Development of a Direct Addition Asphalt Modifier from Coal Liquefaction Residue Based on Rheological Properties

Abstract

To improve the low-temperature performance of direct coal liquefaction residue (DCLR)-modified asphalt, a direct-addition composite modifier (DCLAM) containing DCLR, SBS, crumb rubber, activated nano-CaCO₃, and aromatic oil was developed. A four-factor, three-level orthogonal design was used to optimize the formulation based on dynamic shear rheometer (DSR) and bending beam rheometer (BBR) tests. The optimal formulation contained 5% DCLR, 1% SBS, 15% crumb rubber, 3.5% nano-CaCO₃, and 2% aromatic oil. At 64 °C, the optimized mastic exhibited a complex shear modulus of 18,506.7 Pa and a rutting factor of 19.4 kPa, while its phase angle decreased to 72.4°. At −12 °C, the creep stiffness was 167.3 MPa and the m-value was 0.383. Mixture-level performance was further evaluated through wheel-tracking, low-temperature beam bending, freeze–thaw splitting, and dynamic modulus tests. The DCLAM-designed mixture exhibited a dynamic stability of 7176.0 cycles/mm, a maximum flexural tensile strain of 3087 με, and a freeze–thaw tensile strength ratio of 78.3%. It also showed higher complex modulus and lower phase angle values than the base mixture over the investigated temperature–frequency range. These results demonstrate the potential of the optimized DCLAM formulation for balancing high-temperature deformation resistance and low-temperature cracking resistance in asphalt pavement materials.

1. Introduction

Direct coal liquefaction residue (DCLR) is a complex substance primarily composed of heavy oil, asphaltenes, pre-asphaltenes, and tetrahydrofuran insoluble matter (THFI) [1]. Among these components, THFI significantly deteriorates the ductility of asphalt, while asphaltenes and pre-asphaltenes contribute to increasing asphalt hardness, enhancing high-temperature stability, and raising the softening point [2,3]. Studies have shown that when the DCLR content is 3%, its aging behavior differs markedly from that of the base asphalt. Increasing the content to 6–9% does not lead to a decline in aging performance, and further increases do not induce hardening of the modified asphalt [4,5]. As the DCLR content increases, the high-temperature performance of asphalt improves, but its low-temperature ductility correspondingly decreases [6,7]. To enhance the low-temperature performance of DCLR-modified asphalt, researchers have employed SBS for composite modification. Upon swelling within the asphalt, SBS forms a three-dimensional network structure, which can effectively compensate for the low-temperature performance loss induced by DCLR [8]. However, the dosage of SBS has an optimal range. Excessive SBS can adsorb the light components of asphalt during the shearing process, inhibiting the formation of the network structure and thereby adversely affecting low-temperature properties such as ductility [9]. The modification effectiveness of DCLR is also influenced by the form of its components and the preparation process. When the insoluble matter exists as discrete particles or clusters, its impact on asphalt performance is limited. However, if THFI acts as an active component in the modification, the DCLR content becomes a critical factor [10]. Some studies recommend grinding DCLR into a fine, uniform powder before blending it with the base asphalt, which can improve its adhesion to aggregates and result in performance superior to that of conventional 90# base asphalt [11,12].

Research indicates that crumb rubber from waste tires can synergistically improve the high- and low-temperature performance and aging resistance of asphalt. Building on this, researchers have further combined it with DCLR to prepare DCLR-crumb rubber composite modified asphalt. In this composite system, the incorporation of DCLR significantly enhances the high-temperature stability of the asphalt. Furthermore, it works synergistically with the crumb rubber to collectively improve elastic recovery and low-temperature cracking resistance [13].

While DCLR demonstrates excellent high-temperature stability as a road asphalt modifier, its low-temperature performance is often unsatisfactory [14,15]. Consequently, most research favors its combination with materials such as SBS and crumb rubber to compensate for this deficiency, involving systematic optimization of the dosage for each component. An appropriate amount of DCLR can effectively increase the high-temperature viscosity and hardness of asphalt. However, as its dosage increases, it tends to disrupt the colloidal structure of the asphalt, leading to enhanced heterogeneity within the system, which is detrimental to low-temperature ductility [16]. Therefore, it is recommended that the DCLR content should not exceed 10% of the mass of the base asphalt [17].

The overall performance of DCLR-modified asphalt is comparable to that of asphalt modified with 5% SBS [18]. Consequently, the resulting composite DCLR-modified asphalt mixture exhibits equivalent performance to SBS-modified asphalt mixtures in terms of high- and low-temperature properties and moisture susceptibility [19]. Furthermore, dynamic modulus test results and master curve fitting based on the CAM model demonstrate that this mixture possesses excellent deformation resistance across a broad frequency range. This indicates its potential as an effective pavement modification material, contributing to enhanced service performance and extended service life of road surfaces [20].

Due to the significant differences in molecular structure between SBS, crumb rubber, and base asphalt, their compatibility within the asphalt matrix is poor, making them prone to segregation and phase separation [21]. To address this issue, studies have shown that aromatic oil can serve as an effective compatibilizer, promoting the swelling and dispersion of either crumb rubber or SBS [22]. Aromatic oil is a hydrocarbon rich in aromatic ring structures, characterized by high aromatic content, a high flash point, and environmental friendliness. The aromatic and light components it contains contribute to the formation of a stable colloidal structure, thereby enhancing the compatibility between the modifiers and the asphalt [23,24,25]. Furthermore, the incorporation of aromatic oil can significantly improve the low-temperature performance of DCLR-modified asphalt [26].

While previous studies have demonstrated the beneficial effects of combining DCLR with SBS or crumb rubber, most investigations have focused on single- or binary-modification systems. For a multi-component direct-addition system, the key scientific issue is not merely to identify a formulation with improved performance, but to determine how the relative influence of individual components changes with rheological response and temperature, and how conflicting high- and low-temperature requirements can be reconciled in formulation selection.

Accordingly, the present study treated DCLAM formulation as a multi-factor rheological response problem rather than a simple blend-screening exercise. A four-factor, three-level orthogonal design was employed to identify the response-specific and temperature-dependent influence of DCLR, SBS, crumb rubber, and activated nano-CaCO₃ on complex shear modulus, phase angle, rutting factor, creep stiffness, and m-value. Aromatic oil was maintained at a fixed dosage as a compatibilizing component. The formulation was then selected by considering the trade-off among multiple rheological indicators rather than by maximizing a single property. Finally, the engineering performance of the independently designed DCLAM mixture was evaluated at the mixture scale to examine whether its performance trends were consistent with those identified at the mastic scale.

Therefore, the contribution of this study lies in establishing a multi-component direct-addition formulation strategy, revealing temperature- and response-dependent factor effects through orthogonal analysis, and developing a balanced rheological optimization framework for DCLR-based asphalt modification. The study does not attempt to establish a molecular-level modification mechanism, which requires further microstructural and chemical characterization.

