Structural performance of reinforced concrete beams incorporating recycled hybrid fibers under varying transverse reinforcement configurations

Abstract

This exploratory study examines the material- and member-level response of a waste-derived hybrid concrete containing processed face-mask fibers, waste binding-wire fibers, and scrap-tire rubber. Conducted in Nepal, the work was motivated by waste valorization and the need to assess locally sourced materials for resource-conscious construction. Four recycled-material combinations—BW + M + R, BW + M, BW + R, and rubber only—were screened through compressive, splitting tensile, and flexural strength tests. The BW + M + R mixture, containing nominal batching dosages of 1.5% binding wire, 0.25% mask fiber, and 0.5% rubber, produced the most balanced mechanical response and was selected for beam testing. Four reinforced concrete beams were tested under four-point loading: a conventional reference beam without recycled inclusions and with full stirrups, and three hybrid-concrete beams with full stirrups, localized stirrups in selected shear-span regions, or end stirrups only. The layouts were idealized detailing scenarios used to examine the response of the combined material–reinforcement systems as stirrup continuity decreased; they were not intended to establish an equivalent fiber-for-stirrup replacement. The reference and full-stirrup hybrid beams provide the closest direct comparison of material influence. Relative to the reference beam, the measured ultimate shear force, apparent displacement ductility, and recorded load–deflection energy of the full-stirrup hybrid beam were 13.6%, 24.7%, and 98.3% higher, respectively. The localized-stirrup hybrid beam developed 3.4% greater ultimate shear force, but its strength-normalized shear parameter and energy absorption were only 0.4% and 1.6% higher, respectively. The end-stirrup beam showed 9.5% lower ultimate shear force and 12.1% lower energy absorption, although its apparent ductility increased. The results suggest that the recycled hybrid system can provide supplementary crack-control and post-cracking benefits when used with adequate transverse reinforcement, but the tested system did not compensate for the absence of stirrups in the main shear spans. Because one beam was tested per configuration and matched non-fiber reduced-stirrup controls were not included, the findings are exploratory. The system should therefore be regarded as a waste-derived supplementary crack-control system rather than a validated substitute for conventional transverse reinforcement.

1 Introduction

The increasing generation of solid waste from healthcare, transportation, and construction activities has created an urgent need for sustainable waste management strategies. Among these waste streams, disposable face masks, scrap tires, and leftover binding wires represent materials that are difficult to manage through conventional disposal routes but may offer functional value when incorporated into cementitious composites. The widespread use of disposable face masks during and after the COVID-19 pandemic generated a large quantity of polypropylene-rich waste, while scrap tires remain a persistent environmental concern because of their long degradation period, high volume, and potential fire and pollution hazards. Similarly, short pieces of binding wire produced during construction activities are often discarded despite being metallic materials with potential crack bridging capacity. Reusing these waste-derived materials in concrete can reduce landfill burden while creating a pathway for more sustainable construction materials.

Concrete is widely used because of its availability, economy, compressive strength, and adaptability to different structural forms. However, conventional concrete is brittle in tension and relies on steel reinforcement to resist tensile and shear demands. In reinforced concrete beams, shear failure is particularly critical because it is often sudden and brittle, with limited warning before collapse. Conventional transverse reinforcement, usually in the form of stirrups, is therefore provided to control diagonal cracking and ensure adequate shear resistance. Nevertheless, closely spaced stirrups can cause reinforcement congestion, increase labor demand, reduce construction efficiency, and make concrete placement more difficult, especially in small or heavily reinforced members. These challenges have motivated increasing interest in fiber-reinforced concrete (FRC) as a supplementary crack-control system that can improve post-cracking behavior, ductility, and energy absorption.

Polypropylene fibers have been widely studied for their ability to reduce shrinkage cracking, improve crack distribution, and enhance the toughness of concrete. Disposable medical face masks are primarily composed of polypropylene-based layers, and several studies have shown that mask-derived fibers can improve selected mechanical properties of concrete when used at appropriate dosages. Kilmartin-Lynch et al. (2021) reported that low dosages of polypropylene fibers obtained from single-use face masks improved the compressive and tensile performance of concrete, although excessive fiber content reduced the strength benefits. Ahmed and Lim (2022) further demonstrated that hybridization of mask fibers with basalt fibers enhanced the mechanical performance of recycled aggregate concrete. These studies indicate that processed mask-derived polypropylene can influence tensile response and crack control, although the resulting behavior depends strongly on fiber geometry, dosage, dispersion, and interaction with the cementitious matrix (Kilmartin-Lynch et al., 2021; Ahmed and Lim, 2022).

Waste binding wire fibers offer a different reinforcing mechanism. Because they are steel-based, they can provide relatively high stiffness and tensile resistance compared with polymeric fibers. Previous studies on waste binding wire and short steel fiber reinforcement have reported improvements in tensile strength, flexural behavior, crack-control, and post-cracking resistance when suitable fiber contents are used (Ansari et al., 2021; Tanoli et al., 2014; Umar et al., 2021). In RC beams, steel fibers can also contribute to shear resistance by bridging diagonal cracks, delaying crack widening, and improving stress redistribution after cracking. However, the effectiveness of steel-type fibers depends strongly on fiber aspect ratio, distribution, anchorage, and the bond developed between the fiber and surrounding concrete matrix.

