Performance Assessment of Passive Fire Protection Materials

ARTICLE pubs.acs.org/IECR

Performance Assessment of Passive Fire Protection Materials Mercedes Gomez-Mares,† Alessandro Tugnoli,† Gabriele Landucci,‡ and Valerio Cozzani*,† †

Dipartimento di Ingegneria Chimica, Mineraria e delle Tecnologie Ambientali, Alma Mater Studiorum, Universita di Bologna, via Terracini 28, 40131 Bologna, Italy ‡ Dipartimento di Ingegneria Chimica, Chimica Industriale e Scienza dei Materiali, Universita di Pisa, via Diotisalvi 2, 56126 Pisa, Italy

bS Supporting Information ABSTRACT: The performance of fireproofing materials in providing effective protection from fire strongly depends on the thermophysical properties and on the behavior of the material during fire exposure. Not only active insulators but also nonactive coatings may undergo significant changes in their structure and properties when exposed to high temperatures. The present study focused on the measurement of some key properties for a set of three reference fireproofing materials of different nature. The changes in morphology and in the physical properties during fire exposure were investigated and they were dramatic for the case of active fireproofing materials. For these materials, the time required to reach a steady-state condition in the heat transfer may be significant. The use of a simple heat-transfer model, based on the experimental data obtained for the reference materials studied, demonstrated the importance of accounting for changes in the physical properties, paving the way to further applications in advanced studies of the fireproofing performance in complex geometries and critical scenarios.

’ INTRODUCTION Passive fire protection (PFP) can be defined as a coating, cladding, or free-standing system that provides thermal protection in the event of a fire and that requires any initiation mean, replenishment, or sustenance.1 Fireproofing materials are insulators considered to be a type of PFP. The protection performance of fireproofing materials strongly depends on the thermophysical properties, the structure, and the modifications that such materials undergo during fire exposure. The American Petroleum Institute (API) divides fireproofing materials into active and inactive insulators.2 Active insulators undergo a significant physical or chemical change when exposed to heat (e.g., epoxy intumescent coatings). In the case of such materials, a wide change of the physical and chemical properties during fire exposure is expected. Inactive fireproofing materials include coatings such as cement-based materials and man-made fibers, which are not inherently designed for modification during fire exposure. Nevertheless, all materials actually undergo changes in the insulation properties during exposure to fire. This, in turn, may significantly affect the performance of the PFP and should be carefully considered during rating and design. Current practice in the design of a fireproofing system is based on the fire resistance rating according to standard fire tests, such as UL 1709,3 ASTM E 119,4 ASTM E 1529,5 OTI 95635,6 and ISO 22899-1,7 among others. In such tests, samples are exposed to heat loads according to standardized timetemperature curves and/or heat fluxes depending on the type of reference fire. The results are usually expressed in time periods of resistance, and a given rating is obtained if in the tests the average and/or peak temperature of the material is kept lower than a given threshold value. Although the approach based on the tests and fire resistance rating is applicable in many practical cases, ratings should be used r 2011 American Chemical Society

with judgment because tests are carried out only for specific fire intensities and geometries.8 Advanced analysis of the fireproofing performance in safety critical systems or in complex geometries would require dedicated studies.912 While real-scale tests may be impractical and expensive, model-based simulation can provide effective answers to safety issues concerning the performance of fire-exposed structures and the effectiveness of fireproofing systems.1315 However, these simulations require experimental data and models to properly account for the properties and behavior of the material during exposure to fire, not provided by the standard fire tests. In the present study, an approach is developed to collect and model the key physical properties crucial for PFP materials performance. The approach is demonstrated for a reference set of fireproofing materials. Results from the experimental assessment of the fundamental properties of the selected materials are presented. Data on the thermal and physical properties were obtained at laboratory scale by the integration of several experimental techniques: thermogravimetric (TG) analysis, differential scanning calorimetry (DSC), thermal conductivity analysis by a transient plane heat source (hot disk), and furnace tests. The main regions where sample modifications occurred were identified. Semiempirical models were proposed for the description of key physical properties, such as thermal conductivity, density, and porosity. The thermal effects of the degradation reactions were also assessed. Where appropriate, simplified apparent Special Issue: Russo Issue Received: August 20, 2011 Accepted: October 10, 2011 Revised: October 5, 2011 Published: October 10, 2011 7679

dx.doi.org/10.1021/ie201867v | Ind. Eng. Chem. Res. 2012, 51, 7679–7689

Industrial & Engineering Chemistry Research

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kinetic models were introduced to describe modification of the material properties during fire exposure. The results obtained were applied to the numerical simulation of medium-scale experimental tests reproducing conditions of fire engulfment. Enhancement of the model performance due to an improved knowledge of the physical properties was assessed. In particular, the importance of introducing variable properties into simulations clearly emerged as an element leading to more accurate and more conservative performance predictions of the PFP materials. The results were discussed in light of advancement in the finite element method (FEM) simulation of the real-scale PFP performance.

