High-temperature damage characterization of polymer-modified hybrid fiber-reinforced cement-based composites and underlying mechanisms

Abstract

Polymer-modified hybrid fiber-reinforced cement-based composites (PHFRCCs) have been widely applied in construction, with increasing attention on their applications in thermal engineering. In this study, the thermal strain of PHFRCCs was characterized through high-temperature experiments and digital image correlation techniques. The residual strength and microstructural characteristics of PHFRCCs after high-temperature exposure were investigated. A power-law damage model with an exponent of 0.5 was established. The results show that from 200 to 400 °C, a significant increase in thermal strain and horizontal flexural strain is observed in PHFRCCs. From 400 to 600 °C, an opposite trend occurs, which is attributed to the complete decomposition of most polymers, leading to pore formation and the release of thermal strain. These findings indicate that PHFRCCs exhibit superior thermal resistance. Compared to pure cement, PHFRCCs show higher residual strength and horizontal flexural strain, indicating enhanced flexural resistance due to the reinforcing effect of steel fibers. At 600 °C, PHFRCCs show a relatively low damage rate compared to pure cement. In contrast, between 200 and 400 °C, polymer degradation dominates, resulting in a higher damage rate. These findings provide valuable theoretical insights for the fire safety assessment and design of PHFRCCs.

Graphical abstract

Keywords

High-temperature damage / thermal strain / residual strength / cement-based composites

1 Introduction

Catastrophic events such as fires pose serious threats to engineering structures exposed to high-temperature environments, including residential buildings [1], tunnels, power stations, and bridge decks [2]. These events not only endanger human lives and public property but also disrupt critical infrastructure and result in substantial economic losses [1,3,4]. Elevated temperatures can significantly degrade the mechanical properties of structural materials, compromising structural integrity and occupant safety [5,6], and in extreme cases, may lead to structural collapse due to thermal loading. Consequently, enhancing fire-resistant structural design and improving the post-fire performance evaluation of construction materials are pressing and challenging issues in the development of high-temperature-resistant building materials [7,8]. Cement-based composites exhibit complex behavior under high-temperature exposure, undergoing irreversible chemical and physical transformations that result in internal decomposition and deterioration of mechanical properties [9]. After thermal exposure, structural components often exhibit cracking, debonding due to internal expansion, surface spalling of the cement matrix, and even explosive spalling under severe conditions. In recent years, high-performance concrete and ultra-high performance concrete have garnered increasing attention in the construction industry due to its superior mechanical strength and durability [10–12]. However, the increased density and compactness of these advanced composites—resulting from the inclusion of dense materials, can lead to elevated pore pressure at high temperatures due to water evaporation, aggregate decomposition, and accelerated cement hydration [9,13–15], ultimately undermining their fire resistance. Other studies [16] have reported that geopolymers experience a sudden loss of stiffness and significant deformation when exposed to temperatures around 600 °C. In contrast, polymer-modified hybrid fiber-reinforced cement-based composites (PHFRCCs) exhibit excellent fire resistance and superior mechanical properties after high-temperature exposure.

PHFRCCs have gained widespread application in the construction industry due to their enhanced performance characteristics. The synergistic effects of polymer modification and fiber reinforcement significantly improve various material properties [17,18]. Previous studies have demonstrated that the incorporation of polymer latexes and fibers—such as low-density polyethylene, polypropylene, polyvinyl alcohol (PVA), and polyamide, can substantially enhance spalling resistance under elevated temperatures [19–22]. Furthermore, the combined use of high-melting-point steel fibers and polymer fibers has been shown to reduce strength degradation and improve post-heating ductility [23–26]. Despite these advancements, most existing research has predominantly focused on either polymer modification or fiber reinforcement in isolation. Although several studies—such as those by Zhang et al. [17,18], Li et al. [27], Wang et al. [28], Pachta et al. [29], Lyu et al. [30], Park et al. [31], Ozawa et al. [32], and Constâncio Trindade et al. [33], have investigated the high-temperature behavior of natural fibers, geopolymers, polymer particles, rubber particles, plastic waste and synthetic fibers. Of course, some progress has also been made in studies on hybrid fiber systems incorporating steel and synthetic polymer fibers [34], basalt fiber reinforced mortars [35], as well as polymer-modified cement-based materials [36], under high-temperature conditions. Systematic studies on the high-temperature damage mechanisms of PHFRCCs remain limited. Therefore, a more comprehensive understanding of the thermal damage characteristics of PHFRCCs is essential for advancing fire-resilient structural design and improving fire safety evaluation methodologies.

