NIST Technical Note 1929 

Numerical Simulations of 

Thermoplastic Burning Rate: Effect of 

Property Variations and the Inferred 

Effective Heat of Gasification 

Gregory Linteris 

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NIST Technical Note 1929 

Numerical Simulations of 

Thermoplastic Burning Rate: Effect of 

Property Variations and the Inferred 

Effective Heat of Gasification 

Gregory Linteris 

Energy and Environment Division 

Engineering Laboratory 

This publication is available free of charge from: 

http://dx.doi.org/10.6028/NIST.TN.1929 

September 2016 

U.S. Department of Commerce 

Penny Pritzker, Secretary 

National Institute of Standards and Technology 

Willie May, Under Secretary of Commerce for Standards and Technology and 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 1929 

Natl. Inst. Stand. Technol. Tech. Note 1929, 41 pages (September 2016) 

CODEN: NTNOEF 

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NUMERICAL SIMULATIONS OF POLYMER BURNING RATE: EFFECT OF
 

PROPERTY VARIATIONS AND THE INFERRED EFFECTIVE HEAT OF
 

GASIFICATION*
 

Gregory Linteris
 

Energy and Environment Division
 

National Institute of Standards and Technology
 

100 Bureau Dr. Stop 8631
 

Gaithersburg, MD 20899-8631
 

ABSTRACT 

The mass loss rate of Poly(methyl methacrylate) exposed to known radiant fluxes is simulated with two 

recently-developed numerical codes, the NIST Fire Dynamics Simulator (FDS) and the FAA ThermaKin. 

The influence of various material properties (thickness, thermal conductivity, specific heat, absorption of 

infrared radiation, heat of reaction) on mass loss history is assessed, via their effect on the ignition time, 

average mass loss rate, peak mass loss rate, and time to peak. The two codes predict the influence of 

material parameters on the MLR in the order of decreasing importance: heat of reaction, thickness, 

specific heat, absorption coefficient, thermal conductivity, and activation energy of the polymer 

decomposition. Changes in the material properties also influence the MLR curves by switching the 

sample from thermally thick to thermally thin. The two numerical codes are generally in very good 

agreement for their predictions of the MLR versus time curves, except when in-depth absorption of 

radiation was important. The influence of two of these parameters on the effective heat of gasification 

(extracted from the mass loss rate at differing external heat fluxes) is also assessed by comparing the 

effective heat of gasification to the heat of gasification input into the calculations; the ratio of these is 

highest for calculations in which the radiant energy is absorbed on the surface and the material 

decomposition has a low activation energy, for which it reaches 1.3, but is lower for other cases. 

Modifying the method used to obtain the effective heat of gasification by basing it on the net energy 

delivered to the polymer brings it much closer to the values input into the calculation. 

KEYWORDS: 

Material flammability, heat release rate, heat of gasification, polymer burning rate 

* Official contribution of NIST, not subject to copyright in the United States. 

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NOMENCLATURE 

∆Hg – heat of gasification, kJ kg-1 

∆Hreac – heat of reaction, kJ kg-1 

∆Hv – heat of vaporization, kJ kg-1 

T – temperature, K 

m "– mass loss rate, g m-2 s

-1 

q " – net heat flux, kW m-2 s

-1 

net 

q

" 

ft – heat flux from flame to surface, kW m-2 s

-1 

" -1 qloss –heat flux from surface to ambient, kW m-2 s

" 

qext – externally applied heat flux, kW m-2 s

-1 

k – thermal conductivity, W m s-1 

c – specific heat, kJ kg-1 K-1 

Ea – Arrhenius activation energy of one-step material decomposition 

A – Arrhenius pre-exponential term for one-step material decomposition 

R – Universal gas constant, kJ mol-1 K-1 

S – sample thickness, m 

Greek Symbols 

α – absorption coefficient, m-1 

δ – thermal thickness (with respect to mass loss), m 

δt -thermal conduction length, m 

τign – ignition time, s 

Abbreviations 

MLR – mass loss rate 

MLRav – average mass loss rate 

EHF – external heat flux 

1. INTRODUCTION 

The prediction of fuel generation rates from burning solid materials is an important component of models 

for fire growth in buildings. Validation of these sub-grid models is often done via predictions of the mass 

loss rate for small samples subjected to a known radiant heat flux, in standard devices such as the cone 

calorimeter [1], the FM Global Fire Propagation Apparatus (FPA)[2], and other devices [3]. The mass 

loss rate of thermoplastics in these devices has been predicted analytically and numerically by various 

groups [4,5,6,7,8,9,10,11,12]. 

Accurate prediction of the mass loss rate (MLR) requires the input data for the physical parameters of the 

polymer used in the model. In general, two approaches are being taken to supply these parameters: 

estimating them from the mass loss rate (or temperature) data obtained in the experiments to be modeled 

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themselves [4,7,8] (sometimes with parameter optimization algorithms [13]), or measuring the individual 

parameters with separate experimental devices [5,9,12]. In either case, it is of value to understand the 

sensitivity of the mass loss rate to the individual parameters in the model. In the former case, such 

knowledge can allow one to design specific experimental runs for the most accurate extraction of certain 

parameters. In the latter case, knowing the sensitivity to each of the parameters can allow one to spend the 

limited resources on measuring the most relevant parameters, and to use simpler methods or estimations 

for parameters less important. Finally, mass loss rate curves and ignition times provided here can be used 

as an aid in interpreting cone calorimeter data. The present results are in the spirit of previous work in 

which the variation of mass loss rate with thickness [14], and with heat flux [15] have been illustrated, 

and extend the phenomenological illustration to additional parameters as suggested by Schartel [16,17]. 

Recently, Stoliarov et al. [18] have performed sensitivity calculations and reported the results of input 

parameter variation in terms of their effects on global parameters (such as the peak heat release, average 

heat release, etc.). The present investigation seeks to provide additional information beyond that in ref. 

[18], by presenting the data as the time-dependent mass loss rate. The collection of simulated mass loss 

curves can serve as a database illustrating the effects of changes in the parameters on the burning 

behavior, providing physical insight to developers of less flammable materials. Such calculations are 

easier and faster than the comparable experiments would be, and can be performed with individual 

parameters changed in isolation (in ways not always possible in experiments). In the second part of the 

paper, the simulation techniques are employed to examine the accuracy of one method of extracting the 

effective heat of gasification from experiments done in a radiant heating device. 

While much of the mass loss and ignition behavior of PMMA under constant heat flux is already known, 

often, not all of the physical effects are included in the analysis, and when many are, it is usually not 

possible to obtain closed-form solutions to the equations. Hence, it seemed of value to provide a 

compendium of the effects of multiple parameters on the thermal decomposition of PMMA. In the 

discussion, emphasis is placed on those conditions for which the behavior deviates from that expected 

based on simple models. 

2. APPROACH 

The time-dependent mass loss rate of a thermoplastic material subjected to a known radiant flux has been 

simulated. Two numerical simulation codes describing the solid phase have been used: the NIST Fire 

Dynamics Simulator (FDS 5.3.0. SVN 3193) [19] and the FAA Thermo-kinetic Model of Burning 

(ThermaKin) [20,21]. Poly(methyl methacrylate) (PMMA) was selected as the base polymer for 

simulation, since it is a typical thermoplastic, is nearly a standard material in flammability studies, and its 

properties are relatively well studied. The simulations of the MLR were used in the present work to 

examine how variations in the input parameters affect: 1.) the time-dependent MLR and the ignition time 

(Part I); and 2.) the value of the effective heat of gasification extracted from the mass loss data using the 

method of Tewarson and Pion[4]. 

The input parameters in the models are shown in Table 1. Part I of the present study (parametric analysis) 

uses the nominal parameter values of Rhodes and Quintiere [8] as a base-case (column three), and these 

parameters are varied over the range of values as indicated in the column four. Part II of the present study 

uses the parameter values of Staggs [22] (to allow more direct comparisons with the calculations of 

Staggs) and they are listed in the column five; some were varied over the range indicated in column six. 

The major differences between the two nominal sets are the 36 % lower c and the 58 % lower ∆Hg in 

Staggs analysis. 

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Typically, the parameters in Table 1 are varied over a factor of five, about a factor of 2 to 2.5 above and 

below the nominal value. This is a somewhat larger range than experienced by typical polymers [18] and 

was selected to provide guidance in the event that new, composite polymers are produced with a wider 

range of properties than in the pure polymers. 

The calculated curves of MLR versus time are characterized by the ignition time, average MLR, peak 

MLR, and time to peak MLR. In the present analyses, the ignition time τign is defined as the time at 

which the mass loss rate has first achieved a value of 3 g s-1 m-2 , as suggested in ref. [23]. The average 

mass loss rate (MLRav) is defined as the integral of the MLR versus time curve, between the time at which 

the flux is applied, and the time at which the MLR has decreased back to the mass loss at ignition (3 g s1 m-2 ). The parameters varied are the material properties: thermal conductivity (k), specific heat (c ), 

extinction coefficient (α), heat of reaction (∆Hreac), and the Arrhenius pre-exponential term (A) and 

activation energy (Ea) for the one-step material decomposition reaction. The experimental parameters 

varied are the external heat flux (EHF) and material thickness (S). 

The most important material property affecting the burning rate of a polymer is its heat of gasification 

∆Hg [24]. For a simple vaporizing material the heat of gasification can be described by 

T

∆H = v

c (T )dT + ∆Hv in which the first part is the sensible heat in raising the material from the g ∫Ta 

ambient temperature Ta up to the temperature at which it vaporizes Tv, and the second part is the heat of 

vaporization ∆Hv. This simple model is often used to describe more complicated solid materials which 

undergo endothermic pyrolysis reactions to form gas-phase products, by substituting a pyrolysis 

temperature Tp, and a heat of reaction ∆Hreac (i.e., a heat of decomposition in going from the solid directly 

to the gas-phase species). 

∆ H g = ∫T

Tp 

a 

c T dT + ∆Hreac ( ) 

Eq. 1
 

The assumption is often made that the material undergoes one-step decomposition to gaseous products,
 

with either an infinite rate or a finite rate given by an Arrhenius rate expression. Methods have been
 

developed to estimate the heat of gasification, either from measurements of the steady burning rate of the 

polymer exposed to different external fluxes (and called the “effective” heat of gasification, ∆Hg

(eff) [4]), 

or by differential scanning calorimetry [25,26].
 