Scrap tire rubber has also been investigated as recycled material in concrete. In contrast to steel and polypropylene fibers, rubber particles generally reduce compressive strength, tensile strength, flexural strength, and elastic modulus because of their lower stiffness and weaker bond with cement paste (Girskas and Nagrockienė, 2017; Shu and Huang, 2014; Mendis et al., 2017; Snelson et al., 2009; Pacheco-Torgal et al., 2012; Thomas et al., 2016; Di Mundo et al., 2018; Guo et al., 2018; Youssf et al., 2017; Raffoul et al., 2016). Although rubber incorporation commonly reduces compressive and tensile strength and elastic modulus, rubberized concrete may exhibit enhanced deformability, toughness, impact resistance, energy dissipation, and vibration damping at appropriate rubber contents (Pacheco-Torgal et al., 2012; Bala and Gupta, 2021; Eltayeb et al., 2021). Lower thermal conductivity, improved sound absorption, and enhanced freeze–thaw performance have also been reported, although these responses depend strongly on rubber content, particle size, treatment, and mixture characteristics (Pacheco-Torgal et al., 2012; Thomas et al., 2016; Bala and Gupta, 2021). Accordingly, rubber should not be considered mechanically equivalent to a reinforcing fiber; within a hybrid system, its contribution is more plausibly associated with modification of matrix compliance and deformation response, while metallic or polymeric fibers provide the primary crack-crossing mechanisms.

Hybrid reinforcement strategies seek to combine constituents having different stiffness, geometry, and crack-control functions. Previous studies of steel–polypropylene fiber-reinforced RC beams have reported changes in shear response, post-cracking deformation, and crack distribution, including in members with reduced or absent stirrups (Sahoo and Sharma, 2014; Narayanan and Darwish, 1987; Narayanan and Darwish, 1988; Cucchiara et al., 2004; Kwak et al., 2002; Parra-Montesinos, 2006; de Montaignac et al., 2012; Sahoo et al., 2015). However, these studies predominantly employed manufactured fibers with controlled dimensions and material properties. The present system differs because the binding-wire and rubber constituents are waste-derived and geometrically nonuniform, while the polypropylene component is obtained from processed disposable face-mask material.

Recent literature further demonstrates that hybrid-fiber performance depends on constituent type, dosage, geometry, dispersion, and interaction. Azaz et al. (2026) emphasized these factors in their recent review of hybrid fiber-reinforced recycled aggregate concrete and identified mixture optimization and statistical validation as important requirements for broader application. Arairo et al. (2025) experimentally demonstrated that fiber type, length, and dosage can produce competing effects among compressive, tensile, and flexural properties. Similarly, Khatib et al. (2023) reported property trade-offs when recycled polymeric and other waste constituents were incorporated into concrete. These findings support treating the present material program as a screening investigation of competing property responses rather than as proof of an optimized synergistic formulation.

2 Materials and methods

The experimental program was designed in two stages. In the first stage, concrete mixtures containing different recycled fiber combinations were evaluated using standard material-level tests. The purpose of this stage was to identify a suitable recycled hybrid fiber combination for beam-level investigation. Compressive strength, splitting tensile strength, and flexural strength tests were performed on cylindrical and prismatic specimens prepared with plain concrete and fiber-reinforced concrete mixtures.

In the second stage, reinforced concrete beams were cast using the selected hybrid fiber combination and tested under four-point loading. The beam program was exploratory and nonfactorial, with one specimen representing each transverse-reinforcement configuration. Accordingly, the tests were intended to provide descriptive comparisons of the measured responses and to identify trends and limiting behavior of the combined material–reinforcement systems. They were not designed to establish statistically significant treatment effects, a fiber-for-stirrup replacement ratio, or a permissible reduction in transverse reinforcement. The main response parameters considered were ultimate shear force, load–deflection behavior, displacement ductility, energy absorption, crack propagation, and failure mode.

2.1 Materials

The concrete mixtures were prepared using 43-grade ordinary Portland cement, locally available natural sand, crushed coarse aggregate, water, and recycled materials. The coarse aggregate passed the 20 mm sieve and was retained on the 16 mm sieve. Three waste-derived constituents were incorporated: processed disposable face-mask fibers, scrap-tire rubber strips or chips, and waste steel binding-wire fibers. Before cutting, the face masks were cleaned, disinfected, dried, and manually processed by removing the ear loops and nose strips. The mask layers were then cut into strips approximately 23 mm long and 3 mm wide to reduce contamination risk and obtain reasonably consistent geometry.

Waste steel binding wire was cut into short elements approximately 30 mm long and 1 mm in diameter. The scrap-tire rubber consisted of irregular elongated strips or chips, generally 10–30 mm long and 1–4 mm wide, with some longer pieces. The particles had rough, nonuniform surfaces and were used without chemical treatment. The recycled materials are illustrated in Figure 1. The three constituents were not assumed to provide mechanically equivalent functions. The metallic binding-wire elements have substantially greater stiffness than the polymeric components and may provide tensile restraint when favorably oriented across developing cracks, consistent with established steel-fiber mechanisms (Ansari et al., 2021; Tanoli et al., 2014; Umar et al., 2021; Sahoo and Sharma, 2014; Narayanan and Darwish, 1987; Narayanan and Darwish, 1988; Cucchiara et al., 2004; Kwak et al., 2002; Parra-Montesinos, 2006; de Montaignac et al., 2012; Sahoo et al., 2015). The processed face-mask strips are predominantly polypropylene-based and are considerably more flexible, making distributed crack restraint a more plausible role (Kilmartin-Lynch et al., 2021; Ahmed and Lim, 2022). Scrap-tire rubber differs fundamentally from these crack-crossing constituents and is treated here as a compliant recycled inclusion rather than as structural reinforcing fiber. Its possible contribution is therefore associated primarily with modification of deformation behavior rather than direct steel-fiber-like tensile transfer (Girskas and Nagrockienė, 2017; Shu and Huang, 2014; Mendis et al., 2017; Snelson et al., 2009; Pacheco-Torgal et al., 2012; Thomas et al., 2016; Di Mundo et al., 2018; Guo et al., 2018; Youssf et al., 2017; Raffoul et al., 2016; Khatib et al., 2023).