A comprehensive understanding of the individual roles of polymers, polymer fibers, and steel fibers in PHFRCCs under high-temperature conditions is essential for optimizing their application in fire-resistant structural systems. In recent years, researchers have investigated the high-temperature behavior of cementitious materials using a range of approaches, including theoretical modeling [37,38], macroscopic mechanical performance evaluation [39–41], and microscopic characterization techniques [39,42]. These studies have primarily focused on developing empirical stress–strain relationships for post-fire conditions and on evaluating the effects of polymers, fiber types, coarse aggregates, fly ash, and other heat-resistant additives [43,44]. Widely adopted analytical techniques such as X-ray diffraction (XRD), scanning electron microscopy (SEM), computed tomography, digital image correlation (DIC), and acoustic emission technology [23,45–48] have significantly contributed to the advancement of cement-based composites and have laid a solid foundation for subsequent research. For example, Zawadowska et al. [49] investigated the ductility and compressive strength of cementitious specimens under both transient and steady-state thermal conditions. Rawat et al. [50] and Zhang et al. [51] studied the spalling behavior and thermal damage mechanisms in concrete subjected to elevated temperatures. Despite these valuable contributions, the majority of existing studies have concentrated on material composition, fiber content, and the residual tensile and compressive properties [52–54] after thermal exposure. However, there remains a significant research gap concerning the flexural damage behavior of PHFRCCs under high-temperature conditions and the associated damage mechanisms. Furthermore, a comprehensive understanding of the relationship between high-temperature-induced damage and residual strength in cementitious materials is still lacking.

In this study, the high-temperature thermal strain and residual flexural performance of PHFRCCs were systematically investigated through a series of thermal resistance tests, DIC measurements, and residual flexural strength evaluations of notched beams subjected to elevated temperatures (200 , 400 , and 600 °C). In addition, microscopic characterization techniques were employed to examine the internal microstructure of the composites after thermal exposure. To describe the damage behavior of PHFRCCs under high-temperature conditions, a theoretical damage model was developed based on the residual strength and elastic modulus. The primary objective of this study is to characterize the thermal damage behavior of PHFRCCs and to elucidate the underlying damage mechanisms by constructing a reliable high-temperature damage model for cement-based materials. The investigation of the thermomechanical behavior of PHFRCCs helps to demonstrate their potential applications in thermal energy systems and fire-resistant structural components. It provides strong support for fire safety assessment and structural design, and contributes to a deeper understanding of the intrinsic relationship between the thermophysical and mechanical properties of cement-based composites.

2 Experimental

2.1 Materials and specimen preparation

2.1.1 Experimental materials

The raw materials used in this study include PO 42.5 ordinary Portland cement (manufactured by Conch Cement), which meets the stability requirements specified in the Test Methods of Cement and Concrete for Highway Engineering (JTG E30-2005) [55]. The quality inspection report for the cement is presented in Table 1. The fine aggregate used is medium sand with a fineness modulus of 2.69. The hybrid fibers consist of steel fibers and PVA fibers. The steel fibers are copper-coated micro steel wires (produced by Hengshui Shengying Metal Products Co., Ltd.) with a length of 12–14 mm, a diameter ranging from 0.18 to 3 mm, an aspect ratio of 60, and a nominal tensile strength of 2850 MPa. The PVA fibers (TY high-strength PVA fibers from Tianyi) have a diameter of 18 μm, a length of 6 mm, a tensile strength greater than 450 MPa, and an elastic modulus exceeding 5000 MPa. Styrene-butadiene rubber (SBR) latex is employed as a polymer emulsion for cement modification. Grade I ultrafine fly ash is also used, with its main chemical components listed in Table 2. Additional admixtures include a polycarboxylate superplasticizer, incorporated at 0.44% of the cement weight, and a defoaming agent (tributyl phosphate), added at 0.1% of the total composite weight to minimize air entrainment during specimen preparation. Standard residential tap water is used as the mixing water. Specimens were prepared according to the specified mix proportions to ensure consistency and reliability in the experimental procedures. The careful selection and controlled preparation of these raw materials contribute to the accuracy and reproducibility of the test results.

2.1.2 Specimen preparation

Building on the mix proportions for cement-based materials reported in previous studies [56,57], this research adopts the formulation listed in Table 3, which has demonstrated excellent flexural mechanical properties while promoting sustainable and eco-friendly practices. However, the high-temperature performance of these materials remains insufficiently explored. To address this gap, both pure cement-based specimens (S-type, without polymer or fibers) and polymer-modified hybrid fiber-reinforced cement-based specimens (E-type) were prepared for high-temperature experiment, residual flexural experiment, and DIC analysis. In this study, each test was performed with three parallel specimens, and the average value was used for analysis. The specimen preparation followed the Standard Test Methods for Building Mortar Performance (JGJ/T 70-2009) [58], and the detailed procedure can be referenced from previously published studies [56,57]. The specimens were small-sized beam samples with dimensions of 160 mm × 40 mm × 40 mm. After 28 d of standard curing, prefabricated notches were introduced to facilitate the residual flexural performance tests. These notches were created using an SH-9080A grinder, with dimensions of approximately (17 ± 2) mm in height and (2 ± 1) mm in width, corresponding to a height-to-specimen height ratio of approximately 0.425. For DIC analysis, random speckle patterns were applied to the specimen surfaces using high-temperature-resistant paint (≥ 1000 °C). These speckles enabled accurate tracking of thermal strain during the high-temperature tests and strain analysis during the residual flexural performance evaluation.