’ EXPERIMENTAL SECTION Materials. Three commercial fireproofing materials were selected as a representative reference set for the present study: an epoxy intumescent coating, a lightweight cement-based coating, and a silica blanket. Each of the three materials represents one of the more important categories of PFP materials available, as described by API 2218 concerning fireproofing practices in petroleum and petrochemical plants.16 The epoxy intumescent coating used in the present study is a commercial material, supplied by International Protective Coatings (Akzo Nobel). The main components of the coating are an epoxy resin 2,20 -(4-butylidene)bisphenyleneoxymethylene (DGEBA), ammonium polyphosphate (APP), boric acid, an amine as a curing agent, and fillers (e.g., limestone, magnesium silicates, glass fibers, etc.). The coating is usually spray-applied directly onto the metal surface that has to be protected. The specimens used in the current study were prepared by a conventional procedure but without application to a metal support. Thus, boards of the sole material were obtained, having an average thickness of about 9 mm. The reference lightweight cement material used in experimental runs is a commercial conglomerate constituted principally of Portland cement, vermiculite (a form of mica), and other inorganic fillers, supplied by Cafco. Its density is lower than the one of common cement. The material is usually sprayed on the surfaces to be protected. Also, in this case, nonsupported boards having an average thickness of about 20 mm (typical of several industrial applications) were used. A silica blanket (supplied by Insulcon) was chosen as representative of fireproofing materials based on man-made fibers. The silica glass fibers used were mainly composed of silicon dioxide, aluminum dioxide (97%), and some binders as minor components. Specimens used in the current study were sections cut from commercial rolls, having a nominal thickness of 12 mm. Experimental Techniques. A TGA Q500 thermogravimetric analyzer from TA Instruments was used to study the degradation of the materials as a function of the temperature and heating rate. Constant heating rate runs were used, with heating rates of 5, 10, 25, and 45 C/min. A final temperature of 800 C was selected for the epoxy intumescent material, while final temperatures of 850 and 900 C were used for the vermiculitecement coating and silica blanket, respectively. A purge gas flow of pure nitrogen (100 mL/min) was used. Because the purpose of this paper is to analyze the material property behavior when the material is engulfed by fire, the use of an inert atmosphere was considered because the fire consumes the oxygen available to carry out the combustion, leading to an atmosphere poor in oxygen to be in contact with the analyzed material.

Figure 1. Simplified scheme of the radiant panel setup used in mediumscale fire tests: (a) radiant panel; (b) ignition flame; (c) thermocouple; (d) sample board; (e) IR camera; (f) pyrometer.

TGDSCFTIR tests were performed using a Netzsch STA 409/C thermoanalyzer coupled with a Bruker Equinox 55 spectrometer. The system provided data on the thermal effects and on the volatile substances released during constant heating rate (10 C/min) runs in pure nitrogen (60 mL/min). Further information on the equipment and on the experimental procedures adopted is reported elsewhere (Marsanich et al. 17). A fixed-bed tubular reactor was used to study the behavior of the coatings on a larger scale. A Carbolite HST 12/300 furnace equipped with a W301 controller was used to provide heat to the reactor. The samples were heated using a constant heating rate until a selected final temperature was reached. The tests were carried out under a nitrogen flow, and the samples were weighted and measured before and after each test. Specimens from the furnace were analyzed by several techniques. A scanning electron microscope model JEOL JSM 5600 LV was used for morphological analysis of the materials. A transient plane source instrument (TPS2500S, by Hot Disk AB) was used for measurement of the thermal conductivity. Medium-scale fire tests were obtained by a specifically modified ASTM E162 setup.18 Boards of the sample material (460 150 mm) were vertically exposed to a 304.8 457.2 mm porous gas-operated radiant panel, model RP-1A (Govmark Inc.). The test specimen was located parallel to the panel face. The panel temperature at the beginning of each test was about 700725 C. The temperaturetime profile on the cold side of the board was monitored using an IR camera (Thermovision A40M) from FLIR systems, located orthogonally to the panel at a distance of 1.3 m. A scheme of the experimental setup is reported in Figure 1. Further information on the features of the apparatus and on the preparation of the samples can be found in the original ASTM E162 documentation.18 The following test durations were selected: 15, 25, and 50 min. Samples of the panels were collected at the end of the tests and characterized by the above techniques. Model Developed for the Simulation Medium-Scale Fire Tests. A specific model was developed for simulation of the medium-scale laboratory tests carried out using the modified 7680

