NISTIR 6993

Thermal Performance of Fire Fighters’ Protective Clothing.

3. Simulating a TPP Test for Single-Layered Fabrics

Scott Kukuck

Kuldeep Prasad

Building and Fire Research Laboratory

National Institute of Standards and Technology

Gaithersburg, MD 20899-8663

Sponsored by

Department of Homeland Security

Federal Emergency Management Agency

United States Fire Administration
NISTIR 6993

Thermal Performance of Fire Fighters’ Protective Clothing.

3. Simulating a TPP Test for Single-Layered Fabrics

Scott Kukuck

Kuldeep Prasad

Building and Fire Research Laboratory

National Institute of Standards and Technology

Gaithersburg, MD 20899-8663

January, 2003

Department of Homeland Security

Tom Ridge, Secretary

Federal Emergency Management Agency

Michael B. Brown, Acting Under Secretary of

Emergency Preparedness and Response

U.S. Fire Administration

R. David Paulison, Administrator

U.S. Department of Commerce

Donald L. Evans, Secretary

Technology Administration

Phillip J. Bond, Under Secretary for Technology

National Institute of Standards and Technology

Arden L. Bement, Jr., Director
1 Abstract

Fabrics that are used in firefighter turnout gear are specifically designed to offer protection from various

extreme conditions that may arise during fire exposures. To optimize the design of these garments and

understand the conditions for which protection is offered requires knowing how the transient response of the

fabrics is affected by the many possible factors that arise. A model was recently developed that describes

the heat and mass (moisture) transfer across protective fabrics subjected to incident thermal conditions and

illustrated the importance of including moisture transport effects even under low incident flux conditions.

This model is now used to study the transient response of protective fabrics to high intensity, short duration

heat flux exposures. Numerical simulations of a Thermal Protective Performance test were conducted for

two fabrics used as outer shell materials for turnout garments. Predicted results were found to favorably

agree with experimental measurements, with Thermal Protective Performance ratings from the model within

6% of measured ratings. The ability of the model to describe the temperature across the entire thickness of

the fabric rather than simply present a sensor response during an exposure provides additional information

that can be further utilized to predict the thermal degradation of the materials.

2 Introduction

A review of the statistics for the years 1990 through 2000 shows a general downward trend in the number

of firefighter injuries [1] and fatalities [2] due to burn injuries (note that these statistics include both thermal

and chemical burns). While this trend points to increased safety from burns to the fire service, in the year

2000 there were still 6 deaths and 4500 injuries to personnel where the nature of the injury or fatality was

attributed to burns. Firefighting is inherently dangerous but there is still much work that can be done to

prevent burn injuries and fatalities.

Firefighter protective clothing garments are the shield for the firefighter from insults that could prove

otherwise harmful or fatal. Two items related to protective garments may be suggested to further decrease

the number of burn incidents. The first item is added education for the fire fighter regarding the level of

protection offered by his / her protective garments under fire conditions. The second item is technological advancements in the materials utilized and the design of these garments, advancements that optimize

material properties so as to provide the greatest level of protection. Both items require an enhanced understanding of the response of protective fabrics to thermal and environmental conditions experienced during

fire exposures.

As part of ongoing work aimed at improving fire service technologies and reducing fire loss, a simplified

model was recently developed [3] that describes the transient response of textile materials subjected to

incident radiative and convective heat fluxes. The model incorporates not only the effects of heat transported

across multiple fabric layers but also accounts for the presence and transport of moisture across these layers

as well. The treatment of moisture allows for adsorption / desorption within the fibers as well as water vapor

diffusion through the porous material. Initial results obtained using this model focused upon the response

of fabrics (both wet and dry) that were subjected to low levels of incident heat flux. The results showed

good agreement with measured responses and demonstrated the importance of including the presence and

transport of moisture when analyzing the thermal behavior of protective fabrics.

Firefighter exposures are not limited to simply low incident flux conditions and, therefore, we wish to

expand the range of conditions for which the model will reliably predict the response of protective fabrics.