2. Materials and Methods

2.1. Materials

The materials used in this study included DCLR asphalt, with its performance indices listed in Table 1 and main chemical components provided in Table 2. Before preparation of DCLAM, the collected DCLR was dried and processed into a fine powder form to ensure uniform dispersion during the blending process. A linear SBS modifier, YH-791, with a styrene–butadiene block ratio of 3:7 was employed; its basic properties are summarized in Table 3. A crumb rubber modifier obtained from waste tires was used in this study. The physical and chemical properties of the crumb rubber are summarized in Table 4. Aromatic oil (commercial Grade 4#8 aromatic oil) was added at a dosage of 2% [27,28,29], and its technical specifications are presented in Table 5. It was used as a compatibilizer in the composite modification system. The nano-modifier used was activated nano-CaCO₃, whose technical indicators are given in Table 6. The base asphalt was Haiyun Brand 90# penetration-grade asphalt binder, classified according to the Chinese penetration grading system. Its measured technical properties are summarized in Table 7. Asphalt mastic was prepared using limestone mineral filler at a filler-to-asphalt ratio of 0.8 [27]. The basic technical properties of the mineral filler are listed in Table 8. Limestone coarse aggregates of different size fractions and 0–3 mm limestone fine aggregate were used for asphalt-mixture preparation. Representative basic properties of the aggregates are summarized in Table 9.

Table 1. Performance indicators of DCLR.

Technical Indicator	Density (g/cm3)	Penetration at 25 °C (0.1 mm)	Ductility at 10 °C (cm)	Softening Point (°C)
Test result	1.24	4.8	1.8	175

Table 2. Main components of DCLR.

Component	Heavy Oil	Asphaltene	Pre-Asphaltene & THF-Insoluble Substances
Content (%)	34~37	17~22	43~46

Table 3. Basic properties of SBS.

Grade	Structure	Block Ratio (S/B)	Volatile Content (%)	Ash Content (%)	Tensile Strength (MPa)	Elongation at Break (%)	Permanent Set (%)	Shore A Hardness	Melt Flow Rate (g/10 min)
YH-791	Linear	30/70	≤0.7	≤0.2	≥15	≥700	≤40	68	0.5~2.5

Table 4. Physical and chemical test results of crumb rubber.

Test Item		Unit	Technical Requirement	Test Result
Physical indicators	Sieve residue	%	<10	8
	Relative density	/	1.10–1.30	1.20
	Moisture content	%	<1.0	0.55
	Iron content	%	<0.03	0.02
	Fiber content	%	<1.0	0.04
Chemical indicators	Ash content	%	≤8	6
	Acetone extract	%	≤16	11
	Carbon black content	%	≥28	32
	Rubber hydrocarbon content	%	≥48	56
	Solubility	%	≥16	18

Table 5. Technical specifications of aromatic oil.

Technical Indicator	Kinematic Viscosity (mm2/s)	Open Flash Point (°C)	Ash Content (%)	Aromatic Content (%)	Appearance
Test result	26	220	0.01	80	Yellowish green

Table 6. Technical specifications of active nano-CaCO₃.

Model	CaCO3 Content (%)	Specific Gravity (g/cm3)	Average Particle Size (nm)	Specific Surface Area Provided by Supplier (m2/g)	Oil Absorption (%)	Surface Treatment Agent	Appearance
303	98	1.99~2.01	<100	2.1	28	Composite treatment	White powder

Table 7. Technical specifications of 90# base asphalt.

Test Item	Unit	Technical Requirement	Test Result
Penetration (25 °C, 100 g, 5 s)	0.1 mm	80~100	88
Softening point (Ring & Ball)	°C	≥43	45
Dynamic viscosity (60 °C)	Pa·s	≥140	155
Ductility (5/cm, 10 °C)	cm	≥30	35
Residue after RTFOT
Mass loss	%	≤±0.8	0.018
Penetration ratio	%	≥57	60.9
Ductility (10 °C)	cm	≥8	10

Table 8. Basic technical indicators of mineral filler.

Technical Indicators		Unit	Test Result	Technical Requirement
Apparent Density		kg/m3	2.694	≥2.5
Moisture Content		%	0.54	1
Particle Size Range	<0.6	%	100	100
	<0.15	%	95.4	90~100
	<0.075	%	76.7	75~100
Appearance		/	Free from agglomeration	Free from agglomeration
Hydrophilic Coefficient		/	0.67	<1
Plasticity Index		%	2.2	<4
Heating Stability		/	Free from deterioration	Measured

Table 9. Basic properties of aggregates used in asphalt mixtures.

Aggregate Type	Particle Size	Apparent Relative Density	Bulk Relative Density	Water Absorption/%	Additional Indicator
Coarse aggregate	3–5 mm	3.243	3.104	1.38	/
Coarse aggregate	5–10 mm	3.234	3.165	0.67	LA abrasion 15.4%
Coarse aggregate	10–15 mm	3.252	3.173	0.77	LA abrasion 17.9%
Fine aggregate	0–3 mm	3.111	3.022	/	Sand equivalent 73%; angularity 39 s

2.2. Orthogonal Experiment Design

According to previous studies and compatibility optimization results reported in the literature [30,31], the dosage of aromatic oil was fixed at 2%. This content was selected because aromatic oil mainly functions as a compatibilizer to promote the swelling and dispersion of SBS and crumb rubber rather than as the primary reinforcing modifier. During the high-speed shearing process, the temperature was controlled to avoid excessive thermal oxidation. The shearing procedure was adopted only to improve the dispersion and blending of modifier components, rather than to simulate aging conditions. Moreover, fixing its content avoided introducing additional variables into the orthogonal design and allowed the effects of DCLR, SBS, crumb rubber, and nano-CaCO₃ to be systematically evaluated. An orthogonal experimental design with four factors and three levels was subsequently employed to explore the impacts of DCLR, SBS, crumb rubber, and activated nano-CaCO₃. The factors and their corresponding levels are presented in Table 10, while the specific orthogonal test matrix is shown in Table 11.

Table 10. Factors and levels for the orthogonal experiment design.

Level	DCLR Content (%, by Mass)	SBS Content (%, by Mass)	Crumb Rubber Content (%, by Mass)	Active Nano-CaCO3 Powder Content (%, by Mass)
1	5	1	10	2.5
2	7	2	15	3.0
3	9	3	20	3.5

Table 11. Orthogonal experiment array.

Test No.	Level Combination	A: DCLR Content (%, by Mass)	B: SBS Content (%, by Mass)	C: Crumb Rubber Content (%, by Mass)	D: Active Nano-CaCO3 Powder Content (%, by Mass)
1	A1B1C1D1	5	1	10	2.5
2	A1B2C2D2	5	2	15	3.0
3	A1B3C3D3	5	3	20	3.5
4	A2B1C2D3	7	1	15	3.5
5	A2B2C3D1	7	2	20	2.5
6	A2B3C1D2	7	3	10	3.0
7	A3B1C3D2	9	1	20	3.0
8	A3B2C1D3	9	2	10	3.5
9	A3B3C2D1	9	3	15	2.5

2.3. Preparation of DCLAM

For each formulation, the masses of DCLR, SBS, crumb rubber, activated nano-CaCO₃, and aromatic oil were calculated as percentages by mass of the base asphalt. Taking the optimal formulation based on 1000 g of base asphalt as an example, the modifier consisted of 50 g DCLR, 10 g SBS, 150 g crumb rubber, 35 g activated nano-CaCO₃, and 20 g aromatic oil.