dx.doi.org/10.1021/ie201867v |Ind. Eng. Chem. Res. 2012, 51, 7679–7689


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ASTM device. The model was based on the experimental data obtained from characterization of the PFP materials. The geometry of the experimental setup suggests that heat propagation through the sample exposed to radiant heat takes place mainly in the direction perpendicular to the exposed face. Therefore, a simplified one-dimensional energy balance equation was adopted to model the experiments: ∂T ∂ ∂T q_ 00 ¼ k þ I þ F cp ∂t ∂x ∂x dx

∑j ΔH^ R, j

dξj ω0, j F0 dt

!

ð1Þ

where T is the local temperature, t is the time dimension, x is the spatial dimension through the board perpendicular to the exposed face, F is the local bulk density, cp is the local heat capacity, k is the local thermal conductivity, q00 I is the heat flux entering the local interface, ΔHR,j is the heat effect from reaction j, ξj is the mass conversion for reaction j, ω0,j is the mass fraction of the virgin material consumed by reaction j, and F0 is the bulk density of the virgin material. The contribution to heat transport by the flow of released gases leaving the material was neglected in the model. The constitutive equations for the properties and the reaction kinetics used in the simulation and obtained from the experimental data are reported in the Supporting Information. In materials where swelling occurs, a swelling factor ψ was considered. Accordingly to the one-dimensional model, swelling was considered to occur only in the direction of heat flow and was represented as a local stretching of the spatial dimension of the material: ð2Þ

dx ¼ ψ dx0

where dx is the differential in the spatial dimension used in eq 1, dx0 is the differential corresponding to the same spatial dimension at the initial condition (virgin unexposed material), and ψ is the local swelling factor (ratio between the current swollen length and initial length). Equation 1 was coupled with the relevant boundary conditions. At the initial time, the sample is uniformly at room temperature. Heat flux through the sample interface (q00 I) only occurs at the two faces of the board: at the exposed face, radiative heat exchange occurs with the radiant panel surface, considered to be at a constant temperature over time; at the back face, convective heat transfer occurs with surrounding air, which also is considered to be at a constant temperature over time. These boundary conditions can be expressed as follows: 8 4 4 > for x ¼ 0 < B0 ðεE TE εS T Þ 00 0 for 0 < x < L0 ð3Þ q_ I ¼ > : hA ðT TA Þ for x ¼ L0 Tðx, t ¼ 0Þ ¼ T0 TE ðtÞ ¼ TE, 0

"t

TA ðtÞ ¼ TA, 0

"t

"x

where B0 is the Boltzmann constant, εE and εS are the surface emissivities of the radiant element and board exposed face,

respectively, hA is the convective heat-transfer coefficient with surrounding air, TE and TA are the temperatures of the radiant element and room air, respectively, and L0 is the thickness of the board. The differential equation (1) was numerically solved over space and time by conventional finite-difference methods. Further information on the discretization of heat-transfer equations and verification of the stability criteria can be found elsewhere (e.g., Incropera et al.19).

’ RESULTS AND DISCUSSION Thermal Stability. The thermal stability and thermal degradation of the selected PFP reference sample materials were investigated using TG and TGDSCFTIR analysis. Figure 2 shows the results of constant heating rate runs (10 C/min). Figure 2a reports the results for the epoxy intumescent coating obtained in pure nitrogen. As shown in the figure, two main degradation steps are present. The first (about 12% of the initial weight) corresponds to the dehydration of boric acid and takes place at temperatures lower than 260 C. The second degradation step, from 260 to 600 C, is due to the epoxy resin and APP degradation. In this temperature range, coating expansion also occurs. Weight loss in this step is significant, and the final residue is equal to 35% of the initial sample mass. Results from TG tests at different heating rates can be used to develop a simplified lumped-parameters apparent kinetic model for the thermal degradation of the material.2022 A simplified model obtained for the epoxy intumescent reference material used in the present study is described elsewhere.23 Figure 2b reports the results obtained for the epoxy intumescent coating in air. As is evident from a comparison of parts a and b of Figure 2, the thermal degradation behavior is only slightly influenced by the presence of oxygen, which does not affect the two-step main thermal degradation process. Figure 2c shows the results obtained for the vermiculite cement-based coating. The overall weight loss is equal to 27% of the initial sample weight. Most of the weight loss (21%) takes place at temperatures higher than 500 C, when carbon dioxide (CO2) is released. Water is mainly released at lower temperatures. Figure 2d reports the results for the silica blanket. The material shows a moderate weight loss (14% of the initial weight), mostly due to water evaporation and degradation of binders present in the material. TGDSCFTIR runs were performed on the epoxy intumescent coating and on the vermiculitecement-based coating in order to identify the volatile substances released when exposed to high temperatures. The silica blanket was not analyzed because, as mentioned before, only limited losses take place and do not affect the properties of the material. Also, in the case of vermiculitebased coating and of silica blanket, no relevant changes were detected when air was used as the purge gas in TG runs. Figure 3 reports the results of constant heating rate runs (10 C/min) carried out in pure nitrogen (60 mL/min). As shown in Figure 3a, the water release of the epoxy intumescent coating for temperatures lower than 260 C is confirmed by the FTIR results. With respect to the second degradation region observed in the TG tests, ammonia, CO2, and organic materials (probably phenols), identified by the CH stretching, are released. These substances are possibly formed in the thermal degradation of the epoxy resin and the APP present in the material.2426 In Figure 3b, the results obtained for the vermiculitecement-based coating 7681