FIGURE 1

The direct tensile properties, fiber–matrix bond, pull-out response, orientation, and individual structural contribution of these constituents were not measured in the present program. Consequently, constituent-specific mechanisms discussed later in the manuscript are literature-supported interpretations rather than independently measured contributions.

2.2 Mix proportion

A nominal M20 concrete mixture with a cement:sand:coarse-aggregate proportion of 1:1.5:3 was used as the reference mixture, and the same base proportion was adopted for the recycled-material mixtures. Four combinations were examined during material-level screening to compare the response of rubber alone and selected two- and three-component systems. The selected dosages were informed by ranges reported for mask-derived polypropylene (Kilmartin-Lynch et al., 2021; Ahmed and Lim, 2022), binding-wire and steel-fiber reinforcement (Ansari et al., 2021; Tanoli et al., 2014; Umar et al., 2021; Sahoo and Sharma, 2014; Cucchiara et al., 2004; Sahoo et al., 2015), and rubber-modified concrete (Girskas and Nagrockienė, 2017; Shu and Huang, 2014; Mendis et al., 2017; Snelson et al., 2009; Pacheco-Torgal et al., 2012; Thomas et al., 2016; Di Mundo et al., 2018; Guo et al., 2018; Youssf et al., 2017; Raffoul et al., 2016), together with practical considerations associated with local material availability and hand mixing. The recycled-material dosages were defined on a mass basis relative to the total concrete batch mass. Accordingly, the adopted contents of binding wire, processed mask fibers, and scrap-tire rubber were 1.5%, 0.25%, and 0.5% of the total batch mass, respectively. The adopted combinations are summarized in Table 1.

2.3 Casting and curing

All mixtures were prepared by hand mixing. For reference concrete, cement and fine aggregate were first blended in the dry state. Coarse aggregate was then added and mixing continued until the constituents were uniformly distributed. Water was introduced gradually while mixing continued until a visually homogeneous and workable mixture was obtained. No quantitative analysis of constituent orientation or spatial dispersion was performed; therefore, gradual addition and visual inspection should be regarded as practical measures to reduce clustering rather than verification of uniform dispersion.

The same sequence was used for the recycled-material mixtures, except that the binding-wire fibers, processed fibers, and rubber strips were introduced gradually rather than added at once. This procedure was adopted to reduce fiber balling and local segregation. Attention was given to distributing the three components throughout the mixture because nonuniform dispersion could create local weak zones and alter crack-bridging behavior. Representative stages of dry mixing, preparation of the hybrid mixture, and beam casting are shown in Figure 2.

FIGURE 2

A total of 30 cylindrical specimens and 15 prismatic specimens were prepared for material-level testing. For each of the five concrete mixtures, three cylinders were tested in compression, three cylinders were tested for splitting tensile strength, and three prisms were tested for flexural strength. Thus, each reported mechanical property was based on three replicate specimens (). Concrete was placed in three layers, and each layer was compacted manually with a tamping rod to reduce entrapped air. For the beam specimens, the reinforcement cages were positioned before concrete placement, and a nominal clear cover of 30 mm was maintained.

The specimens were left undisturbed for 24 h before demolding. The beams were cured for 28 days under wet jute coverings, whereas the cylinders and prisms were water-cured for 28 days at approximately 27 °C ± 2 °C. Before testing, the water-cured specimens were brought to a saturated surface-dry condition to remove excess surface water while maintaining internal saturation.

2.4 Test specimens and reinforcement detailing

Four RC beams were prepared. Each beam had an overall length of 1,000 mm, an effective span of 760 mm, a width of 150 mm, and an overall depth of 230 mm. A nominal clear cover of 30 mm was maintained. All beams were doubly reinforced with two 12 mm diameter Fe-500 bars in the tension zone and two 12 mm diameter Fe-500 bars in the compression zone. The two 12-mm-diameter tension bars provided a total tensile reinforcement area of approximately 226 mm². Using mm and mm, the corresponding tension reinforcement ratio, , was approximately 0.78%. Considering both tension and compression reinforcement, the total longitudinal reinforcement ratio was approximately 1.55%.

The control beam, CB, was cast using reference concrete without recycled fibers and contained two-legged 8 mm diameter stirrups at 150 mm center-to-center throughout the span. The other three beams were cast using the selected BW + M + R hybrid mixture. Selection of this mixture was based on the material-level screening results, particularly its balanced compressive, splitting tensile, and flexural strengths.

Three transverse-reinforcement layouts were used for the hybrid-fiber beams. HyFRC-FS had the same full-span stirrup arrangement as CB: two-legged 8 mm diameter stirrups at 150 mm center-to-center. HyFRC-CS contained localized two-legged 8 mm stirrups at 125 mm center-to-center in selected shear-span regions. HyFRC-ES contained end stirrups only and no conventional stirrups within the main shear spans. The layouts were selected to create idealized conditions of decreasing stirrup continuity and to examine the combined effects of the recycled hybrid system and transverse-reinforcement distribution on shear-related response, crack control, ductility, and energy absorption. They do not represent a calibrated sequence of stirrup ratios or a code-approved reduction scheme.

CB established the conventional full-stirrup reference. HyFRC-FS contained the same transverse-reinforcement arrangement as CB and therefore provided the closest direct beam-level comparison for assessing the influence of the selected recycled hybrid mixture. HyFRC-CS and HyFRC-ES differed from CB in both concrete composition and transverse-reinforcement configuration. These specimens were consequently interpreted as combined material–reinforcement systems rather than as isolated evaluations of the contribution of the recycled inclusions. HyFRC-CS represented a localized-stirrup configuration, whereas HyFRC-ES represented a limiting configuration with no conventional stirrups in the main shear spans. Because corresponding non-fiber beams with these two reinforcement layouts were not tested, no fiber-equivalent stirrup contribution, replacement percentage, or allowable reduction in transverse reinforcement was derived from these specimens. The specimen details are summarized in Table 2 and illustrated in Figure 3.