2.2 High-temperature experiment procedure

High-temperature loads were applied to the prepared specimens to simulate fire conditions. Before and after heating, the specimens were weighed and subjected to DIC testing to analyze mass loss and thermal strain. The designated high-temperature experiment temperatures were 200, 400, and 600 °C. The specimens were labeled as follows: pure cement-based (S-type) samples (S-200, S-400, S-600) and polymer-modified hybrid fiber-reinforced cement-based (E-type) samples (E-20, E-200, E-400, E-600). Among these, E-20 represents PHFRCCs samples at room temperature. The specimens were placed in a computer-controlled experimental furnace and heated at a constant rate of approximately 6 °C/min until the target temperature was reached [59]. Figure 1 shows the experimental furnace (left) manufactured by Beijing Jinglu Xinghua Industrial Furnace Co., Ltd., along with the high-temperature loading setup (right). Figure 2 illustrates the temperature-time relationship curve of the furnace. To ensure thermal stability and accurately simulate fire conditions, the furnace temperature was maintained at the target temperature for 2 h [60]. After the high-temperature exposure, the furnace was turned off, and the door remained closed for 1 h. The door was then gradually opened to allow the specimens to cool slowly, preventing significant temperature differences between the furnace and the laboratory environment, which could lead to specimen spalling [61]. Following high-temperature exposure, DIC testing was conducted to measure the thermal strain distribution. Residual flexural performance experiment was then carried out, primarily focusing on three-point flexural experiment. Previous studies [56,57] have demonstrated the excellent flexural properties of PHFRCCs. This residual mechanical performance experiment aimed to further investigate the flexural damage characteristics of the specimens after high-temperature exposure.

2.3 Residual flexural experiment

The flexural mechanical properties of the materials were evaluated using a three-point flexural experiment of notched beams. In this experiment, the loading platform had a span of approximately 144 mm between the lower supporting brackets on both sides, with the loading head precisely aligned at the center of the specimen and the pre-cut notch. The experiment was conducted using an INSTRON electronic universal testing machine with a maximum load capacity of 5 kN. Displacement-controlled loading was applied at a rate of 0.5 mm/min until the specimen lost most of its bearing capacity. Specifically, during the residual flexural experiments, loading was stopped when the load dropped to 0.1 kN after exposure at 200 and 400 °C, and to 0.01 kN after exposure at 600 °C high-temperature experiment, as the specimen had essentially lost its load-bearing capacity. Additionally, DIC technology was employed to observe strain distribution during the flexural process. DIC images were captured using a Panasonic S5 Mark II camera in time-lapse mode. The captured images were processed on a computer to generate strain fields visualizations. Figure 3 illustrates the dimensions of the notched beam, as well as the setup for the residual flexural loading experiment and the DIC testing platform.

2.4 Microstructural experiment

To conduct both qualitative (phase composition) and quantitative (crystallite size) analyses of the PHFRCCs, XRD testing was performed in this study. A small portion of the specimen (approximately 10 g) was collected after high-temperature exposuer, dehydrated in industrial alcohol for 24 h, and then dried in an oven for 1 h. The dried specimen was subsequently ground into a fine powder using a planetary ball mill for XRD analysis. XRD measurements were conducted using a Rigaku-2038 X-ray diffractometer (Japan), equipped with a Cu target (wavelength: 0.15406 nm), operating at a scan speed of 5°/min. The scanning range was set from 5° to 90°, with a tube voltage of 40 kV and a tube current of 35 mA.

Following the residual flexural performance experiment, the specimens failed along the pre-existing cracks. The fracture surfaces of the specimens were then examined microscopically. A ZEISS GeminiSEM 300 scanning electron microscope (Germany) was used to analyze the microstructure of the specimens, providing detailed morphological insights. Additionally, surface scanning energy dispersive spectroscopy (EDS) analysis was conducted using an OXFORD XPLORE30 energy dispersive spectrometer, equipped with a SmartEDX spectrometer. This analysis was performed on gold-palladium alloy-coated PHFRCCs specimens to assess the elemental composition and distribution within the material’s microregions.

3 Experimental results and analysis

3.1 High-temperature experiment results

3.1.1 High-temperature damage mode

The pure cement and polymer-modified hybrid fiber-reinforced cement-based samples were subjected to high temperatures of 200, 400, and 600 °C. After the high-temperature exposure, the specimens were removed, and the observed damage patterns are shown in Fig. 4. At 200 °C, neither the pure cement nor the PHFRCCs samples exhibited significant damage. Both types showed particle detachment on the surface; however, the PHFRCCs samples displayed a denser pattern of surface cavities, likely due to the influence of polymer fibers at this stage. At 400 °C, surface spalling became more pronounced in both types of cement, with the PHFRCCs samples exhibiting more prominent surface cavities. Notably, dark burn-like regions appeared on the surface, likely due to the combustion and decomposition of the polymer latex. At 600 °C, the pure cement samples developed multi-directional surface cracking, whereas no noticeable macro-cracks were observed in the PHFRCCs samples. However, both samples exhibited surface powdering. This suggests that at 200 and 400 °C, the polymers (styrene-butadiene rubber latex (SBRL) and PVA fibers) in the PHFRCCs facilitated the gradual release of pore pressure. At 600 °C, the cavity phenomenon became more pronounced, and the decomposition of burned material in the dark burn regions occurred due to the high temperature. During thermal expansion, the presence of polymers helped mitigate surface cracking and thermal spalling. After observing the surface damage of the samples post-high-temperature exposure, further investigation will be conducted to assess mass loss after heating.