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Figure 2. Results of the constant heating rate (10 C/min) in TG runs on the reference samples of fireproofing materials: (a) epoxy intumescent coating, pure nitrogen (100 mL/min); (b) epoxy intumescent coating, air (100 mL/min); (c) vermiculitecement coating, pure nitrogen (100 mL/min); (d) silica blanket, pure nitrogen (100 mL/min).

Figure 3. Results of TGDSCFTIR runs (10 C/min, 60 mL/min pure nitrogen flow) for the reference materials: (a) TGFTIR results for the epoxy intumescent coating; (b) TGFTIR results for the vermiculitecement-based coating; (c) DSC results for the epoxy intumescent coating; (d) DSC results for the vermiculitecement-based coating;. 7682



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Table 1. Quantities (%) of Selected Substances Released during the TGFTIR Runs Reported in Figure 3a sample

a

water

ammonia

CO2

total

weight loss (TG data)

epoxy intumescent coating, first region (below 260 C)

11.2

N/A

N/A

11.2

12.0

epoxy intumescent coating, second region (above 260 C)

N/A

6.2

2.9

9.1

53.0

vermiculitecement-based coating, first region (below 500 C)

3.7

N/A

N/A

3.7

5.0

vermiculitecement-based coating, second region (above 500 C)

N/A

N/A

16.4

16.4

21.6

Quantitative data were obtained by the procedure described by Marsanich et al.17

Table 2. Calculated Thermal Effects (mJ/mg) Corresponding to the DSC Peaks Identified in Figure 3a material

a

peak ID no.

heat effect

epoxy intumescent coating

1

202 (endothermic)


2

142 (exothermic)

vermiculitecement-based coating

1

191 (endothermic)

vermiculitecement-based coating vermiculitecement-based coating

2 3

8 (endothermic) 281 (endothermic)

Refer to parts c and d of Figure 3 for the peak ID numbers.

are shown. As reported in the literature,27,28 the first step of the weight loss corresponds to water, while at higher temperatures, CO2 is released. The procedure reported by Marsanich et al.17 was used for the quantification of some of the substances released in the TG DSCFTIR runs. Table 1 reports the data obtained for water, ammonia, and CO2 releases, expressed as percent (by weight) of the virgin material. The results show a sufficient agreement with the TG weight losses, in particular for the first degradation region of the epoxy intumescent coating. Parts c and d of Figure 3 report the DSC curves obtained for the epoxy intumescent and vermiculitecement-based coatings, respectively. Figure 3c clearly shows that boric acid dehydration during epoxy intumescent coating heatup is clearly an endothermic process. Epoxy matrix degradation, on the other side, shows a complex system of reactions. However, an exothermic peak can be recognized around 355 C. The average thermal effects calculated for these peaks are summarized in Table 2. For the vermiculitecement-based coating, all of the phenomena that take place during heat exposure are endothermic (water and CO2 release). The first peak can be considered as water release. The second peak is again associated with water release, probably because of the vermiculite components of the sample. Finally, the third peak can be linked to calcination reactions involving the release of CO2. Morphology and Physical Property Modification Following Heat Exposure. Although all of the materials analyzed were designed with the same purpose, i.e., to protect structural elements from fire, the protection mechanisms are different for the different categories of materials. In order to better evidence these differences, tests were carried out using the fixed-bed furnace setup. The samples used were small sections of the reference materials (cylinders of about 20 mm diameter, with thickness equal to realscale applications). Data presented in the following refer to tests carried out using a constant heating rate (10 C/min) and a pure nitrogen flow (80 mL/min) up to different final temperatures. The results presented were confirmed by similar tests performed at higher heating rates (up to 25 C/min), not reported for the sake of brevity. The epoxy intumescent coating is an active insulator and protects surfaces by means of swelling when given