FIGURE 3

2.5 Test setup and loading arrangement

The beams were tested monotonically under four-point loading using a universal testing machine. Supports were positioned 120 mm from each end, giving an effective span of 760 mm. The two loading points were separated by approximately 253.3 mm. This arrangement produced a constant-moment region between the loading points and shear spans of approximately 253.4 mm between each support and the nearest load point. Using the effective depth of 194 mm adopted in the shear calculations, the shear span-to-effective-depth ratio was 1.31. This relatively small ratio is important for structural interpretation because short-span load transfer can involve compression-strut or arching action; this effect is considered in Section 4.2.

The four-point arrangement was selected to generate flexural cracking in the constant-moment region and diagonal or shear-flexural cracking in the shear spans. The machine reading represented the total applied load, P; under the nominally symmetric arrangement, the shear force in each span was taken as P/2. This configuration enabled comparison of load–deflection response, crack development, and failure mechanism across the four beam systems.

Each specimen was aligned carefully to minimize eccentricity, and load was increased monotonically until failure or until the test could no longer be continued safely. Total load and midspan deflection were recorded using the available load-measuring system and a dial gauge. The first visible crack, departure from the initial approximately linear response, peak load, post-peak response, and final failure condition were documented. Crack locations were marked on the beam surfaces during testing; because crack widths were not measured instrumentally, the crack-pattern comparisons are qualitative.

2.6 Material testing procedures

The testing program included material-level strength tests and beam-level structural tests. The material tests determined the 28-day compressive, splitting tensile, and flexural strengths of the reference and recycled-material mixtures. The beam tests evaluated the combined influence of the selected hybrid mixture and the transverse-reinforcement configuration. Before testing, all specimens were inspected for visible surface defects, dimensional irregularities, and casting damage. For the beams, total load, midspan deflection, visible cracking, peak response, post-peak behavior, and failure mode were recorded.

2.6.1 Compressive strength test

FIGURE 4

2.6.2 Splitting tensile strength test

FIGURE 5

2.6.3 Flexural strength test of specimens

Flexural strength tests were conducted on prismatic specimens using a universal testing machine. The supports and loading rollers were cleaned before each test. The prisms were marked 50 mm from each end to establish an effective span of 600 mm and were aligned so that the supports coincided with the marks. Two loading rollers were positioned 200 mm apart on the top surface. Load was applied gradually at 400 kgf/min, following IS 516:1959 (Bureau of Indian Standards, 1959), until failure. The maximum load and the location of the fracture line were recorded. Flexural strength values were converted to MPa for reporting in Table 3. The test setup is shown in Figure 6.

FIGURE 6

The flexural-strength expression depends on the distance, a, between the fracture line and the nearest support. The following IS 516:1959 relations (Equations 3, 4) were used.

2.7 Data analysis and scope of comparison

For the material-level tests, three replicate specimens were tested for each mixture and each mechanical property (). The results are reported as the arithmetic mean, sample standard deviation (SD), and coefficient of variation (COV) to describe within-mixture experimental variability. The sample standard deviation was calculated using degrees of freedom, and COV was calculated as . Because the material testing served primarily as a screening stage and the number of replicates was limited, no inferential significance testing was used to rank the mixtures. At beam level, each transverse-reinforcement configuration was represented by one specimen. Consequently, within-configuration variance could not be estimated, and inferential statistical testing was not appropriate. Beam-level differences in ultimate shear force, normalized shear response, apparent displacement ductility, energy absorption, and cracking behavior are therefore treated as descriptive specimen-level comparisons.

3 Results

3.1 Mechanical performance of recycled fiber concrete mixtures

Table 3 summarizes the mean compressive, splitting-tensile, and flexural strengths together with their standard deviations and coefficients of variation. Across the five mixtures, the COV ranged from 6% to 20% for compressive strength, 8%–14% for splitting-tensile strength, and 6%–12% for flexural strength. The greatest variability occurred in the compressive response of BW + R (COV = 20%) and BW + M (COV = 17%), whereas the rubber-only mixture showed the lowest compressive-strength variability (COV = 6%). Because only three replicate specimens were tested for each property, these statistics are used descriptively to characterize experimental scatter rather than to establish statistical significance.

Among the investigated combinations, BW + M + R exhibited the most balanced measured response. Its mean compressive strength was MPa (COV = 15%), its mean splitting-tensile strength was MPa (COV = 10%), and its mean flexural strength was MPa (COV = 9%). Relative to the corresponding mean values for PC, these represent increases of 6.03%, 27.27%, and 19.51%, respectively. The relatively small 6.03% difference in mean compressive strength should be interpreted cautiously because it is small compared with the observed within-mixture scatter. In contrast, the larger differences in the mean splitting-tensile and flexural strengths contributed more strongly to the selection of BW + M + R as the balanced candidate mixture. Nevertheless, the conventional strength tests do not quantify post-cracking tensile transfer, fracture energy, fiber pull-out, or constituent-specific crack bridging; therefore, the observed response is interpreted at the mixture level rather than as evidence of a quantitatively verified synergistic mechanism.

BW + R produced the highest mean compressive strength, MPa, corresponding to 9.88% above PC. However, it also exhibited the largest compressive-strength variability (COV = 20%), indicating substantial specimen-to-specimen scatter. Its mean splitting-tensile and flexural strengths were 8.64% and 18.34% lower than PC, respectively. Thus, the highest mean compressive strength did not correspond to a consistently balanced response across the three mechanical properties.