The use of cone calorimeter (or FPA) data to estimate the effective heat of gasification has great utility
 

and is widely used [27]; the method is straightforward, and can easily be part of other flammability tests 

for the sample. In the method of Tewarson and Pion [4], the steady mass loss rate is plotted against EHF, 

and the inverse of the slope is taken as ∆Hg

(eff). Nonetheless, the method is based on some simplifications 

in the polymer’s behavior, and it is useful to understand the effects of other parameters on the accuracy of
 

the technique. In particular, the method assumes that at the different levels of the external heat flux, both
 

the heat feedback from the flame to the polymer and the surface heat losses from the polymer are constant
 

(i.e., constant surface temperature). Also, any radiation blocking by gasification products is assumed
 

constant.
 

Using an analytical solution for the steady mass loss rate, Staggs generated synthetic mass loss data, and
 

examined the accuracy for the method of Tewarson and Pion for the case of PMMA and polyethylene
 

(PE). He found that ∆Hg

(eff) was 20 % to 40 % higher than the values of ∆Hg input into the model. Since 

∆H g is the most important parameter affecting the mass loss variation with time, and since it is 

commonly found using the method of Tewarson and Pion , it is of interest to examine the influence of
 

other parameters (besides the few cases examined by Staggs) on the accuracy of the technique.
 

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In the approach taken in Part II, the mass loss rates vs. time are generated numerically for a range of EHF, 

using known input values for c, ∆Hreac, and hence ∆Hg. Then, using these synthetic MLR curves, the 

effective heat of gasification, ∆Hg

(eff), is extracted from these curves, and compared to the values of ∆Hg 

which were input. By repeating the calculations with different values of the material properties, the 

accuracy of the extracted value of ∆Hg

(eff) is assessed, along with its variation with material properties. 

Since there is no flame in the simulations (only radiant heat), the inaccuracies in the method of Tewarson 

and Pion, with respect to variations in the flame heat feedback and radiation blocking, can be avoided. 

Also, influences on the mass loss rate due to charring, bubble formation, and other effects, are avoided, 

since they are neglected in the calculation. 

3. NUMERICAL CALCULATIONS 

In the calculations with FDS or ThermaKin, the time-dependent 1-D heat transfer equation is solved for 

the solid phase, subject to mass (and energy) changes from reaction at any depth. There was only one 

solid-phase component, corresponding to PMMA, with thermal conductivity and specific heat either 

given by a linear correlation [12], or by the average value between the initial temperature and the 

pyrolysis temperature. Thermal decomposition was via a one-step Arrenhius-type reaction (rate = 

A exp(-Ea/RT) ) to a single gas-phase component, and there was no resistance to transport of the reacted 

polymer to the surface. Surface reflection and re-radiation are included. The initial temperature of the 

polymer was set to 293 K, the surface emissivity was set to 0.95 (1 in Part II), and the sample backside 

was insulated (adiabatic) with a reflectivity of 1 (i.e., a foil-wrapped sample). The convective heat 

transfer coefficient of the heated face was set to zero for the parametric analysis in Part I (to allow easy 

comparison of the FDS and ThermaKin results), and to 10 W m-1 K-1 for the analysis in Part II (to 

facilitate comparisons with the analysis of Staggs). 

In the ThermaKin calculations, the calculation was 1-D, the finite element objects were 5 × 10−5 m thick, 

and the time step was 0.025 s. Reducing these by a factor of four had no significant effect on the MLR, 

and only about a 3 % effect on the ignition time (which was more sensitive). In-depth absorption of 

radiation was treated using a random-sampling technique in which the energy is deposited at random 

locations in the 1-D sample, following a Beer’s law attenuation, with energy emission then following 

from the same location. Because of the discrete (and random) treatment of radiation, the calculated mass 

loss (sampled every 2 seconds) had fluctuations, so the data were smoothed (running average) for 8 s. 

In the FDS calculations, all gas-phase reaction was turned off by setting the mass fraction of O2 in the air 

to 0.01 and a minimum gas-phase mesh was used. The time step was set to 0.1 s, except for ignition time 

calculations, for which it was set to 0.005; this value yielded ignition times within 1 % of those with a 

time step of 0.01 s. The parameters used for the solid-phase calculation (which has adaptive gridding) are: 

STRETCH_FACTOR = 2.0, CELL_SIZE_FACTOR = 0.06, and REGRID_FACTOR=1.0, which provide 

the most grid cells, of the most non-uniform size, and which are re-grided the most often. The initial 

gridding gave a smallest grid size of 1.67 x 10-5 m to 3.33 x 10-5 m (depending upon the case). 

The performance of ThermaKin has recently been validated in predictions of the gasification rate of 

poly(methylmethacrylate), high-impact polystyrene (HIPS), and high density polyethylene (HDPE) [12], 

and it is used predominantly in the present work. Calculations with FDS are provided for comparison 

purposes. 

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4. RESULTS 

4.1. PART 1: Parametric Analyses 

4.1.1. Mass Loss Rate vs. Time 

External Heat Flux (EHF) 

The external heat flux is not a material property, but it is worthwhile to first illustrate the effect of this 

parameter on the mass loss rate. Figure 1 shows mass loss rate for PMMA with the nominal properties of 

Table 1 exposed to an EHF of (21, 50, 100, 150, or 200) kW/m2

. Simple analytical models of the 

thermoplastic burning rate [4] show that the steady mass loss rate m " is given by 

dm" qnet " 

MLR = = m "= Eq. 2 

dt ∆H g 

" " " " " " where q is the net external heat flux, and q = q + q − q , in which  is the heat feedback net net fl ext loss q fl 

from the flame to the surface, q

" is the external heat flux, and " is the heat lost from the surface by ext qloss 

convection and radiation. The burning time tburn is given by ρ S ∆H / q " , where ρ and S are the g net 

density and thickness of the1-D sample. Since the heat losses and flame heat feedback are relatively 

" constant as compared to the changes in the external flux qext , the mass loss rate is approximately 

proportional to both the net and incident heat flux (as indicated in Figure 1), and the burning time is 

" inversely proportional to qext . 

The thermal thickness (δ ) of a material under steady surface regression can be given [28] by 

2k∆H g δ = . Eq. 3  " C p qnet 

A sample is thermally thin if S /δ ≤ 1 and thermally thick if S/δ >1. For the nominal PMMA properties, 

the thermal penetration depth is about (45, 13.7, 6.3, 4.1, and 3.0) mm for incident fluxes of (21, 50, 100, 

150, and 200) kW/m2 (assuming a surface temperature (Ts) of 352 °C and no convective losses). 

Including the surface convective heat losses, about 3 kW/m2 for PMMA, would increase the thermal 

penetration depth by about (35, 9, 4, 2, and 2) % for these incident fluxes. The low flux case (21 kW/m2

) 

in Figure 1 is thermally thin, while the others are thermally thick. Another way to show the thin vs. thick 

behavior is to normalized the MLR curves by the steady-burning value (Eq. 2) for the mass loss rate 

( q

" / ∆H ), and for the burning time ρ S ∆H / q " . This is shown in the inset of Figure 1; the curves net g g net 

are reasonably coincident for the thermally thick cases, but deviate for the thinner cases. 

These results illustrate that for 25.4 mm thick PMMA, the sample response for fluxes in the EHF range of 

21 kW/m2 to 75 kW/m2 will readily shift from thermally thick to thermally thin as the various parameters 

are changed. This will become apparent as the mass loss rate curves are discussed below. (Note that the 

absorption coefficient would enter in the denominator of Eq. 3 above, since a low value of α acts like 

high thermal conductivity, sending heat more readily in-depth in the sample.) For all values of the flux, 

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the peak MLR is 38 % to 45 % higher than the average MLR, and the time to peak MLR scales with the 

burning time. 

Sample Thickness 

Sample thickness is also not a material property, but is useful to examine. Figure 2 shows the mass loss 

rate of nominal PMMA at EHF values of 21 kW/m2 (left frame) and 100 kW/m2 (right frame) for sample 

thickness of (2, 4, 8, 15, and 32) mm. At the low flux, the sample is always thermally thin; at the higher 

flux, the behavior becomes thermally thin as the sample thickness decreases to about 8 mm, which is 

consistent with the numbers given above. Both the average and peak MLR are nearly unchanged as S 

decreases, except for the thinnest samples for which both the peak and the average MLR drop off, the 

average somewhat faster, about (30, 40, 48, and 56) % lower at (21, 50, 100, and 200) kW/m2

, for the 

2 mm sample as compared to the 32 mm sample. This is due to insufficient thickness for complete 

absorption of the thermal radiation. For example, the dotted lines in Figure 2 illustrate the results for S = 

2 mm and α = 50000 m-1 (surface absorption), for which the peak and average MLR are restored closer to 

the values at larger S, only (-5, 2, 7, and 16) % lower for the 2 mm sample as compared to the 32 mm 

sample at (21, 50, 100, and 200) kW/m2

, respectively. 

Heat of Reaction, ∆Hreac 

The heat of reaction has the largest effect on the mass loss rate of any material property. Figure 3 shows 

the mass loss rate as a function of time for PMMA with an EHF of a.) 21 kW/m2 and b.) 100 kW/m2

. The 

different curves in the figures show the results for ∆Hreac = (1000, 2000, 3000, 4000, and 5000) kJ kg-1 . 

For higher flux, the behavior is always thermally thick, while for low flux, the behavior is thermally thin 

for all values of ∆Hreac except 1000 kJ kg-1, for which it’s almost thermally thick. At either 21 kW/m2 or 

100 kW/m2

, and the lowest value of ∆Hreac, (upper most curves), the peak heat release is much higher than 

the average heat release as compared to the other cases of ∆Hreac. This is because for low ∆Hreac, the 

sensible part becomes a larger fraction of ∆Hg, and the transient effect from the pre-heating of the indepth layers of the polymer due to thermal diffusion becomes larger relative to the enthalpy change due to 

decomposition. For the nominal polymer properties, the algebraic relationship in Eq. 2 shows 

MLR ~ ∆Hg n , where n=0.72, which is close to the behavior shown in Figure 3, where the value of n 

ranges from 0.64 to 0.80, depending upon EHF. 