Figure 4. Swelling of the epoxy intumescent coating: cylindrical samples heated up to the indicated final temperatures using the fixedbed tubular reactor furnace, a pure nitrogen purge flow of 80 mL/min, and a constant heating rate of 10 C/min. Samples were confined so that swelling occurred only in one direction.

surface temperatures are reached: the coating degradation forms a highly viscous liquid that traps the gases released by a “blowing agent”. This originates a swollen protective foam.2426 Eventually, the material forms a char that has a lower thermal conductivity than the original material, delaying the heating of the protected surface. In the particular formulation of the coating considered here, as shown by TG and TGFTIR analysis, the main transformations occurring are the dehydration of boric acid at temperatures lower than 260 C and degradation of the epoxy resin and APP at higher temperatures. This coating can expand up to 4 times its initial volume. The swelling was studied by analyzing the samples heated in the fixed-bed tubular reactor. Figure 4 shows the swelling process for cylindrical samples heated in a pure nitrogen atmosphere up to final temperatures between 325 and 800 C using a constant heating rate (10 C/min). It can be observed that the main swelling phase occurs at temperatures between 270 and 400 C. The pore fraction is considerably increased by the swelling process, affecting key properties for fire protection such as the thermal conductivity. In Figure 5, scanning electrom microscopy (SEM) pictures of the degraded and nondegraded coating are reported. As the figure clearly shows, the material evolves from a low-porosity coating (Figure 5a) to a high-porosity medium (Figure 5b). The initial pores present in the virgin material were generated during the material spray application. During the swelling process, the porosity grows and eventually results in an open-pore structure, with a significant increase of pore diameters. The final formation of an open-pore structure is believed to prevent further swelling, ending the expansion process. Cement-based materials are characterized by a high porosity, which contributes to the heat insulation effect. As confirmed by the results of TG and TGFTIR analysis, during the heating process, several reactions actually take place in the material: water and CO2 are released, accompanied by dehydroxylation and calcination processes.27,28 However, exposure of the vermiculitecement-based coating in the fixed-bed tubular reactor tests showed only negligible changes in the morphology and volume of this material during exposure to high temperatures. 7683



Figure 5. SEM pictures of the epoxy intumescent coating: (a) virgin material; (b) sample heated up to 800 C using the fixed-bed tubular reactor furnace (pure nitrogen purge flow of 80 mL/min and a constant heating rate of 10 C/min).

A slight embitterment of the material was the only perceptible change observed. SEM images of the virgin material (see Figure 6a) clearly show some mica sheets as well as the openpore structure originating by spray application. Figure 6b shows that a small-scale pore structure appears after exposure to high temperatures, probably because of material shrinkage following water and CO2 losses. Silica blankets consist of a fibrous structure with a high void fraction (over 90%). The silica fiber provides a good chemical resistance as well as a high mechanical strength at temperatures higher than 1000 C. Therefore, no significant modification of the morphological structure nor swelling/shrinkage phenomena was observed for this material in the thermal tests carried out using the above-described conditions (fixed-bed reactor, final temperature of 900 C, purge flow of pure nitrogen 80 mL/min), as confirmed also by the negligible weight loss shown in TG runs. Swelling and weight loss of the fireproofing material may produce a significant change in the apparent density (or bulk density). The variation of the apparent density of the materials with temperature was calculated from experimental data. Figure 7a shows the results obtained for density variation with respect to temperature considering a constant heating rate (10 C/min) heatup to 800 C in nitrogen. TG weight loss and experimental data on the volume variation during the swelling process were used to obtain the curves reported in Figure 7a. Data were validated by direct measurements carried out using the Archimedes principle. Volume variation due to swelling was considered for the epoxy intumescent coating, while the volume was

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Figure 6. SEM pictures of the vermiculitecement-based coating: (a) virgin material; (b) material heated up to 800 C using the fixed-bed tubular reactor furnace (pure nitrogen purge flow of 80 mL/min and a constant heating rate of 10 C/min).

