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[Jonathon Tanks](https://orcid.org/0000-0002-0232-8240), [Kimiyoshi Naito](https://orcid.org/0000-0002-3334-4876), Hisai Ueda

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[Characterization of the Static, Creep, and Fatigue Tensile Behavior of Basalt Fiber/Polypropylene Composite Rods for Passive Concrete Reinforcement](https://mdr.nims.go.jp/datasets/fbc10c84-ae73-454f-a7f5-94d5a3353b4e)

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Characterization of the Static, Creep, and Fatigue Tensile Behavior of Basalt Fiber/Polypropylene Composite Rods for Passive Concrete ReinforcementpolymersArticleCharacterization of the Static, Creep, and Fatigue TensileBehavior of Basalt Fiber/Polypropylene Composite Rods forPassive Concrete ReinforcementJonathon Tanks 1,*, Kimiyoshi Naito 1,2 and Hisai Ueda 3�����������������Citation: Tanks, J.; Naito, K.; Ueda,H. Characterization of the Static,Creep, and Fatigue Tensile Behaviorof Basalt Fiber/PolypropyleneComposite Rods for Passive ConcreteReinforcement. Polymers 2021, 13,3136. https://doi.org/10.3390/polym13183136Academic Editor: Mauro ZarrelliReceived: 23 August 2021Accepted: 14 September 2021Published: 16 September 2021Publisher’s Note: MDPI stays neutralwith regard to jurisdictional claims inpublished maps and institutional affil-iations.Copyright: © 2021 by the authors.Licensee MDPI, Basel, Switzerland.This article is an open access articledistributed under the terms andconditions of the Creative CommonsAttribution (CC BY) license (https://creativecommons.org/licenses/by/4.0/).1 Research Center for Structural Materials, National Institute for Materials Science, 1-2-1 Sengen, Ibaraki,Tsukuba 305-0047, Japan; naito.kimiyoshi@nims.go.jp2 Department of Aerospace Engineering, Tohoku University, 6-6-1 Aramaki-aza-Aoba, Miyagi,Sendai 305-0047, Japan3 Innovative Composite Materials Research and Development Center, Kanazawa Institute of Technology,2-2 Yatsukaho, Ishikawa, Hakusan 924-0838, Japan; h-ueda@neptune.kanazawa-it.ac.jp* Correspondence: tanks.jonathon@nims.go.jpAbstract: Fiber-reinforced polymer (FRP) composites are becoming more frequently adopted asso-called “corrosion-resistant” concrete reinforcement materials due to their excellent mechanicalproperties and formability. However, their long-term reliability must be thoroughly investigatedin order to understand failure mechanisms and to develop service life models. This study is onthe mechanical properties of a prototype basalt fiber-reinforced polypropylene (BFPP) rod underquasi-static and sustained loading. Static strength and modulus at elevated temperatures do notdecrease significantly, but the variability in strength increases with temperature, as shown by aWeibull analysis. Creep behavior is typical of unidirectional FRP, where the creep rupture strengthfollows a power law. Fatigue at various stress ratios R reveals the sensitivity of composite strengthto the matrix damage, which increases at lower values of R (i.e., higher stress amplitudes). Theseresults are discussed in the context of service life and concrete structure design guidelines.Keywords: thermoplastic composite; basalt fiber; fatigue properties; creep properties1. IntroductionReinforced and prestressed concrete is the most common structural system in theworld, given its low cost per unit weight and formability [1]. However, corrosion of thesteel reinforcing/prestressing materials such as bars and strands leads to concrete crackingdue to internal pressure caused by low-density iron oxide byproducts; costly repairs andeven replacement can outweigh the initial construction costs [1–3]. Epoxy-coating carbonsteel and stainless steel reinforcing products are obvious alternatives, but fiber-reinforcedpolymer (FRP) composites are becoming more frequently adopted as so-called “corrosion-resistant” concrete reinforcement materials due to their excellent mechanical properties,low density, and resistance to galvanic corrosion [4–8]. Numerous studies on FRP forstructural reinforcement are reported every year, covering topics such as environmentaldurability [9–16] and material mechanics [16,17].Considering the time scale of service life for a concrete structure, long-term durabilityand reliability of the reinforcing materials are extremely important. Thus, creep andfatigue studies provide crucial data and analysis regarding the long-term performanceof FRP reinforcement and methods for service life prediction. Extensive work has beenconducted on three common types of composites: carbon (CFRP), glass (GFRP), andaramid (AFRP) [18–27]. Extrapolating available experimental data to at least 50-yearservice periods, numerous studies found that, while CFRP exhibits excellent creep andfatigue resistance (less than 20% decrease in strength), GFRP and AFRP tend to show drasticdeterioration of mechanical properties (as much as 90% decrease in strength) [18,19,21–27].Polymers 2021, 13, 3136. https://doi.org/10.3390/polym13183136 https://www.mdpi.com/journal/polymershttps://www.mdpi.com/journal/polymershttps://www.mdpi.comhttps://doi.org/10.3390/polym13183136https://doi.org/10.3390/polym13183136https://creativecommons.org/https://creativecommons.org/licenses/by/4.0/https://creativecommons.org/licenses/by/4.0/https://doi.org/10.3390/polym13183136https://www.mdpi.com/journal/polymershttps://www.mdpi.com/article/10.3390/polym13183136?type=check_update&version=2Polymers 2021, 13, 3136 2 of 13While CFRP is clearly superior in terms of mechanical reliability and durability, its highcost is a hindrance toward its widespread use in infrastructure compared with cheapermaterials such as GFRP [4,8,22].Basalt fibers, which are drawn from basalt rocks and come at relatively low cost,have recently gained more attention as an alternative to glass fibers due to having supe-rior mechanical properties [28–31]. The creep behavior of basalt FRP (BFRP)—especiallycontaining epoxy matrices—has been investigated on cylindrical bars [32,33], and fatiguehas been studied in various geometries with respect to failure mechanisms [34–37], stressratios [38], matrix type [39], and environmental conditions [40]. Several studies foundsignificant reductions in stiffness over the fatigue life caused by increased matrix damageand fiber rupture at longer cycles, with a 107-cycle fatigue strength around 70–75% ofthe initial static strength [37–39]. However, the effect of stress ratio R (i.e., the ratio ofminimum and maximum applied stresses in a sinusoidal cycle) has not been investigatedin great depth, as only R = 0.1, 0.5, and 0.8 were reported [37,38]. These studies found thatthe failure mode of basalt/epoxy laminates changes from interfacial debonding to fiberrupture as R decreases. However, as matrix damage and interfacial debonding are largelymatrix-dominant, a comparison with other matrix resins is needed.Furthermore, polypropylene—a low-cost commodity polymer with excellent moistureresistance—has been used as a matrix for GFRP and BFRP in some studies, showing lowerstrength and modulus than a thermosetting epoxy [15,27,41–45]. The fatigue behavior ofglass fiber-reinforced polypropylene (GFPP) was discovered to improve when