Dynamic Splitting Tensile Behavior of Rubber-Toughened Ceramsite Concrete for Transmission Structure Foundations Under a Wide Range of Strain Rates

Abstract

To address the impact-induced damage to concrete pile foundations of transmission structures caused by nearby blasting vibrations, this study investigates the dynamic splitting tensile behavior of an environmentally friendly lightweight rubberized concrete—Rubber-Toughened Ceramsite Concrete (RTCC)—under impact loading. Quasi-static tests show that the static splitting tensile strength increases first and then decreases with increasing rubber content, reaching a maximum value of 2.01 MPa at a 20% replacement ratio. Drop-weight impact tests indicate that RTCC20 exhibits the highest peak impact force (42.48 kN) and maximum absorbed energy (43.23 J) within the medium strain-rate range. Split Hopkinson Pressure Bar (SHPB) tests further demonstrate that RTCC20 shows the highest strain-rate sensitivity. Overall, RTCC with 20% rubber content provides the best comprehensive performance, achieving a favorable balance between strength and toughness across the entire strain-rate range. These findings offer experimental support for applying RTCC to blast-vibration-resistant transmission structure foundations.

1. Introduction

With the continuous expansion of China’s power grid and the renewal of infrastructure, blasting operations around transmission structures have become increasingly frequent. Although blasting technology plays an important role in mining, tunneling, and urban infrastructure construction, the transient shock waves and ground vibrations generated during blasting inevitably propagate to the surroundings, posing severe challenges to the foundations of nearby transmission structures. The concrete pile foundations of transmission structures are primarily subjected to long-term stable compressive loads; however, the tower swaying and ground vibrations induced by blasting can generate tensile stresses in the underlying piles, leading to microcrack initiation, crack propagation, and even splitting failure within the pile concrete, thereby affecting the overall stability and safe operation of the structures [1,2,3]. In recent years, lightweight aggregate concrete (LAC) has been increasingly applied in pile foundation engineering to reduce foundation self-weight, improve construction efficiency, and lower costs [4,5]. Compared with ordinary concrete, LAC exhibits lower density and a higher strength-to-weight ratio, which can effectively reduce foundation stress and improve construction conditions for pile foundations [6].

Ceramsite aggregate concrete (CAC), a type of LAC using lightweight ceramsite as aggregate, has been widely adopted in recent years due to its low density, high porosity, and excellent seismic performance. It is increasingly used in high-rise structures and foundation weight-reduction projects [7]. In addition, compared with other types of LAC, ceramsite aggregate (CA) as a coarse aggregate not only reduces the self-weight while maintaining adequate strength [8], but also dissipates energy through its microporous structure by attenuating stress waves [9,10,11]. The construction industry has also achieved dual benefits by recycling ceramic waste materials for use in concrete [12], which not only reduces landfill costs but also lowers material expenses. This approach promotes the transition of the construction industry toward greener and higher-performance materials, contributing to the realization of the “dual carbon” goals [13].

However, the incorporation of ceramsite aggregate (CA) cannot overcome the inherent brittleness and low ductility of concrete [14]. Under blasting impact loads, rapid crack propagation occurs, revealing its poor toughness and low energy dissipation capacity. To address this problem, recycled rubber is added into the mortar and combined with ceramsite aggregate to produce Rubber-Toughened Ceramsite Concrete (RTCC), which not only improves toughness but also further reduces self-weight [15]. Rubber is an elastic material with excellent resilience and toughening properties. Researchers have introduced lightweight rubber particles—obtained by crushing and screening waste tires—into concrete to enhance its crack resistance and impact performance [16,17]. Existing studies have shown that the inclusion of rubber can significantly improve the toughness, energy absorption, and crack resistance of concrete, enabling it to exhibit better buffering and ductility under tensile stress at high strain rates [15,18,19,20]. Feng et al. [21] found that under high strain rates, crack propagation tends to bypass rubber particles, forming branched cracks that dissipate large amounts of energy, indicating that rubber particles can effectively enhance the ductility of concrete. Yang et al. [22] reported that rubber concrete exhibits stronger strain-rate sensitivity under high-rate loading. However, due to the irregular shape of rubber particles and their hydrophobicity and low elastic modulus [23], air may be trapped around them [24], thus creating voids in the mortar. This results in a weaker bond between the rubber and the cement matrix, leading to a decrease in the strength of the rubber–cement interface transition zone (ITZ) [25], and ultimately reducing the strength of the concrete [26]. Despite the reduction in compressive strength, the incorporation of rubber promotes crack deflection, branching, and stress redistribution under static loads, thereby improving fracture toughness and energy dissipation [27]. Furthermore, the microporous structure of ceramsite provides internal space for the flow and storage of such micro-air, which may help mitigate this problem [28]. The experimental studies on rubber–lightweight aggregate concrete are summarized in

Table 1. Experimental studies on rubber–lightweight aggregate concrete.