The rubber-only mixture exhibited a mean compressive strength of MPa (COV = 6%) and a mean splitting-tensile strength of MPa (COV = 9%), corresponding to reductions of 22.74% and 13.64% relative to PC. Its mean flexural strength, MPa (COV = 6%), was only 1.95% below PC; given the experimental scatter, this small difference should be regarded as broadly comparable at the descriptive level rather than as evidence of a meaningful reduction. The reductions in compressive and splitting-tensile strength are consistent with previous studies reporting strength penalties associated with tire-rubber incorporation (Girskas and Nagrockienė, 2017; Shu and Huang, 2014; Mendis et al., 2017; Snelson et al., 2009; Pacheco-Torgal et al., 2012; Thomas et al., 2016; Di Mundo et al., 2018; Guo et al., 2018; Youssf et al., 2017; Raffoul et al., 2016). Rubberized concrete has also been reported to exhibit enhanced deformability, toughness, impact-related response, energy dissipation, and vibration damping at suitable rubber contents (Pacheco-Torgal et al., 2012; Bala and Gupta, 2021; Eltayeb et al., 2021); however, these characteristics were not independently quantified in the present material tests.

BW + M exhibited the highest mean splitting-tensile strength, MPa (COV = 14%), corresponding to 36.82% above PC. However, its mean compressive strength was MPa (COV = 17%) and its mean flexural strength was MPa (COV = 11%), representing reductions of 48.09% and 13.46%, respectively. The present tests do not identify the cause of this pronounced property trade-off. Factors such as constituent dispersion, workability, entrapped air, and local clustering may influence fiber-containing mixtures but were not quantified in the present program. The contrasting response nevertheless demonstrates the importance of considering multiple mechanical properties during mixture screening, consistent with the broader fiber-reinforced-concrete behavior discussed by Arairo et al. (2025).

Based on the combined mean-strength response and observed variability, BW + M + R was selected for the beam tests because it provided the most balanced combination of compressive, splitting-tensile, and flexural performance among the mixtures examined. This selection represents a screening-based experimental choice and should not be interpreted as identification of an optimized dosage or a quantitatively verified synergistic formulation.

3.2 Load-deflection behavior of test beams

Figure 7 compares the total load–midspan deflection responses. Interpretation of these curves requires recognition of the experimental design. CB and HyFRC-FS had identical transverse-reinforcement layouts and therefore provide the closest beam-level material comparison, subject to the small difference in measured concrete compressive strength. HyFRC-CS and HyFRC-ES differed from CB in both material composition and stirrup arrangement; their responses consequently represent complete material–detailing systems rather than isolated material effects.

FIGURE 7

HyFRC-FS sustained the highest measured load and the largest recorded deflection range among the tested specimens, with a comparatively gradual post-peak reduction. Because the stirrup layout was identical to CB, this specimen-level difference is consistent with a supplementary contribution from the BW + M + R system. A plausible interpretation is that the continuous stirrups provided anchored resistance across diagonal cracks while favorably oriented inclusions supplied additional local crack restraint. However, fiber orientation, crack-crossing frequency, pull-out behavior, and residual tensile transfer were not measured.

HyFRC-CS reached a peak response close to CB while sustaining a larger recorded deformation. Because both the concrete composition and stirrup distribution changed, this observation applies to the combined BW + M + R/localized-stirrup system and cannot be converted into a fiber-equivalent reduction in stirrups.

HyFRC-ES reached a lower peak load than CB despite its higher apparent displacement ductility index. It therefore provides an important limiting observation that the greater apparent deformability did not preserve the strength level of the conventional full-stirrup reference when conventional stirrups were absent from the main shear spans.

Overall, HyFRC-FS combined the highest peak load with the longest recorded deformation range. When stirrup continuity was reduced, the peak resistance moved toward the CB level in HyFRC-CS and below it in HyFRC-ES, even though both specimens retained higher apparent displacement ductility indices than CB. Thus, deformation capacity and shear resistance did not vary proportionally across the four configurations.

3.3 Ultimate shear force and normalized shear response

For the symmetric four-point loading arrangement, the ultimate shear force in each shear span was calculated as , where is the measured total peak load. The resulting ultimate shear forces were 73.5 kN for CB, 83.5 kN for HyFRC-FS, 76.0 kN for HyFRC-CS, and 66.5 kN for HyFRC-ES. Relative to CB, these correspond to descriptive specimen-level differences of +13.6%, +3.4%, and −9.52% for HyFRC-FS, HyFRC-CS, and HyFRC-ES, respectively. To express the shear response on a common stress basis, the nominal shear stress was calculated using Equation 5.where is the beam width and is the effective depth. Values of mm and mm were used, giving mm². Because all four beams had identical cross-sectional dimensions, conversion from to does not alter the relative ranking of the specimens; rather, it expresses the measured shear resistance in stress terms. The corresponding values were 2.526 MPa for CB, 2.869 MPa for HyFRC-FS, 2.612 MPa for HyFRC-CS, and 2.285 MPa for HyFRC-ES.

The BW + M + R concrete had a mean compressive strength of 28.63 MPa compared with 27.00 MPa for PC. Because concrete compressive strength contributes to shear resistance, comparison based only on or could partly reflect this difference in material strength. A square-root strength normalization was therefore used as a comparative sensitivity measure using Equation 6.where is the measured mean compressive strength of the corresponding mixture. The parameter is used only to examine the influence of the difference in concrete strength on the relative beam responses; it is not presented as a code-defined design resistance or as a direct measure of fiber shear contribution.

Table 4 presents the absolute shear-response parameters. To make the effect of strength normalization more evident, the responses were also expressed relative to CB using and , as shown in Figure 8. A value of 1.0 therefore represents the response of CB, while values above or below 1.0 indicate greater or lower responses, respectively.

