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Article

Mechanical Performance and Microstructure Evolution of High-Ferrite Portland Cement Concrete Under the Coupled Abrasion and Freeze–Thaw Cycling Conditions

by
Xingdong Lv
1,2,*,
Yun Dong
1,2 and
Zeyu Fan
1,2
1
Changjiang River Scientific Research Institute of Changjiang Water Resources Commission, Wuhan 430010, China
2
National Dam Safety Research Center, Wuhan 430010, China
*
Author to whom correspondence should be addressed.
Materials 2026, 19(5), 1044; https://doi.org/10.3390/ma19051044
Submission received: 15 January 2026 / Revised: 13 February 2026 / Accepted: 4 March 2026 / Published: 9 March 2026
(This article belongs to the Special Issue Eco-Friendly and Sustainable Concrete: Progress and Prospects)
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Abstract

This study investigates the performance and microstructure evolution of high-ferrite Portland cement (HFC) concrete under the coupled action of abrasion and freeze–thaw cycles (CAA-FTC). The 3D surface morphology of deteriorated concrete was studied; abrasion depth and volume loss evolution data were collected, while analyzing the abrasion depth fractal dimension. The characteristics of hydration products were determined using mercury intrusion porosimetry and 29Si nuclear magnetic resonance method. The ITZ’s micromechanical properties and thickness were investigated via nanoindentation and SEM-EDS. The results show that under the CAA-FTC conditions, concrete deterioration is significantly exacerbated, leading to increased abrasion depth and volume loss compared to single-factor abrasion. A significant inverse relationship between the abrasion depth fractal dimension and abrasion resistance was revealed. Under CAA-FTC conditions, CG1 and CD1 exhibit increased total porosity with enlarged large pore proportions and reduced medium pores, whereas HFC1 outperforms HFC2-based concrete, showing 8.2–26.4% higher abrasion resistance and 6.5–12.0% greater nanoindentation elastic modulus in the ITZ. Regarding the deterioration factors’ influence weight, abrasion time exhibits a deterioration weight 4.8 times to 10.0 times greater than freeze–thaw cycling, making the former a dominant factor and the latter a secondary contributor. Mechanistically, freeze–thaw cycles reduce the average molecular chain length of C-S-H gel, increase harmful pores and total porosity, and degrade the ITZ’s microstructure, while abrasion causes surface-to-core physical damage and freeze–thaw cycling induces core-to-surface expansive damage. This interaction results in surface scaling, mortar spalling, and structural loosening, significantly reducing physical and mechanical properties of the concrete under study.

1. Introduction

Nearly 70% of hydraulic structures in existing large- and medium-sized water conservancy and hydropower projects, including overflow surfaces, stilling basins, and spillway tunnels, are susceptible to scouring and abrasion damage [1]. In high-altitude cold regions such as Xinjiang and Tibet in western China, the above level reaches 90%, severely compromising the long-term operational safety of these projects. For hydraulic concrete in cold regions, freeze–thaw cycles and abrasion frequently alternate, accelerating the deterioration of both mechanical properties and durability [2,3,4]. Consequently, hydraulic discharge structures in the cold regions of western China are universally subjected to complex multi-factor interactions, including not only scouring and abrasion but also alternating impacts from freeze–thaw cycles [5,6,7].
Researchers have increasingly recognized the critical role of abrasion resistance in practical hydraulic concrete engineering applications, prompting extensive investigations into concrete’s abrasion resistance performance subjected to the coupled action of abrasion and the other deterioration factors [8,9]. Regarding the coupled action of abrasion and freeze–thaw cycles (CAA-FTC), scholars have investigated the degradation patterns of abrasion resistance performance, such as mass loss and abrasion resistance strength, by optimizing the coupled testing system [10,11]. Liu examined the evolutionary dynamics of concrete performance metrics, including relative mass loss rate and abrasion resistance degradation, under simultaneous abrasion–freeze–thaw exposure using a specially designed rotating-water-flow abrasion testing apparatus. Analysis reveals that the synergistic interaction between mechanical abrasion and freeze–thaw cycling accelerates concrete durability deterioration significantly, with degradation rates reaching 200~300% of levels observed under individual deterioration mechanisms [12]. Zhu et al. developed a cyclic testing protocol combining “0.5 h abrasion + 8 rapid freeze–thaw cycles” to investigate concrete deterioration mechanisms under the coupled action of abrasion and repeated freeze–thaw cycles [13]. Employing mass loss and relative dynamic elastic modulus as degradation indicators, the study systematically evaluated progressive damage accumulation in concrete samples. Experimental observations revealed that mass loss under coupled exposure conditions substantially exceeded the additive effects of individual abrasion or freeze–thaw deterioration, indicating a pronounced synergistic degradation mechanism. Furthermore, the relative dynamic elastic modulus of coupled-exposed samples plummeted to 52% of initial values, demonstrating that the combined action significantly accelerates internal microstructure damage accumulation compared to single-factor deterioration scenarios. Jiang et al. designed a stepwise testing methodology: concrete samples were first subjected to freeze–thaw cycling, followed by abrasion resistance testing using the underwater steel-ball method [14]. The results indicated that as the number of freeze–thaw cycles increased, the overall Vickers hardness of the concrete decreased significantly. Under identical abrasion durations, both the abrasion depth and the surface fractal dimension exhibited a marked upward trend, demonstrating that freeze–thaw action significantly exacerbates the deterioration of concrete’s abrasion resistance performance. Horszczaruk established empirical correlations between concrete abrasion resistance and its material properties, proposing that these equations could enable the determination of mass loss based on the compressive strength of concrete [15,16]. Mohebi et al. conducted a comprehensive quantification of concrete abrasion deterioration mechanisms, analyzing parameters such as erosion depth, compressive strength evolution, and curing regimen effectiveness [17]. Thus, it is evident that under the CAA-FTC, a significant synergistic deterioration effect is exhibited. The resulting damage far exceeds a simple summation of the damage caused by the action of either single factor.
High-ferrite Portland cement (HFC) is a novel cement characterized by increased contents of C2S and C4AF (≥18%) and a reduced C3S content [18]. As reported in the literature, HFC demonstrates exceptional durability and superior abrasion resistance performance, attributes that underscore its promising role in advancing sustainable construction materials [19,20,21]. Previous studies have indicated that the degradation progression of concrete under the combined effect of abrasion and freeze–thaw cycling is considerably much more pronounced compared to individual deterioration mechanisms. However, existing research lacks detailed quantification of the relative influence weights between abrasion duration and freeze–thaw cycles, while also omitting critical analysis of microstructural changes in hydration products and the interfacial transition zone (ITZ). Two HFCs with different clinker composites (HFC1:C4AF = 17.75%, C3S = 45.45%; HFC2:C4AF = 15.75%, C3S = 33.80%) were selected. The mechanical properties of two HFCs concrete with added SF were investigated in this paper, including abrasion resistance and impact resistance. The 3D surface morphology of deteriorated concrete was measured using an ATOS scanning system, collecting abrasion depth and volume loss data while analyzing the fractal dimension of the abrasion depth.
The hydration products, pore structure, and microstructure characteristics of HFC were examined via mercury intrusion porosimetry (MIP) and 29Si nuclear magnetic resonance (29Si NMR). ITZ’s microscopic mechanical properties and thickness were analyzed using nanoindentation and scanning electron microscopy–energy dispersive spectroscopy (SEM-EDS), respectively. Based on these multi-scale analyses, this work innovatively establishes quantitative correlations among the fractal dimension of the damaged surface, abrasion volume loss, and residual abrasion strength. Notably, the precise influence weights of abrasion duration and freeze–thaw cycle count are quantified for the first time, revealing their distinct contributions to the coupled deterioration. Furthermore, a detailed and mechanistic relationship between the macroscopic performance degradation under CAA-FTC and the underlying microstructural evolution is rigorously derived and elucidated.