References	Concrete Types	Test Method	Loading Focus
Ghone et al. [20]	Ceramsite lightweight concrete	Static tests	Toughness
Feng et al. [21]	Self-compacting rubberized concrete	SHPB	Dynamic splitting
Yang et al. [22]	Rubberized concrete	SHPB	Dynamic compression
Li et al. [27]	Rubberized concrete	Static tests	Strength, toughness

. Although both ceramsite and rubber can achieve weight reduction and energy dissipation to some extent, their compatibility and synergistic mechanism within the same matrix remain unclear. Therefore, it is necessary to further investigate the interaction between ceramsite and rubber particles in concrete. In general, tensile strength is the weakest property of concrete materials [29]. Hence, understanding the static and dynamic tensile behavior and failure mechanism of high-toughness lightweight concrete used in tower foundations under impact loading is essential. At present, the splitting tensile test is the most commonly used method for evaluating the tensile properties of concrete [30].

In this study, this method is employed to obtain the tensile behavior of RTCC over a wide range of strain rates. The Flat Brazilian Disc (FBD) configuration was adopted to investigate Rubber-Toughened Ceramsite Concrete (RTCC) with rubber particle replacement ratios of 0%, 10%, 20%, 30%, 40%, and 50%. These replacement levels were selected to cover a full range of low, medium, and high rubber contents, enabling a systematic examination of how increasing rubber dosage affects strength, toughness, and strain-rate sensitivity. This range has also been validated in our previous experimental studies [28], where the same intervals effectively captured the mechanical evolution of lightweight aggregate rubber concrete. Therefore, maintaining this replacement range ensures methodological consistency while providing sufficient resolution to identify the optimal rubber content. Through quasi-static loading tests, drop-weight impact tests, and Split Hopkinson Pressure Bar (SHPB) experiments, the splitting tensile strength, energy absorption, failure modes, and strain-rate sensitivity of RTCC were comprehensively evaluated. The objective was to determine the optimal rubber replacement range, reveal the dynamic crack-resistance mechanism, and provide theoretical support for the design of blast-vibration-resistant concrete materials.

2. Experimental Program

2.1. Mix Proportions and Specimen Preparation

The concrete specimens were prepared using ordinary Portland cement (42.5R), water, fine aggregate, and coarse aggregate, with a mass ratio of 398:210:609:1183. Tap water was used for mixing. The fine aggregate was medium sand with a fineness modulus of 2.60 and a bulk density of 1498 kg/m³. Rubber particles, used to replace the fine aggregate by equal volume, were produced by cutting and grinding waste tires. The measured particle size was 20 mesh (approximately 0.85 mm), and the bulk density was 417 kg/m³. The coarse aggregate was spherical ceramsite, used to replace crushed stone, with a bulk density of 1450 kg/m³ and a particle size of 5–10 mm. More detailed mix proportions are shown in

Table 2. Mix proportions of Rubber-Toughened Ceramsite Concrete (kg/m³).

	Mix Proportions
	Cement	Water	Sand	Rubber Particles	Ceramsite
CANC	398	210	609	0	756
RTCC-10	398	210	548	17	756
RTCC-20	398	210	487	34	756
RTCC-30	398	210	426	51	756
RTCC-40	398	210	365	68	756
RTCC-50	398	210	305	85	756

.

To produce RTCC, the specimens were cast in a Flat Brazilian Disc (FBD) configuration with dimensions of 100 mm × 50 mm, as illustrated in Figure 1. The casting process was carried out in the following sequence: first, the coarse aggregate and rubber particles were uniformly mixed in a mixer for approximately 45 s to ensure proper dispersion of the rubber. Then, sand and cement were added, and the mixer was restarted for another 60 s. Next, water was gradually added while mixing continued for 180 s to achieve full homogenization. Finally, the fresh concrete was poured into molds, which were placed on a vibrating table and vibrated for 20 s to ensure proper compaction. The specimens were demolded after 24 h and cured under standard conditions (temperature 20 ± 2 °C, relative humidity > 95%) for 28 days.

Previous studies have shown that for splitting tensile tests using the Flat Brazilian Disc (FBD), the recommended ratio of thickness to diameter is between 0.2 and 0.75. Moreover, when the disc thickness is half of its diameter, the influence of inertial effects on the test results can be effectively eliminated. Therefore, the dimensions of the FBD specimens were designed as 100 mm × 50 mm [31,32]. In addition, both sides of the FBD specimens were ground to form two parallel flat ends. The central angle of the flat ends was set to 2α = 20°, and the grinding depth was 0.76 mm to ensure the initiation of a central crack during testing.

