Escherichia coli inactivation potentiation rates for three antibiotics with different modes of action: ampicillin , ciprofloxacin (inhibits DNA replication), and tetracycline (inhibits protein synthesis). We used concentrations for each antibiotic from 0 to 30× its minimum inhibitory concentration, a single 1-ms electric pulse with amplitude from 0 to 20 kV/cm, and post-pulse pre-dilution incubation either absent (≲1 min) or lasting 60 min, 160 min, or 24 h. Our data show that with incubation, potentiation is significant for all three antibiotics, increases consistently with pulse amplitude, and generally also with antibiotic concentration and incubation time. With incubation, potentiation for ampicillin was rather consistently  superior to both ciprofloxacin and tetracycline: ampicillin was superior to both in 42 of 48 data points, including 7 with significance with respect to both, while at 60- and 160-min incubation, it was superior in 31 of 32 data points, including 6 with significance with respect to both. This suggests that electroporation potentiates wall-targeting antibiotics more than those with intracellular targets, providing motivation for in-depth studies of the relationship between the mode of action of an antibiotic and its potentiation by electroporation. Identification of substances permissible in foods and targeting the cell wall of both Gram-negative and Gram-positive bacteria might provide candidate antibacterials for broad and strong potentiation by electroporation applicable also for food preservation.Antibiotic resistance is a global health threat, and there is ample motivation for development of novel antibacterial approaches combining multiple strategies. Electroporation is among the promising complementary techniques – highly optimizable, effective against a broad range of bacteria, and largely impervious to development of resistance. To date, most studies investigating electroporation as an efficacy potentiator for antibacterials used substances permissible in food industry, and only few used clinical antibiotics, as acceptable applications are largely limited to treatment of wastewaters inherently contaminated with such antibiotics. Moreover, most studies have focused mainly on maximal achievable effect, and less on underlying mechanisms. Here, we compare  Antibiotic resistance is a global health threat associated with increased morbidity, mortality, hospitalization, and healthcare costs . There iElectroporation alone can also cause bacterial death, but the levels of inactivation typically achieved are insufficient for general stand-alone use, providing motivation for development of synergistic treatments . For fooFor any application outside clinical and veterinary medicine, addition of antibiotics is generally problematic due to the involved health risks and/or the resulting environmental burden. However, wastewaters from hospitals and livestock farms are already inherently contaminated with clinical and veterinary antibiotics, respectively , as wellThese problems are increasingly recognized as being critical both for human health and for the environment, and thus approaches to reduce the concentration of viable bacteria, antibiotics, and/or antibiotic resistance genes – preferably all three – in wastewaters are now being proposed and tested. For on-site treatment of wastewaters prior to their release into the communal sewage, various designs have been proposed based on ozonation, ultraviolet light, and/or microalgal biodegradation .With wastewaters already containing various bacteria and antibiotics, methods that render the bacterial envelope permeable, thus facilitating the permeation of antibiotics into the bacteria and potentiating their efficacy, emerge as straightforward candidates for the first “line of attack.” Once the bacterial load is sufficiently reduced, subsequent treatment can focus on degradation/removal of the antibiotics and the bacterial resistance genes. As permeabilization based on additives  would cause additional pollution of the treated wastewater, the acceptable options are limited to green techniques, such as ultrasonication, freeze-thawing, and electroporation. Among these, electroporation is most widely recognized and used due to its general effectiveness, efficiency, and applicability to a broad range of microorganisms . FurtherRhodococcus ruber to five and Escherichia coli to four different antibiotics, at increasing concentrations, and confirmed that such susceptibility increases are generally achievable.To date, the majority of studies that have combined electroporation and antibacterials see  have focAs summarized in Also, where stated in these previous studies, the post-pulse incubation times before diluting out of the antibiotics have been as long as 24 h . ImplemeE. coli inactivation potentiation by electroporation for three antibiotics with different modes of action: ampicillin , ciprofloxacin (inhibits DNA replication), and tetracycline (inhibits protein synthesis). The inactivation rates were investigated at antibiotic concentrations from the MIC to 30-fold the MIC, exposure to a single 1-ms electric pulse with amplitudes from 5 to 20 kV/cm, and a post-pulse pre-dilution incubation from ≲1 min to 24 h.Here, we compared the E. coli K12 ER1821 strain was used , and was propagated according to the protocol of the supplier. The cells were cultured in lysogeny broth rich medium  at 37°C, with agitation. The growth curve was measured from a starting culture with optical density at 600 nm (OD600) of 0.01, and the middle exponential phase was determined at 3.5 h of incubation.To exclude the effect of resistance and virulence genes, the non-resistant and non-pathogenic Three antibiotics with different modes of action were used in this study: (i) ampicillin , which inhibits cell wall synthesis by binding to bacterial penicillin-binding protein transpeptidases, thus preventing them from catalyzing cross-linking of peptidoglycan chains ; (ii) ciE. coli strain used, as the lowest concentration of the antibiotic that inhibited visible growth of the E. coli during the incubation. This is to be distinguished from the minimum bactericidal concentration, MBC, which is higher and is defined as the lowest concentration of an antibiotic that kills at least 99.9% of the bacteria. The standard protocol of agar dilution and overnight incubation was followed for MIC determination  for each antibiotic was determined against the mination , with thE. coli cultures were initiated by inoculation of one colony from lysogeny broth agar plate to 50 mL lysogeny broth, with overnight incubation at 37°C, with agitation. The following morning OD600 was measured, and fresh 250 mL lysogeny broth was inoculated for the starting OD600 of 0.01. The cultures were grown to the middle exponential growth phase, which occurred at 3.5 h at 37°C, with agitation. The cells were then centrifuged, washed with 250 mM sucrose, centrifuged again, and resuspended in 16 mL 250 mM sucrose.Overnight For each antibiotic, the concentrations used were 0 (no antibiotic), MIC, 3 × MIC, 10 × MIC, and 30 × MIC. For the electric pulse, the amplitudes used were 0 (no pulse delivery), 5, 10, 15, and 20 kV/cm. For post-pulse pre-dilution incubation at room temperature with the same antibiotic at the same concentration as in the treatment, the incubation times used were ≲1 min (dilution right after pulse delivery), 60 min, 160 min, and 24 h.For each combination of parameters, the experiments were performed three times on different days, with altered order in which the parameters were varied for each of three repetitions of the experiment. The treatment with no antibiotic and no pulse delivery was considered as the control.A 3-mL volume of prepared washed culture was added to 3 mL of 250 mM sucrose (for control and electroporation-only experiments), or to 3 mL of 250 mM sucrose supplemented with an antibiotic at the final concentration required  as stated above. Electroporation was performed using the HVP-VG square wave pulse generator . The samples (140 μL) were placed between two parallel stainless-steel electrodes with a 1-mm gap, and a single 1-ms rectangular electric pulse was delivered. The voltages applied were 500, 1,000, 1,500, and 2,000 V, for electric pulse amplitude  of 5, 10, 15, and 20 kV/cm, respectively, with the actual time courses of the voltage for each pulse amplitude as measured with LeCroy HDO4104A oscilloscope and HVD3206A voltage probe  shown in E. coli samples were serially diluted in 0.9% NaCl, and for each dilution, three drops of 10 μL were plated on a lysogeny broth agar plate. After the drops dried, the plates were incubated overnight at 37°C. The E. coli counts were recorded for each dilution (colony counts from 3 to 30), and the colony forming units (CFU)/mL were calculated from the mean number of colonies (mean of the three drops). The E. coli inactivation rates were calculated as −log10(N/N0), where N is the E. coli CFU/mL of the sample, and N0 is the E. coli CFU/mL of the control (log10 will henceforth be referred to as log).After the treatment and the post-treatment incubation, the p < 0.05) for each combination of electric pulse amplitude, antibiotic concentration, and post-pulse pre-dilution incubation time. Tukey’s HSD multiple comparison test for evaluation of the difference was used when ANOVA indicated a statistically significant difference (p < 0.05). In The experiments were repeated three times on different days for each antibiotic, and the treatment data were normalized to the control  and expressed as mean ± standard deviation. The data were post-processed in R Commander 2.6 . To compare the effects of the three antibiotics, one-way analysis of variance was used  against E. coli cultures) were not bactericidal even after 24 h of incubation with each of the antibiotics. This was because while the growth of E. coli cells was inhibited throughout the incubation, after the transfer to rich growth medium they recovered , 5, 10, 15, and 20 kV/cm, are presented in The E. coli cells were treated with electroporation alone , and then we focused on the roles of (longer) post-pulse pre-dilution incubation time and of antibiotic concentrations.E. coli inactivation rate was small, particularly for the antibiotic concentrations up to 10 × MIC, and although at 30 × MIC , no significant difference was seen between ampicillin, tetracycline, and ciprofloxacin . With th30 × MIC  ampicillWhen post-pulse pre-dilution incubation time was increased to 60 min, 160 min, and 24 h, relative to the ≲1 min incubation the potentiation also increased for each antibiotic. Specifically, three observed parametric dependences can be inferred.First, the antibiotic potentiation consistently increased with the increase of pulse amplitude . A single exception here was for ciprofloxacin at 24-h post-pulse incubation for the 5 kV/cm pulse, with indication of lower E. coli inactivation for 30 × MIC compared to 10 × MIC, although this did not reach statistical significance .Second, the antibiotic potentiation generally also increased with the increase of antibiotic concentrations at each of these three post-pulse incubation times . Here the exception was tetracycline at MIC and 3 × MIC, with indications of lower E. coli inactivation for 24-h than for 60- or 160-min post-pulse incubations; however, again, none of these reached statistical significance .Third, the antibiotic potentiation generally also increased with the increase of the post-pulse incubation time at each of the four antibiotic concentrations  bacterial membrane is not required, in contrast to both ciprofloxacin and tetracycline that have intracellular targets , for access to which they must permeate through all the layers of the bacterial envelope.E. coli inactivation rates improved by an order of magnitude when the dilution of the antibiotic was delayed by 60 or 160 min, while the 24-h delay resulted in more sporadic further improvements and mostly at the highest pulse amplitudes. This suggests that shortening the post-pulse incubation time with the antibiotic from 24 h (as used in many previous studies) to one or several hours is feasible, with proportionally reduced reservoir