FIGURE 8

HyFRC-FS showed the clearest favorable response. Its relative ultimate shear-force ratio was , corresponding to a 13.6% higher measured shear force than CB. After accounting comparatively for the difference in compressive strength, the ratio decreased to , but remained approximately 10.3% above CB. Thus, the somewhat higher compressive strength of the BW + M + R concrete explains part, but not all, of the measured difference between the two specimens. Because CB and HyFRC-FS had identical geometry and continuous-stirrup detailing, the remaining difference is compatible with a supplementary system-level contribution from the BW + M + R mixture. Mechanically, favorably oriented binding-wire and mask-derived inclusions may provide local restraint across developing cracks while the continuous closed stirrups provide the principal anchored transverse-tie mechanism. This interpretation remains specimen-specific because the beams were not replicated and the individual inclusion contribution was not measured directly.

HyFRC-CS exhibited a distinctly different trend. Its raw relative shear ratio was , suggesting a small 3.4% difference above CB. After compressive-strength normalization, however, the ratio decreased to . The normalized response is therefore essentially the same as that of CB. This indicates that the small difference in raw shear resistance should not be interpreted as evidence that the localized-stirrup configuration was structurally superior to the conventional reference. Moreover, because both concrete composition and transverse-reinforcement distribution changed simultaneously, the measured response cannot be partitioned into separate contributions from the recycled inclusions and the remaining stirrups.

HyFRC-ES provides the clearest limiting response. Its relative ultimate shear-force ratio was , and its strength-normalized ratio decreased further to . Thus, the specimen remained below CB even though the BW + M + R concrete had the higher mean compressive strength. The lower shear resistance therefore cannot be attributed to weaker concrete. Instead, the response is mechanically consistent with the absence of continuous anchored transverse reinforcement in the main shear spans. Distributed inclusions may intersect individual cracks and provide local crack restraint, but their effectiveness depends on orientation, bond, and intersection with the active crack plane; unlike closed stirrups, they do not form a continuous and deliberately oriented tensile-tie system across the shear span.

The short shear span of the tested beams provides additional context for these trends. At , load transfer likely involved substantial diagonal compression-strut or arching action between the loading and support regions. Within this load-transfer mechanism, the concrete compression field, longitudinal reinforcement, transverse stirrups, and distributed inclusions contribute through different mechanical roles. Consequently, the strength-normalized parameter should be interpreted only as a comparative correction for concrete compressive strength and not as a means of separating or quantifying the individual concrete, inclusion, and stirrup contributions. The role of the short-span compression-strut and transverse-tie mechanisms is discussed further in Section 4.2.

3.4 Comparative apparent displacement ductility

A comparative displacement ductility index was calculated as Equation 7.where is the displacement ductility index, is the displacement at the final recorded response point, and is the apparent yield displacement. Longitudinal-bar strains were not measured; therefore, was identified from the first clear departure of the load–deflection curve from its initial approximately linear trend. The resulting parameter is therefore an apparent comparative index rather than a direct reinforcement-yield-based ductility measure.

CB exhibited an apparent displacement ductility index of 2.15. The corresponding values were 2.63 for HyFRC-ES, 2.68 for HyFRC-FS, and 2.67 for HyFRC-CS, representing descriptive differences of approximately +22.3%, +24.7%, and +24.2%, respectively.

The consistently higher indices of the BW + M + R specimens are compatible with a less abrupt post-cracking response, but the mechanism cannot be assigned quantitatively to individual constituents. Metallic binding-wire elements provide a plausible mechanism for tensile transfer across relatively larger cracks when favorably oriented (Ansari et al., 2021; Tanoli et al., 2014; Umar et al., 2021; Sahoo and Sharma, 2014; Narayanan and Darwish, 1987; Narayanan and Darwish, 1988; Cucchiara et al., 2004; Kwak et al., 2002; Parra-Montesinos, 2006; de Montaignac et al., 2012; Sahoo et al., 2015), whereas the more flexible mask-derived polypropylene fibers may contribute to distributed crack restraint (Kilmartin-Lynch et al., 2021; Ahmed and Lim, 2022). Rubber is more appropriately interpreted as a compliant inclusion rather than as a tensile crack-bridging fiber (Girskas and Nagrockienė, 2017; Shu and Huang, 2014; Mendis et al., 2017; Snelson et al., 2009; Pacheco-Torgal et al., 2012; Thomas et al., 2016; Di Mundo et al., 2018; Guo et al., 2018; Youssf et al., 2017; Raffoul et al., 2016; Bala and Gupta, 2021; Eltayeb et al., 2021; Khatib et al., 2023). These constituent-level roles remain literature-supported interpretations rather than directly measured mechanisms.

HyFRC-ES is particularly important because its higher apparent displacement ductility did not translate into preservation of either shear resistance or recorded load–deflection energy. Thus, within the tested configuration, improved apparent deformability alone was insufficient to compensate for the absence of conventional stirrups in the main shear spans.

3.5 Energy absorption

Recorded load–deflection energy was calculated as the area under each experimental load–deflection curve using trapezoidal numerical integration (Equation 8) and the resulting values are summarized in Table 5.where is the energy absorbed over the recorded load–deflection response, and are two consecutive load values, and and are the corresponding midspan deflection values. Because the numerical integration was performed to the final reliably recorded displacement of each specimen rather than to a common post-peak displacement or residual-load criterion, the resulting values are used as comparative measures within the present test series and should not be interpreted as standardized toughness properties.

The energy ranking reflected the combined effects of sustained load level and deformation range rather than apparent ductility alone. HyFRC-FS recorded the largest integrated energy, 1,600.11 kN·mm, followed by HyFRC-CS at 820.20 kN·mm, CB at 806.95 kN·mm, and HyFRC-ES at 709.21 kN·mm. Relative to CB, these correspond to specimen-level differences of +98.29%, +1.64%, and −12.11% for HyFRC-FS, HyFRC-CS, and HyFRC-ES, respectively.