2.2. Basic Principle of the Quasi-Static Test

The splitting tensile test of RTCC under low strain rates was conducted using a compression testing machine. The quasi-static tensile strength was determined by Equation (1) [33]:

$$f_{td}=0.95{\displaystyle \frac{2P_{max}}{\pi DB}}$$

(1)

where P_max is the maximum load, D is the specimen diameter, and B is the specimen thickness. Equation (2) is used to estimate the stress rate and strain rate of the specimen:

$$\dot{\sigma }={\displaystyle \frac{f_{td}}{\tau }}; \dot{\epsilon }={\displaystyle \frac{\dot{\sigma }}{\mathsf{{\rm E}}}}$$

(2)

where $\tau$ is the elastic modulus of the specimen. Previous studies have shown that the sensitivity of the elastic modulus of concrete to strain rate is much lower than that of its compressive and tensile strengths. Therefore, the static elastic modulus of the specimen is adopted. The strain-rate sensitivity of the tensile strength of concrete is characterized by the Dynamic Increase Factor (DIF), defined as the ratio of the dynamic strength to the static strength [34], which can be determined by Equation (3):

$$DIF={\displaystyle \frac{f_{td}}{f_{ts}}}$$

(3)

where f_ts is the static splitting tensile strength of the specimen, which can be calculated using Equation (4):

$$f_{ts}={\displaystyle \frac{2F}{\pi A}}$$

(4)

where F is the ultimate load applied to the specimen, and A is the loaded area of the specimen during the splitting tensile test.

2.3. Basic Principle of the Dynamic Splitting Tensile Test

2.3.1. Drop-Weight Impact Test Setup

The splitting tensile tests of RTCC under medium strain rates were carried out using a CEAST 9350 drop-weight impact tester. A schematic diagram of the drop-weight impact setup is shown in Figure 2. By varying the mass of the hammer head, the additional weights, and the drop height, different levels of impact loads were applied to the specimens to investigate their ultimate tensile strength at various (medium) strain rates. The ultimate tensile strength of the specimen was calculated using Equation (1), where P_max is the maximum impact force applied to the flat end of the specimen. The stress rate and strain rate of the specimen were determined using Equation (2).

2.3.2. Split Hopkinson Pressure Bar (SHPB) Test

The splitting tensile tests of RTCC under high strain rates were conducted using a Split Hopkinson Pressure Bar (SHPB) system, as illustrated in Figure 3. By adjusting the gas pressure (i.e., the initial velocity of the striker bar) and the distance between the striker bar and the barrel, different levels of impact loads were applied to the specimens to investigate their ultimate tensile strength under various (high) strain rates.

The force applied to the specimen can be determined using Equation (5):

$$P\left(t\right)={\displaystyle \frac{E_0{A}_0}{2}}[{\epsilon }_i(t)+{\epsilon }_r(t)+{\epsilon }_t(t)]$$

(5)

where E₀ and A₀ are the elastic modulus and cross-sectional area of the pressure bars, respectively, and ε_i(t), ε_r(t), and ε_t(t) are the incident, reflected, and transmitted strain waves, respectively.

In addition, according to the one-dimensional stress wave propagation assumption [35], this is expressed in Equation (6):

$${\sigma }_i+{\sigma }_r={\sigma }_t$$

(6)

where σ_i, σ_r, and σ_t are the stresses corresponding to ε_i(t), ε_r(t), and ε_t(t), respectively.

Based on Equations (1), (5) and (6), the dynamic ultimate tensile strength f_td obtained from the SHPB test can be determined using Equation (7):

$$f_{td}=0.95{\displaystyle \frac{2{\sigma }_{t,max}{A}_0}{\pi DB}}$$

(7)

where σ_t,max is the maximum value of σ_t.

3. Quasi-Static Test

3.1. Results of the Quasi-Static Splitting Tensile Test

The static compressive and splitting tensile properties of ordinary concrete and RTCC with different rubber volume replacement ratios (0%, 10%, 20%, 30%, 40%, and 50%) are shown in

Table 3. Static Mechanical Properties.

	Compressive Strength [28] (MPa)	Splitting Tensile Strength (MPa)	Splitting Tensile Strength/Compressive Strength
CANC	32.40	1.79	0.06
RTCC10	23.20	1.83	0.08
RTCC20	20.40	2.01	0.10
RTCC30	19.40	1.45	0.07
RTCC40	12.40	2.12	0.17
RTCC50	12.50	1.81	0.14