volume required in applications for wastewater treatment.The data for dilution right after pulse delivery  imply, however, that even for ampicillin, substantial efficacy potentiation requires time, with The small and sporadic further improvements in inactivation rates with the 24-h incubation are most likely due to the physiological uptake of the antibiotics, which would occur even without electroporation by gradual permeation through the intact bacterial envelope. This is reflected in the small but rather consistent improvement of the inactivation rates for all three antibiotics at 0 kV/cm  at 24-h incubation compared to 60- or 160-min incubation, which is detectable also for MIC and 3 × MIC, but is more evident for 10 × MIC and even more so for 30 × MIC. The role of physiological permeation on longer time scales is also consistent with the empirical fact that in medical and veterinary therapies with an antibiotic alone, its concentration must be maintained at a suprainhibitory level for days, and for some infections even for weeks to achieve an effective outcome .E. coli inactivation rates observed here for tetracycline compared to both ampicillin and ciprofloxacin.Regarding the generally weakest potentiation for tetracycline, we note that of the three antibiotics used in our study, tetracycline has the highest molecular weight  and therefore likely requires stronger and/or more extensive electroporation for similarly potentiated permeation into bacteria. However, we stop short of postulating this as the main reason for the relatively inferior E. coli can be potentiated by electroporation to a greater extent – and/or more readily – compared to two antibiotics that target intracellular sites, is also in empirical agreement with findings from a recent study of E. coli (used in the present study) is Gram-negative, while Rhodococcus ruber (used by Kuyukina and colleagues) is Gram-positive, and thus the structure of their envelope differs significantly.Our finding that for a wall-targeting antibiotic, the efficacy against However, there are still some obvious and possibly other unforeseen obstacles for the application of this finding in practice. Adding antibiotics is universally problematic in terms of the resulting environmental burden, and in many applications also from the resulting risks to human health. Conversely, the applications utilizing the inherently present antibiotics, such as treatment of wastewaters from hospitals and livestock farms, are dependent on the persistently fluctuating compositions and concentrations of antibiotics, which are also generally well below their MICs . Thus, aFor use in clinical or veterinary applications, and in general for combining antibiotics with electroporation against bacterial infections of eukaryotic organisms, a major and perhaps largely unsurmountable obstacle lies in the fact that most eukaryotic cells are an order of magnitude larger than bacteria. Since the transmembrane voltage induced by exposing a cell to an electric pulse of a fixed amplitude  is proportional to the cell size , a signiE. coli focus on those among the permissible substances that target the wall. Currently, one such substance widely recognized as targeting the wall is nisin , and theE. coli . Howeverhimurium , while thimurium  and can himurium  or by bihimurium . There iFor the understanding of the dependence of the antibiotic efficacy potentiation by electroporation on the antibiotic’s target site, our results presented above suggest that for antibiotics targeting the bacterial cell wall, this potentiation can be higher than for antibiotics with intracellular targets. For broader testing and a deeper understanding of this thesis, further studies are needed, performed with a broader range of antibiotics and on a broader range of bacteria, including comparisons for antibiotic-sensitive vs. antibiotic-resistant strains, and for bacteria in different growth stages. Identification of substances permissible in foods and targeting the cell wall of both Gram negative and Gram positive bacteria might provide candidate antibacterials for broad and strong potentiation by electroporation applicable also in food industry.The original contributions presented in the study are included in the article/ŽL set up the experimental plan, conducted the experiments, processed, analyzed, interpreted the results, and wrote part of the manuscript. TK set up the concept of the experiments, supervised the work, interpreted the results, and wrote part of the manuscript. AK wrote part of the manuscript. All authors have reviewed and approved the final manuscript.The authors declare that the research was conducted in the absence of any commercial or financial relationships that could be construed as a potential conflict of interest.All claims expressed in this article are solely those of the authors and do not necessarily represent those of their affiliated organizations, or those of the publisher, the editors and the reviewers. Any product that may be evaluated in this article, or claim that may be made by its manufacturer, is not guaranteed or endorsed by the publisher.


The three tested series consisted of three beams with dimensions of 120 × 300 × 4150 mm, with various numbers of stirrups and contents of fiber reinforcement. During the tests, the shear capacity of the elements was determined. The values of support reactions, deflection in the middle of the span of both beam spans, deformations on the surface of the concrete member in the middle of the span in the compressive and tensile zone, and cracking (crack development and crack width) were also measured. The beams were tested using a digital image correlation (DIC) technique. Test results show that shear capacity increases in beams made of concrete with steel (1.87) or basalt fibers (1.23). Moreover, the failure mode changes from shear (brittle) to flexure-shear (less brittle). The experimental shear capacity of beams was compared with the theoretical values predicted by different design codes, i.e., fib Model Code 2010 and RILEM TC 162-TDF 2003. The results show that all the design codes underestimate the contribution of fiber-reinforced concrete beams to shear resistance and greatly overestimate the contribution of shear reinforcement.This study investigates the effects of adding different types of fibers to concrete mixes on the shear behavior of double-span fiber-reinforced concrete beams with or without shear reinforcement. As a part of the experimental study, a total of twenty-seven natural-scale double-span beams were tested. The beams, made of concrete with steel or basalt fiber, with fiber dosages of 78.5 and 5 kg/m Double-span reinforced concrete beams are the most frequently used structural members. They are used in elements with large spans or repetitive structures such as slab beams, roof beams, and bridge structures . They arIn order to understand the shear behavior of beams, it is first important to understand the transfer mechanism of shear forces. Shear failure of reinforced concrete members results from a complicated mechanism related to the size and type of load, dimensions, the geometric shape of the cross-section, and the material properties of concrete and reinforcing steel. A detailed explanation of the theoretical foundations of the shear phenomenon is given in . The basHowever, these mechanisms only apply to normal concrete and are not relevant to fiber-reinforced concrete (FRC), see Matrix fracture and matrix spalling (or fragmentation);Fiber-matrix interface debonding;Post-debonding friction between fiber and matrix (fibre pull-out);Fiber fracture;Fiber abrasion and ductile deformation (or yielding) of fibre.This phenomenon was described by Löfgren . The pro3).The addition of fibers is one of the methods to alleviate the problem of brittleness in concrete. The principal action responsible for the transfer of shear stresses across a crack in plain concrete is aggregate interlock and friction at the crack faces . For fibThe use of steel fibers is a well-known technique ,7,19. FuExtensive work related to the effectiveness of using fibers as a replacement for stirrups concerns steel fibers ,23,24. NThe effect of different fiber materials  on shear cracking behavior.The contribution of steel or basalt fibers to shear resistance, ductility, and changes in the failure mode of reinforced concrete beams.The research presents tests of double-span beams made of innovative concrete with basalt fibers in comparison to beams made of concrete with steel fibers. In this study, the effect of using steel or basalt fibers on the overall shear response and its impact on the ultimate failure mode were examined using the DIC technique. The Digital Image Correlation (DIC) technique is spreading great interest in the research community because of its low costs, effectiveness, and capacity to monitor displacements and strains during experimental tests ,27. Henc3 and the w/c ratio was 0.5. The aggregate used was a mixture of sand with a grain diameter of up to 2 mm (51% of aggregate) and coarse natural aggregate with a grain diameter of up to 8 mm. Steel and basalt fibers were used as reinforcement in concrete. The steel fibers were the hook-shaped type, 50 mm long with a diameter of 1 mm, tensile strength of 800 MPa, and Young’s modulus of 210 GPa; the diameter of basalt fibers was 20 µm, tensile strength 1680 MPa, and Young’s modulus 90 GPa; these were used to modify the concrete mix. The fiber content in the concrete was 78.5 kg/m3 for steel fibers and 5.0 kg/m3 for basalt fibers. The fibers were added by volume as a replacement for part of the coarse aggregate. Tests of concrete properties were performed on three series of reference samples and a series of concrete with steel and basalt fibers.The concrete mixtures were made with CEM I 42,5R Portland cement. The cement content was 300 kg/mckf of concrete were performed pursuant to EN 12390-3:2011 [ctmf on samples with dimensions of 100 × 100 × 400 mm, pursuant to EN 12390-5:2011 [cmE was determined pursuant to EN 12390-13:2014 [Tests of compressive strength 0-3:2011  using cu0-5:2011 ; modulus-13:2014 , using cThe results of tests of the mechanical properties of concrete are presented in The increase in bending tensile strength was 40% for steel fibers introduced into the concrete and 30% for basalt fibers. In the case of concrete specimens with steel fibers, the brittle nature of concrete changed to quasi-ductile. Concrete with basalt fibers exhibited brittle cracking, similarly to plain concrete without the addition of fibers. However, the destructive force was much higher for samples made of concrete with basalt fibers than for plain concrete.Series C-I—shear reinforcement with stirrups; maximum spacing determined pursuant to EN 1992-1-1  is 200 oSeries C-II—stirrups with spacing twice as large and the volume is 150 or 300 mm; pursuant to EN 1992-1-1  Figure c;Series C-III—without stirrups d.Twenty-seven natural-scale double-span beams were tested. Three series of test elements were used, with different types of shear reinforcement. Each series consisted of 3 beams with a dimension of 120 × 300 × 4150 mm each with different numbers of stirrups and three different fiber reinforcement dosages. The series were:3, and WB5 with basalt fibers in the amount of 5 kg/m3.In each beam series, the proportion of longitudinal reinforcement was effl = 2000 mm. Shear rate v/da of the beams was 3.5.The beams’ support points were located at a distance of 75 mm from the outer edge of the beams, which resulted in span axes of During the tests, the shear and/or bending capacity of the element was determined. The following values were also measured: support reactions, deflection in the middle of the span of both beam spans, deformations on the surface of the concrete member in the middle of the span in the compressive and tensile zone, and cracking (crack development and crack width).The beams were preloaded to a force of 30 kN. The time of force buildup was 30 s per 10 kN, and the load was subsequently held constant for 30 s to take the readings of the test values. Then the increment of the force was 10 kN, until failure. Load was applied with a hydraulic