considered to remain constant for the vermiculitecement-based coating and silica blanket. As shown in Figure 7a, the apparent density of the silica blanket remains almost constant in the temperature range considered. The vermiculitecement-based coating only showed a slight density reduction at high temperatures, while the epoxy intumescent coating suffers a dramatic change in the density (about an order of magnitude) due to the swelling and significant weight loss. The real density of the materials (i.e., the actual density of the solid fraction of the material) was measured at different temperatures by a pycnometer, using crushed samples obtained from the fixed-bed tubular reactor (pure nitrogen purge flow of 80 mL/min and a constant heating rate of 10 C/min up to the desired temperature). Table 3 reports the results obtained and the corresponding value of the apparent density of the sample. As shown in the table, the degradation process only marginally affects the real density of the materials. On the basis of the apparent and real densities, the pore fraction (ϕ) can be calculated as follows: j ¼ 1

Fapp Freal

ð4Þ

The results obtained by eq 4 for the three reference materials considered in the present study are reported in Figure 7b. As shown in the figure, the porosity of the silica blanket is almost constant with temperature and very high (around 95%). For the vermiculitecement-based coating, the porosity increases slightly 7684



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only when the sample is exposed to temperatures higher than 500 C, but the variation is fairly small (only 5%), leading to a final pore fraction of 84%. Regarding the epoxy intumescent coating, the variation of the pore fraction with temperature is wide because of the swelling of the material. The pore fraction changes from initial 0.13 to 0.93 when the material is exposed to a final temperature equal to 800 C. The thermal conductivity (k) is a key property for fire protection materials. Thus, the variation of this parameter with heat exposure is crucial to assessing the performance of a PFP material. The thermal conductivity is mainly a function of three parameters: porosity of the solid, thermal conductivity of the solid matrix, and thermal conductivity of the gas inside the pores. A simple model, based on the heat flow through parallel layers, was used in the present study to provide an upper bond estimation of the thermal conductivity:2932 kapp ¼ ð1 jÞksolid þ jkpore

ð5Þ

where ϕ is the pore fraction, kapp is the bulk thermal conductivity, ksolid is the thermal conductivity of the solid fraction of the

Figure 7. Apparent densities (a) and pore fraction (b) calculated for the reference PFP materials as a function of the temperature during constant heating rate (10 C/min) heatup to 800 C.

material, and kpore is the thermal conductivity of the pores, assumed as the thermal conductivity of the gas present in the pores. Figure 8 shows the results obtained assuming that, for the sake of simplicity, the thermal conductivity of air was assumed for that of the gas contained in the pores. Medium-Scale Fire Tests. The fire test device described in the Experimental Section was used to test sample boards of the reference PFP materials. The test provided a link between the laboratory-scale experimental characterization of the fireproofing materials and real-scale applications of fireproofing. In fact, the material thickness and exposure conditions (heat-flow and external temperatures) in these tests are similar to what are expected in real-scale fire tests.9,10 Figure 9 reports an example of the average temperature profile over a section of the back side of the panel (i.e., the side of the panel not exposed to fire radiation) for the three reference materials. As shown in the figure, the profiles are significantly different for the three materials studied. On the one hand, in the case of the silica blanket panel, the temperature rises quickly during fire exposure because of the low heat capacity provided by the highly porous structure of the blanket. On the other hand, the epoxy intumescent coating clearly shows the effect of material modification during heat propagation and temperature rise, evidencing the sacrificial behavior of the material. The swelling both increases the thickness and decreases the thermal conductivity, granting ultimate protection performances similar to those of the silica blanket. The temperaturetime curve in Figure 9 also shows that, besides the higher heat capacity, the presence of endothermic reactions further contributes to delaying the temperature increase in the case of the intumescent coating. In the case of the vermiculitecement material, the high thermal capacity, significant thickness, and presence of endothermic reactions contribute to a slower temperature increase. Several small cracks appear following exposition to radiation, evidencing that also in this case

Figure 8. Thermal conductivity calculated for the reference PFP materials as a function of the temperature during constant heating rate (10 C/min) sample heatup to 800 C.