polypropy-lene was modified with maleic-anhydride (MA) due to the improved interfacial bondingwith the glass fibers, which resulted in a fiber-rupture failure mode rather than interfa-cial debonding (at R = 0.1) [41,42]. In particular, the stiffness degradation was minimalwith MA-modified GFPP in stark contrast to the epoxy- or polyester-based GFRP [42,43].Furthermore, MA-modified PP sizing for BFPP composites showed roughly 20% higherflexural strength compared with neat PP [41,45].Although glass and basalt share some similarities in their chemical composition, simi-lar studies on the long-term mechanical behavior of basalt fiber-reinforced polypropylene(BFPP) were not found in the literature (to the authors’ knowledge). This paper reports thestatic, creep, and fatigue properties of a prototype BFPP rod that is intended for passivereinforcement (i.e., non-prestressed) in the concrete foundation of high-speed railwaysystems. The thermal and mechanical properties of the rapidly produced thermoplasticcomposite were primarily evaluated by mechanical testing (static and fatigue) and electronmicroscopy, supported by an analysis of the material service life considering that theinfluence of stress ratio R is presented.2. Materials and Methods2.1. Materials and PreparationThe material in this study was a heat-resistant basalt fiber (Nakagawa Sangyo Co.,Ltd., Inuyama, Japan) with a matrix made from blended polypropylene (Prima Polymer)and maleic acid-grafted polypropylene (Mitsui Chemical, Tokyo, Japan, MA: 0.25 wt%).Straight rods were manufactured by a pultrusion technique at a rate of 15 m/min bythe following process: m-PP pellets were melted and transferred to a resin bath via ascrew extruder, where the BF rovings were impregnated before being fed through a dieand subsequent water cooling, and finally collected in bundles of seven rods, whichwere twisted to form a stranded cable. The fiber volume fraction was measured as 0.44(by cross-sectional analysis). The low fiber content is due to the prioritization of excesspolypropylene for additional chemical resistance. The cables were slightly twisted (angleof approximately 11◦, measured by digital microscope) to retain the diameter and fiberconsolidation during pultrusion. The diameter of the straight rods was 4.38 mm, and theywere cut to 500 mm lengths for tensile testing. All tensile test specimens (static, creep, andfatigue) were prepared by fixing steel tubes (inner diameter 16 mm and length 200 mm) toboth ends by an expansive grout (Bristar 100, non-explosive demolition grout) and left toPolymers 2021, 13, 3136 3 of 13cure for at least three days before testing (Figure 1a); alignment was ensured by enclosingthe tubes with PVC caps having concentric holes (diameter ~4.5 mm), and the specimenswere secured in steel angles. It was found that the expansive pressure from the grout couldcause early fatigue failure inside the tubes, so a polyurethane coating was applied in thegripping region to more evenly distribute the gripping pressure [46].Polymers 2021, 13, x FOR PEER REVIEW 3 of 12   approximately 11°, measured by digital microscope) to retain the diameter and fiber con-solidation during pultrusion. The diameter of the straight rods was 4.38 mm, and they were cut to 500 mm lengths for tensile testing. All tensile test specimens (static, creep, and fatigue) were prepared by fixing steel tubes (inner diameter 16 mm and length 200 mm) to both ends by an expansive grout (Bristar 100, non-explosive demolition grout) and left to cure for at least three days before testing (Figure 1a); alignment was ensured by enclos-ing the tubes with PVC caps having concentric holes (diameter ~4.5 mm), and the speci-mens were secured in steel angles. It was found that the expansive pressure from the grout could cause early fatigue failure inside the tubes, so a polyurethane coating was applied in the gripping region to more evenly distribute the gripping pressure [46].  Figure 1. (a) Basalt fiber-reinforced polypropylene (BFPP) composite rod (center rod from a seven-wire strand) and tensile specimen diagram (all dimensions in mm), (b) static tensile test setup including localized heater, (c) fatigue tensile test setup, and (d) creep tensile test setup. 2.2. Characterization Methods Static tensile tests were conducted at a displacement rate of 1 mm/min on a universal testing machine (Autograph AGX, Shimadzu, Kyoto, Japan; Figure 1b), and strain was measured by foil gauges and a video displacement system (TRViewX, Shimadzu). In ad-dition to room temperature (~23 °C), higher temperatures of 80 (±2) and 120 (±2) °C were applied through a local heating device (with internal K thermocouple) to avoid slippage in the gripping area. Once the testing temperature was reached, the specimen was allowed to equilibrate for one hour before conducting the tensile test. Ten specimens were tested for all temperatures. Fatigue tensile tests were conducted at room temperature on a servohydraulic testing machine (Servopulser, Shimadzu; Figure 1c) under force control at a frequency of 10 Hz and stress ratios R (=σmin/σmax) of 0.1, 0.3, 0.5, 0.7, and 0.9; this frequency was selected be-cause it allows for faster turnover of the testing equipment without introducing internal heating effects [39,47], and no particular frequency is specified by ASTM [48]. Strain was measured by foil gauges connected to a datalogger. One specimen was tested at each stress level, with run-out (endurance limit) set to Ne = 107 cycles. Creep tests (R = 1) were also conducted at room temperature on a lever-arm creep machine (Shimadzu; Figure 1d) at five different stress levels; 2000 h was chosen as the termination time. To further investigate the thermomechanical properties of the BFPP, dynamic me-chanical analysis (DMA 7100, Hitachi Hi-Tech, Tokyo, Japan) was conducted in flexural mode and differential scanning calorimetry (DSC 7020, Hitachi Hi-Tech) was performed using roughly 7–10 mg of neat polymer blend (called m-PP) over a range of 30–200 °C at a heating rate of 10 °C/min (heat–cool–heat). Fracture surfaces were observed by scanning electron microscope (Quanta 200, FEI, Hillsboro, OR, USA). Localized heaterVideo cameraSpecimenData acquisitionSpecimen(b) (c)500 200200 4.38 20Steel tubeExpansive groutBFPP specimenBFPP Tensile specimen(a)SpecimenWeights(d)Figure 1. (a) Basalt fiber-reinforced polypropylene (BFPP) composite rod (center rod from a seven-wire strand) and tensilespecimen diagram (all dimensions in mm), (b) static tensile test setup including localized heater, (c) fatigue tensile testsetup, and (d) creep tensile test setup.2.2. Characterization MethodsStatic tensile tests were conducted at a displacement rate of 1 mm/min on a universaltesting machine (Autograph AGX, Shimadzu, Kyoto, Japan; Figure 1b), and strain wasmeasured by foil gauges and a video displacement system (TRViewX, Shimadzu). Inaddition to room temperature (~23 ◦C), higher temperatures of 80 (±2) and 120 (±2) ◦Cwere applied through a local heating device (with internal K thermocouple) to avoidslippage in the gripping area. Once the testing temperature was reached, the specimen wasallowed to equilibrate for one hour before conducting the