2.4. Preparation of DCLAM-Modified Asphalt Mastic

Five hundred grams of 90# base asphalt and limestone mineral filler (with a filler—to—asphalt ratio of 0.8, which means 400 g of filler) were pre-heated separately in an oven at 160 °C for 1 h. Subsequently, the mineral filler was added to the asphalt in five installments. The mixture was mechanically stirred at 160 °C and 1500 r/min until it became homogeneous (approximately 15 min). Then, the temperature of the system was increased to 180 °C. The previously prepared DCLAM, with a total addition of 26.5% by mass of the asphalt (that is, 132.5 g), was added in three installments. High-shear mixing was carried out at 3500 r/min for 30 min until a uniform mixture without visible air bubbles was obtained. To minimize possible filler settlement during BBR specimen preparation, the asphalt mastic was cast into the beam molds immediately after completion of high-shear mixing, without prolonged high-temperature storage or static holding before casting. The mastic was maintained in a homogeneous state during specimen preparation, and no visible segregation was observed during casting. The specimens were then allowed to cool and solidify statically for subsequent testing.

2.5. Preparation of Asphalt Mixtures

An AC-16 dense-graded asphalt mixture was adopted for the mixture-level verification tests. The aggregate gradation was designed according to JTG F40-2004, using limestone coarse and fine aggregates together with limestone mineral filler. The same aggregate gradation was used for both the base asphalt mixture and the DCLAM-modified asphalt mixture to ensure comparability.

The optimum asphalt contents were determined using the Marshall method. The base asphalt mixture was prepared with an asphalt–aggregate ratio of 4.6%. For the DCLAM-modified asphalt mixture, the optimized DCLAM formulation determined from the mastic rheological tests was used. The mixture contained 4.0% base asphalt and 1.0% DCLAM by mass of aggregate, as determined from the Marshall mixture design.

The DCLAM-modified asphalt mixture was prepared using a dry-addition process. The aggregates were preheated to approximately 190 °C and placed in a mixing bowl maintained at approximately 180 °C, followed by dry mixing for at least 30 s. DCLAM was then added at room temperature and mixed with the aggregates for at least 30 s. Subsequently, the base asphalt preheated to approximately 160 °C was added and mixed for at least 180 s to promote uniform dispersion of DCLAM and coating of the aggregate surfaces. Finally, the mineral filler preheated to approximately 190 °C was added and mixed for at least 60 s. The total mixing time was not less than 5 min. The resulting mixtures were subsequently compacted and prepared into specimens according to the requirements of the corresponding performance tests.

It should be noted that the two mixtures were independently designed using the Marshall method, and therefore their optimum binder-related contents were not constrained to be identical. Accordingly, the mixture-level comparison in this study was intended to evaluate the engineering performance of the separately optimized mixture systems rather than to isolate the intrinsic effect of DCLAM under an equal total binder content.

2.6. Experimental Methods

2.6.1. Dynamic Shear Rheometer (DSR) Test

The high-temperature rheological properties of the base asphalt mastic and nine groups of DCLAM-modified asphalt mastic were systematically evaluated by means of a DHR-I Dynamic Shear Rheometer (TA Instruments, New Castle, DE, USA) via strain sweep and temperature sweep tests.

(1)

Strain Sweep: Strain sweep tests were carried out at four specific temperatures, namely 46 °C, 52 °C, 58 °C, and 64 °C. A 25 mm parallel-plate geometry was utilized, with a strain range spanning from 0.1% to 100% at a constant frequency of 10 rad/s. Before commencing the tests, the base asphalt mastic and the DCLAM samples were heated to 165 °C and 185 °C respectively, accompanied by gentle stirring to guarantee homogeneity and eliminate air bubbles. The test points were logarithmically spaced so as to ensure uniform data collection throughout the entire strain range. The critical strain corresponding to the limit of the linear viscoelastic region (LVE) was determined based on a 10% reduction in the complex shear modulus (G*) relative to its initial maximum value in the low-strain region. The reference maximum G* value was obtained from the plateau region before significant modulus degradation occurred. Therefore, the 90% G* line shown in Figure 1 represents the criterion used to identify the transition from the LVE region to the nonlinear viscoelastic region.

Figure 1. High-temperature strain sweep curves for the 10 asphalt mastic groups.

(2)

Temperature Sweep: Temperature sweep tests were carried out over a temperature range from 46 °C to 82 °C with an interval of 6 °C. By employing the 25 mm parallel-plate geometry, a sinusoidal oscillatory load was applied at a frequency of 10 rad/s to assess the rheological response of the mastics in relation to temperature.

2.6.2. Bending Beam Rheometer (BBR) Test

According to the AASHTO T313 standard, the BBR test was conducted to evaluate the low-temperature rheological properties of asphalt mastics. The creep stiffness (S) and creep rate (m-value) were selected as the main evaluation parameters. The m-value was determined from the slope of the logarithmic relationship between creep stiffness and loading time, which can be expressed as:

$$S(t)={\displaystyle \frac{\mathrm{P}{L}_3}{4b{h}^3\delta (t)}}$$

(1)

$$m={\displaystyle \frac{\mathrm{d}\left[\lg S\left(t\right)\right]}{\mathrm{d}\left[\lg t\right]}}$$

(2)

where

• P—Constant load;
• L—Beam span length (mm);
• b—Beam width (mm);
• h—Beam height (mm);
• δ(t)—Mid-span deflection at loading time t;
• S(t)—The creep stiffness at loading time t.

2.6.3. Asphalt-Mixture Performance Tests

To evaluate the engineering performance of the independently designed DCLAM-modified mixture and examine whether its mixture-level response was consistent with the rheological trends observed at the mastic scale, the base asphalt mixture and the DCLAM-modified asphalt mixture were evaluated through wheel-tracking, low-temperature beam bending, freeze–thaw splitting, and dynamic modulus tests.

(1)

Wheel-tracking test: The high-temperature rutting resistance of the asphalt mixtures was evaluated according to JTG E20-2011 (T0719). The test was conducted at 60 °C under repeated wheel loading, and dynamic stability (DS, cycles/mm) was used as the primary evaluation parameter. A higher dynamic stability indicates greater resistance to permanent deformation under high-temperature loading.

(2)

Low-temperature beam-bending test: The low-temperature cracking resistance of the asphalt mixtures was evaluated according to JTG E20-2011 (T0715). Beam specimens were tested at −10 °C at a loading rate of 50 mm/min. The flexural tensile strength, maximum flexural tensile strain, and flexural stiffness modulus were determined to characterize the strength and deformation capacity of the mixtures under low-temperature conditions.

(3)

Freeze–thaw splitting test: The moisture susceptibility of the asphalt mixtures was evaluated according to JTG E20-2011 (T0729). The tensile strengths of specimens before and after freeze–thaw conditioning were determined, and the freeze–thaw tensile strength ratio (TSR) was calculated. A higher TSR indicates better retention of tensile strength after moisture and freeze–thaw conditioning.

(4)

Dynamic modulus test: The temperature- and frequency-dependent viscoelastic properties of the mixtures were evaluated using the uniaxial compression dynamic modulus test according to JTG E20-2011 (T0738). Cylindrical specimens were subjected to cyclic compressive loading at five temperatures of 5, 10, 20, 35, and 50 °C and six loading frequencies of 0.1, 0.5, 1, 5, 10, and 25 Hz. The complex modulus magnitude |E*| and phase angle δ were obtained from the stress and strain responses. The complex modulus was used to characterize the overall stiffness response of the mixture, whereas the phase angle was used to evaluate the relative contributions of elastic and viscous deformation under cyclic loading.