[28]. For each mixture, three specimens were tested and the average value was reported. As the rubber replacement ratio increases, the static compressive strength gradually decreases from 32.4 MPa (CANC) to 12.5 MPa (RTCC50). In contrast, the static splitting tensile strength of RTCC shows an initial increase followed by a decrease with increasing rubber content, reaching its maximum value of 2.01 MPa at a 20% replacement ratio and decreasing to 1.45 MPa at higher replacement levels. This is because the inclusion of a certain volume of rubber particles allows the rubber to generate tensile stress during the splitting process, thereby improving the tensile performance of the concrete [36]. At a rubber replacement ratio of approximately 40%, the splitting tensile strength exhibits a slight increase again. This phenomenon is attributed to the fact that, under tensile-dominated loading, the enhanced crack-bridging and stress redistribution effects provided by a larger number of rubber particles can partially compensate for matrix weakening, resulting in a locally improved resistance to crack propagation. However, when the rubber content further increases to 50%, the loss of matrix continuity and stiffness becomes overwhelming. The cementitious skeleton is significantly disrupted, leading to insufficient stress transfer and premature crack coalescence, and consequently, a pronounced reduction in splitting tensile strength. The brittleness index—defined as the ratio of compressive strength to tensile strength [1]—is an important parameter for evaluating the cracking resistance of concrete. Therefore, the tension–compression ratio of rubberized concrete was analyzed to further investigate the variation in the brittleness index after the incorporation of rubber particles. As shown in Figure 4, the tension–compression ratio increases overall with the addition of rubber particles. When the rubber replacement ratio is 20%, the tension–compression ratio increases by 66.7% compared with ordinary concrete, indicating a significant improvement in crack resistance. However, when the rubber content exceeds 20%, the bond between the rubber particles and the cement matrix weakens, leading to a reduction in tensile strength [37] and a subsequent decrease in the tension–compression ratio. Further increasing the rubber replacement ratio to 50% results in a much higher tension–compression ratio, but the compressive strength drops sharply. Therefore, excessive rubber content is not recommended.

3.2. Failure Modes

The failure modes of FBD specimens under quasi-static loading are shown in Figure 5. It can be observed that all specimens experienced typical splitting failure, with a main crack propagating along the loading direction, ultimately causing the specimen to split into two parts. Moreover, different rubber contents had a noticeable influence on crack morphology and failure characteristics. Compared with CANC, the specimens containing rubber particles exhibited slightly narrower crack openings and a more moderate failure pattern, indicating that the inclusion of rubber improved the crack propagation resistance and enhanced the crack-arresting capability of the material to some extent [38].

4. Dynamic Test

4.1. Drop-Weight Impact Test and Discussion

A custom flat-end hammer head with a diameter of 100 mm was used in the drop-weight device. The hammer head was assumed to be rigid. A force sensor with a capacity of 222 kN was employed. The total mass of the impactor was 9.529 kg. Before each test, the FBD specimen was raised until it just contacted the hammer head (the hammer height was 0 m at this moment) to ensure the initial impact velocity. Then, the hammer head was lifted, and the impact test was performed.

4.1.1. Failure Modes

The typical failure modes of FBD specimens under medium strain rates after the drop-weight impact test are shown in Figure 6. All specimens exhibited a main crack along the impact direction, ultimately splitting into two approximately equal parts. Meanwhile, as the strain rate increased, the extent of crack propagation became more pronounced. These failure modes are consistent with the fundamental principles of the Brazilian disc splitting test.

4.1.2. Force–Time Curve and Energy Absorption Capacity

The results of the drop-weight impact tests are summarized in

Table 4. Results of the Drop-Weight Impact Test.

	Peak Force (kN)	ftd (MPa)	Strain Rate (s−1)
CANC-1	24.71	2.99	~0.17
CANC-2	30.78	3.72	~0.23
CANC-3	33.49	4.05	~0.48
RTCC10-1	19.36	2.34	~0.25
RTCC10-2	29.69	3.59	~0.41
RTCC10-3	31.44	3.80	~0.66
RTCC20-1	34.16	4.13	~0.47
RTCC20-2	42.48	5.14	~0.56
RTCC20-3	29.83	3.61	~0.62
RTCC30-1	27.13	3.28	~0.33
RTCC30-2	34.16	4.13	~0.53
RTCC30-3	30.52	4.69	~0.81
RTCC40-1	36.20	4.38	~0.45
RTCC40-2	44.85	5.42	~0.83
RTCC40-3	42.89	5.19	~0.97
RTCC50-1	23.00	2.78	~0.28
RTCC50-2	19.76	2.39	~0.30
RTCC50-3	33.59	4.06	~0.46

. The ultimate tensile strength of the specimens was calculated using Equation (1), shows the typical force–time curve of the RTCC30 specimen under a strain rate of approximately 0.53 s⁻¹, and the stress rate and strain rate were determined using Equation (2). The Dynamic Increase Factor (DIF) of the FBD specimens under medium strain rates was obtained using Equation (3).

