Thermal Protective Performance of Turnout Gear at High Flux Environment

Abstract

Thermal protective performance (TPP) tests are conducted at a heat flux of 84 kW/m2 to evaluate a firefighter’s turnout gear performance. The test criterion used is based on a study conducted by Behnke in 1984. However, an average heat flux of 200 kW/m2 has been documented in the literature over the past decade. Henceforth, experiments are conducted on the turnout gear at a higher heat flux level of 126 kW/m2, which represents a typical heat flux during the initial phase of a fully developed fire. The analytical analysis provides insights into air gaps, the effect of smoke layers on radiative attenuation and the thermal decomposition of the fabric layers. Numerical techniques were employed to introduce a survival curve to effectively gauge the degree of spatial burn at variable levels of heat fluxes, moisture effects and superficial burns. The sustainability of high resistance fabrics in terms of toxicity of flame retardants and recycling of these textiles have been discussed. Bio-composites as flame retardants are being introduced to replace traditional flame-retardant chemicals.

1. Introduction

Turnout gear is worn to safeguard individuals employed in hazardous fields like firefighting, high-speed sports, the oil and gas industry and the military. The protective gear can be either thermal protective clothing for firefighting, chemical resistant clothing for the oil and gas industry or puncture- and cut-resilient clothing for the military [1]. Thermal protective clothing is crucial for firefighters because fire is a common occurrence and poses a significant threat to people’s safety. Firefighter protective clothing can be divided into two types: station wear, which is a single-layer garment worn while at the fire station, and turnout gear, which is a multiple-layer garment worn while responding to a fire incident [2]. The turnout gear is a combination of three fabric layers: an outer shell (OS), a moisture barrier (MB) and a thermal liner (TL).

The outer shell is the primary layer in a multi-layered clothing system and is part of a family of fabrics that are naturally resistant to fire [3, 4]. It has a highly organized structure of aromatic groups attached to a backbone chain, which prevents it from breaking down into smaller molecules and producing less smoke when exposed to fire [5, 6]. When this aramid fabric group is combined with other fire-resistant materials using lamination, coating or bonding, it creates a moisture barrier. This porosity of the membrane is a determining factor in moisture control, that is, it is permeable to sweat vapor and acts as a barrier for water droplets. To make a thermal liner, a batting fabric made from fire-resistant fibers is sewn or quilted onto a face fabric, which is also made using fire-resistant fibers. Its main function is to reduce heat dissipation, provide comfort and release metabolic heat away from the wearer’s body.

The maximum evaluation criterion for thermal protective clothing was proposed by Behnke [7]. He exposed swatches of the protective fabric made of Nomex®IIIA to several incident fluxes. The amount of incident flux that replicated the appearance from damaged clothing was estimated to be $84\mathrm{kW}/{\mathrm{m}}^2\left(\approx 2\mathrm{cal}/{\mathrm{cm}}^2\mathrm{s}\right)$ for 8 seconds of exposure. This estimated value was proposed to be a representation of flashover.

The conditions of a fire, whether it’s inside a building or in a forest, can vary from one incident to another due to the unpredictable nature of the materials that are burning [8]. That is why firefighters’ protective clothing is designed to be effective under a range of heat levels that are representative of several types of fires, including flashovers, backdrafts, fully developed fires, and forest fires. These garments are classified into routine, hazardous and emergency categories [9, 10]. Recent experiments [11, 12, 13, 14, 15, 16, 17, 18, 19] reported an average of 200 $\mathrm{kW}/{\mathrm{m}}^2$ heat flux under extreme conditions in buildings and other enclosed spaces. These are higher levels of fire severity than previously measured. Henceforth, a new upper limit of 126 $\mathrm{kW}/{\mathrm{m}}^2$ is recommended, and the performance of turnout gear is tested at this level.

It corelates with the building data collected by the end of the past decade which emphasize modern construction trends in the housing sector [20, 21]. According to this data, the home area has increased by approximately $232.3{\mathrm{m}}^2$ in $2008$, as compared to $144{\mathrm{m}}^2$ in $1973$. The 26% of homes built in 2008 were bigger than $278.7{\mathrm{m}}^2$; additionally, a growth of 56% in the construction of 2-storey homes (1-level above ground floor, US standard) was reported. Also, the current housing incorporates taller ceilings of 4.3–6.1 m, than the traditional 2.4 m. All these factors contribute to increased area and volume. As a result, the smoke spread during a fire incident remains above the ongoing fire; hence, more oxygen is available, causing the fire to grow rapidly and violently. This fire spread is also linked to the decorations and contents in the accommodation that have transitioned from natural to synthetic materials, such as cotton being replaced by polyurethane foam [22].

The effectiveness of a multi-layered protective garment can be evaluated using either a full-scale manikin test or a laboratory bench-scale test using the TPP rating system [23, 24, 25]. A full-scale manikin test assesses the performance of the entire garment against a convective/radiative thermal load, while a bench-scale test is performed on small patches of the garment. To capture thermal load on the skin, full-scale manikin testing facilities have been developed using a network of over 100 sensors distributed across the manikin’s surface [26, 27, 28, 29, 30]. However, due to the high cost of full-scale tests, bench-scale tests are preferred as they are more affordable and can still replicate real-world room fires. Standards for bench-scale tests have been developed for diverse types of thermal loads, such as convective, radiant or a mix of both. The assessment criteria such as ISO 9151 [31], ISO 6942 [32], ASTM F2702 [33], ISO 17492 [24], ASTM F2703 [34] and NFPA 1971 [35] vary depending on the heat source and are representative of several types of fire incidents. The data from these experiments are then input into analytical models to predict skin burns.