tensile test. Ten specimens weretested for all temperatures.Fatigue tensile tests were conducted at room temperature on a servohydraulic testingmachine (Servopulser, Shimadzu; Figure 1c) under force control at a frequency of 10 Hzand stress ratios R (=σmin/σmax) of 0.1, 0.3, 0.5, 0.7, and 0.9; this frequency was selectedbecause it allows for faster turnover of the testing equipment without introducing internalheating effects [39,47], and no particular frequency is specified by ASTM [48]. Strain wasmeasured by foil gauges connected to a datalogger. One specimen was tested at each stresslevel, with run-out (endurance limit) set to Ne = 107 cycles. Creep tests (R = 1) were alsoconducted at room temperature on a lever-arm creep machine (Shimadzu; Figure 1d) atfive different stress levels; 2000 h was chosen as the termination time.To further investigate the thermomechanical properties of the BFPP, dynamic me-chanical analysis (DMA 7100, Hitachi Hi-Tech, Tokyo, Japan) was conducted in flexuralmode and differential scanning calorimetry (DSC 7020, Hitachi Hi-Tech) was performedusing roughly 7–10 mg of neat polymer blend (called m-PP) over a range of 30–200 ◦C at aheating rate of 10 ◦C/min (heat–cool–heat). Fracture surfaces were observed by scanningelectron microscope (Quanta 200, FEI, Hillsboro, OR, USA).Polymers 2021, 13, 3136 4 of 133. Results and Discussion3.1. Thermomechanical PropertiesThermal analysis of the neat polymer by DSC (Figure 2a) revealed the melting temper-ature Tm to be 166.0 ◦C (onset around 145.7 ◦C) and a crystallinity χc of 35.8%, which aretypical values for m-PP [43]. The storage modulus E’ of the BFPP as measured by DMA(f = 10 Hz) showed typical behavior, with a gradual reduction in stiffness followed by asharp drop at the onset of melting; no influence from the fibers on PP melting was detected.Reductions in E’ of 35.1% and 66.2% were observed at the selected static tensile test tem-peratures of 80 and 120 ◦C, respectively (marked by star symbols). A linear dependence oflog(E’) vs. log(f ) can be seen in Figure 2b, consistent with the literature [49].Polymers 2021, 13, x FOR PEER REVIEW 4 of 12   3. Results and Discussion 3.1. Thermomechanical Properties Thermal analysis of the neat polymer by DSC (Figure 2a) revealed the melting tem-perature Tm to be 166.0 °C (onset around 145.7 °C) and a crystallinity χc of 35.8%, which are typical values for m-PP [43]. The storage modulus E’ of the BFPP as measured by DMA (f = 10 Hz) showed typical behavior, with a gradual reduction in stiffness followed by a sharp drop at the onset of melting; no influence from the fibers on PP melting was de-tected. Reductions in E’ of 35.1% and 66.2% were observed at the selected static tensile test temperatures of 80 and 120 °C, respectively (marked by star symbols). A linear depend-ence of log(E’) vs. log(f) can be seen in Figure 2b, consistent with the literature [49].  Figure 2. (a) Thermal properties of a m-PP matrix and BFPP measured by DSC and DMA, and (b) frequency dependence of the storage modulus of BFPP measured by DMA. 3.2. Static Properties at Elevated Temperatures The static tensile strength and modulus at room temperature were 733.5 MPa and 26.7 GPa, respectively; the constitutive behavior was mostly linearly elastic with a small inelastic portion near failure (Figure 3a). At higher temperatures, the strength shows an insignificant decrease (<2%) even at 120 °C, while the modulus decreases slightly (9.5% and 10% at 80 and 120 °C, respectively) but not with statistical significance. Tensile prop-erties at each temperature are listed in Table 1 and shown in Figure 3b. Despite significant softening occurring in the m-PP matrix at higher temperatures, the fiber-dominant prop-erties of the unidirectional composite do not significantly decline. This means that, alt-hough the tensile strength of the BFPP is not comparable with other standard materials such as GFRP or CFRP, the performance retention at sub-melting temperatures shows promise for thermoplastic composite reinforcing rods. Table 1. Static tensile properties of BFPP rods. Temperature (°C) σu (MPa) EL (GPa) m (-) ~23 733.5 26.7 24.55 80 733.5 23.9 21.67 120 718.8 23.7 17.83 The tensile strengths at each temperature were fit to a two-parameter Weibull distri-bution (total of n specimens, with i from 1 to n) [50]: ln [ln (11 − 𝑃𝐹)] = 𝑚[ln(𝜎𝑢) − ln(𝜎0)] (1) 678910-3000-2500-2000-1500-1000-500050020 40 60 80 100 120 140 160 180Storage Modulus, log(E’) [Mpa]Heat FlowTemperature [C]DSCDMATm = 166 Cc = 35.8%(a) (b)6789100.01 0.1 1 10 100 1000Frequency [Hz]305080120150CFrequency for fatigue testing(f = 10 Hz)Figure 2. (a) Thermal properties of a m-PP matrix and BFPP measured by DSC and DMA, and (b) frequency dependence ofthe storage modulus of BFPP measured by DMA.3.2. Static Properties at Elevated TemperaturesThe static tensile strength and modulus at room temperature were 733.5 MPa and26.7 GPa, respectively; the constitutive behavior was mostly linearly elastic with a smallinelastic portion near failure (Figure 3a). At higher temperatures, the strength shows aninsignificant decrease (<2%) even at 120 ◦C, while the modulus decreases slightly (9.5% and10% at 80 and 120 ◦C, respectively) but not with statistical significance. Tensile properties ateach temperature are listed in Table 1 and shown in Figure 3b. Despite significant softeningoccurring in the m-PP matrix at higher temperatures, the fiber-dominant properties ofthe unidirectional composite do not significantly decline. This means that, although thetensile strength of the BFPP is not comparable with other standard materials such asGFRP or CFRP, the performance retention at sub-melting temperatures shows promise forthermoplastic composite reinforcing rods.Table 1. Static tensile properties of BFPP rods.Temperature (◦C) σu (MPa) EL (GPa) m (-)~23 733.5 26.7 24.5580 733.5 23.9 21.67120 718.8 23.7 17.83Polymers 2021, 13, 3136 5 of 13The tensile strengths at each temperature were fit to a two-parameter Weibull distri-bution (total of n specimens, with i from 1 to n) [50]:ln[ln(11 − PF)]= m[ln(σu)− ln(σ0)] (1)PF =in + 1(2)where PF is the cumulative probability of failure at the applied tensile stress σu, m is theWeibull shape parameter, and σ0 is the characteristic stress or Weibull scale parameter. Thehigher the value of m, the lower the probability of fracture at stresses approaching themean. Figure 3c shows the Weibull distributions for each temperature, which revealedthat m decreases linearly with increasing temperature (Figure 3d). The polymer matrixsoftens at higher temperatures, which ultimately reduces the interfacial properties andincreases the scatter in strength values. This has implications for BFPP rods used at elevatedtemperatures, such as the curing of prestressed concrete, which takes place at ~60 ◦C for6–12 h [51]. While concrete curing temperatures and service temperatures are not likelyto exceed 80 ◦C in most cases and, therefore, the mean tensile strength is not expected todecrease significantly, the decrease in m suggests that failure can occur more frequentlyat stresses well below the statistical mean and should be accounted for when consideringsafety factors in design and