cylinder with a capacity of 500 kN, controlled from the control panel of the PZA machine, through a steel traverse and two rollers with a pad directly on the element. The distance between the forces was 2000 mm. The arrangement of the measuring devices on the beam is shown in Measurement of the support reactions was performed using Wobit-KMM50 strain gauge force sensors with the following ranges: 0-50 kN and 0-100 kN, with a constant 1.5 ± 2% mV/V. They were placed under the supports in pairs: 2 × 50 kN on the extreme supports and 2 × 100 kN on the central support. The value of force was recorded using the ADAE42U Wobit measuring module connected to force sensors.The deflection measurements were performed using inductive sensors from the so-called Megatron-Munchen return spring with measuring ranges of 25, 50, and 75 mm, and an accuracy of up to 0.001 mm. The measurements were recorded continuously using the KSR-32 Sensor diagnostic recorder with a sampling frequency of 2 s. The places where the deflections were measured are marked in Vision measurement systems used in many industries such as: automotive, automotive design, aviation industry, Bbomechanics/medical, sheet metal forming, casting and foundry and materials testing and simulation. The ARAMIS optical measuring system is designed for non-contact measurements of displacements in planar and spatial elements subjected to loading. The configuration used for the tests, i.e., ARAMIS 2M, is a system consisting of two cameras  with a fThe accuracy of the reading of position on the real object depends on the measurement range and the ratio of the actual size of the examined area to the size of the analyzed image on the matrix.The measurement set enabled the recording of up to 5 million readings at a frequency of 15 Hz. The measurement procedure consisted of initial calibration of cameras for a given working area, applying a pattern in the form of black dots with the use of a dedicated preparation, and finally recording the displacement of the measurement points over time, i.e., the so-called facets. The size of the working area was 1000 × 840 mm.The process of bending of each beam was recorded so that two synchronous photographs of the sample were taken every 1 s, together with the value of the corresponding force. After the completion of the study, the images taken were analyzed using the ARAMIS program, which makes it possible to present the results of the analyses in different manners.The analyses were performed on the basis of the DIC measurements obtained from one side of the specimen. However, it should be noted that the profile of the crack is not uniform throughout the width of the member.The maximum measured values of loads, moment capacities, and mid-span deflections; first crack loads, curvature ductility, and failure mode for all the tested beams are compiled in The effect of fiber reinforcement on shear capacity can be determined by comparing the values of destructive forces in each of the series . C-I serThe results of measurements of deflections of reinforced concrete beams are presented in Adding basalt fibers to concrete caused slight deflections of beams in comparison to C-I series reference elements, whereas adding steel fibers to concrete reduced deflection by as much as a half for the same level of load. Deflections for C-II-WB5 and C-II-WS78.5 series beams were reduced by 30–50% in comparison to the C-II-W0 series. In the case of the C-III series , the maximum values of deflection of beams made of concrete with steel fibers were 30% lower, while the maximum deflections of beams made of concrete with basalt beams were 20% lower in comparison to deflections of reference beams.During the test, development of cracks in of all the tested beams were monitored as shown in ultP) in the C-I series, at 90 kN (0.4 ultP) in C-I-WB5 series, and at 86.7 (0.3 ultP) in C-I-WS78.5 series . Later, the beams failed due to clear separation of localized cracks. The number of perpendicular and diagonal cracks as well as the inclination angle of the diagonal crack observed during the test in reference beams were the highest.As expected, failure of the C-I series occurred due to bending. In addition, the presence of steel fibers in concrete resulted in the brittle nature of beams changing to quasi-ductile. First cracks in the beams in this series appeared on the central support at a load of 76.7 kN , in addition to the loosening of the side and upper covers of the reinforcement. Just before failure, diagonal cracks were inclined at an angle of 42° on both sides, while their spacing ranged from 151 to 245 mm. In the C-II-WS78.5 series, failure occurred as a result of bending at the supports. The process of failure was smooth, with significant increase in the deflections in the middle of the spans of the beams at a constant load. During bending, diagonal cracks on the central support reached great widths, with spacings of 151 and 139 mm between them. They did not, however, cause beam failure. In the final stage just before failure, longitudinal cracks appeared along the upper cover. The first crack on the central support was not as destructive in the case of beams made of concrete with basalt or steel fibers in the C-II series.In the C-II series, significant differences in the manner of loss of capacity of the individual elements were observed . Referen support . On thesIn the C-I and C-II series, large deformations were observed stretching along the upper cover, connected with the gradual loss of adhesion of the longitudinal reinforcement to concrete. Just prior to failure, the width of cracks was much larger in beams reinforced with steel fibers, while their numbers were smaller than in the case of reference beams, or beams reinforced with basalt fibers. In fiber concrete beams C-II-WB5 and C-II-WS78.5, unlike in the case of C-II-W0 series beams made of reference concrete, local loss of adhesiveness between the cover and the surrounding concrete did not occur. This resulted from the presence of basalt or steel fibers, which transfer tensile stresses.l from the central support) appear. The first cracks appeared on the central support at the following loads: 0.4 ultP—C-III-W0, 0.3 ultP—C-III-W78.5, and 0.4 ultP—C-III-WB5  to less brittle (shear tension failure) with higher displacements at failure.crV and destructive force ultV. The values shown are for the central support.ult/VcrV ratio on the central support comparable with C-I-W0 series can be obtained.The ratio between the diagonal force causing the cracks and the destructive diagonal force turned out to be the highest in the case of beams made referencing concrete C-II-W0, concrete with basalt fibers C-I-WB5, and concrete with steel fibers C-II-WS78.5 at the central support. Diagonal cracks at the central support appeared last in the following series: C-I-WS78.5 (made of concrete with steel fibers), C-II-WB5 , and C-III-W0 (reference concrete). In the C-II series, with the use of concrete with basalt or steel fibers, diagonal cracks appeared later than in beams made of reference concrete. This points to the effectiveness of using fiber-reinforced concrete to delay the appearance of diagonal cracks. In the C-I and C-III series, at the central support, diagonal cracks appeared faster than in reference beams; this occurred, however, at a higher value of diagonal force. In the C-III series diagonal cracks appeared at the lowest load, at the first stage of beam operation. By using concrete with basalt (C-II-WB5) or steel (C-II-WS78.5) fibers, a The values of shear stresses were determined at three supports for each of the beams. For the C-I series, an increase in shear stresses at the right support was observed. In the case of the left and central support, the stress values were constant. Beams in this series failed as a result of bending in one of the spans, while the other span (less damaged) was still able to transfer loads to the support. In the C-II and C-III series, where beam failure was as a result of shear, the main shear stresses were transferred to the central support. Owing to the use of concrete with basalt fibers, the increase in shear stresses at the central support was 20 and 30% for the C-II and C-III series, respectively.Regarding shear stresses at the central support, it was found out that in the C-I series, the value of shear stresses increased by 24% owing to the used fiber concrete. Stresses at the extreme supports were considerably reduced, due to the appearance of ductile hinges on the spans, which is related to the fact that the beam started to operate only on the part of the central support with two brackets. In the C-II series, the value of shear stresses increased by 10% due to the redistribution of internal forces and the appearance of a ductile hinge in the C-IIWS78.5 series. In the C-III series, stress values were reduced by 35%. When comparing the values of shear stresses and the course of the support reactions in the C-IIIWS78.5 series, the phenomenon can be explained with redistribution of external forces and the appearance of a ductile hinge in C-IIIWS78.5 support on an external support.In the C-II series, where failure was due to shear, the main shear stresses were transferred to the central support. The increase in shear stresses on the central support in the C-II series, owing to the use of concrete with steel fibers, was 17%. In the C-III series, a several-fold increase in shear stresses on the external supports was observed. This resulted from transfer (redistribution) of loads to the external supports, thus relieving the central support.When comparing the values of concrete deformations at the same levels of beam loads, it was concluded that in the C-I series, at the tensile zone, the highest deformations were recorded in the case of beams made of concrete with the addition of basalt fibers during all stages of operation. The difference in deformations reached 56% in comparison to beams from concrete without fibers. The most significant deformations of the compressive zone occurred in the case of reference beams. In the stage before failure, the values were the highest in the compressive and tensile zones in beams made of concrete with basalt fibers.C-II series reference beams and beams made of basalt fibers were characterized by similar increases in values of deformations both in the compressive and the tensile zone. At a load of 180 kN, beams made of concrete with steel fibers reached the highest values of deformations in the compressive zone (30% and more) and the lowest deformations in the tensile zone (lower by 30%) in comparison to beams made of concrete without fibers.A similar distribution of deformations can be observed in the C-III series, where reference beams showed the highest values of deformations of concrete in the compressive zone. In the tensile zone, the greatest deformations, at the same level of load, were recorded on the surface of beams made of concrete with steel fibers—approx. 