Table 3. Apparent Density, Real Density, and Pore Fraction of Selected Fireproofing Materials: Virgin Material and Material Degraded at 10 C/min up to 800 C apparent density (g/cm3) material

virgin

real density (g/cm3)

degraded

virgin

degraded

pore fraction virgin

degraded


1.17

0.10

1.34

1.54

0.13

0.93

vermiculitecement-based coating

0.53

0.39

2.48

2.48

0.79

0.84

silica blanket

0.13

0.11

2.57

2.57

0.95

0.96

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Table 4. Results of Medium-Scale Tests: Thickness and Conversion (as Defined by eq 6) of the Epoxy Intumescent Coating Boards Exposed to Radiant Panels time of exposure 15 min

25 min

50 min

initial thickness (mm)

8.5

8.5

8.5

final thickness (mm)

23

28

31

depth from the exposed face (mm)

Figure 9. Comparison of temperaturetime profiles for the back side of the panel in medium-scale fire tests. Reference exposure time: 15 min.

the material needs to be replaced after fire exposure in real applications. These results suggest that, in defining the thermal protection performance of fireproofing materials, both steady and transitory phases have to be properly considered and modeled because the latter may have a significant duration. In the case of the epoxy intumescent coating, the sample board was analyzed after the end of the test. Samples were taken at different depths and tested for residual degradation potential by TG analysis. The results allowed identification of the overall sample conversion (ξTOT), defined as follows: ξTOT ¼

mV m m 1 ΔWTOT ¼ 1 mV mf mf ΔWTOT

ð6Þ

where m is the mass of the sample to be characterized, mf is the mass of the residue at the end of the TG run on the sample (constant heating rate, 10 C/min, up to final temperature 800 C; pure nitrogen flow 100 mL/min), and mV is the theoretical mass of the virgin material corresponding to the final residue in TG runs. The latter can be easily calculated from the experimental value of mf and from ΔWTOT, the average overall weight loss of the virgin material. The value of ΔWTOT for the epoxy intumescent coating obtained from the TG runs reported in Figure 2 is equal to 65%. Table 4 reports the results obtained. As shown in the table, complete degradation occurred even for exposure times as low as 15 min. Only a few samples on the unexposed side of the board show partial degradation. Thus, the swelling process may be considered almost complete after 15 min. A model simulation of the results of the fire tests was carried out using the model presented above. The model simulations were aimed at identifying the key critical properties of the materials to be considered in simulations. The importance of introducing temperature- and conversion-dependent correlations for these critical properties for the different types of materials was also assessed. It should be noted that model equations are able to account for several phenomena that have been evidenced as of interest for the system studied, e.g., the heat contribution of the reactions and variation of the material properties with position, time, and temperature. As mentioned in the Introduction, this is crucial for the improvement of FEM simulation and of modeling techniques for the advanced assessment of the design and safety performance.14,23 Clearly enough, the use of reliable models

overall conversion (ξTOT)

0

1.00

1.00

1.00

5 10

1.00 1.00

1.00 1.00

1.00 1.00

20

1.00

1.00

1.00

back face

0.20

0.23

0.26

allows the simulation of different fire and equipment scenarios, beyond the standardized cases obtained in fire tests. Figure 10 shows some results obtained for the simulation of two reference sample materials. The two materials showing the different behavior were selected: silica blanket, showing almost no change in the physical properties, and the epoxy intumescing coating, showing a wide change in the properties and chemical structure. As shown in the figure, experimental data and simulations show a fairly good agreement, validating the proposed approach. Robust simulation results require accuracy of the models describing the fundamental properties of the materials: e.g., in Figure 10a, curve 2 shows that neglecting the temperature dependence of the thermal conductivity leads to an underestimation of the protection performance. While these effects can be considered practically negligible for some inactive fireproofing materials, neglecting the temperature and conversion dependency of the properties for active fireproofing materials may dramatically affect the result of the simulations. Figure 10b shows a few examples in the case of the epoxy intumescent boards: assuming fixed values for the material properties (curves 3 and 4) or neglecting the thermal effect of undergoing reactions (curves 2 and 4) leads to unreliable and nonconservative outcomes. Besides a correct modeling of the property dependence, high attention should be devoted also to the quality of the input values. These should be derived as much as possible from specific experimental studies. However, the simulation runs suggest that not all of the properties may require the same accuracy level. Figure 11 shows the results of a simple sensitivity analysis on two key properties for an inactive fireproofing material: thermal conductivity (k) and volumetric heat capacity (Fcp). While both of the parameters affect the transitory phase of the phenomena, the role of the thermal conductivity in defining also the stationary phase calls for a more accurate description of this one property over the others. A further confirmation of the crucial importance of including temperature and conversion dependency of the properties in the case of active insulating materials is provided by the results in Figure 12. The figure shows the simulation results for the transversal section of an epoxy intumescing coating board in a medium-scale fire test at an exposure time of 25 min. Figure 12a evidences the significant expansion of the board and complex profile of the thermal conductivity that needs to be taken into account by the simulation. The results are confirmed by a comparison with experimental data, which evidences the change of the swelling 7686



Figure 10. Simulation results for medium-scale fire tests: (a) silica blanket, 15 min exposure time [1, simulation considering the temperature-dependent properties; 2, simulation considering a constant thermal conductivity (k = 0.134 W m1K1 )]; (b) epoxy intumescent coating, 25 min exposure time (1, simulation considering the temperature- and conversion-dependent properties and thermal effects of the reactions; 2, simulation considering the property dependence but no thermal effects of the reactions; 3, simulation considering constant properties but thermal effects of the reactions; 4, simulation considering constant properties and no thermal effects of the reactions).