2.6.4. Data Processing and Statistical Analysis

To ensure the reliability and repeatability of the experimental results, four parallel specimens were prepared and tested for each experimental condition under identical preparation and testing procedures. The reported values were calculated as the mean of four independent measurements.

The variability of the repeated measurements was evaluated using the standard deviation (SD). The error bars presented in the figures represent the SD values obtained from the four replicate tests, which were used to illustrate the dispersion and repeatability of the experimental results.

The orthogonal experimental results were analyzed using range analysis to evaluate the influence of different factors on the rheological properties. Statistical significance analysis, such as analysis of variance (ANOVA), was not included in this study.

3. Results and Discussion

3.1. Analysis of DSR Tests

3.1.1. Strain Sweep Test

Figure 1 depicts the high-temperature strain sweep results for the ten groups of asphalt mastics. At temperatures of 52 °C, 58 °C, and 64 °C, the strain sweep curves exhibited a relatively flat trend, which suggests that within this temperature range, the mastics retained a linear viscoelastic response throughout the applied strain range from 0.1% to 100%. Nevertheless, at 46 °C, the complex shear modulus in the low-strain region did not decline by more than 10% of its maximum value. Once a certain critical strain was exceeded, the modulus attenuation intensified significantly, indicating the transition of the material from the linear to the nonlinear viscoelastic region.

As the temperature rises, the linear viscoelastic strain range of the asphalt mastics gradually expands. Conversely, under lower temperatures, it correspondingly narrows. This is mainly because the material gradually transitions to a glassy state. Figure 1 shows that at 46 °C, there are considerable variations in the linear viscoelastic strain limits among different asphalt mastics, and the specific values are summarized in Table 12. According to the data in Table 12, the strain limits within the linear viscoelastic range vary significantly across mastics with different formulations. At a strain level of 1%, all mastics remained within or close to the limit of the linear viscoelastic region; therefore, 1% strain was selected as the controlled condition for the subsequent temperature sweep tests.

Table 12. Critical strain limits for each asphalt mastic group within the linear viscoelastic region.

Test No.	Base Asphalt	1	2	3	4	5	6	7	8	9
Strain (%)	5.8	1.7	20.0	4.1	6.8	15.8	36.6	8.7	1.0	17.0

3.1.2. Temperature Sweep Test

Figure 2 illustrates the variations in the complex shear modulus (G*), phase angle (δ), and rutting factor (G*/sin δ) with temperature for both the base asphalt mastic and the asphalt mastics modified with the DCLAM. G* and δ jointly characterize the viscoelastic response of the material, where G* can be decomposed into the storage modulus (G′ = G*∙cos δ) and the loss modulus (G″ = G*∙sin δ). A smaller δ value indicates a relatively larger elastic contribution to the viscoelastic response. However, the absolute magnitude of δ should also be considered when distinguishing between viscous- and elastic-dominated behavior. Therefore, a reduction in δ does not necessarily imply an elastic-dominated response. A higher G* combined with a lower δ signifies enhanced rutting resistance.

Figure 2. Temperature sweep test results for the 10 asphalt mastic groups.

As shown in Figure 2a, G* exhibited a monotonic decreasing trend for all asphalt mastics with increasing temperature. At any given temperature, the G* values of the DCLAM-modified mastics were consistently higher than those of the base asphalt mastic, indicating that the DCLAM effectively increased the stiffness of the mastic. Figure 2b shows that the δ values gradually increase as the temperature rises, indicating an increase in the viscous component and a decrease in the elastic component of the material. The incorporation of DCLAM reduced the phase angle relative to that of the base asphalt mastic at the same temperature, indicating an increase in the relative elastic contribution. It should be noted that the relatively high absolute phase angles, particularly at elevated temperatures, indicate that the mastics remained predominantly viscous. Based on the combined trends of G* and δ, it can be concluded that DCLAM effectively enhances the high-temperature deformation resistance of the asphalt mastic. The mastic modified with the formulation from Group 2 exhibits the most favorable elastic recovery and high-temperature stability.

A higher rutting factor indicates that the asphalt mastic has better high-temperature rutting resistance. As shown in Figure 2c, it is clear that the rutting factor of all groups decreased continuously as the temperature increased, which means that the rutting resistance also declined accordingly. The rate of decrease in G*/sin δ was more pronounced in the lower-temperature range, followed by a sharp decline as the temperature increased further. This phenomenon is mainly ascribed to the weakening of intermolecular forces within the asphalt at high temperatures. This weakening leads to a higher proportion of viscous components and a decreased elastic response. Under high-temperature conditions, the increased fluidity of the material causes a reduction in resistance to permanent deformation, thus significantly impairing rutting performance.

At the same temperatures, the rutting factors of all DCLAM-modified asphalt mastics were higher than those of the base asphalt mastic. This confirms that the incorporation of the DCLAM effectively enhanced the deformation resistance and high-temperature stability of the asphalt. In summary, the results presented in Figure 2 indicate that modification with DCLAM can significantly improve the high-temperature performance of asphalt mastics.

3.2. Analysis of BBR Tests

The BBR test results are presented in Figure 3. The orthogonal experimental results indicated that the effectiveness of different DCLAM formulations in improving the low-temperature performance of the asphalt mastic varied significantly. In the temperature range from −18 °C to −6 °C, the creep stiffness S values of the modified mastics in Groups 9,2,3,6,1,5,7 were all lower than those of the base asphalt mastic, and their creep rate m values were higher. This indicates that these modified mastics have better low-temperature cracking resistance. The creep stiffness index (S_AS) was additionally introduced as a supplementary indicator to evaluate the temperature susceptibility of asphalt mastic. Unlike the standard BBR parameters (creep stiffness S and creep rate m-value) specified in AASHTO T313, S_AS is derived from the temperature dependence of creep stiffness and reflects the sensitivity of stiffness variation to temperature changes. The relationship between creep stiffness and temperature was established according to previous studies [32,33,34]. In this study, creep stiffness was linearly fitted as a function of temperature, and the S_AS value was defined as the absolute value of the regression slope. It should be noted that although the same groups were identified in the comparison of S/m and S_AS, the two sets of indicators represent different aspects of low-temperature rheological behavior. The S and m values characterize the absolute low-temperature response at specific test temperatures, whereas S_AS reflects the temperature sensitivity of creep stiffness over the investigated temperature range.

Figure 3. BBR test results and their fitting with temperature for the 10 asphalt mastic groups.

The results demonstrate that the creep stiffness of all mastics gradually increased as the temperature decreased (Figure 3a). A higher creep stiffness implies that the material is more likely to undergo brittle cracking under low-temperature and loading conditions. Therefore, a lower S value is indicative of better low-temperature performance. An increase in the S value at low temperatures restricts the strain response of the material under constant stress, which is unfavorable for its low-temperature cracking resistance.

To enhance the cracking resistance of pavement, asphalt mastics should display low creep stiffness and a high creep rate. As depicted in Figure 3b, the m-values for all mastics declined as the temperature decreased. A higher m-value indicates better low-temperature performance, and the variation in m-values among the DCLAM-modified mastics at different temperatures also reflects their low-temperature susceptibility.