3.1.2 Mass loss

The weight of the samples was recorded before and after the high-temperature experiment, and the mass loss rate due to high-temperature exposure was calculated using Eq. (1):

(1)\mathrm{\Delta }{m}_{\mathrm{m}}=\frac{m_0-m_n}{m_0}\times 100\mathrm{\%},

where m_0 is the average mass of the sample at room temperature (20 °C) (unit: g). m_n is the average mass of the sample after high-temperature exposure (unit: g), with the subscript n indicating the set temperature. The calculated mass loss rates for both pure cement and PHFRCCs samples are presented in Table 4. As shown, the mass loss increases with higher temperatures, which correlates with the more severe thermal spalling damage observed in the samples. At temperatures of 200 and 400 °C, the mass loss rate of the pure cement samples increased from 8.686% to 10.802%, while the PHFRCCs samples showed an increase from 8.837% to 12.075%. At the highest temperature of 600 °C, the mass loss of the pure cement samples reached 13.172%, while the PHFRCCs samples experienced a mass loss of 14.498%. As shown in the thermogravimetric analysis (TG) curves of pure cement and PHFRCCs in Fig. 5, the mass percentage of the two materials begins to diverge at around 100 °C. During the temperature range of 200–400 °C, the mass loss continues to increase, while in the range of 400–600 °C, a certain degree of moderation is observed. Notably, the PHFRCCs consistently exhibited greater mass loss compared to the pure cement materials. This can be attributed to the burning or decomposition of the polymers incorporated into the PHFRCCs at elevated temperatures, which contributes to the higher mass loss.

3.1.3 High-temperature thermal strain

In this study, DIC tests were conducted on the surfaces of the samples before and after the high-temperature experiment, primarily to characterize the distribution of thermal strain under elevated temperatures. Areas with significant strain changes on the high-temperature speckled surfaces of the cement samples were selected as regions of interest for thermal strain analysis. However, due to thermal spalling of the cement-based surface after high-temperature exposure, the speckle patterns were occasionally obscured, leading to distortions in the strain field calculations. As a result, the calculation regions for certain samples were refined. Figure 6 illustrates the distribution of thermal strain in the horizontal direction during the high-temperature experiment, based on the DIC data. It is evident that after high-temperature exposure, the surfaces of the samples exhibit concentrated strain areas, indicated by red zones, suggesting that the cement-based composites underwent uneven thermal expansion. At temperatures of 200, 400, and 600 °C, the maximum horizontal thermal strains on the pure cement surface were 0.0290%, 0.0306%, and 0.3718%, respectively, representing increases of 5.52% and 1182.07% compared to the strain at 200 °C. For the polymer-modified cement-based composites, the maximum horizontal thermal strains at these temperatures were 0.0493%, 0.3501%, and 0.2722%, showing increases of 610.42% and 452.13% compared to the strain at 200 °C. These results indicate that the thermal strain on the pure cement surface increases continuously with temperature, suggesting a potential for thermal cracking. In contrast, the polymer-modified cement exhibited a trend of increasing strain up to a certain temperature, followed by a decrease, effectively mitigating high-temperature thermal strain. This phenomenon is consistent with the results of the aforementioned TG analysis.

At 200 °C, it is expected that the moisture within both the pure cement and polymer-modified cement matrices evaporates, leading to the expansion of the cement matrix and the formation of microcracks. In the polymer-modified cement matrix, materials such as polymers undergo thermal expansion, resulting in relatively higher thermal strain. However, in this study, the polymer-modified cement matrix reaches its peak thermal strain at 400 °C. At this temperature, the polymer has already softened or largely decomposed, and the expansion characteristics of the polymer, coupled with the interaction between the fibers and the thermal expansion mismatch between the fibers and the matrix, may contribute to the further increase in thermal strain. Despite this, temperatures of 200 and 400 °C do not reach the high temperatures typically associated with standard fire conditions, meaning the full effects of the polymer and fibers have not yet been fully realized. At 600 °C, approaching the high-temperature conditions of a standard fire, the cement matrix undergoes significant physical and chemical changes, such as hydration product decomposition, moisture evaporation, and microcrack propagation. At this point, the decomposition of the polymers (styrene-butadiene rubber latex and PVA fibers) becomes crucial, as the release of internal space helps inhibit the further development of cracks at the phase interfaces and effectively mitigates the continued increase in thermal strain. In this study, the polymer-modified cement matrix begins to demonstrate its beneficial effects at 600 °C, helping the material better resist crack propagation and the increase in thermal strain. It is evident that at temperatures of 600 °C, the challenges faced by the material intensify, and the polymers and fibers play an increasingly important role in enhancing the cement-based composite’s resistance to spalling under fire conditions.

3.2 Residual flexural performance

3.2.1 Load-deflection relationship

To better understand the residual flexural mechanical properties of cement-based materials after high-temperature exposure, the results from the residual flexural experiments were analyzed to obtain the load–deflection curves for each temperature condition, as shown in Fig.7. It is evident that, following high-temperature exposure, the peak load in the residual flexural experiments decreases with increasing temperature for both pure cement and PHFRCCs. Typically, for cement-based materials at a given temperature, the load–deflection curve exhibits the following behavior: initially, the load increases linearly with displacement during the elastic phase; as microcracks form, the curve enters the plastic damage initiation phase, where the load increases non-linearly with displacement until the peak load is reached. This is followed by a significant damage and failure phase, during which cracks propagate unstably, leading to a reduction in load (for pure cement-based materials, this often results in brittle failure with a sharp drop in load, while the PHFRCCs in this study exhibit quasi-brittle failure with some ductility). The PHFRCCs demonstrate larger deflections at the point of complete failure, showcasing excellent ductile failure behavior. Additionally, it is noteworthy that at 600 °C, both pure cement and PHFRCCs samples exhibit load–deflection curves characteristic of ductile failure, rather than the typical brittle failure observed at lower temperatures.