As shown in Figure 7, the force–time curves of specimens with different rubber contents were obtained within a strain rate range of approximately 0.45 s⁻¹ to 0.56 s⁻¹. Overall, the peak force shows a decreasing trend with the increase in rubber content. Among them, the specimen with 20% rubber content exhibited the highest peak force of 42.48 kN, which was greater than those of the other groups. In addition, the incorporation of rubber particles prolonged the impact duration, further indicating that rubber addition helps to restrain crack propagation.

As shown in Figure 8, compared with CANC, the incorporation of rubber significantly enhanced the energy absorption capacity of the material. The total absorbed energy of the CANC group was 28.32 J, while that of RTCC20 reached 43.23 J—the highest among all specimens—demonstrating superior impact energy absorption capability. The RTCC30 and RTCC40 specimens followed, with absorbed energies of 31.19 J and 33.84 J, respectively, whereas RTCC50 showed a slight decline. It is noteworthy that the overall energy absorption curves became smoother, indicating that the inclusion of rubber improved the toughness and energy dissipation mechanism during load transfer and delayed crack propagation [39]. This is because rubber particles are hydrophobic and tend to form a weak ITZ with the cementitious matrix. However, the rubber phase itself has a very low modulus and high deformability. When the main crack approaches a rubber particle, the weak interface and the large mismatch in elastic modulus cause the crack to deflect, branch, or blunt along the rubber–matrix boundary. These interactions redistribute the stress field at the crack tip, increase crack tortuosity, and require additional fracture energy for further crack propagation. Therefore, although the inclusion of rubber reduces the matrix strength, it enhances the fracture toughness and energy dissipation capacity of the composite [40].

4.1.3. Discussion

Considering both the peak impact force and the energy absorption capacity, RTCC20 exhibited the best overall performance. It maintained a relatively high peak impact force while achieving the highest absorbed energy and the longest absorption duration, demonstrating superior energy dissipation capability. Therefore, under the conditions of this study, a 20% rubber replacement ratio provided the optimal impact resistance, achieving a favorable balance between strength and toughness. The main reason is that rubberized concrete shows improved deformation accommodation under higher strain rates, which slows the accumulation of damage and enhances the material’s ability to withstand impact loading. However, as the rubber content continues to increase, a growing number of internal voids are formed, weakening the bonding between aggregates and reducing the overall mechanical performance of the mixture.

4.2. SHPB Test and Discussion

In the SHPB system, the lengths of the striker bar, incident bar, and transmission bar were 1000 mm, 5500 mm, and 3500 mm, respectively. All bars were made of silicon–manganese steel with a diameter of 100 mm, an elastic modulus E₀ = 206 GPa, and a density ${\rho }_0$ = 7710 kg/m³. The longitudinal elastic wave velocity of the bars was $C_0=\sqrt{E_0/{\rho }_0}=5169 \mathrm{m}/\mathrm{s}$. Molybdenum disulfide grease was used to minimize the friction effect between the bar ends and the specimen surfaces [41]. Four semiconductor strain gauges (SB3.8-120-P-2) were used to record the stress wave signals in the bars—two were mounted on the incident bar at a distance of 1435 mm from the specimen end, and two were mounted on the transmission bar at a distance of 2875 mm from the specimen end. In addition, two resistance strain gauges (120-5AA) were attached to the FBD specimen to measure the strain response during failure.

4.2.1. Failure Modes

The failure modes of FBD specimens under different strain rates in the SHPB tests are shown in Figure 9. Similar to the drop-weight test, a main crack propagated along the impact direction after testing, which is consistent with the principles of the Brazilian splitting test and previous research findings [42,43]. Symmetrical triangular compression zones appeared at both flat ends of the specimens, and their areas increased with the strain rate. Moreover, the triangular compression zone near the incident bar was larger than that near the transmission bar. This occurred because the material adjacent to the incident bar needed to generate more cracks to dissipate the higher impact energy. During the development of the central crack, part of the impact energy was dissipated, resulting in a smaller energy demand and, consequently, a smaller compression zone near the transmission bar.

4.2.2. Energy Absorption Capacity

A typical dynamic stress equilibrium curve obtained from the SHPB test is presented in Figure 10. The superposition of the incident and reflected waves closely matches the transmitted wave, confirming the validity of the test results. Therefore, the recorded data can be used to calculate the stress, strain, and strain rate of the specimen during the SHPB test. The results of the SHPB experiments are summarized in

Table 5. Results of the SHPB Test.