construction.Polymers 2021, 13, x FOR PEER REVIEW 5 of 12   𝑃𝐹 =𝑖𝑛 + 1 (2) where 𝑃𝐹 is the cumulative probability of failure at the applied tensile stress 𝜎𝑢, m is the Weibull shape parameter, and 𝜎0 is the characteristic stress or Weibull scale parameter. The higher the value of m, the lower the probability of fracture at stresses approaching the mean. Figure 3c shows the Weibull distributions for each temperature, which revealed that m decreases linearly with increasing temperature (Figure 3d). The polymer matrix softens at higher temperatures, which ultimately reduces the interfacial properties and increases the scatter in strength values. This has implications for BFPP rods used at ele-vated temperatures, such as the curing of prestressed concrete, which takes place at ~60 °C for 6–12 h [51]. While concrete curing temperatures and service temperatures are not likely to exceed 80 °C in most cases and, therefore, the mean tensile strength is not ex-pected to decrease significantly, the decrease in m suggests that failure can occur more frequently at stresses well below the statistical mean and should be accounted for when considering safety factors in design and construction.  Figure 3. (a) Representative stress–strain curves for tensile tests, (b) a summary of the tensile properties at each tempera-ture; (c) Weibull plot of tensile strengths, and (d) Weibull shape parameters at each temperature. Specimens at room temperature exhibited a typical broom-like failure mode, where the fiber twist is clearly visible (Figure 4a). At room temperature, failure mainly consists of cohesive failure, indicated by matrix hackles and a residual matrix adhered to the ex-posed fiber surfaces. Conversely, the failure mode became more localized at higher tem-peratures due to the softening of the polypropylene, which reduces brittle fracture and increases the probability of localized fiber stress concentrations during loading. Micro-graphs of the fracture surfaces (Figure 4b) show more fiber fragmentation and debonding 15202530356006507007508000 25 50 75 100 125 150Modulus, EL[GPa]Strength, u[MPa]Temperature [C]StrengthModulus02004006008000 0.01 0.02 0.03 0.04Stress [MPa]Strain [mm/mm]-3-2-10126.2 6.4 6.6 6.8ln(-ln(1-PF))ln(u)RT80°C120°C10152025300 20 40 60 80 100 120Weibull Shape Parameter, mTemperature [C](c) (d)(b)RT80 C120 C(a)Figure 3. (a) Representative stress–strain curves for tensile tests, (b) a summary of the tensile properties at each temperature;(c) Weibull plot of tensile strengths, and (d) Weibull shape parameters at each temperature.Polymers 2021, 13, 3136 6 of 13Specimens at room temperature exhibited a typical broom-like failure mode, wherethe fiber twist is clearly visible (Figure 4a). At room temperature, failure mainly consists ofcohesive failure, indicated by matrix hackles and a residual matrix adhered to the exposedfiber surfaces. Conversely, the failure mode became more localized at higher temperaturesdue to the softening of the polypropylene, which reduces brittle fracture and increasesthe probability of localized fiber stress concentrations during loading. Micrographs of thefracture surfaces (Figure 4b) show more fiber fragmentation and debonding as temperatureincreases, which is an expected outcome considering the significant softening of the matrixdiscussed in Section 3.1. This corroborates the results of the Weibull analysis regardingthe increase in scatter despite small changes in the mean. Fracture surfaces of individualfibers (Figure 4c) become slightly more angular at higher temperatures, but no significantdifference was noted; this is expected for fibers marketed as heat-resistant.Polymers 2021, 13, x FOR PEER REVIEW 6 of 12   as temperature increases, which is an expected outcome considering the significant sof-tening of the matrix discussed in Section 3.1. This corroborates the results of the Weibull analysis regarding the increase in scatter despite small changes in the mean. Fracture sur-faces of individual fibers (Figure 4c) become slightly more angular at higher temperatures, but no significant difference was noted; this is expected for fibers marketed as heat-re-sistant.  Figure 4. Observation of fractured specimens at higher temperatures: (a) specimen appearance after tensile testing, (b) fiber and matrix condition, and (c) fracture surface of individual fibers (all scale bars are 10 μm). 3.3. Creep Behavior (R = 1) at Room Temperature Creep can be considered a special case of fatigue where R = 1, since there is no fluc-tuation in the applied load but failure still occurs at stresses below the static tensile strength. The creep rupture stress σcr over time is shown in Figure 5a, along with the total strain (elastic + creep) for the stress level σcr/σu = 0.77. The creep rupture data follows a typical power law: 𝜎𝑐𝑟 = 𝑎(𝑡𝑓)𝑏 (3) where a and b are empirical parameters (listed in Table 2), and tf is the time-to-failure. The creep endurance stress level (i.e., terminated at 2000 h) was σe,cr/σu = 0.75, but extrapolating with Equation (3) to 106 h (114 years) yields σe,cr/σu = 0.71. The strain also shows typical creep behavior, with a steady state region in the short-term and a sudden increase shortly before failure. Strain data from other specimens were not recoverable, but this specimen shows the anticipated behavior for FRP. The micrographs in Figure 5b show more damage in the resin for longer creep dura-tions, comparing 10 min with 2000 h. The residual static tensile strength was measured immediately following the termination of the 2000 h creep test, showing no statistically significant change (<0.7%). Creep rupture at short durations despite less matrix damage suggests that internal defects and strength distribution affect the probability of fiber-dom-inated failure at higher stress levels, since more matrix damage appears to be tolerable at (a)(b)(c)RT (~23 C) 80 C 120 CFigure 4. Observation of fractured specimens at higher temperatures: (a) specimen appearance after tensile testing, (b) fiberand matrix condition, and (c) fracture surface of individual fibers (all scale bars are 10 µm).3.3. Creep Behavior (R = 1) at Room TemperatureCreep can be considered a special case of fatigue where R = 1, since there is nofluctuation in the applied load but failure still occurs at stresses below the static tensilestrength. The creep rupture stress σcr over time is shown in Figure 5a, along with the totalstrain (elastic + creep) for the stress level σcr/σu = 0.77. The creep rupture data follows atypical power law:σcr = a(t f)b(3)where a and b are empirical parameters (listed in Table 2), and tf is the time-to-failure. Thecreep endurance stress level (i.e., terminated at 2000 h) was σe,cr/σu = 0.75, but extrapolatingwith Equation (3) to 106 h (114 years) yields σe,cr/σu = 0.71. The strain also shows typicalPolymers 2021, 13, 3136 7 of 13creep behavior, with a steady state region in the short-term and a sudden increase shortlybefore failure. Strain data from other specimens were not recoverable, but this specimenshows the anticipated behavior for FRP.The micrographs in Figure 5b show more damage in the resin for longer creep dura-tions, comparing 10 min with 2000 h. The residual static tensile strength was measuredimmediately following the termination of the 2000 h creep test, showing no statisticallysignificant change (<0.7%). Creep rupture at short durations despite less matrix damagesuggests that internal defects and strength distribution affect the probability of fiber-dominated failure at higher stress levels, since more matrix damage appears to be tolerableat longer durations. Further experimental work is needed to reveal the damage accumula-tion mechanism under creep loading in BFPP.Polymers 2021, 13, x FOR PEER REVIEW 7 of 12   longer durations. Further experimental work is needed to reveal the damage accumula-tion mechanism under creep loading in BFPP.  