1.1 Skin burns

The skin layers of interest in burn degree evaluation are the epidermis, dermis and subcutaneous. The second layer, dermis, is of the most significance as its temperature profile corelates to second-degree burns. The human body maintains a skin temperature of 37°C during daily activities. Pain is induced when the body temperature increases by 8°C, which occurs at a temperature of 44°C. A first-degree (superficial) burn occurs at 48°C, and a second-degree burn occurs with continued exposure to heat at 55°C. At 72°C, the skin tissues cannot sustain thermal equilibrium and are destroyed. Two widely accepted criteria are applied to assess this damage, Stoll’s and Henrique’s.

Stoll’s criterion is applied to evaluate protection against second-degree burns [36]. It involves recording temperature data from a benchtop test, and the intersection of the experimental curve with Stoll’s documented experimental curve predicts tolerance time. This criterion assumes a rectangular heat pulse, and its validity depends on this assumption.

In contrast, Henrique’s is based on a chemical rate process and uses temperature and time to determine skin damage [37]. Skin-simulant sensors are used in modern thermal manikin experiments to predict third-degree burns using Henrique’s burn integral, Ω, as shown in Eq. 1. The values of physical constant have been presented in Table 1. While Stoll’s criterion is easy to use, Henrique integral has shown more accurate predictions for incident heat flux range not covered experimentally by Stoll. Henceforth, it is a preferred choice when assessing superficial burns at heat fluxes above 41 $\mathrm{kW}/{\mathrm{m}}^2$.

Constant	Basel layer	Dermal base	Limit
$\mathrm{\Omega }$	$\le 0.50$, no burn	$\cong 1$, third-degree	$—$
	$=0.53$, first-degree
	$\cong 1.00$, second-degree
P ($1/\text{second})$	$2.185\times {10}^{124}$	$4.32\times {10}^{64}$	$44\le T\le {50}^{°}C$
	$1.823\times {10}^{51}$	$9.39\times {10}^{104}$	$T\ge {50}^{°}C$
$\frac{∆\mathrm{E}}{\mathrm{R}}$(K)	93,534.9	50,000	$44\le T\le {50}^{°}C$
	39,109.8	80,000	$T\ge {50}^{°}C$

Table 1.

Henrique burn model physical constant values [8, 38].

$$\mathrm{\Omega }=\underset{0}{\overset{\mathrm{t}}{\int }}\text{Pexp}\left(-\frac{∆\mathrm{E}}{\mathrm{RT}}\right)\mathrm{dt}$$E1

1. $\mathrm{\Omega },$Henrique integral, second-degree burn occurs when $\Omega$ is unity

2. $∆\mathrm{E}$, the activation energy (J/mol)

3. P, a pre-exponential factor

4. T, time-dependent absolute temperature of the basal layer.

2. Experimental methodology

To conduct novel tests on the thermal protective garment subjected to a fire load of 126 $\mathrm{kW}/{\mathrm{m}}^2$, a purpose-built experimental apparatus was designed by cooperation between the Korean Conformity Laboratories (KCL) and Ulster University (UU), as shown in Figure 1. It comprises a radiant panel, a specimen assembly and a trolley. Two-layered halogen quartz tubes can output a consistent radiant flux of 126 $\mathrm{kW}/{\mathrm{m}}^2$ for more than 60 seconds. More details on the development, configuration and limitations of the test apparatus can be found in the preliminary study of KCL [39].

Figure 1.

Bench-scale apparatus.

The experimental setup consists of an outer shell (OS), a moisture barrier (MB), a thermal liner (TL) and a substrate. The addition of a substrate represents closed-back configuration with variations in the air gap between the skin and the innermost layer [40, 41]. The temperature measurement at the back surface of the thermal liner is vital. Therefore, four type-K thermocouples were prepared to record the thermal liner back surface temperature. Constant contact is needed between the thermocouples and thermal liner back surface, for which sewing of the thermocouple was first implemented, as suggested by the National Institute of Standard and Technology (NIST) [42]. However, it was difficult to keep the bead in contact with fabric; plus, sewing made it demanding to retain the stripped wire away from the fabric surface. As four thermocouples are required to be placed in proper contact with the back surface, sewing proved to be impractical. Additionally, the application of heat-resistant tape is not recommended [42]. To overcome this, a new mechanism was developed based on pressure contact.

Thermocouple beads were soldered to a copper sheet of circular cut with a thickness of 0.35 mm. Thermocouple wires ran across the slits on a rod made of high thermal-resistant plastic. It was then put through circular holes in the substrate. The distance between the copper sheet and the circular rod was kept to greater than 6 mm to ensure proper air circulation. In this manner, four probes were prepared and attached at four distinct points of the fabric back surface. This method of pressure contact can record the spatial variations of the temperature, and it has been validated [43].

Three types of fabric configurations, Type A, B and C as described in Figure 2, were examined at two different exposure levels of 84 and 126 $\mathrm{kW}/{\mathrm{m}}^2$, representative of flashover and post-flashover. Specimens were exposed to the incident radiant flux for three different exposure times, of 10, 20 and 25 seconds. Type-A configuration is of a conventional layup. It is commercially available in the market and, therefore, is treated as a benchmark to access the performance of Type B and C. A 2 mm (2A) air spacing is ensured between the fabric layers and 6.5 mm between TL and substrate as per research standards [29, 41, 42, 44].

Figure 2.

Turnout gear layup for experimental studies.

2.1 Turnout gear characteristics

The functional characteristics of each fabric layer in the turnout gear are distinctive. The OS resists heat; protects the wearer from chemical spills, blood and flames and, additionally, lets vapors from body sweat pass through. The MB also aids in heat resistance and controls moisture transport. The TL provides comfort and heat resistance and allows metabolic heat release to the environment. The underwear (UV) is worn by Formula One drivers for safety from fuel fires during an accident. The honeycomb (HC) is an insulation material utilized to access its suitability in conjunction with the turnout gear. The physical characteristics of each layer are stated in Table 2.