We consider here heat flux exposures of high intensity but short duration and choose to emulate a Thermal

Protective Performance (TPP) test for single-layered fabrics [4]. TPP tests are used as an attempt to quantify

the insulating characteristics of fabrics under conditions that may be observed during a fire. These test are

typically of short duration, less than 30 seconds, and expose the fabric to high incident convective and

1
TO RECORDER

COPPER CALORIMETER

INSULATING

BOARD

SPACER (6.4mm)

WATER

COOLED

SHUTTER

FABRIC

UNDER TEST

RADIANT QUARTZ TUBES

GAS BURNER GAS BURNER

127 mm

∅ 40 mm

32 mm

100 mm square

45o TYP

Figure 1: Schematic of the test apparatus for determining the TPP rating of the single-layered fabric. A

series of radiant quartz tubes and two gas burners supply convective and radiant heat to the fabric under test.

The heat transferred across the fabric is measured with a copper calorimeter positioned 0.0635 m (0.25 inch)

above the fabric.

radiative fluxes.

Two protective fabrics are considered in the following work: Nomex IIIA, a blend of 93% Nomex, 5%

Kevlar and 2% P140 fibers; as well as a 60% Kevlar and 40% Polybenzimidaxazole (PBI) blend1. The

material weight for both fabrics was 0.254 kg/m2(7.5 oz/yd2) and both materials are used as outer shell

fabrics for turnout garments. A series of experimental TPP tests were conducted on each fabric as were

numerical simulations utilizing the simplified model. The simulated transient response of the heat sensor

used in the TPP test, a slug copper calorimeter, is compared with the measured response obtained from the

experiments. From these response curves, the predicted TPP rating also is compared with the rating obtained

from the experiments.

3 The TPP Test

The configuration for the single-layered TPP test utilized in this report follows ISO Standard 17492 [4] and

is presented schematicly in figure 1. A series of radiant quartz tubes and a set of two burners supply incident

heat flux to the test fabric. The orientation of the apparatus is such that the heat sources lie beneath the

test fabric with the heat impinging on a single exposed surface. If the fabric has a preferred orientation,

then it is set in the appratus such that the exterior surface is the side directly exposed to the incident flux.

Heat is transferred across the fabric to the sensor assembly positioned 0.064 m (0.25 inch) above the fabric.

The sensor assembly is composed of a copper slug calorimeter, an insulating board and a series of iron /

constantan thermocouples. The calorimeter face is coated with a flat black paint so as to absorb radiant flux.

Although configuration shown in figure 1 is not standard in the United States, the setup and procedure

is similar to, and does include, elements of both the NFPA standard TPP test [5] and the (withdrawn)

ASTM standard TPP test [6]. Similar to the NFPA standard TPP test, the heat source specified includes a

1Certain commercial materials are identified in this paper in order to specify the experimental procedure 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 materials or equipment identified are necessarily the best available for the purpose

2
combination of two burners and a set of radiant quartz tubes, providing both convective and radiative modes

of heat transfer. With the inclusions of the burners and the radiant quartz tubes, and with proper balancing

(described in the calibration procedure below), the incident heat flux is presumed to have equal contributions

(50%/50%) from the two heat transfer mechanisms. As in the withdrawn ASTM standard, an 0.064 m (0.25

inch) spacer is placed between the fabric under test and the sensor assembly, providing an insulating air gap

for single-layered fabrics.

In the presented configuration, calibration of the test apparatus is performed by first setting the radiant

energy from the radiant quartz tubes to be (40±10) kW/m2, and then measuring the combined flux from

the radiant quartz tubes and the burners with a total calorimeter / radiometer combination. Adjustments

are made to both sources so as to insure the total heat flux of (80±2) kW/m2and the contribution of 40

kW/m2 due to incident radiation. Once the heat sources have been properly set, the total calorimeter /

radiometer combination is removed and the calorimeter inserted at a location corresponding to what would

be the exposed face of the test fabric. The calorimeter is then directly exposed to the heat sources. Incident

heat flux is evaluated from temperature measurements of the calorimeter. For a 10 second sampling period

of calibration, the average temperature rise should be (148±3.7)◦C.