Activation Energy, Ea 

Variations in the activation energy of the PMMA decomposition reaction can increase or decrease the 

MLR, with both the magnitude and direction depending upon the value of the external heat flux. (All of 

the above results are for Ea=209 kJ mol-1.) In the present simulations, the pre-exponential factor A was 

increased as Ea was increased, to maintain a constant value of the rate (0.013 s-1 at 370 °C). Figure 4 

shows the mass loss rate with time for PMMA; each frame a.), b.), c.) and d.) shows the result for (21, 50, 

100, and 200) kW/m2

, respectively, while the different curves on each frame show the mass loss for a 

different value of the activation energy†

. As illustrated, there exists an external heat flux (e.g., 50 kW/m2 

for which the value of activation energy does not affect the mass loss rate curves (and this heat flux value 

changes somewhat as the polymer properties, such as the thermal conductivity, and specific heat are 

varied). At external heat fluxes lower than 50 kW/m2

, lower values of Ea give higher mass loss rates, 

while at higher heat fluxes, the opposite is true. This is also the case for the average values of MLR: with 

† Note: the random absorption algorithm in ThermaKin was turned off for these calculations to remove the noise in 

the output and more clearly show the effects of Ea. The results with the random algorithm turned on are qualitatively 

the same. 

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the highest value of Ea , the average MLR are about 0.8, 1.01, 1.08 and 1.08 those with the lowest Ea, for 

EHF, for (21, 50, 100, and 200) kW/m2

, respectively. This effect of Ea on the MLR is due to changes in 

the temperature profile with mass loss rate: at low mass loss rate (i.e., low EHF), the temperature gradient 

is mild, leading to lower surface temperatures and higher sub-surface mass loss, which give larger mass 

losses at low values of Ea. Conversely, at high mass loss rates (or external heat flux), the high external 

heat flux leads to high surface temperatures (as shown in Figure 12 and Figure 15) which produce 

comparatively higher reaction rates at high activation energies. 

Thermal Conductivity, k 

The effect of changes in the thermal conductivity of the polymer on the mass loss rate as a function of 

time is shown in Figure 5 for an external heat flux of 21 kW/m2 ( frame a.) and 100 kW/m2 (frame b.). 

The results in Figure 5 were calculated using the nominal property values Table 1, and the five curves in 

each frame correspond to the five values of the thermal conductivity (0.1 to 0.5) W m-1 K-1. The MLRav is 

nearly identical regardless of k (at all values of EHF from 21 kW/m2 to 200 kW/m2), as expected from 

Eq. 2, in which k does not appear. The time varying behavior, however, is different. At low EHF, the low 

conductivity case is thermally thick, while the high conductivity case is thermally thin. At low flux, the 

time to peak MLR is about half for the high-conductivity case as compared to the low (reflecting the thin 

behavior); while at high flux, the time to peak MLR is only slightly lower (10 %) for the high value of k. 

(Of course, for the low-flux condition, the time to peak MLR is highly dependent upon the threshold for 

defining the peak. In the discussion here, reaching a value within a few percent of the maximum is 

considered to be reaching the “peak” value.) For either low or higher flux, the peak heat release is only a 

few percent lower for k=0.1 compared to k=0.5. 

Absorption Coefficient, α 

Higher transmission of IR through the polymer (lower values of the absorption coefficient α) creates 

behavior similar to higher thermal conductivity, namely increased thermally thin behavior. Figure 6 a.) 21 

kW/m2 and b.) 100 kW/m2 show the effect for (α = 200, 400, 800, 1200, 50,000) m-1 as different lines in 

each frame. For all values of α, the behavior is thermally thin at low EHF, and thick at higher EHF, and 

for both values of EHF, the behavior becomes thinner as α decreases. At low flux, the peak MLR is about 

13 % lower for surface absorption than for α=200 m-1, whereas for higher flux, the peak MLR is about 

2 % higher for the surface absorption. The average MLR is significantly lower for the low α cases: 

generally about 20 % lower for α=200 m-1 (and 10 % for α=400 m-1) as compared to 50000 m-1 for any 

flux. As with the thermal conductivity, the time to peak MLR is shorter for cases where energy penetrates 

better into the sample (higher k or lower α). The very gradual decrease in the MLR at the end of the 

burning period when α=200 m-1 is like a result of remaining sample being thinner than the characteristic 

thermal radiation penetration depth (1/α, or 5 mm), so that not all of the radiant energy is absorbed, 

decreasing MLR. 

Specific Heat, c 

The effects of variations in the specific heat on the time history of the mass loss rate are shown in Figure 

7 a.) for 21 kW/m2 and b.) 100 kW/m2

. The different curves in each frame correspond to c = (1, 2, 3, 4, 

and 5) kJ kg-1 K-1. The peak MLR is not much affected by c at low EHF (only about 4 % lower at c = 5 

kJ kg-1 K-1 as compared to c = 1 kJ kg-1 K-1); while at high EHF, it’s about 10 % lower. On the other 

hand, the average MLR is about 35 % lower (at all values of EHF except 21 kW/m2

, where it’s 41 % 

lower) at c = 5 kJ kg-1 K-1 as compared to c = 1 kJ kg-1 K-1. This effect is due to the contribution of c to 

∆Hg, and the effect of ∆Hg on the MLR, as described Eq. 2. (Note that in the present calculations, the 

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variation in c from 1 kJ kg-1 K-1 to 5 kJ kg-1 K-1 raises the heat of gasification ∆Hg by about a factor of 

two.) From the average values of MLR calculated from Figure 7, the sensitivity of MLRav to c is 

determined. As described above, MLRav ~ cn with n varying from -0.26 to -0.32, depending upon EHF, 

while the algebraic relation in Eq. 2 predicts that n=-0.30, in good agreement with the detailed numerical 

results predicted here. As for the time-dependent behavior, higher specific heat affects the MLR in the 

same way as lower conductivity or higher absorptivity (see Eq. 3), leading to thermally-thick behavior. 

Results for other heat fluxes (50, 150, and 200) kW/m2 show the same trends, and the thermally-thick 

behavior is accentuated as the flux goes up. 

Constant Values of k or c 

Both the specific heat and the thermal conductivity of polymers vary with the temperature—which 

increases significantly as the polymer heats, melts, and decomposes. In the calculations above, the 

temperature-dependent values were used in the simulations (except when c or k themselves were varied). 

It is of interest, however, to determine how the results would differ when using a single value of k or c, 

evaluated at the average polymer temperature (between ambient and the decomposition temperature). The 

MLR as a function of time was calculated using the average value of either k or c, and these were 

normalized by the results using the temperature dependent values. The error imposed by using the average 

value is not large. For example, for constant c and external heat fluxes of up to 100 kW/m2

, the mass loss 

rate is at most, 2 % to 5 % higher at some times, than with the temperature dependent value (for all values 

of Ea), while for high external heat flux (150 kW/m2 or 200 kW/m2

) using a constant c gives less than 2 % 

error. For a constant value of k, the transient behavior (for example at low flux, or at higher flux and 

towards the end of the mass-loss period) of the mass loss rate can about 5 % higher than the mass loss rate 

calculated with a temperature dependent k. The steady mass loss rates are generally within 1 %, except at 

high flux, where it can be 1 % high or 2 % low, depending upon Ea. The average mass loss rate is also not 

greatly affected by the use of constant k or c. For example, for all values of EHF and Ea, using constant k 

gives average MLRs within 1 % of those using variable k , (except the cases of EHF = 200 kW/m2

, and 

Ea = 97 kJ kmol-1 , and EHF =150 kW/m2 and Ea= 418 kJ kmol-1 , which yields about a 5 % difference. 

Using constant c gives average MLRs within 3 % of the results with variable c, for all values of Ea, 

except at EHF = 200 kW/m2

, for which the results using average c can be 1 % to 10 % low (depending 

upon Ea). These results (here, for a wider range of Ea and incident flux) are consistent with those of 

Steckler et al. who found close agreement between the time-varying mass loss rate using constant values 

of k or c (evaluated at a mean temperature) as compared to using temperature-dependent values (for a 

single case of 40 kW/m2

, surface absorption, and infinite reaction rate). 

4.1.2. Ignition Time 

Simple one-dimensional heat transfer models predict the ignition time for thermally thick or thin materials 

[29]. The characteristic thermal conduction length δt for thick materials is given by δ t = k t / ρ c , in 

which in which t is the exposure time. In the absence of convective and radiative heat losses, thin 

materials, S /δ ≤ 1 , have an ignition time given by: 

(Tig − T0 )

τ ign = ρ cS " Eq. 4 

qext 

and thick materials: 

π (Tig − T0 )

2 

τ ign = kρ c . Eq. 5 "2 4 qext 

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Below, these analytic predictions are compared to the numerical results. 

As noted above, the ignition time is calculated from the numerical MLR predictions as the time at which 

the mass loss rate reaches a critical value (3 g s-1 m-2). The effect of parameter variations on these 

numerically-determined ignition times are summarized in Figure 8 a.), b.), c.), and d.) and Figure 9, which 

show the ignition time as a function of k, c, α, ∆Hreac, and S, respectively. In each figure there is a 

grouping of results for each value of the external heat flux, and the different color lines show the variation 

with Ea. As indicated, the ignition time is strongly dependent upon the external heat flux, and mildly 

dependent upon Ea, except for cases of low EHF (21 kW/m2

), for which Ea can make a difference of a 

factor of three to seven. The variations in magnitude of the parameters k, c , α, and ∆Hg in the figures is 

about a factor two above and two below the nominal value for PMMA. Hence, the slopes of the lines give 

a qualitative estimate of the sensitivity of the ignition time to those parameters. Based on the simple 

thermal conduction analysis above, for the nominal conditions of Table 1, the 25.4 mm thick sample is 

thermally thick with respect to ignition for τign less than 7100 s. Hence, for all conditions shown in Figure 

8 the sample is thermally thick with respect to ignition, except the lowest flux, at which the behavior can 

be thin depending upon the parameter value and the activation energy assumed. 

Effect of q " 

ext 

" In Figure 8a.) and b.) for k and c, τign is proportional to 1/ qext 

2

, following Eq. 5 above for thick materials 

(which is the case for these conditions). For varying α (Figure 8c), the results at α=960 m-1 (the nominal 

value) also give τign ~ 1/ q " 2

, while optically thin material (α=200 m-1) has τign ~ 1/ q " 1.5 and surface ext ext 

" 2.66 absorption shows τign ~ 1/ qext . 

k 

As shown in Figure 8a, a five-fold increase in thermal conductivity generally causes about a two-fold 

increase in the ignition time, with a larger effect (3x) at EHF=21 kW/m2 as compared to (1.5x) at 

EHF=200kW/m2

. That is, τign ~ k 0.25 to τign ~ k 0.67 , a weaker dependence than the linear behavior 

expected form Eq. 5, likely resulting from energy penetration dominated by thermal radiation transport 

for these short times, diluting the effect of thermal conduction. 

c 

In Figure 8b, a five-fold increase in the specific heat is shown to result in a five-fold increase in the 

ignition time for all values of EHF except the lowest (relatively independent of Ea), essentially following
 

the thermally thick prediction (τign ~ c1

) in Eq. 5 above. For the low flux conditions EHF=21 kW/m2, the
 

material is still thermally thick and τign ~ c1.3 (and there is a strong dependence on Ea).
 