Based on the BBR test results at −6 °C, −12 °C, and −18 °C, all asphalt mastics showed a consistent trend: the creep stiffness S increased as the temperature decreased, while the creep rate m-value decreased correspondingly. The relationship between the S value of each mastic group and temperature conforms to the following equation:

$$S=k·T+C$$

(3)

where

• S represents the creep stiffness (MPa).
• T denotes the temperature (°C).
• k is the regression slope of creep stiffness with respect to temperature.
• C is the regression constant. The creep stiffness susceptibility parameter S_AS is defined as the absolute value of the regression slope, i.e., S_AS = |k|, with a unit of MPa/°C.

For each asphalt mastic and test temperature, the creep stiffness used for the temperature-susceptibility analysis was the mean value obtained from four parallel specimens. The fitting results reported in Table 13 are recalculated using these four-replicate mean S values at −6, −12, and −18 °C.

Table 13. Low-temperature susceptibility fitting results based on the four-replicate mean creep stiffness values for the 10 asphalt mastic groups.

Type of Asphalt	Fitted Equation	R2	SAS (MPa/°C)
Base asphalt mastic	S = −35.0136T − 128.7607	0.9995	35.0136
Group 1	S = −28.5169T − 144.9606	0.9874	28.5169
Group 2	S = −29.7045T − 130.3870	0.9893	29.7045
Group 3	S = −29.6675T − 132.1410	0.9893	29.6675
Group 4	S = −35.7972T − 154.4876	0.9638	35.7972
Group 5	S = −28.4587T − 152.7666	0.9218	28.4587
Group 6	S = −28.8340T − 108.8520	0.9903	28.8340
Group 7	S = −26.5183T − 105.2153	0.9754	26.5183
Group 8	S = −35.0204T − 125.2353	0.9801	35.0204
Group 9	S = −30.8511T − 87.3426	0.9985	30.8511

It should be emphasized that the fitting was performed directly on the creep stiffness S in MPa, without logarithmic transformation. Therefore, the S_AS values reported in Table 13 represent the absolute slopes of S versus T and have units of MPa/°C.

The parameter S_AS characterizes the rate of change in creep stiffness S with temperature, thereby reflecting the temperature susceptibility of the asphalt mastic under low-temperature conditions. As shown in Table 13, the fitting results for the low-temperature susceptibility of the ten mastic groups show distinct differences. Among them, the S values for modified mastic Groups 4 and 8 change most rapidly with temperature, indicating the most pronounced low-temperature susceptibility. In contrast, the S_AS values for the base asphalt mastic and the modified mastic Groups 9, 2, 3, 6, 1, 5, and 7 decrease sequentially, reflecting a gradual reduction in their temperature susceptibility.

Comparative analysis reveals that the S_AS values of modified mastic Groups 9, 2, 3, 6, 1, 5, and 7 are all lower than those of the base asphalt. This indicates that these groups exhibit lower temperature sensitivity and better stability under low-temperature conditions. Furthermore, examining the trend of S variation with temperature in Figure 3, the S_AS values for modified mastic Groups 1, 5, and 7 are 28.5169, 28.4587, and 26.5183, respectively. Their rates of change are the most gradual, further verifying that these three modified mastic groups exhibit the optimal low-temperature stability. For the base asphalt mastic, the fitted relationship yielded creep stiffness values of 81.3, 291.4, and 501.5 MPa at −6, −12, and −18 °C, respectively, which were consistent with the temperature-dependent trend shown in Figure 3a.

3.3. Analysis of Orthogonal Experimental Results for Rheology of DCLAM-Modified Asphalt Mastic

In the orthogonal experiment analysis, the mean values of each indicator at different factor levels were first analyzed individually to systematically screen for the optimal proportion scheme. An orthogonal experiment analysis table of means and ranges was constructed by integrating the experimental data, upon which range analysis was conducted. Here, $\overline{\mathrm{X}}$_A1 represents the mean value of the test results corresponding to factor A at level 1, and similarly, $\overline{\mathrm{X}}$_A2 is the mean at level 2.

(1)

Complex shear modulus

Figure 4 depicts the range analysis results of the complex shear modulus in relation to temperature. In the temperature range from 46 to 82 °C, factor C (crumb rubber content) showed the largest range, and its mean value corresponding to level 2 was the highest, which implies that this factor had the most substantial influence on the complex shear modulus. The sequence of influence of each factor was as follows: C > A > D > B. Based on a comprehensive analysis of the ranges and level means, the recommended combination is C₂, A₁, D₁. However, factor B had two effective levels: B₁ showed a more pronounced influence within the 46–70 °C range, suggesting a recommended combination of A₁B₁C₂D₁, whereas B₃ became more prominent within the 76–82 °C range, leading to a recommended combination of A₁B₃C₂D₁.

Figure 4. Range value of complex shear modulus variation with temperature.

(2)

Phase angle

Figure 5 depicts the range analysis results of the phase angle in relation to temperature. The influence of each factor on the phase angle varied across different temperature intervals. Since a lower phase angle indicates a relatively greater elastic contribution to the viscoelastic response, the optimal level was determined according to the minimum mean value rather than the maximum mean value. In the temperature range of 46–58 °C, factor B (SBS content) showed the largest range, indicating that SBS had the greatest influence on phase angle. The order of factor significance was determined as B > A > C > D, with the recommended combination being A₁B₁C₁D₁.

Figure 5. Range value of phase angle variation with temperature.

In the higher-temperature range of 64–82 °C, factor A (DCLR content) showed the greatest range, with level A₁ being the most influential. The significance order shifted to A > B > C > D. Although the influence of factor D (nano-CaCO₃ content) was relatively minor, both levels D₁ and D₃ exhibited certain significance within this interval: the combination A₁B₁C₁D₁ is recommended for 64–70 °C, whereas A₁B₁C₁D₃ is recommended for 76–82 °C.

(3)

Rutting factor

Figure 6 showcases the range analysis results of the rutting factor as a function of temperature. It was discovered that the degree of influence exerted by each factor varied significantly across different temperature intervals. In the 46–58 °C range, factor C (crumb rubber content) had the largest range, and its level C₂ corresponded to the highest mean value, which implies that it had the most significant impact on the rutting factor. The order of factor significance was established as C > A > D > B. Based on the analysis of the ranges and level means, the recommended combinations are A₁B₁C₂D₁ for the 46–52 °C interval and A₁B₃C₂D₁ for the 58–82 °C interval. In this case, the level of factor B changes from B₁ to B₃ as the temperature rises.

Figure 6. Range value of rutting factor variation with temperature.

(4)

Creep stiffness

Figure 7 depicts the range analysis results of the creep stiffness in relation to temperature. As the test temperature dropped from −6 °C to −18 °C, the extent of each factor’s influence on the creep stiffness changed significantly. At −6 °C, factor A (DCLR content) exhibited the largest range, indicating the dominant influence on creep stiffness. Since lower creep stiffness is beneficial for improving low-temperature cracking resistance, the optimal factor level should be selected based on the minimum mean value of S. Therefore, the influence of each factor was evaluated according to the lower-is-better criterion. Specifically, for factor A (DCLR content), level A3 corresponded to the lowest mean creep stiffness at both −6 °C and −12 °C. Therefore, A3 was selected as the recommended level at these two temperatures, consistently following the lower-is-better criterion. At −12 °C, the influence of factor C (crumb rubber content) became the most pronounced. When the temperature was further reduced to −18 °C, factor D (nano-CaCO₃ content) showed the largest range, with level D₃ having the most significant effect. Based on a comprehensive analysis of the ranges and level means across the tested temperatures, the recommended combinations are A₃B₃C₂D₃ for the temperature range from −12 °C to −6 °C and A₃B₂C₂D₃ for −18 °C. This indicates the dynamic adjustment needed for the optimal content of each factor under different low-temperature conditions.