2. Experimental

2.1. Materials

The main chemical compositions of HFC1, HFC2, and silica fume (SF) were analyzed via X-ray fluorescence method. The results are listed in Table 1. The specific surface area of SF is 19,000 m2/kg, with a water demand ratio of 110%. Its specific gravity is 2230 kg/m3, and the 28-day activity index reaches 109%.

2.2. Experimental Methods

2.2.1. Mix Proportions

The concrete mix designs for evaluating surface morphology parameters of degraded concrete, abrasion resistance strength, and impact resistance energy are detailed in Table 2. All formulations maintain a constant water-to-cementitious materials ratio of 0.35, with SF incorporated at a fixed dosage of 5% by mass. The air content of the concrete was controlled within the range of 4.5% ± 0.1%. For cement paste specimens prepared to characterize MIP and 29Si NMR properties, the water-to-cementitious materials ratio was adjusted to 0.50, excluding aggregates (dosages specified in Table 2). The ITZ investigations employed mix proportions with a constant sand-to-cementitious materials-to-water mass ratio of 6:2:1, with SF content variations as per Table 2.

2.2.2. The Coupled Testing Regime for Abrasion and Freeze–Thaw Cycles

The abrasion resistance test involves placing concrete samples into an abrasion testing machine after N freeze–thaw cycles, observing mass loss every 24 h, and calculating the abrasion resistance strength. For the impact resistance test, following N freeze–thaw cycles, a 72 h abrasion test is conducted prior to the impact resistance test. The cylindrical samples subjected to the CAA-FTC had dimensions of Ø 300 mm × 100 mm. According to DL/T 5150-2017 [22], the rapid freeze–thaw resistance test is conducted on concrete samples with strict temperature control: the sample’s core temperature is maintained within the range of −20 °C to +7 °C throughout testing, and each complete rapid freeze–thaw cycle duration is precisely 4 h [23]. The coupled abrasion–freeze–thaw test protocol is detailed in Table 3.

2.2.3. Test Methods for Mechanical Properties

The compressive strength of concrete was tested in compliance with the standard DL/T 5150-2017, employing a WAW-3000 type compression testing machine from Sansi (Shenzhen, China). The result presented is the average of three specimens. Additionally, the concrete abrasion resistance test was performed using the underwater method outlined in the same standard, which aligns with the protocol specified in ASTM C1138 (2012). The concrete abrasion resistance test was performed during a 28-day curing period, with average values determined by testing three concrete specimens. Quality loss for each specimen was evaluated after 72 h of continuous abrasive exposure. The methodology for calculating concrete abrasion resistance is outlined below:
f a = T × A Δ M
where fa (h·m2·kg−1) is concrete abrasion resistance strength, T (h) denotes the abrasion duration (72 h), A (m2) represents the exposed surface area of the concrete specimen (0.07 m2), and ΔM (kg) corresponds to the mass loss of the specimen.
The concrete impact resistance was evaluated using the “drop hammer impact method” specified in ACI 544. Visual documentation of the final cracking state is provided in Figure S2, while Figure S3 illustrates the testing apparatus. The procedure involved a freely falling hammer (mass m = 7.26 kg) dropped from a height h = 1.0 m onto concrete specimens positioned on a rigid platform. Initial cracking was identified by the first observable surface cracks, and the test concluded at the final cracking state, with the total number of impacts (n) required to induce failure recorded. The final cracking impact energy was calculated using Equation (2), and the concrete’s impact resistance performance was assessed based on the magnitude of W:
W = m × g × h × n
where W (J) is the final cracking impact energy, m (kg) is the mass of the heavy hammer, h (m) is the height from which the hammer is dropped to impact the concrete specimens on a platform, g (m/s2) is the acceleration due to gravity, and n is the number of hammer drops. The method is conducted according to the reference [24].