Outer shell (OS)	Material: Aramid fabric 100% (meta-aramid 80%, para-aramid 20%)	Thickness: 0.40 mm
	Weight: 209 g/${\mathrm{m}}^2$	Pattern: Ripstop
Moisture barrier (MB) (substrate + PTFE film)	Material: Meta-aramid 100%(Substrate) + PTFE film
	Weight: 190 g/${\mathrm{m}}^2$	Pattern: Plain weave
	Thickness: 0.347 mm
Felt insulation of thermal liner (TL)	Material: Meta aramid 100%
	Weight: 149 g/${\mathrm{m}}^2$
Face cloth of thermal liner (TL)	Material: Meta aramid 100% (Face cloth)
	Weight: 149 g/${\mathrm{m}}^2$	Pattern: Plain weave
	Thickness: 1.754 mm
Under wear (UW)	Material: Meta aramid 100%	Thickness: 0.61 mm
Honeycomb (HC)	Material: Meta aramid 100%	Thickness: 3 mm

Table 2.

Material properties.

2.2 Experimental results

A permanent skin damage will occur when the temperature of the basal layer exceeds 44°C after a certain exposure duration. Henceforth, to determine the exposure time, experiments are started with an initial guess of 10 seconds. The performance of different layups (Type A, B, and C) was evaluated based on the transmitted flux recorded by the heat flux gauge at the substrate. It was then converted to temperature values using ABAQUS® software. Type A was first exposed to an incident flux of 84 and 126 $\mathrm{kW}/{\mathrm{m}}^2$ for 10 seconds. If the basal layer temperature was estimated to be above 44°C, then types B and C were tested at that exposure time. If not, the exposure time was increased, and the experiment was repeated for type A until the basal layer temperature did not exceed 44°C.

Figure 3(a) displays the skin time-temperature histories for type A, B and C. The basal layer temperature for type A remains below the recommended value of 44°C at 10 and 20 seconds of exposure. However, under an exposure time of 25 seconds, the skin temperature reaches approximately 55°C, which exceeds the threshold value. For type B, the skin temperature remains lower than the threshold value, at around 42°C. In the case of type C, the skin temperature reaches ≈48°C. Based on these results, the three configurations can be ranked as follows: type B performs the best under extreme conditions, followed by type C and then type A.

Figure 3.

Skin basal layer temporal-temperature at two levels of heat flux.

Figure 3(b) illustrates the skin temperature-time history of type A, B and C for an incident flux of 126 $\mathrm{kW}/{\mathrm{m}}^2$. The type A configuration performed well for an exposure time of up to 15 seconds. However, it failed when the exposure time exceeded 15 seconds. Type B and C configurations, on the other hand, were able to maintain lower skin temperatures, 36% and 31% below the maximum estimated temperature of type A for the basal layer, respectively. In terms of the dermal base, type B performed 18% better and type C performed 14% better than type A. These percentages represent a reduction in skin temperature and an improved level of protection.

2.3 Auxiliar protection significance

The significance of auxiliary layers is estimated based on superficial burn injuries utilizing the Henrique integral [45]. Type B and type C provided more protection by prolonging the time to burn injuries. In the case of type A, no burn was predicted for an incident heat flux of 84 $\mathrm{kW}/{\mathrm{m}}^2$ and 126 $\mathrm{kW}/{\mathrm{m}}^2$ up to an exposure duration of 20 and 15 seconds, respectively. For type A, exposed to 84 $\mathrm{kW}/{\mathrm{m}}^2$ for 25 seconds, a first-degree burn was estimated after 35 seconds and a second-degree burn at 37 seconds. This close gap is associated with rapid increase in skin temperature due to outer shell failure.

Similarly, type A exposed to 126 $\mathrm{kW}/{\mathrm{m}}^2$ for 20 seconds, a second-degree burn was predicted at the 24th second and a third-degree burn at the 74th second. In the case of type B, no burn injuries are projected for 84 $\mathrm{kW}/{\mathrm{m}}^2$; except for an incident flux of 126 $\mathrm{kW}/{\mathrm{m}}^2$ for 20 seconds, a first-degree burn is predicted after 35 seconds. Type C underperformed type B with first-degree burn at the 55th second at 84 $\mathrm{kW}/{\mathrm{m}}^2$ incident flux for an exposure time of 25 seconds and a second-degree burn after 30 seconds for an incident flux of 126 $\mathrm{kW}/{\mathrm{m}}^2$ for 20 seconds of exposure time.

At life-threatening conditions of 126 $\mathrm{kW}/{\mathrm{m}}^2$, it is observed that a second-degree burn will ensue rapidly after a first-degree burn if the outer shell fabric undergoes thermal degradation. Since the Henrique integral is a chemical rate process, it is observed that the underlying tissue in the basal layer does not have enough time to react to extreme temperature variations, so skin tissues undergo chemical changes rapidly, causing second-degree burn immediately. Therefore, an auxiliary protective layer is recommended for firefighters involved in severe fire conditions.

Adding an extra layer to the existing protective suit has proven to mitigate the burn injuries, but care must be taken when the thermal comfort of firefighters is assessed. The use of the honeycomb structure between the moisture barrier and the thermal liner was proven to be beneficial. However, because of the difficulties in retaining its shape during use, it can only be applied to protect the front chest or the back area. It will restrict mobility. As a result, it is not recommended. In comparison, the meta-aramid fabric layer is comfortable and can easily be worn inside an existing suit. Based on this work, the application of type B configuration is recommended. However, care must be taken while implementing the result of this study to real-life conditions as the assessment was done under laboratory conditions.

Since the study of high thermal-resistant textiles is destructive in nature, numerical techniques are employed to study heat dissipation rates and evaluate other parameters such as moisture effect and superficial burns at variable heat flux levels. These numerical techniques are accurate and cost effective; henceforth, reliable information can be obtained if boundary conditions are rigorously evaluated and applied.