HyFRC-ES is particularly informative because its apparent displacement ductility was higher than that of CB, yet its recorded energy was lower. Thus, greater relative deformation capacity did not necessarily produce greater energy absorption; the specimen sustained a lower load over its recorded response. Conversely, the much larger integrated energy of HyFRC-FS resulted from the combination of a higher sustained load and a longer deformation range. HyFRC-CS recorded an energy value very close to CB, indicating that its greater apparent ductility was accompanied by only a marginal difference in integrated energy. Because the integration endpoints differed among specimens and only one beam represented each configuration, these energy differences should be interpreted as specimen-level comparisons rather than statistically established material properties.

3.6 Crack propagation, failure mechanisms, and structural implications

Figures 9–12 show the crack development and final failure patterns of CB, HyFRC-FS, HyFRC-CS, and HyFRC-ES. The observed crack patterns varied with the combined material and transverse-reinforcement configurations. Because crack widths were not measured instrumentally, the comparison is limited to crack location, distribution, and qualitative localization. CB developed a dominant diagonal crack from the support region toward the nearest loading point at a total applied load of approximately 140 kN. The concentration of damage into one major diagonal crack indicates a shear-dominated and relatively localized failure response.

FIGURE 9

FIGURE 10

FIGURE 11

FIGURE 12

HyFRC-ES first developed a flexural crack near midspan at approximately 125 kN, followed by a major diagonal crack at approximately 127 kN. Its final response remained shear-dominated. Because crack widths were not measured instrumentally, no quantitative comparison of crack-opening magnitude is made. The lower ultimate shear force nevertheless shows that the tested end-stirrup configuration did not maintain the response of the conventional full-stirrup reference.

HyFRC-CS developed a shear-flexural crack in a shear span at approximately 118 kN, followed by flexural cracks near midspan at approximately 131 kN. Three flexural cracks were documented in the constant-moment region. Compared with CB, the documented cracking was more distributed; however, because both material composition and transverse-reinforcement arrangement changed, their respective contributions cannot be separated.

HyFRC-FS exhibited the most distributed cracking. Initial flexural cracks were observed at approximately 154 kN, and a shear-flexural crack formed at approximately 157 kN. Six cracks were recorded, predominantly flexural, together with shear-flexural and diagonal cracks. The more distributed pattern, together with the larger measured load and recorded energy response, is consistent with a favorable combined response of the BW + M + R system and continuous stirrups. The underlying crack-transfer mechanism, however, was not measured directly.

The crack patterns changed systematically with transverse-reinforcement continuity. CB and HyFRC-ES showed stronger localization into dominant diagonal cracking, whereas HyFRC-FS and HyFRC-CS developed more distributed flexural and shear-flexural cracking. This contrast is compatible with a more distributed stress-transfer response when the BW + M + R system acts together with conventional transverse reinforcement. In HyFRC-FS, continuous stirrups provided anchored crack-crossing resistance throughout the shear spans, while the distributed recycled inclusions may have provided supplementary local crack restraint. When stirrup continuity was reduced in HyFRC-CS, the cracking remained comparatively distributed, but the strength-normalized shear response and recorded load–deflection energy were essentially comparable to CB rather than clearly superior. Moreover, HyFRC-ES provides the contrasting boundary condition. Although its apparent displacement ductility was greater than that of CB, the specimen developed a dominant diagonal shear crack and exhibited lower ultimate shear force, strength-normalized shear response, and recorded load–deflection energy. Thus, greater apparent deformability and the presence of distributed inclusions did not reproduce the structural role of continuous anchored transverse reinforcement in the main shear spans. Because crack widths, fiber–crack intersections, inclusion orientation, and reinforcement strains were not measured, the observed crack patterns cannot be used to quantify the independent contributions of the recycled inclusions and stirrups.

4 Discussion

4.1 Material-to-structural response of the waste-derived hybrid system

The material screening results indicate that the three constituents should be interpreted as an integrated system rather than as independently beneficial additions. Processed mask-derived polypropylene has been reported to influence crack distribution and tensile-related properties at suitable dosages (Kilmartin-Lynch et al., 2021; Ahmed and Lim, 2022), while steel-type fibers can transfer tensile stresses across cracks when their bond, orientation, and anchorage are favorable (Ansari et al., 2021; Tanoli et al., 2014; Umar et al., 2021; Sahoo and Sharma, 2014; Narayanan and Darwish, 1987; Narayanan and Darwish, 1988; Cucchiara et al., 2004; Kwak et al., 2002; Parra-Montesinos, 2006; de Montaignac et al., 2012; Sahoo et al., 2015). In contrast, tire rubber generally reduces stiffness and conventional strength because of its compliant nature and comparatively weak interface with cement paste, although it can modify deformation characteristics (Girskas and Nagrockienė, 2017; Shu and Huang, 2014; Mendis et al., 2017; Snelson et al., 2009; Pacheco-Torgal et al., 2012; Thomas et al., 2016; Di Mundo et al., 2018; Guo et al., 2018; Youssf et al., 2017; Raffoul et al., 2016; Khatib et al., 2023). Recent hybrid-fiber literature similarly emphasizes the importance of constituent geometry, dosage, dispersion, and interaction (Azaz et al., 2026; Arairo et al., 2025).

Within the present screening matrix, BW + M + R avoided the pronounced property trade-offs observed in several two-component mixtures and therefore provided the most balanced measured response. This finding should not be interpreted as proof of synergistic action because the constituent-level mechanisms were not independently measured. Moreover, the CB–HyFRC-FS pair provides the closest member-level material comparison because both specimens had identical geometry and continuous-stirrup detailing. HyFRC-FS exhibited greater measured ultimate shear force, more distributed documented cracking, a higher apparent displacement ductility index, and greater recorded load–deflection energy than CB. These specimen-level differences are consistent with a supplementary contribution from the BW + M + R system acting alongside anchored transverse reinforcement, similar in concept to complementary fiber–stirrup behavior reported for engineered steel and synthetic fibers (Sahoo and Sharma, 2014; Cucchiara et al., 2004; Kwak et al., 2002; Sahoo et al., 2015). However, residual tensile transfer, inclusion orientation, pull-out resistance, crack-opening kinematics, and repeatability were not quantified. Consequently, the present results support system-level interpretation rather than a quantitatively verified constituent mechanism.