2.2.4. Surface Morphology of Deteriorated Concrete

The 3D surface morphology analysis of concrete abrasion-resistant specimens was entrusted to MarSoftware Technology Development (Shanghai, China) Co., Ltd., Chengdu Branch. The measurement system employed an ATOS Q 12M scanning head with MV170 lenses, which were also supplied by MarSoftware Technology Development. The ATOS scanning system comprises two left–right symmetrical industrial-grade CCD cameras, each featuring 8-megapixel resolution, and achieves data acquisition at a 0.1 cm × 0.1 cm sampling interval. The measurement principle involves projecting fringe patterns onto the specimen using a projection device, capturing the resulting images with CCD cameras, and processing the data through specialized software to generate a 3D point cloud of the specimen’s surface. Refer to Figure 1 for the schematic diagram of the 3D scanning measurement principle and camera configuration.

2.2.5. Microanalyses

(1) Sample preparation for microanalyses
Firstly, the hydration of hardened cement pastes was terminated by the absolute ethyl alcohol at a defined curing age. Subsequent sample preparation for microstructural characterization involved mechanical comminution followed by vacuum drying at 40 °C for 72 h. The dried, crushed specimens underwent agate mortar grinding to achieve particle sizes <80 μm, ensuring that NMR analysis is facilitated. For MIP and SEM + EDS analyses, cylindrical specimens with dimensional specifications of 5 mm~8 mm in diameter were carefully selected.
The samples preparation for ITZ’s characteristics analyses was according to the author’s previous study. For the detailed procedure, see the work of Lv et al. [24].
(2) MIP
The pore structure characterization of the hydrated cement paste was conducted using an Auto Pore IV 9500 mercury intrusion porosimeter (Micromeritics Co., Ltd., Norcross, GA, USA). The instrument enables quantification of pore size distribution and porosity parameters across a measurement range spanning 2 to 360,000 nm [25].
(3) 29Si MAS NMR
The 29Si NMR analysis of hydration products was performed using a Bruker AVANCE III spectrometer (Bruker, AVANCE III, Karlsruhe, Germany). Spectral acquisition was conducted at operating frequencies of 79.4 MHz and 104 MHz, corresponding to magic angle spinning (MAS) rates of 8 kHz and 12 kHz, respectively, under a magnetic field strength of 9.4 T. The experimental parameters included pulse durations of 4 μs and 0.5 μs, with recycle delays of 5 s and 1 s for the respective measurement configurations [26].
The 29Si NMR is formed by superposing a series of signals such as Q0, Q1, Q2 (1Al), Q2, Q3, and Q4. Their chemical shifts are approximately −71, −79, −82, −85, −103.4, and −110 ppm. In their anhydrous state, the C3S clinker and C2S clinker exist exclusively as isolated SiO4 tetrahedra, being manifested by a distinct Q0 resonance peak in 29Si NMR spectra [27,28,29]. Upon hydration initiation, the isolated SiO4 tetrahedral units within Portland cement clinker progressively undergo polymerization, transitioning into Q1 and Q2 silicate environments characteristic of C-S-H gel formation. The C-S-H gel phase forms through polymerization into linear silicate chains, where each SiO4 tetrahedron coordinates with calcium ions. Within this structure, Q1 species represent either terminal silicate units (dimers) or isolated tetrahedra within the chain, while Q2 units correspond to bridging silica tetrahedra positioned at mid-chain locations [30,31]. Q3 and Q4 are attributed to vitreous and quartz in fly ash, respectively. The mean chain length (MCL) and the hydration degree (α) for different samples are characterized by Equations (S1) and (S2) in the Supplementary Information [32,33].
(4) Nanoindentation
The nanoindentation of samples was performed using a Hysitron TI980 Tribo Indenter instrument (Bruker Co., Ltd., Karlsruhe, Germany) with a standard Berkovich probe [34]. This study selected two different indentation models: accelerated property mapping (APM) and single indentation. For the detailed procedure, see the work of Lv et al. [24].
(5) SEM + EDS
The JSM-7500F field emission scanning electron microscope (JEOL Corporation, Tokyo, Japan) equipped with an X-Max N80 energy dispersive spectroscopy detector was used to observe the distribution of ITZ. The instrument was operated at an acceleration voltage of 15 KeV. The method employing SEM + EDS to determine the width of the ITZ has been described in Supplementary Materials [24].

3. Results and Discussion

3.1. Surface Morphology Parameters of Deteriorated Concrete

3.1.1. The Surface Morphology of Deteriorated Concrete

The surface morphology of deteriorated concrete is shown in Figure 2. The abraded surface of concrete specimens exhibits a trend of higher elevation in the center and lower elevation at the periphery. This phenomenon occurs because abrasion initially initiates at the specimen periphery and progressively extends toward the center with prolonged abrasion duration. For the same specimen, the abrasion depth and damaged area under the CAA-FTC exceed those caused by abrasion alone. Specifically, this is manifested by a reduction in the central green region (indicating lower damage) and an expansion of the peripheral blue region (indicating higher damage). It can be inferred that the synergistic effect of abrasion–freeze–thaw coupled action exacerbates surface abrasion damage to the concrete [35].

3.1.2. The Abrasion Depth and Volume Loss

As shown in Figure 3a, under the CAA-FTCs, CG1 and CD1 exhibit average abrasion depths increased by 0.52 mm and 0.10 mm respectively compared to single-abrasion conditions. Concurrently, their maximum abrasion depths show increments of 2.87 mm and 1.20 mm under the same coupling conditions. The volume losses under coupling effects increased by 32.5% and 10.5% for CG1 and CD1, respectively. These findings indicate that both abrasion depth and volume loss under the CAA-FTCs are significantly greater than those under single-abrasion conditions, demonstrating that the coupling effect exacerbates the deterioration of concrete’s abrasion resistance [36].
As shown in Figure 3b, when comparing the different samples under single-abrasion conditions, CG1 demonstrates a 1.22 mm reduction in average abrasion depth, a 2.37 mm reduction in maximum abrasion depth, and 53.0% less volume loss compared to CD1. Under the CAA-FTCs, CG1 maintains performance advantages with 1.25 mm lower average abrasion depth, 1.39 mm reduced maximum abrasion depth, and 27.7% lesser volume loss compared to CD1. These findings indicate that HFC1 exhibits a consistently superior durability performance when subjected to various degraded environmental conditions.