3. Boundary conditions

The protective assembly consists of multiple fabric layers, and the thermal exchange across each layer is affected by a combination of conduction, convection and radiation. Heat energy penetrates the system via thermal radiation. A portion of this is reflected towards the ambient as radiation $\left\{\mathrm{\sigma \epsilon }\left(T_{\mathrm{fab}}^4-T_{\mathrm{amb}}^4\right)\right\}$ and convection loss $\left\{\mathrm{h}\left(∆\mathrm{T}\right)\right\}$, while most of it is conducted $\left\{\mathrm{k}\frac{\mathrm{\partial T}}{\mathrm{\partial x}}\right\}$ through the fabric. In between fabric layers, the thermal energy is exchanged via net radiation $\left\{\frac{\mathit{F\sigma }}{\frac{1}{{\epsilon }_1}-\frac{1}{{\epsilon }_2}-1}\times ∆T^4\right\}$ and conduction/convection. External effects on these boundary conditions should be evaluated to predict heat dissipation rates. In addition, at high flux, smoke layers and boundary and cavity convention are the most important. Figure 4 illustrates thermal exchange between the fabric layers and with the environment [46].

Figure 4.

Thermal exchange in a multi-layered turnout gear.

3.1 Effect of smoke layer

Smoke in the path of radiative rays causes a reduction in the intensity of the radiation. An estimate is made to determine the amount of radiative energy absorbed by the smoke layer. If Nomex® burns completely, it produces heteropolar gases such as ${\mathit{CO}}_2$ and $H_{2}O$ and symmetric diatomic molecular gas $N_2$ as per the chemical reaction in Eq. (2). $N_2$ is transparent to electromagnetic waves in the infrared region and does not emit or absorb thermal radiation in its natural state. On the other hand, heteropolar gases absorb and emit radiation in their respective spectral thermal bands. The “effective emissivity” and “effective absorptivity” of heteropolar gases are evaluated using experimental observation charts developed by Hottel [47, 48]. Precise calculations of absorptivity are complicated, so engineering approximations are used [49].

$$C_{14}{H}_{10}{N}_2{O}_2+15.5O_2+58.25N_2\to 14{\mathit{CO}}_2+5H_{2}O+60.27N_2$$E2

The volume fraction of ${\mathit{CO}}_2$ in combustion products is estimated to be 0.18 and that of $H_{2}O$ to be 0.06. These volume fractions are based on complete combustion, which is an engineering approximation and would vary at each experimental execution. Therefore, to account for such variations in an ongoing analysis, an optimistic approach was applied, and volume fractions of 0.3 for ${\mathit{CO}}_2$ and 0.1 for $H_{2}O$ were considered [50]. An atmospheric pressure of 1 atm was also assumed for the estimation of partial pressures $\left(P\right)$. Absorptivity is estimated based on a relation [48] as below,

$${\alpha }_i={\phi }_i\times {\epsilon }_i\left(P_{\mathrm{i}}.L,T_g,1\mathit{atm}\right)\times {\frac{T_g}{T_s}}^{\varnothing }$$E3

where, subscript i refers to the participating gas, $L=2\times ∆x$ is the mean beam length$,{\phi }_i$ is the pressure correction factor, $T_g$ is the gas temperature$,T_s$ is the temperature source emitting radiation normal to the radiative surface and $\varnothing$ is a constant value [47], 0.65 for ${\mathit{CO}}_2$ and 0.45 $H_{2}O$. With the assumption of 1 atmospheric pressure, the correction factor, ${\phi }_i$, reduces to unity. Overlapping waveband correction $∆\epsilon$ for emissivity estimation of a mixture for a maximum gas thickness of $∆x$ = 0.02 m (assumed from experimental observations) was evaluated. If $∆\epsilon$ is found to be of less significance for this case, it is disregarded in extended calculations. ${\epsilon }_{\mathit{mix}}$ was estimated as [51],

$${\epsilon }_{\mathit{mix}}={\phi }_{{\mathit{CO}}_2}.{\epsilon }_{{\mathit{CO}}_2}\left(P_{{\mathit{CO}}_2}.L,T_g\right)+{\phi }_{H_{2}O}.{\epsilon }_{H_{2}O}\left(P_{H_{2}O}.L,T_g\right)-∆\epsilon$$E4

$∆\epsilon$ is evaluated using volume fractions of ${\mathit{CO}}_2$ and $H_{2}O$ as,

$$\begin{array}{c}\frac{P_{H_{2}O}}{P_{{\mathit{CO}}_2}+P_{H_{2}O}}=0.25\\ P_{{\mathit{CO}}_2}.L+P_{{\mathrm{H}}_2\mathrm{O}}.L\approx 0.016\ll 1\end{array}$$E5

From Hottel’s [47, 48] correction charts $∆\epsilon \approx 0$, thus, Eq. (3) gives,

$${\epsilon }_i\left(P_i.L\right)=\underset{0}{\overset{∆x}{\int }}{P}_i\times Ld\left(∆x\right)$$E6

The interference of the smoke layer in front of the outer shell under intense exposure to a radiant heat source is dependent on gas layer thickness $∆x$ and source temperature $T_s$. The gaseous particles attenuate radiative flux passing through, independent of the gas type. The particles of carbon dioxide exhibit stable emissivity, ${\epsilon }_{{\mathit{CO}}_2}$, irrespective of the source temperature, though they have an ascending relationship with increasing smoke layer thickness, $∆x$. The absorptivity, ${\alpha }_{{\mathit{CO}}_2}$, of carbon dioxide demonstrated a descending relationship to source temperature with constant thickness and ascending relationship with thickness, $∆x$, of the smoke layer. These relations for carbon dioxide are presented in Figure 5. The water vapor’s emissivity, ${\epsilon }_{{\mathrm{H}}_2\mathrm{O}}$, decreases as the source temperature increases with a constant water vapor layer thickness; however, higher emissivity is observed with increased water vapor layer thickness, $∆x$. A similar trend is observed in the case of water vapor absorptivity, ${\alpha }_{{\mathrm{H}}_2\mathrm{O}}$, as evident from Figure 6.

Figure 5.

Carbon-dioxide gas emissivity and absorptivity at variable source temperatures.