	Specimen Thickness (mm)	Stress Rate (GPa/s)	Strain Rate (s−1)	ftd (MPa)	DIF
CANC-1	45.45	27.20	1.92	5.03	2.81
CANC-2	46.85	19.34	1.36	5.51	3.10
CANC-3	45.46	30.95	2.18	6.28	3.53
CANC-4	47.24	42.36	2.98	6.48	3.64
CANC-5	44.38	11.30	0.80	3.98	2.23
RTCC10-1	44.12	11.05	0.94	4.00	2.19
RTCC10-2	46.50	30.99	2.65	5.52	3.01
RTCC10-3	47.93	35.28	3.01	7.73	4.22
RTCC10-4	45.63	55.67	4.76	8.13	4.44
RTCC10-5-1	47.00	5.19	0.44	2.61	1.43
RTCC10-5-2	44.95	45.12	3.86	7.49	4.09
RTCC20-2	48.49	49.12	4.20	8.25	4.11
RTCC20-3	50.21	42.97	3.67	7.26	3.61
RTCC20-4	45.63	23.87	2.04	5.80	2.89
RTCC20-5	44.95	15.72	1.34	5.23	2.60
RTCC20-6	42.97	27.51	2.35	6.16	3.07
RTCC30-1	51.79	11.17	0.91	4.67	3.22
RTCC30-2	49.53	9.12	0.74	4.29	2.96
RTCC30-3	47.73	27.47	2.23	5.44	3.75
RTCC30-4	45.76	40.46	3.29	8.17	5.64
RTCC30-5	45.72	49.06	3.99	6.72	4.64
RTCC40-1	49.76	10.39	0.91	4.76	2.25
RTCC40-2	48.54	30.93	2.71	5.23	2.47
RTCC40-3	49.14	27.75	2.43	5.47	2.58
RTCC40-4	48.42	45.10	3.96	8.89	4.19
RTCC40-5	49.18	50.56	4.44	8.04	3.80
RTCC50-1	47.48	18.47	1.71	4.41	2.44
RTCC50-2	50.97	21.51	1.99	4.54	2.51
RTCC50-3	48.30	23.97	2.22	4.75	2.62
RTCC50-4	48.20	39.05	3.62	6.09	3.62
RTCC50-5	47.76	38.01	3.52	6.27	3.47

, where the stress rate, strain rate, dynamic tensile strength, and DIF were calculated using Equation (2), Equation (7) and Equation (3), respectively.

In addition, the peak values between 75 μs and 260 μs in Figure 10 were selected for detailed analysis, and the stress equilibrium factor η, defined in Equation (8), was introduced to evaluate the stress equilibrium condition in the SHPB tests [44]. As shown in Figure 11, η remains almost below 0.1 throughout the time window of 75–260 μs, without pronounced fluctuations. This indicates that the specimen satisfies the assumption of uniform stress distribution during the critical impact loading stage, thereby confirming the validity of the SHPB dynamic splitting test results.

$$\mathsf{\eta }=\left|{\displaystyle \frac{{\epsilon }_i\left(t\right)+{\epsilon }_r\left(t\right)-{\epsilon }_t\left(t\right)}{\left[{\epsilon }_i\left(t\right)+{\epsilon }_r\left(t\right)+{\epsilon }_t\left(t\right)\right]}}\times {\displaystyle \frac{1}{2}}\right|$$

(8)

Based on the principle of energy conservation, the energy dissipation [45] can be defined using Equation (9):

$$\begin{gathered} W_i\left(t\right)=E_0{C}_0{A}_0{\int }_0^t{\epsilon }_i^2(t)dt \\ {W}_r\left(t\right)=E_0{C}_0{A}_0{\int }_0^t{\epsilon }_r^2(t)dt \\ {W}_t\left(t\right)=E_0{C}_0{A}_0{\int }_0^t{\epsilon }_t^2(t)dt \end{gathered}$$

(9)

As shown in

Table 6. Energy Dissipation of Specimens under the SHPB Test.

Specimen	${\mathit{W}}_{\mathit{i}}\left(\mathit{J}\right)$	${\mathit{W}}_{\mathit{r}}\left(\mathit{J}\right)$	${\mathit{W}}_{\mathit{t}}\left(\mathbf{J}\right)$	${\mathit{W}}_{\mathit{s}}\left(\mathbf{J}\right)$	${\mathit{W}}_{\mathit{s}}/{\mathit{W}}_{\mathit{i}}$
CANC-1	94.43	73.22	1.17	20.03	0.21
CANC-2	122.17	95.71	1.61	24.86	0.20
CANC-3	152.37	125.32	1.16	25.89	0.17
CANC-4	221.54	191.13	1.18	29.24	0.13
CANC-5	18.53	12.04	0.83	5.66	0.31
RTCC-10-1	18.53	12.04	0.83	5.66	0.31
RTCC-10-2	109.74	88.34	1.03	20.37	0.19
RTCC-10-3	126.14	96.51	2.04	27.58	0.22
RTCC-10-4	454.42	417.43	1.21	35.78	0.08
RTCC-10-5-1	4.22	2.22	0.54	1.46	0.35
RTCC-10-5-2	217.28	181.16	1.73	34.38	0.16
RTCC-20-2	435.30	397.26	1.42	36.62	0.08
RTCC-20-3	321.04	278.67	1.62	40.76	0.13
RTCC-20-4	85.72	63.16	1.52	21.05	0.25
RTCC-20-5	90.94	70.14	1.30	19.51	0.21
RTCC-20-6	110.52	86.78	0.00	22.47	0.20
RTCC-30-1	51.80	36.24	1.32	14.23	0.27
RTCC-30-2	90.53	71.05	0.99	18.49	0.20
RTCC-30-3	146.11	121.39	0.97	23.75	0.16
RTCC-30-4	311.82	273.63	1.47	36.34	0.12
RTCC-30-5	541.01	516.87	0.80	23.33	0.04
RTCC-40-1	15.74	9.01	1.42	5.31	0.34
RTCC-40-2	116.15	94.00	1.02	21.13	0.18
RTCC-40-4	413.67	373.90	1.47	38.30	0.09
RTCC-40-5	607.05	576.26	1.17	29.62	0.05
RTCC-50-1	91.94	74.26	0.79	16.89	0.18
RTCC-50-2	96.76	77.65	0.87	18.25	0.19
RTCC-50-3	137.82	115.03	0.67	22.12	0.16
RTCC-50-4	455.26	429.11	0.68	25.47	0.06
RTCC-50-5	539.86	515.81	0.80	23.13	0.04