3.1.3. The Fractal Dimension of Abrasion Depth

This study systematically examines the spatial irregularity of erosion surface topography through fractal dimension analysis, employing a cube covering algorithm to quantify the fractal characteristics of concrete specimen degradation under multi-factorial coupled environmental conditions [37,38,39]. The principle of the cube covering method to calculate the fractal dimension of concrete abrasion surface is as shown in Figure 4.
In the planar coordinate system XOY, a square grid ABCD with side length δ is defined, where the four corner points (g, h), (g, h + 1), (g + 1, h), and (g + 1, h + 1) correspond to four spatial points on the concrete abrasion surface. These points have respective heights of L(g, h), L(g + 1, h), L(g, h + 1) and L(g + 1, h + 1) (with constraints g ≥ 1, h ≤ m − 1, where m denotes the number of measurement points along each edge). To quantify surface complexity, a cubic grid with side length δ is employed to cover the abrasion surface within the region bounded by ABCD. The calculation specifically determines the number of δ × δ × δ cubes required to fully encapsulate the topographic variations of the abrasion surface across this domain. The calculation equation is as follows:
Ng,h = INT{δ−1[max(L(g, h), L(g + 1, h), L(g, h + 1), L(g + 1, h + 1)) − min(L(g, h), L(g + 1, h), L(g, h + 1), L(g + 1, h + 1))] + 1}
where INT denotes the integer function.
The total number of cubes required for surface coverage is denoted as N(δ), and its calculation formula is provided in Equation (4).
N ( δ ) = ∑ g , h = 1 n − 1 N g , h
By adjusting the value of δ to cover the rough surface, the total number of required cubes, N(δ), is obtained. As δ decreases, the coverage approaches closer to the actual abrasion surface conditions. However, when δ exceeds a certain threshold, the surface no longer exhibits fractal characteristics. If the abrasion depth of the surface under combined abrasion and multi-factor coupling effects satisfies the fractal condition, the scaling parameter δ and the total cube count N(δ) follow the relationship defined in Equation (5).
ln[N(δ)] = −D × ln(δ) + lnC
where D represents the fractal dimension of the rough surface, and C is the fitting parameter.
Using a 3D scanner to acquire abrasion depth data, the relationship between the total cube count N(δ) and the scaling parameter δ is established through Matlab 2016b version programming calculations based on Equations (3)–(5). By plotting ln(δ) on the horizontal axis and ln[N(δ)] on the vertical axis, linear fitting is performed. The negative slope of the resulting straight line corresponds to the fractal dimension of the surface [40].
The fractal dimensions of concrete under the CAA-FTC are presented in Figure 5 and Table 4. As shown in Figure 5, there exists a robust linear relationship between the total cube count N(δ) and the size scale δ, with correlation coefficients around 0.999. For the same specimen, the fractal dimension under the CAA-FTC is higher than that under single-abrasion conditions. Specifically, the fractal dimensions of CG1-F200A72 and CD1-F200A72 are 0.0122 and 0.0060 higher than those of CG1-F0A72 and CD1-F0A72, respectively. This indicates that the CAA-FTC induces greater surface irregularity on eroded concrete compared to single-abrasion conditions, further exacerbating concrete abrasion damage. The coupling effect results in larger abrasion damage areas, increased surface roughness, and greater abrasion damage depth. A higher fractal dimension corresponds to more severe surface deterioration of the concrete, which aligns with findings from other researchers [41,42].

3.2. Mechanical Properties

3.2.1. The Abrasion Resistance Strength

The abrasion resistance of different concretes subjected to the CAA-FTC was investigated, with detailed results presented in Figure 6. Comparing different cement types under identical freeze–thaw cycles and abrasion durations, the abrasion resistance of CG1 concrete exceeds that of CD1 by values from 1.08 h/(kg/m2) to 10.6 h/(kg/m2). This indicates that concrete prepared with HFC1 exhibits superior abrasion resistance compared to HFC2 concrete under CAA-FTC conditions. For the same sample, the abrasion resistance remains relatively consistent under identical abrasion durations but varying numbers of freeze–thaw cycles. For instance, CG1 samples subjected to 24 h abrasion exhibit abrasion resistance values ranging from 25.8 to 35.4 h/(kg/m2) when exposed to 0~200 freeze–thaw cycles. However, significant variations in abrasion resistance are observed after an identical number of freeze–thaw cycles but different abrasion durations. Specifically, CG1 samples subjected to 200 freeze–thaw cycles exhibited abrasion resistance strength ranging from 5.3 to 25.8 h/(kg/m2) at abrasion durations from 24 to 72 h. These findings suggest that abrasion time has a considerably greater impact on concrete abrasion resistance compared to the number of freeze–thaw cycles. However, the influence weights of both factors need further investigation.

3.2.2. The Impact Resistance Performance

The impact resistance of concrete under the CAA-FTC is presented in Figure 7, with detailed test procedures shown in Figure S3. As illustrated in Figure 7, after 100 freeze–thaw cycles followed by 24 h abrasion testing, the final crack impact energies of CG1 and CD1 decreased by 92.3% and 91.0% respectively, indicating a rapid attenuation of concrete impact resistance performance. Furthermore, the final crack impact energies of CG1 after 50 and 100 freeze–thaw cycles are 2.13 kJ and 0.14 kJ higher than those of CD1, respectively. These results demonstrate that concrete prepared with HFC1 exhibits a superior impact resistance under the CAA-FTC. As shown in Figure S3, following 100 freeze–thaw cycles, CG1 and CD1 specimens exhibit a “brittle” condition. Upon impact testing, CG1 specimen is fractured into three pieces from the center after 13 drops, while CD1 specimens are fractured similarly after 11 drops. This indicates that the concrete experiences significant degradation in structural integrity and impact resistance when subjected to the CAA-FTC.