20% higher than for reference beams.fib Model Code 2010 [In order to assess the effect of the presence of fibers on shear capacity, calculations of the load-bearing capacity of beams with longitudinal and fiber reinforcement in the form of basalt or steel fibers were performed, pursuant to ode 2010  and RILEode 2010 . The theode 2010 .The calculation procedures, taking into account the presence of fibers in concrete, relate to concretes with steel fibers. Due to the lack of guidelines for including basalt fibers in capacity calculations, the calculations employed the same guidelines as that of steel fibers.The basic problem in the case of concrete with basalt fibers is determining the residual tensile strength at bending. Most methods require a crack mouth opening displacement (CMOD) of 3.5 mm. Concrete with basalt fibers made it possible to obtain deformations corresponding to CMOD values of up to 0.8 mm.RdV) is the sum of three components: capacity of concrete member (cdV), effect of stirrups (wdV), and effect of fibers (fdV) [The RILEM TC 162–TDF method  assumes rs (Vfd) .fV is calculated according to the following Equation:Shear capacity that takes into account the addition of fibers fR,4, whose value for CMOD is 3.5 mm, which takes into account residual tensile stresses at bending, determined during testing in conditions of three-point bending of a notched beam.The decisive value is residual strength The shear capacity of beams tested in this study is reported in σ1 cannot be lower than the calculated tensile strength of fiber concrete according to:Ftuf is the characteristic value of residual tensile strength of fiber concrete determined for uw = 1.5 mm.When tensile fiber reinforcement is planned and members without longitudinal or diagonal reinforcement are considered, main tensile stress Equation (3) is applied to steel fiber concrete. The Equation has not been tested for fibers made of other materials and concretes diverging from ordinary ones, such as those made of RPC reactive powders, among others.Rd,FV is assumed to be RdFminV or higher, determined as:Shear capacity The model recommends the following method of calculation of shear capacity:Shear capacity of beams with longitudinal or diagonal reinforcement should be calculated according to the following Equation:The shear capacity of the beams tested in this study is reported in fib Model Code 2010 (MCV) and RILEM TC 162-TDF 2003 (RILEMV) in relation to the value of diagonal force determined experimentally ExpV.C-I series beams failed due to bending, which is why the calculated capacity values were higher than the diagonal force determined experimentally.C-II series beams failed mainly as a result of bending, then due to shear at the central support. The radio between the calculated and the experimentally determined shear capacity at the central support should thus be considered. All calculation procedures correctly forecast the values of diagonal force. In the case of beams with longitudinal or diagonal reinforcement, and/or fibers, shear capacity mainly depends on the shear capacity component as a result of the diagonal reinforcement. The addition of steel fibers increases the value of shear capacity by 18%.fib Model Code 2010 and 3–23% for RILEM TC 162-TDF 2003. In the case of the C-III-WB series, regardless of the calculation method, a reserve of capacity of up to 20% was obtained. In the case of the C-III-WS series, the method that reflected the experimental value of the destructive force was the fib Model Code 2010 procedure.In series C-III, failure occurred at the central support. The calculated values of capacity, regardless of the estimation method, were lower than those obtained experimentally, despite the fact that the safety coefficient was not included in the performed calculations. This means that each of the methods assume a reserve of capacity. The values of forces determined experimentally were higher, depending on the calculation method, by 11–43% for fib Model Code 2010 procedure [rocedure  containsrocedure , the proOwing to the presence of steel fibers, regardless of the calculation method, the component of shear capacity was higher in comparison to concrete with basalt fibers.3 and basalt fibers in the amount of 5 kg/m3 on the improvement of the load-bearing capacity of double-span beams with a length of L = 4.15 m. The effect of the fibers was assessed in comparison to test results for concrete members without fiber reinforcement. Based on the experimental study, the following major conclusions can be drawn:L = 4.15 m increased in beams made of concrete with steel or basalt fibers.The shear capacity of double-span beams The shear stresses at the center support were greater compared to the volume of stresses at the extreme supports. The volume of shear stress depends on the degree of shear reinforcement . The presence of fiber reinforcement in the concrete also increased shear stresses.Beams in all series made of concrete with steel or basalt fibers were characterized by lower deflections compared to reference beams at the same load level. Due to the use of steel fibers in concrete, the values of deflection in all series of beams were three times smaller, while in the case of concrete with basalt fibers in C-III-WB5 beams, they were twice as small compared to beams made of concrete without fibers.Change in the angle of diagonal cracking indicates that the addition of steel or basalt fibers changes the failure mode of reinforced concrete beams from shear (brittle) to less brittle (shear tension failure) with higher displacements at failure.Digital Image Correlation (DIC) analysis of the crack pattern around the central support confirmed the influence of both types of fibers on the crack pattern and shear behavior of beams with reduced reinforcement for shear.fib-MC2010 and RILEM guidelines. RILEM guidelines resulted in closer predictions of test results. The fib-MC2010 recommendations gave a more conservative estimate of the capacities than RILEM provisions.The shear capacity of SFRC (steel fiber-reinforced concrete) beams tested in this study, consistent with literature data, were estimated using In times of rising steel prices they are sought to replace it. The test results confirm the possibility of complete or partial replacement of steel stirrups by concrete with basalt fibers.The study aimed to demonstrate the effect of steel fibers in the amount of 78.5 kg/m


Fiber-reinforced polymer (FRP) bars are increasingly used as a substitute for steel reinforcements in the construction of concrete structures, mainly due to their excellent durability characteristics. When FRP bar-reinforced concrete (referred to as FRP-RC for simplicity) members are used in indoor applications , the fire performance of FRP-RC members needs to be appropriately designed to satisfy safety requirements. The bond behavior between the FRP bar and the surrounding concrete governs the composite action between the two materials and the related structural performance of the FRP-RC flexural member that will be affected when exposed to fire. However, there is a lack of reliable numerical models in the literature to quantify the effect of bond degradations of the FRP bar-to-concrete interface at high temperatures on the fire performance of FRP-RC flexural members. This paper presents a three-dimensional (3D) finite element (FE) model of FRP-RC flexural members exposed to fire and appropriately considers the temperature-dependent bond degradations of the FRP bar-to-concrete interface at high temperatures. In addition, the thermal properties of concrete and FRP bars are considered in the heat transfer analysis to predict the cross-sectional temperatures of the FRP-RC members under fire exposure. In the FE model, the mechanical properties and constitutive laws of concrete and FRP bars at high temperatures in addition to the bond degradations between them have been properly defined, thereby accurately predicting the global and local structural responses of the FRP-RC members under fire exposure. The proposed FE model has been validated by comparing the FE predictions (both temperature and midspan deflection responses during fire exposure) and the full-scale fire test results reported in the literature. The validated FE model is then used to study the effects of bond degradations on the global and local structural responses of the FRP-RC members under fire exposure. It is proved that the temperature-dependent bond degradations need to be considered to achieve accurate predictions of the failure mode and deflection responses. Fiber-reinforced polymer (FRP) bars are increasingly used as a substitute for steel reinforcements in the construction of concrete structures, mainly due to their excellent durability properties. In the literature, a large number of studies have been conducted on the performance of FRP bar-reinforced concrete (referred to as FRP-RC for simplicity) members at ambient temperature ,2,3,4, aExisting studies in the literature mainly focus on the mechanical and bond properties of externally bonded FRP laminates at high temperatures ,26,27,28It is noteworthy that under fire exposure, high thermal stresses may be generated at the interface between the FRP bar and the surrounding concrete cover, mainly due to the different thermal expansion coefficients between FRP and concrete. When the thermal stress increases to the tensile strength of the concrete cover during fire exposure, it will be partially cracked, resulting in the weakening of the bond between FRP and concrete ,54. MoreGiven the above research background, this paper has proposed a three-dimensional (3D) FE model to study the effect of bond degradations of the FRP bar-to-concrete interface at high temperatures on the fire performance of FRP-RC flexural members. The FE model has appropriately considered the constitutive laws of FRP and concrete as well as their bond interface at high temperatures, and therefore, it is expected to accurately predict the thermal and structural responses of FRP-RC flexural members under fire exposure. Then, the proposed FE model is verified by comparing the FE results with the test data of the full-scale fire tests in the literature. The validated FE model is further used to study the effect of the interfacial bond degradations at high temperatures on the fire performance of FRP-RC flexural members. The results have indicated that the consideration of the bond degradations of the FRP bar-to-concrete interface can capture the pull-out failure mode of the FRP bars and yields more accurate predictions of the deflection responses of FRP-RC flexural members under fire exposure.The modeling procedure of the FRP-RC members in fire includes two main steps: heat transfer analysis and mechanical analysis. Heat transfer analysis is used to determine the cross-sectional temperature distributions of the FRP-RC flexural member under fire exposure, while the mechanical analysis takes the cross-sectional temperature distributions as the initial state to determine the structural responses. In other words, the modeling procedure follows a sequential thermo-mechanical analysis in which it is assumed that the results of the mechanical analysis have no effect on the heat transfer analysis of the same FRP-RC member under fire exposure. This sequential thermo-mechanical analysis was also successfully adopted by the corresponding author for predicting the fire performance of RC beams and insulated FRP-strengthened RC beams ,60,61. T3, which is usually assumed to be constant at high temperatures [The thermal properties of concrete have been extensively studied in the literature, including thermal conductivity, specific heat capacity, and density at high temperatures. The temperature-dependent changes of these parameters are provided in current design guidelines, such as EN 1992-1-2 . It proveratures ,61.Limit information is available on the thermal properties of FRP composites at high temperatures. In the heat transfer analysis, the temperature-dependent thermal properties  of FRP bars are determined based on the experimental results reported by Griffis et al. . Such th2·K) is used for the surfaces of the FRP-RC member exposed to fire, and 2·K) for the unexposed surface [−8 W/(m2·K4).The heat exchange between the hot fluxes of fire and the surfaces of the FRP-RC member is realized through heat convection and heat radiation, which is controlled by the Robin boundary conditions as follows: surface ,65. εm r1991-1-2 . Tf is tt = 0 is described as:The heat transfer within the FRP-RC member is realized by heat conduction, which is described by the following equation:The concrete damaged plasticity (CDP) model provided in ABAQUS 6.14  softwareo et al. . Interesψ is the dilation angle and ϵ is the flow potential eccentricity. In this study, ψ = 36° is determined based on an initial trial-and-error analysis using the proposed FE model to obtain excellent consistency with the fire test results. ϵ = 0.1 is determined as the default value suggested by the ABAQUS user manner [In the above equation, r manner . In addir manner .The hardening/softening rule controls the subsequent evolution of the yield surface and failure mode of concrete at ambient and high temperatures, which is related to the uniaxial compressive and tensile stress–strain curves of concrete. Under uniaxial compression at ambient and high temperatures, concrete behavior is defined