Figure 5. (a) Creep rupture strength diagram and creep strain until failure, and (b) comparison of fracture surfaces for short and long creep lives (scale bars are 50 μm). Table 2. Fatigue life parameters of BFPP rods. R a b σe,max/σu 0.1 1211.1 −0.092 0.15 0.3 1054.0 −0.116 0.20 0.5 663.0 −0.115 0.20 0.7 575.6 −0.114 0.25 0.9 391.5 −0.121 0.40 1.0 607.8 −0.011 0.75 * * Extrapolation to 106 h gives 0.71. 3.4. Fatigue Behavior (0 < R < 1) at Room Temperature Figure 6a shows the S–N curves in terms of mean stress σm for each value of R, while Figure 6b shows the S–N curves in terms of stress amplitude σa. Replicates were not tested for each stress level so a statistical analysis could not be performed, but the data appear to follow a power law similar to Equation (3): 𝜎𝑚 = 𝑎(𝑁𝑓)𝑏 (4) where a and b are empirical parameters, and Nf is the cycles to failure; a and b are listed in Table 2, along with the maximum stress at the endurance limit (σe,max/σu). It is clear that a smaller stress amplitude (higher R) results in a higher tolerable mean stress for a given fatigue life while a higher mean stress results in a lower tolerable stress amplitude for a given fatigue life. It is easy to understand that a smaller stress amplitude creates less dam-age and thus correlates to higher mean stress. From a reliability perspective, for a given mean stress, a higher stress amplitude translates to a higher probability that the maximum stress approaches the mean static strength, as shown by the Weibull distribution (Section 3.2). This is especially clear in the fatigue failure diagram in Figure 6c, where the relation-ship between stress amplitude and mean stress are linear for each value of R, and the fatigue endurance (run-out) line is formed by the smallest values from each series. A va-riety of fatigue failure criteria has been developed for metal alloys and applied to compo-site materials [52–54], most notably the Goodman criterion, which is given by the follow-ing: (b) 10 min2000 hrFigure 5. (a) Creep rupture strength diagram and creep strain until failure, and (b) comparison of fracture surfaces for shortand long creep lives (scale bars are 50 µm).Table 2. Fatigue life parameters of BFPP rods.R a b σe,max/σu0.1 1211.1 −0.092 0.150.3 1054.0 −0.116 0.200.5 663.0 −0.115 0.200.7 575.6 −0.114 0.250.9 391.5 −0.121 0.401.0 607.8 −0.011 0.75 ** Extrapolation to 106 h gives 0.71.3.4. Fatigue Behavior (0 < R < 1) at Room TemperatureFigure 6a shows the S–N curves in terms of mean stress σm for each value of R, whileFigure 6b shows the S–N curves in terms of stress amplitude σa. Replicates were not testedfor each stress level so a statistical analysis could not be performed, but the data appear tofollow a power law similar to Equation (3):σm = a(N f)b(4)where a and b are empirical parameters, and Nf is the cycles to failure; a and b are listedin Table 2, along with the maximum stress at the endurance limit (σe,max/σu). It is clearthat a smaller stress amplitude (higher R) results in a higher tolerable mean stress for agiven fatigue life while a higher mean stress results in a lower tolerable stress amplitudefor a given fatigue life. It is easy to understand that a smaller stress amplitude createsless damage and thus correlates to higher mean stress. From a reliability perspective, forPolymers 2021, 13, 3136 8 of 13a given mean stress, a higher stress amplitude translates to a higher probability that themaximum stress approaches the mean static strength, as shown by the Weibull distribution(Section 3.2).Polymers 2021, 13, x FOR PEER REVIEW 8 of 12   𝜎𝑎 = 𝜎𝑤 (1 −𝜎𝑚𝜎𝑢) (5) where σw is an upper bound on the stress amplitude when the mean stress approaches zero—i.e., fully reversed loading. However, this criterion clearly does not fit the experi-mental data for BFPP, as the fatigue limit is overestimated. Other common criteria such as Soderberg and Gerber are also more suitable for metal alloys, which exhibit a yielding behavior, and thus, the fatigue limit is affected by plastic deformation [52]. A new empir-ical criterion is introduced here to more accurately represent the experimental results of this study: 𝜎𝑎 = 𝜎𝑤 + 𝐴(𝜎𝑚)2 + 𝐵𝜎𝑚 (6) where the coefficients σw, A, and B are defined by the following: 𝜎𝑤 =2.5(𝜎𝑢)2𝐸𝐿 (7) 𝐴 =2.5𝐸𝐿=𝜎𝑤(𝜎𝑢)2 (8) 𝐵 = −5𝜎𝑢𝐸𝐿= −2𝐴𝜎𝑢 (9) where EL is the longitudinal tensile modulus. This formulation is expressed in terms of maximum strain energy density at failure for an elastic material. In this context, the fatigue endurance limit σe follows a convex surface where a higher mean stress corresponds to an increasingly lower tolerable stress amplitude. The comparison between the Goodman cri-terion and the proposed quadratic criterion is shown in Figure 6c (enlarged in the inset).  Figure 6. Fatigue life plotted by (a) mean stress σm and (b) stress amplitude σa, (c) fatigue life diagram comparing Goodman and quadratic (convex) failure criterion (Equation (6)), and (d) endurance limit σe and residual strength σr corresponding to different stress ratios R. Figure 6. Fatigue life plotted by (a) mean stress σm and (b) stress amplitude σa, (c) fatigue life diagram comparing Goodmanand quadratic (convex) failure criterion (Equation (6)), and (d) endurance limit σe and residual strength σr corresponding todifferent stress ratios R.This is especially clear in the fatigue failure diagram in Figure 6c, where the rela-tionship between stress amplitude and mean stress are linear for each value of R, andthe fatigue endurance (run-out) line is formed by the smallest values from each series.A variety of fatigue failure criteria has been developed for metal alloys and applied tocomposite materials [52–54], most notably the Goodman criterion, which is given by thefollowing:σa = σw(1 − σmσu)(5)where σw is an upper bound on the stress amplitude when the mean stress approaches zero—i.e., fully reversed loading. However, this criterion clearly does not fit the experimental datafor BFPP, as the fatigue limit is overestimated. Other common criteria such as Soderbergand Gerber are also more suitable for metal alloys, which exhibit a yielding behavior, andthus, the fatigue limit is affected by plastic deformation [52]. A new empirical criterion isintroduced here to more accurately represent the experimental results of this study:σa = σw + A(σm)2 + Bσm (6)Polymers 2021, 13, 3136 9 of 13where the coefficients σw, A, and B are defined by the following:σw =2.5(σu)2EL(7)A =2.5EL=σw(σu)2 (8)B = −5σuEL= −2Aσu (9)where EL is the longitudinal tensile