3.3. Hydration Products Characteristics

3.3.1. MIP Analyses

In this paper, the pores in the hardened cement pastes are divided into gel pores (d < 10 nm), medium pores (10 nm < d < 50 nm), and large pores (50 nm < d < 1000 nm) [43]. Pore sizes below 10 nm are generally believed to be inherent in C-S-H gel. The medium and large pores are residues between the partially hydrated cement grains in a water-filled space. The medium pores belong to the harmless pores [44,45]. Figure 8 presents the MIP test results for concrete subjected to CAA-FTC. As shown in Figure 8a, the total porosity exhibits an increasing trend after CAA-FTC, with CG1 and CD1 showing increases of 0.2% and 3.6%, respectively. Figure 8b reveals that after CAA-FTC, the proportion of large pores (50–10,000 nm) in CG1 and CD1 increases by 5% and 6%, respectively, while medium pores (10–50 nm) decrease by 3% and 7%. The quantity of gel pores (<10 nm) remains relatively unchanged. These changes indicate that the CAA-FTC significantly increases total porosity and the proportion of large pores, with CD1 exhibiting a comparatively more pronounced porosity increase than CG1. This phenomenon confirms the previously observed results that CG1 demonstrates superior abrasion resistance and impact performance compared to CD1 after the CAA-FTC.

3.3.2. NMR

Figure 9 presents the 29Si NMR spectra deconvolution results for concrete after CAA-FTC. In addition, the deconvolution results are detailed in Table 5. The primary distinction in the 29Si NMR results after the CAA-FTC manifests in the Mean Chain Length (MCL), referring to the degree of silicon–oxygen framework polymerization, in contrast to the single-abrasion conditions. The MCL of concrete exhibits a decreasing trend after the CAA-FTC. Specifically, the MCL values in CG1 and CD1 decreased by 1.04 and 0.21, respectively. This reduction in the average MCL of C-S-H gel leads the rapid deterioration of the concrete’s mechanical properties after exposure to the CAA-FTC. However, the coupling effect exerted no significant influence on Al [4]/Si, which corresponds to tetrahedrally coordinated aluminum within a silicon-containing matrix.

3.4. ITZ Characteristics

3.4.1. Nanoindentation

The distribution of microscopic mechanical properties of different samples is presented in Figure 10. Regarding the CG1-F0A72 sample, the indentation modulus of cement pastes mainly ranged from 11.2 to 41.1 GPa, and the indentation modulus of ITZ mainly ranged from 41.1 to 82.2 GPa, while those of aggregate varied from 86.8 to 124.7 GPa (see Figure 10a). Regarding the CG1-F200A72 sample, the indentation moduli of cement pastes mainly ranged from 14.6 to 35.7 GPa, while those of ITZ and aggregate varied from 42.0 to 54.4 GPa and from 57.0 to 92.1 GPa, respectively, as shown in Figure 10b. According to Figure 10c, the indentation modulus in CD1-F0A72 ranged from 14.0 to 39.8 GPa for cement pastes, from 40.5 to 70.8 GPa for ITZ, and from 72.4 to 126.8 GPa for aggregate. As seen from Figure 10d, the indentation modulus in CD1-F200A72 ranged from 17.5 to 36.2 GPa for cement paste, from 39.1 to 55.6 GPa for ITZ, and from 58.4 to 120.7 GPa for aggregate. The significant dispersion observed in the ITZ data primarily results from the inherent heterogeneity of the ITZ in concrete, combined with local variations in aggregate distribution and porosity encountered during micro-measurements. Such variability is commonly observed in the characterization of ITZ.
To analyze the ITZ thickness, contour maps depicting the spatial distribution of indentation modulus and hardness were generated for various samples, as illustrated in Figure 11. The key parameters of nanoindentation in the ITZ are presented in Figure 12. Figure 12a illustrates the evolution of ITZ properties under the CAA-FTC conditions. As shown in Figure 12a, the average indentation modulus of the ITZ decreases after the CAA-FTC, with CG1 and CD1 exhibiting reductions of 5.0 GPa and 1.8 GPa, respectively. This indicates deterioration of ITZ mechanical properties under the coupling effect, contributing to the observed degradation of concrete abrasion resistance and impact performance. In addition, when comparing different cement, the average indentation modulus of the ITZ for CG1-F0A72 and CG1-F200A72 was higher by 6.6 GPa and 3.4 GPa than those of CD1-F0A72 and CD1-F200A72, respectively.
Figure 12b further demonstrates that the concrete subjected to the CAA-FTC exhibited an increasing trend in the average width of ITZ. Specifically, the average widths of ITZ of CG1 and CD1 increased by 1.51 μm and 1.95 μm, respectively. This indicates that the CAA-FTC promoted the ITZ widening. When comparing different cements, average ITZ widths of CD1-F0A72 and CD1-F200A72 exceeded CG1-F0A72 and CG1-F200A72 by 1.03 μm and 1.48 μm, respectively. Consequently, when subjected to CAA-FTC conditions, HFC1 concrete exhibited notably enhanced ITZ microstructural properties.

3.4.2. SEM + EDS

The ITZ width of concrete subjected to the CAA-FTC was investigated using SEM-EDS line scanning mode, with specific results shown in Figure 13, Figure 14, Figure 15 and Figure 16. After the CAA-FTC, the average ITZ width of concrete exhibited an increasing trend. The average ITZ widths of CG1 and CD1 increased by 7.7 μm and 7.2 μm, respectively. When compared with the average ITZ width under single-abrasion conditions, the average width of the ITZ in concrete showed a distinct upward trend. The average ITZ width of CD1-F200A72 was 3.0 μm wider than that of CG1-F200A72. This further demonstrates that HFC1 concrete exhibits superior ITZ microstructure under the CAA-FTC, resulting in better abrasion resistance and impact resistance compared to HFC2.