α
	

The absorption coefficient has very little effect on τign at EHF=21 kW/m2

. At high EHF (100 to 200)
 

kW/m2

, however, τign ~ α−0.72, so that he α=200 m-1 case has an ignition time about ten times longer than 

with α=50000 m-1, as recently discussed by Jiang et al [30].
 

∆Hreac 

The heat of reaction ∆Hreac has almost no effect on the ignition time at a given flux, except at very low 

incident heat fluxes; i.e., very near to the critical heat flux for ignition, where the ignition times are so 

long that the reaction rates (albeit slow) have an effect on the critical mass flux for ignition. 

Thickness 

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The sample thickness affects the ignition time primarily for thinner samples. As shown in Figure 9 a.) for 

α=960 m-1, and b.) α=50000 m-1, for either value of α, and higher values of EHF (100 to 200) kW/m2

, the 

thickness has generally less than a 10 % effect on τign, as long as the sample is 4 mm thick or greater. At 

" 2 mm thickness, the sample is thermally thin, and the dependence of τign on qext follows Eq. 4 above for 

thin materials, for EHF≥50 kW/m2

. For very low EHF (21 kW/m2), the effects of slow reaction and of 

surface re-radiation losses affect the ignition time [31], so Eq. 4 and Eq. 5 are not accurate. 

Effects of Constant c or k 

The ignition time is only moderately affected by using constant values of the thermal conductivity or 

specific heat of the polymer. A constant value of c yields an ignition time about 10 % shorter than using 

the temperature-dependent value, for all values of the external heat flux, and this ratio is only affected 

slightly by variations in Ea. For constant k, the ignition time is again lower, by about 5 %, for all external 

heat fluxes independent of Ea, except at 21 kW/m2

, at which it can be either 6 % higher or lower, 

depending Ea. 

4.1.3. Comparison of ThermaKin and FDS Results 

All calculations in the present work were performed with both ThermaKin and FDS; however, due to 

space limitations, only the former are presented in Part 1 (Part 2 includes both). For the nominal values 

of k, c, S, and ∆Hreac, and surface absorption of thermal radiation, the entire time-dependent MLR curves 

obtained with the two codes are in within 3 % of each other for the range of EHF and Ea in Table 1, 

except the case of Ea=837 kJ/mol and EHF=200 kW/m2

, for which the difference was 6 %. Using the 

nominal value of the absorption coefficient (960 m-1) , the general behavior is the same, as evidenced by 

very close values of average and peak MLR, but the shape of the curves differ slightly. Examination of 

other solutions with lower values of α indicates that the disagreement becomes worse as α decreases. For 

the ignition time, when surface absorption is assumed, and over the range of conditions of Table 1, τign 

using FDS is always within 10 % of that from ThermaKin. However, when α=960 m-1, the FDS 

prediction is about 46 % higher than the ThermaKin at EHF=21 kW/m2

, and 22 % lower at 

EHF=200 kW/m2

. The different treatment of the in-depth absorption of radiation by the two codes is 

likely responsible for the differences observed, and work is continuing to understand the reasons. 

4.1.4. Conclusions (Part 1) 

The effect of the parameters ∆Hreac, k, α, and c on τign and MLRav are provided in 

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Table 2, which gives the power-law dependence of τign or MLRav on each of the parameters. For example, n τIgn ~ ∆H reac , with n=0.02 at EHF=200 kW/m2

, and increasing to n=0.14 at EHF=50 kW/m2; at very low 

EHF (21 kW/m2

), n=1.70. That is, the ignition time is typically not dependent upon ∆Hreac, except at low 

flux. Also given in Table 2 are the expected values of n based on the simple algebraic model in Eq. 2. 

(Note that the expected dependence of τign and MLRav on α are not given in Table 2. The effect of indepth absorption of radiation shows up as changes to the surface temperature, and hence the radiative heat 

loss term. While analytic solutions are available [30], they do not have a simple form. ) The ignition 

time is somewhat dependent upon k, increasingly so at lower flux, with n=0.28 or 0.71 at EHF=(200 and 

21) kW/m2

, respectively. The ignition time is approximately proportional to c, with n close to unity, 

except at EHF=21 kW/m2

, for which n=1.3. The effect of α on τign has the opposite trend, with little 

dependence at low flux (n=-0.03), but increasing importance at higher flux (n=-0.54 and -0.77, at 

EHF=50kW/m2 and 200 kW/m2

). For comparison, the simple thermal conduction model predicts n=1 for 

variation of τign with k and c, and n=0 for ∆Hreac. That is, the dependence calculated here is greater for 

∆Hreac, especially at low flux, about as expected for c (and somewhat greater at low flux), and 

significantly less for k, increasingly so at lower flux. Variation in α can have substantial effect on τIgn at 

high flux. 

As described above, MLRav is roughly inversely related to ∆Hreac, with n=-0.64 to -0.80, mildly related to 

c or α, with n=-0.27 to -0.32, or n=0.016 to -0.14, respectively, and nearly unrelated to k with n=-0.004 to 

-0.010. Expected values of n for MLRav based on a simple heat balance model (Eq. 2) are -0.72, -0.30, 

and 0 for ∆Hreac, c, or k. Hence, for these three parameters, the simple relationship gives a good estimate 

of their relative influence on MLRav. The effects of in-depth absorption on MLRav can be significant. 

Numerical values of the peak MLR and the time to peak MLR were not calculated since these values are 

highly dependent upon both the threshold for defining the peak (e.g., region of 95 % of peak, etc.), as well 

as the thermally thick or thin behavior, which switches readily as the physical properties are changed. 

Instead, the qualitative behavior with respect to these metrics are provided. 

Table 3 summarizes the results of the variation in each of the material properties on the mass loss. The 

influence is characterized by the shape of the mass loss rate (MLR) curve, average MLR, peak MLR, time 

to peak MLR, and ignition time. In the table, double check marks indicate a large effect on that metric, 

while single check marks indicate a moderate effect; and a gray single check, even less effect. A blank 

means no significant effect. Subscripts indicate the conditions (e.g., LF: low flux, HF, high flux) to which 

the importance is limited. (Note that these qualitative rankings supplement the quantitative results in 

Table 2.) As shown, the heat of reaction ∆Hreac is the most important parameter, followed by the 

thickness S and specific heat c. The absorption coefficient and the conductivity behave similarly, showing 

an effect on the shape of the MLR, as well as the time to peak at low flux, and on the ignition time for 

moderate and high fluxes (α also affects the average MLR at high flux). For the conditions assumed here, 

the activation energy of the decomposition step is the least important parameter, mildly affecting the 

shape of the MLR curve at high or low flux, and the average MLR and ignition time at low flux. In future 

research, it would be of value to examine the effect of the overall reaction rate on the MLR, as well as 

char layer properties. 

For the polymer modeled here, the behavior is often close to intermediate between thermally thick and 

thin. As a result, varying the material properties often leads to changes in the time-varying MLR curve 

which are due to switching of the behavior from thick to thin (or visa versa). This has been discussed by 

Delichatsios et al. [32] in the context of ignition. It is of value to keep this in mind when interpreting the 

mass loss rate data for materials for which some property change has been made, for example in 

determining the mode of action of polymer fire retardant additives. 

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4.2. PART 2: Estimation of Heat of Gasification from Cone Calorimeter Mass Loss Data 

4.2.1. Introduction 

The heat of gasification has a large effect on the mass loss rate. A common method of estimating ∆Hg is 

the method of Tewarson and Pion [4], in which the steady mass loss rate is plotted against the external 

heat flux, and the inverse of the slope is the effective heat of gasification ∆Hg

(eff). As described by Staggs 

[22], even in the absence of the complicating effects (e.g., changing flame heat feedback, cone radiation 

trapping by the flame, bubble formation in the polymer and mass transfer limitations to the surface, etc.), 

the extracted value of ∆Hg

(eff) can differ from the actual value substantially (30% to 40 %). Hence, it is of 

value to explore the sensitivity of the results of Staggs for a wider range of material properties. 

First, the numerical simulations are used to repeat the analysis of Staggs for his input conditions. This 

also allows comparison of the analytical prediction of the steady mass loss rate of Staggs with the 

numerical results calculated using ThermaKin or FDS. Next, the activation energy and absorption 

coefficient of the PMMA (which were expected to have the largest effect) are varied in a series of 

calculations, allowing assessment of the effect of these parameters on the inferred effective heat of 

gasification. 

4.2.2. Comparisons with Analytical Results 

The burning behavior as a function of time of a sample of PMMA exposed to a known radiant heat flux 

was calculated assuming the material properties and conditions of ref. [22] as listed in Table 1. Staggs’ 

analysis is based on simulations for a mild range of external heat flux (20 kW/m2 to 70 kW/m2

), and that 

range is used here. (Staggs includes a convective heat loss term in his calculations (10 W m-2 K-1) so this 

term is added in the FDS and ThermaKin simulations.) As an overview, Figure 10 shows the predicted 

MLR (solid lines, left axis) and surface temperature (dotted lines, right axis), for an external heat flux of 

(20 to 70) kW/m2

. As expected (Eq. 2), the average burning rate and the inverse of the burning time are 

roughly proportional to the external heat flux. At each value of the external heat flux, the mass loss rate 

increases from zero at early times, as the energy at the surface penetrates the cold polymer. At longer 

times, the heat conduction and surface regression rates find an equilibrium which leads to a relatively 

constant temperature profile through the polymer, and which gives a relatively steady burning rate [22] 

(for EHF<20 kW/m2

). The higher EHF conditions have higher surface regression rates, which makes the 

steady burning rate assumption more accurate (due to more thermally-thick behavior). 