modulus. This formulation is expressed in terms ofmaximum strain energy density at failure for an elastic material. In this context, the fatigueendurance limit σe follows a convex surface where a higher mean stress corresponds toan increasingly lower tolerable stress amplitude. The comparison between the Goodmancriterion and the proposed quadratic criterion is shown in Figure 6c (enlarged in the inset).3.5. Implications for Service Life and DesignThe endurance limit for each stress ratio is shown in Figure 6d, where σe increaseswith R according to the following:σe = 1.5σw((1 − R) + 1.5σw(1 + R2σcr,e))−1(10)where σe,cr is the creep endurance strength (i.e., σe at R = 1). The residual static strength σrwas measured from specimens after reaching run-out (N = 107 cycles) and was found toincrease with R according to the following:σr = 5σw((1 − R) + 5σw(1 + R2σu))−1(11)which is similar to Equation (10) except that the static tensile strength is used as the upperbound instead of the creep endurance strength. As mentioned above, more damage is accu-mulated at higher stress amplitudes (lower R), so the residual static strength is significantlyreduced. Conversely, the post-creep (R = 1) residual static strength is unchanged. This issupported by the micrographs in Figure 7, which show significant resin damage for lowervalues of R while higher values do not differ from static tensile fracture surfaces. Thisdeviates from results for basalt/epoxy composites [38,39], which is assumed to be causedby the difference in stiffness and ductility between epoxy and polypropylene. However,these results may extend the observations of the effect of limited ductility for toughenedvinylester-based BFRP [39]. No significant reduction in stiffness (E/E0) was observed forany of the loading conditions in this study due to the fiber-dominant behavior of unidi-rectional composites. Further experimental and analytical investigation is needed to fullycharacterize and quantify damage under sustained loading.Existing standards and design guides for FRP concrete reinforcement do not includeBFRP and nearly all referenced data come from brittle thermoset matrix composites, so wereference the guidelines made for GFRP as it is the most similar to BFRP. The AmericanConcrete Institute (ACI) published the ACI 440.1R-15 “Guide for the Design and Con-struction of Structure Concrete Reinforced with Fiber-Reinforced Polymer (FRP) Bars”,which is currently the most comprehensive document on the topic [55]. Section 7.4 of ACI440.1R-15 addresses creep rupture and fatigue limits, where a maximum long-term stressof σmax = 0.2σu is recommended for GFRP. Referring to the endurance limits for each R inTable 2, BFPP exhibits similar values ranging from 0.15σu to 0.40σu (for R = 0.1 and 0.9,respectively) and 0.71 for creep (R = 1). Although ACI 440.1R-15 does not mention stressratios, in-service structures experience variable loading scenarios that make life predictioncomplex, which is why conservative stress limits are needed for safe designs. More ex-Polymers 2021, 13, 3136 10 of 13perimental data and theoretical analysis are needed to better understand the fatigue andcreep behavior of basalt fiber/thermoplastic composites and to develop accurate servicelife prediction models.Polymers 2021, 13, x FOR PEER REVIEW 9 of 12   3.5. Implications for Service Life and Design The endurance limit for each stress ratio is shown in Figure 6d, where σe increases with R according to the following: 𝜎𝑒 = 1.5𝜎𝑤 ((1 − 𝑅) + 1.5𝜎𝑤 (1 + 𝑅2𝜎𝑐𝑟,𝑒))−1 (10) where σe,cr is the creep endurance strength (i.e., σe at R = 1). The residual static strength σr was measured from specimens after reaching run-out (N = 107 cycles) and was found to increase with R according to the following: 𝜎𝑟 = 5𝜎𝑤 ((1 − 𝑅) + 5𝜎𝑤 (1 + 𝑅2𝜎𝑢))−1 (11) which is similar to Equation (10) except that the static tensile strength is used as the upper bound instead of the creep endurance strength. As mentioned above, more damage is ac-cumulated at higher stress amplitudes (lower R), so the residual static strength is signifi-cantly reduced. Conversely, the post-creep (R = 1) residual static strength is unchanged. This is supported by the micrographs in Figure 7, which show significant resin damage for lower values of R while higher values do not differ from static tensile fracture surfaces. This deviates from results for basalt/epoxy composites [38,39], which is assumed to be caused by the difference in stiffness and ductility between epoxy and polypropylene. However, these results may extend the observations of the effect of limited ductility for toughened vinylester-based BFRP [39]. No significant reduction in stiffness (E/E0) was ob-served for any of the loading conditions in this study due to the fiber-dominant behavior of unidirectional composites. Further experimental and analytical investigation is needed to fully characterize and quantify damage under sustained loading.  Figure 7. (a) Micrographs of fatigue run-out specimens showing more resin damage at lower R values and (b) fracture surfaces of individual fibers (scale bars are 10 μm). Existing standards and design guides for FRP concrete reinforcement do not include BFRP and nearly all referenced data come from brittle thermoset matrix composites, so R = 0.9, max = 0.7u R = 0.7, max = 0.5u R = 0.5, max = 0.6u(a)(b)Figure 7. (a) Micrographs of fatigue run-out specimens showing more resin damage at lower R values and (b) fracturesurfaces of individual fibers (scale bars are 10 µm).4. ConclusionsThis paper reports the static and fatigue tensile behavior of a novel basalt fiber/polypropylene composite rod for concrete reinforcement. In particular, the elucidation ofthe effect of stress ratio on fatigue life of BFPP, and the proposal of a failure criterion andthe relationship between endurance limit and residual strength are major contributions ofthis study. To summarize, the static tensile properties at elevated temperatures (T < Tm)decreased slightly but the fiber-dominant nature of unidirectional composites resulted in asmaller decrease than expected based on the neat resin’s thermal properties. Rather, themost notable change was an increase in the variability in strength as temperature increased,as indicated by a decrease in the Weibull shape parameter. Fatigue behavior was similar toother FRP (particularly GFRP) in terms of the general relationship between stress level andfatigue life; however, the endurance limit deviated from standard failure criteria such asthe Goodman equation, instead being better described by a strain energy density-basedquadratic (convex) function. Additionally, we found that a higher stress ratio R (i.e., lowerstress amplitude) resulted in a higher residual static strength after run-out, with no changefor creep loading (2000 h run-out). A lower stress amplitude corresponds to less damageaccumulation in the m-PP matrix, although the damage mechanism for creep (R = 1) isyet unclear. Several equations were introduced to describe the fatigue endurance limitand residual strength, showing good agreement with experimental data. The strength andstiffness of this prototype cable are insufficient for prestressed concrete applications, but