Figure 6.

Water vapor gas emissivity and absorptivity at variable source temperatures.

During pyrolysis, the volume fraction of carbon dioxide is significant compared to water vapor in standard humid conditions, so in theory, its effect on irradiance at the outer shell boundary is more important than water vapor. The expected interference of carbon dioxide particles due to absorption or emission of incident radiative flux is in the range of $0.03–0.055$, estimated from Figure 5, and that of water vapor is in the range of $0.015–0.035$, projected from Figure 6. When the heat source was radiating at flux densities of magnitude greater than $41\mathrm{kW}/{\mathrm{m}}^2,$ the combined effect of heteropolar gases would attenuate the maximum radiative flux by $8.5\%$ and the minimum by $4.5\%$. An insignificant portion of incident flux density can be safely ignored for the rest of the analysis.

In conclusion, during pyrolysis, smoke layers up to $20\mathrm{mm}$ in thickness have little to no interference to irradiance at the outer shell boundary layer for extreme fire conditions.

3.2 Air cavity

The thermal exchange through the air cavity in between fabric layers dictates the dominant mode of heat transfer, that is, convective or conductive. The cavity thickness, ${\delta }_s$, as per the research prospective is 1–6.35 mm. Based on relation ${\mathit{Ra}}_{{\delta }_s}=\frac{\mathit{g\beta }\left(∆T\right){\delta }_s}{v^2}\times \mathit{Pr}$, for varying separation distance ${\mathrm{\delta }}_{\mathrm{s}}=2,6.35,10,20\text{and}50\mathrm{mm}$, where $∆T$ is the temperature difference at bounding layers of the air cavity, a dominant mode of heat exchange is estimated. The slave boundary, $T_{\mathit{\text{cold}}}$, is assumed to be at room temperature of 25°C. It is estimated that when cavity thickness is below 6.35 mm, the interface media is purely conductive. Figure 7 can be utilized to estimate flow behavior at other varying thicknesses.

Figure 7.

Flow characteristics at varying cavity thicknesses.

3.3 Thermal decomposition

The energy released during thermochemical reactions and mass loss rate at the specific temperature ranges were established for each fabric layer, utilizing thermogravimetric analysis (TGA) and High Differential Scanning Calorimetry (HDSC). Linseis model STA PT1000 [52] is utilized. The purpose of TGA and HDSC is to estimate density change and specific heat of the fabric layers at elevated temperatures. During simulation, the algorithm calls in temperature-specific user-defined material properties for each layer, depending on the solution phase (heating or cooling).

From Figure 8, for outer shell, the density is estimated to change linearly over time from 522.5 to 458 $\mathrm{kg}/{\mathrm{m}}^3,$ valid in a temperature range of 25–125°C, due to evaporation of moisture from the fabric. From 150 to 480°C, it is 442–380 $\mathrm{kg}/{\mathrm{m}}^3,$ and from 480 to 720°C, it is 380–27 $\mathrm{kg}/{\mathrm{m}}^3$. For moisture barrier, the temperature ranges as per density change are evaluated as 550.7–522.5 $\mathrm{kg}/{\mathrm{m}}^3$ between 25 and 125°C and 522.5–467 $\mathrm{kg}/{\mathrm{m}}^3$ between 250 and 418°C. In the decomposition temperature range of 418–719°C, it is estimated to be 467–27.5 $\mathrm{kg}/{\mathrm{m}}^3$, linearly. For thermal liner, the density change is estimated to be (−) 15%, as 170–144 $\mathrm{kg}/{\mathrm{m}}^3$ between 25 and 432°C. For Nomex® under layer, the density is taken as a constant value of 138 $\mathrm{kg}/{\mathrm{m}}^3$ as the temperature is not expected to reach above 300°C.

Figure 8.

Thermal gravimetric analysis (TGA) of the turnout gear fabric layers.

The concept of apparent specific heat requires the knowledge of temperature ranges and respective specific heat values at those limits. The HDSC analysis for the outer shell is not performed, as the composition is of Nomex®IIIA, and extensive literature is available on the apparent specific heat of it.

In case of the outer shell, the latent energy released due to moisture evaporation between 75 and 125°C and the thermochemical reaction energy released during decomposition between 420 and 700°C is added to the initial value of specific heat during the heating phase. On cooling, the normal specific heat path is followed. Similarly, the density is varied as per mass loss analysis performed on TGA and its implementation to account for decomposition. Henceforth, detailed material properties are shown in Table 3. Skin thermal properties have been documented in Ghazy and Bergstrom [53].

Characteristic	Outer shell	Moisture barrier	Thermal liner	Underlayer
Thickness (m)	0.4 × ${10}^{-3}$	0.34 × ${10}^{-3}$	1.75 × ${10}^{-3}$	0.6 × ${10}^{-3}$
*Density (kg/m^3)	521–380	170–144	550–27	138–110
*Conductivity (W/m °C)	0.047–0.255	0.047–0.255	0.036	0.047–0.255
*Specific heat (J/kg °C)	1300–2515	1300–2500	1088–2300	1300
Emissivity	0.9	0.9	0.9	0.9

Table 3.

Fabric material thermal properties.

^*

The thermal properties are dependent on temperature ranges and thermal phase as per TGA and HDSC.

4. Numerical model

Numerical models and their validity have been well documented for the prediction of heat dissipation in a single or multi-layered turnout gear for bench-scale studies [42, 49, 50, 54, 55, 56, 57, 58]. Figure 9 represents thermal exchange in between fabric layers in a bench-scale setting. Henceforth, an overview is presented in this section to understand basic mechanics of thermal exchange.

Figure 9.

Thermal exchange between fabric layers in turnout gear.