Figure 7. Range value of creep stiffness variation with temperature.

(5)

Creep rate

Figure 8 depicts the range analysis results of the creep rate in relation to temperature. Within the temperature range of −12 °C to −6 °C, factor A (DCLR content) showed the largest range, and level A₁ attained the highest mean value, thus identifying it as the primary influencing factor. The significance order of the factors was A > C > D > B, resulting in a recommended combination of A₁B₂C₃D₁. When the temperature dropped to −18 °C, factor A still had the greatest range, and level A₁ continued to show the optimal effect. At this temperature, the recommended combination is modified to A₁B₁C₃D₂. The analysis reveals that factor A consistently exerted a dominant influence on the creep rate under low-temperature conditions, while the optimal levels of the other factors changed correspondingly as the temperature decreased.

Figure 8. Range value of creep rate variation with temperature.

Taken together, the orthogonal analysis reveals that the dominant factor was response- and temperature-dependent rather than fixed across all rheological indicators. Crumb rubber exerted the strongest influence on complex shear modulus and rutting factor over most of the investigated high-temperature range, whereas the dominant factor controlling phase angle shifted from SBS at 46–58 °C to DCLR at 64–82 °C. For low-temperature creep stiffness, the dominant factor shifted from DCLR at −6 °C to crumb rubber at −12 °C and nano-CaCO₃ at −18 °C, while DCLR remained the primary factor affecting the m-value. These results demonstrate that no single component governed all aspects of rheological performance and that the optimal formulation necessarily involved a trade-off among different response variables and temperature ranges. Thus, the orthogonal design provided not only formulation screening but also a factor–response map for the multi-component DCLAM system.

3.4. Analysis of Optimal Formulation Based on DCLAM-Modified Asphalt Mastic

Based on the analysis of the data presented in the aforementioned figures, the orthogonal experimental results for each performance indicator of the DCLAM-modified asphalt mastic are summarized in Table 14. A comprehensive analysis of Table 14 shows that C₂ was the favorable level for complex shear modulus and rutting factor, whereas B1 was favorable for phase angle. The creep-stiffness response was strongly influenced by factors A, C, and D depending on temperature, whereas factor A exerted the greatest influence on the m-value based on the range analysis. Therefore, the key factor levels influencing the high- and low-temperature performance of the DCLAM-modified asphalt mastic are identified as A_1–3, B₁, C₂, and D₃.

Table 14. Summary of orthogonal test results for various test indicators of DCLAM.

Test Indicator	Temperature (℃)	Order of Influence Based on Range Analysis	Recommended Combination
G*	46~70	C > A > D > B	A1B1C2D1
	76~82	C > A > D > B	A1B3C2D1
δ	46~58	B > A > C > D	A1B1C1D1
	64~70	A > B > C > D	A1B1C1D1
	76~82	A > B > C > D	A1B1C1D3
G*/sin δ	46~52	C > A > D > B	A1B1C2D1
	58~82	C > A > D > B	A1B3C2D1
S	−6	A > C > D > B	A3B3C2D3
	−12	C > A > D > B	A3B3C2D3
	−18	D > C > A > B	A3B2C2D3
m	−6~−12	A > C > D > B	A1B2C3D1
	−18	A > C > D > B	A1B1C3D2

Based on the analysis of factor significance, two candidate formulations were obtained: A₁B₁C₂D₃ and A₃B₁C₂D₃. A comparison of the rheological properties corresponding to levels A₁ and A₃ revealed that, under high-temperature conditions, A₁ exhibited a higher complex shear modulus and rutting factor, together with a lower phase angle. In contrast, A₃ provided lower creep stiffness under the investigated low-temperature conditions, whereas A₁ showed an advantage in creep rate. Therefore, neither A₁ nor A₃ was optimal for all individual rheological indicators. Considering the overall balance between high-temperature deformation resistance and low-temperature cracking resistance, A₁ was selected as the preferred level for the final formulation. Although some factor combinations showed advantages for individual low-temperature indicators, the final formulation was not selected based on a single performance parameter. Considering both high-temperature deformation resistance and low-temperature cracking resistance, A₁B₁C₂D₃ was selected as the optimal formulation because it provided a balanced improvement in G*, G*/sin δ, S, and m. Although A₃ provided the lowest creep stiffness under certain low-temperature conditions, the final formulation was determined by considering the comprehensive balance among high-temperature deformation resistance and low-temperature cracking resistance rather than optimizing a single indicator. Therefore, the optimal formulation is determined to be A₁B₁C₂D₃. This selection therefore represents a multi-response compromise derived from the factor–response relationships rather than the simple selection of the best-performing individual orthogonal trial. For creep stiffness (S), the recommended combination was determined based on the lower-is-better criterion. Therefore, the factor levels corresponding to lower mean S values were considered favorable for low-temperature performance.

The optimal formulation combination (A₁B₁C₂D₃) was further verified through temperature sweep and BBR tests by comparing its rheological properties with those of the base asphalt mastic and representative orthogonal groups. The verification considered both high-temperature deformation resistance indicators (G*, δ, and G*/sin δ) and low-temperature cracking resistance indicators (S and m), thereby providing a comprehensive evaluation of the combined modification effects of the different components. The temperature sweep and BBR tests data indicators for the optimal formulation combination are shown in Table 15 and Table 16.

Table 15. Temperature sweep test results for the optimal formulation combination (A₁B₁C₂D₃).

Test Indicator	46 °C	52 °C	58 °C	64 °C	70 °C	76 °C	82 °C
G*/Pa	99,780.2	48,921.5	28,289.6	18,506.7	14,148.3	10,981.4	9946.6
δ/°	68.3	69.1	70.5	72.4	74.5	76.3	77.6
G*/sin δ/kPa	107.4	52.4	30.0	19.4	14.7	11.3	10.2

Table 16. BBR test results for the optimal formulation combination (A₁B₁C₂D₃).

Test Indicator	−6 °C	−12 °C	−18 °C
S	36.625	167.312	374.236
m	0.7115	0.3825	0.2647

As depicted in Figure 9, the optimal formulation exhibited superior rheological performance compared with the base asphalt mastic and Groups 1–4. At 64 °C, the optimal formulation exhibited a higher complex shear modulus than the base asphalt mastic. The corrected rutting factor calculated from the measured G* and δ values also showed an improvement compared with the base asphalt mastic, confirming the enhanced high-temperature deformation resistance of the modified asphalt mastic.

Figure 9. Comparison of temperature sweep and BBR test results between the optimal formulation combination and other asphalt mastic groups.