according to the stress–strain relationship specified in EN 1992-1-2 . It is wSome tensile tests were carried out on FRP bars at ambient and high temperatures, and the results showed that the mechanical properties such as the tensile strength and elastic modulus of FRP bars decreased at moderately high temperatures ,69,70. TThe bond interaction between the FRP bar and the surrounding concrete governs the load transfer between the two materials, which is usually described by the local bond stress–slip relationship at the FRP bar-to-concrete interface. Due to the bond action, the FRP bars located between two adjacent concrete cracks have a stiffness contribution to the tensile behavior of the concrete between the two cracks. This phenomenon is termed the “tension-stiffening” effect ,73, and At high temperatures, some pull-out tests on FRP bars have indicated that the decreases in the bond strength are more severe than the corresponding reductions in the tensile strengths of the same FRP bars ,12,58. Ifpm,0). That is, Equations (11) and (13) are used for the cases of 500 ≤ fpm,0 ≤ 1000 MPa, while Equations (12) and (14) are adopted for the cases of 1000 < fpm,0 ≤ 1500 MPa. The parameter In above equations, et al.’s  model. and the two-node heat transfer link element (DC1D2) provided by the ABAQUS software were used to model the concrete section and the GFRP bars, respectively. In other words, the loading plates and the supporting plates were not modeled in the heat transfer analysis. For the mechanical analysis, the eight-node solid element with reduced integration and hourglass control (C3D8R) was adopted to model the concrete section, the loading plates and the supporting plates, while the two-node linear 3D beam element (B33) was used in modeling the GFRP rebars. The nonlinear spring element was used to simulate the bond–slip behavior along the two tangential directions of the FRP bar-to-concrete interface in the mechanical analysis. For the normal behavior of the spring elements, the interaction between the GFRP bar and the concrete cover was defined by setting a large stiffness almost equal to the elastic stiffness of the concrete. The mesh size of all the solid elements was determined as 50 × 50 × 25 mm based on the convergence study, and the corresponding element size was adopted for the beam and spring elements. It should be noted that the configurations of all the nodes and elements between the heat transfer analysis and the mechanical analysis are the same. Therefore, the results of the cross-sectional temperature distributions generated by the heat transfer analysis can be used as an initial condition that can be properly applied to the corresponding time step of the mechanical analysis.The loading plates were tied to the top surface of the flexural member, and the service loads acting on the member during the fire exposure were applied to the reference points on the loading plates. The supporting plates were attached (tied) to the bottom of the flexural member to ensure that there was no slip at the interface between the supporting plates and the concrete surface during the loading and fire process. The pinned boundary conditions were assigned at the reference points at the bottom of the supporting plates during the entire fire test.−6/°C and −6/°C, which are determined according the suggestion provided by Hajiloo et al. [Two series of fire tests conducted by Hajiloo et al.  and Nigry Aslani , as exply Aslani  conducteut tests  are direy Aslani . It is wresearch ,78 if tharch db.  and Nigrarch db. , respecto et al. .Hajiloo et al.  conducteThe advantage of the proposed FE model is that it can reflect the local behavior of the GFRP bar-to-concrete interface at high temperatures, which is usually not easily measured with instruments during the fire test. Such local behavior is beneficial for explaining the anchorage failure mode of GFRP bars in the tested slab. Nigro et al.  carried The midspan deflection responses of the slabs S1, S2, and S3 measured during the fire tests are compared with the FE predictions in For slab S2, the consideration of temperature-dependent bond degradations has a significant impact on the deflection responses, mainly due to the larger applied load of S2 compared with S1. In the fire test, it was observed that two main cracks appeared at the bottom of slab S2, located below the two loading points. The FE predictions of the interfacial slips in (1)The proposed FE model has good accuracy in predicting the thermal and structural responses of FRP-RC flexural members under fire exposure. The constitutive models used to define the thermal and mechanical properties of concrete and FRP bars at high temperatures are reliable. The temperature predictions of the proposed FE model are very similar to the measured results, with a maximum deviation of about 10% during the entire fire test.(2)The perfect bond consideration of the FRP bar-to-concrete interface at high temperatures often yields less accurate predictions of midspan deflection responses. Thus, proper consideration of the local bond–slip behavior of the FRP bar-to-concrete interface at high temperatures is required for accurately predicting the midspan deflection responses of FRP-RC flexural members exposed to fire.(3)The consideration of the local bond–slip behavior also gives detailed information on the interfacial slip responses at the FRP bar-to-concrete interface at high temperatures during a fire, which can reveal the concrete cracking pattern and failure mode of the tested member under fire exposure. The predicted maximum interfacial slips of the FRP bars near the end of the fire test are almost 20 mm.(4)The proposed FE model can accurately predict the anchorage failure of FRP bars in the FRP-RC flexural member during fire exposure, which is a typical failure mode in the existing fire tests in the literature. The previous numerical studies based on a perfect bond consideration cannot provide a reliable prediction for this failure mode.This paper has proposed a 3D FE model to predict the fire performance of FRP-RC flexural members with appropriate considerations of the constitutive models of FRP bars, concrete, and their bond interface at high temperatures. The latter issue has not been accurately simulated by the previous numerical studies. The proposed FE model has been validated by comparing the FE results with the test data of the full-scale fire tests of FRP-RC slabs in the literature. The following conclusions can be drawn based on the results presented in the paper:Further research is needed to develop more reliable constitutive models of FRP bars and bond interface at high temperatures, which are necessary for more accurate predictions of the deflection responses of the FRP-RC flexural members under fire exposure.


Studies on allogeneic demineralized dentin matrix (Allo-DDM) implantation in the 1960s and 1970s provided the most reliable preclinical evidence of bone formation and antigenicity in an extraosseous site. Recently, applications of Allo-DDM at skeletal sites were studied, and have provided reliable evidence of bone-forming capacity and negligible antigenicity. However, the osteoinductivity and antigenicity properties of Allo-DDM in extraskeletal sites have not yet been investigated due to the lack of follow-up studies after the initial research. The clinical applications of autogenous DDM (Auto-DDM) have been standardized in some countries. Long-term clinical studies have reported the development of several shapes of Auto-DDM, such as powders, blocks, moldable forms, and composites, with recombinant human bone morphogenetic protein-2. For the development of Allo-DDM as a reliable bone graft substitute next to Auto-DDM, we reviewed preclinical studies on the bone induction capacity of allogeneic dentin at extraskeletal as well as skeletal sites. Electronic databases were screened for this review in January 2020 and searched from 1960 to 2019. This review aims to provide a foundation on the preclinical studies of Allo-DDM, which could enable future researches on its osteogenic capability and antigenicity. In conclusion, Allo-DDM showed great potential for osteoinductivity in extraskeletal sites with low antigenicity, which neither adversely affected osteogenic capability nor provoked immunologic reactions. However, the risk of viral disease transmission should be researched before the clinical application of Allo-DDM. Dentin is a cell-free matrix without vascularization, while bone includes osteocytes and vessels. The organic and inorganic components of dentin and bone consist of similar components, such as biological apatite (HAp: 70%), collagen 18%), non-collagenous proteins NCPs, 2%), and body fluid (10%) in weight by volume [8%, non-c%, and boThe general method for producing DDM, which is an acid-insoluble type I collagen showing a high degree of cross-linking with collagenous and matrix-binding proteins such as transforming growth factors (TGFs), insulin growth factor, fibroblast growth factor (FGF), and bone morphogenetic proteins (BMPs), involves crushing dentine, after removing the cementum and enamel, into a powder and demineralizing it ,4,5,6.Demineralization is a necessary process for DDM to act as a bone substitute because HAp inhibits the release of growth factors, and carrying out this process does not lead to the degradation of these growth factors .Major osteoinductive growth factors, similar to dentin-matrix-derived BMPs, have been identified in rats , bovinesAccordingly, several forms of autogenous DDM (Auto-DDM)  have been developed and their clinical safety and effectiveness in implant dentistry have been tested . An AutoThe applications of Auto-DDM, as an alternative to autogenous bone grafts, have shown promising clinical and histological results for SP and GBR in implant dentistry owing to its inherent osteoinductive and osteoconductive capacity ,15,16,17Allo-DDM was conceptualized from the demineralized bone matrix (DBM), which was largely developed and defined for the bone induction principle (BIP), which states that a protein macromolecule in dentin and bone induces the differentiation of mesenchymal cells into osteoblasts; this was postulated by Urist in 1965 ,20. The Although Auto-DDM is commonly applied in some countries such as Korea, India, and Japan, very few studies have investigated the application of Allo-DDM as bone substitutes for bone graft surgery with regard to its osteoinductivity and antigenicity. This narrative review aims to summarize the preclinical evidence on the osteoinductivity and antigenicity of Allo-DDM and to provide future directions for research on the clinical safety and efficacy of Allo-DDM in maxillofacial bone regeneration.Google Scholar, Scopus, PubMed, and the Cochrane Library databases were screened for this review in January 2020. The years searched were from 1960 to 2019, using the keywords “Demineralized Dentin Matrix” AND (“Allogenic” OR “Allogeneic”) AND . The relevant full-length articles were obtained from the electronic databases, and the authors read and selected relevant studies for review according to the following inclusion criteria: (1) articles written in English in a peer-reviewed journal, (2) any in vivo  studies that included any of the search keywords, and (3) articles that focused on osteoinduction or antigenicity. The exclusion criteria were: (i) articles for which full-text was not available, (ii) articles that did not pertain to the allogeneic application of demineralized dentin matrix, and (iii) classifications described in a textbook. The final articles selected were classified based on the application of the extraskeletal and the skeletal sites after the inclusion and exclusion criteria were applied .Among the selected articles, twelve in vivo studies evaluated the osteoinductivity of allogeneic dentin at extraskeletal sites of the abdominis muscle or subcutaneous pockets in rabbits, guinea pigs, and rats mainly during the 1960s and 1970s  23,24,2,32,33,34Most of the studies evaluated the osteoinduction property of dentin with regard to histological, radiological, and biochemical outcomes at the extraskeletal sites .In 1967, Bang and Urist  first reThe