The temperature curves in Figure 10 show that at early times, the surface temperature increases as the 

sample heats, until a relatively constant value is obtained. The apparent rapid surface temperature rise at 

the very end of the burning period is a result of the limited time step and grid size resolutions in the 

calculation (so the temperature data are truncated for the last few points). Further refinement of the time 

stepping and mesh size was deemed unnecessary since that portion of the burning corresponds to a small 

fraction of the total burning history. As shown in the figure, the higher external heat flux leads to higher 

surface temperature, about 104 K higher at 70 kW/m2 as compared to 20 kW/m2

, an effect which was 

shown previously in experiments [33] and numerical calculations [5]. To estimate the average surface 

temperature for each value of external heat flux in Figure 10, the central quarter of the surface 

temperature-time history was used (which also corresponds to the time very near to the average mass loss 

rate). 

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In order to extract the effective heat of gasification from the mass loss rate data, one plots the steady mass 

loss rate with the external heat flux, and takes the inverse of the slope as ∆Hg

(eff) [4]. In our analysis using 

the synthetic MLR curves, the calculated mass loss as a function of time is not purely steady, so we used 

the average burning rate for the time period during which the sample is exposed to the external flux as the 

steady burning rate; that is, the preheat period during which the sample is heating but not yet losing mass 

was also used in calculating the average. The logic is that the higher mass loss rate at the end occurs 

because the sample has been preheated by conduction which occurred earlier, and conceptually, this is 

offset by the energy input at early times which added to the energy in the sample, even though the mass 

loss rate was zero or low. The end of the integration period in defining the average was the time at which 

the MLR decreased to 0.01 g m-2 s

-2; this gives nearly identical values of MLRav as would using a 

criterion of 3 g s-1 m-2 since the MLR drops very rapidly at the end of the mass loss period for the present 

simulations. These calculated average values of the mass loss rate are indicated by the short horizontal 

lines near the center of each curve in Figure 10. As indicated, these average values are very close to the 

steady burning region (when one exists), and to the inflection point in the MLR versus time curves. 

The average values of the mass loss rate are plotted against the external heat flux in Figure 11, for the 

results from the ThermaKin and FDS numerical models, as well as from the analytical model used by 

Staggs [22]. As shown, the two numerical models (which use the same input material properties, but 

different solution techniques) give results with about 1 % of each other (except at the lowest mass loss 

rates, which differ by 3 %). Both numerical results are about 8 % higher than the analytic results of Stagg. 

The reason for the lower mass loss rates in the analytic model results may be related to the estimated 

surface temperatures, as shown in Figure 12 (right scale, bottom curves). Because the analytic model only 

has mass loss at the surface (while the numerical models allow mass loss at all positions in the sample), it 

needs a higher surface temperature for the PMMA to lose the required mass (determined by the energy 

balance at the surface). This higher surface temperature in the analytical prediction leads to higher 

radiation heat losses, and lower resulting mass loss rates (less net energy makes it into the polymer). 

Another possible reason for the discrepancy between analytical and numerical predictions of the mass loss 

rates shown in Figure 11 is the use of the average burning rates from Figure 10. An alternative choice 

would be to use the inflection point in the time-varying mass loss rate shown in Figure 10. Doing so leads 

to minor changes in the values of the “steady” mass loss rate shown in Figure 11 (-17 %, - 5 %, and 0 % 

for 20, 45, and 70 kW/m2

, respectively), which leads to an increase in the slope of the mass loss vs. 

external heat flux curves in Figure 11 of about 4 %. 

Since the slopes in Figure 11 are in reasonable agreement with each other, the inferred effective heats of 

gasification ∆Hg

(eff) (1/slope) are also in reasonable agreement. Figure 12 shows the effective heat of 

gasification ∆Hg

(eff) and the heat of gasification ∆Hg (upper curves, left scale) from the Staggs, FDS, and 

ThermaKin models (note: ∆Hg

(eff) is obtained from the slope of the calculated mass loss vs. EHF curve, 

while ∆Hg is calculated from Eq. 1 above, using the calculated surface temperatures and the assumed c of 

(eff) the PMMA), together with the value of ∆Hreac from Table 1. As discussed in ref. [22], value of ∆Hg

extracted from Staggs analytical prediction of the mass loss rate is 1.875 kJ/g, while that extracted from 

either the ThermaKin and FDS predictions is (1.72 ± 0.02) kJ/g, which is about 8 % lower. Hence, given 

the differences in the assumptions in the calculations leading to the mass loss curves, this agreement in 

∆Hg

(eff) is considered acceptable. The FDS- and ThermaKin-calculated values of ∆Hg agree within 3 %, 

differing slightly due to the mildly different calculated surface temperatures, also shown in Figure 12. 

As a final step in the comparison, we examine the ratio of the value of the heat of gasification extracted 

from the numerical experiment ∆Hg

(eff), with the values of ∆Hg which were input into the models. Ideally, 

the value of ∆Hg

(eff) extracted from the mass-loss versus external heat flux curves would reproduce the 

value of ∆Hg which was used as input into the numerical model that generated the mass loss curves. The 

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advantage here is that since mass loss curves were calculated, only those physical effects included in the 

model can affect the value of ∆Hg

(eff) obtained. For example, in an actual experiment, as the external heat 

flux to the polymer is varied by changing the cone heater temperature, variations in the heat flux reaching 

the polymer surface can be influenced by differences in the heat feedback from the flame (from different 

flame size and shape), as well as from flame blockage of the radiation from the cone [8]. Since the 

external heat flux in the numerical calculations is specified exactly, no assumptions are required 

concerning the constancy of the heat flux from the flame (or radiation absorption by the flame). Hence, 

one can separate out any differences between the input ∆Hg and the extracted ∆Hg

(eff) without concern for 

effects from the flame. 

Figure 13 shows the ratio of the effective heat of gasification ∆Hg

(eff) (extracted) to the heat of gasification 

∆Hg (input) for the calculated mass loss rates using the model of Staggs, ThermaKin, and FDS. The two 

numerical models give results within 1 % of each other, and these are about 6 % lower than the ratio 

calculated analytically by Staggs. The difference between the analytical and numerical results occurs 

because, as Figure 11 shows, the Staggs model and the two numerical models predict somewhat different 

variation of the MLR with EHF. Nonetheless, the basic trends and conclusions are the same: the heat of 

gasification extracted from the mass loss vs. flux curve will be about 30 % higher than the actual value 

based on the physical properties of the material, and this difference occurs even the absence of other 

complications. It should be noted that for accurate numerical prediction of the mass loss rate of several 

polymers in the NIST gasification device, ThermaKin simulations required the actual physical parameters 

of the materials, not the effective gasification values obtained through the method described above [12]. 

As pointed out by Staggs [22], this result has implications for the fire community, which routinely uses 

cone measurements to obtain effective heats of gasification. The surface temperature variation, caused by 

the finite rate of chemical reaction of the polymer, leads to errors in the effective heat of gasification 

extracted from mass loss versus flux curve. The above results were demonstrated, using several different 

calculation schemes, for a given set of physical parameters of the polymer. It is of interest to determine 

the extent to which different polymer properties affect this result. The parameters expected to have the 

biggest effect on the ratio of ∆Hg

(eff) to ∆Hg are the activation energy of the one-step polymer 

decomposition reaction Ea, and the absorption of infrared light by the material (which provides in-depth 

heating), characterized by the absorption coefficient α. In the discussion below, we vary these parameters 

in numerical simulations using ThermaKin, and repeat the above procedure to assess the effect of these 

(eff) changes on the ratio of ∆Hg to ∆Hg.

(eff) 4.2.3. Parametric Analyses of the Effects of Ea and α on ∆Hg

The activation energy and the absorption coefficient were varied over a limiting but reasonable range. 

Values of Ea of (97, 209, 418, and 1050) kJ/mol were used (while changing the pre-exponential factor A 

to maintain the overall rate at 0.013 s-1 as described above). The value of 97 kJ/mol is that of Staggs [22] 

and is at the lower end of Ea reported; other values in the literature are close to 200 kJ/mol [,12] , and 

1050 kJ/mol represents a value approaching infinite reaction rate. For the absorption coefficient, while 

930 m-1 is close to recent measurements for black Polycast PMMA [30], some measurements indicate that 

polyethylene might be close to 400 m-1 [34], and 50000 m-1 is a large value representing surface 

absorption. 

Figure 14 shows the results of the calculations of the average mass loss rate as a function of the external 

heat flux; each of the three frames shows the results for values of α of (400, 930, and 50,000) m-1 , while 

the different curves within each frame show the results for the four values of Ea (97, 209, 418, and 

1050) kJ/mol. As the figures indicate, all of the curves are quite linear, and the surface absorption case 

(α=50000 m-1) shows the largest effect of Ea, on the magnitude of the slope. From these curves one can 

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extract the inverse of the slope to provide the values of the effective heat of gasification ∆Hg

(eff), which are 

shown in Figure 15 (dotted lines, left axis) on the upper part of the frames. As expected from Figure 14, 

the values of ∆Hg

(eff) are quite constant over the range of fluxes tested here, and the variation with Ea is the 

largest for the case of surface absorption of the energy. 

The numerical results also predict the expected surface temperature as a function of external heat flux, 

and these are also shown in Figure 15 (dashed curves, bottom part of frames, right axis). The surface 

absorption case gives the largest variation in Ts with the incident flux, and the variation is the largest with 

the lowest value of Ea. As in the discussion above, the values of Ts allow one to calculate the heat of 

gasification ∆Hg which was essentially input into the numerical calculations, and these also are shown in 

Figure 15 (solid curves, left axis). Following the calculated surface temperature results (via Eq. 1 above), 

the variation in ∆Hg with the incident flux is largest for the surface absorption case with low activation 

energy. 

The ratio of the ∆Hg

(eff) to ∆Hg, are shown in Figure 16. As indicated, the surface absorption case with low 

activation energy (as in Figure 13) gives the largest ratio of ∆Hg

(eff) to ∆Hg. The ratio approaches unity as 

the activation energy increases, and the activation energy becomes less important as the polymer becomes 

more transparent to infrared light. Thus, the overestimate of ∆Hg

(eff) becomes small as Ea approaches 

infinity, and is also reduced as the polymer becomes more transparent. 