3.2.2 Residual flexural strength

The flexural strength of the notched beam with a rectangular cross-section at a specific temperature can be calculated from the load–deflection curve using Eq. (2):

(2)\sigma =\frac{3F_{\mathrm{m}}L}{2b(h-a{)}^2},

where F_{\mathrm{m}} is the peak load (unit: N), L is the span (unit: m), b is the width (unit: m), and h is the thickness (unit: m), a is the notch depth (unit: m). The variation in residual flexural strength with temperature is shown in Fig. 8. As the high-temperature experiment from 200 to 400 °C and then to 600 °C, the residual strength of the pure cement sample decreases from (4.83 ± 0.49) at 200 °C to (3.58 ± 0.13) MPa at 400 °C, and further to (0.65 ± 0.21) MPa at 600 °C, representing reductions of approximately 25.90% and 86.47%, respectively. For the polymer-modified cement sample, the residual flexural strength decreases from (5.07 ± 0.55) at 200 °C to (3.89 ± 0.24) MPa at 400 °C, and further to (1.28 ± 0.18) MPa at 600 °C, with reductions of approximately 23.29% and 78.18%, respectively. The residual flexural strength of the cement-based materials decreases as the temperature rises during high-temperature exposure, with the polymer-modified cement sample consistently outperforming the pure cement sample. At 400 and 600 °C, the strength reduction in the polymer-modified cement is somewhat less severe. As the temperature approaches standard fire conditions, the polymer-modified cement in this study demonstrates excellent post-fire flexural performance, enhancing the safety of structural inspection and use after a fire. However, it should be noted that compared to room temperature, at medium to low temperatures (approximately 200–400 °C) during high-temperature exposure, the residual flexural strength of the polymer-modified cement decreases significantly. The polymer components (SBRL and PVA fibers) may begin to soften or decompose, leading to a rapid decline in the composite material’s strength. At high temperatures (around 600 °C), as the concrete matrix severely degrades, the residual strength primarily derives from the bridging effect of the fibers and other microstructural features. Due to the fibers, interfaces, and residual structures, the polymer-modified system may more effectively retain some ductility, resulting in higher residual strength and better ductility. Therefore, while the polymer-modified cement exhibits less favorable performance in the medium-to-low temperature range, it shows excellent flexural performance as it approaches fire temperatures.

3.2.3 Residual flexural strain

The strain field images obtained from the DIC tests were processed to generate strain maps in both the horizontal (ε_XX) and vertical (ε_YY) directions, as shown in Figs. 9(a) and 9(b), respectively. An interesting phenomenon is observed in Fig. 9(a): the residual flexural strain of the PHFRCCs in the horizontal direction reaches a maximum of 2.3351% (second image from the right in Fig. 9(a)) at 400 °C, followed by 1.7431% (rightmost image) at 200 °C, and only 0.9356% (third image from the right) at 600 °C, which is an anomaly compared to pure cement. By combining the strain data from the DIC tests with the TG analysis, this phenomenon can be well explained by the softening and decomposition state of the polymer at the corresponding temperatures, as well as the internal structural changes within the material. Since the polymer latex and fibers have not fully decomposed or been consumed, these factors provide the PHFRCCs sample with some adhesive or tensile force, allowing for greater deformation capacity and strain accommodation. Therefore, although the residual flexural strength significantly decreased and the thermal strain sharply increased after exposure at 200 and 400 °C, the residual bending strain remained relatively high, indicating better strain adaptation ability after exposure to high temperatures (fire conditions). Figure 9(b) shows that the residual flexural strain in the vertical direction of the PHFRCCs increases as the temperature rises during the experiment. At 200 °C, the residual flexural strain is approximately 0.024%, which increases by 79.58% and 732.50% at 400 and 600 °C, respectively. A similar trend is observed for the residual flexural strain of pure cement in both the horizontal and vertical directions. Furthermore, in both the horizontal and vertical directions, the residual flexural strain of the PHFRCCs samples is greater than that of pure cement, exhibiting better post-fire ductility, which is advantageous for building safety assessments following high-temperature exposure. The experimental results and calculations presented above are on a macroscopic scale. To further understand the contributions of polymer latex and hybrid fibers to high-temperature resistance and residual flexural performance, the study will continue with microstructural observations.