qualitative trends of cellular sequences after the implantation of Allo-DDM could be as follows: (1) inflammation, (2) vascularized connective tissue formation, (3) erosion, (4) recalcification, and (5) bone formation . Nilson The sequences of cellular transformation following the implantation of Allo-DDM suggest that the graft is invaded by the vascular “mesenchyme” with a brief inflammatory reaction. Some of the mesenchymal cells became multinucleated giant cells that proceed to erode tunnels in the matrix and enlarge the pre-existing cavities . The matrix around the eroded chambers is then re-calcified, presumably due to the diffusion of mineral ions from the new blood vessels. Osteoblasts then replace the multinucleated cells on the eroded and calcified surfaces, which start to deposit the bone matrix and cement line ,45. In eIn general, osteoinduction is a surface-oriented reaction that does not involve the deep, relatively non-available structures of the matrix ,46. In aGomes and his colleagues ,37,38,42Bone morphogenetic activity in the DDM indicates that BMPs reside in or on the quaternary structures of collagen fibrils, or the protein core of proteoglycans ,43. DuriA few studies assessed the levels of antigenicity by immunologic reaction markers such as histocompatibility, second-set reaction of skin grafts in extraskeletal sites, and white blood cell (WBC) count in skeletal sites .At the extraskeletal site, histocompatibility antigens in Allo-DDM were first investigated in 1968 . Weaker In an experimental model with the rejection reaction in skin allografts in 1972, Bang  reportedWhen used in skeletal defects, Allo-DDM showed no or low antigenicity at the tissue level ,36,37. ESome BIPs are lost in DBM with lyophilization, irradiated, or heating processes . SeveralIn many studies, complete demineralization of the dentin matrix, until a calcium-free state, is achieved using 0.25–0.5 M EDTA and 0.2–0.6 N HCl ,32,33,34The minerals from the dentin matrix insulate the BMPs and interfere with the transmission of the bone morphogenetic property ,24,53. DWith regard to partial demineralization, a study from 1998 on human DDM as a carrier for recombinant human BMP-2 reported that partially demineralized dentin matrix  did not cause osteoinduction on allogeneic transplantation into the muscle of mice . On the Many researchers found that undemineralized allogeneic dentin matrix (Allo-MDM) did not induce alkaline phosphate activity and cartilage or bone formation in the extraskeletal sites ,29,31,32However, in rabbit skeletal defects, Allo-MDM acted as a three-dimensional osteoconductive scaffold contrary to the results obtained in the extraskeletal sites ,30,39,44Therefore, MDM appears to act as an osteoconductive scaffold; however, it has poor bone formation capacity or is rejected in extraskeletal sites, which requires the activation of inducible osteogenic precursor cells (IOPCs). According to Friedenstein et al.  and OwenIn summary, the demineralization with 0.2–0.6 N HCl showed the most effective osteoinductivity of Allo-DDM. With regard to the degree of demineralization, partial-DDM was superior for bone-forming outcomes in comparison with complete-DDM, since many endogenous growth factors could be lost during complete demineralization.3 [3 of dentin used by Urist and colleagues [3 introduced by Reddi et al. [The osteoinductivity of Allo-DDM at extraskeletal sites was not affected by its various shapes and sizes, including pieces of 2 × 2 × 1 mm3 ,30,32, w3 ,31, and 3 . Other glleagues ,24,25,26i et al.  and Pinhi et al.   in 1970 and 1 × 1 × 1 mm in 1990 ,34. Most in 1990 ,62,66. D in 1990 . In the  in 1990 . Recentl in 1990  reported in 1990 . Nam et  in 1990  compared3, approximately 3 µm diameter) are a unique spatial nanoporous structure that can be enlarged to microporous geometric structures by the demineralization process, resulting in increased porosity from 3% to 20% [Dentinal tubules in dentin (20,000–60,000/mm% to 20% ,54,55,69% to 20% ,54,70.In 2018, Tanoue et al.  reportedThis article represents the first comprehensive review focused on the osteoinductivity and antigenicity of Allo-DDM. Allo-DDM has demonstrated a great potential for osteoinductivity in extraskeletal sites. Allo-DDM showed low antigenicity, which neither adversely affected osteoinductivity nor provoked immunologic reactions. Owing to the limited amount of research related to Allo-DDM and the lack of follow-up studies after the initial research, there has been no clear evidence to support the free antigenicity of Allo-DDM due to the potential immune components from vital dentin, which might be removed during demineralization process. Further, even though acellular and avascular dentin matrix cannot be a carrier for a virus, safety from the risk of viral disease transmission has not been mentioned so far in the condition of in vivo. Future studies should investigate the optimization of the processing methods and the geometry of Allo-DDM, which plays an important role in its osteoinductivity and osteoconductivity. Furthermore, the risk of viral disease transmission should be researched in vivo before the clinical application of Allo-DDM.


Polymer-modified cement has found numerous applications due to its excellent tensile strength. When cement was modified with a single polymer, its tensile strength and performance marginally increased. However, when a small amount of the flexible polymer latex was added to cement, the mechanical performance of cement considerably improved. In this study, single-graded copolymer-modified pervious concrete was prepared by mixing styrene-butadiene rubber (SBR) with different acrylate polymers, and its mechanical performance and durability were systematically studied through mechanical tests and theoretical analyses. The main findings are as follows: (1) When a waterborne emulsion was added, the freeze resistance of cement concrete increased, and its mass loss rate decreased. Cement concrete with two latexes had higher sulphate corrosion resistance and substantially lower dry shrinkage strain than ordinary cement concrete. (2) Through scanning electron microscopy, the microstructural changes in the cement binder, ordinary polymer-modified concrete, and the copolymer-modified cement concrete used in this study were observed, and the findings were compared with those reported in the literature. (3) Single-graded copolymer-modified pervious concrete exhibited excellent shrinkage performance. This study showed that single-graded copolymer-modified pervious concrete satisfied the performance requirements for use as a paving material for special cases. Various methods have been developed to improve the functional performance of concrete. With the development of polymer material science, polymer latex has been proven to reduce its brittleness ,2,3,4. PIn addition to the polymer, PMC contains cement paste, mortar, and concrete . PMC andWhen PMC is prepared using a single polymer, such as rigid styrene-acrylic latex, it retains its high strength and shows a marginal improvement in compressive strength and tensile strength However, even if a small amount of flexible polymer latex is used to prepare PMC, its mechanical performance improves considerably. Such polymer-modified cement has an aggregate structure and shows excellent mechanical performance improvement in relation to the performance of the original components because of an agglutination reaction of two polymer latexes with cement, which is similar to the reaction occurring in metal alloys ,11. Ther(1)The mechanical performance and durability indexes, such as freeze resistance, sulphate corrosion resistance, and dry shrinkage performance, of concrete samples with different amounts of the polymer latex were measured.(2)The microstructure of polymer-blended concrete samples was observed using scanning electron microscopy (SEM), and the effect of the microstructure on cement performance was studied.However, Sofi et al.  reportedThe findings of this study showed that single-graded copolymer-modified cement concrete can play an important role in structural materials due to its resistance to permeation, cracking, corrosion, and shaking. Consequently, it can find extensive applications in the construction of roads and bridges, airport runways, and marine structures.In this study, cement, gravel, and compound polymer latex were used. The sample used in this paper is the same as that used in our previous study .(1) Cement: Cement is the most important binding material in concrete. Ordinary silicate cement is utilised in polymer-modified mortar and PMC. In China, road cement is required to meet the technical specifications listed in the Chinese standard JTG D40-2011 , accordi(2) Coarse aggregate: The selected raw material was stone provided by Jiangchuan Building Material Factory, Chongqing, China. After treating the raw material using several processes—such as screening in a screener, rotating and cleaning in a mixer, and drying in a dry box—single-graded stone with a grain size of 4.75–9.5 mm was obtained.(3) Polymer: The copolymer latex was prepared by blending styrene-butadiene rubber (SBR) and the acrylate polymer, and the modification effect of a single polymer latex SBR on cement was studied at the early stage . In a pr(4) Water: The selected raw material was ordinary tap water.The concrete mixture used contained water, cement, and aggregate in a ratio of 0.41:1:4.7 by weight, based on our previous study ,13,14. VAll the samples were prepared according to the Chinese Standard GB/T 50081-2019 , demouldThe selected latexes were mixed with the base polymer latex according to the predetermined proportions, and the compound latex was well stirred and completely blended. Cement and stones were weighted according to the predetermined proportions and stirred with an SWJ-60 double-horizontal-shaft forced concrete mixer . The dry mixing time was 2 min. After cement and stones were properly mixed, the completely blended compound latex was slowly added, and the new mixture was further stirred to obtain well-stirred copolymer-modified concrete. The stirring time was 120 s. The stirring process is shown in Due to the large porosity and loose structure of single-graded copolymer-modified cement concrete, if a vibrator is used, cement paste and stone can be separated. Therefore, a handheld drill rod  was used to tamp concrete. In the sample modelling process, concrete pouring was done in two steps. First, half of the concrete was poured and tamped with the drill rod. Then, the other half was poured, and all concrete was further tamped to ensure the overall compactness and uniformity of the samples. Finally, the surface of the samples was smoothed.In the sample preparation process, the standard curing method was adopted, followed by plastic film covering. This method was adopted to avoid water loss in the curing process to enable the participation of both polymer and water of the compound latex in the hydration process. This enables the proper formation of polymer cement paste and helps avoid rapid water evaporation, which can lead to incomplete cement hydration and thus adversely affect the concrete strength. Therefore, after the completion of moulding, all the samples were placed in a curing room .. The samples were demoulded after curing them for 24 h. They were then covered with a plastic film before further curing at room temperature. The sample dimensions are listed in Dry shrinkage is structural shrinkage with drying, which is caused by gradual water evaporation during the hydration or hardening of cement concrete, carbonisation resulting from the chemical reactions of hydration products  with airborne carbon dioxide at certain humidity levels, and temperature changes.Standard for Test Methods of Long-Term Performance and Durability of Ordinary Concrete [t is the dry shrinkage rate at a certain age (accuracy: 0.001%), L0 is the initially read value (mm), Lt is the measured value at a certain age (mm), and 250 is the effective length of the samples.The method for the dry shrinkage test was designed in view of the Chinese Concrete   [(1)The samples were cured in the standard curing room for 24 days and then soaked in water at 20 °C ± 2 °C for 4 days. In the soaking process, the water surface was kept 20 mm higher than the top surface of the sample. The freeze–thaw cycle test was started when the sample age reached 28 days.