In order for the method of Pion and Tewarson to be accurate for extracting the heat of gasification from 

the mass loss rate data, the changes in the net heat flux to the polymer must be approximated well by 

changes in the applied heat flux from the radiant heater. That is, as discussed above, the flame heat 

transfer to the polymer and the flame (or pyrolysis products) blockage of cone heater radiation must be 

constant at all applied heat fluxes. These are taken care of in the numerical calculation above, in which 

the imposed heat flux (purely radiant) is specified, and there is no flame. The net heat flux from the 

radiant source, however, must also account for heat losses due to re-radiation from the hot polymer 

surface. Correction for this effect is easily made using the numerical results since the surface temperature 

is calculated for all conditions, and the surface radiation heat losses are then given by: 

q 

.

"= σ (T 4 − T 4 ) in which σ (5.67 x 10-11 kWm-2) is the Stefan-Boltzmann constant, Ts is the surface r s a 

temperature, and Ta is the ambient temperature. It is of interest to determine if the results given in Figure 

14 through Figure 16 above are modified if one uses the net heat flux (i.e., the external heat flux minus 

the energy reradiated from the surface) as the independent variable, instead of the imposed heat flux from 

the cone as used above [35]. Figure 17 shows the ratio of the ∆Hg

(eff) to ∆Hg plotted against the net heat 

flux , for the same three values of the absorption coefficient, and four values of the activation energy of 

(eff) decomposition. As the figure indicates, much of the deviation from unity of the ratio ∆Hg / ∆Hg has 

been eliminated. Curves for the highest value of Ea are always very close to unity (around 1.03), and those 

for the lowest value of Ea are only slightly higher (1.06); the results for the most transparent case (α = 400 

m-1) are about a percent higher than the more opaque cases. 

For the polymer conditions simulated here, the effective heat of gasification which can be extracted from 

the mass loss rate data at different incident heat fluxes is only 2 % to 6 % higher than the true values if the 

heat flux is corrected for the radiant heat losses from the polymer surface. Nonetheless, in practice, one 

usually does not know the polymer surface temperature a priori—especially how it changes with heat 

flux, so this is a difficult correction to make. The advantage here is that since all of the mass loss rates 

were calculated, we can estimate the radiant heat losses directly to assess the importance of this parameter 

on the result. 

16
 
 

 

 

 

 

 

 

 

 

 

 

 

 

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Higher values of the activation energy and lower values of the absorption coefficient lead to better 

agreement between ∆Hg

(eff) and ∆Hg , with the ratio ∆Hg(eff)/∆Hg 1.04 to 1.14 for α=400 m-1 and 1.03 to 

1.22 for α=50000 m-1, for EHF of 20 kW/m2 to 70 kW/m2

. Correcting the external heat flux for the 

surface re-radiation (yielding a net external heat flux) lead to values of ∆Hg

(eff)/∆Hg of 1.04 to 1.07 α=400 

m-1 and 1.02 to 1.04 for α=50000m-1. A method for estimating the pyrolysis temperature of charring 

materials has recently been reported [36]. If a similar analysis could be developed for thermoplastics 

(eff) which would predict the surface temperature as a function of EHF, it could improve estimates of ∆Hg . 

Interestingly, the surface temperature measurements of Rhodes [37] for this material for a range of EHF 

show an increase in the surface temperature from about 635 K at 20 kW/m2 to 650 K at 60 kW/m2 , which 

from the middle frame of Figure 15, implies an activation energy between 209 kJ mol-1 and 418 kJ mol-1 . 

4.2.4. Conclusions (Part 2) 

The influence of sub-surface reaction on the effective heat of gasification, ∆Hg

(eff), inferred from the mass 

loss vs. flux data was determined. Previous analytical predictions of mass loss rate of PMMA exposed to 

external heat fluxes of 20 kW/m2 to 70 kW/m2

, were validated using ThermaKin and FDS simulations. 

These two codes predicted effective heat of gasification, ∆Hg

(eff) within about 0.06 % of each other, and 

these were about 6 % smaller than those of Staggs. In the ThermaKin simulations, this lead to a 

discrepancy between ∆Hg

(eff) (inferred) and ∆Hg (input) of 20 % to 30 %, which was somewhat lower than 

30 % to 40 % value estimated analytically by Staggs. Using ThermaKin, that analyses were extended to 

other conditions of PMMA properties. 

Higher values of the activation energy and lower values of the absorption coefficient tended to drive the 

ratio of ∆Hg

(eff) / ∆Hg towards unity. Plotting the mass loss rate against the net external heat flux (i.e., 

correcting for differing surface re-radiation due to varying surface temperature with imposed heat flux) 

makes ∆Hg

(eff) and ∆Hg within 2 % to 6 % of each other. Hence, devising a means to estimate Ts could 

improve somewhat the utility of the method of Tewarson and Pion to estimate ∆Hg. In future work, it 

would be interesting to numerically investigate the influence of other parameters, such as char layer 

(eff) formation, on the ratio of ∆Hg to ∆Hg.

5. CONCLUSIONS 

The time varying mass loss rate for PMMA with a range of physical properties has been predicted 

numerically. The variations in the average mass loss rate due to changes in the heat of reaction, specific 

heat, and thermal conductivity are well predicted by the simple algebraic relations based on energy 

balance at the surface. The ignition time is influenced by heat of reaction and the specific heat as 

expected based on simple thermal conduction models, although their effect is somewhat greater than 

expected at low flux. The variation in the ignition time with changes to the thermal conductivity is about 

one third the expected value based simple theory, most likely due to the competing effect of in-depth 

absorption as a mechanism for heat transfer into the sample. The material’s absorption coefficient for IR 

can influence the MLRav significantly, and the ignition time at substantially. 

" An insight gleaned from Part I of the present work is that while most of the properties studied here ( qext , 

k, c, α, ∆Hreac, and S, ) have varying effect on the mass loss rate and ignition time, they all affect the 

thermal thickness of the material and hence, have the potential to switch the behavior from thermally 

thick to thermally thin, and consequently change the MLR curves both qualitatively and quantatively. 

When making changes to a polymer to promote fire-safe behavior, care should be taken in interpreting 

experimental data to insure that the material property being changed is not just changing the thermal 

17
 
 

 

 

 

thickness of the material. The heat of reaction is the most important parameter of those examined, and 

needs to be determined most accurately. 

In Part II, the predictions of Staggs with regard to the differences between the actual heat of gasification 

of a material and that extracted from experimental mass loss vs. heat flux data was verified. For the 

conditions simulated here, varying the polymer properties to include more optically transparent materials, 

and those with higher activation energy of decomposition made the discrepancy between the two numbers 

less, and correcting for the surface re-radiation losses virtually eliminated the discrepancy. 

6. ACKNOWLEDGEMENTS 

The author is grateful to Kevin McGrattan and Simo Hostikka for aid in using FDS, and 

Stanislav Stoliarov and Richard Lyon for aid in using ThermaKin. Takashi Kashiwagi at NIST originally 

suggested this line of research, and helpful conversations with him contributed throughout. 

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18
 
 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

7. REFERENCES
 

This publication is available free of charge from: http://dx.doi.org/10.6028/NIST.TN.1929 

[1] Standard test method for heat and visible smoke release rates for materials and products using an 

oxygen consumption calorimeter. 2008;(ASTM E 1354-08a):1-31. 

[2] Standard test methods for measurement of synthetic polymer material flammability using a fire 

propagation apparatus. 2006;(ASTM E2058-06): 

[3] Austin, PJ, Buch, RR, and Kashiwagi, T. Gasification of silicone fluids under external thermal 

radiation part I. Gasification rate and global heat of gasification. Fire and Materials 1998;22:221­

37. 

[4] Tewarson, A and Pion, RF. Flammability of Plastics .1. Burning Intensity. Combustion and
 

Flame 1976;26:85-103.
 

[5] Vovelle, C, Delfau, JL, Reuillon, M, Bransier, J, and Laraqui, N. Experimental and numerical 

study of the thermal degradation of PMMA. Combustion Science and Technology 1987;53:187­

201. 

[6] Steckler, KD, Kashiwagi, T, Baum, HR, and Kanemaru, K. Analytic model for transient 

gasification of non-charring thermoplastic materials London: Elsevier Applied Science; 1991. p. 

895-904. 

[7] Hopkins, D and Quintiere, JG. Material fire properties and predictions for thermoplastics. Fire 

Safety Journal 1996;26:241-68.
 

[8] Rhodes, BT and Quintiere, JG. Burning rate and flame heat flux for PMMA in a cone 

calorimeter. Fire Safety Journal 1996;26:221-40.
 

[9] Staggs, JEJ and Whiteley, RH. Modelling the combustion of solid-phase fuels in cone 

calorimeter experiments. Fire and Materials 1999;23:63-69.
 

[10] Nicolette, VV, Erickson, KL, and Vembe, BE. Numerical simulation of decomposition and 

combustion of organic materials. In: Proceedings of Interflam 2004, 10th InternationalConference 

on Fire Science and Engineering London: Interscience Communications Ltd. 2004. 

[11] Lautenberger, C and Fernandez-Pello, AC. Approximate analytical solutions for the transient 

mass loss rate and piloted ignition time of a radiatively heated solid in the high heat flux limit. In: 

Fire Safety Science-Proceedings of the Eighth International Symposium: International 

Association for Fire Safety Science; 2005. p. 445-56. 

[12] Stoliarov, SI, Crowley, S, Lyon, RE, and Linteris, GT. Prediction of the burning rates of noncharring polymers. Combust Flame 2009;156:1068-83. 

[13] Kim, E, Lautenberger, C, and Dembsey, NA. Property estimation for pyrolysis modeling applied 

to polyester FRP composites with different glass contents. In: 11th International Conference and 

Exhibition London: Interscience; 2009. p. 87-102. 

[14] Paul, K. unpublished data, RAPA Technology, Shawbury, England, as cited in: SFPE Handbook, 

Babrauskus, Chapter 3-1. 2002; 

19
 
 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

This publication is available free of charge from: http://dx.doi.org/10.6028/NIST.TN.1929 

[15] Hirschler, MM. Heat release from plastics. In: Heat release in fires Amsterdam: Elsevier; 1992. p. 

375-422. 

[16] Schartel, B, Bartholmai, M, and Knoll, U. Some comments on the use of cone calorimeter data. 

Polymer Degradation and Stability 2005;88:540-47. 

[17] Schartel, B and Hull, TR. Development of fire-retarded materials - Interpretation of cone 

calorimeter data. Fire and Materials 2007;31:327-54. 