4.2 Role of stirrup continuity and the short shear-span configuration

The localized- and end-stirrup specimens help define the boundary of the tested material–reinforcement system. HyFRC-CS remained close to CB in both ultimate and strength-normalized shear response; however, its concrete composition and stirrup arrangement changed simultaneously. In the absence of a matched non-fiber beam with the same localized-stirrup layout, this response cannot be translated into an allowable reduction in transverse reinforcement. HyFRC-ES provides a clearer limiting observation. Although its apparent displacement ductility was greater than that of CB, its ultimate shear force, strength-normalized shear response, and recorded load–deflection energy were lower. Thus, improved deformability did not compensate for the loss of continuous transverse reinforcement in the main shear spans.

The short shear span provides an important mechanical context for these results. With , load transfer was likely governed by a combination of diagonal compression-strut or arching action and transverse tensile resistance rather than by ordinary sectional shear behavior alone. A diagonal compression field can transfer a substantial portion of the applied load from the loading region toward the support, while closed stirrups crossing the potential diagonal-crack path act as anchored transverse ties. In CB and HyFRC-FS, this tie mechanism was available throughout the shear spans. In HyFRC-CS, the transverse-tie system was only locally available, whereas in HyFRC-ES it was essentially absent from the main shear spans. The recycled inclusions may provide local crack restraint when favorably oriented across an active crack, but they do not form the continuous, anchored tensile load path provided by conventional stirrups. This load-transfer interpretation is consistent with the strut-and-tie concepts used for discontinuity regions in ACI CODE-318–25 and EN 1992-1–1:2023 (ACI Committee 318, 2025; European Committee for Standardization CEN, 2023). Accordingly, the measured responses should be regarded as configuration-specific short-span behavior rather than directly compared with ordinary slender-beam shear expressions or used to derive an equivalent fiber shear resistance. Testing at larger ratios is required to determine whether the same material–reinforcement trends persist as direct strut action diminishes and conventional beam action becomes more dominant.

4.3 Practical relevance, scalability, and sustainability potential

The use of processed disposable face-mask material, binding-wire offcuts, and scrap-tire rubber provides a potential waste-valorization pathway that may be relevant to resource-constrained construction environments such as Nepal. However, successful implementation of alternative materials depends on more than measured mechanical performance. Mattar et al. (2026), for example, identified material performance, quality control, technological capability, regulatory conditions, and project-level factors as interconnected considerations for sustainable-material adoption in a developing-country construction context.

The present study therefore demonstrates experimental material reuse and selected structural responses rather than environmental superiority. No life-cycle assessment, cost analysis, durability program, or field-scale production study was conducted. Waste diversion alone does not establish net sustainability because collection, cleaning, processing, quality control, service life, maintenance, and end-of-life effects must also be considered. Accordingly, sustainability benefits of the proposed system should be regarded as potential rather than demonstrated.

5 Conclusion

6 Limitations and future recommendations

The beam program was exploratory and nonfactorial. Only one specimen represented each transverse-reinforcement configuration; therefore, specimen-to-specimen variability and statistical reliability could not be quantified at the beam level. Matching non-fiber beams were also unavailable for the localized- and end-stirrup configurations. Replicated paired specimens with systematically varied transverse-reinforcement ratios are therefore required to distinguish material effects from reinforcement-layout effects and to establish the repeatability of the observed structural trends.

At the material level, three replicate specimens were tested for each mixture and each mechanical property. Experimental variability was characterized using the sample standard deviation and coefficient of variation; however, the limited replication () supports descriptive assessment of scatter rather than strong inferential conclusions regarding differences among mixtures. Accordingly, the material-level comparisons were used primarily for mixture screening rather than statistical ranking or optimization.

The recycled-material dosages were defined on a mass basis relative to the total concrete batch mass. Because verified constituent densities were not measured during the original experimental program, equivalent volume fractions were not reconstructed retrospectively. Future studies should report constituent contents in both kg/m³ and verified volume fraction and should characterize inclusion length distribution, aspect ratio, tensile properties, spatial distribution, orientation, fiber–matrix bond, and pull-out response.

Residual flexural strength, direct tensile response, fracture energy, and constituent pull-out behavior were not measured. Longitudinal-bar and stirrup strains and quantitative crack-width measurements were also unavailable. Consequently, the apparent yield displacement and the constituent-level mechanisms associated with crack bridging and post-cracking stress transfer could not be verified directly. Future studies should incorporate residual flexural or direct-tensile testing, reinforcement strain measurements, LVDTs, and quantitative crack-monitoring techniques such as digital image correlation.

The tested shear-span-to-effective-depth ratio was ; therefore, compression-strut or arching action likely contributed to load transfer. The findings are consequently specific to the tested short-shear-span geometry and should not be generalized directly to slender RC beams. Additional testing over a wider range of ratios and beam sizes is required to determine whether the observed material–reinforcement trends persist as the response transitions toward conventional beam action.

Finally, recorded load–deflection energy was integrated to the final reliably recorded point of each test rather than to a common post-peak displacement or residual-load criterion. The reported energy values should therefore be interpreted as comparative measures within the present test series rather than standardized toughness properties. Future studies should adopt a consistent termination criterion and report energy to peak load and to defined post-peak thresholds. Practical implementation also requires assessment of workability, mixing reproducibility, bond to conventional reinforcement, shrinkage, creep, durability, cost, life-cycle impacts, and field-scale performance.

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