At 64 °C, the optimal formulation exhibited a phase angle of 72.4°, compared with approximately 84.1° for the base asphalt mastic. Although a phase angle of 72.4° still indicates a predominantly viscous response in absolute terms, its lower value relative to that of the base asphalt mastic indicates a relatively greater elastic contribution to the overall viscoelastic response. Therefore, the improvement should be interpreted as a relative increase in elastic contribution rather than a transition to elastic-dominated behavior. At low temperatures, it showed lower creep stiffness and a higher creep rate, leading to overall superior performance in comparison with the other groups. Therefore, A₁B₁C₂D₃ is verified as the optimal formulation, corresponding to the following composition: 5% DCLR, 1% SBS, 15% crumb rubber, 3.5% nano-CaCO₃, and 2% aromatic oil.

Based on the creep stiffness values of the optimal formulation at −6, −12, and −18 °C, linear fitting of S against temperature yielded a regression slope of −28.1343 MPa/°C. Therefore, the corresponding S_AS value of the optimal formulation was 28.13.

3.5. Performance Evaluation of Asphalt Mixtures

To further evaluate the engineering performance of the optimized DCLAM mixture and compare its response with that of the independently designed base asphalt mixture, conventional pavement performance and dynamic modulus tests were conducted. Because the two mixtures were separately optimized using the Marshall method, the following comparison should be interpreted as a comparison between two independently designed mixture systems rather than as a controlled isolation of the DCLAM effect at identical binder content.

3.5.1. Conventional Pavement Performance

The conventional pavement performance of the base asphalt mixture and the DCLAM-modified asphalt mixture is summarized in Table 17. The dynamic stability of the DCLAM-modified asphalt mixture reached 7176.0 cycles/mm, compared with 2732.2 cycles/mm for the base asphalt mixture, corresponding to an increase of approximately 162.7%. The substantially higher dynamic stability of the DCLAM-modified mixture indicates that the independently designed DCLAM mixture exhibited greater resistance to permanent deformation at high temperature than the base mixture. This trend was consistent with the higher G* and G*/sin δ values observed for the optimized DCLAM-modified mastic. Thus, the mixture-scale rutting response was consistent with the rheological trend observed at the mastic scale.

Table 17. Conventional pavement performance of the base asphalt mixture and DCLAM-modified asphalt mixture.

Test Parameter	Base Asphalt Mixture	DCLAM-Modified Asphalt Mixture
Dynamic stability (cycles/mm)	2732.2	7176.0
Flexural tensile strength (MPa)	5.39	7.86
Maximum flexural tensile strain (με)	2039	3087
Flexural stiffness modulus (MPa)	2643	2546
Tensile strength before freeze–thaw conditioning (MPa)	8.8	8.3
Tensile strength after freeze–thaw conditioning (MPa)	6.6	6.5
Freeze–thaw tensile strength ratio (%)	74.6	78.3

The independently designed DCLAM-modified mixture also exhibited more favorable low-temperature flexural performance than the base mixture. The flexural tensile strength increased from 5.39 MPa to 7.86 MPa, corresponding to an increase of approximately 45.8%, while the maximum flexural tensile strain increased from 2039 με to 3087 με, representing an increase of approximately 51.4%. Meanwhile, the flexural stiffness modulus decreased slightly from 2643 MPa to 2546 MPa. The simultaneous increase in flexural strength and ultimate tensile strain, together with the slight reduction in stiffness modulus, indicates that the DCLAM-modified mixture had a greater capacity to accommodate tensile deformation at low temperatures. This finding is consistent with the BBR results, in which the optimized DCLAM-modified mastic exhibited lower creep stiffness and a higher m-value than the base asphalt mastic.

With respect to moisture susceptibility, the tensile strength before freeze–thaw conditioning decreased slightly from 8.8 MPa to 8.3 MPa, while the tensile strength after freeze–thaw conditioning changed from 6.6 MPa to 6.5 MPa. Nevertheless, the freeze–thaw tensile strength ratio increased from 74.6% to 78.3%, corresponding to a relative increase of approximately 5.0%. The higher strength-retention ratio indicates that the DCLAM-modified mixture exhibited improved resistance to strength loss induced by moisture and freeze–thaw conditioning, although the absolute tensile strength was not increased.

Overall, the conventional mixture tests indicated that the rheological advantages of the optimized DCLAM formulation were generally reflected in the mixture-level performance, particularly in terms of high-temperature rutting resistance and low-temperature deformation capacity.

It should be noted that the total binder-related content was not identical between the two mixtures. The base asphalt mixture was prepared at an asphalt–aggregate ratio of 4.6%, whereas the DCLAM-modified mixture contained 4.0% base asphalt and 1.0% DCLAM by mass of aggregate. Because both mixtures were independently optimized using the Marshall method, the observed performance differences reflect the combined effects of mixture design and the incorporation of DCLAM. Therefore, these results should not be interpreted as isolating the intrinsic contribution of DCLAM under equal binder-content conditions. Future controlled studies using equivalent total binder content are needed to quantify the independent contribution of the modifier.

3.5.2. Dynamic Modulus Test

The dynamic modulus test was conducted to characterize the temperature- and frequency-dependent viscoelastic response of the base asphalt mixture and the DCLAM-modified asphalt mixture. The test results were evaluated in terms of the complex modulus magnitude |E*| and phase angle δ, which respectively reflect the overall stiffness response and the relative contribution of elastic and viscous components under cyclic loading.

The complex modulus results are presented in Table 18. In general, |E*| decreased with increasing temperature at a given loading frequency and increased with increasing loading frequency at a given temperature, reflecting the typical temperature- and frequency-dependent viscoelastic behavior of asphalt mixtures.

Table 18. Complex modulus of the base asphalt mixture and DCLAM-modified asphalt mixture at different temperatures and loading frequencies.

Complex Modulus of the Base Asphalt Mixture/MPa
Frequency/Hz	5 °C	10 °C	20 °C	35 °C	50 °C
0.1	5709	2850	1056	326	63
0.5	8847	5623	2095	493	132
1	10,277	6860	2794	635	181
5	13,500	10,028	5010	1253	380
10	14,986	11,594	6216	1390	523
25	16,025	12,857	7950	2232	797
Complex modulus of the DCLAM-modified asphalt mixture/MPa
Frequency/Hz	5 °C	10 °C	20 °C	35 °C	50 °C
0.1	10,740	7350	4097	1716	1399
0.5	17,202	12,304	6545	2714	1897
1	20,648	15,127	8416	3280	2532
5	28,855	22,088	14,560	6285	3779
10	33,435	26,350	17,950	8416	4913
25	37,380	29,230	22,144	12,248	6353

Taking a loading frequency of 25 Hz as an example, the complex modulus of the base asphalt mixture decreased from 16,025 MPa at 5 °C to 797 MPa at 50 °C, corresponding to a reduction of approximately 95.0%. For the DCLAM-modified asphalt mixture, the modulus decreased from 37,380 MPa to 6353 MPa, corresponding to a reduction of approximately 83.0%. Despite the reduction with increasing temperature, the DCLAM-modified mixture maintained a substantially higher modulus than the base asphalt mixture throughout the investigated temperature range. At 25 Hz, the modulus of the DCLAM-modified mixture was approximately 133.3% higher at 5 °C and 697.1% higher at 50 °C than that of the base asphalt mixture.