4. Discussion

4.1. Relationship Among Fractal Dimension, Abrasion Volume and Abrasion Strength

The relationships linking fractal dimension, abrasion volume loss, and abrasion strength were investigated, with specific results shown in Figure 17. As depicted in Figure 17a, a positive correlation exists between abrasion volume loss and fractal dimension, with a correlation coefficient of 0.891. As abrasion volume loss increased, the fractal dimension exhibited a gradual upward trend. As shown in Figure 17b, it is evident that fractal dimension and abrasion resistance demonstrate a negative correlation, with a correlation coefficient of 0.924. Specifically, as the fractal dimension increased, the abrasion resistance showed a corresponding gradual decrease. Consequently, in practical engineering applications, when direct assessment of concrete abrasion conditions and abrasion resistance is not possible, the fractal dimension and abrasion volume loss can serve as indirect indicators to estimate the abrasion resistance performance of concrete.

4.2. The Influence Weights of Abrasion Time and Freeze–Thaw Cycle

To better analyze the influence weights of abrasion time and freeze–thaw cycle on the abrasion resistance of concrete, a nonlinear surface Poly 2D function was employed to fit the relationship between the combined factors of abrasion time and freeze–thaw cycle and the concrete abrasion resistance, with the fitting equation presented as Equation (6) [46].
fa = Z0 + a × X + b × Y + c × X2 + d × Y2 + f × X × Y
where fa (h·m2·kg−1) represents the abrasion resistance strengths, Z0 (h·m2·kg−1) is a constant, a (h·m2·kg−1) represents the influence weight for the single factor of freeze–thaw cycles, X is the normalized factor for freeze–thaw cycles, b (h·m2·kg−1) represents the influence weight for the single factor of abrasion time, Y is the normalized factor for abrasion time, c (h·m2·kg−1) is the influence weight for the dual-factor interaction between freeze–thaw cycles, d (h·m2·kg−1) is the influence weight for the dual-factor interaction between abrasion times, and f (h·m2·kg−1) is the influence weight for the dual-factor interaction between abrasion time and freeze–thaw cycles.
The fitting equations for abrasion resistance strengths are presented as Equations (7) and (8), with the fitted surface plots shown in Figure 18. As seen in Figure 18, the abrasion resistance strength of concrete is jointly influenced by abrasion time and freeze–thaw cycle, exhibiting a three-dimensional decreasing trend with increases in both factors. When the abrasion time is held constant, the abrasion resistance strength remains relatively stable with the increase in the number of freeze–thaw cycles. However, when freeze–thaw cycle count is maintained, the abrasion resistance strength shows a rapid decline as abrasion time increases [47].
According to Equation (7), the correlation coefficient reaches 0.999, indicating that the Poly 2D function provides an excellent fit for the variation in abrasion resistance strength under the CAA-FTC. The influence weight of the single factor of freeze–thaw cycles is −14.1 (h·m2·kg−1), while that for abrasion time is −140.3 (h·m2·kg−1). This quantitative relationship confirms that abrasion time is the dominant factor, and freeze–thaw cycles play a secondary role, under the CAA-FTC.
CG1: fa = 73.6 − 14.1 × X − 140.3 × Y + 1.96 × X2 + 76.5 × Y2 + 7.8 × X × Y, R2 = 0.999
CD1: fa = 52.3 − 19.3 × X − 92.4 × Y + 3.57 × X2 + 47.6 × Y2 + 12.9 × X × Y, R2 = 0.994
It indicates that under the CAA-FTC, the deterioration weight of abrasion time is 4.8 times to 10.0 times greater than that of freeze–thaw cycles. The abrasion time is the primary factor affecting concrete performance degradation, while freeze–thaw cycles are a secondary factor. These primary and secondary factors interact and jointly contribute to the degradation of concrete performance.

4.3. Degradation Mechanism of Concrete Under the CAA-FTC

Under the CAA-FTC, the performance degradation mechanism of HFC concrete is as follows: the freeze–thaw cycles induce a reduction in the average MCL of C-S-H gel in HFC, accompanied by an increase in harmful pores and total porosity, as well as rapid microstructure degradation in the ITZ. The abrasion behavior causes surface-to-core physical damage, while freeze–thaw cycles result in core-to-surface expansive damage, leading to phenomena such as surface scaling, mortar spalling, and structural loosening, which collectively cause a significant reduction in the physical and mechanical properties of concrete. Subsequent abrasion events further exacerbate this concrete damage. It demonstrates that concrete prepared with HFC1 exhibits superior comprehensive performance under the CAA-FTC. Specifically, the abrasion resistance of HFC1 is 8.2~26.4% higher than that of HFC2, and the nanoindentation elastic modulus of the ITZ in HFC1 is 6.5~12.0% greater than that in HFC2. The total porosity exhibits an increasing trend after the CAA-FTC, with HFC1 and HFC2 showing increases of 0.2% and 3.6%, respectively [48].