3.3 X-ray diffraction test results

Figure 10(a) presents the XRD patterns of the PHFRCCs at room temperature (20 °C) and after high-temperature exposure at 200, 400, and 600 °C. At room temperature and 200 °C, the diffraction peaks are relatively broad, indicating a higher content of amorphous phases in the material. As the temperature increases to 400 and 600 °C, the diffraction peaks in the 20° to 30° range (e.g., SiO₂) become sharper, and their intensity significantly increases, suggesting enhanced crystallization of mineral phases such as C-S-H, ettringite, and calcite. Additionally, at 200 and 400 °C, the diffraction peak near 35°, corresponding to Ca(OH)₂, remains prominent. However, at 600 °C, Ca(OH)₂ in the cement matrix undergoes significant decomposition, leading to the formation of a new crystalline phase, CaO. This observation indicates that at moderate temperatures (200–400 °C), the cement matrix does not undergo substantial decomposition, resulting in a relatively simpler phase composition compared to that at 600 °C. This finding helps explain the gradual reduction in residual flexural strength as the temperature increases. Specifically, at 200 and 400 °C, the cement matrix has not yet reached a high-temperature decomposition state, whereas at 600 °C, the decomposition of several hard phases in the cement matrix occurs, further elucidating the trend of decreasing residual flexural strength with rising temperature.

Furthermore, a quantitative analysis was performed based on the XRD test results. Assuming that the increase in peak width is primarily attributable to a reduction in crystallite size, the crystallite size was estimated using Eq. (3) [62]:

(3)d=k\lambda /\beta \cos \theta ,

where k is the Scherrer constant, typically taken as 0.89; \lambda represents the wavelength of the incident X-ray, which is 0.15406 nm in this study; \beta denotes the full width at half maximum of the diffraction peak, which needs to be converted to radians; and \theta is the Bragg angle corresponding to the diffraction peak. As shown in Fig. 10(b), the relationship between the crystallite size of the PHFRCCs and temperature indicates that the crystallite size fluctuates within the range of 15–60 nm. At 400 and 600 °C, the crystallite size of the material phases is relatively large, while at 200 °C, it is smaller. This trend aligns with the observation that high temperatures promote crystallization within the cement matrix. Additionally, the crystallite size of phases such as SiO2 in the 20°–30°, 50°, and 60° regions increases with temperature. In contrast, the crystallite size of phases such as C-S-H and calcium silicate (in the 30°–35° region) does not show significant changes at temperatures below 400 °C. However, at 400 and 600 °C, C-S-H undergoes dehydration, structural collapse, and even decomposition, resulting in a continuous decrease in crystallite size. This further suggests that as the temperature increases, the crystallite size of the cement matrix increases, and some phases may decompose or undergo dehydration, leading to a reduction in strength. On the other hand, this also explains the gradual increase in thermal strain of the pure cement matrix. However, due to the presence of polymers and fibers, an anomalous change in the residual flexural strain of the PHFRCC is observed in the horizontal direction, which can be further analyzed using SEM-EDS testing.

3.4 Scanning electron microscopy-energy dispersive spectroscopy test results

3.4.1 Scanning electron microscopy analysis

Figure 11 presents the SEM images of the PHFRCCs after testing at 200 °C. Figure 11(a) shows that the bond between the steel fibers and the matrix remains intact, with the cement matrix tightly wrapping the steel fibers. However, microcracks are visible at the interface between the steel fibers and the matrix. Figure 11(b) illustrates the bonding between the sand particles, PVA fibers, and the matrix. After the high-temperature experiment at 200 °C, some of the PVA fibers (with a melting point of approximately 230 °C) have been consumed, creating voids or channels. Microcracks are also present at the interface between the sand particles and the matrix. Figure 11(c) shows that part of the SBRL has undergone shrinkage and decomposition due to moisture evaporation, but it still maintains a film-like structure, with the matrix remaining bonded to the polymer film and filling the pores of the cement matrix. Some areas of the matrix surface, where the polymer film does not cover, exhibit microcracks. Figure 11(d) reveals numerous microcracks in the PVA fibers. Due to partial decomposition and consumption of the polymer (SBRL, PVA fibers) at 200 °C, a small number of voids and cracks have formed within the structure, leading to misalignment in thermal expansion. However, the bonding and tensile forces of the polymer are still retained, which helps explain the relatively high residual flexural strain in the horizontal direction.

Figure 12 presents the SEM images of the PHFRCCs after testing at 400 °C. Figures 12(a) and 12(b) depict the bonding between the steel fibers, PVA fibers, sand particles, and the matrix. After the 400 °C high-temperature experiment, the bonding condition is similar to that at 200 °C, with many microcracks at the phase interfaces, and the PVA fibers have decomposed, forming voids or channels. Figure 12(c) illustrates the bonding between the SBRL and the matrix. At approximately 400 °C, the polymer film has decomposed, releasing numerous pores, which loosens the matrix, consistent with the significant thermal strain observed at this temperature. After the 400 °C high-temperature experiment, the PVA fibers have largely lost their functionality. Figure 12(d) shows the broken form of the remaining, partially decomposed PVA fibers, with signs of explosive failure on their surface. At 400 °C, the polymer is nearly consumed, weakening the bonding capacity at the interface, increasing the number of microcracks and pores, and further reducing the residual flexural strength. The pore pressure, thermal expansion mismatch between the fibers and the cement matrix, and physical and chemical reactions within the cement matrix further increase the thermal strain, explaining why the thermal strain is highest at this temperature.