, the table presents the incident, reflected, and transmitted energies of specimens with different rubber replacement ratios under various strain rates, where Wi, W_r, and W_t represent the maximum values of W_i(t), W_r(t), and W_t(t), respectively. In addition, the ratio of the dissipated energy to the incident energy, W_s/W_i, is defined as the energy dissipation ratio.

As shown in Figure 12, under the same impact energy level, the energy dissipation ratio of the material first increases and then decreases with the increase in rubber replacement ratio. Among all mixtures, the specimen with a 20% rubber replacement ratio exhibited the highest energy dissipation ratio, indicating that the concrete with 20% rubber content possesses superior crack resistance under high strain rates, which is consistent with the previous results. Moreover, all six curves show that as the incident energy increases—that is, as the strain rate rises—the energy dissipation ratio decreases. This occurs because, at high strain rates, the failure duration becomes extremely short, causing the material to behave in a more brittle manner.

4.2.3. Dynamic Splitting Tensile Behavior

As shown in Figure 13, under high strain rates, the dynamic splitting tensile behavior exhibited a trend similar to that observed in dynamic compression. The splitting tensile strength of rubber concrete showed a clear strain-rate effect [46], demonstrating significant strain-rate sensitivity across different rubber replacement ratios. It can also be observed that the slopes of the fitted lines for RTCC specimens were greater than that of CANC, indicating that the incorporation of rubber particles enhanced the strain-rate sensitivity of the concrete at high strain rates.

Moreover, as the rubber content increased, the slope of the fitted lines first increased and then decreased, while the intercepts generally showed a downward trend (for example, the intercept of RTCC10 was 2.3512, whereas that of RTCC50 was 1.57083). This indicates that in the low strain-rate region ($ln\dot{\epsilon }$ < 0), the DIF values of rubberized concrete were generally lower than those of ordinary ceramsite concrete; however, in the high strain-rate region ($ln\dot{\epsilon }$ > 0), the DIF of high rubber replacement concrete increased more significantly, showing a stronger dynamic enhancement effect than ordinary ceramsite concrete. For RTCC20, the slope of the fitted line reached 1.3182 and the intercept was 1.8993, which ensured a relatively high baseline strength while exhibiting superior dynamic enhancement under high strain rates.

4.2.4. Discussion

Combined with the results of the quasi-static compression and drop-weight impact tests, it can be concluded that the 20% rubber replacement ratio provides the best strain-rate effect and overall performance under both static and dynamic loading conditions. The quasi-static splitting tensile test results show that the mixture with 20% rubber content achieves the highest tensile strength. The drop-weight impact tests further indicate that RTCC20 reaches the highest peak impact force and absorbs the greatest amount of impact energy among all mixtures. The SHPB results also demonstrate that RTCC20 exhibits the strongest strain-rate sensitivity and attains the highest energy absorption ratio, indicating that it can dissipate a larger portion of the input energy under high strain rates. Moreover, all three tests consistently show that the overall performance of RTCC first increases and then decreases as the rubber content increases. This trend can be explained that a small amount of rubber can provide limited deformation accommodation and help reduce localized stress concentration, allowing the material to sustain higher loads before brittle failure occurs. However, as the rubber content continues to increase, the presence of more low-stiffness regions and weak interfaces begins to dominate the structural response, reducing the effective load-bearing capacity and resulting in a decline in strength.