Figure 11. Sensitivity analysis on the simulation results for the silica blanket: effect of a 30% variation in the thermal conductivity (k) and in volumetric heat capacity (Fcp).

factor and porosity as a function of the distance from the side of the board exposed to fire (Figure 12b). Again, the results validate that the swelling is important only at temperatures higher than 270 C.

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Figure 12. Simulation results for an exposed board of the epoxy intumescent coating at time t = 25 min: (a) temperature and thermal conductivity profile (1, position of the exposed face at t = 0 min; 2, experimental position of the exposed face at t = 25 min); (b) pore fraction and swelling factor (1, experimental position of the appreciable material discoloration; 2, experimental position of the appreciable increase in the pore size).

All of these results confirm the need to effectively include in FEM simulations detailed submodels for the temperature- and conversion-dependent key physical properties of PFP materials. These advanced models have enhanced performances and a higher accuracy in the prediction of temperaturetime profiles. The extension of these results to real-scale simulations14,23,33 may potentially improve the robustness of the models assessing the actual protection performance on specific setups.

’ CONCLUSIONS The results obtained confirm that the performance of fireproofing materials in providing effective protection from fire strongly depends on the thermophysical properties and on the behavior of the PFP material during fire exposure. The use of conventional experimental techniques allowed the assessment of the key physical properties (thermal conductivity and density, porosity) for reference samples representing the three more widely used categories of PFP materials (intumescent coatings, lightweight cement-based coatings, and man-made fibers). The results evidenced that not only active insulators but also nonactive coatings may undergo significant changes in their structure and properties when exposed to high temperatures. As shown in the simulation of medium-scale fire tests, accounting for changes in the key physical properties is of crucial importance for the correct simulation of the PFP performance, at least when active fireproofing materials are adopted. This knowledge can contribute 7687


Industrial & Engineering Chemistry Research to a more accurate calculation of the time to failure of the protected material, providing more reliable information on the PFP expected performance.

’ ASSOCIATED CONTENT

bS

Supporting Information. Constitutive equations used in the heat-transfer model, material property data used in the simulations presented, and a Nomenclature section. This material is available free of charge via the Internet at http://pubs.acs.org.

’ AUTHOR INFORMATION Corresponding Author

*Tel.: (+39)-051-2090240. Fax: (+39)-051-2090247. E-mail: [email protected].

’ NOMENCLATURE B0 Boltzmann constant [W/(m2 K4)] cp specific heat capacity [J/(kg K)] convective heat-transfer coefficient [W/(m2 K)] hA k, kapp bulk thermal conductivity [W/(m K)] thermal conductivity of the pores [W/(m K)] kpore thermal conductivity of the solid [W/(m K)] ksolid initial thickness of the board [m] L0 m mass of the sample [kg] mass of the final residue in TG runs [kg] mf theoretical mass of the virgin material corresponding to mV mf [kg] interface heat flux [W/m2] q00 I t time [s] T temperature [K] temperature of the environment [K] TA temperature of the radiant element [K] TE x position through the board, perpendicular to the faces [m] position through the board at the initial conditions x0 (virgin unexposed material) [m] ^ R,j heat effect from reaction j [J/kg] ΔH ΔWTOT overall fractional weight loss surface emissivity of the radiant element εE surface emissivity of the board exposed face εS mass conversion for reaction j ξj overall mass conversion ξTOT F, Fapp bulk density [kg/m3] initial bulk density (virgin material) [kg/m3] F0 real density [kg/m3] Freal ϕ pore fraction ψ swelling factor mass fraction of the virgin material consumed by ω0,j reaction j ’ REFERENCES (1) DET NORSKE VERITAS (DNV) Offshore Standard DNV-OSD301 Fire Protection, 2008. (2) Recommended Practice for Fire Prevention and Control on Fixed Open-type Offshore Production Platforms. API Recommended Practice 14g, 4th ed.; American Petroleum Institute: Washington, DC, April 2007; API14G-2007. (3) UL Standard for Safety for Rapid Rise Fire Tests of Protection Materials for Structural Steel, 3rd ed.; Underwriters Laboratories Inc.

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