Figure 13 presents the SEM images of the PHFRCCs after testing at 600 °C. Figure 13(a) shows thermal delamination between the steel fibers and the matrix, with cement material also peeling off the surface of the steel fibers. The cement matrix exhibits decomposition, with a large dispersion of particles and poor bonding ability. Figure 13(b) reveals that a few remaining PVA fibers are still attached to the cement matrix but have essentially lost their functionality. Microcracks at the phase interfaces have expanded, showing a trend of large areas of the cement matrix delaminating along the cracks. In Fig. 13(c), large scale decomposition of SBRL occurs, leading to the formation of large voids or cavities that interconnect, creating noticeable defects and further loosening the cement matrix. Figure 13(d) shows that the remaining PVA fibers, with little to no bonding to the matrix, exhibit cracks on their surface, forming wrinkles and contraction. By 600 °C, the polymers and the matrix have undergone extensive decomposition and degradation, causing significant structural collapse, increased porosity, and numerous through-cracks. While this helps alleviate thermal strain and prevent thermal cracking, the residual flexural strength has significantly decreased. Comparatively, the microstructural evolution of PHFRCCs from 200 to 600 °C reveals a clear transition from mild physical changes to severe thermal degradation. At 200 °C, the composites maintain a relatively dense structure, with intact bonding between the steel fibers and the cement matrix, and only limited microcracks caused by partial polymer softening. Similar phenomena can also be observed in the related literature [53,63]. At 400 °C, extensive polymer decomposition occurs, producing numerous voids and microcracks at the fiber–matrix interfaces. The SBRL films largely decompose, loosening the matrix and resulting in the highest thermal strain and a noticeable reduction in flexural strength. When the temperature reaches 600 °C, severe interfacial delamination, matrix decomposition, and large interconnected cavities appear, accompanied by the complete loss of polymer adhesion and bridging effects. Consequently, the structure becomes highly porous and fragile, leading to significant mechanical deterioration despite partial relief of internal thermal stress.

3.4.2 Energy dispersive spectroscopy analysis

To further investigate the microstructural changes, an EDS analysis was conducted based on the SEM observations to examine the elemental distribution and overall spectra of the samples subjected to high-temperature experiments at 200, 400, and 600 °C, as shown in Fig. 14. The elemental mapping of selected microregions revealed a significant presence of carbon (C) and oxygen (O), which are abundant in polymers such as SBRL and PVA fibers. From the overall spectra (Figs.14(b), 14(d), and 14(f)), it is evident that the atomic percentage of carbon (C) decreases with increasing temperature, from 33.92% to 29.33%, and then to 27.00%. This reduction is primarily attributed to the thermal decomposition of PVA, which lowers the carbon content, while organic components like SBRL undergo melting, carbonization, or complete combustion at higher temperatures, further diminishing the carbon levels. The decrease in iron (Fe) content, from 0.58% to 0.29% and then to 0.24%, is likely due to the oxidation of steel fibers at elevated temperatures, leading to the formation of iron oxides such as Fe₂O₃ or Fe₃O₄. Conversely, the increase in oxygen (O) content, from 49.47% to 54.56% and further to 55.99%, can be attributed to the decomposition of polymers and the dehydration of phases like Ca(OH)₂ and C-S-H gels, which release oxygen that subsequently reacts with other elements. Additionally, the C-S-H gels within the cement matrix may undergo structural rearrangement, potentially resulting in a relative increase in the concentrations of calcium (Ca) (from 10.36% to 11.24% and then to 12.25%.).

The elemental composition analysis further confirms that the varying degrees of polymer (SBRL and PVA fibers) decomposition and cement matrix degradation with increasing temperature contribute to the anomalous thermal strain at high temperatures and the residual flexural strain observed in the horizontal direction. These findings provide valuable insight into the microstructural mechanisms underlying the thermal and mechanical responses of the PHFRCCs under high-temperature conditions.

4 Theoretical model

4.1 Damage theory model

It is well established that cement-based materials predominantly exhibit brittle or quasi-brittle failure, often leading to sudden and catastrophic damage. The damage theory model describing catastrophic failure has been widely applied in various fields, such as ceramic coating systems [64–66] and batteries [67]. Building upon this model, we further developed a theoretical framework to characterize high-temperature damage behavior in cement-based composites. Analysis of the residual flexural performance reveals that, with increasing temperature, both the residual flexural strength and elastic modulus generally decline to varying extents, suggesting a damage-dependent degradation relationship. To analyze the variation in the damage rate with temperature following high-temperature exposure, a damage variable can be used:

(4)D=1-\sigma (T)/\sigma (T_0),

where \sigma (T) represents the residual flexural strength under damage at a specific temperature, corresponding to the stress intensity at peak load (unit: MPa). \sigma (T_0) represents the flexural strength at room temperature (unit: MPa). Additionally, the normalized elastic modulus can be used as a control variable:

(5)\varphi =1-E(T)/E(T_0),

where E(T) represents elastic modulus, formulated as a temperature-dependent function (unit: GPa). E(T_0) represents the elastic modulus at room temperature (unit: GPa). At room temperature, the damage variable is zero. When the specimen reaches the high-temperature spalling state, the following equation can be obtained, where the damage reaches 1.