Following the calibration procedure, the fabric under test and the spacer blocks are placed within the

apparatus. The water cooled shutter is then opened to expose the fabric to the incident heat flux. Temperature

measurements are obtained from the thermocouples within the calorimeter. The TPP rating is established by

plotting on a single set of axes the time versus temperature response of the calorimeter along with a curve

that represents the onset of a second-degree burn for applied heat flux versus time [7]. The intersection of

the two curves yields the time, tb for which the fabric and air gap would have provided protection from the

applied exposure. This time is then used in the relation,

TPP rating = F ×tb, (1)

where F is the applied heat flux, the specific value of which is obtained from the calibration procedure, and

tb is the time to second-degree burn, obtained from the plot. The units of the TPP rating are thus either kJ/m2

in the SI system or, as is common in the United States, cal/cm2(the minimum NFPA Standard TPP rating

of 35 cal/cm2for protective garments would correspond to an SI rating of 1460 kJ/m2and a 17.5 second

exposure until the onset of a second-degree burn).

4 Simulating the TPP Test

The computer code utilized to simulate the thermal and mass transport across textile materials is called

the Protective Clothing Performance Simulator (PCPS) [3, 8]. PCPS is specifically designed to solve a

simplified system of model equations and calculate the response of one-dimensional textile fabrics subjected

to conditions that may be observed during fire exposures. The numerical model incorporates separate layers

of a composite fabric so as to evaluate the response of both the individual materials and the entire ensemble.

Heat is supplied to the fabric by applying a combination of two boundary conditions. The first is a specified

external radiant flux and the second is an ambient gas temperature. The presence of moisture within the

fabric, the ambient environment and any air gaps are included in the model. Transport of water vapor across

the system is through diffusion and is computed during a simulation. The model also allows adsorption and

desorption of moisture within the pore spaces.

Material properties for the individual fabrics are drawn from a database coded into the software with

values taken from the measurements of Lawson and Walton [9]. As the material properties were only measured over finite temperature ranges up to 100◦C, they are taken to be constant above this limit. The material

properties do not account for any thermal degradation that may occur at sufficiently high temperatures.

3
SPACER

FABRIC

UNDER TEST

COPPER

CALORIMETER

INSULATING

BOARD

EXTERNAL RADIATION

Thot gas Tcold

Figure 2: The model of the ISO Standard 17492 TPP test apparatus used for predicting the TPP rating of a

single layer fabric.

While developed for porous materials, the governing equations for PCPS were formulated utilizing a

continuum approximation. Therefore, in addition to considering heat and mass transfer through woven

fabrics, PCPS also will simulate the transfer of heat through nonporous materials as well. The sensor

assembly of the TPP test is simulated as a layer of copper and a layer of insulating board. The entire TPP

test apparatus is presented as a series of individual layers of a multi-component “fabric.”

A schematic representation of the configuration utilized for simulating the TPP test is presented in

figure 2. Heat is supplied to the left boundary of the fabric under test from a hot gas ambient and an

externally applied radiant flux. Between the fabric and the copper calorimeter is an air gap 0.064 m in

width. Beyond the copper, without an air gap, is a layer of insulating material which is exposed to the cold

ambient.

4.1 Calibration of the Model

As previously stated, the net heat flux (convective and radiant fluxes) supplied to the fabric under test is

set during the calibration procedure by measuring the temperature rise in the calorimeter over a 10 second

sampling. The need for the time averaging is due to the fact that the instantaneous flux is constantly changing

during the test. As the material warms, the temperature difference between the fabric surface and gas

decreases, diminishing the convective heat flux to the fabric. Similarly, the net incident radiation decreases

due to the re-radiation of the fabric surface as it warms. Thus, although we may initially supply 80 kW/m2 of

heat energy to the fabric, at the end of the 10 second exposure the value may vary significantly. To determine

the combination of externally supplied radiant energy and the hot ambient temperature necessary to achieve

the 148◦C rise in temperature, it is necessary to perform a calibration procedure similar to that employed

4
0

100

200

300

400

500

600

700

800

900

0 20 40 60 80 100

q'' ext (kW / m2

)

Externally applied radiant flux

Ambient Temperature Difference

DT (°C)

Calorimeter Temperature rise of 148°C

D T =792°C

q' ext=87.5 kW / m2

Figure 3: Required ambient temperature difference versus external radiant flux that produces a temperature

rise of 148◦C for a 10 second exposure to bare copper calorimeter during calibration of TPP test.

during the experiment.

To mimic calibration, a series of simulations were performed in which the fabric and air gap were

removed from the model, as in the physical test, and the copper was exposed directly to the heat sources.

Simulations were run for 10 seconds with the average temperature of the copper obtained at the end of the

simulation. The incident radiant flux and the hot gas temperatures were adjusted until the desired 148◦C rise

was obtained.

Results of the calibration procedure are shown in figure 3. For the required 148◦C temperature rise,

the difference between the ambient environment temperature and the initial temperature of the material

is plotted as a function of the external radiant flux. For zero externally supplied flux, and thus entirely

convective heating, the temperature difference between the ambient and the initial temperature of assembly

must be 792◦C The inital temperature difference then decreases as the level of externally applied radiant

flux increases. For an ambient temperature equal to the initial temperature of the fabric, and thus heating

strictly from the applied radiation, a flux value of 87.5 kW/m2is required. As expected the applied flux

must be greater than the specified average of 80 kW/m2 of the ISO standard to account for the convective

and radiative losses that occur as the material warms.

4.2 Results and Discussion

Experimental measurements for the TPP results presented were conducted in the thermal lab of DuPont’s

Spruance Plant facility. Raw data was then transferred to and analyzed at NIST. Numerical simulations of

the TPP tests were also conducted at NIST.

The results of the study are presented in figure 4. Figure 4a is for the Nomex IIIA simulations and

experiments while figure 4b is for the PBI / Kevlar blend. These plots present the temperature of the copper

calorimeter as a function of time. For each material three experimental tests were performed with the

5
5.0

0.0

20.015.010.0

0 4 8 12

Time (s)

Calorimeter Temperature Rise (°C)

0 8 16 24

TPP value

Model, Qe=0Model, Qe=41.5Model, Qe=87.5Experiment 1Experiment 2Experiment 32nd degree burnTPP rating

Nomex IIIA

5.00.0

20.015.010.0

0 4 8 12

Time (s)

Calorimeter Temperature Rise (°C)

0 8 16 24

TPP value

Model, Qe=0Model, Qe=41.5Model, Qe=87.5Experiment 1Experiment 2Experiment 32nd degree burnTPP rating

PBI / Kevlar Blend

(a) (b) Figure 4: Calorimeter temperature response for (a) Nomex IIIA and (b) PBI / Kevlar blend. Solid lines denote model simulations and discrete points experimental measurements. The TPP criteria curve for onset of second-degree burn is also presented.

6
measured results presented as the discrete points. The TPP rating obtained from the experiments is presented

as the larger solid dot overlaying the second-degree burn criteria. Results of the numerical simulations

are presented as the solid and dashed lines for three unique incident flux conditions. The nominal 50%

convective, 50% radiative incident flux is presented as the solid line. The short dashed line corresponds

to zero externally applied radiative flux and represents strictly convective heating. The long dashed line

corresponds to an externally applied radiative flux equal to 87.5 kW/m2.

The experimental results for each of the fabrics show that the temperature of the calorimeter rises monotonically for the duration of the test. There is, however, for each of the fabrics a slight inflection that occurs

shortly, approximately 2 seconds - 3.5 seconds, into the test. An overall TPP rating for each of the fabrics

was calculated by averaging the results of three independent tests. For Nomex IIIA, the experimental TPP

rating was measured as (640±50) kJ/m2((15.4±1.2) cal/cm2) and for the PBI / Kevlar blend the TPP rating

was measured as (650±50) kJ/m2(15.5±1.2) cal/cm2). The uncertainty in the experimental TPP rating is

presented as the error bars of figure 4. This uncertainty was calculated using the statistical variation of the

three independent measurements (Type A) as well as the allowable limits of the test configuration (Type B).

The type B uncertainty was evaluated by considering rectangular probability distributions for the the tolerance limits presented [4]. Using a coverage factor of two, for each of the materials the expanded uncertainty

was evaluated as plus or minus 7.5%.

The model results for Nomex IIIA show that for the idealized 50% radiative and 50% convective case the

temperature is underpredicted for a range of approximately 2 seconds to 3.5 seconds, and then overpredicted

as time increases and the calorimeter temperature approaches the second-degree burn criteria. Unlike the

experimental results, there is not an inflection in the curve. At no time during the length of the test does the

temperature predicted by the model vary by more than 1◦C from the experimental measurements. For the

high radiative flux simulation, the curve rises at a much steeper rate and intersects the second-degree burn

curve more than one second before the 50/50 case. For the zero radiative flux simulation, and thus pure

convection, the curve rises at a shallower rate and intersects the second-degree burn curve more than one

second later than the 50/50 case.

Similar results are observed for the PBI / Kevlar blend. Again, early in the simulation, the model

underpredicts the temperature for the 50/50 case. Unlike the previous results, however, this trend continues

for the entire length of the experiment although the results are converging near the end of the test. The model

does not do quite as accurate a job of predicting the temperature for the PBI / Kevlar blend test as compared

to the Nomex IIIA test; however, the predicted temperature is always within 2◦C of measured values. Similar

to the Nomex IIIA results, the higher level of radiation results in a steeper rise in the calorimeter temperature

and for no radiative flux, a shallower rise.

As observed from the results, higher amounts of radiative flux result in a faster rise in the calorimeter

temperature. This expected result is due to the penetrating nature of the incident radiation. Unlike convection, which is treated as a surface phenomena, radiation can penetrate into and across the entire width of the

fabric. Thus the temperature of the outer surface of the fabric, the surface exposed to the incident heat flux,

is greater for the strictly convective case, but the heat transferred across the fabric and to the calorimeter

is greater for the case of pure radiation. This result is clearly dependent upon the optical properties of the

material under test. For highly reflective surfaces, such as aluminized outer shells in proximity garments,

this behavior would not be expected. The results demonstrate the importance of identifying the nature of the

incident heat flux when ascertaining the amount of protection afforded by protective fabrics.

A comparison of the TPP ratings achieved from the experimental measurements and model predictions

is presented in table 1. Although the Nomex IIIA value is slightly underpredicted and the PBI / Kevlar

blend slightly overpredicted in both cases the predictions lie within 6% of the measured values. Further,

the predicted values lie within the calculated uncertainty for the measured results. As described above,

the predicted TPP ratings for both materials under pure radiative heating is significantly less than for the

idealized 50/50 case, while for strictly convective heating the TPP rating is significantly greater.

7
Condition Nomex IIIA PBI / Kevlar

kJ/m2cal/cm2 kJ/m2cal/cm2

Measured 640±50 15.4±1.2 650±50 15.5±1.2

50% Convective / 50% Radiative 607 14.5 682 16.3

Radiative limit 510 12.2 582 13.9

Convective limit 732 17.5 803 19.2

Table 1: Measured and predicted TPP ratings for Nomex IIIA and PBI / Kevlar blend.

Mentioned previously, an interesting feature of the experimental data is the presence of a slight inflection

for both of the materials. For Nomex IIIA, this occurs at approximately (2 - 2.5) seconds, whereas for the

PBI/Kevlar blend it occurs at approximately (2.5 - 3) seconds. Neither inflection is very sharp, but each

corresponds to a change in the rate of temperature rise in the calorimeter. The model did not reproduce these

results; however, it may provide a explanation for the observed inflections.

Figure 5 plots the surface temperatures of both the exposed and unexposed faces for each material.

Also denoted on the figure are corresponding oxidation and decomposition temperatures for each of the

materials. Nomex IIIA is composed primarily (93%) of Nomex fiber which begins to undergo oxidation at

approximately 427◦C [10]. For the PBI / Kevlar blend, the Kevlar fibers (60%) begin to thermally degrade

between 427◦C and 482◦C [11]. Both of these reactions are endothermic, absorbing some of the incident

heat energy and slowing the rate of heat transfer across the fabric and into the calorimeter. The Nomex

fibers also swell as the oxidation reaction occurs, reducing the porosity of the material and further altering

the heat transfer mechanisms. Kevlar fibers do not swell in this manner. From figure 5 we observe that at the

approximate location of the inflection in the experimental data, the outer surface temperature of each of the

fabrics has reached a temperature that is near either the oxidation point (for Nomex) or the degradation point

(for Kevlar). The heat transfer across the fabric is thus alterred perhaps resulting in the measured inflection

for each fabric.

5 Conclusion

We have summarized a series of simulations that have been performed to predict the response of a single

layer fabric to the high intensity, short duration convective and radiant heat fluxes that may arise during fire

exposures. The predictions have been compared with experimental measurements for two different fabrics

utilized in firefighter protective garments. To compare predicted behavior to measured behavior we have

chosen to simulate a Thermal Protective Performance test. The results show that the predicted response of

the calorimeter sensor from the model system of equations closely matches with experimental results for the

case of 50% convective and 50% radiative heat transfer mechanisms.

Predicted TPP ratings for the fabrics also yield good agreement with experimental results. For Nomex

IIIA, the predicted value of 607 kJ/m2(14.5 cal/cm2) is within 6% of the experimental result of (640±50)

kJ/m2((15.4±1.2) cal/cm2). For the PBI / Kevlar blend the predicted value of 682 kJ/m2(16.3 cal/cm2) was

also within 6% of the the experimental result of (650±50) kJ/m2((15.5±1.2) cal/cm2). For both materials

the predicted value fell within the range of uncertainty for the experimental measurements.

A possible explanation for observed inflections in the calorimeter response during the experimental

measurements was also presented. Model results showed that approximately 2 seconds into the exposure, the

outer surface temperature of the Nomex IIIA fabric had reached a temperature sufficient to begin oxidation

of the Nomex fibers. Likewise, approximately 3 seconds into the exposure, the outer surface temperature of

the PBI / Kevlar blend had reach a temperature sufficient to begin decomposition of the Kevlar fibers. Both

of these times roughly correspond to the observed inflections of the experimental data, although more tests

8
(a) (b)

0

750500250

0 2 4 6 8

Time (s)

Surface Temperature (°C)

Exposed SurfaceUnexposed Surface

OxidationTemperature

Nomex IIIA

0

750500250

0 2 4 6 8

Time (s)

Surface Temperature (°C)

Exposed SurfaceUnexposed Surface

Decomposition

Temperature

PBI / Kevlar Blend

Figure 5: Exposed and unexposed surface temperatures for (a) Nomex IIIA and (b) PBI / Kevlar blend. For Nomex IIIA the temperature at whichthe Nomex fiber undergoes oxidation is presented and for the PBI / Kevlar blend the approximate decomposition temperature of the Kevlar fiber isdenoted.

9
showed be made to confirm this conclusion.

Typical simulations utilized in this work were performed in less than 10 minutes on a 933 MHz Pentium

III desktop computer with 1GB of RAM. This performance implies a large potential for conducting more

complete parametric analyses of fabric responses to various conditions. Physical parameters such as thickness and weight; thermal properties such as thermal conductivity and specific heat; and optical properties

such as absorptivity and transmissivity can be systematically varied so as to determine the most efficient

means of providing maximum protection while balancing undesireable quantities.

In addition to studying the physical characteristics of the fabrics utilized in protective clothing, the

model can be used to determine how fabrics respond to the way in which the thermal energy is applied

as well as how they respond to other environmental factors. Previous discussions on materials utilized for

protective garments have centered on issues such as the effect of moisture upon the insulating characteristics

of fabrics, the presence of stored heat from previous exposure, and the shape of the incident heat flux curve

versus time [12, 13]. Any of these items may affect the performance of a firefighter’s protective garment.

6 Future Research

Of particular interest is the effect of moisture on the insulating characteristics of protective fabrics. Prior

results obtained with the model [3] have shown that the transport of moisture through a composite fabric

and the adsorption / desorption of water within the fabric weave can have a dramatic effect on how heat is

transported. While the added mass and the evaporative effect of water may slow the transport of heat across

a fabric layer, the thermal conductivity of water compared to the base material can also increase the rate in

which heat energy is transported across the fabric. An additional concern for wet fabrics is the transport

of hot water vapor which may condense further into the fabric at say the surface of the wearer’s skin. The

possibility of not only dry burns but also scalding burns therfore must be considered for the case of wet

fabrics.

While this work has restricted simulations to single layers of protective fabrics, similar work is needed

for multilayer fabrics such as are utilized in complete firefighter turnout garments. These garments are

composed of not only the outer shell, as utilized in this work, but also include a moisture barrier and thermal

liner as well.

Acknowledgements

Appreciation is extended to Robert T. McCarthy of the United States Fire Administration for assistance

provided during discussions related to firefighter burn injuries and protective clothing performance. The

authors acknowledge the assistance of Dr. Richard Young of DuPont Advanced Fiber Systems for informed

discussions and the willingness to share experimental data from DuPont to help advance and evaluate the

model. We also wish to extend their appreciation to Dan Madrzykowski, Nelson Bryner, Dave Evans, Randy

Lawson, Anthony Hamins and Bill Grosshandler for their support and guidance throughout this work.

10
References

[1] National Fire Protection Association, “United States Firefighter Injuries,” NFPA Journal, 85-95 No. 6

November / December (1991-2001).

[2] United States Fire Administration, “Firefighter Fatalities in the United States,” Annual Analysis Reports, United States Fire Administration, Federal Emergency Management Agency, Emmitsburg, MD

(1991-2001).

[3] K. Prasad, W. Twilley and J. Randall Lawson, “Thermal Performance of Fire Fighters’ Protective

Clothing. 1. Numerical Study of Transient Heat and Water Vapor Transfer,” Technical Report NISTIR

6881, National Institute of Standards and Technology, Gaithersburg, MD (2003).

[4] International Organization for Standardization, Clothing for protection against heat and flame - Determination of heat transmission on exposure to both flame and radiant heat, ISO Standard 17492,

Geneva, Switzerland (2003).

[5] National Fire Protection Association, Protective Clothing for Structural Firefighting, NFPA Standard

1971-2000, Quincy, MA (2000).

[6] American Society of Testing and Materials, Thermal Protective Performance of Materials for Clothing

by Open Flame Method, ASTM Standard D4108-1987 (withdrawn), West Conshohocken, PA (1987).

[7] A.M. Stoll and M.A. Chianta, “Method and rating system for evaluation of thermal protection,”

Aerospace Medicine, 40, pp. 1232-1238 (1969).

[8] K. Prasad, H.M. Ho Dac and S.R. Kukuck, “Thermal Performance of Fire Fighters’ Protective Clothing. 2. Protective Clothing Performance Simulator - User’s Manual,” Technical Report NISTIR 6901,

National Institute of Standards and Technology, Gaithersburg, MD (2003)

[9] J.R. Lawson and W.D. Walton, “Estimates of Thermal Properties for Fire Fighters’ Protective Clothing Materials,” Technical Report to be published, National Institute of Standards and Technology,

Gaithersburg, MD (2003).

[10] E.I. du Pont de Nemours and Company, Technical Guide for NOMEX Brand Fiber, H-52720, Richmond, VA (2001).

[11] E.I. du Pont de Nemours and Company, Technical Guide for KEVLAR Aramid Fiber, H-77848, Richmond, VA (2000).

[12] J.F. Krasny, J.A. Rockett and D. Huang, “Protecting Firefighters Exposed in Room Fires: Comparison

of Results of Bench Scale Test for Thermal Protection and Conditions During Room Flahsover,” Fire

Technology, 24 No. 1, pp. 5-19 (1988).

[13] R.D. Peacock, J.F. Krasny, J.A. Rockett and D. Huang, “Protecting Fire Fighters Exposed in Room

Fires, Part 2: Performance of Turnout Coat Materials Under Actual Fire Conditions,” Fire Technology,

26 No. 3, pp. 202-222 (1990).

11