(2)The samples  were rem(3)The samples were inspected at regular intervals. For each group, the freeze–thaw cycle test was stopped when the mass loss reached 5%.(4)The samples were taken out after 50 freeze–thaw cycles and wiped dry to remove any surface water. Their compression strength was then measured.The method for the freeze–thaw cycle test was designed by referring to the slow freezing method outlined in 82-2009) . The samSulphate corrosion is a cyclic occurrence of internal swelling caused by sulphate solution corrosion and internal shrinkage resulting from water content loss during high-temperature drying in cement under the alternating conditions of high temperature and sulphate solution. The cyclic swelling–shrinking phenomenon persists till the alternating high temperature–sulphate solution environment exists, and internal cracks form in cement when its ultimate stress is exceeded.Characteristics of sulphate attack on concrete: the concrete surface turns white, all edges and corners begin to be damaged, cracks appear, and then concrete peeling occurs. The concrete becomes loose and fragile from compactness. The essence of sulphate erosion damage to concrete is that when the concrete is in the sulphate rich area, sulphate ions enter the concrete and react with the components in cement stone to form refractory salt minerals. These substances can form easily expandable chemical substances such as ettringite and gypsum. The expansion of ettringite and gypsum will cause micro cracks in the concrete, leading to cracking, peeling and the disintegration of concrete .Standard for Test Methods of Long-Term Performance and Durability of Ordinary Concrete [The test method for the sulphate corrosion test was designed by referring to Concrete . An actu4 solution for 14 and 28 days, respectively, and then observed for surface damage and their compression strength was measured, on the basis of which the corrosion coefficient of compression strength of cement was calculated. The formula for the coefficient is as follows:f is the corrosion coefficient of compression strength (%), f0 is the mean value of the compression strength of the samples without sulphate corrosion (MPa), and fn is the mean value of the compression strength of the samples with sulphate corrosion (MPa).To study the impact of sulphate corrosion on the durability of ordinary cement concrete and single-polymer-modified cement concrete with different amounts of the polymer latex, the original components of cement should be free from sulphate, thereby ensuring that any sulphate corrosion is external. The test had 17 groups, each with 6 samples, or 102 samples in total. The samples, designed as 100-mm-edge cubes, were divided into seven types. The samples were cured in the standard curing room for 28 days, soaked in clear water and 5% NaSOTo observe the modification of concrete caused by the copolymers, the morphologies, microstructural characteristics of samples are observed by Zeiss Gemini Sigma 300 VP SEM machine . The working distance of the machine’s lens and samples is 8.5 mm and the general test steps are conducted according to study .There are three types of samples: ordinary cement concrete, single-polymer-modified cement concrete, and copolymer-modified cement concrete. Single-polymer-modified cement concrete contains only one type of polymer, i.e., 100% latex SBR. Copolymer-modified cement concretes that contains three type of polymers, respectively, i.e., 70% latex SBR + 30% latex P1, 70% latex SBR + 30% latex P2 and 70% latex SBR + 30% latex P3 were selected for the analysis.Sample preparation: first, five disposable paper cups are used and labelled for this test corresponding to the test piece one by one. Second, samples are taken during the 28-day compressive strength test of cement-based materials. A large number of small fragments are generated during the compressive test. The size of the sample selected from the fragments shall not exceed 3.5 mm * 3.5 mm. The sample surface shall be flat for observation during scanning. Third, dry the sample in an electric blast drying oven  at 60 °C for 24 h. Fourth, put the dried specimen in the gold plating instrument  for gold plating. Gold plating can improve the conductivity of the specimen and improve the clarity of the scanned image of the specimen. The experiment was conducted in the materials laboratory of Chongqing Jiaotong University. The scanning process is shown in The results of the dry shrinkage test for the various groups of cement paste samples at various ages from day 1 to 28 are listed in (1)For all groups of the copolymer-modified cement concrete samples, the dry shrinkage rate first gradually increases and then stabilises when the sample age increases. Before 7 days, the dry shrinkage rate increases at the fastest rate, reaching approximately 85% of the dry shrinkage rate on the 28th day. After 7 days, the dry shrinkage rate increases at a steady rate.(2)After the polymer latex is added, the dry shrinkage performance of the cement paste samples substantially improves with the latex content. After P1 and P2 are added, the dry shrinkage rates of the copolymer-modified cement paste are decreased by 6.68–12.76% and 7.78–12.21%, respectively. By contrast, after P3 is added, the dry shrinkage rate is decreased by only 1.23% and increased by 1.23% when the latex content is 30%. These findings indicated that the dry shrinkage performance of copolymer with SBR and P1 or P2 is better than that of the copolymer with SBR. Furthermore, the dry shrinkage performance of the copolymer with SBR and P3 decreases when the P3 content increases.The following conclusions can be derived from the strain values given in The main reasons for the improvement in the shrinkage performance of polymer-modified cement are as follows: Protective films are generated by the polymer on the surface of the hydrated cement products. These films prevent internal water evaporation and promote cement hydration reaction. Furthermore, solid components in the polymer become dispersed in the microcracks of the cement stone, which increases cement stone compactness and reduces the number of water evaporation paths in the cement stone. In a dry environment, the water pressure in the pores in the cement stone is maintained, which reduces the dry shrinkage rate of the cement paste.(1)The surface of the cubic samples of the copolymer-modified cement binder is smooth and damage-free.(2)As the freeze–thaw cycle test proceeds, the sample colour changes from grey to white, and many tiny pores gradually form on the sample surface.(3)As latex loss in the samples due to water soaking continues, the surface pit area and the number of pores increase in the samples, and extensive arc-shaped fall-off occurs at the edges and corners in some samples.The results of this test showed the compression failure of the samples after the freeze–thaw cycles as gradual damage starting from the sample surface. The damage process of the samples can be generally divided into three parts as shown in (1)After 50 freeze–thaw cycles, a minor mass loss occurs in all the samples, and damage appears on the sample surface to varying degrees. The test results indicated that the freeze resistance performance of cement binders with various polymer combinations is significantly better than that of the ordinary cement paste.(2)The mass loss rates of the copolymer-modified cement binder are lowest for the latex content of 30% P1 and 30% P2, which show the mass loss rates of 0.04% and 0.03%, respectively, or 7 and >9 times that of ordinary cement binder, respectively. The freeze–thaw failure develops from the sample surface and travels inside. Due to the restrictions of test conditions, only 50 freeze–thaw cycles were conducted; consequently, freeze–thaw failure from the surface to the inside of the samples could not be observed, and only edge and corner damages and arc-shaped damages were detected in some samples.The following conclusions can be derived from the test data given in (1)The corrosion coefficients of the compression strength of various groups of the copolymer-modified concrete samples are smaller than those of ordinary skeleton-pore cement concrete and ordinary polymer-modified skeleton-pore cement concrete to varying degrees. Among the samples, the corrosion coefficient of compression strength of ordinary skeleton-pore cement concrete is the highest, reaching 45.32%, and its sulphate corrosion resistance is the lowest.(2)After the latex is added, the sulphate corrosion resistance performance of the samples improves. Furthermore, the resistance performance improvement increases with the latex content. The sulphate corrosion resistance performance of copolymer-modified cement concrete blended with latex P3 is the optimum, which is higher than that of ordinary cement concrete by approximately 26.14%. However, its after-test compression strength is still lower than that of other types of copolymer-modified cement concrete because of its low before-test compression strength.(3)In comparison, the sulphate corrosion resistance performance of copolymer-modified concretes blended with latexes P1 and P2 is optimum, and the difference between these is considerably small. Protective films are properly formed by the polymer latex on the surface of the hydration products of cement, thereby effectively reducing the sulphate content in cement, as well as ettringite and gypsum contents, which are materials with expansibility.The following conclusions can be derived from the values given in (1)As shown in (2)As shown in The SEM magnification was adjusted to 500 times, and the cross-sectional SEM images of the ordinary cement and polymer-modified cement binders are shown in The comparative analysis of the SEM images presented in In this study, the mechanical performance of copolymer-modified cement concrete in a corrosive environment was studied. Several durability performance indexes of various concrete samples were determined by testing parameters, such as freeze resistance, corrosion resistance, and dry shrinkage. The microstructures of the samples were observed through SEM, and the effect on cement performance was analysed.(1)The dry shrinkage coefficient of polymer modified cement is less than that of ordinary cement concrete, and the shrinkage of blended polymer cement is better than that of polymer modified cement. Among them, the blending of P1 and P2 has the best improvement effect.(2)The frost resistance of copolymer-modified cement concrete was analyzed by compressive strength loss and mass loss. The test shows that the compressive strength of the concrete with the addition of polymer emulsion is lower than that of the ordinary cement concrete, but the compressive strength and the mass loss rate are better than those of ordinary cement concrete, that is, its frost resistance is better than that of the ordinary cement concrete.(3)A comparative analysis of the loss rates of compression strength after soaking in 5% sulphate solution for 14 and 28 days showed that copolymer-modified cement concretes with P1 and P2 have the highest sulphate corrosion resistances, and is better than a simple polymer modified cement concrete and ordinary cement concrete.The main conclusions are as follows:In this study, the mechanical performance and durability of copolymer-modified cement concrete were studied. The effects of various copolymers on the mechanical performance of concrete at the same polymer-to-cement ratio were analysed. Furthermore, some preliminary conclusions for future research on new paving materials were proposed. However, further study in some areas is required. For example, the durability of materials such as fatigue performance needs to be further studied, and there is still the problem that the amount of polymer latex is too high. It is necessary to further improve the performance of latex to reduce the amount of latex to reduce the cost, so as to be applied to more fields.