[18] Stoliarov, SI, Safronava, N, and Lyon, RE. The effect of variation in polymer properties on the 

rate of burning. Fire and Materials 2009;submitted: 

[19] McGrattan, KB, Klein, B, Hostikka, S, and Floyd, J. Fire dynamics simulator (version 5) user's 

guide. 2007;NIST Special Publication 1019-5:1-176. 

[20] Stoliarov, SI and Lyon, RE. Thermo-kinetic model of burning. 2008;Federal Aviation 

Administration Technical Note DOT/FAA/AR-TN08/17:1-32. 

[21] Stoliarov, SI and Lyon, RE. Thermo-Kinetic model of burning for pyrolyzing materials. In: 

Proceedings of the Ninth International Symposium on Fire Safety Science London: Interscience 

Communications, Ltd.; 2008. p. 1141-52. 

[22] Staggs, JEJ. The heat of gasification of polymers. Fire Safety Journal 2004;39:711-20. 

[23] Panagiotou, T and Quintiere, JG. Generalizing Flammability of Materials. In: Interflam 2004 

London: Interscience Communications Ltd.; 2004. p. 895-905. 

[24] Linteris, GT, Gewuerz, L, McGrattan, KB, and Forney, GP. Modeling Solid Sample Burning 

with FDS. 2004;NISTIR 7178:36 p. 

[25] Stoliarov, SI and Walters, RN. Determination of the heats of gasification of polymers using 

differential scanning calorimetry. Polymer Degradation and Stability 2008;93:422-27. 

[26] Stoliarov, SI, Walters, RN, and Lyon, RE. Determination of heats of gasification of polymers 

using differential scanning calorimetry. 2007; 

[27] Tewarson, A. Generation of heat and chemical compounds in fires. In: SFPE Handbook of Fire 

Protection Engineering Quincy, MA: National Fire Protection Association; 1995. p. 3-53. 

[28] Quintiere, JG and Iqbal, N. An approximate integral model for the burning rate of a 

thermoplastic-like material. Fire and Materials 1994;18:89-98. 

[29] Drysdale, D. An introduction to fire dynamics. Chichester, England: John Wiley and Sons; 

1998. 

[30] Jiang, FH, de Ris, JL, and Khan, MM. Absorption of thermal energy in PMMA by in-depth 

radiation. Fire Safety Journal 2009;44:106-12. 

[31] Mikkola, E and Wichman, IS. On the Thermal Ignition of Combustible Materials. Fire and 

Materials 1989;14:87-96. 

20
 
 

 

 

 

 

 

[32] Delichatsios, MA, Panagiotou, T, and Kiley, F. The Use of Time to Ignition Data for 

Characterizing the Thermal Inertia and the Minimum (Critical) Heat-Flux for Ignition Or 

Pyrolysis. Combustion and Flame 1991;84:323-32. 

[33] Kashiwagi, T and Ohlemiller, TJ. A study of oxygen effects on flaming transient gasification of 

PMMA and PE during thermal irradiation. Proceedings of the Combustion Institute 1982;19:815­

23. 

[34] Linteris, GT. Absoprtion of thermal radiation by burning polymers. Fire Safety Journal 2009;to 

be submitted:1-30. 

[35] Janssens, ML. Personal communication. 2008; 

[36] Park, WC, Atreya, A, and Baum, HR. Determination of pyrolysis temperature for charring 

materials. Proceedings of the Combustion Institute 2009;32:2471-79. 

[37] Rhodes, BT. Burning rate and flame heat flux for PMMA in the cone calorimeter. 2009;NISTGCR-95-664:1-118. 

This publication is available free of charge from: http://dx.doi.org/10.6028/NIST.TN.1929 

21
 
 

 

 

 

 

 

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Table 1– PMMA model input parameters. 

0Parameter Units 

Part 1 

Parametric Analysis 

Nominal 

Value Range 

Part 2 

∆Hg(eff) Determination 

Nominal 

Value Range 

Material Bulk Properties: 

Thermal Conductivity 

Specific Heat 

Absorption Coefficient 

W m-1 K-1 

kJ kg-1 K-1 

m-1 

0.235 

2.22 

960 

0.1 to 0.5 

1 to 5 

200 to 50000 

0.22 

1.42 

∞ 400 to 50000 

Material Decomposition 

Properties: 

Heat of Reaction 

Activation Energy 

Pre-exponential 

kJ kg-1 

kJ mol-1 

s-1 

2000 

209 

1.29 x 1015 

1000 to 5000 

97 to 837 

1.02 x 106 

to 1.25 x 1066 

840 

97 

1.02 x 106 

97 to 1050 

1.02x106 

to 1.24 x 1083 

Experiment Properties: 

Incident Heat Flux kW m-2 50 21 to 200 20 to 70 

Thickness mm 25.4 2 to 32 15 

Not varied: 

Density 

Ambient Temperature 

kg m-3 

K 

1190 

293 

1190 

293 

Convective Heat Transfer 

from Surface W m-2 K-1 0 10 

Surface Emissivity 0.95 1 

22
 
 

 

 

 

 

 

 

 

 

 

 

Table 2 – Power-law (y=Axn) fit parameter n for y=ignition time, τign or average mass loss rate MLRav, with 

x=k, c, α, or ∆Hreac. 

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Power-Law Parameter n 

Parameter: ∆Hreac c α∗ k 

Flux τign MLRav τign MLRav τign MLRav τign MLRav 

21 1.70 -0.72 1.30 -0.32 -0.03 0.016 0.71 -0.005 

50 0.140 -0.745 1.00 -0.26 -0.54 -0.079 0.54 -0.004 

100 0.046 -0.69 0.91 -0.26 -0.67 -0.12 0.33 -0.010 

150 0.044 -0.64 1.00 -0.26 -0.75 -0.14 0.29 -0.009 

200 0.020 -0.80 0.97 -0.27 -0.77 -0.12 0.28 -0.010 

Expected value 

0.0 -0.72 1.0 -0.30 n.a. n.a. 1.0 0.0 

* the data range for α was limited to 200 m-1 to 1200 m-1 to give a good fit. 

n.a. not available. 

Table 3 – Influence of model input parameters on mass loss rate and ignition time. 

Parameter Mass Loss Rate (MLR) Ignition Time 

Shape Average peak tpeak 

∆Hreac     LF 

S  HF

Lα  LF 

HF c   LF 

HF α  LF not LF 

k  LF 

E HF a LF LF LF 

23
 
Mass Loss Rate / g m-2 s-1 MLR / MLRnorm 

21 

200 80 

1.5 

50 

100 

150 

200 

1.0 60 150 

0.5 

100 40 

0.0 

0.0 0.5 1.0 t / tnorm 

50 20 

21 

0 

0 2000 4000 6000 Time / s 

Figure 1 – Calculated mass loss rate of 25.4 mm thick black PMMA subjected to incident radiant fluxes of 

(21, 50, 100, 150, and 200) kW/m2

. (Inset shows same data with MLR and t normalized.)

F:\Home\Greg\Current Research\ConeFDS\Excel Files\bcase1compare0 sheet: Bcase0Stas,F21-200vAEa

24

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4 

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- MLR / g m-2 s -1 MLR / g m-2 s 1 

6 40

30 

20 

2 2 4 8 16 32 32 mm 

16 10 8 4 

2 

0 

0 5000 Time /s 10000 0 500 1000 Time /s 

0 

Figure 2 - Mass loss rate versus time for PMMA at external heat fluxes of a.) 21kW/m2 and b.) 100 kW/m2

 . 

Different curves on each frame show the effect of thickness of (2, 4, 8, 16, and 32) mm. Dotted line shows the 

result for 2 mm case with surface absorption.

File: vThPstas0.xls sheet: Flux 21 vary Th, Flux100 VaryTh, K=ramp, c=ramp, Ext=960 normal,. 

-1 -1 12 MLR / g m-2 MLR / g m s -2 s 80 

1000 

1000 

10 

60 

8 

2000 2000 

6 40 

3000 3000 

4 

4000 4000 

20 5000 5000

2 

0 0 

0 5000 10000 Time / s 15000 0 1000 2000 Time / s 

Figure 3 – Mass loss rate versus time for PMMA at external heat fluxes of (21and 100) kW/m2, frames a. ) 

and b.), respectively, with a value of Ea=209 kJ/mol. Different curves on each frame show the effect of ∆Hreac= 

(1000, 2000, 3000, 4000, and 5000) kJ kg-1. File: varyHgPStas0.xls sheet: Flux21varyHg 25.4 mm K=normal 

Ext=normal C=normal right, sheet: Flux100varyHg

25
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-1 -1 20 MLR / g m-2 s MLR / g m-2 s

6 

15 

97 

4 

1050 10 

2 

5 

0 0 

0 2000 4000 6000 Time / s 0 1000 2000 Time / s 

a.) b.)

/ -2 s-1 MMLRLR /g mg m-2 s-1 MLR / g m-2 s-1 

1050 

1050 

30 60 

97 97 

20 40 

10 20

0 0 

0 500 1000 Time / s 0 200 400 Time / s

c.) d.)

Figure 4 – Mass loss rate versus time for PMMA at external heat fluxes of (21, 50, 100, and 200) kW/m2, 

frames a. ) through d.), respectively. Different curves on each frame show the effect of Ea = (97, 209, 418, and 

1050) kJ/mol. File: RampsPMMA6Stas.xls sheet: RampKandC,F21-200vAEa fluxes of 21 50 100 200 each 

color is a different Ea 25.4 mm K-0.2 Ext=normal C=normal

26
0.1 

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7

MLR / g m-2 s-1 MLR / g m-2 s-1 

40
 

6
 0.5 

0.5 

5
 

30 0.1 

4 

20 3 

2 

10 

1 

0 0 

0 2000 4000 6000 Time / s 0 500 1000 Time / s 

Figure 5 - Mass loss rate versus time for PMMA at external heat fluxes of (21and 100) kW/m2, frames a. ) and 

b.), respectively, with a value of Ea=209 kJ/mol. Different curves on each frame show the effect of k = (0.1, 0.2, 

0.3, 0.4, and 0.5) W m-1 K-1. File: vKPStas0.xls sheet: Flux21varyK Ea=50 each color is a different K=0.1 to 

0.5 25.4 mm Ext=normal C=normal right, sheet: Flux100varyK

MLR / g m-2 -1 s-1 MLR / g m-2 s

6
 

40
 50000 200 200 

50000 

30
 

4
 

20 

2 

10 

0 

0 5000 Time / s 10000 0 500 1000 Time / s 1500 

0 

Figure 6 - Mass loss rate versus time for PMMA at external heat fluxes of (21and 100) kW/m2, frames a. ) and 

b.), respectively, with a value of Ea=209 kJ/mol. Different curves on each frame show the effect of alpha = 

(200, 400, 800, 1200, and 50000) m-1. vExtPStas1.xls sheet: Flux21varyExt Ea=50 each color is a different 

Ext=00200 to 50000 25.4 mm K=normal C=normal right, sheet: Flux100varyExt

27
4 

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6 

MLR / g m-2 s-1 MLR / g m-2 s-1 

1 

40 

1 

5 

30 

5 20

2 

10 

0 0 

0 5000 Time / s 10000 0 500 1000 Time / s 1500

Figure 7 - Mass loss rate versus time for PMMA at external heat fluxes of a.) 21kW/m2 and b.) 100 kW/m2, 

with Ea=209 kJ/mol. Different curves on each frame show the effect of c = (1, 2, 3, 4, and 5) kJ kg-1 K-1 . m-1. 

vCStas0.xls sheet: Flux21varyC Ea=50 each color is a different C=1 to 5 25.4 mm Ext=normal K=normal 

C=normal right, sheet: Flux100varyC

28
 

 

 k: file: 

vKPStas0.xls ; sheet: Flux 021

10000 

10000 

τign 

21 

50 

100 

150 

200 

k Flux 

(kW/m2

) 

------------­

Ea = 837 kJ/mol 

418 

209 

97 

τ ign 

21 

50 

100 

150 

200 

Flux, n 

(kW/m2

) 

-----------­

c

 840 

420 

210 

97 

1000 

1000 

100 100 

10 10 

1 1 

0	 0.1 0.2 0.3 0.4 0.5 0.6 0 1 2 3 4 5
 

k / W m-1 K-1 Specific Heat / kJ mol-1 K-1
 

10000 τ ign 

21 

50 

100 

150 

200 

Flux 

(kW/m2

) 

-----------­

α 

837 

418 

209

 97 

τ ign 

10000 

21 

50 

100 

150 

200 

∆Hreac 

Flux 

(kW/m2

) 

-----------­

840 

420 

210

 97 

1000 1000 

100 100 

10 10 

1 1 

100 1000 10000 100000 0 2000 4000 6000 

Absorption Coefficient / m-1 Heat of Reaction / kJ mol-1 

Figure 8 – Ignition time variation for values of the external heat flux of (21, 50, 100, 150, and 200) kW/m2, 

and values of the PMMA decomposition activation energy of (97, 209, 518, and 1050) kJ/mol: a.) thermal 

conductivity, k; b.) specific heat, c; c.) absorption coefficient, α; and d.) heat of reaction, ∆Hreac.

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29
 

6 
 

 

vThPstas0.xls sheet: Flux021

10000 10000 S, α =960 m-1 

Flux 

1000 1000 21 21 

50 

100 

150 

200 

Flux 

S, α =50000 m-1 

τign 

τ

100 

ign 

50 

100 

150 10 

100 

10 

200 

1 1 

0 10 20 30 0 10 20 30 

Thickness / mm Thickness / mm 

Figure 9 - Ignition time variation for values of the external heat flux of (21, 50, 100, 150, and 200) kW/m2

, and 

values of the PMMA decomposition activation energy of (97, 209, 518, and 1050) kJ/mol and a.) α=960 m-1, b.) 

α=50000 m-1.

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30
 
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80 

70 

400 

60 

70 kW/m2 

50 60 300 

40 50 

200 

30 

40 

20 30 

100 

20 

10 

0 0 

0 500 1000 1500 2000 2500 3000 3500 

Time / s 

Figure 10 – Calculated mass loss rate (solid lines, left axis) and surface temperature (dotted lines, right axis) 

as a function of time for 25.4 mm thick PMMA samples exposed to external radiant fluxes of (20, 30, 40, 50, 

60, and 70) kW/m2

. Horizontal bars are the average burning rate. File: StaggsPMMA-4_sm03.xls (Stas calc) 

Sheet: Ext50000,F20-70vAEa Sheet: Ext50000,F20-70vAEa TS

31

Mass Loss Rate / g m-2 s-1

Surface Temperature / oC 
 

 File: 

StaggsHgdata.xls Sheet: Staggs mdot

40
 

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"Steady" Mass Loss Rate / g m -2 s-1 

30
 

20
 

10
 

Staggs 

Thermakin 

FDS 

1/Slope: 

1.875 

1.755 

1.754 

Model: 

0
 

20 30 40 50 60 70
 

External Heat Flux / kW m-1 

Figure 11 – Calculated mass loss rate as a function of external heat flux for 25.4 mm thick PMMA, from FDS, 

ThermaKin, and analytic model of Staggs[22] (points: data; dotted lines: linear curve fit). 

32
 
 

 

 File: StaggsHgdata.xls sheet: Staggs mdot

1.0

1.2

550

600

650

700

750

800

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∆

Hg(eff) or ∆ 

Hg / kJ g-1 

1.8 

1.6 

1.4 

350 

400 

450 

500 

T surface / o C 

Staggs 

Thermakin 

FDS 

∆Hg 

(eff) 

∆Hg 

T 

10 20 30 40 50 60 70
 

External Heat Flux / kW m-2 

Figure 12 - Calculated heat of gasification ∆Hg and effective heat of gasification ∆Hg

(eff) (left scale), and 

surface temperature (right scale) as a function of external heat flux for 25.4 mm thick PMMA, from FDS, 

ThermaKin, and analytic model of Staggs[22]. 

33
 
 

 File: 

StaggsPMMA-4_sm03.xls sheet: Hgeff_hg plot simple

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1.4 

1.3 

(eff) / ∆ 

Hg ∆ 

Hg1.2 

Staggs 

ThermaKin 

FDS 

20 30 40 50 60 70 80
 

External Heat Flux / kW m-2 

1.1 

1.0 

Figure 13 - Ratio of effective heat of gasification ∆Hg

(eff) to the heat of gasification ∆Hg as a function of 

external heat flux for 25.4 mm thick PMMA, from FDS, FAA, and analytic model of Staggs[22]. 

34
 
 

 

 File: 

StaggsPMMA-4_sm03.xls Sheet: SumMdots

0.0

10.0

20.0

30.0

40.0

50.0

0.0

10.0

20.0

30.0

40.0

50.0

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50 Absorption Coef. = 400 m-1 930 m-1 

1050
 

418
 

209
 40
 97
 

30
 

20
 

10
 

0
 

20 30 40 50 60 70 20 30 40 50 60 70 20 30 40 50 60 70

External Heat Flux / kW m-2 

Figure 14 – Calculated mass loss rate as a function of the external heat flux for 25.4 mm thick PMMA, with
 

values of the absorption coefficient of (400, 930, and 50000) m-1. In each frame, the different curves 

correspond to PMMA decomposition activation energies of (97, 209, 418, and 1050) kJ/mol.
 

50000 m-1 

Ea = 

Mass Loss Rate / g m-2 s-1 

35
 
 

 

 

 File: StaggsPMMA-4_sm03.xls Sheet: 

Sum Hgeff,Hg,Tsurf

1.0

1.2

1.4

1.6

1.8

320

370

420

470

520

570

620

670

720

770

1.0

1.2

1.4

1.6

1.8

520

570

620

670

720

770

320

370

420

470

520

570

620

670

720

770

1.8 

This publication is available free of charge from: http://dx.doi.org/10.6028/NIST.TN.1929 

T 

930 m-1 

Surface Temperature C

T

 50000 m-1 

20 40 60 80 

T 

209 

418 

1050

 Ea = 

97 

Absorption Coef. = 400 m-1 

∆Hg 

(eff) 

∆Hg 

97 

1050 

1050 

97 

1.6 

(eff) or ∆ Hg ∆Hg / kJ kg-1 

1.4 

1.2 

420
 

370
 

1.0 320
 

20 40 60 80 20 40 60 80
 

External Heat Flux / kW m-2 

Figure 15 – Calculated effective heat of gasification ∆Hg

(eff) , heat of gasification ∆Hg , and polymer surface 

temperature as a function of the external heat flux for 25.4 mm thick PMMA, with values of the absorption
 

coefficient of (400, 930, and 50000) m-1. In each frame, the different curves correspond to PMMA
 

decomposition activation energies of (97, 209, 418, and 1050) kJ/mol.
 

36
 

470 
 

 

 

 

 File: StaggsPMMA-4_sm03.xls Sheet: Hg_Hg Plot (change cell 

B11 to 0 in sheet: Hg,eff)

0.9

1.1

1.3

0.9

1.1

1.3

This publication is available free of charge from: http://dx.doi.org/10.6028/NIST.TN.1929 

(eff) / 

∆

Hg ∆

Hg 

1.4 Absorption Coef. = 400 m-1 

1.3 

1.2 

1.1 

1.0 

0.9 

209 

418 

1050

 Ea = 

97 

930 m-1 50000 m-1 

20 40 60 80 20 40 60 80 20 40 60 80

External Heat Flux / kW m-2 

Figure 16 – Calculated ratio of effective heat of gasification ∆Hg

(eff) to the heat of gasification ∆Hg as a 

function of the external heat flux for 25.4 mm thick PMMA, with values of the absorption coefficient of (400,
 

930, and 50000) m-1. In each frame, the different curves correspond to PMMA decomposition activation
 

energies of (97, 209, 418, and 1050) kJ/mol.
 

37
 
 

 

 File: StaggsPMMA-4_sm03.xls 

Sheet: Hg_Hg Plot (change cell B11 to 1 in sheet: Hg,eff).

0.9

1.1

1.3

0.9

1.1

1.3

This publication is available free of charge from: http://dx.doi.org/10.6028/NIST.TN.1929 

1.4 Absorption Coef. = 400 m-1 

209 

418 

1050

 Ea = 

97 

930 m-1 50000 m-1 

1.3 

1.2 

1.1 

1.0 

0.9
 

20 40 60 80 20 40 60 80 20 40 60 80

External Heat Flux / kW m-2 

Figure 17 – Calculated ratio of effective heat of gasification ∆Hg

(eff) to the heat of gasification ∆Hg as a 

function of the net heat flux for 25.4 mm thick PMMA. In each frame, the different curves correspond to
 

PMMA decomposition activation energies of (97, 209, 418, and 1050) kJ/mol.
 (eff) / ∆Hg ∆Hg 

38