An energy balance for pure heat exchange between multiple fabric layers can be modeled as per,

$$\rho C_p\frac{\partial T}{\partial t}=\frac{\partial }{\partial x}\left(k\frac{\partial T}{\partial x}\right)+{\omega }_b{\left.\left(\rho C_p\right)\right|}_b\left(T_a-T\right)+Q_m$$E7

Conduction across air gap is treated as per Eq. (8), where $k_{\mathit{eff}}$ is the ratio of thermal conductivity of air and its thickness t.

$$k\frac{\partial T}{\partial x}=k_{\mathit{eff}}\frac{\partial T}{\partial x}=\frac{k_{\mathit{air}}}{t_{\mathit{air}}}\frac{\partial T}{\partial x}$$E8

In between fabric layers, the amount of radiation being transmitted is treated as $q_{\mathit{rad}}=C\left(T_1^4-T_2^4\right)$ and is dependent on view factor F and emissivity $\epsilon$, for radiation constant C. Where $\mathrm{C}=\frac{\mathit{F\sigma }}{\frac{1}{{\epsilon }_1}-\frac{1}{{\epsilon }_2}-1}$, ${\omega }_b$ is the blood perfusion rate, $\mathit{kg}/m^{3}s$, ${\left.\left(\rho C_p\right)\right|}_b$ is the volumetric heat capacity (blood), $J/m^3.{}^{°}\mathrm{C}$, $T_a$ is the arterial temperature/core body, and $Q_m$ is the metabolic heat generation/tissue heat generation rate, $\mathrm{W}/{\mathrm{M}}^3$. Torvi rigorously tested this model and found that the profusion term {${\omega }_b{\left.\left(\rho C_p\right)\right|}_b\left({\theta }_a-\theta \right)$} had no impact on thermal conductivity of skin in different states [59]. Correlating with Lipkin and Hardy, it takes 20 seconds for the skin to react by increasing blood flow and thus the minimal effect of heat diffusion through it [60]. The emissivity of 0.94 has little effect on skin temperature [61]. Moisture evaporation and carbonization of the skin ensue second- and third-degree burns and can be ignored in the study of burn predictions at high exposures levels.

For thermal manikin, it is necessary to model the incompressible Navier-Stokes equations to simulate large air cavities of more than 10 mm. Limited studies [30, 62, 63] have been done on this as the analysis is time costly with large data points to process. Commercially available software such as ABAQUS® package can be utilized to achieve this task, where air is treated as an incompressible Newtonian fluid. The continuity, momentum and energy equations [64] are all used in this scenario. Detailed study on numerical models have been presented in a recent study [46].

ABAQUS® is employed to model time-dependent thermal behavior. The practical implications influencing numerical solutions are variant. Each layer is distinctive, materially and structurally. The outer shell has a different thickness, texture and material characteristics as opposed to the moisture barrier and thermal liner. The garment interacts with the outside environment and wearer’s skin via thermal conduction, convection and radiation. Therefore, moisture significance, burn degree and survival time are studied utilizing numerical approximations based on boundary conditions defined in Section 3.

5. Discussion

5.1 Moisture effect

Moisture can accumulate in firefighter garments during active use, either from an outside source during operation or through perspiration. The latent heat of vaporization, $∆h_{\mathit{er}1}$, is dependent on the mass fraction of moisture, $\varphi ,$present within the outer shell. The initial moisture mass fraction, $\varphi ,$ of 0.05 (5%) [65] is varied to 0.5 (50%), representing partial saturation of the outer shell. The safety offered by convectional protective garment assembly (type A) is then estimated at radiant flux levels of 41, 84 and 126 $\mathrm{kW}/{\mathrm{m}}^2$ for an exposure duration of 25, 25 and 20 seconds as shown in Figure 10. The effect of moisture permeation on burn degree with each successive increment of heat flux intensity is presented. The red markers indicate an initial mass fraction of moisture of fabric, that is, 5%, due to a relative humidity of 65%. At a radiant flux of 41 $\mathrm{kW}/{\mathrm{m}}^2$, no burn damage is predicted. However, the temperature of the skin escalates to the threshold level of pain when the moisture content is 5% and 10%, transcribed as “time to pain.”

Figure 10.

Moisture effect on first-degree burns.

5.2 Superficial burns

Burn degree damage for the basal layer and the dermal base is estimated as shown in Figure 11. Subjected to an incident heat flux of 84 $\mathrm{kW}/{\mathrm{m}}^2$ for 25 seconds (Model 1), first-degree burn is observed after 36 seconds, and a second-degree burn is detected after 38 seconds. This minor difference between first and second-degree burns indicates physical failure, also evident with a sharp rise in burn integral Ω. Also, skin temperature T is an exponential function of Ω; as a result, a second-degree burn ensues from a first-degree burn in close approximation. Under exposure of 126 $\mathrm{kW}/{\mathrm{m}}^2$ for 15 seconds (Model 2), no superficial degree burn is estimated by Henrique integral evaluation. Under radiant exposures, by the current estimate, a third-degree is highly unlikely, as no third-degree burns for models 1, 2 and 3 are predicted. It is apparent from the present study that garment assembly with an added protective layer protected the wearer from all types of burn injuries and increased thermal efficiency by absorbing additional thermal energy within the auxiliary layer.

Figure 11.

Approximation of superficial burn degree.

5.3 Survival curve

Radiant exposure survival curves are proposed for a heat source ranging from 2.5 to 126 $\mathrm{kW}/{\mathrm{m}}^2$. They represent the modern thermal environment. These curves would be beneficial to estimate the survival time for firefighters wearing an existing multi-layered protective suit. The tolerance time of a firefighter wearing special protective clothing varies depending on the subjected thermal environment. It is argued that those conditions would only occur under circumstances where a firefighter is involuntarily centred due to falling off a burning ceiling or roof [66]. A time-temperature curve was presented in Stoll’s work by accessing the damage to naked skin [36]. This curve represented superficial burn damage, and in concurrence with the fabric test, tolerance time from exposure was established by Stoll. Limitations of Stoll’s tolerance exposure curve are,

1. Rectangular heat fluxes ranging from 4.2 to 16.8 $\mathrm{kW}/{\mathrm{m}}^2$ were used to establish this criterion, and extrapolation techniques are required for exposure level not considered by Stoll.

2. Assessment of fabric was strictly dependent on the assumption of rectangular heat pulse/source; discrepancy from this shape invalidates the data.

3. A feasibility study on Stoll’s curves indicates that it may fail when an extra layer of protection is added, that is, from single fabric to multiple fabrics [67].

The proposed radiant exposure curve addressed these limitations. It utilizes Henrique integral [68] for superficial burn degree evaluation at each successive increment in exposure level. The linear incremental exposure levels selected for this study are “2.5, 10, 20, 41, 60, 84, 100, 115, 126” $\mathrm{kW}/{\mathrm{m}}^2$. An incident radiant heat flux of 2.5 $\mathrm{kW}/{\mathrm{m}}^2$ represents routine working conditions; a value of 10 $\mathrm{kW}/{\mathrm{m}}^2$ represents hazard working conditions; above 10 $\mathrm{kW}/{\mathrm{m}}^2,$ they are categorised as emergency working conditions [10, 69, 70]. An incident heat flux of 20 $\mathrm{kW}/{\mathrm{m}}^2$ signifies floor level heat flux for flashover fires [71, 72, 73, 74]. The values of 41 and 84 $\mathrm{kW}/{\mathrm{m}}^2$ represent industrial flashfire lower and average heat flux values [7]. They are recommended by international standards for performance evaluation [24, 32, 35]. The value of 126 $\mathrm{kW}/{\mathrm{m}}^2$ is the newly proposed limit for performance evaluation from the current study.

The established radiant exposure curves are plotted in Figure 12. The black squared markers represent first-degree burn limit, the red solid line to represent second-degree burn limit, and black circle markers show third-degree burn limit. It indicates a time to burn damage under several incident heat fluxes for a convectional garment layup.

Figure 12.

Radiant exposure survival curve.

It is observed from Figure 12 that the time to skin burn injuries is an exponential function of time, with a linear increase in incident heat flux; the time to burn injuries will proceed at a 100 times faster rate. This behavior is analogous to the rate of injury reported for the naked skin [61]. Effective use of this graph is an extension of Stoll’s curve with its limitations addressed. The information extracted from the proposed curves are,

1. A first-degree burn under exposure of current flashfire conditions of 84 $\mathrm{kW}/{\mathrm{m}}^2$ is estimated at 33.6 seconds. For a newly proposed limit of 126 $\mathrm{kW}/{\mathrm{m}}^2,$ it is approximated at 27.6 seconds, a 6-second difference.

2. Similarly, a second-degree burn for 84 $\mathrm{kW}/{\mathrm{m}}^2$ and 126 $\mathrm{kW}/{\mathrm{m}}^2$ is estimated at 34.4 and 28.5 seconds, respectively.

3. Irreversible damage to the skin will occur when the exposure duration exceeds 56.2 seconds for 84 $\mathrm{kW}/{\mathrm{m}}^2$ and 47.9 seconds for 126 $\mathrm{kW}/{\mathrm{m}}^2$.

In practical conditions, it is unfeasible for firefighters to gauge their skin temperature without the aid of a thermal sensor. However, the severity of fire can be assessed before engaging it. During operation, the initial indicator of skin burn is the sensation of pain, felt when skin temperature reaches 44${}^{°}\mathrm{C}$. The exponential nature of burn damage in skin tissue demonstrates that once the pain is felt, superficial burn injuries will ensue shortly, and in extreme conditions, it will be within seconds. To escape from such a situation without harm, it is important to have knowledge of “survival time.” The survival time is defined as a point between the pain threshold and the second-degree curve (blister curve) [36]. In a physical situation, it indicates when pain is felt but blister is not produced. In the current study, this point is determined by considering the first-degree curve as a limitation. Hence, a survival time is estimated by subtracting the time when pain is felt from when a first-degree burn is observed. In Table 4, the survival time for a linear increase in incident heat flux for a multi-layered garment is presented.

Incident flux $\left(\mathrm{kW}/{\mathrm{m}}^2\right)$	**Time to pain (sec)	*Time to blistering (second-degree burn) (sec)	***Survival time (sec)
2.5	200.0	435.4	203.2
10	89.3	122.8	28.2
20	57.8	76.8	16.4
41	39.2	50.3	9.7
60	32.4	40.9	7.5
84	27.6	34.4	6.0
100	25.4	31.6	5.5
115	23.9	29.7	5.1
126	23.0	28.5	4.9

Table 4.

Survival time estimate under extreme radiant exposure.

^*

Time to second-degree burn from Henrique integral and blistering defined by ASTM C 1055-99.

^**

Time skin temperature reached a value of 44°C defined by ASTM C 1055.

^***

Survival time defined by Stoll and Chianta [36].

5.4 Fire-resistant fabrics and sustainability

Textile fabrics, irrespective of their application, will ignite and burn. Fire-resistant fabrics will burn either via flaming combustion (flame contact) or via smoldering (radiant source). Henceforth, protective fabrics are treated with various flame retardants to survive both types of burn. Flame retardants can be categorized into two types: additive and reactive. Additive flame retardants are incorporated into a material or product without forming any bonds or chemical reactions with the product. On the other hand, reactive flame retardants undergo a chemical reaction with the raw materials during the manufacturing process of the final material or product intended for commercial use [75].

In terms of environmental impact, the toxicity of smoke generated due to the burning of additives has implications on its end user, that is, firefighters or military personnel. The different routes of exposure, such as dermal absorption, ingestion (oral) and inhalation, play a role in determining the level of risk posed by a hazardous substance. US National Research Council Committee on Toxicology summarized 16 key flame retardants or flame-retardant classes for textiles by surveying industries making and/or using flame retardants [76]. They have been detailed in [75], with their relative hazard index concluding that it is less than 1 for all types of exposure routes to public or professional.

Inherently fire-resistant fibers are manufactured by chemically bounding flame retardants to the fiber-forming polymeric molecules or as an additive in the structure. Henceforth, the potential toxicity is limited during manufacturing and during disposal at the end-of-life cycle. Below are such fibers with their sustainability towards the environment [77],

Polyester: The manufacturer of these products claims that they possess environmentally friendly characteristics. During their manufacturing process, no emissions or waste products are generated, and they are cost-effective to clean. Trevira fiber and filament yarns have the potential to meet the criteria for the Oeko-Tex 100 Certificate/eco-label, which signifies that they do not contain any harmful substances.

Modacrylics: They have been in existence for over half a century and are categorized as polymers capable of forming fibers. These polymers are derived from resins that consist of copolymers of acrylonitrile. Acrylonitrile units make up a minimum of 35% but less than 85% of the resin’s weight (±CH₂CH[CN]±), while another comonomer like vinyl chloride, vinylidene chloride or vinyl bromide is also present. Vinylidene chloride is a comonomer, and to enhance flame resistance, antimony trioxide may be incorporated into the resin before it undergoes wet or dry spinning. Alternatively, vinyl bromide can be utilized as a comonomer. At present, there are only a limited number of commercially available modacrylic fibers, with Kanecaron (manufactured by Kaneka Corporation) being a notable example, which is based on vinylidene chloride/antimony oxide. However, concerns have been raised regarding the presence of both halogen and antimony in these fibers, casting doubt on their future economic viability.

Polypropylene: Prior to the melt spinning of polypropylene fiber, additives, including both halogen and non-halogen types, are introduced into the molten mixture. Examples of these additives include Sandoflam 5072 (manufactured by Clariant) and Ciba® Flamestab® NOR™ 116.

Aromatic polyamide: Two prominent commercial examples of para-aramids and meta-aramids are manufactured by DuPont as Kevlar® and Nomex®. The primary environmental concerns associated with these materials arise from the challenges involved in disposing of both manufacturing waste and end-of-life waste. However, their high value encourages recycling efforts. During the manufacturing process, the potential toxicity of intermediates has also posed a significant challenge. Specifically, Kevlar® is synthesized using the monomers 1,4-phenylene-diamine (also known as para-phenylenediamine) and terephthaloyl chloride through a condensation reaction that produces hydrochloric acid as a by-product. This reaction results in liquid-crystalline behavior and a mechanical drawing, causing the polymer chains to align along the fiber’s direction. Initially, hexamethylphosphoramide (HMPA) was used as the solvent for polymerization. However, toxicology testing revealed that HMPA caused tumors in the noses of rats. As a result, DuPont discontinued the use of HMPA and switched to a different solvent, which is a mixture of N-methyl-pyrrolidone and calcium chloride.

Polymelamine fiber: Basofil® (manufactured by BASF) is a heat and flame-resistant fiber that utilizes patented melamine technology. According to the manufacturer, Basofil® fiber complies with all environmental regulations pertaining to its processing and use. It possesses a Limiting Oxygen Index (LOI) value of 32, has low thermal conductivity, exhibits excellent heat dimensional stability and does not shrink, melt, or drip when exposed to flames. Although Basofil® fiber is relatively weak, it can be processed using standard textile machinery to create woven, knit and nonwoven fabrics. It is also suitable for producing batting or fabric fire barriers for mattresses and upholstered furniture. Like aramids, the disposal of Basofil® fiber after use presents an environmental challenge.

Non-toxic phosphorus fire resistant (FR) solutions such as EDA-DOPO (ED) and DOPO-PEPA (DP) were applied to polyethylene terephthalate (PET), replacing halogenated FR solutions [78]. The DOPO-based FR solution passed the vertical burning test, making it a suitable candidate for future application. Other studies have focused on developing carbon fiber-reinforced composites engineered with renewable compounds [79]. These bio-based organic compounds are combined with a renewable monomer, the triglycidyl ether of phloroglucinol (TGPh), with hexahydro-4-methylphthalic anhydride (HMPA), resulting in exceptional intrinsic flame resistance. Bio-composite applications to produce eco-friendly FR with renewable materials have proven to have a significant impact in lowering environmental footprint. Such impact has been discussed in a review study published by Madyaratri et al. [80]. Several recent studies have also proven this concept such as for cotton fabric [81, 82], aerogels linked with biobased cationic amylopectin derivative for extreme temperature conditions [83, 84] and glass fiber composite [85].

The utilization of flame retardants offers advantages to both society and the environment. While it is true that not all flame retardants, past or future, can be deemed completely safe from all perspectives, the application of proper scientific knowledge and regulatory measures can effectively prevent the distribution of hazardous substances to consumers. Nonetheless, it is crucial to recognize that flame retardants, as a group, significantly enhance fire safety by reducing the likelihood of ignition, minimizing heat release and decreasing the levels of smoke, combustion by-products and harmful environmental toxins.

6. Conclusion

The chapter highlights the significance of fire impact on the individual such as firefighters or military personnel. Emerging new evidence suggests that there is a need to redefine the upper limit of heat flux, which is $84\mathrm{kW}/{\mathrm{m}}^2$, which has been the same for the last four decades. Evident from modern construction trends and experimental studies on compartment fires in the past two decades reported an increasing heat flux trend with an average of $150\mathrm{kW}/{\mathrm{m}}^2$. Presently, the convectional turnout gear does not offer adequate thermal protection at 126 $\mathrm{kW}/{\mathrm{m}}^2$; henceforth, auxiliary protection measures and a realistic increase of $126\mathrm{kW}/{\mathrm{m}}^2$ in the upper limit for testing is recommended. Fire-resistant fabric implications on sustainability in terms of toxicity and end-of-life cycle are a challenging research area. A balance is needed to meet the demand for functionality along with the protection offered by flame retardants. Nevertheless, biobased composites and renewable monomers are being researched to replace conventional halogenated and toxic flame retardants.