5. Conclusions

The main conclusions are as follows:
(1) The CAA-FTC conditions significantly exacerbate concrete deterioration, increasing both abrasion depth and volume loss compared to single-abrasion conditions. Furthermore, HFC1 consistently demonstrates superior durability performance over HFC2, showing lower abrasion depths and volume losses under both single and coupled degradation conditions. Under CAA-FTC conditions, the total porosity of CG1 and CD1 increased by 0.2% and 3.6%, respectively, accompanied by a rise in large pore proportion of 5% and 6% and a concurrent decrease in medium pores of 3% and 7%.
(2) The fractal dimension of abrasion depth and abrasion resistance exhibit a significant inverse relationship, with abrasion resistance showing a progressive decline as fractal dimension increases. For practical engineering applications assessing concrete abrasion conditions, where abrasion resistance is not directly measurable, fractal dimension can serve as an indirect indicator to estimate abrasion resistance performance.
(3) Regarding the influence weights of abrasion time and freeze–thaw cycle, under the CAA-FTC conditions, the deterioration weight of abrasion duration is 4.8 times to 10.0 times greater than that of freeze–thaw cycles. The abrasion time constitutes the dominant factor influencing concrete performance degradation, with freeze–thaw cycling acting as a secondary contributor. These primary and secondary factors synergistically interact, collectively driving the deterioration process of concrete properties.
(4) Regarding the degradation mechanism of concrete under the CAA-FTC, the freeze–thaw cycles reduce the average MCL of C-S-H gel, increase harmful pores and total porosity, and rapidly degrade the ITZ’s microstructure. Under CAA-FTC conditions, HFC1 exhibits superior overall performance. Specifically, it demonstrates 8.2% to 26.4% higher abrasion resistance and 6.5% to 12.0% greater nanoindentation elastic modulus in the ITZ compared to concrete prepared with HFC2. Simultaneously, abrasion causes physical damage progressing surface-to-core, while freeze–thaw cycles induce expansive damage propagating core-to-surface. This interaction leads to surface scaling, mortar spalling, and structural loosening, resulting in a significant decline in the concrete’s physical and mechanical properties, which is further accelerated by subsequent abrasion events.

Supplementary Materials

The following supporting information can be downloaded at: https://www.mdpi.com/article/10.3390/ma19051044/s1, Figure S1: Element distribution in ITZ measured by SEM-EDS spot scanning mode; Figure S2: Appearance of final crack state of concrete; Figure S3: Concrete impact resistance testing instrument; Figure S4: The impact resistance test process of normal abrasion concrete.

Author Contributions

X.L.: Conceptualization, Investigation, Funding acquisition, Methodology, Writing—Original draft preparation. Y.D.: Funding acquisition, Investigation. Z.F.: Investigation. All authors have read and agreed to the published version of the manuscript.

Funding

This research was funded by the Hubei Provincial Young Science and Technology Talents Cultivation Program [No. 2025DJA075], the National Natural Science Foundation of China [No. 52109147 and 52179122] and the Fundamental Research Funds for Central Public Welfare Research Institutes [Grant No. CKSF20241023/CL and CKSF2023312+CL].

Data Availability Statement

The original contributions presented in this study are included in the article/Supplementary Material. Further inquiries can be directed to the corresponding author.

Conflicts of Interest

The authors declare no conflicts of interest.