We proposed an empirical correction formula, as expressed using Equation (10), to fit the strain rate effect for different rubber admixtures of RTCC.

$$DIF=a{\left(\ln \dot{\epsilon }\right)}^2+b\left(\ln \dot{\epsilon }\right)+c$$

(10)

As shown in Figure 14, it illustrates the relationship between the DIF and the strain rate for RTCC. Compared with ordinary concrete, all RTCC mixtures exhibit an enhanced strain-rate effect after the incorporation of rubber particles. Among them, RTCC20 shows the most pronounced strain-rate sensitivity, indicating that the influence of strain rate on the dynamic splitting tensile strength is most significant at a rubber replacement ratio of 20%. As shown in

Table 7. The fitting results of DIF and $\dot{\epsilon }$ of RTCC.

	a	b	c	R2
CANC	0.18898	0.81189	2.55607	0.93
RTCC10	0.23063	1.00359	2.32953	0.94
RTCC20	0.61331	0.36208	2.14161	0.82
RTCC30	0.28672	1.03703	3.14107	0.85
RTCC40	0.42514	0.2861	2.28215	0.85
RTCC50	0.20008	0.67249	2.12061	0.89

, when the rubber content exceeds 20%, the strain-rate effect on the dynamic splitting tensile strength of RTCC gradually weakens. This phenomenon can be attributed to the fact that rubber particles may carry a small amount of entrapped air. When the rubber content exceeds 20%, the internal void content of the concrete increases, which weakens the aggregate–matrix bonding interface. As a result, the bonding interface is more susceptible to damage under impact loading, leading to a reduction in deformation capacity and strain-rate strengthening efficiency. Nevertheless, the microporous structure of ceramsite aggregate provides internal space for the migration and storage of micro-air, which partially mitigates the strength-reduction effect induced by rubber incorporation. Overall, the results indicate that RTCC exhibits a stronger strain-rate effect than ordinary concrete, with the most significant enhancement observed at a rubber content of 20%.

5. Conclusions

This study systematically investigated the static and dynamic splitting tensile behavior of Rubber-Toughened Ceramsite Concrete (RTCC) with different rubber fine aggregate replacement ratios (0–50%). The experimental program covered quasi-static loading, drop-weight impact, and SHPB high strain-rate tests, focusing on the mechanical properties, crack failure modes, and energy dissipation capacity of the material. Based on the results of the static and dynamic splitting tensile tests, the following conclusions can be drawn:

(1)

The static compressive strength of Rubber-Toughened Ceramsite Concrete (RTCC) was lower than that of ordinary concrete, while the static splitting tensile strength first increased and then decreased with increasing rubber replacement ratio. In particular, when the rubber replacement ratio was 20%, the tensile strength reached its maximum value of 2.01 MPa, and the tension–compression ratio increased by 66.7%, significantly enhancing the crack resistance of the material. A moderate amount of rubber effectively improved toughness, whereas excessive rubber content weakened the interfacial bonding performance.

(2)

For the dynamic tests (with strain rates ranging from approximately 0.16 s⁻¹ to 4.76 s⁻¹), the dynamic splitting tensile strength of Rubber-Toughened Ceramsite Concrete (RTCC) exhibited a pronounced strain-rate effect, similar to its dynamic compressive strength behavior [46]. In addition, existing studies [47,48] have shown that the strain-rate range associated with blasting-induced vibration acting on structural components typically falls within 10⁻²–10¹ s⁻¹, which is close to the strain-rate interval adopted in this study. This further supports the engineering relevance and appropriateness of the selected strain-rate range.

(3)

Under medium strain rates (drop-weight tests), the peak impact force generally decreased with increasing rubber content, while the energy absorption capacity first increased and then decreased. The specimen with a 20% rubber replacement ratio showed the best overall performance, with a peak impact force of 42.48 kN and an absorbed energy of 43.23 J.

(4)

Under high strain rates (SHPB tests), Rubber-Toughened Ceramsite Concrete (RTCC) exhibited pronounced strain-rate sensitivity, with the DIF increasing significantly as the strain rate rose. As the rubber replacement ratio increased, the slope of the DIF-$\dot{\epsilon }$ curve gradually decreased, indicating that the incorporation of rubber reduced the strain-rate sensitivity of the concrete but improved its initial strength under low-velocity impacts. Moreover, at higher strain rates, due to the superior energy absorption capability of rubber, RTCC demonstrated greater energy dissipation capacity than ordinary ceramsite concrete. However, the overall energy dissipation capacity tended to decrease with further increases in rubber content.

(5)

RTCC can generally be produced using conventional concrete mixing and casting procedures without special equipment, indicating good construction adaptability. Although rubber incorporation may reduce static compressive strength, its enhanced deformation capacity and energy dissipation are advantageous for structures subjected to vibration or dynamic loading. In addition, the use of waste rubber contributes to resource recycling and environmental sustainability, suggesting acceptable economic potential when RTCC is applied in vibration-sensitive or non-primary load-bearing components.