(6)\sigma (T_f)=E(T_f)=0,

where the subscript f denotes the temperature at which high-temperature spalling occurs. \sigma (T_f) and E(T_f) represent the flexural strength and elastic modulus of the specimen at the point of high-temperature spalling failure. At the point of abrupt change, corresponding to the high-temperature spalling state, the damage rate tends to infinity:

(7)\underset{\varphi \to {\varphi }_f}{lim}\frac{\mathrm{d}D(\varphi )}{\mathrm{d}\varphi }=\mathrm{\infty },

or

(8)\underset{D\to D_f}{lim}\frac{\mathrm{d}\varphi (D)}{\mathrm{d}D}=0,

where the subscript f denotes the temperature at which high-temperature spalling occurs. Assuming that \varphi is continuous, it can be written as the sum of the derivatives of D:

(9)\varphi ={\varphi }_f+{\varphi }^{\prime }(D_f)\cdot (D-D_f)+\frac{{\varphi }^{″}(D_f)}{2}\cdot (D-D_f{)}^2+o(D-D_f{)}^2.

Ignoring higher-order terms (second order and above), and considering that:

(10)\underset{D\to D_f}{lim}\frac{\mathrm{d}\varphi (D)}{\mathrm{d}D}={\varphi }^{\prime }(D_f)=0.

Equation (9) can be written as:

(11)D=1-\gamma (1-\varphi {)}^{1/2},

then the damage coefficient can be expressed as:

(12)\gamma ={\left[\frac{-{\varphi }^{″}(D_f)}{2}\right]}^{-1/2},

where \gamma is the damage coefficient, which can be preliminarily calculated by \gamma ={\displaystyle \frac{\left(1-D_1\right)}{\sqrt{\left(1-{\varphi }_1\right)}}}. Among them, D_1, {\varphi }_1 represent the normalized damage and elastic modulus at the initial temperature of 200 °C, corresponding to the design. According to Eq. (2), the flexural strength is derived from the load–deflection curve, while the elastic modulus is determined based on the secant modulus in the linear elastic region [68]. The relevant parameters are summarized in Table 5. Additionally, the damage rate can be expressed as follows:

(13)R=\frac{\mathrm{d}D}{\mathrm{d}\varphi }=\frac{\gamma }{2}{\left(1-\varphi \right)}^{-1/2}.

Consequently, the damage rate increase sharply as the temperature approaches the thermal spalling threshold.

4.2 Application and results

Figures 15 (a) and 15(b) present the relationships between normalized damage, damage rate, and normalized elastic modulus, showing experimental data points at different temperatures together with the fitting curves of theoretical model. It is evident that the proposed high-temperature damage model is consistent with the experimental results, confirming its general applicability to both plain cement and PHFRCCs. It shows that as the experimental temperature increases, the degree of damage to the cement-based composite intensifies, transitioning from a gradual increase to a more pronounced acceleration. Notably, at 200 and 400 °C, the PHFRCCs experience more severe damage than pure cement, primarily due to the thermal degradation of the polymer components. However, at 600 °C, the damage rate of pure cement becomes more significant, highlighting the superior performance and resilience of the PHFRCCs under extreme high-temperature or fire conditions.

5 Conclusions

This study investigated the high-temperature damage of PHFRCCs through a combination of high-temperature experiments, DIC analysis, and notched beam residual strength experiments. A theoretical model was eatablished to characterize the damage. Microstructural characterization techniques, including XRD and SEM-EDS analysis, were also employed to elucidate the damage and failure mechanisms at the microscale. Based on the research findings, the following conclusions can be drawn.

1) As the exposure temperature increases, the damage intensifies and the residual strength decreases. The thermal strain of pure cement increases with temperature, indicating a higher risk of thermal spalling. In contrast, the PHFRCCs show an initial increase in thermal strain followed by stabilization as the temperature rises. Compared to pure cement, the PHFRCCs demonstrate superior flexural strength and horizontal flexural strain after high-temperature exposure. From 200 to 400 °C, the PHFRCCs show a reduction in residual strength and a sharp increase in thermal strain. Nevertheless, they maintain greater horizontal flexural strain after high-temperature exposure, indicating enhanced resistance to deformation. At 600 °C, the PHFRCCs exhibit reduced thermal strain compared to pure cement, demonstrating better high-temperature resistance.

2) Microstructural characterization through XRD and SEM-EDS analyses reveals that the phase composition of the cement matrix progressively decomposes with increasing temperature, leading to a decline in the residual flexural strength of the cement. For the PHFRCCs, at 200 and 400 °C, the polymers (SBRL and PVA fibers) do not fully decompose, and the mismatch in thermal expansion between the fibers and the cement matrix contributes to increased thermal strain and horizontal flexural strain after high-temperature exposure. At 600 °C, the large scale decomposition of polymers effectively reduces thermal strain due to the formation of increased pores. Throughout the high-temperature exposure, the steel fibers serve as essential reinforcements, helping to resist crack propagation and maintain the structural integrity of the material.

3) A high-temperature damage power-law model with an exponent of 0.5 was developed. The results reveal that as the experimental temperature increases, the extent of damage of the cement-based composite intensifies, transitioning from a gradual increase to a more pronounced acceleration. At 200 and 400 °C, the PHFRCCs experience more significant damage compared to pure cement. However, at 600 °C, the PHFRCCs exhibit a relatively lower damage rate than pure cement.

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