The influence of loading frequency was also evident. At a given temperature, increasing the loading frequency generally increased the complex modulus of both mixtures. For example, at 20 °C, the complex modulus of the base asphalt mixture increased from 1056 MPa at 0.1 Hz to 7950 MPa at 25 Hz, while that of the DCLAM-modified mixture increased from 4097 MPa to 22,144 MPa. A higher loading frequency corresponds to a shorter loading duration, which limits the development of time-dependent deformation and consequently increases the apparent stiffness of the mixture. Conversely, increasing temperature enhances the viscous contribution of the asphalt binder phase and results in a reduction in mixture stiffness.

Under all tested temperature–frequency combinations, the DCLAM-modified asphalt mixture exhibited a higher complex modulus than the base asphalt mixture. This indicates that the independently designed DCLAM-modified mixture exhibited higher load-bearing stiffness and greater resistance to deformation than the base mixture over the investigated loading conditions. This trend is consistent with the higher G* observed in the mastic-scale DSR test and the higher dynamic stability obtained from the wheel-tracking test.

The corresponding phase-angle results are listed in Table 19. In general, the DCLAM-modified asphalt mixture exhibited lower phase angles than the base asphalt mixture under all tested temperature–frequency combinations. Since the phase angle reflects the relative contributions of the elastic and viscous components in the overall viscoelastic response, a lower phase angle indicates a relatively greater elastic contribution.

Table 19. Phase angle of the base asphalt mixture and DCLAM-modified asphalt mixture at different temperatures and loading frequencies.

Phase Angle of the Base Asphalt Mixture/°
Frequency/Hz	5 °C	10 °C	20 °C	35 °C	50 °C
0.1	62.7	68.5	77.3	80.2	79.4
0.5	58.5	64.6	75.1	79.1	71.2
1	53.9	59.1	69.6	72.4	69.0
5	52.6	56.7	66.6	73.4	67.6
10	51.9	55.1	64.6	74.2	66.9
25	47.9	53.9	63.9	75.3	66.1
Phase angle of the DCLAM-modified asphalt mixture/°
Frequency/Hz	5 °C	10 °C	20 °C	35 °C	50 °C
0.1	46.8	50.5	48.6	74.5	56.2
0.5	42.6	45.8	48.4	71.5	58.3
1	40.6	43.5	47.6	69.5	59.9
5	33.9	36.7	41.7	60.3	50.7
10	31.4	33.6	36.7	56.0	45.7
25	20.1	24.4	29.2	32.7	34.2

The variation in phase angle with temperature was not strictly monotonic, particularly in the intermediate- and high-temperature ranges. For the base asphalt mixture, the phase angle generally increased with temperature up to approximately 35 °C, followed by a slight decrease at several loading frequencies. A similar non-monotonic trend was observed for the DCLAM-modified mixture under some low- and intermediate-frequency conditions. This behavior reflects the combined influence of the temperature-dependent response of the asphalt binder and the load-bearing contribution of the aggregate skeleton.

Nevertheless, the DCLAM-modified mixture consistently exhibited lower phase angles than the base mixture. For example, at 25 Hz and 50 °C, the phase angle decreased from 66.1° for the base asphalt mixture to 34.2° for the DCLAM-modified mixture. At the same loading frequency, the phase angle decreased from 47.9° to 20.1° at 5 °C, from 63.9° to 29.2° at 20 °C, and from 75.3° to 32.7° at 35 °C.

Overall, the dynamic modulus test results show that the DCLAM-modified asphalt mixture exhibited a higher complex modulus and a lower phase angle than the base asphalt mixture under all investigated temperature–frequency combinations. These results indicate a higher stiffness response together with a relatively greater elastic contribution, thereby providing mixture-level evidence for the improved resistance to load-induced deformation.

4. Conclusions

This study developed a direct-addition DCLR composite modifier and systematically evaluated its performance through orthogonal formulation optimization, asphalt–mastic rheological tests, and asphalt-mixture performance tests. The optimal formulation was determined based on DSR and BBR results, and the engineering performance of the independently designed DCLAM-modified mixture was subsequently evaluated through conventional pavement-performance and dynamic modulus tests. The main conclusions are as follows:

(1)

DSR tests demonstrated that the DCLAM improved the high-temperature rheological performance of the asphalt mastic. At 64 °C, the optimal formulation (5% DCLR, 1% SBS, 15% crumb rubber, 3.5% nano-CaCO₃, and 2% aromatic oil) exhibited a complex shear modulus of 18,506.7 Pa, representing an increase of approximately 72.1% compared with the base asphalt mastic (10,752 Pa), and a rutting factor of 19.4 kPa. Its phase angle was 72.4°, indicating a predominantly viscous response at this temperature. Nevertheless, the lower phase angle relative to that of the base asphalt mastic indicated a relatively greater elastic contribution. Combined with the increases in G* and G*/sin δ, these results demonstrate improved resistance to high-temperature deformation.

(2)

BBR test results demonstrated that the DCLAM can effectively enhance the low-temperature cracking resistance of asphalt. At −12 °C, the optimal formulation showed a creep stiffness of 167.3 MPa, which was lower than that of the base asphalt mastic (291.4 MPa), and an m-value of 0.383, which was higher than that of the base asphalt mastic (0.298). These values were comparable with the commonly used Superpave evaluation criteria and indicated improved low-temperature rheological behavior. Low-temperature susceptibility analysis indicated that the optimal formulation had an S_AS value of 28.13, lower than that of the base asphalt mastic (35.01), suggesting lower temperature sensitivity and better low-temperature stability.

(3)

The orthogonal analysis revealed clear response- and temperature-dependent factor effects within the DCLAM system. Crumb rubber predominantly controlled high-temperature stiffness and rutting response, whereas the dominant factors governing low-temperature creep stiffness shifted with temperature. On this basis, the optimal formulation was identified as 5% DCLR, 1% SBS, 15% crumb rubber, 3.5% nano-CaCO₃, and 2% aromatic oil. This formulation represents a balanced multi-response solution rather than the optimum of any single rheological indicator.

(4)

Mixture-level tests showed that the independently designed DCLAM-modified mixture exhibited favorable engineering performance relative to the base asphalt mixture. The dynamic stability increased from 2732.2 to 7176.0 cycles/mm, corresponding to an increase of approximately 162.7%. The flexural tensile strength and maximum flexural tensile strain increased by approximately 45.8% and 51.4%, respectively, while the freeze–thaw tensile strength ratio increased from 74.6% to 78.3%. In the dynamic modulus test, the DCLAM-modified mixture exhibited a higher complex modulus and a lower phase angle than the base asphalt mixture over the investigated temperature–frequency range. These mixture-level trends were generally consistent with the rheological behavior observed at the mastic scale. However, because the two mixtures had different total binder-related contents, the observed differences cannot be attributed exclusively to DCLAM itself.

Nevertheless, the present study primarily focused on the macroscopic rheological and mixture-level performance of unaged materials. Microstructural and chemical characterization techniques, such as SEM, AFM, and FTIR, were not employed in this study. Therefore, the proposed synergistic effects among the DCLAM components were inferred mainly from the observed macroscopic rheological and mixture-performance responses and have not yet been directly verified at the microstructural or molecular level. In addition, the difference in total binder-related content between the base asphalt mixture and the DCLAM-modified mixture should be considered when interpreting the mixture-level performance improvements observed in this study. Future studies should incorporate microstructural and chemical characterization, together with short- and long-term aging tests, to further clarify the modification mechanism and long-term performance of DCLAM-modified asphalt.