The present research study aims to investigate numerically the behavior of steel fiber-reinforced high-strength concrete (SFRHC) beam–column joints (BCJs) under seismic action. Based on the plastic damage constitutive model of concrete and elastic–plastic mixed-strengthen constitutive model of steel material, the finite element software ABAQUS was utilized to establish the 3D finite element (FE) model of BCJs. Additionally, the feasibility and accuracy of the numerical simulation were verified by comparing the computed results and experimental observations in terms of the hysteresis curves, skeleton curves, and failure mode. Furthermore, based on the validated FE modeling approach, load vs. displacement hysteresis curves of SFRHC–BCJs during the loading process were analyzed in detail; the failure process was also investigated. Furthermore, the effect of various parameters on the seismic behavior of BCJs was analyzed comprehensively, including the concrete strength, the volume ratio of steel fiber, and the stirrup ratio in the core area. Finally, parametric studies illustrated that increasing the concrete strength helps in enhancing the ultimate load, while the ductility decreased noticeably. Both adding the steel fiber and increasing the stirrup ratio can significantly improve the seismic performance of BCJs. Beam–column joints (BCJs) are the critical parts of the frame structure, and their failure can easily lead to the collapse of the whole structure. Previous tests ,3,4,5,6 Over the past years, experimental studies have always been a basic analysis method for investigating the seismic behavior of RC structures. However, with the advancement of computing techniques to conduct research more efficiently and economically, numerical analysis is becoming more and more popular to simulate the behavior and predict RC structures’ response because it can simulate the case that is difficult and/or complex . Bui et Against this background, this study aims to explore the seismic behavior of BCJs. Based on the experimental and numerical studies in this team, the main contents of this paper are as follows: Firstly, the finite element software ABAQUS was utilized to establish the 3D finite element model of BCJs. Then, the feasibility and accuracy of the numerical simulation were verified by comparing the calculation results with the experimental observation results, and the skeleton curve and failure process were analyzed and clarified in detail according to the simulation results. In addition, based on the stress nephogram, the development of the stress distribution in the concrete and reinforcement during the loading process was clarified comprehensively. Finally, based on the verified FE model, the effect of various parameters on the seismic behavior of BCJs was analyzed in detail and fully evaluated, including concrete strength, steel fiber volume ratio, and stirrup ratio in the core area. In this study, the experimental results and the specifications of the SFRHC as BCJs reported by Shi et al.  were usecd) and tension (td). The elastic behavior of concrete material was defined using the elastic isotropy option. The steel fiber-reinforced concrete constitutive model [Thus far, different materials models have been utilized in the numerical modeling of concrete cracking, such as concrete damaged plasticity (CDP), discrete crack model (DCM), and smeared crack model (SCM). In this study, the concrete damage plasticity (CDP) model was adopted to simulate the nonlinear behavior of the concrete elements under multiaxial stress state conditions. That is because the CDP model can capture the degradation of the elastic stiffness in the strain-softening branch of the stress–strain curve under compression εcf,cfkf is the standard value of axial compressive strength of SFRHC; tf,rf and tf,rε are the axial tensile peak stress and corresponding peak strain of SFRHC; tfα is the shape parameter of the descending section of the stress–strain curve of SFRHC under uniaxial tension; tf,rf is the representative value of tensile strength of SFRHC; tf,rε is the peak tensile strain of SFRHC corresponding to uniaxial tensile strength tf,rf; tfd is the damage evolution parameter of SFRHC concrete under uniaxial tension; cfE is the elastic modulus of SFRHC.In tension ,(2)ftfkcd)and tension (td) available in the concrete damage plasticity option. Since cyclic loading was applied to the FE model, the degradation of the concrete elastic stiffness was considered for the unloading process. The damage parameters in compression (cd) and tension (td) were calculated using Equations (3) and (4) [cw) and tension (tw) were considered in the concrete damage. and (4) . Furthervc) is set to 0.2; the eccentricity is 0.1; the ratio of initial equal biaxial compressive yield stress to initial uniaxial compressive yield stress (fb0/fc0) is 1.16; the ratio of the second stress invariant on the tensile meridian to that on the compressive meridian is 2/3; the viscosity parameter is taken as 0.005; the dilation angle is 30°. Meanwhile, the default values of the recovery factors of compressive and tensile stiffness wc = 0.8 and tw = 0.2 were used in the calculations, respectively. Furthermore, the compressive damage variable and tensile damage variable [The relevant parameters in ABAQUS were defined as follows: the Poisson’s ratio , which is extensively adopted to analyze the RC structures, was used to establish FE models. In those models, the 8–node reduced integral format 3D solid element (C3D8R), and the 2–node linear 3D truss elements (T3D2) were employed to model the concrete and steel reinforcement, respectively ,31. In aRigid gaskets were provided at the top and bottom of the column to prevent stress concentration at the loading point. A tie constraint may couple two separate surfaces so that no relative motion would occur. Therefore, the tie option is adopted for the constraint between the rigid gasket and the column end as well as the bottom. All the rigid gaskets were modeled as rigid bodies. The structured meshing technique was adopted, as shown in The same boundary conditions and loading modes were adopted in the FE models as the pseudo-static tests. Firstly, the top end of the upper column was constrained in X and Z directions at reference point 1, and the axial force was in the Y direction to enable the free movement of the component. Then, the bottom end of the lower column was constrained in X, Y, and Z directions at reference point 2. This simulates the hinge support system at the top and bottom ends of the tested columns. Finally, the free ends of the beam were constrained in the Y direction for imposing the displacement at reference points 3 and 4. The FE modeling is shown in 0N was applied to the top of the column in the form of pressure and kept until the end of loading. In the second step, the corresponding cyclic loading was applied to the beam end. In addition, the cyclic displacement loading system was the same as the test. Two loading steps were set in the FE pseudo-static analysis. In the first step, the initial constant axial load According to the material constitutive models, interaction, and boundary condition as mentioned above, the accuracy of numerical analysis results from established FE models are standardized against the corresponding experimental results in this section in terms of the load vs. displacement hysteresis curves, skeleton curves, ultimate bearing capacity, and failure modes.yP, peak load mP, failure load uP, the relative errors (RE), and coefficient of variation (COV). The load values for all the specimens are listed in In summary, the FE modeling approach and the material constitutive model adopted in this study are proved to be reasonable and adequate and can be used to investigate the seismic behavior of BCJs further.In this paper, there were two main failure modes of SFRHC: beam-end bending failure  and joint shear failure  in the core area. BFF mainly occurred in specimens with large stirrup ratios and volume ratios of steel fiber. This type of failure was characterized by the crushing of concrete at the compression zone of the beam after yielding tension beam bars, for instance, BCJ1 and BCJ3. JSF mainly occurred in specimens with small stirrup ratios and volume ratios of steel fiber, such as BCJ0, BCJ2, BCJ4, BCJ5, and BCJ6. This type of failure happened due to insufficient strength of stirrups inside the joint to resist the combined shear and tensile forces.Based on the validated FE modeling approach, the FE models are developed to further investigate the seismic performance of beam–column joints in terms of the predicted load vs. displacement hysteric curves during the loading process and failure modes.Elastic stage (0–1): At the initial stage of loading, all the specimens were in the elastic phase, indicating that the applied load increased linearly with the increase of the displacement. In this stage, the load increased faster than other stages, whereas the increment of the displacement was very limited. In addition, no obvious deformation of all the specimens was observed, because the vertical load was very small, and the area surrounded by the hysteretic curve was very narrow. The slope of the hysteretic curve changed little, and the residual deformation after unloading was also minimal, indicating that the specimen was in the elastic working state.Elastic–plastic stage (1–2): When the load reached point 1, many vertical cracks appeared in the concrete at the beam end, which led to the degradation of the stiffness of beam–column joints. With the increase of cyclic load, the concrete in the core area of the joint began to crack, and more cracks appeared at the root of the beam end. The displacement no longer increased linearly with the load, and the specimen entered the elastoplastic stage. The slope of the skeleton curve started to decrease and deviate from the displacement axis. Meanwhile, the stiffness degradation phenomenon was obvious. The concrete strain gradually increased and reached the yield state, and the strain of reinforcement increased accordingly. At this time, the area enclosed by hysteretic curves gradually increased and became fuller, indicating that the energy dissipation capacity of the specimen was increasing.Failure stage (2–3): After reaching the peak load (point 2), the curve entered the descending stage, and the bearing capacity of the specimen began to decline. The crack at the beam end developed rapidly, while the crack in the core area developed slowly. The strength and stiffness of SFRHC deteriorated seriously until it lost its bearing capacity and was destroyed (point 3). The load–displacement hysteretic curve of the whole beam–column joint was full without the pinch phenomenon, revealing that the steel bar and concrete can work in harmony and deliver good seismic performance.Based on the validated FE modeling approach, 12 numerical models were established to extensively investigate the effect of some key parameters on the seismic performance of beam–column joints, including the concrete strength, steel fiber volume ratio, and stirrup in the joint area. The above analysis can be expected to provide a reliable reference for design and application in RC structures.fV) ranged from 0 to 2%, while others were kept the same. As shown in The reasonable agreement between experimental and corresponding FE results is achieved generally, and therefore, the FE modeling approach can be adopted to further investigate the seismic behavior of SFRHC–BCJs.Based on the predicted load vs. displacement hysteresis curves, the specimens are considered to experience three stages during the loading process: elastic stage, elastic–plastic stage, and failure stage. In addition, the failure modes and processes were analyzed and clarified in detail, including the beam-end bending failure (BFF) and joint shear failure (JSF) in the core area.Based on the stress nephogram, the development of the stress distribution in the concrete and reinforcement during the loading process was clarified comprehensively.Increasing the concrete strength can slightly improve the ultimate load of BCJs, while the ductility decreased noticeably. Additionally, adding the steel fiber into concrete can clearly improve the seismic performance of the structure. Moreover, increasing the stirrup ratio in the core area can effectively improve the seismic performance of BCJs, and change the failure mode of BCJs.Based on the validated FE modeling approach, this paper aimed to present a numerical investigation on the seismic behavior of SFRHC–BCJs in terms of complete curves analysis, failure mode, parametric study, etc. The main conclusions can be drawn based on the limited research work, as follows: