, gel permeation chromatography27 was applied to all of 

the melt samples and to the original resin. This technique reveals the molecular weight 

(MW) distribution of a polymer sample. It showed that all of the melt samples were 

essentially indistinguishable though clearly lower in mean MW that the original resin. 

These results imply that the degradation kinetics are proceeding slowly at the surface 

temperature – slow enough to obscure any pronounced effect of sample scale on the melt 

viscosity-temperature relationships derived here. 

Interestingly, at this 20 kW/m2 flux, the viscosity of the polystyrene melt has dropped by 

only a factor of ten and that of the polypropylene by a factor of thirty. Since the weightmean molecular weight (MW) is expected to have a 3.4 power effect on polymer melt 

viscosity [12], this implies the weight-mean molecular weight is dropping by only a 

factor of two or three. For a typical polymer, the ratio of weight-mean MW to numbermean MW is greater than one and this ratio is expected to increase in a thermal 

degradation process driven by random chain scission, as we have here. This means that 

the change in the number-mean MW will be less than the factor of two or three. In other 

words, the melt flowing off the samples at 20 kW/m2

 is still a long chain polymer. The 

melts from higher fluxes shown in Figs 8-11 drop by factors ranging up to a maximum of 

30,000 (6523 polypropylene at 40 kW/m2

). This corresponds to the weight-mean MW 

being cut by a factor of 21 and a number-mean MW cut by a lesser factor. Since these 

polymers likely start out with number-mean molecular weights in the range of a few 

hundred thousand [12], we are still dealing with long chains (> 1000 monomer units) in 

the melt and limited polymer degradation, on average (even though some volatile 

molecules are being generated)28. One implication of this is that the amount of energy 

required to break the polymer down to the point where it can readily flow (as it does from 

 

27 These GPC results were obtained by B. Cipriano of the University of Maryland. 28 When the number-mean MW drops below a critical value, the 3.4 power no longer holds. We are 

probably entering that domain at the higher heat flux levels. 
20

the sample face, especially at the higher radiant fluxes) is small. i.e., it is a small fraction 

of the heat given off during the polymerization process. 

The “melt” from polyurethane foam is the result of more extensive polymer degradation, 

at least in one sense. The foam is originally a cross-linked polymer, i.e., its molecular 

weight is effectively infinite and it cannot flow. These cross-links must be broken before 

any flow can occur. The cross-links take several forms and the proportions depend on the 

particular foam formulation. They include urethane, urea, biuret and allophanate bonds. 

There seems to be some question as to whether these last two bond types actually exist to 

any significant extent in a typical polyurethane foam [13]; in any event they are easily 

broken at temperatures well below 200 oC [14]. The urethane and then the urea bonds 

begin to break at temperatures somewhat over 200 oC by several routes [15], one of 

which (de-polymerization) can regenerate the isocyanate and active hydroxyl group on 

the original polyol. This last fact seems particularly relevant here, as will be seen below. 

The polyol itself, which typically is based on polymers of either ethylene oxide or 

propylene oxide (or co-polymers of both), degrades only at substantially higher 

temperatures, reported to be up to 375 oC [15]29. As a result of this dichotomy in ease of 

bond breakages (which leads to the separate weight loss peaks in the thermogravimetry 

discussed above), one expects the “melt” from a flexible polyurethane foam to be 

dominated by largely intact polyol; Ref. 15 shows several lines of evidence to support 

this. 

Polyurethane “melt” was captured from several foam compositions either in flaming 

experiments or in radiative heating experiments (and sometimes in both). Unlike the melt 

from the full-density thermoplastics discussed above, the captured material for most of 

the foams was a liquid at room temperature30. It was, however, significantly 

contaminated with bits of the blackened “skin” that tends always to form on the face of a 

degrading polyurethane foam. The samples were centrifuged to separate what was 

typically a light brown liquid from the (semi-)solid black particles. Figure 12 shows an 

example of the viscosity versus temperature for this liquid (here, for melt from the Z0 

flame-retarded foam). Two curves are shown – one for a melt sample collected from the 

radiation-driven melt experiments in the cone calorimeter (at a radiant flux of 40 kW/m2

) 

and one for a melt sample collected during a three-dimensional fire growth experiment 

(with flames as the only heat source)31. Much of the noise in the data is a result of the 

low viscosity levels here, though some residue of solid particles may be contributing as 

well. The behavior in the neighborhood of 200 oC is unexpected – the viscosity, which 

decreases with increasing temperature above room temperature in a more or less normal 

 

29 Note that the temperature range of significant decomposition will be pushed upward in experiments such 

as those here in which the heating rate of the polymer is higher than those in the typical polymer 

decomposition study. 

30 The captured samples were kept frozen in sealed containers in the time between the experiment that 

generated them and the subsequent rheometry. The two viscoelastic foams gave a melt that turned into a 

rubber-like material at room temperature. Rheometry on these materials could only be achieved by an 

oscillatory method. It showed the effective viscosity of these “melts” at room temperature to be ca. 10,000 

times higher than those from the other foams. 

31 The viscosity of the melt from flaming was, for some of the foams, up to ten times lower than that from 

the 40 kW/m2

 cone calorimeter radiation-driven melt experiments. 
21

manner, suddenly reverses the trend and increases to near or above the original level 

before decaying once more at still higher temperatures. 

Figure 13 is an annotated viscosity curve for the melt from a different foam (NFR, nonretarded foam). There are two unexpected phenomena seen there. First, the sample 

spreads from under the rotating “tool” that is applying the controlled shear. Such 

spreading is not unusual as a consequence of simple expansion of the sample with 

increasing temperature. What is unexpected is that by 180 oC the volume that has 

escaped from under the tool appears to be comparable to the volume that the tool can 

capture beneath itself. One might conclude that the subsequent increase in measured 

viscosity is a result of this sample migration. However, at a temperature around 250 oC 

another phenomenon that evidently was already in process becomes evident – arc-shaped 

“islands”, which stand somewhat above the surrounding liquid level, begin to break away 

from the meniscus region surrounding the edge of the rotating tool. These “islands” 

appear distinctly non-fluid and as the temperature goes still higher they appear to “melt” 

back into the liquid surrounding the tool32. In short, the increased viscosity (and its 

subsequent decrease) appears to be caused by a transient “solidification” or, at least, a 

partial gelling, of the polyurethane melt. 

As noted above, Ravey and Pierce [15] give evidence that de-polymerization of the 

urethane bonds is one degradation pathway. Breakage of both urethane bonds on a crosslink would regenerate toluene diisocyanate (TDI) which boils at about 250 oC. The TDI 

would normally tend to escape from the liquid phase; according to Ref. 15 it then is 

consumed in the vapor phase via reaction with amines generated during the polymer 

degradation. In the present tests, there is not necessarily enough time for all of the TDI to 

escape from the melt which flows from the sample and then cools. In Ref. 15 it is shown 

that residual TDI in the melt depends on the pyrolysis conditions. The above behavior 

suggests that urethane bonds are re-forming in the melt sample at about 200 oC and then 

breaking again as the sample temperature increases further. This anomalous viscosity 

behavior was seen with the melt samples from all of the polyurethane foams though the 

onset temperature and curve shapes showed some variation with foam composition. All 

except Foam L viscoelastic foam (Foam K was not measured and, of course the graphited 

foam could not be measured since it gave no melt) gave very noisy data at the upper 

temperature limits of the measurement (ca. 300 oC) but were roughly the same in 

viscosity (ca. 0. 1 Pa-s, comparable to PD 702N polypropylene at 400 oC). Foam L, 

measured in an oscillatory manner, gave results at this temperature that were 1000 to 

10,000 times higher. Since this foam (and Foam K) did manage to flow off the sample 

surface at temperatures not too much higher than this we infer that the oscillatory method 

does not provide a pertinent measure of viscosity, at least for these foams.33

 

32 Note that a few bubbles are being generated beneath the tool as the viscosity is peaking but serious 

bubbling that would definitely affect the viscosity measurement is delayed to well over 300 oC. This 

presumably reflects accelerating degradation of the polyol. 33 The oscillatory method does not directly provide a melt viscosity. One has to apply the Cox-Merz [18] 

rule to infer a viscosity and that rule is most accurate when the viscosity is low and the fluid is Newtonian. 
22

The practical implications of this behavior for modeling of fires involving polyurethane 

foam are not yet entirely clear. In particular we note that we have not obtained a measure 

of the foam melt viscosity at temperatures higher than about 300 oC and the results 

below this are odd, to say the least. Such lower temperature results are relevant mainly to 

how far the melt can spread on a catch surface and this, in turn, affects the diameter of a 

pool fire and also its plume height. It is possible that the complications seen here are 

largely irrelevant, if the re-polymerization reactions are kinetically-limited. In the above 

rheometric measurements, the heating rate was 1 oC/min; in a real fire heating and 

cooling processes are at least an order of magnitude faster and there may not be enough 

time for re-polymerization. It may be more profitable to study the viscosity-temperature 

relation of a few representative polyols in isolation since we have inferred that the melt is 

mainly composed of polyol. 

Another real complication in a polyurethane melt fire is that the melt is frequently seen to 

be bubbling, especially in the flaming areas. The flow behavior of a two phase fluid like 

this may not be readily predicted from our existing rheometric measurements. This 

remains an area in need of further study; we make further observations below on the 

behavior of polyurethane melts in burning pool fires. In any event, it is worth noting that 

the foam melts remain much more fluid at low temperatures than do the melts from the 

full-density thermoplastics discussed above. 

 

Two-Dimensional Radiation-Driven Melt/Gasification/Flaming Experiments. As 

noted above, these experiments were mainly done to provide data to test intermediate 

stages of models in the process of developing a full three-dimensional model of fire 

growth on a mass of a thermoplastic material. These experiments are essentially twodimensional in nature and, though they are initially transient in character, they yield 

steady-state results for mass loss rate by melt flow and by gasification along with the 

corresponding surface temperature. All of these data are to be compared to the Reaction 

Engineering International (REI) model of the same basic set-up. These experiments also 

provide a context in which to develop a greater understanding of the factors affecting the 

behavior of these melt/gasification systems. The processes occurring in these simplified 

experiments are the same as those occurring locally on any surface element during threedimensional fire growth over the same material. 

Appendix 1 presents a one-dimensional, steady-state model of these melt/gasification 

experiments. The model, though it is a simplified description of what is happening in the 

present experiments, succeeds in capturing the correct trends of behavior and provides 

some useful insights into what would be required to alter this behavior appreciably. 

Figures 14a, 14b and 14c show an example of the direct experimental outputs from a 

radiative melt/gasification experiment, i.e., the sample assembly weight, the weight of the 

melt flowing off of the front face of the sample and the output from the surface 

thermocouple probe. These results are all for a 10 cm by 10 cm cross-section slab of 

PS666D polystyrene heated on one face by a radiant flux of 20 kW/m2

. In Fig. 14a the 

weight of the entire apparatus on top of the scale is shown but the change is due only to 

weight loss from the sample (after the slight downward step at ca. 35 s indicating the 
23

removal of the shield on the face of the sample)34. The sample soon begins to show some 

weight loss (combined gasification and melt flow). At about 500 s it rather abruptly 

begins losing weight more rapidly as the melt layer begins flowing smoothly from its 

entire face. Soon thereafter, the rate of weight loss is essentially steady for the rest of the 

experiment. Fig. 14b shows an essentially steady rate of melt flow onto the catch pan 

after about 600 s. (When melt was being captured for later rheometry, two different 

catch pans were used. One was used for the transient period of changing weight loss and 

one was used for the steady-state period. Rheometry was done only on the melt from the 

steady-state portion of the test.) 

Figure 14c shows the rather problematical nature of the surface temperature 

measurement. As described above, the thermocouple probe is brought into contact with 

the molten surface and worked in and out slightly to wet it fully with the melt; the 

movement is then continued in an attempt to find the maximum temperature in the 

surface region. The movement is also intended to counteract the difficult visual situation 

presented to the test operator in which it is not easy to tell exactly where the probe is with 

respect to the sample surface (because of reflections from an uneven surface and 

refraction below the surface). We are interested in the maximum temperature and assign 

it to the surface.35 How many of the temperature peaks in a plot like Fig. 14c to include 

in the surface temperature estimate is a judgment call and here it gives an 8 oC ambiguity, 

depending on whether one averages just the three highest peaks or averages all of the 

peaks. This is in addition to the other sources of surface temperature measurement error 

discussed above. Note that this interacts with the extrapolation of the melt viscosity to 

the sample surface temperature, as well. 

The results of the various radiative melt/gasification experiments for four full density 

thermoplastic polymer resins are summarized in Table 5; virtually all test conditions were 

run at least in duplicate. Results for PD702N polypropylene and for PS666D polystyrene 

are plotted in Figures 15a, 15b, 15c and 16a, 16b, 16c, respectively, since these are used 

in the Appendix 1 to compare with the simplified steady-state model presented there. 

Figures 15 and 16 show essentially the same trends with increasing radiant flux on the 

sample face, though quantitatively, of course, the two polymer resins yield differing 

results. As expected, the total mass loss rate (due to combined gasification and melt flow 

from the sample face) increases with increasing heat flux, as does the surface temperature 

and the melt flow rate (not plotted). What one might not a priori anticipate is that both 

resins show a decreasing fraction of material lost as melt when the heat flux is increased. 

This is a result of a shifting competition between gasification and melt flow as mass loss 

mechanisms. Evidently the increasing surface temperature has a stronger effect on the 

gasification process than on the melt flow process. All of these trends are correctly 

 

34 The stepped nature of the weight signal is due to the rather low resolution of the available analog output 

board for this scale. The occasional upward spikes are due to the operator moving the sample forward to 

keep its face in the outlet plane of the sample holder tube. 

35 Actually, since the radiation is absorbed in-depth and there is heat loss from the surface, the maximum 

temperature is slightly below the surface [9, 10] but this tends to be at a depth that is smaller than the 

thermocouple probe diameter and would not be resolved here. 
24

predicted by the simplified steady-state model presented in Appendix 1, where there is 

more discussion of these issues. 

Table 5 shows that the two resins of differing “Melt Flow Index” from each polymer type 

did not yield greatly differing behavior. Fig. 17 shows a comparison of the total mass 

loss rates for the two polystyrenes. Inspection of Figs. 10 and 11 shows that, at 300 oC, 

PS663 has a melt viscosity that is about 2.5 times higher than that of PS666D at the same 

temperature. In Fig. 17, at 30 kW/m2

, this yields a total mass loss rate that is reduced by 

only about 20%. The limited data for the two polypropylenes indicate practically no 

change in total mass loss rate between the two resins. This resistance to change in the 

amount of melt flow is predicted by the model in Appendix 1 and is discussed further 

there. 

 

An issue raised above was whether these experiments, when done in an open air 

atmosphere, are affected by oxygen attack on the molten polymer surface. To get some 

information on this point, two tests were run on PD702N polypropylene at an incident 

flux of 30 kW/m2

, using a nitrogen purge around the sample area. This was not an easy 

set-up and it was probably not fully effective in precluding all access of air to the sample 

face. It required surrounding the space between, and generally below, the sample face 

and cone heater with an enclosure and feeding a well-dispersed flow of nitrogen into the 

base of that enclosure so that it flowed upward past the sample face. The set-up appeared 

moderately successful at its goal; the nitrogen flow certainly tended to somewhat disrupt 

the smoky boundary layer on the sample face, especially at the higher of the two flow 

rates tested. The resulting data points fall into the scatter seen in Figs. 15 at 30 kW/m2

. 

This implies that the effect of oxygen attack on the sample face was a minor factor. The 

enclosure was very difficult to work with, blocking access to the melt capture area, so it 

was not generally adopted. 

Experiments were also done with these resins in which the sample face was burning; 

Table 6 summarizes those results. The heat release rate (HRR) was measured since these 

tests were conducted in a cone calorimeter; the accuracy of HRR measurements in this 

device are within ± 10 %. The burning of the melt as it left the sample face was 

suppressed as quickly as possible though it tended to continue on a triangular “drip lip” 

of about 3 cm to 4 cm length that channeled the melt into a catch bottle. For the 

polypropylenes, surface temperature measurements here were even more challenging 

since the maximum reading of the thermocouple was not necessarily the value at the 

surface – the thermocouple was in the flame zone if it emerged from the surface. 

Generally, the temperature measurement results showed substantially more scatter. For 

the polystyrenes the technique used in all of the preceding measurements failed 

completely – it was impossible to see the thermocouple through the visually opaque 

flame so its position in the surface could not be determined. Instead, a smaller sheathed 

thermocouple (0.25 mm dia.) was placed in the melt downflow from the sample surface 

as it passed through the semi-funnel-shaped lip immediately at the bottom of the sample 

face. The thermocouple lead was approximately parallel to the melt flow. This was only 

moderately successful since it was difficult to position reproducibly in a newly-shaped 

aluminum foil funnel lip in each test. Also, since the thermocouple was now within the 
25

(still-burning) melt, it tended to measure some average of the melt flow layer rather than 

the surface temperature. In two tests with polypropylene, where both types of 

temperature measurements were possible, there was not a clear indication that one 

method gave results much different than the other though the tests were done in the most 

difficult conditions (flame flux plus 20 kW/m2

 from the cone heater) and both results 

were correspondingly noisy. In any event, this is what is reported for the flaming 

polystyrenes under the “Surface Temperature” heading in Table 6. 

Fig. 18 shows an example of the effect of the flame on the overall mass loss rate from the 

sample, in this case, PS 666D polystyrene. The net effect of the flame for this polymer 

and this configuration appears to be the equivalent of adding an additional 25 kW/m2 to 

the sample face; it does not otherwise change the physics of the problem. This heat flux 

differential is slightly lower than what one might expect here [11] and may indicate that 

this very sooty polystyrene flame is not entirely transparent to the radiant flux from the 

cone heater. The equal slopes of the flaming and non-flaming cases imply that the heat 

of gasification is unchanged by the presence of the flame; this concept is discussed 

further in Appendix 1. 

 The heat release rate (HRR) measurement results are reported in Table 7. These are data 

for comparison with a first stage flaming model, when it is available (REI model of the 

condensed phase, coupled with the NIST Fire Dynamics Model of the gas phase 

flaming). Most of the heat release rates were fairly steady and for these an average value 

is reported. A few showed a tendency to rise over the test duration; for these we report 

beginning and ending values. All of these values are roughly 20% higher than they 

would have been if there had not been burning melt on the drip lip, in addition to the 

burning on the sample face. 

In Table 7 the limited data (6523 polypropylene and PS666D polystyrene) on the effect 

of doubling the radiant heat flux from10 kW/m2

 to 20 kW/m2 show a minimal effect for 

the polypropylene and a moderate effect (roughly 50% increase) for the polystyrene. 

Note the comparison with HRR results on the same materials obtained in the normal cone 

calorimeter configuration with the sample oriented horizontally. Here the values reported 

are for the bulk of the test, i.e., before the sample is reduced to a boiling melt pool that 

gives a spike in HRR. The range shown indicates a slow increase with time; the earlier 

values are probably more relevant here since they correspond to thermally thick behavior. 

The 6523 polypropylene shows a clear dichotomy between the two types of experiment – 

the normal cone calorimeter gives more than double the HRR level seen in the 

experiments where the melt is allowed to escape by flowing off the sample face. 

Although we lack the complete data to prove it, the above observation about the effect of 

the radiant flux on HRR for polypropylene, coupled with the results in Table 7, implies 

that a similar dichotomy holds for PD702N polypropylene. This HRR ratio may also be 

the case for PS663 polystyrene (though < 2X) but it appears less clear for the case of 

PS666D polystyrene36. Of course, we expect there to be a lower HRR for all cases in 

 

36 When the results from the cone melt experiment are reduced by 20% to account for the extra flaming on 

the drip lip, there does appear to be a significantly lesser HRR in the cone melt experiments with 
26

which the melt flow escapes – it is taking a substantial amount of sensible heat with it. 

The amount of heat removed depends on the melt flow rate which, in turn, depends on the 

melt viscosity. The melt viscosity depends on the surface temperature which depends on 

a heat balance (see Appendix 1); we cannot a priori estimate which polymer will come 

out with the greatest or least heat loss via melt flow. 

Table 7 also includes some results for two mildly flame-retarded formulations. This was 

one small step in a direction that needs much further work since it is likely that 

minimizing the flammability hazards of these materials will call for an optimum 

combination of both viscosity modifiers and flame retardants. Magnesium hydroxide 

decomposes to give off water, absorbing heat as it does so. In the quantity used here it 

has virtually no effect on polymer melt viscosity. It also appears to have no significant 

effect on the HRR here, probably because the add-on level is much lower than the levels 

used in commercial FR polymers. 

Radiation-driven melt experiments were also carried out on most of the polyurethane 

foams in Table 237. The goal was to determine the differences in ease of melting among 

these foams which, as will be seen below, exhibited widely differing fire growth 

behavior. Recall that the thermal analysis curves of these foams did not differ greatly, 

suggesting that they all share similar controlling steps, as least as far as their weight loss 

is concerned. As noted above, in an overall sense, the foams first lose their isocyanaterelated constituents and then their polyol-related constituents. This does not tell us when 

they become fluid, nor how fluid. The experiments above (in the rheometer oven) 

provided information about the temperatures at which the foams collapse to liquids. The 

radiation-driven melt experiments provide direct information on how much fluid is 

produced by the collapse. 

Figure 19 summarizes the melt flow rate results for the six foam formulations which were 

tested. Reported there is the steady-state value of absolute flow rate of melt from the 

sample face (2.5 cm by 2.5 cm). The data at 40 kW/m2

 are probably most relevant to 

what one could expect in actual flaming38. Note that there seem to be substantial 

differences in response to a varied incident radiant flux among the foam formulations 

(though the data are rather sparse in this respect). The most striking result, however, is 

the roughly 10× variation in the rate of mass lost as melt among these six formulations 

(compared at 40 kW/m2

). 

Figure 20 shows the percent of mass lost as melt at 40 kW/m2

. This removes the issue of 

the differing foam densities. Now the difference among foam formulations is about a 

factor of six. There seems to be a dichotomy in behavior with the two viscoelastic foams 

 polystyrene than in the normal cone experiments. This is to be expected since fuel is escaping from the 

burning zone. 

37 The graphited foam was not tested in this mode since it showed no tendency to melt in the fire growth 

experiments described below. It also left a char-like residue on the sample surface (from the expanded 

graphite) which would have precluded a steady-state regression of the type being examined here. 

38 Very limited work was done with flaming foams in this cone calorimeter-based apparatus. Considerable 

difficulty was encountered with uneven regression of the sample face (faster at the top). It is not clear, at 

present, if these problems are inherent or can be overcome. 
27

and the high melamine foams losing by far the greatest mass as melt. The other three 

foams (which come from the same manufacturer and are probably similar in formulation) 

yield roughly half as much melt. The non-retarded foam is rather strikingly resistant to 

melting. 

In Fig. 20, for the three foams on the left, mass which was not lost as melt was lost as 

gas. For the other three foams (on the right), some mass also went into the formation of a 

black surface “skin” which was periodically removed. The mass going into this skin was 

highest for the non-flame retarded foam (NFR) and amounted to 20% of the total mass of 

the foam. 

Since we lack any details regarding the formulation of these foams, we cannot pin down 

the reasons underlying their widely varying behavior. We can point out some aspects, 

however. The “conventional” polyurethane foam studied in Ref. 15 was about 2/3 polyol 

by mass and 1/3 isocyanate. Thermogravimetry on their foam looked generally similar to 

that seen here (e.g. Figs 6 and 7). They report that the second DTG peak is primarily due 

to gasification of the polyol and it accounts for about 2/3 of the total mass lost. Those 

proportions (and, presumably the underlying reasons) are also roughly true for all of our 

foams. If the thermogravimetry behavior carries over into the radiative heating 

experiments we then expect, as a first approximation, that we will have 1/3 mass lost due 

to gasification of the isocyanate-related species and 2/3 mass lost due to the more 

thermally stable polyol left behind and then flowing off of the sample face. A glance at 

Fig. 20 shows that the foams fall either distinctly higher or distinctly lower than this 

expected value. 

In fact the picture underlying this expectation is too simple. It ignores the issue of 

whether and how fast the melt (whatever it is) can get off of the sample surface and how 

hot that surface is. Temperature measurements were made on these sample surfaces 

though the contact between thermocouple and surface material was less effective than it 

was for the full-density thermoplastics discussed above. Table 8 summarizes the 

measurement results. Focusing on the results at 40 kW/m2

, one sees that the surface 

temperatures did vary by more than 100 oC with the P (high melamine) foam giving the 

lowest result and the non-flame-retarded foam giving the highest. Comparing the 

temperature values to those seen in thermogravimetry for the same foams (Table 4) and 

even allowing that the temperature on these higher heating rate experiments are shifted 

upward somewhat, one infers that, with the probable exception of Foam P, the surface 

temperature is high enough to be partially degrading/gasifying the polyol as the “melt” 

tries to flow off of the sample surface. This might explain the lower melt yields seen 

with the three foams on the right in Fig. 20. It does not explain the higher than expected 

melt yields for the three foams on the left in this figure. Particularly anomalous there are 

the two viscoelastic foams which give high viscosity “melts” that do not easily flow off 

of the sample face. These three foams have the lowest “collapse” temperatures in Table 

4. Foams K and L also have the highest peak temperatures for the second weight loss 

peaks in thermogravimetry (Table 4), possibly implying more stable polyols. These 

combinations seem to help explain the trends in Fig. 20, but it is not clear that they fully 

explain them. 
28

We will see below that this diverse range of melt/gasification behaviors with foam 

formulation carries over into the fire growth behavior. 

Preliminary Fire Growth Experiments. A limited number of experiments has been 

done on melt-pool assisted fire growth over sheets of full-density thermoplastic (a 

commercial polypropylene with no flame retardant). Much more experimentation has 

been carried out on polyurethane foams. Here we discuss the full density thermoplastic 

results first. 

The goal here was simply to explore the phenomena and obtain some quantitative data 

that can be used to compare with the REI or particle method models when either reaches 

full three-dimensional capability. 

Polypropylene. Figure 21 sets the stage, showing a melt-pool assisted fire for a case in 

which the polymer sheet is very close to the horizontal, Marinite P, catch surface (4 cm to 

5 cm), thus bringing the pool fire plume into good contact with the sheet. Note that the 

sample is 30 cm tall and 60 cm long; the grid lines on the sample are 5 cm apart. The fire 

was initiated by igniting the full left vertical end of the polymer sheet. In the figure the 

fire is propagating to the right with the flames on the vertical sheet being assisted by the 

fire plume of the pool fire created by the polymer melt. 

Inspection of Fig. 21 shows several interesting features. The pool fire is centered near 

the rear edge of the sample, not under the leading edge of the flames on the sheet. This is 

because it is being fed flaming polymer melt most rapidly from an area several 

centimeters behind the forward-most portion of the sample flame front where the shape 

of the trailing edge of the sheet curves rapidly from near vertical toward the horizontal. 

There is a flow separation region there that tends to dump nearly all of the melt flow 

accumulated from higher up on the trailing edge of the sheet. From this flow 

impingement area on the catch surface, the melt tends to flow radially at first. That 

portion of the melt that is going forward (in the direction of flame spread) under the 

leading edge of the flames on the base of the sheet encounters a cold catch surface that 

extracts heat from the melt, lowers its temperature and rapidly raises its viscosity. This 

nearly halts the flow in this direction, which, in fact, greatly slows the potential rate of 

fire spread39. Because much of the “forward” flow of melt is inhibited and, because the 

catch surface in the opposite direction has been pre-heated by the pool fire in its march 

 

39 The forward rate of flow is limited by the rate at which the melt can heat the catch surface and so it could 

vary strongly with the thermal inertia of that surface (as can the heat release rate of the pool fire [3]). This 

“reluctance” to flow onto the cold catch surface in the direction of fire spread can be readily overcome by 

tilting the catch surface downward just a few degrees in the direction of fire spread. The result is erratic, 

however, since the front of the moving melt still tends to freeze and thus create a local dam blocking the 

melt behind it. The dam is then bypassed to one side or the other and this can move the center of the pool 

fire away from the polymer sheet causing very erratic behavior. Too much tilt will move the pool fire 

rapidly in the direction of tilt and preclude enough local pool fire dwell time to ignite (or melt) the sheet 

above it. Finally, with a sheet-like fuel such as this, a tilt in any direction away from the long axis of the 

sheet will remove the pool fire from under the sheet and defeat the self-feeding mechanism. 
29

forward, there is a preferential melt flow backward, away from the direction of fire 

spread and toward the rear end of the pool fire. This tends to somewhat disengage the 

pool fire from the overall forward fire spread process. The melt flow on the catch surface 

appears to be driven by the small hydrostatic head that develops due to the finite 

thickness of the melt layer on this horizontal surface. Near the foot of the pool fire 

flames, the flow is also driven outward, away from the pool fire center, by the surface 

tension gradient that is large in this region. (One can often see a step up in melt layer 

thickness beneath the flame foot.) Note that the region of the pool directly beneath the 

flames is bubbling, indicating in-depth generation of gaseous degradation products from 

the polymer melt. Also note that, on the left (just to the left of the flame foot), the pool 

fire has burned out by locally consuming all of the melt, leaving a dry central area40. 

Around this area, however, there is a substantial amount of melt that has not burned and 

is left by the fire. Evidently, this residue has cooled sufficiently (and remains sufficiently 

heat-sunk to the catch surface) that it will not allow flame spread onto its surface (in 

effect, its temperature cannot be raised to the point where it will ignite). 

Note that the polymer sheet itself is somewhat wavy on its rear edge (warped out of the 

plane defined by the cold portion of the sample sheet). This is a consequence of the heatinduced softening (and, perhaps, expansion) of the sheet before it actually begins to melt 

and flow at an appreciable rate. This aspect of the sample behavior was not reproducible. 

It interacts with the location of the separation point on the rear edge of the sample and 

thus influences where the bulk of the melt gets deposited in relation to the leading edge of 

the fire on the base of the sample. This appeared to be a major source of scatter in the 

evolution of the heat release rate from the fire, as described further below. 

The above processes could conceivably produce an essentially steady-state, propagating 

fire after some initial transient. Interestingly, while the flame spread rate along the 

bottom edge of the sample sheet is nearly steady in all cases, other aspects of the fire, 

including the heat release rate, are not steady. 

Figures 22a, 22b, 22c and 22d show the results as a function of time from one of the tests 

(in this case, a 3 mm thick polypropylene sheet placed 4.3 cm above the Marinite catch 

surface); this is the same condition as that for the photo in Fig. 21. 

Fig. 22a shows the overall heat release rate (HRR) which includes the heat released from 

all surfaces – the sample and the pool.41 Clearly there is a large transient effect that 

increases the HRR strongly about half-way through this test. Fig. 22b shows just a hint 

as to what is causing this. In this Figure, one sees that the progress of the flame front 

along the bottom edge of the sample is fairly steady. Progress of the flame front near 

 

40 The Marinite P is somewhat porous and some of the melt penetrates into its pore structure. Since these 

Marinite boards were used in repeated tests, the “infused” melt was burned out with a torch after every test. 41 An initial goal was to attempt to separately measure the HRR from the sample and its pool. Further 

thought reveals that this can only be done by separately measuring the mass loss rates due to flaming from 

these two burning areas. That , in turn requires that the flow of mass between the sample and its pool be 

interrupted periodically. This would be, however, highly disruptive of the overall burning process since the 

disruptions would have to last tens of seconds (given the response time of the calorimeter). Thus, given 

present technology, it is impractical to separately resolve the HRR from the sample and its pool fire. 
30

the top of the sample (along the top grid line visible in Fig. 21) seems to proceed at two 

distinct rates. The early progress is much slower than that along the bottom edge of the 

sample, but, just after 600 s, the rate of flame spread near the top accelerates to a rate 

faster than that along the sample bottom42. Figure 22c shows a related phenomenon 

which is directly tied to the increased HRR – the length of the pool fire (in a direction 

parallel to the long axis of the sample) increases strongly at about the same time as the 

top edge flame spread rate accelerates and the HRR increases. It is the increase in pool 

fire size (somewhat roughly measured by this pool fire length) which causes the increase 

in HRR. 

Fig. 22d shows the mass of the sample and the pool as a function of time. A notable 

feature here is the amount of mass still in the pool at the end of the experiment43. This 

fire spread process is more efficient at melting the sample onto the catch surface than it is 

at burning up the melt. (Again, since it is due in large measure to heat lost to the catch 

surface, this “inefficiency” will vary with the thermal inertia of the catch surface 

material.) 

The phenomenon underlying the increase in pool fire length/size is somewhat complex 

(and poorly reproducible). Recall that the fire spread process is initiated by igniting the 

full height of one edge of the vertical sheet. This purely vertical flame front is not stable. 

The bottom edge portion of the front tends to move forward faster that that above it; the 

reason appears to be related to the sensible heat content carried by the polymer melt 

flowing down the burning area44 of the sample. The bottom-most portion of the sample 

has flames on both sides, on the rear-most edge and also on the bottom edge of the 

sample. This slightly greater heat input in this location (versus any higher location), 

coupled with the heat input this region receives from down-flowing melt from regions 

above, tends to create a slight slope on the bottom edge, going downward in the direction 

of flame propagation. This slope feeds melt in this direction and its heat content 

accentuates the rate of flame propagation on the bottom edge. The amount of melt being 

fed toward the leading edge of the propagating flame front on the bottom edge of the 

sample is rather small45. As noted, above, the bulk of the melt flow from higher portions 

of the sample is lost at a separation point part way up the now curved trailing edge of the 

sample. This flow separation is presumably occurring at a location where the gravity 

vector (corrected for the local slope of the surface on which the melt flow is occurring) 

just exceeds the surface tension force that keeps the melt attached to the sample surface. 

This location is not necessarily stable since the out-of plane curling of the heat-softened 

sheet varies with time, altering the local slopes. In any event, as soon as the flame begins 

to move faster on the bottom edge46, it alters the heat transfer to those positions of the 

 

42 The reported spread rates are for the rate of movement along a horizontal line. The flame front is always 

actually moving normal to itself. If the front is nearly parallel to the horizontal line of interest (which 

happened in some tests near the top of the sample) the movement rate along that line can appear to be 

exceptionally fast. 

43 The experiment was typically ended after the sample flames had ended and the pool fire was dying down 

to a fraction of its peak size. 

44 The sample burns on its trailing edge surface and on a length of 2-3 cm ahead of this. 

45 It is the melt that is generated below the separation point. 

46 This can take ten or more minutes. 
31

sample sheet above it since that flame begins to impinge on those areas (for a limited 

height above the bottom). The heat transfer is increased in these areas due to greater 

flame contact. This can have two effects: increasing the flame spread rate at these areas 

above the bottom and increasing the melt flow rate off of the sample. The flow 

separation point also moves somewhat further back (and higher up) on the sample with 

the result that the pool fire gets somewhat longer (growing in the direction opposite that 

of the fire propagation)47. The greater area of the pool fire tends to increase its heat 

release rate. This curvature-inducing effect can be self-reinforcing so that the curved 

trailing edge of the sample becomes still more curved and all of these trends become 

more pronounced. Countering this are at least two non-reproducible factors: (a) as noted, 

the out-of-plane curling of the sample seemed random and it tends (when the curling is 

partially upward) to disengage the flames from the sample surface above the curl (b) 

random flow disturbances seen under the three meter calorimeter hood tended to move 

the flames around and also disengage them from the sample surface. The net result was 

that two nominally identical tests produced substantially differing HRR vs time curves 

such as those shown in Fig. 23. It should be noted that both of these tests were conducted 

with the sample elevated high above the pool fire on the Marinite catch surface so 

interactions with its fire plume were not a factor in the propagation process. 

Figures 24, 25 and 26 summarize the results of the six tests performed with this 

polypropylene material. The variables were sheet thickness and height above the catch 

surface. All conditions were tested in replicate; these are shown as gray and black bars 

for each test condition. 

Figure 24 shows the spread rate of the fire along the bottom edge of the sample. It is 

denoted there as the “dominant” spread rate because the flame front on the bottom always 

led that elsewhere on the polymer sheet. Note first the absolute numbers – this is not a 

rapid fire spread process; it required times on the order of 20 s to 30 s to move 1 cm. It 

appears to be dominated by the rate at which heat from the flames and from the flowing 

melt can raise the temperature of the polymer at the bottom edge of the sample. Because 

of this slow spread, both sample thicknesses are essentially thermally thin so that the 

thicker sample yields only a slightly slower spread rate. One would expect further 

thickness increases to slow the spread rate significantly (assuming all other parameters 

are fixed). There is a modest but significant effect of moving the sample far above its 

pool fire suggesting that, when it is close to that pool fire, the additional 

radiative/convective heat from the pool fire plume does accelerate the heating of the 

bottom edge somewhat. The effect is less than one might expect. The limitation appears 

to be an inability of the pool fire to move further under the sample because, as noted, 

above, the melt cools and becomes too viscous to flow in this direction on a horizontal 

surface. An implication here is that the forward progress of the melt fire on the catch 

surface is largely determined by the forward progress of the flame front on the bottom of 

the sample sheet (and not the other way around). (In effect, the pool fire is being pushed 

forward by the forward movement of the separation point on the trailing edge of the 

 

47 The pool fire may also get somewhat wider. The camera view did not provide useful data on this issue. 

The actual pool area is the most pertinent parameter in HRR. The length values were sometimes rather 

poor representatives of area since the pool shape, on the trailing end, could be quite irregular. 
32

sample sheet that dumps melt into the pool.) In turn this means that, for this particular 

situation, the fire spread process is dominated by the spread of flames on the bottom edge 

of the polymer sheet. At the same time, the fire size (as measured by its HRR) is 

determined by more stochastic factors affecting the area of the pool fire behind the 

separation point. 

 

This particular balance of factors can be skewed readily by tilting the catch surface 

slightly downward in the direction of propagation. A 3o tilt produced the erratic behavior 

described in a footnote above. A lesser tilt could be more effective if it simply 

encouraged the melt pool to move somewhat more rapidly under the unburned portion of 

the sheet at a rate that allowed the pool fire to feed itself more rapidly; then control of 

spread could shift to the leading edge of the pool fire. These are the sorts of issues one 

can explore extensively when a verified model of this rather complex fire spread situation 

is available. 

Figures 25 and 26 provide a somewhat unclear story. In Fig. 25, the 6 mm thick sheet 

clearly gives the highest HRR. The result in Fig. 24 implies that this doubled thickness 

sheet is dumping melt onto the catch surface at nearly twice the rate of that from the 3 

mm thick sheet. Fig. 26 does not confirm that the resulting peak pool size is larger, 

however. As noted previously, this peak pool length measurement was not necessarily a 

reliable measure of pool area. At this point no other measures of pool area are available 

since we did not have a camera looking along the length of the sample. Fig. 25 shows 

that the thinner polymer sheet, when far removed from its pool fire, gave the smallest 

peak HRR. Figures 24 and 26 suggest that this is due, at least in part, to a slower melt 

flow rate and a smaller pool fire; it is very likely also due in part to a reduced HRR from 

the two fire zones (object and pool fire) since they were too far apart to reinforce each 

other by radiative/convective interactions. (In fact, it is rather surprising that a material 

such as polypropylene, with its high heat of combustion and very mobile melt, gives such 

a weak fire in the case where the source object is far removed from its pool fire.) 

Polyurethane Foam. We turn next to the preliminary fire growth experiments on 

polyurethane foam. Here the goals were: (1) to explore the fire behavior phenomena and 

how they vary with foam properties; (2) to explore possible configurations for a test 

method which will give an assessment of how a foam would behave when placed within 

a real product. 

Before describing the actual fire growth experiments, it is useful to introduce data on the 

cone calorimeter heat release rate (HRR) behavior of these foams. These were standard 

cone calorimeter tests with the irradiated surface of the sample facing upward. All of the 

foams were tested at an incident heat flux of 35 kW/m2

; the 5.1 cm thick samples were 

contained in aluminum foil so that their melt could not escape. Table 9 summarizes the 

results. Figure 27 shows the HRR vs time for the NFR foam, which is qualitatively fairly 

typical in its shape (note the two exceptions in Table 9). The HRR process is divided into 

two stages. In the first the HRR is roughly constant; it corresponds to a burning process 

above a collapsing, receding foam surface. When the foam has “melted” to the bottom of 
33

the foil container, the melt proceeds to burn more vigorously as the liquid is effectively 

pre-heated48. 

Note that we expect the burning process on the foam surface of a sample in the 

experiments below to reflect (see caveat, next paragraph) the HRR given for the plateau, 

not the peak (except for the graphited foam). Table 9 reports the approximate average 

HRR value of this initial plateau and the value of the peak HRR. Also included is the 

range of computed heats of combustion obtained by dividing the rate of heat release by 

the rate of mass loss from the sample. The sample mass measurement data had to be 

smoothed and differentiated to obtain the heat of combustion. This process gave rather 

noisy results (Fig. 27 shows the raw data points and a running average through them) but 

it was clear that, in general, the heat of combustion always increased with time over the 

course of the burn. The values given are the range as the burning progressed from the 

start of the initial HRR plateau to the end of significant combustion. 

The primary purpose of these standard cone calorimeter experiments was to assess how 

the heats of combustion varied with foam composition. The incident radiant flux of 

35 kW/m2

 was chosen to assure complete burning of the flame-retarded samples. This 

flux is much higher than the samples see in the tests below so the HRR values do not 

directly apply. We assume here that the relative HRR values (for the plateau) apply, but 

the experimental situations with their much lower heat fluxes may actually accentuate 

these relative differences. 

Table 9 shows that there are significant differences in the plateau HRR as a function of 

foam formulation. The flame retardant(s) in Z0 and Z” lower both the initial heat of 

combustion and the plateau value of HRR that goes with it (compared to NFR, which is 

probably a similar, non-retarded composition). In this table, it is difficult to distinguish 

the behavior of these two FR foams though Z” contains more retardant materials (Table 

2). The viscoelastic foams appear to be slightly lower in their plateau HRR than NFR but 

their heats of combustion are essentially the same. The P foam, with its high melamine 

content, resists ignition for more than 300 s at this heat flux. This is much more than the 

time necessary to melt the foam down to a pool, so we infer that ignition requires that 

much of the melamine be driven out of this pool during continued heating. Melamine has 

a high melting point (354 oC) but sublimes at temperatures above 200 oC; the 

endothermicity (230 cal/g) and oxygen dilution effects of this sublimation are part of the 

mode of action of this retardant. When it does finally burn, the HRR peak is not 

particularly low and the heat of combustion of the residue is typical of the non-retarded 

foams in the Table. The graphited foam exhibits a HRR vs. time curve qualitatively 

similar to charring materials such as wood – an early HRR peak is followed by a 

monotonic decay. This is consistent with its physical behavior – the heated surface does 

not melt or recede but rather turns black and fuzzy as the graphite expands. Note that the 

peak is quite low compared to the other foams. 

 

48 This “pool” burning is sensitive to heat loses to the substrate below the sample. In the cone calorimeter, 

this substrate is a ceramic fiber insulation. In the fire growth experiments below, the pool fire occurs on the 

Marinite P surface which is a significant heat sink, slowing the HRR. See also the caveat in the text about 

the differing heat fluxes. 
34

 

The medium scale fire growth experiments on these foams were carried out in the same 

apparatus as that used for the polypropylene sheet experiments described above (shown 

in Fig. 2). As described above, they included a few different configurations. The 

ultimate configuration of principal interest for its test method potential is the 30 o Vee 

that uses a pair of foam slabs. Some interesting results were obtained with other 

configurations, however, and they are summarized here. Note that all of the foam 

samples were thermally thick (7.6 cm or 10 cm thick) in contrast to the thin sheets of fulldensity polypropylene. 

The manner in which melt flows from a vertical polyurethane foam surface is very 

different than that seen with the full density thermoplastics. This was evident in the cone 

calorimeter, radiation-driven steady feed experiments, but it is much clearer here since it 

is stretched out in space over the sample surface. Figure 28 shows a close-up view of the 

area ahead of and behind a laterally-advancing flame front on a single vertical slab of Z’’

foam (there is a melt pool fire below and to the left of the area shown in the figure; the 

overall fire is moving from left to right in the photo). Again, the grid lines on the 

unburned portion of the sample surface are 5 cm apart. The flame front appears to be 

attached to the sample face along a line roughly connecting the points where the 

horizontal grid lines disappear. To the right of this there is some darkening of the foam 

surface which could be due to soot deposition or to a limited amount of foam cell 

collapse due to pre-heating from the flickering flame (or both). Immediately to the left of 

the flame attachment line the foam surface is covered by a pattern of fine dots. These 

dots grow and their numbers decrease as one moves further to the left behind the flame 

front where the burning process is penetrating further into the depth of the foam. This 

trend continues until it is apparent the dots are coagulating melt. When the local melt 

drop size is sufficiently large, the melt begins to flow down the face of the foam, forming 

the dark, continuous rivulets seen there. The melt can be seen dripping from the foam at 

two closely-spaced locations about 7 cm behind the flame attachment line (near the 

bottom of the photo.). 

The behavior seen in Figure 28 is largely a consequence of the very low density of the 

foam (it is roughly 97 % to 98 % open space). When it is heated by the flame front, the 

foam cell structure collapses yielding a much smaller volume of “melt”. As the flame 

heating continues and the local temperature increases, a portion of this melt is gasified; 

that portion varied strongly with foam composition in the cone calorimeter, radiation 

driven, steady feed experiments described above (Fig. 20). The non-gasified portion of 

the melt accumulates in the outer portion of the remaining cell structure. One might 

expect that this melt, being fully chemically compatible with the foam surface, would wet 

it and be pulled into the foam structure by capillary forces, disappearing below the 

sample surface. While wetting seems to occur, the capillary forces seem insufficient to 

pull in the melt49. This may be a result of the rapid decrease in temperature below the 

 

49 In a configuration in which the melt is accumulating on top of a burning layer of polyurethane foam, one 

does see a tendency for melt to ultimately flow down through the foam and out the bottom. This probably 

requires the “pre-heating” of the lower foam that occurs as the thermal wave approaches the bottom 

boundary of the foam mass so that the melt can stay fluid enough to pass through this hot foam layer. 
35

foam surface; the increased melt viscosity may make it essentially unresponsive to 

capillary action so the melt stays very near the surface. It is essentially stuck there until 

the melt droplets grow by agglomeration with their neighbors and with more droplets 

being formed by the local recession of the foam surface as the heating continues to force 

it to melt and/or gasify. Finally the melt droplets reach a size where gravity overcomes 

their attachment to the outer portion of the foam structure via surface tension and they 

begin to run downward. That size appears to be roughly 3 mm for the foam in Fig. 28. 

The rivulets form and flow on a very uneven surface consisting of the partially and fully 

collapsed foam cells50. This is a much more difficult flow situation to model than the 

simple laminar shear flow used for the full density polymers in the Appendix 1. Note 

that, for this foam, Fig. 28 shows only about half of the foam surface is covered by the 

flows, adding further complexity to a description of the heat transfer to and within the 

condensed phase51. 

Other foams, with their differing melt vs. gasification tendencies and differing melt 

viscosities, look qualitatively similar but quantitatively different from that in Fig. 28. 

Figure 29 provides an overall view of the same sample as is seen in Fig. 28; Fig. 30 

shows a similar view but the sample there is Foam L, a viscoelastic formulation with no 

flame retardant (Table 2). Both are burning in a configuration fundamentally similar to 

that used for the full-density thermoplastics – a vertical slab (here both are 7.6 cm thick), 

close to the Marinite P catch surface, ignited on one vertical end. Both have fires 

propagating to the right with the accompaniment/assistance of a melt pool fire. The Z’’

foam, with its halogen-based flame retardant, is showing signs of influence of its flameinhibiting effects. The foam is burning preferentially on the top surface of the vertical 

slab where the residence time for gas phase oxidation reactions is greater than on the 

sides of the slab. The flames on the sides probably require the constant re-ignition 

provided by the pool fire in order to overcome (in a weakened manner) the gas phase 

flame retardant. 

It is apparent that the pool fire in Fig. 30 is larger (ca. 50 %) than that in Fig. 29 even 

though more of the sample has been consumed in Fig. 29. This is a result of both the 

higher density of the foam in Fig. 30 and the higher melt fraction for that foam as 

measured in the cone-based radiation experiments described above (see Fig. 20). The 

reach of the pool fire is higher, as well. Despite this and the lack of flame retardants in 

Foam L of Fig. 30, there is very little flame attached to the sample itself (see the lower 

trailing edge of the foam). Its shape suggests that it is reacting in a mostly passive 

manner to heat from the pool fire plume. Both viscoelastic foams (K and L) showed this 

type of behavior in further tests described below. 

The results for the above type of test configuration, using varied polyurethane foam 

formulations, are summarized in the first two pages of Table 10 in terms of peak HRR 

and average flame spread rates along the top and bottom surfaces of the foam slab. All 

 

50 The cells in most of these foams were on the order of 1⁄2 mm in diameter. 51 Some appropriate, larger-scale continuum description (probably with empirical parameters) of this 

finely-detailed flow behavior is needed. 
36

tests with this configuration gave transient behavior, i.e., the fire growth process was not 

steady and the reported spread velocities are averages over a 25 cm length away from the 

sample ends. The overall HRR peaked (and that is the value reported in Table 10) but the 

peak was not associated with early (ignition-dominated) or late (end of sample) behavior 

(NFR foam was an exception, as described below). 

In the early tests, the intent was to cover the Marinite P catch surface with a thin layer of 

aluminum foil for two reasons: (1) it would halt any polymer melt infusion into the pores 

of the Marinite52 and (2) it would greatly facilitate clean-up between tests. The first two 

tests in Table 10 were done with foil covered Marinite. Two unrealistic effects of the foil 

cover were seen. It tended to form a low “ridge” around the pool fire due to thermal 

expansion of the foil; this inhibited the free movement of the polymer melt. The foil also 

greatly boosted the tendency of flame to spread on the lower surface of the foam sample 

(and the speed of that spread); this had a substantial effect on the peak HRR, as can be 

seen in Table 10. Evidently the high infrared reflectivity of the foil boosted the net heat 

flux impinging on the bottom of the sample, overcoming much of the inhibiting effects of 

the flame retardant in the Z0 foam. Since both of these effects are unrelated to the 

conditions the foam would encounter in the real world, the use of the foil was halted53. 

All of the polyurethane foams in Table 2 were subjected to this test of lateral spread on 

an up-standing foam slab with a maximum chance for pool fire interaction facilitated by 

the close spacing above the Marinite. As Table 10 shows, the results varied widely. Two 

of the foams (P and Graphited) would not sustain post-ignition flame propagation in this 

configuration. Foam P readily yielded a melt but neither it nor the sample itself sustained 

burning after igniter flame removal (recall its long ignition delay in the standard cone 

calorimeter tests described above). The melt pool on the Marinite would not sustain 

ignition after repeated applications of a propane torch flame (in a situation where it could 

not flow away from the igniter flame as it had flowed off of the sample). The graphited 

foam sample sustained local flames for a few tens of seconds after igniter removal, but 

these died out without spreading onto unheated portions of the sample. This foam 

produced no visible melt and simply appeared to retain its original shape while becoming 

black and “fuzzy” with expanded graphite granules (exactly the same behavior as in the 

standard cone calorimeter tests above, although there the continued radiant flux forced 

continued burning)54. 

At the opposite extreme of behavior was the NFR foam, a conventional formulation with 

no flame retardants. Table 10 shows somewhat variable flame spread rates but some 

values more than quadruple the spread rate seen on the Z0 foam (similar density and 

same manufacturer but with a halogen-phosphorus retardant). Flame spread over the 

 

52 There is apparently no high temperature-resistant, medium thermal inertia material available which is not 

somewhat porous. 

53 The flammability-enhancing effects of aluminum foil could prove useful in other contexts to boost the 

severity of a test, e.g., by simulating radiative feedback from above in a room fire environment. It is a very 

inexpensive substitute for external radiation, though, of course, it can only make the burning process more 

adiabatic. 

54 This foam may have a pronounced tendency to smolder after it has been flaming. This issue emerged 

more clearly in other tests described below. 
37

surfaces of the NFR foam was so fast that it left the melt/flow process roughly 15 cm 

behind the flame front55. As a result, the melt pool fire had no participation in the flame 

spread process over the foam surfaces. This rapid surface flame spread has an important 

corollary – it increases the burning surface area on the sample. This, in turn, raises the 

HRR of the sample. This is the one foam for which the HRR peak would have been 

higher if the sample were longer; flame spread reached the end of the sample about 15 s 

before the HRR peaked. Clearly, a longer sample would have allowed more area to be 

burning at a given time. Since the burn-out process at the ignited end had yet to be 

reached (the sample itself was gone but its pool fire on that end was still substantial), the 

peak HRR for a sample of unlimited length could have been perhaps 50% higher. At the 

time of the reported HRR peak, a visual estimate indicates that 3⁄4 of the burning area was 

on the sample and 1⁄4 was in the pool fire56. 

The Z0 and Z” foams both have flame retardants, with Z” having slightly more of the 

same halogen-phosphorus compound as Z0 plus a modest amount of melamine. Their 

strong effect on reducing the flame front velocity is much more than one might expect 

given their modest effects in the standard cone calorimeter tests (at 35 kW/m2

)

 above. 

Opposed flow flame spread of this sort is simply a process of successive ignitions of 

surface elements as a result of heat input from the close approach of the flame foot 

(especially in a foam with its poor thermal conductivity). The halogen (and perhaps the 

phosphorus) content of the retardant inhibit the oxidation kinetics of the flame. Evidently 

they increase the flame stand-off distance significantly and thereby help slow the flame 

spread rate. The small amount of melamine adds modestly to the endothermicity in the 

condensed phase at the flame foot. 

As noted above in connection with Fig. 29, gas phase flame retardants such as those in 

Foams Z0 and Z”, by slowing the oxidation kinetics in the gas phase57, inhibit flame 

spread onto the vertical surfaces of the foam slab. A major difference between the flame 

spread on the NFR foam and on the Z foams is the ready spread on all surfaces of the 

NFR foam and the predominance of spread only on horizontal surfaces (preferably the 

top58) for these flame-retarded (FR) foams. The boundary layer residence time of gas 

phase species is evidently insufficient for the FR foams to anchor a flame on surfaces 

where the upward buoyant flow velocity is high. This is a Damkohler number59 effect 

and it presents a challenge in correct modeling of the gas phase behavior during the fire 

growth process. Another implication of the above behavior is that these foams do not 

have enough of this gas phase retardant to squelch flames in a wake zone such as that on 

 

55 Recall that this foam produced the lowest melt fraction in the radiation-driven melt/gasification 

experiments. This clearly helps slow the melt accumulation process on the surface, separating the flame 

front from the melt run-off location. 

56 We cannot directly translate this to the assertion that the HRR contributions were similarly distributed. 

The burning rate (and HRR) of heat-sunk pool fires tends to be suppressed [16] and, in any event would not 

be expected to equal that on the thermally-thick burning area of the foam slab. 

57 They do this presumably by inhibiting free radical reactions, thereby cutting the net heat release rate and 

the local temperature upon which the overall kinetics depend. 

58 Spread on the bottom of the FR foams is enabled by the pool fire as a pilot source. 

59 Ratio of a flow time to a chemical reaction rate time. 
38

top of the slab (or in the pool fire). If they did, they would resist post-ignition flame 

spread in this configuration60 as well as the P and Graphited foams. 

The observed effects of these two retardants on the overall fire growth processes here do 

not follow the expectation one would have from the relative amounts of flame retardants. 

Note that Z” has twice the peak HRR in Table 10 as does Z0. This higher peak HRR was 

a result of a seemingly more facile spread of flames on the bottom of the sample for the 

Z” as compared to the Z0 foam (though the actual spread velocities are essentially the 

same in Table 10). Note that the flame spread on top of Z” was slower than that for Z0 as 

one would expect from the relative flame retardant levels. The pool fire was definitely 

larger for Z” and this aided the amount of sample area burning on the bottom of the 

sample. Though the available melt viscosity data for these foam melts are poor in quality 

at the temperatures of interest, the viscosity for the Z” foam appears to be somewhat 

lower than that for the Z0 foam and this may help explain the larger pool fire it generated. 

Whether the melamine in Z” is responsible for this is not clear. It is interesting to note 

that, by slowing the surface flame spread process on the sample, the retardants in Z” and 

Z0 do appear to enhance the relative role that the pool fire plays in driving the fire growth 

process forward. This implies (not unexpectedly) that one must deal with both aspects of 

fire growth in attempting to control it. 

The viscoelastic foams (L and K) yield a fire growth process in this configuration that is 

clearly dominated by the melt pool fire. They burn vigorously in their self-generated 

melt pools and attach flames weakly on the foam surface, with the aid of the heat input 

from the pool fire plume. (This is demonstrated more convincingly below with the Vee 

tests.) The entire shape of the sample is dominated by the fact that the pool fire leads the 

lateral spread process. (The higher top velocities in Table 10 are deceptive in this regard; 

they are a result of the top playing catch-up with the bottom when the pool fire flames got 

large enough to really boost flame spread near the top.) Unlike the polypropylene sheet 

discussed above, there is not necessarily a concentrated melt separation location on these 

foams; they can feed melt to the pool fire over a 20 cm to 25 cm length (this grew with 

time). Of course this means that most of the melt is fed to the pool well behind its 

leading edge, away from the lateral fire growth process (as was the case with the 

polypropylene sheet). Close inspection of the fire spread process strongly suggests that 

for most of the test the forward movement of the melt pool fire is prodded by flaming 

melt dripping into the pool at the leading edge of the flames on the pool (a non-burning 

melt zone extends about 2 cm ahead of this). That is, the two fronts are very closely 

coupled in their spread. As the fire grows, the rate of melt flow into the pool grows and, 

near the end of the test, it appears that the overall melt flow to the pool is starting to push 

the pool and its leading edge flame front ahead somewhat faster than the flame front on 

the base of the sample (in the hydrostatic flow manner inferred above for the 

polypropylene sheets)61. 

 

60 They might still allow spread in more adiabatic configurations such as the Vee studied below. 

61 Two factors complicate the pool burning process more than with the polypropylene or polystyrene: (1) 

The melt, especially for these viscoelastic foams, seems to swell in depth as it burns suggesting a foaming 

process. This would de-couple the upper portion of the melt from the heat sink provided by the Marinite 

below and, conceivably could boost the burning rate. (2) Burning of the melt leaves a swollen char residue. 
39

Two other single slab fire propagation situations were explored to see how they 

compared with the results above. Foam Z0 gave faster flame spread and a substantially 

enhanced peak HRR when it was laid flat rather than oriented vertically as in the 

preceding tests. This orientation puts more of the mass close to the catch surface and 

allows a stronger interaction with a larger melt pool. It also puts more of the slab surface 

on top where flame spread is more readily sustained on this FR foam. However, the 

overall behavior was also hard to reproduce and very obstructive of viewing the pool fire 

dynamics. The NFR foam, which did not appreciably interact with its own pool fire 

when vertical, did much the same here giving comparable flame spread and peak HRR 

results. 

Foam P was tested here (and above, see Table 10) with a wrap of a medium weight 

polypropylene fabric because the present senior author had seen high melamine foams 

give serious fires when combined with non-flame-retarded materials in prior projects. 

Polypropylene fabric is generally regarded as quite challenging to a foam in furniture 

flammability tests such as California Technical Bulletin 133. The horizontal orientation 

for a fabric-wrapped Foam P cushion clearly gave a more serious fire than did the vertical 

cushion, aided by the flaming melt drip from the fabric onto the pool surface but, overall, 

the spread and burning processes were very erratic. 

Foam Z0 was also examined with the slab having its long axis vertical and the igniter at 

the bottom of one large face. This gave a very rapid (igniter assisted) vertical fire growth 

process on the ignited face of the sample (with no melt pool having time to develop). 

After this, however, the burning process became quite erratic, interacting strongly with 

exactly how the sample was held. The results were not reproducible. 

Finally, three of the foams were subjected to external radiation from a large area, electric, 

radiant panel. The flux was not highly uniform (ca. ± 12 % about an average of 

7.5 kW/m2 ; it was higher on the upper half of the sample) since these preliminary tests 

were aimed mainly at determining how the foam would respond to a steady radiant flux 

that preceded the flame front. (Would this flux level62, which is about equal to the 

minimum flux for ignition of some unretarded foams, cause the foam to melt and shrink 

before flames reached it?) 

As one would expect, the presence of this external heart flux heavily biased the fire 

growth process to occur fastest on the irradiated side of the sample. In no case did the 

foam surface show any degradation or shrinkage before the flame front reached it but that 

front was strongly accelerated by the flux as Table 10 shows (compare Z0 in tests 2, 3 

and 5 with test 49; compare Z” in tests 23 and 24 with test 51. Note also that flame 

 The two issues may be related by the chemistry that is transforming the melt as it burns. Any relation to 

the complex viscosity-temperature curves discussed above is unclear. None of this is apparent near the 

leading edge of the pool fire as it aids in the fire spread; however, anything that affects the pool size and 

burning rate affects the plume height and thus its contact with the sample which is feeding the pool. This 

may pose a substantial modeling challenge. 

62 This level of flux is relevant to the type of radiative interchange that can occur among surfaces on a piece 

of somewhat fire resistant upholstered furniture. 
40

spread was sustained on the graphited foam in the presence of this rather low heat flux 

though the resultant HRR was quite low. This graphited foam had flame spread over the 

entire radiated surface followed by some weak sustained flames on the top of the sample 

during a ten minute exposure. The flames ultimately died back but the sample went into 

a smolder mode which was difficult to extinguish, apparently because the expanded 

graphite limited access of the water used for extinguishing to the smolder zone beneath it. 

These results did confirm, as expected, that relatively low levels of radiant heat flux can 

substantially enhance the flammability behavior of flame-retarded foams. These flux 

levels could be introduced into a foam flammability test method with some difficulty and 

expense via an electric-powered radiant panel such as that used here or they can be 

produced by configurations that provide radiative feedback among surface elements in 

much the manner that occurs in such polyurethane foam applications as upholstered 

furniture. With this in mind we focused further tests on a configuration that can provide 

spatially-varying levels of radiative feedback – two foam slabs arranged in a Vee (see 

Fig. 3). The point of a non-parallel arrangement such as this is the same as in ASTM 

1354 (Lateral Ignition and Flame Spread test, known as the LIFT test). The configuration 

provides a decreasing feedback flux as flames move outward from the base of the Vee 

where ignition occurs; this provides one basis for differentiating the performance of 

differing foam formulations which have differing responses to an added heat flux. Unlike 

the LIFT test and unlike the radiant panel tests above, this two slab configuration63

provides self-generated feedback. This is more like the actual situation for an object such 

as an upholstered chair. This system generates and responds to its own radiation rather 

than responding to a rigidly fixed external flux. 

Note that a two-slab configuration of this type could have any angle between the slabs, 

from 0 o to 180 o. An angle of 0 o (parallel flaming surfaces at some spacing) gives a 

nearly non-variant (except near the slab edges) feedback flux which is at a maximum. 

The sample is effectively rendered adiabatic because all radiation from the flames and the 

burning surface is returned. (This does not correspond to any fixed level of external 

radiation; the returned flux depends on the surface temperature and the 

temperature/emissivity of the flames.) Here we have chosen an angle of 30 o as one 

which is relatively harsh in terms of feedback but still open enough to allow following 

the progress of the flames on the inner surfaces of the slabs.64 The angle can be viewed 

as a potential parameter of any foam test method along these lines. 

Another parameter of such a test is the spacing between the bottom of the sample and a 

melt catch surface. In the tests here this was varied at the same extremes as were used for 

the polypropylene sheet tests, i.e., either very close or very far. Still another parameter is 

the nature of the catch surface; this has been shown to have a potentially strong effect [3]. 

We have used Marinite P throughout this test series because it is inert and comparable to 

 

63 In principle the second slab could be replaced by a front surface mirror but combustion deposits on its 

surface would very quickly inhibit its function. 

64 An angle of 0 o would render the test more harsh but also more go/no-go in nature. It is unlikely that 

foam in furniture would be subjected to such a configuration. The 30 o angle is probably already 

conservative in this regard. 
41

wood in its thermal properties. The implicit assumption is that the residential soft 

furnishing fires of primary interest here occur over wood floors. We avoid wood itself 

because it would introduce a need for a standard material that would be difficult to 

obtain. 

Figures 31 and 32 show examples of two foam behaviors in the 30 o Vee test with the 

sample close to the catch surface. The foam in Fig. 31 is NFR (non-flame-retarded); that 

in Fig. 32 is L (viscoelastic, also without flame retardant). Fig. 31 illustrates one of two 

non-idealities which showed up in these preliminary tests. The flame has surged ahead 

on the right hand slab, especially on its top. This occurred on at least three tests in the 

series but only on one test day and it always occurred on the right hand slab65. We infer 

that it was due to some intermittent asymmetry in the hood flow, but the cause could not 

be identified. The other non-ideality that was more common, especially in the earlier 

tests of the series, was collapse of the foam slabs, sometimes before the HRR had peaked. 

This was simply the result of a loss of support when the material at the small end of the 

Vee burned away (the slabs were impaled on spikes at that end). As the series went on, 

various added supports were explored. The three wires between the two slabs seen in 

Figs 31 and 32 were reasonably successful and they did not in any way interfere with 

flame spread.66 

Examination of Figs. 31 and 32 shows that, as in the single slab tests with these two 

foams, there are large differences in the role of the melt pool fire. The NFR foam has a 

minor, trailing pool fire which has little to do with the forward spread of flames over the 

slabs. The L foam has a substantial pool fire which is advancing at the same rate as the 

fire on the bottom of the slabs. This suggests that the same growth controlling 

mechanisms inferred above for lateral spread on a single slab apply here as well though, 

in most cases, the absolute behavior is modified by the new configuration. 

The flame spread behavior exhibited in these tests (as measured by the rate of movement 

of the front on the inner surface, upper left edge) covered nearly the full gamut of 

possibilities: constant, decelerating, decelerating then accelerating and accelerating. 

Figures 33, 34, 35 and 36 show these variations, which were a result of the varied foam 

composition. 

Figure 33 shows a result for the NFR foam. The jumpiness of the velocity values is 

mainly due to the use of two point derivatives to infer velocity from flame front position; 

the average result is essentially a constant spread rate. Here the flame front tended to 

race ahead of the pool fire, as noted, and in both tests with this foam the flames surged 

forward on the right hand slab, probably causing some artificial speed-up in this left hand 

side result as well. A second test gave a similar but noisier result. Comparison of the 

spread velocity in Fig. 33 with those in Table 10 (single slab) shows, rather surprisingly, 

no significant difference (though both sets of data are noisy). 

 

65 We report only flame spread data for the left hand slab but, of course, there is some influence when the 

flame surges ahead on the other slab. 

66 Further development of this approach into a potential test method requires improved sample holding. 
42

Figure 34 shows a result for the Z” foam (flame retarded). There is a clear slowing of the 

spread velocity as the flame front proceeds away from the narrow end of the Vee. A 

replicate gave somewhat lower initial numbers and still greater decrease with distance out 

from the corner. In the replicate, the velocity dropped to the levels seen in Table 10 

(single slab) as the flame front reached the outer end of the Vee; in Fig. 34, the velocity at 

the outer end of the Vee is still appreciably higher than that for the single slab case. 

Figure 35 shows a similar result for the Z0 foam (flame retarded). Note that the foam 

slabs used here were longer than the Z” foam slabs in Fig. 34. Comparing Figs. 34 and 

35 (in the latter, only the two curves marked “close to catch surface”), one sees similar 

rates of flame spread and similar decreases in the flame spread rate, out to about 40 cm. 

Beyond this, the longer Z0 samples show an acceleration of the flame spread rate. This is 

a result of the increasing role of the pool fire, trumping the effect of decreased radiative 

interaction between increasingly separated surfaces. Recall from Fig. 20 that these two 

foams yielded comparable melt mass fractions with Z” being slightly greater. Here the 

somewhat greater amount of melt from Z” is reflected in the fact that the two melt pools 

(one from each slab forming the Vee) merged into one roughly 20 s to 25 s sooner for this 

foam, as compared to Z0. (That growth differential was reflected also in a somewhat 

greater mass of Z” melt on the catch surface at any given time.) However, Foam Z” 

contains more flame retardants and this presumably weakened the effect of this somewhat 

more rapid growth of the lateral dimensions of the pool fire. Tests with both foams gave 

increasingly rapid forward spread of the pool fire (as the test progressed) and it is quite 

probable that, had the Z” samples been longer67, the same reversal in flame spread rate 

would have been seen. 

 

A “hidden” factor affecting the radiative feedback between the two slabs in this type of 

Vee test (particularly for foams with a moderate amount of flame retardant such as the Z 

foams above where, as noted, the flame zone tends to be predominant on upward facing 

surfaces) is the shape of the actively burning area of the foam. This is not simply an 

exchange of radiation between two vertical surfaces at a 30 o angle; the hottest surfaces 

(and flame attachment regions) are on the receding portions of the slab (top and trailing 

edge, which tended toward a continuous, upward sloped surface). The shapes were not 

well-monitored in these preliminary tests and they are not easy to record through the 

semi-opaque flame. An available photo shows that the burning surfaces open out sharply 

behind the advancing flame fronts thereby greatly diminishing their mutual radiative 

view factor (and pointing them more toward the pool fire plume between the slabs). 

Figure 36 shows the same type of flame spread data but here it is for the L viscoelastic 

foam. The entire trend of the spread rate with increasing spread toward the open end of 

the Vee is the opposite of what one would expect from a dominance of radiative 

interchange between the two slabs. Results for the other viscoelastic foam (Foam K) 

were very similar. Comparison to Foam L results in Table 10 (single slab) shows that 

here the spread velocity starts at the single slab level and goes monotonically upward as it 

progresses. The spread process here is dominated early on by the pool fire which grows 

rapidly due to the strong propensity of this foam to “melt”. In contrast to the NFR and Z 

 

67 Longer samples were not available. 
43

foams above, the leading portion of the flame spread on the foam slab surfaces here is on 

the bottom, reflecting control by the pool fire spread process (see below for a more 

definitive proof of this pool fire control). 

Foam P was not expected to yield a propagating fire by itself in this configuration. 

However, an ad hoc test was made with this configuration, including added materials of 

the type found in soft furnishings. Thus the two foam slabs were wrapped with the same 

medium weight polypropylene fabric as was used above in the single slab test of this 

foam. In addition, the area on the Marinite, beneath the Vee, on which a pool fire would 

tend to form was covered with a double layer of cotton-polyester shoddy pads68 (33 cm x 

38 cm by ca. 4 mm thick). The pads are themselves flammable and tend to char; they 

form a porous layer that can hold and pilot the burning of any polyurethane foam melt. 

The resulting test gave full fire propagation along the interior of the Vee surfaces largely 

without burning through the sample slabs themselves. This and the subsequent burning 

strongly suggest a fire in which the flames on the polypropylene fabric and on the top of 

the shoddy pad melted foam onto the shoddy pad where a significant fraction of the melt 

did burn. Roughly half of the original mass of organic materials remained after the fire. 

The fire was modest in size – 35 kW at its peak. This result does make a point, however: 

quite fire-resistant foams placed into the context of other flammable materials, can make 

a significant contribution to the overall burning process69. In assessing the flammability 

of such foams, it is necessary to factor in this sort of response. This issue is discussed in 

more detail below. 

The graphited foam did show a limited amount of flame propagation (ca. 22 cm), 

especially along the top of the foam slabs while the igniter was on (this is more than four 

times that seen with the single slab configuration) but it spread very little more after the 

igniter was turned off. The weak flaming on the inner surfaces of the Vee (a few kW) 

provided little radiative interchange with which to force further flame front progress (and, 

again, there was absolutely no melt flow from the samples). The sample proceeded to 

weakly flame and then smolder. The smolder became so intense after about 15 min 

(presumably because it was boosted by the radiative interchange) that it caused a brief 

episode of renewed flaming. This alternation between smolder (predominantly) and 

flaming appeared to be capable of proceeding cyclically until the bulk of the sample was 

consumed. Note that this smoldering tendency is not a real problem in the aircraft 

context for which this foam was developed, but it could prove problematical for 

residential soft furnishings applications. 

A few tests were run with this Vee configuration in which the sample was elevated 

sufficiently far above the catch surface (77 cm) as to remove it from the heat of any pool 

fire plume. The goal was to clarify the role that the pool fire was playing in the 

preceding tests where it was much closer to the sample. Flame spread rate results for the 

Z0 foam are shown in Fig. 35. Given the rather noisy data, it appears that prior to about 

the 40 cm point the pool fire had no effect on the fire growth rate along the top edge of 

 

68 These shoddy pads are commonly used in mattress construction, immediately above the inner spring. 

69 It also makes another point: such material assemblies are very likely always going to burn to some 

extent; the goal must be to limit the peak burning to a level that poses a minimal life threat [17]. 
44

the sample. The behavior at the bottom of the sample was substantially different, 

however, with the large-spaced configuration slowed to about 1/3 the spread rate of the 

small-spaced configuration. All of this had a dramatic effect on the peak HRR as seen in 

Table 11. The roughly factor of five decrease in peak HRR is a result of both a slowed 

rate of overall fire growth and a substantially decreased contribution from the pool fire. 

Although we could not get a separate measure of the HRR contributions from the sample 

and pool fire, there were very evident visual differences as to the fate of the melt for the 

two separation differences. When the sample was close to the catch surface, the melt fell 

directly below it, establishing a local melt pool fire under the end of each foam slab 

which sustained burning within seconds70. By the time that the flames on the foam slabs 

were reaching the open end of the Vee, the pool fire clearly accounted for well over half 

(probably 3⁄4) of the total HRR. When the same Z0 foam was elevated far above the catch 

surface, the melt tended to scatter over a much wider area, and being thereby much more 

effectively heat-sunk to the Marinite, most of it failed to sustain burning. Visually, it 

appeared that the pool fire never yielded more than roughly 10 % of the HRR and nearly 

all of the mass that fell to the catch surface remained unburned. The results for the Z” 

foam (Table 11) were essentially the same. Clearly then, the pool fire can not only affect 

the sample burning, but the opposite is also true. 

The only other spacing comparison made in these tests was with the L viscoelastic foam. 

Recall that it had a very strong (dominant) pool fire in tests where the sample was close 

to the catch surface. When the catch surface was far away, the fire on the sample went 

out 2 s after that igniter flame ended. The pool fire reported in Table 11 was generated 

by the igniter and it persisted only long enough to burn itself out (since it lacked a 

continuing supply of melt after the igniter was turned off). Although it was not tested 

here with a large spacing above the catch surface, it is probable that the other viscoelastic 

foam (K) would behave the same way. 

The two viscoelastic foams seem to have a somewhat paradoxical ability to shed melt at a 

temperature below the ignition temperature of that melt. This behavior in itself is not 

paradoxical, any low molecular weight polymer can do this. The paradoxical part comes 

from the fact that the melt, as seen in our radiation-driven melt tests and as captured from 

Vee tests above, is not a liquid at room temperature but, in fact, seems very rubber-like. 

In the radiation-driven experiments, even at 40 kW/m2

, the “melt” seemed not to flow 

easily/smoothly off the sample surface. The viscosity of the “melt”, which could only be 

measured in oscillatory shear71 rather than rotational shear, was 10 to15 times lower (at 

350 o C) for the flaming sample than for the 40 kW/m2

 experiments but, in absolute 

terms, when compared to all the other melts in this study, it was high (ca. 50 Pa-s). Thus 

it seems quite surprising that it flows off the sample surface at a low enough temperature 

 

70 The rapid pool fire establishment in this case was assisted significantly by the proximity of the igniter 

flame and the flames on the sample engendered by the igniter. 

71 Recall the previous caveat about oscillatory testing for viscosity of these materials. It is, however, the 

only measure available at present. 
45

that it is not burning (except weakly, in areas of direct pool fire flame impingement, in 

the closely-spaced single and double slab experiments)72. 

The melt from Foam P manages the same trick of not igniting as it flows off the foam 

surface but it has a low melt viscosity (comparable to the NFR and Z foams) plus a high 

level of melamine to help keep that melt from igniting. As Table 11 shows, the 

combination led to essentially no post ignition fire after removal of the igniter, even for 

the closely-spaced configuration. We did not pursue the large-spaced configuration for 

this foam on the expectation that it would again yield no fire. 

Summary of Significant Observations. Although this study has not been definitive in 

many regards, it has produced a number of relevant points which we summarize briefly 

below before turning to possible implications for a polyurethane foam flammability test 

method. 

A melting material poses a new mode of fire spread – flaming melt can flow downward 

to new locations, extending the burning area of fuel in new directions and thereby 

increasing the heat release rate. In worst case circumstances, the pool of burning melt 

that forms on some catch surface can augment itself by increasing the melt supply to the 

pool in a self-accelerating manner. When such a self-generated pool fire participates in 

the growth of fire on a thermoplastic material, the result is quite complex. For this 

situation to occur, the object must be close enough to the pool to exchange a significant 

amount of heat with it. When this happens, both the fire on the object and the pool fire 

itself are strengthened and the fire growth process is accelerated. The potential for pool 

fire-object fire growth is actually significantly diminished by the difficulty that the pool 

fire, on a horizontal surface, has in moving onto the cool area of the catch surface in a 

direction that will spread the fire on the object. This difficulty is caused by the heat loss 

experienced by the spreading melt and the subsequent increase in melt viscosity which 

this causes. The net movement of the pool front is closely coupled to the movement, on 

the lower surface of the melting object, of the point or region from which the melt is 

flowing onto the pool. The pool fire plume, which tends to trail the flame front 

significantly, helps drives the melt generation from the object, even as the pool fire adds 

significantly to the overall heat release rate of the fire. The flow within the melt pool is 

affected by simple hydrostatic pressure differentials, by surface tension gradients near the 

flame foot and by blockages set up around the pool periphery by cooled, immobile melt. 

With a highly mobile melt such as is seen with polypropylene (and perhaps various 

polyurethanes), the reluctance of the melt to spread forward onto a cold surface can be 

easily overcome by just a few degrees of downward tilting of that surface in the spread 

direction (thereby introducing a hydrostatic gradient that can trump the other forces). 

 

72 Conceivably such behavior could reflect the fact that a melt resists ignition because it has a higher 

ignition temperature than other melts. Thermogravimetry does in fact show that the K foam had the highest 

temperature for onset of weight loss but in absolute terms that temperature is still low (ca. 230 o C to 

250 o C) so that would seem to contribute little here. 
46

Modeling the above fire growth processes requires several parameters and/or functions. 

The most basic of these is a description of the melt viscosity as a function of melt 

temperature. At the temperatures of interest, the melt viscosity has achieved a value that 

is a result not only of thermal loosening of inter-chain forces but also of a pyrolytic 

degradation of the length of those chains. For polypropylene and polystyrene we have 

been reasonably successful in obtaining an empirical description of melt viscosity which 

appears to be only a function of temperature. However, the viscosity in the temperature 

range of interest involves some substantial extrapolations which limits its accuracy. For 

polyurethane foams, the melt exhibits unexpected non-monotonic behavior as a function 

of temperature and we have yet to obtain a suitable description for use as a model input. 

Further work is indicated on whether the unexpected results will actually carry over to a 

pool fire situation. Exploration of the behavior of pure polyols, as a simple alternative, is 

also indicated. Another complication is the complex sequence of interactive chemical 

and physical steps that ultimately result in polyurethane melt flowing off of a burning 

foam surface; this will require empirical approximation. The possible roles of foaming 

and charring in polyurethane melt pool fires also need investigation. 

Fuel geometry and the direction of fire propagation have, as expected, been shown to 

have strong effects on the growth of fire on polyurethane foams. Because of their low 

density, even thick slabs of these materials tend to collapse in a fire test making sample 

holding (retention of the original sample geometry to the end of propagation over the 

foam) a special challenge. An inability to do so rendered some of the configurations 

examined highly non-reproducible, though some of these could probably be improved 

with further work. The configuration of greatest interest, however, is a Vee with lateral 

fire spread from the narrow end of the Vee outward. This provides a level of radiative 

interchange that should be comparable to that in upholstered chairs. The Vee can be kept 

close enough above a horizontal catch surface to stimulate substantial interactions 

between the pool fire and the foam fire making this a potentially severe test. 

Several polyurethane foam formulations tested in this Vee configuration exhibited a wide 

variety of fire growth rates and peak heat release rate values. The results were generally 

in accord with expectations based on the amount and type of flame retardant(s) in the 

foam formulation. For non-FR foams, however, there was a strong dichotomy in melt 

and pool fire behavior depending on whether the foam was of a conventional formulation 

or a viscoelastic formulation. The behavior of the former was little dependent on the pool 

fire that it left well behind its spreading flame front. The behavior of the latter was 

entirely dependent on its pool fire, being unable to even sustain flaming when separated 

from that fire73. These differences apparently follow from the unknown nature and 

proportion of the isocyanates and polyols that went into these foam formulations. In a 

model they would have to be captured by the differing kinetic constants for the two 

overall degradation steps that gasify a foam and by the differing melt viscositytemperature relationships for the foams. All of these are expected to dictate whether the 

 

73 The tendency of some materials to “shrink away” from a small applied flame has sometimes been touted 

as a safety feature and, indeed, it can make a material pass some flammability tests. The viscoelastic 

foams can be seen as illustrators of the two-sided nature of such materials; in some applications the melting 

may simply be contributing to a serious pool fire. 
47

melt is generated at a low enough temperature and with low enough viscosity that it can 

or cannot escape from a fire-exposed foam surface without actually igniting. 

The Z foams, with their gas phase flame retardants, revealed another aspect of fire growth 

that presents a modeling challenge – prevention of flame spread onto vertical foam 

surfaces where the slowed gas phase oxidation rates cannot support a flame. Proper 

description of this type of kinetic effect is a major challenge for existing gas phase flame 

models. 

Overall it appears that the path to a model of fire growth on full density thermoplastics is 

reasonably clear and much of the data for testing the model from Reaction Engineering 

International (REI) is in hand (though there is certainly room for more experimental data 

on both coupled pool fire-object fire spread and on pool fire behavior alone). The path to 

the more relevant problem of fire growth on polyurethane foams is less clear and more 

extensive work is clearly needed.
48

Implications for Flammability Testing of Polyurethane Foam. 

Proposed Foam Flammability Test Configuration. Recall that the goal of the 

flammability test method we wish to develop is to determine whether a foam is likely to 

perform well in the context of an upholstered chair or mattress in which it comprises the 

majority (but not all) of the flammable material. By good performance we mean that the 

foam does not substantially increase the peak heat release rate74 that results from the 

burning of the other materials in the chair or mattress assembly nor does it substantially 

shorten the time to that peak. In its end product application, the foam is assumed to be 

exposed to the flaming of a potentially wide variety of non flame-retarded materials 

(cover fabrics, insulating pads, wood structure, etc.) which will most likely keep burning, 

once ignited, threatening to spread flames over and onto the foam. Ideally, in spite of 

these “external” fire insults and radiative feedback among chair surfaces and the potential 

for a self-feeding pool fire, the foam would respond only weakly, yielding a minimal fire 

threat. To assure this, the foam flammability test method must subject the foam to 

circumstances that are comparable to those it sees in the burning of the end product. This 

is a tall order since those circumstances have never really been quantified (and could be 

expected to vary considerably). 

The upholstered chair end product application, in particular, poses an extremely wide 

range of potential fabric associations for a polyurethane foam. These in turn impose a 

wide variety of external fire insults on the foam, assuming that the fabric is ignited first. 

Here we want to think in terms of worst-case fabrics; we assume that these are based on 

fibers that are thermoplastic and have a high heat of combustion. These tend to open up 

to expose the foam while at the same time depositing vigorously flaming material on its 

surface. Such fabrics also drip flaming materials onto lower surfaces and/or the floor 

where they can serve as pilot flames for melt pool fires. The ideal test would reveal the 

likely behavior of a foam in these worst case conditions. We would then expect that the 

foam would be better behaved with less energetic and/or charring fabrics. (This is an 

assumption, based on some experience, but it needs checking. The real world of fabric 

blends can harbor an extraordinary range of behavior.) 

The test configuration we are currently focused on derives from the Vee configuration 

studied above, with some enhancements to capture fabric effects (see Fig. 37). The 

motivation derives mainly from the application to upholstered furniture. The two vertical 

slabs present opportunities for radiative interactions (quantified further below) similar to 

those on the interior seating surfaces of a chair or sofa. Strong, full-height ignition in the 

narrow end of the Vee deliberately provokes a large area of the sample into a possible 

flaming condition to simulate a large area ignition threat to a chair (rather than a small, 

 

74 This is, of course, somewhat vague due to the fact that the proximity of flaming material will force some 

unknown amount of heat release from even a very good polyurethane. Ultimately it is up to a regulatory 

agency to decide what constitutes acceptable performance for the product in which the foam is used. For 

example, the Consumer Product Safety Commission has set an upper limit of 200 kW from a mattress 

tested in accord with CFR 1633 (simulating exposure of the mattress to burning bedclothes). Ref. 17 gives 

a methodology for assessing the relation between peak HRR from a product and the hazard this represents. 
49

localized flame). The increasing distance between foam slabs decreases the radiative 

interaction as fire growth attempts to progress out of the narrow end of the Vee. This 

provides one basis for differentiating foam performance since foam formulation should 

affect its susceptibility to “external” radiation. The configuration (and the response of the 

foam to it) determines the level of that radiation, although it is supplemented (see below). 

The proximity of a catch surface deliberately accentuates the threat posed by flaming 

melt material by having a lesser spacing above the “floor” than is the case in a typical 

chair. Flaming melt flow has the potential to enhance fire growth in a chair configuration 

even before it reaches the floor so we, in effect, bring the floor closer to accentuate any 

such tendency and give it the opportunity to show its worst characteristic – development 

of a self-feeding pool fire. The choice of Marinite P as the catch surface puts an inert and 

reproducible material with thermal properties comparable to wood in the role of a “floor” 

surface. 

We add two further elements as a substitute for the additional impacts of a flaming fabric. 

It is not practical to use a fabric itself to fulfill this role even though it could come closer 

to reality. This would call for a standard fabric, something which has proven to be 

notoriously elusive. Furthermore, fabrics tend to burn in a very poorly reproducible 

manner. Instead we use additional gas burners to impose two effects of a flaming fabric. 

The first is piloting of a potential pool fire on the catch surface. This is achieved by 

adding small multi-jet gas burner tubes just above the catch surface, arranged diagonally 

under each foam slab and pointing their flames downward onto the catch surface. 

Second, we add another multi-jet gas burner tube as a (mainly radiative) heat source right 

in the plane of symmetry of the Vee, with its several flame jets pointing upward75. The 

objective is to operate this burner at conditions such that it provides a substantial 

additional heat flux to the inner surfaces of the foam slabs at the narrow end of the Vee 

(and then decaying as the fire spreads outward on the Vee). This is a substitute for the 

additional heat input that a burning fabric could impose.76

The foams that have been tested in this study would be expected to rank the same in this 

tougher test as they did here, although most would be expected to burn somewhat more 

vigorously and/or extensively. The two foams which did the best here (Foam P and 

Graphited) would be expected to do well in the enhanced test also. The additional heat 

sources are still needed, in general, to better simulate the real conditions in chair burning 

to which still other foam compositions might behave differently. Note that the proposed 

test particularly “punishes” flaming melt (even if it requires piloting to sustain that 

flaming) and a strong response to external radiation. These are two areas in which the 

two best foams in this study excel. 

 

75 The ultimate incarnation of this modified Vee test (to be published) places a solid body radiator in this 

role. 

76 A concern, which can only be addressed experimentally, is whether the added flames in this augmented 

Vee configuration are sufficient to capture the interaction between a foam with a gas phase flame retardant 

and the piloting effects of flaming thermoplastic fabrics on vertical foam slab surfaces. This configuration 

might give an overly optimistic prediction of the upholstered chair performance of such a foam if it 

underestimates the fire growth contribution from such surfaces. It should be possible to compensate for 

such a shortcoming, but it would probably make the test more complex. 
50

This proposed foam flammability test is plausible based on the available information 

from this study. However, it needs to be assessed against more realistic conditions. 

Plans are being made to test a set of foams in isolation in the configuration proposed 

above and also in the form of four cushion California Technical Bulletin 133 mock-ups to 

check the correlation between the two. As a practical matter, it would be highly desirable 

if the measure of fire intensity (peak heat release rate) could be replaced with reasonable 

accuracy by peak rate of mass loss since the required instrumentation is much reduced (at 

the cost of some accuracy in HRR assessment). This issue will also be examined. 

Scaling Issues in a Foam Flammability Test Method. The proposed Vee test 

configuration is close to full-scale and thus demands rather large slabs of polyurethane 

foam. A major planned use of the test method, if it is successful, is as a basis for 

assessing new foam formulations using non-traditional flame retardant systems (such as 

nano-composites). A natural question is whether that same information about the 

flammability hazard level of a foam could be obtained with much less material, since this 

would facilitate more rapid assessment of alternative formulations. This leads us to a 

closer look at the relative roles of various processes during fire growth in the Vee 

configuration. A smaller scale test method can be expected to produce similar ranking 

results of varied foam compositions only if it brings approximately the same balance of 

elements that contribute to fire growth. The relevant processes include (1) the rate of 

flame spread over foam surfaces, (2) the radiative interchange between the hottest surface 

elements of the two slabs or between the fire plumes attached to them and those surfaces, 

(3) the radiative interchange between any pool fire and the slab surfaces and their flames, 

(4) susceptibility to a self-feeding pool fire. 

To get an estimate of the roles of these processes we have focused on one of the Z foams 

used in the preceding study; this represents a moderately flame-retarded foam. We 

consider various processes for a foam with its characteristics both on the scale tested in 

the present study and on 1/3 to 1⁄4 of this scale (in terms of the length and height of the 

foam slabs, thickness is taken as unchanged)77. 

As noted previously, a significant effect of the gas phase flame retardant in the Z foams is 

its tendency to inhibit flaming on the inner vertical surfaces of the slabs. As a result, the 

rapid spread over these surfaces, seen in the unretarded NFR foam, is halted and flames 

spread stably only on surfaces that face upward (i.e., have a surface normal that has an 

appreciable upward component, thereby forming a flame-holding wake region for a 

buoyant fire plume).78 Such a wake region exists on the top of each slab. Another 

flame-holding region is formed on the “trailing edge” of each slab by the burner ignition 

process when the initially vertical slab surface at the narrow end of the Vee develops a 

finite slope (due to faster propagation on the top surface). It is then this burning “trailing 

 

77 Test samples of polyurethane foam on this smaller scale can be made in a laboratory bench-top apparatus 

generating a few liters of foam, though two batches would be needed to make two slabs. 78 This has an interesting consequence. The normal flame spread process of successive ignitions can 

proceed along the top of the sample but the vertical sides are engulfed only as fast as the inward (in-depth) 

mass burning process can progress on the upward facing surfaces. That mass burning rate process is 

augmented by any “external” radiation it sees. 
51

edge” surface on each slab that is the source of radiation impinging on the same trailing 

edge surface of the other foam slab. Another source of such radiation is the fire plume 

attached to each trailing edge. 

One can estimate the radiative view factors between these elements using a few plausible 

assumptions that reflect the behavior seen in the Vee tests reported above. As noted 

above, the burning, trailing edge surfaces of the foam slabs form a very obtuse angle with 

respect to each other and this greatly reduces their mutual radiative view factor; available 

data indicate an angle of the order of 120o. The fire plumes attached to them have a 

diameter roughly equal to the width of a foam slab and a height roughly equal to 1.5X the 

slab height. Using standard methods [19, 20] we can then estimate various view factors 

such as between the two solid burning surfaces or between an attached fire plume on one 

slab and the solid burning surface on the other slab. The results are shown in Table 12. 

Note that the values in Table 12 all correspond to the condition in which the two foam 

slabs (on either scale) are half consumed by the fire growth process. To give some idea 

of how these numbers can vary as the fire progresses from the narrow end of the Vee 

outward, note the following for the large scale version of the radiation between the 

trailing edge fire plume of slab 1 and the trailing edge surface of slab 2: 

 Fractional Length Consumed View Factor 

 1⁄4 0.175 

 1⁄2 0.13 

 3⁄4 0.074 

Table 12 shows that solid surface to solid surface radiation is likely to be minimal for 

these foams; the emissivity could be high but the temperature is roughly 400 oC and the 

view factor is quite small. Fire plume radiation, both between slabs and from the slabs to 

the pool fire is much more favored by the geometry, by the plume temperatures (expected 

to be roughly 1200 oC) and the moderate sootiness of the flames which can boost their 

emissivity. Furthermore, once a merged pool fire becomes fully established in the space 

just to the rear of the trailing edges of the two foam slabs, its view factor with respect to 

those trailing edges approaches unity at mid-height and 0.5 at the top and bottom of the 

trailing edged surfaces. For the Z foams in the present large scale study, this did not 

occur until about 2/3 of the slabs were consumed. For the viscoelastic foams it occurred 

much sooner. 

Table 12 seems to imply that there is good promise in the scaling down from the 

dimensions used in the above Vee tests to 1/3 of this size. There are differences in view 

factors but they are not extreme. There are further complications, however, as described 

below. 

Consider first the issue of susceptibility to a self-feeding pool fire. Appendix 2 addresses 

this issue in some detail and presents example calculations of the distance the bottom of 

the foam slabs would have to be above the surface on which the pool fire occurs in order 
52

to begin initiating a self-feeding pool fire. The results are given in terms of this 

separation distance versus the heat release rate from the object (here, the total from the 

two foam slabs). If we assume that the heat release rate from the object on the two scales 

of interest here is proportional to the foam slab areas on which flames are attached, then 

we have a factor of three difference for the Z foams. A change in peak HRR from 75 kW 

(seen for the Z” foam) to 25 kW in Fig. A2.1 implies that the separation distance should 

go from 0.22 m down to 0.15 m (no radiative boost curve) or from 0.42 m down to 0.27 

m (pool fire burning rate doubled by radiation). Again, these are separation distances that 

would just begin the self-feeding pool fire process at the noted values of peak HRR. All 

of these numbers are much larger than the 0.04 m separation used in the actual 

experiments where we deliberately minimized the separation distance in order to 

maximize the chances of an interactive pool fire. To maintain the severity of the test at 

its current level (which we regard as necessary in order to focus the test on one of the 

most troublesome aspects of polyurethane foam fire behavior), we evidently need to keep 

the 0.04 m separation distance for the larger scale and decrease it for the smaller scale79

in accord with the model in Appendix 2, though reduced to much smaller fires 

(something more like 1 kW for the larger scale test). A problem here is that this pushes 

the scale right into the domain of laminar plume pool fires where it may be difficult to 

find plume height data for small fires (the calculations in Appendix 2 are for turbulent 

fire plumes). In principle, at least, the separation distance issue is addressable and would 

lead to an equivalent separation distance, less than 0.04 m, which could be used for the 

smaller scale test. 

All of the preceding arguments about scaling have essentially been static, comparing 

view factors, etc. for two fixed geometries. The actual test behavior is quite dynamic 

with no pool material at all in the beginning, followed by a period of pool material 

accumulation then ignition of that pool and growth to the point where it begins to interact 

with the base of the foam slabs before going on to dominate later. When one views a 

video of a Vee test on one of the Z foams while focusing on the one third of the foam 

material that would be included in a one-third scale test, it is apparent that there is a 

problem: the flames spread to the end of that section of the sample before the melt pool 

fire becomes significant. While the surface flame spread is forced in this case by the 

nearby igniter flame, so also is the accumulation of a melt pool from the taller sample. 

The implication is that the role of a pool fire in a potentially 1/3 scale test is suppressed 

by a mismatch between the flame spread time over the shorter sample and the 

development time for a significant pool fire. 

The pool fire development time is strongly affected by the thermal properties of the catch 

surface to which it is losing heat. Right from the start, the small scale and large scale test 

configurations flow melt onto the catch surface at differing rates because of the differing 

sample heights impinged by the igniter flames. We can estimate the differing results for 

the pool fire initiation process as follows. Assume that the melt pool can ignite only 

when it reaches a sufficient temperature and doing so in turn requires getting the surface 

 

79 We could not go much below 0.04 m (4 cm) in the tests reported here due to the need to keep the sample 

separate from the catch surface since the two were weighed separately. We would drop this requirement, 

thus allowing the separation in a smaller scale test to be less. 
53

temperature of the catch surface beneath it up to some value (which does not enter into 

the calculation). Then to a first approximation the melt ignition process is controlled by 

the transient heating of the catch surface by what we will take to be a constant 

temperature melt flowing from the sample. The only thing that changes with the test 

scale is the flow rate of that melt onto the catch surface and that flow rate in turn affects 

the heat transfer coefficient between the melt and the surface. At the stagnation point 

where the flow impinges, the heating of the catch surface is approximately onedimensional. The temperature-time solution for this purely convective heating situation 

can be found in Ref. 21. The temperature of the top surface of the catch pan varies as 

follows: 

 ( )[1 exp( ) ( )] 2 T T T T t erfc t surf = amb + melt − amb − h α h α

where h is the heat transfer coefficient in the neighborhood of the stagnation point and α

is the thermal diffusivity of the catch surface material; t is time from the start of 

impingement of a constant flow of constant temperature melt. The heat transfer 

coefficient in the neighborhood of a stagnation point is proportional to the square root of 

the Reynolds number [22]. For convenience we consider the situation in which the melt 

flow changes by a factor of four for the two test scales under consideration; the heat 

transfer coefficient, h , thus varies by a factor of two. Applying the above relation to the 

two scales, we can find a relation between the time to reach the critical catch surface 

temperature on the small scale, given the time to reach it on the large scale. 

 exp( ) ( ) exp(4 ) ( 2 ) 1 1

2

2 2

2 h α t erfc h αt = h α t erfc h αt

where h is the value for the smaller scale (estimated in accord with Ref. 22), as is the 

time, 2t . For the value of 1

t we insert the 50 seconds seen in the full scale tests with a Z” 

foam. The value of thermal diffusivity, α , is given by the manufacturer of the Marinite 

P used here. Once 1

t is chosen the right hand side takes on a fixed numerical value and 

the resultant equation can be solved for the value of 2t , the requisite time to get to the 

same temperature given the lower melt flow in the smaller scale test. One finds that the 

result is 4 1t = 200 s, which is obvious from the nature of the above expression. In 

general the result is whatever the mass flow ratio is so, for our original factor of three 

scaling this implies that it would take 150 s to get pool fire ignition (and longer to get to 

the self-feeding pool fire condition). 

The flame spread process for the Z foams had consumed the corresponding 1/3 of the 

sample length in about 30 s. Note that as it did so, the melt flow location moved along 

the sample. This implies an even less effective catch surface heating process than 

assumed in the above calculation, which implicitly assumes that it stays in the same 

location. There is thus a large mismatch in the pool fire development process between 

the two scales, relative to flame spread time over the sample. The only compensating 

factor that could bring this more in line on the two scales would be to use a catch surface 

with a lower thermal diffusivity than the Marinite P. The above calculation implies a 
54

need for a factor of three reduction but, as we have just noted, this is not sufficient. A 

factor of ten or more reduction may be necessary to allow the pool fire to develop as 

rapidly as it does in the larger scale test. It is not evident at this point whether such a 

material exists. Most non-metallic, thermally inert materials have thermal diffusivities on 

the order of 10-3 cm2

/s and we are asking for an inert, preferably non-porous material 

with a thermal diffusivity on the order of 10-4 cm2

/s. An alternative might be to substitute 

a thermally-thin material as the catch surface. This may be worth further investigation.80 

It is worth noting that useful information regarding new flame-retarded foam 

formulations can be obtained from other techniques, such as the radiation-driven 

melt/gasification experiments described above and this does require a much more limited 

amount of foam per test. Thus it is likely that the issue of how to reduce the melt fraction 

and/or increase the viscosity of the melt will be pursued in seeking less flammable foam 

formulations. Radiation-driven tests at 40 kW/m2

 can provide the melt fraction on 

samples of about one liter in volume. One can then capture the melt and measure its 

viscosity. 

Summary of Issues Related to Foam Flammability Testing. We have proposed a 

modified version of the Vee test, on the scale used in this study, as the basis for assessing 

whether a polyurethane foam is likely to perform well in an upholstered chair subjected 

to a large flaming ignition source. The modifications (piloting of any pool fire, an added 

heat flux to the foam surfaces) attempt to mimic important aspects of the complexities 

associated with flaming thermoplastic fabrics. The viability of the test has yet to be 

assessed; a series of tests comparing foam behavior in the test versus that of fabricwrapped foam in a CTB 133 mock-up is planned. If the test is a reasonably good 

predictor of foam fire behavior in the context of chair mock-ups, it will become a useful 

procedure in the development of polyurethane foams of lesser flammability. 

A valid test of this nature would be of even greater use in the search for improved foam 

formulations if it could be significantly reduced in scale so as to require a lesser amount 

of each new foam. While reducing the geometry by a factor of three does not introduce 

any large disparities in the important radiative interchange aspects of the fire growth in 

the Vee configuration, it does introduce a large mis-match between flame spread time and 

pool fire initiation time. Use of a thermally-thin catch surface in a smaller scale test 

appears may be a potentially viable means of countering this mis-match but this is 

requires further study. It is worth noting that the fire growth model which this study 

seeks to foster can, when validated, be used to help assess the viability of smaller scale 

test methods.

 

80 There is another possibility and that is to heat the catch surface, as first suggested by W. Grosshandler. 

This idea has been exploited in subsequent developments on both the bench scale and intermediate scales. 
55

References. 

1) Kashiwagi, T., Omori, A. and Brown, J., “Effects of Material Characteristics on 

Flame Spreading,” Fire Safety Science – Proceedings of the Second International 

 Symposium, International Association of Fire Safety Sciences, Hemisphere 

 Publishing, New York (1989) p.107 

2) Zhang, J., Shields, T. and Silcock, G., “Effect of Melting Behavior on Upward 

Flame Spread Behavior of Thermoplastics,” Fire and Materials, 21, (1997) pp 1-

6 

3) Ohlemiller, T., Shields, J., Butler, K., Collins, B. and Seck, M. “Exploring the 

Role of Polymer Melt Viscosity in Melt Flow and Flammability Behavior,” 

Proceedings of the Fall Conference of the Fire Retardant Chemicals Association, 

Ponte Vehdra, Florida, October, 2000 

4) Sherratt, J. and Drysdale, D., “The Effect of Melt-Flow Process on Fire Behavior 

of Thermoplastics,” Proceedings of Interflam 2001, Interscience 

Communications, London (2001) 

5) Butler, K., Ohlemiller, T. and Linteris, G., “A Progress Report on Numerical 

Modeling of Experimental Polymer Melt Flow Behavior,” Proceedings of 

Interflam 2004, Vol. 2 (2004) p. 937 

6) Kissinger, H., “Reaction Kinetics in Differential Thermal Analysis,” Analytical 

Chem. 29, (1957) p. 1702 

7) Rogers, F. and Ohlemiller, T., “Pyrolysis Kinetics of a Polyurethane Foam by 

Thermogravimetry; A General Kinetic Method,” J Macrmol. Sci – Chem.A15

(1981), p. 169 

8) Rein, G., Lautenberger, C., Fernandez-Pello, C., Torero, J. and Urban, D., 

“Application of Genetic Algorithms and Thermogravimetry to Determine the 

Kinetics of Polyurethane Foam in Smoldering Combustion”, Comb. Flame 146

(2006) p. 95 

9) Carslaw, H. and Jaeger, J, Heat Conduction in Solids, Oxford University Press, 

Oxford, UK (1959) p. 80 
56

10) Ohlemiller, T., “Radiative Ignition of Polymeric Fuels in an Oxidizing Gas”, 

Ph.D. Thesis, Princeton University, Dept. of Aerospace and Mechanical Sciences 

(1969) 

11) Ohlemiller, T. and Cleary, T., “Upward Flame Spread on Composite Materials,” 

Fire and Polymers II, Materials and Tests for Hazard Prevention, (G. Nelson, 

ed.) American Chemical Society Symposium Series 599, Washington, DC (1995) 

p. 422 

12) Fried, J., Polymer Science and Technology, Prentice-Hall, Upper Saddle River, 

NJ (1995) p. 18 

13) Herrington, R., Broos, R. and Knaub, P., “Flexible Polyurethane Foams”, Chap. 4 

in Polymeric Foams and Foam Technology (D. Klempner and V. Sendijarevic, 

eds.) Second Edition, Hanser Publishers, Munich (2004) pp. 60-61 

14) Levchik, S. and Weil, E., “Thermal decomposition, combustion and fireretardancy of polyurethanes – a review of the recent literature,” Polymer 

International 53, (2004), p. 1585 

15) Ravey, M. and Pierce, E., “Flexible Polyurethane Foam. 1. Thermal 

Decomposition of a Polyether-Based, Water-Blown Commercial Type of Flexible 

Polyurethane Foam,” J. Appl. Poly. Sci. 63, (1997) p. 47 

16) Babrauskas, V., “Heat Release Rates,” Section 3 Chapter 1 in the SFPE 

Handbook of Fire Protection Engineering, 3rd Edition, National Fire Protection 

Association, Quincy, Mass (2002) pp.3-25 ff 

17) Ohlemiller, T. and Gann. R., “Estimating Reduced Fire Risk From an Improved 

Mattress Flammability Standard,” National Institute of Standards and 

Technology Technical Note 1446, August, 2002 

18) Fried, J., op cit, p. 206 

19) Anon., NFPA 555 Guide on Methods for Evaluating Potential Room Flashover, 

1996 Edition, National Fire Protection Association, Quincy, Massachusetts 
57

20) Siegel, R. and Howell, J., Thermal Radiation Heat Transfer, Hemisphere 

Publishing Corp, New York (1981), Appendix B 

21) Carslaw, H. and Jaeger, J., ibid, pp. 72 ff 

22) Kays, W., Convective Heat and Mass Transfer, McGraw-Hill, New York (1966), 

p. 210 

23) Sundstrom, B. (ed.), CBUF, Fire Safety of Upholstered Furniture – the final 

report on the CBUF research programme, European Commission Measurements 

and Testing Report EUR 16477 EN, Interscience Communication, London 

(undated, ca. 1995) 

24) Ohlemiller, T. and Shields, J. R., “Behavior of Mock-Ups in the California 

Technical Bulletin 133 Test Protocol: Fabric and Barrier Effects,” National 

Institute of Standards and Technology NISTIR 5653, May 1995 

25) Butler, K., Ohlemiller, T. and Linteris, G., “A Progress Report on Numerical 

Modeling of Experimental Polymer Melt Flow Behavior,” Proceeding of the 

Interflam 2004, Interscience Communications, London (2004) pp. 937-948 

26) Butler, K. and Ohlemiller, T., “Some Difficult Problems in the Modeling of Fire 

Spread,” paper presented to the 7th World Congress on Computational 

Mathematics, Los Angeles, CA (July 2006) 

27) Butler, K., Onate, E. and Rossi, R., “Modeling Polymer Melt Flow Using the 

Particle Finite Element Method,” paper presented at Interflam 2007, Edinburgh 

Scotland ( July 2007) 

28) FAR 25.853 (a) Appendix F Part I (a) (1) (ii) Vertical Burn Test. Flame 

application of 12 s results in less than 15 s of burning and less than 20 cm of 

flame propagation.
58

Acknowledgements

Most of the polyurethane foams were provided by the Consumer Product Safety 

Commission. Richard Harris and Michelle Donnelly performed the polymer viscosity 

measurements. Michael Smith and Richard Harris performed the thermogravimetric 

analyses. Michael Smith performed the standard cone calorimeter measurements. 

Bani Cipriano of the University of Maryland provided the gel permeation 

chromatograms on polymer melts. 
59

Table 1. Properties of Full Density Thermoplastics 

Resin Type Designation 

Melt Flow 

Index81

Density 

(g/cm3) Appearance82

Polypropylene 6523 4.0 0.90 White 

Polypropylene PD702N 35.0 0.90 White 

 

Polystyrene PS 663 1.5 1.04 Clear 

Polystyrene PS 666D 8.0 1.04 Clear 

Table 2. Properties of Flexible Polyurethane Foams83

Designation 

Density (kg/m3) 

(lb/ft3) Comment 

NFR 23.2 (1.45) Non-flame retarded; 30-36 lb IFD84; standard 

mattress quality foam 

Z0 FR Foam 22.4 (1.4) 6 % of bromine/phosphorus flame retardant 

Z’’ FR Foam 22.4 (1.4) 7 % of bromine phosphorus compound 

 + ca. 3 % melamine 

K Visco Foam 43.2 (2.7) Viscoelastic foam; no flame retardants 

L Visco Foam 36.8 (2.3) Viscoelastic foam; no flame retardants 

P FR Foam 48.1 (3.0) 28 % melamine + ca. 3 % tris dichloro 

isopropyl phosphate (TDCP) 

Graphited Foam 56.0 (3.5) Contains expandable graphite; gives positive 

qualitative test for presence of a halogen 

compound; aircraft rated 

 

81 Melt Flow Index is measured in accord with ASTM D 1238 and is reported by the manufacturer. It gives 

the flow rate of the melt through a standard orifice at a standard temperature and pressure condition. The 

polypropylene and polystyrenes are not reported at the same conditions. 82 The white appearance of the polypropylenes reflects the presence of a crystalline phase dispersed in an 

amorphous phase; the crystals melt as the polymer is heated and it then becomes clear. The clear 

appearance of the polystyrenes indicates a lack of crystallinity. 

83 Information on flame retardant content provided by the Consumer Product Safety Commission. 

84 IFD is Indentation Force Deflection, measured in accord with ASTM D3574 and reported by the 

manufacturer. It is a measure of the hardness or stiffness of the foam under standardized conditions. 
60

Table 3. Summary of Kinetic Parameters for Gasification of Full Density 

Thermoplastic Resins as Inferred from Thermogravimetry 

Polymer Resin 

Reaction 

Order 

Pre-Exponential 

Factor, Atga (s-1)

Activation Energy 

Etga (cal/mol) 

PD702N Polypropylene 1.0 2.18 · 1012 48, 400 

“ 0.85 3.68 · 1012 49, 500 

 

6523 Polypropylene 1.0 2.24 · 1013 51, 570 

PS 663 Polystyrene 1.0 7.60 · 1012 46, 610 

PS 666 Polystyrene 1.0 1.24 · 1012 44, 260 

Table 4. Polyurethane Foam Degradation Behavior85

Foam 

Designation 

Temp. (oC) at 

2 % Wgt Loss 

Temp. (oC) of 

1st Wgt Loss 

Peak 

Temp. (oC) of 

2nd Wgt Loss 

Peak 

Temp. (oC) 

Range of 

Foam Collapse

 

NFR 227 281 360 295 - 312 

 K 252 277 (301)86 377 280 - 308 

L 254 292 384 280 - 330 

Z0 198 281 362 287 – 306 

Z” 191 283 354 290 - 309 

P 204 260 (307)60 359 270 – 284 

Graphited 209 250 374 

272-285 

(Graphite 

Expansion) 

 

85 The first three columns of results are from thermogravimetry in nitrogen at 5 oC/min. The last column of 

results is for the same heating conditions but used the NIST Paar-Physica rheometer oven as a hot stage. 86 Parenthetic temperature refers to a secondary peak 
61

Table 5. Summary of Results from Radiative Melt/Gasification Experiments With F

Thermoplastics Using Cone Heater 

Resin Test # 

Heat 

Flux 

(kW/m2) 

Avg. 

Surface 

Temp 

(oC) 

Avg Total 

Mass Loss 

Rate (g/s) 

Avg. Melt 

Flow 

Rate (g/s) 

% of 

Mass 

Lost as 

Melt 

C

PP702, Low Visc. PP702-10 20 338 ± 2 0.116 0.094 81.0 Sheet-lik

“ PP702-11 20 332 ± 5 0.117 0.091 85.5 Li“ PP702-1 30 N. A. (0.14) 0.11 (79) Unfused p“ PP702-2 30 375 ± 4 0.15 0.11 74.7 Bubbl“ PP702-3 30 373 ± 3 0.15 0.12 80.0 “ PP702-4 30 358 ± 3 0.173 (0.12) (69) Nitroge“ PP702-5 30 384 ± 3 0.179 (0.12) (67) Nitrogen “ PP702-6 30 375 ± 3 0.182 (0.11) (60) Air Pu“ PP702-7 46 448 ± 2 0.244 0.131 53.7 No visible “ PP702-8 46 450 ± 1 0.24 0.128 53.3 OK until va“ PP702-9 46 460 ± 3 0.238 0.120 50.4 Very m

PP6523, Hi Visc. PP6523-1 30 400 ± 4 0.15 0.073 48.7 Bubbl

“ PP6523-2 30 395 ± 3 0.156 .068 43.6 

PP6523-FR 

(Br/Sb) 

PP6523FR1 30 366 ± 2 0.180 0.139 77 Very fluid m
62

Resin Test # 

Heat 

Flux 

(kW/m2) 

Avg. 

Surface 

Temp 

(oC) 

Avg Total 

Mass Loss 

Rate (g/s) 

Avg. Melt 

Flow 

Rate (g/s) 

% of 

Mass 

Lost as 

Melt 

C

PS 666, Low Visc PS666-3 20 339 ± 6 0.147 0.126 86 Some lar

“ PS666-4 20 340 ± 4 0.159 0.132 83 Very few b

layer 

“. PS666-1 30 389 ± 3 0.228 0.179 78.5 Minima

“ PS666-2 30 392 ± 3 0.230 0.183 80 “ PS666-5 46 427 ± 4 0.331 0.248 74.9 Few bubble“ PS666-6 46 423 ± 6 0.319 0.236 74.0 Low densi

PS 663, High 

Visc. 

PS663-5 20 332 ± 3 0.123 0.102 83 Larger su

“ PS001 25 355 ± 4 0.132 00.123 93 Shee“ PS663-2 25 350 ± 0 0.152 0.122 80 Machin

behav

“ PS663-3 30 403 ± 3 0.193 0.130 67 Somewha

“ PS663-4 30 397 ± 1 0.188 0.144 77 
63

Table 6. Summary of Full Density Thermoplastic Tests with Flaming 

Sample and Cone Heater

Resin Test # 

Heat 

Flux 

(kW/m2) 

Avg. 

Surface 

Temp 

(oC) 

Avg 

Total 

Wgt 

Loss 

Rate 

(g/s) 

Avg. Melt 

Flow Rate 

(g/s) 

% of Wgt 

Lost as 

Melt 

 

PP6523 MBC-1 20 N.A. 0.23 0.15 65 

“ MBC-2 10 (490 ± 

10) 

0.17 0.12 70 

“ 

“ -B1 10 N. A. 0.179 0.122 68.2 

“ -B2 10 466 ± 20 

457 ± 14 

46 6± 14 

0.159 0.113 71 

“ -B3 10 453 ± 13 0.154 0.105 68.2 

“ -B4 20 477 ± 10 

477 ± 10 

0.212 0.151 71.2 

“ -B5 20 479 ± 4 

495 ± 7 

0.207 0.148 71.5 

FRPP6523 

Mg(OH)2 as 

FR 

-B3 20 463 ± 10 

439 ± 10 

0.210 0.141 67.1 

FRPP6523 

Mg(OH)2 as 

FR 

-B4 20 440 ± 19 

449 ± 18 

0.209 0.145 69.4 

PP702 -B1 10 482 ± 19 

494 ± 14 

0.172 0.125 72.7 

“ -B2 10 452 ± 13 

450 ± 13 

0.175 0.129 73.7 
64

 Table 6. Continued 

Resin Test # 

Heat 

Flux 

(kW/m2) 

Avg. 

Surface 

Temp 

(oC) 

Avg 

Total 

Wgt 

Loss 

Rate 

(g/s) 

Avg. Melt 

Flow Rate 

(g/s) 

% of Wgt 

Lost as 

Melt 

 

PS 663 -B1 10 (375) 

Funnel Base 

0.196 0.115 58.7 

“ -B2 10 405 ± 4 

Melt in 

Funnel 

0.20 0.11 55.0 

 

PS 666 -B1 10 394 ± 3 

410 ± 1 

0.22 / 

0.29 

0.14 / 0.19 63.6 / 65.5 

“ -B2 10 357 ± 3 0.249 0.174 69.9 

“ -B3 20 442 ± 4 0.292 0.209 71.6 

“ -B4 20 457 ± 4 0.321 0.198 61.7 

 

FR PS 666 -B1 20 400 ± 5 0.250 0.142 56.8 

“ -B2 20 456 ± 3 0.261 0.138 52.9 
65

Table 7. Average Heat Release Rate (HRR) from Cone Melt Experiments 

 and from Normal Horizontal Cone Tests 

Sample Test # 

Heat Flux 

(kW/m2) 

HRR (kW/m2

) 

Cone Melt 

Expt. 

HRR (kW/m2) 

Normal Horiz. 

Cone 

 

6523 

Polypropylene MBC-1 10 150 -200 

“ MBC-2 “ 130 

“ 6523 B1 “ 155 

“ 6523 B2 “ 160 

“ 6523 B3 “ 165 

“ 6523 B4 20 175 

“ 6523 B5 “ 160 

400-525 

400-550 

 

6523 PP + 

10% Mg(OH)2

FR6523 B3 20 175 

“ FR6523 B4 “ 175 

PD702N 

Polypropylene 

PP702 B1 10 140 

“ PP702 B2 “ 135 

“ 20 375-520 

PS666D 

Polystyrene 

PS666 B2 10 220 

“ PS 666 B3 20 250-380 

“ PS666 B4 “ 300-330 

330-475 

 

PS 666 + 

10% Mg(OH)2

FR666 B1 20 310 

“ FR666 B2 “ 320-380 

PS663 

Polystyrene 

PS663 B1 10 191-210 

“ PS663 B2 “ 190-250 

“ 20 410-540 
66

Table 8. Summary of Surface Temperature Measurements 

On Polyurethane Foams 

In Cone Melt/Drip Experiments 

Foam 

Incident Heat 

Flux (kW/m2

) 

Surface Temperature (0

C) 

Results 

 

K Viscoelastic 30 412 ± 20 (Poor Wetting?) 

349 ± 7 

 40 392 ±12, 366 ± 15 

L Viscoelastic 30 312 ± 11, 295 ± 0 

 40 340 ± 14, 340 

P High Melamine 20 294 ± 14, 287 ± 24 

 30 353 ± 9 

 40 300 ± 11, 306 ± 6 

Z0 (P-Br-1) 20 346 ± 1 

 30 370 ± 9, 379 ± 9 

 40 372 ± 31, 388 ± 8 

Z’’ (P-Br-2) 30 359 ± 13, 374 ± 10 

 40 378 ± 8, 379 ± 12 

NFR 30 377 ± 21 

 40 410 ± 8, 412 ± 12 
67

Table 9. Summary of Cone Calorimeter Results for Polyurethane Foams 

at a Radiant Flux of 35 kW/m2

(Average of Three Replicates) 

Sample 

Designation 

Initial HRR 

“Plateau” 

(kW/m2) 

Peak HRR 

(kW/m2) 

Heat of Combustion 

Variation 

(kJ/g) 

 

NFR 290 ± 30 570 ± 75 ca. 25 to ca. 31 

L Viscoelastic 240 ± 10 445 ± 25 ca. 24 to ca. 30 

K Viscoelastic 260 ± 10 385 ± 50 ca. 24 to ca. 30 

 

Z0 210 ± 12 590 ± 6 ca. 19 to ca. 31 

Z” 225 ± 5 530 ± 115 ca. 19 to ca. 31 

 

P one peak only 400 (2 repl.) ca. 25 to ca. 30 

Graphited one peak only 175 ± 15 ca. 20 to 25, constant with 

time but very noisy 
68

Table 10. Summary of Peak HRR and Average Flame Spread Rates in Single Foam Slab T

Test # Sample 

Designation 

Sample Dimensions Test Configuration 

Peak 

HRR 

(kW) 

1 Z0 30 cm x 60 cm 

by 7.6 cm 

Lateral spread along long dimension after ignition 

of vertical 30 cm high end; close to Marinite88

 (with Al foil covering Marinite) 

72 

4 “ “ “ 54 2 “ “ Same as above but no Al foil 8 3 “ “ “ 14 5 “ “ “ 12 

23 Z” “ “ 25 24 “ “ “ 27 21 P “ “ 0 post-ig22 “ “ As above but wrapped with polypropylene fabric 8 

87 Averaged over 25 cm length well away from both ends of sample. 88 For various reasons, “close” varied from about 3.8 cm to 5.1 cm in different tests. There was no indication that this variation in heigh

a significant effect. 
69

Table 10, Cont’d. Summary of Peak HRR and Average Flame Spread Rates in Single Foam S

Test # Sample 

Designation 

Sample Dimensions Test Configuration 

Peak 

HRR 

(kW) 

8 NFR 30 cm x 60 cm 

by 10 cm 

Lateral spread along long dimension after ignition 

of vertical 30 cm high end; close to Marinite 

56 

9 “ “ “ 56 6 L 30 cm x 45 cm 

by 7.6 cm 

“ 48 

17 “ “ “ 50 19 “ “ “ 47 

7 K 30 cm x 45 cm 

by 7.6 cm 

“ 57 

18 “ “ “ 59 20 “ “ “ 50 29 Graphite 29 cm x 59 cm 

by 10 cm 

“ ca. 0 

post-ign
70

Table 10, Cont’d. Summary of Peak HRR and Average Flame Spread Rates in Single Foam S

Test # Sample 

Designation 

Sample Dimensions Test Configuration 

Peak 

HRR 

(kW) 

10 Z0 30 cm x 60 cm 

by 7.6 cm 

Sample laid flat on wires, close to Marinite; ignited 

at base of 30 cm wide by 7.6 cm high face 

50 

11 “ “ “ 53 14 “ “ “ 50 16 “ “ “ 57 25 P “ As above but sample wrapped in polypropylene 

fabric as in Test 22 

30 

12 NFR 30 cm x 60 cm 

by 10 cm 

Sample laid flat on wires, close to Marinite; ignited 

at base of 30 cm wide by 7.6 cm high face 

57, 54 

26 Z0 30 cm x 60 cm 

by 7.6 cm 

Sample vertical with bottom close to Marinite; 

ignited at base of 30 cm wide x 60 cm high face 

50 

27 “ “ “ 36 28 “ “ “ 21 
71

Table 10, Cont’d. Summary of Peak HRR and Average Flame Spread Rates in Single Foam S

Test # Sample 

Designation 

Sample Dimensions Test Configuration 

Peak 

HRR 

(kW) 

49 Z0 30 cm x 56 cm 

by 7.6 cm 

Lateral spread along long dimension after ignition 

of vertical 30 cm high end; close to Marinite; 30 cm 

tall by 60 cm long face irradiated with 7-8 kW/m2 

41 

51 Z” 31 cm x 46 cm 

by 7.6 cm 

“ 27 

50 Graphite 30 cm x 60 cm 

by 10 cm 

“ ca. 5 
72

Table 11. Peak HRR Prior to Loss of Configuration Geometry in 30 o Vee Tests 

Of Polyurethane Foams 

 

89 Especially in earlier tests with this configuration, the Vee geometry was lost before the flames had spread 

fully on the inner surfaces of the foam slabs. Cases where this appeared to have lessened the reported peak 

HRR are indicated here by an asterisk. 

Foam 

Designation 

Test 

Number 

Spacing Above 

Marinite (cm) 

Peak HRR89

(kW) 

 

Z0 32 4 100 

“ 33 4 110 * 

“ 37 4 120 

Z0 43 77 21 

“ 44 77 22 

Z” 35 4 80 * 

“ 36 4 74 * 

Z” 45 77 15 

“ 46 77 15 

P 31 4 0 

P + PP wrap + 2 

shoddy pads 

40 4 32 

NFR 41 4 125 

“ 42 4 150 * 

L 38 4 120 * 

“ 39 4 120 * 

L 47 77 Visual estimate 

30 kW, pool only 

K 34 4 125 * 

 
73

Table 12. Estimates of Radiative View Factors in Vee Test Configuration 

(Small and Large Scale) 

From Radiator to 

Receiver 

Small Scale 

20 cm long by 10 cm high 

Large Scale 

60 cm long by 30 cm high 

 

Trailing Edge Surface of 

Slab 1 to Trailing Edge 

Surface of Slab 290

0.04 0.03 

 

Trailing Edge Fire Plume of 

Slab 1 to Mid-Height on 

Trailing Edge Surface of 

Slab 289 

0.15 0.13 

 

Trailing Edge Fire Plumes 

(2 slabs) to Pool Fire 

Midway Between89

0.40 0.26 

 

90 Assumes that slabs are at a 300

 angle and the trailing edges are at a 1200 angle. Also assumes that, on 

each scale, both slabs are half consumed and the trailing edge separations correspond to this condition. 
74

 Figure 1. Schematic side-sectional view of cone calorimeter, 

 radiation-driven melt apparatus. 

SAMPLE

SCALE

SCALE

HRR

(WHEN 

BURNING)

MELT

HEATER

CONE

{ MELT 

NOT

BURNING

INSUL.

INSUL.

SS TUBE (SQUARE)

SLIDE

TC

MOTORIZED SLIDE
75

Figure 2. Overall view of the apparatus used for three-dimensional, melt pool fire-assisted fire 

 growth tests. A thin sheet of polypropylene is shown in its high position above the 

 Marinite catch surface (white sheet). 
76

Figure 3. Front view of a pair of foam slabs in the Vee test configuration. The wire supports under 

the foam are weakly visible near the open end of the Vee. Brace wires between the tops of the slabs 

are also visible. The vertical igniter at the narrow end of the Vee is slightly visible (mainly its lower 

half). 
77
78
79
80
81
82
83
84

Figure 9. Development of approximate melt viscosity-temperature 

 relationship for PP6523 polypropylene. 

 
85
86
87

Figure 12. Viscosity of polyurethane foam melt captured from two types 

 of experiment. Foam is Z0. 
88
89

Figure 14 a. Example of sample assembly weight during a polymer 

 melt/drip test. Sample is PS 666 polystyrene 

 subjected to 20 kW/m2

; non-flaming.
90

 Figure 14 b. Example of weight seen on catch pan during a 

 polymer melt/drip test. Sample is PS666 polystyrene 

 subjected to 20 kW/m2

; non-flaming. 
91

Figure 14 c. Example of surface thermocouple reading as it is moved into 

 and out of surface melt layer. Sample is PS 666 polystyrene 

 subjected to 20 kW/m2

; non-flaming. 
92
93
94
95
96
97
98
99
100
101
102

Figure 21. Pool-assisted fire growth on a vertical, 3 mm thick sheet of 

 polypropylene. 
103
104
105
106
107
108
109
110
111
112

Figure 28. Close-up of burning zone on side of polyurethane foam slab 

showing melt bead coalescence and flow; Z’’ foam. Flame zone is moving 

from left to right in photo. See Fig. 29 for overall view. 
113

Figure 29. Side view of single vertical slap of flame-retarded Z’’ foam 

 burning from left to right. Note preferred spread on top surface 

 where flow wake can create longer residence time for reactions. 
114

Figure 30. Side view of single vertical slab of L foam, burning from left to 

 right, showing dominance of pool fire for this viscoelastic foam. 
115

Figure 31. Vee test of NFR polyurethane foam showing minimal pool fire 

 contribution to first half of test. Stabilization wires added to 

 outer end of Vee are evident in picture. 
116

Figure 32. Vee test of L viscoelastic polyurethane foam, showing strong 

 contribution from melt pool fire. Compare Fig. 31. Note 

 differing shapes of flame fronts on interior surfaces of foam 

 slabs. 
117
118
119
120
121

 Figure 37. Potential foam flammability test configuration. 

FOAM

 SLAB

o

o

o

o

o

o

o

o

o

o

o

o

o

FOAM

 SLAB

RADIATION 

 SOURCE POOL FIRE

 PILOTS

(JUST ABOVE

 MARINITE P

 CATCH SURFACE)

POOL 

 FIRE

 PILOTS

PROPANE 

 IGNITER
122

Appendix 1 

A Simplified Steady-State Model of a Thermoplastic 

Undergoing Gasification and Melt Flow 

It is of interest to see to what extent the behavior of a thermoplastic material, subjected to 

a constant incident heat flux, can be predicted by a steady-state model based on 

simplified heat and mass balances. The specific problem is a vertical slab of 

thermoplastic heated uniformly on one vertical face with the opposite face being 

sufficiently far away that the material behaves as if it is thermally-thick. This is the 

intermediate problem addressed by REI in the sequence leading to a full 3-D fire growth 

model. Here, the heat flux is assumed to come from radiation (though it could, as well, 

come from a flame attached to the heated face). We are modeling experiments performed 

on both polypropylene and polystyrene in the Cone melt/drip apparatus, as discussed in 

the main body of this report. There a steady-state is assured by using a thick sample that 

is fed forward as it melts and gasifies so that its heated face is essentially stationary and 

the flux it sees is thus constant. REI is modeling the full transient problem from initially 

cold solid to a steadily regressing surface as a result of gasification (following thermal 

degradation of the initially high molecular weight polymer) and melt flow downward off 

of the heated face. Here, for simplicity, we focus only on the ultimate steady-state 

regression process that the sample achieves when it is fed forward at the same rate as its 

surface regresses (due to melt flow off the vertically-oriented face and simultaneous 

gasification from that face). 

A virtual steady-state is achieved experimentally in all of the cone calorimeter-based 

melt/gasification experiments that have been done, thus this aspect of the model posed 

here is not a major simplification of reality. However, there are definite two-dimensional 

aspects to the sample behavior which this model ignores or simplifies. For example, the 

sample face tends to recede somewhat faster at the top than elsewhere and the melt flow 

layer varies somewhat in temperature, velocity and thickness with height on the sample 

face. The goal of the simplified model posed here is to see if the major behavioral trends 

observed in the experiments can be captured. If so the model may produce useful 

insights as to controlling factors. 

Model Description. Figure A1.1 is a sketch of the system being considered. The 

radiant flux on the heated surface is spatially uniform and constant in time. The thermal 

wave does not reach the back of the sample in the time scale of interest. As a result of 

the heating on one face, the polymer there increases in temperature, causing its melt 

viscosity to decrease. The onset of thermal degradation lowers the mean molecular 

weight and this lowers the viscosity still further. Gravity causes the lowest viscosity 

outer layers to flow downward. Eventually the temperature is high enough to degrade the 

polymer to the point where small molecules can escape into the gas phase. Thus both 

melt flow and gasification remove mass and heat from the system. Surface re-radiation 

and convective heat losses also remove a part of the incoming heat. 
123

Figure A1.2 is a sketch of this system as it will be treated here. The system in Figure 

A1.1 will achieve a temporary steady-state when the incoming radiant heat flow is 

matched by the sum of the several heat loss paths shown in Fig. A2.2. This steady-state 

behavior will persist until the thermal wave in the solid begins to reach its rear surface. 

The melt flow rate and the gasification rate will increase after the onset of the radiant flux 

until they rise to their steady-state values. In Fig. A2.2 the solid thermoplastic is fed into 

the system from the left at a mass flow rate that matches the sum of the steady melt flow 

and gasification rates. It is this steady-state behavior which will be approximated with 

the following assumptions. 

1) The polymer melt viscosity is only a function of temperature and that function 

can be described by the empirical results described in the main body of in this 

report. 

2) There is a finite melt layer whose temperature decreases with depth below the 

heated surface. The thickness of this layer can be prescribed by the assertion that 

the back of the layer corresponds to the melt viscosity rising to a fixed multiple 

of that viscosity that exists at the surface temperature. (This viscosity rise is 

caused by the temperature decrease across the layer depth.) 

3) Gasification occurs from the outer surface of this falling melt layer (with no 

bubble formation). 

4) The mean velocity of this falling melt layer is given by a balance between gravity 

and viscous drag. 

5) The behavior of a large heated surface can be captured by considering a typical 

unit area on that surface (which we will take to be at mid-height). 

6) The melt flow layer is thin compared to the overall thermal wave. That overall 

wave is therefore taken to follow the one-dimensional steady-state temperature 

distribution. 

The meaning of some of these assumptions is made clearer when the model equations are 

stated. 

Overall energy balance at the sample surface (Eqn. A1.1): 

( ) ( ) ( ) ( exp( / ))( ( )) 0 0 0

4

0

4

αqin − εσ Ts −T − hconv Ts −T = mmeltCp Tavg −T + As −Es RTs Qg +Cp Ts −T

This is taken on unit area of the sample surface in a “representative” location. Here qin is 

the radiant flux (cal/cm2

s) incident on the sample face. This flux is assumed to be 100% 

absorbed (α = 1.0) and that absorption is taken to occur at the sample surface. In the 

absence of melt flow the presence of a steady-state would make it irrelevant whether the 

incident flux was absorbed exactly at the surface or below it. In actuality the flux is 

absorbed over a depth of the order of 1 mm or more (at least for polypropylene, as 

measurements have shown; see Appendix 4) and this may interact somewhat with the 

temperature profile in the melt layer; here any such interaction is ignored91. The second 

 

91 This in-depth absorption can be included in this model; its only effect is to make Eqn (A1.3) somewhat 

more complex since, for polypropylene, at least, some of the incoming radiation passes through the thin 
124

term in the above equation describes the radiative loss from the heated surface which is at 

the temperature Ts (Kelvin), interacting with the room at temperature T0 . The surface 

emissivity, ε , is also taken to be unity. The third term describes the natural convection 

heat loss from the sample face; the convective heat transfer coefficient, hconv , is obtained 

from a standard handbook heat transfer correlation (the value used here is 2.5 x 10-4

cal/cm2

 s K). The sum of the three terms on the left hand side of Eqn (A1.1) is the net 

heat flux into the sample surface. That net heat goes into the melt flowing from the 

sample face and into gasifying a portion of the polymer melt. The melt heat content is 

the product of its mass flow rate, mmelt (g/s), its heat capacity, Cp (cal/g K) and its 

average temperature, Tavg (above room temperature). Here the average temperature of 

the melt, over its depth, is taken to be the average of the surface temperature and the 

temperature at the back of the melt layer, defined below. The last term on the right 

describes the heat absorbed by the gasification process; Qg is the heat required to gasify 

unit mass of the melt at the surface temperature (here taken as 300 cal/g for all cases 

examined). This gasification process is also, in reality, occurring over some finite 

(though very small) depth in the polymer melt below the surface. Because the activation 

energy of that reaction (as measured with thermogravimetry, as described in the main 

body of this report) is large, we can apply the following approximation, which derives 

from laminar flame theory [Ref. A1.1]: 

 

0.5

0

2

0

2

[( / )( / )(( exp( / ))/( ))]

/ ( / ) exp( / ( ))

C RT E A E RT T T

Net Gasification Rate Unit Area g cm s A E RT x dx

s p s tga tga tga s s

tga tga

≅ − −

= − ∫

−∞

λ

 

Here λs is the thermal conductivity of the polymer melt, R is the universal gas constant, 

Atga and Etga are the pre-exponential factor and activation energy, respectively, for 

polymer gasification, derived from thermogravimetric measurements. Note that the 

reaction is taken to be zero-order in polymer concentration because bubbling, typically a 

minor aspect of the gasification processes examined in this study, is ignored. In light of 

the above approximation, the gasification kinetic factors in Eqn (A1.1) can be expressed 

in terms of their thermogravimetric counterparts as follows: 

 

0.5

0

2 [( / )( / )( /( ))]

/ 2

A C RT E A T T

E E

s s s s tga s tga s

s tga

= −

=

λ ρ

 melt layer. It will shift all of the solutions somewhat but should not introduce any new qualitative trends. 

Since we are not trying to make a quantitative model here, the in-depth absorption is ignored. 
125

Here ρ s is the polymer melt density, taken here as equal to the original polymer resin 

density. Note that while the activation energy for the surface reaction is simply one half 

that for the volumetric reaction, the pre-exponential factor is modified considerably (in 

effect, accounting for the very small depth over which the reaction is actually occurring). 

The second model equation derives from the assumption that the melt layer has a finite 

thickness specified by making the viscosity at the back of the melt layer a fixed multiple 

of that at the front of the melt layer. 

 ) μ(Tδ ) = φ μ(Ts ) (A1.2

Here ) ( μ Tδ is the viscosity at the back of the layer where the temperature is Tδ , φ is a 

constant (and a parameter of the model), ) ( μ Ts is the melt viscosity at the surface of the 

melt layer. The value of φ has been taken to be 25 for most of the model cases 

examined. In actuality the viscosity of the melt is a continuous function of temperature 

and thus there could be finite flow at virtually any depth in the polymer. This equation 

recognizes the fact that increasing viscosity in depth implies a decreasing flow velocity 

and the flow at some depth has a negligible impact on the problem over the time scale of 

interest (essentially the time for the surface to regress through one thermal layer 

thickness). Equation (A1.2) is an ad hoc expression of this fact. It is informed by the 

fact that an isothermal, gravity-driven boundary layer has an average velocity that varies 

inversely with the first power of the fluid viscosity [Ref. A1.2]. 

The next model equation can be derived from a steady-state heat balance on the flowing 

melt film alone but it arises also from the assumption that the overall thermal wave in the 

polymer is largely unchanged by a thin melt layer flowing down the sample face. 

 ) (Tδ − T0 )/(Ts − T0 ) = exp(−(mmeltCp / λs ) δ a ) (A1.3

Here δ a is the average melt layer thickness on the sample face (taken as the value at the 

mid-height of the face). One can easily show that the one-dimensional temperature 

profile in a gasifying material follows this exponential decay with a characteristic depth 

that is the inverse of ( / ) mmeltCp λs . Here, as noted above, we assume that this profile 

holds also for this problem where there is a thin melt layer flowing down the sample face. 

Then Eqn (A1.3) says that the depth of the melt layer is related to the temperature at the 

back of that layer by the above relation. 

The next model equation simply expresses an overall mass balance on the flowing melt 

film at the mid-height of the sample face: 

 ) ρ s Vm δ a = (h / 2)(mmelt − mgasif ) (A1.4
126

Here Vm is the average downward velocity of the flowing melt layer, h is the height of 

the sample face and mgasif is the flux of mass from unit area of the face due to the 

gasification reaction. This equation brings in information on the actual sample size and 

accounts approximately for the fact that the melt layer varies in thickness with height on 

the sample face. 

The final equation in the model relates the average melt flow velocity to the melt 

viscosity. Here we take advantage of an available solution that applies to a fluid layer 

with a depth-dependent viscosity analogous to that here (but not precisely the same). Ref. 

A1.3 gives an expression for average fluid velocity in a falling film for a fluid whose 

viscosity is an exponentially-decaying function of temperature. Most polymer melts can 

be described by such a temperature function over at least a limited part of the temperature 

range of interest; here we have used fitted polynomials instead, but we adopt this 

available solution nonetheless as an approximation. The solution is the same as that for 

an isothermal film except that it includes a correction factor that is a function of the melt 

viscosity at the inner and outer surfaces of the film. In the present context it is re-cast to 

give the following relation: 

 ) [( ( ) /(2 )( )] ( 1.5 2 0.333 δ a = μ Ts h ρ s g f mmelt − mgasif A

Here g is the acceleration due to gravity and f is the noted correction factor for the 

temperature dependence of the melt viscosity over the depth of the melt layer. It is a 

complex function of the ratio ) ( μ Tδ / ) ( μ Ts (i.e., 1/φ , from Eqn (A1.2); see Ref. A1.3). 

This factor is typically of the order 0.04 in the model solutions described here; it thus has 

about a factor of three effect on the value of δ a . 

Model Solution. This set of algebraic equations is solved for the unknowns Ts , Tδ , 

δ a and Vm . The equations are non-linear and require an iterative solution that involves 

two Newton-Raphson iterative solution loops for individual equations within the overall, 

multi-equation iteration process. As a first step, the non-melting solution is obtained for 

Ts to provide an estimate for the Ts value for the more general case with melting. The 

more accurate two-segment polynomial fits to the temperature-dependent melt viscosity 

(described elsewhere in this report) are problematical in the iteration scheme (particularly 

for temperatures near the patching of two segments) and therefore they are replaced by 

somewhat less-accurate single polynomial fit equations (as appropriate to each polymer 

of interest). 

Using the available input data for both temperature-dependent viscosity and for 

gasification kinetics, it is found that model solutions are not always available (i.e., the 

iteration scheme does not converge) without some adjustment of at least one parameter. 

The model has been applied to both PP702N polypropylene and to PS 666D polystyrene. 
127

It is found that the solutions do not exist for the polystyrene unless the value of As is cut 

by a factor of 4 or more. For the polypropylene, on the other hand, solutions do exist for 

the nominal parameter set; lack of solutions was encountered only when the value of 

As was increased to more than four times the value derived from the thermogravimetry 

experiments. For both polymers there was an incentive to increase the value of 

As because this brought the fraction of mass lost as melt more in line with the 

experimental results (see below). As discussed in the main body of the report and 

somewhat further below, both the kinetic parameters and the viscosity are used in a 

domain that is a considerable extrapolation beyond the conditions under which they were 

obtained. This may be the source of some of these problems; this is explored somewhat 

below. 

Model Results and Discussion. First consider the effect of the model parameter φ in 

Eqn (2) which was somewhat arbitrarily (but plausibly) set at a value of 25. Varying this 

value from 20 to 30 over the range of incident heat fluxes of interest (20 kW/m2 to 45 

kW/m2

) gives very little change in the solutions for such variables as the surface 

temperature and the total mass flow (sum of melt flow plus gasification rate). The 

greatest sensitivity is at the low heat flux end where this range of values of φ causes the 

surface temperature to vary by about 1.2 K and the total mass flow to vary by about 3%. 

At the high end of the heat flux range the variations in both variables are about half as 

much. Thus the solutions are rather insensitive to the value of φ and we retain a value of 

25 throughout the rest of the study. 

Figures A1.3a and A1.3b show a set of model solutions for PP702N polypropylene. Note 

that the mass loss rate values are for a 100 cm2

 sample face (10 cm tall by 10 cm wide, 

the same as the experimental samples), obtained by multiplying the model results for unit 

area by a factor of 100. For the results in these figures the value of As , the preexponential factor for the gasification reaction, was increased by a factor of 4 above that 

corresponding to the thermogravimetric experiments, as mentioned above. This increase 

was made to bring the fraction of mass lost as melt flow somewhat closer to the 

experimental values. Comparison of this prediction with the actual experiments (Figs. 

15) shows that the melt fraction is still somewhat higher than experiment. The model 

gives no convergence if the value of As is raised any further. (This issue is addressed 

further below.) Overall the solutions are not dramatically changed by increasing the 

value of As by a factor of 4. For example, at an incident heat flux of 30 kW/m2

, the 

factor of 4 increase lowers the total mass loss rate by 12% and lowers the surface 

temperature by 18 K; the fraction lost as melt goes down to 0.80 from 0.93. The overall 

comparison with experiment is thus not radically affected. 

Comparing Figs. A1.3a and A1.33b with Figs. 15a, b, c, one sees that the model gets all 

of the trends with varied heat flux correct. Thus, as expected, an increase in heat flux 

drives up the surface temperature and this drives up the overall mass loss rate from the 

polymer sample. The mass lost via melt also increases as the increased surface 

temperature lowers the melt viscosity and allows the melt to flow away faster. Note that 
128

the fraction of total mass that is lost via melt flow decreases; evidently the temperature 

sensitivity of the gasification process trumps that of the melt flow process (both are 

affected by chemical kinetic processes so this is not a foregone conclusion and it may not 

be a universal trend for all polymers). 

Further comparison of Figs. A1.3a and A1.3b with Figs. 15 shows that, for this 

polypropylene, the quantitative agreement between model and experiment is surprisingly 

good given the fact that the model is one-dimensional and the real experiment twodimensional. There may be a fortuitous element to this level of agreement but, overall, it 

adds an element of confidence that the model may be of some use in evaluating the 

behavior of such systems. The melt layer thickness (varying from about 1⁄2 mm at the 

highest flux to about 2 mm at the lowest flux) is of the same order as was inferred during 

the experiments (no direct measurements were made). The average melt flow velocity 

(varying from 1⁄4 mm/s at the lowest flux to 1 mm/s at the highest flux) is lower than the 

measured surface values for this polymer (e.g., ca. 2 mm/s at mid-height for 30 kW/m2

). 

The measured values are the maximum and the predicted values are the average over the 

depth of the melt layer; Ref. A1.3 indicates the difference between the average and the 

peak flow velocity for an analogous, variable-temperature melt flow is a factor of two. 

This observation partially explains the difference but implies that the model prediction of 

flow velocity is still low by roughly a factor of two. 

The predicted surface temperatures in Fig. A1.3b (mixed melting and gasification case) 

tend to be not far from experiment (ca. 20 oC low; Fig. 15b) at the low and mid fluxes but 

are about 45 oC below experiment at the highest flux. The experimental measurement is 

most difficult at this highest flux, so its accuracy is somewhat questionable, as discussed 

in the main body of this report. 

Figure A1.3b includes surface temperature predictions for two limiting cases of the 

steady-state model, pure gasification and pure melting. Figure A1.3c repeats the limiting 

case surface temperature results and also includes the limiting case total mass flow 

predictions. We, of course, have no comparable experimental data for these situations 

since these processes cannot be switched off and on. Fig. A1.3b shows that when both 

routes for heat dissipation are available, the surface temperature is lowered significantly, 

as one might intuitively expect. Figure A1.3c shows that, in spite of the very comparable 

surface temperatures for these two limiting cases, the pure melting case yields a 

substantially greater mass loss rate from the sample surface since it is an energetically 

less costly path. Comparison of Figures A1.3a and A1.3c shows that the mixed case, 

with both melting and gasification, does not yield still greater mass loss because the two 

paths are competing for the available energy – a relatively small amount of gasification 

removes a substantial amount of energy that could have gone into a relatively large 

amount of melt flow . Thus the mass loss rate in the mixed case falls somewhat below 

the pure melting case, in keeping with the large melt fraction shown in Fig. A1.3a. 

Polystyrene 666D is a polymer resin that is both significantly less thermally stable than 

the above polypropylene and that has a significantly higher viscosity over the relevant 

temperature range. Figs A1.4a, b and c show the model predictions for this resin. Here 
129

the value of the pre-exponential factor in the gasification reaction, As , had to be cut by a 

factor of four below that corresponding to the thermogravimetry measurements on this 

polymer in order to get convergent solutions. The results in Fig. A1.4, when compared 

with experimental results (figs. 16), are again better than one might expect given the 

simplified nature of the model. The trends are again in agreement with experiment and 

the absolute values of mass flows are comparable to experiment (ca. 50% lower). The 

surface temperatures are about 40 K higher than experimental values across the heat flux 

range. The fraction lost as melt is fairly close to experiment. We have no experimental 

values for layer thickness and flow velocity with which to compare the model but the 

predicted values of these are of the same order as with polypropylene and thus are not 

implausible. 

The same observations made above about competition between heat dissipation paths 

apply to this polystyrene as well. 

The noted limitation on the domain of available model solutions set by the value of the 

pre-exponential factor, As was examined further. Note that the solutions are based on 

parameters for both melt viscosity and for gasification kinetics that could not be 

measured in conditions most relevant to the actual experiments being modeled. Thus the 

melt viscosity was measurable only to temperatures between 300 and 350 oC whereas the 

model solutions give surface temperatures from 350 oC to 450 oC. The viscosity 

extrapolation to these higher temperatures is somewhat uncertain, especially given that 

the extrapolation is occurring on a semi-log plot. Also, the heating rates at which the 

gasification kinetics were measured (0.5 to 5 oC/min) are well below those in the 

experiments (25 to 150 oC/min, calculated from the steady-state model solutions). 

Increased heating rate pushes a reaction to higher temperatures where the controlling 

mechanism could possibly change; here it is assumed (unavoidably) that there is no 

change. These two sources of uncertainty imply not only that the absolute values of the 

model parameters are uncertain but also that the relative temperature dependencies of the 

two competing processes at the sample surface (mass loss by melt flow and by 

gasification) may be somewhat incorrect and that competition thereby skewed. This 

skewing of the competition may, in turn, limit the range of model solutions.92 To test 

this idea, the viscosity –temperature relation for PD702N polypropylene was somewhat 

modified to increase its rate of decay with temperature. Thus the experimental relation 

was multiplied by a linear function of temperature that went from unity at 230 oC to 0.33 

at 427 oC. Examination of the process by which the actual melt viscosity data are 

extrapolated to higher temperatures indicates that this high temperature reduction by a 

factor of three is not unreasonable. When this modified viscosity-temperature function is 

inserted in the model, the range of values for As for which steady-state solutions can be 

found is increased by 50%. This supports the idea that the limitations on the range of 

 

92 Interestingly, it is found that there is no difficulty in finding solutions to either of the limiting cases of the 

model (pure gasification or pure melting). The difficulty comes when the two processes compete. The 

gasification process cannot be so strong that it strips away the melt layer but the limitation on its magnitude 

seems to come well before this. 
130

model solutions are, at least to some extent, a result of deficiencies in the input 

parameters. 

It is of interest to note the distribution of incoming energy among the various sinks, as 

computed from the steady-state model. We do this for the polypropylene since it seems 

to agree somewhat more closely with experiment. Fig. A1.5 shows the distribution at the 

low and high end of the incident heat flux range. Note that re-radiation always accounts 

for a significant fraction of the energy budget, as does melt flow. Increased incident heat 

flux (up into the range provided by flames) markedly increases the fraction lost via 

gasification of the polymer, at the expense of the “passive” paths of re-radiation and 

convection. (Of course, if the sample were actually burning, there would be no 

convective heat loss.) 

The concept of a heat of gasification is frequently used to model the rate of gasification 

of a solid subject to a steady heat flux on its surface. This is the amount of heat needed to 

convert unit mass of the solid at room temperature to unit mass of gas at the surface 

temperature. The concept does not apply rigorously to charring solids since they are not 

capable of steady gasification, though it is sometimes used regardless. From a surface 

energy balance on the steadily regressing surface, one has: 

 qnet= qin − qloss = m L

 

where qin is the incident heat flux on the regressing surface and qloss is the net loss from 

the heated surface via re-radiation and convection; m is the steady mass loss rate from 

the surface via gasification (equal to the gasification reaction rate at the surface) and L is 

the heat of gasification. From its definition, L is given by the following 

 [ ( ) ] L = m Cp Ts − T0 + Qg

Comparison of the above with Eqn (A1.1) shows that they are basically the same except 

for the heat that goes into the melt flow. If the above is re-arranged to show the relation 

between m and the incident heat flux, the result is the equation for a straight line with a 

slope that depends on L : 

 m = (qin − qloss )/ L

Inspection of Fig A1.3c or Fig. A1.4c for the pure gasification cases shows that a linear 

relation like this tends to hold. However, note that the pure melting case also follows a 

linear relationship with a greater slope. The mixed gasification plus melting cases follow 

yet another line with an intermediate slope. As soon as melt flow is allowed, the 

response of sample mass loss to incident heat flux is no longer a direct measure of heat of 

gasification. This implies, for example, that thermoplastic samples will not behave the 

same in a device such as the cone calorimeter when tested in horizontal and vertical 
131

modes. This is yet another facet of how and why the flammability characteristics of 

thermoplastic materials are uniquely difficult to assess. 

The pure melting case is one extreme example of what one might do to a thermoplastic 

material to alter its flammability – shift the competing melt and gasification paths so far 

toward melting that all of the solid melts away from a heat source without ever getting 

hot enough to generate ignitable gases. There are real materials that behave in this 

manner. Unfortunately, this is no general guarantee of fire-safe behavior since the melt 

inevitably accumulates on some horizontal surface where it can be ignited by any falling 

bits of flaming material (not necessarily the thermoplastic material) typically found in 

any real product – the melt can no longer escape when it is laying on a horizontal surface 

and it may burn vigorously, possibly even yielding a self-feeding pool fire. 

An alternative fire-resisting treatment of a thermoplastic might be to greatly increase its 

melt viscosity to at least keep molten fuel from enlarging the burning area and spreading 

flames to new locations (at the expense of somewhat enlarged flames at the original fuel 

location). Melt viscosity can be increased by the addition of fine particle solids to 

polymer resins.93 The model can give some information on how responsive this type of 

system is to an increase in melt viscosity. Here again convergence limitations show up. 

For the polypropylene, the value of the pre-exponential factor was reduced to that which 

corresponds to the value deduced for this resin from thermogravimetry (rather than 4 

times this as in the above solutions). Only a factor of 9 increase in the melt viscosity (at 

all temperatures) could be accommodated before convergence limitations re-surfaced, but 

the results still show how this system would respond to attempts to suppress melt flow by 

increasing melt viscosity. The results at the lowest and highest incident heat fluxes are 

summarized in Table A-1; the reference case is that with the nominal, measured melt 

viscosity-temperature relation. 

Table A1-1. Effects of a Factor of Nine Increase in Melt Viscosity 

Incident Heat 

Flux 

(kW/m2

) 

Surface 

Temperature 

Change (oC) 

Change in Avg. 

Melt Flow 

Velocity 

Change in Melt 

Layer 

Thickness 

Change in Melt 

Flow Rate 

 

20 Up 3.1 Down 37 % Up 56.5 % Down 1 % 

45 Up 6.8 Down 50 % Up 80 % Down 6.8 % 

The picture that emerges from Table A1-1 is that of a system that is quite resistant to 

changes in viscosity; the effect of an order of magnitude increase in viscosity on melt 

flow rate is practically negligible. Surprisingly, the resistance to change comes more 

through adjustments in the melt flow layer than in the viscosity itself. One would expect 

 

93 As a practical matter, thermoplastics are usually formed by molding processes that require a low to 

moderate melt viscosity, so a very large increase in viscosity in the temperature range used for 

thermoforming (e.g., 200 oC to 250 oc) would be impractical. Some more subtle (presumably chemical) 

means would be needed to increase melt viscosity at temperatures above the thermoforming range, in 

response to impending ignition conditions. 
132

the system to have resisted change by having the surface temperature go up and thereby 

drive the local melt viscosity back down; here this seems to be a secondary effect. In any 

event an implication of these results is that it would take several orders of magnitude in 

viscosity change to substantially suppress the melt flow from a thermoplastic material. 

Thus this approach to modifying the flammability behavior of thermoplastics is not an 

easy one. 

Summary. The complex, time-dependent behavior of a vertical slab of a thermoplastic 

material subjected to a constant heat flux on one face, tends toward a quasi-steady mass 

loss behavior if the sample is thermally thick. That behavior is still inherently twodimensional but, if the polymer yields a low viscosity melt and, thereby, a thin downward 

flow of melt over its face, the behavior can be approximated as one-dimensional. It is 

then approximately describable by a set of algebraic equations which determine the 

relevant temperatures and mass loss paths. The model can capture the competition 

between mass loss by melt flow and mass loss by gasification. It reveals that any 

attempts to shift this competition call for large changes in model parameters and thus the 

flammability of thermoplastics is not easily altered via such an approach. 

Appendix 1 References. 

 A1.1) Glassman, I., Combustion, Academic Press, New York (1977), p. 74 

A1.2) Bird, R., Stewart, W. and Lightfoot, E., Transport Phenomena, Wiley and 

 Sons, New York (1960), p. 40 

A1.3) Bird, R., Stewart, W. and Lightfoot, E., ibid, p. 327 
133

MELT FLOW

Temp.

Temp.

MELT 

 FLOW

m

TOT

RADIANT

 FLUX

RADIANT

 FLUX

GASES

GRAVITY

GRAVITY

RADIANT

 FLUX

RADIANT

 FLUX

RE-RADIATION

CONVECTION

RE-RADIATION

GASES

Figure A1.1 (top). Schematic of 2-D thermoplastic slab 

 heated on one face. 

Figure A1.2 (bottom). Schematic of the energy and mass 

 flows at the steady-state condition.
134
135
136
137
138
139
140
141

Appendix 2 

Condition for a Self-Feeding Pool Fire 

It was noted in the main text that a worst case condition for a material which sheds 

flaming melt material is that its pool fire plume grows to the point where it strikes the 

object and thereby substantially increases the melt flow rate to itself because such a 

situation is potentially subject to an exponential growth in fire size.94 We call this a 

self-feeding pool fire situation. Note that a very low density material like a 

polyurethane foam will also tend to regress away from the impinging pool fire plume 

even as it feeds melt to it and this can mitigate the self-feeding tendency to a certain 

degree. Here we do not attempt to factor in all of the possible complications, but 

rather, we want to infer how the self-feeding condition varies with the size of the fire 

on the object. By confining our analysis to the condition at which self-feeding just 

begins to occur (taken as the condition when the pool fire plume tip [Ref. A2.4] just 

reaches the bottom of the object), we can infer results that should hold for both full 

density and low density thermoplastic materials. 

Consider a burning thermoplastic object that is at some height directly above its pool 

fire and, as a result of its burning, is losing some fraction of its mass as flaming melt 

feeding that pool fire. It feeds the pool fire at the rate 

 

 

/ .2.1 m feed = γ mobj = γ Q& obj Hc,obj A

Where mfeed is the mass flow rate (kg/s) of melt into the pool fire, γ is the fraction of 

melt that the object creates as it burns, mobj is the total mass loss rate of the burning 

object, Qobj & is the heat release rate of that burning object and Hc,obj is its heat of 

combustion. This creates a pool fire below the object whose quasi-steady diameter is 

given by a simple mass balance 

 [4 ( / ) / ] .2.2 0.50 d pool = γ Q& obj Hc,obj π m& pool A

where d pool is the diameter (m) of the pool fire and mpool & is the mass loss rate per 

unit area (kg/m2

s) of the pool fire due to its own burning. The burning rate of a pool 

fire can be estimated from [Ref A2.1] 

 3 m 0.015 f [1 exp( k d )] A.2. pool = − − β pool &

 

94 We are considering a geometry (pertinent to chairs or mattresses at some distance above a floor) in which 

the object which is the source of the melt is directly over the horizontal catch surface on which the pool fire 

exists. Other possibilities exist such as a thermoplastic object sitting on a horizontal catch surface with a 

pool fire next to it. In this case there is always an interaction and a tendency to self-feed. 
142

where f is a fraction whose base value is about 0.20 for pool a thin pool fire on a 

solid surface [Ref A2.2] and which we want to adjust for possible feedback between 

the pool fire and the object fire. The parameters kβ are based on a mean beam 

length estimation of feedback from the pool fire plume to the pool; we set their 

product equal to the value (3 m-1) given in Ref. A2.3 for polymethylmethacrylate (a 

somewhat oxygenated molecule that comes closest to our complex, somewhat sootforming flame gases). This pool fire has a plume height (flame tip height) given by 

the following turbulent fire plume correlation [Ref. A2.4] 

 1.02 0.235 ( ) .2.4 0.40 Lf ≡ hobj = − d pool + Q&

pool A

 

where Lf is the flame tip height and it is set equal to hobj , the height of the bottom of 

the burning object above the surface on which the pool fire is occurring; Qpool

& is the 

heat release rate (kW) of that pool fire. That heat release rate can be expressed in 

terms of the other parameters from above 

 

 ( / 4)( ) .2.5 ,

2 Q&

pool = m& pool π d pool Hc pool A

where Hc, pool is the heat of combustion of the melt in the pool, taken here as being 

equal to the heat of combustion of the object itself, Hc,obj . 

Given the above set of relations, one can pick a value for the heat release rate from 

the object, Qobj & , solve for the pool fire diameter (this requires a Newton-Raphson 

iteration to solve Eqns A.2.2 and A.2.3 together), then find the object height at which 

the pool fire plume tip is just reaching it. 

Figures A2.1 and A2.2 show the results of such calculations for a material with a heat 

of combustion of a typical polyurethane (30 MJ/kg). Recall that the melt fraction 

from polyurethanes was found to vary substantially. Here we have used values of 

0.50 and 0.75. In the main text it was also shown that the fire plume on the object can 

send a significant amount of radiation onto the pool fire boosting its burning rate. 

Thus we have also shown calculations for the case in which this effect doubles the 

pool fire burning rate. 

As expected, both Figures show that the larger the fire in the object the larger is the 

objects height above the catch surface at which the pool fire can interact with, and 

begin to boost, that object fire. Doubling the pool fire burning rate roughly doubles 

the reach of the fire plume. Increasing the object melt fraction by 50% has only a 

weak effect on the reach of the fire plume. Note that all of the separation distances 

fall between values used in the foam tests reported in the main text. 

The implications of these plots for scaling down a potential test method for 

polyurethane foam are discussed in the main text. 
143

Appendix 2 References 

A2.1 Gottuk, D. and White, D. “Liquid Pool Fires” Section 2, Chapter 15 in the SFPE 

Handbook of Fire Protection Engineering, 3rd Edition, National Fire Protection 

Association, Quincy, Massachusetts, p. 2-310 

A2.2 ibid, p. 2-311 

A2.3 Babrauskas, V., “Heat Release Rates”, Section 3, Chapter 1, ibid, p. 3-26 

A2.4 Heskestad, G., “Fire Plumes, Flame Height and Air Entrainment,” Section 2 

Chapter 1, ibid, p. 2-3 
144
145
146

Appendix 3 

A Limited Exploration of an Alternative Approach to TG Kinetics 

In the review process for this report, one of the reviewers (M. Nyden) suggested an 

alternative approach to inferring kinetic parameters for gasification of a polymer or its 

melt from weight loss data obtained by thermogravimetry. Its attraction is that it is 

capable of using (nearly) all of the thermogravimetry data obtained at several heating 

rates rather than just the data at the peak rate of weight loss as in the Kissinger method 

utilized in the main body of this report. 

If the thermogravimetry sample is assumed to be a collection of molecules that, at any 

given point in the linear heating process, are at the same temperature and are reacting to 

that temperature by degrading in an effectively single, first order, overall reaction step in 

a manner independent of any mass transfer processes95, then at all times during the 

weight loss process, regardless of heating rate, we can write: 

 dm dt k (m m ) (A3.1) = − R

which says that the rate of sample mass loss is proportional to the first power of the 

remaining mass minus some non-participating residual mass, mR (if any). Here the 

proportionality constant is the kinetic rate constant whose parameters ( A and E ) we 

seek to characterize: 

 k = A exp ( − E / RT ) (A3.2)

If the above assumptions about the sample hold true, we can solve for k in Eqn (A3.1) to 

get: 

 k (1/(m m )) (dm dt) (A3.3) = − R

In this manner, one can use the derivative of the entire sample mass loss curve to 

calculate a running value of the kinetic rate constant for all mass loss data at all available 

heating rates. If the sample is “well-behaved” the resulting set of values for k can be 

plotted versus the inverse of the absolute temperature and one will find unique values for 

the pre-exponential factor, A , and the effective activation energy, E . Nyden [Ref. 

A3.1] found this approach to work for a polyisobutylene sample at four heating rates, 

obtaining an essentially unique pair of values for the two kinetic parameters for weight 

loss rate from that polymer. 

Here we investigated this approach for a sample of polypropylene melt generated at 

 

95 Such processes might limit the rate of escape of volatile molecules formed by the degradation reactions. 

These assumptions are common to any kinetic analysis of thermogravimetry data. 
147

20 kW/m2

 with a gas-fired panel as a radiant heater. Thermogravimetry data for this 

degraded polymer were obtained at three heating rates, 0.5, 2 and 5 oC/min. The mass vs 

time data were differentiated and used in Eqn A3.3 to calculate values of k versus 

temperature at all three heating rates. 

A semi-log plot of these results, in terms of k vs. (1/T) is shown in Fig A3.1. In this 

form, the slope of the curve is proportional to the effective activation energy and the 

intercept yields the pre-exponential factor. Also shown on the graph is a linear fit to the 

total data set (all three heating rates). 

It is clear in Fig A3.1 that all three curves have significant curvature, tending toward a 

higher slope at higher temperatures (left side of the graph). For the three heating rates, 

the change in slope is not clearly related purely to temperature, nor, one can show, is it 

related purely to the extent of the reaction. Generally, however, it occurs in the 

temperature range from 370 oC to 405 oC. 

In general, a one step kinetic rate law has one more available parameter, the reaction 

order. That is, a more general version of Eqn 3.1 is 

 ) dm/ dt k (m m ) ( A3.4 n = − R

where n is the reaction order. Here, however, we have no data manipulation procedure 

that yields a value of reaction order. A trial and error process was therefore used to 

investigate the issue of whether values of n other than unity (larger or smaller) had an 

effect on the curvature seen in Fig A3.1, i.e., could a different choice of this additional 

parameter yield a straighter line in this type of plot? Thermogravimetry results from a 

heating rate of 2 oC/min were used for this purpose. The result was that values from 0.50 

to 1.50 only made the plot of k vs. (1/T) look worse. Values less than unity tended to 

introduce a new, reverse curvature in the plot at high temperatures. Values above unity 

increased the overall upward curvature seen in Fig A3.1. We infer that unity is the most 

appropriate value for this parameter and that the curvature in Fig A3.1 is fundamental. (It 

should be noted that a reaction order very near unity was reported for polypropylene 

degradation in Refs A3.2 and A3.3) 

If the curvature is real then the reaction process is not as simple as Eqn A3.1 assumes. 

Either the effective activation energy varies (evidently more so with temperature than 

with extent of reaction96) or the system requires two or more sequential reactions to 

describe its mass loss behavior. Either would complicate the overall melt fire modeling 

problem. 

The use of the average line fitted over all three heating rates shown in Fig A3.1 yields a 

unique value for both the activation energy (170 kJ/mol) and the pre-exponential factor 

(1.19·1010 s-1). This value of activation energy falls within the wide range of values 

 

96 There are data reduction schemes in the thermal analysis literature which can yield an effective activation 

energy as a function of extent of reaction [e.g., Ref. A3.4]. 
148

reported in the literature for polypropylene (A3.2, A3.3). This pair of values does predict 

weight loss curves that fall in the approximately correct temperature range (350 oC to 450 

oC). Recall, however, that the shift of the reaction with heating rate is a function of 

activation energy. Since this simple description uses an average activation energy, it is 

not quantitatively correct with regard to the shift in the position of the curves with 

increasing heating rate. This is not particularly troublesome when matching just the 

thermogravimetry results. However, the kinetic parameters are to be used in a melt fire 

model in which the heating rates will be substantially higher (a factor of ten or more 

higher than in thermogravimetry). Given this, it is possible that the balance of 

gasification of a melt and the flow of the melt could be seriously miscalculated if the 

activation energy for gasification is substantially incorrect. One possible way of 

avoiding this, short of using a more complex, multi-step gasification description, would 

be to use the activation energy one would infer from Fig A3.1 (or its analog for another 

material) at the lower temperature, lower conversion end of the weight loss process. This 

has more of a chance of putting the gasification in proper competition with other 

processes in a model of burning, though it may still incorrectly estimate the amount 

gasified. 

Overall, it appears that this alternative approach to inferring kinetic parameters is more 

informative than the Kissinger method used in the main body of this report. However, 

the additional information it brings can point to a need for increased complexity in the 

modeling description. 

Appendix 3 References 

A3.1 Jee, C., Guo, Z., Stoliarov, S. and Nyden, M., “Experimental and Molecular 

Dynamics Studies of Thermal Decomposition of a Polyisobutylene Binder,” Acta 

Materialia 54, pp. 4803-4813 (2006) 

A3.2 Gersten, J., Fainberg, V., Hetsroni, G. and Shindler, Y., “Kinetic Study of the 

Thermal Decomposition of Polypropylene, Oil Shale and Their Mixture,” Fuel 79, 

 pp. 1679-1686 (2000) 

A3.3 Bockhorn, H., “Kinetic Study of the Thermal Degradation of Polypropylene and 

Polyethylene,” Journal ofAnalytical and Applied Pyrolysis 48, p. 373 (1999) 

A3.4 Friedman, H. 1964–1965. J. Polym. Sci. C 6: 183– 95 
149
150

Appendix 4 

Cone Radiation Absorption In-Depth in Polymer Resins 

The following measurements were carried out after the main body of the text of this 

report was written. It was noted there that, in the various cone calorimeter-based polymer 

melt/gasification experiments, in-depth absorption of the cone radiation had a potential 

influence on the results. However, the only data available on this issue that appeared 

relevant to the polymers used here were obtained with a different apparatus (a larger 

diameter cone heater), operated at a different temperature, and the polymer resins came 

from a different source. To rectify this, we carried out a measurement of radiation 

penetration on the specific polypropylene and polystyrene resins used throughout the 

bulk of this report and we did this in conditions directly relevant to the conditions 

reported in the main body of this report. 

The exact geometry and source temperature affect the extent of radiation penetration 

below a polymer surface. The geometry determines the range of angles over which the 

radiation strikes the polymer surface. Thus, a ray striking the surface at, for example, 

45o, penetrates a lesser distance in the direction normal to the sample surface than does a 

ray striking the surface at 90o. The cone heater provides radiation over a range of angles 

and that range depends on the spacing between the cone and the irradiated sample 

surface. Furthermore, the cone heater can be regarded as a gray-body radiator. The peak 

of the radiation spectrum it emits then roughly follows Wien’s Displacement Law, 

moving to shorter wavelengths as the cone temperature is turned higher; this alters the 

amount of radiation interacting with peaks in the infrared absorption spectrum of a 

polymer and thus the amount of radiation absorbed up to any given depth. In the 

experiments reported in the main text, the distance between the base of the cone heater 

and the polymer sample surface was fixed at 7.6 cm and the radiant flux was varied by 

changing the cone heater temperature. Here we used the same cone heater distance and 

we focused on the cone heater temperature (ca. 680 oC) that provided an incident heat 

flux of 30 kW/m2

, the middle value in the preceding experiments. 

The physical set-up for these measurements was simple. The cone heater, oriented 

horizontally, was placed 7.6 cm above a thin sheet of the polymer resin of interest. Just 

below (2 mm) the back surface of the sample, a Schmidt-Boelter total heat flux gage (6 

mm diameter) was placed (on the cone heater centerline). This type of flux gage has a 

time response of 0.1 to 0.2 sec. according to the manufacturer and was previously 

calibrated to a heat flux accuracy of approximately ± 3 %. The high incident heat flux 

meant that the sample was heated rapidly. The 2 mm spacing above the flux gage 

minimized the conductive flux from the warming sample to the gage but it did nothing to 

halt sample re-radiation. To establish the position of what was therefore an upward 

shifting baseline, the cone radiation was chopped at ca. 1⁄2 Hz using a handheld piece of 

sheet metal. This permitted acquiring 10-15 seconds of absorption data before the sample 

began to degrade. 
151

The sheet samples were made by hot pressing the polymer beads. This did not provide a 

very flat sample and it left a few bubbles in random locations. Since circles of only about 

6 mm diameter were needed for each measurement (on a sheet that was 7.6 to 10.2 cm 

across), clear, bubble-free areas were selected and marked so that they could be placed 

over the flux gage. The local thickness of these areas was measured with a micrometer 

(0.01 mm resolution). Duplicate sheets of each resin with nominal thicknesses of 0.8 

mm, 1.6 mm, 3.2 mm and 6.4 mm were made; each was tested in at least two locations. 

The resins were PD702N and 6523 polypropylene plus 663 and 666D polystyrene. 

The raw data in this experiment are affected by reflection of radiation at two surfaces – 

the front and back of the resin sample. That is, the amount of radiation reaching the heat 

flux gage is diminished both by the reflection process and by the absorption process that 

is the real point of interest. A model of the heating of the polymer by radiation needs 

information on both the loss of incident radiation by reflection at the front surface of the 

sample and on the deposition of the remaining radiation in the depth of the sample. 

These polymer resins are dielectric materials whose reflectivity is dictated by their 

refractive index (A4.1). In general, the reflection of light at the surface of a dielectric 

material is a function of the angle of incidence and that function, in turn, depends on 

whether the light is polarized. Here the light is unpolarized and the range of incidence 

angles (0 to ca. 45o) is well within the range (0 to ca. 60o) for which the reflectivity of 

unpolarized light is independent of angle. Both polystyrene and polypropylene have 

refractive indices of about 1.4 to 1.697; based on this we used a value of single-surface 

reflectivity of 4%. The raw data from the measurements are thus divided by a factor of 

0.92 (i.e., (1 – r)2

) to obtain the amount of radiation actually deposited within the polymer 

sample.98

The main interest here is the deposition of radiant heat within the polymer bulk. This is 

obtained by differentiating the measured attenuation function obtained here (and 

multiplying by the heat flux that penetrates the outer surface of the material). This 

becomes the radiant source term in a transient energy equation for a polymer sample. 

The results for these measured attenuation functions for the four resins examined here are 

shown in Figures A4.1 to A4.4. 

In all cases it was possible to fit the data very well with a double exponential decay 

expression, as shown on each plot. A simple interpretation99 of this form comes from the 

nature of the near to middle infrared spectra of these types of polymers. Both 

polypropylene and polystyrene are weak absorbers out to about 5 1⁄2 microns except for a 

strong absorption band centered at about 3 1⁄2 microns. The cone heater radiation at 680 

oC can be expected to peak at about 3 microns. Thus a substantial portion of the radiation 

just above 3 microns plus that portion above about 5 1⁄2 microns is strongly absorbed and 

 

97 We ignore the anomalous behavior of refractive index in the neighborhood of strong absorption lines. 

98 We also ignore the higher order terms that come from considering the possibilities for multiple internal 

reflections. 

99 We note that this fitting function also has four adjustable constants so the good fit is also enabled by this 

flexibility. 
152

everything else is very weakly absorbed. The figures show that about half of the energy 

is deposited in the first half millimeter and the rest is spread over the next several 

millimeters. 

Not surprisingly, the absorption curves for the two polypropylenes are similar to each 

other. The same can be said for the two polystyrenes. The real surprise is that the 

polypropylenes and the polystyrenes are so similar. We do not have infrared spectra on 

these specific resins but the available spectra for these types of resins (cf., Sadtler 

Standard Spectra) indicate that polystyrene has a more complex and broader absorption 

band around 3 1⁄2 microns, compared to that for polypropylene. Evidently such 

differences have a relatively small impact for the conditions studied here. 

A concern for polypropylene is that at room temperature up to 165 oC, it has two phases, 

one of which is crystalline. This renders it a scattering medium, at least in the visible 

portion of the spectrum. If it scatters in the infrared also, its attenuation function could be 

expected to change appreciably from that measured here for temperatures above the 

crystalline melting point. We had no data on the crystallite size in our samples but we 

note that if they were of the order of 1 micron or less, scattering in the infrared would be 

negligible (even though appreciable in the visible). A search of available polypropylene 

IR spectra (Sadtler Standard Spectra) showed no signs of the kind of inclined baseline in 

non-absorbing, short wavelength regions (1-2 microns) that scattering tends to cause. An 

attempt was made to perform our cone radiation attenuation experiments at temperatures 

above 165 oC by laying the sample on top of a layer of zinc selenide (an infrared 

transparent support surface) to prevent it from sagging onto the top of the flux gage. The 

samples always underwent severe physical distortions and dimensional changes by the 

time the crystalline phase melted so no useful data could be obtained in this manner. At 

this point we have only the rather weak inference from the available infrared spectra that 

scattering in not an issue with polypropylene. 

Reference for Appendix 4 

A4.1 F. Jenkins and H. White, Fundamentals of Optics, 3rd Edition, McGraw-Hill Book 

Company, New York (1957), p. 511 
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