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Materials 19 01044 g001
Figure 1. The principles of 3D scanning measurement and camera appearance.
Figure 1. The principles of 3D scanning measurement and camera appearance.
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Materials 19 01044 g002
Figure 2. 3D scanning pictures for different samples: (a) CG1−F0A72; (b) CG1−F200A72; (c) CD1−F0A72; (d) CD1−F200A72.
Figure 2. 3D scanning pictures for different samples: (a) CG1−F0A72; (b) CG1−F200A72; (c) CD1−F0A72; (d) CD1−F200A72.
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Materials 19 01044 g003
Figure 3. The abrasion depth and the abrasion volume for different samples: (a) The abrasion depth of CG1 and CD1; (b) the abrasion volume of CG1 and CD1.
Figure 3. The abrasion depth and the abrasion volume for different samples: (a) The abrasion depth of CG1 and CD1; (b) the abrasion volume of CG1 and CD1.
Materials 19 01044 g003
Materials 19 01044 g004
Figure 4. Schematic diagram of the cube covering method to calculate the fractal dimension of concrete abrasion surface [13].
Figure 4. Schematic diagram of the cube covering method to calculate the fractal dimension of concrete abrasion surface [13].
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Materials 19 01044 g005
Figure 5. Fractal dimension of the abrasion resistance concrete: (a) CG1−F0A72; (b) CG1−F200A72; (c) CG1−F0A72; (d) CG1−F200A72.
Figure 5. Fractal dimension of the abrasion resistance concrete: (a) CG1−F0A72; (b) CG1−F200A72; (c) CG1−F0A72; (d) CG1−F200A72.
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Materials 19 01044 g006
Figure 6. The abrasion resistance strengths of concrete under the CAA-FTC: (a) CG1; (b) CD1.
Figure 6. The abrasion resistance strengths of concrete under the CAA-FTC: (a) CG1; (b) CD1.
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Materials 19 01044 g007
Figure 7. The impact energy of concrete after different numbers of freezing–thawing cycles.
Figure 7. The impact energy of concrete after different numbers of freezing–thawing cycles.
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Materials 19 01044 g008
Figure 8. MIP test results on concrete after CAA-FTC: (a) Total porosity; (b) the percentages of pore size distribution.
Figure 8. MIP test results on concrete after CAA-FTC: (a) Total porosity; (b) the percentages of pore size distribution.
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Materials 19 01044 g009
Figure 9. 29Si NMR test results of concrete after coupling abrasion and freezing–thawing cycles: (a) CG1; (b) CD1.
Figure 9. 29Si NMR test results of concrete after coupling abrasion and freezing–thawing cycles: (a) CG1; (b) CD1.
Materials 19 01044 g009
Materials 19 01044 g010
Figure 10. Distribution of microscopic mechanical properties of different samples: (a) CG1-F0A72; (b) CG1-F200A72; (c) CD1-F0A72; (d) CD1-F200A72.
Figure 10. Distribution of microscopic mechanical properties of different samples: (a) CG1-F0A72; (b) CG1-F200A72; (c) CD1-F0A72; (d) CD1-F200A72.
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Materials 19 01044 g011
Figure 11. Contour map of indentation results in ITZ of different samples: (a) CG1-F0A72; (b) CG1-F200A72; (c) CD1-F0A72; (d) CD1-F200A72.
Figure 11. Contour map of indentation results in ITZ of different samples: (a) CG1-F0A72; (b) CG1-F200A72; (c) CD1-F0A72; (d) CD1-F200A72.
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Materials 19 01044 g012
Figure 12. Key parameters of the ITZ nanoindentation: (a) Indentation modulus; (b) ITZ’s width.
Figure 12. Key parameters of the ITZ nanoindentation: (a) Indentation modulus; (b) ITZ’s width.
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Materials 19 01044 g013
Figure 13. The ITZ width of CG1-F0A72 measured by SEM-EDS line scanning mode.
Figure 13. The ITZ width of CG1-F0A72 measured by SEM-EDS line scanning mode.
Materials 19 01044 g013
Materials 19 01044 g014
Figure 14. The ITZ width of CG1-F200A72 measured by SEM-EDS line scanning mode.
Figure 14. The ITZ width of CG1-F200A72 measured by SEM-EDS line scanning mode.
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Materials 19 01044 g015
Figure 15. The ITZ width of CD1-F0A72 measured by SEM-EDS line scanning mode.
Figure 15. The ITZ width of CD1-F0A72 measured by SEM-EDS line scanning mode.
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Materials 19 01044 g016
Figure 16. The ITZ width of CD1-F200A72 measured by SEM-EDS line scanning mode.
Figure 16. The ITZ width of CD1-F200A72 measured by SEM-EDS line scanning mode.
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Materials 19 01044 g017
Figure 17. Relationship among fractal dimension, abrasion volume loss and abrasion strength: (a) Abrasion volume and fractal dimension; (b) fractal dimension and abrasion strength.
Figure 17. Relationship among fractal dimension, abrasion volume loss and abrasion strength: (a) Abrasion volume and fractal dimension; (b) fractal dimension and abrasion strength.
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Materials 19 01044 g018
Figure 18. The fitting curve of the abrasion resistance strengths: (a) CG1; (b) CD1.
Figure 18. The fitting curve of the abrasion resistance strengths: (a) CG1; (b) CD1.
Materials 19 01044 g018
Table 1. The main chemical composition of HFC, FA, and SF (wt.%).
Table 1. The main chemical composition of HFC, FA, and SF (wt.%).
SampleCaOSiO2Al2O3Fe2O3MgOSO3Na2OK2OIg. Loss
HFC162.1321.324.655.841.872.010.13 0.39 0.78
HFC261.2522.934.175.283.142.240.16 0.32 0.61
SF0.3896.640.390.050.360.120.22 0.57 1.15
Table 2. The mix proportions of concrete for measuring mechanical properties.
Table 2. The mix proportions of concrete for measuring mechanical properties.
SampleWater/Cementitious Materials RatioConcrete Material Consumption (kg/m3)The Measured
Parameters
WaterCementSFSuperplasticizerFine AggregateCoarse Aggregate
CG10.35140402213.386061240Abrasion resistance strength,
impact resistance energy
CD1
Table 3. The CAA-FTC testing regime.
Table 3. The CAA-FTC testing regime.
SampleThe CAA-FTC Testing Procedure
① The Number of Freeze–Thaw Cycles② Abrasion Duration/Hours
F0A024, 48. 72
F50A5024, 48. 72
F100A10024, 48. 72
F200A20024, 48. 72
Table 4. Fractal dimension of different surfaces after abrasion.
Table 4. Fractal dimension of different surfaces after abrasion.
SampleFitting FormulaFractal Dimension
CG1-F0A72ln[N(δ)] = −2.0460 × ln(δ) + 11.3352.0460
CG1-F200A72ln[N(δ)] = −2.0582 × ln(δ) + 11.3832.0582
CD1-F0A72ln[N(δ)] = −2.0554 × ln(δ) + 11.3282.0554
CD1-F200A72ln[N(δ)] = −2.0614 × ln(δ) + 11.3252.0614
Table 5. 29Si NMR spectra deconvolution results for concrete after CAA-FTC.
Table 5. 29Si NMR spectra deconvolution results for concrete after CAA-FTC.
SampleQn Relative Intensity (%)MCLAl [4]/SiHydration Degree (%)
Q0Q1Q2(1Al)Q2Q3 + Q4
CG1-F0A7215.16.39.227.042.35.480.1173.8
CG1-F200A7214.49.811.225.738.94.440.1276.5
CD1-F0A7231.27.810.123.627.34.630.1257.0
CD1-F200A7231.610.812.227.817.64.420.1261.7
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Lv, X.; Dong, Y.; Fan, Z. Mechanical Performance and Microstructure Evolution of High-Ferrite Portland Cement Concrete Under the Coupled Abrasion and Freeze–Thaw Cycling Conditions. Materials 2026, 19, 1044. https://doi.org/10.3390/ma19051044

AMA Style

Lv X, Dong Y, Fan Z. Mechanical Performance and Microstructure Evolution of High-Ferrite Portland Cement Concrete Under the Coupled Abrasion and Freeze–Thaw Cycling Conditions. Materials. 2026; 19(5):1044. https://doi.org/10.3390/ma19051044

Chicago/Turabian Style

Lv, Xingdong, Yun Dong, and Zeyu Fan. 2026. "Mechanical Performance and Microstructure Evolution of High-Ferrite Portland Cement Concrete Under the Coupled Abrasion and Freeze–Thaw Cycling Conditions" Materials 19, no. 5: 1044. https://doi.org/10.3390/ma19051044

APA Style

Lv, X., Dong, Y., & Fan, Z. (2026). Mechanical Performance and Microstructure Evolution of High-Ferrite Portland Cement Concrete Under the Coupled Abrasion and Freeze–Thaw Cycling Conditions. Materials, 19(5), 1044. https://doi.org/10.3390/ma19051044

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