itis a promising candidate material for passive concrete reinforcement due to its durabilityand low cost. Although further investigation is needed to thoroughly characterize fatiguePolymers 2021, 13, 3136 11 of 13damage mechanisms and to accurately predict fatigue life, these results suggest that theBFPP material in this study is comparable with other FRP and seems conforms to ACI440.R-15.Author Contributions: Conceptualization, K.N. and H.U.; methodology, K.N. and J.T.; investigation,J.T. and K.N.; resources, K.N. and H.U.; data curation, J.T. and K.N.; writing—original draft prepara-tion, J.T.; writing—review and editing, J.T. and K.N.; visualization, J.T.; supervision, K.N.; projectadministration, K.N.; funding acquisition, K.N. and H.U. All authors have read and agreed to thepublished version of the manuscript.Funding: This research was funded by COI program “Construction of next-generation infrastructureusing innovative materials—Realization of safe and secure society that can coexist with the Earth forcenturies” supported by Japan Science and Technology Agency (JST) grant number JPMJCE1315.Data Availability Statement: Requests for experimental data may be considered on a case-by-casebasis.Conflicts of Interest: The authors declare no conflict of interest.References1. Koch, G.; Brongers, M.; Thompson, N.; Virmani, Y.; Payer, J. Corrosion Costs and Preventative Strategies in the United States; FederalHighway Administration Report FHWA-RD-01-156; Federal Highway Administration: Washington, DC, USA, 2003.2. Val, D.V.; Stewart, M.G. Life-cycle cost analysis of reinforced concrete structures in marine environments. Struct. Saf. 2003, 25,343–362. [CrossRef]3. Cheung, M.M.S.; So, K.K.L.; Zhang, X. Life cycle cost management of concrete structures relative to chloride-induced reinforce-ment corrosion. Struct. Infrastruct. Eng. 2012, 8, 1136–1150. [CrossRef]4. Zdanowicz, K.; Kotynia, R.; Marx, S. Prestressing concrete members with fibre-reinforced polymer reinforcement: State ofresearch. Struct. Concr. 2019, 20, 872–885. [CrossRef]5. Balaguru, P.; Nanni, A.; Giancaspro, J. FRP Composites for Reinforced and Prestressed Concrete Structures: A Guide to Fundamentals andDesign for Repair and Retrofit; Taylor: New York, NY, USA, 2009.6. Meier, U. Composite materials in bridge repair. Appl. Compos. Mater. 2000, 7, 75–94. [CrossRef]7. Maissen, A. Concrete beams prestressed with CFRP strands. Struct. Eng. Int. 1997, 7, 284–287. [CrossRef]8. Gudonis, E.; Timinskas, E.; Brigniak, V.; Kaklauskas, G.; Arnautov, A.; Vamulenas, V. FRP reinforcement for concrete structures:State-of-the-art review of application and design. Eng. Struct. Technol. 2013, 5, 147–158. [CrossRef]9. Tanks, J.D.; Sharp, S.R.; Harris, D.K. Kinetics of in-plane shear degradation in carbon/epoxy rods from exposure to alkaline andsaline environments. Compos. Part B 2017, 110, 204–212. [CrossRef]10. Ali, A.H.; Mohamed, H.M.; Benmokrane, B.; El Safty, A. Theory-based approaches and microstructural analysis to evaluatethe service life-retention of stressed carbon fiber composite strands for concrete bridge applications. Compos. Part B 2019, 165,279–292. [CrossRef]11. Bazli, M.; Zhao, X.-L.; Jafari, A.; Ashrafi, H.; Bai, Y.; Raman, S.; Khezrzadeh, H. Mechanical properties of pultruded GFRP profilesunder seawater and concrete environment coupled with UV radiation and moisture. Construct. Build. Mater. 2020, 258, 120369.[CrossRef]12. Tanks, J.D.; Kubouchi, M.; Arao, Y. Diffusion kinetics, swelling, and degradation of corrosion-resistant C-glass/epoxy wovencomposites in harsh environments. Compos. Struct. 2018, 202, 686–694. [CrossRef]13. Lu, Z.; Xie, J.; Zhang, H.; Li, J. Long-term durability of basalt fiber-reinforced polymer (BFRP) sheets and the epoxy resin matrixunder a wet-dry cyclic condition in a choride-containing environment. Polymers 2017, 9, 652. [CrossRef]14. Hashim, U.R.; Jumahat, A.; Jawaid, M.; Dungani, R.; Alamery, S. Effects of accelerated weathering on degradation behavior ofbasalt fiber reinforced polymer nanocomposites. Polymers 2020, 12, 2621. [CrossRef] [PubMed]15. Tang, C.; Xu, F.X.; Li, G. Combustion performance and thermal stability of basalt fiber-reinforced polypropylene composites.Polymers 2019, 11, 1826. [CrossRef] [PubMed]16. Mohamed, O.A.; Al Hawat, W.; Keshawarz, M. Durability and mechanical properties of concrete reinforced with basalt fiber-reinforced polymer (BFRP) bars: Toward sustainable infrastructure. Polymers 2021, 13, 1402. [CrossRef]17. Ricciardi, M.R.; Papa, I.; Coppola, G.; Lopresto, V.; Sansone, L.; Antonucci, V. Effect of plasma treatment on the impact behaviorof epoxy/basalt fiber-reinforced composites: A preliminary study. Polymers 2021, 13, 1293. [CrossRef] [PubMed]18. Berardi, V.P.; Perrella, M.; Feo, L.; Cricri, G. Creep behavior of GFRP laminates and their phases: Experimental investigation andanalytical modeling. Compos. Part B 2017, 122, 136–144. [CrossRef]19. Wu, Z.; Wang, X.; Iwashita, K.; Sasaki, T.; Hamaguchi, Y. Tensile fatigue behavior of FRP and hybrid FRP sheets. Compos. Part B2010, 41, 396–402. [CrossRef]20. Noel, M. Probabilistic fatigue life modelling of FRP composites for construction. Construct. Build. Mater. 2019, 206, 279–286.[CrossRef]http://doi.org/10.1016/S0167-4730(03)00014-6http://doi.org/10.1080/15732479.2010.507474http://doi.org/10.1002/suco.201800347http://doi.org/10.1023/A:1008919824535http://doi.org/10.2749/101686697780494536http://doi.org/10.3846/2029882X.2014.889274http://doi.org/10.1016/j.compositesb.2016.10.092http://doi.org/10.1016/j.compositesb.2018.11.083http://doi.org/10.1016/j.conbuildmat.2020.120369http://doi.org/10.1016/j.compstruct.2018.03.078http://doi.org/10.3390/polym9120652http://doi.org/10.3390/polym12112621http://www.ncbi.nlm.nih.gov/pubmed/33172162http://doi.org/10.3390/polym11111826http://www.ncbi.nlm.nih.gov/pubmed/31698868http://doi.org/10.3390/polym13091402http://doi.org/10.3390/polym13081293http://www.ncbi.nlm.nih.gov/pubmed/33921019http://doi.org/10.1016/j.compositesb.2017.04.015http://doi.org/10.1016/j.compositesb.2010.02.001http://doi.org/10.1016/j.conbuildmat.2019.02.082Polymers 2021, 13, 3136 12 of 1321. Demers, C.E. Fatigue strength degradation of E-glass FRP composites and carbon FRP composites. Construct. Build. Mater. 1998,12, 311–318. [CrossRef]22. Hollaway, L.C. Key issues in the use of fibre reinforced polymer (FRP) composites in the rehabilitation and retrofitting of concretestructures. In Service Life Estimation and Extension of Civil Engineering Structures; Karbhari, V.M., Lee, L.S., Eds.; WoodheadPublishing: Cambridge, UK, 2011; p. 3.23. Yang, Y.; Fahmy, M.F.M.; Guan, S.; Pan, Z.; Zhan, Y.; Zhao, T. Properties and applications of FRP cable on long-span cable-supported bridges: A review. Compos. Part B 2020, 190, 107934. [CrossRef]24. Yamaguchi, T.; Kato, Y.; Nishimura, T.; Uomoto, T. Creep rupture of FRP rods made of aramid, carbon and glass fibers. InProceedings of the Third International Symposium on Non-Metallic (FRP) Reinforcement for Concrete Structures (FRPRCS-3), V.2, Tokyo, Japan, 14–16 October 1997; Japan Concrete Institute: Tokyo, Japan, 1997; pp. 179–186.25. Saadatmanesh, H.; Tannous, F.E. Relaxation, creep, and fatigue behavior of carbon fiber-reinforced plastic tendons. ACI Mater. J.1999, 96, 143–153.26. Wicaksono, S.; Chai, G.B. A review of advances in fatigue and life prediction of fiber-reinforced composites. J. Mater. Design Appl.2013, 227, 179–195. [CrossRef]27. Admjadi, M.; Fatemi, A. A fatigue damage model for life prediction of injection-molded short glass fiber-reinforced thermoplasticcomposites. Polymers 2021, 13, 2250. [CrossRef] [PubMed]28. Wang, X.; Shi, J.; Wu, G.; Yang, L.; Wu, Z. Effectiveness of basalt FRP tendons for strengthening RC beams through the externalprestressing technique. Eng. Struct. 2015, 101, 34–44. [CrossRef]29. Benmokrane, B.; Elgabbas, F.; Ahmed, E.A.; Cousin, P. Characterization and comparative durability study of glass/vinylester,basalt/vinylester, and basalt/epoxy FRP bars. J. Compos. Construct. 2015, 19, 04015008. [CrossRef]30. Ali, A.H.; Mohamed, H.M.; Benmokrane, B. Bar size effect on long-term durability of sand-coated basalt-FRP composite bars.Compos. Part B 2020, 195, 108059. [CrossRef]31. Ali, A.H.; Mohamed, H.M.; Benmokrane, B.; El Safty, A.; Chaallal, O. Durability performance and long-term prediction models ofsand-coated basalt FRP bars. Compos. Part B 2019, 157, 248–258. [CrossRef]32. Wang, X.; Shi, J.; Liu, J.; Yang, L.; Wu, Z. Creep behavior of basalt fiber reinforced polymer tendons for prestressing application.Mater. Des. 2014, 59, 558–564. [CrossRef]33. Sokairge, H.; Elgabbas, F.; Rashad, A.; Elshafie, H. Long-term creep behavior of basalt fiber reinforced polymer bars. Construct.Build. Mater. 2020, 260, 120437. [CrossRef]34. El Refai, A. Durability and fatigue of basalt fiber-reinforced polymer bars gripped with steel wedge anchors. J. Compos. Construct.2013, 17, 04013006. [CrossRef]35. Colombo, C.; Vergani, L.; Burman, M. Static and fatigue characterization of new basalt fibre reinforced composites. Compos. Struct.2012, 94, 1165–1174. [CrossRef]36. Dorigato, A.; Pegoretti, A. Fatigue resistance of basalt fibers-reinforced laminates. J. Compos. Mater. 2012, 46, 1773–1785. [CrossRef]37. Zhao, X.; Wang, X.; Wu, Z.; Zhu, Z. Fatigue behavior and failure mechanism of basalt FRP composites under long-term cyclicloads. Int. J. Fatigue 2016, 88, 58–67. [CrossRef]38. Zhao, X.; Wang, X.; Wu, Z.; Keller, T.; Vassilopoulos, A.P. Effect of stress ratios on tension-tension fatigue behavior and micro-damage evolution of basalt fiber-reinforced epoxy polymer composites. J. Mater. Sci. 2018, 53, 9545–9556. [CrossRef]39. Zhao, X.; Wang, X.; Wu, Z. Experimental study on effect of resin matrix in basalt fiber reinforced polymer composites under staticand fatigue loading. Construct. Build. Mater. 2020, 242, 118121. [CrossRef]40. Shi, J.; Wang, X.; Wu, Z.; Zhu, Z. Fatigue behavior of basalt fiber-reinforced polymer tendons under a marine environment.Construct. Build. Mater. 2017, 137, 46–54. [CrossRef]41. Van den Oever, M.; Peijs, T. Continuous-glass-fibre-reinforced polypropylene composites, II: Influence of maleic-anhydridemodified polypropylene on fatigue behavior. Compos. Part A 1998, 29, 227–239. [CrossRef]42. Gamstedt, E.K.; Berglund, L.A.; Peijs, T. Fatigue mechanisms in unidirectional glass-fibre-reinforced polypropylene. Compos. Sci.Technol. 1999, 59, 759–768. [CrossRef]43. Bureau, M.N.; Denault, J. Fatigue resistance of continuous glass fiber/polypropylene composites: Temperature dependence.Polym. Compos. 2004, 25, 622–629. [CrossRef]44. Bureau, M.N.; Denault, J. Fatigue resistance of continuous glass fiber/polypropylene composites: Consolidation dependence.Compos. Sci. Technol. 2004, 64, 1785–1794. [CrossRef]45. Ignaczak, W.; Skob, A.L.; El Fray, M. Interfacial polarization in thermoplastic basalt fiber-reinforced composites. Polymers 2020,12, 1486. [CrossRef]46. National Institute for Materials Science (NIMS); Komatsu Matere Co., Ltd. Gripping Device for Fiber Reinforced Plastic,Manufacturing Method for the Same. Japan Patent Office JP 2020-153001, 11 September 2020. (In Japanese).47. Barron, V.; Buggy, M.; McKenna, N.H. Frequency effects on the fatigue behaviour of carbon fibre reinforced polymer laminates. J.Mater. Sci. 2001, 36, 1755–1761. [CrossRef]http://doi.org/10.1016/S0950-0618(98)00012-9http://doi.org/10.1016/j.compositesb.2020.107934http://doi.org/10.1177/1464420712458201http://doi.org/10.3390/polym13142250http://www.ncbi.nlm.nih.gov/pubmed/34301008http://doi.org/10.1016/j.engstruct.2015.06.052http://doi.org/10.1061/(ASCE)CC.1943-5614.0000564http://doi.org/10.1016/j.compositesb.2020.108059http://doi.org/10.1016/j.compositesb.2018.08.065http://doi.org/10.1016/j.matdes.2014.03.009http://doi.org/10.1016/j.conbuildmat.2020.120437http://doi.org/10.1061/(ASCE)CC.1943-5614.0000417http://doi.org/10.1016/j.compstruct.2011.10.007http://doi.org/10.1177/0021998311425620http://doi.org/10.1016/j.ijfatigue.2016.03.004http://doi.org/10.1007/s10853-018-2260-1http://doi.org/10.1016/j.conbuildmat.2020.118121http://doi.org/10.1016/j.conbuildmat.2017.01.063http://doi.org/10.1016/S1359-835X(97)00089-4http://doi.org/10.1016/S0266-3538(98)00119-5http://doi.org/10.1002/pc.20057http://doi.org/10.1016/j.compscitech.2004.01.016http://doi.org/10.3390/polym12071486http://doi.org/10.1023/A:1017576725885Polymers 2021, 13, 3136 13 of 1348. ASTM D3479-02 Standard Test Method for Tension-Tension Fatigue of Polymer Matrix Composite Materials; ASTM International: WestConshohocken, PA, USA, 2002; p. 4.49. Kehrer, L.; Wicht, D.; Wood, J.T.; Bohlke, T. Dynamic mechanical analysis of pure and fiber-reinforced thermoset- andthermoplastic-based polymers and free volume-based viscoelastic modeling. GAMM-Mitteilungen 2018, 41, e201800007. [Cross-Ref]50. Weibull, W. A statistical distribution function of wide applicability. J. Appl. Mech. 1951, 18, 293–297. [CrossRef]51. Yazdani, N.; Filsaime, M.; Islam, S. Accelerated curing of silia-fume concrete. J. Mater. Civil. Eng. 2008, 20, 521. [CrossRef]52. Marguitu, D.B.; Diaconescu, C.I.; Ciocirlan, B.O. Mechanics of materials. In Mechanical Engineer’s Handbook; Marguitu, D.B., Ed.;Elsevier: Amsterdam, The Netherlands, 2001; pp. 119–188.53. Lu, Z.; Feng, B.; Loh, C. Fatigue behavior and mean stress effect of thermoplastic polymers and composites. Fract. Struct. Integ.2018, 12, 150–157.54. Bohm, M.; Glowacka, K. Fatigue life estimation with mean stress effect compensation for lightweight structures—The case ofGLARE 2 composite. Polymers 2020, 12, 251. [CrossRef] [PubMed]55. ACI 440.1R-15 Guide for the Design and Construction of Structural Concrete Reinforced with Fiber-Reinforced Polymer Bars; AmericanConcrete Institute: Farmington Hills, MI, USA, 2015; Volume 24, pp. 11–12.http://doi.org/10.1002/gamm.201800007http://doi.org/10.1002/gamm.201800007http://doi.org/10.1115/1.4010337http://doi.org/10.1061/(ASCE)0899-1561(2008)20:8(521)http://doi.org/10.3390/polym12020251http://www.ncbi.nlm.nih.gov/pubmed/31973174 Introduction  Materials and Methods  Materials and Preparation  Characterization Methods  Results and Discussion  Thermomechanical Properties  Static Properties at Elevated Temperatures  Creep Behavior (R = 1) at Room Temperature  Fatigue Behavior (0 < R < 1) at Room Temperature  Implications for Service Life and Design  Conclusions  References